IC 9258 D U I INF' : Mine Power Systems fHf^ YEAMS Information Circular 9258 Mine Power Systems By Lloyd A. Morley UNITED STATES DEPARTMENT OF THE INTERIOR Manuel Lujan, Jr., Secretary BUREAU OF MINES T S Ary, Director Library of Congress Cataloging in Publication Data: c <3 o ' Morley, Lloyd A. Mine power systems. (Information circular: 9258) Includes bibliographies. Includes index. Supt. of Docs, no.: I 28.27:9258. 1. Electricity in mining. I. Title. II. Series: Information circular (United States. Bureau of Mines); 9258. TN295.U4 [TN343] 622 s [622'.48] 87-600213 For sale by the Superintendent of Documents, U.S. Government Printing Office Washington, DC 20402 PREFACE The application of electricity to the mining industry is a distinctive area of both mining engineering and electrical engineering. The difficult environment, the dynamic power loads, the cyclic and mobile operation and stringent safety requirements that characterize mining, all place unique demands on the mine power system. No other industry makes such extensive use of portable extensible equipment or has such com- plex grounding problems. Mine power systems can range from relatively simple in- stallations for small surface mines to complex underground systems where the harsh environment of dust, humidity, and cramped spaces stretches the ingenuity and crea- tivity of the engineer to provide reliable service. At the present time there is no up-to-date engineering text available that deals specifically with mine power systems. This has created extensive difficulties for edu- cators, industry engineers, and regulatory agency personnel. The need for a suitable reference for students in mining engineering provided the main impetus for this book, since the technician-level material that was in existence proved unsuitable for teach- ing young engineers who have little practical experience. The objective in preparing this manuscript was to assemble a single engineering reference on mine electrical power systems that is as comprehensive as possible. Ear- lier drafts of this material have been used successfully to instruct university students in courses ranging from basic electrical engineering through power-system design. It is felt, however, that the usefulness of this material extends beyond that of a student text. While not intended to replace other electrical or mining references, this publica- tion is also an indexed, reasonably comprehensive reference handbook for industry engineers and training personnel, and a source of material for electrical engineers who wish to expand their education into industrial power-system applications. Obviously, there will be some omissions; to include all aspects of mine electrical systems in one volume would approach an impossibility, but an attempt has been made to collect together the most significant information, thereby providing the tools needed to con- tinue a knowledgeable involvement in mine electricity. This reference work is divided into three general content areas. Chapters 1 through 5 contain information considered elementary, chapters 6 through 11 deal with power- system components, and chapters 12 through 17 contain specifics on mine power sys- tems. A person familiar with electrical principals can use the earlier chapters as re- view material, but all chapters contain material relevant to mining and discuss the necessary combinations of equipment and components that should be contained in the mine power system. Emphasis throughout is placed on coal mining systems, although much of the material pertains to all mining operations. Both surface and underground power systems are discussed, the latter in more detail since these are the more com- plex systems and encounter the most problems. This publication is a thoroughly upgraded and extensively revised edition of Bureau of Mines Open File Reports 178(l)-82 and 178(2)-82, prepared under Bureau contract J0155009 by The Pennsylvania State University. It contains new chapters, new illustra- tions, and example problems that were not included in the original report. The assembly of this material has been a major undertaking. Many industry, academic, and Government agency personnel helped to review and critique practically every stage of draft preparation. The original report version was made available to students taking the mine power-systems courses at The Pennsylvania State Univer- sity, and their involvement was critical input to manuscript preparation. The author is grateful to all the companies and individuals who contributed or cooperated in this effort; so much information could not have been gathered without their help. A special thanks is owed to the late Robert Stefanko. He originally perceived the need for this text and provided guidance and encouragement throughout the proj- ect that produced the original report version. Others deserving special mention are A. M. Christman, R. H. King, J. A. Kohler, G. W. Luxbacher, T. Novak, J. N. Tomlinson, F. C. Trutt and D. J. Tylavsky. Each contributed directly to the text while on the fac- ulty or staff at The Pennsylvania State University; acknowledgements for their con- tributions are made in the individual chapters. 11 The author is grateful to the several individuals and companies that supplied noncopyrighted material for use in this publication. This material is noted by the vari- ous courtesies given throughout the text. Its incorporation does not constitute an en- dorsement by the author, The Pennsylvania State University, the University of Ala- bama, or the Bureau of Mines. Reference to specific products, equipment, or manufacturers does not imply endorsement by the Bureau of Mines. CONTENTS in Page Page Preface i Abstract 1 Part I: Fundamentals Chapter 1.— Electrical power in mining 2 Mine electrical history 2 Underground mine history 2 Surface mine history 4 Mine power equipment 4 Substations 5 Switchhouses 5 Power centers 5 Distribution equipment 5 Basic distribution arrangements 5 Radial system 5 Primary-selective system 6 Primary-loop system 6 Secondary-selective system 6 Secondary-spot network 7 Utility company power 7 Surface mining 8 Power systems in surface mines 8 Main substations and subtransmission ... 8 Surface mine distribution 9 Underground coal mining 11 Room-and-pillar mining 11 Longwall mining 12 Power systems in underground mines 13 Regulations 13 Underground mine distribution 13 Surface facility power requirements 17 Basic design considerations 17 References 19 Chapter 2.— Electrical fundamentals I 20 Basic electrical phenomena 20 Coulomb's law 20 Voltage and current 20 System of units 21 Experimental laws and parameters 21 Ohm's law 21 KirchhofPs voltage law 22 Kirchoffs current law 23 Series circuits 24 Parallel circuits 25 The magnetic field 26 Inductance 26 Capacitance 28 Electric field 28 Instantaneous power 29 Idealization and concentration 29 Direct current circuits 30 Direct current and circuit elements 30 Series and parallel resistance 30 Wye-delta transformations 33 Circuit and loop equations 36 Node equations 38 Network theorems 40 Time-varying voltages and currents 45 Steady alternating current 48 Effective alternating current 50 Phasors 51 Phasors and complex quantities 52 Impedance transforms 53 Steady-state analysis 55 Chapter 3.— Electrical fundamentals II 59 Average power and power factor 59 Complex and apparent power 59 Resonance 63 Series resonance 63 Parallel resonance 64 Transformers 64 Ideal transformer 66 Actual transformers 68 Conductor loss 68 Leakage reactance 68 Core losses and exciting current 69 Power-transformer construction 70 Transformer models 71 Determination of transformer parameters 72 Transformer efficiency and regulation ... 73 Autotransformers 74 Multivoltage transformers 74 Current and potential transformers 75 Chapter 4.— Power-system concepts 76 Basic power circuit 76 Three-phase circuits 76 Balanced three-phase circuits 76 Three-phase system voltages 77 Load connections 78 Line and phase currents 79 Equivalent delta and wye loads 80 Three-phase power 80 Three-phase transformers 82 Balanced three-phase circuit analysis 83 One-line and three-line diagrams 85 IV Page Page Circuits containing transformers 90 Per-unit system 93 Transformer impedance 94 Three-winding transformers 95 Per-unit method in system analysis 95 Unbalanced three-phase circuits 97 Fault types 98 Fault analysis 98 Symmetrical components 98 Sequence components 99 Sequence-quantity combinations 99 Symmetrical-component relationship .... 100 Symmetrical-component impedance 101 Fault calculations 101 Power terminology 102 References 103 Chapter 5.— Basic solid-state devices and instrumentation 104 Semiconductors 104 Diodes and rectifiers 104 Diode equations 105 Rectifier circuits 105 Cooling 106 Overloads 107 Three-phase rectification 107 Rectifier circuits 108 Parallel rectifier operation 109 Transistors 109 Transistor operation 109 Bipolar-transistor amplifiers 110 Field-effect transistors 112 Silicon-controlled rectifiers 113 Integrated circuits 114 Basic instrumentation 114 Basic meter movements 115 Meter-movement applications 116 Wattmeters 117 Varmeters 118 Power-factor meters 118 Power-system instrumentation 118 Instrument transformers 118 Single-phase connections 119 Three-phase connections 120 Special instruments 122 Watthour meters 122 Demand meters 122 Bridges 122 Megohmmeters 123 Phase-sequence indicators 124 Recording instruments 124 Electronic instruments 125 Electronic meters 125 Oscilloscopes 125 Tape recorders 126 Transducers 126 Instrument installations 127 Part II: Power-System Components Chapter 6.— Motors and motor control 129 Alternating current generation 129 Principle of generator operation 129 Generator construction 129 Three-phase generation 131 Direct current generators 131 Motor basics 133 Torque 133 Speed-torque relationships 133 Standardization 134 Motor type 135 Three-phase squirrel-cage induction motors 136 Elementary three-phase motor 136 Motor construction 138 Motor behavior 138 Insulation 139 Design characteristics 139 Induction-motor starting 141 Wound-rotor induction motors 142 Three-phase synchronous motors 143 Synchronous-motor starting 144 Synchronous-motor torque 145 Generated voltage 146 Power factor 146 Applications 147 Direct current motors 147 Elementary motor 147 Actual motor construction 148 Torque 148 Motor connections and performance .... 148 Ward-Leonard system 152 Mine motors 153 Applications 153 Actual equipment operation 153 Single-phase motors 156 Rotating stator field 156 Split-phase starting 157 Capacitor-start motors 157 References 158 Chapter 7. — Grounding 159 Grounding systems 160 Ungrounded neutral 160 Page Solidly grounded neutral 160 Low-resistance grounded neutral 160 High-resistance grounded neutral 160 Electric shock 161 Characteristics of mine grounding systems . . 162 Ground beds 162 Grounding in underground mining 164 Grounding in surface mines 166 Ground-bed construction 166 Ground resistance 166 Electrode configuration formulas 167 Two-layer earth structures 170 Soil-heating effects 170 Control of potential gradients 171 Ground-bed resistance measurement 172 Measurement method 172 Ground test instruments 173 Ground-bed resistivity 174 Factors affecting resistivity 174 Resistivity measurements 175 Effect of chemical treatment of soils .... 177 Ground-bed corrosion 177 General ground-bed guidelines 178 Grounding equipment 179 Grounding resistor 179 Grounding transformers 179 Summary 180 References 180 Chapter 8.— Distribution 182 Nature of cable distribution 182 Cable components 183 Conductors 184 Insulation 185 Cable jacket 185 Cable shielding 186 Cable types 186 Cable terminations 191 Cable couplers 191 Coupler contacts 192 Coupler insulation 192 Coupler housing 192 High-voltage couplers 193 Low-voltage couplers 194 Cable selection 194 Cable length 195 Conductor selection 195 Cable installation and handling 202 Borehole cables 203 Feeder cable installation 204 Recommended handling practices 204 Cable failures and repairs 206 Page Cable testing 206 Failure location 207 Splicing 207 Trolley systems 211 Trolley wire 211 Trolley feeder 211 Supports, lubrications, and turnouts 211 Rails and bonds 215 Overhead lines 216 Overhead-line design 217 Overhead-line electrocutions 218 References 222 Chapter 9.— Protective equipment and relaying 224 Switching apparatus 224 Arcs and circuit interruption 225 Switches 226 Circuit breakers 226 Circuit breakers for low and medium voltage 227 Molded case circuit breakers 228 Power circuit breakers 232 High-voltage circuit breakers 232 Typical ratings 232 Oil circuit breakers 232 Minimum-oil circuit breakers 233 Vacuum circuit breakers 234 Fuses 235 Low-voltage fuses 235 Non-time-delay fuses 236 Time-delay fuses 236 Dual-element fuse 236 Current-limiting fuses 236 Standard fuses 236 Nonstandard fuses 237 High-voltage fuses 237 Expulsion types 237 Current-limiting high-voltage fuses 238 Load-break switches 239 Relays 240 Relay terminology and types 240 Thermal relays 240 Electromagnetic-attraction relays 241 Electromagnetic-induction relays 242 Basic relay connections 244 Alternating current direct relaying 244 Alternating current potential relaying .... 246 Alternating current differential relaying . . 247 Direct current connections 247 Kinds of protection 248 Control wiring 248 VI Page Phase protection 248 Ground overcurrent 249 Ground-check monitoring 251 Advantages and disadvantages 253 Arrangements for mining 254 Zones of protection 254 Coordination 254 Ground-fault protection 254 Overloads and short circuits 255 Surface mines 255 Underground mines 256 References 259 Chapter 10.— Sizing protective devices 260 Fault current 260 Fault-current sources 260 Source equivalent circuit 260 Fault calculations for three-phase systems . . 261 Short-circuit calculation procedures 261 Three-phase calculation example 264 Computer fault analysis 268 Ground-fault current calculations 268 Direct current system faults 269 Device settings 270 Relay pickup settings 270 Short-circuit protection 270 Overload protection 271 Ground-fault protection 271 Current transformer matching 272 Current transformer accuracy 272 Accuracy calculations 273 Low-voltage circuit breaker trips 274 Overload protection 274 Short-circuit protection 275 Low-voltage power circuit breakers 275 Fuses 276 Coordination 276 References 278 Chapter 11.— Transients and overvoltages 280 Transient sources 280 Lightning phenomena 280 Switching transients 281 Capacitance switching 282 Current chopping 284 Prestrike 285 Direct current interruption 286 General switching transients 287 Other transient phenomena 287 Traveling waves 287 Electromagnetic phenomena 290 Transient-induced failures 290 Page Winding response 290 Coupling through transformers 291 Transient protection 292 Surge arresters 292 Surge arrester applications 293 Capacitors and system capacitance 295 Other suppression devices 298 Faraday shields 298 Circuit arrangements 298 Protection of overhead lines 298 Impulse performance of ground beds .... 300 References 301 Part III: Mine Power Systems Chapter 12.— Mine power centers 302 Equipment specifications 302 Mine power centers 303 High-voltage cable coupler 304 Interlock switches 305 Disconnect switch 305 High-voltage fuses 306 Surge arrestors 306 Transformers 307 Specifications 307 Transformer construction 311 Faraday shields 311 Grounding resistor 311 Busway 312 Outgoing circuit breaker 312 Ground-fault protection 314 Single-phase transformers 316 Metering circuits 316 Outgoing cable couplers 317 Ground-check monitors 317 Power-factor correction 319 Direct current utilization 320 Rectifier transformer 321 Rectifier 322 Direct current ground-fault protection schemes 323 Direct current control circuitry 324 Direct current interrupting devices 324 References 325 Chapter 13.— Switchhouses and substations . . . 326 Switchhouses 326 Switchhouse internal components 326 Switchhouse protective relaying 328 Power circuit breakers 329 Switchhouse control circuits 329 Switchhouse design 331 vu Page Substations 332 Basic substation arrangements 332 Single-ended substations 333 Double-ended substations 334 Substation transformers 334 Substation switching apparatus 335 Reclosers 335 Disconnect switches and fuses 336 Protective relaying in substations 336 Lightning and surge protection in substations 337 Substation grounding 338 Substation ground mat 339 Ground-fault protection 340 Additional mine substation loads 340 Portable substations 342 Utility voltage as mine distribution 343 Additional substation design considerations 344 References 345 Chapter 14.— Solid-state control and relaying Motor control Simple motor control Control systems Physical characteristics of thyristors . . . Direct current applications Alternating current applications Static protective relaying Operation of simplified solid-state and hybrid relays Static and electromechanical relay comparison Static relay mining applications Sensitive earth-leakage system Phase-sensitive short-circuit protection . Solid-state relays in the future Summary References 346 346 348 349 349 350 351 356 356 359 361 362 363 364 364 365 Chapter 15.— Batteries and battery charging . . . 367 Basic battery and battery-charging theory . . . 367 Battery maintenance 370 Chargers 370 Charging stations 372 Battery-box ventilation 374 Page Battery surface leakage and faults 375 Battery-charging hazards 377 References 381 Chapter 16.— Permissibility and hazard reduction 382 Terminology 382 Hazard-reduction methods 383 Explosion-proof enclosures 383 Explosion transmission 384 Enclosure joints 385 Enclosure mechanical strength and internal pressures 388 Enclosure hazards 389 Permissible equipment 391 Permissible equipment schedule 391 Maintenance of permissible equipment . . . 392 Coal dust hazards 393 Classifications of dust locations 393 Reducing dust hazards 394 Hazardous locations in preparation plants 394 References 395 Chapter 17.— Maintenance 396 Mine maintenance program 397 Economic justification 397 Preventive maintenance program implementation 397 Techniques of preventive maintenance 398 Basic electrical measurements 398 Insulation measurements 398 Megohmmeter tests 400 Mechanical measurements 404 Continuous-monitoring systems 406 Corona 406 Corona behavior 408 Corona detection 409 Partial-discharge problems in mining .... 410 Intermachine arcing 411 Ground direct current offsets 412 Summary 413 References 414 Bibliography 415 Appendix.— Abbreviations and symbols 416 Index 420 ILLUSTRATIONS 1.1. Simple mine electrical system arrangement 1.2. Simple radial distribution system 1.3. Power-center type of radial distribution . . Mil ILLUSTRATIONS-Continued Page 1.4. Primary-selective distribution system 6 1.5. Primary-loop distribution 6 1.6. Secondary-selective system 7 1.7. Secondary-spot network technique 7 1.8. Representative utility transmission and distribution 7 1.9. Subtransmission for surface mine 8 1.10. Radial strip mine distribution system 9 1.11. Secondary-selective distribution in strip mining 9 1.12. Primary-loop design for strip mining 9 1.13. Radial distribution for strip mine with overhead poleline base line 10 1.14. Radial distribution for strip mine with all-cable distribution 10 1.15. Surface mine distribution system using two base lines 10 1.16. Open pit power system 11 1.17. Layout of underground coal mine 11 1.18. Plan view of retreating longwall 12 1.19. Subtransmission for underground mine 13 1.20. Radially distributed underground power system 14 1.21. Secondary-selective distribution in underground mines 15 1.22. Utilization in continuous mining section 15 1.23. Power-system segment with longwall equipment 16 1.24. Diagram of electrical-system segment for longwall 16 1.25. Parallel-feed haulage system 17 1.26. Representative expanded radial distribution for preparation plant 18 1.27. Representative secondary-selective distribution for preparation plant 18 2.1. Circuit element illustrating voltage polarity and current flow direction 22 2.2. Simple series circuit 22 2.3. Ideal and actual voltage sources 23 2.4. Circuit for example 2.1 23 2.5. Demonstration of Kirchhoff s current law 23 2.6. Simple parallel circuits 24 2.7. Ideal and actual current sources 24 2.8. Parallel circuit for example 2.2 24 2.9. Simple series circuit and equivalent 24 2.10. Simple parallel circuit 25 2.11. Series-parallel circuit for example 2.3 25 2.12. Series-parallel circuit for example 2.4 26 2.13. Magnetic flux in a straight conductor and in a long coil 26 2.14. Demonstration of induced current 26 2.15. Two coils demonstrating mutual inductance 27 2.16. Long-coil inductance and inductor symbols 27 2.17. Toroidal coil 28 2.18. Charge, voltage, and current relationships of capacitor 28 2.19. Electric lines of force between two parallel charged plates 28 2.20. Resistor used to demonstrate instantaneous power 29 2.21. Simple example of idealization and concentration 30 2.22. Modeling of load center, trailing cable, and shuttle car 30 2.23. Basic elements of resistance, inductance, and capacitance 31 2.24. Simplification of dc circuit 31 2.25. Simple circuit reduction 31 IX ILLUSTRATIONS-Continued Page 2.26. Circuit for example 2.5 32 2.27. Circuit for example 2.6 32 2.28. Series-parallel conductances for example 2.7 32 2.29. Series-parallel circuit for example 2.8 33 2.30. Two-terminal and three-terminal networks 33 2.31. Wye and delta circuit configuration 34 2.32. "T" and "tt" circuit configurations 34 2.33. Common bridge circuit 34 2.34. Circuit reduction of bridge circuit 35 2.35. Parts of circuit 36 2.36. Circuit demonstrating two independent loops 36 2.37. Two-loop circuit for example 2.11 37 2.38. Bridge circuit demonstrating loop analysis 37 2.39. Three-loop circuit for example 2.12 38 2.40. Simple two-node circuit 39 2.41. Three-junction circuit 39 2.42. Three-junction circuit with grounds 39 2.43. Voltage-source circuit demonstrating node analysis 39 2.44. Circuit for examples 2.13, 2.15, and 2.16 39 2.45. Circuit for example 2.14 40 2.46. Circuit for demonstrating superposition theorem 41 2.47. Circuit in figure 2.44 with sources turned off 41 2.48. Demonstration of reciprocity theorem 42 2.49. Practical voltage-source model 42 2.50. Practical current-source model 42 2.51. Source transformation 43 2.52. Circuit in figure 2.44 with current sources transformed to voltage sources 43 2.53. Thevenin's theorem 43 2.54. Norton's theorem 44 2.55. Comparison of Thevenin's and Norton's circuits 44 2.56. Circuit for example 2.17 44 2.57. Active circuit for example 2.18 45 2.58. Circuits illustrating solution steps to example 2.18 45 2.59. Some time-varying electrical waves 46 2.60. Sinusoidal ac waveform 46 2.61. Steady ac showing phase shift 46 2.62. Steady ac through resistance 46 2.63. Steady ac through inductance 47 2.64. Steady ac through capacitance 47 2.65. Simple series RL circuit 48 2.66. Simple series RC circuit 48 2.67. Simple series RLC circuit 48 2.68. Graphical representation of complex number 49 2.69. Trigonometric or polar representation of complex number 49 2.70. Sinusoid versus time and as phasor 51 2.71. Phasor representation of current and voltage 51 2.72. Other expressions for phasors 51 2.73. Voltage-current phasor relationships for circuit elements 53 2.74. Steady sinusoid analysis of simple RL series circuit 53 2.75. Steady sinusoid analysis of simple RC series circuit 54 ILLUSTRATIONS-Continued Page 2.76. Steady sinusoid analysis of simple RLC series circuit 54 2.77. Circuit for example 2.21 56 2.78. Circuit for example 2.22 56 2.79. Two-loop circuit for example 2.23 57 2.80. Active circuit for example 2.24 57 3.1. Power represented as real and imaginary components 60 3.2. Illustration of leading and lagging power factors 61 3.3. Circuit demonstrating sum of complex powers 62 3.4. Simple series RLC circuit for resonance 63 3.5. Plot of impedance magnitude versus frequency for series RLC illustrating resonance 63 3.6. Circuits that exhibit parallel resonance 64 3.7. Magnetic coupling between two conductors 64 3.8. Magnetic coupling between two coils 65 3.9. Demonstration of coil winding sense 65 3.10. Dot convention for mutal inductance sign 65 3.11. Demonstration of impedance transfer in transformers 67 3.12. Ideal transformer with winding resistance included 68 3.13. Accounting for transformer leakage flux 69 3.14. Transformer magnetizing current 70 3.15. Eddy current and magnetic hysteresis creating power loss in core 70 3.16. Equivalent circuit of practical transformer 70 3.17. Common power-transformer construction techniques 71 3.18. Movement of exciting components to input 71 3.19. Transferring secondary components to primary 71 3.20. Final simplification of pratical circuit model 71 3.21. Transformer parameter test series 72 3.22. Circuit for example 3.8 73 3.23. Comparison of two-winding transformer and autotransformer 74 3.24. Two-winding transformer as an autotransformer 74 3.25. Examples of transformers for multivoltage applications 75 3.26. Two types of CT's 75 3.27. Examples of CT and PT placement in circuit 75 4.1. Basic power circuit 76 4.2. Applications of basic power circuit 76 4.3. Elementary three-phase generation 77 4.4. Three-phase voltage sources 77 4.5. Wye-connected source demonstrating line-to-line and line-to-neutral voltages 77 4.6. Balanced three-phase load connections 78 4.7. Four-wire wye-to-delta system 79 4.8. Balanced delta load illustrating phase and line currents 79 4.9. Comparison of equivalent delta and wye loads 80 4.10. Three-single-phase transformers connected for three-phase operation 82 4.11. Three-phase diagrams for the transformers of figure 4.10 82 4.12. Open-delta three-phase transformer operation 83 4.13. Per-phase reduction of wye-to-wye system 84 4.14. Per-phase reduction of delta-to-delta system 84 4.15. Three-line diagram 86 4.16. One-line diagram of circuit shown in figure 4.15 86 XI ILLUSTRATIONS-Continued Page 4.17. Commonly used symbols for one-line electrical diagrams 87 4.18. Symbols for relay functions 89 4.19. One-line diagram for example 4.7 91 4.20. Three-phase diagram of figure 4.19 91 4.21. Per-phase diagram of figure 4.19 91 4.22. One-line diagram with delta-delta transformer 92 4.23. Per-phase diagram of figure 4.22 92 4.24. One-line diagram with delta-wye transformer 92 4.25. One leg of three-phase transformer from figure 4.24 92 4.26. Approximate per-phase equivalent circuit for 750-kVA load-center transformer; impedance referred to high side 94 4.27. Transformer of figure 4.26 with impedance referred to low side 94 4.28. Simplified equivalent circuit of transformer expressed in per-unit 94 4.29. Approximate equivalent circuit of three-winding transformer expressed in per-unit 95 4.30. One-line diagram of small mine power system 95 4.31. Impedance diagram of system in figure 4.30, expressed in per-unit on a 1,000-kVA base 97 4.32. Basic fault descriptions 98 4.33. Positive-sequence, negative-sequence, and zero-sequence vector sets 99 4.34. Symmetrical component addition to obtain unbalanced three-phase set 100 4.35. Equivalent delta-connected and wye-connected loads 100 4.36. Three-phase system with line-to-neutral fault 101 5.1. Symbol and operation of a p-n junction device 104 5.2. Bias conditions and current flow for a diode 104 5.3. Diode or rectifier characteristic curve 105 5.4. Half-wave rectifier circuit and waveforms 106 5.5. Single-way full-wave rectifier waveforms 106 5.6. Bridge rectifier circuit and waveforms 106 5.7. Example of filtering a rectifier output 106 5.8. Heat sink cooling 107 5.9. Heat sink thermal relationships 107 5.10. Three-phase half -wave rectifier circuit and output voltage waveform 108 5.11. Three-phase full-wave rectifier circuit with input and output voltage waveforms 108 5.12. Parallel operation of rectifiers using paralleling reactors 109 5.13. An n-p-n junction transistor 109 5.14. A p-n-p junction transistor 109 5.15. Current relationships for p-n-p and n-p-n devices 110 5.16. Common-base amplifiers 110 5.17. Common-emitter amplifier Ill 5.18. Common-emitter characteristic curves Ill 5.19. Bias techniques for common-emitter amplifiers 112 5.20. Common-collector amplifier arrangment 112 5.21. Model and symbols for junction FET's 112 5.22. Example of a junction-FET application 113 5.23. Model and symbols for MOS-FET devices 113 5.24. SCR model and symbol 113 5.25. SCR equivalent model and circuit 113 5.26. General characteristic curve for SCR 114 5.27. Sketch of simple monolithic IC cross section 114 5.28. Top view of an actual IC 114 xu ILLUSTRATIONS-Continued Page 5.29. Examples of symbols employed for IC's 115 5.30. Permanent-magnet moving coil movements 116 5.31. Shunting d'Arsonval meter for high-current tests 116 5.32. D'Arsonval meter used to measure dc potentials 116 5.33. External shunts used for high-current measurements 117 5.34. Simple ohmmeter circuit 117 5.35. Rectifier ammeter 117 5.36. Dynamometer connected as wattmeter 117 5.37. Power-factor movement 118 5.38. Simple instrument-transformer connections 118 5.39. Voltmeter, ammeter, and wattmeter arranged as single-phase system 119 5.40. Use of transducers with standard d'Arsonval movements 119 5.41. Three-phase wattmeter connections 120 5.42. Two-wattmeter method 120 5.43. Three-phase power measurement with transducer 120 5.44. Balanced three-phase measurement of voltage, current, and average power 121 5.45. Line current measurements with two or three CT's 121 5.46. Line-to-line voltage measurements with three or two PT's 121 5.47. Simplified sketch of watthour meter induction mechanism 122 5.48. Wheatstone bridge circuits 122 5.49. Kelvin double bridge 123 5.50. Megohmmeter testing insulation resistance 123 5.51. Internal components of megohmmeter 123 5.52. Phase-sequence indicator 124 5.53. Strip-chart recorder 124 5.54. Input circuits on electronic voltmeter 125 5.55. Digital display 126 5.56. Cathode-ray tube 127 5.57. Semiconductor illustrating Hall effect 127 6.1. Production of voltage from magnetic field 129 6.2. Demonstration of ac generation 130 6.3. Cross section of machine with salient poles on stator and nonsalient poles on rotor 130 6.4. Cross section of machine with nonsalient poles on stator and rotor 130 6.5. Simplified sketch of electromechanical machine illustrating physical components 130 6.6. Elementary four-pole, single-phase ac generator 131 6.7. Elementary two-pole, three-phase generator 131 6.8. Elementary four-pole, three-phase generator 131 6.9. Demonstration of dc generation 131 6.10. Dc generator with two armature windings at right angles 132 6.11. Separately excited dc generator 132 6.12. Series dc generator 132 6.13. Shunt dc generator 132 6.14. Compound dc generator 132 6.15. Current-carrying conductor in a magnetic field 133 6.16. General speed-torque motor characteristic 134 6.17. Examples of three frame number dimensions 134 6.18. Demonstration of induction-motor operation 136 6.19. Elementary three-phase induction motor 136 6.20. Squirrel-cage rotor winding 136 XU1 ILLUSTRATIONS-Continued Page 6.21. Rotating magnetic field in elementary three-phase, two-pole induction motor 137 6.22. Induced rotor potential by rotating flux 137 6.23. Lapped windings of three-phase motor stator 138 6.24. Characteristic curves of three-phase induction motor 138 6.25. Typical torque-speed characteristic for general-purpose induction motor 138 6.26. Phasor diagrams of rotor and stator flux density for induction motor 139 6.27. Typical torque-speed characteristics for NEMA-design three-phase squirrel-cage motors 140 6.28. Other rotor-conductor designs 140 6.29. Across-the-line magnetic starter 141 6.30. Starting methods for induction motors 142 6.31. Schematic of wound-rotor induction motor showing external resistance controller 143 6.32. Torque-speed characteristics for wound-rotor motor with stepped-resistance controller 143 6.33. Simplified step starter using individually timed magnetic relays 143 6.34. Sketch showing construction of salient-pole synchronous motor 143 6.35. Simplified diagram of synchronous motor using generator for field excitation 144 6.36. External solid-state supply used to provide field excitation 144 6.37. Schematic of low-speed cylindrical-rotor synchronous motor 144 6.38. Controller used to demonstrate general starting method for synchronous motor 145 6.39. Typical torque-speed characteristic for synchronous motor with damper winding 145 6.40. Effect of load on rotor position 145 6.41. Equivalent per-phase circuit of a synchronous motor and phasor diagrams for underexcited and overexcited field winding 146 6.42. V-curves for synchronous motor 146 6.43. Plan view of typical mining shovel showing m-g set 147 6.44. Elementary two-pole dc motor 147 6.45. Elementary four-pole dc motor 147 6.46. Cross-sectional sketch of dc motor showing interpole and compensating windings 148 6.47. Interaction between armature and main-field flux to produce main-field distortion 148 6.48. Four connections for dc motors 149 6.49. Typical characteristics for shunt, series, and compound motors of equal horsepower and speed ratings 149 6.50. Simplified dc motor schematics with starting resistances 149 6.51. Faceplate manual starter 150 6.52. Multiple-switch starting 150 6.53. Drum-type starter 150 6.54. Simplified diagram of dynamic braking applied to shunt motor 151 6.55. Two-step resistance starting of series-wound motor 151 6.56. Forward-reverse switching of series-wound motor 152 6.57. Dynamic braking applied to series-wound motor 152 6.58. One-step starting of compound-wound motor 152 6.59. Basic Ward-Leonard system 153 6.60. Typical characteristic curves for each motor in traction locomotive 155 6.61. Stator field of two-pole, single-phase induction motor 156 6.62. Rotor field of stationary two-pole, single-phase induction motor 156 6.63. Phase relationships between stator and turning rotor 156 6.64. Starting and running stator windings 157 6.65. Centrifugal switch to remove starting winding 157 6.66. Capacitor-start motor 157 7.1. Illustration of electrical shock hazard 159 7.2. Capacitance coupling in ungrounded system 160 XIV ILLUSTRATIONS-Continued Page 7.3. Solidly grounded system 160 7.4. Resistance-grounded system 160 7.5. Effect of frequency on let-go current for men 162 7.6. Simplified one-line diagram of substation 163 7.7. Step potentials near grounded structure 163 7.8. Touch potentials near grounded structure 163 7.9. Line-to-earth fault resulting in current flow through safety ground bed 163 7.10. Lightning stroke to equipment causing current flow through safety ground bed 164 7.11. Lightning stroke current through system ground bed causing elevation of safety ground bed 164 7.12. One-line diagram of simplified mine power system 164 7.13. Mixed ac-dc mine power system; dc load energized from trolley system 165 7.14. System grounding with current-limiting resistors 165 7.15. Diode grounding of machine frame 165 7.16. Resistance of earth surrounding electrode 166 7.17. Decrease in earth resistance as electrode penetrates deeper soil horizons 167 7.18. Calculated values of resistance and conductance for 3/4-in rod driven to depth of 25 ft 167 7.19. Calculated values of resistance and conductance for 3/4-in rod driven to depth of 100 ft 167 7.20. Nomogram to provide resistance of driven rod 168 7.21. Resistance of one ground rod, 3/4-in diameter 168 7.22. Resistance of parallel rods when arranged in straight line or circle with spacing equal to rod length 168 7.23. Variation of earth resistance as number of ground rods is increased for various spacings between rods 168 7.24. Values of coefficient V. x as function of length-to-width ratio of area 169 7.25. Values of coefficient k 2 as function of length-to-width ratio of area 169 7.26. Influence of first-layer height of potentials 171 7.27. Potential on ground surface due to rod 6 ft long and 1-in diameter buried vertically at various depths 172 7.28. Potential on ground surface due to strips, 1 in by 0.1 in, of various lengths buried horizontally at depth of 2 ft 172 7.29. Measuring resistance of grounding system 173 7.30. Concentric earth shells around ground connection being tested and around current electrode 173 7.31. Correct spacing of auxiliary electrodes to give true resistance within 2.0% 173 7.32. Resistivity range of some rocks, minerals, and metals 174 7.33. Variation in soil resistivity with moisture content 175 7.34. Typical resistivity curves of solutions 175 7.35. Diagram for four-electrode resistivity survey showing lines of current flow in two-layer earth 176 7.36. Connections for Wenner four-terminal resistivity test using megohmmeter 176 7.37. Typical curve of resistivity versus electrode separation 176 7.38. Reduction in ground mat resistance by soil treatment 177 7.39. Seasonal resistance variations attenuated by soil treatment 177 7.40. Trench model of soil treatment 177 7.41. Voltage gradients in earth during ground-fault conditions 178 7.42. Delta secondary with zig-zag grounding 180 7.43. Delta secondary with wye-delta grounding transformer 180 8.1. Cable distribution in underground coal mines 182 8.2. Cable distribution in surface coal mines 183 8.3. Shield types 186 8.4. Cross sections of round unshielded mining cables 188 XV ILLUSTRATIONS-Continued Page 8.5. Cross sections of flat unshielded mining cables 188 8.6. Cross sections of some shielded mining cables 188 8.7. Round unshielded mining cables 189 8.8. Flat unshielded mining cables 189 8.9. Round shielded mining cables 189 8.10. Cable types for typical distribution systems in underground coal mines 190 8.11. Cable types for typical distribution systems in surface coal mines 190 8.12. Cable terminations for applications up to 15 kV 191 8.13. Coupler components 193 8.14. Simplified one-line diagram for situation described in example 8.4 201 8.15. Allowable short-circuit currents for insulated copper conductors 202 8.16. Representative end-suspension termination for borehole cable 203 8.17. Messenger wire supports for mine power-feeder cable 205 8.18. Splice layout using template for staggered connections 208 8.19. Effective method for removing unwanted insulation 208 8.20. Staggering splice connections 209 8.21. Examples of popular connectors and connections used in splices 209 8.22. Reinstating power conductors with soft rubber tape 210 8.23. Typical taped splice in high-voltage shielded cable 211 8.24. Trolley-wire cross sections 212 8.25. Typical trolley-wire and feeder-cable supports 214 8.26. Trolley-wire semicatenary suspension 214 8.27. Trolley system accessories 215 8.28. Theoretical resistance of bonded joint 216 8.29. Pole strength calculations 217 8.30. Guy and log-anchor calculations 218 8.31. Typical arrangements and pin-insulator spacings on wooded poles 218 9.1. Typical system fault current 225 9.2. Steps in circuit interruption 225 9.3. Arc between two contacts 225 9.4. Load-break switch 226 9.5. Extinguishing arc by increasing the length 227 9.6. Metal-barrier arc chute assists in arc deionization 227 9.7. Insulated-barrier arc chute used with magnetic field 227 9.8. Molded-case circuit breaker components 228 9.9. Magnetic-trip relay 230 9.10. Adjustable instantaneous setting 230 9.11. Thermal-magnetic action of molded-case circuit breaker 230 9.12. Time-current characteristics for thermal-magnetic circuit breakers 230 9.13. Shunt-trip and undervoltage-release accessories 231 9.14. Construction and operation of dead-tank OCB 233 9.15. Turboaction are chamber for OCB's 233 9.16. Cross section of minimum-oil breaker 234 9.17. Cross section of VCB 234 9.18. Operating mechanism for vacuum interrupter 235 9.19. VCB assembly incorporating a load-break switch 235 9.20. Common cartridge fuses 236 9.21. Inside view of dual-element fuse 236 9.22. Current-limiting action of fuses 237 XVI ILLUSTRATIONS-Continued Page 9.23. Energy-limiting action of fuses 237 9.24. High-voltage power fuse and support 238 9.25. Fusible element under spring tension in high-voltage fuse 238 9.26. Cross section of boric acid power fuse refill 238 9.27. Disassembled refill unit for boric acid fuse 238 9.28. Load-break switch with interlocked high-voltage fuses 239 9.29. Relay contact symbols 240 9.30. Temperature-monitoring protector 240 9.31. Electromechanical-thermal relays 240 9.32. Solenoid and clapper relays 241 9.33. Polar relay 242 9.34. Common induction-disk relay 242 9.35. Front view of induction-disk relay removed from case 242 9.36. Inverse-time curve compared with definite-time curve 243 9.37. Various time characteristics of induction units 243 9.38. Family of inverse-time characteristics 244 9.39. Cylinder directional relay 244 9.40. Directional overcurrent relay using induction-disk relay and cylinder relay 245 9.41. Direct relaying in ac system 245 9.42. Potential-relaying connections 246 9.43. Differential-relaying connections 247 9.44. Dc direct-relaying connections 247 9.45. Typical control wiring for UVR 248 9.46. Typical control wiring for shunt-tripping element 248 9.47. Three-phase overcurrent and short-circuit connections 248 9.48. Two CT approaches 249 9.49. Neutral-resistor current-relaying scheme 249 9.50. Neutral-resistor potential-relaying scheme 250 9.51. Zero-sequence ground relay connections 250 9.52. Ground relay in residual connection 250 9.53. Broken-delta protection 251 9.54. Series loop ground-check monitor 251 9.55. Transmitter loop ground-check monitor 252 9.56. Bridge-type ground-check monitor 252 9.57. Pilotless ground-check monitor 252 9.58. Some difficulties associated with ground-check monitoring in mining 253 9.59. Pilot interlocking circuit using ground-check monitor 254 9.60. Simple surface mine power system illustrating protective relaying 255 9.61. Typical schematic for three-phase molded-case circuit breaker with ground-overcurrent and ground-check protection 256 9.62. One-line diagram of simple underground mine power system illustrating protective circuitry 257 9.63. Diode-grounded system with possible fault indicated 257 9.64. Basic grounding-conductor system 258 9.65. Relayed grounding-conductor system 258 9.66. Neutral-shift system 258 9.67. Current-balance dc ground-fault relaying using saturable reactor 259 9.68. Current-balance dc ground-fault relaying using saturable transformer 259 10.1. Fault current waveform illustrating asymmetry 262 10.2. Multiplying factors applied to three-phase faults to obtain momentary ratings for switching apparatus . . . 264 xvu ILLUSTRATIONS-Continued Page 10.3. Multiplying factors applied to three-phase faults to obtain close-and-latch ratings for switching apparatus 264 10.4. One-line diagram for fault calculations 264 10.5. Impedance diagram for one-line diagram of figure 10.4 266 10.6. Simplification of figure 10.5 266 10.7. Simplification of figure 10.6 267 10.8. Further reduction of example network 267 10.9. Equivalent circuit of figure 10.6 267 10.10. Example problem with motor contribution neglected 267 10.11. Network to calculate momentary or close-and-latch current duties 267 10.12. Fault current in dc system 269 10.13. Available fault current versus distance of fault from rectifier on typical trolley systems 269 10.14. One-line diagram for pickup setting example 271 10.15. Model of CT and its burden 272 10.16. Typical set of saturation curves for 600/5 multiratio bushing-type CT 273 10.17. Example of one-line diagram for preparing a coordination curve plot for one path 277 10.18. Coordination curve plot for figure 10.17 showing various protective-device characteristics 277 11.1. Schematic representation of lightning stroke discharge 280 11.2. Distribution of crest currents in lightning strokes 281 11.3. Map showing average number of thunderstorm days per year in United States 281 11.4. Striking distances for negative and positive strokes 281 11.5. Crest voltages induced on transmission lines by nearby strokes 281 11.6. Simple circuit to illustrate capacitance-switching voltage transients 282 11.7. Voltage and current waveforms before and after current interruption 282 11.8. Voltage and current transient waveforms occurring with capacitance switching and restrike 283 11.9. Per-phase diagram of 4,160-V pump-motor circuit 283 11.10. Voltages and current wavesforms resulting from multiple restrikes after capacitance switching 284 11.11. Graphic example of current chopping by breaker interruption 284 11.12. Equivalent circuit of power-system segment with lumped components per phase, neglecting resistance . . 284 11.13. Graphic example of chopping voltage transients 285 11.14. Segment of mine power system 285 11.15. Circuit to demonstrate voltage transients in dc system 286 11.16. Transient overvoltage resulting from current interruption on dc system 286 11.17. An undergrounded system, showing capacitive-current flow 287 11.18. An undergrounded system, with fault on phase A 287 11.19. The distributed inductance and capacitance of two-wire line shown as incremental sections 288 11.20. Demonstration of traveling wave on overhead line 288 11.21. Incident waves being reflected and refracted at discontinuity 289 11.22. Electric field between conductors 290 11.23. A 1.2 x 50 wave test used for BIL measurement 290 11.24. Equivalent circuit of multiturn winding showing distribution inductance and capacitance 291 11.25. Initial voltage distribution across uniform winding from step function 291 11.26. Capacitive coupling of transient voltage through two-winding transformer 291 11.27. Basic valve surge arrester 292 11.28. Surge arrester with nonlinear resistance grading to equalize each gap structure 293 11.29. Surge approaching surge- arrester-protected equipment 294 11.30. Typical surge protection of rotating machinery and dry-insulated transformers 295 11.31. Simplified sketch of mine power-system segment 296 Will ILLUSTRATIONS-Continued Page 11.32. Capacitance for 2,300-V induction motors 297 11.33. Capacitance for 2,300-V synchronous motors 297 11.34. Overhead ground-wire shielding for low and high distribution towers 299 11.35. Static-wire-protection designs of wooded support structures using 30 protective angle 299 11.36. Ratio of impulse to 60-Hz resistance as a function of peak impulse current, for driven rods 300 11.37. Impulse breakdown of sand for two moisture conditions using spherical electrodes 300 11.38. Impulse characteristics of spherical electrode, with seven attached pointed protrusions of various lengths 300 12.1. Typical power centers used in underground coal mines 304 12.2. Schematic illustrating major components in power center 304 12.3. Top view of mine power center showing placement of many internal components 305 12.4. Interconnections between input and feedthrough receptacles 305 12.5. Graph illustrating transient crest voltage caused by ribbon-element current-limiting fuse operation 306 12.6. Comparison of transformer withstand characteristic and surge arrester withstand characteristic 307 12.7. Typical primary winding taps on power cable transformer 308 12.8. Zig-zag grounding transformer 309 12.9. Delta-wye connection for deriving a neutral , 309 12.10. Technique for measuring transformer impedance 309 12.11. Typical X/R ratio versus transformer capacity 310 12.12. Typical mine power-center transformer under construction 311 12.13. Completed transformer prior to installation 311 12.14. Typical bus work in power center under construction 312 12.15. Typical conductor connection to molded-case circuit breaker 313 12.16. Zero-sequence relaying on outgoing circuit with control connections to breaker 314 12.17. Zero-sequence relaying with jumper in relay case 314 12.18. Neutral relaying applied to grounding-resistor current as backup protection 315 12.19. Backup protection devices associated with mine power cables 315 12.20. Typical test circuit for zero-sequence relaying 315 12.21. Simple control circuit incorporating one ground-fault relay and one ground-check relay 316 12.22. Simple convenience-outlet circuit for 120- or 240-V single phase 316 12.23. Fuse mountings 316 12.24. Typical metering circuit for line-to-line voltages 317 12.25. Typical metering circuit for line currents 317 12.26. Typical impedance monitor circuit 318 12.27. Block diagram of continuity monitor connected in pilotless mode 318 12.28. Block diagram of continuity monitor wired for pilot operation 318 12.29. Application of power-factor correction in mine power center 320 12.30. General arrangement of dc components for combination power center 320 12.31. Full-wave bridge rectifier 321 12.32. Series reactance to reduce available short-circuit current 321 12.33. Separate transformer to increase impedance of dc circuit 321 12.34. Typical full-wave bridge rectifier with two diodes in parallel per leg 322 12.35. Diode with RC snubber protection 322 12.36. Diode-grounded system 323 12.37. Basic grounding-conductor system 323 12.38. Relayed grounding-conductor system 323 12.39. Neutral-shift system 323 12.40. Differential current scheme 323 12.41. Representative control circuit for rectifier 324 12.42. Cross section of dc contactor 324 XIX ILLUSTRATIONS-Continued Page 13.1. Diagram for typical single switchhouse 326 13.2. Control circuitry for single switchhouse using battery tripping 326 13.3. Diagram for typical double switchhouse 327 13.4. Control circuitry for double switchhouse using capacitor tripping 327 13.5. Typical family of curves for inverse-time relay 328 13.6. Illustration of fault location for adjusting selectivity 328 13.7. Typical control circuit for double switchhouse using capacitor tripping 330 13.8. Typical control circuit for single switchhouse using battery tripping 331 13.9. Overall view of main substation serving mine 332 13.10. Radial distribution applied to underground mine and its surface facilities 333 13.11. One-line diagram for single-ended substation with fuse-protected transformer 333 13.12. One-line diagram for single-ended substation with circuit-breaker-protected transformer 333 13.13. Simplified one-line diagram for doubled-ended substation 334 13.14. Typical liquid-immersed transformer in substation 334 13.15. Dead-tank OCB in substation 335 13.16. Standard percentage-differential relaying system for transformer protection 336 13.17. One-line diagram of substation with percentage-differential relaying 337 13.18. Insulation characteristic of liquid-immersed transformer compared with the characteristic of valve surge arrester 338 13.19. Plan view showing locations of system and safety ground beds 338 13.20. Typical system ground bed for large substation 340 13.21. Typical system ground bed for small substation 340 13.22. Substation feeding both surface and underground loads (no grounding conductor) 341 13.23. Substation feeding both surface and underground loads 342 13.24. Typical portable substation to service small mine 343 13.25. Providing mine ground and protective relaying from utility substation 344 13.26. Use of isolation transformer with utility substation 344 14.1. Model and circuit symbol for thyristor 346 14.2. Typical characteristics curve for thyristor 346 14.3. Thyristor half-wave rectifier 347 14.4. Alternating current thyristor control 347 14.5. Three-phase control with bidirectional thyristor arrangement 348 14.6. Full-wave thyristor bridge rectifier 348 14.7. Three-phase thyristor-controlled rectifier 348 14.8. Simplified chopper control 348 14.9. Basic control-system block diagram 349 14.10. Simplified block diagram of a motor controller 349 14.11. Common thyristor configurations 349 14.12. Heat sinking of disk-type thyristors 349 14.13. Block diagram of ac-dc shuttle car 350 14.14. Block diagram of ac-dc continuous miner 351 14.15. Simple variable-frequency control 351 14.16. Elementary inverter circuit 351 14.17. Use of variable-frequency drive on production mining shovel 352 14.18. Simplified diagram of current-regulated static belt starter 353 14.19. Simplified diagram of linear-acceleration static belt starter 353 14.20. Types of thyristor firing pulses 355 14.21. Thyristor protection for static belt starters 355 14.22. Protective-relay connections 356 14.23. Simple electromechanical relay 357 XX ILLUSTRATIONS-Continued Page 14.24. Simple static relay 357 14.25. Transistor used as relay : . . . 357 14.26. Optical transistor as relay 357 14.27. Thyristor used as relay 357 14.28. Triac used as relay 358 14.29. Hybrid static relays 359 14.30. Simple overcurrent static relay 359 14.31. Simplified sketch of the SEL system 362 14.32. Simplified sketch of the multipoint SEL system 363 14.33. Diode-bridge phase-sensitive protection 364 14.34. Equivalent model of figure 14.33 364 14.35. Electronic-comparator method of phase-sensitive protection 364 14.36. Digital-controlled continuous static relay used for timed overcurrent 365 15.1. Composition of lead-acid storage battery in various states of charge 368 15.2. Voltage per cell of a typical lead-acid battery with varying continuous rates of discharge 369 15.3. Typical charging process of cell from 18-cell, 725-Ah battery 369 15.4. Simplified schematic of saturable-reactor charger 371 15.5. Simplified schematic of single-phase thyristor charger 371 15.6. Two-winding transformer model 371 15.7. Representation transformer magnetization curve 371 15.8. Ferroresonant transformer model 372 15.9. Ferroresonant transformer 372 15.10. Ferroresonant battery charger 372 15.11. Plan of underground charging station 373 15.12. Circuit for detecting faults in batteries 376 15.13. Curve of relay current for various fault positions on battery 376 15.14. One-line diagram of desired charger features 378 16.1. Cross-sectional sketch of typical explosion-proof enclosure 384 16.2. Typical plane-flange joint 385 16.3. Typical step-flange joint 385 16.4. Threaded joint 385 16.5. Tongue-and-groove joint 385 16.6. Blind screw hole 386 16.7. Pressure vent limiting pressure buildup during internal explosion 387 16.8. Pressure vent assembly using metal-foam material 387 16.9. Typical slip-fit straight stuffing box and packaging-gland lead entrance 387 16.10. Typical slip-fit angle stuffing box and packing-gland lead entrance with hose clamp 387 16.11. Typical slip-fit angle stuffing box and packing-gland lead entrance 388 16.12. Typical plug for spare lead-entrance hole 388 16.13. Typical threaded straight stuffing box and packing-gland lead entrance with provision for hose conduit 388 16.14. Prototype trailing cable entry with polyurethane grommet 388 16.15. Insulated-stud lead entrance 388 16.16. Decision flow chart of class II, division 1 and 2 hazardous locations 394 17.1. Circuit modeling a dielectric 399 17.2. Current-voltage characteristics in a dielectric 399 17.3. Graph relating approximate insulation resistance variation with temperature for rotating machines 400 XXI ILLUSTRATIONS-Continued Page 17 A. Insulation resistance versus application time of test voltage 400 17.5. Megohmmeter test connections for checking cable insulation in line A .... 401 17.6. Megohmmeter test connections for ac motor 401 17.7. Megohmmeter test connections for dc motor 401 17.8. Spot resistance curve for normal motor 401 17.9. Spot resistance curve showing effects of dust and moisture 401 17.10. Spot resistance curve for detective motor 402 17.11. Megohmmeter test connections for transformer 402 17.12. Time-resistance curve 402 17.13. Three time-resistance curves for deteriorating motor 402 17.14. Time-resistance curves showing polarization for hypothetical motor 403 17.15. Polarization factor curve for deteriorating motor 403 17.16. Multiple voltage curves for deteriorating motor 403 17.17. Circuit for harmonic tests 404 17.18. Power-factor versus voltage curves showing tie-up 404 17.19. Mounting techniques for two vibration transducers 405 17.20. Four typical vibration measurement points 405 17.21. Typical vibration severity chart 405 17.22. Comparison of acoustic-emission techniques for detecting failing roller bearings 406 17.23. Conceptual diagram of generalized mine monitoring and control system 406 17.24. Conduction in gas 407 17.25. Discharge sequence in an ionizing field 407 17.26. High-stress geometries 408 17.27. Typical dielectric voids in cables 409 17.28. Block diagram for corona-detection system 409 17.29. High-voltage cable terminations 410 17.30. Major insulation void sometimes found in high-voltage coupler terminations 411 17.31. Possible stress site in high-voltage coupler insulators 411 17.32. Power-conductor transposition on three-conductor type G cable 412 17.33. Application of diode-suppression bridges in power center 412 17.34. Typical saturable-reactor characteristic 412 TABLES 2.1. SI symbols and units 21 2.2. Resistivity of some common materials at 20 C 22 4.1. IEEE device numbers and functions 90 4.2. Device numbers and letters common to mining 90 6.1. Motor voltage ratings common to mining 135 6.2. Motor insulation classes 135 6.3. NEMA class A standard starters for three-phase induction motors 141 6.4. Common motors for mining equipment 153 7.1. Current range and effect on a typical man weighing 150 lb 161 7.2. Typical resistances for various contact situations 162 7.3. Approximate resistance formulas for various electrode configurations 170 7.4. Comparison of grounding grids with other types of electrodes 172 7.5. General resistivity classification 174 7.6. Variations in resistivity with geologic age 174 7.7. Typical values of resistivity of some soils 174 7.8. Variation in soil resistivity with moisture content 175 7.9. Typical potentials of metals in soil measured from a copper and copper sulfate reference electrode .... 178 XX11 TABLES-Continued Page 8.1. Conductor sizes and cross-sectional areas 184 8.2. Letters used in alphabetic cable code 187 8.3. Codes for typical cables used in mining 187 8.4. Typical diameters for round portable power cables 193 8.5. Typical diameters for flat portable cables 193 8.6. Specifications for trailing cables longer than 500 ft 195 8.7. Ampacities for portable power cables 196 8.8. Ampacities for three-conductor mine power cables 196 8.9. Correction factors for ampacities at various ambient temperatures 196 8.10. Ampacity derating factors for 60 C-rated trailing cables operated on drums 197 8.11. Australian specifications for ampacity derating factors for trailing cables operated on drums 197 8.12. Some estimated power factors and load factors for various underground coal mining equipment in good operating conditions 198 8.13. Intermittent-duty ratings for trailing cables 199 8.14. Resistance and reactance of portable power cable 200 8.15. Resistance and reactance of mine-power-feeder cable 201 8.16. Solid-wire breaking strength 202 8.17. Recommended minimum bending radius, unshielded or unarmored cables 204 8.18. Recommended minimum bending radius, shielded and armored cables 204 8.19. Trolley-wire specifications 212 8.20. Characteristic data for solid copper feeder cable 213 8.21. Characteristic data for stranded copper feeder cable 213 8.22. Trolley-wire support spacings on curves 215 8.23. Resistance of steel rail at 20 C 215 8.24. Data for rail-bond cable 216 8.25. Minimum vertical conductor clearances as specified by the NESC, applicable to mining and mining-related operations 220 8.26. Minimum distances from overhead lines for equipment booms and masts 221 9.1. Ratings for mining-service molded-case circuit breakers 228 9.2. Interrupting-current ratings versus system voltage 229 9.3. Maximum instantaneous-trip settings 230 9.4. Commonly available magnetic-trip ranges for mining- service molded-case breakers 230 9.5. Some typical ratings for low-voltage power circuit breakers 232 9.6. Typical minimum-oil circuit breaker ratings 234 9.7. Ratings of high-voltage power fuses 239 9.8. Common current ratings of induction-disk overcurrent relays 243 9.9. Standard burden for current transformers 246 9.10. Standard ratings for potential transformers 261 10.1. Sample reactances for synchronous and induction motors 261 10.2. Three-phase transformer per-unit impedances for liquid-immersed transformers 262 10.3. Three-phase transformers impedances for distribution transformers, including load centers 262 10.4. Sample applications of fault calculations 263 10.5. Impedance of cables in figure 10.4 265 10.6. Burdens of relay elements and ammeter connected to CT's 273 10.7. Recommended instantaneous trip settings for 480-, 600-, 1,040-V three-phase trailing-cable protection . . 275 10.8. Recommended instantaneous trip settings for 300- and 600- Vdc trailing-cable protection 276 11.1. Recommended station and intermediate surge arresters for resistance-grounded mine power systems to protect oil-immersed transformers 294 XX111 TABLES-Continued Page 11.2. Recommended distribution-class, RM-type, surge arresters for resistance-grounded mine power systems to protect rotating machinery and dry-type transformers 294 11.3. Commonly used surge capacitors for limiting voltage rate of rise on rotating machinery and dry-insulated transformers 295 11.4. Typical capacitances per phase of power-system components, for shielded power cable SHD, SHD-GC, and SHD+GC 297 11.5. Typical capacitances per phase of power-system components 297 11.6. Protective angle versus structure height 299 12.1. Typical current ratings of 400-A load-break switch 305 12.2. Typical ratings for combination power centers 321 13.1. Standard impedance for liquid-immersed three-phase transformers 335 13.2. Standard BIL's for oil-immersed power transformers 337 14.1. Typical electromechanical and static relay characteristics 358 14.2. Time-margin comparison between electromechanical and static relays 360 14.3. Comparison of induction-disk and static time-overcurrent relay burdens to a current transformer 361 15.1. Formulas to estimate hydrogen evolution 373 16.1. Structural gap dimensions for explosion-proof enclosures as specified by 30 CFR 18 385 16.2. Minimum autoignition temperatures versus layer thickness for bituminous coals 393 17.1. Common causes of vibration 405 MINE POWER SYSTEMS By Lloyd A. Morley 1 ABSTRACT This Bureau of Mines publication presents a comprehensive review of mine elec- trical power-system theory and practice. It discusses fundamental theory and the vital aspects to be considered in planning and designing mine electrical power systems. The report is divided into three major sections. The first presents the history of electricity in mining and the fundamentals of electrical phenomena and components. The second focuses on power-system components: motors, grounding systems, cables, and protec- tive equipment and devices. The final section includes mine power-center equipment, switchhouses and substations, batteries, and mine maintenance. 'Professor of mining engineering, The Pennsylvania State University, University Park, PA (now professor and department head, mineral engineering, University of Alabama, Tuscaloosa, AL). CHAPTER 1 .—ELECTRICAL POWER IN MINING Probably no other mining area has grown so rapidly yet been as little understood by the average mine worker or operator as the mine electrical power system. Traditionally, the field has held little interest for the mining engineer, who has tended to avoid it, or for the electrical engineer, who has given it scant attention. But today's mine power system is both complex and subject to numerous legal con- straints, and it is no longer possible to treat it with the indifference of the past. Underground mining machines are among the most compact and rugged equipment over designed, and individ- ual units can have up to 1,000 total horsepower. Mining equipment is usually mobile and self-propelled; most is powered electrically through portable cables and, for safety, must be part of an elaborate grounding system. The ma- chines and power-distribution equipment are seldom sta- tionary, must be adapted to continuous cyclic operation, and must resist daunting levels of dust and vibration. Surface mining can involve the largest earth-moving equipment built, where one piece can have 12,000 or more connected horsepower— the largest today is over 30,000 hp. The electrical loads created by this machinery are cyclic and extremely dynamic: the largest excavator, for exam- ple, can require electrical loads that range from 200% motoring to 100% generating every 50 to 60 s, under the most exacting physical conditions. In the ever-moving min- ing operation where distribution of power must be con- stantly extended and relocated, subjected to abuse by machine and worker alike, the potential for safety hazards is always present. Engineering and maintaining such an electrical system is demanding and challenging. It requires a specialist with knowledge of both mining and electrical engineering. Yet conversely, the effective management of a mine requires that anyone responsible for production and safety also be conversant with the mine electrical system. Management should understand the advantages and disadvantages of one system over another, for if the power system is poorly de- signed, not only will safety be compromised but the mine operator will pay for the resulting conditions with high power bills, high-cost maintenance, and loss of production. Too often, a new mine is designed to use the type of power system employed in the preceding mine, without a comprehensive power study to determine the system needs and examine the alternatives available. Problems arise in existing mines when new mining equipment has been adopted without due regard for its impact on the operating power system; these problems haunt the mine electrical engineer who must frequently cope with a system that is a mongrel, bred from diverse inheritances from the past combined with recent changes and additions. New laws, standards, and safety requirements must frequently be accommodated by power systems not originally designed to meet their specifications; new and unfamiliar equipment must be grafted to the existing network, and the result can be a hybrid of considerable complexity. This text has been produced to assist the power engineer and the student in understanding these complexities and the principles that lie behind them. The material presented here is structured so an indi- vidual unfamiliar with electrical engineering can first develop the necessary fundamentals before embarking into mine electrical design. A basic physics and calculus knowl- edge is necessary to understand the content completely. The goal has been to assemble the most significant information required for comprehension of mine power systems so that the reader may then progress to more specialized topics. But first, a brief review of the development of electrical usage in mines is given, in order that the reasons for some of the peculiarities of mine power systems can be appreciated. MINE ELECTRICAL HISTORY Electricity was first introduced into coal mines shortly before the beginning of the 20th century in the form of di- rect current (dc) for rail haulage. This form of current was used because at that time most systems were powered by dc generators. It had a number of advantages for haulage; the most outstanding was that the dc series-wound motor had (and has) excellent traction characteristics. Speed con- trol was a simple matter of placing a resistance in series with the motor armature or field circuits. Batteries served as the first power source, and hence the vehicle was extremely mobile even though constrained on rails. However, keeping the batteries charged was both- ersome, so trolley wires were soon introduced in several mines. Allowing the trolley wire to act as one conductor and the rail as the other provided the simplest form of power distribution yet known to the mining industry. Available haulage machinery of that period was low in horsepower and the mines were relatively small so the increased resis- tance that reduced voltage and power supplied to the motors was still acceptable. Thus, the dc system at a voltage of 250 or 550 V became firmly entrenched in coal mines. Underground Mine History Underground, the first electrically driven coal mining machine, the coal cutter, was installed in the early 1920's. Although dc offered no special advantage, it was readily available; hence, the machine was equipped with a dc motor and added to the system. The cutter was followed almost immediately by the loader, and it too was driven by dc motors. If there was rail haulage in the mine, trailing cables supplied power from the trolley wire and the rail to the machines. The next significant increase in power consumption came with the introduction of the shuttle car, almost 20 yr after the coal cutter. Actually, when the shuttle car was first invented in 1937, it was battery powered. The addi- tion of an automatic reeling device to handle a trailing cable came later, in an attempt to overcome battery deficiencies. These trailing cables were also connected to the haulage power system, and this equipment, when combined with the cutters and loaders, placed additional stress on the dc distribution system. At that time, the horsepower required to operate each piece of electrical mining equipment was quite small and no individual machine used a large amount of current. However, when all machines were combined, significant power was required, and because all the conductors offered resistance, voltage drops and transmission losses in the distribution system were extensive. Alternating current (ac) would have been more practical because it could have been distributed easily at a higher voltage, thereby reducing cur- rent, voltage drops, and transmission losses. But many States had stringent limitations on maximum voltage levels, usually around 300 V, and with this restriction ac had no advantage over dc. Hence, dc continued to be used to operate the successful combination of cutters, loaders, and shuttle cars. Development in ac-to-dc conversion equipment played an important role in underground coal mine power utiliza- tion throughout this period. Motor-generators or synchron- ous converters were originally employed for conversion pur- poses, but in addition to being heavy and bulky, they could not be operated effectively in a gassy and dusty atmosphere, and maintenance requirements were substantial. As a re- sult, most conversion installations were placed on the sur- face with borehole connections to the underground mine. This was acceptable as most mines were then relatively shallow. In the 1930's, the same decade that saw the inception of the shuttle car, mercury-arc-ignition rectifiers began to be employed to provide dc underground. The arc tubes al- lowed more efficient use of electricity in deeper and larger mines than had previously been possible. As the tubes had no moving parts, maintenance was lower, efficiency was higher, and portability was improved. These rectifiers were usually centrally located in the mines because a liquid heat exchanger made them heavy and bulky. In this way, dis- tribution to the mine rectifier was ac, but distribution throughout most of the mine electrical system was still dc. At about the same time, some mines found that haulage of materials by conveyor could be more efficient than haul- age by rail. The conveyors were also powered by dc motors, and stress continued to be added to the electrical system. In the late 1940's, when continuous mining machines first began to be used extensively, dc was again expected to provide the power. However, the continuous miners nor- mally needed more energy input than the sum of the various conventional mining equipment they replaced, and because the required horsepower created high current demand, dc was found to be entirely unsatisfactory in most cases. The attendant current demand created enormous voltage drops in the distribution system. As a possible solution, the dc supply system was separated from the haulage system, but eventhis was unable to improve voltage regulation. During peak operation periods, voltages at the machines were so far below the values called for that even moderate efficiency was impossible. The increasingly large cable sizes required to supply the needed power created difficult cable-handling problems. The use of three-phase ac distribution and motors was an obvious answer, but for at least a decade some min- ing companies were reluctant to make the change. In many instances this was because the laws in some States limited maximum voltages in the mine. Lawmakers were convinced that high voltages were synonymous with high safety risks. Some State laws were not updated until the mid 1960's. When higher voltages were finally permitted, the de- sirable economics of ac employment could be realized and there was a swift transformation from dc to ac for both distribution and high-horsepower loads in underground coal mines. Unfortunately, many mine electrical systems were at least partially modified without concern for the compat- iblity of these changes with the remainder of the system, and various safety and production problems arose. As a result of conversions, mine power systems gener- ally had two voltage levels, one for distribution and one for utilization. The simplified mine electrical arrangement shown in figure 1.1 illustrates the results. Here, the sub- station transforms the utility voltage down to distribution levels, which are most often at high voltage greater then 1,000 V. Power at this voltage is transmitted or distributed through conductors from the substation to the power center; hence, this system is called the distribution system. The power center or load center, actually a portable substation, transforms the voltage to utilization levels, which are typically at low voltage of 660 V or less, or medium vol- tage of 661 to 1,000 V. At this level, or face voltage, power is normally delivered to the equipment. Despite this ref- erence to voltage levels, it should be noted that distribu- tion and utilization describe functions of a power system segment, not specific voltage ranges. Originally, primary ac distribution was made at 2,300 or 4,160 V. In most mines, these levels were later increased to 7,200 V. Some operations recently increased the voltage to 12,470 or 13,200 V for both longwall and continuous- mining applications. Each new distribution voltage, it may be noted, is a factor of \pi times the previous value (\/ r 3(2,300) = 4,160). The principal reason for increasing the voltage was that, for the same load, current would be cor- respondingly smaller, and lower distribution losses would result even though the same conductor sizes were used. From the beginning, 440 Vac was the most popular voltage for utilization, despite the fact that when the con- tinous miner proved so successful its horsepower was pro- gressively increased, following the sometimes misguided notion that a directly proportional increase in coal produc- tion would follow. As with dc, the additional horsepower resulted in an increase in trailing-cable sizes, until the weight of the cables was almost more than personnel could handle. To compensate, the most common move was to raise the rated motor voltages to 550 Vac. More recently, manu- facturers have produced machines with 950-V (550 \/15) motors to further overcome the trailing-cable problems. While these changes were being made to ac for machine operation and distribution, the use of dc for haulage con- tinued to be advantageous. In the early 1960's, silicon diode rectifiers with large current capabilities became available. Simple, efficient, and small, these rectifiers were ideally suited for use underground and made ac distribution possi- ble for the entire electrical system except rail haulage. Through the use of rectifiers, the benefits of dc for traction and of ac for distribution and utilization on high power loads could be realized. For example, while continuous miners normally used ac, part of the supply at the power center was rectified to dc, primarily for powering the shuttle cars. These underground electrical systems appeared to be simple, and as a result they did not become the focus of at- tention for some time. Systems were frequently designed and maintained by a "seat-of-the-pants" approach, to the point that ac distribution and equipment were installed source Substation Switchhouse Power center o- Distribution voltage To the loads Utilization voltage Transmission Distribution Utilization Figure 1.1. —Simple mine electrical system arrangement. employing dc concepts. However, ac systems are more com- plicated than dc systems and call for meticulous planning; if wrong decisions are made, the results can be extremely costly in terms of safety, production, and economics. A great deal of effort is needed to maintain an electrical power sup- ply within the requirements of the individual pieces of mining equipment, and mixing ac and dc in the same mine has added greatly to the problems. This brief review of the development of electrical sys- tems in underground coal mines has shown that the mines went from minor electrical usage with the introduction of rail haulage to almost total dependency on electricity in a period of 50 yr. In the same period, surface coal mining underwent changes that were as substantial if less numer- ous. They were centered around the enormous growth in equipment size. Surface Mine History The first mechanization of strip mining occurred in 1877 with the application of an Otis-type steam shovel in a Pitts- burg, KS, mine (5). 1 This early attempt was somewhat un- successful, but it served as an important step in the evolu- tion of strip-mining machinery. Several successful attempts to use steam shovels and draglines were made in the next 30 yr, and these proved that the surface mining of coal was completely practical. In time, the advantages of electricity over steam became more apparent, and the first significant introduction of electric-powered shovels was made in the early 1910's. Whereas dc series motors were universally employed in underground rail haulage, the first large motors used in surface mining were dc shunt wound because of their constant-speed characteristics. These motors almost directly replaced the single-speed steam engine found on practically all shovels prior to that time and allowed an immediate reduction in work force requirements. Before long another important advance in shovel design occurred: the applica- tion of separate steam engines to power the shovel motions of hoist, crowd, and swing. This change gave increased flexibility through the individual control of each operation. In a short time, the two major shovel manufacturers of that era, Marion and Bucyrus, began to produce both steam and electric multimotor shovels with similar characteristics (5). Since series-wound dc motors had speed-torque relationships similar to those of steam engines when they were used for this type of duty, they were employed to drive each shovel motion. The initial distribution for electric shovels was dc be- cause of the nature of the power generation, but technolog- ical advances soon made ac power systems superior, and ac motors were tried with some success. However, by 1927, ac-dc motor-generator (m-g) sets and the invention of the Ward-Leonard control concept caused these efforts to be abandoned. The new control system enabled the motor characteristics to be modified as desired within the motor and generator commutator limits, and as a result, separately excited dc motors became more attractive than series-wound motors. The m-g sets functioned as on-board power-conversion units, thereby establishing the use of ac distribution in surface mines. Motor-generator sets driven by synchronous or induc- tion ac motors, Ward-Leonard control, and separately ex- 'Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. cited dc motors established the standard, and even now the combination is used on most mining excavators, especially the larger varieties (2). On smaller machines, some single ac electric-motor drives with either mechanical-friction or eddy-current clutch systems have evolved, but these are often driven by diesel engines. Present excavating equipment is generally classified into three size groups, although actual capacity ranges are normally not assigned. Small shovels are used primarily in general excavation, while the intermediate types work at bench mining and coal production, and large shovels han- dle overburden stripping. Draglines of all sizes are used only for stripping. Small and intermediate equipment originally ran on rails, but crawler mountings that give improved mobility made their first appearance in 1925 (5). Today, small and intermediate-sized shovels and draglines are mounted on two crawlers, while large shovels have eight crawlers (2). Large draglines and some intermediate sizes are usually walking types that feature a circular base or tub that provides low ground-bearing pressure and a walk- ing device for mobility. The design of surface mine drilling equipment paral- leled excavator development. Initially, most drills were pneumatic-percussion types, but because electricity was readily available in mines, some machines were designed with internal motor-driven compressors. By the early 1950's, large rotary drilling equipment was necessary to satisfy the blasting requirements of thick, hard overbur- den (5). This drilling equipment was again electrically powered and was very successful. The most outstanding change that has taken place in electrically powered surface mining equipment has been in connected horsepower. For example, a 25-yd 3 dragline or stripping shovel that had a maximum total load of around 2,000 hp was considered enormous in the late 1940's (5, 9). By 1955, 50- to 70-yd 3 excavators were being manufactured with maximum horsepower at 4,650 hp. Five years later, shovels had reached a 140-yd 3 capacity wth 12,000 hp of main drive motors (7). In 1976, the largest excavator in service had 20,000 hp in m-g set drive motors (4). Distribution and utilization voltages also increased to keep pace with the peak load demands of this machinery. Sometimes the mine distribution and machine voltages for these excavators remained the same. Until the mid-1950's, 4,160 and 2,300 V were the usual mine levels (9). Then, with the advent of larger concentrated loads, 7,200 V was con- sidered advisable (10). However, this level was found to be unsatisfactory for the newly introduced machines with a capacity larger than 100 yd 3 , and so 13,800-V mine and excavator voltage became a standard. With machines hav- ing greater then 200-yd 3 capacity, 23,000-V utilization was established (4), but even with these substantial increases in distribution, some loads up to 1,000 hp continued to be driven at 480 V (10). Production shovels with loads up to 18 yd 3 commonly stayed at 4,160 and 7,200 V, while in general, 4,160 V became standardized for machinery with 1,500 hp or less. As a result, more than one voltage level could be required at a mine when excavators of different sizes were employed. MINE POWER EQUIPMENT A few pieces of power equipment have already been mentioned but only to the extent necessary to describe the concepts of distribution and utilization. The evolution of mine systems has resulted in major items of power appa- ratus, each serving a specific function (1,9). In general, they can be listed as • Generation, • Main substations, • Portable and unit substations, • Switchhouses, • Distribution transformers and power (or load) centers, and • Distribution (conductors and connectors). The following paragraphs explain these components only in sufficient detail for their inclusion in system ar- rangements to be understood. More detailed descriptions of substations, switchhouses, and power centers are pre- sented in chapters 12 and 13, while chapter 8 is devoted to distribution. Power generation is beyond the scope of this text, but Ehrhorn and Young (13) provide a thorough discus- sion of generation related to mining. Substations It is common mining practice to purchase all or most power from utility companies if it is available. As utility voltages usually range from 24 to 138 kV, a main (primary) substation is required to transform the incoming levels down to a primary distribution voltage for the mine. In ad- dition to having the transformer, substations contain a complex of switches, protection apparatus, and grounding devices, all having a function in safety. Main substations are often permanent installations. The nature of the min- ing operation and its power needs dictate how many main substations are required and where they should be placed. They may be owned by the utility or the mining company; the decision of ownership is commonly dependent on eco- nomics. However, if the total connected load is greater than 1,000 hp, mine ownership is often more favorable (13). Portable and unit substations are similar in operation to main substations except they serve to transform the primary distribution voltage to a lower distribution level. The term "unit" means that the substation and power equipment are designed and built as a package. In a typi- cal strip-mining deployment, a large dragline may require 24 kV while the production shovels and other mining equip- ment need 4,160 V. Switchhouses Switchhouses are portable equipment that protect the distribution circuits. Their internal components are chiefly protection devices, with circuit deenergization performed by disconnect switches, oil circuit breakers, or vacuum cir- cuit breakers. The switchhouse may contain more than one complete set of devices, for instance, a double switchhouse, which can independently protect two outgoing circuits. This category encompasses disconnect switches, which are power equipment containing only manual devices, with the prime function of allowing mine power to be removed from the main supply. Power Centers At the outermost distribution points there are power centers and distribution transformers, which transform and convert the distribution voltage to utilization levels. In- cluded here are ac to dc conversion equipment or rectifi- ers, which convert the distribution voltage to dc for use on rail trolley and other systems. The power center, also termed a load center, usually implies an internal bus, which is defined shortly, in the section, "Radial System." In essence, these are all portable substations, and as with switchhouses, each outgoing circuit has its own set of in- ternal protection components. However, an individual unit may supply from 1 to as many as 20 machines. Power centers can be considered the heart of an underground min- ing section power system. In surface mines, power centers supply power to low-voltage auxiliary machinery and loads; there may be no need for this equipment with the primary mining machinery. Distribution Equipment This category of major power equipment is often referred to as the weakest link in the mine power systems. It en- compasses all the overhead powerlines, cables, cable cou- plers, and trolley lines used to carry power and grounding between the power equipment and eventually to the loads. The conductors are usually called feeders when they are part of distribution; at utilization, when connected to portable mining machines, they are called trailing cables. BASIC DISTRIBUTION ARRANGEMENTS The basic distribution arrangements available for in- dustrial applications are radial, primary selective, primary loop, secondary selective, and secondary-spot network (6). Radial systems are the most popular arrangements in min- ing, though other configurations can be found where special circumstances call for greater system reliability (3). Sur- face mines have, of course, greater flexibility than underground mines and employ a wider range of distribu- tion arrangements. Secondary-spot networks, which are the most popular system for large facilities in other industries, are uncommon but could be applied to preparation and mill- ing plants. The following descriptions of the main distribu- tion patterns are based on the Institute of Electrical and Electronics Engineers (IEEE) definitions (6). This institute, the leading national professional electrical organization, sets standards and recommended practices that are inter- nationally renowned for their correctness. Radial System Figure 1.2 shows radial distribution in its simplest form. Here, a single power source and substation supply all equip- ment. The single vertical line represents one connection point for all feeders, or all connecting lines, and is termed a bus. Voltage along the bus is considered to be constant. Radial systems are the least expensive to install as there is no duplication of equipment, and they can be expanded easily by extending the primary feeders. A prime disad- vantage is tied to their simplicity; should a primary com- ponent fail or need service, the entire system is down. An expanded radial system, the load-center radial, is illustrated in figure 1.3. As in figure 1.1, two or more volt- age levels are established, but the feeders form a treelike structure spreading out from the source. This system has the advantages of the simple system and several others too. If the load centers or distribution transformers are placed as close as practical to the actual loads, most distribution Bus Substation O Utility source Switchhouse Switchhouses Utilization equipment Cable or powerline O O o Source Figure 1.2.— Simple radial distribution system. Substation Circuit breakers Utilization equipment Figure 1.3.— Power-center type of radial distribution. will be at the higher voltage. This allows decreased con- ductor investment, lower electrical losses, and better voltage regulation. Primary-Selective System The primary-selective system (fig. 1.4) adds downtime protection through continuity of service. Each substation can receive power by switching from either of two separate primary feeders. Each feeder should have the ability to carry about 80% of the load, so that one feeder can accept a temporary overload (10) and provide continued operation if one source should fail. During normal service, each feeder should handle one-half of the load. The system is simple and reliable but costs are somewhat higher than for the radial system because of the duplication of primary equipment. Primary-Loop System Though found in some mines, the primary-loop system (fig. 1 .5) is not considered good practice. It offers the advan- tages and disadvantages of the primary-selective system and the cost can be slightly less, but this configuration can Main Source substation Main Source su bs totio n O Unit substations Bus — *- Circuit breakers — *- To the loads Figure 1. 4.— Primary-selective distribution system. a Switchhouse Switchhouse Sources Substations O- Loop feeder Figure 1. 5.— Primary-loop distribution. result in dangerous conditions when a primary feeder fails. For instance, a failed portion can be energized at either side, creating an extreme hazard to maintenance personnel. Secondary-Selective System In a secondary-selective system, a pair of substation secondaries are connected through a normally open tie cir- cuit breaker as shown in figure 1.6. The arrangement al- lows greater reliability and flexibility than do the preced- ing techniques. Normally, the distribution is radial from either substation. If a primary feeder or substation fails, the bad circuit can be removed from service and the tie breaker closed either manually or automatically. Mainten- ance and repair of either primary circuit is possible with- out creating a power outage, by shedding nonessential loads for the period of reduced-capacity operation. Other methods that can be used to provide continuity of service include oversizing both substations so that one can carry the total load, providing forced-air cooling to the substation in serv- ice for the emergency period, or using the temporary over- load capacity of the substation and accepting the loss of com- ponent life (6). Economics often justify this double-ended arrangement if substation requirements are above 5,000 kVA. Note that the substation capacity or ability to trans- form power is rated in kilovoltamperes. Secondary-Spot Network In the secondary-spot network, two or more distribution transformers are supplied by separate primary-distribution feeders as illustrated in figure 1.7. The secondaries are tied together through special circuit breakers, called network protectors, to a secondary bus. Radial secondary feeders are tapped to the bus and feed the loads. This arrangement creates the most reliable distribution system available for industrial plants. If a failure occurs in one distribution transformer or primary circuit, perhaps by acting as a load to the bus, its network protector can quickly sense the reverse power flow and immediately open the circuit. Total power interruption can occur only with simultaneous mis- haps in all primary circuits or a secondary-bus failure. However, this type of system is expensive, and the relia- bility gain is not warranted for the majority of mining applications. It may be obvious that these basic distribution tech- niques can be combined into hybrid systems. When this is done, there can be confusion about what is primary or secondary. Ordinarily, the subsystems are defined by the specification of the substations. This will be demonstrated in the next two sections. UTILITY COMPANY POWER As utility companies are the principal power source for mines, an understanding of utility system power transmis- sion and distribution is important. Often this system greatly affects the power available to the mine, including voltage regulation, system capacity during power failures in the mine, and overvoltage occurrences. o Sources Main substations Switchhouses Normally open tie breaker o Switchhouses T TT In a nearby substation, power from a generating sta- tion is transformed up to a transmission voltage, commonly 69,000 V or more (6). This power is carried on transmission lines to major load areas, either supplying large industrial users directly or powering the utility's own distribution substations. Distribution substations step the voltage down, this time to a primary-distribution level ranging from 4,160 to 34,500 V, but most often at 12,470 or 13,200 V (6). These stages are illustrated in figure 1.8. The utility service, therefore, can be any of the follow- ing standard values, in kilovolts: 138, 115, 69, 46, 34.5, 23, 13.8, 12.47, 6.9, 4,800, 4.16, and 2.4 (13). Generally, the delivered voltage ranges from 23 to 138 kV, but other values such as 480, 2,300, and 7,200 V are also found. What is available to the mine depends on whether the possible con- nection is to the power company transmission system, a primary-distribution system, or a distribution transformer. It is the responsibility of the mining company to select that voltage best suited to its needs. Primarily, the choice depends on the amount of power purchased. It is not safe to assume that the power company has the capability to serve a large mine complex from existing primary- distribution lines or even from the transmission system. The problem stems from the fluctuating nature of mine loads. For example, large excavators in surface mines can require high peak power for a short time, followed by regenerative peak power, cycling within the span of 45 s. The fluctuat- ing load may create voltage and frequency variations be- yond the limit set for other utility customers. Accordingly, Sources o o o o Substations Network protectors Bus Switchhouse To the loads Figure 1. 7.— Secondary-spot network technique. Generating station C^ Transmission line Regulated primary / distribution system Substation Substation Secondary distribution Distribution transformer Figure 1. 6.— Secondary-selective system. Branch circuit Utilization equipment Figure 1.8.— Representative utility transmission and distribution. most large draglines and shovels require power from 69- to 138-kV transmission systems to get adequate operational capacity, and the construction of several miles of trans- mission equipment can result in a sizable cost for the mine budget. Regardless of where the main mine substation is tied into the utility complex and who owns that equipment, its outgoing circuits will here be termed the mine primary- distribution system or just distribution. The incoming power will most often be referred to as a transmission system. The following sections identify the main types of mines in the United States, classify the major equipment em- ployed, and describe the power-distribution arrangements that are found in them. Of necessity this can be only a very brief overview, but it is designed to indicate the problems and complexities that can arise in mine power distribution and utilization. Individual topics mentioned here are ex- panded in detail in later chapters. SURFACE MINING Surface mining methods are selected over underground methods when the overburden, the earth above the coal seam, can be removed economically to expose the coal. Productivity, safety, and economics usually favor surface mining of seams less than 150 ft deep. Surface coal mining consists of four basic operations: overburden removal to ex- pose the coal, coal loading, haulage, and reclamation. The mining method is generally classified according to such physical characteristics as topography or land contour, over- burden thickness, coal thickness, number of coal seams, type of overburden, fragmentation characteristics of the over- burden, climate, and hydrology. The mining method is also affected by Federal and State requirements. The mining method selected must protect the health and safety of the workers and minimize environmental disturbance and be designed for the specific set of prevailing physical condi- tions. The major surface mining methods for coal are con- tour mining and area mining. Contour mining methods are commonly used in rolling or mountainous terrain; they are called contour mining because overburden removal progresses around the hillside at the coal seam horizon such that the pit resembles a con- tour line. There are many varieties of contour mining, but in all methods overburden is fragmented by drilling and blasting, and removed to expose the coal seam. The over- burden may be removed by small diesel or electric drag- lines, or by diesel-powered front-end loaders and trucks. In soft overburden and for topsoil removal, scrapers and bull- dozers may be used. Area mining is the predominant stripping method in more level terrain. As its name implies, area mining can cover an extensive region, using various box-cut or strip pit and benching techniques. It may be used to mine both thick and thin seams, or multiple seams; where these seams are dipping, area mining is modified to approximate the open pit methods common in metal mining. In all cases, over- burden handling and reclamation are an integral part of the process. Equipment varies, depending on the scale of the operation, from small draglines and dozers to massive equipment that has more than 30,000 connected horsepower. In general, the magnitude of electrical distribution and utilization is greater in area mining than in contour min- ing. Combination of equipment employed in large multi- seam operations may include tandem draglines, dragline and shovel, pan scrapers with attendant dozers, and drag- line and bucket-wheel excavators. Bucket-wheel excavators can be very effective where overburden is soft and does not require drilling and blasting. Front-end loaders, electric and diesel shovels, ripping dozers, and tracked highlifts can all be combined with truck haulage for coal removal. POWER SYSTEMS IN SURFACE MINES Mine power systems can be divided into three categories, depending upon the purpose of a specific portion: 1. Subtransmission, 2. Primary distribution (or distribution), and 3. Secondary distribution (or utilization). Often, if a subtransmission system is needed, it will have the same general arrangement in any mine. At distribu- tion and utilization, power-system installations can vary greatly, but in some mines distribution and utilization can be the same system. Electrical installations in surface coal mines are regulated under 30 CFR 77 (14). Main Substations and Subtransmission Main substations may range from 500-kVA capacity, supplying 480 V for only pumps and conveyors, to 50,000 kVA, servicing a large strip-mining operation and prepara- tion plant (10). The substation location is usually an eco- nomic compromise between the cost of running transmis- sion lines and power losses in primary distribution. From the main substation, power is distributed to the various centers of load in the operation. However, individual loads or complexes, such as preparation plants and other surface facilities, may have large power requirements or be so iso- lated that primary-distribution operation is not practical. In these cases, or for safety reasons, incoming utility trans- mission should be extended close to the load. The extension is designated a subtransmission system, and the conductors are usually suspended as overhead lines (13). As shown in figure 1.9, subtransmission commonly re- quires a primary switchyard of high-voltage switching ap- paratus for power tapping. Branch circuits are fed through Incoming transmission lines Y Primary switchyard t Normally open L j J Dual tie breaker / bus i — i Subtransmission Preparation plant substation To preparation plant 1^ Second -< — \y subtransmission if primary selective or secondary selective desired on major load concentration To mine distribution Figure 1. 9.— Subtransmission for surface mine. circuit breakers to protect the subtransmission line and the utility's system. Dual-bus configurations are employed if primary-selective or secondary-selective distribution is de- sired on major load concentrations to provide high reliabil- ity. This additional subtransmission circuitry is illustrated in figure 1.9 by dashed lines. Subtransmission circuits, primary switchyards, and main substations are almost always permanent installa- tions located in areas unaffected by the mining operation. The main substation is where the grounding system for the mine is established. This ground is carried along the powerlines through overhead conductors or in cables and is connected to the frames of all mobile mining equipment. Surface Mine Distribution Mine power distribution, in its simplest radial form, has already been shown to consist of a substation, distribution, and a power center feeding the mining equipment. The ar- rangement is very common in small surface operations where the distribution voltage is commonly 4,160 V but can be 2,300 V in older equipment. In the smallest mines, power is purchased at low-voltage utilization (often 480 V) and fed to a distribution box to which motors and equip- ment are connected. At times, simple radial distribution is employed in large surface mines where only one machine must be served or an extensive primary-distribution net- work cannot be established, as in some contour operations. The great majority of strip mines employ radial distribu- tion, but secondary-selective and primary-loop designs can also be found. Simplified examples of the three systems are provided in figure 1.10 to 1.12. In all configurations, a por- tion of the primary distribution is established at a base line or bus. The base line is usually located on the highwall, paralleling the pit for the entire length of the cut. Its loca- tion is typically maintained 1,500 ft ahead of the pit, and it is moved as the pit advances (3). Distribution continues from the base line to the mining equipment, with the con- nections maintained at regular intervals. As the machines move along the pit, the base-line connections are changed to another convenient location. The base line can consist of overhead polelines or a cable-switchhouse configuration, figures 1.13 and 1.14 (3). It can be seen that cable distribution brings power into the pit area, where shielded trailing cables connect to the ma- chines. The overhead poleline plus cable arrangement is common in older mining operations, especially when utiliza- tion is at 7,200 V or less (3). Typical spacing between poles, or line span, is 200 ft. Drop points are noted in figure 1.13 by triangles. These are terminations between the overhead conductors and the cables, mounted about 8 ft above the ground on poles spaced at regular intervals of around 1,000 to 1,500 ft. Cables connected to the drop points deliver power to skid-mounted switchhouses located on the highwall or in the pit. The switchhouses may contain manual disconnect switches, which are commonly termed switch skids or disconnect skids, automatic circuit-protection devices or breaker skids, or a combination of both. The skids can either be boat design with flat bottoms or have fabricated runners, depending upon the allowable bearing pressure of the mine terrain. Couplers or plug-receptacle pairs are commonly used for both feeder and trailing-cable connections. Discon- nect and circuit-protection functions are required for each distribution load, and double switchhouses (two-breaker skids) are frequently employed for two loads. Unit substa- tions often contain internal circuit protection on the incom- ing side, and thus do not require a breaker skid. Trailing cables are usually 1,000 ft in length, although lengths to 2,000 ft can be found. When longer cables are necessary to reach a breaker skid, in-line coupling systems can be used, and these are commonly mounted on small skids for easy movement. Trailing-cable handling for strip- ping equipment is often assisted by cable reels mounted on skids or self-propelled carriers. Large excavators can require the self-propelled variety. Switchhouses 1_ Unit substation Base line ■*■ Pit highwall 9~Q h^H -T] [I ►Base line Dragline Production shovel Pump, lighting Figure 1.10.— Radial strip mine distribution system. Main substation I Base line — a— Lp Normally open tie breaker / I Main substation . Switchhouses Pit highwall Production shovels Other pit power Figure 1.11.— Secondary-selective distribution in strip min- ing. Main substation Figure 1.12.— Primary-loop design for strip mining. 10 Utility company metering 7. 2- kV overhead poleline ( base line ) 1,000 ft /1,000 ft 1,000 ft 1,000 ft •\ 1 — \ — i *\/ rAr Power company supply, 69 kV Substation 69kV/7.2kV) V V V Lateral cables Trailing cable Dragline To auxiliary equipment Production shovel Drill 7.2-kV to 4,160-V trans- former^ TBS Htbs r— ' Trailing cable Dragline Production shovel Spare KEY DS Disconnect switch TBS 2- breaker skid Lateral 'Cable 7.2-kV to \4,160-V ■ trans- 1 former v TBS Drill Figure 1.13.— Radial distribution for strip mine with overhead-poleline base line. Utility company supply, 138 kV Substation (!38kV/7.2kV) Additional J distribution -^ if pullback mining method is used 7.2-kV cable by incline to pullback machine 138-kV powerline ] Utility company metering Overload circuit breaker, 138 kV Substation (138kV/25kV) 25-kV cable Skid-mounted triple - switch skids Single-breaker skid Dragline 25 . kV \ (transformed trailing cable i — i to 7,200V - — — I \~ 1 1 on machine ) Lateral cables/ 2-breaker skid r? t Dragline ! Single-breaker skid Spare ■* 25 kV/ 440 V Auxiliary 440-V coble tl 1,500 ft 25-kV ^Of'Kcoble Production shovel equipment 2-breaker skid Drill- Spare- 1 o- Lateral cables-^ 1 7.2-kV cables 25-kV/ 7.2-kV transformed Drill -. •1,500 ft 1,500 ft Trailing cables Production shovel — — 2-breaker skid Figure 1.14.— Radial distribution for strip mine with all-cable distribution. The layout for an all-cable mine distribution, figure 1.14, is very similar to that just described. In this case, how- ever, the base line is assembled using cable-interconnected switchhouses. As noted in the illustration, the common approach is to use disconnect skids with three internal switches in the base line and to have separate breaker skids in line with the cables feeding the machinery. Another ap- proach is to combine the single-breaker skids into the base- line switchhouses. When a secondary-selective configuration is used, as shown in figure 1.11, a normally open tie circuit breaker is placed in the base line in a location approximately equi- distant from the main substations. In some operations, the two substations and the tie circuit breaker may be in the same location, with two feeders running from the substa- tion area to the base line. More than two main substations may be established in very large operations. Primary-loop systems have occasionally been used in strip mining. It can be noted from figure 1.12 that the sub- stations actually operate in parallel, considering the base line to be a bus. Here the substations can be smaller than those needed for a radial system. Notwithstanding, certain precautions should be taken with this configuration (9). For example, the substations must be identical if they are to share the load, but as an unbalanced load distribution is probable on any system, it is likely that the two substa- tions will not be equally loaded. Regardless, because of the safety hazards, primary -loop distribution is considered un- satisfactory and is not recommended. Distribution voltage for the surface mine may be 7.2, 13, or 23 kV, and to a lesser extent 4.16 kV. Regardless of the level, drills and production shovels usually operate at 7,200 or 4,160 V. Therefore, when higher distribution levels are needed, portable unit substations are commonly used in the pit. One instance would be when the load cre- ated by a large machine is several times that for auxiliary machines. Another method is to establish two base lines on the highwall for two distribution voltages, as shown in figure 1.15. Here, a large unit substation interconnects the two base lines. Even in this situation, as can be seen in the preceding illustrations, low-voltage unit substations Utility Main substation CbFD — a — — D Switchhouse for 23-kV base line Switchhouse for 7,200-V base line Dragline 10,000-ft pit Figure 1.15.— Surface mine distribution system using two base lines. 11 or power centers are often required for 480-V auxiliary equipment. The primary purpose of any primary-distribution scheme in a surface mine is to provide a flexible, easily moved or modified power source for the highly mobile mining equipment. System designs must also be considered as an integral part of the total mine operation. Those described have these objectives in mind. As will be seen, the distribution system in any surface or underground mine that serves portable equipment is subject to damage from the mining machinery itself, and as a result, the system must be designed with optimum flexibility and considera- tion for personnel safety. Open pit power systems are quite similar to those in stripping mines but with one main exception: primary dis- tribution typically establishes a ring bus or main that partially or completely encloses the pit. Radial ties to the bus complete the circuit to switchhouses located in the pit, and portable equipment again uses shielded trailing cables. An example is shown in figure 1.16. Distribution voltage is normally 4.16 kV, but 7.2 or 6.9 and 13.8 kV are some- times used. Unit substations are employed if equipment voltages are lower. Primary distribution is almost invari- ably through overhead lines. UNDERGROUND COAL MINING Figure 1.17 is a plan view of a typical U.S. underground coal mine. A system of main entries, each 16 to 20 ft in width, is developed from the coal seam access point to the production areas, which are called panels or sections. Pil- lars of coal are left during mining to support the overburden above the entries. Crosscuts are mined between the entries. The main entries may remain standing for several years while coal is being extracted from the panels. Haulage of Utility Substation Disconnects \ Ultimate pit limit Substation Utility or subtransmission Overhead line ring main Poles Substation Utility or subtransmission personnel, supplies, and coal, together with provision of ventilating air and dust-suppression water, and electrical distribution are necessary functions of the main entries throughout the life of the mine. The mining method is defined by the configuration of the open workings and by the classification of equipment used. The important underground coal mining methods are room and pillar, which may be either conventional or con- tinuous, and longwall. To the miner, the type of mining machinery used is implied by each category. The room-and- pillar method remains dominant in the United States, although there has been a recent substantial increase in longwall mining. The choice of a specific mining method is frequently dictated by such natural conditions of the mine as the characteristics of the overburden, roof, and floor, plus the seam dip, water, methane, and seam height (11). Es- sentially, the method and equipment selected are based on the combination that will provide the safest and most prof- itable extraction within the given set of geologic conditions, while complying with State and Federal health, safety, and environmental regulations. Room-and-Pillar Mining Room-and-pillar mining is named for the regular pat- tern of openings made in the coal seam and was the earliest form of underground coal workings. Conventional Mining The conventional mining method represents a direct evolutionary link with the early mining techniques. It is based on the original loading machine, which came into use Pillars 84 ft Previously mined longwall panel, roof caved (Gob) Room-and-pillar panels developed Panel entries' Crosscuts ' Longwall face Submain entries Main entries Figure 1.16.— Open pit power system. Figure 1.17.— Layout of underground coal mine. 12 in the early 1920's. Modern conventional mining consists of six distinct operations: 1. Undercutting the coal face, 2. Drilling holes in the face for blasting, 3. Blasting, 4. Loading the broken coal onto a face haulage system, 5. Hauling the coal from the face area to a subsequent haulage system, and 6. Providing roof support. In order, these steps comprise a mining cycle; after roof sup- port is completed, work begins again at step 1. Ventilation, although essential, is not included as a separate step in the cycle as it must be provided continuously. Other safety pro- cedures include careful examination of the face and roof after blasting and before each job begins at the face. Mobile self-propelled mining equipment performs most of the operations in conventional mining. The cutting ma- chine, basically an oversized chain saw, is employed to cut a slot at floor level, called the kerf, which allows coal ex- pansion during blasting. A self-propelled face drill follows the cutter and makes several holes in the face with its carbide-tipped auger-type drill bits. Blasting is carried out either by chemical explosives approved as permissible by the U.S. Mine Safety and Health Administration (MSHA) or to a lesser extent by high-pressure air. Permissible ex- plosives will not ignite methane and coal dust when used correctly. A crawler-mounted loading machine loads the broken coal onto the face haulage vehicle, typically a shuttle car. The car is equippped with a chain conveyor that moves the coal from the load end to the discharge end and subse- quently unloads it from the vehicle. Shuttle cars almost invariably work in pairs and move the coal to rail cars or a conveyor belt, which makes up the next stage of materi- als handling in the mine. The roof bolter, sometimes called a roof drill, is a rubber-tired vehicle that secures the roof by first drilling vertical holes and then emplacing roof bolts, which secure the roof either by clamping thin roof layers together to form a thick beam, or by hanging weak strata to a more competent upper layer. Drilling is usually ac- complished by rotary action with auger-type bits. The re- sulting dust is collected through the bit and hollow drill rod by vacuum. With few exceptions, all these machines are electrically driven, powered via trailing cables from the mine power system. Since the mining equipment is continually moved among several faces in a coordinated plan designed for maximum production efficiency, the handling of trailing cables is a significant part of the mining cycle. The result of coal removal is a system of open rooms divided by coal pillars that support the roof as mining advances toward the property boundaries. When the equipment approaches the property limit, the operation is turned around and retreat mining takes place. If surface subsidence is permitted, the pillars are removed in an organized extraction plan and the roof is allowed to cave. The broken material that then fills the mined void is known as gob. Continuous Mining The heart of the continuous coal mining method is the continuous mining machine, which replaces the conven- tional room-and-pillar unit operations of cutting, drilling, blasting, and loading. The mining functions of haulage and roof support remain, although some continuous miners also perform roof bolting. The term "continuous" is actually a misnomer because of legal constraints that mandate inter- ruptions in the mining process for safety checks and ven- tilation requirements. The most common form of face haulage in continuous mining is again the shuttle car. One of the main problems associated with continuous mining is the intermittent nature of the shuttle car haulage system, which causes fre- quent delays at the face. As a result, various types of con- tinuous haulage systems have been developed to alleviate this problem. Mobile chain and belt conveyors are the most popular of these systems, and these are applicable to min- ing low coal. Continuous haulage systems have not been without problems, and some designs have been hampered by poor reliability and lack of maneuverability. Hydraulic systems have shown great promise; they operate by pulver- izing and slurrying the coal immediately behind the miner, then pumping the slurry to the surface. Longwall Mining Longwall mining is the most popular underground coal mining technique in Europe, and it is growing rapidly in the United States. In contrast to room-and-pillar mining, longwall is capital intensive rather than labor intensive. Longwalls are usually 300 to 600 ft wide, and the direction of mining with respect to the main entries classifies them as either advancing or retreating longwalls. The latter is the most frequent in the United States. A typical retreating longwall is shown in figure 1.18. The section of coal to be mined, the longwall panel, is first delineated by two room-and-pillar entries or headings driven perpendicular to the main entry. These two head- ings, the headgate and the tailgate, handle haulage equip- ment and ventilation. The longwall panel is then mined back and forth, retreating toward the main entry. The roof is allowed to cave immediately as the longwall equipment moves, as is shown by the gob area on the diagram. The longwall equipment consists of an interconnected system of cutting machine, roof support equipment, and Direction of mining V Headgate entries Longwall panel Barrier pillar Tailgate entries Main entries Figure 1.18.— Plan view of retreating longwall. 13 conveyor haulage. The cutting machine moves along the face on a conveyor that also carries away the mined coal. Behind the face conveyor, and connected to it, is the roof support equipment, which supports a protective metal can- opy or shield that extends over the face area. These roof support units provide both the protection and the forward mobility of the system. The typical face conveyor is a flexible armored-chain conveyor powered by motors at the headgate and tailgate. Mined material moves toward the headgate, where it dis- charges to the panel belt via an elevated intermediate haulage unit, the stage loader. Shortwall mining is a less common mining method; it is very similar to longwall mining except that the short- wall panel is normally 150 to 200 ft wide. From the stand- point of equipment, shortwall can be considered as a com- promise between room and pillar and longwall in that the extractive and face haulage systems are identical to those in continuous mining, while the roof support equipment is similar to that used in longwall mining. POWER SYSTEMS IN UNDERGROUND MINES Regulations Underground mine power systems have different char- acteristics from those for surface mines, and these two basic mining operations are regulated by separate codes and standards. For instance, although 30 CFR 77 covers elec- trical installations of surface coal mines and surface facili- ties of underground coal mines, Part 75 regulates the under- ground installations and Part 18 specifies standards for electrically powered face machinery (14). Part 77 illustrates an overlap between surface and underground legal de- mands, which is logical because the surface electrical counterparts of both mine types are similar; examples in- clude substations and subtransmission. Figure 1.19 can be compared with figure 1.9 to see the similarity between sur- Incoming transmission lines Primary switchyard 6 6 6 6 6 Note: Dual- bus configuration can be used if second source desired To fan power system To surface loads Mam substation 1 Switches and circuit breakers 1 To preparation plant Main substation 2 Subtransmission Preparation plant substation Borehole 1 to underground Borehole 2 to underground Figure 1.19.— Subtransmission for underground mine. face mine and underground mine subtransmission. As a general situation, however, the mine distribution system is related to the mining method. Hence, underground mine systems become different from surface mines at the point where the circuits leave the substation and go underground. Underground Mine Distribution As shown in figures 1.20 and 1.21, underground mine power systems are somewhat more complicated than those for surface applications. Because of the nature of the mine and its service requirements, distribution must almost always be radial (fig. 1.20); the freedom in routing distri- bution enjoyed by surface mines is not available under- ground. For increased reliability, secondary-selective main substations are employed (fig. 1.21). The secondary-selective operation is defined by the use of two substations and mine feeders with a normally open tie breaker. Primary-distri- bution voltage is most commonly 7,200 V; however, older 4,160-V systems can still be found, and 12,470 V has in- creased in popularity in recent years, especially for long- wall operations. The grounding system for the underground mine distribution must be separated from that used for sur- face equipment. Power and mine grounding are fed underground in in- sulated cables, either through a shaft or borehole or a fresh-air entry. The cables terminate in disconnect switches within 500 ft of the point of power entry into the coal seam. These switches allow total removal of underground power in an emergency. From the disconnects, which may be a part of a switchhouse, the power is distributed through cables to power centers or rectifiers located as close to the machinery as practical. All the cables on high-voltage cir- cuits, usually involving only distribution, must have shield- ing around each power conductor. The prime load concentrations in underground mining are created by the mining sections. Distribution terminates at the section power center, which is a transformer com- bined with a utilization bus and protective circuitry. From this, several face machines are powered through couplers and trailing cables. Power-system segments for typical con- tinuous and longwall operations are given in figures 1.22, 1.23, and 1.24. Rated machine voltage for most installations is 550 Vac, but 250- Vdc and 440-Vac equipment is used extensively, and 950 Vac has become quite popular for high-horsepower continuous miners and longwall shearing machines. In the longwall system, power is fed through controls to the various motors. On conventional or con- tinuous equipment, the utilization approximates the radial system shown in figure 1.22. If belt haulage is used, distribution transformers are located close to all major conveyor belt drives and are re- ferred to as belt transformers. After transformation, power is supplied through starter circuitry to the drive motors. With rail haulage, distribution terminates at rectifiers that contain a transformer and rectifier combination. The rec- tifiers are located in an entry or crosscut just off the rail- way. As shown in figure 1.25, dc power is then supplied through circuit breakers to an overhead conductor or trolley wire and the rail, with additional rectifiers located at reg- ular intervals from 2,000 to 5,000 ft along the rail system. For further protection, the trolley wire is divided into elec- trically isolated segments. The typical rectifier supplies the ends of two segments of trolley wire and each feeder has its own protective circuitry to detect malfunctions. Each trolley-wire segment is called a dead block. This loop-feeding 14 arrangement is continued throughout most of the haulage system except for the most inby segment, which is dead- ended. In some mines, dc face equipment and small dc motors are powered from the trolley system through a fused connection (or nip) to the trolley conductor and rail. The dc distribution can also serve large motors directly through switchgear; however, this practice is rare in underground coal. All power equipment used underground must be rugged, portable, self-contained, and specifically designed for in- stallation and operation in limited spaces. In addition, all equipment and the cables connecting them must be pro- tected against any failures that could cause electrical haz- ard to personnel. This is primarily provided by protective relaying built into each system part, with redundancy to maximize safety (4). Utility company I I metering UJ Main substation Ground level Rectifier Cool seam To other distribution Miscellaneous loads (shop, pump, etc.) CONTINUOUS MINING SECTION IJUUUUUUUU /|\ /Tv /K /T\ /K /K /K /t\ 3 Oj — (M ° s t « •t- •« o a S\ /\ /\ S\ /\ 01 o • /f\ /*\ y^ /K /K Master control e e - O -O TJ W >* >s z r Figure 1.20. -Radially distributed underground power system. 15 MAIN SUBSTATION AREA Ground level Coal seam r rrr t ~*— Substations '"# 1 h r 4r " Normally open tie breaker Shafts or boreholes To other portable ^_ switchhouses and loads (1/2 total) -» Disconnect switchhouses ^>-r-»- UJ UJ To other portable switchhouses and loads (1/2 total) Major load concentrations Figure 1.21. -Secondary-selective distribution In underground mines. Low -voltage couplers High-voltage couplers Feeder cable / C Trailing cable iter Continuous miner "5 r Bolter Shuttle car Shuttle car s Feeder L pow Mine /er co Figure 1.22. -Utilization in continuous mining section. 16 / 2 3 4 5 6 7 8 Self-advancing supports Stage loader KEY Motor, 125 hp 1,000- kVA power center 125 -hp stage- loader starter Dual 125-hp face-conveyor starter Dual 75- hp pump and 230 -hp shear starter Pump, 75-hp Pump, 75-hp Master control Figure 1.23. — Power-system segment with longwall equipment. High-voltage feeder cable High-voltage coupler Medium-voltage cables Permissible medium-voltage couplers Power center Medium- voltage outlets mm Face -conveyor starter ^ un uu^ Medium- voltage couplers Medium-voltage cables Master control Cables for control of starters / d_d u~u II II D_a UUu" Shearer and stage- loader starter d_a Medium-voltage cables if \ i UUP" IMMI Permissible medium -voltage couplers Hydraulic- pump starter Medium-voltage cables 950 -V face conveyor motors 950-V shearing- ,950-V stage- 950-V hydraulic- machine motors loader motor pump motors Figure 1.24. -Diagram of electrical-system segment for longwall. 17 High- voltage feeder Switchhouse Switchhouse Power from substation «n-«r UJ IT '/f\~ High-voltage ac input i 1 «-r-«- LaJ "4\ A r /" Low -voltage dc output Rectifier 2 positive outputs to trolley conductor, each protected by a separate rectifier A ■*• To other downstream switchhouses Rectifier Dead-block segment Figure 1.25. -Parallel-feed haulage system. Dead - ended segment SURFACE FACILITY POWER REQUIREMENTS The surface activities of any mine, which may include shops, changing rooms, offices, ventilation fans, hoisting equipment, preparation plants, and so forth, can have large power requirements. For safety, these facilities should at least have an isolated power source and at times a separate substation. In preparation plants, the distribution arrangements are almost always expanded radial or secondary selective (8). Representative system layouts are shown in figures 1.26 and 1.27. In both, distribution is at 2.4 to 13.8 kV, with 4,160 V the most common level. Power is distributed at one of these levels to centers of electric load. This power may be used directly for high-voltage motors, but usually the voltage is stepped down to supply groups of motors or single high-horsepower motors. The power centers must be in an elevated location or totally enclosed. The rooms used for these and other electrical components may also be pressur- ized to exclude coal dust. The most popular voltage for preparation plant utiliza- tion is 480 V. This voltage is used to drive all motors throughout the plant except those with high-horsepower demands, such as centrifugal dryers and large fan drives, where 2,300 or 4,160 V is commonly employed. These higher voltages may also be preferred for any motor that requires continuous service or independence from the power-center loads. Note that 240-V motors are unsatisfactory for typi- cal preparation plant demands. Most modern preparation plants use group motor control instead of individually housed control units, since this method facilitates main- tenance and enables the interlocking of the various motor functions required for semiautomatic facilities. All manual controls, indicating lights, and so on are grouped in one central operating panel to allow easy access and visual indication of plant operation. The panel is often called a motor control center (MCC), as shown in figure 1.27. BASIC DESIGN CONSIDERATIONS The goal of the power engineer is to provide an effic- ient, reliable electrical system at maximum safety and for the lowest possible cost. The types of information made available to the power engineer include the expected size of the mine, the anticipated potential expansion, the types of equipment to be used, the haulage methods to be em- ployed, and whether or not power is available from a util- ity company. The amount of capital assigned for the elec- trical system will also be designated. 18 From utility to subtransmission system 24 to 13.8 kV Substation [;] □ — a □ a i I — 1 — y Power center Y ^ I T Power centers Thermal dryer Loadout station Slope conveyor _mrtL pry " T T I T T T T I To 480-V preparation plant loads Figure 1.26.— Representative expanded radial distribution for preparation plant. The designed system must meet certain minimum criteria. IEEE (12) has defined these basic criteria for in- dustrial electrical systems that must be applied to mines: • Safety to personnel and property, • Reliability of operation, • Simplicity, • Maintainability, • Adequate interrupting ability, • Current-limiting capacity, • Selective-system operation, • Voltage regulation, • Potential for expansion, and • First cost. Of these, safety, reliability, and simplicity are closely re- lated. All are dependent on good preventive maintenance. In the cramped uncompromising environment of an under- ground mine, these are of vital concern. Since continuous operation is the aim of every mine operator, planned main- tenance should be held to a minimum. Most routine main- tenance should be capable of being performed by unskilled personnel, since it will be done by the miners themselves. Training for these tasks must be provided. Adequate interrupting capacity, current-limiting capa- bility, and selective-system operation are projected at safety through reliability. The first two areas ensure protection during a disturbance. Current limiting, when applied to grounding, is perhaps the most significant personnel safety feature of mine electrical systems. Selective-system opera- tion is a design concept that minimizes the effect of system Utility metering Motor control center Motor control Motor control center •i/ \l ~ i/ in w i in 480-V motors 4,160-V crusher motors 480-V motors Motor control center ^ vl/ V \b nil About 2,500 -hp, 4,160-V motors Motor control center \b 4> 4< ™ IIII About 2,500-hp, 4,160-V motors Motor control center -J-j-j-J Motor control center _\|/ \|/ \ l \L \l/ -is in in 480-V motors 480-V motors Row-coal circuit Coarse- coal circuit Pumps, conveyors, blowers, filters, fans, jigs, etc. Fine-coal c i re u i t Auxiliaries Figure 1 .27. — Representative secondary-selective distribution for preparation plant. 19 disturbances. Voltage regulation is a limiting factor in system design, particularly underground, and is often the main constraint to system expansion. It should be antici- pated that when the size of the mine is increased, this might involve augmenting the power-system supply through ad- ditional power sources. While first cost is important, it should never be the determining factor, since high-cost equipment, projected at maximizing safety and reliability, can easily offset the in- creased first cost through the reduction in operating costs. At times, this fact appears to elude some company pur- chasing agents. Using the data available, the task of the power engi- neer is to select one combination of power equipment over another, provide power or circuit diagrams, estimate the equipment, operating and maintenace costs, set the speci- fications for the system, and receive and assess the proposals from suppliers. For success, the engineer requires a firm knowledge of mine power systems, but this understanding cannot be based on a "standard mine electrical system" because such a standard does not exist: no two mines are exactly alike. The engineer must resort to the fundamen- tal concepts, an awareness of what has worked in the past, and a clear understanding of the legal constraints. This in- formation is provided in the subsequent chapters. REFERENCES 1. American Standards Association. American Standard Safety Rules for Installing and Using Electrical Equipment in and About Coal Mines (M2.1). BuMines IC 8227, 1964. 2. Bergmann, R. W. Excavating Machinery. Ch. in Standard Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 3. Bucyrus-Erie Co. (South Milwaukee, WI). Surface Mining Supervisory Training Program. 1976. 4. Cranos, J. C, and D. E. Hamilton. Portable Substations for Mine Power Systems. Ind. Power Syst, v. 19, Mar. 1976. 5. Hollingsworth, J. A., Jr. History of Development of Strip Mining Machines. Bucyrus-Erie Co., South Milwaukee, WI, 1967. 6. Institute of Electrical and Electronics Engineers (New York). Recommended Practice for Electric Power Distribution for In- dustrial Plants. Stand. 141-1986. 7. Jackson, D., Jr. Coal Mines. Ch. in Standard Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 8. Lordi, A. C. Electrification of Coal Cleaning Plants. Mechanization, v. 20, Oct. 1956. 9. Trends in Open-Pit Mine Power Distribution. Coal Age, v. 66, Jan. 1961. 10. Rein, E. C. Electrical Apparatus for Surface Mining Opera- tions. Ch. in Surface Mining. Soc. Min. Eng. AIME, 1968. 11. Robinson, N., II. Underground Coal Mining Equipment. Ch. in SME Mining Engineering Handbook. Soc. Min. Eng. AIME, v. 1, 1973. 12. Stefanko, R. Coal Mining Technology Theory and Practice. Soc. Min. Eng. AIME, 1983. 13. Thuli, A. J. Power. Sec. in SME Mining Engineering Hand- book, ed. by J. M. Ehrhorn and D. T. Young. Soc. Min. Eng. AIME, v. 2, 1973. 14. U.S. Code of Federal Regulations. Title 30 -Mineral Resources; Chapter I -Mine Safety and Health Administration, Department of Labor; Subchapter - Coal Mine Health and Safety; Part 18 -Electric Motor-Driven Mine Equipment and Accessories; Part 75 -Mandatory Safety Standards, Underground Coal Mines; Part 77 -Mandatory Safety Standards, Surface Coal Mines and Surface Work Areas of Underground Coal Mines; 1981. 20 CHAPTER 2.— ELECTRICAL FUNDAMENTALS I The technique used to solve problems in complex elec- tronic circuits or mine power systems is called circuit an- alysis. It involves calculating such circuit properties as cur- rents, voltages, resistances, inductances, and impedances. Circuit analysis serves as the knowledge base on which an understanding of mine electrical systems can be built. This chapter will diverge from classical circuit-analysis presentations by not covering transient effects in circuits. From experience, studying currents and voltages existing in a circuit immediately after a change in circuit configur- ation can be confusing and clouds understanding of the most used segments of circuit analysis. Therefore, although some necessary statements will be made, the subject of transi- ents is delayed until chapter 11, where it can be combined with practical examples. This chapter commences by introducing electrical phe- nomena and continues through to a presentation of steady- state ac circuit analysis. Chapter 3, "Electrical Fundamen- tals II," continues the coverage of basic electrical subjects and starts with the basics of electrical power consumption. Numerous excellent circuit-analysis textbooks have been produced over the years. Many can be employed ef- fectively to cover the subject, and some of these are provided in the bibliography at the end of this book. Because practic- ally all fundamental electrical relationships are considered common knowledge, the concepts introduced in this chapter will seldom be referenced other than by giving credit to the discoverer. BASIC ELECTRICAL PHENOMENA The nature of electricity is not yet fully understood, but it is well known as a form of energy that can be conven- iently converted into and utilized as light, heat, and me- chanical power. Like all science, knowledge about electricity has been developed from observation and experimentation. The generalization of this experimental evidence combined with information about the nature and behavior of electrons and electron flow forms the basis of electron theory. The atoms of each element consist of a dense nucleus around which electrons travel in well-defined orbits or shells. The subatomic particles, the building blocks out of which atoms are constructed, are of three different kinds: the negatively charged electron, the positively charged pro- ton, and the neutral neutron. The negative charge of the electron, e~, is of the same magnitude as the positive charge of the proton, e*. No charges of smaller magnitude have yet been concretely observed. Thus the charge of a proton or an electron is taken as the ultimate natural unit of charge. It is these two particles that are of principal interest in electricity. Coulomb's Law The force, F, between two charges, q and q', varies directly as the magnitude of each charge and inversely as the square of the distance (r) between them. This relation- ship, known as Coulomb's law, is represented mathemat- ically by F = k^-, r 2 If force is in newtons, charge in coulombs, and distance in meters, then k = 9 x 10 +9 N-nrVC 2 . The unit of charge, the coulomb (C), can be defined as the quantity of charge that, when placed 1 m from an equal and similar charge, repels it with a force of 9 x 10 +9 newtons (N). The charge carried by an electron or by a proton is e = 1.602 x 10- 19 C. Voltage and Current A proton in the nucleus of an atom can hold only one electron in orbit around it. When an atom contains fewer than the normal number of electrons that the protons can attract, the atom has an excess of positive charge and is said to be positively charged. Atoms with an excess of elec- trons are said to be negatively charged. The net amount of these charges is termed potential or electromotive force (emf) and is measured in volts. The separation of opposite charges of electricity may be forced by physical motion or may be initiated or complemented by thermal, chemical, or mag- netic causes or even by radiation. The potential difference or voltage existing between two points can be measured by the work necessary to transfer a unit charge from one point to the other. The volt is the potential between two points when 1 joule (J) of work is re- quired to transfer 1 coulomb (C) of charge. In other words, 1 V = 1 J/C. In some metals or conductors, electrons in the outermost orbit of the atoms are rather loosely bound to their respec- tive nuclei. These are called conduction electrons, since they can leave the atom upon the application of a small force and become free to move from one atom to another within the material. In some materials, however, all the electrons are tightly bound to their respective atoms. These are called insulators, and in these materials it is exceedingly difficult, if not impossible, to free any electrons. Conductors and in- sulators are the principal materials used in electrical systems. The application of a voltage across a conductor causes the free electrons within the conductor to move. Electrical current is defined as the motion of electrical charge. If the charge in the conductor is being moved at the uniform rate of 1 coulomb per second (C/S), then the constant current existing in that conductor is 1 ampere (A), the unit of elec- trical current. The amount of current in a conductor can also be measured as the rate of change of the charge flow. Such changing current at any point in time is called in- stantaneous current or (the rate of change of charge) = -rf, (2.2) where i (2.1) where k = proportionality constant that depends on units used for force, charge, and distance. instantaneous current, A, q = flow of charge, C, t = time, s. When electricity was first discovered, it was erroneously thought that it was the flow of positive charges. Since the laws of attraction and repulsion were known, the movement was assumed to be from positive to negative. This theory 21 was accepted until the discovery of the radio tube, when it was recognized that the flow was movement of electrons from negative to positive. However, the concept of positive- charge flow was firmly entrenched and has remained stand- ard in the United States, and so it will be used here. SYSTEM OF UNITS Most material contained in this text is given in the International System of Units (SI); exceptions are calcula- tions that are more conveniently expressed in terms of the English or American engineering systems. A listing of the basic symbols, units, and abbreviations that are used is given in table 2.1. The decimal system is used to relate larger and smaller units to basic units, and standard pre- fixes are given to signify the various powers of 10; for example: pico- (p-, 10" 12 ) nano- (n-, 10~ 9 ) micro- (^-, 10" 6 ) milli- (m-, lO' 3 ) kilo- (k-, 10 3 ) mega- (M-, 10 6 ) giga- (G-, 10 12 ) Voltage, current, and power variables are represented by the letter symbols V, I, and P in both uppercase and lowercase letters. Uppercase letters represent voltage, cur- rent, and power when the variable is constant, as in dc circuits. In ac circuit work, uppercase V and I represent effective values and uppercase P represents average power. Lowercase v, i, and p depict voltage, current, and power when these quantities are varying with time. Where needed, double-subscript notation is used to de- scribe current and voltage. V AB represents the voltage of point A with respect to point B. I CD represents the current flowing through a circuit element from C to D. Note that in the circuit shown in figure 2.1, the voltage V^b causes the current l AB to flow. These meet with the standard for electrical current, which is positive-charge flow from posi- tive to negative. EXPERIMENTAL LAWS AND PARAMETERS It is remarkable that the entire theory of electrical circuits is based on only six fundamental concepts. One is Ohm's law, two are named for Kirchhoff, two relate to in- ductance and capacitance, and one has to do with power. To understand any electrical system, comprehension of these relationships is mandatory. Ohm's Law Georg Simon Ohm (1789-1854) discovered that the elec- trical current through most conductors is proportional to the voltage (potential) applied across the conductors. This phenomenon is known as Ohm's law and is expressed mathematically as = Ri, (2.3) where v = applied potential, V, i = current through the conductor, A, R = proportionality constant known as resistance of conductor, Q. Table 2.1— SI symbols and units Quantity Variable symbol 1 SI unit Unit Identical symbol unit C As A V W/A V V Q V/A S A/V V/A S A/V Q V/A S A/V F C/V H Wb/A J N-m W J/s VA var Q-m S/m C C/m 2 V/m F/m Wb Vs A/Wb H- 1 Wb/A H T Wb/m 2 A/m H/m Charge Current Voltage Electromotive force . Potential difference . Resistance Conductance Reactance Susceptance Impedance Admittance Capacitance Inductance Energy, work Power (active) Power — apparent . Power — reactive . . . Resistivity Conductivity Electric flux Electric flux density, displacement . Electric field strength Permittivity Relative permittivity Magnetic flux Magnetomotive force . Reluctance Permeance Magnetic flux density.... Magnetic field strength . Permeability (absolute) . Relative permeability.... I V.E...U V V,* R G X B Z Y C L W P S...P„ Q...P q p * D E * F...F R...R P...P B H V- Mr coulomb . ampere- volt ...do ...do ohm Siemens. ohm Siemens. ohm Siemens. farad henry joule watt , voltampere. var , ohm-meter Siemens per meter coulomb coulomb per square meter. volt per meter farad per meter (numeric) weber , ampere (amp turn) i ampere per weber reciprocal henry weber per ampere henry , tesla , ampere per meter henry per meter (numeric) V,E indicates alternative symbols;. ..U indicates reserve symbols. 22 No restriction is placed on the form of v and i. In dc cir- cuits they are constant with respect to time, and in ac circuits they are sinusoidal. For metals and most other conductors, R is constant. In other words, its value is not dependent on the amount of current, i. In some materials, especially in crystalline materials called semiconductors, R is not constant, and this characteristic is useful in diodes, amplifiers, surge arresters, and other devices. Further experiments by Ohm indicated that the resist- ance of a piece of metal depends on its size and shape. How- ever, the resistivity, q, of the metal depends only on its composition and physical state. This is an inherent prop- erty that opposes current through the conductor just as the frictional resistance of a pipe opposes the flow of water through it. Resistivity is defined as the resistance of a unit cube of homogeneous material; hence, resistivity can be thought of as a property of the material at a point. Its value remains the same at all points in a homogeneous conduc- tor, but if the material is not homogeneous, its resistivity can vary from point to point. The value may also vary greatly for different conductors. The concept of resistivity is often used in the grounding and distribution aspects of mine electrical systems. Using the definition, practical resistivity units would be ohm-centimeter (Q-cm) and ohm-inch (S2-in). However, resistivity is usually expressed in ohm-meter (Q-m) (SI) and ohm-circular-mill-foot (English). The ohm-meter is the re- sistance of a material 1 mm 2 in cross section with 1 m length. Likewise, the ohm-circular-mill-foot (usually abbre- viated to Q-cmil-ft) refers to the conductor resistance for a volume 0.001 in (1 mil) in diameter and 1 ft long. For calculating the resistance in this latter case, the cross- sectional area of the conductor is measured in circular-mills, which can be found from A = d 2 , (2.4) where A = cross-sectional area of circular conductor, cmil, and d = conductor diameter, 10" 3 in. Resistivity values of some common conductors are given in table 2.2. Table 2.2— Resistivity of some common materials at 20° C Material Aluminum, commercial Copper, annealed Iron, annealed Lead Nichrome Silver Steel, mild Tin Tungsten Temperature Resistivity (p) coefficient (a) 10- 8 Q-m Q-cmil-ft 0.0039 2.824 17.1 .00393 1.724 10.5 .005 9.50 57.4 .0034 21.83 132.31 .04 100 606.1 .0038 1.63 9.85 .002 11.91 72.17 .0042 11.50 69.7 .0045 5.50 33.2 If I is in meters and A is in meters squared, then q must be given in units of ohm-meters. Electrical resistivity does not remain constant if the temperature is permitted to change. For most materials, the resistance increases as the temperature increases; car- bon is an exception to this rule (negative temperature co- efficient, 0.005). If the temperature coefficient is known, the resistance of a given conductor at a given temperature is R = R [1 + a (t - to)], (2.6) where R = resistance at temperature t, Q, Ro = resistance at reference temperature t , usually 20° C, Q, a = temperature coefficient, Q/°C, t = given conductor temperature, °C, t = reference temperature, °C. At very low temperatures (about - 200 ° C for copper) or as the melting point is reached, the temperature coefficient is no longer constant and changes with temperature. As a result, equation 2.6 is not valid for very high or low temperatures. The symbol illustrated in figure 2.1 portrays a resistor in a circuit, and often its resistance is stated. Again by definition, R (2.7) Sometimes, the element's conductance, G, is referenced and is defined as the reciprocal of resistance: G = i = l. R v (2.8) In circuit analysis, it is occasionally more convenient to use conductance than resistance. Later, the explanation of this symbol will be generalized. Kirchhoff's Voltage Law In the simple series circuit shown in figure 2.2, three resistors are connected in tandem to form one single closed loop. Kirchhoff has shown that when several elements are r^WV Figure 2.1. - Circuit element Illustrating voltage polarity and current flow direction. The resistance of any specific conductor can be calcu- lated from the material resistivity using the formula R where R = resistance, Q, I = conductor length, A = conductor cross-sectional area, and q = material resistivity. (2.5) Figure 2.2. — Simple series circuit. 23 connected in series, the current in the circuit will adjust itself until the sum of voltage drops in the circuit is equal to the sum of voltage sources in the circuit. This can be restated as the "sum of all voltages around any closed cir- cuit is zero," which is called Kirchhoff s voltage law. For the circuit shown in figure 2.2, or or v oi + v ic + v cd + v da = v afc + v bc + v cd - v ad = v 1 + v 2 + v 3 - v s = 0. (2.9) (2.10) (2.11) Obviously, some of these potential differences could be negative and some positive. This circuit shows only resist- ances and a voltage source, but the network could contain other kinds of elements and might be as complicated as desired. However, Kirchhoff found that the sum of the volt- ages around any closed loop in a circuit, such as a-b-c-d, is always zero. The symbol shown in figure 2.2 beside v s represents an ideal voltage source. Such a source maintains a given voltage across its output (terminals) regardless of the load, but actual voltage sources cannot supply an infinite cur- rent if the terminals are short-circuited; that is, they are tied together so the resistance approaches zero. Therefore, actual sources are usually considered to be ideal voltage sources with an internal resistance connected in series with the source and the output terminals. The assumption is illustrated in figure 2.3. EXAMPLE 2.1 Find the current I flowing in the single-loop cir- cuit in figure 2.4. SOLUTION. Adhering to the assigned clockwise direction for current, Kirchhoff s voltage law produces the following equation: -50 + V, + 100 +V 2 = 0, where V, and V 2 are the voltages across the l-£2 and 2— Q resistances. From Ohm's law, V, = II, V 2 = 21. Inserting these expressions into the voltage law equa- tion produces or -50 + II + 100 + 21 = 31 = -50, I = -16.7 A. The negative sign states that the actual current flow is in the opposite direction from that shown in fig- ure 2.4. It should be noted that when writing the voltage- law equation, voltages that oppose the assigned cur- rent flow are considered positive, otherwise negative. Therefore, the 100-V source is positive, and the 50-V source is negative. The positive signs for V! and V 2 assumed opposition by the convention shown in fig- ure 2.1. Kirchhoff's Current Law The other law attributed to Kirchhoff specifies that "the sum of all electrical currents flowing toward a junc- tion is zero." In figure 2.5, five wires are soldered together at a common terminal and the current in each wire is measured. If current flowing toward the junction is called positive (the direction shown in the figure) and the current outwards is negative (against the arrows), then the sum of the five currents is zero: i, + i, + i, + L + i. = 0. (2.12) As was the case for equation 2.9, this equation implies that some currents must be positive, some negative. If two or more loads are connected between two com- mon points or junctions, these elements are said to be in parallel, as shown in figure 2.6A. The same is true for figure 2.6B, and moreover, the two circuits illustrated in figure 2.6 are identical, just drawn differently. It is important to t/W- I O Resistance A Ideal — — source Output terminals ,X Ideal voltage source Actual voltage source Figure 2.3. — Ideal and actual voltage sources. Figure 2.4. -Circuit for example 2.1. Figure 2.5. — Demonstration of Kirchhoff's current law. 24 note that the lines in these and all circuit diagrams usu- ally show no resistance. Each line is only a connection be- tween elements or between an element and a junction. The similarity in the diagrams can be shown using KirchhofFs current law. In both, there are only two independent junc- tions, a and b, and for either point, ii = i 2 + i 3 + U (2.13) The circuit symbol next to i x in figure 2.6 represents an ideal current source, and a similar situation exists for all practical current sources as was mentioned for practical voltage sources. However, the internal resistance is effec- tively connected in parallel across the ideal current source. Both ideal and actual current sources are shown in fig- ure 2.7. EXAMPLE 2.2 Verify that KirchhofFs current law holds for junc- tion x in figure 2.8. SOLUTION. The three resistances in figure 2.8 are in parallel, and the 100 V produced by the voltage source exists across each. Therefore, by Ohm's law, the current through each resistance is T = 100 = 4 a ■•■2S * ■**! 25 I 50 = 100 = 2 A, 50 I = 100 - i a i 100 — -■-"" — ± t\. 100 KirchhofFs current law states that for junction x, I25 + I50 + lioo = 7 A. Accordingly, 4 + 2 + 1 + = 7 A. Series Circuits To restate the earlier definition of a series circuit, elements are said to be connected in series if the same cur- rent passes through them. Such is the situation for the four resistors shown in figure 2.9. It would be convenient to find a resistance, R, that could replace all series resistors. This equivalent resistance can be found by returning to the Ohm and Kirchhoff voltage laws. By Kirchoff s law, V = v, + v 2 -I- v 3 + v„, but by Ohm's law, v, = iRj, v 2 = iR 2 , v 3 = iR 3 , - Therefore, For the circuit in figure 2.9, if the voltage, v, produces the same current, i, through the circuit, then but or v = iR, v = i (R x + R 2 + R 3 + R 4 ), iR = i (R, + R 2 + R 3 + R 4 ), R = R x + R 2 + R 3 4- R 4 . (2.14) Here R is said to be the equivalent resistance for the previous series circuit. In other words, R is the series resistance of that circuit. The same logic applies to all elec- trical elements in series. a * •- Q) R. »" i\" ^ A B Figure 2.6. -Simple parallel circuits. (Italic letters are cited in text.) Ideal current source 1 . Ideal S\ 5!*"* Resistance current source Output terminals Actual current source Figure 2.7. — Ideal and actual current sources. IOOV 25/1 <>5on. -t/W — ♦ 28A (T V $2S §1S Figure 2.28. -Series-parallel conductances for example 2.7. 33 EXAMPLE 2.8 For the circuit shown in figure 2.29, find the total circuit current, I a , with the components as shown, with the 5-Q resistor short-circuited, and with the 5-Q resistor open-circuited. SOLUTION. For the circuit as shown, the two 10-Q resistors between points c and d are in parallel and _ mm _ 5a 10 + 10 This resistance is in series with the 5-Q resistance and these three elements are in parallel with both the 15-Q and 10-Q resistances between points b and d. Thus, R.C = 5 + 5 = 10 Q and or R bd is in series with the 7-Q resistance and Raw = 7 + 3.75 = 10.75 Q, and R abd is in parallel with the 10-Q resistance be- tween points and a and d; both are across the 120-V source. Therefore, 1 Rid = _1 15 10 1 •Kicd 1 Rid = _1 15 + -1+ 10 1 10 R >d = 3.75 Q. R, (10.75) (10) 10 + 10.75 and total circuit current is V = _120 R,„ " 5.18 Ix = = 5.18 Q, 23.2 A. When the 5-Q resistance in figure 2.29 is short- circuited or replaced with zero resistance, points b and c are electrically the same. Four resistances are now in parallel between points b and d; three 10-Q and the 15-Q resistance. Following the same procedure as before, the equivalent resistance becomes K'„ = 4.93 Q, and total circuit current is 120 V 4.93 24.3 A. For the case of an open-circuited 5-Q resistance, the resistance between b and c is assumed to be infinite, and the two 10-Q resistances between points c and d are disconnected from the circuit. The equivalent resistance is now R4' = 5.65 Q, and the total circuit current is 120 Ix" 5.65 21.2 A. This example illustrates an important concept. When an element in a circuit is short-circuited, the equivalent resistance of the circuit will decrease, and total circuit current will increase. Conversely, with an open-circuited element, the equivalent resistance of the entire circuit will increase while total circuit current decreases. Wye-Delta Transformations Any of the circuits now covered can be reduced to a two- terminal network, as seen in figure 2.30A. The circuit receives power from an external source and can contain resistance, inductance, and capacitance. Such networks are called passive. For dc, only resistance is of interest, and it can be found from the terminal voltage and current by R V I' Numerous circuits can be represented by a two-terminal arrangement. Other circuits, including several in mine power systems, cannot be represented in this way, but many of these can be resolved into the three-terminal network given in figure 2.30B Even though with three terminals there now exist three voltages and three currents, the con- cept of circuit equivalence still holds; that is, voltages and currents are identical and the circuits are equivalent. Ii 120V a 7H t — mh 6 isn 1 */VW- lon ion Figure 2.29. - Series-parallel circuit for example 2.8. '12 I 1 i— __1 1 « o ^ Figure 2.30.— Two-terminal (A) and three-terminal (B) net- works. 34 For three-terminal networks, there are two basic circuit configurations: the wye (y) and the delta (A) (fig. 2.31). The wye is sometimes called a star, but the term y is standard. It is sometimes advantageous to replace or substitute the three wye-connected resistances with another set that is delta-connected, or vice versa. By using equivalence of input currents and voltages for wye and delta circuits, delta-wye (or delta-to-wye) and wye- delta transformations can be derived. Thus for equivalence of the circuits in figure 2.31, and R.» = R»c = R„ = R„R(, + R(,R C "+■ RcR Re R<.R(> + R R C + R C R„ R i R Rt + R 6 R C + R C R„ R* (2.45) (2.46) (2.47) Figure 2.31. -Wye (A) and delta (8) circuit configurations. In other words, the delta is equivalent to the wye if the resistances of the delta are related to the wye by equations 2.45, 2.46, and 2.47. Accordingly, with three terminals a, b, c, containing wye-connected R„, R„, R c , the circuit perform- ance is unaffected by replacing them with a delta-connected R a6 , R bc , R ca . Likewise, for equivalence of delta-to-wye sets, and R„ R t = Rc = RotRtc R afc + R bc + Rj RoiRtc R o6 + R ic + R c „' RacRic R„t + Rl,,. + R„ (2.48) (2.49) (2.50) © R ab -WV\A- R ac Ri be These transformations are useful in allowing three-terminal circuit reduction because they allow substitution when a network does not contain either series or parallel elements. The circuit or circuit portion may not outwardly appear as three-terminal, and common examples, n and T, are given in figure 2.32. These are actually delta and wye circuits drawn in a slightly different fashion. It will be shown in chapter 4 that delta and wye circuits are the two most im- portant configurations for power systems. These trans- formations will be called upon again at that point. It has already been shown that when circuit elements are neither all in series nor all in parallel, but have some other series-and-parallel arrangement, the elements can be handled in groups to reduce the circuit to an equivalent resistance. This important kind of circuit analysis has been called circuit reduction. Now that delta-wye and wye-delta transformations have been introduced, the substitution process can be employed to solve networks that contain elements neither in series nor parallel. A prime instance is the common bridge circuit shown in figure 2.33. The bridge is one of the most used configurations in electrical instrumentation. The objective here is to find all available currents and voltage drops in the network, and an overall solution approach is illustrated in the following example. Figure 2.32. — "T" and V circuit configurations. ■ab ao V, ab bo VWV Figure 2.33. — Common bridge circuit. 35 EXAMPLE 2.9 Consider that the resistances shown in figure 2.33 are as follows: R x = 5 Q R 2 = 10 Q R 3 = 15 Q R 4 = 20 Q R 5 = 25 Q R 6 = 0.4 Q Find the equivalent resistance of the circuit between points a and b. SOLUTION. The original circuit has been redrawn in figure 2.34A, in which a delta configuration is clearly defined by points a, c, d. The first step for cir- cuit reduction is to convert the delta to a wye. From equations 2.48, 2.49, and 2.50, R„ RadRc R ad + R C(J + R (10) (5) 10 + 15 + 5 = 1.67 Q, and (5H15) = 25 30 Q5H10) = 5£2 30 This conversion results in the simple series-parallel circuit in figure 2.345. Combining the series elements and parallel branches in the center of the circuit fur- ther reduces the circuit to that shown in figure 2.34C: R c + R 4 = 2.5 + 20 = 22.5 Q, R a + R 5 = 5 + 25 = 30 Q, = (22.5) (30) = 22.5 + 30 The equivalent resistance of the total circuit is then simply R c , = R„ + R x + R 6 = 1.67 + 12.9 + 0.4 = 15 £2. The total circuit current can now be found using Ohm's law; for instance, if V ai = 30 V, then ^ = 30 = 2A. R.. 15 Finally, Kirchhoff s current and voltage laws and the voltage and current distribution formulas can be employed to find currents through and the voltage drops across each circuit element. For example, if I 4 is the current through R 4 , then I 4 = U = 2(- Rd + R 5 R c + R 4 + Rd + R5 -) 30 22.5 + 30 -)=1.14A. It should be noted that the current through R c is also I 4 , but R c does not exist in the original circuit of figure 2.33. Thus, a problem exists in finding the currents through R 1( Rj, and R 3 . One solution would be to solve for the three potentials among points a, c, d and use Ohm's law to find the three currents in the assigned delta connection. , R 6 b )b A B Figure 2.34. — Circuit reduction of bridge circuit. Ft Rc 36 EXAMPLE 2.10 Consider that the resistances shown in figure 2.33 are R t = 15 Q R 2 = 15 Q R s = 15 Q R 4 = 20 Q R 5 = 25 Q R 6 = 10 Q Find the equivalent resistance of the circuit between points a and b. SOLUTION. Three identical resistances form a delta configuration in the circuit, or R t = R 2 = R) = 15 Q. Following the same processes as in example 2.9 for figure 2.34B, R„ (15) (15) = 5 2, 15 + 15 +15 R„ = 5 Q and R t = 5 Q. Now, the center resistance of figure 2.34C is (5 + 20) (5 + 25) R, 13.64 Q, 5 + 20 + 5 + 25 and the equivalent resistance of the circuit is R e , = 5 + 13.64 + 10 = 28.6 Q. It should be noted that an important situation is es- tablished where all resistances in a delta or a wye configuration are equal. If R 4 is each resistance in the delta and R y is that in the wye, then from equa- tions 2.48, 2.49, or 2.50, R v = Ra Ra or Rv = Ra + Ra + Ra Ra The majority of delta or wye configurations used in power systems consist of identical elements in each leg. Much of circuit analysis can be handled by circuit re- duction, but as circuits become more complex this process becomes cumbersome. Nevertheless, circuit reduction should always be used when it produces results more eas- ily than other methods. There are solution approaches that are more systematic, and the next two sections discuss two of these. Circuit and Loop Equations Before more general solution methods can be identified, the meanings of some words need to be clarified. A node is the position or point in a circuit where two or more elements are connected. When three or more elements ex- tend from a node, the node is called a junction. A branch is a circuit portion existing between two junctions and may contain one element or several in a series. A loop is a sin- gle closed path for current. Figure 2.35 illustrates all these circuit parts. The following technique, loop analysis, is based entirely on Ohm's law and Kirchhoff s voltage law. The analysis principle produces n simultaneous equations requiring the solution of n unknowns, and the unknowns are currents. In loop analysis it is only necessary to determine as many different currents as there are independent loops; that is, the equations are constructed by defining independent- loop currents. For example, in figure 2.36, the current 1^ flowing out of source V t and through R„, will be around loop 1. The current flowing from source V 2 through R c will be around loop 2. Although not essential, these directions follow the general convention of assigning all reference loops clockwise. It is sometimes more desirable to use other directions, for instance with currents flowing out of a source positive terminal, but it is imperative that the use of cur- rents within a specific loop remains consistent after the loop is assigned. Notice in figure 2.36 that both I t and I 2 flow through R fc . Depending on the loop direction, that is, the direction defined by I t or I 2 , the total current through R A is either I t — 1 2 or I 2 — Ii- Thus if L_ and I 2 can be found, the current through each circuit element can be determined. The first task in loop analysis is to use Kirchhoff s volt- age law to write equations about each current loop, stating that the sum of voltages about each loop equals zero. For loop 1, -V x + RJ, + RA-U (2.51) Notice that RJ, equals the voltage drop across R„, and Rftdj-Lj) equals that across R b . In the latter case, the voltage can be taken as RJi — RJ 2 , considering that the voltage produced by I 2 opposes that produced by I x . Likewise, for loop 2, = -V 2 + R C I 2 + R^-U Junction - Branch *-n^-v*-n^f-n^ V— *-Hh Source- ^-Nodes /^X ,r ±: /Loop) <=} Branch J=- '7 (2.52) Passive elements Figure 2.35. — Parts of circuit. Figure 2.36. — Circuit demonstrating two independent loops. 37 By rearranging equations 2.51 and 2.52, and (R„ + RJIi - RJ 2 = V, -RJi + (R b + RJI 2 = V 2 , (2.53) (2.54) which are two simultaneous equations with two unknowns, Ij and I 2 . These can be solved easily by simultaneous methods. It should be noted that one additional loop equation could be written, that for the loop containing both V! and V 2 . However, this will not provide another independent equation. Information concerning the maximum number of independent equations available will follow shortly. EXAMPLE 2.11 Find the current through the 1.5-Q resistor in figure 2.37 using loop analysis. SOLUTION. Two loops are defined in the figure where the current through the 1.5-Q resistor is Ii + I 2 . Applying Kirchhoffs voltage law to loop 1, 0.51, + 1.5^ + I 2 ) + 1.0I X = 250 and for loop 2, 0.5I 2 + 1.5(1! + I 2 ) + 1.0I 2 = 300. By simplifying these equations, 3I X + 1.51a = 250, 1.5Ii + 3I 2 = 300. Simultaneous solution of these results is I x = 44.4 A, I 2 = 77.8 A, and the current through the 1.5-Q resistor is Ii + I 2 = 122.2 A. To further enforce the concept of loop analysis, again consider the common bridge circuit, which is redrawn in figure 2.38 to include current loops. Three loop equations can be written because there are three possible indepen- dent loops. For loop 1, RA - I 2 ) + RA - U + RJ! - V B4 = 0; (2.55) for loop 2, RA - U + RA +RA - I 3 ) = 0; (2.56) for loop 3, 0.511 o.sn A/WV- 250 V 300 V Figure 2.37.— Two-loop circuit for example 2.11. Figure 2.38. — Bridge circuit demonstrating loop analysis. Again, rearranging, (R x + R 4 + R^ I x - RJ. 2 - R 4 I 3 = V„ t , (2.58) -Rjlj + (R, + R 2 + R 3 ) I 2 - R 3 I 3 = 0; (2.59) -R 4 l! - R 3 I 2 + (R 3 + R 4 + R 5 )I 3 = 0; (2.60) which are three simultaneous equations with three un- knowns, I u I 2 , and I 3 . The proper combination of these cur- rents will yield the current through each branch of the cir- cuit. The process was again to employ Kirchhoffs voltage law for the purpose of finding the unknown currents. Other loops about the bridge could be assigned and will produce the same valid results. Generally, the particular choice of loops can enhance a desired result. For instance, if only the current through R 3 of figure 2.38 is desired, establishing one loop current through that resistor would create a more direct solution. R 4 (I 3 - Ii) + R 3 (I 3 - I*) + R 5 I 3 = 0. (2.57) EXAMPLE 2.12 Using loop equations, solve for each branch cur- rent in the circuit shown in figure 2.39. SOLUTION. Applying Kirchhoffs voltage law to loops 1 and 2 respectively, 2(l! - I 2 ) + 5(Ii - I 3 ) + 2I X = 56, 2(I 2 - I x ) + 10I 2 + 1(I 2 + I 3 ) = 0. 38 As the assigned current for loop 3 passes through an ideal current source, I, = 6 A. Therefore, the equations for loops 1 and 2 become 2^ - I 2 ) + 5(1, + 6) + 21, = 56, 20, - I,) + 101, + 1(I 2 + 6) = 0, or 91, - 2I 2 = 26, -21, + 13I 2 = -6. Solution of the last two simultaneous equations yields I, = 2.9 A, I 2 = A. Each branch current can now be resolved from the loop currents. For the branch containing the 2-Q resistor and the 56-V source, I = I, = 2.9 A; for the other 2-Q resistor, I = I, - I 2 = 2.9 A, and the resistors in the other branches, J-10O = ±2 = V ■"■» 1,2 = 1 3 — 1 2 = D A, I 4a = I 3 = 6 A, I sa = I, + I 3 = 8.9 A. A loop equation could have been written for loop 3, but it can only state that the voltage drops across the 1-Q, 5-Q, and 4-Q resistors in that loop are equal to the voltage across the 6- A ideal current source, which is unknown. Such an equation would only complicate the solution to the problem. As circuits become more complex and the number of possible loops increases, a method for determining the number of required equations is useful. By counting the number of branches and junctions in the circuit, the follow- ing expression provides the necessary number of loop currents: number of equations = branches — (junctions — 1). ion ^vwv- 2n C I 2^ in o — vwv- J V\M " 56V p Q i 5n (S(t)6A 2il 4X1 Figure 2.39. -Three-loop circuit for example 2.12. For figure 2.38, there are six branches and four junc- tions; therefore, the number of equations needed equals 6 - (4 - 1) = 3. Node Equations In the preceding analysis, Kirchhoff s voltage law established the method of loop equations. Kirchhoff s cur- rent law did not receive any attention, yet it was satisfied. This can be demonstrated with figure 2.38 by taking any junction and summing the currents through it. Consider- ing that the currents through R, and R 2 flow from junction or hence, li l«i 1r 2 — I, - (I, - I 2 ) - I 2 = 0; I, - I, + I 2 - Ij = 0. Kirchhoff s current law is used directly in node analysis, and the unknowns are voltages across branches. The tech- nique by which these voltages are referenced or measured provides a simplifying procedure for a circuit being ana- lyzed. Each junction or principal node in a circuit is assigned a number or letter. Voltages can then be measured from each junction to one specific junction, called the reference node. In essence, the reference node is dependent on all other nodes in the circuit. Node analysis consists of find- ing the voltages from each junction to the reference node. The procedure can be demonstrated easily with the simple two-junction circuit shown in figure 2.40, in which I„, R A , and R c are known. The existing junctions are A and O, and O is taken as the reference. The voltage from A to O is then Y AO , and Kirchhoff s current law can be used to write an equation for junction A: L - L - L = 0. (2.61a) By Ohm's law, VylO — IfcRfc = IX Therefore, la Rt R c since 1/R = G, h - VaoG„ - V = 0; G C = 0. (2.616) (2.62) Equation 2.62 can be further solved for unknown, V ao . During the process, an equation was written for each junction, excluding the reference node. The number of re- quired equations for node analysis is therefore always one less than the number of junctions in a circuit. To illustrate node analysis further, consider the three- junction circuit in figure 2.41. If junction is taken as the reference node, V AO and V BO are the unknown voltages. The reference node, which establishes a reference potential across the bottom of the circuit, is normally assumed at zero potential. Accordingly, the double-subscripted voltages are unnecessary and unknown values can be simply called V^ and V B . Further, as zero potential is often referenced to earth or ground, a most convenient reference, figure 2.41 can be redrawn as shown in figure 2.42. These circuit elements are still connected to a reference node through the ground symbols, as shown. Hence, each of the circuit elements is said to be grounded. 39 Now, applying KirchhofFs current law to junctions A and B, L - Ie - l b = 0, (2.63a) -h-U+Ic = 0. (2.636) By Ohm's law, and I b = V A G b , h = V B G d l = (v„ - V B )G C . (2.64) The last expression is evident because, by KirchhofFs voltage law, the voltage across G= is the potential at junc- tion A minus that at junction B. Therefore, I. - Wa - V*)G C - V A G b = 0, (2.65) -I, - V B G d + )I 1(1 , The second part of the double subscripts is used only to signify that the currents are caused by source 1. The neg- ative sign in the last expression is caused by the current direction assumed in the illustration. The next step is let- ting V! = 0, thus restoring V 2 (fig. 2.46C), R,2 — Rj + 1:2191 = I3U] = (: Iim = v R1R3 Ri + R 3 = V, R*2 Ri + R R, ")l2[2]> I )I2[2) Rj + R 3 Finally, the sums of steps 1 and 2 yield II = IlUl + Il[21> *2 = *2[1] + l2[2] > I3 = J-311] + J-3I2] » (2.69) (2.70) (2.71) which are the currents with both sources in operation as in figure 2.34A. The process is adaptable (and perhaps more useful) for circuits having more than two voltage or cur- rent sources. As with current sources in node analysis, the unknowns in each step are voltages. Nevertheless, super- position allows many sources to be considered separately, and it is of great benefit in the analysis of circuits. EXAMPLE 2.15 Use the superposition theorem to find the voltage across the 0.5-Q resistance in figure 2.44. Note that this is the same circuit used for example 2.13. SOLUTION. Following the first step of the superposi- tion theorem, the 1,000- A current source on the right side of the circuit will be turned off. The circuit is now operating as shown in figure 2.47A. Only L Bm need be known to solve the problem. Using current divi- sion for the parallel branches, l ABll] = 1,500(-^-) = 429 A. 0.0 Figure 2.47J3 shows the second step in the problem solution, where the 1,500- A source is turned off. Now the current through the 0.5-Q resistor is L 1,000(— ) = -571 A. 3.5 Summation of these two findings produces the cur- rent from A to B with both sources operating. •MB = 1>1B[1) T 1/1B[2]> l AB = 429 - 571 = 143 A. Thus, Vab = (-143X0.5) = -72 V. It is obvious that this technique produces the answer faster than the process given in example 2.13. How- ever, node analysis may give a more efficient solu- tion with other problems. Reciprocity The reciprocity theorem states that in a linear passive circuit, if a single source in one branch produces a given result in a second branch, the identical source in the sec- ond branch will produce the same result in the first branch. IL Ri r 2 Jl IH!) r i R? fed ! '(2) Rj R2 fe 2 > 1+ Figure 2.46. — Circuit for demonstrating superposition theorem. 1.5QOA A Iab( 1 ) b A jABg) B 1 '^ )00A ^ww- 05A I 0.5SI I- 1,000- A 1,500- A. 1.500A(t) I A (i)J|in2ni'^e A £urce|in I B(2 )||2n ©'.000A off off A B Figure 2.47. - Circuit In figure 2.44 with sources turned off. 42 This reciprocal action is demonstrated in figure 2.48. In figure 2.48A, if V, produces I t in the branch that goes through R 5 , moving V, to the R 5 branch will produce I 2 in the original location of V\ (fig. 2.48B). The currents Ij and I 2 will be equal. The dual form of reciprocity has a similar function in relating a current source to the voltage pro- duced. The great advantage of this theorem is that a source may be moved to another location that is more convenient to analyze. Source Transformation and Maximum Power Transfer Before defining the theorems associated with source transformation and maximum power transfer, it is advis- able to expand the topics of ideal and practical sources. An ideal voltage source has been defined as a device whose terminal voltage is independent of the current that passes through it. Although no such device exists in the practical world, it is convenient to assume a resistance in series with an ideal source as a datum, against which the performance of an actual voltage source can be measured. This is shown in figure 2.49 where the performance of a 12-V automotive storage battery is plotted against an ideal voltage source. The internal resistance, R v , compensates the output volt- age, V x , for varying load currents, I L . These currents are obtained by changing the load, R L . It will be found that with small current the practical source approximates the ideal one. But under heavy duty where there are high current and low load resistance, the output voltage drops substan- tially. Using the Ohm and Kirchhoff voltage laws, (- V, Rv + R ■JRl, Il = V s R v + R/ (2.72) (2.73) V, equals the voltage of the ideal source, which can be found by measuring the terminal voltage with no load resistance. The internal resistance, R v , can then be determined by applying a known R L and measuring V L . Similarly, figure 2.50 models a practical current source where R, is the internal shunt resistance. The graph illus- trates the effect of this resistance: as the load resistance increases, terminal current decreases. Using Kirchhoff s current law, it can be shown that V £ R, + R L and Ir = C Rz R, + R L )I. (2.74) (2.75) The output of the ideal current source, L, can be found by short-circuiting the output terminals and measuring the resulting current. Then R, can be calculated by measuring Vi and l L with a known load, R L . Actually, shorting the ter- minals of a source is usually unwise because it could dam- age the real-world source, not to mention being an unsafe practice. I, can also be determined through source transforma- tion, which uses the fact that two sources are equivalent if each produces identical terminal voltage and current in any load. Therefore, for equivalence of practical voltage and Ri R 3 i— M/V- A B Figure 2.48. - Demonstration of reciprocity theorem. O.Olil > 12 1 l ' Ideal i ' i ' i source — ilO - - fin Practical UJ source £ 6 - h 4 - o „ > 2 / , i , i,i,i 0.02 0.04 0.06 0.08 0.1 LOAD (R L ), XI Figure 2.49. - Practical voltage-source model. t Is UJ S3 §G5I S |- 9 I Ideal source - '- \^ Practical ^ source : , i i R, 2R; 3Rj L0AD(R L ) Figure 2.50. — Practical current-source model. practical current sources, equations 2.72 and 2.73 must equal 2.74 and 2.75, respectively. It is obvious that both sets are interrelated. In other words, for load current, Ix R.-L R, + Rz R, + Ri (2.76) If equation 2.76 is valid for any load, R i( it must hold that R, = R„ (2.77) V. = R,I„ (2.78) where R s = the internal resistance for either equivalent practical source, V 5 = output voltage of ideal voltage source, and I, = output current of ideal current source. This relationship is shown in figure 2.51. The two circuits shown will be named shortly. Source transformation states that if one source is known, it can be replaced with the other. Note however that even if two practical sources are equivalent, the power that the two internal ideal sources supply and the internal resistances absorb may be quite ■■ ^^■^M 43 different. Notwithstanding, this substitution is helpful in writing network equations because constant-current sources are more convenient for node equations, and constant- voltage sources are best for loop equations. In addition, the exchange of particular sources may permit direct cir- cuit reduction. EXAMPLE 2.16 Solve the problem in example 2.13 using only source transformation. SOLUTION. Two practical current sources exist in figure 2.44 between junctions A and O and between junctions B and O. Applying equation 2.78 for the left- hand source, L, = 1,500 A , R rf = 12, V*. = 1,500(1) = 1,500 V; and for the right-hand source, L, = 1,000 A , R^ = 2 S2 , V* = (1,000X2) = 2,000 V. Ri4 V^ and Rsb V^ describe two practical voltage sources that can replace the current sources between junctions A and O and junctions B and O, respectively. Figure 2.52 shows the results of this transformation, where the circuit becomes a simple loop. The current from A to B is now 1ab= 1,500- 2,000 = _ 143A> 3.5 and the voltage between is V AB = (-143X0.5) = -72 V. Source transformation also produced results quicker than node analysis, but again, this might not occur with other circuit configurations. In the above solution, practical current sources were replaced by practical voltage sources. By com- paring figure 2.44 with figure 2.52, it can be seen that points A, B, and O exist in both. Caution should always be taken to ensure that a desired node is not lost after the transformation. Since load resistance can vary from zero to infinity, some value of resistance must exist that will receive the maxi- mum power available from a particular source. It can be proven, using the concepts just presented, that an indepen- dent voltage source in series with a resistance, R„ or an independent current source in parallel with a resistance, R„ delivers maximum power to a load resistance, R L , when Rx. = R,. This is called the maximum power transfer theorem. Theveain's and Norton's Theorems These theorems are closely related to source transform- ation. They can be illustrated by considering the active net- work (one that delivers power) with two output terminals shown in figure 2.53. Here, the internal configuration is unimportant, but the elements must be linear. The sources can be either ideal voltage or ideal current. Thevenin's theorem states that if an active network (fig.2.53A) is attached to any external network (fig. 2.53B), it will behave as if it were simply a single ideal voltage source, V„, in series with a single resistance, R„ (fig. 2.530. In other words, the active circuit will appear as a practical voltage source. Values for V„ and R„ can be found as follows. When all internal sources are operating normally and no loads are connected, the open-circuit voltage across the out- put terminals equals V„. With all the ideal sources turned off, a resistance, R,,, can be measured at the terminals. This is because when an ideal current source is turned off, it ap- pears as an open circuit (an infinite resistance). An ideal voltage source that is not operating acts as a short circuit, thus having zero resistance. This theorem is important because it means that any linear circuit where the internal components are unknown can be considered as a constant-voltage source in series with a resistance. Any circuit reduced to this form is called a Thevenin circuit. Norton's theorem is the corollary to Thevenin's theorem. Norton relates that if such an active network is attached to any external network, it will behave as a single ideal current source, I , in parallel with a single resistance, R„. The values for V„ and R„ can be determined by consider- ing the same linear active network, this time as showrrin figure 2.54A, with internal sources operating normally. The -Wh o Re v s =i s Rs Figure 2.51. -Source transformation. : AB in a — ► b 2n -A/WY— *- J WV\r-«— vwv- 0.5 n i + 1,500 V O O 2,000V Figure 2.52. - Circuit In figure 2.44 with current sources transformed to voltage sources. + Active network 1 Active network External network A B Figure 2.53. -Thevenin's theorem 44 output terminals are short-circuited, and a terminal cur- rent is measured to give the value for t,. R„ is found in ex- actly the same way as in Thevenin's theorem. The combina- tion of these elements gives the practical current source shown in figure 2.54C, which is also known as a Norton circuit. The Thevenin and Norton circuits are obviously related by source transformation so that if one is known, the other can be constructed. The equations relating the two are shown in figure 2.55. These theorems are usually employed when a series of calculations involves changing one part of a network while keeping another part constant. This manipulation helps to simplify complex computations such as power-system short-circuit currents. EXAMPLE 2.17 Find the Thevenin and Norton equivalents for the circuit shown in figure 2.56. SOLUTION. Applying Thevenin's theorem, the equivalent resistance of the circuit between a and b with the internal source off is R„. When the 50-V source is off, it acts as a short circuit, shorting out the 50-Q resistance in parallel with it. Thus, R = R ab = 2 + (10) (10) 10 + 10 7 Q. The voltage across a and b with the internal source operating is V„. Using circuit reduction, the equiva- lent resistance as seen by the 50-V source with no load across the terminals a and b is R.,= 50 (10 + 10) 50 + 10 + 10 = 14.3 Q. (Note that this resistance is not R„.) The current delivered by the source is Ix = 50 14.3 = 3.5 A, and from current division, 50 I 2 = (3.5) — = 2.5 A. 70 As no current is flowing between terminals a and b, V 2 shown in figure 2.56 is equal to V ai , which is equal to V„. Thus, V = V a6 = V 2 = (2.5) (10) = 25 V. V„ and R„ describe the Thevenin equivalent, and L and R represent the Norton equivalent where L = ^ = 25 = 3.6A. Ro 7 ALTERNATIVE SOLUTION The definition for R, in Norton's theorem is the same as in Thevenin's, again, R„ = R„6 = 7 Q. However, Norton states that if the terminals a and b are short-circuited, the current through that short circuit is X,. The short circuit is noted by the dashed line in figure 2.56. Using circuit reduction, the 2-Q resistance connected to terminal a is in parallel with the 10-Q resistance connected to terminal b, or (10) (2) 10 + 2 = 1.67 Q. The equivalent resistance as seen by the 50-V source is (50) (10 + 1.67) = 50 + 10 + 1.67 and the current from the source is Ix = 50 9.46 From current division, I 2 = (5.28) ( = 5.28 A. 50 50 + 10 + 1.67 ) = 4.28 A, and the current through the shorted terminals is 10 I. (4.28) (- -) = 3.6 A. 10 + 2 R„ and !«, again describe the Norton equivalent. Active network X K Active network External network ioCpRoi External network ° _J ° ~^ B C Figure 2.54. — Norton's theorem. VioRo I =^ l o EL 0_J Active network i = v, = Io(D i R o Figure 2.55. - Comparison of Thevenin's and Norton's circuits. 2n AAM — ?a 50 V Figure 2.56. - Circuit for example 2.17. 45 EXAMPLE 2.18 Determine the Thevenin's and Norton's equiva- lents for the circuit in figure 2.57. SOLUTION. In the branch containing the 900-V source, the two 5-Q resistances are in series. If these are combined into one 10-Q resistance, it should be quite obvious that two practical voltage sources ex- ist between junction 1 and the junction connected to terminal b. Source transformation can be employed to solve the problem. The resistance and magnitude of the ideal current source of the Norton equivalent to the 900-V and 10-Q source are R t = 10 Q, I 1 = ^0 = 90A. 10 For the Norton equivalent of the 2,250-V and 15-Q source, Rj = 15 S, j = 2,250 = 15Q A 15 Figure 2.58A shows the voltage sources transformed to practical current sources. Notice that junction 1 and the junction associated with terminal b still exist. Be- tween these two terminals, the 90- and 150-A sources are operating in parallel, and the 10- and 15-Q resistances are connected in parallel. Combining these ideal current sources and resistance results in the cir- cuit of figure 2.58S. Again, notice that the afore- mentioned junctions are retained. Converting the 60-A and 6-Q current source to its Thevenin equiva- lent produces the circuit in figure 2.58C. The 6- and 4-Q resistances in series with the 360-V are combined in figure 2.58D. The 360-V source and 10-Q resistance form a practical voltage source between terminals a and b, and this is converted to its Norton equivalent in figure 2.58.E. Here, simple combination of the two parallel 10-Q resistances yields one answer to the original problem and is shown in figure 2.58F. The remaining answer, the Thevenin equivalent, is in figure 2.58G, obtained by source transformation of figure 2.58F. 511 1 411 -tyW\r f ° a To summarize the preceding sections, the fundamental laws and parameters were first applied to circuits under the influence of dc. Expanding upon these laws, several circuit-analysis techniques and theorems were covered. Be- cause only dc was considered, resistance was the only cir- cuit element of interest. As will be shown shortly, most of this theory is also valid for circuits acting under current forms other than dc, where inductance and capacitance may also enter into the picture. TIME-VARYING VOLTAGES AND CURRENTS As the name implies, the magnitude of time-varying voltages and currents may not be constant with time. Con- sequently, the instantaneous values of the voltage and cur- 900 V | Figure 2.57. -Active circuit for example 2.18. 1 4A -WVW -OO -od I)9oa |ion (t)i50A fisn §ion ^b T)36A |io.n. lion I 1 1 — ob i 4n -awv- 60 A (f 36A(t) | 5 n ^>b 6n 1 4A 360 V 180 V( sn -WA od ^>b 360 V Figure 2.58. — Circuits illustrating solution steps to example 2.18. rent waveforms, v and i, must be considered. Both v and i are functions of time, as they were when originally intro- duced in this chapter, and they can assume any form from constant to the most complex. Figure 2.59 presents just a minor sampling of time-varying waveforms to illustrate their general characteristics. As with dc, the method for analyzing circuits that have time-varying current and voltage is first to form a model of the circuit, then to apply the fundamental laws and rela- tionships. Unlike dc circuits, a differential equation usu- ally results. To demonstrate the effect of time-varying and current on circuit elements, this section will first consider a special waveform, steady alternating current (ac). 46 An example of a steady-state ac waveform is provided in figure 2.60. The repetitive nature of this sinusoidal func- tion can be expressed mathematically as i = I„cos(cut) (2.79) where = current at any time, t, = crest or maximum value of current, a constant = radian frequency, rad/s. The term sinusoid or sine wave is used collectively to in- clude cosinusoidal or cosine-wave expressions. The above equation could also use a sine function, but the cosine is employed for convenience when referring to current. It can be seen in figure 2.60 and equation 2.79 that the instantaneous value of current repeats itself every 2n rad or 360°; that is, the waveform goes through one complete cycle every 2n rad. The number of cycles per second is cu/2tt which is defined as the frequency, f, of the waveform or f = or 2n CD = 2nf. (2.80) (2.81) The units of frequency are hertz (Hz). One hertz is equal to 1 cycle-per-second (cps), an expression whose use is now obsolete. The common power frequency in the United States is 60Hz, for which go = 377 rad/s, or just simply co = 377. A more general form of ac is where i = I m cos(cut + 6), 6 = phase angle. (2.82) Instead of expressing the phase angle in radians, such as rt/6, angular degrees, 30°, are customarily used. By adjusting 9, the sinusoid can be moved left (increasing 0) or right (decreasing 0. Such movement is illustrated in figure 2.61. Using the earlier technique of developing differential equations through circuit analysis, steady ac can be applied to pure resistance, inductance, and capacitance to observe what happens. Alternating Current Through Resistance Figure 2.62A shows a resistor of resistance R. From equation 2.79, if the current through this element is i = I m cos(cut), by Ohm's law, the voltage developed across the resistor is v = Ri or v = R(I m cos(cut)) = V m cos(o)t), (2.83) where V m = RI„ = maximum or crest value of voltage waveform, V. Figure 2.62B shows both voltage and current as functions of time. At every instant, v is proportional to i, and v and i are said to be in phase. When two sinusoidal waves are compared for phase in this manner, both must be sine waves or cosine waves; both must be expressed with positive amp- litude and have the same frequency. f 7T 27T 2 u 2 | ; 2 v 2 -cut V ■cut 7T 2tt 37T Figure 2.59. - Some time-varying electrical waves. -^m a / !b \ \ / ; \ T/2 / T t in seconds ; i W 1 277" cut in electrical radians \ 180° / 360° cut in electrical degrees Figure 2.60. - Sinusoidal ac waveform. i = I m cos cut Figure 2.61. -Steady ac showing phase shift. + A A B Figure 2.62. — Steady ac through resistance. 47 Alternating Current Through Inductance Suppose that current through the pure inductance of figure 2.63A is again as in equation 2.79. The voltage across the element is T di v = L — dt or or v = L — (I m cos(o)t)) dt v = LI, dcos(a)t) ' dt Differentiating, v = -a>LI m sin(cot) or v = wLI m cos(cot + 90°) = V m cos(ot + 90°), (2.84) where V m = coLI m = maximum or crest voltage. The term o>L is used so frequently that it is provided with a special name, inductive reactance, and is designated "X," where and X = coL =2nfL V m = L.X. (2.85) (2.86) Figure 2.63B compares equations 2.79 and 2.84, with i and v as functions of time. Here, it can be seen that the current crest is reached at a later time than the crest voltage. The current waveform is said to lag the voltage waveform by 90°. The phase angle is called lagging. Alternating Current Through Capacitance Consider the capacitance shown in figure 2.64A, and let the voltage across it be v = V m cos(a>t). (2.87) The current through the capacitor is then ,dv 'dt i = C^ or i = C-^-(V m cos(a)t)). dt Differentiating, i = -a>CV„,sin(o)t) or i = o)CV m cos(cot + 90°) = I m cos(cot + 90°), (2.88) where I m = coCV m = maximum or crest current through the capacitor. As with the inductive resistance, coC is also provided a special name, capacitive susceptance, and symbol, "B." Thus, and B = coC = 2TtfC L= BV m . (2.89) (2.90) The relationship between the current and voltage wave- forms (fig. 2.64B) is the reverse of the inductance situation; the current waveform is now leading the voltage waveform A B Figure 2.63. -Steady ac through inductance. Figure 2.64. — Steady ac through capacitance. by 90°. The phase angle is also called leading. The impor- tance of current and voltage waveforms being compared for lagging and leading phase angles will be brought out later in this and the next two chapters. Time-Varying Equations The preceding discussion considered voltage and cur- rent to be steady sinusoids. But what if they are allowed to have any form? To illustrate the consequences, the fun- damental laws and parameters can be applied to the sim- ple series RL, RC, and RLC circuits shown in figures 2.65, 2.66, and 2.67, respectively. For the series RL circuit, using Kirchhoff s voltage law V = V* + Vt. Substituting in the relationships for voltages across resist- ance and inductance, v = iR + L 'dt' (2.91) Now for the series RC circuits, v = v R + v c . Applying the elementary laws, v = iR + — /' idt + V . C o (2.92) The differential equations 2.91 and 2.92 are valid for any voltage and current, no matter what form. As before, V is the initial charge on the capacitance. Considering figure 2.67, which shows the series RLC combination, thus, v = v* + Vi + v c ; v = iR + L^- + - /' idt + V . dt C » (2.93) 48 To arrive at an equation that is easier to handle mathe- matically, equation 2.93 can be differentiated once: dv = R di_ + L d^ + 1 Q dt dt dt 2 C (2.94) This equation again describes or models the circuit for all electrical situations, as no restrictions have been placed on voltage and current. The preceding has shown that when voltages and cur- rents represent any form, the application of circuit rela- tionships results in a differential equation. Through clas- sical differential-equation methods, such equations can provide the required solution, but these techniques will not be shown because it can confuse the understanding of the vital aspects of electrical fundamental methods. 6 Figure 2.65. — Simple series RL circuit. V Q Transients and Circuit Response Solution of these equations for all situations yields the complete response of the circuit. For linear circuits, the solu- tion will have two parts: forced response and natural response. The forced or steady-state response can be attrib- uted directly to the applied source or forcing function. This is the action of voltage and current within the circuit if no changes or disturbances are made. The natural or transi- ent response is a characteristic of the circuit only, not a result of the sources. Such action occurs when a circuit is disturbed by a change in the applied sources or in one of the circuit elements. After the change, the circuit currents and voltages undergo transition from their original state to the point where their action is again steady state. The time period involved is normally very short, and the occur- rence within the transition is called a transient. For simplicity, the forcing functions mentioned earlier in this chapter were dc, and in network analysis the study was devoted only to resistive circuits and dc sources because here only the forced response is present. When both induc- tance and capacitance are circuit elements, both forced and transient responses can be encountered. However, knowl- edge of circuit transients is not required when considering steady-state voltages and currents, as was seen in the case of steady ac. By far the majority of mine power problems only require knowledge of steady-state circuit currents and voltages, and it will be shown that even though inductance and capacitance might be present, as long as only the steady-state response is considered the solution of differen- tial equations is not needed. However, transient circuit responses are an extremely important input in the design of mine power systems, and they will be explained in de- tail in_chapter 11. It has been shown in this section that any resistor, in- ductor, or capacitor carrying a sinusoidal current has a sinusoidal voltage developed across it. Furthermore, the sum or difference of two sinusoidal waveforms with the same frequency is another sinusoid. From these concepts, it can be shown that for a steady-state circuit, if voltage or current at any part of a linear circuit is sinusoidal (alternating at a particular frequency), voltages and cur- rents in every part of the circuit are sinusoidal with the same frequency. Figure 2.66. — Simple series RC circuit. STEADY ALTERNATING CURRENT The form of steady ac has already been shown and used in the analysis of simple ac circuits, but here the concepts of steady-state ac circuit analysis are introduced. This necessitates a review of a familiar but easily forgotten sub- ject, complex algebra. Real numbers such as 2, 4, and n are easy to understand in terms of physical things. Any mathematical operation on these numbers always results in another real number, except when the square root of a negative real number is taken. The term \/ -1 cannot be satisfied by any real number. Therefore, the square root of any negative num- ber is called an imaginary number. Mathematicians dis- tinguish imaginary numbers by writing "i" in front of them, but to avoid confusion with the symbol for current, electri- cal engineers use the symbol "j" where Figure 2.67. -Simple series RLC circuit. 49 Addition or subtraction of imaginary numbers yields another imaginary number. Yet, when an imaginary num- ber is added to a real number, a complex number is created. These have the rectangular form, x + jy (for instance, 3 + j4), where x is the real part and y the imaginary part or if then Z = x + jy, Re[Z] = x Im[Z] = y. (2.95) Complex numbers can be represented graphically by a pair of perpendicular axes as shown in figure 2.68. The horizontal axis is for real quantities, the vertical one for imaginary. Considering x + jy, if y = 0, the complex num- ber is a pure real number and falls somewhere on the real axis. Similarly, if x = 0, the complex number (now be- ing purely imaginary) exists on the vertical axis. Hence, complex numbers encompass all real and all imaginary numbers. In the case of the rectangular forms Z = x + jy, W = u + jy, the following common definitions and mathematical opera- tions of complex algebra are applied. 1. Two complex numbers are equal if and only if the real components are equal and the imaginary components are equal: Z = W, IFF x = u, y = v. 2. To sum two complex numbers, the real and imaginary parts are summed separately: Z ± W = (x ± u) + j(y ± v). 3. The product of a real and an imaginary number is imaginary: x(jy) = j(xy). 4. The product of two imaginary numbers is a negative real number: (jy) (jv) = -yv. 5. The multiplication of two complex numbers follows the rules of algebra (note, an easier way to perform the multi- plication will be shown): (x+jy) (u+jv) = xu -I- jxv + juy - yv = (xu - yv) + j(xv + uy). 6. By definition, the conjugate of a complex number is formed by changing the sign of the imaginary part. An asterisk denotes the conjugate: Z = x + jy becomes x - jy. 7. For division, the numerator and denominator are multiplied by the conjugate of the denominator (again, an easier method exists): x +jy = x + jy (u + jv) = ( xu - yv j + j( uy + xv } u — jv u — jv (u + jv) (u 2 + V 2 U 2 + V 2 Besides the rectangular, there are three other general forms of complex numbers: trigonometric, polar, and ex- ponential. Figure 2.69 illustrates the conversion of rectan- gular to trigonometric or polar forms where Z = x + jy. (2.96) The absolute value of Z is represented by "r," and x = rcos0, y = rsin0, = tan" 1 (£), x r = (x 2 + y 2 ) 1/2 . where x (2.97) Hence, the trigonometrical form of the complex number is Z = r(cos0 + jsin0), (2.98a) with the conjugate Z* = r(cos0 - jsin0). (2.986) The polar form is widely used in circuit analysis and is sim- ply written as Z = r|0 and the conjugate, Z* = r|H? Euler's theorem states that cos0 + jsin0 = e' 9 . (2.99a) (2.996) (2.100) J3 — J2 - 1 I J1 i , I 1 1 -3 -2 -1 -J1 -J2 -i3 1 2 3 Figure 2.68. -Graphical representation of complex number. REAL (y) Figure 2.69. —Trigonometric or polar representation of com- plex number. 50 This expression allows a complex number to be written as an exponent, the exponential form, Z = r(cos0 + jsin0) = re> 8 (2.100a) and Z* = re ■*. (2.1006) All four complex forms are therefore identical or Z = x + jy = Kcos0 + jsin0) = r|0 = re' 8 . The form should be selected that gives the easiest mathematical manipulation of complex numbers. For ad- dition or subtraction, the rectangular expression is best, but multiplication and division are much more convenient when the number is in exponential or polar form, the latter be- ing the most used. For instance, in polar, Z,Z 2 = rjfl r a |+ = r x r g | fl + , Z 2 r 2 |j> r 2 and in exponential, ZiZ 2 = (r,e^)(r ,e*) = = r 1 r 2 e /(8 -♦) z, r^ 8 r 2 e" £,(»-.)_ It will be shown shortly that circuits containing resist- ance, inductance, and capacitance can be represented by complex numbers, and that the solution of these circuits under steady ac will use complex algebra. This can be done with almost as much ease as the dc circuit analysis pre- sented earlier. EXAMPLE 2.19 Find the answer to the following expression in polar and rectangular form: (2 + j6)(18 |21°) . (1.63JM2.6 + jl) SOLUTION. Both the numerator and denominator of the above expression are multiplications of complex quantities. For ease of solution, the rectangular term should be converted to polar. This results in (6.32|71.6°)(18|21°) (1.63|90°)(2.79|21 o ) (6.32) (18) or or (1.63) (2.79) 71.6° + 21' 90° - 21° 251-18.4°. Effective Alternating Current The power available in the outlets of U.S. homes is a very familiar quantity: it is sinusoidal, having a frequency of 60 Hz and a voltage of 115 V. But what does 115 V actu- ally stand for? The voltage waveform, being a sinusoid, is not constant with time. Therefore, the voltage is certainly not instan- taneous. If a measuring device could be connected to an outlet in order to visually observe the waveform, it would be found that "voltage" is not the maximum value, V„„ be- cause this waveform crest is HSyT or 162.6 V. "Voltage" does not describe an average value either, because the average of a sine wave is identically zero. As another resort, the average throughout one positive or one negative half- cycle of the waveform could be calculated, but the result gives a measurement of 0.637 V m or 103.5 V. To discover the meaning of the term voltage, the reason for measuring the voltage must be considered. In any system, current and voltage are defined in terms of what they will do. Conse- quently, the voltage is the effective value of the sinusoidal waveform. It is a measure of the effectiveness of the volt- age source in delivering power to a resistive load. The effective value is called root-mean-square (rms). In order to understand rms measurements, it is neces- sary to return to the concept of instantaneous power, where If the power was being developed across a resistance, R, it was shown that and p = i 2 R v 2 P= R- These equations have little practical value for ac as they represent the value of power for a particular instant and in ac this is ever changing. A more effective measure for the value of power is based on the fact that power is the rate of doing work. A reasonable measure is then the ave- rage rate or average power. For average power, P, consumed by the resistance, R, P = ave(p) = ave(i 2 R) = (ave i 2 )R and Pv = ave — R (ave v 2 ) Average power is then an effective way to measure or quantify ac voltage and current. It has already been seen that the units of voltage and current in dc are easy to com- prehend; the magnitudes are constant with time, and their ability to deliver power is constant. Therefore, it is appro- priate to equate ac and dc rates of work, P oc and P<, c , re- spectively, in order to determine an effective measurement for alternating voltages and currents: p dc =PR = P oc = (ave i 2 )R or I 2 = (ave i 2 ) or I = V (ave i 2 ) = ™s current. Employing the same procedure, V = \/(ave v 2 ) = rms voltage. (2.101a) (2.1016) Current and voltage in ac are therefore expressed as the square root of the mean-square values or rms. They are sometimes written I rmj and V rTO . It can be shown from the voltage and current waveforms (that is, substituting 51 I m cos(cut + 9) into equation 2.101a and similarly for voltage) that I™ = —= or L = \T~2~ !"-. (2.102a) and V rms = v„ or V. (2.1026) Root-mean-square currents and voltages are used so often that they are directly implied when referring to an ac magnitude. They are almost always used in calcula- tions. For simplicity, the subscripts of V rmj and l rms are eliminated in practice, and just V and I are written to in- dicate rms voltages and currents. All common ac voltmeters and ammeters are also calibrated to read rms values. The preceding analysis of average power concepts ap- plies only to resistance. Average power in the steady state supplied to either a theoretically pure inductance or pure capacitance is identically zero. This can be proved by inte- grating instantaneous power to these elements to obtain an average. The results show that the energy received dur- ing one-half cycle is stored and then transferred back to the source through the balance of the cycle. The stored energy in the capacitance is greatest at the maximum of the volt- age wave, while in the inductance it is maximum at the current-wave crest. Phasors A steady-state sinusoidal current or voltage at a given frequency is characterized by only two parameters: ampli- tude and phase angle. This can be seen in figure 2.70A, which shows two voltage waveforms separated by a phase angle. An ac quantity may also be represented graphically by a phasor, illustrated in figure 2.70B. The phasor is a continually rotating line that shows magnitude and direc- tion (time). In this figure, the phasor is assumed to have a length representative of V m , rotation about point 0, and an angle increasing with time according to 9, = cot + 9. The figure shows the line as if a snapshot had been taken, freez- ing action. The alternating quantity, V m cos(a>t + 0), is the projection of the phasor on the horizontal axis. In other words, as the phasor in figure 2.70B rotates, a plot of its projection on the horizontal axis with time reproduces the waveform in figure 2.70A. The phasor length shown here represents crest voltage but does not necessarily need to be equal to it. It is common practice to draw phasors in terms of effective (rms) values. Although voltage has been employed as an example, phasors can also represent sinu- soidal current, among other things. Voltage and current phasors are both illustrated as rotating lines in figure 2.7 1A, where v = V m cos (cot + <{>), For example, in figure 2. 7 IB the phasor is shown where the current phasor angle is zero. Here, the current phasor is termed a reference phasor, and all other phasors are drawn relative to it. Either voltage or current can be selected as the reference. A phasor may be expressed in several ways. To illus- trate the most used expressions, consider figure 2.72A, which shows one phasor displaced from the horizontal by an angle, cot + 9. Recalling complex algebra, the horizon- tal axis can be assigned as a real_axis and the vertical as the imaginary axis. The phasor, V, is then the sum of the real and imaginary components, or V re and V lm , V = V„ + V im . (2.103) V m cos(cut+0) \ \ J I \ \ I I K Phase angle A B Figure 2.70.— Sinusoid versus time (A) and as phasor (B). \+8 B Figure 2.71.— Phasor representation of current {A) and voltage (B). i = I„cos cot. To show both current and voltage, two phasors can be drawn, with one of them advanced by the phase angle, <{>. Both lines rotate indefinitely about the axes, and one line will always lead the other in the same relative position; therefore, the axes are superfluous and need not be drawn. Since it is necessary to orient the phasors at a specific point in time, a convenient instant is selected as a reference. Vj m =Vsin(a>t + 6) A B Figure 2.72.— Other expressions for phasors. 52 Figure 2.12B clearly illustrates the rectilinear form of equation 2.103. The real and imaginary components of the phasor are V„ = Vcos(cot + 0), (2.104a) V )m = jVsinM + 0), (2.1046) Thus V = Vcos(a)t + 0) + jVsin (cot + 0) (2.104c) or V = V(cos(a>t + 0) + jsin (cot + 6)). (2.104 V = Ve^'e". (2.105a) (2.1056) The factor, e""', is superfluous, as it contains no unique information about the phasor, and it can be suppressed: Ve". (2.106) This is called the exponential form of the phasor. Thus equa- tion 2.105c can be expressed in polar form, V 10. (2.107) These phasor forms are very useful in solving ac cir- cuit problems. The terms phasor and vector are often interchanged. Phasors and Complex Quantities When introducing the action of time-varying sinusoids, certain voltage-current phase-angle relationships were found to exist for pure resistive, inductive, and capacitive circuit elements. In general, if a steady-state sinusoidal current has the time-domain form and voltage, I = I m cos(cot + 0), v = V m cos(cot -I- <}>), (2.108a) (2.1086) current is said to be lagging voltage by the phase angle, $— (or conversely, leading voltage by the phase angle, 0—$). Using the exponential and polar phasors, this current and voltage can also be stated I = Ie""' + •>, I = Ie", or I = I|0 (2.109a) and V = Ve""' + •», V = Ve", or V = I|$ (2.1096) where I and V = rms values of current and voltage, respectively. Before steady-state circuit analysis can be performed, pure circuit elements must again be considered, this time to analyze the voltage-current relationships using complex quantities. Equations 2.108a and 2.1086 are assumed to represent the general current through and voltage across each element. EXAMPLE 2.20 A circuit has the following voltage and current waveforms applied across and through its terminals: v = 282.8 cos (377t - 20°), i = 42.4 cos (377t + 25°). Write the phasor expression for voltage and current. What is the phase angle between current and voltage? SOLUTION. The two given expressions are in the time domain, where for the voltage, V„ = 282.8 V, 4> = -20°, and for the current, I m = 42.4 A, = 25°. The phasors for voltage and current are then, respectively, v v - I = i = 200 1-20° V, /T = 30 125° A. The current waveform is leading the voltage wave- form, and the phase angle between current and voltage is 4> - = -20° - 25° = -45°. If sinusoid current is applied to a resistance, R, the voltage across it is v = Ri. Applying the general time-domain expressions, V m cos(cot + 4>) = RI m cos(cot + 0) or in exponential form, y e ir(»rt») _ RIe><"" + <». Suppressing e""', Ve* = Rle", or in polar form, V 14 = RI |0. 53 In phasor form, V |£ and I |0 are the phasor polar representations, V = RI. (2.110) This is the same relationship that exists for time-varying waveforms and dc. It is apparent that angles and $ are equal and that voltage and current are in phase (fig. 2.73A). Suppose the same general forms of current and voltages were applied to a pure inductance where, as before, dt then, using the general exponentials, Ve"»<+*> = L — (Ie""" + '"). dt Differentiating (e"' is a constant with time), Ve/ (■«♦•> _ jajLIe"""* 9 ', and suppressing &"", Ve* = juLIe". Thus, in phasor form, V = ja>LI. (2.111) The imaginary operator, j, denotes a +90° displacement of voltage from current; such as illustrated in figure 2.73J3. In general, if the current phasor has an angle, 0, the voltage phasor angle, <(>, is + 90° for a pure inductance. For a pure capacitance, dv i = C dt Employing the same process to find equation 2.111, I = jwC V. (2.112a) or v = (t^-)! (-£D. (2.1126) In this case, -j indicates a —90° displacement of the voltage phasor from current, as shown in figure 2.73C. Now that the phasor relationships of the fundamental elements have been covered, the stage is set for impedance transforms. A very important quantity, impedance, signified by Z, is defined as the ratio of the phasor voltage to the phasor cur- rent for a circuit or Z = (2.113) This expression is often called Ohm's law for ac circuits. Impedance is a complex quantity with dimensions of ohms, but it is not a phasor. Therefore, the impedance of the pure passive circuit elements, resistance, inductance, and capacitance, are respectively Zr — R, jcoL, Z c = jcoC (2.114) These can be applied directly to circuit analysis when a cir- cuit is in steady state. In other words, element impedances are employed to convert or transform a time-domain circuit model into a form in which the circuit can be analyzed us- ing only complex algebra. Hence the expressions of equa- tion 2.114 are called impedance transforms, and the transformed mathematical model is then in the impedance (or joj) domain. As a result, no differential equations are used to solve a steady ac circuit. All previous fundamental theorems, laws, and circuit- analysis techniques also apply to steady ac circuit analysis using impedances. Thus, an ac circuit representation in the impedance domain is analogous to a dc circuit model. On the other hand, the concept of impedance has no meaning in the time domain with time-varying voltages and currents. To demonstrate these concepts, consider the simple RL circuit in figure 2.74A, now with a complex voltage source, that is, a steady-state sinusoid defined as a phasor. Here, the current through the resistance and inductance is the phasor, I; therefore, v R = Tr, V £ ^Tjo)L. VI I v 90° Pure resistance : V and I in phase Pure inductance: I lags V by 90° 90° Vt Pure capacitance: I leads V by 90° Figure 2.73.— Voltage-current phasor relationships for cir- cuit elements. Impedance Transforms The current-voltage relationships for the three fun- damental elements have been found using phasors, as V = RI, V = jcuLI, V = I jo>C - These can be rewritten as voltage-phasor to current-phasor ratios: y-R, V • T I j«C" Vr V l V L =jIX L Vr I B Figure 2.74.— Steady sinusoid analysis of simple RL series circuit. 54 By KirchhofTs voltage law, V = V R + V, or V = IR + IjwL = I(R + jcoL). The impedance (equivalent) of the entire circuit is then »-? R + jcoL. (2.115) Because impedance is a complex quantity, it also has a polar form: Z = |Z| If, (2.116) where |Z| = (R 2 + (wL) 2 )^ mag nitude of impedance, Q, and = tan-H— ). R A phasor diagram for the circuit current and voltages is given in figure 2.74B. Note that as current is common to both elements, it could be used as the reference phasor. Here, voltage across the resistor, V R , is in phase with cur- rent, while that across the inductor, V £ , leads current by 90°. The total circuit voltage, V, can be resolved noting that v = v* + v L = v |e, where V = ( \ R + V £ ) ^ magnitude of source voltage, V, and = tan-K— ). V This last angle is identical to that found for the impedance. It should be noted that the current and voltage relation- ships for the inductor are as those found previously when time-domain voltages and currents were considered. Now consider figure 2.75A, which shows a simple RC series circuit in which V* = IR, v c =T(^-)=-j-L, V= IR--^= I(R - -l_). and V = IR - -$- The impedance becomes T ojC R + jcoC (2.117) Figure 2.755, the circuit phasor diagram, shows the current- voltage phase-angle relationships with the voltage across the capacitor now lagging that across the resistor. Continuing the process for an RLC series circuit (fig. 2.76A), the voltage across each element is V R = TR. V, = IjwL, v c V R =IR I Vc=-jix<;t Figure 2.75.— Steady sinusoid analysis of simple RC series circuit. j V R \^_ V C -VWW^ou — 1( — | ■: R L C A B Figure 2.76.— Steady sinusoid analysis of simple RLC series circuit. and across the entire circuit, V = % + V L + v c or V = IR + IjwL + Kt^tt), jcoC (2.118) with the circuit impedance, or Z = R + jcoL + Z = R + j(coL jo)C o>C (2.119) The foregoing gives the essence of impedance transforms. Each impedance shown in equations 2.115, 2.117, and 2.119 is the equivalent impedance of that circuit and has the general form Z = R + jX = |Z| \9, where R = resistance component, Q, X = reactance component, £2, |Z| = (R 2 + X 2 ) % , magnitude of impedance, Q, and = tan'HX/R). Here, depending on the pure circuit elements, the reactive component is X = ojL = inductive reactance, Q, X = — — = capacitive reactance, Q, o)C X = ojL = reactance for series LC elements, Q. o>C 55 From this equation, it can be seen that resistance is con- stant while reactance is variable with frequency. The time-domain expression found for a general series RLC circuit can be used to clarify the transformation process: v = iR + L i + i / .' idt + v - (2.93) It has been demonstrated in the impedance domain for steady ac that V = TR + TjcoL + T— — . jcoC (2.118) Accordingly, time-domain differential equations can be changed to the impedance domain when the circuit is under steady ac by 1. Replacing v with V (in rms), 2. Replacing i with I (in rms), j 3. Replacing — with jco, dt 4. Replacing / . . . dt with — , and 5. Letting V = 0. However, it is a much more efficient approach to ac circuit analysis to assign the impedances directly using equation 2.114, and soon more will be stated regarding this. Admittance Admittance, which is given the symbol Y, is defined as the reciprocal of impedance, Z, and Y = l V (2.120) The units are now Siemens, replacing the previous designa- tion, mhos. Admittance is therefore a complex quantity, the real part being conductance, G, and the imaginary compo- nent susceptance, B, or Y = G + jB. (2.121a) It should be noted that conductance is not the reciprocal of resistance unless reactance is zero, likewise for suscep- tance, reactance, and resistance. In general form, through equating Y and Z, R R 2 + X 2 and B = -X R 2 + X 2 (2.1216) Admittance affords basically the same convenience in steady ac circuit analysis that conductance provides for parallel dc circuits. Steady-State Analysis As previously stated, all circuit-analysis techniques that were covered for dc circuits still apply to steady ac circuits in the impedance domain. These include network reduction, Kirchhoff s laws, loop and node^analysis, network theorem, plus delta-wye transforms. Impedances simply replace resistances in the concept, and steady ac sources replace dc. Even with dc, the impedance domain can be used; in other words, dc sources can be thought of as steady-state sinusoids with w = 0. Therefore, with dc, reactance has no effect. A summary of circuit relationships follows, this time including impedance. 1. Impedances in series. A single equivalent impedance, Z, is Z = Z x + Z 2 + Z 3 + + Z„. (2.122) 2. Impedances in parallel. A single equivalent here is (2.123) 1 = JL + JL + JL + .... + .L /j Zi} Zi 2 Zi 3 Zj„ 3. Admittances in parallel, Y = Y, + Y 2 + Y 3 + . . . Y„. (2.124) 4. Voltage distribution of series impedances, (2.125) where V is the input voltage, V x is across Z„ and so on. ZJi Zin V = —V V = —V V x Z V, V 2 z V, 5. Current distribution through parallel admittances, (2.126) _ Y,- - Y 2 _ I 1 = Y I ' l2 = Y 1 ' where I is the total circuit current, I x is through Y„ and so on. Or parallel impedances, Z_ - Z _ (2.127) The overlines are removed on the above impedances and admittances simply for convenience, but it should be remembered that all are complex numbers. In essence, ac circuits in the steady state can be solved almost as easily as dc circuits employing only resistance. The major addi- tion is that the solution now uses complex algebra. 56 EXAMPLE 2.21 Consider the circuit shown in figure 2.77, where v = 5,880 cos (377t + 53.1°), i = 141.4 cos 377 t. The circuit is under steady-state conditions. What are the values of R and L? SOLUTION. The phasor representations for voltage and current are V = ^HP 1 53.1° = 4,158 1 53.1° V, VT I= 14JL4 |()C \T2 - 100|0° A. The total impedance of the circuit is then = 41.58 153.1° Q V = 4,158 |53.1' T 100 |(P or in rectangular form, Z = 25 + J33.25 Q. The real part of this impedance must be the circuit resistance and the imaginary part equal to total reac- tance. Thus, R = 25 Q, X = 33.25 Q, but X = coL - 0.3. Therefore, as a> = 377 rad/s, 33.25 + 0.3 L = 377 0.09 H. — vwv vtO L ^h -jO.3 Figure 2.77.— Circuit for example 2.21. EXAMPLE 2.22 Find the voltage, V, across the 2-Q resistance in figure 2.78. SOLUTION. Circuit reduction appears to be the easiest way to solve the problem. Noting that u> = 377 rad/s, the reactances of the impedance and capacitance are X L = coL = (377)(0.12 x 10- 3 ) = 0.045 Q, 1 1 Xc = coC (377) (3,535 x 10" 6 ) = 0.75 Q. The impedance of the branch containing the im- pedance is Z t = R x + jXc = 1 + jO.045 Q for the branch with the capacitance Z 2 = R 2 + jX c = 1 - J0.75 Q. Combining these two parallel impedances in polar form, Z X Z 2 (1.0 1 2.58 "X 1.25 1 -36.87°) Z x + Z 2 2-J0.705 0.59 1-14.9° Q, and the equivalent impedance seen by the 1,000-V source is Z ea = Z + 0.59 1-14.9° = Z + 0.57 - jO.15 = 2.57 - J0.5 = 2.57 1 -3.4° Q. Assigning the source voltage as the reference phasor, the total circuit current is - V = 1,000 |01 = 388 | 3 . 4 o A Z„ 2.57 1 -3.4° The voltage across the 2-S2 resistance is then V = 21 = (2)(388|3.4°) = 777|3.4° V. V l — vwv i 2n 1,000 Vf A 60Hz \ty Figure 2.78.— Circuit for example 2.22. 57 EXAMPLE 2.23 Calculate the current, I, through the branch indi- cated in figure 2.79 using only loop equations. SOLUTION. Two loop currents have been assigned in the figure. Using Kirchhoffs voltage law, (5 - j5)I x - 5I 2 = 1,000|(T, -5I X + (5 + j5)I 2 = 800|0°_. The solution to these simultaneous equations gives l x = 360 + j200 A, I 2 = 360 - jl60 A, The current through the 5-Q resistance is then I = Ix - I 2 or I = 360 + J200 - 360 + J160; thus, I = j360 A or I = 360190° A. EXAMPLE 2.24 What are the Thevenin's and Norton's equivalents for the circuit shown in figure 2.80? SOLUTION. Applying either Thevenin's or Norton's theorem, the equivalent impedance of the circuit be- tween a and b with the internal source off is Z s . When the steady-state voltage source is off, it acts as a short circuit, and the 4-Q and 12-Q resistances are effec- tively in parallel, and (4) (12) 4+12 3 Q. This combined resistance is in series with the jlO-Q reactance, and the series combination is in parallel with the — j6-Q capacitance, and z = z (3 + jl0)(-j6) . ° b 3 + jlO - j6 ' thus, Z, = 12.5| -69.8° = 4.3 - jll.7 2. If the terminals a and b are shorted out according to Norton's theorem, a short circuit exists across the capacitance, and the jlO-Q impedance and 12-Q resist- ance are placed in parallel. The equivalent impedance of the circuit under this shorted condition as seen by the ideal voltage source is then Z„ = 4 + (12)( J 10) = 10.7| 33.5° Q. 12 + jlO The circuit delivered by the ideal source is V 25|0 C Ix = 10.7133.5' = 2.33-33.5 A, and the current through the jlO-Q reactance and the shorted terminals is I 2 = I„ or I, = 2.33 1 33.5 (- I = Ix L„ 12 .., 12 15.62 39.8° 12 + jlO ) = 1.8|-73.3° A. Z s and I s define the components of the Norton equivalent for the circuit in figure 2.80. By source transformation, V s = I S Z S = (1.8| -73.3° )(12.5l -69.8° ) = 22.5| -143.1° V, which is identical to V* = -22.51 36.9° V. Z s and V s define the components of the Thevenin equivalent. It can be noted in figure 2.80 that the ideal volt- age source is in series with the 4-Q resistance in one branch. Therefore, source transformation alone could be employed to solve the problem. The use of subscripts in this example did not follow the format previously used in the chapter. In other words, Z„ I„ and V £ described the equivalent circuits rather than Z , I , and V . The reason is that the subscript zero has a special meaning in three-phase ac circuits, which will be discussed in chapter 4. j5n or vf©2) iji 5 n (V ©| 80 °I0-° V I— H( * 1 Figure 2.79.— Two-loop circuit for example 2.23. 4fL jion 25 loivf© Figure 2.80.— Active circuit for example 2.24. 58 Chapter 2 has introduced the concepts of electrical cepts are fundamental to electrical engineering, regardless circuit analysis. The fundamental laws were covered first, of application. Thus, comprehension of the contents of this followed by numerous circuit analysis techniques, which chapter is vital to understanding the following chapters, were applied to dc circuits. Steady ac was then presented, The next chapter will continue the study of electrical fun- and the chapter concluded with examples of circuit analy- damentals, with emphasis on power consumption in ac sis on ac circuits under steady-state conditions. These con- circuits. 59 CHAPTER 3.— ELECTRICAL FUNDAMENTALS II The measures of instantaneous power, p, and average power, P, were introduced in chapter 2. Instantaneous power does not have application in steady ac circuit analysis, so the concept of average power has been devel- oped to gauge the rate at which electricity does work. This chapter continues to build the foundations for mine power fundamentals that will be expanded into full comprehen- sion in chapter 4. There, the discussion will focus on three-phase power; here, the purpose is to introduce single- phase power and transformers. AVERAGE POWER AND POWER FACTOR To find the average power consumed by a circuit, the resistance of each element can be examined and all the individual power consumptions computed. Reactance, ei- ther capacitive or inductive, does not affect average power. When all the average powers have been determined, their sum yields the total average power delivered to the circuit. Obviously, if the circuit elements are numerous, the pro- cess can be time consuming, but this approach is some- times necessary. If the average power needs to be determined for the total circuit, it would be more desirable to perform only one calculation by computing average power in terms of the terminal current and voltage in the circuit. Yet, when complex or imaginary components exist in the circuit, can they be ignored, as this implies? In other words, the voltage and current waveforms might not be in phase, and when a phase angle is involved, the product of effective voltage and current no longer equals average power. However, instantaneous voltage and current can be used to calculate average power and to demonstrate what occurs if a circuit has reactance. Assume that the following current and voltage are monitored at the terminals of a circuit: i = ImCOS wt, v = V m cos(a>t + 0). Current is taken as reference, and the phase angle by which voltage leads current is 0. The instantaneous power consumed is then p = vi = V m I m cos(wt + 0)cos tot. From the trigonometric identity for the product of two cosines, V L. p = -son (CO80 + cos(2wt + 0)) The first term of equation 3.1 is constant, while the second is a sinusoid. Thus, taking the average to find average power results in V I P = av(p) = -^ cos 0. (3.2) Realizing that V„ becomes = V2V and I m = V2I, average power P = VI cos (3.3) or P = V^ cos + V„Xn co8(2wt + 0). (3.1) in which V and I are root-mean-square (rms) voltage and current at the circuit terminals and is their phase angle. If the voltage and current had been dc values, the average power would just be the product of voltage and current. However, when voltage and current are sinusoidal, equa- tion 3.3 specifies that the average power entering any circuit is the product of the effective voltage, effective current, and the cosine of the phase angle. The function cos is called the power factor (pf). For a purely resistive load, the phase angle is zero and the power factor is unity. Unity power factor may also exist when inductance and capacitance are present, if the effects of reactive elements cancel. If the circuit is totally reactive (either inductive or capacitive), the phase angle is a positive or negative 90°, the power factor is zero, and average power must be zero. COMPLEX AND APPARENT POWER When there is reactance in a circuit, a component of circuit current is used to transfer stored energy. The energy is periodically stored in and discharged from the reactance. This stored energy adds to circuit current but not to average power because average power to reactive elements is zero. In such cases, the power factor is not unity. Thus, as no work is performed by the added current, the power factor can be considered to be a measure of circuit efficiency or its ability to perform work, and aver- age power, defined by equation 3.3, is often called active power or real power. Power calculations can be simplified if power is de- fined by the complex quantity shown in figure 3.1, which is expressed mathematically as S = P + jQ, (3.4) where S = complex power, P = real power, as before, and Q = reactive power or imaginary power. Imaginary power accounts for the energy supplied to the reactive elements. If P = VI cos0, 60 s i 1 >■ ir < -z. o < ^ /\ e I i REAL P Figure 3.1. —Power represented as real and imaginary com- ponents. then the magnitude of complex power, S, called apparent power, is S = VI, and imaginary power is Q = VI sine. (3.5) (3.6) Voltage and current are again rms, and is the phase angle. Therefore, or S = P + jQ = VI cos0 + jVI sine S = VI(cos0 + jsin0) = Vie* = VI|0. (3.7) Complex power is then simply the product of terminal rms voltage and current magnitude acting at a phase angle. Applying dc concepts, the product, VI, is the power appar- ently absorbed by the circuit, hence the term apparent power. Apparent power, real power, and imaginary power are di- mensionally the same, but to avoid confusion with real power (units of watts), apparent power has units of voltam- peres, and reactive power uses voltamperes reactive. When sinusoidal voltage and current have general form, as in V = V|0, I = I|0, instead of using equations 3.4 and 3.7, the following expression is more convenient for computing complex power: S = VI*, where V = complex voltage, V, and I* = conjugate of complex current, A. (3.8) Accordingly, or S = Ve^e*-* = Vie**-**, where - 4> = phase angle between voltage and current. EXAMPLE 3.1 When operating under normal conditions, an induction motor has been found to draw 100 A when 440 V is across its terminals. Current is lagging voltage by 36.87°. Find the average, reactive, appar- ent, and complex powers for this load. SOLUTION. From equation 3.3, the average power is P = (440X100) cos 36.87° = 35,200 W. Using equation 3.6, the reactive power is Q = (440X100) sin 36.87° = 26,400 var. Equation 3.5 defines the apparent power as S = (440X100) = 44,000 VA, and equation 3.4 yields the complex power as S = 35,200 + j26,400 VA. ALTERNATIVE SOLUTION. If voltage is assigned as the reference phasor, then V = 440|0°^ V, I = 100| -36.87° A. From equation 3.8, the complex power is S = (440|0°X100| -36.87° )* or S = (440|0°X1001 36.87° ) = 44,0001 36.87° VA, where the magnitude is the apparent power, or S = 44,000 VA. Converting the polar expression for complex power to a rectangular form, S - VI* = V\0I\-4> = VIl0-4> which yields S = 35,200 + j26,400 VA, P = 35,200 W, Q = 26,400 var. It should be noted that the above solutions are only two of the many possible. 61 EXAMPLE 3.2 A load consumes 1,250 kW at 0.6 lagging power factor when 4,160 V at 60 Hz is across it. The load is connected in series with a (0.71 + jO.71) fl impedance to a constant source. Determine the voltage and power factor at the source. SOLUTION. From the stated conditions, the aver- age power is P 1 = 1,250 kW. From equation 3.3, the current through the load is Pi Ii = V x cosd 1 ' where P x , V 1( and cos^! relate the conditions for the load, or I _ 1 > 250 > Q0Q _ sol A x i - (4,160) (0.6) " 0U1 A - For convenience, the voltage across the load can be assigned as the reference phasor, then V\ = 4,160 V, I = 501|^L = 5011 -53.1° A. The load current also flows through the series im- pedance. Using polar expressions, the voltage drop across this impedance is V 2 = I X Z 2 or V, = (501| -53.1 o Xl|45°) = 501| -8.1° V. The voltage at the source is then v 8 = V\ + v 2 = 4,160 + (496 - j71) = 4,656 - j71 = 4,6571 -0.9° V. The power factor at the source can be found by first calculating the phase angle between current and voltage at the source with the current phasor taken as reference. Here, 8 = Figure 3.3.— Circuit demonstrating sum of complex powers. The power-factor angle is = Tan- 1 ^ (P ! /85.42 - Tan ' ( :: ^) = 48.72°, and the power factor of the combination is pf = cos = cos 48.72° = 0.66. EXAMPLE 3.4 The maximum capacity of a piece of power equip- ment is rated by apparent power at 500 kVA. The unit is being loaded by 300 kW at 0.6 lagging power factor. The power factor must be improved to 0.8 lagging by adding capacitance in parallel with the equipment. Find the required capacitance in kilovoltamperes reactive. With the capacitance in place, find the reserve capacity that is available from the power equipment. SOLUTION. For the load on the equipment without the capacitance, P, = 300 kW, Pi 300 = 500 kVA, 1 cos0, 0.6 Q x = SiSinflx = 500(0.8) = 400 kvar. It can be said from S x that the equipment is fully loaded. When pure capacitance is added, average power remains constant, and only reactive power and apparent power change. For the desired power factor, cos0 2 , EL- gi COS0Q 300 0.8 = 375 kVA, Q 2 = S 2 sin0 2 = 375(0.6) = 225 kvar. 63 Consequently, the added capacitance causes the total reactive power to decrease. The difference between the reactive power without and with the capacitance must be the amount inserted by the capacitance. In other words, Q c = - Q = -(400 - 225) = -175kvar. The negative sign is used here to indicate that the capacitance adds negative reactive power to the system. Finally, the difference between the apparent power without and with the capacitance yields the reserve capacity available from the equipment, or = 500 - 375 = 125 kVA. It can be noted that additional average power can now be added to load on the equipment without exceeding its maximum capacity. For instance, con- sider that average power P will load the equipment so that the equipment is again operating at full capacity. Then, the total average power is P T = 300 kW + P. Reactive remains constant, Q T = Q 2 = 225 kvar, and apparent power changes to St = 500 kVA. Therefore, St = (P T 2 + Q T 2 ) 1/2 , 500 = [(300 + Pf + 225 2 ] 1/2 . Solving for the new average power, P = 146.5 kW. At w , the circuit is said to be in resonance, and 2 1 1 u ° = LC 0r '° o = (LC7 75 - Since a) = 2irf, the resonance frequency, f , is given by 1 f„ = 21KLC) 1 (3.13) For a series RLC circuit in resonance, it can be shown that 1. The applied voltage, V, and the resulting current, I, are in phase, 2. The power factor of the circuit is unity, 3. The impedance, Z, is minimum, and 4. The current, I, is maximum. At all other frequencies that are significantly higher or lower than f , the series RLC circuit appears as a high impedance. With frequencies below resonance, capacitive reactance is greater than inductive reactance, so the angle of impedance is negative (total reactance is negative). Above resonance, the situation reverses and the imped- ance angle is positive. This can be seen clearly in figure 3.5 where circuit impedance versus frequency is plotted. The energy stored in a resonance circuit is essentially constant, yet the energy level within the circuit may be many times higher than the energy being supplied from an external source during any period. The source itself does not supply any reactive power, only active power. The reactive power transfers energy back and forth between the resonant-circuit inductance and capacitance. The re- sult of this energy transferral can be very high voltages, several times the terminal voltage, existing across the inductance and capacitance within the resonant circuit. o— WV^TP — 1( — o RLC Figure 3.4.— Simple series RLC circuit for resonance. RESONANCE Series Resonance Earlier, the impedance for the simple series RLC circuit shown in figure 3.4 was found to be Z = R + jcoL + j-q = R + j( u L - ^ ). (2.119) A special circuit phenomenon can now be demonstrated with this equation. There exists one specific frequency, w , where total circuit reactance is zero and the circuit imped- ance is purely resistive, or w„C w„L - — = 0or w L = — p (3.11) 1_ w„C UJ o < Q LU Q. CD R x L = ojL Z = t/r 2 h ■X2 x = ojL- J_ x= 1 and Z„ = R. (3.12) FREQUENCY (co), rad/s Figure 3.5.— Plot of impedance magnitude versus frequency for series RLC illustrating resonance. 64 This situation can be the cause of some severe overvoltages in mine power systems, and the concept will be explored further in chapter 11. The amount of energy stored, compared with that dissipated by the resistance, is related to the shape of the curve representing impedance magnitude, as shown in figure 3.5. This curve is an example of a response curve. The quality factor of a circuit is a measure of the sharp- ness of the response curve and is expressed as a ratio: Q„ = 2* maximum energy stored per period total energy lost per period (3.14) where the period is one complete cycle of the resonant frequency. By finding the ratio of the energy stored in either of the circuit's reactive components to the energy dissipated in the resistance, it can be shown that Q = reactance resistance R „CR (3.15) The quality factor normally has greater application in the communications aspects of electrical engineering than in the power aspects. For instance, the width of the response curve is also related to Q and has great relevance to the tuned circuits used in radio and television. Parallel Resonance The resonance of the simple parallel RLC circuit shown in figure 3.6A is very similar to that just discussed. This circuit is obviously idealized, but its performance is of general interest. The admittance can be written as Y = G + juC + j^l = G + j( w C - ^ ), (3.16) and the circuit is in resonance when susceptance B is zero. Hence, the circuit exhibits low admittance and high impedance at resonance, while the series RLC circuit had low impedance and high admittance: B = a. C - — j- = or co C = — j- (3.17a) or Y = G. (3.176) On the other hand, the resonant frequency is again 1 f„ = Therefore, Q of this parallel resonant circuit is the dual of equation 3.15 or Q = susceptance « C R conductance G «„L (3.18) The concept also relates to many fundamentals covered in chapter 2. For example, two circuits are called duals if the loop equations for one have the same forms as the node equations for the other. Because figure 3.6A is idealized (as actual inducting elements must have associated resistance), figures 3.6S and 3.6C are presented to show practical circuits that exhibit parallel resonance. TRANSFORMERS Early in chapter 2, the concept of mutual inductance was introduced. To review, Faraday found that a time- varying current in one circuit would induce a voltage in a nearby circuit. If the adjacent circuits are simply conduc- tors and are labeled 1 and 2, as in figure 3.7, this statement means that • i x in circuit 1 produces v 2 in circuit 2, • v 2 in turn causes i 2 to flow (if circuit 2 is part of a complete loop), then • i 2 induces v x in circuit 1. These interrelated phenomena can be thought of as mag- netic coupling between the two circuits, and it has been shown that di 2 di x = L 12 — andv 2 = L 21 — where L 12 = L 21 = M = mutual inductance, H. C£ C~ 27KLC) 1 The statements previously given for series circuits also apply, except that current replaces voltage and voltage replaces current. This is an example of duality. Anything stated about a series resonant circuit applies to its dual, the parallel resonant circuit, if each word in the left column below is replaced by its opposite word shown in the right column: Series Parallel Voltage Current. Impedance Admittance. Resistance Conductance. Reactance Susceptance. Inductance Capacitance. Figure 3.6.— Circuits that exhibit parallel resonance. Flow of current causes magnetic field that cuts other conductor U fet V, Magnetic flux lines Figure 3.7.— Magnetic coupling between two conductors. 65 Because of the equality, M is used to represent mutual inductance. These equations are true only for straight wires, and magnetic coupling exists only if voltage and current are time varying. The circuits considered previously were loops or meshes composed of passive and active elements, and these were conductively coupled by common branches or nodes. The following paragraphs develop the concept of magnetic coupling further and introduce the fundamen- tals behind one of the more important components of ac mine power systems, the transformer. Transformers are prime examples of magnetic cou- pling. They are often designed to optimize this coupling, and their operation is based inherently on mutual induc- tance. Transformers are employed to increase the magni- tude of voltage for more economical power transmission or, conversely, to decrease the level to provide voltage more suitable for electrical equipment operation. In essence, these changes can be made with either total isolation or direct conduction between circuits. Instead of straight conductors, assume that two coils are situated side by side, and their magnetic action is passing through any environment (fig. 3.8). The current in coil 1 is then partly the result of self-inductance in coil 1 and mutual inductance from coil 2, and vice versa for coil 2. Expressed mathematically: _ d^ _ , di 2 dii , dio = « — o o =- i : ) < , Z o • A B Figure 3.9.— Demonstration of coil winding sense. Ml Li N 2 L 2 _ i, . • • i 2 f N U|IC N 2 | v 2 Polarity V,[ L 5 t|_ | v 2 change is '[ I 2 for equation M 3.20 only A Actual winding sense B Dot notation Figure 3.10.— Dot convention for mutual inductance sign. 66 The equations just presented are valid for any voltage or current waveform. If the currents are sinusoidal and have a radian frequency, w, transforms can be employed so that for equations 3.20c and 3.20d, Vi = juL^ - jcoMI 2 , V 2 = -jwMlj + jwL 2 I 2 . (3.21a) (3.216) These relationships can be used to analyze circuits con- taining magnetically coupled elements. It should be stated that equations 3.20 and 3.21 relate only to the magneti- cally coupled elements; equations for complete circuits containing these devices will follow. IDEAL TRANSFORMER The level of mutual inductance, M, depends upon the spacing and orientation of the coils and the permeability of the medium between them. In other words, M is a function of the magnetic flux linking between the coils. More will be said about this phenomenon later in the section. In figure 3.10A, by comparing the power entering L x of the circuit with that stored or available in L 2 , it can be proved from flux-linking concepts that M < (I^L,) 1 (3.22) Consequently, M has an upper limit defined by the geo- metrical mean of the two inductances involved. The ratio of M to its theoretical maximum is called the coefficient of coupling. This is by definition k = M (L^) 1 (3.23) where k can range from zero to unity. Coils having a low coefficient of coupling are said to be loosely coupled. Here the coils could be far apart or have an orientation such that little magnetic flux interacts between them. Loosely coupled circuits may have a k that ranges between 0.01 and 0.10. For tightly coupled circuits, such as air-core coils, k can be around 0.5. A power transformer is a device having two or more tightly coupled coils or windings on a common iron core. The coils are wound and oriented to provide maximum common magnetic flux and can have a coefficient of coupling very close to 1.00. The usual range is 0.90 to 0.98. Resistance and other power losses are small. The winding receiving power is called a primary; that delivering power is called a secondary. In the circuit in figure 3.10, h 1 is the primary and L 2 is the secondary. An ideal transformer is an idealized form of transformer where k = 1 and losses within the device are zero. Hence, an ideal transformer can deliver all the power it receives. Many useful relation- ships for real transformers can be obtained by assuming the ideal transformer case. The self-inductance of a coil has been shown to be proportional to the square of the number of turns forming the coil (N), provided that all the flux, created by the current in the coil, links all the turns (see chapter 2, "Inductance"). If a sinusoidal current, I, flows in a coil of N turns, then the voltage produced across an N-turn coil must be N times that caused in a 1-turn coil. Further, for a sinusoidal voltage, V, which is constant across an N-turn coil, the current allowed through must be 1/N times that caused in a 1-turn coil. Both these statements can be proved by magnetic field concepts, again assuming that all magnetic flux produced in a coil links all turns. It follows that for an ideal transformer with two windings: L, Nf N5 N, Vx N/ I 2 N x * (3.24) (3.25) (3.26) where N x = number of turns in primary winding, N 2 = number of turns in secondary winding, L 1; l lf V x = primary winding inductance, rms cur- rent, and rms voltage, respectively, and L 2 , I 2 , V 2 = secondary winding inductance, rms cur- rent, and rms voltage, respectively. For this two-winding arrangement, the voltage and cur- rent can be complex sinusoids. The turns ratio of the transformer, a, is defined as the ratio of the number of turns in the secondary winding to the turns in the primary winding: a = N2 N x (3.27a) Hence, for an ideal transformer, W y, i 2 (3.276) In other words, the sinusoidal voltages across the primary and secondary windings are in direct proportion to the number of turns of the windings, and the currents are related inversely to the turns. In addition, the last equa- tion shows that the apparent power at the primary and secondary windings is indeed equal: VA = v 2 i 2 . The magnitude of this power in voltamperes is specified for the maximum allowable or rated capacity of power transformers. Another useful transformer relationship can be deter- mined through a demonstration of steady ac circuit anal- ysis with magnetically coupled circuits. Consider figure 3. HA, where a sinusoidal voltage source, V s , with an internal impedance, Z g (the combination is the Thevenin equivalent for a source), is connected to the primary of an ideal transformer. The secondary delivers power to a load impedance, Z L . The vertical lines between the transformer windings indicate that the core is made of iron lamina- tions. The turns ratio above the transformer symbol, l:a, relates a convention of N 2 to N x . A very useful relationship is the ideal-transformer input impedance with the load connected, that is, the load 67 Ideal transformer B Figure 3.11.— Demonstration of impedance transfer in transformers. or rearranging, 7 _ 7 J^LiZ L ^-^Z^jcoa 2 ^ Now allowing L x to tend toward infinity, the input imped- ance for the voltage source becomes Z in = Z B + -| (3.28) that the source sees through the transformer. Loop equa- tions can be used to solve the problem. Two loops, \ x and I 2 (both express complex currents), are available in the circuit; the loops are magnetically coupled through the transformer. Employing Kirchhoff s voltage law for loop 1, V s = I^g + IJuL,! - LjuM, and for loop 2, O = -IjjwM + I 2 Z L + LjcoL. M is again the mutual inductance. Notice that current enters the dot of the primary and leaves the dot on the secondary, making the sign of M negative. Rewriting these into standard loop-equation form gives V s = IjfZg + j
    L 2 ). Solving for I ls V 8 = IiCZ. + j«L x ) + U ) Z L + jwL, Therefore, the impedance seen by the source, Z in , is the ratio of the source voltage to terminal current, or V w 2 M 2 Ii Z L +jwL 2 but then M 2 = L X L 2 Z in = Z_ + jwLi + w L X L 2 Z L + jcoL 2 There must be total coupling between primary and sec- ondary windings for an ideal transformer; thus, the self- inductances, L x and L 2 , have no effect in the circuit, and their value can be considered infinite. Notwithstanding, the ratio is still finite, as specified by the turns ratio: L 2 = a L x . For this reason, primary and secondary inductances are conventionally not specified on ideal transformers. When this is related to the input impedance expression, Z in = Z g + jcoLi + co 2 a 2 L 1 2 ZL+joa 2 !^ ' Equation 3.28 is significant as it shows that the source sees the load impedance, Z L , through the trans- former as Z L /a 2 . This means that an ideal transformer has the capability to change an impedance magnitude. There- fore, to assist in circuit analysis, the circuit in figure 3.11A can be redrawn to its equivalent, shown in figure 3.1 IS. Here, the impedance connected to the secondary is trans- formed to the primary. Obviously, the reverse process, primary to secondary, also holds, but the impedance is multiplied by a 2 . The impedance angle remains constant in either situation. EXAMPLE 3.5 A 60-Hz single-phase transformer has a rated capacity of 250 kVA and a turns ratio of 15:1. Assuming that the transformer is ideal, find the primary voltage if the secondary voltage is 480 V. What are the magnitudes of primary and secondary currents with these voltages applied and the trans- former operating at full capacity? SOLUTION. For the turns ratio of 15:1, a = 15 As the turns ratio specifies the secondary voltage to the primary, a = Y2 and the primary voltage is V Vj = — = 15(480) = 7,200 V. a The primary current for 250 kVA at 7,200 V is _ 250,000 1 " 7,200 " 6 ° A ' and the secondary current for 250 kVA at 480 V is _ 250,000 _ 12 " 480 " ° 68 EXAMPLE 3.6 Consider that the circuit shown in figure 3. HA has the following parameters: Z g = 6 + j3 n, Z L = 1 + jO.5 «, V 8 = 7,200 V, 60 Hz, Turns ratio = 12:1. Find the value of the load impedance (Z L ) referred to the transformer primary, the complex power at the source, the transformer secondary voltage and cur- rent, and the required transformer capacity. SOLUTION. For the specified turns ratio, 1 a = 12" Transferring the load impedance to the primary, \ = (12m + jO.5) = 144 + J72 «, which is the impedance referred to the transformer primary. The total impedance seen by the source is then Ze, = Z g + = 6 + j3 + 144 + j72 = 150 + j75 «. Assigning the source voltage as the reference pha- sor, the transformer primary current is Ii = 7,200 1 0J = 42.91 -26.6° A. The transformer secondary current is I 2 = - = 12(42.91 -26.6° ) = 515| -26.6° A, a and the secondary voltage is V 2 = I 2 Z L = (515| -26.6° X1 + jO.5) or V 2 = (5151 -26.6° X1-12| 26.6° ) = 576|0^ V. The complex power delivered to the load is then S = V 2 I 2 * or S m (576|0°X5151 26.6° ) = 296| 26.6° kVA. This may also be found from S = I 2 Z L = (515m + jO.5) = 265 + J133 kVA or S = 296126.6° kVA. Finally, the apparent power demanded by this load is the required transformer capacity, 296 kVA. ACTUAL TRANSFORMERS In actual transformers, a source must furnish the power dissipated by the secondary load plus the power needed to operate the transformer. The additional power is created from losses within the transformer circuit. The transformer capacity, the amount of power it can handle, is dependent upon the character of these losses, which are dissipated as heat in the core and the windings. Because excessively high temperatures are destructive to insula- tion, the capacity is limited by this rise in temperature, usually specified as an allowable temperature rise above ambient conditions. The major losses in an iron-core transformer are winding resistance (conductor loss), leakage reactance, eddy-current loss, and hysteresis loss. This section will expand upon the ideal-transformer concept to produce a transformer equivalent circuit that accounts for these losses and is a good approximation for real-world trans- former performance under any condition. Conductor Loss As the conductors used for the transformer windings have resistance, current flowing in the primary and sec- ondary produces an I 2 R power loss that creates heat. The loss is minimized by conductors with larger cross sections, but if the resistance is too large to be neglected, primary resistance, R l7 and secondary resistance, R^, can be placed in series with the ideal-transformer windings as shown in figure 3.12. Leakage Reactance For the ideal transformer, all the flux produced by the primary must link with the secondary winding. In the real world, however, a small percentage of the total flux pro- duced fails to link all the secondary turns; this is called leakage flux. Leakage flux can be reduced by placing the primary and secondary windings very close together, per- haps interleaving them. Further reduction comes from Figure 3.12.— Ideal transformer with winding resistance in- cluded. 69 winding the coils tightly on the core and providing a short magnetic path between them, thus creating a low- reluctance path between the coils. Nevertheless, even with the best transformer designs, leakage is significant and cannot be neglected. Inductance is the ratio of flux linkage to the current producing the flux, or L = Nd = flux linkage of circuit, Wb, and di = current producing flux, A. For transformers with iron or ferromagnetic cores, current and flux do not have a linear relationship, and differen- tials must be used. Consider the time-varying primary current, i lt in figure 3.13A, where the changing current produces the magnetic flux, <^> 1 , and Li = Nid0i dix (3.30a) The part of 1 that links the secondary is 12 ; that which only links the primary (or is lost in terms of magnetic coupling) is 0l X , where Similarly, although not shown in the figure, N 2 d L 2 = N 2 d<£ 21 N 2 d<^. 1 di.. die (3.31a) (3.316) (3.31c) Interestingly, the coefficient of coupling is also related to flux by k _^12 01 021 02 (3.32) The first term in equations 3.30c and 3.31c is the transformer mutual inductance, and the second terms are the primary and secondary leakage inductances, L L1 and Ll2> respectively, or M = Nid^ = N 2 d = (712.5X0.6) = 427.5 kvar. The difference between this new or improved reac- tive power and that without the capacitance is the reactive power less the capacitance, or Qc = (Qt " Qnew) = (661 - 427.5) = 233.5 kvar. It can be noted that this example is much like example 3.4. The concept of power-factor improve- ment has been repeated here to show the similarity of most power problems, be they single phase or three phase. THREE-PHASE TRANSFORMERS Considerable background information about trans- formers was presented in chapter 3, and most of that theory is also applicable to three-phase transformers. The prime purpose is the same as with single-phase systems, to provide the different voltages required for distribution and equipment operation. The transformer can be constructed as either a single three-phase unit or a bank of three single-phase units. The only difference between the two is that the three-phase unit has all windings placed on a common core. The connections can best be described by considering a bank of three single-phase two-winding transformers. Every coil is insulated from the rest, and there are three primary and three secondary windings, all of which can be interconnected independently. The primary and secondary windings can be delta or wye connected while complete electrical separation is retained between all the primary and secondary windings. The possible connections are wye to wye, delta to delta, delta to wye, or wye to delta. Figure 4.10 illustrates the physical connections of each combina- tion, and figure 4.11 shows the corresponding symbols used in the three-phase circuit diagram. Any one of these combinations can be found in or about mine installations, but mine power transformers are typically delta to wye. Delta-to-wye connections are popular in mine power systems because of the load advantages of the delta connection of the primary, which is in essence the load for the incoming power. The neutral of the wye-connected secondary provides a good grounding point for the outgo- ing system from the transformer and does not shift poten- tial under unbalanced load conditions. The delta-wye winding combination does not generate third-harmonic (180 Hz for 60-Hz systems) voltages and currents that hamper delta-delta and wye-wye connections. The second most popular transformer configuration in mines is the delta to delta. Although system s requiring a grounding neutral point create some difficulty for the delta secondaries, the delta-delta connection has one sub- stantial advantage. If one of the single-phase transformers fails, operation can be continued by removing the defective unit and operating the two remaining transformers as Wye- to- wye Delta -to -delta ? a' ) b Delta -to -wye Wye -to -delta Figure 4.10.— Three single-phase transformers connected for three-phase operation. Phase Wye -to -wye Delta- to- delta Neutral Delta- to- wye Wye -to -delta Figure 4.11.— Three-phase diagrams for the transformers of figure 4.10. open delta. This open-delta or V connection can be illus- trated by the two single-phase transformers shown in figure 4.12. Although it is an unsymmetrical connection, it does provide a symmetrical three-phase power input and output. However, using the two transformers in this man- ner reduces capacity to 57.7% of the three-transformer kilovoltampere rating. Nevertheless, it is an effective emergency measure. The open-delta configuration is some- times used as a temporary circuit; for example, when the completion of delta is postponed until load conditions warrant a third unit. 83 Figure 4.1 2.— Open-delta three-phase transformer operation. Calculations with delta-to-delta and wye-to-wye trans- formers are straightforward and easy to comprehend. With delta to delta, primary line-to-line voltages and phase currents are transformed to secondary line-to-line and phase values, while for wye-to-wye transformers, line- to-neutral voltages and line currents transfer directly. Delta-to-wye and wye-to-delta combinations are different. With a delta-to-wye configuration, primary line-to-line voltages become secondary line-to-neutral, and primary phase currents transform to secondary line currents. Through this, the current and voltage for all three phases shift in phase by 30° across the transformer. EXAMPLE 4.3 The main substation at a mine contains a delta- delta connected transformer bank composed of three identical single-phase transformers. With rated volt- age applied, a 6,000-kW load at 0.8 lagging power factor is causing the transformer bank to be fully loaded. The rated primary and secondary voltages of each single-phase transformer are 36 kV and 7.2 kV, respectively. 1. Find the capacity of each single-phase trans- former in the bank. 2. What are the magnitudes of the primary and secondary currents in each single-phase transformer? 3. What are the magnitudes of the primary and secondary line currents to and from the transformer bank? SOLUTION: The problem states that the trans- former bank is fully loaded by an average power, P T . Thus, the capacity or apparent power load, S T , of the bank is available from P T = St cos0, The capacity of each single-phase transformer, S P , one-third the total bank capacity, or fc>T = OOp IS and Sp = L 5 ™ = 2,500 kVA. If the transformer is assumed to be ideal, current and voltage in the primary or secondary are related to apparent power by Sp = Vplp. Hence, the primary and secondary currents in each single-phase transformer are V V pl 2,500 36 = 69 A and ±.± m m m ,; v. p2 7.2 It can be noted that these currents also correspond to the transformer turns ratio, a, which is 1/5 or 0.2. Because the transformer bank is delta-delta con- nected, the currents in each transformer are also the phase currents in the bank. Therefore, the line currents to and from the bank are, respectively, and I L = V3 I P I L1 = V3(69) = 120 A, I L2 = V3(347) = 601 A. BALANCED THREE-PHASE CIRCUIT ANALYSIS By definition, any element in one phase of a balanced (or symmetrical) system is duplicated in the other two phases. In other words, currents and voltages for the other phases are equal in magnitude but displaced symmetri- cally in phase position. Therefore, the analysis of voltage, current, impedance, and power in one phase can provide complete knowledge about the entire three-phase system. In addition, reactions between phases, such as phase curren ts, line-to-lin e voltage s, or line-to-line connected impedances, may be represented by an equivalent line or line-to-neutral value by using delta-wye transformations. The solution technique is called per-phase or single-phase analysis. The technique has wide application because almost all three-phase power systems that are operating normally are approximately balanced. As a simple demonstration of the concept, consider figure 4.13A, which illustrates a wye generator connected through line resistance to a wye-connected load. In figure 4.13.B, one phase of this circuit is extracted, and here • V is one leg of the wye-connected source, • R is the line resistance per phase, • Z is a "single-leg" impedance of the wye-connected load, and • The unconnected points, n and n', are the neutrals of the source and load, respectively. For the balanced system, the vectorial sum of all three line currents is zero. Hence, the current between n and n' is zero, and the potential at n equals that at n'. Accord- ingly, the two neutrals can be joined as shown in figure 4.13C This last diagram is the single-phase equivalent 84 circuit, single-phase diagram, or per-phase reduction of figure 4.13A. It should be noted that figure 4.13C is indeed a basic power circuit, similar to figure 4.1. Reduction of circuits containing delta-connected sources and loads is almost as easy, but one additional step is involved: the application of delta-wye transformation. Figure 4.14 demonstrates a simple example. Here, all sources and loads must be wye connected. The aim is to convert delta connections to wye using equation 4.9, and line-to-line voltages and phase currents to line-to-neutral and line, respectively. Thus, figure 4.14S is the per-phase equivalent of figure 4.14A. The simplified representation of the balanced three- phase circuit can now be analyzed, employing all the single-phase techniques previously discussed. When the solution is found, the three-phase parameters can be determined by reversing the reduction. This need only be performed when line-to-line or phase values are required; no changes are necessary with line-to-neutral and line parameters, as can be seen in figure 4.13. EXAMPLE 4.4 A load has a balanced delta-connected imped- ance of 5 1 45° per leg. This load is connected through three balanced line impedances of 1 + jl Q to a three-phase source that has a line-to-line voltage of 500 V. What is the magnitude of line current deliv- ered to the load? SOLUTION. This problem is basically the same as that for the circuit in figure 4.14, except here a line impedance exists between the source and the delta- connected load. As a per-phase solution is called for, the delta load must be transformed to an equivalent wye: Z A = 5|45^ n, Z Y = ^.5l^l.i.67|4610, or Z Y = 1.18 + jl.18 Q. The per-phase equivalent impedance as seen by the source is simply the sum of the line impedance and the equivalent wye impedance of the load, or Z eq = Z L + Z Y = 1 +jl + 1.18 + jl.18 = 2.18 + J2.18 = 3.08 |45_1 Q. The magnitude of this impedance divided into the line-to-neutral voltage across any one phase yields the answer. Il " IZ Ln eql 288 3.08 500 V3 3.08 = 94 A. |v an | = |v bn |=|v cn | = Rq = Rb = Re = R z a = z b = z c = z Figure 4.13.— Per-phase reduction of wye-to-wye system. Ra " Rb " Re " R Z a = Z b = Z c = Z Figure 4.14.— Per-phase reduction of delta-to-delta system. EXAMPLE 4.5 A three-phase 200-hp induction motor has a full-load efficiency of 90%, power factor of 0.85 lagging, and a rated terminal voltage of 950 V line- to-line. Find an equivalent delta-connected imped- ance for the motor when it is operating at full load under rated voltage. SOLUTION. Perhaps the best way to start this solution is to find the per-phase average power consumed by the motor under the stated conditions. The total average power input to the motor can be calculated from P T = (0.746) hp where hp is the motor horsepower and r\ is its efficiency. Thus, P T = (0.746X200) 0.9 = 165.8 kW. As single-phase analysis is desirable, the power consumed by each element of the equivalent wye- connected load is needed: 165.8 P -^ = ^ p " 3 3 = 55.3 kW 85 Equation 4.12 can now be used to find the line current to the motor, or Pp = V Ln I L cos0, II V t „ cos0 55£60 (950/VF) (0.85) = 118.5A. The line-to-neutral voltage divided by this line cur- rent is the magnitude of each leg of the equivalent wye-connected impedance for the motor, and the impedance angle is identical to the power-factor angle. Therefore, z Y = ^ \e 950/V3 118.5 cos _1 0.85 = 4.63131.8° fi; as the equivalent delta-connected impedance is re- quested, J Y Z A = 3Z V = 13.9|31.8 C or Z A = 11.8 + J7.3 a EXAMPLE 4.6 A production shovel, operating at full load, uses 1,200 kW at 0.9 lagging power factor with 3,750 V line-to-line at the machine. The shovel is supplied through a trailing cable that has an impedance of 0.04 + jO.03 fl per phase. If the voltage at the source side of the trailing cable is maintained constant, what is the voltage regulation at the machine? SOLUTION. The per-phase equivalent circuit for this problem is again similar to figure 4.145. The equivalent impedance of the shovel is not necessary, but the line currents and line-to-neutral voltage conditions for full load and no load are. Those for full load are p p = -2 = 400 kW, p 3 Vfl = •M?T — ' fe = 2,165 V line-to-neutral, 400,000 (2,165X0.9) = 205.3 A. The line-to-neutral voltage of the constant source, Vnl, can be found by computing the voltage drop across the trailing cable, V^, then adding it to the voltage at the machine. The full-load voltage at the machine can be assigned a zero phase angle. Thus, based on the given power factor, V FL = 2,165|0°_ V, I FL = 205.31 -25.84° A. The voltage drop across the trailing cable is V te = I FL Z te = (205.31 -25.84° ) (0.05| 36.9° ) = 10.271 11.06° V. The voltage at the constant source is = 10 + v FL j2 + 2,165 = 2,175 + j2 = 2,175|0O^ V. The subscript for this voltage is used to signify that for these conditions it is the no-load voltage at the machine. In other words, under no load, line current through and the voltage drop across the trailing cable will be zero, and the voltage at the machine will be the same as at the source voltage. Conse- quently, the voltage regulation is VR = V FL 2,175 - 2,165 (100%) 2,165 (100%) = 0.5%. The solution may be difficult when there are delta- wye or wye-delta transformers in the system because of the 30° phase shift of voltage and current between the wind- ings. In other words, a line-to-neutral transformer second- ary voltage transfers to a line-to-line primary value and vice versa. Obviously, when the three-phase system is not balanced, the per-phase reduction method cannot be ap- plied, and other more specialized techniques are required. These are discussed at the end of this chapter. One-Line and Three-Line Diagrams It is now apparent that practically all three-phase circuits consist of three conductors, three transformers, and so forth. When all these components are shown in a schematic, the drawing is called a three-phase diagram, as given in figure 4.15. Such diagrams can be extremely helpful, especially when the circuits are concentrated in a piece of machinery or power equipment, because they allow a complete view of component interconnections. They are imperative in manufacturing or troubleshooting. However, when the circuits are large, as in a complete mine power system, three-line diagrams are not only 86 cumbersome to draw but also exceedingly difficult to read and interpret. Furthermore, if the power system is nor- mally balanced, a three-line diagram is unnecessary since the system is always solved as a single-phase circuit composed of one line conductor and a neutral return. In these cases, the three-line diagram is replaced by a one- line diagram in which the interconnections or conductors between components are represented by single lines plus conventional symbols. This is a further simplification of the per-phase diagram because the completed circuit through the neutral is omitted. One-line diagrams are an invaluable tool in analysis, in designing new power sys- tems, or in modernizing existing ones. Furthermore, since circuit diagrams of coal mine power systems are required by Federal law (30 CFR 75, 77), a one-line diagram is the most practical way to comply. Figure 4.16 shows a one-line diagram designed to convey in concise form the significant information about the system shown in figure 4.15. In such diagrams it is usually implied that all information is per-phase, unless stated otherwise. Hence it is vital to remember that each device shown is actually installed in triplicate. The con- ventional notations are line or line-to-neutral impedance, line current, and line-to-neutral voltage. If line-to-line values are listed, they should be stated as such. Every line, symbol, figure, and letter has a definite meaning. Conse- quently, when a one-line diagram is constructed, specific conventions (1-2) 1 must be followed to ensure that the result will be complete, accurate, and correctly interpreted by anyone. The following guidelines are recommended. Relative geographic relationships for the power- system components should be maintained as far as prac- tical. The typical mine map provides an excellent layout guide for mining applications. Because of the shorthand form and definite meaning of every entry, duplication must be avoided. Standard numbers and symbols are mandatory, and those commonly used in mining are shown in figures 4.17 and 4.18 and tables 4.1 and 4.2. Many of the devices listed have not yet been covered but are included here for completeness. All known facts about the power system should be shown on the diagram, including on), • Apparatus ratings (volts, amperes, power, and so Ratios and taps of current and potential transformers, Power-transformer winding connections, Relay functions, and Size and type of conductors. The amount of information shown depends on the purpose of the diagram. For instance, if the diagram is to assist in studying the power flow to loads throughout the system (a load-flow study), the location of circuit breakers and relays is unimportant. However, for the solution of other power problems, complete knowledge of these de- vices can be mandatory. It is important that the one-line diagram contain only known facts; implications and guesses can lead to disastrous results. On many one-line diagrams, knowledge of future electrical plans is very helpful, and this information can be entered either diagrammatically or through explana- tory notes. Finally, the diagram should include correct title data so that the installation is clearly identified and cannot be confused with another or portion thereof. Z, =0.6 + j0.6 Z 2 =0.07 + j 0.05 Y-Y Transformer Figure 4.15.— Three-line diagram. O Line 0.6 tj 0.6 l:a Line 0.07 + j 0.05 1,385/277 Z=0.1+j0.3 referred to high side V=277V If machine, could be circle symbol 1 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. Figure 4.16.— One-line diagram of circuit shown in figure 4.15. 87 ■«■ ■» Air circuit breaker, removable type DH Circuit breaker, nondrawout type (oil or vacuum) «-0^ Air circuit breaker, drawout type r\ Air circuit breaker, nondrawout type, series trip cHh «-o^ Magnetic starter Current-limiting breaker, drawout type -j&- Disconnecting fuse, nondrawout «d> Drawout fuse Disconnecting switch, nondrawout ^^^ Disconnecting switch, drawout type -f^t Current transformer tt Potential transformer t> o-)|l Pothead Ground Surge arrestor Surge capacitor II > Battery Cable terminations Rectifier bank Reactor, nonmagnetic core Reactor, magnetic core Power transformer 3-phase power transformer connected delta- wye Y Wye A Delta Figure 4.17.— Commonly used symbols for one-line electrical diagrams. 88 Overvoltage Undervoltage Overcurrent •'HW- Ground overcurrent > < Undercurrent « X ► Differential current ■■i N X ► Differential ground current I A Balanced current Directional overcurrent ■I -W— ► Directional ground overcurrent Directional power GP Gas pressure PW N Pilot wire (differential current) PW Pilot wire (directional comparison) 4 Z ► Distance -i n z >: Ground distance ^ 1r+ Directional distance "HI Z ► Directional ground distance CC Carrier directional comparison (phase and control) J\ CC^ ^Sh> Carrier phase comparison Synchro check 4 S ► Auto synchronizing d fc < A > Phase balance -0— ► Phase rotation 4 r ^u ► Overfrequency « T ► Overtemperature Note: For directional relays, arrow points toward fault that will cause tripping Figure 4.17.— Commonly used symbols for one-line electrical diagrams— Con. 89 4 1 Wye, neutral ground <^ Zig-zag ground j/ CT ' Meter Current transformer with ammeter; letter indicates instrument type Hr^ ■&S-45 -0 Relay Relays connected to CT's and PT's. Number indicates relay type function 1 Ground RELAY FUNCTIONS ± Overcurrent Differential O Induction motor or general source O Synchronous motor Instrument transfer switch. Letter indicates type < > Dummy circuit breaker. Removable type -< >- Future breaker position. Removable type -WW Resistor Figure 4.18.— Symbols for relay functions. 90 Table 4.1. —IEEE device numbers and functions (1) Device Function 1 Master element. 2 Time-delay starting or closing relay. 3 Checking or interlocking relay. 4 Master contactor. 5 Stopping device. 6 Starting circuit breaker. 7 Anode circuit breaker. 8 Control power-disconnecting device. 9 Reversing device. 10 Unit sequence switch. 1 1 Reserved for future application. 12 Overspeed device. 13 Synchronous-speed device. 14 Underspeed device. 15 Speed, or frequency matching device. 16 Reserved for future application. 17 Shunting or discharge switch. 18 Accelerating or. 19 Starting to running transition contactor. 20 Electrically operated valve. 21 Distance relay. 22 Equalizer circuit breaker. 23 Temperature control device. 24 Reserved for future application. 25 Synchronizing or synchronism check. 26 Apparatus thermal device. 27 Undervoltage relay. 28 Reserved for future application. 29 Isolating contactor. 30 Annunciator relay. 31 Separate excitation device. 32 Directional power relay. 33 Position switch. 34 Motor-operated sequence switch. 35 Brush-operating or slip-ring short-circuiting device. 36 Polarity device. 37 Undercurrent or underpower relay. 38 Bearing protective device. 39 Reserved for future application. 40 Field relay. 41 Field circuit breaker. 42 Running circuit breaker. 43 Manual transfer or selector device. 44 Unit sequence starting relay. 45 Reserved for future application. 46 Reverse-phase-balance current relay. 47 Phase-sequence voltage relay. 48 Incomplete sequence relay. 49 Machine or transformer thermal relay. 50 Instantaneous overcurrent or rate-of-rise relay. Device Function 51 Time overcurrent relay — ac. 52 Circuit breaker— ac. 53 Exciter or dc generator relay. 54 High-speed dc circuit breaker. 55 Power-factor relay. 56 Field application relay. 57.. Short-circuiting or grounding device. 58 Power rectifier misfire relay. 59 Overvoltage relay. 60 Voltage balance relay. 61 Current balance relay. 62 Time delay stopping or opening relay. 63 Liquid or gas pressure level or flow relay. 64 Ground protective relay. 65 Governor. 66 Notching or jogging device. 67 ac directional overcurrent relay. 68 Blocking relay. 69 Permissive control device. 70 Electrically operated rheostat. 71 Reserved for future application. 72 Circuit breaker — dc. 73 Load resistor contactor. 74 Alarm relay. 75 Position-changing mechanism. 76 Overcurrent relay — dc. 77 Pulse transmitter. 78 Phase-angle measuring or out-of-step protective relay. 79 Reclosing relay — ac. 80 Reserved for future application. 81 Frequency relay. 82 Reclosing relay— dc. 83 Automatic selective control or transfer relay. 84 Operating mechanism. 85 Carrier or pilot-wire receiver relay. 86 Locking-out relay. 87 Differential protective relay. 88 Auxiliary motor or motor generator. 89 Line switch. 90 Regulating device. 91 Voltage directional relay. 92 Voltage and power directional relay. 93 Field-changing contactor. 94 Tripping or trip-free relay. 95 Reserved for future application. 96 Do. 97 Do. 98 Do. 99 Do. Table 4.2.— Device numbers and letters common to mining (2) Device Function 1 Control switch. 3 Plus interlock relay. 37 Ground-continuity check undercurrent relay. 49 D Diode thermal relay. 49 GR Grounding resistor thermal relay. 49 T Transformer thermal relay. 50 Instantaneous overcurrent relay— ac. 50 G Instantaneous overcurrent relay — dc — connected in ground wire. 50 N Instantaneous overcurrent relay— ac — connected to neutral. 50 Z Instantaneous current-balance relay — ac— zero sequence. 51 Time delay overcurrent relay — ac. 51 N Time-delay overcurrent relay— ac — connected to neutral. 51 Z Time-delay current-balance relay— ac— zero sequence. 52 Circuit breaker — ac. 59 GR Ground protective relay — dc — unbalance relay. 72 Circuit breaker — dc. 76 Overcurrent relay — dc. A Ammeter. D Demand meter. GD Ground detector. PF Power factor. V Voltmeter. VA Volt-ammeter. VAR Varmeter. W Wattmeter. WH Watthour meter. The remainder of the text will employ one-line and per-phase diagrams almost exclusively. The main thing to remember is that practically every item represents three identical or corresponding items in the actual system. Even when the normally balanced system becomes unbal- anced through component failures, the same diagram is used, the only change being the notation for the specific failure. Circuits Containing Transformers As previously stated, solving a balanced three-phase system problem by per-phase analysis is as simple as the single-phase techniques covered in chapters 2 and 3. However, the solution is not so clear when delta-wye or wye-delta transformers are involved. Perhaps the simplest way to demonstrate the approach is first to illustrate a problem solution where the per-phase reduction is of a straightforward wye-wye transformer, then change to a delta-delta or delta-wye transformer and show the compli- cations that might arise. 91 EXAMPLE 4.7 Consider figure 4.19, which shows a one-line diagram of a substation supplying power through about 1 mile of overhead line (power conductors on poles) to a three-phase wye-wye transformer bank, then through a trailing cable to a three-phase induc- tion motor. The motor consumes an average three- phase power of 150 kW, operating at 0.8 lagging power factor. The problem is to find the rms voltage needed at the substation output to provide the rated motor terminal voltage of 480 V line-to-line. A three-phase diagram of the circuit is shown in figure 4.20 for reference. The first step in the solu- tion is usually to develop a per-phase circuit as shown in figure 4.21. Although the solution could be Substation Load center Utility Overhead line Z = 0.6+jO6 Trailing cable ' Z=0.07 + j0.5 c (t) -ir o -»- ^a 1,385 /277 V Z = 0.1+jO3 referred to high side 50 kW at 0.8 lagging pf 277 V line-to-neutral Figure 4.19.— One-line diagram for example 4.7. = Z 2 = 0.07 +j 0.05 Transformer Figure 4.20.— Three-phase diagram of figure 4.19. 5M (r) Zol=0.6+j0.6 Z X p=0.1+j0.3 Z tc = 007 +j 0.05 -VWV V-r VWV 1 r—n WW- Vt V, V L =277V Figure 4.21.— Per-phase diagram of figure 4.19. performed directly from the one-line diagram, the per-phase diagram allows direct application of single-phase techniques. The following should be noted in figure 4.21. • The line and trailing-cable conductor imped- ances are now illustrated as circuit elements. • Only one phase of the transformer bank is shown, represented as an impedance in series with the primary of an ideal transformer. The trans- former turns ratio is computed from a = Vg _ 277 1 Vj " 1,385 " 5 ' where these rated voltages are line-to-neutral rms. • The induction motor is represented by its single-phase equivalent, P , where 150 Fp " 3 3 - 50 kW. The solution can now follow a stepwise process. 1. Assume the motor terminal voltage, V L , is the rated 277-V rms line-to-neutral (480/V3). 2. Compute motor line-current magnitude, I L , using or P P = V L I L (load pf) P P 50,000 II V L (load pf) (277X0.8) = 225.5 A. If the motor terminal voltage is taken as the refer- ence phasor, this current has a phase angle deter- mined by the load power factor. Therefore the motor current phasor is I L = It.I -cos-^O.S) = 2261 -36.9° A. 3. I L is then employed to find the voltage drop across the per-phase equivalent of the trailing cable, Z^. When this is added to the motor terminal voltage, the voltage at the ideal transformer second- ary, V 8 , is V 8 = I L Z te + V L or V 8 = (226| -36.9° X0.09| 35.5° ) + 277|0°_ = 20.301 -1.4° + 277|0^ = 20.29 -J0.50+ 277 = 297.29 -j 0.50 + 297| -0.1° V. 4. For the ideal transformer with a turns ratio, a, of 1/5, the voltage across the primary is v -i p a or V P = 5(2971 -0.1° ) = I486] -0.1° V, 92 with the primary current being I P = ai L or I P = (|X226| -36.9 ) = 45| -36.9° A. Notice that there is no change in current or voltage phase angle across the transformer. 5. This primary current can now be used to find the voltage drop resulting from transformer and overhead-line impedances. Summing this voltage drop with the transformer primary voltage gives the desired answer to the problem, the substation out- put voltage, V t : v t = I P (Z OL + Z xp ) + Vp, where Z OL , Z xp = overhead-line and transformer Then, impedances, respectively. V t = 451 - 36.9° (0.6 +J0.6 + 0.1 +.J0.03) + 1,486|-0.1° or V t = 45| -36.9° (0.94 | 41.99° ) + 1,486| -0.1° = 42.39|5J^ + 1,486 | -0.1° = l,528| +0.04° V. Because the analysis is per phase, the result is obviously line-to-neutral voltage. If line-to-line val- ues are required, the above answer need only be multipled by V3. It is interesting that in this exam- ple the phase angle of the substation voltage is practically the same as that at the motor. This may not be the case in actual mine power systems. When the transformer is delta-delta connected, problem solutions are practically the same as in example 4.7. The one-line diagram of figure 4.22 provides an illustration. When the system is represented per phase (fig. 4.23), the only additional concern is delta-wye transformation of the trans- former impedance; then the solution proceeds as before. However, the process may not be as simple when delta-wye or wye-delta transformer connections are involved. Consider the one-line diagram in figure 4.24 that shows a delta-wye transformer supplying the same motor as that shown in figure 4.19. Figure 4.25 illustrates one leg of the three-phase transformer. From this, it can be seen that secondary line currents appear as phase cur- rents on the primary, and line-to-neutral secondary volt- ages become line-to-line primary voltages. In other words (fig. 4.25), for primary voltage in terms of the secondary (the ideal transformer with turns ratio, a, only): V AR = _an | 30 c a — (4.18) where V_ab = line-to-line primary voltage, V, and V an = line-to-neutral secondary voltage, V, and for primary and secondary current, I AB = a IJ301, where I^p = primary phase current, A, and I a = secondary line current, A. (4.19) Substation Z = 0.6+j0.6 ] £ Z=0.07 + j0.05 *-; " r» — < s — • — —- :t) A J ""A Bank of 3-10 XFMRS, each 2.400/480V Z=0.3+j0.9 referred to high side o 150 kW total at 0.8 lagging pf 480 V line-to-line Figure 4.22.— One-line diagram with delta-delta transformer. ^-xp ^-xp'^ •-''' (t) Z 0L =0.6+j0.6 =Ql+j0.3 Z tc = 0.07 +j 0.05 -VWV r*7 WvV 1 i — r*7 VWv- V L =277V Figure 4.23.— Per-phase diagram of figure 4.22. Load Substation center Utility -\ r Overhead line > r Z = 0.6 + j0.6 Trailing cable Z =0.07 tj 0.5 (t) -O "la 1,385 /277 V Z=0.1+j0.3 referred to high side 50 kW at 0.8 lagging pf 277 V line-to-neutral Figure 4.24.— One-line diagram with delta-wye transformer. Primary Secondary Figure 4.25.— One leg of three-phase transformer from figure 4.24. However, when performing balanced three-phase analysis, the parameters of interest are line-to-neutral voltages and line currents. Thus, to continue the analysis in a fashion similar to that used in example 4.7 (the wye-wye transformer), V^ and Ij^ must be converted to their respective per-phase equivalents. Recalling that V ab = V3 V„J+30°, (4.20) and applying this concept to equation 4.18, the primary line-to-neutral voltage, V An , is _ V V - *£. VAn_ aV3- (4.21) 93 Employing equation 4.76 to convert equation 4.19, pri- mary line current, I A , is voltamperes. The mathematical interrelations of the bases are as follows: I A = a L (4.22) Equations 4.21 and 4.22 simply state that the phase shifts that occur across delta-wye or wye-delta-connected trans- formers do not interfere with the analysis when this is performed per phase. Analysis can be enhanced by chang- ing the delta primary or secondary to an equivalent wye connection, thus enabling the construction of a per-phase diagram for the entire system. Concerning the actual per-phase analysis, it has been shown that the three-phase circuit is reduced by a process no more difficult than the single-phase work covered in chapters 2 and 3. The next section will present a technique that further simplifies power-system analysis. PER-UNIT SYSTEM Problems related to electrical circuits should be solved in terms of volts, amperes, voltamperes and ohms. The answers to mine power problems, and indeed any electrical problem, almost always require these terms, but in the process of computations it is often more convenient to express these quantities in percent or per-unit (pu), re- ferred to some arbitrarily chosen base. For example, if a base voltage of 100 kV is selected, voltages of 90, 120, and 125 kV have percent representations of 90%, 120%, and 125%, or per-unit values of 0.9, 1.2, and 1.25 pu, respec- tively. Both percent and per-unit values express a ratio of a specific quantity to the base quantity. Per-unit is given as a decimal, whereas the ratio in percent is 100 times the per-unit value. These expressions, especially per-unit, are becoming standard for equipment specifications. There is a definite advantage in using per-unit nota- tion over percent. Per-unit multiplication or division yields a result in per-unit. However, the product of two percent quantities must be divided by 100 to obtain a percent answer. For example, if two quantities are both 0.9 pu or 90%, then and but (0.9 puX0.9 pu) = 0.81 pu (90%) (90%) * 8,100% 8,100 „ (90%X90%) = 100 Consequently, per-unit notation will be employed almost exclusively, the only exception being where conventions dictate otherwise. Voltage, current, voltamperes, and impedance are obviously interrelated for any specific circuit or system. As a result, the selection of any two determines the base values for the remaining two. For example, if the current and voltage bases are specific, the base impedance and base voltamperes can be found. Since three-phase circuits are usually solved as a single line with a neutral, base quantities in the per-unit system are line current, line- to-neutral voltage, per-phase impedance, and per-phase v b = i b z b , KVA b = kV b I b , T kVA b b " kV„ V b (kV^l.OOO by. — - — Iv kVA, (kVb) 2 MVA h (4.23) (4.24) (4.25) (4.26) where V b = base line-to-neutral voltage, V, kV b = base line-to-neutral voltage, kV, kVA b = base per-phase voltamperes, kVA, MVA b = base per-phase voltamperes, MVA, I b = base line current, A, and Z b = base per-phase impedance, fi. All these formulas are adaptations of the fundamental Ohm's law and power material; the last three are ex- pressed in kilovolts and kilovoltamperes because of the levels normally found in power systems. It should be remembered that line-to-line voltages and total power (kilovoltamperes or megavoltamperes) are customarily specified; these must be changed to line-to-neutral volt- ages (dividing by V3) and per-phase power (dividing by 3) before equations 4.23 through 4.26 can be applied. To apply the per-unit system to power problems, base values for kilovoltamperes and kilovolts are normally selected first in order to minimize calculations as much as possible. Base values for impedance and current are then found. Next, all the actual voltages, currents, impedances, and powers in the power system or system segment are expressed as a ratio to the base quantities; these are the per-unit quantities. Problems are then solved in per-unit, with the answers converted back to actual parameters. The actual values and per-unit quantities are related by Za - Z pu Z b , (4.27) Ia = Ipu lb, (4.28) v A = v pu v b , (4.29) VA A = VAp U VA b) (4.30) where Z A , I A , V A , VA A = actual values of impedance, current, voltage, and power, re- spectively, fi, A, V, VA, and Z pu , Ip U , V pu , VAp U = per-unit values of impedance, current, voltage, and power, re- spectively, pufi, puA, puV, puVA. It is important to note that all impedances in a problem are referenced to the same base impedance, whether they are pure resistance or pure reactance. The same holds for all average, reactive, or apparent powers, which are refer- enced to the base kilovoltamperes. 94 Often, per-unit impedances or percent impedances of a system component have already been assigned to a base referenced to the component or power-system segment in which that component is located. These impedances can be changed to another base impedance by = Z Due kVA b /kV^ 2 pu kVA„ IkV, (4.31) where Z ie = per-unit impedance of specified component, puft, kV e , kVA e = base kV and kVA used to reference Z pue , V, VA, kV b , KVA b = base kV and kVA to which new per-unit impedance is to be referenced, V, VA, and Z pu = new per-unit impedance referenced to kV b and kVA b , pu ft. Transformer Impedance Transformers are the most common devices in power systems where the component impedance is referenced to the rated power and voltage of the component. Convention- ally, percent impedance is specified, but this can be converted to per-unit simply by dividing by 100. A major advantage of using per-unit computations is seen when circuits are connected through transformers. Through the proper selection of voltage bases, the per-unit impedance of the transformer is the same no matter which winding is used. Consequently, if exciting and magnetiz- ing currents are ignored, as they often can be in power systems, the transformers become a simple series imped- ance in per-unit calculations. In other words, the ideal transformer is not needed in the equivalent circuit. Exam- ple 4.8 explores this advantage. EXAMPLE 4.8 Consider a 750-kVA power-center transformer, the approximate per-phase equivalent circuit for which is shown in figure 4.26. The per-phase ratings are 250 kVA, 5,000/1,000 V, and 5-ft reactance re- ferred to the high side. Following convention, the base power kVA bl is 250 kVA, and the base voltage for the high side kV bl is 5 kV. One kilovolt is selected as the low-side base voltage, kV b2 . With these, the high-side base quantities can be calcu- lated using equations 4.23 through 4.26: kVA bl = 250 kVA, kV b Ik, and = 5kV, = 50 A, Zb! = 100 ft. The per-unit impedance of the transformer is thus rj Z^l 100 = 0.05 pu. Now consider the actual transformer impedance as it would appear referred to the low side, as in figure 4.27. /N V, or Z A2 - Z A1 \^J - Z A1 ( y * The base quantities on the low side are kVA bl = 250 kVA, kV b2 = 1 kV, I, and = 250 A, = 4 ft. Notice that the base power does not change and the low-side base voltage defines base current and im- pedance. The per-unit transformer impedance as seen from the low side is -7 ^A2 V = -7 ^b2 0.2 = 0.05 pu. Therefore, the per-unit impedance of the trans- former is the same, regardless of the side it is viewed from, and the per-unit equivalent circuit is simply the series circuit shown in figure 4.28. Here, the input and output voltages are now expressed in per-unit since the transformer is operating at rated voltage. Z A1 =j5n V,= 5,OOOV V 2 = 1,000 V Primary Secondary Figure 4.26.— Approximate per-phase equivalent circuit for 750-kVA load-center transformer; impedance referred to high side. Z A2 =j0.2il V,= 5,000 V V 2 = 1,000 V Figure 4.27.— Transformer of figure 4.26 with impedance referred to low side. Z=j0.05pu V, = 1.0pu V 2 = 1.0pu Figure 4.28.— Simplified equivalent circuit of transformer expressed in per-unit. 95 Three-Winding Transformers In chapter 3 and to this point in chapter 4, equivalent circuits have been shown only for two-winding transform- ers, those having only one primary and one secondary winding. However, many transformers in mine power systems have three windings, with the third winding termed the tertiary or second secondary. These include power-center transformers supplying two different utiliza- tion voltages, such as 950 and 550 Vac to face equipment or 550 and 250 Vdc to machinery. The latter case not only uses a three-winding transformer but also three-phase rectification, which will be described in chapter 5. Both the primary and secondary windings of the two-winding transformer have the same kilovoltampere capacity or rating, but the three windings of a three- winding transformer may have different kilovoltampere ratings. The impedance of each winding may be given in percent or per-unit, based on each winding rating. The three impedances can also be measured by the following short-circuit test, where rated voltage is applied to the primary for Z ps and Z pt and to the secondary for Z 8t : Z ps = leakage impedance measured in primary (or first winding), with secondary (or second wind- ing) short-circuited and tertiary (or third wind- ing) open, Q, Z pt = leakage impedance measured in primary with tertiary short-circuited and secondary open, Q, Z st = leakage impedance measured in secondary with tertiary short-circuited and primary open, The impedances of the primary, secondary, and tertiary windings are found from or where Z p , Z s , Z t = impedances of primary, secondary, and tertiary, respectively, Q. they must have the same kilovoltampere base. Further, voltage bases for the circuits connected through the trans- former must have the same ratios as the turns ratio of the transformer windings; that is, primary to secondary, pri- mary to tertiary, which are actually the same as the ratios of the related winding voltages. Per-Unit Method in System Analysis As mentioned earlier, the use of per-unit equivalents in the analysis of power-system problems can greatly simplify the work involved, especially when the system contains transformers and different voltage levels. How- ever, as per-unit calculations require the change of famil- iar parameters (ohms, volts, amperes, and so on) into values representing a ratio, this advantage is often diffi- cult to comprehend. Example 4.9 will illustrate the per- unit method of analysis using the one-line diagram pro- vided in figure 4.30, which could represent a mine power system in the early stages of development. All power levels listed are given per-phase; those shown for the mining equipment represent consumption. The voltages are all line-to-neutral. EXAMPLE 4.9 The information required could be the voltage or current level at any point. Regardless, solution by the per-unit method first requires translation of the impedance of all components to the same base. The base selection is arbitrary, but for convenience, the largest kilovoltampere capacity of a system compo- nent is usually taken. In this case a good base would Z p8 = Z p + Z s , (4.32a) Z pt = Z p + z t) (4.326) Z 8t = z 8 + z t , (4.32c) Po 3 C C OS Zp = 2 (Z ps + Z pt - Z 8t ), (4.33a) *5t C-<- Z 8 = 2 (Z ps + Z 8t - Zpt), (4.336) 6 t Z t = 2 ( Z pt + Z at - Zps), (4.33c) Hine symbol pa- -A/WV- ^vwv- -oS z» -AAMr -ot Neutral Equivalent circuit ^n Figure 4.29.— Approximate equivalent circuit of three- winding transformer expressed in per-unit. In equations 4.32 and 4.33, all impedances (Z p8 , Z pt , Z st ) must be referred to the primary winding voltage. If Z 8t is obtained from the described measurements, the imped- ance is referred to the secondary- winding voltage, hence it must be transferred. The approximate per-phase equivalent circuit for a three-winding transformer with the winding impedances of Z p , Z 8 , and Z t is provided in figure 4.29. Magnetizing and exciting currents are ignored. The terminals p, s, and t are the primary, secondary, and tertiary connections; the common point is unrelated to the system neutral. The three impedances must be in the per-unit system, as was the case for the equivalent circuit in figure 4.28. Hence Substation center Feeder cable ^ y H Z=0.13 + j0.06 H ^^il — — it Trailing cable Miner Load p O.Q3 + j 0.01 p = 57 k W ~^WQ=47kvar V=320V 1,000 kVA 40/7.2 kV Z = 7% 150 kVA 7.2/350kV Z = 4.5% ^ Trailing cable Z = 0.1+ jO.02 _ Shuttle • Q Bus P=4kW Q=5kvar Figure 4.30.— One-line diagram of small mine power system. 96 be 1,000 kVA, corresponding to the per-phase capac- ity of the substation. But two base parameters are needed in order to define the four base quantities; as the nominal voltage for each voltage level can be or can approach a constant, the system voltages are an excellent choice for the second base parameter. For figure 4.30, these would be the line-to-neutral volt- ages of 40 kV at the utility, 7.2 kV at mine distri- bution, and 350 V at mine utilization. Note that the system voltages are usually given as line-to-line in one-line diagrams, so they must be changed to line-to-neutral values to employ the formulas given here. In any event, base voltages must correspond with the turns ratio of any interconnecting trans- former. The ones selected do. With base kilovoltamperes and base kilovolts specified, the base quantities can be calculated using equations 4.23 through 4.26. 1. For the utility: kVA b = 1,000 kVA, kV bl = 40 kV, _kV^_ i^oo J-bl - UT7 - An - S.O J\, kV bl 40 Zbi J^l^OO m (40^,000 m 1600 kVA, 1,000 2. For mine distribution: kVA b = 1,000 kVA, kV b2 = 7.2 kV, Zb2 ~ 2 ~ 7.2 (7.2)2 1,000 1,000 = 52 fl. 3. For mine utilization: kVA b = 1,000 kVA, kV b3 = 0.35 kV, 1,000 „ „_, . Ik, = -:h^ = 2,857 A, Zb3 - 0.35 (0.35) 2 1,000 1,000 = 0.12 Q. The per-unit representations for all components of the mine system can now be found, and the needed formulas are equations 4.27 through 4.31. 1. For the substation: percent reactance is 7% or 0.07 pu based on the transformer rated kilovoltam- peres, referred to the high side, 40 kV. z = Z DUe kVA b £V^ 2 kVA e \kV b where Z pue = J0.07 pu, kVA b = kVA e = 1,000 kVA, kV e = kV b = 40 kV; thus, Z pue = J0.07 pu. 2. For the feeder cable: actual impedance is given, Z A = Ra + JX A = 0.13 + J0.06 Q, Za _ 0.13 + J0.06 'b2 and Zp U - 7 52 = (0.0025 + jO.0012) pu. 3. For the load center: percent reactance is 4.5% or 0.045 pu based on the transformer rated kilovolt- amperes, referred to the high side, 7.20 kV. _ Z Due kVA b / kV e , 2 pu kVA„ IkVu where and Zpu = Z pue = J0.045 pu, kVA b = 1,000 kVA, kVA e = 150 kVA, kV b2 = kV e = 7.2 kV, (j0.045Xl,000) 150 = jO.3 pu. 4. For the trailing cables: actual impedances are again given. Miner. Z A = 0.03 + J0.01 fi, Za _ 0.03 + jO.Ol pu " Z b3 " 0.12 = (0.25 + jO.083) pu. Shuttle car. Z A = 0.1 + J0.02 12, 7 - --^- - pu ~ Z ~ ^b3 0.1 + jO.02 0.12 = (0.833 + jO.167) pu. 97 5. For the machines: consumption is given in terms of average and reactive power. Miner. P = 57 kW, Q = 45 kvar or kVA A = (57 + j45) kVA; thus, kV V - kYA - - (5 ^^ ) kVA, 1,000 = (0.057 + jO.045) pu. Shuttle car. kVA A = (4 + j5) kVA, hence, kVA„ u = (0.004 + jO.005) pu. At this point, the entire system of figure 4.30 may be redrawn into the impedance diagram in figure 4.31. Figure 4.31, when compared with a per-phase diagram in the impedance domain such as figure 4.21, illustrates the advantage that simplified per- unit computations lend to power-system analysis. The circuit shown is merely a series-parallel ar- rangement of basic electrical elements, and obvi- ously it may be further simplified if desired, say into an equivalent per-unit impedance. This is only one example; an actual appreciation of power-system analysis by per-unit techniques can come only through experience. The impedance diagram can be used for the solution of most power problems. Suppose currents under normal operation are desired. One method is to apply known voltages at system points and calcu- late the resulting currents and voltages. For in- stance, suppose the line-to-neutral at the miner is 320 V (about 555 V line-to-line). The per-unit equiv- alent of this is tv pu . ^ kV b3 0.32 0.35 = 0.91 pu. The line current through the miner trailing cable is then or _ /kVA^ ^ u \ kV. (0.057 - jO.045) nntin . nnAn ^ = 091 = 0063 " j0049 pu - The conjugate of power is employed because voltage is implied as the reference phasor following the conventions stated earlier. The process is then con- tinued through the entire system. When the desired per-unit values are obtained, they are simply con- verted to actual values. Considering the current in the miner trailing cable, the actual line current is Ia = Ipu Ib3 or I A = (0.063 - J0.049X2857) = 180 - J141.3A or I A = 229 1 -38.1° A. (0.25 + j 0.083) pu Utility j0.07pu j0.3pu (0.0025 +j 0.0012) Miner load kVA "> (0.057 + jO045)pu (0.833 + j 0. 167 )p U Shuttle car kVA ' (0.004 + j 0.005) pu Figure 4.31.— Impedance diagram of system in figure 4.30, expressed in per-unit on a 1,000-kVA base. UNBALANCED THREE-PHASE CIRCUITS The solution of balanced three-phase circuit problems is usually accomplished by converting the constants, cur- rents, and voltages to per-phase values. Because symme- try determines the magnitude and phase position of all currents and voltages, actions occurring between phases can be represented by equivalent impedances. Further- more, currents and voltages for the other phases are equal in magnitude to those in the per-phase solution but are simply displaced symmetrically in phase position. This is extremely important because normally operating three- phase power systems can usually be approximated as balanced. However, the solution of unbalanced three-phase circuits or balanced circuits with unbalanced loads does not permit the same simplification. Mine power systems are designed to have a high degree of reliability and therefore to operate in a balanced mode. But at times, equipment failures and unintentional or inten- tional disturbances from outside sources can result in an unbalanced operation. The most common sites for mine power-system unbalance are equipment trailing cables. The consequence of unbalance is abnormal currents and volt- ages, and if safeguards are not designed into the system to protect against these anomalies, the safety of personnel as well as equipment can be compromised. The protective circuitry within the mine power system serves as the safety valve for such hazardous m alfunctions. Power-system unbalance can occur either from open circuits or from faults. A fault occurs whenever electricity strays from its proper path. Faults can be visualized as the connecting together of two or more conductors that nor- mally operate with a voltage between them. The connec- tion that creates a fault can be from physical contact or an arc caused by current flow through a gaseous medium. A short circuit is one type of fault. Currents in the power system resulting from an open circuit or a fault can be exceedingly large. An overload is not a fault. The term overload merely implies that currents exceed those for which the power system was designed. Such currents are usually much smaller than fault currents. Nevertheless, overloads can create equipment failures by exceeding the thermal design limits of the system. If not corrected, the overload can result in a hazard to personnel. However, such problems occur only with unattended overload operation for an extended time period, whereas faults can create an unsafe condition almost instantly. Both circuit breakers and fuses are used to protect circuits from excessive current flow, be it a result of faulting or overloading. The circuit-interrupting operation consists of parting a pair of contacts, and since an arc is 98 drawn between the contacts, the process must also extin- guish the arc. The interruption is handled mechanically in the circuit breaker, and the excessive current is monitored electrically or thermally. Fuse operation is based on sim- ple thermal operation. The fusible element is responsive to the heat of an overload or fault current and melts open. The fuse jacket is employed to extinguish the subsequent arc. A complete discussion of interrupting devices and the associated protective circuitry is presented in chapter 9. Fault Types The fault type often seen in literature is called a bolted fault, which can be described as a zero-resistance short circuit between two or more conductors. In reality most faults are not dead shorts but have some finite value of resistance. Faults may be classed as permanent or temporary. A permanent fault is one where equipment operation is impossible and repairs are mandatory. A temporary fault is intermittent in nature. For instance, two closely spaced overhead conductors may cause trouble only on windy days, when they can be forced into contact or close prox- imity by the wind. A very sinister fault type is the arcing fault. This condition is now believed to be the most common fault. When two conductors of different potential are in very close proximity, the intervening space between them can be considered as a spark gap. If the two conductors are part of an ac power system, the insulating material between the conductors may break down when the sinusoidal waveform reaches a certain value. Fault current will then flow. The potential drop across the conducting gap, which is actually an impedance, remains at a nearly constant level. It is this energy source, releasing terrific quantities of heat, that causes the devastation that is typical of an arcing fault. Soon after the sinusoid reverses polarity, the arc quenches until the spark-gap breakdown voltage or restrike potential is reattained. This repetitive arcing is almost always self-sustaining at ac voltage levels of 480 V and above. Depending on how the fault occurs, it may be de- scribed as three-phase, line-to-line, or line-to-neutral. In mining, the cable shields and the grounding system of the equipment are at the same potential as the system neu- tral, and line-to-neutral faults are the most prevalent. Line-to-line faults and line-to-neutral faults are unbal- anced or unsymmetrical, but the three-phase fault is balanced or symmetrical. These three basic fault descrip- tions are illustrated in figure 4.32. The impedance, al- though very small, is shown to signify its finite value. -1 1 1 — — » Zq r $ Zg $ Zq i — 1 1 1 1 . — Line to neutral Line to line 3 phase Figure 4.32.— Basic fault descriptions. Fault Analysis Fault analysis is a desirable and often mandatory part of any mine power-system analysis. As faulting of some nature can occur at any time, knowledge of how faults affect currents and voltages is necessary to design proper protection and to ensure personnel protection. Although faults usually occur in mine-system trailing cables, the actual fault location and time of occurrence is unpredict- able. Consequently, fault analysis is frequently effected on a trial-and-error basis, searching for a worst case solution. It is necessary to assume a fault location, the configura- tion of power-system components prior to the fault, and sometimes the system loads. Such an effort can result in numerous calculations, to the point where digital comput- ers can be extremely advantageous. Nevertheless, the results provide invaluable information on which to base the design of the mine power system. Though not particularly common, fault analysis using three-phase faults has distinct advantages. Using this method, balanced faults, like balanced loads, can be inves- tigated on an equivalent per-phase basis and therefore become as simple as faults on single-phase lines. In the majority of cases, bolted three-phase faults cause larger fault currents than line-to-line or line-to-neutral events. Unsymmetrical faults are often of high interest in mine power systems. Instances include finding a mini- mum fault current or current flowing in the system neutral conductors. When an unsymmetrical fault is placed on the system, the balance is disrupted. It is possible to solve an unbalanced power system by using a three-phase diagram with symbols assigned to the quan- tities in each phase and carrying the phase solutions simultaneously. This complicates the problem enormously, but it can be simplified by applying the method of sym- metrical components, which reduces the solution of such problems to a systematic form. The reduction is particu- larly applicable to balanced systems operating under unsymmetrical faults. SYMMETRICAL COMPONENTS The method of symmetrical components provides a means for determining the currents and voltages at any point of an unbalanced three-phase power system. In this method, the unsymmetrical phasors representing the un- balanced voltage or current are expressed as the sum of three symmetrical phasor sets. These phasor sets or sym- metrical components are designated as the positive se- quence, negative sequence, and zero sequence. The first two consist of three balanced phasors with equal magni- tude, set 120° apart. The zero-sequence set has three phasors equal in magnitude but operating in the same time. The components are illustrated in figure 4.33, where the instantaneous values may be determined by projection upon the X-axis. The positive-, negative-, and zero- sequence components are then employed to solve the unbalanced-system problem. These sequences are so com- mon in power-system terminology that they are often used to describe the quality of system operation. It might be asked why the resolution of three phasors into nine phasors simplifies the solution of unbalanced power systems. The answer is straightforward. The resolu- tion results in three symmetrical systems in which each 99 V a0 V b0 ) v//- v c0 Figure 4.33.— Positive-sequence, negative-sequence, and zero-sequence vector sets. balanced phasor set can be treated separately, just as in balanced three-phase systems. In other words, the power system can be reduced to per-phase values, then analyzed separately for each symmetrical component. This analysis hinges on the fact that currents and voltages of different sequences do not react upon each other: currents of the positive sequence produce only positive-sequence voltage drops; the same is true for the negative and zero sequences. In addition to aiding analysis, the method of symmetri- cal components separates electrical parameters into parts that can represent better criteria of the controlling factors for certain phenomena. For example, the presence of negative- sequence current or voltage immediately implies that the system is unbalanced, and this can be utilized to detect malfunctioning power systems. Grounding phenomena are other good criterion examples; neutral current is very closely related to zero-sequence components. Sequence Components The positive sequence for voltage is composed of three symmetrical phasors, V al , V bl , and V cl for phases a, b, and c, respectively (fig. 4.33). The quantities have equal mag- nitude but are displaced by 120° from each other. There- fore, following equation 4.26, V al = V bl |12po = V cl |240^, or rewriting in exponential form, » al = »al> V - pi 240 V v bl - ^ v al> V cl = e* 120 V al . The unit vector, e* 120 , is used so frequently that it is given the symbol "a" (not to be confused with the transformer turns ratio), where and a = e^ 120 = 1 | 120° & 2 _ ^120^120 _ eJ240 (4.36a) (4.366) Equations 4.35 and 4.37 also relate to the standard practice of symmetrical-component calculations; equations are al- ways expressed in terms of the phase a quantities. There are several mathematical properties of the unit vector a that are useful in computations: 1 = e* = 1.0 + J0.0, a = e* 120 = -0.5 + jO.866, a 2 = (J 240 = -0.5 - jO.866, a 3 = e* 360 = 1.0, a 4 = e* 480 = e* 120 = a, a 5 = e* 600 = e j240 = a 2 ; and for specific calculations: 1 + a 2 + a = 0, a - a 2 = V3e*° = jV3, a 2 - a = V3e " j9 ° = -jV3, 1 - a = 3e - j30 = 1.5 - jO.866, 1 - a 2 = V^ 30 = 1.5 + jO.866. These allow easy conversion to simpler forms when sym- metrical components are being manipulated mathemati- cally. For the negative and zero sequences (fig. 4.33), the symmetrical voltage sequences can be written V a2 = V b2 1 -120° = V c2 1 120' and V„„ = V hn = V„ n . (4.38) (4.39) Rewriting these equations in terms of the unit vector, a, it is found that for the negative sequence, V^ = V^, Thus the positive-sequence vectors (equations 4.35) are customarily written as (4.40a) (4.406) (4.40c) (4.41a) (4.416) (4.41c) It is important to note that in all three sequence systems, the subscripts denote specific components in each phase (a, b, or c). Furthermore, the subscripts, 1, 2, and signify whether that component is part of the positive-, negative-, or zero-sequence set. Using the same reasoning, symmetrical-component equations can also be written for current. Vt.ot; V b2 = aV a2 , *c2 = a *a2> (4.35a) (4.356) and for the zero sequence, (4.35c) * a o = »aO> ' bO = "aO> s given iformer V c0 = V a0 . 'al — *al> V bl = a 2 V al , V cl = aV al . (4.37a) (4.376) (4.37c) Sequence-Quantity Combinations The total voltage or current of any phase is equal to the vectorial sum of the corresponding components in that 100 phase. Figure 4.34 illustrates this concept for three unbal- anced voltage phasors, V a , V b , and V c . Expressed mathe- matically, V a = V a0 + V al + V^, (4.42a) V b = V b0 + V bl + V b2 , (4.426) V c = V c0 + V cl + V c2 . (4.42c) Substituting in the equivalent values given by equations 4.37, 4.40 and 4.41, equations 4.42 become V a = V a0 + V al + V^, (4.43a) V b = V a0 + a 2 V al + aV^, (4.436) V c = V a0 + aV al + a 2 ^. (4.43c) These equations state that an unbalanced system can be defined in terms of three balanced phasor sets. In other words, positive-, negative-, and zero-sequence components of phase a can be added together to obtain the unbalanced phasors. Following convention, the equations are ex- pressed only in phase a quantities. Similarly, three unbalanced voltages or currents may be resolved into their symmetrical components. Consider the zero sequence first. By adding equations 4.43a, 4.436, and 4.43c together, V a + V b + V c = 3V a0 + (l + a 2 + a)V al + (l + a + a 2 )^. Since 1 + a 2 + a = 0, V a0 = I(V a + V b + V c ). (4.44a) If equation 4.436 is multiplied by a and equation 4.43c by a 2 and these results are added to equation 4.42a, V a + aV b + a 2 V c = (l + a + a 2 )V a0 + 3V al + (l + a 2 + a)V a2 . Therefore, V al = |(V a + aV b + a 2 V c ), (4.446) which relates the positive-sequence component of phase a to the unbalanced vectors. Finally, for the negative se- quence, if equation 4.436 is multiplied by a 2 and equation 4.43c by a, V a + a 2 V b + aV c = (1 + a 2 + a)V a0 + (1 + a + a 2 )V al + 3V a2 . Then, V a2 = |(V a + a 2 V b + aV c ). (4.44c) Equations 4.44a, 4.446, and 4.44c are therefore the reverse of equations 4.43a, 4.436, and 4.43c; they allow the sym- metrical components to be written in terms of the unbal- anced phasors. Symmetrical-Component Relationships Currents in equivalent delta-connected and wye- connected loads or sources form a good basis to illustrate the existence of symmetrical components in three-phase circuits. Consider the two loads shown in figure 4.35, where I ab , 1^, and I ca are the three phase currents and I a , I b and I c are the line currents. These may all be assumed to result from an unbalanced condition. At the three terminals of the delta load, the following relationships are satisfied by Kirchhoff s current law: •'■a = ■'■ab — ^ca> (4.45a) lb — Ibc - Iab> (4.456) ■*-c A ca -4>c' (4.45c) The zero-sequence currents of the wye-connected load are (using equation 4.44a): IaO = 3 da + lb + U (4.46) Substituting equations 4.45 into equation 4.46 it is found that IaO = IWca + Iab + Ibc) " (Lb + Ibc + lj\ = 0. (4.47) This shows that the zero-sequence current of a three-phase circuit feeding into a delta connection is always zero. In addition, the currents to a three-phase wye-connected load with a floating neutral (fig. 4.35B) can have no zero- sequence component. Simply, a neutral-return circuit must be available for zero-sequence currents to flow. However, zero-sequence current may circulate in a delta connection without escaping into a neutral conductor (see figure 4.35A, note directions of I ab , 1^, and I ca ). Figure 4.34.— Symmetrical component addition to obtain unbalanced three-phase set. Figure 4.35.— Equivalent delta-connected {A) and wye- connected (B) loads. 101 For transforming positive-sequence line currents to phase currents, it can be shown from J ai = 3 da + alb + a 2 I c ). (4.48) which applies equation 4.446 to current, that J V3 Using a similar process for negative-sequence currents, Iab2=^Ia2- (4-50) Iabi = -75lai- (4-49) 73 When the foregoing is applied to line-to-neutral and line- to-line voltages for figure 4.33, the transformation equa- tions are V ab0 = 0, Vabi=jV3V al , V M = -W3 V a5 (4.51a) (4.516) (4.51c) where V ab0 , V abl , V ab2 = zero-, positive-, and negative- sequence line-to-line voltages, V, V a0 , V al , V^ = zero-, positive-, and negative- sequence line-to-neutral volt- ages, V. These equations demonstrate another general relation- ship of zero-sequence components: a line-to-line voltage, however unbalanced, can have no zero-sequence compo- nent. Line-to-neutral voltages, on the other hand, may have a zero-sequence value. Symmetrical-Component Impedance Before the solution of unbalanced system problems can be discussed, the concepts of impedance under the influence of symmetrical components need to be covered. Impedance relates the current in a circuit to the impressed voltage. Symmetrical-component impedance behaves in a similar manner, except that it is sometimes affected by additional parameters. There are three likely cases for a power system: an unbalanced static network, a balanced static network, and the balanced nonstatic network. All these impedance values are created by the fact that positive- and negative-sequence currents produce only positive- and negative-sequence voltage drops, respec- tively. The flow of zero-sequence currents in a neutral can result in an impedance that is apparently greater than the actual impedance. In an unbalanced static network, the sequence imped- ances in a particular phase are equal, but not necessarily equal to those in another phase: Z a o - Z al - Z^, (4.52a) Zbo = Zn,! = Zbg, (4.526) Zco = Z cl = Z c2 , (4.52c) where Z a0 , Z b0 , and Z c0 are symmetrical-component imped- ances for the zero sequence; Z al , Z bl , and Z cl are symmetrical-component impedances for the positive se- quence; and Z^, Z b2 , and Z c2 are symmetrical-component impedances for the negative sequence. The balanced nonstatic network is given by Z a o = Z b o = Z c0 , (4.53a) Z a l = Z b i = Z cl , (4.536) Z a 2 = Z b2 = Z c2 . (4.53c) This states that in a balanced nonstatic network the imped- ances in a sequence are equal, but not necessarily equal to the other sequence component impedances. Cables and pow- erlines are included in this case, and specifically, Z a l ~ Za2 =£ Z a0 . (4.54) The last likely case is the balanced static network, where Z = Z K = Z„ (4.55) It should be obvious that this is a situation where sym- metrical components would not normally be applied. As a general rule, positive- and negative-sequence impedances of a power system are on the same order of magnitude, but the system zero-sequence impedance may vary through a very wide range. This range is dependent upon the resistance-to-reactance ratio as seen by the zero-sequence current. Fault Calculations One of the most significant uses for the method of symmetrical components is the computation of voltages and currents resulting from unbalanced faults. The three- phase diagram in figure 4.36 represents a simple power system with a four-wire wye-connected source. The imped- ance of each phase conductor is Z, while Z is the neutral- conductor impedance. A bolted line-to-neutral fault is occurring on a phase a (an x signifies the fault). The resulting current in the fault, If, is of interest, and the following shows how symmetrical components can be used to find its value. V V f ■ / -o VWV o X Fault Figure 4.36.— Three-phase system with line-to-neutral fault. 102 It is apparent from figure 4.36 that the line currents under the fault condition are I a = ^ I b = 0, I c = 0, (4.56) where If = current in fault, A, and I a , I b , I c = unbalanced line currents, A. Applying equation 4.43, the symmetrical components of these currents are the current is 3I . Zero-sequence impedance, Z a0 , is there- fore greater than Z. As was implied in the preceding section, the quantification of Z a0 , or just simply Z , is not an easy matter. However, in order to limit the amount of fault current flowing in mine power-system neutrals, large resistances are placed in the neutral circuit. With this in mind, the resistance-to-reactance ratio of the neutral im- pedance, Z n , is very large, and in this instance for mine power systems under line-to-neutral faults, the impedance seen by the zero-sequence current can be approximated as therefore, I al = | (I a + al b + a 2 I c ) = I If, (4.57a) ia 2 = |(Ia + a 2 I b + ai c ) = |if, (4.576) Ial=ia2 = IaO = 5lf (4-58) Z = Z + 3Z, (4.60) To define the fault completely, it must be known whether the fault between line a and the neutral is a dead short or exhibits an impedance. Although all faults have a finite impedance, the faulting assumption states that it is bolted. Therefore ,Jault impedance is zero and the voltage across the fault, V fn , is also zero. With this, the current through the fault, If, can be described. However, to perform the required computations, it is necessary to know the force driving the fault current and the impedance existing between this driving potential and the fault location. The source, V, is the driving potential, and it can be assumed as purely positive sequence. It is also assumed from figure 4.36 that the source has negligible internal imped- ance (in practical situations, however, the source impedance is of great importance). Therefore, the source line-to-neutral potentials are equal to the terminal voltages: V =VW.=VwV =V 'an v a> v bn v b> v cn v c where V^V^V^ = terminal line-to-neutral voltages and V Q , V b , V c = corresponding ideal-source poten- tials. The impedances involved are simply the line impedance of phase a (Z) and the neutral impedance (Z n ). With these parameters known, the process is now to convert the unbalanced system to symmetrical components, solve the problem in terms of these balanced vectors, and then reconstruct the result for the fault current. Following convention, all work is performed in phase a quantities. Notwithstanding, phase a contains the line- to-neutral fault; thus, it is the only phase involved. Since only positive-sequence voltage is supplied by the source, the symmetrical components of the driving potential are 'anl — *an> *an2> — U, *anO — "■ (4.59) The impedance to positive-sequence or negative- sequence current in any of the three lines is equal; thus, for phase a, Z a i = z, Z a2 = z. Zero-sequence current follows a different path from posi- tive or negative sequence. From the source to the fault, zero-sequence current, I , exists in each line, but from the fault back to the source (through the neutral conductor) Loop equations for each sequence current can now be expressed for figure 4.36. If a voltage is assumed to exist across the fault, for phase a, 'anl = Z al I al + V n , v an2 = z a2 I a2 + V^ = 0, V a „0 = Z I a0 + V ro = 0, (4.61a) (4.616) (4.61c) where V anl , V an2 , V^q = sequence voltages for source, V, I al , I a2 , I a0 = sequence components of fault current, A, Vfi, Vf2, VfQ = sequence components of volt- age across fault, V, and Z al , Z a2 , Z = sequence impedance seen by fault current, fl. Equation 4.59 generalizes the fault condition and is prac- tical because of fault impedance. However, a bolted fault has been assumed to exist; thus, v n = V» = V ro = 0. (4.62) All input to the problem is now available, and simulta- neous solution of equations 4.57 through 4.62 shows that Van = V anl = § ZIf + § ZIf + | Z If (4.62a) or but then If = Va (2Z + Z ) ' (4.626) Z = Z + 3Z r If = Z + Z„ (4.62c) Consequently, symmetrical components have been em- ployed to solve this unbalanced faulting problem. This work can easily be expanded to cover other unbalanced faulting problems, and the process can be employed to solve any unbalanced three-phase or even polyphase condition. Because fault analysis is imperative in protective-device sizing, additional discussion can be found in chapter 10. POWER TERMINOLOGY If the sum of the electrical ratings is made for all equipment in a power complex, the result will provide a 103 total connected load. The measure could be expressed in horsepower, but the electrical quantities of kilowatts, kilovoltamperes, or amperes are more suitable units. Note that the connected horsepower can be converted to con- nected kilowatt simply by multiplying by 0.746. Many loads operate intermittently, especially mining production equipment, and other equipment operates at less than full load. Accordingly, the demand upon the power source is less than the connected load. This fact is important in the design of any mine power system, as the system should be designed for what the connected load actually uses, rather than the total connected load. Obviously, these consider- ations have great impact on power-system investment or the capital required to build the system. Because of the importance of assessing equipment power demands, the Institute of Electrical and Electronics Engineers (IEEE) has standard definitions for load combi- nations and their ratios. The important ones follow (3). • Demand is the electrical load for an entire complex or a single piece of equipment averaged over a specified time interval. The time or demand interval is generally 15 min, 30 min, or 1.0 h, and demand is generally expressed in kilowatts, kilovoltamperes, and amperes. • Peak load is the maximum load consumed or pro- duced by one piece or a group of equipment in a stated time period. It can be the maximum instantaneous load, the maximum average load, or (loosely) the maximum connected load over the time period. • Maximum demand is largest demand that has occurred dur ing a specif ied timejseriod. • Demand factor is the ratio of the maximum demand to the total connected load. • Diversity factor is the ratio of the sum of the individual maximum demands for each system part of subdivision to the complete system maximum demand. • Load factor is the ratio of the average load to the peak load, both occurring in the same designated time period. This can be implied to be also equal to the ratio of actual power consumed to total connected load in the same time period. • Coincident demand is any demand that occurs si- multaneously with any other demand. All these definitions may be applied to the units of average power, apparent power, or current. Thus they are invaluable in power-system design. A few examples are in order to illustrate their versatility. Consider a feeder cable supplying several mining sections in an underground mine. The sum of the con- nected loads on the cable, multiplied by the demand factor of these loads, yields the maximum demand that the cable must carry. When applied to current, this demand would be the maximum amperage. Good demand factors for mine power systems range from 0.8 to 0.7 depending upon the number of operation sections. The lower value is used when there are fewer producing units, that is, from two to four. The demand factor can be extended to include esti- mates of average load. For instance, the sum of the average loads on the cable, multiplied by the demand factor, provides the average load on the cable. A prime applica- tion here is for approximating the current that a conductor is expected to carry. If, for example, 10 identical mining sections draw 53 A each; the conductors feeding all these sections would be expected to carry (total average loadXdemand factor) = (average load) or (10X53 A) (0.8) = 424 A. The demand factor and the diversity factor can be applied to many other mine electrical areas, such as estimating transformer capacities, protective-circuitry continuous ratings, and the load that a utility company must supply. The load factor can be used to estimate the actual loads required by equipment. Here, the total connected load multiplied by the load factor is an approximation of the actual power consumed. It should be noted that the average load factor in underground coal mining tends to be rather low, mainly because of the cyclic nature of equip- ment operation but also because of the employment of high-horsepower motors that are needed to perform spe- cific functions but only operate for a small fraction of the possible running time. For instance, when cutting and loading, a continuous miner will have all motors operat- ing, thus have a total connected load of 385 hp or (0.746)385 = 287 kW. The average load factor might be 0.6; therefore, the actual power consumed is (0.6)287 or 172 kW. The load factor can also be applied to equipment combinations. The maximum power demand normally forms one basis that utility companies use to determine power bills; most often, 1 month is the specified time period. Demand meters are often installed at the utility company metering point. Chapter 4 has covered a broad range of fundamentals projected towards three-phase power systems in mining. Items have included balanced three-phase circuit analysis, the per-unit system, the method of symmetrical compo- nents, and specific terminology to describe power-system operation. Comprehension of this material is vital in order to understand many chapters that follow. REFERENCES 1. American National Standards Institute (New York). ANSI Standard Device Numbers C37.2. 1978. 2. Graphical Symbols for Electrical Diagrams Y32.2. 1970. 3. Institute of Electrical and Electronics Engineers (New York). Recommended Practice for Electric Power Distribution for In- dustrial Plants. Stand. 141-1986. 104 CHAPTER 5.— BASIC SOLID-STATE DEVICES AND INSTRUMENTATION Through the advancement of technology, the motor- generator (m-g) sets and Ignitron rectifiers for power conver- sion used in early mining have been all but replaced by semiconductor devices, except for m-g sets and synchronous rotary converters in specific surface mining equipment. Equipment employing semiconductors exclusively is often termed solid state or static. In mine power systems the use of semiconductors has grown from simple rectification (the conversion of power to direct current (dc)) to include such areas as motor and equipment control, protective relaying, and lighting power supplies, not to mention extensive use in communications and instrumentation. Since the topics of solid-state devices and basic instru- mentation are closely related, they are introduced to- gether in this chapter. The discussion will be primarily informative rather than theoretical. SEMICONDUCTORS Semiconductors are nonmetallic elements that are characterized by relatively poor conductivity. Silicon is the most popular and germanium the second most important semiconductor in electrical or electronic applications. Semiconductors are useful in electrical circuits because they can pass current in two different conduction modes when impurities or imperfections exist in their crystal lattices. The process of carefully adding impurities to a pure or intrinsic semiconductor crystal while it is being grown is called doping. The impurities are selected for their size, so they will fit into the crystal lattice and provide either an excess or a deficiency of electrons. For example, when a few parts per million of arsenic atoms are added to germanium, or antimony atoms are added to silicon in the crystal structure, an overabundance of free electrons is created. The result is a net negative effect, and the crystal is termed an n-type semiconductor. If a potential is placed across the impure crystal, conduction occurs primarily through an apparent drift of these free electrons. On the other hand, if indium or gallium is used to dope germanium or silicon, a deficiency of electrons exists, and an excess positive charge is created in the doped crystal. Thus, it is called a p-type semiconductor. If a potential is applied, the atoms conduct current by an apparent movement of electron sites or holes. These holes are places in the crystal lattice where an electron can be held temporarily. When there is an abundance of holes, free electrons generated within the crystal can quickly recombine with available atoms. The free electrons in the n-material and the holes in p-material are known as majority carriers. However, be- cause of thermal and other energies, free electrons are also found in a lesser amount in the p-type and a few holes exist in n-type semiconductors. These are called minority carri- ers. Nevertheless, even with the excess charge, both semi- conductor types are electrically neutral. DIODES AND RECTIFIERS The operation of most semiconductor devices is depen- dent upon a p-n junction, which is the boundary formed when a piece of p-type material is joined with a piece of n-type. In the actual production, a single semiconductor crystal (or monocrystalline material) is grown so that part is doped to create a p-type region, with the balance doped for n-type. A solid-state diode or rectifier has one p-n junction; it is a device that readily passes current in one direction but does not permit appreciable current in the opposite direction. The symbol for a diode or rectifier is given in figure 5.1 A. Figure 5.15 is a simple model of a diode that can be used to explain p-n junction electrical operation. When the two semiconductor materials are joined, a charge redistribution occurs. Both the p-region and the n-region contain a high concentration of majority carriers. Elec- trons from the n-material diffuse across the junction to the p-material; similarly, holes migrate from the p-material into the n-material. The net result of this diffusion is a depletion region with negatively charged (acceptor) ions on the p-side and positively charged (donor) ions on the n-side of the junction. The electric field across the deple- tion region is established and opposes further majority- carrier diffusion, but the field creates a minority-carrier flow across the junction in the opposite direction. Current caused by majority-carrier diffusion is called injection current, I ls and that from minority carriers, saturation current, I s . If no external voltage is applied to the p-n device (fig. 5.2A), the junction is in equilibrium because the net hole and electron flow across the junction is zero. In other words, injection current equals saturation current. However, if an external voltage is applied with a polarity such that the p-region is positive with respect to the n-region (fig. 5.25), the depletion-region electric field is decreased, and a large number of majority carriers are able to cross the junction and diffuse toward the device terminals. Hence, injection current is substantially in- creased, and because saturation current remains constant, the result is current flow in the external circuit. In this case, the external voltage polarity is called forward bias, p-type June tion n-type L±J Ld L+J Ld 69 eeeee © ® © © ® © © ++++++ — — — — — — 0990969 ©©©©©©© + + + ++ + Depletion region Figure 5.1. —Symbol (A) and operation (8) of a p-n junction device. Js -M- Js I s =l! Is <:[ I I S >I I A No external voltage B Forward bias C Reverse bias Figure 5.2.— Bias conditions and current flow for a diode. 105 and the current is forward current. Conversely, reverse bias, an applied voltage of reverse polarity (fig. 5.2C), opposes majority-carrier diffusion by enforcing the depletion-region electric field, and current is greatly re- duced. As saturation current is still constant, the external reverse current is primarily a result of minority-carrier diffusion and is therefore very small. Because there are many more majority carriers than minority carriers, the injection current, under forward bias, is orders-of-magnitude greater than the constant saturation current. As external circuit current is the algebraic sum of injection and saturation currents, for- ward current is significantly greater than reverse current. Furthermore, to enhance this one-directional current phe- nomenon, junctions are manufactured in which one side of the junction is more heavily doped than the other. Forward current is then mainly a result of majority carriers from the more heavily doped region. The arrow portion of the diode symbol (fig. 5.1A) points in the same direction as forward current. As a carryover from vacuum-tube terminology, the side symbol- ized by the arrow is also called an anode (the p-region), with the opposite terminal, the cathode (the n-region). the external current is about equal to the saturation current. Therefore, by placing a reverse bias across the device and measuring the resulting reverse current, the forward current can be predicted. The foregoing equations result in the theoretical curve, termed a characteristic curve, which is given in figure 5.3. This curve diverges from that for an actual diode in one main aspect, the breakdown of the p-n junction noted at point c. Here, the external voltage meets the limit capabilities of the junction, and a greater reverse voltage will create an avalanche current that can destroy the device. As a result, p-n junctions normally require a rating for maximum reverse voltage or peak inverse volt- age (PIV). Zener diodes are of special interest as they operate in this avalanche current area to regulate an applied dc voltage. As long as the p-n junction is operated within the limits of its reverse voltage and forward current, the device can be represented by a very low resistance for forward- bias conditions and a high resistance during reverse bias. Ideally, and for the majority of applications, a diode can be assumed to have zero resistance under forward bias and infinite resistance under reverse bias. Diode Equations Rectifier Circuits The number of minority carriers is dependent upon temperature and the difference in energy levels between the p- and n-regions. If the energy difference is constant, the concentration of minority carriers plus the saturation and reverse currents will vary exponentially with temper- ature. Therefore temperature is a limiting factor in diode operation, and the maximum rated current of a given device is determined by the heat-dissipating properties of the device mounting system. The formula relating external and saturation current with the energy difference and temperature is I = -I s (e KT - 1), (5.1) where I s = saturation current, A, q = charge of one electron, 16 x 10 ~ 20 C, V = voltage across junction Qess than external volt- age, but approximately equal to it), V, qV = energy difference between p- and n-materials, K = Boltzmann constant, 1.38 x 10 ~ 23 J/K, T = junction temperature, K. and At room temperature (300 K), I = -I s (e 39V - 1), (5.2) or at other temperatures, I = -I s (e (39VT ^ - 1), (5.3) where T 1 = 300 K, and T 2 = other temperature, K. The negative sign for the saturation current denotes it as flowing in the opposite direction to forward current. The equations relate that if voltage is 0.1 V or more negative, A rectifier can be considered as a diode specifically designed or applied to convert power to dc. The principal application in mining is to use the unilateral properties of tbe rectifier for direct alternating current (ac) power conversion. With single-phase ac, there are three basic rectifier circuits to perform this function: half-wave, full wave, and bridge. Figure 5.4A illustrates the circuit of a simple half- wave rectifier in which a transformer magnetically cou- ples the source to the rectifier. This could also be direct, unisolated source connection. With a sinusoidal voltage input (fig. 5.4B), the rectifier acts as a switch. When forward biased (positive anode with respect to the cath- ode), the load, R, is electrically connected to the source, but during reverse biasing it is disconnected. In other words, low and high resistances to current exist with respect to the bias condition. These resistances create a pulsating dc waveform across the load, as shown in figure 5.4C. This variation of voltage is often termed ripple. ~]~ Theoretical ,'N / Actual / b 1 c i i i i i d < VOLTAGE, ICf'V u cr_ => v Figure 5.3.— Diode or rectifier characteristic curve. 106 Only the positive portions of the input sinusoid ap- pear in the pulsating dc output, and as a result, the conversion efficiency of the half- wave rectifier leaves much to be desired. The single-way full-wave rectifier is a method of rectifying both the positive and negative por- tions of a sinusoidal voltage input, and it can be analyzed as two half-wave rectifiers. The circuit shown in figure 5.5A utilizes a center-tapped transformer secondary. When referenced to ground, the V 2 and V 2 ' waveforms (fig. 5.5J3) are then 180° out of phase. Therefore, one rectifier is conducting current (forward biased) while the other is not (reverse biased). The consequence is pulsating dc power to the load during both the negative and positive portions of the ac input (fig. 5.5C). Conversion efficiency is greatly improved over half-wave circuits. Full-wave rectification can also be obtained with the bridge rectifier. As shown in figure 5.6A, the circuit employs a transformer with a single secondary and four rectifiers. During either the positive or negative portions of the input waveforms, two of the rectifiers are effectively in series with the load resistance. For instance, when the top secondary transformer rectifiers D 2 and D 3 are forward biased but D x and D 4 are reverse biased, current flows from the top seco ndary ter minal thro ugh D 2 , R L , and D 3 back to the transformer. The rectifier biasing condition reverses with the transformer secondary polarity (figure 5.6B, bottom), but the current through the load has the same direction. Hence, the same full-wave pulsating dc waveform in figure 5.6C appears across the load with only half of the secondary turns needed for the single-way full-wave rectifier. Although the output of these three basic rectifier circuits is effectively dc and the current flow is in only one direction, the voltage fluctuation or ripple is often too great to be useful. Consequently, filtering is required to change the pulsating voltage to a relatively ripple-free potential. This filtering action is provided by inductors in series with the load, or capacitors shunting (in parallel with) the load, or both. Each of these methods will smooth the voltage output. An example of this filtering is shown in figure 5.7. It will be shown later that such filtering is not needed for dc mining equipment. I W T ° c t L ♦ out ^d- A r\ AT v, out A B C Figure 5.4.— Half-wave rectifier circuit and waveforms. V 2 J v out V2 I /YY^ A B C Figure 5.5.— Single-way full-wave rectifier waveforms. zf\f ^V rYY\ lv, 'out Figure 5.6.— Bridge rectifier circuit and waveforms. Cooling It was stated earlier that the operation of a p-n junction is highly dependent upon temperature. It follows that there exists a maximum temperature beyond which the device will be destroyed if operated. Such a point is called the maximum junction operating temperature. For silicon semiconductors, this temperature is usually around 175° to 200° C, for germanium, 85° to 110° C, but the maximum varies according to the individual device and manufacturer. The temperature at which the junction operates is dependent upon the power dissipated in the junction, the ambient temperature, and the ability of the device to transfer heat to the surrounding environment. Devices designed and operated for small currents usually do not need cooling assistance. However, adequate external cool- ing is required in p-n junctions dissipating 1 or more watts. The simplest method is to mount the semiconductor case securely on a heat sink, which is commonly metal with a large surface area. Thermally conductive washers, silicon compounds, and correct bolting pressure allow good heat transfer from the device to the heat sink, and air Series Rectifier ^ inductor output \ \ Capacitor shunting load Figure 5.7.— Example of filtering a rectifier output. convection transfers heat to the surrounding atmosphere. In high-power applications, forced-air cooling of the heat sink is sometimes employed to increase heat dissipation further. Figure 5.8 illustrates a rectifier using a heat sink for this purpose. The diagram in figure 5.9 represents the typical relationships in all solid-state devices between the 107 Heat sink Figure 5.8.— Heat sink cooling. Collector junction temperature _ Case temperature, T c _ Heat sink temperature, T s Ambient temperature, T a JL*. Absolute-zero temperature Figure 5.9.— Heat sink thermal relationships. solid-state device, its heat sink, and the surrounding envi- ronment. The following equation relates these parameters: T j = 0jaPd + T a» (5.4) where Tj = junction temperature, °C, T a = ambient temperature, °C, P d = power dissipated by junction, W, and ja = ambient-to-junction "thermal resistance," °C/W. The last item, thermal resistance, is actually composed of three parts, as shown in figure 5.9, 0ja = 0jc + Acs + C (5-5) where jc = junction-to-case thermal resistance, °C/W, flea = case-to-heat-sink thermal resistance, °C/W, and flga = heat-sink-to-ambient thermal resistance, °C/W. The junction-to-case and the heat-sink-to-ambient thermal resistances are almost always available from manufactur- ers. The thermal resistance between the device case and the heat sink can be neglected if the mounting is carried out correctly as described here. Junction power can be found by the relationship *d — Imax » f» where I,,,^ = maximum forward current, A, and V f = junction forward voltage drop, V. (5.6) The junction forward voltages normally range from 0.5 to 0.75 V for silicon and from 0.2 to 0.3 V for germanium, but typical values for specific devices are also available from manufacturers. When the total thermal resistance, ja , is known, the operating junction temperature can be calcu- lated and compared with the maximum allowed. Overloads The thermal relationship of figure 5.9 shows three capacitances, Cj, C c , and C s , which are the thermal capac- itances of the p-n junction, the device case, and the heat sink, respectively. Thermal capacitance resists changes in temperature in the same way that capacitance restricts voltage change. For the p-n junction, Cj is usually very small; hence, its time constant is also small. This means that the semiconductor must not be overloaded (excessive power dissipation) for more than a few milliseconds; other- wise, the device will be destroyed. For this reason, high- speed overload protection must be applied to semiconduc- tor devices. For rectifiers, the protection takes two forms: against excessive overloads and short circuits in load currents, and against failure in the rectifier itself (over- temperature or excessive voltage). THREE-PHASE RECTIFICATION Large amounts of dc power at either 250 or 500 Vdc are required for locomotives and face equipment in many mining operations. When more than a fractional kilowatt of dc power is needed from an ac source, a polyphase rectifier circuit is employed. The direct voltage is derived from three-phase ac power, most often from distribution voltages. There are specific advantages to using polyphase rectifier circuits for dc power. As the number of ac phases driving the rectifier is increased (say, above single-phase ac), the frequency of output ripple is increased, the inter- val between rectifier conduction is decreased, and the ripple magnitude in the dc voltage and current waves decreases. Transformers are almost always used between the ac source and the rectifiers. The rectifier transformer per- forms one or more of the following functions: • To transform the available ac supply voltage to a value needed for the desired dc voltage; • To provide the number of phases required to obtain the desired waveshapes of dc voltage, dc current, and ac supply current; • To isolate the dc circuit from the ac source; and • To limit, through transformer impedance, damag- ing overcurrents that might flow during malfunctions. It is important to note that the decrease in the rectifier- conduction interval also increases the required trans- former rating. The transformer utilization factor can be defined as the ratio of dc power delivered to the required transformer secondary voltampere rating. The utilization factor has been found to have a maximum value of 0.520 when three-phase ac input is used. This implies that from a transformer utilization standpoint, the most economic rectifier-conduction angle is 120°. When power rectifiers are mentioned today, the refer- ence is almost invariably to solid-state units using silicon 108 rectifiers as the rectifying elements. Indeed, the silicon rectifier is virtually the only type considered for mine power installations. While there are many possible recti- fier circuits, only two or three types are found in mining equipment. Circuits for silicon rectifiers are selected to make the most efficient use of the transformer, and the results usually are the single-phase full-wave bridge or the three-phase full-wave bridge. The next section will discuss three fundamental three-phase rectifier circuits, and it will be apparent why the full-wave bridge is popular. Phase Phase A C t Phase B Rectifier Circuits Rectifier circuits can be classified as single way or double way. The phase currents of the transformer secondary (also termed the dc winding) are unidirectional in a single- way circuit but alternating in the double-way circuit. The simplest three-phase rectifier circuit is the three- phase half- wave shown in figure 5.10A, where a delta-wye transformer is used, with each leg connected to a rectifier anode. The three rectifier cathodes are tied together to form the positive dc bus. The neutral point of the trans- former winding serves as the negative connection for the load, in this case resistance, R. Being a single-way recti- fier, each leg of the transformer secondary conducts cur- rent unidirectionally. If the load is pure resistance, the relationship of output voltage (that across the load) versus time is as shown in figure 5.105. Each rectifier conducts for the cycle portion in which its anode has a higher positive value than the anodes of the other rectifiers. Therefore, each rectifier passes current for 120° of the input three-phase cycle. Since the current through the load is directly proportional to the output voltage, the load current has the same waveform as voltage. Inspection of the three-phase half- wave output voltage shows that the ripple voltage is much lower than the single-phase full-wave rectifier circuit. Actually, the rms value of the ripple voltage waveform is only 18% of the average output voltage (this average voltage is the average dc load voltage, V dc ). If rectifier losses are ignored, since they are very small for silicon diodes, the dc output voltage and the transformer secondary voltage are related by A B C A B V V i 1 i 180° out, deg 360° Figure 5.10.— Three-phase half-wave rectifier circuit (A) and output voltage waveform (6). Phase Phase A C D, D 3 D 5 D 2 D 4 D 6 cut, deg V dc = 0.827V max = 1.17V™, (5.7) where V dc = average dc output voltage, V, V max = peak value of voltage applied to rectifier circuit, V, and Vj^g = rms value of voltage applied to rectifier cir- cuit, V. Both V max and V^g are line-to-neutral voltages. Note that the fundamental frequency is three times the ac line frequency. As a result, any filter components required to lower the ripple voltage further can be much smaller than in single-phase rectifiers. The relationships presented here for the three-phase half-wave rectifier apply only to ideal transformers and rectifiers. In actual circuits, the voltage drop caused by dc current and the transformer secondary-winding resistance creates a dc component that pushes transformer magnetic operation toward saturation. Consequently, this simple three-phase rectifier circuit is seldom used. Output ripple can be further reduced by a three-phase full-wave rectifier circuit, connected as shown in figure 5. 11 A. This circuit is also called the three-phase bridge or 180° C cut, deg Figure 5.11.— Three-phase full-wave rectifier circuit (A) with input (0) and output (C) voltage waveforms. a six-phase rectifier. Being a two-way rectifier, the mag- netic saturation problem in the transformer is not present. Furthermore, this configuration retains the advantage of 120° conduction for transformer economy, plus a funda- mental ripple frequency of six times the ac source fre- quency. These characteristics make this double-way recti- fier circuit of great practical value, and it is the most popular configuration for dc power in mining. The trans- former dc winding may be either wye or delta connected. In the full-wave rectifier circuit, each terminal of the transformer secondary is connected to two diodes, one at 109 the anode and the other at the cathode. The cathodes of three rectifiers are common and form a positive dc voltage bus, while the common anode connection of the other three rectifiers represents the negative dc voltage bus. The load is connected between these two common points. Each rectifier conducts for 120° of one input cycle, and current alternates in each transformer winding. However, current flows through a specific combination of rectifiers for only 60° of the input cycle. This combination could be D x and D 4 with transformer secondary terminals A and B. Therefore, the peak-to-peak voltage across the load resis- tance appears as six-phase ripple as shown in figure 5.11C. Analysis of figure 5.11C shows that the rms value of the fundamental component of the ripple voltage is now only 4.2% of the average dc output voltage. In addition, the average dc output voltage for ideal rectifiers is V„„ = 0.95V„„ = 1.34V. dc (5.8) where V dc = average dc output voltage, V, V max = peak line-to-line voltage applied to rectifi- ers, V, and Vrmg = rms value of line-to-line voltage applied to rectifiers, V. The foregoing circuits are typical of most polyphase rectifier circuits, but many additional configurations are available. Because mining almost always employs full- wave rectifier circuits, coverage of more circuits is beyond the scope of this text, but the bibliography can be con- sulted if desired. Parallel Rectifier Operation The current requirements of a rectifier circuit are often too large to be handled by one rectifier for each circuit element. Two or more rectifiers must then be connected in parallel. Direct operation of two silicon rectifiers in parallel is very difficult, because unbalance between the parallel paths can be caused readily by unequal rectifier characteristics (mainly the forward volt- age) and by unequal impedance in the bus bars or cables. The result is that the rectifier with the least forward voltage can be destroyed by overcurrent. To eliminate this problem, the parallel rectifiers must be forced into sharing the current equally. The method used almost exclusively in mining equip- ment to force current-sharing employs paralleling reac- tors, sometimes called current-balancing transformers. Figure 5.12 shows how several rectifiers can be paralleled using these reactors. The combination acts as one rectify- ing element in a rectifier circuit such as in figure 5. 11 A. In figure 5.12, each reactor is a laminated magnetic core linked in opposing polarity by the anode currents of two rectifiers, and the cores are designed not to saturate at the highest expected current. If the two rectifier currents become unequal, the current difference excites a magnetic flux that induces an aiding voltage. This voltage is in- duced in the rectifier leads in a direction that will equalize the currents. TRANSISTORS The principal tool of the electronics industry is the amplifier, a device that can increase the power level of an input waveform or signal. An amplifier is actually an energy converter in which energy from a power supply is converted by the amplifier to signal energy. The most common device used in amplifiers is the transistor. A bipolar transistor is formed in a manner similar to that of the junction diode, but it consists of two junctions in close proximity and parallel to each other in the same crystal. When a p-region is sandwiched between two n-regions, the device is termed an n-p-n transistor, the model and symbol of which are given in figure 5.13A. Similarly, if a thin portion of n-material is bounded by two p-regions, the transistor is termed p-n-p, as shown in figure 5.14A. As illustrated, each semiconductor region is given a name: emitter, base, and collector. Transistor Operation The operation of the transistor is dependent upon the bias voltages across the junctions. If voltages are applied to an n-p-n device as shown in figure 5.13B, the emitter- base junction is forward biased, and the collector-base Reactor ac input dc output Figure 5.12.— Parallel operation of rectifiers using paral- leling reactors. Base Symbol e C Emitter Collector B V CI A B Figure 5.13.— An n-p-n junction transistor. Base Emitter Collector E C Veb B V CB LVhI t±|.lpJ A B Figure 5.14.— A p-n-p junction transistor. 110 junction is reverse biased. These are the normal bias conditions. Electrons will flow into the base region, caus- ing an excess of majority carriers there. Because the base region is thin and the potential existing across the two n-regions is much higher than the base-to-emitter poten- tial, most electrons from the emitter region diffuse across the base and are accelerated into the collector region. The electrons drift across the collector and cause current flow in the collector circuit. However, a small percentage (typ- ically, 5% or less) flows out from the base connection because of recombination with holes in the base region. This process can be considered amplification since the small base current controls the much larger collector current. A p-n-p transistor operates on the same princi- ple, but here it is hole flow rather than electrons that causes the amplification. Consequently, the bias condi- tions are reversed from that for an n-p-n (the normal conditions are shown in figure 5.14B). From the preceding discussion, it would appear that either end of the transistor could be called an emitter because either hole flow or electron flow creates the current amplification, but this is generally not the case. Heat dissipation is much larger in the collector-base junction because of the greater difference in potential. Therefore both p-n-p and n-p-n transistors are designed so this heat can be diffused through the collector region. As might be expected, a saturation current resulting from thermally-generated minority carriers flows across the reverse-biased collector-base junction. In the diode, the current is designated "I s ;" in a transistor, it is termed "Icbo" I n the same manner as for diodes, the increase of saturation current with temperature sets the maximum operating temperature for a transistor. Heat sinks are commonly used in high-power transistor applications to diffuse collector-base junction heat and maintain temper- ature below critical levels. The same calculations that were presented in the preceding section on rectifiers can also be applied to transistors to determine a safe operating temperature. The fraction of constant emitter current that reaches the collector is called alpha, a, and the collector circuit itself can be considered to be the output circuit. Since as much emitter current as possible should be collected, alpha should be as close to 1 as possible. When combined with Icbo' th e collector current, i c , can be expressed in terms of emitter current, i E , as in = ai, + Ir (5.9) Figure 5.15 shows the relationship of these currents. However, in practical applications, I CB o * s often so small that it can be neglected. Since base current controls collector current, an im- portant expression can be obtained from figure 5.15 using Kirchhoff s current law on either transistor: 1r = lp - i c> or i B = i E - (ai E + I CBO ) = (1 - a)i E - Icbo- In terms of collector current, it can be shown that Icbo (5.10) Ir- = 1 - a Ir + 1 - (5.11) The term, a/(l - a), is called beta, /3, and also the dc current amplification factor, and i c = /3i B + (1 + 0)1 CBO- (5.12) This last equation shows the significant effect of temper- ature on transistor operation; that is, the temperature- sensitive I CBO is multiplied by (1 + /3). Even though a is less than 1, ft may range from 20 to 200 for amplifying transistors. Bipolar-Transistor Amplifiers Bipolar transistors can be operated with any one of the terminals common to the input and output, thus there are three basic circuit arrangements: common-base, common-emitter, and common-collector. The most popular is common-emitter. Illustrated in figure 5.16, the common-base or grounded-base configuration employs the emitter and base terminals as input, with the collector and base terminals supplying output. Current gain, which is the ratio of output to input, is usually just less than 1. Because the emitter-base junction is forward biased, the circuit has low input impedance as viewed from the input terminals. Because the collector-base junction is reverse biased, the output impedance is high in comparison to the input. Hence, voltage and power amplification can be realized. Two different circuits, signal and bias, are necessary for the operation of either of the two common-base ampli- fiers shown in figure 5.1. The bias voltage sources, often termed the amplifier power supply, fix the dc level for proper operation of the two junctions. If the signal input and output are not separated electrically from the bias source, as seen in figure 5.16A, the circuit is called a dc amplifier. Although it is beneficial in applications such as amplifying dc voltages for instrumentation, a signal with E C 'C E C L CB0 'B L CB0 A B Figure 5.15.— Current relationships for p-n-p {A) and n-p-n (B) devices. ^* p-n-p Ri 'i(D IR 'BB =L V CC A B Figure 5.16.— Common-base amplifiers. cc =b- Ill dc content or offset can interfere with correct transistor biasing. Figure 5.16B illustrates a popular method of removing this problem: the use of capacitors to isolate the amplifier. The capacitors exhibit high impedance to dc but low impedance to ac signals, thus they block input and output dc. As the circuit now reacts only to ac signals, it is called an ac amplifier. It can be noted that transformers can perform a similar function. With either the dc amplifier or the ac amplifier, a small change in input voltage causes significant variation in the injection current across the emitter-base junction. As previously discussed, most majority carriers diffuse to the collector, causing collector current, i c . If the load resistance, R L , is small with respect to the transistor output impedance, i c is approximately equal to i E . The collector current creates voltage variations across the load resistance that can be much larger than the input voltage. In the common-emitter transistor arrangement, the source signal only supplies current to the base. Because base current is much smaller than either the emitter or collector current, current amplification or gain, Gj, is high. Neglecting Icbo in equation 5.12, the gain is approx- imately equal to G _ L c _ & { B _ (5.13) which can be from 10 to several hundred. The input impedance is also higher than in common-base amplifiers. Figure 5.17 shows a simple common-emitter ampli- fier. The control action of the base current can be demon- strated by assuming that the base-emitter forward bias is increased. This increase creates a corresponding increase in emitter-base junction current; thus, collector current is raised substantially. Because the base current is approxi- mately proportional to but usually much less than collec- tor current, base current is the controlling parameter of the amplifier. The concept of characteristic curves has already been introduced in figure 5.3 in the section on diodes and rectifiers. Characteristic curves are an extremely useful tool for the graphical design and analysis of transistor circuits. Four independent transistor parameters control the number of necessary curves. When figure 5.17 is used, these parameters are as follows: • a and /3 increase with V CE , the collector-to-emitter voltage. • i B is dependent on i c and V CE . • i B is not a linear function of i c . • When V CE is zero, i c is approximately zero, regard- less of i B . Consequently, two sets or families of curves are needed: 1. Collector or output characteristics, i c versus v CE for varying values of i B , and 2. Common-emitter input characteristics, V BE versus i B for varying values of V CE . Figures 5.18A and 5.18B show typical output and input characteristics for an n-p-n transistor connected for common-emitter operation. The nonlinear and propor- tional properties of the four independent transistor param- eters are evident in the graphs. These curves can be employed for design and analysis purposes. The analysis often uses a load line (the straight line in figure 5.18A) to observe dynamic variations of voltage and current. The dashed line in figure 5.18A is very important as it delineates the safe operation boundary. Manufacturers specify maximum permissible collector voltage, current, and power dissipation, since outside this area damage to the transistor will probably result. As noted earlier, allow- able power dissipation must be reduced as temperature is increased. 'B t v CE H(-s 'B> h| t—M' Figure 5.17.— Common-emitter amplifier. J ^W | V CE X 1 ^ t0 ySMfi E K U I ll M III. 1 4 s Safe operation boundary 10 20 COLLECTOR-TO-EMITTER VOLTAGE (V CE ), V A Output Dynamic input 20 V QL characteristics ^-15 V LU> 0.5 >^10V t w 4 ^.^5V LU> 1 w 3 *-<2 1 IS .2 <> 1 r i i i v CE =o 30 60 90 120 150 BASE CURRENT (I B ). M B Input Figure 5.18.— Common-emitter characteristic curves. 112 Figure 5.17 illustrates an amplifier circuit with two batteries supplying dc for transistor bias, but single dc source for all bias voltages is more desirable in practical applications. Three bias techniques are frequently used for common-emitter amplifiers, and these are shown in figure 5.19. Each circuit uses resistors to supply dc bias to the base for a center bias condition about which the transistor operates. The center condition is termed the quiescent point of the amplifier. Of the circuits illus- trated, the stabilized bias circuit (O gives the best thermal stability, maintaining the quiescent point within a desired or specified range regardless of the normal operating temperature. The bypass capacitor, shown across the emit- ter resistor of the stabilized bias circuit, establishes a constant base bias bypassing or acting as a low impedance to time-varying voltages. The two preceding amplifier configurations employed the collector circuit for output. In the common-collector arrangement, the output is obtained across a load resis- tance in the emitter circuit, as illustrated in figure 5.20. Because the source and output voltages are now in series but have opposing polarities, the circuit gives high input impedance and approximately unity voltage gain, yet current gain is about the same as in common-emitter amplifiers. A main advantage of the common-collector is that the output impedance is about equal to the load resistance, which is lower than the preceding two connec- tions. This allows the circuit to be adjusted to fit the output needs precisely; hence, this circuit can be used for impedance matching the output of a source signal to the input of another amplifier. Field-Effect Transistors The n-p-n and p-n-p junction transistors just covered contained two junctions. Field-effect transistors (FET's) have effectively only one junction but still can operate as amplifiers. These devices are voltage controlled, whereas bipolar transistors can be considered as current-controlled devices. There are two general classifications: junction FET's and metal oxide semiconductor FET's. Both have very high input impedances, much higher than bipolar transistors and approaching the input impedance of vac- uum tubes. To demonstrate the amplifying action available with FET's, consider the cross-sectional model of an n-channel junction FET, illustrated in figure 5.21A. The gate-to- channel junction is reverse biased by placing the voltage V GS between the gate and source terminals as shown. The level of V GS establishes a specific size of depletion region about the gate semiconductor and within the channel. Changing this reverse bias increases or decreases the size of the depletion region and decreases or increases the available conduction area remaining in the channel. Therefore, voltage changes between the gate and source terminals can control the allowable current through the channel from the drain to the source terminals. The action can be employed to amplify voltages or currents. The conduction channel in the junction can be either n-type or p-type semiconductor, with the gate being p- or n-material, respectively. Figures 5.21B and 5.21C give the symbols for either junction FET type. An important ad- vantage of FET's over junction transistors is that the source-to-drain channel is resistive without a diode effect. In essence, this allows FET's to be operated as electrically controlled resistors. r-AM/V- neW ,JL A Fixed bias B Self-bias Bypass capacitor • » — o * C Stabilized bias Figure 5.19.— Bias techniques for common-emitter amplifiers. oV r Figure 5.20.— Common-collector amplifier arrangement. p-semiconductor gate q .Jj3 ^ gA d Vest DS B n-channel symbol Ig n-channel semiconductor A Simple bar model V SG { D SD C p-channel symbol Figure 5.21.— Model and symbols for junction FET's. 113 As an application example, figure 5.22 shows a junc- tion FET used in a typical amplifier circuit. The input signal is applied across the gate to the source, with output taken from drain to source. R s is employed to set the proper dc quiescent point bias for the gate, and the capacitor in the source circuit bypasses ac, thus maintain- ing the bias level. In metal oxide semiconductor FET's (or MOS-FET's), the depletion region used in the junction FET is replaced by a thick layer of silicon oxide, a good insulator, and the semiconductor employed for the gate is replaced by a metal conductor, thus forming a high-quality capacitor. A model of a MOS-FET, including the symbols, is given in figure 5.23. The operation of these transistors is similar to that of junction FET's but much more complex. The preceding information on transistors is intended as just an introduction to a few important devices. For complete information, the bibliography must be consulted. The coverage here is justified because transistors are an extremely important, but often hidden, segment of mine power systems. The next section will cover another device that has revolutionized the control of electrical machinery. SILICON-CONTROLLED RECTIFIERS In past few years, the use of solid-state power equip- ment in mining has accelerated. One primary reason has been the introduction and acceptance of static or solid- state starting of conveyor-belt drive motors. The heart of these starters is the silicon-controlled rectifier or SCR. SCR's have many other applications; among these, the most common is in dimmers for home lighting. SCR's, also called thyristors, are three-terminal semi- conductor devices having a four-layer p-n-p-n material combination. Figure 5.24A shows a model of the SCR construction. The outer two layers act as a p-n junction and the inner layers serve as an element to control that junction. The symbol for the SCR is given in figure 5.24.B, and figure 5.25 illustrates how the operation of the three- junction combination can be equated to two transistors connected as shown. The equivalent circuit is represented by one n-p-n and one p-n-p transistor. When the bias on the gate, the n-p-n transistor base, is negative with respect to the cathode, the n-p-n transistor cannot conduct appreciable current. In other words, it is cut off. As no n-p-n transistor collector current can flow, the p-n-p transistor is also cut off. There is high impedance between the anode and cathode for this bias condition, and the SCR operating condition is called OFF. However, if the gate bias is made positive so that the n-p-n transistor conducts, current will flow into the n-p-n collector from the p-n-p transistor base. This p-n-p base current in turn causes collector current in the p-n-p transistor. The action between the two transistors has a positive feedback effect because an increase in current in one transistor creates an increase in the other. Therefore, once conduction in the SCR is estab- lished, the gate no longer has any controlling effect, and Metal Source Gate(-)/ ^ Drain S1O2 AsfetfSdg D, drain p (substrate)x\\y^ j ga t te^^' bs ^ ate gate Model S, source n- channel symbol D, drain SS, substrate S, source p-channel symbol A Depletion mode operation D, drain D, drain G,o-(fs)-° SS, g, »-(jt)-^ SS, gate ^n substrate g a t e ^^ substrate S, source n-channel symbol S, source p-channel symbol B Enhancement mode operation Figure 5.23.— Model and symbols for MOS-FET devices. i> Anode Gate n — -o V 7- Cathode SCR A B Figure 5.24.— SCR model (A) and symbol (8). -K- €) Input \P r :Ri R sl 4c c 4-% ^^— Output DD l fl e c P b n n — o P P n C Figure 5.22.— Example of a junction-FET application. Figure 5.25.— SCR equivalent model and circuit. 114 the SCR is latched ON; that is, anode-to-cathode imped- ance becomes very low. The gate cannot turn the SCR conduction OFF. Cessation of current requires a negative gate bias and an essentially zero anode-to-cathode voltage. This allows the p-n-p transistor to cut off. The OFF and ON characteristics are apparent in the typical curve provided in figure 5.26. The breakover voltage noted here is the anode-to-cathode potential at which the SCR will turn itself ON. There are many applications for the SCR or thyristor. Some of the devices and system components that the thyristor replaces include • Thyratrons, • Mercury-arc rectifiers, • Saturable-core reactors, • Relays and contactors, • Rheostats and motor starters, • Constant-voltage transformers, • Autotransformers, and • Mechanical speed changers. Thyristor applications are a major subject in chapter 14. INTEGRATED CIRCUITS of this layer to interconnect the different regions. The top view of an actual IC is provided in figure 5.28. These devices can contain hundreds of transistors but can be small enough to pass through the eye of a needle. Except for high-power applications, IC's are preferred over discrete-component assemblies because they add reli- ability to equipment while reducing both size and cost. Consequently, IC's are employed where specific circuits require many transistors, diodes, and resistors. In circuit diagrams, it is accepted practice to show only the symbol for the specific application; some of these are given in figure 5.29. The use of IC's is extremely widespread in recently manufactured mining equipment, especially in control, monitoring, and communications applications. BASIC INSTRUMENTATION Much has been said in the preceding chapters about electrical parameters and their quantification: voltage, current, power factor, power, and so on. Instruments that measure these quantities are necessary to monitor and troubleshoot the operation of a power system and can be used to ensure optimum operation and to find malfunc- tions. The devices can be indicating instruments or record- ing instruments that are permanently installed in major The semiconductor devices discussed so far are termed discrete components if they are manufactured as single units, for example, one diode or one transistor. They must be combined with other electrical and electronic compo- nents to perform any required function. Manufacturing processes have been refined so that several transistors, diodes, and resistors can be made in a single circuit, or in other words on one single semiconductor chip. Such de- vices are termed integrated circuits (IC's), and their study in electrical engineering is known as microelectronics. Today, many circuits requiring numerous individual tran- sistors, such a complete amplifiers and digital computers, are packaged in a single semiconductor chip or microcir- cuit. When employing only one semiconductor chip, the IC is called monolithic; when the unit is created by intercon- necting more than one microcircuit, the device is a hybrid IC. The structure illustrated in figure 5.27 represents the cross section of a simple monolithic IC. The device is fabricated on a chip of p-type semiconductor, termed a substrate, by forming a number of junctions. The three sections shown are electrically isolated by reverse-biased p-n junctions, and the silicon surface is protected by a silicon oxide layer. A thin film of metal is deposited on top Avalanche breakdown ON characteristics U^ / Breakover / voltage OFF characteristics npn transistor Diffused resistor KEY B Base E Emitter C Collector n+, n 2 , n 3 Various n-type regions Silicon dioxide layor, metal film on top to interconnect components. Figure 5.27.— Sketch of simple monolithic IC cross section. B 30:1 enlargement of 1 circuit Figure 5.26.— General characteristic curve for SCR. A Silicon wafer slice Figure 5.28.— Top view of an actual IC 115 30 kO Vjn — WW R NOR Amplifier within circuit Special-purpose circuit Figure 5.29.— Examples of symbols employed for IC's. AND NAND Logic circuit symbols equipment or they can be self-contained and portable. It is not unusual for every piece of power equipment in or about the mine to have some form of enclosed instrumentation. The devices can range from basic meter movements to transducers connected to on-line computers that monitor the status of the entire power-system complex. The word "meter" is often used as a suffix or part of a compound word that describes the function of the instru- ment. Of all the instruments designed to measure electri- cal quantities, the voltmeter and ammeter are the most basic. Voltmeters measure the potential difference or volt- age between two points and must present a very high impedance to the circuit so as not to interfere with normal circuit operation. Ammeters measure current flow and must have a near-zero impedance. The dc voltmeters and ammeters sense average quantities, while their ac coun- terparts usually provide rms voltage and current values. Instrument current inputs are normally at 5 A, with potential inputs at 120 V. The following section will explore the various instru- ments available to the mining industry, commencing with a description of the basic instrument or meter types and then showing how the devices are employed to monitor system quantities. BASIC METER MOVEMENTS A meter movement is an electromechanical device that provides the mechanical motion to an indicator in response to an applied electrical signal. Regardless of the type of meter movement, opposing magnetic fields are employed to activate the indicator or pointer. These move- ments can be classified as electrostatic, dynamometer, moving iron vane, and permanent-magnet moving coil. An electrostatic movement is the only type that measures voltage directly as opposed to a voltage-produced current. This meter is basically a variable capacitor with a restoring resistor connected between a fixed and a movable plate or vane. When a difference in potential exists between the plates, the opposing charges produce a mutual attraction and the movable vane will move toward the fixed vane with the deflection proportional to the applied voltage. Upon removal or change in potential, the resistor discharges the capacitance. Thus any current through the movement is merely incidental to the opera- tion. Electrostatic instruments can measure either ac or dc potentials; they have true rms response to ac regardless of waveform shape. Full-scale readings (maximum meter deflection) range from 100 V to 10 kV depending on the movement, with a measurement precision of 0.5% to 2%. A dynamometer movement consist of two coils, one fixed and the other movable. The movable coil rotates in the magnetic field produced by current through the sta- tionary coil. If the current being measured flows through both coils, (that is, they are in series), the resulting torque is proportional to the current, and the displacement is proportional to the square of current. Thus the pointer deflection indicates the rms value of current. The move- ment can be designed to measure dc or ac very precisely to within 0.1%. However, the dynamometer is not commonly employed as an ammeter. Its prime application is as a wattmeter, which, will be described shortly. Moving-iron-vane movements are similar to the dyna- mometer, except the moving coil is replaced by a soft iron vane with no permanent magnetization. Here, current through the fixed coil produces a magnetic field that induces magnetism in the soft iron vane. The magnetic fields oppose each other, producing torque that deflects the vane with a force proportional to the square of the current. The instru- ment can therefore measure dc or the rms value of ac, but with less precision (1% to 2%) than the dynamometer. The last basic type of meter movement is the permanent-magnet moving-coil or d'Arsonval meter, which is a dc ammeter. The moving element is a coil of fine wire suspended so that it is free to rotate in the field of a permanent magnet. Sketches of typical movements are provided in figure 5.30. When dc flows in the coil, a torque is produced that tends to rotate the coil. The rotation is opposed by some form of spring restraint, usually a helical spring, so that coil motion and thus pointer position is proportional to the coil current. If the dc through the coil is varying so fast that the pointer cannot follow the fluctuations, the pointer will assume a position relative to the average torque, and therefore indicate the average value of current. However, if the current is a sinusoid, the average of moving-coil torque is zero, and the pointer will not be deflected. Nevertheless, d'Arsonval movements can obtain a precision of 0.1%. For measuring current, both dynamometer and moving-iron-vane movements are often restricted to fre- quencies less than 200 Hz. Yet both these yield true rms readings within their frequency range. Electrostatic in- struments can be extremely precise for observing voltage, but they are often very delicate and are applicable only for laboratory use. Even though d'Arsonval movements mea- sure only dc, they are the most common type in use for both direct dc measurements and ac measurements using rectification. 116 Spring Pointer Coil Pivot External magnet Moving-coil construction Figure 5.30.— Permanent-magnet moving-coil movements. Internal magnet Meter-Movement Applications When a d'Arsonval meter is used as an ammeter, it is inserted in series with the circuit being measured. The current range for this direct application is obviously restricted by the maximum scale reading or maximum current of the movement. D'Arsonval meters can have full-scale limits from 1.0 /zA to 50 mA, although the basic movement is considered to be 1.0 mA, which allows measurement from zero to 1.0 mA. For higher current requirements, the meter is shunted with a low resistance as shown in figure 5.31. Such shunts can be tapped to provide several current ranges, or several shunts might be available, each selected by a switch to provide a specific current range. Commercially available ammeters of this type offer up to a 50-A full-scale reading. To measure dc voltages, a d'Arsonval movement is simply placed in series with a selected high resistance, and the combination is connected between the two points where a voltage measurement is desired (fig. 5.32). Be- cause meter deflection is still proportional to current, the meter scale can be calibrated to read the voltage required to produce a specific current. The sensitivity of such voltmeters is stated in ohms per volt. For instance, if a meter has a range of to 200 /zA and if the movement is to be used to measure to 200 V, the total meter resistance must be R = 200 V 200 M A = 1.0 Mfi. As moving-coil resistance, R m , is generally on the order of 50 to 100 12, it can be neglected in this case. Sensitivity of the combination is therefore 1,000,000 fl 200 V = 5,000 Q/V. A higher value of sensitivity for a specific meter implies higher quality. Presently, the upper limit for the commer- cially available d'Arsonval voltmeter is 50 k(2/V The standard d'Arsonval movement of to 1 mA has a coil resistance of 100 Q; hence, it can be employed to read to 100 mV directly. External shunts are utilized for a desired maximum current when the current is higher than measurable by Permanent magnet RM = 100 n, 0-100 mV, 0-1 mA Load line /^\ 1mA Shunt Figure 5.31.— Shunting d'Arsonval meter for high-current tests. • Permanent magnet RM=100il 0-200 uA Figure 5.32.— D'Arsonval meter used to measure dc poten- tials. normal instruments with internal shunts. Figure 5.33 provides a couple of typical constructions where terminals are available for circuit as well as meter connections. These are simply standard resistance units, designed to be used with either 50-mV (0- to 50-mA) or 100-mV (0- to 100-mA) movements, in which a current through the shunts is indicated by a specific voltage drop across the shunt. For example, if a shunt is designated 100 mV, 600 A, a reading of 50 mV across the shunt signifies that 300 A is flowing in the circuit. Any time that metering or instrumentation is part of dc mine power equipment, it can almost be assumed that external shunts are involved. 117 lb this point, only the measurement of circuit opera- tion has been considered. A d'Arsonval meter can also be used to measure resistance by the addition of a dc source in the dc voltmeter circuit. Consider the circuit shown in figure 5.34, which has a dc movement in series with a dc source (usually a battery) and one or more resistors, one of which is usually variable to be used for calibration. The unknown resistance to be measured completes the loop. Meter deflection is still proportional to dc through the loop and is therefore a function of the unknown resistance. Using known resistances, the meter scale can be cali- brated to read resistance directly, and different fixed resistors or multipliers can be used to extend the single scale. The combination is easily calibrated before each use by adjusting the pointer to zero using the variable resis- tance. The resistance desired could be a simple component or a complex circuit, but the ohmmeter should never be used on an energized circuit because of the internal source. Combining the d'Arsonval movement with a half- wave or full-wave rectifier allows the reading of ac values in terms of dc through the coil. The full-wave or rectifier- ammeter circuit shown in figure 5.35 is the most common. Here, current through the movement is I d , and thus, meter deflection is proportional to the average of I d . This reading is the half-cycle average if the ac is symmetrical (that is, the dc scale of the meter will read the half-cycle average sinusoidal current). As the rms value of current is usually desired, the scale is calibrated in rms by multiplying the average current by 1.11. This is the rms value for a sinusoidal waveform only; for any other waveshape, rely- ing on the rectifier circuit can produce large errors. Moving-iron and dynamometer movements record rms current automatically, and many permanent meters built into power equipment to measure ac voltage and current are moving-iron types. However, the d'Arsonval meters are often preferred because of their greater sensitivity. For ac measurements of voltage or high current, the concepts of high series resistance and low parallel resistance also can be applied to the rectifier, moving-iron, and dynamometer movements, but such practices are not common except in small portable test equipment. It can be seen in the foregoing that the d'Arsonval meter is used to measure ac or dc voltage or current as well as resistance. An instrument incorporating all these func- tions is called a multimeter. The selection of a specific parallel or series resistance combination provides the needed measurement function and parameter range. Line terminal Instrument terminals Line terminals ^Instrument terminals Copper blocks Line terminal Figure 5.33.— External shunts used for high-current measurements. Unknown resistance Figure 5.34.— Simple ohmmeter circuit. i„A A ( Q A A r. VWYY Figure 5.35.— Rectifier ammeter. Wattmeters As mentioned earlier, the main application for dyna- mometer movements is in wattmeters. Figure 5.36 illus- trates the wattmeter connection. Typically, the fixed coil carries circuit current while the moving coil is connected in series with a high resistance and is attached across the terminals of the circuit (the moving coil can itself be of high resistance). Circuit current flows through the fixed (or current) coil, and the current through the moving (or potential) coil is proportional to circuit voltage. Therefore, the movement torque is proportional to the product of instantaneous voltage and current, with the indication relative to the produce average or average power. The dynamometer connected as such will measure correctly the average power of a dc or ac circuit of any waveform, even when a power factor is involved. Current coi Current coil r o Loads 6 Figure 5.36.— Dynamometer connected as wattmeter. 118 Varmeters In addition to being used for measuring watts, the dynamometer movement has wide application in measur- ing reactive power or vars. This is done in single-phase instruments by shifting the phase of the voltage coil by 90°. The voltage coil flux is then in phase with the flux produced by the reactive-current component in the current coil. Varmeters are installed in the same manner as wattmeters are. Power-Factor Meters A power-factor meter shows the power factor continu- ously and indicates whether the current is leading or lagging the voltage. The movement resembles a single- phase wattmeter but has no control spring and has two moving potential coils mounted on the same shaft 90° apart. One potential coil (B of figure 5.37) is in series with a noninductive resistor so that it produces torque propor- tional to the line voltage and in phase with the real component of line current. The other coil (coil A) is in series with a higher quality inductance, so its torque is proportional to the line-current reactive component. The fixed coil (coil C) is again the current coil. With unity power factor, the average torque between coils A and C is zero since the currents are 90° apart, but the currents through coils B_an_d C are in phase, so the torque p roduced aligns their axes, and the pointer indicates unity power factor (1.0 pf). For leading or lagging power factors, the net torque created by currents in coils A, B, and C will swing the moving coils to the right or left, aligning the pointer in a position relative to the power factor. Meter scales are therefore calibrated so that the center position is unity power factor, and to the left and right of center are lagging and leading power factors from unity to zero. This section has presented some direct applications for basic meter movements. Some concepts shown here apply to all electrical parameter measurements, but for ac power systems, additional components are normally employed. POWER-SYSTEM INSTRUMENTATION In chapter 3, the subject of current transformers (CT's) and potential transformers (PT's) was introduced. These devices actually fall under the general category of instru- ment transformers and serve two main functions: • To isolate instruments, relays, and meters from line voltage, and • To transform line currents and voltages into values suitable for measurement by standard instruments. Thus, the normal ratings of instrument transformer sec- ondaries are 5.0 A for CT's and 120 V for PT's. This measurement implies not only metering or actual visual readings but also sensing for such purposes as protective relaying. The following material will cover specifics of CT's and PT's as they apply to instrumentation of mine power systems. Chapter 9 will discuss the application of these transformers to protective relaying. Instrument Transformers Instrument transformers are connected in the power system in a manner related to the function they monitor. The primary winding of a CT is placed in series with the Load Figure 5.37.— Power-factor movement. 9- Source H 1 H 2 Load o £^ i o O- Output Current input % CT 0-5.0 A Current transducer O-l mA Output 0-1 m A meters Figure 5.40.— Use of transducers with standard d'Arsonval movements. 120 a higher voltage than normal to appear on the secondar- ies. Without grounding, the transformer insulation could fail. The transformer case should also be grounded for the same safety reason. Three-Phase Connections When the measurement of average power in a three- phase system is required, it seems obvious to place one dynamometer wattmeter in each phase and add the re- sults together. This is shown in figures 5.41A and 5.41S for a four-wire wye load and a three- wire wye or delta load. The sum of the meter readings is total power for either connection, for any waveform, and whether the system is balanced or not. The common connection of the three wattmeter potential coils may be placed at any potential without affecting the total power readings. If the potential is that of one phase conductor (see figure 5.42), one wattmeter becomes inoperative and thus may be omitted. The result is the two-wattmeter method of three-phase power measurements. Commercially available transduc- ers can be used instead of the two wattmeters. The transducer inputs are two line-to-line voltages and two line currents, and the single output, which is proportional to total power as before, can be used with a standard d'Arsonval movement. A circuit arrangement for this method is shown in figure 5.43. Under balanced conditions, the readings from the two-wattmeter method can be used not only for total power but also to determine the power-factor angle. It can be shown that Current Wattmeters connection _ / Voltage coil — j=r tM Figure 5.41.— Three-phase wattmeter connections. Wattmeter Clirrpnt rnil Potential coil Source » • Load Potential coil Wattmeter Current coil Figure 5.42.— Two-wattmeter method. tan = V3 P2-P1 P 2 + Pi ' (5.14) where P 1; P 2 = two power readings, corresponding to ar- rangement in figure 5.42, and 6 = load power-factor angle. If P x represents a measurement of phase a current, equa- tion 5.14 provides the correct sign for the power-factor angle, thereby specifying whether the load is capacitive or inductive. At times, phase sequence is hard to distinguish in practice, but the equation yields the angle magnitude and this is often sufficient information since the reactive characteristics of the load are usually known. If the system is balanced or can be approximated as such, the circuit shown in figure 5.44 can be employed to measure the line-to-line voltage, line current, power factor, and total average power. The two-wattmeter approach calls for two PT's and two CT's. One PT supplies the voltmeter and one CT provides information to the amme- ter, while the remaining PT and CT supply the power- factor meter so that the transformer burdens are balanced. It is often useful to observe each line current or line-to-line voltage for major power equipment. Figure 5.45A provides an economical method for the line currents in which only two CT's are needed. If one CT secondary is measured, the current will correspond to the CT phase (that is, phase a or phase c), but if both CT secondaries are in parallel, the current reading is for the phase without the CT (that is, phase b). This metering is theoretically correct only for balanced voltages, but on most systems the voltage is close enough to balance that the two-CT ap- proach gives acceptable precision. If greater accuracy is Source Load To standard 0-1 mA meter Figure 5.43.— Three-phase power measurement with transducer. needed, three CT's should be used as shown in figure 5.45S. It is possible to connect the CT secondaries in delta or wye, but the burden impedances should always be wye connected. To observe all three line-to-line voltages, three potential transformers can be used as in figure 5.46A. The open-delta arrangement shown in figure 5.46S is not as accurate but gives satisfactory precision and uses only two PT's. For current or voltage with two or three instrument transformers, power-equipment metering is performed with a voltmeter or ammeter or both. The required phase is switch selected by connecting the transformer combina- tion to the meter. Source Yyyyyy^ ^npnpn Neutral 6 High Med Low • 6 • 6 oHigh • +4+3+2 + Low Med -o Low Med High • +4 + 3+2+1 Low< Med t ti=± u Watts + t t Power-factor meter output High Voltage i C C "»7 In 1 u '»8 2"—j J I -10 4*'- I t t < C * r— »I1 5>" C C ' "12 6"-£-£ ' Load 4 3 2 1 Fl Current 4 3 2 1 m nxt Average - power meter output Voltage Current meter meter output output Figure 5.44.— Balanced three-phase measurement of voltage, current, and average power. 121 Source Meters n^\ n •-Load *- Load Source Load ■*- Load Meters B Figure 5.45.— Line current measurements with two or three CTs. Figure 5.46.— Line-to-line voltage measurements with three or two PT's. 122 SPECIAL INSTRUMENTS Several special, if not very common, instruments are available to perform measurements on specific electrical quantities. These include but are not limited to watthour meters, demand meters, bridges, megohmmeters, and phase-sequence indicators. Each of these is described in the following paragraphs. Watthour Meters The watthour meter is a common power instrument, used in nearly every building to measure consumed elec- trical energy. The typical watthour meter consists of a small induction motor with an aluminum disk that is rotated by a torque proportional to voltage times current at every instant. The principle of operation is similar to that of the dynamometer wattmeter, except the disk is allowed to turn continually with a speed proportional to average power. The number of turns is counted by a train of clocklike gears. The counter thus indicates the product of power and time, or energy, which is measured in kilowatthours. A simplified sketch of the induction mech- anism is shown in figure 5.47. Demand Meters Demand meters are usually of two types (although there are others): integrated demand or lagged demand. The readings may be indicating or recording. Integrated- demand meters consist of an integrating meter element, such as the watthour meter just described, that totals the energy used over the demand interval and drives a maxi- mum indicating device, which can be a passive pointer, display, or chart. The meter can be reset manually, or a timing device can be used to return the drive to zero at the end of the recording period, thus leaving an indication of maximum demand. Lagged-demand meters provide a maximum demand indication that can be subjected to a characteristic time lag by either mechanical or thermal means, but usually the exponential heating curve of electrical equipment is followed. The demand interval is then defined as the time required to indicate 90% of the maximum value of a suddenly applied steady load; thus, maximum demand can be observed. Demand meters, whatever the type, can provide input to the power-system studies. Bridges Bridge circuits yield the most precise measurements of impedance, be it resistance, capacitance, or inductance, for two reasons: the measurements rely on null methods, and comparisons are made directly with standardized impedances that are precisely known. The term null method means that a zero reading or null indicates the correct value. The Wheatstone bridge is the most widely used of these circuits. Shown in figure 5.48, the bridge is dedi- cated to measuring resistance, capacitance, or inductance depending on its internal components. When the Wheatstone bridge is intended to measure resistance (figure 5.48A), the circuit consists of two fixed precision resistances, R 1 and R 3 , which are known as the ratio arm; a variable precision resistance, R 2 ; and the , Transformers Dial Line Permanent magnets Aluminum disk Simple schematic Line Load Eddy currents produced Voltage coil by voltage coil (highly reactive) Disk plan view Figure 5.47.— Simplified sketch of watthour meter induction mechanism. Galvanometer Unknown A Wheatstone bridge for resistance Audible device or meter Unknown B Impedance measurements with a Wheatstone bridge Figure 5.48.— Wheatstone bridge circuits. Unknown 123 unknown, R x . A dc source supplies current to the arrange- ment, and a galvanometer, G, is located at the center of the bridge across points b and d. The galvanometer is simply a very sensitive ammeter with a center-scale zero-reading pointer and the ability to read very small currents in either direction. R 2 is adjusted to provide a null reading on the galvanometer, which means the potential between b and d must be zero. With this balanced condition, the unknown resistance can be calculated by R x = ^~ Ra (5.15) In commercially available bridges, R x , R 2 , and R 3 are all variable and the value of each is readily determined by calibrated dials. Thus, the bridge can measure resistances precisely over a broad range. To measure impedance, R 3 of the resistance bridge is replaced by Z 3 , and the unknown is now Z x . An ac source is used, together with some means of measuring the potential between points b and d. This could be a sensitive ac ammeter or an audible device such as a set of head- phones. R x and R 2 are then adjusted to provide a null, and the balanced condition means that z, = z s | Megohmmeters The preceding resistance-measuring devices can be ineffective when resistance is in the many millions of ohms. An important factor here is the resistance of insu- lation, such as that around conductors (fig. 5.50). One problem in these and other high-resistance measurements is to provide sufficient potential so the resulting current can be detected by an indicating device that provides resistance readings. The instrument designed to perform these tests is called a megohmmeter (fig. 5.51), where the unknown resistance is R x , and R x and R 2 serve as current- limiting resistors to protect the meter from damage. Galvanometer (5.16) Obviously, the values of Z x and Z 3 depend upon the frequency of the ac source. The most typical value used is 1,000 Hz. If very low resistances in the order of 10 jtQ to 1.0 mfl must be measured, the Kelvin double bridge shown in figure 5.49 can be used. The circuit consists of ratio arms R A , R B and R a , R^; a connecting link or conductor, R f ; a known resistance, R s ; the unknown, R x ; an adjustable dc source; and a null indicator. The indicator could again be a galvanometer. The resistances r x , r 2 , r 3 , and r 4 are those of the connecting leads between the four-terminal bridge and the resistances to be compared (R x and R 8 ). These lead resistances should be in the same ratio as the bridge arms to which they are connected; otherwise, the ratio unbal- ance will cause incorrect measurements. A small adjust- able resistor can be used to balance the lead resistances. The balance equation is thus Figure 5.49.— Kelvin double bridge. R_x R„ R, R T R s \R a + Rg + R f / VRg R«/ Instrument test leads Conductor insulation Indicating scale shows resistance Megohmmeter Figure 5.50.— Megohmmeter testing insulation resistance. When R x and R s are so small that R f is comparable, the term in equation 5.17 involving R f However, if can be significant. R, Rfl then the R e term becomes zero. The source is adjustable so that current through R x , R e , and R 8 (the series resistance of which is small in comparison to the bridge) is large enough to allow a measurable milling current through the indicating device, G. An application for the Kelvin double bridge is in the measurement of cable and conductor resistances. Generator- C_Hand crank Figure 5.51.— Internal components of megohmmeter. 124 The most evident difference between the megohmme- ter and the preceding instruments is the hand-driven generator, which supplies the needed dc potential for measurement. The generator applies from 500 to 2,500 V depending on the instrument and is tied to the resistance range desired (the higher the measured resistance, the higher the required voltage). Typically, a friction clutch is employed to restrict the generator to rated output voltage. In some megohmmeters, the potential is from batteries via an electronic power supply located within the instrument. As shown in figure 5.51, the meter has two coils mounted over a gapped core. The movement is similar to the d'Arsonval, but there are no restraining springs, so the indicator is free to move when there is no output from the generator. If the instrument terminals are open (that is, R x is infinite) when the generator is operated, current will flow through R 2 and coil A x , and the torque produced will force the pointer counterclockwise to the infinite scale reading. When the terminals are shorted (R x is zero), the torque produced by coil B is greater than that from coil A and this moves the pointer to a zero reading. For measur- ing an unknown resistance, the pointer location is depen- dent upon the opposing torque from the two coils, and the position is a function of R x . Another prime application for megohmmeters is the measurement of ground-bed resistances. These specialized testing procedures are covered in chapter 7. Phase-Sequence Indicators In order to prevent damage or incorrect operation, all conductors in a three-phase distribution system must be properly connected so they will provide the same phase sequence to all equipment. Correct interconnections can at times be difficult to accomplish in mine power systems, especially with feeder and trailing cables. At present there is no standard color coding for phase conductors. The phase-sequence indicator illustrated in figure 5.52 can be used to determine the phase relationship of energized three-phase conductors. It falls in the simplest class of testing devices: indicating instruments; other examples include a light bulb with leads to test for the presence of potential, or a battery in series with a light bulb with leads to check continuity by completing the series circuit. The phase-sequence indicator consists of two light bulbs and a capacitor connected in wye, and the lamps are labeled in the two possible phase combinations. Because of this arrangement, one lamp will burn brighter than the other depending on the connections to the power system. The foregoing has provided information on several devices that are helpful in measuring mine electrical systems. Other instruments that are equally useful for specific applications include the split-core ac ammeter, a handheld ac ammeter that has its own CT; the synchro- scope, which measures proper phase connections and the correct speed of parallel ac generators; and a frequency meter, which indicates the frequency of an electrical supply in hertz. Often there is also a need to obtain a continuous record of an electrical parameter, and the next section discusses the popular recording devices. RECORDING INSTRUMENTS Many of the direct-reading indicating instruments just presented are also available as recording instruments. Some of these are very similar to their indicating counter- parts in that they can use the same electrical movements; they differ because the pointer is also used to provide a graphic record on a chart. These are termed chart record- ers; one popular class is strip-chart recorders, so named because the electrical parameter is recorded on a strip of paper. The similarity between the movement of the strip- chart recorder and the indicating instruments is illus- trated in figure 5.53. The strip-chart recorder movement is actually a d'Arsonval type. The pen can trace on paper in several ways. • Inking. The pen is a capillary tube through which ink flows from a well to the chart. This is perhaps the most used system. • Inkless. The tip of the pen is a stylus that impacts the paper like a typewriter key with a regular force supplied by a cam, leaving a series of dots. • Thermal. The pen tip contains a heating element that leaves a trace by heating specially treated paper. Feeder line Figure 5.52.— Phase-sequence indicator. Input circuitry to condition input voltage and establish sensitivity of the recorder Comparator, compares input with reference, outputs a voltage in proportion to the needed position of the servomotor Amplifier, amplifies Circuitry to condition pen- sensor output and establish reference, including to zero setting Paper strip chart, driven at varbus constant speeds y comparator output to drive servomotor Servomotor, drives mechanical pen system Indicator and inking pen Figure 5.53.— Strip-chart recorder. 125 The simplest unit provides a curved recording as the pen swings in an arc, but articulated pen arms are also available that produce linear or rectilinear traces. The paper chart moves past the pen at a predetermined speed driven by an electric motor or a mechanical-spring clock- work mechanism. This recorder provides a continuous record of the average or rms value of the electrical param- eter of interest, which is advantageous in obtaining records of equipment operation, for example, the electrical performance of a mining machine. A variation of these recorders uses a round chart, driven like a disk on a record player but at very slow speed. These charts can be built into major equipment to provide permanent records. Sometimes recordings of the actual electrical wave- forms are needed to study power systems. This calls for an instrument that can resolve instantaneous values of elec- trical parameters. Electromechanical instruments that have this resolution are called oscillographs, and the movement in most of these is a sensitive galvanometer of low mass. Two types of writing systems are normally available: • Direct writing. This is similar to either the inking or thermal strip-chart recorder types. The pen has high inertia, and instrument response is about 0.5 to 100 Hz (some to dc). • Optical. Instead of a pen, the movement drives a low-mass mirror that deflects a light beam that exposes a light-sensitive paper. Developing is required to obtain the record, but the system can have resolution to 10,000 Hz. For many applications, magnetic tape recorders and oscil- loscopes, both electronic instruments, find favor over oscil- lographs. However, oscillographs still have some practical use, especially where an extended-time hard copy is needed immediately. An example would be in measuring neutral currents existing on three-phase equipment, which can have dc as well as ac components. ELECTRONIC INSTRUMENTS The employment of complex and sophisticated control equipment in the mining industry is continuing to in- crease. Instances include solid-state motor starters, elec- tronic protective relaying, computer logic circuits on min- ing machinery, and so on. These types of systems require precise voltage, current, and waveform measurements that are not possible with the preceding instruments. Certain phenomena existing on power systems, such as transients, require precise measurements with frequency response into the megahertz. Electronic measuring equip- ment answers this need. This section will introduce only the more popular instruments. Electronic Meters These instruments use many of the basic circuits that have been described for multimeters; that is, series resis- tances for voltage (fig. 5.54A), voltage-drop for resistance (fig. 5.545), and shunts for current. The prime difference is that a scaled-down dc voltage, which is proportional to the actual circuit voltage, current, or resistance, is amplified by electronics. When the parameter is sinusoid, the ac is rectified before amplification. The amplified signal then drives the indicating device. In the past, vacuum tubes Range switch (attenuator), allows selection of sensitivity Filter, removes any ac superimposed on dc Amplifier, amplifies dc input drive meter Multiplier resistors Meter Feedback, stabilizes amplifier characteristics Full-scale sensitivity of voltmeter A dc voltmeter Range, value per division on meter scale -\ Resistance being measured ^1,000 n Multiplier resistors IOO n 10fl$. B %V^# c 0° 60° 120° 180° 240° 300° 360° Figure 6.21.— Rotating magnetic field in elementary three- phase, two-pole induction motor. where s = motor slip, expressed as a per-unit decimal or a percent, and n,. = actual motor speed, r/min. Losses will occur in actual motors because of electrical and mechanical inefficiencies. Those prominent in induc- tion motors are • Rotor winding loss, related to I 2 R; • Stator winding loss, also an I 2 R loss; • Stator core loss, caused by eddy currents and hys- teresis in the core iron; and • Friction and windage (rotational or mechanical) losses. These are almost pure active powers; therefore they are often expressed in watts. The losses in both windings of induction motors vary as the square of line current, core loss is nearly constant, and unless motor speed varies considerably rotational losses are nearly constant (11). Knowledge of machine losses allows the determination of motor heating and efficiency. The efficiency of motor operation is a measure of the ability to convert input power to mechanical power: Efficiency = output input - total losses input input (6.6) which may be expressed as a per-unit decimal or a percent. Slip is also related to motor efficiency, being numerically equal to the ratio of winding loss in the rotor to the total rotor power input: Rotor conductor Figure 6.22.— Induced rotor potential by rotating flux. s = rotor winding loss rotor power input or s = rotor winding loss motor power input - stator losses ' (6.7a) (6.76) Stator losses in equation 6.76 include friction and wind- age. Equations 6.4 through 6.7 can be employed to calcu- late the synchronous and actual motor speeds and also the possible power and torque output, realizing that (8) power output (watts) = 746 (horsepower output), (6.8a) 138 and in which hp = 5250 T = KI^Rr (6.86) where hp = horsepower output of motor, n,. = actual motor speed, r/min, T = motor torque, ft-lb, K = a torque constant, ft-lb/V, I,. = rotor current, A, and R,. = rotor resistance, Q. ^ ^ ^ C B A Figure 6.23.— Lapped windings of three-phase motor stator. Motor Construction The elementary salient-pole motor of figure 6.19 is undesirable from the standpoint of the ineffective use of material and space, as well as its overall inefficiency. The main disadvantage is coupled to the distinguishable stator poles. To overcome this problem, actual induction motors have lapped stator coils, as shown in figure 6.23, where several coils make up a stator winding that can be either delta or wye connected. The flux directions of each coil are illustrated as c , B , 4> A , and each coil contributes to the rotating flux development of the entire stator. The coils and windings are arranged to have the same effect as salient poles, but the poles are not physically distinguish- able. An induction motor is assigned a specific pole num- ber if at any given instant the stator windings set up the same number of magnetic pole fields. The rotor core and squirrel-cage conductors are usu- ally not insulated from each other, because the induced current is effectively contained within the conductors owing to their significantly lower resistance. The rotor core is pulled magnetically toward the stator core across the air gap. If the force is uneven when the rotor turns, the result is vibration. This is detrimental in several ways as it can lead to structural insulation failures, premature bearing failures, and misalignments with the motor load. Vibration does not occur if the magnetic effect about the rotor periphery is equal. An additional method for pre- venting vibration is to place rotor conductors in slots skewed to the stator slots so that a rotor slot passes gradually under a stator slot rather than abruptly. This practice also prevents "dead spots," or positions of near- zero or minimum magnetic influence. Another method of eliminating dead spots is to construct the motor so that the number of rotor slots plus the stator slots sums to a prime number. Motor Behavior Figure 6.24 is a graph of the speed, efficiency, power factor, power input, and current load of a typical three- phase induction motor found in mining applications. Fig- ure 6.25 shows a representative torque-speed characteris- tic for a similar machine. These curves can be used to describe the electrical and mechanical operation of induc- tion motors under loading. From the typical torque-speed curve, the torque at locked rotor is approximately 150% of rated. The level increases steadily with rotor acceleration to the maximum or breakdown torque. With applied power input, the rotor continues to accelerate until the slip reduction reduces the 5 (£* UJ 5 < t- LJ => £E 15 O H o 2 180 i- 150 - uj 120 - ft 90 - 60 30 1 1 ■ t— L.^ rpm — i— I — i — 1—1 — eff j 90 - ^^^pT 80 70 / X ° I / ° 60 50 1 / *3 J LL. 40 30 A / s 20 10 1 I I I i i i i i 1,200 1,150 1.100 10 20 30 40 50 60 70 80 90 I00 IIO OUTPUT, hp Figure 6.24.— Characteristic curves of three-phase induc- tion motor. f Rated torque 50 1 00 SYNCHRONOUS SPEED. % I I 1 00 50 SLIP. % Figure 6.25.— Typical torque-speed characteristic for general-purpose induction motor. rotor current to a point where torque is equal to the load torque. Consider the motor running with no load. As the motor is loaded, slip increases, causing an increase of induction in the rotor. Hence, rotor current rises, resulting in a stronger rotor magnetic field and motor torque. Torque continues to increase with the increased shaft load 139 until breakdown torque is reached. Any further load results in a slip value that decreases torque. If the high load is sustained, the rotor will stop. Because the induction motor operates basically as a transformer, its electrical characteristics, as seen by the power source, will be a reflection of those occurring in the stator winding. Figure 6.26 shows phasor diagrams for rotor current and voltage during three operation points; these are referenced to the flux-density phasor of the stator, B stator (11). The rotor bars are embedded in the steel core so they have a high reactance (3). At locked-rotor conditions (rotor stationary), the stator magnetic field rotates past the motor at synchronous speed, and the induced voltage in the rotor conductors has the same frequency as the stator (or line frequency). The result is a high ratio of rotor reactance to resistance, and stator current lags stator voltage by a large amount (fig. 6.26A). During rotor acceleration, slip decreases, which also lowers the fre- quency of rotor current and voltage according to the following relationship (8): The torque developed by a three-phase induction mo- tor varies as the square of the stator supply voltage, or f r = sf, (6.9) where f r = frequency of sinusoidal voltage and current induced in rotor bars, Hz, s = slip, expressed as a decimal, and f = frequency of voltage and current in stator, Hz. Thus, inductive reactance drops, increasing the power factor (fig. 6.265). Theoretically, if the motor could obtain synchronous speed, the rotor power factor would reach unity (fig. 6.26Q. However, as this cannot happen in actual squirrel-cage motors, the maximum power factor is seldom greater than 0.85 (fig. 6.24) and never greater than 0.95. Because the output torque increases with slip, motor speed decreases slightly as the load increases from no load to full load. Yet efficiency and power factor drop rapidly on low load conditions. Hence, an induction motor should not be operated at much below rated load for any length of time. It is apparent from figure 6.24 that efficiency dimin- ishes when motor load increases above a given value. Consequently, an induction motor should not be over- loaded for any extended period. Power-factor and efficiency curves normally follow roughly the same path; thus, power factor can be considered as an estimate of motor operating efficiency. Torque oc V^ (6.10) Therefore, a 10% reduction from rated stator voltage will cause a 19% reduction in available torque output. Insulation Insulation in motors normally has five forms: strand, turn, lead, crossover, and ground (15). Since the rotor conductors are uninsulated, the insulation of the stator winding conductors is the critical concern. The primary insulating system is that between the windings and the stator core or ground, and the secondary insulation is in strands, turns, leads, and crossovers. Copper magnet wire, and to a much lesser extent aluminum magnet wire, is used to construct the stator winding or coils. Strand insulation is most frequently a resinous coating on the wire. Turn insulation is applied after strands are wound into coils (or the actual windings), and this may be a resinous coating, resinous-film taping, paper taping, or a fibrous wrapping. These types of turn insulation are utilized for applications of 6,600 V and less; for higher voltages, additional layers of mica or varnished cloth tape can be used. Crossover insulation is employed to protect wires that cross each other. The crossovers are often the weakest point in winding construction; thus, they require additional protection. Lead insulation is simply insulation about the conductors leading to the windings. Lastly, ground or ground-wall insulation is the major insulation system of the motor and isolates the windings from the core. This insulation is always sub- jected to the highest potential difference and requires the most attention. Design Characteristics Figure 6.27 illustrates the standard NEMA torque- speed characteristics for squirrel-cage induction motors. The shapes of these curves depend primarily on the ratio of rotor conductor resistance to reactance. For instance, to obtain a greater locked-rotor torque, as well as a greater slip over the unable load range, rotor conductor resistance may be increased by decreasing the conductor cross- sectional area, or inductive reactance may be decreased by placing the bars closer to the rotor surface. On the other i cos e *-v rotor 1 rotor B i cos e rotor rotor I cos e ^r v,0, °' . i. s to tor A *rotor Figure 6.26.— Phasor diagrams of rotor and stator flux density for induction motor. 140 hand, an increase of conductor resistance will decrease overall motor torque and the stator current drawn during locked-rotor conditions. Single-cage rotors, as previously described, are the most rugged and the most used. Double-cage rotors use two conductors, one over the other, per rotor slot (fig. 6.28A) and provide higher starting torques with higher load efficiency and lower running slip than the single cages (14). Here, the higher conductor would have high resistance and low reactance, while the lower set would have low resistance and high reactance. Double-bar rotor conductors are often susceptible to damage on loads with long accelerating times, lb overcome this problem, deep- bar rotors (fig. 6.28S) can be used. These have a thermal advantage in that the full conductor area is available for heat dissipation, but the design still approximates the performance of the double bar. Regardless of the design, the torque-speed curves are matched to the squirrel-cage rotor construction, which is fixed for a specific motor. In addition to rotor design changes, the actual values of breakdown and locked-rotor torque vary with the horse- 300 £ 250 - 2 200 DC o H Q 150 < O i U. 1 --^^^ ^--Design D - ^-Design C N. / \ . ^" >, "**** ,, "-w x\ y^ \ \ - Design A — ^±.^ZS^\ \\ \ Design B ^^"^ ^\ \\| - ^*-— " -""^^ Design F \\\\| - ' 1 100 - 50 - 50 100 SYNCHRONOUS SPEED. % Figure 6.27.— Typical torque-speed characteristics for NEMA-design three-phase squirrel-cage motors. power, frequency, and speed ratings of the motor. Although the operating characteristics are a function of rotor imped- ance, the horsepower rating is mostly dependent upon the power (or kilovoltampere) capacity of the stator and rotor windings. As rotor losses are constrained to the rotor cage, rotor thermal capacity is limited. Therefore, motor designs that create large rotor currents, such as high-torque high-slip, may have intermittent time ratings or a limited number of allowed successive starts. Unless these con- straints are heeded, improper operation will burn out the rotor winding. The different rotor designs have led to a variety of speed-torque characteristics. To distinguish among the various types, NEMA uses a code letter system that signifies specific rotor constructions (8). Design B serves as the comparison basis for the motor performance of other designs and is often called the general-purpose motor. This design has relatively high efficiency even at light loads and a reasonably high power factor at full load. It has single rotor bars located rather deep in the core but with large-area slots for good heat dissipation. Starting cur- rents range from 4.5 to 5 times the rated full-load current. The design B motor has the broadest industrial applica- tion field. Design A has characteristics similar to those of design B, except that it has a higher breakdown torque. The rotor conductors are shallower, which decreases rotor reactance but increases the starting current, being five to seven times rated current. As a result, design B motors are often preferred over design A for large motor applications. As shown in figure 6.27, design A motors have the best speed regulation, as evidenced by the steep curve portion be- tween synchronous speed and breakdown torque (8). Design C motors have a double-cage rotor construc- tion that results in higher locked-rotor torque and lower breakdown torque than those of design B. Starting cur- rents are about 3.5 to 5 times rated current (8). These characteristics are well suited for conveyor belt drives and other applications that have sudden large load increases, but low or normal starting inertia. The motors are not suited for heavy high-inertia loads because the thermal dissipation is limited and high rotor current tends to concentrate in the upper bars (8). Accordingly, frequent starting of these motors can cause rotor overheating. Very high locked-rotor torque and high slip are found with design D characteristics. Design D's principal appli- cation is for high-inertia loads. The rotor is of high- resistance design with bars located close to the surface (8). bars uiah-reoctonce. ToVres-'Stonce bo< s / // '/' '/' I. ■ II ill a 1 TTT, /'I' Mi n. I \fe£' Rotor bor A Double-cage rotors B Deep-bar rotor Figure 6.28.— Other rotor-conductor designs. 141 Starting currents range from three to eight times rated load current. The motor is suited for heavy-duty starting, but again, the poor heat dissipation of the rotor design means that starting cannot be frequent. Design F has lower locked-rotor and breakdown torques than does design B. Design F motors also use a double-cage rotor with high resistance in both conductors, which reduces both starting and running current (8). The locked-rotor current is the lowest of all motor designs. Thus, design F motors are applied when starting-current limitations are severe and both starting and maximum torque requirements are low. The design, however, has poor speed regulation, low overload capacity, and usually low full-load efficiency. Induction-Motor Starting From the foregoing it can be seen that if an induction motor is started by directly connecting it to a power system, the momentary starting current can range from three to eight times the full load current. While this will not damage the motor, the high current can cause a significant disturbance on the power system, and, in some cases, activate overcurrent protection devices. However, most induction motors in mining applications are started by directly connecting them to the power system, espe- cially those within mining machines such as continuous miners. The system usually has enough impedance that protective devices can be set above the in-rush current to prevent nuisance tripping. This, however, is a major prob- lem, which is further discussed later in this chapter and in chapter 10. Full-voltage starting can usually be performed on 440- to 550-V motors up to 1,600 hp. NEMA standard magnetic starters for this range are shown in table 6.3 (3). The jogging service listed in the table refers to frequent stop-start or plugging (reversing under load) applications. As shown in figure 6.29, the across-the-line starter is simply three contacts driven by a solenoid, also called a contactor. Pressing the start button energizes the solenoid, which closes the M contacts. An auxiliary contact set (M b ) simultaneously closes and bypasses the start switch. Pressing the stop button deenergizes the solenoid. Above 1,600 hp (but sometimes lower), full-voltage starting becomes impractical even when the load con- nected to the motor can withstand the stress. Common methods for starting these large induction motors are shown in figure 6.30. In basic terminology, all these methods can be called reduced-voltage starting. In figure 6.30A, an autotransformer is used to start the motor at reduced voltage (50% to 80% of rated), thus limiting starting current and torque. When almost at full speed, contactors quickly change the motor from the autotrans- former to the full-voltage supply. Primary resistor or reactor starting (fig. 6.30S) inserts fixed or variable im- pedances in series with the motor; these are shorted out after acceleration. For the wye-delta technique (fig. 6.30O, the motor is started as a wye connection, which places about 58% of the rated delta terminal voltage across the windings, limiting line current to 58% and torque to 35%. After acceleration, motor operation is with a delta connec- tion. Part-winding starting requires that the motor have two identical stator windings (fig. 6.30D). Starting uses only one winding and limits starting current to about 65% of normal, torque to 45%. After acceleration, the second winding is switched in. There are many systems that cannot take the shock of full-voltage starting. One instance is a conveyor belt drive Table 6.3.— NEMA class A standard starters for three-phase induction motors Continuous Maximum horsepower Maximum horsepower Controller current for normal service for jogging service size rating, A 220 V 440-550 V 230 V 460-575 V 00 9 1.5 2 NAp NAp 18 3 5 1.5 2 1 27 7.5 10 3 15 2 45 15 25 10 30 3 90 30 50 20 60 4 135 50 100 30 150 5 270 100 200 75 300 6 540 200 400 150 NAp 7 810 300 600 NAp NAp 8 1,215 450 900 NAp NAp 9 2,250 800 1,600 NAp NAp NAp Not applicable. conductors 3-phase motor -1 Start I Stop_|_ 1 2^ O.L. L2 -®-WM 3-phase diagram Control circuit Figure 6.29.— Across-the-line magnetic starter. 142 3 -phase supply 1-2 L 3 l-l p A Autotransformer 3- phase supply Li L 2 l 3 o o o / / / I *<\ /« 1 \ T 2 < Tl Run Start # Primary reacter 3 -phase supply 3 -phase supply C Wye-delta Run D Part winding Figure 6.30.— Starting methods for induction motors. where the horsepower limit for full-voltage starting is perhaps as low as 50 hp. The wound-rotor motors described in the next section provide an alternative. WOUND-ROTOR INDUCTION MOTORS As mentioned earlier, the starting and running char- acteristics of an induction motor may be adjusted by varying the resistance-to-reactance (E/X) ratio of the rotor conductors. Instead of rotor bars and end rings, the wound- rotor motor has insulated windings much like the stator, with the same number of poles and windings placed in the rotor slots. The windings are usually connected in wye with the ends connected to three slip rings mounted on the rotor shaft. The brush and slip-ring circuit is completed through a wye-connected set of variable resistances, as shown in figure 6.31. Thus, the external resistance can be used to vary the speed-torque characteristics by changing the rotor R/X ratio. The stator of the motor is the same as for a squirrel-cage machine. A typical family of wound-rotor motor characteristics is illustrated in figure 6.32 (11). As external resistance is increased, the starting current is decreased and starting torque is increased. For a given shaft load, the reduction in rotor current will result in a speed decrease. Thus when starting a wound-rotor motor, a maximum resistance is inserted in the rotor circuit (R 9 curve). As the rotor accelerates, the resistance is reduced until the desired speed is obtained, or if full speed is required, the resis- tance is brought to zero (R x curve). Therefore the wound- rotor motor can be considered a variable-speed machine. Thermal considerations do place a lower speed limit on it, and for self-ventilated motors, continuous rated torque operation below 70% of rated full speed is not recom- mended (15). This lower limit may be reduced to 50% if the motor load is 40% of rated. Applications for wound-rotor motors include loads that require constant-torque, variable-speed drives or for which a sequence of slow-speed steps is needed to limit motor current during acceleration, such as for high-inertia or high-torque loads. Since they are suited to high-torque loads. Since they are suited to high-torque loads, these motors have found extensive used in the mining industry to operate crushers, grinders, ball and roller mills, con- veyor belt drives, and hoists. The automatic starting method for these motors uses definite-time acceleration where a series of fixed resistances are shorted out one at a time on a predetermined schedule (12). This step starter is shown in a simplified schematic in figure 6.33. When the starting sequence is initiated, all resistors are in series with the rotor winding; then the relay 143 3-phase winding on rotor External variable resistor Figure 6.31.— Schematic of wound-rotor induction motor showing external resistance controller. contacts 1A, 2A, and 3A are sequentially closed, resulting in four speed-torque characteristics. The last effectively shorts out the rotor winding. Since the sequence proceeds regard- less of motor speed, the method requires close coordination with motor characteristics {IS). The actual operation of the relays is discussed in chapter 9. Whether started automatically or manually, the wound-rotor motor continues to find application for the functions previously mentioned. However, for conveyor belt drives, these motors are now tending to be displaced by squirrel-cage induction motors equipped with solid- state starters (see chapter 14). Reasons for this change involve maintenance problems and a desire to eliminate the failures inherent with brushes, slip rings, and relay contacts. THREE-PHASE SYNCHRONOUS MOTORS 250 25 50 75 SYNCHRONOUS SPEED, % 100 Figure 6.32.— Torque-speed characteristics for wound-rotor motor with stepped-resistance controller. The three-phase synchronous motor has a stator and rotor and is similar to the induction motor. The stator and stator winding have the same basic construction and purpose: to receive the power to drive a load (15). However, in this motor, the rotor consists of field poles connected in series, parallel, or series-parallel combinations and termi- nated at slip rings. The field windings are excited by an external dc source, the exciter. The number of field- winding poles equals the number of magnetic poles present in the stator. A sketch of a typical large synchro- nous motor is shown in figure 6.34 (12). Rotor field excitation is often supplied from a small dc generator mounted on the same rotor shaft, as dia- grammed in figure 6.35. Alternatively, dc supply can be obtained from a three-phase full-wave bridge rectifier, as illustrated in figure 6.36, or by a separate m-g set. Pure synchronous motors are not self-starting and are generally accelerated in the same manner as inductor motors. Salient-pole rotors commonly have a squirrel-cage winding (fig. 6.34) to produce the necessary induction motor action. Low-speed cylindrical rotors closely resem- ble a wound-rotor induction motor, but with five slip rings T3A T2AT1A <^f-Os.o--« LB JCtC TS1A Atc TS2A CH "°$ -, (6.14) where n = armature speed, r/min, and $ = magnetic flux per main field pole, Wb. As the armature accelerates, the cemf rises, and the armature current drops. Yet, according to equation 6.12, motor torque decreases. A final speed is reached when the cemf is almost equal in magnitude to the supply voltage. If the motor is unloaded, the difference between the terminal voltage and the cemf will allow only enough armature current to overcome friction, winding, and core losses. Under motor loading, the armature slows down, cemf decreases, and more current enters the armature. However as shown in figure 6.49, the speed of the shunt motor remains relatively constant from no-load conditions up to 100% rated and slightly beyond. The speed can be easily adjusted by changing a resistance in series with the field winding. From equation 6.14, weakening the field flux by decreasing field current increases motor speed. Yet, for a constant field flux, torque varies linearly with armature current (that is, T oc I a ). If across-the-line starting was attempted with the shunt motor shown in figure 6.48, the cemf would proba- bly not build up fast enough to limit armature current to a safe value, and hence, damage to the commutator, brushes, and the armature winding could result. For this reason, a starting resistance is used in series with the armature (fig. 6.50A) for all dc motors except those of fractional horsepowers. The resistance is usually selected to limit armature current from 150% to 250% of rated current depending on the starting torque required. The shunt winding is always connected across full line voltage when starting so less armature current is needed to develop the rated torque. Variable starting resistor Speed- adjusting rheostat Shunt field Variable starting resistor Variable starting resistor + A Shunt motor B Series motor C Compound motor Figure 6.50.— Simplified dc motor schematics with starting resistances. 150 In mining, manual controllers are found on many dc machines. These are available in three general forms: faceplate, multiple-switch, and drum controllers. Sche- matics for these are shown in figures 6.51, 6.52, and 6.53. -Holding coil Figure 6.51.— Faceplate manual starter. Circuit breaker Overload trip ±l_ine Disconnect with cover interlock Shunt fields Both switches must be in off position to open cover w Series field Figure 6.52.— Multiple-switch starting. I 2 3 4 Shunt field Figure 6.53.— Drum-type starter. The faceplate starter is often used with small station- ary dc motors. The level is advanced (to the right) in steps, momentarily stopping at each position to allow the motor to accelerate, until the resistance is removed. A holding coil then maintains the lever in the last position. A spring is used to return the lever to the off position during a power failure or if the lever is left in an intermediate position. One method of multiple-switch starting, shown in figure 6.52, uses two double-pole, single-throw switches. The upper switch is closed first, energizing the shunt field through a 100-B resistor. This allows the main field flux to build up to some extent before the armature is connected to the line. Initial inrush current is thus reduced, which helps to prevent brush arcing and the possibility of com- mutator flashover. The lower switch is then closed, ener- gizing the armature through a second resistor, and the motor accelerates. The armature resistor remains con- nected during running. The two line switches are mechan- ically interlocked so the upper switch must always be closed first. A variation of this technique is to use relay contacts or contactors to supply main field excitation, insert the starting resistance, then bypass the resistance. Drum controllers (fig. 6.53) are frequently used on mine locomotives but are also found on some dc mining machines. A handle-controlled rotary shaft is connected to the switch segments indicated by dark lines in the figure. These segments are of various lengths so contact with the stationary contacts can be made at different intervals. When starting, the Ml and M2 contacts engage first, energizing the shunt field and inserting all resistors in series with the armature. The resistors are then removed one at a time by advancing the controller. Although not shown, an additional drum or reversing controller is usually available to reverse armature current and thus motor direction. The use of fixed resistance starting has widespread application in mining. Here the starting resistance re- mains in series with the armature for running. An in- stance would be a small dc motor, such as a pump in a remote location. The resistance gives poor speed regula- tion, but the motor can be started unattended. Dynamic Braking If a shunt motor is running under load and the armature circuit is opened, the inertia of that load will drive the machine as a dc generator. Dynamic braking simply connects a resistance across the armature to dissi- pate the available energy and decelerate the load (fig. 6.54). The braking action is most effective at high arma- ture speeds, becoming negligible at low speeds. The value of resistance, R, is selected from (12) R _ V c - I a R a (6.15) where V c - I a R a = armature voltage at start of braking, V, and I = dynamic braking current, depending upon desired braking level, A. The normal value for I is 150% of rated motor current but I may be as high as 300% for quick stopping. 151 Series Motor The armature and main field winding are connected in series and both carry load current in a series motor. The magnetic flux, $, now produced in the main field winding, is proportional to the armature current. Thus, motor torque varies as the square of armature current (T Starting winding > >4 and poles » N r^ f^ ^ Rotor •it s it r •/ < < S y / y J y Line Running winding and poles \ ' Figure 6.64.— Starting and running stator windings. Starting Centrifugal switch closed on start B Running I jj Running II y Line j windin ^ , J Starting Switch opens at 75% of speed Figure 6.65.— Centrifugal switch to remove starting winding. Centrifugal switch Capacitor Running Starting wind^ S winding Figure 6.66.— Capacitor-start motor. 158 REFERENCES 1. Allis-Chalmers (Milwaukee, WI). Motor Control-Theory and Practice. 1955. 2. Bergmann, R. W. Excavating Machinery. Ch. in Standard Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 3. Fitzgerald, A. E., C. Kingsley, Jr., and A. Kusko. Electric Machinery. McGraw-Hill, 3d Ed. 1971. 4. Hardie, R. C. Mine Hoists. Ch. in Standard Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 5. Hugus, F. R., J. A. Buss, and E. L. Parker. Mining Machine Motor Characteristics. Min. Congr. J., v. 41, May 1955. 6. Mining Machine Motor Identification. Min. Congr. J., v. 41, Mar. 1955. 7. Joy Manufacturing Co. (Franklin, PA). Direct Current Min- ing Machinery. 5th ed., 1971. 8. Kosow, I. L. Electric Machinery and Transformers. Prentice- Hall, 1972. 9. Lloyd, T. C. Electric Motors and Their Applications. Wiley- Interscience, 1969. 10. Marion Manufacturing Co. (Marion, OH). 191-M Mining Shovel. Doc. Specification 542-5, 1979. 11. Marklekos, V. E. Electric Machine Theory for Power Engineers. Harper and Row, 1980. 12. Millermaster, R. A. Harwood's Control of Electric Motors. Wiley-Interscience, 4th ed., 1980. 13. Morley, L. A. Utilization and Efficiency in Underground Coal Mine Electrical Systems. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pitts- burgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 14. Oscarson, G. L. The ABC's of Large Induction Motors. E-M Synchronizer. Electrical Machinery Manufacturing Co., Min- neapolis, MN, 1955. 15. Smeaton, R. W. (ed.). Motor Application and Maintenance Handbook. McGraw-Hill, 1969. 16. Stefanko, R. Coal Mining Technology Theory and Practice. Soc. Min. Eng. AIME, 1983. 17. Tsivitse, P. J. Mining Motors. Ch. in Motor Application and Maintenance Handbook, ed. by R. W. Smeaton. McGraw-Hill, 1969. 159 CHAPTER 7.— GROUNDING A vital part of any mine power distribution system is the connection to earth or ground, which is referred to as the mine grounding system. It consists of grounded or grounding conductors, extending from ground beds to equipment. A grounded conductor is a power conductor tied to the grounding system; a grounding conductor is separate from the power conductors and is used only to ground exposed metallic parts of the power system. A ground bed, also termed a ground mesh or grounding electrode, as well as other names, is a complex of conduc- tors placed in the earth to provide a low-resistance con- nection to "infinite" earth. The grounding system serves to protect personnel and machinery from the hazards associated with electrical equipment that is operating improperly. The protection afforded can be divided into the following four functions, which are the main purposes behind grounding the system. First, the grounding system must limit potential gradients between conducting materials in a given area (38). 2 During a ground fault, for instance, a phase conduc- tor comes into contact with a machine frame, and current flows through the equipment; subsequently, the potential of the equipment tends to become elevated above ground potential by an amount equal to the voltage on the conductor. If a person touches the machine, while being simultaneously connected to ground in some manner, the body's potential can become elevated, possibly to a lethal extent. The maximum potential to which a person could be exposed when touching a machine frame is equal to the voltage drop along the grounding conductors. Thus, the grounding system must provide a low-resistance path for the fault current to return to the source, and the ground conductors should have low resistance so they can carry the maximum expected fault current without excessive voltage drop. An example of the exposed potential in a surface mining situation is illustrated in figure 7.1 (38). Second, the grounding system should limit the energy available at the fault location. Heavy arcing or sparking can ignite nearby combustible material. The air itself can become ionized, making it capable of carrying tremendous amounts of current. A high-energy fault can vaporize breakers, switchgear, and phase conductors, and protec- tive enclosures may be blown apart with explosive force (21). Controlling the maximum allowable fault current significantly reduces the danger of fire and holds equip- ment damage to a minimum. Third, the control of overvoltages is essential. An overvoltage condition may occur by accidental contact of equipment with a higher voltage system, or from transient phenomena due to lightning strokes, intermittent ground faults, autotransformer connections, or switching surges (4). The maximum ratings for cable insulation, trans- former windings, relay contactors, and so forth may be temporarily exceeded in these cases. This does not usually result in an immediate breakdown of equipment, but component parts of the electrical system are successively overstressed and weakened by repeated exposure (see 1 The author wishes to thank Alan M. Christman, who prepared the original material for many sections of this chapter while he was a graduate student at The Pennsylvania State University. 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. chapter 11). This leads to premature failures, reduced component life, and mysterious "nuisance trips," which can occur without apparent reason. By providing a path between the transformer neutral and ground, most of the sources of transient overvoltages can be reduced or possi- bly eliminated. Last, a grounding system should isolate offending sections by selective relaying of ground faults (44). The sensitivity and time delays of the protective circuitry should be adjusted so a fault in a certain area will cause the local breaker to sense the malfunction and quickly remove power from only the affected section. If the relative tripping levels and speeds are not established correctly, nearby breakers may not trip when they should, and a small problem could escalate into a large calamity. Con- sequently, power to half a mine may go out because of poor relay coordination, and much time could be lost in the effort to trace and locate the trouble spot. Thus, the relaying system must be arranged so, even at the lowest level of the power-distribution chain, sufficient fault cur- rent can flow to enable the protective circuitry to sense it and take remedial action. Chapters 9 and 10 cover the protective circuitry used to provide the function of section isolation, while chapter 11 describes the devices employed with the grounding 3- phase diagram Phase-conductor resistance -AAM VWV — 31 Neutral resistor Total fault current Current through grounding conductor . Total fault Line-to-ground j current fault on shovel ^* f Shovel frame is at this potential ^VWV -WJV- 1 Grounding-conductor resistance Safety ground-bed R s resistance T . Current through earth _ -2^- Potential to A which operator may be p § subjected U Im Circuit for line-to-ground fault Figure 7.1.— Illustration of electrical shock hazard. 160 system for transient overvoltage control. As an introduc- tion, this chapter looks mainly at the first three purposes and presents the common methods of system grounding, the effect of electric shock on human beings, mine ground system characteristics, and ground-bed construction. Ex- tensive information about grounding is contained in prac- tically all subsequent chapters. GROUNDING SYSTEMS figure 7.3, now has its neutral solidly referenced to ground (20). The hazards of this system are due to the magnitude of the fault current. Detection equipment must be sensi- tive enough to detect low-level fault currents and fast enough to disconnect bad circuits before heavy faults can disrupt system integrity. Large fault currents, typically several thousand amperes, can explode protective enclo- sures, destroy equipment, and start fires, which is an excellent reason for not using this technique in explosive atmospheres. Over the past few decades, several different grounding philosophies have held sway in the electrical industry, each with its own advantages and disadvantages (20). These methods of grounding are discussed below. Note that reactance-grounded systems are not presented in the following paragraphs, as they are not normally used in industrial power systems. Ungrounded Neutral The ungrounded system was probably the first to be used because of its simplicity. Here there are no inten- tional ground connections in the system whatsoever. How- ever, a perfect ungrounded system cannot exist, since any current-carrying conductor may be coupled to ground through numerous paths, including the distributed capac- itance of its wiring, or through motor windings (49). This phenomenon is shown in figure 7.2 (20). The first line- to-ground fault on such a system will have very little effect (27) because there is no way for the fault current to find a complete circuit back to the source, and its magnitude will be very small or nil. Very low fault current means no flash hazard and no equipment damage. Circuit operation con- tinues normally with no interruption of power, an impor- tant consideration in industries where downtime is criti- cal (60). The first fault is often hard to locate because its effects are negligible. Often no repair effort is made until a second fault occurs, with its concomitant hazards of arcing, heavy current flow, and equipment damage. Since the entire system is "floating," there is no control of transient overvoltages. Except for the problem of acciden- tal contact with a higher voltage system, all the other overvoltage sources mentioned previously are enhanced because of distributed capacitance to ground (20). Solidly Grounded Neutral An alternative is the solidly grounded neutral (20). The first ground fault produces a substantial neutral current flow, which may be quickly sensed by protective circuitry, thereby shutting down the bad section. Overvolt- ages are controlled since the system, as illustrated in Supply transformer Ground "'^Ja^^ balanced Figure 7.2.— Capacitance coupling in ungrounded system. Low-Resistance Grounded Neutral The low-resistance grounded-neutral system is estab- lished by inserting a resistor between the system neutral and ground. The resistance is such that ground-fault currents are limited from 50 to 600 A, but are commonly about 400 A (20). Transients are controlled by the ground connection, and ample fault current is available for actu- ating protective relays. The flash hazard is not as serious as in the solidly grounded neutral system, but a current flow of 400 A can still do considerable damage. To limit damage, the least sensitive ground relay should respond to 10% of maximum ground-fault current. A schematic dia- gram of this method is shown in figure 7.4 (20). High-Resistance Grounded Neutral Perhaps the best technique, and that required by law in coal-mining applications on portable or mobile equip- ment, is the high-resistance grounded system, often re- ferred to as the safety ground system. The neutral ground- ing resistor is sized according to the system voltage level, Supply transformer secondary Ground Figure 7.3.— Solidly grounded system. Supply transformer secondary Figure 7.4.— Resistance-grounded system. 161 in general to limit ground-fault current at 50 A or less. Where the line-to-neutral potential is 1,000 V or less, the grounding resistor must limit fault current to 25 A or less; above 1,000 V, the voltage drop in the grounding circuit external to the resistor must be 100 V or less under fault conditions, With this system, sensitive relaying must detect faults on the order of a few amperes to provide fault isolation and facilitate quick location of the trouble spot (60). The level of fault current is also low enough to practically eliminate arcing and flashover dangers. The ground connection also serves to limit the amplitude of overvoltages. However, loads cannot be connected line to neutral, as the grounding conductor must not carry any load current. ELECTRIC SHOCK For a safe grounding system to be efficiently and economically designed, voltage and current levels that are harmful to human beings must be determined. With the trend toward larger and more powerful mining machinery, distribution voltage and current levels have risen propor- tionately. Constant vigilance is required when using elec- tricity if the hazard of electrocution is to be avoided. Even if a shock is nonlethal, involuntary movement caused by the shock may lead to serious injury or death. As an example, a man standing upon a ladder may come into contact with a live wire and fall from his perch (12). Physiologically speaking, the muscles of the body are controlled by electrical impulses transmitted from the brain via the nervous system. These pulses occur at a rate of about 100 per second and may be of positive or negative polarity. From this, it can be seen that the human "in- ternal power supply" operates at about 50 Hz, which is exactly the frequency of the electric power generated in Europe, and is only 10 Hz removed from the U.S. power generation frequency of 60 Hz. This is an unfortunate coincidence, for tests have shown that the most dangerous frequencies to which a person can be exposed are power frequencies in the range of 50 to 60 Hz (12). How sensitive are human beings to the flow of elec- tricity? Tests have indicated that for an average male holding No. 7 AWG (American Wire Gauge) copper-wire electrodes in his hands, 60-Hz ac is first perceived at a level of about 1 mA (12). By intermittently touching or tapping an electric conductor, currents of only 1/3 mA can be felt. In the case of dc, the threshold of perception for the average male is 5.2 mA. Sensitivity levels for women in the cases mentioned above can be found by multiplying the male values by a factor of two-thirds (13). It is gener- ally agreed that the magnitude and duration of the cur- rent are the important shock parameters, rather than the potential difference or voltage (12), as can be seen in table 7.1 (42). As current magnitude is increased above the level of perception, many test subjects have reported a tingling sensation, the intensity of which increases as the current rises. Generally, muscles in the vicinity of the current path start to contract involuntarily, until finally a point is reached where the subject being tested can no longer release his grip on the conductor (14). The maximum current magnitude that a person can withstand while still able to release the live conductor through the use of muscles stimulated directly by the current, is called the let-go current (fig. 7.5) (14, 16). Tests performed on hun- dreds of volunteers have shown that the maximum let-go current for a healthy adult male is 9.0-mA ac and 60-mA dc. The corresponding values for women are 6.0-mA ac and 41-mA dc. These safe-limit values apply to 99.5% of the sample population (11). The value of a specific individual's let-go current is virtually constant, even with repeated exposures to that current level. In addition, these multiple exposures can be tolerated with no ill effects (16). It has been stated that human tissue possesses a negative resistance characteristic. In other words, an increase in current magnitude or contact duration leads to a decrease in the value of skin resistance (17). In any case, if a person has grasped a live conductor and realizes that he/she cannot let go, fear-induced perspiration will cause a lowering of the body's resistance, and more current will flow. For ac, when the current level across the chest reaches more than 18 to 22 mA, the chest muscles tighten involuntarily and breathing ceases. Although circulation of blood by the heart is unimpaired, death by asphyxiation can occur within minutes (43). If an individual's initial contact with a live wire causes a current flow ranging from about 50 to 500 mA, ventricular fibrillation may result (48). Under normal conditions, the heart beats with a strong, coordinated rhythm. However, a current passing through the heart when the ventricles (the heart's two large pumping cham- bers) are just starting to relax after a contraction, can cause the various fibers of the heart muscle to beat weakly in an uncoordinated manner (43). In this condition, known as ventricular fibrillation, the heart is almost totally incapacitated and blood circulation decreases practically to nothing. Within 2 min, the brain begins to die because of oxygen deficiency. Once initiated, ventricular fibrilla- tion almost never stops spontaneously, and treatment by trained medical personnel must be secured if the victim is to survive. Obviously, people cannot be used as test subjects in ventricular fibrillation experiments because of the high risk involved. Numerous tests have been carried out on several species of animals and the results extrapolated, Table 7.1. —Current range and effect on a typical man weighing 150 lb Current Less than 1 mA.. 1 mA 3 mA 10 mA 30 mA 75 mA 4 A Greater than 5 A. Physiological phenomena Effect on man None Imperceptible. Perception threshold Mild sensation. Pain threshold Painful sensation. Paralysis threshold of arms and hands Person cannot release hand grip; if no grip, victim may be thrown clear. Tighter grip because of paralysis may allow more current to flow; may be fatal. Respiratory paralysis Stoppage of breathing, frequently fatal. Fibrillation threshold (depends on time) Heart action uncoordinated, probably fatal. Heart paralysis threshold (no fibrillation) Heart stops on current passage, normally restarts when current interrupted. Tissue burning Not fatal unless vital organs are burned. 162 Ld X o 9 I 30 r 20 10 _L J L J 5 10 50 100 5001,000 5,000 FREQUENCY, Hz Figure 7.5.— Effect of frequency on let-go current for men. based upon body weight, to cover human beings (15). It was found that fibrillating current is proportional to body weight and inversely proportional to the square root of the shock duration. Using 50 kg (110 lb) as a body weight, it has been proposed that the value of current that can be safely endured by 99.5% of normal adults without causing ventricular fibrillation is (16) I = 116 5s 8.3 ms (7.1) where I = rms ac, mA, and t = shock duration, s. As noted, this equation is valid for values of time between 8.3 ms and 5.0 s (15). It may be seen from the above equation that for a 1-s contact time, the ventricular fibrillation threshold current is about 116 mA. Since a normal person has a pulse rate between 60 and 80 beats per minute, the critical phase of the heartbeat (when a person is vulnerable to ventricular fibrillation) occurs about once each second. Therefore during a shock lasting for 1 s or more, the heart must pass through this critical phase (48). As a result, it is thought that ventricular fibrillation is the leading cause of death by electric shock. Higher currents on the order of a few amperes will freeze both the chest and heart muscles, thereby prevent- ing the onset of ventricular fibrillation. Generally, the heart will restart upon the cessation of current flow (48). These current magnitudes are less dangerous statistically than the lower values where fibrillation is prevalent. Further increases in current level, to 5A and above, may produce serious burns leading to shock and possible death, while current levels that substantially elevate body tem- perature produce immediate death (16). In an electric-shock situation, the victim's electrical resistance plays an important role in determining how much current will flow, as indicated by Ohm's law: I = (7.2) For a human being, at least three components of resis- tance have been isolated: contact resistance, skin resis- tance, and internal resistance (43). Contact resistance, as illustrated by table 7.2, depends upon the degree of skin moistness and the area of contact with the live conductor (42). Values of 40,000 to 50,000 fl/cm 2 are given for dry skin and 1,000 ft/cm 2 for wet skin (13). Skin resistance depends upon the physical condition of the tissues: A person who does rough, heavy outdoor work may have a skin resistance of 10,000 fi, while a value of 1,000 Q is typical of a sedentary office worker (43). Internal resis- tance is the resistance of the body's interior and is gener- ally accepted to be about 500 ft between major extremities (25). Table 7.2.— Typical resistance for various contact situations, ohms Contact 1 Dry skin Wet skin Finger touch 500,000 20,000 Hand on wire 50,000 10,000 Finger-thumb grasp 20,000 5,000 Hand holding pliers 20,000 2,000 Palm touch 10,000 1,000 Hand holding 1.5-in pipe 2,000 500 2 hands holding 1.5-in pipe 500 NA Hand, immersed NAp 200 Foot, immersed NAp 100 NA Not available. NAp Not applicable. 1 Skin surface only; resistance may be lower when skin is cut, blistered, or abraded. Voltage magnitude has some effect upon the body's reaction to electric shock, although current is by far the most important parameter. Potentials greater than about 240 V simply puncture the skin, thereby negating the effects of skin resistance (12). There is also some evidence that overall body resistance varies inversely with the applied voltage, although this is subject to disagreement. The relationship is given by (43). R oc E (7.3) where R = resistance, 12, E = potential, V, and n = 1.5 to 1.9. Above about 2,400 V, tissue damage due to burning becomes the major cause of electric-shock injury (42). Thus it can be seen that the body's response to electricity is extremely complex, and currents on the order of a few milliamperes can be fatal if long continued. CHARACTERISTICS OF MINE GROUNDING SYSTEMS The concept of protecting mine electrical equipment and personnel against the consequences of ground faults by suitable grounding has existed since electricity was first introduced into coal mines. As early as 1916, the Bureau of Mines recommended equipment frame ground- ing as a means of preventing electrical shock to miners working on or around electrical equipment (6). For the coal mining industry, a suitable grounding system has always been a difficult problem, more complex and difficult than in other industries. Ground Beds For mine usage, the electrical distribution cables and overhead transmission circuits carry into the mine one or 163 more grounding conductors in addition to the phase con- ductors. Each piece of ac equipment has its frame solidly connected via these grounding conductors to a safety ground bed commonly located near the main surface substation and consisting of buried horizontal conductors or driven rods, or a combination of both. The neutral of the substation transformer secondary is also connected to the safety ground bed through the neutral grounding resistor, as shown in figure 7.6. It should be noted that many important components are missing from this diagram, and chapter 13 covers substation circuitry in detail. The substation actually requires two ground beds, maintained some distance apart. Lightning discharges and other transformer primary surging conditions are directed to the system or station ground. The system and safety grounds must be kept separate so current flow intended for one will not enter the other. It is essential for the safe operation of the mine power system that the resistance of the beds be maintained at 5.0 U or less (3, 39, 44). A ground bed with this resistance range is often termed a low-resistance ground bed. To demonstrate one reason for a low-resistance bed, consider a situation where lightning strikes the substa- tion, and 10,000 A is discharged through the surge arrest- ers into the system ground bed. If the ground bed is of 5.0-fi resistance, a potential of 50,000 V is developed, and the grounding grid of the ground bed becomes elevated to 50 kV above infinite earth. Depending upon the physical extent of the grid, a person walking through the area underlain by the grid could bridge a lethal potential gradient with his or her feet (2). Metallic objects within the potential gradient field can also be elevated to danger- ous potentials and become lethal to the touch. Typical step and touch potentials are illustrated in figures 7.7 and 7.8 (2). These step and touch potential hazards are applicable to both the system and safety ground beds. However, the dangers of a high-resistance safety ground bed are not found close to the bed but at the mining equipment. The most insidious feature of the safety ground system is that the equipment connected to it is maintained not at earth potential, but at the safety ground-bed potential. Unless the bed has low resistance, any safety ground-bed current flow can render every piece of mine equipment potentially lethal. The flow can be created by faults to earth, coupling from lightning strokes to the system ground, lightning strokes to safety grounded machinery, and stray currents from dc haulage systems. Three such cases are illustrated in figures 7.9, 7.10, and 7.11 (9). Consequently, with high-resistance ground beds, an elevated frame potential is a problem not just on the machine where it occurs, but everywhere (10). Rf? $ f "flfc 1 — wwwwwww IR 1 "2 Figure 7.7.— Step potentials near grounded structure. Potential rise above remote earth during fault 0.5 R f I A/WVWWVWV 1 Figure 7.8.— Touch potentials near grounded structure. Power center Load Incoming power > Surge arresters 9 System ^. ground = Substation transformer Neutral resistor -± Safety - ground 3-phase power to equipment Grounding conductor Figure 7.6.— Simplified one-line diagram of substation. to machine frame W Safety ground bed ys Fault to Figure 7.9.— Line-to-earth fault resulting in current flow through safety ground bed. 164 Substation Figure 7.10.— Lightning stroke to equipment causing current flow through safety ground bed. Mining machine Safety ground bed Figure 7.11.— Lightning stroke current through system ground bed causing elevation of safety ground bed. Grounding in Underground Mining Early practice in underground coal mining was to drive a metal rod into the mine floor and use that as a ground. In almost every case this arrangement proved to be totally unacceptable, with test measurements indicat- ing 25-fi or more resistance (28). With the exception of pumps, the contact resistance of mining machinery with the mine floor also proved to be too high for adequate grounding. Rail haulage track systems, even though often poorly bonded, showed much lower resistance to ground than most metallic rods driven specifically for that pur- pose. As a solution, Griffith and Gleim (28) in 1943 stated that ". . . consideration should be given to a grounding circuit carried to the outside of the mine." Present coal mine practice does just that. A simple form of the bipower (mixed ac-dc) system in use in underground coal mines today is illustrated in figure 7.12. After transformation, three-phase ac power enters the mine to supply the various three-phase ac loads. Some of the ac power is converted to dc at rectifier stations to power the locomotive system and, occasionally, dc face equipment. More often, any dc face machinery is powered from rectifiers located in the mine section. Except for the trolley system, all dc as well as the ac equipment frames are connected to a common junction, which is tied to the surface safety ground bed. In order for the system to be effective, grounding conductors must be continuous and ac loads Figure 7.12.— One-line diagram of simplified mine power system. this continuity must be verified. Ground-check monitors ensure this. Trolley locomotives generally utilize the overhead trolley wires as the positive conductor and the tracks as the negative. Neither of these is tied to the rectifier- station frame ground. However, because the track is in contact with the mine floor, the negative conductor for the trolley system is grounded. The dc system that supplies power to face equipment normally employs trailing cables that have neither the negative nor positive conductor grounded. Thus, this subsystem is often ungrounded un- less the supply is obtained from the trolley system. Note that in diode-grounded systems, the negative conductor is grounded. At each transformation step within the power system, such as in a power center, an additional neutral point must be established on the transformer secondary. The neutral is tied through a grounding resistor to the equipment frame and thus via the grounding conductors to the safety ground bed (an exception will be discussed later). Even with all these grounding points, the ac ground- ing system must be isolated from separate dc power systems. If it is not, dc may appear in the ac grounding system, thus elevating it above true ground potential. If an ac ground current is present, it will be offset by the dc level. The principal concern is with trolley installations, where isolation is achieved by having no common points between the ac and dc systems. Various techniques have been tried to maintain separation or to eliminate dc offsets while grounding dc face equipment frames. 165 Face Equipment Grounding When a working section utilizes an ac continuous miner energized from a section power center and dc shuttle cars powered from the trolley system, the ground potentials of the dc and ac equipment frames are not necessarily equal, because of the voltage drop in the track. Jacot (33) suggested that this problem could be solved by isolating the low-voltage ac neutral point from the power- center frame and also the high-voltage grounding system, and connecting it via an insulated cable to the track, as shown in figure 7.13. The low-voltage neutral point re- mains connected to the ac face-equipment frames. This technique should make the low-voltage ac and dc equip- ment frame potentials the same, thus eliminating dc offset problems. Difficulties can still arise with this method. If any track rail bonds are bad between the ac and dc low-voltage ground points, the dc frame potentials might be elevated with respect to the ac frames. Further, the power center must be constantly maintained at a safe distance from the tracks to preserve isolation between the track and high-voltage grounding systems. Another method is shown in figure 7.14A. Here, a section power center supplies power to ac face equipment and also, through a rectifier, to dc machinery (usually shuttle cars). The rectifier is isolated within the section power center, but the dc output is grounded through a center-tapped current-limiting resistor. All dc equipment frames are then grounded by trailing-cable grounding conductors, which in turn are connected to the center tap of the grounding resistor. The latter point is connected to the high-voltage grounding system. This has been consid- ered a very safe dc ground protective system because it permits the use of protective circuitry to trip the rectifier breaker in case of a dc ground fault (see chapter 9). However, the use of the center-tapped resistor has been criticized (46). On such a system, any failure to maintain grounding-conductor conductivity or accidental connec- tion of a wrong conductor when splicing cables may lead to a hazard. Nevertheless, an important advantage of the method is that the dc and ac frame potentials can be the same. A more recent method for limiting dc ground-fault current is similar to high-resistance ac grounding and is illustrated in figure 7.145. In 1963 the Bureau of Mines accepted the use of silicon diodes as a means of grounding dc face equipment frames. When a diode is used, the grounding resistors are not needed because the frame is grounded through the diode to the negative conductor, as illustrated in figure 7.15. The diode circuit also includes a ground protective device, which will interrupt the power if a current flows from a positive power conductor to an equipment frame (again, see chapter 9). According to Jones (37), diode grounding should ensure good ground continuity since the same conductor acts as both a dc negative conductor and the grounding connection. However, a grounding diode only protects the dc system against ground faults within the equipment frame. Current leakage to ground or faults within trailing cables can still present hazards. A Resistors between dc line conductors and grounding conductor Grounding resistor From safety ground bed Ground conductor from power-center frame to safety ground bed Power center Continuous miner(ac) J7 Neutral ^ resistor § Lifted from ' frame Insulated grounding > length), then (50) R = p 22 (In-), n27if (7.8) where n is the number of rods and the other variables are as previously defined. If the rod spacing-to-length ratio is small (length > > spacing) (50), R = 2^ (ln A ) ' (7.9) where A = (a S 12 S 23 S 34 ) l and S 12 = spacing between electrodes 1 and 2, and so forth. Resistivity, Resistance, XI rlOO r p5,000 100 -80 1 co 100,000 r3,000 ^ with mutual effects Rod length, 10,000 ^20 - 100 1 20 \ \ \ \ ft ^s. r 100 - 50 . ft o p 2 i < "> 1- o CO " CO LU en 800 Figure 7.21.— Resistance of one ground rod, %-in diameter. A, 0.5 ft B C 20 Jyy 25 /// /^ /// s' £",100 ft 30 'Wy/' ^^r^^' 40 JS^^^^ 50 70 ^y/ — A, B, C, D, E are various // rod spacings between /^ 0.5 and 100 ft 1 00 / i i ii.i 2 4 6 8 NUMBER OF GROUND RODS 10 Figure 7.23.— Variation of earth resistance as number of ground rods is increased for various spacings between rods. 169 A formula for determining the resistance of grounding meshes is given by (53) ""SKlS- + K u 1.0 .9 I C \ i i,i, KEY Curve A ■ For depth z = Curve B : For depth z = iSEgU . (area) l/2 Curve C '■ For depth z = 13 5 7 LENGTH-TO-WIDTH RATIO Figure 7.24.— Values of coefficient k, as function of length- to-width ratio of area. UJ y u. UJ O o KEY Curve A : For depth z = ^ o . i- j xl. (area) l/2 Curve a . For depth z - — ^ — Curve C ■ For depth z _ (area)' 72 13 5 7 LENGTH-TO-WIDTH RATIO Figure 7.25.— Values of coefficient k 2 as function of length- to-width ratio of area. where L = total length of buried bare conductor, and r = equivalent radius of the system. The equivalent radius of a grounding system varies de- pending upon the exact configuration, but a safe estimate is one-half of the length of the longest diagonal line contained by the system (55). Contact resistance between the surface of the elec- trodes and the soil is not normally a significant factor if the bed has been in existence long enough for the soil to settle and compact, but in new beds it may amount to 20% of the total resistance (57). In summary, the best way to achieve a low-resistance ground is to maximize the periphery or areal extent of the grounding system. Conductor diameter has little effect upon resistance, and mechanical strength requirements should be the primary consideration. Because of wide seasonal variations in the soil resistivity of surface layers, deeply buried meshes or deeply driven rods are often preferable. This is also advisable if lower resistivity layers are known to exist at depth. Driven rods are usually preferred over buried meshes for three reasons (39): • The expense of earth removal to bury the mesh is avoided. • Rods do not require the packing of earth around the buried electrodes to ensure good earth contact. • The use of rods can give a desired resistance more easily than using any other ground-bed form. Note that although formulas are excellent for calculating the theoretical resistance of a grounding bed, the actual resistance should always be measured with an earth tester to ensure system integrity. 170 Table 7.3.— Approximate resistance formulas for various electrode configurations Electrode configuration Description One ground rod; length L, radius a *-£:»"*-« Two ground rods; spacing s>L = 7 £ r(ln «:_ 1)+ _£_ (1 .j4 + l4. . ., 4TTL 41TS 1 2 c 4 3s 5s • • Two ground rods; spacing s h- 00 co LU cr 1,000 500 100 0.20 0.08 0.12 SOLUTION, % Figure 7.34.— Typical resistivity curves of solutions. periods may heat the soil to the point where most of its moisture will evaporate. When this condition is reached, soil resistivity increases drastically. Different soils are characterized by various resistivity levels (table 7.7). To a large extent, this is due to the previously discussed effects of structure as it pertains to conductivity. Loams and clays possess a low resistivity, while shales, sandstones, and crystalline rocks occupy the high end of the scale (50). The nature of the particles making up the soil or rock is another aspect of rock structure, which influences conductivity through the rock's ability to trap and retain water. Surface tensions cause water to cling to large soil particles or grains; with small-grained substances, mois- ture simply fills up the multitude of pore spaces between individual particles. The range of particle sizes and their packing determines how much of the volume occupied by a particular soil will be void space and thus available for filling by water. If most of the grains are the same size, total pore volume may range from 26% to 46%, depending upon the manner in which the grains are packed (19). If a particular rock structure or formation is confined to a small geographical area, then it probably has a fairly uniform resistivity, excluding areas of subsequent igneous activity. Should the formation be widespread, however, chances are that variable resistivities will be noted de- pending upon location. This is due to the differences in local conditions that may have prevailed over a small area during actual deposition or formation of the rock strata. This may also be caused by variations in the ground water properties from place to place within a large region (5). Resistivity Measurements The basic procedure for measuring soil resistivity involves the determination of the potential gradient on the earth's surface caused by flow of a known current through the area. To illustrate the basic technique, assume an earth structure composed of two horizontal layers, the top one of 176 high resistivity, p lt and the lower one of low resistivity, p 2 , as shown by figure 7.35 (19). The thickness of the upper layer is given by h. A power source forces current flow through the ground between the two outer electrodes. At very small electrode spacings, the apparent resistivity will approximate p 1 since most current flow would be confined to the upper layer. At very wide spacings (much larger than h), the apparent resistivity will be about the same as p 2 , because the majority of the current would flow through the deeper layer. Many methods are available for measuring earth resistivity, such as the techniques of Gish-Rooney, Lee, and Schlumberger. Most of these procedures are based on the arrangement described by Wenner (58), which is shown in figure 7.36 (35). Four uniformly spaced electrodes are used, and a current source is connected across the two outer terminals while the potential drop is measured across the inner terminals. When the electrode length b is small compared with the spacing a, then the resistivity is (51) p = 2iraR, (7.18) where p = resistivity, fl-m or Q-ft, a = spacing between electrodes, m or ft, and R = resistance = V/I, fl. Some problems that may arise from the use of this method are • Stray currents due to leakage as from motors, • Natural currents due to electrolysis of nearby min- erals, • Polarization due to use of a dc source, • Inductance between the lead conductors, and • Leakage from the conductors and the instrument when in wet areas. The first three problems are circumvented through the use of an ac source operating at the nonpower frequency of an instrument that generates the equivalent of a square wave. The use of a well-insulated instrument and conduc- tors solves the latter two difficulties. The megohmmeter has all these features and is an excellent apparatus for use in work of this type. To perform a resistivity survey, the megohmmeter is set up as shown in figure 7.36, the instrument is operated, and a resistance value R is read from the built-in meter. The procedure is then repeated at different electrode spacings. A graph may be made comparing the resistivity, p, with the electrode spacing, a, as shown in figure 7.37 (55). For each value of electrode spacing, there is a corre- sponding value of resistivity, p a , seen by the instrument. This apparent resistivity is equal to the resistivity that a semi-infinite homogeneous earth would display at an equal electrode spacing and an identical value of R. In the example shown, the apparent resistivity decreases as electrode spacing increases. The overall shape of the curve indicates that the soil here is composed of two horizontal layers, with the overlying horizon having a higher resis- tivity then the lower one. As the electrode spacing, a, is increased, more and more of the current flow between the outer electrodes occurs in the deeper layer of the soil, and this is reflected in the continuous decrease in the apparent resistivity (5). In a case like the one just described, a grounding grid composed of deeply driven vertical rods would be best, since the rods would penetrate into the underlying layer of Electrodes ', >P 2 Figure 7.35.— Diagram for four-electrode resistivity survey showing lines of current flow in two-layer earth. Megohmmeter Figure 7.36.— Connections for Wenner four-terminal resistivity test using megohmmeter; distance a should be at least 20 times b. in UJ or UJ or < a. a. < ELECTRODE SEPARATION Figure 7.37.— Typical curve of resistivity versus electrode separation. 177 higher conductivity and thus provide a more effective ground. Additionally, soil horizons near the surface are usually subject to wide seasonal variations in resistivity due to changes in ambient temperature and moisture (40). Tagg (55) presents several methods whereby an accu- rate interface-depth determination may be calculated. Values are read from a standard graph, and multiple calculations are then performed, followed by another graph construction from which the correct depth is read. Core drilling has verified that values derived in this manner agree closely with the actual conditions. Effect of Chemical Treatment of Soils The natural resistivity of some soils is so high that it is virtually impossible to construct a ground bed with a satisfactorily low value of resistance. By injecting into the earth a substance whose resistivity is very low, the local soil resistivity can be effectively reduced, thereby lowering the resistance of a grounding grid. Such chemical treat- ment acts to increase the apparent dimensions of the metallic electrodes (7). The result of chemical treatment is to reduce ground resistance by a considerable amount, often as much as 15% to 90%. Figure 7.38 shows an example of this effect (36). Generally, the percentage improvement is greater for a very high resistance ground. Substances traditionally used as chemical additives include sodium chloride, calcium chloride, copper sulfate, and magnesium sulfate (36). Newer additives include gels composed of acrylamide, silicic acid, or copper ferrocya- nide. In the past, electrodes were sometimes surrounded by a bed of coke, not a true chemical treatment but rather a partial soil substitute (24). The effectiveness of most treatments in lowering ground-bed resistance is about the same, with the ultimate selection depending upon the criteria of cost, availability, and corrosive properties. A prime disadvantage shared by most chemical treat- ments is the fact that they will corrode most metals (7). Magnesium sulfate has little or no corrosive effect, and graphite is also innocuous. Other additives generally speed up the decay of grounding electrodes. Another disadvantage is that chemical treatments are dissipated and carried away by neutral drainage through the soil (36). Acrylamide gel, which is not water soluble, is an exception (34). The rate at which chemical additives are washed away depends upon the soil type and porosity as well as the amount of rainfall. Useful life may range from 6 months to 5 or more years. The cost of chemical treatment may be higher than the price of driving longer ground rods to reach deeper, lower resistivity soil layers, but in some instances it is not feasible or desirable to increase penetration depth. As shown in figure 7.39, the seasonal variations in resistance that are exhibited by grounding grids because of temper- ature and moisture fluctuations, are attenuated in those cases where chemical treatment has been applied (36). The best method of application, illustrated in figure 7.40, is to dig a circular trench about 1 ft deep and with an inside diameter of 18 in around each ground rod (36). The additive is placed into the trench and then covered with earth. The area is then flooded with water to initiate the solution process. In this manner, the solution can perme- ate a greater volume of soil, while any corrosive action is minimized. 1,600 r # Before treatment July July Jan. MONTH July Jan. Figure 7.38.— Reduction in ground mat resistance by soil treatment. Untreated Treated j i I L Jan. MONTH Figure 7.39.— Seasonal resistance variations attenuated by soil treatment. Ground rod Figure 7.40.— Trench model of soil treatment. GROUND-BED CORROSION Corrosion is a phenomenon that must be considered in the design of a ground bed. There are three basic ways by which underground corrosion can occur (52): • Dissimilar metals connected together electrically and surrounded by an electrolyte such as soil, 178 • Dissimilar electrolytes in close proximity to the same piece of buried metal, and • Stray electrical current leaving a buried metal structure. In the first mode, variations in electrochemical poten- tial provide the key to the dilemma. The standard half- cell, upon which most corrosion work is based, consists of a copper rod bathed in a saturated copper sulfate solution. When measured with reference to the copper and copper sulfate half-cell, each metal displays a certain character- istic potential, as shown in table 7.9 (61). If two metals are joined and immersed in soil, the one whose potential is more negative will discharge current and be corroded, but the more positive (noble) species will collect current and be protected. When only one metal is used, corrosion can still occur because of differences in soil composition. Metal in an oxygen-rich zone will be protected, while metal in a relatively oxygen-poor soil horizon will be attacked. For- eign metallic structures in the grounding-grid vicinity, such as pipes, cable sheaths, and building frames, may also act in conjunction with the ground bed to form an anode-cathode corrosion situation. Table 7.9.— Typical potentials of metals in soil measured from a copper and copper sulfate reference electrode Metal Potential, V Magnesium -2.5 Aluminum -1.3 Zinc -1.1 Iron -.7 Copper - .2 The engineer designing a ground-bed system is faced with the problem of solving two conflicting sets of de- mands. For safe grounding, a very low resistance is desired between the soil and the buried metallic grid. To eliminate potential-gradient hazards, all metal structures should be tied together. However, protection requires that under- ground metallic structures be insulated from the corrosive effects of the soil. Similarly, the soil and metallic structure should be isolated from one another (61). This seeming paradox may be remedied by making the correct choice of ground-bed conductor and by applying suitable preventive techniques. Copper makes an ideal ground-bed conductor since it is corrosion resistant, has a high electrical conductivity, and is easy to clamp or weld (61). However, a good all-copper system is often ruined by tying it together with noncopper structures in the same locale, thereby leading to the corrosion of the less noble species (52). If the ground bed must be located in an area where steel or lead are present, two options are available. First, an insulating coating may be applied to the base metals. If this is not feasible, an all-steel grounding system is preferable, or one composed of steel rods connected with insulated copper wire (8). The idea here is to minimize the exposed surface area of the more noble metal. Normally, steel electrodes can be improved by applying a heavy zinc coating or by driving zinc electrodes in addition to the steel. Known as sacrificial anodes, the zinc conductors will be preferen- tially attached, thereby protecting the steel members. For extra protection, magnesium may be used instead of zinc. In highly corrosive soils, it may be necessary to utilize an external power source that supplies dc to the soil in order to nullify the natural corrosion currents. This is known as cathodic protection (41). For externally driven anodes, zinc or magnesium may be used; graphite and high-silica cast iron are also suitable. It may be seen that judicious choice of grounding materials and the use of corrosion-prevention techniques such as cathodic protection can provide a ground bed that is both low in resistance and high in longevity. GENERAL GROUND-BED GUIDELINES The primary objective of a grounding system is "to limit the potential rise above ground that appears on the frame and enclosures of the equipment connected to the power system" (30). Consequently, the station ground and safety ground beds should be spaced at least 50 ft apart, even though the law presently permits only a 25-ft sepa- ration (21). A typical voltage-gradient representation is shown in figure 7.41 (30). The two ground beds must be far enough apart so current surges through the station ground bed will not cause the safety ground bed to rise to more than 100 V above infinite earth. Once the site has been selected, the excellent guide developed by King (39) can be used for the design and construction of low-resistance driven-rod ground beds. The simplified procedure consists of the following four steps. 1. Using the Wenner array, earth resistivity is mea- sured along the two lines at right angles to each other, centered across the proposed ground-bed site. Two mea- surements, with 6-ft and then 18-ft spacings, are taken along each line. 2. Depending on the magnitude and homogeneity of the resistivities measured, the rod length, number, and arrangement are selected from tables. These tables (39) are based on the same information presented earlier in this chapter but are too extensive to be reproduced here. L±J o > 100 V System ground Safety ground Figure 7.41.— Voltage gradients in earth during ground-fault conditions. 179 3. The selected rod configuration is driven, and the rods are interconnected with flexible, bare copper conduc- tors. Recommended size is 4/0 AWG, and connection should be clamped or brazed, but never soldered (1). 4. The completed bed is measured by the fall-of- potential method to check that its resistance is below 5.0 Q. If it is more, a new resistivity, p n , is calculated by P» = Po|. (7-19) where p = old resistivity, and R = measured resistance. Again, using the tables, additional rods are selected and then driven. Afterward, the resistance is again measured. Whatever the procedure used to construct the bed, the resistance should be checked not only when it is installed but periodically thereafter to ensure that it is still func- tioning properly. GROUNDING EQUIPMENT The basic resistance-grounded system consists of a resistance inserted between the power-system neutral point and ground. Specific concerns when selecting the grounding resistor are resistance, time rating, insulation, and connection. A problem also exists if there is no available connection to the power-system neutral. Grounding Resistor The ohmic value of the resistor is determined by the line-to-neutral system voltage and the maximum ground- current limit. As stated earlier, when portable or mobile equipment is involved, the maximum limit on low-voltage and medium-voltage systems is 25 A, and the upper current limit on high voltage is set by the grounding- conductor resistance, because flow through this conductor cannot cause any machine frame potential to be elevated more than 100 V above earth potential. However, high- voltage limits are typically chosen at 25 or 50 A (50 A is the maximum allowed in some States). For instance, if grounding-conductor resistance is 3.3 Q, and maximum allowable ground current is 30 A, then 25 A is normally chosen. When the resistance-grounded system is feeding only stationary equipment, there is no specified maximum ground current, but industry practice sometimes specifies low-resistance grounding with a 400-A limit. For all appli- cations, sizing the ohmic value of the grounding resistor is simply performed by dividing the line-to-neutral voltage by the selected ground-current limit; conductor impedance is neglected. The technique is justified by the method of symmetrical components for a line-to-neutral fault. Ground current can be limited at a level less than the restricted maximum, but for high-resistance grounding the smallest value chosen has two concerns: ground-fault relaying and charging current. For maximum safety, ground protective circuitry should sense ground current at a fraction of current limit (see chapter 9). Hence, reliable relay operation with electromechanical devices can be a problem if maximum current is less than 15 to 20 A. The other limitation is that ground-fault current should al- ways be greater than the system-charging current (59), the current required to charge system capacitance when the system is energized (see chapter 11). When very low ground-relay settings are used, the charging current may itself cause tripping. The second main concern in selecting a grounding resistor is its time rating, or the ability to dissipate heat. A grounding resistor carries only a very small current under normal system operation, but when a ground fault occurs, the current may approach full value. The high current exists until the circuit breaker removes power from the faulted circuit, which may take from a fraction of a second to several seconds depending upon the protective circuitry used. With correct fault removal, the physical size of the resistor can be small, as very little heat is produced. However, protection devices have been known to malfunction, and in these instances ground current might continue to flow until the power is removed manually. Thus, the resistor must be able to dissipate the power produced from full ground current for an extended time when portable or mobile equipment is involved. If not, the resistor can burn open and unground the system. Two ratings that ensure safety are continuous and extended time. These are essentially the same, since the extended- time rating refers to a heat-dissipation ability for 90 days per year (32). To provide a safety margin, the transformer-neutral side of the resistor (often called the hot side) must be insulated from ground at a level to withstand the line- to-line system voltage. Both resistor ends are at ground potential with normal operation but under a ground fault, the transformer end can approach line-to-neutral poten- tial. To afford good insulation, it is recommended that the resistor frame be placed on porcelain insulators, not tem- porary supports such as wooden blocks. Furthermore, for wye-connected secondaries, the transformer-neutral bush- ing must be insulated to at least line-to-neutral voltage. The last concern is the resistor connection. The grounding resistor is installed between the transformer neutral and the safety ground bed. In substations it is important to use insulated conductors, because bare con- ductors can easily compromise the required separation between the system and safety ground beds. Grounding conductors must extend from the ground-bed side of the resistor. Finally, to minimize resistor conductor lengths, the resistor must be located on the power-source end of distribution, as close as possible to the source power transformer. Distances greater than 100 ft are usually too long. Grounding Transformers Delta-wye, wye-delta, and delta-delta power trans- formers are extremely important in mine power distribu- tion because they offer very high impedance to zero- sequence currents. As a result, a ground fault existing on the secondary will do no more than raise primary line current. However if the transformer has a delta secondary, there is no neutral point to which the grounding system can be connected. Another case where this occurs involves mines where the utility company owns the substation and supplies ungrounded delta power. For both these situa- tions, a separate grounding transformer is needed to obtain an artificial neutral. The two types of grounding transformers in general use are the zig-zag and wye-delta, with the former being more popular. As shown in figure 7.42, the zig-zag is a special three-phase transformer designed for deriving the neutral. 180 The transformer winding interconnections are such that a very high impedance is shown for positive-sequence and negative-sequence currents but a very low impedance is exhibited during zero-sequence flow. A wye-delta grounding-transformer bank uses three identical single-phase transformers (fig. 7.43). The pri- mary windings, rated at line-to-neutral voltage, are con- nected in wye among the power-transformer secondary terminals and the grounding-resistor hot side, and the secondaries are connected in delta. Any secondary voltage rating can be used. Normally, no secondary current will flow, but during a ground fault, current will circulate in the secondary. This will cause the ground to be shared by the three transformers such that the neutral point will remain at constant potential. Grounding-transformer capacity only needs to be large enough to carry the maximum ground-fault current. Grounding transformers' primaries cannot be fused, as an open fuse will essentially unground the system, creating a dangerous situation. Main transformer delta secondary Zig-zag - transformer To ac circuit breakers Grounding resistor 1 WW Figure 7.42.— Delta secondary with zig-zag grounding. Incoming power Wye-delta grounding transformer System ground bed Grounding resistor Safety -= ground bed Figure 7.43.— Delta secondary with wye-delta grounding transformer. SUMMARY Several basic grounding methodologies exist, and each has its merits. The resistance-grounded neutral sys- tem is superior for mining applications involving portable or mobile equipment. The design of ground beds is a complex field, and many variables must be examined in an attempt to derive an optimum configuration. A low value of resistance is of primary importance so dangerous poten- tials are not developed on machine frames. High potential gradients in the ground-bed area must also be avoided to prevent injury to personnel. A study of electric shock and its effects on humans is helpful in further delineating this subject. Formulas have been presented that may be used to predict the earth resistance of a particular metallic array or to determine how much buried metal is needed to achieve a desired value. In order to verify the ground-bed earth resistance, a description of ground test instrumen- tation, its utilization, and data interpretation was also included. When designing a ground bed, corrosion effects and soil-heating phenomena, caused by current flow in the ground system, must be considered. The resistivity of the soil in which the ground bed is immersed has a significant effect upon its earth resis- tance. Resistivity in turn is influenced by other factors such as earth composition, temperature, and moisture, and a thorough understanding of these relationships will be of use in metallic grounding-network design. Instru- mentation was again discussed, as well as practical appli- cations such as the determination of the best location for a ground bed. Chemical treatment of soils to increase con- ductivity and attenuate seasonal resistivity variations was reviewed. Correct selection and coordination of protective cir- cuitry is essential to gain the full benefits of a low- resistance ground bed. Protective circuitry must be in- stalled to monitor current flow in the ground conductors or the potential drop across the neutral grounding resistor. When properly coordinated, this protective circuitry will quickly shut down faulty sections of the electrical system. In the event of a fault or short circuit on a piece of mine machinery, its frame may become hot or elevated above ground potential. An unsuspecting miner could be seriously injured or killed if the machine is touched. Fast-acting relays and circuit breakers will minimize the length of time during which this shock hazard exists, and the bad circuit will be isolated from the remainder of the system. These protective devices form the subject of chap- ters 9 and 10. The grounding conductors that tie equip- ment frames to the safety ground bed are discussed in the next chapter, "Distribution." REFERENCES 1. American Institute of Electrical Engineers Committee. Ap- plication Guide on Methods of Substation Grounding. Trans. Am. Inst. Electr. Eng., Part 3, v. 73, Apr. 1954. 2. Voltage Gradients Through the Ground Under Fault Conditions. Trans. Am. Inst. Electr. Eng., Part 3, v. 77, Oct. 1958. 3. American Standards Association. American Standard Safety Rules for Installing and Using Electrical Equipment in and About Coal Mines (M2.1). BuMines IC 8227, 1964. 4. Brereton, D. S., and H. N. Hickok. System Neutral Ground for Chemical Plant Power Systems. Trans. Am. Inst. Electr. Eng., Part 2, v. 74, Nov. 1955. 5. Card, R. H. Earth Resistivity and Geological Structure. Trans. Am. Inst. Electr. Eng., v. 54, Nov. 1935. 181 6. Clark, H. H., and C. M. Means. Suggested Safety Rules for Installing and Using Electrical Equipment in Bituminous Coal Mines. BuMines TP 138, 1916. 7. Clark, R. J., and B. 0. Watkins. Some Chemical Treatments To Reduce the Resistance of Ground Connections. Trans. Am. Inst. Electr. Eng., Part 3, v. 79, Dec. 1960. 8. Coleman, W. E., and H. G. Forstick. Electrical Grounding and Cathodic Protection at the Fairless Works. Trans. Am. Inst. Electr. Eng., Part 2, v. 74, Mar. 1955. 9. Cooley, W. L. Design Consideration Regarding Separation of Mine Safety Ground and Substation Ground Systems. Paper in Conference Record -IAS 12th Annual Meeting (Los Angeles, CA, Oct. 1977). IEEE, 1977. 10. . Evaluation of In-Mine Grounding System and Codification of Ground Bed Construction and Measurement Techniques (grant G0144138, WV Univ.). BuMines OFR 20-77, 1975; NTIS PB 263 119. 11. Dalziel, C. F. Effect of Wave Form on Let-Go Currents. Trans. Am. Inst. Electr. Eng., v. 62, Dec. 1943. 12. Electric Shock Hazard. IEEE Spectrum, v. 9, Feb. 1972. 13. . The Threshold of Perception Currents. Trans. Am. Inst. Electr. Eng., v. 62, Dec. 1943. 14. Dalziel, C. F., J. B. Lagen, and J. L. Thurston. Electric Shock. Trans. Am. Inst. Electr. Eng., v. 60, Feb. 1941. 15. Dalziel, C. F., and W. R. Lee. Lethal Electric Currents. IEEE Spectrum, v. 8, Sept. 1971. 16. , Re-evaluation of Lethal Electric Currents. IEEE Trans. Ind. and Gen. Appl., v. 4, Sept./Oct. 1968. 17. Dalziel, C. F., and F. P. Massoglia. Let-Go Currents and Voltages. Trans. Am. Inst. Electr. Eng., Part 2, v. 75, May 1956. 18. Dawalibi, F., and D. Mukhedkar. Optimum Design of Substation Grounding in a Two-Layer Earth Structure; Parts I, II, and III. IEEE Trans. Power Appar. and Syst, v. 94, Mar./Apr. 1975. 19. Dobrin, M. B. Introduction to Geophysical Prospecting. McGraw-Hill, 1952. 20. Dornetto, L. D. The Importance of Grounding Systems in the Protection of Personnel and Equipment. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 21. Dunki-Jacobs, J. R. The Effects of Arcing Ground Faults on Low-Voltage System Design. IEEE Trans. Ind. Appl., v. 8, May/- June 1972. 22. Dwight, H. B. Calculations of Resistances to Ground. Trans. Am. Inst. Electr. Eng., v. 55, Dec. 1936. 23. Eaton, J. R. Grounding Electric Circuits Effectively; Parts I, H, and III. Gen. Electr. Rev., v. 44, June 1941. 24. Fawssett, E., H. W. Grimmitt, G. F. Shorter, and H. G. Taylor. Practical Aspects of Earthing. J. Inst. Electr. Eng. (Lon- don), v. 87, Oct. 1940. 25. Friedlander, G. D. Electricity in Hospitals: Elimination of Lethal Hazards. IEEE Spectrum, v. 8, Sept. 1971. 26. Giao, T. N., and M. P. Sarma. Effect of a Two-Layer Earth on the Electric Field Near HVDC Ground Electrodes. IEEE Trans. Power Appar. and Syst., v. 91, Nov./Dec. 1972. 27. Gienger, J. A. Fourteen Years of Data on the Operation of One Hundred Ungrounded 240 and 480 Volt Industrial Distribution Systems. IEEE Trans. Ind. and Gen. Appl., v. 2, Mar./Apr. 1966. 28. Griffith, F. E., and E. J. Gleim. Grounding Electrical Equip- ment in and About Coal Mines. BuMines RI 3734, 1943. 29. Gross, E. T. B., B. V. Chitnis, and L. J. Stratton. Grounding Grids for High-Voltage Stations. Trans. Am. Inst. Electr. Eng., Part 3, v. 72, Aug. 1953. 30. Hamilton, D. E. Mine Power Systems: What's Your Ground Practice IQ? Coal Age, v. 66, Feb. 1961. 31. Higgs, P. J. An Investigation of Earthing Resistances. J. Inst. Electr. Eng. (London), v. 68, 1930. 32. Institute of Electrical and Electronics Engineers (New York). Requirements, Terminology and Test Procedures for Neutral Grounding Devices. Stand. 32-1972. 33. Jacot, H. D. Grounding Practice in Coal Mines. Pres. at Min. Electro-Mech. Maint. Assoc, and IEEE Min. Ind. Comm. Joint Meet., Greensburg, PA, Apr. 1, 1961; available from H. D. Jacot, North American Coal Co., Bismark, ND. 34. James G. Biddle Co. (Plymouth Meeting, PA). Getting Down to Earth. Booklet 25T, 1970. 35. Megger Ground Tester; Special Instructions. Booklet 25-J-l, undated. 36. Jensen, C. Grounding principles and Practice; II Establishing Grounds. Electr. Eng., v. 64, Feb. 1945. 37. Jones, D. R. Frame Grounding D-C Mining Machines With Silicon Diodes. Min. Congr. J., v. 51, May 1965. 38. Kaufmann, R. H. Important Functions Performed by an ef- fective Equipment Grounding System. IEEE Trans. Ind. and Gen. Appl., v. 6, Nov/Dec. 1970. 39. King, R. L., H. W. Hill, Jr., R. R. Bafana, and W. L. Cooley. 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Appl., v. 6, May/June 1967. 182 CHAPTER 8.— DISTRIBUTION The distribution system within a mine consists of various types of cables that connect equipment to power supply, the conductors that form the trolley system used in many underground mines, and the overhead lines that distribute power in some surface mines. The character of the mining operation imposes constraints on the distribu- tion system unlike those of other industries and magnifies its importance within the overall power system. Mining is by definition constantly mobile; hence, the distribution system must be handled and extended frequently and can be susceptible to damage from mobile equipment. The mobility in turn necessitates efficient methods for joining cables and repairing them in order to minimize production downtime and operating costs. In all mines there is the potential for electric shock when handling distribution components. In the hazardous environment of an under- ground coal mine, damaged systems can be a potential fire and gas-ignition source. Proper installation and correct handling practices are essential if these hazards are to be minimized. This chapter's purpose is to introduce the various distribution components used in mine power systems, as well as to discuss their construction, installation, and maintenance. Cable systems are covered first and com- prise the majority of chapter content because of their uniqueness to mining. Typical trolley-system arrange- ments are then presented, and the chapter is concluded with a brief introduction to overhead lines. NATURE OF CABLE DISTRIBUTION It was shown in chapter 1 that cables can carry the electricity from the substation, where the power is taken from utility company lines, to the point of utilization by a mining machine, pump, conveyor belt, or other piece of equipment. There are many possible variations in mine distribution, and several types of cables can be put to a similar use. Only the most typical schemes are covered in this chapter, but some notable exceptions are included. Representative systems are depicted for underground coal mines in figure 8.1 and for surface coal mines in figure 8.2. Obviously, the circuits shown in the figures are only simplified examples of actual mine systems. In prac- tice, an underground coal mine would not have one long- wall, one continuous mining section, and one conventional section, but several continuous mining sections or several conventional sections in addition to one or more longwall units. Surface mines would usually have more than one dragline and one stripping shovel, not necessarily all electrically powered. As might be supposed, the kind of cable is tied to the application. Examination of figures 8.1 and 8.2 indicates that some cables remain in stationary locations for several years, while others are moved frequently. The cables that are connected to mining machines are termed portable by the Insulated Cable Engineers Association (ICEA) stan- 1 The author wishes to thank Robert H. King, who prepared original material for many sections of this chapter. Thanks are also extended to James N. Tomlinson, who assembled the original section on splicing, and to George Luxbacher, who assembled the original material on conductor ampacities and cable derating. dards (19-2 1). 2 The Code of Federal Regulations uses the term trailing cables for the specific variety of portable cables used in a mine (38). Trailing cables are flame- resistant flexible cables or cords through which electrical energy is transmitted to a machine or accessory. In underground mines, trailing cables are generally attached to the inby end (toward the face) of the power center or distribution box. The portable cables that feed the power center or are attached to the outby end (toward the portal or shaft) have to be moved when the power center is advanced and retreated (perhaps once every 2 weeks), but they are not moved as often as the trailing cables. The most stationary cables are those that bring power into the mine, for instance down the borehole and from the borehole to the portable switchhouses. These are the feeder cables. A special type, designated mine power feeder, can be used for installations that may not be moved for several years. However, the use of the word feeder here is to denote a cable type rather than a function in distribution. Both feeder and portable cables can be used for feeder applications, where the cable supplies two or more major loads (38). 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. Entry, shaft, or borehole | I Switchhouse KEY Feeder or borehole cable Feeder cable Portable cable Trailing cable CONVENTIONAL UNIT 4 Shuttle car 4 Shuttle car 4 Water pump 4 Belt feeder 4 Roof bolter 4 Coal drill 4 Loading machine 4 Cutting machine CONTINUOUS UNIT 4 Shuttle car 4 Shuttle car 4 Water pump 4 Roof bolter 4 Belt feeder 4 Continuous miner LONGWALL UNIT 4 Hydraulic pump 4 Hydraulic pump 4 Face conveyor Master control 4 Face conveyor 4 Stage loader 4 Shearer Figure 8.1. —Cable distribution in underground coal mines. 183 KEY / Feeder cable 2 Portable cable 3 Trailing cable Switchhouse 3 Dragline 1" 2 \ V 3 3 Shovel , r 2 A V / Switchhouse V- -o= Water pump Lighting 1 l« 3 Power center Figure 8.2.— Cable distribution in surface coal mines. Similarly, in surface mines, the cables that feed from the switchhouses or unit substations to mobile equipment are trailing cables. Those moved only occasionally, which are not connected directly to a machine, are portable cables. Stationary (or near so) cables can be feeder or portable types. Moving the cable is a constant task both under and above ground. Some trailing cables are placed on reels or spools to facilitate moving. Prime instances of reeled cables are cables associated with the reeling devices on board shuttle cars and with mobile cable reels used in conjunction with many draglines. Trailing cables without reels are usually termed drag cables. Regardless of the application, cables are heavy and cumbersome and must often be manipulated by hand. Although the most fre- quent personnel injuries are strains, bruises, and frac- tures, cable handling is always potentially hazardous, and investigations in mines have indicated that exposed "live" conductors are a too-common occurrence. Indeed, most fatalities in cable-handling accidents are a result of rou- tine handling of unshielded cable (25). Constant handling also imposes considerable stress on the cables. While cable life is rated by manufacturers at up to 20 yr for other industrial applications, in an under- ground mine the actual cable life does not even approach this. Mine personnel have estimated the life of continuous miner cables at 8 months; roof bolter cables at 7 months; and shuttle car cables at 3 months (25), and within this lifespan the cable usually requires frequent repair. It has been estimated, for example, that 75% of the total ma- chine downtime for shuttle cars is cable related. CABLE COMPONENTS Cables are made up of three basic components: the conductor, the insulation, and the jacket, although there may also be fillers, binding, shielding and armor. In basic cable construction, the conductors are surrounded by in- sulation and the jacket covers the insulation. The design of these components is heavily dependent upon the physical stresses that the cable must withstand in the mine envi- ronment, including tension, heating, flexure, abrasion, and crushing. Hence, a discussion of typical stresses is helpful prior to describing component specifics, cable types (the various component assemblies into cables), and cable coding. High cable tensions are characteristic of both drag and reeled cables. When combined with other stresses such as flexure and twisting, tension can be very harmful to cable life. Drag cables are pulled around pillar corners, through mud, and over jagged rocks where the drag resistance is high. Consequently, a considerable force can be required to drag the cable and, thus, high tensions can develop. Machinery that utilizes cable-storage reels also fre- quently causes excessive cable tensions (13). For instance, the stored cable on the shuttle car is either payed out of the reel or spooled up into the reel as the machine is trammed. The tension required is dependent upon mine conditions, machine type, and cable size, but must be sufficiently high to prevent running over or pinching slack cable. However, if tensions become too high as a result of sudden jerks on the cable, cable and splice failures can become excessive. In addition, instantaneously high cable tensions can result in cable whipping. This is common with shuttle cars and also occurs on other machines that utilize cable reel storage devices, such as roof bolters, coal drills, and cutting machines. This whipping action is a hazard to mine personnel, who may be struck by the cables as they handle the cables or work nearby. In addition to excessive cable tensions, high cable temperatures frequently occur on machinery that utilizes cable-storage reels (6). The cable is wound on the reel, layer upon layer. Such layering prevents the cooling action of circulating airflow, and heating occurs. Consequently, the cable jacket and insulation may become softened and more susceptible to damage from cutting, tearing, and abrasion. If excessive temperatures occur, the cable jacket and insulation can actually blister or crack, becoming brittle. Thus, the physical damage caused by heating poses another hazard to mine workers who must handle the cable, especially in a wet mine environment. Another common cable stress prevalent in all mining cables is cable flexure. As with any material that is bent, internal tension and compression occur in flexed cables. These stresses cause relative movement of individual wire strands, abrading one wire against another and gradually deteriorating the conductors. Stresses fatigue the conduc- tors, making them brittle and more susceptible to further damage. Abrasion is also deleterious to cables and can have severe consequences. Cutting or tearing can occur when the cable becomes snagged or caught on rock, nails, and so on (6). Ripping or tearing of the cable jacket and insulation often results. Such damage can cause immediate cable failure, but more often than not, the damage goes unno- ticed. In a wet environment, water penetration can create a current path to the outer surface of the cable. An individual could come in contact with the wet cable several feet from a damaged area and still receive a shock that might be fatal. Another important cause of failure is cable crushing (6). This is usually the result of runovers or pinching the 184 cable with a machine frame. Here, the conductors are compressed against one another or against the machine, causing the insulation and jacket to split, as well as damaging the conductors. Even if there is no immediate failure, line-to-line or line-to-neutral faults that result in nuisance tripping of the protective circuit breakers can occur later. Water penetrating into damaged areas of the jacket can eventually work into areas of damaged insula- tion causing short circuits or a safety hazard. Conductors Line and ground currents are carried by either copper or aluminum conductors, depending on the specific char- acteristics required. Copper has high conductivity, is heavier and more flexible, but also more expensive. Be- cause of its greater flexibility, it is used in all portable mining cables. Copper cable conductors are usually composed of many fine wires combined into strands. Varying numbers of strands form the conductor. At the cable manufacturing plant, a cold-drawing process is used in which the copper rod passes through successively smaller dies to reduce its diameter (5). This process hardens the copper and makes it less flexible, so that if a soft-temper copper (strength about 24,000 psi) is required, the wire must be annealed. Con- ductors that require a high tensile strength but are not bent frequently use medium- to hard-temper copper; medium-hard is rated at 40,000 psi. Copper conductors can become annealed in service if they are used at high operating temperatures for long periods of time. In fact, copper can lose 5% of its original tensile strength in 10,000 h at 70°C (5). Cable manufac- turers should always be consulted about the capability of their products to resist annealing when installed as bore- hole or high-tension overhead cables. To prevent corrosion by insulation vulcanizing agents, copper strands are usu- ally coated or tinned with lead or tin alloys, though this reduces the surface conductivity. Aluminum conductors are also used in mines. Alumi- num is cheaper, lighter, and less flexible, and has lower conductivity than copper. Aluminum conductivity is 61% that of copper; therefore, an aluminum conductor must have a cross-sectional area 1.59 times that of copper to have an equivalent dc resistance. However, copper conduc- tors weigh 3.3 times as much as aluminum; so even though the cross-sectional area of an aluminum conductor is greater, the total weight of an equivalent-resistance alu- minum conductor is less. Poor flexibility eliminates the use of aluminum in trailing cables. Aluminum is some- times used for feeder cables because of its lower cost, but problems can arise in jointing. An improperly constructed joint can allow the formation of aluminum oxides, which increase resistance and cause heating at the connection. Extreme care must also be taken to exclude moisture from any copper-to-aluminum joints because of the potential for electrolytic corrosion of the aluminum. Conductor Sizes The cross-sectional area of conductors is important for mechanical strength and is closely related to current- carrying capacity. Since the proper capacity is both a legal requirement and a desirable practice for safe operation, an understanding is needed of the methods commonly used to specify cross-sectional areas and ampacities. In the United States, both the American Wire Gauge (AWG) (or Brown and Sharpe Gauge) and circular-mil designations (MCM) are used (1 cmil is the area of a circle that is 1 mil in diameter). The AWG specifies 38 steps or sizes between No. 36, which is 0.0050 in. in diameter, and No. 4/0, which is 0.4600 in. in diameter (5). These sizes closely conform to the steps of the wire-drawing process. Table 8.1 specifies the cross-sectional areas and equivalent circular-mil sizes for some of the AWG designations. The 38 intermediate sizes are calculated in a geometric progression relating the ratio of any diameter to the next smaller or larger by: 39 fo. 4600 0050 = 1.1229. (8.1) Table 8.1. —Conductor sizes and cross-sectional areas Conductor size Cross-sectional area cmil AWG: 22 640 0.000502 20 1,020 .000804 19 1,290 .00101 18 1,620 .00127 17 2,050 .00161 16 2,580 .00203 15 3,260 .00255 14 4,110 .00322 13 5,180 .00408 12 6,530 .00513 11 8,230 .00646 10 10,380 .00808 9 13,090 .0102 8 16,510 .0129 7 20,820 .0164 6 26,240 .0206 5 33,090 .0260 4 41,740 .0328 3 52,620 .0413 2 66,360 .0521 1 83,690 .0657 1/0 105,600 .0829 2/0 133,100 .1045 3/0 167,800 .1318 4/0 211,600 .1662 Conductor size Cross- sectional area, in 2 MCM: 250 0.197 300 .236 350 .274 400 .315 450 .353 500 .392 550 .432 600 .471 650 .510 700 .550 750 .589 800 .628 900 .707 1,000 .786 1,100 .863 1,200 .942 1,250 .981 1,300 1.02 1,400 1.10 1,500 1.18 1,600 1.26 1,700 1.33 1,750 1.37 1,800 1.41 1,900 1.49 2,000 1.57 Short-cut conductor-size approximations can be made by applying some simple rules if a table is not available. For example, the diameter will be doubled or halved by moving six sizes up or down the table. The weight, area, and dc resistance is doubled or halved by moving three gauge sizes, and they are changed by a factor of 10 over 10 gauge sizes. A convenient reference point from which to apply these rules is the No. 10 wire since its diameter is about 0.1 in., its dc resistance is nearly 1 fl per 1,000 ft, and it weighs 10^ lb per 1,000 ft. In applying these rules, it should be remembered that the outer diameter and weight of conductors depends on the stranding configura- tion, which is described below. Federal regulations require grounding conductors to have at least one-half of the cross-sectional area of the power conductors (38). When the power conductor is a No. 8 AWG or smaller, the grounding conductor should be the same size as the power conductor. The ground-check (pilot) conductor must not be smaller than a No. 10 AWG (38). 185 Conductor Stranding In order to obtain the required flexibility, mining cable conductors are made with numerous small wires rather than a single solid copper rod. These small wires are wound or laid together in strands, which are wound together in a rope in specific patterns. In a shuttle car cable, 37 wires are wound or bunched together, then 7 of these strands are spiraled together to form the conductor. Consequently, the total number of wires in this case is 259 and equals the number of strands multiplied by the number of wires in each. The cross-sectional area of a stranded conductor is defined as the sum of the area of its component wires. In the simplest terms, conductor flexibility is greatest when the largest number of small-diameter wires is used. However, a certain amount of tensile strength is also required in mining cable conductors, and the tensile strength is greatest when a small number of larger wires is used. The design of a specific cable must therefore optimize these opposing factors, while taking into account the effects of twisting and bunching. Different applica- tions obviously necessitate different configurations. The engineer must examine cable stranding specifications carefully and select the one that best suits the application. Where historical information is not available, several types should be tried to find the best performer. Flexibility is also influenced by the method of insulating the power and ground-check conductors and applying the overall jacket. Insulation Insulation of mining cables is required to withstand stress from heat, voltage, and physical abuse. The insula- tion must be specially designed not only to protect mine personnel from electric shock, but also to separate power and grounding circuits effectively. Heating affects insulating materials in different ways, depending on their chemical composition. Heating either softens insulation, causing it to lose physical strength, or causes it to age or become brittle. Conse- quently, heat can make insulation lose its original shape, tensile strength, cut resistance, elongation, and effective- ness as an insulator. The main sources of heat are the environment, related to the ambient temperature, and power (PR) loss in the cable conductors. Hence, cable heating is directly connected to the maximum current the conductors can carry safely. Cable manufacturers usually prefer to use a thermo- setting insulation. After being extended over the conduc- tors, this insulation changes chemically by vulcanizing into a material that softens very little within the rated temperature range. The most common insulating com- pounds in this group are neoprene, styrene butadiene (SBR), ethylene propylene (EPR), and crosslinked polyeth- ylene (XLP). These compounds are usually mixed with other materials to achieve improved physical and electri- cal properties. SBR is used in 600-V trailing-cable insulation. It has a high modulus of elasticity, good flexibility, and a 75°C temperature rating, and resists damage by crushing from runovers and rock falls. EPR has replaced SBR in many trailing cables because it allows the cable rated voltage to be increased to 2,000 V and the temperature rating to 90°C, while maintaining the same insulation thickness as SBR and neoprene. The EPR emergency-overload rating is 130°C, and the short-circuit rating is 250°C. XLP is also rated at 90°C for normal operation and is used in high- voltage (> 1,000 V) mine-feeder and portable strip-mining cables. XLP is a rather stiff material, however, and is not recommended for reeling applications. The cable voltage rating is closely associated with the maximum anticipated operating voltage. The most com- mon ratings for mining cables are 600 V, 2 kV, 5 kV, 8 kV, 15 kV, and 25 kV. The 5-kV, 8-kV, 15-kV, and 25-kV ratings are used primarily for stationary feeder cables and are generally not connected to mining equipment, except in surface mines. Usually, 4.16-kV distribution requires 5-kV rated cables, 7.2 kV requires 8 kV, and 12.47 kV and 13.2 kV require 15-kV ratings. The utilization voltages of 250 Vdc, 440 Vac, and 550 Vdc usually call for 600-V or 2-kV cables, and medium-voltage applications (661 to 1,000 V) need 2-kV insulation. The voltage rating of an insulation is actually based on its ability to withstand a test voltage that is many times the anticipated operating voltage, for a specified period of time. The test procedure and specifications are published in ICE A standards (19-21). Insulating com- pounds have different voltage ratings, which are usually expressed as the amount of voltage they can withstand per mil of thickness. Consequently, higher voltages can be used with any compound by increasing its thickness. Insulation thicknesses are also specified by ICEA. Insulation must resist damage from corona, particu- larly in high- voltage applications, as discussed in detail in chapter 17. The term partial discharge describes the type of corona stress imposed on cables. Partial discharges deteriorate insulation by ion bombardment and chemical action from ozone, nitrogen oxides, and nitric acid, which can occur in such voids as found between a stranded conductor surface and the insulation. Hence, insulation voids must be minimized and the insulation must resist the formation of this type of corona. ICEA standards specify corona-extinction voltage levels for insulation (19- 21). Ozone resistance is important for high-voltage cable insulation and sometimes for low voltage, and standards are again given by ICEA. Ozone is formed when electrical discharge is present in air, and it attacks compounds containing double carbon bonds, by splitting the carbon chain and deteriorating the material. Radiating cracks are a physical symptom of this occurrence. Insulation must withstand cold temperatures as well as heat, particularly in surface operations: some of the open-pit iron mines in Minnesota and Michigan, for exam- ple, have experienced temperatures as low as -50°C. Cables stored on the surface at underground minesites are also exposed to extremely low temperatures. Most prob- lems occur when a cold cable must withstand mechanical stress, such as bending or impact. Cable Jacket The main purpose of the jacket is to provide protection for the inner components and hold the assembly in the designed configuration. Jackets are not required to pass ICEA voltage withstand or insulation resistance tests, but tests for tensile strength, elongation, and aging are man- datory. Ozone and discharge-resisting jackets must also pass surface-resistivity and partial-discharge tests. Min- ing cable jackets must withstand an extensive tempera- ture range, maintaining their physical properties through- out, and furthermore, they must not deteriorate when 186 exposed to direct sunlight. Obviously, resistance to abra- sion, crushing, tearing, and impact are extremely impor- tant. Cable jackets must also be resistant to the chemical action of acid or basic mine water and hydraulic fluids, and underground coal mine cable jackets must be flame resistant. Finally, jackets must exclude moisture and be very flexible. One of the most commonly used materials for cable jackets is neoprene, a chloroprene polymer. Nitrile buta- diene and polyvinyl chloride (NBR/PVC) is also used, particularly where jacket coloring is desired. Chlorosul- fonated polyethylene (CSP) or Hypalon synthetic rubber is also used extensively, especially in combination with 90°C EPR insulation. EPR is used where extreme cold is en- countered and flame resistance is not essential. Armored cables are used in some borehole applications. Here the jacket is a heavy metallic covering that affords extra protection to the conductors. Cable Shielding The ICEA defines the practice of shielding an electri- cal power cable as confining the electric field to the inside of the cable insulation or assembly with a grounded conducting medium called a shield (19-21). Two shield types are used in practice: the conductor shield and the insulation shield. Shown in figure 8.3, the conductor shield is placed between the conductor and the insulation, and the insulation shield surrounds the insulation. Two distinct types of materials are employed in con- structing cable shields: nonmetallic and metallic. Nonme- tallic shields may consist of a conducting tape or a layer of extruded conducting compound. The tape may be made from conducting compound, be a conducting fibrous tape, or be a fibrous tape faced or filled with conducting com- pound. A typical conducting compound is carbon- impregnated rubber, which is commonly referred to as a conductive-rubber, semiconducting, or semicon shield. Me- tallic shields are nonmagnetic and may consist of a thin metal tape, wire-woven braid, or concentric serving of wires. Copper-braided shields may be made entirely of copper wires or have nylon twine in combination with Insulation Conductor shield Conductor Conductor copper wires. Nonmetallic and metallic elements may be juxtaposed to form the shield. Conductor shields are made of nonmetallic materials and are used only in high-voltage cable. The roles of this shield type are to eliminate air spaces or voids between the conductor and the insulation and to present a smooth electrode to the inner insulation surface. To be effective, it must adhere to or remain in intimate contact with the insulation under all conditions. This can substantially reduce the number of sites where partial discharge can form and helps reduce electrical stress on the insulation by uniformly distributing the electrical field about the con- ductor. The use of conductor shields becomes critical at higher operating voltages, especially 12.47 kV and above. Insulation shields can perform three principal func- tions. If placed directly over individual conductor insula- tions, along with confining the electric field caused by conductor current within the insulation, the shield helps to maintain a symmetrical radial distribution of voltage stress within the dielectric. The possibility of partial discharges is minimized by precluding tangential and longitudinal stresses, and insulation is utilized to its greatest efficiency and in the direction of highest strength. This again becomes critical at higher operating voltages. Insulator shields also provide a continuous capacitance to ground for the conductor along its entire length. The uniformity is important in terms of transients on the power system, and this is discussed in chapter 11. The third function of insulation shields is the most important for mining in view of the extensive handling of cables: reducing the hazard of electric shock. A major cause of electrical fatalities in mining has been workers' cutting into energized unshielded cables, for instance, during repair. Another source has been handling of ener- gized unshielded cables with damaged jacketing and insu- lation or splices (the spot where a cable has been repaired). An insulation shield can be thought of as a safety barrier to penetrating metallic objects. If the percent of coverage of the shield over the insulation is high enough and its impedance is low enough, any metallic object compromis- ing the conductor insulation will establish a fault between the power conductor and the grounded shield, with suffi- cient current to trip the ground-fault protective circuitry. Damage to insulation and jacketing, such as a pinhole, that would cause a handling danger to unshielded cable also creates a probable ground fault in cables with insu- lation shields. An individual touching the penetrating metallic object or handling the damaged shielded cable should be safe from electrocution. Insulation shields are usually metallic. Recently, how- ever, semicon insulation shields for trailing cables have found application in the United Kingdom, Australia, and to a lesser extent, the United States. This is to take advantage of semicon flexibility, especially in reeled-cable situations. Insulation #5|r Conductor «titi_ _ Insulation shield Insulation Figure 8.3.— Shield types. CABLE TYPES An identifying code, related to standard specifications designated by ICEA, is embossed on the cable throughout its entire length. The code includes any approval number for flame resistance by the Mine Safety and Health Administration (MSHA) and approval by the Common- wealth of Pennsylvania (indicated by the letter P preced- ing the MSHA approval number). MSHA approval is mandated for cables in underground coal mines, and the 187 Pennsylvania approval is necessary for cables used in underground coal mines in that State. The code includes the term nJc where n is the number of power conductors in the cable, an approved voltage designa- tion, and letters describing the cable type. Table 8.2 summa- rizes the meaning of the letters used in the code, and table 8.3 presents the codes for typical cable types used in mining. Figures 8.4, 8.5, and 8.6 correspond to table 8.3 for un- shielded round, unshielded flat, and shielded cable configu- rations, respectively, and detail the cable components as seen in cross section. Photographs of actual mining cables are provided in figures 8.7, 8.8, and 8.9 and show both side and cross-sectional views. Figures 8.10 and 8.11 are similar to figures 8.1. and 8.2 and show common applications of cable types in mine power systems. Figure 8. 7 A is a single-conductor cable insulated for use at 600 V. This specific cable is not widely used. However, it has found application on twin-reel dc shuttle cars and small locomotives with reels; therefore, it must be highly flexible. Single-conductor cable similar to that shown is used extensively for connections inside power equipment, and a typical voltage rating is 15 kV for system voltages less than that level. The most common dc shuttle-car cables are types W and G, figures 8.8A and 8.8.B, respectively. The flat con- figuration is used since it allows an increased length on cable reels and is less susceptible to runover damage than round cables. The type W is used where diode grounding is allowed in lieu of a separate grounding conductor. Because shuttle car cables are damaged frequently, type W is preferred by some mine operators since it is easier to repair (splice). Flat cable types employed for ac shuttle cars are shown in figure 8.8C and 8.8D. The three power conduc- tors are separated by two grounding conductors in figure 8.8C and by one grounding and one ground-check conduc- tor in figure 8.8D. These cables are also used on other equipment with reels, such as cutting machines and drills. Table 8.2.— Letters used in alphabetic cable code Code Meaning Comments G Contains uninsulated grounding conductor(s) Common on low-voltage ac systems but used on dc systems where grounding conductors are needed. W Without uninsulated grounding conductor(s) Typical on dc diode-grounded systems but 1 insulated power conductor may be used as a grounding conductor. GC Includes insulated ground-check (pilot) conductor. Used where pilot-type ground-continuity monitoring is required, usually replaces 1 grounding conductor of type G cable. SH Shielded cable None. D Multiple insulation shields Shields surround each individual power-conductor insulation. C 1 insulation shield 1 shield surrounds entire cable assembly just inside jacketing. MP Mine power feeder None. Table 8.3.— Codes for typical cables used in mining. Code Components W Contains 2, 3, or 4 insulated power conductors G Contains 2 or 3 insulated power conductors and 1 to 3 uninsulated grounding conductors. G-GC Contains 3 insulated power conductors, 1 or 2 uninsulated grounding conductors, and 1 insulated ground-check conductor. G + GC Contains 3 insulated power conductors, 3 uninsulated grounding conductors, and 1 insulated ground-check conductor. SH-D Contains 3 shielded insulated power conductors, 2 or 3 uninsulated grounding conductors. SH-C Contains 3 insulated power conductors, 2 or 3 uninsulated grounding conductors, assembly shielded. SHD-GC Contains 3 shielded insulated power conductors, 1 or 2 uninsulated grounding conductors, and 1 insulated ground-check conductor. SHD + GC Contains 3 shielded insulated power conductors, 3 uninsulated grounding conductors, and 1 insulated ground-check conductor. SHC-GC Contains 3 insulated power conductors, 1 or 2 uninsulated grounding conductors, 1 ground-check conductor, assembly shielded. MPF Contains 3 shielded insulated power conductors, 3 uninsulated grounding conductors. MP-GC Contains 3 shielded insulated power conductors, 2 uninsulated grounding conductors, and 1 ground-check conductor. 1 Although not presently available, 2/C cable design for dc systems is possible. Comments See table 8.2. Flat or round cross section. Grounding conductors are placed in the interstices between the power conductors. Flat or round cross section. Ground-check conductor replaces 1 grounding conductor of type G cable. Flat or round cross section. Presently, for ac systems only. 1 Similar to round 3/C type G cable but has ground-check conductor in cable center. Insulation shields about each individual conductor, grounding conductors contact shields. High-voltage cables usually have conductor shields. Round or flat cross sections. Presently for ac systems only. 1 A flexible portable cable. Shielding encloses all conductors and is located just under . jacketing. Grounding conductors should contact shield. Round or flat cross sections. Presently for ac systems only. 1 A flexible portable cable. Ground-check conductor replaces 1 grounding conductor of type SH-D cables. Round or flat cross section. Presently for ac systems only. 1 A flexible portable cable. Similar to round 3/C type SH-D cable but has ground-check conductor in cable center. Ground-check conductor replaces 1 grounding conductor of type SH-C cables. Round or flat cross section. Presently for ac systems only. 1 A flexible portable cable. Similar to round SH-D cable. Designed for relatively stationary high-voltage feeder applications. Similar to round SHD-GC cable. Designed for relatively stationary high-voltage feeder applications. 188 Insulated ground -check conductor Type W Type G Type G-GC Uninsulated grounding conductors Insulated ground -check conductor Type G+GC Figure 8.4.— Cross sections of round unshielded mining cables. Fillers, may not be needed if conductor insulation fills voids Jacket 2/C type W 2/C type G 3/C type G Figure 8.5.— Cross sections of flat unshielded mining cables. 3/C type G-GC Conductor insulation Conductor shield, copper braid if SHD-GC, metallic tape if MP-GC Grounding conductor, contacts shield Jacket Type SHD-GC or MP-GC Grounding conductor may contact shield Type SH-C Conductor shield Grounding conductors contact shield Jacket Flat type SH-D Jacket , Grounding conductors may contact shield Flat type SHC-GC Figure 8.6.— Cross sections of some shielded mining cables. 189 (A) 2/C typeW, 600V (A) 1/C.600V (B) 2/C type G, 600 V (B) 3/C type G-GC, 2,000 V (C) 3/C type G, 600 V (C) 3/C type G+GC, 2,000 V Figure 8.7.— Round unshielded mining cable. (Courtesy Anaconda Ericsson Co.) (D) 3/C type G-GC, 600 V Figure 8.8.— Flat unshielded mining cables. (Courtesy Anacon- da Ericsson Co.) (A) 3/C type SHD-GC, 2,000 V (C) 3/C type SHD-GC, 15 kV (B) 3/C type SHD+GC, 2,000 V (D) 3/C typeMP-GC, 15 kV Figure 8.9.— Round shielded mining cables. (Courtesy Anaconda Ericsson Co.) 190 KEY Borehole cable: 3/Ctype MP-GC; 5, 8, 15, or 25 kV 3/C type MP-GC, SHD-GC, or SHD+GC; 5,8, 15, or 25kV 3/C type SHD-GC or SHDtGC; 5,8, !5,or25kV 3/C type G,G-GC,orG + GC; 2kV 3/C typeGorG-GC, flat, 2kV 2/Ctype WorG, flat, 2 kV Entry, shaft, or borehole I Switchhouse k Switchhouse CONVENTIONAL UNIT 250-Vdc Shuttle car 250-Vdc Shuttle car 550-Vac Water pump 550 -Vac Belt feeder 550-Vac Roof bolter 550-Vac Coal drill 550-Vac Loading machine 550-Vac Cutting machine CONTINUOUS UNIT (550 Vac) Shuttle car Shuttle car Water pump Roof bolter Belt feeder Continuous miner LONGWALL UNIT (550 Vac) 4 Hydraulic pump 4 Hydraulic pump 4 Face conveyor ^ ° W ° 2 kV 2 kV 5 kV 8 kV AWG: 8 0.838 0.878 0.034 — — — 7 .665 .696 .033 — — — 6 .528 .552 .032 0.038 0.043 — 5 .418 .438 .031 .036 .042 — 4 .332 .347 .031 .035 .040 0.043 3 .263 .275 .031 .034 .039 .042 2 .209 .218 .029 .033 .038 .040 1 .165 .173 3 .030 .033 .036 .039 1/0 .128 .134 .029 .032 .035 .037 2/0 .102 .107 .029 .031 .034 .036 3/0 .081 .085 .028 .030 .033 .035 4/0 .065 .068 .027 .029 .032 .034 MCM: 250 .055 .057 .028 .030 .031 .033 300 .046 .048 .027 .029 .031 .032 350 .039 .041 .027 .029 .030 .032 400 .035 .036 .027 .028 .030 .031 500 .028 .029 .026 .028 .029 .030 600 .023 .024 .026 .027 .028 .030 700 .020 .021 .026 .027 .028 .029 800 .018 .019 .025 .026 .028 .029 900 .016 .017 .025 .026 .027 .028 1,000 .014 .015 .025 .026 .027 .028 1 Criteria: a. Sizes 8 to 1 based on tinned copper 94.16% conductivity. b. Sizes 1/0 AWG and larger based on tinned copper 96.16% conductivity. c. Resistance increased by increments per ASTM B-172, Note 7 (3), to compensate for stranding factor. d. Skin effect calculated according to Arnold's Table, National Bureau of Standards Monograph 125 (29). e. Nominal cross-sectional areas. 2 Criteria: a. Based on conductor dimensions given for class-H rope-lay conductors in table 2.5 of ICEA S-19-81 (21). b. Extruded-strand shield thickness, 0.015 in. c. Insulation thickness according to nominals given in Interim Standard 6 to ICEA S-68-516 (79). d. Diameter adder of 0.075 in to allow for semiconducting tape and metal-braid shield. 3 Deviation from normal progression due to changes in insulation. NOTE.— Dash indicates cable is not made. 0.044 — .042 0.046 .040 .044 .039 .043 .038 .041 .036 .040 .036 .039 .035 .038 .034 .037 .033 .036 .032 .035 .032 .034 .031 .033 .030 .033 .030 .032 .030 .032 201 Table 8.15.— Resistance and reactance of mine-power-feeder cable Conductor size R (ac), 1 Q/Mft, 90°C X L (60 Hz), 2 Q/Mft MP-GC, 5kV MP-GC, 8kV MP-GC, 15 kV AWG: 6 5 4 3 2 1 1/0 2/0 3/0 4/0 MCM: 250 300 350 400 500 600 700 800 900 1,000 1 Criteria: a. b c 0.510 .404 .321 .254 .201 .160 .127 .101 .080 .063 .054 .045 .039 .034 .027 .023 .020 .017 .016 .014 0.041 .040 .038 .037 .036 .035 .034 .033 .032 .031 .030 .029 .029 .029 .028 .028 .027 .027 .027 .026 0.044 .042 .041 .039 .038 .037 .035 .034 .033 .032 .031 .031 .039 .030 .029 .029 .028 .028 .027 .027 0.042 .041 .039 .038 .036 .035 .034 .034 .033 .032 .031 .031 .030 .030 .029 .029 Based on bare copper 100% conductivity. Nominal cross-sectional areas. Resistance increased by increments per ASTM B-8, Note 3, to compensate for stranding factor, d. Skin effect calculated according to Arnold's Table, National Bureau of Standards Monograph 125 (29). 2 Criteria: a. Based on conductor dimensions given for class B concentric stranded conductors in table 2.2 of ICEA S-19-81 (21). b. Extruded strand shield thickness, 0.015 in. c. Insulation thickness according to nominals given in Interim Standard 5 to ICEA S-68-516 (79). d. Diameter adder of 0.033 in to allow for semiconducting tape and copper-tape shield. NOTE.— Dash indicates cable is not made. EXAMPLE 8.4 Distribution cables for a segment of an under- ground coal mine must be sized. A sketch of the situation is provided in figure 8.14 where the loads are two continuous mining sections. Voltages given are line to line. In-mine measurements and analysis of identical section equipment working in similar condi- tions have shown an effective current demand of 58 A with 0.8 lagging power factor at the power-center primary, when the continuous miner is cutting and loading. Maximum ambient temperature is 20°C. In a detailed study, the substation transformer impedance must be included. For the sake of demonstration, however, the 7,200-V line-to-line voltage at the substa- tion secondary will be assumed constant. The recom- mendation for allowable voltage drop is 10% across the distribution system. As the impedances of the feeder and portable cables must be known to make the calculation, a good place to start is to estimate line currents and make an initial cable selection by am- pacity. From the given information, Ii = I 2 = 53 A. I 3 is related to I x and I 2 but is not necessarily equal to their sum, because of the diversity of mining Substation 5MVA 7% reactance 69 kV : 7.2 kV 10,000-ft feeder cable / 1,000-ft portable cable It Bus representing double-breaker switchhouse ( Power centers for continuous mining sections, J T y ZQQ-M primaries 1,000-ft portable cable Figure 8.14.— Simplified one-line diagram for situation described in example 8.4. operations. Chapter 4 presented the concept of de- mand factor (DF) where using a value from 0.7 to 0.8 is considered reasonable for mining sections: 0.8 corresponding to two sections and 0.7 to four or more sections. Therefore, or I 3 = I 3 = DFCIi + I 2 ) (8.5) (0.8X53 + 53) = 84.8 A. A 7,200-V system requires the use of 8-kV shielded cables, and the corrected ampacity for No. 6 AWG from table 8.7 or 8.8 and table 8.9 is ampacity = (93X1.18) = 110 A. This means that on a current basis the size is ade- quate for all distribution cables. Considering the pref- erence of the coal mining industry for using only portable cables for flexibility, ground-check conductors for ground-continuity monitoring, and 90°C insula- tion, an SHD-GC cable is indicated. Table 8.14 can be consulted for its impedance. It can be seen in the table that No. 4 AWG is the smallest 8-kV SHD-GC porta- ble cable readily available. Hence, a No. 4 AWG will be tried. Its impedance per 1,000 ft is Z cable = 0.347 + J0.043 Q = 0.35 \TA°_ fl. Referring to figure 8.14, the voltage drop across the distribution line conductors to either power center is (taking the power-center voltage as the reference phasor): V d = I 3 [10(Z cable )] + \ [l(Z cable )] V d = (84.8[ -36.9° X3.5|7.1°) + (53|-36.9°X0.35|7.1°) or V d = 296.81 -29.8° + 18.6 | -29.8° = 3151-29.8° V As per-phase analysis is required to compare this drop with that allowed, the line-to-neutral voltage of the distribution system is used, or V ln = 7,200 = 4,160 V. The allowable voltage drop is V d allowable = 0.1(4,160) = 416 V. 202 Therefore, the 315-V drop using No. 4 AWG SHD- GC cables is tolerable. If the voltage drop were not acceptable, an increase in cable size would lower the impedance and the drop. This simple example had equal cable lengths to the loads, and currents operating at the same phase angle. It should be noted that typical mining sys- tems have many more loads, varying cable length, varying load power factors, and so forth, and the complexity of hand calculations will increase sub- stantially. Per-unit techniques are a tremendous help, but computer analysis is a much more efficient way to solve such problems. Nonetheless, the tech- niques shown here are useful for partial sizing or spot-checking distribution cables. strength because copper begins to elongate at that point. Federal regulations acknowledge the problem of exceeding the cable mechanical strength and mandate a minimum trailing-cable size for underground coal mine face equip- ment: No. 4 AWG for two-conductor dc cables and No. 6 AWG for three-conductor ac cables (38). Short-Circuit Currents The emergency-overload currents that copper conduc- tors can withstand without serious insulation damage are shown in the graph in figure 8.15 (5). If the anticipated short-circuit currents are greater than those shown in the graph for the initial selection of conductor size, a larger conductor or a better grade of insulation should be chosen. Chapter 10 covers the calculation methods. Cable Mechanical Strength CABLE INSTALLATION AND HANDLING The tensile load on the cable should be determined from measurements in the mine, bearing in mind the problems discussed at the beginning of this chapter. The power-conductor breaking-strength data in table 8.16 should then be consulted to assure that the conductor size is large enough to carry the tensile load (5). Two things must be considered when using this table. First, ground- ing and ground-check conductors should not support any of the tensile load, so the overall cable breaking strength should include only the sum of the power-conductor val- ues. Second, the working tension, especially in reeling applications, should not exceed 10% of the breaking Table 8.16.— Solid-wire breaking strength Conductor Hard— Medium— Soft- size, 65,000 psi 55,000 psi 40,000 psi AW G lb kg lb kg It) kg 4/0 8,143 3,693.6 6,980 3,166.1 5,983 2,713.8 3/0 6,720 3,048.1 5,666 2,570.1 4,744 2,151.8 2/0 5,519 2,503.4 4,599 2,086.1 3,763 1,706.9 1/0 4,518 2,049.3 3,731 1,692.4 2,985 1,354.0 1 3,688 1,672.8 3,024 1,371.7 2,432 1,103.1 2 3,002 1,361.7 2,450 1,111.3 1,928 874.5 3 2,439 1,106.3 1,984 899.9 1,529 693.5 4 1,970 893.6 1,584 718.5 1,213 550.0 5 1,590 721.2 1,265 573.8 961.5 436.1 6 1,280 580.6 1,010 458.1 762.6 345.9 7 1,030 467.2 806.7 365.9 605.1 274.5 8 826.1 374.7 644.0 292.1 479.8 217.6 9 660.9 299.8 513.9 233.1 380.3 172.5 10 529.3 240.1 410.5 186.2 314.0 142.4 11 423 191.9 327 148.3 249 112.9 12 337 152.9 262 118.8 197 89.4 13 268 121.6 209 94.8 157 71.2 14 214 97.1 167 75.7 124 56.2 15 170 77.1 133 60.3 98.6 44.7 16 135 61.2 106 48.1 78.0 35.4 17 108 49.0 84.9 38.5 62.1 28.2 18 85.5 38.8 67.6 30.7 49.1 22.3 19 68.0 30.8 54.0 24.5 39.0 17.7 20 54.2 24.6 43.2 19.6 31.0 14.1 21 43.2 19.6 34.4 15.6 24.6 11.2 22 34.1 15.5 27.3 12.4 19.4 8.80 23 27.3 12.4 21.9 9.93 15.4 6.99 24 21.7 9.84 17.5 7.94 12.7 5.76 25 17.3 7.85 13.9 6.30 10.1 4.58 26 13.7 6.21 11.1 5.03 7.94 3.60 27 10.9 4.94 8.87 4.02 6.33 2.87 28 8.64 3.92 7.02 3.18 4.99 2.26 29 6.97 3.16 5.68 2.58 4.01 1.82 30 5.47 2.48 4.48 2.03 3.14 1.42 Cables must be installed and handled correctly in order to minimize damage from tension, bending, twist- ing, physical wear, cold, heat, and chemical reaction. Cable maintenance costs can be reduced, cable life im- proved, and safety enhanced by proper installation and handling. In other words, the considerable amount of 100 80 60 50 40 30 20 1 0.8 .6 .5 .4 .3 .2 / / - A i &£>/ i> J #&m y>7 \ V ovr^V y°/l^/T &X C >/ /*■ /:^/»%/°/^/ C e- / Sj/&£ v /<&- y. 'WJ5TSS \yf N %Y*9, "\ 7 °X r ,o/& ?/ yf w* * • iyl l£v\ \ T^ «V^^ D— ^ '"'^Am x^V\\ ^fSucJi ■*^-m/.^T iH ^^ ^S^-. *'& n W§M \\ffi K^ Figure 8.19.— Effective method for removing unwanted insulation. 209 Modified crowsfoot Figure 8.20.— Staggering splice connections. Connectors A variety of connectors (fig. 8.21) and connector crimp- ing tools are available. It is generally recommended that lapped-joint connections be used where maximum tensile strength is desirable, as in shuttle car cables. Research has shown that the modified crowsfoot connection, when properly installed, can restore 80% to 100% of the original tensile strength for Nos. 6, 4, and 2 AWG conductors, the smaller conductors being easier to restore. The modified crowsfoot connection offers additional advantages of axial symmetry (no mechanical couple) and a small profile (an important consideration with multiconductor cables). The lap joints, being shorter than butt joints, are better for reeling applications since repeated flexure on a long connection might accelerate fatigue failures. The lap joints generally outperform butt connections in tensile strength, and Bureau-of-Mines-sponsored research has shown that restoring tensile strength is probably more important than restoring high flexibility to shuttle car cables. Either way, the lap connections are superior. A major consideration in obtaining high tensile strength is the use of the proper crimping tool for a given connector. Furthermore, tools that reduce or eliminate operator judgment tend to provide the best repeatability, since overcrimping as well as undercrimping can reduce tensile values. The lap connection has also been recommended for restoring the grounding conductors. In this case, it has been suggested that the connection be a little forgiving and allow the grounding conductors to slip slightly inside the connec- tor should the cable undergo excess tension. This would cause the power conductors to take all the tension and would perhaps prevent the grounding conductors from being ten- sioned, so that they would be the last to fail. Although this concept has not been verified, it may have some merit (assuming of course that a good electrical connection is maintained and that otherwise the grounding conductor might not extend sufficiently under tension). An important consideration in selecting and install- ing connectors in reeled cables is awareness of the connec- tor profile after installation. Bulky connectors with abrupt edges are more difficult to insulate effectively, simply because they tend to cut through the insulation materials with repeated cycling under normal operations. These connectors can also cause excess pressure and fatigue on adjacent grounding conductors, which are uninsulated and somewhat less protected from mechanical abuse. Although generally unsatisfactory for related applica- tions, the butt connection is effective for larger portable Parallel Slide connector into place and crimp Full crowsfoot Butt (where tensile strength is not most important consideration ) Figure 8.21.— Examples of popular connectors and connec- tions used in splices. cables such as those used for continuous miners, because it offers the least bulk. Here it does not need to withstand the repetitious flexing so often experienced by the smaller size cables. Reinsulating Because of the repeated bending stresses, reinsulat- ing procedures require special attention in portable ca- bles. The key is to provide a flexible joint and seal where the new insulation contacts the original cable insulation. As shown in figure 8.22, this is best accomplished using soft rubber tape that completely fills the volume and laps over the original insulating material. The lap is important since a tape fill that only butts to the insulation is almost sure to separate after very little flexing. Where it is desirable to use slit tubes as part of the reinsulating procedure, soft tape is recommended underneath and over the tubing. Soft rubber tape alone will not hold up under repeated cable flexing. Therefore it is further recommended that tougher vinyl tape be applied over the rubber tape. The vinyl tape accomplishes two objectives: it restrains the soft tape, thus preventing it from squeezing and extruding from its intended area, and it allows the reinsulated connections to slide relative to one another and the 210 Cable insulation Connector Vinyl electrical tape Figure 8.22. — Reinstating power conductors with soft rub- ber tape. grounding conductors with minimum wear. The vinyl tape can also be used to bind the multiple conductors together for maintaining positioning and limiting excess relative motion. A single-width wrap of tape near the middle of the splice area is generally sufficient. Care should be taken not to use too much vinyl tape over the splice area, since the final splice covering is generally intended to bond to the inner parts and the vinyl can in some cases make the subsequent adhesive bond less effective. In the case of heat-shrink splices, the conductor insu- lations are also made of heat-shrinkable tubing, and the tubes must be slipped onto the conductors before the connector is applied. When shrinking the tubes with a heat source, care must be taken to avoid overheating or rupturing the insulation on the sharp connector edge, and so forth. After heating, the installer should inspect the work to ensure that the adhesive has sealed the sleeve to the original insulation material. This is especially impor- tant for flat cables where the insulation cross sections are not always smoothly continuous. The heat-shrink insula- tion tubes provide a generous lap over the original insu- lation and are usually tough and resistant to rubbing wear inside the splice. Shielded cables require complete shield replacement over the conductor insulation. This process is similar for all cables but requires more care in high-voltage splices and will be covered later. Rejacketing The outer splice covering provides protection for the more delicate inner splice components and serves basi- cally the same purpose as the rugged cable jacketing. It is important that it be tough and flexible and at the same time maintain an acceptable bond to the original jacket- ing material. Of principal concern is a splice condition generally termed end lipping, the result of the splice- covering ends' pulling away from the cable jacket. When this occurs, contaminants such as fine solids and water can enter the splice and contribute to failure or an unsafe condition. The causes of end lipping are combinations of poor adhesive bonds, discontinuities and dissimilar mate- rials, or simply physical wear as a result of the normal mining process. The amount of end lipping will vary depending on the types of covering used and the conditions to which it is exposed. Various attempts have been made to provide splicing products that resist end lipping, with varying degrees of success. The general recommendation is to prevent occur- rence by making every effort to clean the cable surface where the adhesive bond is to be made. As a minimum, any soiled surfaces should be wiped with a suitable solvent and abraided with nonconductive emery material to reveal a fresh bonding surface. It should be noted that newer cable jackets can be more difficult to bond simply because waxes from the manufacturing process are often on the jacket surface. In general, the heat-shrink sleeves are good abrasion- resistant coverings. However, it should be noted that they are usually stiffer and sometimes require more attention to obtain a good and lasting bond to the cable jacket. Furthermore, a heat-shrink sleeve can take on a thermal set, for example, if it is allowed to cool in a curved position on a reel and then is later unreeled while still cool. The cold splices are generally quite flexible, but end lipping can result from bending and scuffing on various machine parts. Major cable and splice wear usually occurs during contact with the machine and its spooling mechanism when the relative motion is at a maximum. It is normal practice to tape down the ends of the splice coverings. This can help to reduce end lipping and can also prevent foreign matter from entering an already lipped end. Regular inspection and renewal of the end taping is a must, since abrasive wear and cutting on machine parts can quickly destroy even well-applied end tapes. The use of exposed soft rubber tapes is considered poor cable repair practice. The softer rubber tapes can provide good moisture seals but should be protected with an overcovering of tough vinyl tape. This vinyl tape will help contain the rubber tape, and the lower friction will give better wear characteristics. High- Voltage Cable Splices When splices are required on high-voltage cables in underground mines or in surface mines, problems are introduced by the presence of shields and semiconducting layers. The high voltage means that care must be taken to achieve an excellent splice, that is, one that closely ap- proximates the qualities of the original cable. The splicing procedure is basically the same as that just covered, but the cable insulation and jacket are usually tapered as shown in figure 8.23. Tapering is performed to improve the bond, increase the leakage path length, and lessen the chance of a direct vertical path to a ground plane. Extra care and skill are necessary as any damage to the insulation during splicing, such as a small cut, will cause more rapid dielectric failure at higher voltages. In the same context, a small protrusion such as a sharp edged connector or loose wire will be a more noticeable failure initiator as the voltage increases. The presence of semiconductive tape and braid or tape shielding in cables requires extra caution. The shielding system must be separated from the conductor insulation in such a way that residue on the insulation from the semiconducting tape is completely removed before the conductor insulation is reapplied. In addition, the wires from a braid shield must not protrude into the insulation. The shield must be replaced completely, and the grounding conductor must be placed in intimate contact with the shield. Splice Inspection A recommendation for improving splice performance is to inspect splices on a regular basis and use the information to institute new procedures or even new splice kit designs. An opportune time for doing this is before shipping an extensively damaged cable to a repair shop for vulcanized repairs. When the cable is idle and quite 211 Outer protective cover tapes Cable shielding Grounding lead if necessary Figure 8.23.— Typical taped splice in high-voltage shielded cable. accessible, perhaps stored in a supply yard, old splices can be cut out and scrutinized. Just the simple process of slitting the old splice lengthwise using a sharp linoleum knife can provide good information regarding insulation procedures, wear characteristics, effectiveness of bonds, and so forth. Electrical tests and tensile evaluations can be made on the insulations. Samplings of this type can readily provide extensive data on splices with varying amounts of in-service time. TROLLEY SYSTEMS The conductors that provide power for electric track haulage systems form a major part of the power- distribution system in many underground mines. The trolley system is a potential hazard for fires, ignition of methane, and shock since it utilizes uninsulated conduc- tors. The danger in underground coal mines is greater than that in surface mines because of limited space and the presence of methane, However, all mines that utilize trolley conductors can benefit from proper design, selec- tion, and installation of the system components. Several conductors are used in the trolley circuit: trolley wire, feeder cable, rail-bond cable and steel track rails. The trolley wire supplies power directly to a rail- mounted vehicle, such as a mine locomotive, through a collector called a shoe or harp. The trolley wire and collector connection can cause frequent severe arcing, which may damage either part and cause an obvious ignition hazard. Proper positioning of the trolley wire, particularly at curves and switches, correct holding force on the collector, and the required amount of lubrication are necessary to minimize arcing. A feeder cable supplies power to the trolley wire. Consequently, both must be sized properly to provide enough current-carrying capacity yet minimize heating and voltage drop. In addition, rectifiers must be positioned at adequate intervals to supply the proper voltage to the feeder. The current return path utilizes the steel rails, which must have adequate conductivity to minimize the total system resistance. Rails are laid in segments, and the connections between them can loosen or the rails could break; hence, rail-bond cable is installed to maintain continuity. Rail-bond cable is attached at each rail joint, and as a further precaution, between the two rails at specified intervals (cross bonds). Trolley Wire The trolley-wire conductor used in mines is hard- drawn copper, but brass is available for high-speed surface transportation. Round, grooved, figure 8 and figure 9 (deep-section) wire shapes, shown in figure 8.24, are avail- able (31). At one time, round wire was prevalent, but the clamps necessary to support it caused the collector to jump and arc, so it was replaced with the figure 8 shape. Additional problems occurred with the figure 8 because it twisted and kinked when being reeled and unreeled dur- ing installation, and it frequently pulled out of hangers on curves. Consequently, the grooved type was developed and, together with the figure 9, has almost completely replaced the round and figure 8 shapes. Figure 9 and deep-grooved shapes are almost mandatory with a 350-MCM size and above, because these sizes require large splices and fit- tings and the widths are too large for proper tracking of the collector. The upper section of the wire, to which the support clamp attaches, has the same width dimension whether the wire is grooved, figure 8 or figure 9. Table 8.19 provides the necessary specifications for correct wire size selection (31). The most common wire is 350 or 400 MCM (both often called 6/0) figure 9. Trolley Feeder In order to reduce voltage drop and supply the neces- sary current, a feeder cable, which is uninsulated and stranded, is hung alongside the trolley wire. Both alumi- num and copper feeders are used, and their size depends on the load drawn by the track vehicles and the voltage regulation desired. Common sizes are 1,000 MCM copper or 1,590 MCM aluminum. Tables 8.20 and 8.21 specify copper feeder data (31 ). As noted in the tables, feeder can be purchased with a weatherproof jacket. Supports, Lubrication, and Turnouts As shown in figure 8.25, the feeder cable and trolley wire can be hung side by side to gain additional support clearance. The feeder can also be used as a messenger to increase the support-bracket spacing, as shown in figure 8.26. In this configuration, a cushioning effect is provided for the trolley wire since the wire is free to flex under pressure. Typical brackets for supporting trolley and feeder are shown in figure 8.25 and 8.26. The amount of deflection or sag between supports can be calculated by D = 3WL 2 2T (8.9) where D = sag, in, W = weight, lb/ft, L = distance between supports, ft, and T = tension, lb. Since the figure 9 350-MCM conductor has a breaking strength of 12,000 lb, it can safely be tensioned to 1,200 lb, which is 10% of the breaking strength. This will reduce sag and keep the wire straight and level. The dead-end hooks and turnbuckles shown in figure 8.27 are used to install tension in the wire. The maximum spacing recommended for roof- mounted support for a semicatenary installation (fig. 8.26) is 20 ft. Direct suspension (fig. 8.25) spacing should be less than 15 ft. Table 8.22 gives support spacings on curves (31). When selecting proper support types and spacings, 212 ROUND 0.325" 0.365" 5 I/O AWG 2/0 AWG 3/0 AWG 4/0 AWG 0.548" 300 MCM GROOVED FIGURE 8 0.106 0.312 1/0 AWG FIGURE 9 DEEP- SECTION GROOVED m 0.429 3/0 AWG 0.196 0.108'i 0.222 O.BO 1 "*^ 0.400" 3/0 AWG •« -*- 0.496 350 MCM COPPER 0.150' L ^ Figure 8.24.— Trolley-wire cross sections. Table 8.19.— Trolley-wire specifications Nominal Cross-sectional Weight dc resistance (volts drop per amp Minimum Minimum Elongation Type of wire size, AWG or at 20 >C) tensile strength, breaking load, within 10 in, MCM Nominal Actual Ib/Mft lb/mi fiorV QorV psi lb % MCM MCM in 2 /Mft /mi Round, hard-drawn copper, 97.16% conductivity 1/0 105.6 105.6 0.0829 319.5 1,687 0.1011 0.5339 54,500 4,518 2.40 2/0 133.1 133.1 .1045 402.8 2,127 .08021 .4235 52,800 5,519 2.80 3/0 167.8 167.8 .1318 507.8 2,681 .06362 .3359 51,000 6,720 3.25 4/0 211.6 211.6 .1662 640.5 3,382 .05045 .2664 49,000 8,143 3.75 300 300.0 300.0 .2356 908.0 4,794 .03558 .1879 46,400 10,930 4.50 Grooved, hard-drawn copper, 97.16% conductivity 2/0 133.1 137.9 .1083 417.6 2,205 .07741 .4087 50,200 5,437 2.80 3/0 167.8 167.3 .1314 506.4 2,674 .06380 .3369 48,500 6,373 3.25 4/0 211.6 212.0 .1665 641.9 3,389 .05035 .2659 46,600 7,759 3.75 300 300.0 299.8 .2355 907.6 4,792 .03560 .1880 44,200 10,410 4.50 350 350.0 351.2 .2758 1063.0 5,612 .03040 .1605 42,800 11,800 4.50 Figure 8, hard-drawn copper, 97.16% conductivity 1/0 105.6 105.6 .0329 319.5 1,687 .1011 .5340 51,800 4,294 2.40 2/0 133.1 133.1 .1045 402.8 2,127 .08021 .4235 50,200 5,246 2.80 3/0 167.8 167.8 .1318 508.0 2,682 .06361 .3359 48,500 6,392 3.25 4/0 211.6 211.6 .1662 640.5 3,382 .05044 .2663 46,600 7,745 3.75 350 350.0 350.1 .2750 1 ,060.0 5,597 .03049 .1610 42,800 11,770 4.50 Figure 9 deep- section, hard-drawn copper, 97.16% conductivity 350 350.0 348.9 .2740 1,056.0 5,576 .03060 .1616 42,800 11,730 4.50 400 400.0 397.2 .3120 1 ,202.0 6,347 .02687 .1419 41,300 12,890 4.50 213 Table 8.20.— Characteristic data for solid copper feeder cable Conductor size, AWG Section area Overall diameter, in Weight, Ib/Mtt Bare wire br strength, Baking lb cmil in 2 m 2 Bare Weatherproof Bare Insulated Hard drawn Annealed 0000 211,600 0.1662 107 0.4600 0.6163 641 767 8,143 5,320 oon 167,800 .1318 85.0 .4096 .5659 508 629 6,722 4,220 00 133,100 .1045 67.4 .3648 .5211 403 502 5,519 3,340 105,500 .08289 53.5 .3249 .4812 320 407 4,517 2,650 1 83,690 .06573 42.4 .2893 .4456 253 316 3,688 2,100 ? 66,370 .05213 33.6 .2576 .3826 210 260 3,003 1,670 3 52,640 .04134 26.7 .2294 .3544 159 199 2,439 1,325 4 41,740 .03278 21.2 .2043 .3293 126 164 1,970 1,050 5 33,100 .02600 16.8 .1819 .3069 100 135 1,591 880 6 26,250 .02062 13.3 .1620 .2870 79 112 1,280 700 7 20,870 .01635 10.6 .1443 .2693 63 NA 1,030 550 a 16,510 .01297 8.37 .1285 .2535 50 75 826 440 NA Not available. Table 8.21.— Characteristic data for stranded copper feeder cable „ .. . Number of Overall diameter, ... . . . ,. ,,,,. Conductor Cross-sectional area wjres jn Weight, Ib/Mft size in cmil in 2 m 2 strand Bare Weatherproof Bare Insulated MCM: 2,000 2,000,000 1.571 1,014 91 1.630 1.880 6,175 7,008 1,750 1,750,000 1.374 887 91 1.526 1.776 5,403 6,193 1,500 1,500,000 1.178 760 61 1.411 1.661 4,631 5,380 1,250 1,250,000 .9817 633 61 1.288 1.538 3,859 4,508 1,000 1,000,000 .7854 507 61 1.152 1.402 3,088 3,674 900 900,000 .7069 456 61 1.094 1.313 2,779 3,332 800 800,000 .6283 405 61 1.031 1.250 2,470 2,992 750 750,000 .5890 380 61 .998 1.217 2,316 2,822 700 700,000 .5498 355 61 .964 1.183 2,161 2,650 600 600,000 .4712 304 37 .893 1.112 1,853 2,235 500 500,000 .3927 253 37 .813 1.001 1,544 1,894 450 450,000 .3534 228 37 .772 .960 1,389 1,724 400 400,000 .3142 203 19 .726 .914 1,235 1,553 350 350,000 .2749 177 19 .679 .867 1,081 1,345 300 300,000 .2356 152 19 .629 .817 926 1,174 250 250,000 .1963 127 19 .574 .762 772 985 AWG: 0000 211,600 .1662 107 '7,19 .528 .684 653 800 000 167,800 .1318 85.0 1 7,19 .470 .626 518 653 00 133,100 .1045 67.4 7 .414 .570 411 522 105,500 .08289 53.5 7 .368 .524 326 424 1 83,690 .06573 42.4 7 .328 .484 258 328 2 66,370 .05213 33.6 7 .292 .417 205 270 3 52,640 .04134 26.7 7 .260 .385 163 206 4 41,740 .03278 21.2 7 .232 .357 129 170 5 33,100 .02600 16.8 7 .206 .331 102 140 6 26,250 .02062 13.3 7 .184 .309 81 115 7 NA NA NA NA NA NA NA NA 8 16,510 .01297 8.37 7 .146 .271 51 78 NA Not available. 1 Sizes AWG 0000 and 000 cable are usually made of 7 strands when bare and 19 strands when insulated. Bare wire breaking Resistance, n/Mft strength, lb at 20°C, standard Hard drawn Soft annealed annealed 87,790 43,830 0.005289 77,930 38,350 .006045 65,840 32,870 .007052 55,670 27,390 .008463 45,030 21,910 .010578 40,520 19,270 .011753 36,020 17,530 .013223 34,090 16,430 .014104 31,820 15,340 .015112 27,020 13,150 .017631 22,510 10,960 .021157 20,450 9,860 .023508 17,560 8,765 .026447 15,590 7,669 .030225 13,510 6,574 .035262 1 1 ,260 5,478 .042315 9,617 4,637 .04999 7,366 3,677 .06304 5,926 2,916 .07949 4,752 2,312 .10024 3,804 1,834 .1264 3,045 1,525 .1594 2,433 1,209 .2009 1,938 959 .2535 1,542 761 .3195 1,228 603 .4029 NA NA .5080 777 379 .6406 214 >^^^^ TT~TT i— © 578, 1976; NTIS PB 283 493. 29. National Bureau of Standards. Thermocouple Reference Tables Based on the IPTS-68. NBS Monogr. 125, 1974. 30. Neher, J. H., and M. H. McGrath. The Calculation of the Temperature Rise and Load Capability of Cable Systems. Trans. Am. Inst. Electr. Eng., Part 3, v. 76, Oct. 1957. 31. Ohio Brass Co. (Mansfield, OH). Haulage Product Informa- tion and Design Drawings. 1978. 32. Stefanko, R., and L. A. Morley. Mine Electrical Systems Evaluation (grant G0133077, PA State Univ.). Mine Power System Performance. BuMines OFR 76(4)-75, 1974; NTIS PB 245 930. 33. Stefanko, R., L. A. Morley, and A. K. Sinha. Evaluation of Mine Electrical Systems With Respect to Safety, Technology, Economics, and Legal Considerations (grant G0101729, PA State Univ.). Volume 1. Text, Tables, and Analyses. BuMines OFR 70(l)-73, 1973; NTIS PB 225 476. 34. Tomlinson, J., T. Rusnak, R. H. King, and L. A. Morley. Splice Testing Using a Figure-S Machine and a New Shuttle Car Simulation (grant G0188036, PA State Univ.). BuMines OFR 80-80, 1979, NTIS PB 80-210222. 35. Trutt, F. C, J. W. Robinson, L. A. Morley, and P. M. Zahn. Electrical Materials Analysis- Arcing (grant G0155197, PA State Univ.). BuMines OFR 90-78, 1977; NTIS PB 284 946. 36. Tsivitse, P. J. Mining Motors. Ch. in Motor Application and Maintenance Handbook, ed. by R. W. Smeaton. McGraw-Hill, 1969. 37. U.K. National Coal Board (London). Flexible Trailing Cables for Use With Coalcutters and for Similar Purposes. N.C.B. Spec. 188/1971. 38. U.S. Code of Federal Regulations. Title 30 -Mineral Resources; Chapter I -Mine Safety and Health Administration, Department of Labor; Subchapter O-Coal Mine Health and Safe- ty; Part 18 -Electric Motor-Driven Mine Equipment and Ac- cessories; Part 75 -Mandatory Safety Standards, Underground Coal Mines; Part 77 -Mandatory Safety Standards, Surface Coal Mines and Surface Work Areas of Underground Coal Mines; 1981. 223 39. U.S. Mine Safety and Health Administration. Coal Mine 41. Woboditsch, W. Belastbarkeit von Aufgetrommelten Bzw. Safety Electrical Inspection Manual, Underground Coal Mines. Sonnenbeschienenen Leitungstrossen (Current Carrying Capacity Apr. 1979. of Reel-Wound and Sun-Exposed Cables). Elektrie, v. 18, Dec. 40. Westinghouse Electric Corp. (East Pittsburgh, PA). Elec- 1974. trical Transmission and Distribution Reference Book. 4th ed., 1964. 224 CHAPTER 9.— PROTECTIVE EQUIPMENT AND RELAYING Even the best designed electrical systems occasionally experience faults and overloads, or disturbances that cause abnormally high currents. These currents can exist in the ground system or in the phase conductors. Wherever the occurrence, the situation is likely to precipitate a hazard to either equipment or personnel. Of the basic design criteria that underlie all mine power systems, three are of critical importance in protec- tive equipment and relaying: adequate interrupting capac- ity, current-limiting capacity, and selective system opera- tion. The first two provide protection to the system during a disturbance, while the third is designed to locate the problem, then minimize its effect. In chapter 7, current limiting and selective relaying were designated as two prime purposes of grounding. It was shown that ground- fault currents can be limited by inserting a resistance in series with the neutral conductor. However, not much has been presented about selective system operation, other than its need. Protective circuitry and protective relaying are the tools behind selective system operation and are the main topics of this chapter. The protective circuitry associated with the power system consists of transducers, relays, and switching ap- paratus. Its role of safeguarding personnel and equipment can be effected manually or automatically. An instance of manual utilization would be removing power from a sys- tem portion for maintenance. An example of automatic operation would be a situation in which protective cir- cuitry first senses then clears each hazardous current resulting from a disturbance. As might be expected, the process of clearing is disconnecting the affected circuit from the power source safely and as quickly as possible, with minimum interference to the system balance. In other words, protective circuitry must isolate a malfunc- tion at a given location with minimum damage to circuits and equipment and minimum operation downtime. The function of protective circuitry to provide detection and isolation is termed selective relaying. All the devices that comprise the protective circuitry in the mine power system thus play a vital role in safety. In fact, protective circuitry is probably the most important component of the power system and forms a major portion of all power equipment. For example, a switchhouse, which has the principal function of protection, is simply a complex of protective devices. The basic concepts of overloads and faults are intro- duced in chapter 4. Although the removal of destructive overloads is important, the main concern is the clearing of faults, since their occurrence can be catastrophic. Because of the preponderance of cables, cable shield- ing, and grounded equipment in mine power systems, line-to-neutral faults are the most common, and most of these are arcing with relatively short length and con- trolled distance. Ground-fault current is predominantly limited by neutral grounding resistors, whereas in other industrial applications, ground-fault currents are often limited by fault impedance. Line-to-line and three-phase faults can also occur, as when a mobile machine severs a cable during a runover. Extremely large line currents can result, which can be limited in the mine system only by transformer and line-conductor impedances. System components, such as couplers, cables, transformers, bus bars, and disconnect switches, must be capable of withstanding the momentary mechanical and thermal stresses created by the flow of fault current through them. Interrupting devices, such as circuit breakers, must be able not only to withstand these momentary fault-through stresses, but to interrupt or terminate these anomalous currents. The maximum magnitude of possible fault currents existing in line conductors must be known in order to select adequate ratings of protective equipment. Indeed, this knowledge is required to coordinate protective- circuitry operation for the entire complex. It may also be necessary to know the minimum sustained fault current that is available in the system in order to determine the sensitivity requirements of the current-responsive protec- tive devices. These fault magnitudes, both maximum and minimum, are usually estimated by calculation, and the equipment is selected using the calculated results. Because of the many hazards that can occur, the system must be capable of detecting overloads, short circuits Qine faults), undervoltage, and ground faults, as well as any compromise in grounding-conductor continuity. With the use of resistance grounding in mine power systems, the protec- tive relaying or sensing device associated with ground faults or zero-sequence currents is usually handled separately from that for line faults causing only anomalous positive-sequence or negative-sequence currents. In addition, the relaying for overloads may be separate from that for faults. Except for fuse applications, the sensing devices for each function will normally cause the activation, or tripping, of the same circuit-interrupting device no matter what the protection requirements are for an individual location. The sensing devices may be an integral part of the interrupting appara- tus or be separated from it and connected only through control wiring. This chapter builds upon the material covered in chapter 4, beginning with the main protection compo- nents, switching apparatus and sensing devices. Basic relay connections, relay terminology, and different kinds of protection follow. Finally, typical assemblies and com- binations of protective circuitry are discussed. Essentially, this chapter sets the stage for chapter 10, where fault calculations, device sizing, and coordination are outlined. SWITCHING APPARATUS A switching apparatus is defined as a device for making (closing), breaking (opening), or changing connec- tions (6). 1 There are three basic types of apparatus in this classification: switches, circuit breakers, and fuses. All switching devices are given certain design ratings, which are a measure of the electrical stresses they can withstand (6). Obviously, the ratings must be correlated with the intended use or duty. A listing and definition of these ratings follows but is restricted to those terms having direct application to the development of the topic in this and the subsequent chapter. Further concepts will be added in the discussion of transients and overvoltages in chapter 11. 1. Voltage. The maximum nominal system voltage at a specified frequency (usually line-to-line for ac devices) on which the device may be installed. 1 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. 225 2. Continuous current. The maximum continuous cur- rent that the apparatus may carry. 3. Short-circuit current. Usually, the maximum cur- rent the device is capable of interrupting. This may be further qualified by an interrupting-current or interrupting-capacity rating. 4. Close-and-latch or momentary current. The maxi- mum short-circuit current that the device can withstand during the few cycles after the fault occurs without expe- riencing severe mechanical damage. The ratings of switching apparatus are based on the maximum possible values of fault currents. To help visu- alize the importance of these ratings, consider that a three-phase fault has occurred on a power system. Figure 9.1 illustrates the resulting line current versus time, created by the flow of energy from the source or sources to the fault (7). This asymmetrical waveform is made up of two components: dc and a symmetrical ac. At any instant after the fault occurs, the total fault current equals the sum of the two. The dc component decays to zero in a short time, with the total current gradually changing from asymmetrical to symmetrical. Switching-apparatus ratings, as a measure of the stresses involved during faulting, are based on the sym- metrical rms value. Asymmetry is accounted for by taking the basic symmetrical value and applying multiplying factors. These concepts are presented in detail in chapter 10. However, figure 9.1 does provide a useful visualization of rating magnitudes, and these will be discussed in following sections, along with each switching device. Total asymmetrical current dc component Symmetrical ac TIME Figure 9.1. —Typical system fault current. Wh -Contacts closed, circuit energized Load or short circuit ARCS AND CIRCUIT INTERRUPTION After a switching apparatus receives a message that circuit current is to be interrupted, the device proceeds through definite steps to terminate the current (4). These are illustrated in figure 9.2. Under normal operation, the contacts of the apparatus are closed, current flows through the interface, and the outgoing circuit is thus energized. To terminate current, the contacts begin to separate and an electric arc is drawn. The arc is composed of free electron and free positive- ion flow, as shown in figure 9.3. To initiate this arc, free electrons and/or free positive ions must exist between the contacts. Their availability depends upon the following environmental conditions: • In air or gas, the conductive elements are generated prior to the initiation of the arc by radiation and cosmic rays, which knock off electrons from neutral gas molecules. • In a liquid such as oil, the conductive elements exist as impurities. • In a vacuum, they can be emitted from the cathode by a high-strength electric field with the process known as high field emission. The last case can add free electrons to any environment. Even though the voltage between the cathode and anode is low immediately after separation, the free electrons are attracted to the anode, and the positive ions toward the cathode. The electron flow accounts for about 90% of arc current (4). If the voltage across the arc remains large enough, the movement of charge between the contacts initiates the mechanisms that can increase and sustain the arc. This / Contacts parting, arc drawn 1-711-71 between contacts Load or short circuit •Contacts open, arc extinguished -0 LZh Load or short circuit deenergized Figure 9.2.— Steps in circuit interruption. Region B positive column Region A Electrons emitted in great number thermionically or from cathode spot Figure 9.3.— Arc between two contacts. 226 again depends upon the environment. In a gas, the free electrons can collide with neutral gas molecules, producing additional free electrons and positive ions, termed ionization by collision. In any atmosphere, the collision of heavy posi- tive ions on the cathode produces heat, which can augment field emission in low-melting-point materials such as copper, creating intense electron discharge from a small area, called a cathode spot, and can cause thermionic emission in high- melting-point substances, such as carbon, where electrons are boiled out by high temperature. Once the arc is established, processes must be brought into play to extinguish it. In general, the greater the arc current and the higher the voltage of the circuit, the more difficult is the problem of arc extinction. The situation is easier in ac systems than in dc systems because the current waveform passes through zero in ac systems. However, the arc can restrike when the voltage rises again if the ionic conditions across the contacts permit. For dc, the arc is readily maintained because a normal cur- rentzero does not exist. Whatever the extinction process, the switching device can open the circuit successfully, provided that the current to be interrupted is within the rated value. However, if the switching device is required to terminate a current well above the design value, the arc between the parting contact may not extinguish or may continue to restrike, and the apparatus could be destroyed by the gas pressure built up within it (5). When a device is designed to interrupt fault current, it is often called an interrupting device; otherwise, it is commonly called a switch and is designed only to open and close a circuit. While some switching apparatus are in- tended to serve only one of these functions, others can do both. SWITCHES A switch has exactly the same definition as switching apparatus, with the qualification that it is a manual device (6); in other words, its operation is a normal or intended occurrence on the power system. The switch types common in mine systems are the disconnect and the load interrupter. Both have the prime function of isolating outgoing circuits from the power source. A disconnect switch is not intended to interrupt circuit current and can be operated only after the circuit power has been removed. Interlocks must be provided to prevent manual operation under load, and latches may be needed to prevent opening from the stresses resulting from fault circuits. Consequently, disconnect switches do not have an interrupting rating; but beyond a continuous- current rating, they may need a momentary-duty or close- and-latch rating for handling fault-through currents. An interrupter or load-break switch differs from a disconnect in that it has an interrupting rating. The device has the capability of terminating currents that do not exceed the continuous-current rating, although this is not its normal operation. Interrupter switches usually have a quick-make, quick-break mechanism, which pro- vides a fast switch-operation speed independent of the handle speed. The illustration in figure 9.4 shows a three-pole device; the mechanism on the right side of the connecting shaft provides the fast operation. Some units can be motor driven, thus allowing remote or automatic operation. In most mining applications, load-break switches need a close-and-latch rating. Where interlocks Figure 9.4.— Load-break switch. (Courtesy Line Power Manufac- turing Corp.) are not employed, this rating indicates the margin of safety when the switch is closed into a faulted circuit. Switches are normally used as disconnects in mining systems regardless of their ratings; in fact, some States require load-break switches with interlocks for all discon- necting applications. These interlocks cause interruption of source power prior to contact separation, and the operation is usually performed through the ground- monitoring circuitry. Load-break switches, used in con- junction with fuses, are employed as interrupters in cer- tain circumstances. CIRCUIT BREAKERS A circuit breaker is primarily an interrupting device, but in some cases it is also used as a switch (6). A circuit breaker can be defined as a device designed to open and close a circuit manually and to open the circuit automat- ically at a specific current level without injury to itself when properly applied. It is available as a single pole, double pole, or triple pole. Manual operation, be it me- chanically or electrically actuated, is again intended where the circuit current is not in excess of rated contin- uous current. Automatic operation is dictated by a system abnormality, such as a fault or an overload. In this case, the device may be called upon to interrupt current in excess of the rated continuous current. Circuit breakers in the role of interrupting devices must be used with sensing devices to perform their intended function. In medium-voltage and low-voltage mining appli- cations, the operation may be internally controlled by self- contained current-responsive elements, external protective relays, or a combination of both. In high-voltage situations, the sensing devices are always separate, with interconnec- tions only through control wiring. Circuit breakers can generally be broken into two classifications: those intended for systems over 1,000 V, and those for 1,000 V and below. Devices in the first class are called power circuit breakers, while the second class is divided into power circuit breakers and molded-case cir- cuit breakers. Following mining standards, circuit break- ers for systems below 661 V are called low voltage; for 661 to 1,000 V, medium voltage; and above 1,000 V, high voltage. It should be noted that IEEE Standards define above 1,000 V to 72,500 V as medium voltage and below 1,000 V as low voltage. Low- voltage and medium-voltage circuit breakers are usually considered together and can 227 find ac and dc service. High-voltage breakers involve only ac circuits. The next paragraphs look at typical apparatus and operation. CIRCUIT BREAKERS FOR LOW AND MEDIUM VOLTAGE The term air circuit breaker is often used when referring to molded-case and power circuit breakers de- signed for low-voltage and medium-voltage systems (7). Air circuit breakers employ the simplest method of inter- rupting current: extinguishing the arc in normal atmo- sphere by increasing its length (4). Several different pro- cesses can be used to force the arc to lengthen. lb illustrate one arc-lengthening technique, consider figure 9.5, where two circuit breaker contacts, a and b, have just separated. The horn-like arrangement of the contacts shown in the figure can be considered an arc chute, which is a barrier that confines, cools, and extin- guishes the arc (5). By the ionization of the air between the contacts, an arc is drawn and heat is generated. The arc extinction action then commences; this is also called deionization because it serves to reduce the free electrons and positive ions in the gas (4). Air currents, created by the heat and confined by the arc chute, force the arc upward to form a loop. Electromagnetic forces within the loop further encourage the lengthening. As a result of cooling by radiation or convection, the longer arc requires a higher arc voltage to sustain current flow, and thus, the arc is extinguished. As noted earlier, arc interruption in an ac circuit occurs much more easily than in a dc circuit. All voltages and currents in an ac system go through cyclic changes, and consequently, ion-producing effects for the arc are variable too: falling as current becomes smaller, ceasing at current zero (4). Deionizing effects in the arc chute remain steady. To take advantage of this situation, circuit break- ers for ac systems are often designed around the minimum voltage required to establish a cathode spot. Because there is no natural current zero in dc systems, the circuit breaker must force the current to zero. For this to happen, the arc voltage must be greater than the system voltage (14). An enormous amount of heat can be generated in all circuit breakers while the arc exists, and an important function of the circuit breaker assembly is to dissipate this heat safely. The foregoing simple arc-lengthening technique works well for 240-Vac applications. Conventional practice is to use a single-pole breaker for 120 Vac and a double- pole breaker for 240-Vac single-phase circuits. The latter employs one pole of the circuit breaker in series with each power conductor. For circuits 250 V and above, the direct arc-lengthen- ing approach is not enough; special arc chutes, quenchers, or deionizing chambers are needed to assist in arc termi- nation (5). Figure 9.6 illustrates one approach, where the arc is forced into metallic barriers by magnetic attraction and broken into a series of smaller arcs. Each of these arcs is subjected to lengthening, cooling, and the problem of reestablishing a cathode spot if low-melting-point materi- als are used (4). Another approach is depicted in figure 9.7. Because the arc establishes its own electromagnetic field, an external magnetic field can enhance arc lengthening. The process is termed magnetic blowout, and breakers using this principle are called air magnetic. Coils carrying the circuit current in series with the arc can provide the Arc rises on horns V b Arc drawn here Figure 9.5.— Extinguishing arc by increasing the length. a b Metal barriers Arc Arcing contacts Main contacts Figure 9.6.— Metal-barrier arc chute assists in arc deioniza- tion. Barriers of insulating material Arcing contacts Figure 9.7.— Insulated-barrier arc chute used with magnetic field. 228 magnetic field. As shown in the figure, the magnetic field forces the arc into insulated barriers or fins, creating further lengthening; recombination and cooling at the barrier surfaces accelerates deionization (4). In dc mine power circuits below 660 V, air-magnetic breakers are used extensively, especially on trolley sys- tems. With very few exceptions, molded-case breakers are employed for ac circuits below 1,000 V. In addition, molded-case units are often used to protect low-voltage dc face equipment. Molded-Case Circuit Breakers The molded-case circuit breaker is the most explicit example of interrupting apparatus with self-contained current-responsive elements. It is defined as a breaker that is assembled as an integral unit in a supporting and enclosing housing of insulating material (5). Depending upon the amount of protection desired, these devices can sense internally and then clear undervoltage, overcurrent, and short-circuit conditions. Some tripping elements, that is, the actual components that cause the contacts to start separating, are also externally accessible through control wiring. Hence, other circuit protection can be added. Except for some power circuit breakers of low-voltage and medium-voltage design, all the circuit breakers that will be discussed in this chapter rely solely on outside infor- mation to perform their prime function. Molded-case ap- paratus will be presented first so that many important terms can be introduced. The application of molded-case circuit breakers in mining began in the 1950's with the conversion from low-voltage dc power distribution to ac power distribution and face rectification, expanding further with the trend toward ac face equipment. In fact, Wood and Smith (21) have attributed the introduction of low-height, solid-state rectifier units in underground mines (which permitted the use of ac distribution) to molded-case circuit breakers, citing the lack of high-speed dc circuit breakers of the proper height as the previous limitation. Molded-case breakers placed between the transformer and the rectify- ing bridge lowered the height limitation to that of the transformer, allowing a unit design complementary with the mining environment. The largest mining application is trailing-cable pro- tection in underground face areas. The breakers are lo- cated in power centers and provide cable protection on each outgoing circuit, as required by 30 CFR 75.900, in addition to functioning as switching devices. The typical molded-case breakers, however, are not designed for repet- itive switching. Mining use subjects them to many more operations than found in other industries, and regular or standard circuit breakers generally cannot hold up to the stress. Several manufacturers, recognizing this problem, have produced a special line of mine-duty molded-case breakers, which have stronger construction to withstand the punishment of mine use. Except for external adjustments, molded-case devices sometimes do not allow field maintenance; many are sealed to prevent tampering. Although some manufactur- ers offer a complete line of replacement components, repairs other than an exchange of easily removable parts, such as arc chutes or trip units, should be made only by qualified repair facilities. This is critical, given the impor- tance of the molded-case circuit breaker in personnel protection. All component parts of these circuit breakers are built into one insulated housing, the molded case. These parts are the operating mechanism, arc extinguishers (arc chutes), contacts, trip elements, and terminal connectors, as shown in figure 9.8 (19). Additional accessories may be included. The molded case is made of a glass polyester or similar synthetic material that combines ruggedness and high dielectric strength with a compact design. Each type and size of molded case is assigned a frame size or designation for easy identification. This coding, loosely based on an old Underwriters' Laboratories standard, refers to a number of breaker characteristics, including maximum allowable system voltage, maximum allowable continuous current, interrupting capacity, and the physi- cal dimensions of the molded case. Several trip units may be available for a particular frame size, so a specific assembled breaker may have a lower continuous-current rating than the current designation of the frame. Table 9.1 lists the continuous ratings considered to be standard for mining service. The currents in parentheses are the lower current settings available in that frame size from certain manufacturers. Unfortunately, manufacturers have vary- ing design criteria and hence size their units to dissimilar Table 9.1. —Ratings for mining-service molded-case circuit breakers Frame size, ' A Continuous-current ratings, 2 A 100 100 (70, 50, 30) 225 225 (175, 150, 125, 100) 400 400 (300, 225, 175, 150, 125, 100) 600 600 (500, 400) 800 800 (600) 1,200 1,200 (1,000, 800, 600) 1 Regular-duty breakers also available in 1,600-, 2,000-, and 2,500-A frames. 2 Currents in parentheses are lower settings available in the frame size. Molded case (frame) \ Operating mechanism Arc extinguishers Contacts Trip elements Figure 9.8.— Molded-case circuit breaker components. (Courtesy Westinghouse Electric Corp.) 229 specifications. For example, a 225-A, 600-V breaker sup- plied from two separate manufacturers may have different physical dimensions so that direct interchanging is diffi- cult, if not impossible. The circuit breakers rated in table 9.1 are generally available as two-pole or three-pole units at 600 Vac or 300 Vdc, but only as three-pole devices at 1,000 Vac. The two-pole breakers are intended for dc face equipment or single phase ac applications. By convention, one pole is used for each ungrounded conductor in a circuit (5). The arc chutes define the interrupting-current capac- ity of the assembly in conjunction with the insulating and heat-dissipation properties of the molded case. The chutes assist arc deionization by the principle discussed for figure 9.6. They are also termed arc extinguishers or arc quench- ers by some manufacturers. The breaker case must be mounted vertically with the arc chutes at the top for correct arc-extinction operation. Circuit breakers designed for 1,000 V and below are capable of clearing a fault faster than those constructed for high voltage (6). The contacts often begin to part during the first cycle of fault current, and consequently, the breaker must be capable of interrupting the maximum allowable first-cycle asymmetrical current. Thus, for lower voltage breakers, the close-and-latch and interrupting ratings are usually the same, a characteristic not found with high-voltage breakers. The rating of these units is carried out on a symmetrical basis, so multipliers account- ing for the dc offset need not be applied as long as the system X/R ratio does not exceed 6.6 (6) (see chapter 10). Table 9.2 lists typical interrupting ratings versus the system voltages for mine-duty circuit breakers; the ac system values are based on the symmetrical rating. Some manufacturers offer both standard-duty and high- interrupting-capacity breakers for mining service. The table values presented parenthetically indicate the supe- rior construction, which incorporates sturdier contacts and mechanism plus a special high-impact molded casing. Table 9.2 shows that typical molded-case circuit breakers constructed for 1,000- Vac mine systems have only a 10,000-A symmetrical interrupting rating. This presents a concern, as available short-circuit currents on high- power 1,000-Vac systems can be greater. Instances include longwall mining equipment, which needs a power-center capacity of 1,500 kVA or more. To overcome the problem, a manufacturer has introduced molded-case breakers with a 24,000-A asymmetrical interrupting rating at 1,000 Vac and continuous-current ratings of 600, 800, 1,000, or 1,200 A. The asymmetrical rating is used to provide more flexibility for designing the breaker into power systems. The function of the operating mechanism of a typical molded-case circuit breaker is to provide a means of opening and closing. It is a toggle mechanism of the quick-make, quick-break type, meaning that the contacts snap open or closed independent of the speed of handle movement. The breaker is also trip-free; that is, it cannot be prevented from tripping by holding the breaker handle in the ON position during a fault condition. In addition to indicating whether the breaker is ON or OFF, the operating-mechanism handle indicates when the breaker is tripped by moving midway between these positions. To reactivate the tripped breaker, the handle must first be moved from the central position to OFF, which resets the mechanism, and then to ON. This distinct trip point is particularly advantageous where molded-case breakers are grouped, as in a power center, because it clearly indicates any faulty circuits. The function of the trip elements is to trip the operating mechanism in the event of prolonged overload or short-circuit current. Two common types of trip elements are used in mining, magnetic and thermal magnetic. When the circuit being protected involves portable or trailing cables, the thermal-magnetic combination is strongly recommended and is mandated by some States. The magnetic trip protects against short circuits, and an electromagnet in series with the load current provides the trip action (19). This type of short circuit is actually a line-to-line or three-phase fault on ac, or a line-to-line fault on dc systems. When a short occurs, the high fault current causes the electromagnet in the breaker to attract the armature, initiating an unlatching action, which in turn causes the circuit to open (fig. 9.9). The action takes place within 1/2 s (usually within 1 cycle or 16 ms), instanta- neously tripping the breaker. Since tripping takes place with no intentional delay, the magnetic trip is often called the instantaneous-trip element. Screwdriver slots, located on the front of the trip unit, are used in adjusting the sensitivity (fig. 9.10A). By law, the maximum setting is established by the protection of the minimum conductor size in the circuit (16-17). Table 9.3 lists these maximum settings applied to trailing cables. Figure 9.10B illustrates a family of time-current curves resulting from the adjust- able range; to the left or below each curve, the breaker will not be tripped magnetically. Typical instantaneous-trip ranges versus frame sizes for mining-service breakers are given in table 9.4. Note that this is not a rigorous listing, since some manufacturers will provide any desired trip range with most frame sizes upon request. The other common molded-case breaker type is the thermal-magnetic variety. In addition to providing short- circuit protection, the thermal-magnetic breaker also guards against long-term current overloads existing longer than roughly 10 s, by incorporating thermal trip elements (fig. 9.11). The thermal action is accomplished through use of a bimetal strip heated by load current (19). The strip consists of two pieces of metal bonded together, each with a different coefficient of thermal expansion. A sustained overload causes excessive heating of the strip, resulting in deflection of the bimetal, which in turn causes the operating mechanism to trip the breaker. Because the Table 9.2.— Interrupting-current ratings 1 versus system voltage, amperes Frame size, A 240 Vac 100 18,000 (65,000) 225 25,000 (65,000) 400 42,000 (65,000) 600 42,000 (65,000) 800 42,000 (65,000) 1,200 42,000 (65,000) 1 Parenthetical ratings are for typical premium-duty circuit breakers. 2 Actual dc interrupting current dependent upon system inductance. 480 Vac 600 Vac 1 ,000 Vac 300 Vdc 2 14,000 (25,000) 14,000(18,000) 10,000 10,000 (20,000) 22,000 (35,000) 22,000 (25,000) 10,000 10,000(20,000) 30,000 (35,000) 22,000 (25,000) 10,000 10,000(20,000) 30,000 (35,000) 22,000 (25,000) 10,000 10,000(20,000) 30,000 (35,000) 30,000 (25,000) 10,000 10,000(20,000) 30,000 (35,000) 30,000 (25,000) 10,000 10,000 (20,000) 230 Table 9.3.— Maximum instantaneous-trip settings Maximum Maximum Conductor allowable Conductor allowable size instantaneous setting, A size instantaneous setting, A AWG: AWG: 14. 50 1 .... 1,000 12. 75 1/0. 1,250 10. 150 270. 1 ,500 8... 200 3/0. 2,000 6... 300 4/0. 2,500 4... 500 MCM: 3... 600 250 to 2,500 2... 800 Magnetic element Load Magnetic element closes gap and opens contacts on short circuit Latch ^Contacts closed Latched ^ Contacts open Tripped Figure 9.9.— Magnetic-trip relay. Table 9.4.— Commonly available magnetic-trip ranges for mining-service molded-case breakers Frame size, A 100 225 400 600 800 1,200 Magnetic-trip range, A Range of allowable conductor sizes 50- 180 150- 500 300- 700 500-1,000 300-1 ,000 500-1 ,000 800-1 ,600 500-1,500 900-3,000 750-1 ,500 1,000-2,000 1 ,500-3,000 2,000-4,000 1,500-3,000 2,000-4,000 2,500-5,000 14-10 AWG 10-4 AWG 6-3 AWG 4-1 AWG 6-1 AWG 4-1 AWG 2-2/0 AWG 4-2/0 AWG 1 AWG - 500 MCM 2-2/0 AWG 1-3/0 AWG 2/0 AWG - 500 MCM 3/0 AWG - 500 MCM 2/0 AWG - 500 MCM 3/0 AWG - 500 MCM 4/0 AWG - 500 MCM Low Intermediate High ABC CURRENT A Adjustment knob B Characteristics Figure 9.10.— Adjustable instantaneous setting. bimetal deflection is dependent upon current and time, the thermal unit provides a long-time delay for light overloads and a fast response for heavy overloads. A representative current-time curve for the thermal unit alone is shown in figure 9.12A; later in this chapter, it will be described as an inverse-time characteristic. In compar- ison, figure 9.12B shows the circuit breaker response when both thermal and magnetic trip elements are incorpo- rated. The shaded area for each curve represents a toler- ance between the minimum and maximum total clearing time. The thermal-magnetic unit shown in figure 9.11 is ambient-temperature sensitive. Assuming the circuit breaker, cable, and equipment being protected are in the same ambient temperature, the circuit breaker trips at a lower current as the ambient temperature rises in corre- spondence to safe cable and equipment loadings, which vary inversely with ambient temperature (19). Thermal- magnetic trip elements are available that automatically compensate for ambient-temperature variations. The am- bient compensation is obtained through an additional bimetal strip, which counteracts the overload bimetal. Such trip units are recommended whenever the protected conductors and the circuit breakers are in different ambi- ent temperatures (19). Most mining-service molded-case breakers with 225-A frame sizes and above have interchangeable trip units. For straight magnetic elements these allow different instantaneous-trip ranges per frame size. However, thermal-magnetic units can be used to establish a lower continuous-current limit for the breaker. The National Electrical Code (13) is used as a guide to define the current Magnetic element / Load Load K Contacts open V 7 , Contacts closed Tripped Figure 9.11.— Thermal-magnetic action of molded-case cir- cuit breaker. 1,800 10 1a 8 135 500 CURRENT, % of breaker rating A Thermal only Figure 9.12.— Time-current characteristics for thermal- magnetic molded-case circuit breakers. 60 ''^A — I — Thermal action 0.016 i|jb - Magnetic action ; , 250 4,000 CURRENT, % of breaker rating B Thermal magn etic 231 at which the long-time-delay thermal element must ini- tiate the circuit-clearing operation and specifies a point that is 125% of the rated equipment or conductor ampac- ity. As seen in figure 9.12A, the circuit breaker will take no action below this current. Hence, the thermal portion defines the continuous-current rating of the breaker, speci- fied as 100% at 40° C for conventional (non-compensating) thermal-magnetic elements. Obviously, the thermal element current rating cannot exceed the frame rating. Because of the connection, some manufacturers recommend that the continuous current through the breaker be limited to 80% of the frame size. This topic will be continued in chapter 10. Electromechanical magnetic and thermal-magnetic trip elements have been replaced by solid-state compo- nents in some molded-case breakers. Although the solid- state counterparts may become popular in the future, they have not yet achieved wide acceptance in the mining industry. Nevertheless, these breakers are discussed in chapter 14. The last basic breaker components are the terminal connectors. Their function is to connect the circuit breaker to a desired power source and load. They are usually made of copper and must be constructed so that each conductor can be tightened without removing another. The terminal connectors shown in figure 9.8 are for direct connection of one cable connector per terminal. Many molded-case breakers also have provisions for threaded-stud terminals. These studs can be used not only for connection of more than one conductor per terminal, but also for breaker mounting. It should be noted that the type of terminal used on a breaker may change its heat dissipation proper- ties and thus lower its interrupting rating. In addition to the basic components, several accesso- ries are available, of which the most common are the terminal shield, shunt trip, and undervoltage release (UVR). Terminal shields protect personnel from accidental contact with energized terminal connections and are sim- ply plates that shield (guard) the terminals. The other two accessories are used to trip the operating mechanism. A shunt trip is employed to trip a circuit breaker electrically from a remote location. It consists of a momentary-rated solenoid tripping device mounted inside the molded case that activates when control power is applied across the solenoid coil. The magnetic field created by the solenoid moves a plunger, which in turn activates a trip bar. At the same time, a series cutoff switch removes power to the solenoid coil, preventing it from burning up under continuous load. A typical shunt-trip assembly is shown in figure 9.13. The shunt trip can remotely trip the breaker but cannot remotely operate it. To reclose the breaker, the handle must first be moved to the reset position, then to the ON position. The purpose of the UVR is to trip the breaker when- ever control voltage to the UVR falls below a predeter- mined level, usually 35% to 70%. This device is also mounted inside the breaker frame and consists of a spring and a solenoid. The spring is cocked or precharged by the operating mechanism when the breaker is closed and is held in the cocked position by the solenoid after closure. If the voltage drops below the required level, the solenoid releases the spring, causing the circuit breaker to trip. The breaker cannot be turned on again until the voltage returns to 80% of normal. The importance of the shunt trip and UVR is far ranging, as they allow the protection capabilities of circuit breakers to be extended. The molded-case breaker alone can provide overload and short-circuit protection in an outgoing circuit. The UVR adds undervoltage protection; in fact, undervoltage protection is normally required at most breaker locations. Note that undervoltage protection is required for all equipment, but it is not required on all circuit breakers as long as all equipment downstream from the breaker has undervoltage protection. The under- voltage protection provided by a UVR is actually "loss- of- voltage" protection since the dropout level is well out- side the recommended operating range of most motors (see chapter 6). Through a specific combination of relays and sensing devices, additional types of protection can be applied through shunt or UVR tripping. With a shunt trip, the relay completes the circuit between the control-power source and the solenoid coil. When a UVR is used, the relay removes the control voltage across the solenoid coil. This circuitry will be discussed in detail later in the chapter. The molded-case circuit breaker is the most widely used breaker in mining, even though its employment is restricted to low-voltage and medium-voltage systems. The principal application is on ac, where it provides high B Figure 9.13.— Shunt-trip (A) and undervoltage-release (8) accessories. (Courtesy General Electric Co.) 232 interrupting capacity for short circuits in minimum space. On ac or dc systems, it is often the first protection device called upon to handle electrical problems existing on trailing cables and mining machinery. A clear understand- ing of the construction and rating of these breakers is required to assure adequate protection. The operating characteristics must be closely matched with those of the trailing cable to minimize hazards to personnel. Power Circuit Breakers listed values, frame sizes are available up to 6,000-A continuous ac current and 12,000-A continuous dc current (5). These frame sizes are rated to carry 100% of the continuous-current rating inside enclosures at 40°C. In power breakers with low current ratings, arc interruption can utilize arc-chute arrangements similar to those used in molded-case breakers. The full air-magnetic arrange- ments described for figure 9.7 are employed for high- current-interruption power breakers. Some mining-industry engineers have found that molded-case circuit breakers cannot handle the available short-circuit currents in certain low-voltage applications, such as the outgoing dc circuits of trolley rectifiers and dc face equipment. The low-voltage power circuit breaker provides an alternative in these cases. Power circuit breakers for applications of 1,000 V and below are of open construction assembly with metal frames. They are designed to be field maintained under planned periodic inspection, and all parts are accessible for ease of maintenance, repair, and replacement (6-7). The design enables higher endurance ratings and greater repetitive-duty capabilities than are available from molded-case devices. However, power circuit breakers are intended only for service inside enclosures with "dead- front" construction, that is, not accessible to unauthorized personnel. Electromechanical units are available for long-time tripping, but mechanical-displacement dashpot types are normally used for this function and provide the same overcurrent protection as does the bimetal thermal trip- ping in molded-case breakers. Although long-time charac- teristics are not adjustable with bimetal strips, the dash- pots allow the long-time-delay "pickup" current and operation time to be changed. This extends the capabili- ties of the power circuit breaker over the molded case by providing not only short-circuit but also overload tripping adjustments, thereby allowing a broader range of applica- tions (7). Low-voltage power circuit breakers are available with or without direct-acting instantaneous units and with or without long-time-delay units. Furthermore, most manufacturers offer three different separately adjustable long-time-delay operation bands as well as three different short-time-delay operation bands. As with molded-case breakers, power breakers are available with either shunt- tripping or UVR units or both. Solid-state devices are also manufactured for all tripping arrangements. Some typical ratings for low-voltage power circuit breakers are provided in table 9.5 (7). In addition to these Table 9.5.— Some typical ratings for low-voltage power circuit breakers Ac system Rated nominal maximum voltage, V voltage, V Frame size, A 3-phase short-circuit current rating, symmetrical, A Range of trip-device current ratings, A 600. 480. 635 508 225 14,000 40- 225 600 20,000 40- 600 1,600 42,000 200-1 ,600 2,000 42,000 200-2,000 3,000 65,000 2,000-3,000 4,000 85,000 4,000 225 22,000 40- 225 600 30,000 100- 600 1,600 50,000 400-1,600 2,000 50,000 400-2,000 3,000 65,000 2,000-3,000 4,000 85,000 4,000 HIGH-VOLTAGE CIRCUIT BREAKERS The power circuit breakers used in high-voltage min- ing applications include air-magnetic, oil, minimum-oil, and vacuum types. Vacuum circuit breakers or VCB's are by far the most popular because of their small size and high efficiency. Oil circuit breakers or OCB's once were the most common, but their use has dropped substantially in recent years, since the interrupting sizes needed for min- ing are not available. Air-magnetic types are normally limited to surface installations. The next few paragraphs will examine typical apparatus ratings, and then the operation of oil, minimum-oil, and vacuum types will be described; air-magnetic breakers are excluded as their operation is the same as that presented previously for lower voltage breakers. Typical Ratings The typical nominal voltage ratings corresponding to nominal system voltages are 4,160, 7,200, and 13,800 V, with 23,000 V used in some strip mines. The system portions of interest are obviously ac. Common continuous- current ratings are 400, 600, 800, 1,200, and 2,000 A. The majority of mine systems do not call for current greater than 600-A continuous, which has become the most used rating. Interrupting and close-and-latch ratings are very im- portant high-voltage parameters (6). For low-voltage and medium-voltage circuit breakers, the two ratings are usu- ally the same. As high- voltage circuit breakers rarely terminate current flow until a few cycles after the first- cycle peak, the close-and-latch rating must be higher than the interrupting rating. A typical interrupting rating for high-voltage circuit breakers found in mining is 12,000-A rms symmetrical, while the typical close-and-latch rating is 20,000-A rms asymmetrical. The asymmetrical close- and-latch rating is often found by multiplying the sym- metrical interrupting rating by 1.6 (see chapter 10) (6). High-voltage circuit breakers can also be given an interrupting-capacity class, which is an identifying group- ing rather than a rating. It is expressed in megavoltam- peres, such as 250, 350, 500, and 750 MVA. The interrupt- ing capacity is related to the interrupting-current rating by (5) MVA = V3 kV^kA^a (9.1) where MVA = interrupting capacity, MVA, kV rated = rated system voltage, kV, and kAra^j = rated rms interrupting current, kA. Oil Circuit Breakers Even though their popularity has been dropping, OCB's are still used extensively in surface installations, 233 especially substations. The common type of construction is the dead tank, shown in figure 9.14A. This steel tank is partly filled with oil and has a cover with porcelain or other composition bushings or insulators through which the conductors are carried (4-5). The breaker contacts are located below the bushings and are bridged by a conduct- ing crosshead supported by a lift rod. In most designs, two contacts and the crosshead provide two interruptions per pole. The majority of OCB's in mining have three such poles in one tank. The tank has an insulated liner to prevent the arc from striking the tank walls. The entire assembly is oiltight; a vent with oil-separating properties permits the escape of any gases generated but prevents the escape of entrained oil. Arc interruption in high-voltage circuit breakers em- ploys the cathode-spot phenomenon combined with arc lengthening and deionization of the arc path. In the case of the OCB, oil is vaporized as an arc is established between the parting contacts, and this produces a bubble around the arc. The gases within the bubble are generally not conducive to ionization, but in most modern OCB's, an oil-filled insulating chamber surrounds the parting con- tacts (fig. 9.145). When the moving contact is lowered, the gas generated by the arc portion within the chamber forces oil out through the chamber throat (4). The blast of oil comes into intimate contact with the arc, accelerates the cooling and ion recombination process (fig. 9.14C), and carries away available ions. A different arc-chamber ap- proach is shown in figure 9.15. Here the chamber throat is made of laminations so that during interruption, the oil can move radially into the arc path. This is sometimes termed a turbo action. In high-interrupting capacities, the gases developed within the chamber can be used to blast oil horizontally across the arc path. Whatever the specific design, the chambers are intended to contain the devel- oped high gas pressures and reduce any pressure on the main oil tank (5). After being effectively cooled, the generated gases are allowed to pass through the vent into open air. The result of OCB construction and operation is a very effective arc interrupter. However, beyond availability, there are inherent disadvantages that discourage use of OCB's (4-5). The oil presents a fire hazard, particularly if the tank is ruptured because of unexpected pressure; this has led some States to prohibit OCB application in under- ground coal systems above 10,000 V. The oil is bothersome to handle and creates maintenance problems including cleanliness problems. Finally, the inertia of the heavy operating mechanism severely limits operational speed, causing a time delay in opening the arc. Despite these problems, other advantages, which are discussed in chap- ter 11, still make the OCB desirable to many industry engineers. When used underground, the physical size of three- pole units usually limits the interrupting capacity to 100 MVA or less, with continuous-current ratings of 400 A. The operating mechanism on these small OCB's is typi- cally spring-gravity and manual-reset; a handle-driven mechanism (quick break, quick make) is used to close the breaker manually while at the same time automatically tensioning an opening spring. With the breaker engaged, the spring becomes armed, allowing a shunt-trip or UVR device to trigger the breaker opening by releasing the spring. A motor-driven system is also available to close the breaker, but the tripping method is the same. The motor- driven OCB's can thus be electrically engaged as well as Moving contact Oil level Arc chamber Insulating lift rod Stationary contact Moving contact Insulating chamber Throat Figure 9.14.— Construction and operation of dead-tank OCB. r~\ • Moving contact Oil flows into throat between laminations Figure 9.15.— Turboaction arc chamber for OCB's. tripped. Larger OCB's such as those used in substations are typically motor driven. Minimum-Oil Circuit Breakers Minimum-oil circuit breakers, also termed low- volume oil or live tank, enclose each pole in its own small-diameter tank (5). In modern versions, the tank is made of insulated high-strength, high-resistance material, and the top and bottom covers are high-dielectric-strength insulators (fig. 9.16). Contacts consist of a movable vertical rod and a stationary contact in the tank bottom. Oil volume is about 1 L, and the top surface of the oil is at atmospheric pressure. Arc extinguishing is assisted by oil blast, and resulting gases are vented to outside air. The operating mechanism can be either manual-reset and spring-trip or motor-reset and spring-trip. Some typical ratings of these breakers are listed in table 9.6. The arrangement of a three-pole minimum-oil unit with moving contacts mechanically interconnected results in a smaller overall package than comparable dead-tank breakers. The smaller mass of moving parts (operating 234 Moving contact Stationa contact -Operating mechanism Figure 9.16.— Cross section of minimum-oil breaker. Table 9.6.— Typical minimum-oil circuit breaker ratings Rated voltage, V Interrupting capacity, MVA Continuous current, A 5,000.... 173 478 680 1,000 15,000 1,000 25,000 630 mechanism and rods) enables higher operating speeds, while the advantages of oil interruption are maintained. However, the low volume of oil is such that after about five operations, the oil level must be checked. Even though oil-level indicators are available, this can create a main- tenance problem in mining. Vacuum Circuit Breakers With all the circuit breaker types covered so far, a gaseous atmosphere exists between the parting contacts. The gas is ionized by many processes and thus provides free electrons, which move to the anode, and positive ions, which are attracted to the cathode (4). As the positive ions arrive at the cathode, they can cause thermionic or high- field emission of electrons, which has a negative effect on arc interruption. Almost all these phenomena cease to exist if the gas between the breaker contacts is removed; in other words, if the arc is drawn in a vacuum. For this reason, vacuum is considered an extremely good medium for switching, and circuit breakers have been developed to take advantage of this feature. Figure 9.17 shows a sketch of a VCB, again with one pole. The assembly is sometimes called a bottle. The main advantages of VCB's are • Interruption usually occurs at the first zero current; • There are no blind spots in their interrupting range; • They have extraordinarily long life; • They are relatively maintenance free; and • Recovery of dielectric strength (between the parting contacts) following interrupting is extremely fast. Moving contact \ Insulation 1 Metal bellows ^M U Stationary contact High vacuum Figure 9.17.— Cross section of VCB. These all result from the fact that the vacuum totally discourages ionization. An important aspect of VCB's is in the long service. For instance, if a unit fails to clear a short circuit beyond its interrupting range, but another unit down the line does, the exceeded VCB can be employed again up to the full rating without difficulty. Because of their efficient ratio of size to capacity, they are extremely well suited to underground mining use. Their interrupting capacity for large currents is such that they can be utilized anywhere on high-voltage distribution, usually without reservation. This flexibility has made the VCB the most popular high-voltage interrupter for distribution systems in min- ing today. Added to these advantages is the fact that the VCB does not have any physical orientation problems. This is a consid- erable constraint with OCB's, where the tanks must always be vertical. Vertical placement of VCB bottles is sometimes necessary, however, to minimize dust accumulation. Ironically, the high efficiency of vacuum interrupters, which has favored their wide application, is the same property that can lead to severe transients. If care is not taken with VCB installation, switching transient-related problems can occur throughout the mine electrical com- plex. A detailed discussion of this important problem is deferred until chapter 11 because of related phenomena. The operating mechanism, which includes the mount- ing structure for the vacuum bottles, is an important factor in proper VCB operation. As a result of the small contact travel distance, usually on the order of 1/4 in, four criteria are mandatory: 1. Rugged construction to withstand the shock and stress of equipment movement; 2. A firm, smooth closure motion to prevent contact bounce; 3. Forceful opening of contacts in the case of contact welding; and 4. Clean, smooth opening motion to prevent contact bounce and subsequent arc restriking. In most cases, manufacturers rely on a spring-reset and spring-trip mechanism to meet items 2 through 4, and figure 9.18 illustrates one approach. The closing opera- tion, also termed resetting or reclosing, may be manually or motor driven. The trip solenoid can be a shunt-trip or UVR device, and in some cases, both are used. In VCB applications, the compact size of the operating mechanism and mounting structure has made possible a substantial reduction in overall power-equipment dimen- sions. Manufacturers have even incorporated a disconnect 235 Contact opening spring (extended) Contact pressure spring (compressed Main contacts (closed) Bounce latch (disengaged) v.. Trip latch engaged ) V v Trip coil (deenergized) Reset latch (engaged) Mechanism rec losing lever Main contacts closed Contact opening spring (relaxed) Contact pressure spring (relaxed) Bounce latch (engaged) ^-Trip coil (energized) Motor Main contacts -— ^== 1 (open) ^Cp/ ! ff / _. . . . v ^s£W — Mechanism Reset latch reclosing lever (disengaged) Contact travel just completed after tripping Contact opening spring (relaxed) Contact pressure spring (relaxed Main contacts (open Bounce latch (disengaged) Trip coil (deenergized) J Motor turns lever to rec lose contacts Reset latch (disengaged ) Main contacts open, ready for reclosing Figure 9.18.— Operating mechanism for vacuum interrupter. (Courtesy McGraw Edison) switch in their designs (fig. 9.19). The operating mecha- nism for the switch is mechanically interlocked with the circuit breaker mechanism. If the switch is opened when the breaker is closed, the interlock trips the circuit breaker prior to switch-contact parting. FUSES The fuse is the simplest and oldest device for inter- rupting an electrical circuit under short-circuit or excessive-overload current (5, 7). Fuses are installed in series with the protected circuit and operate by melting a fusible link. The response is such that the greater the current, the shorter the time to circuit opening, that is, an inverse-time characteristic. Fuses may be used in ac or dc circuits, and there is such variation in their time-current Figure 9.19.— VCB assembly incorporating a load-break switch. (Courtesy Ensign Electric) characteristics that they are suitable for many special purposes. While circuit breaker contacts rely on external sensing, the fuse acts as both the sensing device and the interrupting device. Unlike circuit breakers, fuses are "one-shot," as their fusible element is destroyed in the circuit-protection process. Fuses are available with interrupting-current ratings up to 200,000-A symmetrical rms, much higher than the capacity of circuit breakers. Fuses are also available with current-limiting abilities to provide maximum protection for all circuit components. Fuses are normally classified as low voltage or high voltage: The low-voltage types are intended for service in systems 600 V and below, while the high-voltage varieties are suitable for installations 2.3 to 161 kV (7). LOW-VOLTAGE FUSES Plug fuses and cartridge fuses are the two principal categories of standard low-voltage fuses, and they are classified as non-time-delay, time-delay, dual-element, or current-limiting (13). There are also miscellaneous and nonstandard fuse classes. As with circuit breakers, there are three general fuse ratings (7); 1. Current. The maximum dc or rms ac, in amperes, which the fuses will carry without exceeding a specified temperature rise limit (available range: milliamperes to 6,000 A). 2. Voltage. The maximum ac or dc voltage at which the fuse is designed to operate (usual low-voltage ratings are 600, 300, 250, or 125 V ac or dc or both). 3. Interrupting. The assigned maximum short-circuit current that the fuse will safely interrupt (typical ratings are 10,000-, 50,000-, 100,000-, or 200,000-A symmetrical rms). Special ratings are also given to current-limiting fuses to specify the maximum current and energy the device will let through to the protected circuit when clearing a fault (7). 236 Plug fuses are rated at 125 V and are available with current ratings up to 30 A. Their use is thus limited to circuits with this voltage rating or less, except that they may be employed on systems having a grounded neutral where the maximum potential to ground of any conductor does not exceed 150 V (7). As a result, plug fuses have limited application in mine power systems (although an extensive popularity still exists for homes). Cartridge fuse applications, on the contrary, are widespread, to the point where mention of a fuse implies a cartridge. Figure 9.20 shows the three standard low-voltage cartridge-type fuses (7). Non-Time-Delay Fuses As the name implies, these have no intentional built- in delay. They have a very simple construction, consisting of two end terminals joined together by a copper or zinc fusible element. The link is more current sensitive to melting than to time. Non-time-delay fuses are available as one-shot (or nonrenewable) and renewable; the former is the oldest cartridge fuse type in use today (7). With the one-shot, the link is in a sealed enclosure and the entire cartridge must be replaced after interruption. The renew- able fuse can be disassembled, and the link replaced. The lack of intentional time delay and a limited interrupting rating of around 10,000 A have substantially reduced the popularity of these fuses in recent years. Time-Delay Fuses The metal alloy used in time-delay fusible links is not only sensitive to current but also to the time period involved. In other words, a specific current existing for a specified time period is necessary to cause the heat- melting energy of the alloy. Such an arrangement permits harmless high-magnitude, short-duration currents to ex- ist, which are oftentimes necessary for proper system operation, as in motor starting. Dual-Element Fuse Originally designed primarily for motor-circuit pro- tection, the dual-element fuse (fig. 9.21) combines the features of non-time-delay and time-delay units. The time- delay or thermal cutout is provided for overload protection, while two fuse link elements give short-circuit protection, blowing in a fraction of a cycle on heavy currents. The thermal cutout will allow the passage of currents as high as five times its continuous rating for up to 10 s. Hence, these fuses may be matched closely to protect the actual motor running current and at the same time be sized to protect wiring and other equipment, and provide both these functions without nuisance blowing. In fact, prop- erly sized dual-element fuses are required on all fuse- protected trailing cables. They are available with up to a 200,000-A symmetrical rms interrupting-current rating, and for further protection, most dual-element fuses also have a current-limiting feature. Current-Limiting Fuses Short-circuit protection requires that a fuse limit the energy delivered by the short circuit to a faulted compo- nent. Obviously, the energy any interrupting device lets through under fault conditions cannot exceed the pro- fl-rnn ^=2ZX^ Q) (8 T¥T Ferrule type Knife-blade type 0-60 A 70-600 A Figure 9.20.— Common cartridge fuses. <2> Bolt type 601-6,000 A Multiple-bridge short-circuit link Fiber tube Quartz sand filler Alloy time-delay element Figure 9.21.— Inside view of dual-element fuse. tected components withstand rating. Current-limiting fuses provide this protection by restricting or cutting off fault currents before damaging peaks are reached. With very high fault currents, they are extremely fast, limiting current in less than one-quarter cycle, with current inter- ruption occurring within the first one-half cycle. Only a portion of the destructive short-circuit energy that is available is let through. By this, the current-limiting fuse allows the use of lower momentary and interrupting ratings by cutting off current within equipment ratings (7). Figure 9.22 illustrates how the fuse operates: the large waveform represents the available short-circuit current on a faulted system, and the performance of the fuse is superimposed. Restricting energy is a means of limiting the mechan- ical and thermal stress imposed on equipment that is carrying fault current. To illustrate this energy, consider figures 9.22 and 9.23 and the peak let-through current, L. It has been found that the magnetic forces during a fault vary as the square of fault current, I 2 , (7). These forces translate to mechanical stress, which could damage trans- former frames, bus structures, or cable supports. The let-through energy, I 2 t, represents a measure of the heat- ing effect or thermal energy of the fault with or without the fuse (with the fuse, the value is I 2 ,). I 2 t actually equals ji 2 dt, the time integral of the current squared for the time under consideration (8). Both I 2 , and I^t can be consider- ably reduced when current-limiting rases are used (7). Furthermore, equipment with an I 2 t withstand rating can be matched with the energy let-through limit of the fuse. Standard Fuses As implied by the foregoing, cartridge fuses come in a wide range of types, sizes, and ratings. Various classes for 237 Peak available current h— I — *» II I i / '/ 2 cycle \ 1 A , jL ,*-— Arc \ r fy/x \ Peak let-through -fcH-Clearing current t j me Melting _| [_ Arcing time time Figure 9.22.— Current-limiting action of fuses. I 2 t Time rms current Figure 9.23.— Energy-limiting action of fuses. low- voltage units have been standardized (15), and a listing of general-purpose fuses follows (the first value listed is the range of continuous currents): Class G: to 60 A, 300 V to ground maximum, 100,000-A symmetrical rms interrupting, current limit- ing, fit only class G fuse holders. Class H: to 600 A, 250 and 600 V, interrupting capacity up to 10,000 A, either one-time or renewable construction, commonly termed the "old NEC fuse." Class J: to 600 A, 600 V, 200,000-A symmetrical rms interrupting, current limiting, fit only a class J fuse holders. Class K: to 600 A, 250 and 600 V, 50,000-, 100,000-, or 200,000-A symmetrical rms interrupting, have the greatest current-limiting effect of all low-voltage fuses (available as straight current limiting, dual-element cur- rent limiting, and dual-element time-delay current limit- ing), fit class H fuse holders. Class L: 601 to 6,000 A, 600 V, 200,000-A symmetrical rms interrupting, current limiting, bolt-in mounting. Class R: to 600 A, 250 and 600 V, 200,000-A symmetrical rms interrupting, current limiting similar to class K level 5 fuse, fit only class R fuse holders. Class T: to 600 A, 250 and 600 V, 200,000-A symmetrical rms interrupting, current limiting but effect less than class J fuses, fit only class T fuse holders. Nonstandard Fuses Nonstandard fuses receive their name because of their special dimensions or use in special applications; they are not general-purpose fuses (7). Of the many available, four have important applications in mining: Cable Limiters. These fuses are for use in multicable circuits (paralleled cables) and are placed in series with each cable in parallel. They are designed to provide short-circuit protection to each cable, removing it from power in case of failure. Cable limiters are rated according to cable size (AWG 4/0 and so forth). Semiconductor Fuses. These devices are available in two types: semiconductor-protection fuses or semiconductor- isolation fuses. Both are used in series with the applica- tion. Protection fuses are employed where solid-state de- vices are to be protected rather than isolated after a failure; they have lower let-through characteristics than other current-limiting fuses. A specific application is pro- tecting a rectifier or thyristor in case of an overload current. Isolation types are high-speed fuses, used to isolate a defective solid-state device in case of its failure. These are mandatory fuses for individual power diodes paralleled in large rectifier banks. Capacitor Fuses. Capacitor fuses are applied in series with power-factor (pf) correction (or other type) capacitors and are used to isolate a failed component by clearing short-circuit current before excessive gas is generated in the capacitor. Welding Fuses. These are current-limiting fuses for use in welder circuits only. The time-current characteris- tics are such that these fuses allow a longer intermittent overload than general-purpose fuses, but still provide short-circuit protection. HIGH-VOLTAGE FUSES High-voltage fuses provide usable protection for 2.3- to 161-kV systems and fall into two general categories: distribution fuse cutouts and power fuses (7). Distribution fuse cutouts were designed for overhead distribution cir- cuits, such as the protection of residential distribution transformers. Even though their employment in utility- type systems is extensive, their use in mining is limited and in some cases restricted. Power fuses are another matter, as certain types offer extremely practical protec- tion in mine power systems. They can be applied to substation, distribution, and potential transformers (in series with the primary) and occasionally to distribution circuit conductors. For surface mine systems, the fuses are often equipped with contacts, arranged so that the fuse and its mounting act as a disconnect switch (fig. 9.24). There are two basic power fuses, expulsion and current- limiting types, and the next few paragraphs will discuss their operation, ratings, and application. Expulsion Types As with low-voltage fuses, high-voltage types start the current-interruption process by the melting of a fusible link, but as might be expected, deionization of the atten- dant arc becomes the most substantial part of current termination. To help the process, as shown in figure 9.25, the link is held under tension by a coil spring; upon melting, the spring pulls the contacts apart, lengthening the arc (4). In expulsion fuses, gases are liberated from the lining of the current-interrupting chamber by the heat generated from the arc. Both the earliest form of expulsion 238 attachment Figure 9.24.— High-voltage power fuse and support. (Courtesy S&C Electric Co.) Fiber tube -, Boric acid Plunger Strain element Main fuse Gap -, Disk Figure 9.26.— Cross section of boric acid power fuse refill. B- Fusible link Spring Glass tube / * E Flexible lead Figure 9.25.— Fusible element under spring tension in high- voltage fuse. fuse and distribution fuse cutouts use a liner of organic material to deionize the generated gases by expelling them from the fuse holder tube to the surrounding air. The problem with this operation is the attendant flame expul- sion and loud noise. Hence, expulsion fuses are suitable only for outdoor usage, generally in substations remotely located from human habitation (7). The limited interrupting capacity (table 9.7) and unsuitability for indoor use of early expulsion fuses led to the development of the boric acid or solid-material fuse (7). Here, the interrupting chamber is made of solid boric acid. When exposed to arc heat, the material liberates steam, which can be readily condensed to liquid by venting the gas into a cooling device. The result is an operation with negligible or harmless flame and gas emissions and noise levels. The range of voltage, continuous current, and interrupting ratings is also greatly expanded. High-voltage boric acid fuses are manufactured in two styles (7): the fuse unit (nonrenewable), where the fusible unit, interrupting element, and operating element are all combined in an insulated tube; and the refill unit or fuseholder (renewable), where only the refill unit is re- placed after interruption. Figure 9.26 shows the internal components of a refill unit, while figure 9.27 illustrates the construction of the entire fuse. Table 9.7 provides a list of typical ratings for both styles. The fuse-unit style is intended for outdoor use at system voltages of 34.5 to 138 kV, while the refill unit can be used indoors or outdoors on the surface at 2.4 to 34.5 kV. KEY A Fuseholder B Spring-and-cable assembly (copper cable carries load current ) C,D Fuseholder upper contacts and latch E Fuseholder lower contacts and latch F Refill unit Figure 9.27.— Disassembled refill unit for boric acid fuse. (Courtesy S&C Electric Co.) Current-Limiting High-Voltage Fuses High-voltage current-limiting or silver-sand fuses have the same advantages as previously discussed for low- voltage fuses and are of two different forms (7): those to be used with high-voltage motor starters for high-capacity distribution circuits at 2,400 and 4,160 V and those for use with potential, distribution, and small power transformers from 2.4 to 34.5 kV. The operation of either form is such that the arc established by the melting of the fusible element is subjected to mechanical restriction by a powder or sand filler surrounding the fusible element. The tech- nique provides three important features: • Current is interrupted quickly without arc-product or gas expulsion. This allows use indoors or in small-size enclosures on the surface or underground. There is no noise from the operation, and since there is no gas or flame discharge, only normal electrical clearances need by met. Table 9.7.— Ratings of high-voltage power fuses 239 .- , . fc , Boric acid fuse, Boric acid fuse, Nominal Expulsion-type fuse 1 -shot type renewable rating, Maximum Maximum Maximum Maximum Maximum Maximum kV continuous interrupting continuous interrupting continuous interrupting current, A rating, MVA 1 current, A rating, MVA 1 current, A rating, MVA 1 2.4 — — — — 200,400,720 155 4.16 — — — — 200,400,720 270 4.8 — — — — — — 7.2 100, 200, 300, 400 162 — — 200, 400, 720 325 14.4 100,200,300,400 406 — — 200,400,720 620 23 100,200,300,400 785 — — 200,300 750 34.5 100,200,300,400 1,174 100,200,300 2,000 200,300 1,000 46 100,200,300,400 1,988 100,200,300 2,500 — — 69 100, 200, 300, 400 2,350 100, 200, 300 2,000 — — 115 100,200 3,110 100,250 2,000 — — 138 100,200 2,980 100,250 2,000 — — 161 100, 200 3,480 — — — — 1 3-phase symmetrical rating. NOTE.— Dashes indicate that standard fuses are not available in the specific voltage rating. Current-limiting fuse Maximum Maximum continuous interrupting current, A rating, MVA 1 100, 200, 450 155, 210, 360 450 360 100, 200, 300, 400 310 100, 200 620 50, 100, 175,200 780-2,950 50, 100 750-1,740 40,80 750-2,600 • Very high interrupting ratings are available so these fuses can be applied on systems with very high short-circuit capacity (within their voltage rating). • All of the advantages of current-limiting action are available for high voltage. Table 9.7 provides a listing of typical ratings for current-limiting fuses. Instead of being rated by current, these fuses can also be "E-rated" (for instance, 100 E instead of 100 A), "C-rated," or "R-rated." The specifica- tions for E and C ratings are as follows: • E-rated fuses: 100 E and below, open in 300 s at an rms current within the range of 200% to 240% of the continuous rating of the fuse element; above 100 E, open in 600 s at an rms current within the range of 220% to 264% of the continuous (or E) rating; • C-rated fuses: open in 1,000 s at an rms current within the range of 170% and 240% of the C ratings. E-rated fuses are considered as general-purpose or backup fuses, while R-rated devices are intended for use with high-voltage motor starters (7). Load-Break Switches Fused load-break switch It is possible that after the occurrence of a short circuit on a fuse-protected three-phase system, only one of the three fuses could open. Here current through the remaining two fuses might be reduced so that they do not open. The system then becomes single phased, which can cause serious damage to equipment. In a low-voltage circuit, dual-element fuses that are closely matched to the overcurrent point can usually handle the situation. On high-voltage systems, the problem is much more difficult when protection is by fuses alone. However, to take advan- tage of the lower cost of fuses and load-break switches versus the cost of a high-voltage circuit breaker, some manufacturers produce load-break switches with incorpo- rated high-voltage fuseholders. An example is shown in figure 9.28 where the fuses are interlocked to trip the operating mechanism of the switch if one or more of the fuses fail. Interlocking is usually accomplished with spe- cial high-voltage fuses that contain a spring-loaded plunger. Fuse activation releases the plunger, which trips h j J< I Switch -»-» XT\ Large fuse Actuator fuse Remote signaling circuit or switch- opening circuit Schematic showing interlocks Figure 9.28.— Load-break switch with interlocked high- voltage fuses. (Courtesy Line Power Manufacturing Corp.) 240 the switch mechanism. Precautions must be observed when using or considering these devices, and these are discussed in chapters 12 and 13. RELAYS Relays perform a major role in power-system protec- tion, where their purpose is to detect voltage and current anomalies. They normally receive information about sys- tem conditions through transformers or resistors, which reduce system parameters down to levels that the relays can handle. Upon detection of a problem, a relay operates to supply or remove control power to the shunt or UVR tripping elements of the switching apparatus. Because of their function, relays are sometimes called sensing devices. While transformers might also be consid- ered sensing devices, their function in protective relaying is solely as transducers. There are four basic relay types: thermal, electromag- netic attraction, electromagnetic induction, and static. (D'Arsonval movements are actually considered another relay type, but their operation is completely covered in chapter 5). The first three are electromechanical devices, and the following paragraphs will present their operation. Static or solid-state relays are discussed in chapter 12, because of related content. Relay Terminology and Types When a relay operates, it is said to close or open its contacts (9). Most relays are restrained by spring control and assume a specific position, either open or closed, when deenergized: hence there is a normally closed or NC contact and a normally open or NO contact. Symbols for both situations are shown in figure 9.29. When a relay operates to open NC contacts or close NO contacts, it is said to pick up the contacts, and the smallest actuating quantity to cause contact operation is referred to as the pickup value. When a relay operates to close NC contacts or open NO contacts, it is said to reset or drop out, and similarly, the largest actuating quantity to cause reset is the reset value of the relay. When the relay is deenergized to reset, the reset value is almost always greater than zero and is often specified as a percentage of normal operation. Most relays have adjustments or tap settings to adapt them to as wide an operating range as possible. The word describing relay operation has a formal meaning; for example, overvoltage relays, overcurrent re- lays, overtemperature relays, and so forth. Here the suffix refers to the actuating source (voltage, current, etc.), and the prefix "over" means that the relay picks up to close a set of NO contacts (or open NC contacts) when the actuat- ing quantity exceeds the magnitude at which the relay is adjusted to operate. Similarly, undervoltage, undercur- rent, and undertemperature relays reset to close NC contacts (or open NO contacts) when the actuating quan- tity decreases below a predetermined level. Some relays have both "over" and "under" functions (7, 10). Even with these definite meanings, common usage of relay terminology is rather straightforward. Pickup is used to refer to the point where the relay changes from its normal state to indicate a malfunction, while reset implies that the relay returns to its normal position. The normal position may occur when the relay is energized or deener- gized and depends on the application. Relays designed for protective circuits are usually provided with some means of visual indication that a specific relay has operated to trip a circuit breaker. These operation indicators or targets are often brightly colored and are operated mechanically or electrically. Specific relay types have been developed to meet special or general system-protection needs. Thermal relays serve directly or indirectly to measure power-system tem- peratures. Electromagnetic-attraction relays are used to instantaneously detect voltage and current changes. Electromagnetic-induction relays allow a time delay be- tween relay detection and contact action. Directional re- lays can sense the direction of current flow. Thermal Relays Thermal relays most commonly employ bimetallic- driven contacts with an operation similar to that described for the molded-case circuit breakers. Another approach is to use ambient temperature, as in the temperature- monitoring protector shown in figure 9.30. This is a sealed bimetallic thermostat that opens or closes at a specific temperature; it can be used, for example, to sense motor overtemperature if mounted against the end turns of a motor winding. Yet another bimetallic approach is to employ a heater element within the relay enclosures, connected in series with the circuit under consideration, as illustrated in figure 9.31A. The relay trip point for opening or closing the contacts is expressed in amperes, but is also a function Normally open Normally closed Figure 9.29.— Relay contact symbols. Insulation Device i^ss Bimetallic strip Figure 9.30.— Temperature-monitoring protector. Heater Thermal relay unit —•-To motor To • magnet coil Normal position A Bimetallic relay B Melting-alloy relay Figure 9.31.— Electromechanical-thermal relays. 241 of time and is determined by the heater rating. The trip setting is commonly based on a 40° C ambient tempera- ture, but the relay may be ambient or nonambient com- pensating. Most relays of this type must be manually reset after tripping. An electromechanical-thermal device not using bime- tallics is the melting-alloy or eutectic-alloy relay, figure 9.31B. Being shock resistant and having high contact force, this is considered one of the most reliable thermal relays available, but because of its cost, it is not nearly as popular as the bimetallic type. The alloy melting point is extremely precise and is again related to a specific current-time characteristic. The relay can be reset after tripping and alloy resolidification. Two other thermal devices, resistance or thermistor types and thermocouples, operate with associated elec- tronic equipment to provide very precise temperature sensing and relaying. Here, for example, a probe can be inserted or embedded in a transformer or a motor winding to provide a spot temperature response. This type of device is very popular especially where large horsepower or capacity is involved. Electromagnetic-Attraction Relays There are three electromagnetic-attraction relays in common use: the solenoid, the clapper, and the polar (20). Although their operational speed might vary, all are considered instantaneous relays, since there is no built-in delay for pickup or reset. The solenoid and clapper types are available for ac or dc and are voltage or current actuated. Coil impedance is high for voltage and low for current. Polar units are dc sensing only, but may be used on ac circuits through rectification. All electromagnetic relays are available with NO contacts, NC contacts, or both. In solenoid units, the relay contact movement is initiated by a plunger being drawn into a cylindrical solenoid coil. Typical operating times are 5 to 50 ms, with the longer times associated with operation near the min- imum pickup value (20). A cross-sectional sketch of a solenoid relay is given in figure 9.32A. Four different clapper relays are shown in figure 9.32B. These have a magnetic frame with a movable armature and operate by the attraction of the armature to Adjusting core screw Coil area I- Magnetic frame Nonmagnetic ring - Plunger Contacts Helical spring A Solenoid-type relay Mainspring — M\ Core- ^f^r Magnetic frame Coil Residual pin Target - Armature Moving contact Indicating contact switch (ICS) ^-Normally closed (break) contact "-Normally open (make) contact Contact multiplier Adjustable core - Target ■-Magnetic frame M ^ in 9 contact i ji r ■■ Stationary__^f-~^n /Armature 011 contact jpfer__ ^j|g~ ~ Armature adjusting screw jydfcLaTloop Insulating card- " p__^_ 52 Armature Residual plating Coil- Mi! J) Moving contact Lag loop ^zy Indicating instantaneous trip(IIT) Core ^--Magnetic frame High speed B Clapper-type relay Figure 9.32.— Solenoid and clapper relays. (Courtesy Westinghouse Electric Corp.) 242 an electromagnetic pole {20). The armature controls the pickup or reset of contacts. As illustrated in figure 9.33, polar relays have a hinged armature in the center of the magnetic structure, which is here shown as an electromagnet but may be a permanent magnet. The relays operate when dc is applied to the actuating coil, and the polarity of the actuating source determines armature action, be it stationary or movement in either direction (10). In some units there is no retaining spring, and through a combination of con- tacts, the relays can sense actuating current through the coil in either direction (20). The pickup and reset values of clapper units are less precise than those of solenoid and polar relays; thus, clapper relays are used often as auxiliary or go, no-go devices (20). A common use for polar relays is in dc circuit protection where the actuating source is obtained from a shunt or directly from the circuit (10). A characteristic that should be considered when ap- plying any electromagnetic-attraction relay is the large difference that can exist between pickup and reset values. When an attraction relay picks up, the air gap is short- ened, and a smaller coil current is needed to retain pickup. Thus, the reset current may be much lower than the pickup current. The disparity is usually expressed as a percent ratio of reset current to pickup current, and is less pronounced in ac than dc relays. The ac relays can have a reset up to 90% or 95% of pickup, but dc ratios range from 60% to 90% (10). This is no problem in overcurrent appli- cations where relay coil current drops to zero after pickup, but it is a concern where reset values are important. Electromagnetic-Induction Relays Contacts Stop o Control < spring < Actuating coiU umx >■■■ Stop Movable armature J Polarizing I COil v Pivot r Polarizing magnet To actuating quantity Figure 9.33.— Polar relay. Butt joint Main coil Lag coil Keeper - Disk air gap Figure 9.34.— Common induction-disk relay. Electromagnetic-induction relays are of two general types: induction disk and cylinder (20). Depending on the design, the induction-disk unit can be either a single- quantity or directional relay, whereas cylinder relays are intended to be directional. A single-quantity relay, as might be supposed, is actuated by and compares two sources (10). The most commonly used time-delay relays for system protection employ the induction-disk principle (7). Single Quantity Single-quantity time-delay relays of the induction- disk type use the same principle of operation that was described for induction motors in chapter 6, but the physical construction is quite different (20). A sketch of an elementary induction-type device is shown in figure 9.34, and most time-delay relays in use today have this arrange- ment. The disk, made of aluminum, is mounted on a rotating shaft restrained by a spring, and a moving contact is attached to the shaft (fig. 9.35). On one side of the disk is a three-pole electromagnet; the other side has a common permanent magnet or keeper. The operating torque on the disk is produced by the electromagnet, and the keeper provides a damping action or restraint after the disk starts to rotate. The retarding effect of the keeper creates the time delay or desired time characteristic of the relay. Figure 9.35 is a front-view illustration of an actual induction-disk relay removed from its drawout case; all important components are indicated. The unit pictured is for overcurrent, but overvoltage and undervoltage relays are also available and are identical in construction except for the electromagnet coil rating. nstonforssous unit CQiibrotion plate Control spring Crodfe Chassis contact block Figure 9.35.— Front view of induction-disk relay removed from case. (Courtesy General Electric Co.) The control spring carries current for the moving con- tact. If the actuating quantity driving the electromagnet is of sufficient magnitude and is sustained for enough time, the disk will rotate until the moving contact touches the station- ary contact. (Some relays use a lever on the moving disk that forces a pair of stationary contacts to close, so that no current 243 flows through the control spring and disk.) Pickup of these main contacts triggers the seal-in or time-delay element, which is an electromagnetic-attraction relay with its coil in series and contacts in parallel with the main contacts. When activated, this relay picks up and seals in, thus lightening the current-carrying duty of the main contacts as well as operating a target indicator. After pickup, it usually must be reset manually. The tap block at the top of figure 9.35 is to allow different tap settings on the electromagnet coil. Table 9.8 lists the tap settings generally available in overcurrent relays (7), but some relays have wider ranges than those shown. Each range represents a different operating coil. Voltage relays have a narrower range of adjustment, because they are usually expected to operate within a limited change from the normal magnitude of the actuat- ing quantity (10). Be it a voltage or current relay, the coil and its tap settings are normally selected with respect to the ratios of the potential or current transformer used. Table 9.8.— Common current ratings of induction-disk overcurrent relays Time-delay elements Typical * instantaneous sl©m6nts Coil ran 9 e ' A Tap settings, 1 A adjustment range, A 0.5 to 2.5 0.5, 0.6, 0.8, 1.0, 1.2, 1.5, 2.0, 0.5- 4.0 2.5 2-16 1.5 to 6.0 1.5,2,2.5,3.0,3.5,4.0,5.0,6.0 10 -80 20 -160 4.0 to 16 4.0, 5.0, 6.0, 7.0, 8.0, 10, 12, 16 (^) 1 Tap settings will vary slightly according to manufacturer. Additional units are available for each time-delay range. 3 Not adjustable. separate and relatively independent adjustment of the relay inverse-time characteristics. They are preset by the manufacturer, and the common responses are "inverse," "very inverse," "extremely inverse," "short time," and "long time," the first three being the most popular in mining. A comparison of these responses is given in figure 9.37. The need for a specific response depends upon the application, and a few thoughts in terms of overcurrent relays follow (6). When the available fault-current magnitudes vary considerably, faster overall protection is usually gained with an inverse-time response. Very inverse curves provide the best overall protection where fault current remains MAGNITUDE OF ACTUATING QUANTITY Figure 9.36.— Inverse-time curve compared with definite- time curve. As shown in figure 9.35, overcurrent disk relays often have a second (auxiliary) ac-operated instantaneous ele- ment, which is a clapper-type relay (7). The unit is contin- uously adjustable over a calibrated range, and table 9.8 lists some of these representative values. This relay oper- ates in series with the time-delay operating coil and is usually set to operate instantaneously at a current pickup value higher than that of the time-delay element. How- ever, since the same actuating source drives both ele- ments, the instantaneous-relay setting must be coordi- nated not only with the same source but also with the timed element. The instantaneous contacts can be in parallel with the time-delay contacts or can be connected to separate terminals. The unit also has a target indicator, which normally requires manual reset after tripping. The operational characteristic produced by the induction-disk principle is termed inverse time. Although mentioned earlier in this chapter, the inverse response is illustrated again in figure 9.36 to emphasize that the operating time becomes less as the magnitude of the actuating quantity is increased (10). The more pronounced this effect becomes, the more inverse the curve is said to be. All relay time curves are actually inverse, with the exception of a theoretical definite-time response. By defi- nition, definite-time characteristics imply that the operat- ing time of the relay is unaffected by the magnitude of actuating quantity. In reality, an actual definite-time curve is very slightly inverse (fig. 9.36). Regardless, the term definite time is normally applied to all fixed-time relays that approach this response. The control-spring tension, the damping magnet, and the magnetic plugs (A and B of figure 9.34) provide 100.00 60.00 30.00 10.00 6.00 UJ 1 1- Extremely inverse 10 20 50 MULTIPLES OF TAP VALUE CURRENT Figure 9.37.— Various time characteristics of induction units. 244 constant (detection of the fault, as seen by the relay, is mainly a function of fault location). Extremely inverse relays are designed to coordinate rather closely with power fuses and distribution cutouts and are also used in systems that have large inrush currents. The actual application of these characteristics in the mine is given in chapter 13. The operating time of an induction relay can usually be adjusted by selecting the distance of rotor travel from the reset to the pickup position (20). This is accomplished by adjusting the rest position of the moving-contact stop. The time dial, with evenly divided markings, facilitates positioning. When the response of the relay for different time dial settings is plotted, the result is a family of curves, an example of which is shown in figure 9.38. Current is plotted in terms of multiples of pickup, which enables the curves for a specific relay to be used with any tap setting. Directional The basic ac directional electromagnetic-induction relay or cylinder unit in common use is sketched in figure 9.39. Its operation is similar to that of an induction motor that has salient poles for the stator, except that here the rotor iron is stationary and only the rotor conductor is free to rotate (10, 20). The rotor conductor is a thin- walled aluminum cylinder, and the two actuating quantities, causing I x and I 2 , independently produce torque on the cylinder. The cylinder drives a moving contact whose travel is restricted to a few degrees by the stationary contact and stops. Reset torque is established by a spiral spring. The ac directional relays are used to distinguish between current supplied in one direction or the other in an ac circuit, by recognizing phase-angle differences be- tween the two actuating quantities (10). (Conversely, a dc directional relay, or polar unit, recognizes differences in polarity.) To perform the ac comparison, one actuating value is used as a reference or polarizing quantity. There- fore, the polarizing quantity phase angle must remain fixed while the phase angle of the other fluctuates widely. One application of this technique is in power relays where the unit is polarized by circuit voltage, with circuit current being the other actuating value. Through this, the cylin- der detects power flow in one direction or the other. Another important application is an ac directional relay combined with an overcurrent relay, as shown in figure 9.40. Here, tripping occurs only when the current has a specific relationship to the voltage, and power flow is in the tripping direction. BASIC RELAY CONNECTIONS In order to sense a malfunction and then supply tripping energy to the appropriate circuit breaker, a relay must be attached in some manner to the power system. Circuit connections for protective relaying are basically not too different from those discussed for instrumentation in chapter 5. Here, however, the relay coil receives the input information, and its contacts pick up or reset, thus affecting the control power to the circuit breaker. Direct relay connections to the monitored circuit are often re- stricted to low- voltage, low-power circuits because most relay current or voltage coils are designed to operate in the vicinity of 5 A or 120 V (4). Obviously, if power-system values exceed these levels, some interface is needed be- UJ »- 3 3 \ \ Tie c ettir 10 K iia g 1 \ \ \ Ti \. s 1 1 i .5 .3 _,, I "■-- ££;;; .2 1 i/T 1.5 2 3 4 5 ( 5 7 8 9 10 15 20 1.5 2 MULTIPLE OF PICKUP Figure 9.38.— Family of inverse-time characteristics. t 1 2 1 . n V . i 2 1 j^ / )V *, \ \ 1 ^ 'I! ,1 , — JA^ r Cylinder • > Inner core @ @ ti 2 v F 3 lug Laminations Figure 9.39.— Cylinder directional relay. tween the monitored circuit and the relays. Again, instru- ment transformers for ac and resistors for dc are used, a subject also introduced in chapter 5. There are five basic relay connections used for protec- tive relaying in the mining industry. For ac systems, these are direct, potential, and differential; and for dc work, direct and potential are used. Differential relaying is also available for dc, but the circuitry is not considered basic. Although some of the techniques are employed much more frequently than others, this section serves to introduce all these connections. Alternating Current Direct Relaying Direct relaying is used to sense the magnitude of current flow. As shown in figure 9.41 A, its simplest form consists of a current transformer (CT) secondary connected to a relay operating coil. Relay pickup current is thus a 245 function of line current. For instance, consider that the transformer ampere-turns ratio or current rating is 50/5 A or 10/1 and the relay pickup setting is at 0.5 A. This relay would theoretically pick up its contacts when line current is (10X0.5) or 5 A. The purpose of this connection is therefore to provide protective relaying for current in any conductor. The important items to consider in direct relaying are concerned with matching the performance of the CT with that of the relay. IEEE standards provide most of these (7). 1. Ratios. As an obvious starting point after the foregoing example, standard ratios are listed below: Single-ratio CT, amperes: Coil terminal 10/5 200/5 2,000/5 15/5 300/5 3,000/5 25/5 400/5 4,000/5 40/5 600/5 5,000/5 50/5 800/5 6,000/5 75/5 1,200/5 8,000/5 100/5 1,500/5 12,000/5 3-ratio CT with centered-tapped secondary, amp 25/50/5 400/800/5 50/100/5 600/1,200/5 100/200/5 1,000/2,000/5 200/400/5 1,500/3,000/5 300/600/5 2,000/4,000/5 Multiratio CT with multitapped secondary, amperes (cur- rent ratings higher than those shown are also available): 600/5 Rating 1,200/5 2,000/5 Taps 50/5 100/5 150/5 200/5 250/5 300/5 400/5 450/5 500/5 600/5 100/5 200/5 300/5 400/5 500/5 600/5 800/5 900/5 1,000/5 1,200/5 300/5 400/5 500/5 800/5 1,100/5 1,200/5 1,500/5 1,600/5 2,000/5 Figure 9.40.— Directional overcurrent relay using induction- disk relay and cylinder relay. CT Line to be monitored / L, H, Relay operating coil A Circuit connections '2 x.4 n 2 ■ • X? B Instantaneous current Figure 9.41.— Direct relaying in ac systems. The double-ratio and multiratio types provide flexibility through secondary taps. These values are for bushing-type or window-type CT's, which are the most popular in the industry. All these have the standard 5-A-rated secondary current. 2. Secondary Current. The continuous-current rating of the secondary should be at least equal to the actual drain, but a full-load secondary current of 3 to 4 A is normal practice. An oversized CT is bad practice, as the percent error is much greater than with a correctly rated CT 3. Short-Time Ratings. Both thermal and mechanical ratings should be considered. The thermal short-time value relates to the maximum symmetrical rms primary 246 current that the CT can carry for 1.0 s without exceeding its maximum specified winding temperature. The mechan- ical rating refers to the maximum asymmetrical rms current the CT can withstand without damage. In both cases, the rating is made with the secondary short- circuited. 4. Voltage Rating. Standard voltage ratings are 600, 2,500, 5,000, 8,700, and 15,000 V, and are the same as insulation classes found in mine systems. The CT will operate continuously at 10% above rated voltage without insulation failure. 5. Burden. As defined in chapter 5, burden is the load connected to the CT secondary; expressions used are volt-amperes at a given power factor or an impedance with a power factor. The power factor is that of the burden. Table 9.9 lists standard values for CT's at 60 Hz. Relay burdens are so varied they cannot be listed, but chapter 10 shows how CT burden and relay burden can be compared. 6. Accuracy. Accuracy of a CT relates to its transfor- mation ability. In protective-relaying applications, accu- racy is not only important at normal circuit currents but also at fault-current levels. The problem in CT's is that core saturation leads to poor accuracy or ratio errors. Accuracy class designations use a C or T identifying letter followed by a classification number. C states that percent ratio error can be calculated, whereas T means that the value has been found by testing. The classification number relates to a standard secondary voltage of 10, 20, 50, 100, 200, 400, or 800 V. At this voltage, the CT will deliver to a standard burden, 20 times normal secondary current with 10% ratio error or less, and it will not exceed 10% with any current from 1 to 20 times rated current with a lesser burden. (For example, C200 relates that for a 2.0-Q burden, (20X5) or 100 A can be delivered from the CT without exceeding 10% error. This error can also be calculated.) 7. Polarity. Polarity relates to the correct phasing of primary and secondary currents, and figure 9.4 IB shows the relative instantaneous directions of current as per standard markings. This allows correct connections when more than one transformer is used, which is imperative in three-phase systems. As can be seen in the foregoing listings, actual man- ufacturer specifications should always be consulted before attempting to match CT's with relays for direct-relaying applications. Alternating Current Potential Relaying Potential relaying is as simple as direct relaying and enables circuit voltage to be monitored. Figures 9.42A and 9.42S show two applications: sensing voltage across a resistance and between two conductors. There is a poten- tial transformer (PT) between the circuit and the relay. Figure 9.42C gives the polarity correspondence of instan- taneous voltages between the primary and secondary windings as well as conventional transformer markings. Standard PT's are single-phase, two-winding units con- structed so that the primary and secondary voltages always have a fixed relationship (7). To visualize the operation, consider figure 9.42A. The transformer is rated 2,400/120 V or has a 20/1 ratio, an overvoltage relay is used, and the relay coil is rated at 120 V with the contacts set to pick up at 80% of rated. The contacts will therefore pick up when 1,920 V exists across the resistor. IEEE standards also provide guidelines for PT utili- zation, and a summary of these follows (7). In general, they are less rigorous than those for CT's. 1. Voltage. Standard voltage ratios are available in table 9.10. When applied to sense voltage between two conductors, the nominal system voltage should be within + 10% of the transformer nameplate rating. When used in three-phase mining systems supplying portable or mobile equipment, primary connections must be line to line. Special ratings, providing other than the standard 120-V secondary, are usually available. Table 9.9.— Standard burden for current transformers „ , . , ._.. Characteristics for 60 Hz and Standard Genera characteristics _ . . , 2 oianaara 5.^ secondary current 2 burden designation 1 Resistance Inductance Impedance Apparent power pf (R), fi (L), mH (Z), fi (S), VA B-0.1 0.09 0.116 0.1 2.5 0.9 B-0.2 .18 .232 .2 5.0 .9 B-0.5 45 .580 .5 12.5 .9 B-1 .5 2.3 1.0 25 .5 B-2 1.0 4.6 2.0 50 .5 B-3 2.0 9.2 4.0 100 .5 B-4 4.0 1^4 aO 200 .5 1 B-0.1, B-0.2, and B-0.5 are usually applied for metering purposes; B-1 through B-4 are usually applied for relaying. 2 At 5 A, S = l 2 2; for example, for B-2, S = 5 2 2 = 50 VA. Table 9.10.— Standard ratings for potential transformers Primary, V 120 (Seconds Ratio 1/1 2/1 4/1 5/1 20/1 35/1 40/1 60/1 iry, 120 V) Primary, V 8,400 Ratio 70/1 240 12,000 100/1 480 14,400 120/1 600 24,000 200/1 2,400 .. 36,000 300/1 4 200 48,000 400/1 4 800 72,000 600/1 7,200 •Line monitored , Relay coil Lines monitored A Resistance B Between conductors Instantaneous direction of ■ Voltage- drop resistor Direct relaying Potential relaying Figure 9.44.— Dc direct-relaying connections. 248 KINDS OF PROTECTION Several relaying terms describe the protection re- quired in many mine power systems: 1. Undervoltage, 2. Overload (sometimes called overcurrent), 3. Short circuit, 4. Ground overcurrent (or ground fault), 5. Ground continuity, and 6. Overtemperature. Classifications such as these are known formally as kinds of protection. The first five are necessary protection on all portable and mobile mining equipment, although excep- tions are provided within Federal regulations (17). This section expands the basic relaying material by describing how each kind of protection is used in the mine power system. The content is mainly pointed at high-voltage, three-phase ac mining systems and in general is restricted to relaying external from circuit breakers. Accordingly, these kinds of protection imply the following parameters: • Line-to-line voltages for undervoltage, • Line overcurrent for overload, • Three-phase or line-to-line faults for short circuit, • Faults causing zero-sequence current for ground overcurrent, and • Grounding-conductor resistance for ground continuity. Even though overtemperature is listed as item 6, it is usually applied to protect a specific component; thus, it will be discussed in chapters 12 and 13. Control Wiring Figures 9.45 and 9.46 show simplified diagrams of typical control wiring interconnections among the power source, relay contacts, and circuit breaker tripping ele- ments. In both diagrams, a potential transformer supplies 120 Vac with its fused primary connected line to line. Figure 9.45 illustrates cases where the tripping ele- ment is a UVR. The contacts can either reset to remove power from the coil (contacts in series with coil) or close to short it out (contacts parallel the coil). In the latter case, it can be seen that a resistance is placed in series with the contacts. In fact, the UVR itself will trip the breaker if control voltage is decreased in the range of 40% to 60%. Basic shunt-tripping connections are given in figure 9.46, where power is supplied to the element to cause tripping. Here the contacts for the various protective relays are paralleled, and the combination is in series with the trip coil; closure of any contact trips the breaker. It should be obvious in figure 9.46A that the power causing tripping is ac, while in figure 9.46B, it is dc. The capacitor in the dc circuit is employed for energy storage to augment tripping if there is a drop in the PT primary voltage when a relay contact closes. Phase Protection Phase protection by protective relaying can be over- load, short circuit, or both, depending upon the relays used. (Molded-case circuit breakers afford this same flexi- bility depending upon the internal tripping element used.) Time-delay relays are employed for overload, with instan- taneous units for short circuit. Figure 9.47 illustrates the combined protection for three line conductors, using an Power to load Relay contacts open to trip Removing power to UVR Fuses Power to load I Limit PT resistor ' UVR w» — t — *~0 — I T X Relay contacts close to trip Shorting out UVR Figure 9.45.— Typical control wiring for UVR. Line-to- line Manual U * ip Line-to- Ime PT Manual trip H f f oo— » . Shunt O trip coil N Shunt O trip coil A Simple B Capactor tripping Figure 9.46.— Typical control wiring for shunt-tripping ele- ment. Time-trip contact Instantaneous trip contact To trip circuit • Instantaneous-element current coil Time-element current coil 1 Protected equipment Figure 9.47.— Three-phase overcurrent and short-circuit con- nections. 249 induction-disk relay (7). The three current transformers are placed in wye, driving the wye-connected operating coils. The time-delay element is set on as low a tap setting as practical, enabling protection for sustained moderate overloads. The instantaneous units, however, are set to pick up on a current value slightly higher than the maximum peak load, thereby affording protection against short circuits or enormous overloads. The device numbers that were presented in chapter 4 are used extensively to describe the relay function. The number 51 signifies time-delay relays for ac overcurrent, and 50 is used for instantaneous devices. A combination instantaneous and time-delay ac overcurrent relay is often noted by 50/51. If the connections are as shown in figure 9.47 and transformer phase-angle errors are ignored, the secondary currents of each CT are in phase with the primary cur- rents, and each relay responds to abnormal conditions for its respective line (10). This also applies to figure 9.48A where the 50 elements are omitted. If line currents are approximately balanced, short-circuit protection for all three lines can also be provided with an open-delta con- nection, as in figure 9.485 (10). As might be expected, this approach is not as precise as the straight wye connection, but a third overcurrent relay may be inserted in the CT common connection for backup protection (see chapter 5 for a similar discussion on instrumentation). An advan- tage of the wye connection that is lost when the open-delta approach is use is the ability to sense zero-sequence currents through residual relaying. Thus, two current transformers are rarely applied as the only means of circuit protection. Ground Overcurrent To this point, the chapter has basically considered power-conductor protective relaying, and the extremely important subject of ground-fault protection has received only terse reference. Various relay configurations may be utilized to provide ground-overcurrent protection, some of which are quite elaborate. However, nearly all these techniques fall into one of five broad classifications (7): direct relaying, potential relaying, residual connection, zero sequence, and broken delta. Direct relaying, potential relaying and zero sequence are frequently used in resistance-grounded mine power systems, with zero- sequence relaying being the most popular. Unless other- wise noted, the following discussion will assume that the system is resistance grounded. The point of application for direct and potential ground-fault relaying is usually restricted to the system neutral point or grounding resistor, whereas the other three techniques can provide protection anywhere in the system. Usually a combination is needed for complete assurance of clearing all ground faults. Direct Relaying The simplest form of ac ground-fault protection is direct or neutral relaying. A current transformer is placed about the grounding conductor and located between the neutral point of the source transformer and the grounding resistor, as shown in figure 9.49. The grounding conductor acts as the primary winding of the CT, while the secondary winding is connected to the ground-overcurrent relay (5 IN). If the current through the grounding conductor Line A Line B Line C H :r >^ A Wye connected B Open-delta connected Figure 9.48.— Two CT approaches. ^Pb Neutral grounding resistor Grounding conductor Figure 9.49.— Neutral-resistor current-relaying scheme. exceeds a predetermined value, the relay acts to trip the circuit breaker. In many situations, some ground-current flow is nor- mal, due to system unbalance, capacitive-charging cur- rents, or inductive-coupling effects, and so the circuitry must be adjusted to pick up only when the normal level is exceeded. As will be seen, the pickup point should always be less than the system current level. The major disadvantage with this direct relaying method is that, should the grounding resistor or the grounding conductors become open, it will never detect any ground-current flow. The system will continue to operate with no abnormal indication, and then the system can become essentially ungrounded, posing a personnel hazard especially where resistance grounding is manda- tory (12, 14, 18). Accordingly, although the technique does find application on some portions of the ground system, some States do not allow its use on substation grounding resistors, even for a second line of defense. Potential Relaying Potential relaying, as shown in figure 9.50, is often used as a sole means of ground-fault protection at the surface substation and can also be used as a backup to other protection schemes at a unit substation or power center. With this method, the primary winding of a PT is connected across the neutral grounding resistor, while the 250 secondary winding is connected to a voltage-sensing ground-trip relay (59G). If current flows through the grounding conductor, a voltage is developed across the grounding resistor. When the voltage rises above a preset level, the ground-trip relay causes the circuit breaker to trip. Unlike direct relaying, potential relaying has the advantage of being able to detect a ground fault with the neutral grounding resistor in an open mode of failure. However, if the grounding resistor fails in a shorted mode, potential relaying is rendered inoperable. Zero-Sequence Relay Zero-sequence relaying, also termed balance-flux re- laying, is the most reliable first defense against ground faults in mine power systems. As shown in figure 9.51, the circuitry consists of a single window-type CT; the three line conductors are passed through the transformer core, forming the CT primary (6). On four-wire systems that have line-to-neutral loading, the neutral conductor must pass through the window but the grounding conductor must not. However, such loading is not allowed in mining, so only the line conductors can be used. In a symmetrical phase set, the vector sum of the three currents in the primary circuit will be zero, and no current will flow in the secondary. During a line-to-neutral fault, the phase unbalance will induce a current flow in the CT secondary, proportional to zero-sequence current. If the secondary current exceeds the relay pickup setting, circuit breaker tripping will be initiated. This phenomenon can be easily demonstrated. In terms of symmetrical components, the three phase cur- rents can be written as la = ^al + *a2 + ^aO> lb = Ibl + lb2 + IbO> Ic = Icl + Ic2 + IcO- The primary current, L rim , for the CT can be considered as the vector sum of the three line currents, which is also the current flowing through the grounding conductor (or that external to the CT window). Therefore, I P x. = la + lb + Ic = (a 2 + a + l)I al + (a 2 + a + 1)1^ + 3I a0 . Because a 2 + a + 1 = 0, Iprim = 3I a0 - Thus, for an unfaulted or balanced condition, Iprim = 3I a0 = 3I b0 = 3I c0 = 0, and no current is induced in the CT secondary. However, for a line-to-neutral fault involving phase a, the primary current equals the ground-fault current, If, or Iprim = If = 3I a0 , and a current is induced in the CT secondary to initiate tripping. Zero-sequence relaying is not affected by CT error, and therefore gives very sensitive tripping. The scheme is widely applied in mining at all voltages. Residual Relaying Residual relaying (fig. 9.52) is used in conjunction with CT's placed about the phase conductors. This relay- ing technique is used primarily on high-voltage distribu- tion circuits that require CT's and inverse-time relays for overcurrent protection. The CT's and the phase- overcurrent relays are both connected in a wye configura- tion. The ground-fault or residual relay is connected be- tween the neutral points of the CT's and the relays. PT and voltage relay (59G) Grounding conductor Figure 9.50.— Neutral-resistor potential-relaying scheme. Ckt bkr ;^Rb ;^^ :^Pf 'ih Phase relays Zero-sequence relay (51G) Figure 9.51.— Zero-sequence ground relay connections. Ckt bkr P^Pt ;=^R ■ih Phase relays J L@J Residual relay (51G) Figure 9.52.— Ground relay in residual connection. 251 As the current flowing through the residual relay is in proportion to the sum of the line current, the principle of operation of the residual method is similar to that of zero-sequence relaying. However, because of errors due to CT saturation and unmatched characteristics, residually connected relays are often subjected to nuisance tripping. Hence, they cannot have sensitive or low pickup settings. This arrangement will not always provide consistent re- petitive tripping at the required tripping levels for mine power systems. Grounding conductor Broken-Delta Relaying Broken-delta ground-fault protection is somewhat similar to the residual method, except that the three CT's are wired in series, as shown in figure 9.53. The resulting output voltages from the transformers form a closed delta if the load is balanced (fig. 9.53A). An unbalanced condi- tion, such as a line-to-neutral fault, will cause the forma- tion of an open delta (fig. 9.535), and the resulting voltage causes current through the relay operating coil. The broken delta is sensitive to any unbalance, but the zero- sequence relay operates only on faults causing ground- current flow (6). Ground-Check Monitoring The effectiveness of all the ground-relaying methods depends upon the integrity of the grounding system. A ground-check monitor is the device used to continuously monitor the grounding connections to verify continuity (2-3, 9). If conductivity is inadequate, it is the function of the monitor to trip the circuit breaker that feeds power to the system experiencing defective grounding. As shown in chapter 7, the grounding conductor is not essential for mining machine operation, but it is impera- tive for personnel safety. The ground-check monitor en- hances safety by making sure, via the ground connections, that the equipment frames are at near-neutral potential. Again, the maximum allowable frame potential to earth is 40 V on low and medium voltage and 100 V for high- voltage systems. Ground-check monitoring is an extensive subject, which can only be outlined here; the references listed at the end of the chapter can be consulted for more detail, particularly references 2-3, 9, and 11 . Although there are potentially numerous ways of monitoring ground continuity, only a few are considered practical to construct or have the required high reliability (9). These techniques can be divided into two general classifications: pilot monitors and pilotless monitors. Mon- itors in the mining industry use these techniques but are also referred to by different names, which will be discussed later. Pilot Pilot monitors use a pilot or ground-check conductor (see chapter 8) to perform the task and are of three general kinds: series loop, transmitter loop, and bridge (9). In the most common series loop circuit, a power supply, the relay operating coil (instantaneous contacts or a minimal time delay), the pilot conductor, and grounding conductors are connected as shown in figure 9.54 (9). If the pilot or grounding conductor breaks the loop or if the power supply fails, the relay contacts will reset. The circuit can be either ac, using 60-Hz line frequency, or dc. A Normal operations IbA U Ia "Ib w I 'b+g firam v b v=o B Ground fault Figure 9.53.— Broken-delta protection. To circuit breaker trip Pilot •* Machine frame Voltage source Ground \ J Relay 1 To power-center frame ground Figure 9.54.— Series loop ground-check monitor. 252 The transmitter loop concept is basically the same as the series loop, except the voltage source is installed in the machine (fig. 9.55) (9). Here, the source must receive its power from the machine and the relay cannot pick up until the circuit is energized; therefore, the monitor must be temporarily bypassed in order to close the circuit breaker. Bridge-type monitors use the series combination of the pilot and grounding conductors as one leg of a Wheat- stone bridge. Figure 9.56 shows a general circuit, where Z 3 is used to balance the bridge for a specific pilot and grounding-conductor impedance (9). Bridge output is sometimes amplified, but with or without amplification, the relay resets if a preset impedance level is exceeded. Bridge input can be 60-Hz ac, dc, or an audio frequency such as 5,000, 2,500, or 900 Hz (2, 9). Pilotless Pilotless monitors, as the name implies, do not use the pilot conductor. Instead, as shown in figure 9.57, an audio signal is placed on the phase conductors through a filter and removed at the machine through filters, completing its path back to the source in the grounding conductor (9). Instead of the filters, some models use coils similar to CT's to send and receive the audio signal. Between the ground- ing conductor and the power-center frame is a saturable reactor, which shows high impedance to the monitoring frequency. Its purpose is to restrict the monitoring signal to the intended path. This presents problems in coupler grounding, as the coupler metallic shell is commonly grounded to the grounding conductor as well as physically connected through its receptacle to the power-center frame. The grounding conductor must be isolated from the shell ground so that the reactor will not be bypassed. Problems and Requirements All of the basic techniques are plagued by some disadvantages, and the attempt to achieve a reliable monitor has been a perplexing experience for the mining industry. The reason is tied to the basic character of the mine power system: figure 9.58 provides a conceptual view of some difficulties that can arise (2). One of the more pronounced problems is the parallel ground paths estab- lished by contact through the mine floor or through grounding conductors on other machines. The alternate ground paths may have a resistance as low as the ground- ing conductor, but in the majority of cases these are very temporary in nature and thus cannot be relied upon. Stray ac and dc and induced ac are an ever-present problem in many mines. With dc rail haulage, for example, substan- tial direct current can stray from the rail when a poor bond is present and end up flowing in the ac ground system. If the cable has a G-GC or SHD-GC configuration (chapter 8) or the system current is unsymmetrical, ac can be induced in the grounding conductors. At times, this cur- rent can be of significant magnitude, not only on under- ground loads but especially on surface excavating machin- ery. Another problem results from trailing-cable deterioration; in a splice, for instance, there is a chance the ground-check conductor could short to the grounding conductors. Power-system transients, occurring from light- ning or switching surges and wiper contacts on reeled units, present additional problems, and all of these situa- tions can affect ground-check monitors. MSHA has established several guidelines for low- voltage and medium-voltage monitors that must be met before a monitoring device is approved. In these regula- tions, two monitors are recognized: a continuity type and an impedance type. A continuity monitor is one that meets the general requirements of 30 CFR 75.902 (17). It moni- tors only the grounding-conductor continuity and does not measure impedance; pilotless techniques fall into this To circuit breaker trip Machine frame r\ Pilot J7^ Relay Ground 1 Transmitter I To power-center frame ground Figure 9.55.— Transmitter loop ground-check monitor. To circuit breaker trip t-n-» Relay Preset level Voltage source Pilot L/^i v\ Machine frame > V^3 Ground Figure 9.56.— Bridge-type ground-check monitor. Line conductors Ground Saturable reactor To power-center frame ground Figure 9.57.— Pilotless ground-check monitor. 253 Machine Ground Grounding conductor Pilot conductor Cable -v coupler i Possible X fault Grounding conductor Safety § Earth contact % Earth contact ground bed Wiper contact if ree Earth Earth Figure 9.58.— Some difficulties associated with ground-check monitoring in mining. class. An impedance monitor requires a pilot conductor and monitors any change in the impedance of the loop formed by the pilot and grounding conductors. The bridge technique is therefore an impedance monitor. The relevant requirements for both monitor types are as follows (2, 9, U): 1. The monitor must be "fail-safe"; in other words, the failure of any component, other than the relay contacts, must make the trip-circuit contacts reset. (The relay must pick up its contacts when in normal operation.) 2. The monitor must not trip when (a) input voltage is varied by + 15% or - 20%, or (b) 5.0 V minimum to 25 V maximum, 60 Hz, or 10 A dc is introduced in the grounding circuit. These conditions are intended to ensure that the unit stays operational, even when under the influence of power- line fluctuations, stray currents, or induced currents. 3. The open-circuit monitor voltage cannot exceed 40 V rms. 4. When detecting grounding-conductor continuity, (a) continuity monitors must trip the circuit breaker if the grounding conductor is broken at any point regardless of low-impedance parallel paths (75 Q is considered an open connection), and (b) impedance monitors must trip the circuit breaker if the impedance of the grounding circuit, external to the grounding resistor, increases to cause a 40-V drop under fault conditions (or, by Ohm's law, 1.6 Q for a 25-A limit). 5. Filters must not cause a personnel hazard during normal operation or when the grounding conductor is opened. 6. The maximum time delay for contact reset, after an inadequate ground is detected, cannot exceed 250 ms. 7. When two or more monitors are operated in paral- lel, no interference can occur to cause incorrect tripping. At this writing, similar guidelines for high-voltage ground-check monitors have yet to be established. How- ever, by 30 CFR 75 and 77 (17), the maximum open-circuit monitor voltage is established at 96 V rms. Furthermore, continuity monitors must adhere to item 4a above, with impedance monitors conceivably tripping if the grounding circuit impedance causes a 100-V drop. Advantages and Disadvantages As mentioned, each basic ground-check technique has inherent advantages and disadvantages. A listing of these is rather informative (2-3, 9). Other than simplicity, the advantages of series loop circuits are minimal. Designs using a dc source are im- mune to stray ac but not dc. Further, ac monitoring can be subject to nuisance tripping by stray dc current that offsets the signal current. However, when the relay operating coil is isolated by a blocking capacitor, an immunity to stray dc is gained. In any case, the relay coil must have a very low impedance to be sensitive to grounding-conductor imped- ance. Two disadvantages of the circuit are substantial: Parallel paths and grounded pilot conductors easily negate its operation. The advantages and disadvantages of the transmitter loop technique are basically the same as for series loop, except the circuit can detect pilot-to-ground shorts. Simple bridge monitors have the same problems as series loop models, but they are very sensitive to changes in grounding-conductor and pilot-conductor impedance. The more elaborate designs can be made immune to ac and dc stray currents, yet even these sometimes cannot distin- guish between a sound grounding conductor and an illegal parallel path. In general, pilotless monitors are superior to pilot designs, except they are obviously more expensive. Be- cause the ground-check conductor is not needed, all asso- ciated problems are removed. The most elaborate models can distinguish parallel paths and are immune to stray currents; however, the simple designs are vulnerable to both. Some pilotless designs also have the advantage of being adaptable to pilotless or pilot use, depending on the need. Pilot-conductor ground-check monitors can serve the very important function of safety interlocking. This fea- ture is required on many portions of the mine power system and is used on almost all high-voltage systems. An example of interlocking is shown in figure 9.59. The loop circuit sensed by the monitor not only includes the pilot and grounding conductors, but can also involve a series of contacts and switches. Whenever one of these is opened, 254 Upstream I power I equipment Top and side cover interlocks Figure 9.59.— Pilot interlocking circuit using ground-check monitor. the continuity of the ground-check circuit is compromised, the monitor resets its rely, and the circuit breaker trips the power to that section. Safety devices incorporated in the interlocks include top and side cover switches, which open if a cover is removed, emergency-stop switches, the pre- break contacts on couplers, and pilot-break monitors on disconnect switches. When such sophisticated interlocking is required, pilot-type monitors must almost invariably be used. For instance, pilotless models cannot be employed when in- line cable couplers are on the circuit. When there are gear-mounted couplers on the power equipment in the monitored circuit, pilotless monitors may be used if they are wired into the ground-check and grounding contacts of the couplers, so that they will trip the circuit breaker prior to plug removal. This applies to almost all low-voltage and medium-voltage monitored circuits in mine distribution. ARRANGEMENTS FOR MINING There are two groups of protective-relaying equipment within a typical power system: primary and backup. Primary relaying has the goal of clearing all faults and overloads, and aims to isolate offending power-system segments with minimum interruption to the system bal- ance. Backup relaying operates only in the event of a primary relaying failure; its action is only for uncleared faults. In mining, be it overload, short-circuit, or ground- fault relaying, both groups are used extensively to the point of redundance. Zones of Protection Protection to the entire system is principally related to primary relaying and is accomplished by establishing zones of protection. Each zone has an associated circuit breaker and fusible disconnect, or fuses with the required sensing devices, and adjacent zones overlap (10). If a failure occurs within an individual zone, only the switch- ing apparatus within that zone should open. If there were no overlap, there could be an unprotected region in the system in some situations; a failure within this area would produce no safety tripping. Although such a situation seems unlikely in theory, in practice it does occur and can be caused either by oversight or ignorance on the part of an engineer. It results in broad outages to the system, rather than restraining the problem within a specific zone. The coordination of protective relaying between the zones is extremely important. The aim is to isolate faults down- stream from the power source without disturbing up- stream zones. Unfortunately, obtaining this coordination is perhaps the most outstanding problem of relaying. By introducing the general arrangements of protec- tive relaying in resistance-grounded mine power systems, this section actually serves as a transition between the basic relaying principles and chapter 10, where there is a more detailed and specific analysis. The objective here is to show how primary and backup relaying, zones of protection, and coordination are utilized in both surface and underground mines. At all levels of the mine power system, protection against short circuits, overloads, and ground faults holds priority. Because the majority of failures in mining involve line-to-neutral faults, ground-fault protection commands special interest, but this does not negate the need to establish adequate line-conductor relaying. Coordination The objective of coordination is to determine the optimum characteristics, ratings, and settings for the protective-relaying devices (7); consequently, fault analysis of the system must be involved. The two common coordi- nation schemes that are utilized are pickup setting and time. With the first, relay pickup settings for a specific actuating quantity are set at progressively higher values from the loads to the power source, such that a higher level of the actuating quantity is required to trip the circuit breaker in an upstream zone. For the time coordination of a specific actuating quantity, pickup settings throughout the system will generally be the same, but the operating times to achieve contact closure at or above pickup are set progressively longer toward the source. One technique or the other may be applied to provide coordination between zones in the mine power system; at times, a combination might be needed. The design of this protective-relaying system can be a substantial problem, as exemplified by the fact that many engineers consider protective relaying more an art than a science. Ground-Fault Protection Except for capacitive ground current (see chapter 11), ground-fault current magnitudes are limited by grounding resistors. As mentioned in chapter 7, the maximum cur- rent is 25 A at low and medium voltages, but the level is seldom limited below 15 A, whereas it is very rare for a 50-A limit to be exceeded on high-voltage grounding systems, with 25 A seen on many systems. Thus, ground- fault current is probably close to constant throughout the system. The use of delta-wye, delta-delta, or wye-delta transformers enhances this nearly constant current situ- ation. At each transformer step, a new or separately derived ground system is produced, each with its own ground resistor. Because the transformer configuration blocks zero-sequence components, a ground fault on cir- cuits connected to the secondary raises primary current, but the vectorial sum of the line currents is zero (1). It can thus be seen that coordination of ground-fault relaying across transformers is unnecessary, and both primary and backup protection must be established for each derived grounding system. Selective coordination at each voltage level by pickup setting alone is normally 255 impossible, and time settings must be relied upon when multistage protection is used (1). It must be remembered that resistance grounding is used to reduce fault energy, frame potentials, or system potentials. If the first ground fault is not cleared, an alternate ground fault in another machine will be in a ground condition and not be limited by the grounding resistor (1). The result is the possibility of dangerous frame potentials and powerful intermachine arcing. Conse- quently, ground-fault relays must always be arranged to trip circuit breakers; fuses cannot be used. Overloads and Short Circuits The occurrence of a line-to-line or three-phase fault anywhere in the system can cause substantial outages if it is not cleared downstream. Because they are not restricted like ground faults, the anomalous positive-sequence and negative-sequence currents can be passed across trans- formers to higher levels of the system. The resulting wide-range problems are especially evident on radial dis- tribution. System protection is usually coordinated by using instantaneous relays to adjust pickup, and the effort could involve fault currents from the machines to the substation. A maximum setting within an individual zone would be the minimum fault current at which thermal or mechanical damage to a protected device could occur. The problems resulting from inadequate overload protection are just as widespread. In most cases this protection must be applied to the specific zone that the fuse or circuit breaker is protecting. For example, the pickup setting of a time-delay relay could be determined by the ampacity rating of the smallest conductor within the zone of the switching apparatus, whereas time settings might be employed to coordinate between zones. Overload protection can be critical in large underground mines with numerous sections; when several machines are operating simultaneously, the maximum current demand could over- load an upstream conductor. KEY 27 Undervoltage relay 37 Ground-continuity relay 50 ac instantaneous overcurrent relay 51 ac time-delay overcurrent relay 51G,51N ac time-delay ground overcurrent relay 52 ac circuit breaker SA Surge arrester Frame, ground Grounding conductor Ground-check conductor 52 C U — (5 og) C } — fcoe) C D — (500) rr« £j- rr SJ6 Excavator 3f" T-©-! 6 Production shovel SWITCHHOUSE POWER CENTER Grounding^ resistor [ — 6oX 5 Sr — '.so'.si; )52 )52 CD— |g) 6 Drill FT 66 Pumps Figure 9.60.— Simple surface mine power system illustrating protective relaying. Surface Mines To illustrate the application of these general consid- erations, consider figure 9.60, a one-line diagram (in- cluding grounding) of a simple surface mine power system (1). The diagram is drawn to show the various combina- tions of protective circuitry that can be found in surface operations. Each protective device is installed to trip the closest circuit breaker (52). In terms of ground-fault protection for the low-voltage machines, zero-sequence relaying (50G, usually instanta- neous) establishes primary protection. Backup protection could be placed here by adding timed direct relaying (5 IN) about the grounding conductor between the grounding resistor and the transformer neutral. For high voltage, remembering that the transformer configuration blocks zero-sequence current, primary ground-fault protection is provided in the switchhouse, again by zero-sequence relay- ing (50G, instantaneous or minimum-time dial setting), establishing a zone of protection for each outgoing circuit (the excavator, loader, and ac power center). The time- delay zero-sequence relay (51G) in the substation can be considered to give both backup protection for the down- stream 50G relaying and primary ground-fault protection for the zone between its location and the switchhouse. In any event, backup protection is the principal duty of the 5 IN direct relay about the neutral conductor of the sub- station. (N is used to signify neutral relaying, but G can be used.) All this relaying is normally coordinated by time settings, the pickup setting being specified as a percent- age of the ground-current limit. The use of the grounding-conductor direct relaying, shown in figure 9.59, is restricted to situations where the main circuit breaker that the relay trips sees total ground current. In the ac power center, this means that the backup relaying must cause tripping of both circuit break- ers. Because here there is no main breaker, some States require potential relaying of the grounding resistor for backup. Potential relaying does give more safety than direct relaying, as previously mentioned. Ground-check monitoring is illustrated in its most extensive form for surface mining in figure 9.59, where every grounding conductor in the system is measured. Ground-check conductors are shown connected to each monitor. In instances where pilotless devices are applied, ground-check conductors are not needed. In cases where conductors from the substation form an overhead ring bus, such as in some open pit operations, ground-check moni- tors are often not used because of the circular nature of the distribution. In any overhead distribution arrangement, monitors can experience extensive failures because of lightning strokes. 256 Overload and short-circuit protection in the high- voltage portion is provided by line-conductor CT's usually connected in wye to time-delay relays that are normally induction-disk types with both instantaneous (device 50) and time-delay (51) elements. Each 50/51 combination estab- lishes a zone of protection downstream from its location. As shown in the power center, the same kinds of protection can be provided at utilization by low-voltage power circuit break- ers in conjunction with external CT's and relays. In figure 9.60, these protection devices are indicated as numbers within dashed circles. Normally, they would be protected by molded-case units. Note that all 50/51 phase protection could also be done by fuses or fusible disconnects, but circuit breakers or power-driven load-break switches still are needed for ground-fault protection. Figure 9.61 shows a three-line diagram of a typical molded-case arrangement. These components replace the 50, 51, 50G, and 27 devices associated with each low- voltage breaker in figure 9.60. The molded-case circuit breaker provides short-circuit protection through its mag- netic trip units (not shown), with overload tripping given by the internal thermal elements. (Overload protection may not be required on low- and medium-voltage circuits; see chapter 10 for discussion.) The external protective circuitry is commonly a zero-sequence relay and pilot-type ground-check monitor, and the contacts of both trip the undervoltage release. Notice that here the ground-fault relay shunts the trip coil, while the ground-trip relay is in series with the trip coil. The UVR itself provides under- voltage protection. In this way all the essential kinds of protection are available for a resistance-grounded trailing- cable installation. An undervoltage relay (device 27) is also shown in the switchhouse of figure 9.60, but this would not be required at this or any location provided that all equipment down- stream has undervoltage protection. Window (zero-sequence CT Receptacle -plug combination Figure 9.61.— Typical schematic for three-phase molded- case circuit breaker with ground-overcurrent and ground-check protection. Underground Mines For comparison, figure 9.62 illustrates a simple one- line diagram of an underground mine power system. The ac portion is very similar to the surface circuitry of figure 9.60, and the reasoning and arguments behind the protec- tive devices are the same. There are differences, however, in the disconnect switch, the relaying in the substation, and possible extra outgoing circuits from the substation, but again, all these features are also possible in surface mines. By law, some means of visible disconnect is required within 500 ft of the point at which power enters the underground workings, and a separate switch is shown in the diagram for this purpose. Actually, it is also advisable, if not required, to place visible disconnect switches on incoming high-voltage (distribution) circuits within all power equipment. (These have not been included in figures 9.60 and 9.62 merely to maintain clarity.) Consequently, it is not uncommon to find the first disconnect in the mine as part of a switchhouse. Ground-fault backup relaying located in this substa- tion is mainly by the potential technique. Here, the relay operating coil is placed directly across the resistor, and in addition to relaying, the transformer is used for resistor impedance matching for current limit. Although not shown, the outgoing circuit to the surface equipment has basically the same relaying as for underground (see chap- ter 13 for the implications of such loads). Direct Current Only one form of relaying for dc equipment is provided in figure 9.62; this is short circuit for rail haulage systems and consists of a dc overcurrent relay (device 76) driven by the voltage drop across an in-line shunt. For off-track dc machinery, the protective-relaying arrangement is di- rectly tied to the power source. Five basic systems of dc ground-fault protection are presently being used in the United States: • Diode grounding, • Basic grounding conductor, • Relayed grounding conductor, • Neutral shift, and • Differential current. The grounding philosophy for all but the last was intro- duced in chapter 7. The neutral-shift and differential- current systems can be used only when the machine is powered from the output of a rectifier, such as that contained in a section power center, whereas the first three systems are more commonly employed when a trolley system is the dc source. When dc equipment is powered from the trolley sys- tem, short-circuit and overload protection is normally provided by a dual-element fuse. The device is mounted in a holder or nip, which is clipped to the trolley wire. The other trailing-cable power conductor is connected to the rail. Fuses are rarely applied to trailing-cable protection in load-center rectifiers. Here, air-magnetic power break- ers, molded-case devices, or dc contactors (see chapter 12) are employed. In practice, these are usually tripped only for short-circuit conditions, with overload protection not used. 257 -ft-Chf--' h" ] " r «j. £ © .'51k .»V / SURFACE AREA 77777777777777777777777777777777777777777777777% UNDERGROUND KEY 27 Undervoltage relay 37 Ground-continuity monitor 50 ac instantaneous overcurrent relay 51 ac time-delay ground- overcurrent relay 51G.5IN ac time-delay ground- overcurrent relay 52 ac circuit breaker 59G ac ground overvoltage relay 72,72T dc circuit breaker 76 dc overcurrent relay Machine frame MAIN SUBSTATION vf v Line-to-cable termination ^7777777777777777777777777, SWITCHHOUSE MINER Figure 9.62.— One-line diagram of simple underground mine power system illustrating protective circuitry. Diode-Grounded The diode-grounded system is often found with dc vehicles that employ cable reels, because it permits a two-conductor cable to be used, which is less expensive and takes less space on the cable spool than type G cables. These features make the system attractive, but the diode- grounded system has deficiencies from a safety standpoint. A simplified diagram of the system is shown in figure 9.63. The machine frame is tied to the grounded negative conductor by means of the grounding diode (DJ. The grounded conductor is connected to the power-center frame or trolley system rail, depending on the power source. In series with the diode lead is a ground-fault device, G, a mercury-magnetic switch or the operating coil of an elec- tromagnetic attraction relay. The pickup setting of this device can be no greater than 25% of the forward-current rating of the diode. Figure 9.63.— Diode-grounded system with possible fault in- dicated. If a positive-conductor-to-machine-frame fault occurs (location 2 of figure 9.63), current flows through the grounding diode and the ground-fault relay. When the fault current exceeds the pickup setting of the relay, the fault is then isolated by deenergizing the machine contac- tor (M x ) with opening of the G contacts. Actually, this sequence of events occurs only if the fault exists between the load side of the contactor and the motor. This leads to some of the basic drawbacks of the diode-grounded system since location 2 is the only safe area for a ground fault to occur. Location 1 of figure 9.63 includes the length of un- grounded conductor from its entrance point on the ma- chine to the main contacts of the M 1 contactor. Since these machines normally utilize cable reels, slip-ring assemblies are required for connecting the incoming power conductors to the machine circuit. Hence, it is difficult to locate a circuit-interrupting device at the immediate cable en- trance of the machine. If a fault occurs at location 1, the fault current flows through the machine frame, causing diode D x to become forward biased, and current then passes through the ground-fault relay. When the relay operates, the M x contactor resets, and the motor circuit is deenergized. However, this does not isolate the fault; isolation is solely dependent upon the opening of the M 1 interrupting device, and the fault is located on the line side of its contacts. In this case, fault current can be terminated only by the switching apparatus on the source side of the trailing cable (dual-element fuses for a trolley nip, or fuses or a circuit breaker for a power center). Location 3 of figure 9.63 includes the entire length of the grounded conductor within the machine. A fault here renders the diode-grounded system unreliable, since it effectively provides a parallel path from the frame to the grounded conductor. The purpose of diode D 1 is to block the voltage drop across the grounded conductor from the machine frame. Therefore, if D x is bypassed, the frame potential is raised. A fault at location 3 can easily go undetected. As a result, a simultaneous fault at any location of the ungrounded circuit must again rely on isolation of the source side of the trailing cable. Failure of the grounding diode is a common problem associated with the diode-grounded system. With L\ in the shorted mode, the situation is the same as for a location 3 fault. However, when the grounding diode fails in an open mode, all ground-fault protection is lost. Faults occurring in location 1 or 2 can cause the frame to become raised to a value equal to the supply voltage. The D 2 diode is referred to as the polarizing diode. Its purpose is to protect against installing the positive and 258 negative conductors in switched positions, which can occur when adding a new trailing cable or while making a cable splice. If the polarity of the supply voltage is inadvertently reversed, the polarizing diode will block the current flow, and the motor circuit cannot be energized. If D 2 failed in the shorted mode when the dc circuit was under reversed conditions, the diode-grounded system would be rendered totally inoperative, but the machine would still operate. Grounding Conductor Other methods of grounding dc machines require the use of a separate grounding conductor and type G cable. The basic grounding-conductor system (fig. 9.64) and the relayed grounding-conductor system (fig. 9.65) are the simplest of these techniques and are commonly used when the source is a trolley system. With the basic system, a ground fault is detected only by the overload protection on the source side of the trailing cable, and the result is extremely poor sensitivity with respect to ground faults. In the relayed system, the ground-fault sensitivity is improved by adding a ground-fault relay in series with the grounding conductor, as shown in figure 9.65. However, the current flowing through the relay may not be the total ground-fault current, since a parallel path may be estab- lished through the earth. Neutral-Shift System The neutral-shift system is illustrated in figure 9.66. The resistors Rj and R 2 create a dc neutral point for the system as well as limiting ground-fault current. A ground fault on either the positive or negative conductor will cause the neutral point to shift, which results in detection by the voltage-sensitive relays M 1 and M 2 . These relays are usually given the device number 64, a potential relay, also termed a dc unbalance relay. The primary limitation of this system is that it cannot discriminate between ground faults for individual pieces of equipment fed by the same rectifier. If a ground fault occurs on one machine, all circuits are interrupted. Differential Current As shown in figures 9.67 and 9.68, differential-current dc ground-fault protection utilizes the neutral of the wye-connected secondary feeding the rectifier. With a delta-connected secondary, a neutral connection can be provided through a grounding transformer. A grounding resistor is inserted in the neutral circuit to limit dc ground-fault current, typically to 15 A. Ground faults are then detected using either a saturable-reactor or saturable-transformer system. Actually, this results in an alternative method for grounding dc off-track equipment to that presented in chapter 7. With the saturable-reactor system (fig,. 9.67), a voltage-sensitive ac relay is placed in series with the winding of a reactor that encircles the positive and nega- tive outgoing conductors. The relay and reactor are excited by a 60-Hz control voltage, and under normal conditions, the current through the positive conductor equals that through the negative conductor. The magnetic fields about both conductors tend to cancel each other, and thus the reactor winding exhibits maximum impedance and pro- hibits the ground-fault relay from operating. However, when a ground fault occurs, the currents in the positive and negative conductors become unequal in magnitude Machine frame IX - v Grounding conductor Figure 9.64.— Basic grounding-conductor system. s Machine frame (+) \ // ' Ground-fault relay Figure 9.65.— Relayed grounding-conductor system. To rectifier M< i -+• To circuit breaker -*• trip circuit (+) Mo) i GND To circuit breaker trip circuit Figure 9.66.— Neutral-shift system. because current is flowing in the grounding conductor. The magnetic fields about each power conductor are no longer equal, and the resultant magnetic field causes the reactor core to saturate. This in turn reduces the impedance of its winding, causing the ac control voltage to be impressed across the relay. Typical relay pickup is from 4 to 6 A of ground-fault current. The saturable-transformer system (fig. 9.68) operates on the same principle. Here, the ac control voltage excites the primary (bias) of a two-winding toroidal transformer. The transformer secondary is connected to a rectifier bridge whose output feeds an adjustable resistor in series with a dc relay. The relay has NO contacts when deener- gized. Under normal operation, ac is induced in the transformer secondary, and the rectified ac causes the relay to pick up its contacts. During a ground fault the core again saturates, but in this case the saturation stops transformer action. In turn, the dc is removed from the relay, which resets its contacts to cause circuit breaker 259 To circuit breaker trip circuit Grounding resistor To ac ■►control circuit Figure 9.67.— Current-balance dc ground-fault relaying using saturable reactor. To circuit breaker trip circuit *■ Positive »■ Negative Ground To machine Figure 9.68.— Current-balance dc ground-fault relaying using saturable transformer. tripping. The resistor allows adjustment of the pickup setting. Another feature of differential current relaying in addition to sensitive detection of dc ground faults is that the technique senses dc unbalance only in the conductors that pass through the core. By using a detection device for each outgoing dc circuit, the individual relay system is sensitive only to ground faults existing downstream from its reactor or transformer. Thus, selective dc ground-fault relaying can be realized, an advantage not available with the neutral-shift system. The material in this chapter has served to introduce the complexities of protective equipment and relaying, culminating with figures 9.60 and 9.62, where it was shown that the protective relaying in a mine power system can be divided into zones of protection for the various power equipment. The application of this circuitry is the subject of chapters 12 and 13, where additional variations of these basic techniques are discussed. But first, a careful study of the overall mine power system is needed to emphasize the correct coordination of the various protec- tive devices. This is the topic of chapter 10. REFERENCES 1. Chumakov, W. V. Protective Relaying for Mining Applica- tions. Pres. at 3d Annu. West Protective Relay Conf., Spokane, WA, Oct. 18-21, 1976; available from BBC, Allentown, PA. 2. Cooley, W. L., and R. L. McConnell. Current State-of-the-Art in Ground-Check Monitoring. Paper in Conference Record -IAS 11th Annual Meeting (Chicago, IL, Oct. 1976). IEEE, 1976. 3. Cooley, W. L., and R. L. Rinehart. Grounding and Ground Check Monitors. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 4. Eaton, J. R. Electrical Power Transmission Systems. Prentice-Hall, 1972. 5. Fink, D. G., and J. M. Carroll (eds.). Standard Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 6. Institute of Electrical and Electronics Engineers (New York). Recommended Practice for Electrical Power Distribution for In- dustrial Plants. Stand. 141-1986. 7. Recommended Practice for Protection and Coordina- tion of Industrial and Commercial Power Systems. Stand. 242-1986. 8. Kaufman, R. H. The Magic of Pt. IEEE Trans. Ind. and Gen. Appl., v. 2, Sept./Oct. 1966. 9. King, R. L. Development of an Electrical Engineering Course for Mining Engineers. M.S. Thesis, Univ. Pittsburgh, Pittsburgh, PA, 1977. 10. Mason, C. R. The Art and Science of Protective Relaying. Wiley, 1956. 11. Mason, R. H. Enforcement Begins for Ground Check Monitoring Underground. Coal. Min. & Process., v. 14, June 1977. 12. Myers, W. P. Current-Limited Ground Fault Relaying. Min. Congr. J., v. 59, Apr. 1970. 13. National Fire Protection Association (Quincy, MA). National Electrical Code. NFPA 70-1981 (ANSI Cl-1981). (Updated every 3 yr-) 14. Shimp, A. B., and D. A. Paice. Application of Molded-Case Breakers on DC Electrical Systems in Coal Mines. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 15. Underwriters' Laboratories, Inc. Fuse Standards. UL 198.1, 1973 et seq. 16. U.S. Bureau of Mines. Schedule 2G, Electric Motor-Driven Mine Equipment and Accessories. Federal Register, v. 33, No. 54, Mar. 19, 1968. 17. U.S. Code of Federal Regulations. Title 30-Mineral Resources; Chapter I -Mine Safety and Health Administration, Department of Labor; Subchapter O-Coal Mine Health Safety; Part 75 - Mandatory Safety Standards, Underground Coal Mines; Part 77 -Mandatory Safety Standards, Surface Coal Mines and Surface Work Areas of Underground Coal Mines; 1981. 18. Wade, E. C. Ground Relaying for Mining Distribution Systems. Coal Age, v. 71, July 1966. 19. Westinghouse Electric Corp., Low-Voltage Breaker Div. (Beaver, PA). Breaker Basics. 1973. 20. Westinghouse Electric Corp., Relay-Instrument Div. (Newark, NJ). Applied Protective Relaying. Silent Sentinels Publ., 1976. 21. Wood, R. J., and H. D. Smith. Low-Profile Semiconductor Equipment for Mine Application. IEEE Trans. Ind. and Gen. Appl., v. 4, May/June 1968. 260 CHAPTER 10.— SIZING PROTECTIVE DEVICES 1 In chapter 9, it was shown that circuit breakers, fuses, and switches are rated in terms of the nominal circuit voltage, the continuous currents they may carry, the sort-circuit currents they may interrupt, and the fault- through currents they must withstand, lb ensure that these interrupting devices disconnect faulted equipment promptly and correctly, it is necessary to have a separate protective system that recognizes the presence of a fault, determines what is faulted, and supplies energy to the mechanisms that will terminate current flow. It then becomes necessary to calculate the maximum fault cur- rents and, in many cases, the minimum sustained overcur- rent values in the mine system in order to determine the sensitivity requirements for the current-responsive protec- tion devices (3). 2 For multistage time-delay devices, the operating time of each device and the time relationships between devices must be found. These parameters are mandatory for the successful selection, installation, and coordination of protective equipment and relaying. FAULT CURRENT The fault current at any point in the mine complex is limited by the impedance of circuits and equipment from the source or sources to the fault point. The level is not directly related to the load on the system (5). When a mine is in development, system additions are often made to increase the capacity to handle the growing load. While these changes will usually not change the normal load on preexisting system portions, they may substantially in- crease short-circuit current during three-phase and line- to-line faults. Nevertheless, ground-fault currents, except for some special cases, remain relatively constant in high-resistance grounded applications. Whatever the situ- ation, the available fault currents must be predetermined to ensure adequate protective-device operation. Fault-Current Sources To find the available fault current correctly, all sources of fault current should be known. The main sources in mining are electrical utility systems, synchronous motors, induction machines, and capacitors (and capacitance). Syn- chronous generators are also a significant source, but these are only a concern if the mine generates its own power. Of the others, some can be eliminated because of their negligi- ble contribution to fault current, but with caution, depend- ing upon the installation. In most mine systems, the generators of the electric utility system are the principal source of short-circuit current. Because these generators are often remote from the mine, the current that results from a fault within the mining operation usually appears as merely a small increase in the generator load current, and the contribu- tion remains fairly constant (3-4). In this case, the electric 1 The author wishes to thank Frederick C. Trutt, professor and chairman of electrical engineering, the University of Kentucky, for his assistance in preparing the fault-calculation and computer-analysis sections of this chapter. 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. utility can be represented at the mine as a Thevenin's equivalent source— in other words, a single-value equiva- lent impedance driven by a constant voltage and referred to the point of connection. Even in proximate locations such as mine-mouth electric plants, this approximation usually gives adequate calculation results. Synchronous motors can be a substantial contributor to fault current, but because their prime coal mining application is as drive motors for excavators, they are mainly of concern to surface mines. The current supplied to a fault can be described as follows (3—4). After the occurrence of a fault, system voltage tends to decrease: the motor receives less power to rotate its load, and simulta- neously the inertia of the motor and load, as the "prime mover," causes the motor to act as a generator. The motor thus supplies fault current, which diminishes as the motor field excitation and kinetic energy decay. Induction motors can also contribute fault current to the system when inertia drives the machine as a generator (3-4). Here, however, the presence of field flux in the rotor is produced by induction from the stator; this decays rapidly with motor terminal voltage and disappears com- pletely after a few cycles. Accordingly, the effect of induc- tion machines, except for large horsepower ratings, can sometimes be neglected but should always be considered. Because of the widespread use of shielded cables, power-factor correction capacitors, and surge capacitors, capacitance is found in great abundance in mine power systems. Although the most destructive power-frequency fault currents come from rotating machinery (including the utility), capacitance can produce very high transitory overvoltages and fault currents (4). Fortunately, these are usually of short duration and of a natural frequency much higher than the power frequency, but they can lead to insulation damage. This subject is discussed in chapter 11. The stored charge of the capacitance acts as the prime mover, but because the fault-current contribution is of such short duration (less than 1 cycle), it can often be neglected when calculating line current (4). The main problem with fault conditions lies in the area of ground overcurrent during line-to-neutral faults. For instance, most high-voltage resistance-grounded distribution sys- tems have a 50-A ground-current limit or less, and some mines use a bank of surge capacitors across the terminals of all high-voltage loads. In 12.47-kV systems, the stan- dard 0.25-^F surge capacitor adds about 5-A capacitive ground-charging current, directly translating to ground- fault current during a line-to-neutral mishap. Adding to this current is the charging current of high-voltage feeder cables, typically 0.2 A per 1,000 ft. With just 10 capacitor banks on the distribution system in a moderately sized mine, a 50-A limit is exceeded. If a line-to-neutral fault occurs, the capacitance could discharge and feed the fault with capacitive current in excess of the ground-current limit. As this could happen within the distribution cir- cuitry, there is a possibility that the ground resistor would not see it, and because the duration could be less than 1 cycle, the protective circuitry might not react. Source Equivalent Circuit The representation of equivalent circuits is the prin- cipal difficulty in the calculation of fault currents. An 261 equivalent circuit for the most important typical source, the utility system connection, has already been presented and is usually straightforward. Rotating machinery is less straightforward because the fault current contributed by each machine is limited by the machine impedance, which is unfortunately a rather complex and time-dependent variable. To simplify calculations of fault currents, Thevenin's equivalent sources are again assumed. Three specific values of reactance are used to estab- lish the fault current delivered at three points in time (3): 1. The subtransient reactance, X^', for current during the first cycle after the fault occurrence; 2. The transient reactance, X^, for current after sev- eral cycles at 60 Hz; and 3. The synchronous reactance, X d , which determines current flow for the steady-state region. Substransient reactance is assumed to last about 0.1 s, after which the machine impedance increases to the transient reactance value. After 0.2 to 0.5 s, the value again increases to the synchronous reactance. With each increase, current contribution decreases. Depending on the maintenance of field excitation, synchronous motors could use all three values, but X d and sometimes X d are not needed in mining applications since the values approach infinity. Because of rapid current decay, X d is the only value used for induction motors, as these motors do not contribute significant fault current beyond the first cycle after the fault. Table 10.1 lists some typical motor reactances, given in per-unit based on the machine kilovoltampere rating (3). Following the previous statement and considering the type of rotating machinery, calculations on underground mining systems sometimes ignore the motor contribution, but computations for sur- face mines and surface facilities, especially where thermal dryers are involved, generally cannot. Table 10.1.— Sample reactances for synchronous and induction motors Subtransient Transient Motor type reactance reactance Comments (X"d) (X'J SYNCHRONOUS Motors: 6 poles 0.15 0.23 ) 8 to 14 poles .... .20 .30 > None 16 poles or more .28 .40 J Converters: 60O-V dc output .20 01 0/ None. 250-V dc output .33 INDUCTION Individual motors .17 p) Motors generally above 600 V. Groups of motors .25 ( 1 ) Motor voltage usually 600 each less than 50 hp. V and below. Subtransient reactance is increased slightly because of very rapid current decay. Lower value of X" d would be in order for groups of larger motors. transient reactance not usually needed in calculations. NOTE. — The above values are in per-unit referenced to the motor kilovolt- ampere base. Approximate power bases can be determined from the motor horsepower rating: 0.8 pf motor, kVA base = hp rating; 1 .0 pf motor, kVA base = 0.8 (hp rating). FAULT CALCULATIONS FOR THREE-PHASE SYSTEMS In specifying equipment or checking protective- circuitry performance, certain simplifying assumptions are normally made to calculate fault currents. An impor- tant assumption when finding line currents on three- phase systems is that the fault is three-phase. This type of fault generally causes the maximum short-circuit current to flow. Perhaps the only exceptions are on utility systems and similar industrial applications where a line-to-neutral fault can cause current up to 125% of a three-phase fault (3), but these fault currents are substantially limited in most mining systems. The significance of assuming a three-phase fault is that symmetrical-component methods are not needed for the solution of routine fault calcula- tions, although detailed investigations may sometimes require looking at asymmetrical problems. Because mine power systems are normally radial or operated radially, calculating a three-phase fault is a rather simple task. Basically, all that is required is Ohm's law and an equivalent circuit. Even during a line-to-line fault, positive-sequence and negative-sequence imped- ances are essentially equal. A good estimation can thus be made by applying a fixed fraction of the three-phase case, 3 which has been found to be about 0.87 (3). A second assumption is that the fault is customarily assumed to be bolted; that is, it has zero impedance (3-4). This not only simplifies calculations but also applies a safety factor, since the results provide a value greater than maximum and equipment is rarely stressed beyond its full rating. For instance, analytical studies on low-voltage systems have shown that minimum arcing fault currents, expressed as a factor of bolted faults, are typically 0.89 at 480 V for three-phase arcing faults, and 0.74 at 480 V for line-to-line arcing faults (3). As the voltage level is in- creased, the arcing current level approaches but is always less than that of the bolted case. IEEE standards 141-1976 and 242-1975 detail exten- sive fault-calculation recommendations, many of which are directly adaptable to mine power systems (3-4). A summary of these follows, after which an example will be discussed. In both, specifics about sizing high-voltage switchgear will be presented to illustrate the method, but the same line of reasoning can be applied to low voltage and medium voltage. Short-Circuit Calculation Procedures The calculation procedure for finding currents result- ing from faults between power conductors, often termed short-circuit currents, can be divided into a series of steps. By assuming a bolted three-phase fault condition, the power-system parameters remain symmetrical regardless of the neutral conditions or even delta-wye transformer connections. The balance is especially close in high- resistance grounded systems, as line-to-neutral loads are not allowed. Therefore, the balanced fault currents can be calculated from a single-phase equivalent circuit, which has only per-phase impedance and line-to-neutral voltage. The procedure is to find the Thevenin's equivalent imped- ance of the entire system (Z), then solve for fault current 3 Personal communication from E. K. Stanek, West Virginia University, Aug. 1977. 262 using Ohm's law. Calculations may use real, percent, or per-unit values. As shown in chapter 4, per-unit methods have the advantage of greatly simplifying the work when the system has different voltage levels. The first calculation step is always the preparation of a good one-line diagram that shows all fault current sources and all significant elements. All major imped- ances, both resistance and reactance, should be included: the utility, transformers, conductors, cables, and rotating machines. The collection of this information is perhaps the principal difficulty in fault calculations. Impedances for cables and conductors can be assembled from chapter 8. If transformer impedances are not known, the values can be estimated from the standard values given in tables 10.2 and 10.3 (other impedances are given in chapters 12 and 13) (3). Table 10.2.— Three-phase transformer per-unit impedances 1 for liquid-immersed transformers, 501 to 30,000 kVA High-voltage Low-side rating, Low-side rating, rating, V 480 V 2,400 V and up 2,400 to 22,900 0.0575 0.055 26,400, 34,500 .0625 .060 43,800 .0675 .065 69,000 NAp .070 115,000 NAp .075 138,000 NAp .080 NAp Not applicable. 1 Actually a reactance, the value is normally expressed as a percentage. The per-unit values given use the transformer self-cooled kilovoltampere and voltage ratings as a base. Transformer resistance is usually well below 0.01 pu. Table 10.3.— Three-phase transformer impedances for distribution transformers, including load centers High-voltage kVA Per-unit rating, V rating impedance 1 2,400to 13,800 112.5- 225 >0.02 300 - 500 >.045 750 -2,500 .0575 22,900 All .0575 34,500 AH .0625 Actually a reactance, the value is normally expressed as a percentage. The per-unit values given use the transformer self-cooled kilovoltampere and voltage ratings as a base. Transformer resistance is usually well below 0.01 pu. In terms of system impedance, fault current has been shown to be mainly dependent upon the reactance be- tween the sources and the fault. This holds true except where there is substantial resistance such as with the extensive use of cables, overhead lines, and buses, which is the case in mining. However, if the reactance-to-resistance ratio, X/R, for the entire system from the source to the fault is greater than 5.0, negligible error is introduced by ignoring resistance. In fact, omitting resistance actually provides a small safety factor and has become common practice in other industrial applications. For most cases, it is recommended that the per-unit system be used in calculation. Therefore after the one-line diagram is complete, all parameters must be converted to per-unit values on a set of consistent bases. Normally, the base power, kVA b , is selected first. The current base, I b , and impedance base, Z b , are then derived, using the nominal voltage at each system level as the other base, V b . Accordingly, the base voltages must be related to the turns ratios of the interconnecting transformers. The power base may be any convenient level, but the main-substation kilovoltamperes is often selected. The next step in the calculation procedure is to reduce the system to its Thevenin equivalent by combining all impedances. The result of the reduction is a single driving voltage in series with a single impedance and the fault. Ohm's law can then be used to compute the per-unit fault current, or I -^ pu Z„ (10.1) where V pu = driving voltage of circuit, pu V, Z pu = equivalent impedance from sources to fault, including source impedances, pu fi, and I pu = symmetrical rms fault current, pu A. The prefault voltage, the level existing just prior to the fault occurrence, is assumed to be the nominal system voltage at the fault location. With this assumption, short- circuit current will approach maximum. Furthermore, when using the per-unit system, if the voltage bases are equal to the system nominal voltages, the driving voltage, V pu , is simply equal to 1.0 pu. The per-unit current can be converted to amperes using the base current multiplier. This calculated fault current is an alternating symmetrical quantity, because the sources are rms voltages, and can be used to compare equipment ratings that are expressed in symmetrical rms currents. However, the fault calculations must also recognize the asymmetry of typical fault currents and account for it. Fault current waveforms are discussed in chapters 4 and 9, but the typical asymmetrical type is again reproduced in figure 10.1 for convenience (3). The compensation for asymmetry consid- ers current composed of two components: 1. The ac symmetrical component, taken as the calcu- lated symmetrical value, and 2. The dc component, with its initial maximum magni- tude taken as the peak of the ac symmetrical component. Total asymmetrical current dc component Symmetrical ac TIME Figure 10.1. metry. ■Fault current waveform illustrating asym- 263 The time period in which the dc component decays is related to the reactance-to-resistance ratio, (X/R), of the system. Hence, the first cycle maximum fault current is estimated as 1.6 times the symmetrical rms value. De- pending on the X/R ratio, various other multiplying fac- tors are used to approximate maximum asymmetrical levels throughout the dc decay. These result in estimates of asymmetrical rms current that can be used to compare with ratings based on total (asymmetrical) rms current. Because fault current varies with time, several differ- ent results are commonly desired from the fault calcula- tions. To obtain these might require carrying out simulta- neous impedance reductions that account for X^', X^, and X d of the rotating machinery. However, when the system has just induction motors as loads, these usually contrib- ute fault current only within the first cycle, and at most only two reductions are necessary: one with the motors as sources, the other with just the utility system. In marginal situations, the motor contribution could be the extra current that results in equipment destruction during line-conductor faulting. Various applications of fault computations are shown in table 10.4. When aimed toward line-current applica- tions in mining, the first-cycle maximum symmetrical current is always needed. The maximum value that occurs between 1.5 and 4.0 cycles after the fault is used for sizing high-voltage circuit breakers. Fault current levels, possi- bly existing beyond 6.0 cycles, are needed to estimate the performance of time-delay relays and fuses. The calcula- tion result listed in table 10.4 refers to the ac symmetrical value of fault current. Figures 10.2 and 10.3, mentioned in the table, relate the system reactance-to-resistance ratio to the circuit-breaker contact parting time to obtain multi- plying factors for momentary or close-and-latch rating comparisons (3). To assist in using these curves, typical breaker speeds are 3, 5, and 8 cycles, with respective contact parting times of 2, 3, and 4 cycles (a 2-cycle circuit breaker has 1.5-cycle contact parting). As a summary of the foregoing procedures, the calcu- lation steps could be listed as follows: 1. Express all impedances between the sources and the fault in per-unit on a set of consistent bases. (In rare cases where the sources and fault are at the same voltage level, such conversion is not necessary.) 2. Reduce the system to one equivalent impedance. by 3. Calculate the three-phase bolted-fault current, I s V T - P u T Z PU (10.2) where V pu = driving voltage of circuit, 1.0 pu, if nominal voltages are taken as the prefault level, pu V, Z pu = magnitude of equivalent impedance, pu fi, I b = base current for system portion in which fault exists, A, and I sc = symmetrical rms fault current, A. 4. Apply appropriate multiplying factors to account for fault current asymmetry. 5. If line-to-line fault current is desired, I P = I.-V3 = 0.87 L (10.3) where I p = approximate symmetrical rms fault current resulting from bolted line-to-line fault, A. (Note that this is derived from the fact that positive- sequence and negative-sequence impedances are basically equal for phase-to-phase faults). For line-to-line-to-neutral faults, equation 10.3 also approximates the line-current level if the system is high-resistance grounded. Because of the high mobility of mine power equip- ment, any specific unit, say a switchhouse, could be installed anywhere, and the location is controlled more by the operation personnel than engineering personnel. It is often desirable that portable apparatus be able to with- stand or interrupt the worst fault conditions, and equip- ment should have a certain uniformity in design. To size this equipment correctly, repetitive fault calculations must be performed, assuming various fault locations throughout the mine system, to arrive at a worst case short-circuit current. More repetitions are usually needed at distribution levels than at utilization. Such extensive fault computations sometimes call for computer analysis, which will be discussed shortly. Table 10.4.— Sample applications of fault calculations Application Operation Fault current Comments 1 . Fuses and low-voltage circuit breakers. 2. High-voltage switching apparatus, excluding fuses. Contact parting and momentary withstand. Momentary rating or close-and-latch rating. 3. High-voltage circuit breakers.... Interrupting duty. 4. Time-delay relays Operation under short-circuit currents. First cycle maximum... Machine subtransient reactance, X" d , must be included. If ratings are symmetrical rms current, symmetrical fault current values are used directly. ..do Machine subtransient reactance, X" d , must be included. Use values given in table 10.1, except use 1.2 X" d for individual induction motors 50 to 250 hp, neglecting those below 50 hp. Multiply symmetrical fault current values by 1 .6 to obtain total short-circuit current. From 1.5 to 4 cycles Multiply all motors X" from table 10.1 by 1.5, except use 3.0 X" d after fault. for individual induction motors 50 to 250 hp, neglecting those below 50 hp. Determine X/R ratio of Thevenin's equivalent impedance. Select appropriate multiplying factor from figure 10.2 or 10.3. Beyond 6 cycles All motor contributions omitted. Procedure same as application 1. 264 3-phnsp 140 x| CD 00 O O o O ' /I s f >»l CD 1 REACTANCE o o ' 7 ' o^/ - 140 1 120 / I ] 100 ~~1 J1 1 7 < 80 J // 60 / t a >l 40 20 \rf>^ 'co ^ 1.4 1.6 1.0 1.2 MULTIPLYING FACTOR 1.4 5 2 or more transformations (remote generation) 4 Not more than 1 transformation (local generation) Figure 10.2.— Multiplying factors applied to three-phase faults to obtain momentary ratings for switching apparatus. -^n c o k f I I 'I ' C7 o s 7 4j 3-CYCLE - CIRCUIT - BREAKER 1 — t\j- f 1 a; >\ a> E Cn c 1 / / f 1 2-C - CIRC BRE rCLE :uit ■ iKER 1.0 1.1 1.2 1.3 1.4 1.0 1.1 1.2 1.3 MULTIPLYING FACTOR A Not more than I transformation ( local generation 1.0 1.1 1.2 1.3 1.01.1 1.0 I.I 1.2 1.2 1.3 1.4 MULTIPLYING FACTOR B 2 or more transformations (remote generation) Figure 10.3.— Multiplying factors applied to three-phase faults to obtain closeand-latch ratings for switching ap- paratus. Three-Phase Calculation Example lb illustrate an example of bolted three-phase fault computations, consider the one-line diagram in figure 10.4, which approximates an underground coal mine in its early stages of development. All cables are given by their conduc- tor size; transformers by their voltages, capacity, and percent reactance; utility supply by its ability to deliver short-circuit current (a kilovoltampere capacity); and motors by their horsepower. Transformer resistance, power-equipment bus work impedance, and interrupting-device impedance are assumed to be negligible. Three possible fault locations are indicated, but for this example suppose only fault 1 has occurred. To ensure correct results, the contribution of rotat- ing machinery is included. The initial direction is to compute the first-cycle maximum fault current. Following the recommended pro- cedure, the first concern is to convert all impedances to per-unit. For calculation convenience, a base power, kVA b , of 5,000 kVA is chosen, a level not too large to make the per-unit values of any component insignificantly small. Prefault voltages are taken as nominal system voltages and define the second required base quantities, or for 7,200-, 600-, and 480-V line to line converted to line to neutral, kV b69 = 39.838 kV, kV b72 = 4.16 kV, kV b 6 = 0.346 kV, kV b48 = 0.277 kV. Utility line 1,000,000 kVA short-circuit level 69 kV | A 3,000 kVA Z = 7.0% 7.2kv" 7.2 kvT M000 kVA ucuj Z = 5.0 /o 5 g?°up P 100h P l00h P 50h P O O Q Q SURFACE . '= u ", h P FACILITIES induction SWITCH HOUSE 1 600 750 kVA v - l - u ^ 7,200 I A 1 2,000 "3/C 4/0 MPF SWITCHHOUSE 2 -D — »- LOAD CENTER a 2,000' 3/C 1/0 MPF -a— * >■ ; Fault 2 Figure 10.4.— One-line diagram for fault calculations. 265 Calculations then provide the base current and base impedance for each base voltage. Thus, the base quantities for the 7,200-V system are At 600 V, At 480 V, kVA h = 5,000 kVA, kV h7? = 4.16 kV, ^b7.2 = 401 A, Zb7.2 = = 10.37 Q. kVA h = 5,000 kVA, kV b6 = 0.346 kV, h.e = 4,811 A, Zb.6 - 0.072 Q. kVA h = 5,000 kVA, kV h48 = 0.277 kV, Ib.48 = 6,014 A, Zk 4.R = = 0.046 fi. The base impedance and current at 69 kV will not be used. The next process in the calculations is to reference the power-system parameters to these base quantities. The short-circuit level of 1 million kVA shown in figure 10.4 relates to 1.0-pu impedance on the utility system at 69 kV. It is common practice to assume that this impedance is pure reactance, X pu , so the utility system base values can be taken as high side, and 69 kV, and is based on the transformer rated kilovoltamperes. The conversion is thus X % kVA b /kV e x 2 7(5,000) x - 2 = Took^ (kvj = Ir3o(3\oooj (1) = °- 117 pu - The 1,000-kVA load-center transformer has a percent reactance of 5% based on the rated capacity of 1,000 kVA and referred to the high side, 7.2 kV. Therefore, Xpu3 - 5 /5,000 100 \1,000 (1) = 0.250 pu. The 750-kVA load center has a 4.5% reactance similarly referenced, so 4.5 /5,000\ n nnn x p- = io1)(w) = - 300 p u - The resistance and reactance for each cable can be found from tables in chapter 8, and each impedance is easily changed to per-unit with Z -^ V - 7 (10.5) kVA e = 1,000,000 kVA, kV e = 39,838 kV, Xpue = 1-0 pu. But the calculation base values at 69 kV are kVA b = 5,000 kVA, kV b69 = 39.838 kV. The utility per-unit reactance must therefore be converted using or _ X pue kVA b /kV e pu kVA b IkV (1.0X5,000) /39.8 (10.4) ^ = lLW00b-feJ =00 ° 5PU - Likewise, the per-unit reactance of transformers in the power equipment must be referenced to the calculation base quantities. The main substation contains a 3,000- kVA transformer, which has a 7% reactance referred to the Table 10.5 provides the results of all these cable computa- tions. The reactances in table 10.1 can be used to approxi- mate the variable impedance of each rotating machine. As a time saver, it is often best to determine all reactances at this point and carry them through the subsequent calcu- lations. However, because there are only induction motors in this example, subtransient reactance, Xj, is the only value defined, with transient and synchronous reactances approaching infinity. The load in the 480-V system is a 150-hp group of induction motors with each motor less than 50 hp. Here the subtransient reactance is 0.25 pu, based on the motors combined with kilovoltamperes which is approximately the combined horsepower rating, or 150 kVA. This reactance referenced to the base power for the calculations is thus xs = (0.25X5,000) 150 = 8.33 pu. The 600-V system contains a 500-hp motor group, two 100-hp motors, and one 50-hp motor. Each motor in the 500-hp group is about 100 hp. Hence, from table 10.1, this Table 10.5.— Impedance of cables in figure 10.4 System, V Base Cable Actual impedance, ft Per-unit impedance impedance, fi Length, Size, ft AWG Type Voltage, kV Resistance (R) Reactance (X) Resistance (R) Reactance (X) 7,200. 480.... 600.... Borehole. 10.37 .058 .0752 500 2,000 2,000 1,000 500 500 500 4/0 4/0 1/0 4/0 4/0 1/0 2 MP + GC 1 MPF MPF G + GC G + GC G + GC G + GC 0.032 .126 .254 .068 .035 .064 .109 0.016 .064 .070 .027 .0135 .0145 .0145 0.0031 .0122 .0245 1.1724 .4514 .889 1.514 0.0015 .0062 .0068 .4655 .1875 .201 .201 266 subtransient reactance is about 0.17 pu on the combined kilovoltampere base, 500 kVA, and on the new base power Xa = (0.17X5,000) 500 = 1.7 pu. emphasizes the need to be cautious about neglecting the motor contribution. Nevertheless, this calculated value should be very close to the symmetrical component of fault current and can be used for time-delay relaying compari- sons (see application 4 of table 10.4). Continuing the process for the single-motor loads, the reactance for each 100-hp motor is X d ' = 0.17 pu on a 100-kVA base, X d ' = 8.5 pu on the 5,000-kVA base, and for the 50-hp motor, X d ' = 0.17 pu on a 50-kVA base, 17 pu on the 5,000-kVA base. X d ' = Motor resistance could also be assinged, but since the X/R ratio for induction machines is about 6, this is neglected. The next step is to draw an impedance diagram as in figure 10.5 to show the calculated per-unit resistance and reactance values. The diagram should then be simplified as much as possible to show clearly the data required for calculations, as in figure 10.6. Even though this example concerns fault 1 only, all three fault locations are shown. All rotating machines, including motors and any genera- tors, are represented by their per-unit reactances, X d , X d , and X d , and are connected to an equivalent source bus (the dashed line in figure 10.6). Thus the utility supply equiv- alent reactance is in parallel with the rotating-machine reactances, and the combined potential driving the fault is 1.0 pu. Further simplifications can now be made for each specific fault location. Figures 10.7 through 10.9 show the process graphically for fault 1. The operations involve no more than combining parallel and series impedances to obtain a single impedance between the source bus and the fault, which is the Thevenin's equivalent of the faulted system (fig. 10.9). The single equivalent impedance can now be used to calculate the first-cycle symmetrical rms current of the bolted three-phase fault, or ■ V = 1.0pu 0.005 pu . 0.117pu Only used to denote separate components : 0.0031 pu ! 0.0015 pu : 1.478 pu ; 0.587 pu ■ 0.25pu X^' = 8.33pu \ fO.OI22pu i 0.0062 pu ;; Fault! To signify that this is indeed a rotating machine 0.0245 pu VWV 0.0068 pu Fault 2 0.300 pu Fault 3 Figure 10.5.— Impedance diagram for one-line diagram of figure 10.4. I -S»I 1.0 0.1132 (401) = 3,542 A. With the fault located in high-voltage distribution, this value can be used to compare fuse ratings based on symmetrical rms currents. If the fault had been in a low-voltage portion, the current would also apply to circuit breakers. It is obvious that in this example, resistance could have been omitted with negligible error. Figure 10.10A is a simplification of figure 10.5 to demonstrate the error that could be introduced by neglect- ing the motor fault current contribution. As shown, the only driving source is the utility, and the equivalent circuit in figure lO.lOfi can easily be found by series combination. The first-cycle symmetrical rms current in this situation is 1.0 0.1235 (401) = 3,247 A, which is 295 A less than the previous value. Thus by omitting the motors, the result is an 8.3% error, which Shows source of driving potential X Fault 2 0.0245 pu 0.3068 pu Figure 10.6.— Simplification of figure 10.5. 267 If fault calculations are to be applied to high-voltage switchgear, the momentary or close-and-latch current du- ties must be calculated. The next step in the calculations would therefore be to account for fault current asymmetry by employing multiplying factors that adjust for the dc component decay. Here the subtransient reactance of all induction motors greater than 50 hp is modified by mul- tiplying by 1.2. The contribution of motors less than 50 hp is neglected. These adjustments can be compared to a decrement factor, allowing for fault current decay with time. Figure 10.1 LA reflects these changes and also omits resistance (compare with figure 10.6). Figure 10.1 IB shows the reduced single equivalent reactance, and using the formula, V pu Isc(mom) — rj lb d-6) (10.6) 0.122 0.0031 0.0015 1.478 . 9.167 X d ; 0.0367 ,( 0.313 Fault -Note : fault 2 removed • 0.486 > 0.444 1.888 X d " C 4.35 X d " C 17.201 X d Figure 10.7.— Simplification of figure 10.6. or Isc(mom) = OHM (401X1.6) = 5,559 A. This current represents the maximum asymmetrical rms value in the first cycle and can be compared with close- and-latch capabilities of high-voltage switching appara- tus. These are normally circuit breakers, but any device that needs to be rated on total fault current also applies. Interestingly, if subtransient reactances were not modified V = 1.0pu Figure 10.9.— Equivalent circuit of figure 10.6. V=1.0pu J0.I22 j 0.0031 . j 0.0015 Fault A Utility, substation, and borehole cables V = 1.0pu Figure 10.10.— Example problem with motor contribution neglected. ■=> ■ 0.0033 0.1219 : 0.2824 X Fault 1.543 Figure 10.8.— Further reduction of example network. A Complete network V-I.Opu J0.II54 B Equivalent circuit Figure 10.11.— Network to calculate momentary or close- and-latch current duties. 268 in this example, the calculations would produce an asym- metrical rms current of 622 A, thus creating a small safety factor for switching-apparatus comparison. To illustrate a further application of fault calculations (table 10.4, item 2), the circuit of figure 10.6 can again be adjusted to determine the interrupting duties of high- voltage circuit breakers that have minimum contact part- ing times between 1.5 and 4.0 cycles. Again, to allow for decrement of the dc component, subtransient reactances are multiplied by (3) 1.5, for synchronous motors, 1.5, for 3,600-r/min induction motors over 250 hp and 1,800-r/min induction motors over 1,000 hp, and 3.0 for induction machines of 50 hp and larger (item 2 supersedes this item where there is any overlap in coverage). Smaller motors are considered insignificant. To obtain the necessary multiplying factor from figures 10.2 or 10.3, the X/R ratio of the equivalent impedance or fault-point im- pedance is needed. This can be found by reducing resistance-only and reactance-only networks, which elim- inates handling complex numbers. The formula for inter- rupting duty is then where Iscont) = if* h (multiplying factor), L (10.7) j. sc( tl = total rms current interrupting duty, A, and multiplying factor = value from figure 10.2 or 10.3. For comparison with circuit breakers rated in megavolt- amperes, the interrupting capacity can be converted to an interrupting-current capability by (10.8) MVA Isdint) - ^3 y , v (- o UJ CL (/) O rr 10 - LL i i i i 1 i i i i 1 i i i i 5 10 DISTANCE FROM SUBSTATION. 10 3 A 15 Figure 10.13.— Available fault current versus distance of fault from rectifier on typical trolley systems. 270 but rather the possibility of having small undetected faults occur on trolley systems. In Pennsylvania, during 1978 alone, two extensive mine fires were initiated in this way. A trolley-wire-to-rail fault, for example, could occur as the result of a roof fall but be of high enough resistance so the fault current would be less than normal load currents; hence, the fault would not be detected by conventional overcurrent relaying. Several methods have been proposed to detect such illegitimate loads, including discriminating relaying and rate-of-current-rise detection (11). At this writing, these are still in the demonstration stage. DEVICE SETTINGS Coordination of a mine power system entails complete organization of time settings and/or current settings for all protective devices from the loads to the sources. This necessitates a comprehensive coordination study of the entire system to determine the range of correct values for all instrument transformers, pickup and time settings, fuse ratings, and circuit breaker trip ratings, which will provide effective coordination and selectivity and ensure that the minimum of unfaulted load is disturbed when protective devices isolate a fault. The prime concern is overcurrent, since the circuitry must provide simulta- neous overload, short-circuit and ground-fault protection without causing nuisance tripping. The application of this "art" is perhaps the most perplexing problem facing practicing engineers. A system fault analysis is vital input for any comprehensive coordination study and should address not only maximum values but also minimum values, together with the normal operation and maximum allowable currents. The balance of this chapter is broken into the major aspects of the coordination study: relay pickup settings, CT matching, circuit breaker trip settings, fuse character- istics, and overall coordination. Emphasis is placed on radial ac systems that are high-resistance grounded. RELAY PICKUP SETTINGS Pickup has already been defined as the minimum value of the actuating quantity that will cause a relay to operate its contacts. Whether the application is at a trailing cable, feeder cable, or overhead conductor, the requirements for overload, short-circuit, or ground-fault protection usually translate into current pickup settings. The values used in this chapter are generally in line with those contained in 30 CFR 75 and 77, which are in effect at this writing (16). The quantities that define pickup may change in the future, but the techniques presented here for establishing relay pickup settings should not. Because of their widespread use, induction-disk relays are implied in most of the following pickup applications; however, pickup techniques for other relays are basically the same. To avoid confusion, molded-case circuit breaker trip set- tings will be covered later. Establishing a pickup setting for an ac relay involves selecting a CT ratio and operating-coil current. For mar- ginal short-circuit currents, the accuracy of the combina- tion might require verification. The following material describes pickup settings for a single zone in a mine power system, but it must be remembered that the overall goal is to obtain coordination, and the settings at any location can be affected not only by requirements and regulations but also by other upstream and downstream relaying. Short-Circuit Protection Short-circuit protection can be obtained with an in- stantenous element (no intentional delay) or an inverse- time overcurrent relay using the minimum time dial setting (maximum time delay here is often restricted to no more than 0.6 s). The general requirements can be deter- mined by selecting the lower value calculated from the following: 1. 115% of the maximum starting current or 115% of the peak load current, whichever is higher, for the equip- ment being protected; or 2. 60% of the smallest bolted three-phase symmetrical rms fault current for any point of the zone protected by the relay. The first value is designed to have pickup above the normal operating current to prevent nuisance tripping. Usually, the bolted fault current value is higher than the first value. If the maximum starting currents of the motors are not known, each can be approximated by L = 1.25 x;') If (10.12) where I s = approximate maximum starting current, A, = per-unit subtransient reactance of motor or motor group, pu Q, and I n = motor full-load current, A. X3 The motor full-load current can be estimated from 746(hp) lfl " V3 Vi; (pf) ' (10.13) where hp = rated machine horsepower, V = rated line-to-line voltage of motor, V, rj = motor efficiency, and pf = full-load power factor, which can be assumed to be 0.85. Inrush currents for transformers usually range from 8 to 12 times the full-load current rating for a duration of 0.1 s; typical values for inrush currents should be available from the transformer manufacturer. To show how short-circuit pickup is selected, consider that a production shovel in a surface mine has 2,000-hp connected load rated at 4,160 V line to line. As shown in figure 10.14, the shovel is powered through 1,000 ft of 4/0 AWG cable. The shovel induction motor is 85% efficient at full load and operates at 0.8 pf. The minimum value of bolted three-phase fault current has been found to be 6,130 A. CT ratios are 400:5 A, and the instantaneous element has a pickup range from 10 to 50 A. The procedure is to find the minimum current according to the above criteria. First, for the fault current, the pickup would be 60% of the bolted value divided by the CT turns ratio: (0.6X6,130) AC . pickup = oq = 4b A. This value must now be compared with the pickup needed to slightly exceed maximum starting current, which may 271 be estimated using equations 10.13 and 10.12. Here, the full-load current of the shovel motor is about _ 746(2,000) ^ ~ V3 (4,160X0.85X0.8) = 305 A. From table 10.1, the per-unit subtransient reactance of individual induction motors is 0.17, and the estimated maximum starting current is then I s = 1.25 [——J (305) = 2,243 A. 50/51 o > — (f 4,160-V, 6,130-A symmetrical rms fault current 52 -D- y^K( I.OOOft 4/0 SHD-GC «r o 2,000-hp shovel Figure 10.14.— One-line diagram for pickup setting example. Allowing 115% to prevent nuisance tripping, the pickup setting is then • i (2,243X1-15) QQ pickup = on = 32 A. As this is less than the value for the fault current, 32 A is the selected pickup setting for short-circuit protection. Overload Protection The prime purpose of overload pickup settings is to protect conductors and insulation from damage by excess temperature. Temperature here is a function of the ambi- ent temperature and the I 2 R power loss in the conductors; temperature rise also involves time; therefore, inverse- time overcurrent relays are used on high-voltage systems to provide this protection. (The thermal overloads and fuses commonly used for low-voltage and medium-voltage systems will be discussed later.) The general overload requirement is that relay pickup should occur whenever 125% of the ampacity of the smallest power conductor in the protected zone is exceeded. When this requirement is directly applied to compo- nents such as transformers, the overload value could be 125% of the rated full-load current. However, the percent- age recommended for transformer overload protection var- ies depending upon the protection scheme associated with the transformer. For transformers rated greater than 600 V, the National Electrical Code allows 300% of rated primary current with circuit breaker protection, and 150% for fuses when there is no protection at the transformer secondary (10). With transformers rated less than 600 V, overload protection for the primary winding is 125% when there is no overload protection at the secondary and 250% if overload protection for the secondary is set at 125% (8). On the other hand, IEEE standard recommendations for transformer overload protection state that time-relay pickup should be set at 150% to 200% of the primary full-load current (3). Considering figure 10.14 as an example for overload pickup, the trailing cable is 4/0 AWG, three-conductor, 5-kV, with 90°C rated insulation, and the relay operating coil (for the 51 contacts) has a pickup range of 4.0 to 12 A. The maximum ambient temperature at this location is 30°C. From the data given in chapter 8, the cable ampac- ity at 40 °C is 321 A, and when applying the temperature correction factor for 30°C, ampacity = (1.1X321) = 353 A. Allowing for 125%, the minimum conductor current that defines overload is I = 1.25(353) = 442 A. The CT ampere-turns is 400:5 A, so the required pickup setting is 442 pickup = -qtt = 5.5 A. A tap setting corresponding to this pickup would allow the time-current characteristics of the relay to determine an overload condition. If the next tap setting available was 6.0 A, this would relate to 480-A conductor current as an overload, which is too high. If an instantaneous element with a 32-A pickup is available in the relay, short-circuit protection is also provided with one CT per line conductor. Ground-Fault Protection Most experts suggest that ground-fault relays should pick up at no more than 30% of the maximum current limit of the grounding system to provide an ample margin of safety in high-resistance grounded systems (9, 13, 19). For a 25-A current limit, this represents a line-conductor unbalance producing a zero-sequence current of about 8 A. However, such a demand is lower than present protective- circuitry detection capabilities when electromechanical relays are used. For instance, the optimum arrangement with induction-disk relays is zero-sequence circuitry with a 25:5 ampere-turns CT in which the most reliable pickup performance is not less than 12 A. One reason for this limit is connected to the fact that the window-type CT needs a large opening in order to pass the three line conductors through. Here, zero-sequence currents less than 12 A do not generate enough capacity from the CT to drive the relay adequately. Another problem is that induction-disk relay opera- tion is not reliable when the magnitude of actuating current is only slightly above the tap setting. This is because the net actuating force is so low that any addi- tional friction in the rotating-disk mechanism can prevent operation or increase the operating time. Even if the relay does close its contacts, the contact pressure may be so low that contamination of the contact surface can prevent electrical contact. To minimize this problem, it is common practice to apply induction relays in such a way that their actuating quantities are at least 1.5 times the tap setting (6). In fact, time-current curves are rarely shown for less than this amount. On the other hand, an induction relay is 272 most effective if its pickup is selected so it will operate on the most inverse part of its time curve. Thus, the mini- mum value of actuating current should be only slightly higher than 1.5 times the tap setting. Often a pickup corresponding to 2.5 times is selected for very inverse relays, the type usually applied to ground-fault protection. It might be thought that requiring relay operation below the ground current limit is too stringent and that relays should only be required to operate at the fault current available on the system. This is a logical deduc- tion, which relates that on a properly installed grounding system no potential greater than that allowed can ever exist between any metallic object and earth under fault conditions (40 V on low-voltage and medium-voltage sys- tems and 100 V on high voltage). Nevertheless, in terms of personnel safety, setting a relay pickup at less than the limited fault current provides viable backup protection. In other words, machine frames would remain well below the current allowed during any ground fault. Any trend to- ward hazardous conditions could also be detected, al- though such trends involving the system neutral are infrequent on high-voltage distribution systems. Considering these thoughts and in-mine practice, ground current pickup should not be greater than 50% of the current rating of the grounding resistor. This level should provide reliable repetitive operation of zero- sequence protective circuitry. As technology improves, 30% pickup should be the goal, as is suggested in the literature (9, 13, 19). To give an example of pickup settings, consider that ground-fault protection is provided by zero-sequence relay- ing A very inverse induction-disk relay is used and has a tap-setting range of 0.5 to 4.0 A for the time element coil. The CT has a 25:5 ampere-turns ratio, and the ground current limit is set at 25 A. Applying the 50% recommen- dation, zero-sequence current flow in the three line con- ductors cannot be greater than 12.5 A for relay pickup; thus, the maximum tap setting of the operating coil is 12.5/5 or 2.5 A. However, the selected tap setting should be lower to allow for effective time-delay operation. Using 2.5 times the minimum value of actuating current, the tap setting for the relay should be 2.5/2.5 or 1.0 A. On solidly or low-resistance grounded systems, the available ground-fault current is substantial enough that there are usually no relay pickup problems, even if resid- ual ground-fault relaying is used. Accordingly, the pickup level in this case could be defined by 30% of the minimum bolted line-to-neutral fault current. Similar logic could also apply to ungrounded systems, but here the maximum ground current requirement or even a ground-fault re- quirement is not easy to define. As a result, ground-fault pickup could, if necessary, be related to a decrease in the line-to-neutral potential of any power conductor and per- haps be 30% of the nominal system voltage. This line of reasoning could apply to dc as well as ac systems. CURRENT TRANSFORMER MATCHING From the foregoing, it appears that the selection of CT ratios and relay pickups to provide protection against excessive currents is quite a straightforward process. How- ever, CT's are magnetic devices and can give inaccurate results when improperly applied. CT performance is an important factor in protective-relay design because relays are only as accurate as the CT's that energize them. The main CT problem is core saturation caused by excessive primary current, incorrect secondary burden, or a combi- nation of both these factors {4). When a CT is in satura- tion, its accuracy is very poor, secondary current is actu- ally less than it should be, and relays tend to operate more slowly than intended. One danger is the loss of relay coordination. Current Transformer Accuracy A model of a CT and its burden is shown in figure 10.15. The dependent current source delivers a current equal to the primary current (I p ) divided by the turns ratio (N) of the CT. The remaining impedances, voltages, and currents are defined as Z e = secondary excitation impedance, Q, Z s = secondary winding impedance, ft, Z b = burden impedance, ft, E s = secondary excitation voltage, ft, V t = secondary terminal voltage, V, I e = secondary excitation current, A, and I s = secondary current, A. The burden refers to the impedance of the external load applied to the CT secondary, including the impedance of the relay, its associated wiring connections, and any meters. As shown in figure 10.15, a portion of the gener- ated current (I e ) is consumed in exciting the CT core. The remainder of the generated current (I s ) is the true value of the secondary current. The percent ratio correction error (IJIJ is defined as that factor by which the nameplate ratio of a CT must be multiplied to obtain the true ratio (6). It is apparent that this error will remain small as long as the excitation current is small. The magnitude of the excitation current is a function of the excitation voltage (E s ) as illustrated in the typical secondary-excitation characteristics of figure 10.16 (4). It can be noted that the secondary-excitation curves are linear until the saturation point is reached. Beyond the saturation point, a small increase of E s results in a large increase of I e , which in turn causes the percent ratio error to increase dramati- cally. The magnitude of E s is primarily a function of the secondary current (I s ) and the burden impedance (Z b ). There- fore, to minimize inaccuracies of the relay system, the burden impedance should be kept as low as possible. The pickup value for the relay must always result in the excita- tion current's lying in the linear portion of the secondary- excitation characteristic. It should also be noted that the impedance of electromechanical relays is not constant. Since they are magnetic devices, their impedance decreases as the secondary current increases because of saturation. Thus, the Z c X, Figure 10.15.— Model of CT and its burden. 273 < O > o u X UJ ig^ 500/5 ^?-400/5 300/5 O^/R 0.005 001 0.05 0. 5 10 30 SECONDARY EXCITING CURRENT A Figure 10.16.— Typical set of saturation curves for 600/5 multiratio bushing-type CT. relay impedance should be considered over the entire opera- tion range of currents when matching the relay to the CT. Beyond keeping secondary burden as low as possible, there are other recommended guidelines for protective relaying that help to minimize the effects of saturation. CT turns ratios should be maintained as high as practical. Usually, saturation is only a problem on low-ratio current transformers, such as 50:5 versus maybe 300:5 ampere- turns. CT secondaries with a higher ratio for a specific application develop higher voltage and are less likely to be saturated under normal burden. Hence the lower the ratio, the greater is the chance that a fault will not be cleared within the intended zone. Operation of an upstream cir- cuit breaker could then cause substantial outages. Since underutilizing a CT also produces inaccuracies, the maximum anticipated load current should be as close to the CT current rating as possible without exceeding it. On CT's with 5-A secondaries, IEEE suggests that the CT be operated at 3 to 4 A during normal full-load currents (4). It has also been suggested that the CT and relay values be chosen such that the relay tap setting is at least one-half the current rating of the CT secondary (5). If this cannot be achieved, the performance of the CT should be checked. In applications where saturation must be completely prevented, the CT should be sized to carry twice the peak flux associated with the symmetrical ac fault current (5). In most cases, it is not necessary to prevent saturation totally in order to provide adequate relaying. For instance, immediately after the initiation of a fault, the fault current might contain a substantial dc component. This component could cause the CT to become saturated, but in most mining applications, the dc component decays in less than x /i cycle. As high-voltage circuit breaker trip times are usually 3 to 4 cycles, the saturation during the first l /z cycle would be insignificant in terms of protection. If operation is in the linear curve portion, the CT would be functioning satisfactorily. Nevertheless, the maximum available fault current should be not than 20 times the CT current rating, and CT performance should be checked if the maximum is above this limit (6). Accuracy Calculations For a CT under a specific burden, the accuracy of its operation (transformer performance) can be calculated easily using manufacturer data, provided that the CT has a "C" accuracy class. As an example, suppose a 600:5 multiratio CT is connected through 50 ft of No. 12 AWG conductor to an induction-disk relay and an ammeter. Both overload and short-circuit protection are intended. Figure 10.16 gives the typical set of CT saturation char- acteristics, and burden data are contained in table 10.6 (4). The primary current (system line conductor) has 24,000-A symmetrical rms of available fault current. The objective of the calculation procedure is to find the percent ratio error. Table 10.6. -Burdens of relay elements and ammeter connected to CT's Element Relay, timed element, 4-12 A pickup. Relay, instantaneous element, 10-40 A pickup. Ammeter Conductor (interconnecting) Transformer secondary resistance... Burden 2.38 VA at 4 A at 0.375 pf; 146 VA at 40 A at 0.61 pf. 4.5 VA at 10 A; 40 VA at 40 A at 0.20 pf. 1 .04 VA at 5 A at 0.95 pf. 0.08 fiat 1.0 pf. 0.298 Q at 25°C. Although the error should not exceed 10% under the most extreme circumstances (20 times the secondary cur- rent), satisfactory protective-relay operation often calls for smaller error, on the order of 2% or less. To compute the 274 error, CT secondary burden and relay operating voltage must be known. The voltage shown in figure 10.16 is related to the CT secondary exciting current, the source of the secondary current error. Burden should be calculated at maximum secondary current. In most cases this is the pickup current for the element providing short-circuit protection. Thus, the in- stantaneous element of the relay is assumed to be set at 40 A, which is 8 times rated secondary current or 4,800 A primary current. It was stated in chapter 5 that a direct summation of the burden voltampere ratings will produce adequate results, and while this is generally true, for precise work the impedance of each CT load should be found and then summed. The time-element burden is 146 VA at 40 A and 0.61 pf, or an impedance, Z t = rnlcos I 2 l!pf = :^|52< (4or Therefore, the percent ratio error correction can be found by % ratio error = -^ (100), (10.9) or at a secondary current, I s , of 40 A, % ratio error = -^-(100) = 0.1%. Obviously, the accuracy for this application is extremely good. Only one accuracy calculation example Sas been shown because as long as the data are available, this technique can be used for any CT application at any pickup current. or Z = 0.091 |53^ = 0.0557 + jO.0723 Q. For the instantaneous element, S = 40 VA at 40 A 0.20 pf or Z= = 40 5 |78^ = 0.005 + jO.025 fi. (40) For the ammeter, negligible saturation will occur at 8.0 times rated current (essentially an air-core magnetic cir- cuit), and its impedance at 5.0 A is about the same as at 40 A, Z a = (1.04) (5) 2 18° = 0.039 + jO.012 U. The interconnecting wire resistance (R w ) and CT resis- tance (R ct ) are 0.08 and 0.298 fi, respectively. Conse- quently, as all burdens are in series, the total burden impedance at 40 A is Z = Z t + Z ; + Z a + R w + R ct = 0.477 + jO.110 = 0.489 | 12.9° Q. The magnitude of CT secondary voltage needed to produce 40 A is then V = IIZI = (40X0.489) = 19.6 V. From figure 10.16, the secondary exciting current, I e , at this voltage is L = 0.03 A. LOW-VOLTAGE CIRCUIT BREAKER TRIPS In the preceding discussion, overload and short-circuit pickup settings were primarily related to external relays driven by CT's. The principles of pickup can also be applied to low-voltage power circuit breaker settings and the instantaneous magnetic settings of molded-case units. In this context, pickup would be the minimum current within a specified tolerance (usually + 10%) that would cause the trip element to activate the operating mecha- nisms (4). Although not entirely accurate, pickup can be loosely used to describe overload current as specified by the thermal trip elements of molded-case circuit breakers. However, as element tripping is a function of stored heat, tripping times rather than pickup are often employed. This section will focus on overload and short-circuit trip settings with molded-case breakers and then briefly out- line trip levels for low-voltage power circuit breakers. Overload Protection Overload protection by thermal-magnetic molded-case breakers is mentioned in chapter 9. To review, most manufacturers calibrate their breakers to carry 100% of the continuous current at 40° C; thus, maximum contin- uous current is the rating of the thermal-trip unit. Derat- ing is only necessary in noncompensating devices when the ambient temperature exceeds 40° C. With these thoughts, the overload procedure given for relay pickup protection can be adapted to this continuous current. The thermal element itself defines overload at 125%. For instance, the minimum trip time for a typical unit is around 10 s for 500% or more of the current rating, with time increasing inversely to about 1,800 s at 135% of the current rating (20). Sections 75.900 and 77.900, 30 CFR (26), define breaker use according to Federal law for short-circuit and overcurrent (overload) protection of low-voltage and medium-voltage power circuits serving three-phase ac equipment. No further references are made to overcurrent protection except in section 75.900-2(c), which specifies that a circuit breaker protecting more than one branch circuit must be sized to afford overcurrent protection for 275 the smallest conductor. Despite these references to over- current protection, Federal regulations are regularly in- terpreted and enforced to mean that overload protection is not required at each circuit breaker protecting a trailing cable (17). Perhaps the lack of mandatory cable overload protec- tion can be best explained in the following quotation from a design engineer working with ac power centers (21). Breaker thermal trip ratings should be applied to insure complete continuity of operation through all overload current peaks rather than to give complete cable protection. Many applications are being made with no breaker overload protection. Short circuit protection only, by means of the adjustable magnetic trip units, takes care of line to line cable faults. Motor control equipment on the machine takes care of motor overloads and may give some overload protection to the cable. Some States have more stringent requirements than the Federal law for overload protection. For example, Pennsylvania requires overload protection on all breakers serving trailing cables (12). However, there are no specific requirements for sizing this overload protection, and so the engineer must select an appropriate size for the thermal- magnetic circuit breaker. Short-Circuit Protection The most common short-circuit protection is for trail- ing cables and is provided by the magnetic-trip elements of molded-case breakers. For surface mines, the pickup re- quirements listed for relay pickup protection are used; however, for underground coal mines, a set of maximum instantaneous settings based on conductor size are man- dated in section 75.601-1, 30 CFR (16). These settings, listed in chapter 9, table 9.3, apply to both ac and dc systems and are based on (2): 1. An ideal 250-Vdc source feeding a bolted fault at the extremity of a 500-ft two-conductor trailing cable, and 2. A 50% safety factor to allow for circuit breaker tolerance, system impedance, and so forth. These values are presently in effect for all underground coal mine trailing cables. Fesak (2) has expressed concern that the dc basis of these values does not provide an adequate safety margin for resistance-grounded three-phase cables and has recom- mended a new set of maximum allowable instantaneous settings, which are shown in table 10.7. Here, a specific setting is selected not only by conductor size, but also by cable length and system voltage. The table stops with 4/0 cables because this is the maximum practical trailing- cable size for underground mines, but the recommenda- tions extend to 1,000 MCM. The calculations on which these values are based assume an arcing line-to-line fault to find the minimum short-circuit current. The multipliers to obtain the minimum values from bolted three-phase fault currents are 0.85 at 480 V, 0.9 at 600 V, and 0.95 at 1,040 V. Similarly, Vilcheck (18) has recommended a refined set of maximum instantaneous settings for dc trailing Table 10.7.— Recommended instantaneous trip settings for 480-, 600-, 1 ,040- V three-phase trailing-cable protection Conductor size, AWG Cable length, ft Maximum instantaneous circuit breaker setting, A 480 V 600 V 1,040 V 14.. 12.. 10.. 8.... 6.... 4.... 3.... 2.... 1.... 1/0. 2/0. 3/0.. 4/0.. 0- 500 0- 500 0- 500 0- 500 0- 550 0- 500 501- 600 0- 500 501- 650 0- 500 501- 600 601- 700 0- 500 501- 601- 0- 501- 601- 751- 600 750 500 600 750 800 0- 500 501- 600 601- 750 751- 850 0- 500 501- 600 601- 750 751- 900 0- 500 501- 600 601- 750 751-1,000 75 125 200 300 400 700 600 850 700 1,000 900 750 1,200 1,050 850 1,400 1,250 1,050 1,000 1,600 1,400 1,200 1,100 1,900 1,700 1,450 1,300 2,050 1,850 1,650 1,350 100 150 250 400 550 850 750 1,050 850 1,200 1,050 950 1,350 1,200 1,050 1,550 1,400 1,200 1,150 1,700 1,550 1,350 1,250 1,950 1,800 1,600 1,450 2,100 1,950 1,750 1,500 NAp NAp NAp NAp 850 1,200 1,050 1,350 1,150 1,450 1,350 1,250 1,600 1,450 1,350 1,750 1,650 1,500 1,450 1,800 1,750 1,650 1,550 1,900 1,800 1,700 1,600 1,950 1,900 1,800 1,650 NAp Not applicable. cables, based on minimum expected short-circuit current. These pickups are given in table 10.8 and correspond to conductor size, cable length, system voltage, and the method for grounding. Two types of grounding are in- cluded in the table: by means other than a grounding conductor, such as diode grounding, and via a grounding conductor. Where there is no grounding conductor, only line-to-line faults are of concern and the t-i values in table 10.8 are used. With grounding conductors, line-to-line ground faults can also occur, and if no ground-fault protec- tion is used, a lower circuit breaker setting is needed for cables No. 6 AWG and larger because of the smaller grounding conductor ((-g values in table 10.8). Low-Voltage Power Circuit Breakers Trip settings for low-voltage power circuit breakers can involve three direct-acting elements (4). The require- ment for long-time delay elements is the same as for the overload discussed in relay pickups, and the unit is usu- ally set at 100% of the rating, regardless of tolerance. In general applications, the short-time-delay element is set at five times the overload current point, with the instan- taneous element set at nine times the overload level. However for mining applications, except for trailing-cable protection, the short-circuit requirements covered in relay pickups also apply to the instantaneous device. Trailing cables demand the same maximum short-circuit settings discussed in the preceding section. 276 Table 10.8.— Recommended instantaneous trip settings for 300- and 600-Vdc trailing-cable protection (19) Conductor size Cable length, ft 300-V maximum instantaneous setting, A i-e 1-9 600-V maximum instantaneous setting, A t-i t-g AWG: 14 0- 500 50 50 50 50 12 0- 500 75 75 100 100 10 0- 500 75 75 200 200 8 0- 500 100 100 300 300 6 0- 500 200 100 450 350 4 0- 500 400 250 650 500 501- 600 300 150 550 450 3 0- 500 500 350 700 550 501- 650 400 250 600 450 2 0- 500 600 400 800 650 501- 600 500 350 750 600 601- 700 450 300 700 550 1 0- 500 850 600 1,350 1,050 501- 600 700 500 1,200 900 601- 750 600 400 1,050 750 1/0 0- 500 1,050 750 1,550 1,200 501- 600 900 600 1,400 1,050 601- 750 750 500 1,200 900 751- 800 700 450 1,150 850 2/0 0- 500 1,250 900 1,750 1,400 501- 600 1,100 750 1,600 1,250 601- 750 900 600 1,400 1,100 751- 850 800 550 1,300 1,000 3/0 0- 500 1,500 1,100 1,950 1,600 501- 600 1,300 950 1,800 1,450 601- 750 1,100 750 1,600 1,250 751- 900 950 650 1,450 1,100 4/0 0- 500 1,750 1,350 2,150 1,800 501- 600 1,550 1,150 2,000 1,650 601- 750 1,350 950 1,800 1,450 751-1,000 1,050 750 1,550 1,200 MCM: 250 0- 500 1,950 1,500 2,300 1,950 501- 600 1,750 1,300 2,150 1,800 601- 750 1,500 1,100 1,950 1,600 751-1,000 1,200 850 1,700 1,350 f-g Line to ground (with grounding conductor). l-l Line to line (without grounding conductor). FUSES Overload and short-circuit protection can be provided by dual-element fuses on low-voltage systems or by power fuses on high-voltage systems. The sizing of fuses is basically the same as that for relays or tripping devices in conjunction with circuit breakers. In cable overload pro- tection, the continuous-current rating of a fuse cannot exceed the lowest ampacity of any conductor in the pro- tected system. Beyond this, as a general protection guide, fuses should always be sized to the very smallest continuous-current rating that will safely carry any load that should be maintained (4). Transformer inrush currents, if not heeded, can be a problem with high-voltage fuses and for relay pickup as well. For instance, upon energizing, a power transformer can draw 8.0 to 12 times its full-load current rating for a duration up to 0.1 s (4). Expulsion and boric acid fuses can be obtained that have inverse-time characteristics to with- stand this, thereby affording both overload and short- circuit protection for the transformer. High-voltage current-limiting fuses, because of their fast response, must often have a continuous-current rating to account for inrush. In these instances, current-limiting fuses can provide only short-circuit protection. The use of fuses to provide ground-fault protection is more complex. In high-resistance-grounded systems, fuses cannot be used to protect against ground faults; in fact, their use in series with grounding conductors can create a serious personnel hazard if they open. During line-to- neutral faults, the current in these systems is limited to a level lower than typical load current in the line conduc- tors; line-to-line-to-neutral faults cause about the same current levels. As there are no neutral-connected loads in these systems, fault-current contribution is only from generating stations and capacitance; that is, motors have no effect. However in ungrounded or solidly grounded systems, ground-fault currents can be substantially greater; thus, line-interrupting devices including fuses can provide ground-fault protection. For instance, in an ungrounded system, fuses sized to the short-circuit requirements dis- cussed in this chapter would probably cause interruption during two simultaneous line-to-neutral faults on different power conductors. COORDINATION The basic principles of coordination have already been discussed in chapter 9 and earlier in this chapter, but here the specific procedures used in coordination studies will be demonstrated. References 1 -2 are highly recommended for more detailed coverage of this topic. A coordination study is basically a comparison of the times-to-operate for individual protection devices under both normal and abnormal current flows (3). A prelimi- nary study should always be made in the early stages of mine power-system design, since it could indicate that transformer sizes or impedances require modification or that cable sizes need to be changed. After final selection of devices and components, a second study must be made to confirm the tentative study. A new coordination investiga- tion is required whenever the original system is changed, whenever new loads are added, or whenever existing equipment is replaced with higher rated components. An additional study should always be made if a fault causes a major shutdown of an existing system. Coordination has two conflicting objectives: protection and selectivity (3). Each relay, tripping unit, or device is designed to protect a specific circuit segment and for maximum safety must isolate each faulted circuit as fast as possible. At the same time, continuity of service must be provided, and this requires selectivity. Again, selectiv- ity is the concept of isolating only that portion of the system experiencing failure. This demands successively longer fault-clearing times or higher pickup settings for each protection zone from the loads to the source. It is the role of the coordination study to optimize these require- ments without compromising personnel safety. A coordination-curve plot is one of the best methods for achieving these objectives. For resistance-grounded systems, two plots can be made, one for line-conductor protection and the other for ground faults. The protection described here refers mainly to line-conductor protection plots but also applies to grounding. Before a curve plot can be constructed, a good one-line diagram of the system must be prepared. In addition to the 277 normal diagram features discussed in chapter 4, it should include the following information {3-4): 1. Maximum and minimum levels of short-circuit current that are expected to flow through each protective device (obtained from a system fault analysis); 2. Normal and maximum load currents for each sys- tem portion (from a load-flow analysis); 3. CT ratios; and 4. Relay, circuit breaker, and fuse ratings and adjust- ment ranges (time-current characteristics must also be known). With these data available, as shown in figure 10.17, the coordination study can proceed to plotting. Coordination curve plots are normally drawn on stan- dard log-log graph paper, using the vertical scale for time and the horizontal axis for a common current source (3). For good results, this scale usually corresponds to currents at the lowest voltage level, while currents at higher voltage levels are plotted on the same scale as equivalents. For instance, if utilization is at 600 V with distribution at 7,200 V, the current scale is calibrated for 600 V. Currents in those system portions are plotted directly, but distribu- tion currents are multipled by 7,200/600 or 12, then plotted. However, note that this is not a firm rule, as any convenient scale can be used. Important current levels such as transformer inrush, motor starting, and fault current can be placed on the graph as points. A critical protection path is then chosen from the one-line drawing by selecting a critical load or safety hazard and following the protection path back to the source. Fortunately, because of radial distribution, this selection is an easy matter in most mine power systems. The time-current characteristics of all protective devices along the path or affecting the path are then plotted on the graph. As shown in figure 10.18 (4), such a plot is invaluable as it allows protection-device action to be visualized throughout the system on the same current basis. This is an advantage similar to that of per-unit analysis. It also provides for the normalization of the characteristics for each protection device, which rarely have the same shape or even scale. The plot simplifies the selection of current and time settings that will provide the best possible protection and safety while also giving selectivity. One way to do this is by specifying the characteristics of the most downstream protection, then sequentially position- ing each protection step toward the source (or vice versa). As is evident in figure 10.18, operating time intervals must be maintained between protective devices in order to 100 E 34.5-kVfuse 34.4-kV <^_Lkj A 3,750-kVA,Z = 6% 4,160-V ^ry-^ _£_ 800/5 300/5 } (tt 100-A ) molded-case circuit breaker 40 100 400 CURRENT, A at 4,160-V multiply by 10 " 480 -V " " 87 " 34.5-kV " " 1.21 4,000 10,000 Figure 10.17.— Example of one-line diagram for preparing a coordination curve plot for one path. Figure 10.18.— Coordination curve plot for figure 10.17 showing various protective-device characteristics. 278 achieve correct sequential operation. When applied to relay-activated circuit breakers, this time margin must allow for circuit breaker interrupting time, overtravel time, and a safety factor (4). All these times are additive. The time required by a circuit breaker to interrupt current once it receives the trip signal is equal to its speed in cycles divided by 60. A typical power circuit breaker has an operative speed of 5 cycles, corresponding to an 0.08-s operating time. Overtravel is important in electromechan- ical time-delay relays, especially inverse-time, induction- disk types. Owing to moving-part inertia, these relays will continue to close their contacts even after the fault is removed, and overtravel can thus create a relay operation delay. Typical overtravel time for inverse-time, induction- disk relays is 0.10 s. The safety factor mainly allows for variations in relay characteristics due to manufacturing differences, tolerance, aging, and dust; the commonly assigned range is 0.12 to 0.22 s. Summing these factors results in a necessary time margin range of 0.3 to 0.4 s (4). While the typical case is considered to be 0.4 s, the margin can be reduced to 0.3 s for carefully tested systems. Note that static relays eliminate overtravel and allow the use of the minimum safety factor, so that the typical time margin can be reduced to 0.2 s (see chapter 14). The typical time interval of 0.4 s is used for relay- to-relay coordination in figure 10.18, and the time selec- tions for line-conductor relays can be seen. Although this time interval considered only external relay tripping, the 0.4-s time margin is often utilized even when direct-acting low-voltage circuit breakers are coordinated together or with relayed interrupters. The same is true when a circuit breaker is coordinated with an upstream fuse if the total fuse clearing time is greater than 1.0 s. With less than 1.0 s, the fuse-to-circuit-breaker time margins can be reduced as low as 0.1 s (4). For an additional example of coordination, consider ground-fault protection in a high-resistance-grounded dis- tribution system. A separate plot can be made for the ground-fault conditions. If the induction-disk relay at the most downstream switchhouse is set to pick up with the minimum time dial setting, the time dial on the ground- fault relay in the next upstream switchhouse would be set 0.4 s higher. At each subsequent upstream switchhouse, the ground relay time would be set progressively higher until the neutral bushing of the source transformer is reached. These time allowances also involve the current equivalent of any potential relaying about the grounding resistor. Because of the need to maintain the time inter- vals, many engineers have found the coordination of more than five relays in one ground-fault current path nearly impossible, since the maximum practical time setting is about 2.0 s. Although coordination curve plotting has been discussed here, in practice the manufacturer characteris- tic curve could be used in this case. The reason is that available ground-fault current in a high-resistance- grounded mine distribution system remains fairly con- stant, so that the same very inverse relays could be used for all ground-fault protection, with the possible exception of resistor potential relays. The material presented in this chapter and the last has perhaps given the impression that in order to create a safe mine power system all circuits should have overload, short-circuit, and ground-fault protection. While this is generally true, it is not quite the entire picture. Worker safety must always come first; hence, equipment protec- tion is not applied in those cases where its use can cause hazards to personnel. These exceptions almost always involve unit-designed equipment, such as elevator motors or the swing, hoist, and propel motors of surface excava- tors. The only other exceptions to mandatory equipment protection are cases where there is no possibility of com- promising personnel safety if it is omitted. This chapter has introduced the important parame- ters used in the selection and adjustment of protective devices. This information together with the material in chapter 9 serves as one basis for chapters 12 and 13, where the components and techniques are applied to the equip- ment used in the mine system. But first another type of protection must be discussed, protection against the some- times hidden dangers of transients and overvoltages. REFERENCES 1. Crites, W. R., and A. G. Darling. Short-Circuit Calculating Procedure for D-C Systems With Motors and Generators. Trans. Am. Inst. Electr. Eng., Part 3, v. 73, Aug. 1954. 2. Fesak, G., W. Helfrick, W. Vilcheck, and D. Deutsch. Instan- taneous Circuit Breaker Settings for the Short Circuit Protection of Three Phase 480, 600 and 1040 V Trailing Cables. Paper in Con- ference Record -IAS 12th Annual Meeting (Los Angeles, CA, Oct. 1977). IEEE, 1977. 3. Institute of Electrical and Electronics Engineers (New York). Recommended Practice for Electrical Power Distribution for In- dustrial Plants. Stand. 141-1986. 4. Recommended Practice for Protection and Coordina- tion of Industrial and Commercial Power Systems. Stand. 242-1986. 5. Kiefer, J. A., and J. L. Kohler. Ground Fault and Overcurrent Protection Criteria for Coal Mine AC Distribution Systems (con- tract J0395035, Ketron, Inc.). BuMines OFR 158-81, 1980; NTIS PB 82-137001. 6. Mason, C. R. The Art and Science of Protective Relaying. Wiley, 1956. 7. Morley, L. A., F. C. Trutt, and R. A. Rivell. Coal Mine Elec- trical System Evaluation. APL Mine Electrical System Load-Flow Program (grant G0155003, PA State Univ.). BuMines OFR 61(7)-78, 1977; NTIS PB 283 496. 8. Coal Mine Electrical System Evaluation. Extended APL Mine Electrical System Load-Flow Program (grant G0155003, PA State Univ.). BuMines OFR 55-78, 1977; NTIS PB 283 497. 9. Myers, W. P. Current-Limited Ground Fault Relaying. Min. Congr. J., v. 59, Apr. 1970. 10. National Fire Protection Association (Quincy, MA). National Electrical Code. NFPA 70-1981 (ANSI Cl-1981). (Updated every 3 yr.) 11. Paice, D. A., and A. B.. Shimp. Discriminating Protection for Trolley Wires. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 12. Pennsylvania Department of Environmental Resources. Bituminous Coal Laws of Pennyslvania, Sec. 333. 1961. 13. Poker, L. R. Economical Ground Fault Protection Available With a Standard Low-Voltage Tripping System. IEEE Trans. Ind. and Gen. Appl., v. 6, Mar./Apr. 1970. 14. Shimp, A. B., and D. A. Paice. Application of Molded-Case Breakers on DC Electrical Systems in Coal Mines. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 15. Stagg, G. W., and A. H. El-Abiad. Computer Methods in Power System Analysis. Mc-Graw-Hill, 1968. 16. U.S. Code of Federal Regulations. Title 30-Mineral Resources; Chapter I -Mine Safety and Health Administration, 279 Department of Labor; Subchapter O-Coal Mine Health and Safe- ty; Part 18 -Electric Motor-Driven Mine Equipment and Ac- cessories; Part 75 -Mandatory Safety Standards, Underground Coal Mines; Part 77 -Mandatory Safety Standards, Surface Coal Mines and Surface Work Areas of Underground Coal Mines; 1981. 17. U.S. Mine Safety and Health Administration. Coal Mine Safety Electrical Inspection Manual, Underground Coal Mines. Apr. 1979. 18. Vilcheck, W., G. Fesak, and W. Helfrich. Instantaneous Cir- cuit Breaker Settings for the Short-Circuit Protection of Direct Current 300 and 600 Volt Trailing Cable. Paper in Conference Record -LA.S 13th Annual Meeting (Toronto, Ontario, Canada, Oct. 1978). IEEE, 1978. 19. Wade, E. C. Ground Relaying for Mining Distribution Systems. Coal Age, v. 71, July 1966. 20. Westinghouse Electric Corp., Low-Voltage Breaker Div. (Beaver, PA). Breaker Basics. 1973. 21. Youel, V. H. The Underground A-C Mine Power Center. Mechanization, v. 24, Aug. 1960. 280 CHAPTER 1 1 .—TRANSIENTS AND OVERVOLTAGES The term electrical transient can mean different things to people of different interests. However, there are some common ideas. As defined by Greenwood (19), 1 "an electrical transient is the outward manifestation of a sudden change in circuit conditions, as when a switch opens or closes, or a fault occurs on a system." The circuit parameters of inductance and capacitance are found in any power system to some extent. When the system is changed, the quantities of current, voltage, magnetic flux, and so on, do not instantly assume new values. Rather, they go through a transition to reach the new steady-state condition (33). It is this transition period that gives rise to the transient voltages and currents. Transients are of short duration, and the time in which they occur is an extremely small percentage of the total operating time. But it is during these short periods that some of the greatest electrical stresses can occur, mainly because of excessive currents or voltages. The excessive voltage is often critical in the design of mine power systems. In extreme cases, system parts are dam- aged and equipment failure follows. Safety can be further compromised because anomalous voltages may exist at machine frames. Transients are a fact of life on every power system, yet careful design within the technical and economic con- straints of a system can result in a reduction of transient- related component failures. However, mine power systems are particularly vulnerable to transient-induced failures because their operation and arrangements are extremely dynamic. As a mining activity advances, the electrical system is expanded, often on a weekly basis. Although this expansion is normally designed into the system, circum- stances in the mine can call for additional modifications. In some cases, these on-the-spot modifications may not be in line with sound engineering principles. A lack of knowledge about transients has resulted in power-system requirements that could be misunderstood by some mine power engineers. Furthermore, in the pro- cess of maintaining the system, faulty components are sometimes replaced with devices having different specifi- cations or even total incompatibility. These two factors also increase the vulnerability of mine power systems to transient occurrences. TRANSIENT SOURCES There are several discrete sources of transient over- voltages. Although slight differences exist in classifying these events, IEEE standards catalog seven types (23): references 19, 33, and 37 are invaluable sources of tran- sient information, and these should be consulted for de- tailed coverage of transient phenomena and problems. LIGHTNING PHENOMENA On a global scale, the earth and the lower part of the ionosphere may be considered as conductive bodies sepa- rated by a rather poor conductor, the atmosphere. The entire system is analogous to an enormous capacitor with a leaky dielectric. Any charge unbalances that accumu- late are dissipated by sudden breakdowns in the dielectric (atmosphere), and the resulting current (lightning) is of short duration but high amplitude. At any given time, electrical storms are taking place somewhere on earth, and the average current flow between air and earth is more or less constant at a level of 1,500 A. During fair weather, the normal voltage gradient in the air near the earth's surface is about 3.0 V/cm, but this rises to around 500 V/cm beneath a thundercloud. The potential differ- ence needed to initiate a lightning stroke is on the order of 50 million V (12). In the continental United States, a typical stroke lasts for only a fraction of a second, yet releases a tremendous amount of energy, approximately 200 million J. The total quantity of charge involved in the average stroke is about 20 C, and the peak current value is around 20 kA (28). As shown in figure 11.1, the shape of the discharge starts with an extremely short-duration voltage pulse, whose crest may reach 5 MV or more (29). This is closely followed by the current waveform, which rises more slowly to its peak value and lasts longer. About 82% of the strokes occurring in the United States are negative in polarity; that is, they transfer electrons from the clouds to earth. Figure 11.2 illustrates the wide distribution of current magnitudes that may be expected in various lightning strokes (27). Figure 11.3 shows the expected number of thunder- storm days per year for any of the 48 contiguous States (27). The probability that a particular object will be hit by lightning depends upon the cloud charge intensity, the geographical locale, and the height of the object in relation to its surroundings. Charge intensity is a parameter because an increased charge in the tip of the stroke generates a proportionally higher electric-field gradient. This determines the attractive range of the stroke, in other words, the horizontal distance from the tip of a downcoming leader to an object receiving the stroke. As 1. Lightning, 2. Switching surges, 3. Static, 4. Contact with a higher voltage system, 5. Line-to-ground faults, 6. Restriking ground faults, and 7. Resonant conditions. The following paragraphs introduce these sources briefly in light of mining applications. It should be noted that 1 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. VOLTAGE : Rise rate from 1 to 3 MV///S / Duration on the order of a few microseconds HIGH CURRENT : Maximum value about 100 kA Duration, lOtolOO^s / LOW CURRENT : Value, 100 to a few thousand amperes \ Duration, 0.001 to 0.01 s II I I Figure 11.1.— Schematic representation of lightning stroke discharge. 281 200 - 160 - LU 01 en LU cc o 0.02 0.1 0.2 12 5 10 20 50 100 STROKES WHERE CURRENT EXCEEDS ORDINATE, % Figure 11.2.— Distribution of crest currents in lightning strokes. shown in figure 11.4, the striking distance for a typical 20-kA stroke is about 30 m or 100 ft (27). The exposed surface portions of a mine electrical distribution system are particularly vulnerable to light- ning. For instance, overhead lines on wooden poles are often used to carry three-phase power from substations to surface facilities of underground mines. In surface mines, the overhead lines can extend close to the shovels and other electrically operated equipment in the pit. Since these poles can also carry the grounding conductor from the safety ground bed, it is doubly important that they be adequately protected from lightning. Three separate failure modes are recognized whereby an electrical distribution system can be damaged by lightning. If the impedance to ground of a supporting structure is high and the structure is struck by lightning, its potential may become elevated above ground to the point that a flashover occurs from the tower to a power conductor. This is known as backflashover. If a high- intensity stroke to earth occurs nearby, the resulting electric field may be great enough to induce excessive potentials on conductors; a power conductor can then arc over to the tower. In a shielding failure, lightning hits a power conductor directly, raising its potential to the point where arcing takes place between the conductor and the support structure (27). Induced potentials on grounding conductors, including equipment frames, can elevate the potential of the grounding system above that of earth and cause damage to protective devices or create a personnel hazard. With nearby strokes to earth, the magnitude of the induced voltage depends upon stroke current, distance, and height of the conductor, as illustrated in figure 11.5 (25). SWITCHING TRANSIENTS Switching operations account for the majority of all transient phenomena on mine power systems, and except for direct strokes by lightning, the most destructive power- system transients are initiated by this source. Every time a switch is operated, a transient can occur, and these transients can be classified into two types: normal and Figure 11.3.— Map showing average number of thunder- storm days per year in the United States. 240 - Positive Negative LU o < I- co a o \- 40 80 120 LIGHTING-CURRENT AMPLITUDE, kA Figure 11.4.— Striking distances for negative and positive strokes. DISTANCE A FROM STROKE CHANNEL, ft 1,000 600 400 200 150 100 75 200 400 600 400 800 1,200 800 1,600 CREST VOLTAGE, kV 800 1,000 /?=25ft /?-50ft /?= 100ft Figure 11.5.— Crest voltages induced on transmission lines by nearby strokes. 282 abnormal (33). Normal voltage and current transients are considered a characteristic of proper operation, and their maximum outcome is a transient voltage or current no greater (but usually less) than two times the peak steady- state voltage or current of the system. Transients exceed- ing this level are termed abnormal, and these are the result of incorrect component operation or design. Abnor- mal switching transients can occur in several ways, but all involve the release or exchange of trapped energy in the power-system inductance or capacitance (19, 33). Interest- ingly, a normal transient can be generated on a quiet power system, but if energy from this transient is stored and afterwards a second transient is created, the latter can be abnormal. Transient phenomena caused by switching can be divided into those created on ac systems and those associ- ated with dc. The ac switching events can be further subdivided into capacitive switching, current chopping, and prestrike. Capacitance Switching Because of the extensive use of cables, surge capaci- tors, and power-factor correction capacitors, capacitance plays a major role in most mine electrical systems. Abnor- mal transients can occur when ac current to capacitance is interrupted (6, 19). 2 The problem is that current leads voltage by nearly 90°. Capacitive currents are usually rather small so there is a good chance that current flow is stopped at a very early current zero. Thus, the load-side capacitance would be charged to the peak voltage of the system, and this energy becomes trapped. A simple circuit to demonstrate capacitance switching is shown in figure 11.6, where an ac source is feeding cable or line lumped capacitance through a circuit breaker. Source voltage and current waveforms are given in figure 11.7A; the point of current interruption is signified by the dashed vertical line. Figure 11. IB shows the voltage across the capacitance, and figure 11. 7C relates the poten- tial or recovery voltage across the circuit breaker contacts, which is the difference between the source and load sides. System inductance and stray capacitance are neglected in these waveforms to illustrate the effects of the trapped charge (19). Because early interruption is possible, the recovery voltage may attain two times the system peak when the contact gap is quite small. Hence the arc may reignite or restrike. With restrike, an inductance-capacitance circuit is formed with a resonant frequency of Source L © Breaker ' l J- Lumped capacitance Figure 11.6.— Simple circuit to illustrate capacitance- switching voltage transients. A Source -side voltage B Load -side voltage C Recovery voltage Figure 11.7.— Voltage and current waveforms before and after current interruption. The basic equation for the per-phase voltage after restrike is f„ = (11.1) T dl (11.2a) where f = resonant or natural frequency of system, Hz, L = system inductance, H, and C = system capacitance, F. 2 Personal communication from E. K. Stanek, West Virginia University, Aug. 1977. where v = V m cos(wt) = source voltage, V, V m = peak source voltage, V, v c = V c(0) + p \ idt = capacitance voltage, V, V c(o) = voltage across capacitance at restrike, V, and i = current through circuit breaker, A, or di L dt + h J idt = Vm cos (wt) " v °- (11 - 26) 283 lb investigate the current transient in the short time interval after restrike, and thus, cos(wt) ~ 1, L oi + ^ idt = v ~- v °- (11.2c) Solving this equation for the time-domain current, it is found that i(t) = (V m - V ) (£) ' sin ( Wo t). (11.3) Voltage Current 60-Hz current approaches Figure 11.8.— Voltage and current transient waveforms oc- curring with capacitance switching and restrike. Accordingly, the voltage across the capacitance would be 1 ft /C\ - 5 v c = V + -j (V m - V )(-) Bin(« t)dt or or v. = V - V ft = V Q+ ™ C)0 .5° ) sin(o> t)dt v c = V + V ^ C)0 J° [1 - cos (co t)]. (11.4) 1.0 mH 2,300 V Contactors 25-ft SHD cable / 100 kvar (50^F) h (\ 500-hp Figure 11.9.— Per-phase diagram of 4,160-V pump-motor cir- cuit. For the worst case, V = -V * o * m or V - V = 2V , T m o m> and the transient current is i(t) = 2V m Q a5 sin (« t), (11.5a) with transient voltage per phase for the load-side capaci- tance being v c = V m - 2V m cos ( Uo t). (11.56) Both these sinusoidal waveforms are illustrated in figure 11.8; the oscillations are at the system natural frequency (19). Equation 11.56 is of special interest as it indicates that the transient voltage can reach three times the peak system voltage. The decay shown in the waveforms is due to damping by system resistance. Obviously, system resis- tance is neglected in the foregoing derivation, but its effect is considered small compared with capacitance and induc- tance in this type of switching. To show the significance of the current transient, consider figure 11.9, which is a per-phase diagram of a three-phase motor circuit. Here, three-phase 4,160-V power is being fed through a circuit breaker (used as a motor starter) through 25 ft of SHD cable to a 1, 500-hp wound-rotor motor. A 300-kvar capacitor bank is used for power-factor correction and is placed at the load side of the breaker. Assume that the motor is drawing very little current, the breaker is opened, and interruption occurs at the first current zero. If the motor current is very small compared with the current drawn by the capacitor, the stage is set for restrike. Peak restrike current would be approximately i P = 2 Vp g»-= or Ip = 2(3,396)i 5 x 1Q- 10~ 3 = 1,520 A. Since the steady-state current through the capacitor is around 45 A, this transient-current peak is about 34 times the normal capacitance current flow, approaching a level expected for motor-starting inrush. Even though this is significant, the serious problem is not the current transient but the possible overvoltage produced. It can be seen from equation 11.56 that, after one restrike, the per-phase voltage can approach 3 V m , 3(3,396) V, or about 10 kV. This is substantial but still within the insulation capabilities of most 4,160-V systems. However, if the capacitor voltage is still around 3 V m when current becomes zero again, a current interruption will trap 3 V m across the capacitor. Now, the circuit breaker recovery voltage may approach 4 V m , causing a second restrike, which can then cause a capacitor voltage of - 5 V m . This situation can continue, developing even higher voltages. Such a multiple-restrike process is shown graph- ically in figure 11.10 (33). The voltages created can cause insulators to flashover. 284 The problem here is large capacitance as seen by the switching apparatus, with a very small impedance be- tween that capacitance and the load-side contact. Exam- ples of this situation could include the charging current to an unloaded cable or distribution line. Yet perhaps the most drastic instance would be interrupting current to a static capacitor bank, involving one line of a grounded bank or two lines of an ungrounded bank. Current Chopping Current chopping is the phenomenon of forcing cur- rent to zero before a natural current zero. This can occur when small currents, such as transformer magnetizing current, are interrupted by switching apparatus, as illus- trated graphically in figure 11.11 (19, 26). Current chop- ping can trap magnetic energy in the power-system seg- ment being interrupted, and the result can be severe transient voltages. In recent years, current-chopping transients have re- ceived more attention than any other type. The concern has been connected with the increased use of vacuum circuit breakers (VCB's), perhaps to the point where these transients have become associated with the use of VCB's. Through their high efficiency, these interruptors can eas- ily chop small current flow. However, it should be noted that all circuit breakers can cause current chopping, as well as certain fuses, especially the current-limiting types. To appreciate the magnitude of the created overvolt- ages, consider the simplified equivalent circuit of a power- system segment shown in figure 11.12. The components can be considered per phase for the three-phase system. If the circuit breaker chops the current at magnitude I, magnetic energy is stored at that instant in the system inductance (mainly the transformer) at a level (33) W L = LI 2 (11.6) where W L = energy stored, J, L = transformer magnetizing and cable induc- tance, H, and I = magnitude of chopped current, A. This stored energy is then transferred to the capacitance and charges the capacitance to W„ = CV 2 (11.7) where V = voltage produced across capacitance, V, and C = lumped capacitance of cable and transformer, F. Neglecting losses, both energies are equal, producing a voltage Source voltage Source voltage and current Capacitor voltage Recovery voltage Figure 11.10.— Voltage and current waveforms resulting from multiple restrikes after capacitance switching. Figure 11.11.— Graphic example of current chopping by breaker interruption. -• •- Transformer and cable Source r\; Circuit breaker contacts Ts' Figure 11.12.— Equivalent circuit of power-system segment with lumped components per phase, neglecting resistance. Vp-I = iz„ (11.8) ferral back and forth causes oscillations with a frequency equal to where V p = peak transient voltage, V, and Z = system surge (or characteristic) impedance, U. Following the capacitance charging, the energy is trans- ferred back to the system inductance. The effect of trans- Ry f = i[J__(A) ° 2tt LLC \2L/ (11.9) where f = oscillation frequency, Hz, and R = system resistance, 12. 285 Equation 11.9 shows the damping effect of system resistance, but for very small resistances (as are usually the case in power systems) the equation reduces to approximately f„ = (11.10) This is again the natural frequency of the power system. The preceding has considered the theoretically pure world and, of course, actual systems exhibit losses. Present-day dry-transformer hysteresis losses limit energy storage to about 40% (20). Therefore, the peak transient voltage is restricted to about 63% of equation 11.8 or V p = 0.631 ^ = 0.613IZ o (11.11) 0.63 IZ n - Transient recovery voltage Transient waveform TIME / s ■-- < Normal waveform if no interruption Figure 11.13.— Graphic example of chopping voltage tran- sients. An illustration of a voltage transient resulting from chop- ping (by trapping energy) is given in figure 11.13. Consider the power system shown in figure 11.14, where a switchhouse is connected to a load center by a short cable. A typical switching procedure in mining is to interrupt or switch out an unloaded load center, for example, during a maintenance shift. Even though there are no connected loads on the secondary, the transformer still draws a small amount of magnetizing current, about 0.03 to 0.05 pu of rated current for most mining applica- tions. Assume that the cable is so short that it has negligible capacitance and inductance as compared with the transformer. Thus, the equivalent circuit would be as shown in figure 11.12. A typical load center in underground coal mining is 750 kVA, and common unloaded transformer values for a 7,200-Vac system are C = 3,000 pF and L = 15 H. If 1.0 A is chopped by the circuit breaker, the per-phase crest voltage that can be produced on the system is 750 kVA 7,200 Vac No loads ' Circuit Short I/O SHD- cable ■GC connected breaker Fe edtf trough Figure 11.14.— Segment of mine power system. In addition to system losses, there are other phenom- ena that can reduce the transient voltage. After chopping, with only a small VCB contact separation, the dielectric in the circuit breaker is often unable to support the transient voltage, and an arc restrikes (19). Sometimes the circuit breaker will make successive attempts to clear the circuit in this manner, reaching progressively higher voltages as the contact gap increases, until isolation is finally estab- lished. But the maximum voltage attained may not be as great as when switching is clean; the stored energy is not allowed to accumulate, and the effect can reduce V p by as much as one-half (20). Nevertheless, the sequential re- striking and clearing of a circuit breaker can create dangerous overvoltages, and the voltage escalation can be much more rapid (19). If the cable length between the switchhouse and the load center is increased, the cable capacitance will be proportionally increased. This will reduce the surge im- pedance and therefore the possible severity of the chopping transient. Yet too large a capacitance can result in other transient phenomena, not only from restrikes as previ- ously covered, but also from prestrikes. Prestrike During the initial energization of system capacitance, an arc ignition or prestriking may occur across the contacts of a circuit breaker, prior to (final) mechanical closure (32). This phenomenon appears to be enhanced when the load- side capacitance is considerably in excess of the source-side capacitance. These events seem dependent upon the capaci- tive inrush current, and the magnitude of inrush current is mainly controlled by the amount of load-side capacitance. 286 Prestrike transients have usually been associated with vac- uum interrupters, and overvoltages have been found to reach substantial levels if the initial inrush current caused by the prestrike is momentarily stopped, then followed by a subse- quent prestrike or contact closure. Situations can involve energizing one line of grounded capacitance or two lines of ungrounded capacitance. Hypotheses that have been advanced to account for prestriking include the following (10): • Contact "whiskers." When exposed to a high elec- trostatic stress in a vacuum, extremely fine filaments can grow from the surface of metals. Such filaments could cause an arc ignition. • High electric-field strength. Stress created by a high electric-field strength could cause breakdown of the dielec- tric. • Contact bounce. Contact bounce is the deflection of the moving contact after impact with the fixed contact of the interrupter. This is caused primarily by a weak oper- ating mechanism, and the resulting separation could allow arc extinction. Whatever the source, if the load-side capacitance is uncharged and a prestrike occurs, the voltage on the immediate load side of the breaker will collapse to zero. This can cause voltage and current traveling waves to radiate on cable and lines. As will be discussed later, the waves can reflect and refract at discontinuities in the system characteristic impedance, and the typical frequen- cies of the resultant waveforms can be 60, 600, 6,000, 60,000 and 600,000 Hz, all superimposed. The combina- tion of the harmonics obviously has many current zeros, and when these zeros occur between the prestrike and the mechanical contact closing, the circuit breaker can easily interrupt the inrush-current flow. In the same manner as in capacitance switching, energy can be trapped on the capacitance being energized. Subsequent arc reignitions followed by interruptions can create an unusually rapid escalation of high overvoltage. Pflantz (32) found that prestrike transients can have crest voltages up to 7.0 pu of the system peak, with an oscillating frequency band that spans from 60 Hz into the megahertz region. The combination of capacitance switching, chopping, and prestrike creates many problems in the design of high-voltage distribution. If capacitance on the down- stream side of an interrupter is small, destructive voltage transients can occur from current chopping. When this load-side capacitance is large, overvoltages may result from capacitance switching or a prestrike event. Conse- quently, the placement of capacitance within a mine power system is critical, and this will be discussed later in this chapter. It is interesting to note from the foregoing discussion of ac switching transients, that the resulting overvoltage in each case can crest at about 7.0 pu of the system peak voltage or more. The oscillation frequency provides the key to distinguishing among the three types: the system natural frequency indicates capacitance switching or chopping (generally less than 10 kHz), and much higher frequencies indicate prestrike (often these have frequency components greater than 100 kHz). Direct Current Interruption The dc systems are also subject to abnormal voltage transients from circuit openings. Unlike ac, where inter- C Breaker Figure 11.15.— Circuit to demonstrate voltage transients in dc systems. Current through breaker System voltage Fault Contacts occurs part Voltage across breaker Figure 11.16.— Transient overvoltage resulting from current interruption on dc system. ruption generally occurs at a current zero, dc must be forced to zero. Large amounts of magnetic energy can be stored in the system inductance, and any sudden current decrease can result in an overvoltage. The energy might be transferred to capacitance as in ac circuits, but the energy can be dissipated in an arc. 3 The arc voltage tends to drive current in the direction opposite to the source, forcing the circuit current to zero. Figure 11.15 shows a simple dc circuit to illustrate the overvoltage phenomenon associated with arcing. The over- voltage could be caused by a switching operation to clear a short circuit. The relationship for voltage and current is L^ = V s - v A , (11.12) where L = system inductance, H, i = system current, A, di/dt = rate of current change, A/s, V s = source voltage, V, and v A = arc voltage, V. To stop current flow, the arc voltage must be greater than the source potential (fig. 11.16). As the source is constant, the peak transient voltage is proportional to the rate of current decrease, that is, the faster the decrease, the higher the voltage produced. 3 Personal communication from E. K. Stanek, West Virginia University, Aug. 1977. 287 The same reasoning can also be applied to interrupt- ing dc motor operation as well as other current flows. For instance, if a dc motor current is terminated, there exists a high rate of current change, which can develop a voltage across the system inductance between the source and the motor contactors, equal to Ldi/dt. A dramatic example of this would be dropping the contactors on a large trolley locomotive while drawing full-load current. General Switching Transients Switching transient problems are not restricted to main power components. In either ac or dc systems, small tripping and relay coils, when coupled with a very small capacitance, can exhibit high voltages even though they operate in low-voltage circuits (33). Silicon diodes and thyristors can create considerable overvoltage by current chopping, sufficient to destroy themselves (19). In fact, solid-state conversion equipment is constantly in a tran- sient state. OTHER TRANSIENT PHENOMENA Because of the existing protection methods in mine power systems, transients resulting from line-to-ground faults or accidental contact between two lines of differing voltage are minor, although destruction from localized heating may be severe. However, if the protective circuitry malfunctions, as might occur in the harsh mining envi- ronment, the problem can become critical. Local heating at the fault site can cause the conductors to melt. As this happens, the potential gradient across the conductors can be sufficient to strike an arc. The arc will extinguish and reignite, all things remaining equal, at each zero current crossing. The random fluctuation of arc impedance, as well as phenomena related to arc reignition, results in large transient overvoltages (33). Resonant conditions can result in transients exceed- ing 10 times the nominal line voltage (33). Since the power system contains inductance and capacitance that are nor- mally much greater than system resistance, resonance can occur at the natural frequency of the system. Resonance may result from the presence of line-frequency harmonics (harmonic overvoltage) or from the frequency components of other transients (dynamic overvoltage). Although dy- namic overvoltage is probably not a frequent form of transient voltage, the number of failures resulting from this mode are believed by some engineers to be significant. When transformers are involved, transient resonant con- ditions creating dynamic overvoltages have traditionally been given the term ferroresonance. Considering circuit component values, the occurrence of these transients in mine power systems should be rare, but a possible problem area could be a pole-mounted transformer powered through a cable feeder. Here, the series-resonant condition could easily be corrected by simply changing the cable length. Restriking ground faults, which are found primarily on ungrounded systems, can cause transients several times greater than the nominal line voltages. Some sys- tems are operated ungrounded because when a line-to- neutral fault occurs, little or no fault current flows and the system remains operational. Yet some current will exist because of system capacitance. To illustrate the effect of this capacitive current during faulting, figure 11.17A shows a normal ungrounded system and figure 11.17B Figure 11.17.— An ungrounded system, showing capacitive- current flow. A B VCG = ^CA Vbg "^ba > >Ib+Ic Figure 11.18.— An ungrounded system, with fault on phase provides the voltage and current phasors. Consider that a line-A-to-ground fault occurs as shown in figure 11.18A, where I F is a small fault current but sufficient to support an arc. The arc may extinguish itself at a current zero, but when this happens line-to-line voltages are trapped on lines B and C. The system voltage will thus be offset on these lines (fig. 11.18B). In the oscillatory return to steady state, it is possible to get restrikes of the fault current and further self interruption: in other words, a large voltage can build up. This is a major reason why systems should not be operated ungrounded, unless stern overvoltage precautions are heeded. Since coal mine power systems are required to use resistance grounding for portable and mobile equipment, restriking ground-fault transients are also rare, except when the protective circuitry malfunctions. A specific case would be an open grounding resistor, where uncleared faults could cause transients. TRAVELING WAVES The discussion of transients has considered only cir- cuits that have lumped resistance, capacitance, and induc- tance, except for circuits in prestrike conditions. In many circuits, transient behavior can be accurately predicted even though these parameters are distributed (19). But there are other power-system portions where circuit- element concentration will result in too large an approxi- mation. An outstanding example is a transmission or distribution line, be it overhead or cable (16). Fortunately, 288 these exhibit a certain resistance, inductance, and capac- itance per unit length of line, respectively, R, L, and C. For analysis, the line can be divided into small but finite elements, as shown in figure 11.19 (resistance and leakage are neglected) {16). The voltages created by transients can have a rise time in microseconds; in other words, they rise from zero to peak in that time. When very fast voltage changes occur, it is often best to analyze the system in terms of traveling waves, rather than by conventional methods (16). To show why this is so, consider figure 11.19, given that the switch has just closed. Conditions existing at the source end are not immediately observed at the load end because it takes time for the voltage-current conditions to pass through each LC segment. The movement of these conditions with time is known as traveling waves, and a characteristic feature of a circuit with distributed impedance is its ability to support these waves of voltage and current (19). From analysis of this circuit, where Ax is made very short, it can be shown that the voltage-current relation- ship for each incremental section is (19) dv = VF di or ■ w = z i, where v = voltage existing in incremental element, V, i = current existing in that element, A, L = total inductance of line, H, C = total capacitance of line, F, and Z = surge impedance or characteristic imped- ance of line, U. The propagation velocity, U, of voltage and current is (19) U = dx dt 1_ LC (11.14) For open lines, the propagation velocity is approximately the speed of light, 1,000 ft//xs; with solid-insulation cable, the speed is about 500 ft//xs (22). To illustrate the time behavior, suppose an overhead line with 400-ft surge impedance, as depicted in figure 11.20A, is hit with an 100-kV step function at time t = (16). At this instant, voltage and current exist only at the source end and nowhere else on the line (fig. 11.20.B). At t = 1.0 /is, the voltage-current conditions have propagated down the line for 1,000 ft, and between zero and 1,000 ft, the voltage is 100 kV with the current v z: 100,000 400 = 250 A. Beyond 1,000 ft, voltage and current are zero (fig. 11.200. At t = 4.0 ^s, the surge has moved 4,000 ft, voltage and current being 100 kV and 250 A to the left of that point and zero to the right (fig. 11.20D). If the line shown in figure 11.20A was of infinite length, the bundle of energy would theoretically travel forever. Because conductors have finite length, problems occur at the line end or at a discontinuity, which is a point where the surge impedance changes (16, 19). At either location, the AL| Al_2 AL3 AC, ± AC 2 ;L AC3J; AL5 AL 6 ^5± ALf LAx AC 4 =CAx Figure 11.19.— The distributed inductance and capacitance of two-wire line shown as incremental sections. So A 100 kV - 1,000 2,000 3,000 4,000 5,000 DISTANCE, ft (11.13) C D m W////MWW//, .. nn W/////////£W^WM F Figure 11.20.— Demonstration of traveling wave on overhead line. strict proportionality between the voltage wave and the associated current wave must be satisfied by natural adjust- ment. Reflected and refracted waves are the result. The reflected wave propagates back down the line and is super- imposed on the original or incident wave, whereas the refracted wave travels beyond the discontinuity. The re- flected and refracted amplitudes are such that the voltage- to-current proportionalities are preserved, as dictated by equation 11.13 and the surge impedances of the lines on which they travel (19). Expressed mathematically, + v, = v. + 1, = u (11.15a) (11.156) where vi = Z^, (11.15c) v 2 = -Z^, (11.15d) v 3 = Z 3 i 3 , (11.15e) 289 and v^iiiZi = conditions for incident wave, V, A, Q, v 2 ,i 2 ,Z 2 = conditions for reflected wave, V, A, Q, v 3 ,i 3 ,Z 3 = conditions for refracted wave, V, A, fl. The assumption is that energy is conserved. The reason for the minus sign indicated in equation 11.15d is that i 2 is traveling in a minus-x direction, and thus has a sign opposite to v 2 . By combining the above equations, expres- sions for the reflected and refracted voltage-wave magni- tudes in terms of the incident wave can be obtained: -z x + Z, 2Z„ lz 3 + zj v (11.16a) (11.166) Traveling-wave behavior under reflection and refraction can be demonstrated using the preceding example and terminating the overhead line by either an open circuit, short circuit, or a line with a different surge impedance. Figure 11.20.E illustrates the conditions if the line end is open-circuited. At the instant the incident waves reach the end, the current at that point must be zero, as equation 11.156 relates, i, + i 9 = or lo = -u but, by equation 11.15d, v 2 = -Zjia = Z^ = v x . In other words, the reflected current wave will have a magnitude of - i x , and the reflected voltage wave will have v x -. Moving to the left, both will superimpose on the incident waves. Therefore at t = 6.0 /xs, as shown in the figure, the voltage from 4,000 to 5,000 ft is 200 kV and the current is zero. Instead of an open circuit, consider that the line end is short-circuited. When the incident waves reach the short circuit, the voltage at that point must be zero, or from equation 11.15a, Vo = -v,, yet, by equation 11.15d, 2 " Z, " Z x ~ h In this case, the reflected waves superimposed on the incident waves will result in zero voltage, but at two times the incident current magnitude. Figure 11.20F shows the conditions at t = 60 ^s, with current to the right of 4,000 ft as 500 A. Now suppose that the overhead line is terminated by a cable with a 50-Q surge impedance, a typical value for feeder cable. The situations before and after the incident waves reach the junction are shown in figure 11.21 (19). Zi 4 "1 -k- i, 2, Z 2 Junction Before -^— X- z, Junction After Figure 11.21.— Incident waves being reflected and refracted at discontinuity. For the reflected wave (equation 11.16a), /50 - 400\ 450 100 kV = -78 kV and (equation 11.15 a 7 ), (-78,000) la = 400 = 195 A. For the refracted wave (equation 11.166), /2 (50)\ Vo = 450 -J 100 kV = 22 kV and (equation 11.15e), 22,000 A - x 3 " "SOT- = 445 A " Therefore, the surge that penetrates the cable has a marked reduction of voltage magnitude. This benefit is employed in power systems to protect equipment from surges coming down connected lines (19); instances in- clude machine trailing cables in surface mines and feeder cables in underground mines. It should be noted that for this case, the refracted waves propagate at 500 ft//xs, whereas the reflected waves are traveling at 1,000 ft//*s. The foregoing discussion can be expanded to any line termination, including those with more than two connec- tions. Within a specific line, reflections can occur at both ends, causing the traveling waves to move back and forth. In this theoretical model, resistance and leakage have been neglected so the energy of the traveling wave is maintained. In practical circuits, the presence of resis- tance and leakage means losses are incurred when current flows (19). These losses serve to attenuate the magnitude of the waves as they propagate. Even though the preceding demonstrations were simplified through use of a step function and by ignoring losses, the concepts can be applied to any waveform with a very fast voltage rise time. Two important points can be extracted from this brief outline of traveling waves (16). First, if a line is open ended, any terminal equipment on the line may experi- ence a potential up to two times higher than that of the traveling wave that produced it. Next, it is a common 290 practice to protect personnel working on exposed power- lines by placing protective grounds on each power conduc- tor. Consider what would happen if a traveling wave caused by a lightning stroke was on the line and a person was touching the line between the surge source and the protective ground. The person would be exposed to the incident voltage for the time it took the wave to travel from that location to the protective ground and back. It is then obvious that protective grounds should always be installed right at the worksite, or at least between probable surge locations and the workers. ELECTROMAGNETIC PHENOMENA A 2 charged conductors B 3d conductor presence Figure 11.22.— Electric field between conductors. It should now be apparent that electrical transients are in essence electromagnetic phenomena. According to Faraday's law, transients existing on a specific circuit can produce electrostatic and electromagnetic induced volt- ages on nearby or associated circuits. It was shown in chapter 3 that whenever two charged conductors are separated, an electric field and a potential difference exist between them. The charge is related to the potential difference by the proportionality of constant capacitance. The physical arrangement between the con- ductors is shown in figure 11.22A (19). As shown by figure 11.22.B, when another conductor is placed in the same space, it distorts the original electric field and there will be a charge separation on the third conductor surface. The conductor will assume a potential somewhere between the original two conductors. The relationship will also estab- lish capacitances among the three conductors (19). Suppose that conductors 1, 2, and 3 are, respectively, high-voltage power, grounding, and ground-check conduc- tors. If the potential difference between the power and grounding conductors suddenly changes, an electromotive force is induced in the ground-check conductor, which can cause current flow. This redistribution of charge could operate a relay, thereby nuisance-tripping a circuit breaker. If the third conductor is for some other control, monitoring, communication, or grounding circuit, the re- sult, beyond false relay operation, could be incorrect ma- chine operation, erroneous meter readings, communica- tions noise, or injury to personnel. TRANSIENT-INDUCED FAILURES Deteriorating electrical insulation affects the entire mine power system and may jeopardize its safe operation. Whether the dielectric is in a transformer winding, motor winding, portable cable, or rectifier, it is a critical factor in the safe, economical, and reliable operation of any mine power system. Disturbances that threaten to compromise the integrity of the power system must be eliminated at the source. This is why attention must be paid to the causes of electrical transients as well as to their elimina- tion. Consequently, it may be helpful to examine the effect of abnormal voltages on dielectrics. Each type of insulation or dielectric is designed for a safe maximum applied voltage and a transient overvolt- age. The transient overvoltage rating is given in terms of BIL, the basic impulse insulation level. The most common BIL measurement is the 1.2 x 50 wave test (fig. 11.23), where the voltage impressed across the dielectric reaches its peak in 1.2 us (22). Thus a dielectric with a 95-kV BIL rating can safely withstand a 1.2 x 50 pulse of 95 kV 100 - 1.2 20 30 40 TIME, fis Figure 11.23.— A 1.2 x 50 wave test used for BIL measure- ment. peak. The peak voltage in the 1.2 x 50 test is considered more severe than the transients actually found in mine power systems. Dielectric deterioration is created in large part by the rise time of the transient, as well as the crest magnitude. Furthermore, the greater the overvoltage pulse width (duration), the greater the probability for failure, since more energy is involved. These voltage anomalies break molecular bonds in the dielectric, which reduces its effec- tiveness. If the overvoltage contains sufficient energy, the dielectric can fail immediately; however, this is usually not the case. Instead, the insulation is progressively weakened until it finally fails, resulting in a line-to- ground or line-to-line fault. If the weakened insulation is in a portable cable or splice, a considerable personnel safety hazard arises since the insulation appears to be functional when in reality a lethal potential may exist on the cable surface. Although the deteriorating dielectric may not present a direct safety hazard, a complete failure can, because it may cause an explosion or arcing. In either case, the equipment faces serious problems in terms of repair, replacement costs, and downtime. Winding Response The physical construction of equipment may increase its susceptibility to transient failure; for instance, motor and transformer windings often fail at the end of a coil because of increased electrical gradient. Figure 11.24 can be used to explain why this can happen (19). It shows the distributed nature of the winding inductance and capaci- tance, where capacitance is assumed to be uniformly divided among the windings and to consist of capacitance to ground and to adjacent turns; resistance is neglected. The winding neutral may be grounded depending upon the position of the switch. 291 From an analysis of the circuit, the following general equation can be obtained: u at 2 " * dx 2 dt 2 d 4 V 1 d 2 V LdX 2 (11.17) where C = winding capacitance to ground, F/m, K = winding series capacitance, F/m, L = winding inductance, H/m, and V = voltage applied across the winding, V. By simplification, this can be used to solve for the response of the winding to a surge. Consider that surge waveform is a step function with amplitude V. If the neutral is grounded, at the instant the step function hits the wind- ing, the initial voltage distribution across the winding is V. = V sinh(ax) sinh(af) (11.18) Line end 'o^wmmmwm^ Neutral i >J L, lJ L^ U C >J L, J L, >J L> iJ L> .-J U i i Figure 11.24.— Equivalent circuit of multiturn winding show- ing distribution inductance and capacitance. and for the ungrounded neutral, cosh(ax) V = V cosh(af) ' (11.19) where a = VI I = winding length, m, x = distance between neutral and a point on the winding, m, and V x = voltage at that distance x, m. For a specific a, the initial voltage distribution across the winding in response to the surge can be plotted as shown in figure 11.25 (19). This figure relates that as a increases, the distribution becomes nonuniform. When a = 10, 60% of the voltage is initially impressed across the first 10% of the winding, with 75% across the first 20%. Therefore, under surge conditions, very high stresses can occur on the first few turns, and if precautions are not taken, trans- formers and motors can fail by breakdown of the turn- to-turn insulation in this region. Coupling Through Transformers When a fast rise-time voltage surge hits a trans- former, as just shown, the parameter of concern is the winding capacitance. Series capacitance and capacitance to ground exist not only in the primary, but in the secondary winding as well. Figure 11.26A approximates this situation for a two-winding transformer, and the equivalent circuit shown in figure 11.26B forms a crude capacitive voltage divider (19). When a change of voltage is applied to the primary, the voltage divides inversely with capacitance (15, 35): V, = Cx Ci + C. V. (11.20) where V p = magnitude of transient voltage impressed on primary, V, V 3 = magnitude of transient voltage transmitted to secondary, V, Cj = primary-to-secondary winding capacitance, F, and C 2 = secondary-to-ground winding capacitance, F 100 Q 80 ^ D O 60 LC O o 40 H 20 ^(=0 ,3=5 a=iN a=i0 1 0.8 0.6 0.4 0.2 Grounded neutral too 80 60 40 20 a = \ ^ a=i \ aT^ a-\0 s<£=5 1 0.8 0.6 0.4 0.2 Ungrounded neutral Figure 11.25.— Initial voltage distribution across uniform winding from step function. Secondory 'mm. WW/r Core- B Figure 11.26.— Capacitive coupling of transient voltage through two-winding transformer. Obviously, the voltage transmitted from the primary to the secondary is not tied to the transformer turns ratio. Typical values are 35% to 40% of the impressed primary voltage (15). If the transformer is a step down, for example, from distribution to utilization, the result can be low-side per-unit voltage levels much greater than those that existed on the high side. 292 TRANSIENT PROTECTION Ideally, the elimination of transient-voltage problems begins with an excellent power-system design applying time-tested principles. The basic goal is to avoid the situations covered in this chapter. However, even this ideal situation can address only normal conditions; unpredict- able abnormal conditions can still arise, producing de- structive transient overvoltages (19). To address this prob- lem, additional overvoltage control must be placed in the mine power system through use of such protective devices as surge arresters, surge capacitors, shielding, and circuit arrangements. The role of these protective schemes is to ensure that equipment dielectric strengths are not exceeded, so that if a transient attempts to raise the voltage above an insula- tion withstand level, the protective device will exert a clamp or restraint to maintain the voltage within accept- able limits (19). Effective transient voltage suppression is basically the dissipation of transient energy. Surge Arresters The simplest form of overvoltage device is the spark gap, which is essentially two conductors separated by air, as in the tips of two rods where one side is connected to a powerline, the other to the neutral or earth. The spacing between the conductors establishes a specific dielectric strength so that voltage above that level will cause the gap to spark over. The spark gap has no effect on normal system operation, but its main disadvantage is that once an arc occurs, a fault is created on the system, which remains until the gap is deionized. Often, the attendant current flow or follow current can only be interrupted by a circuit breaker or a fuse. Surge arresters, formerly called lightning arresters, use this spark-gap principle to clip the peak of a voltage and divert the excess current to ground. They also contain a device to interrupt the follow current (31, 34). There are two common surge-arrester types, expulsion and valve; they differ in the scheme they use for interruption. The expulsion surge arrester extinguishes the arc in a manner similar to an expulsion fuse (see chapter 9); that is, an overvoltage establishes an arc across the spark gap and also across a gas-evolving material (usually organic). Ignition of the material causes the expulsion of gas, which blows out the arc. The operation has three disadvantages. First, some gas-producing material is destroyed during each operation, and only a limited number of interrup- tions are available. Second, because of the gaseous dis- charge, care must be taken in placement and installations are restricted to outdoors. Lastly, the arrester has an assigned current-interrupting rating and cannot be used on circuits that have a greater available fault current than this rating. In valve surge arresters, the spark gap is in series with a nonlinear resistor or valve block, as shown in figure 11.27. A property of the block, which is commonly made of silicon carbide, is that the resistance diminishes sharply as the voltage across it increases. Tb increase gap effi- ciency, a number of short gaps are used because these spark over more consistently and in less time than one long gap. On an overvoltage, the gaps spark over and the valve block operates in its highest conductance to pass the surge current safely to ground. After the surge voltage diminishes, the block changes to a low-conductance mode Line Spark gap Valve block — Ground Figure 11.27.— Basic valve surge arrester. to limit the follow current, such that the gaps can provide an interruption. The valve surge arrester has none of the disadvantages of expulsion types, and it is used exten- sively for equipment protection, especially on distribution systems. The balance of this section will thus cover only the valve units. Four important parameters are connected with the proper application of surge arresters (31 , 34): 1. Voltage Rating. The power frequency sparkover voltage is the lowest rms 60-Hz ac voltage across the arrester at which it will perform the operating cycle. This level is 1.5 times the arrester voltage rating for arresters rated at 60 kV and below. 2. Sparkover Voltage. The highest crest voltage at which arcs will form across the spark-gap electrodes, initiating the operating cycle. 3. Discharge Current. The current through the arrester created by the overvoltage immediately after sparkover. 4. IR Discharge Voltage. The voltage formed across the arrester during the discharge of surge current. Ideally, gap sparkover should occur on any dangerous system overvoltage and ignore all minor and harmless transients. Proper sparkover requires high-speed response to fast rise-time wave fronts (as in lightning surges) and consistent response to slower rates of voltage rise (as in system-generated surges). Both requirements are satisfied by electrical grading of the spark-gap structure, which consists of shunting each gap with a high resistance. Figure 11.28 shows the technique in simplified form (31). After sparkover, the IR discharge voltage occurs, being equal to the product of the discharge current (I) and the arrester discharge-path resistance (R). As discharge cur- rent may be very large, the discharge voltage may equal or exceed the sparkover voltage. Thus, protected equipment is exposed to both the sparkover and IR discharge voltages, and the system insulation withstand ability must be safely above both. To establish an IR discharge voltage, it is important to recognize the magnitude of possible discharge currents. A surge arrester is likely to be exposed to a wide discharge range, but experience from field measurements has shown that discharge currents typically range from 1,000 to 2,000 A, that 5.0% exceed 9,000 A, and that only 1.0% exceed 20,000 A (31). Even though 20,000 A is rare, it is often used as a worst case to estimate the discharge voltage. 293 Gap electrodes Gap electrodes Nonlinear grading resistor Nonlinear grading resistor Gap electrodes Nonlinear grading resistor Valve block Figure 11.28.— Surge arrester with nonlinear resistance grading to equalize each gap structure. There are three classes of valve-type arresters avail- able (34): 1. Distribution-class arresters have the lowest cost and are satisfactory for general equipment protection pur- poses. Typical voltage ratings range from 1,000 V to 18 kV, and these arresters can withstand a 65-kA current surge. (Surge arresters are generally used on high-voltage sys- tems in mining, but these devices are available with ratings as low as 50 V.) When protecting rotating machin- ery and dry-type transformers, the arresters must be the low-sparkover rotating-machinery (RM) type because of low BTL's. 2. Intermediate-class arresters have lower sparkover and IR discharge voltages than do distribution-class ar- resters. Available voltage ratings are from 3,000 V to 73 kV, and the arresters can typically withstand a 65-kA surge current. The cost is about five times that for a distribution arrester. 3. Station-class arresters can handle discharge cur- rents up to 100,000 A and are available for almost any distribution or transmission voltage application. They are considered to provide the best possible protection, but cost about twice as much as intermediate-class arresters. Both station-class and intermediate-class arresters have pressure-relief systems; if stressed by a surge beyond their capability, the internal pressure is vented to prevent housing rupture. Surge Arrester Applications The factors to consider when selecting an arrester class are the degree of transient exposure and the impor- tance of the equipment being protected. This is basically an economic question, but in general, intermediate or station arresters are justified for surface substations, with distribution-class arresters being suitable for distribution and utilization equipment protection. An arrester voltage rating of a certain class has associated sparkover and IR discharge voltages. System voltage, as well as the method of system grounding, affects the voltage that the arrester is exposed to and therefore the selection of the arrester voltage rating. Consequently, once an arrester voltage rating is set, the system insula- tion withstand ability must be coordinated with it. This is most often related as a BIL for the equipment being protected. The voltage-rating selection is affected by the system grounding categories: effectively grounded or noneffectively grounded. The coefficient of grounding can be employed to find which category is being used. The coefficient of grounding is defined as the percent ratio of the highest rms line-to-ground voltage existing during a line- to-ground fault to the nominal line-to-line voltage (31). If the ratio does not exceed 80%, the system is termed effectively grounded. Solidly grounded and typical low- resistance-grounded systems are in this class (the neutral potential remains rather constant during line-to-neutral faults). However for high-resistance and ungrounded sys- tems, the occurrence of a line-to-neutral fault can shift the neutral to near the faulted line, with the potential to ground of the other two lines approaching line-to-line system voltage. The coefficient of grounding here can be from 80% to 100%. These systems are termed noneffec- tively grounded. Resistance-grounded mine power systems for portable and mobile equipment are included in this category. In either grounding case, the arrester voltage rating should be above the possible exposed crest voltage; if not, a disruptive discharge might occur. Therefore, on effectively grounded systems, the ar- rester is sized to maximum expected line-to-ground volt- age, whereas maximum line-to-line voltage is used for noneffectively grounded systems (2). To allow for the expected increase due to voltage-regulation compensation, the arrester voltage rating should be 5% to 10% above these values. Tables 11.1 and 11.2 list the recommended sizing for resistance-grounded mine power systems (3, 34). The first table refers to station and intermediate arresters, the second to the low-sparkover distribution class. Note that the transformer BIL's specified are according to ANSI C57. 12. 00-1973 (1) and may be too low for some mining applications (see chapters 12 and 13). The equipment protection from voltage surges can be verified for any arrester selected. Full coordination re- quires checking the arrester performance over a full time range for an expected surge (1). However the following quick guidelines will ensure safe protection (34). • The insulation BIL ratings of equipment should be at least 20% greater than the arrester sparkover voltage. • The BIL rating should be above the IR discharge voltage of the arrester. As mentioned earlier, a discharge current of 20,000 A to establish IR discharge voltage may be used as an approx- imation of worst case conditions. Tables 11.1 and 11.2 also provide typical sparkover and IR discharge voltages for arresters used in mining service (34). However, manufac- turer catalog values should be consulted for actual appli- cations, as differences exist among products. Additional details of surge-arrester equipment protection are pro- vided in chapters 12 and 13. Another important point about maximum exposed surge voltage concerns the arrester connections to line and 294 Table 11.1.— Recommended station and intermediate surge arresters for resistance-grounded mine power systems to protect oil-immersed transformers Insulation class. kV System voltage, 1 V Transformer BIL, 2 kV Arrester rating, Front-of-wave sparkover, 3 kV IR discharge voltage 3 (20,000 A) kV Maximum 3-phase line-to-line Nominal Maximum Power Distribution l\V SC IC SC IC voltage, ' 2,400 2,540 60 45 3 13 12 7.8 12 3,000 4,160 4,400 75 60 6 19 21 15.5 20 6,000 4,800 5,080 75 60 6 19 21 15.5 20 6,000 7,200 7,620 95 75 9 30 31 23 27 9,000 12,470 13,200 110 95 15 51 51 39 45 15,000 13,200 13,970 110 95 15 51 51 39 45 15,000 13,800 14,520 110 95 15 51 51 39 45 15,000 14,400 15,240 110 95 18 NA 61 NA 54 18,000 14,400 15,240 110 95 21 70 NA 54 NA 21 ,000 2.4. 5.0 8.7. 15., NA Not available. 1 Maximum system voltages are from ANSI C84.1-1970 (4). 2 BIL ratings are from ANSI C57. 12.00-1 973 for oil-immersed transformers (1). 3 SC, station class; IC, intermediate class. Table 1 1 .2.— Recommended distribution-class, RM-type, surge arresters for resistance-grounded mine power systems to protect rotating machinery and dry-type transformers Insulation class, kV System voltage, 1 V Nominal Maximum Transformer BIL, 2 kV Arrester rating, kV Front-of-wave sparkover, kV IR discharge voltage (5,000 A), kV Maximum 3-phase line-to-line voltage, V 2.4. 5.0. 8.7. 15.. 2,400 4,160 4,800 7,200 12,470 13,200 13,800 14,400 2,540 4,400 5,080 7,620 13,200 13,970 14,520 15,240 20 25 25 35 50 50 50 50 3 4.5 6 3 7.5 15 15 15 3 15 13 17 22 24 44 44 44 44 10 15 20 25 50 50 50 50 3,000 4,500 6,000 7,500 15,000 15,000 15,000 15,000 1 Maximum system voltages are from ANSI C84.1-1970 (4). 2 BIL ratings are from ANSI C57. 12.00-1 973 for oil-immersed transformers (7) 3 Arresters may occasionally be exposed to voltage above their rating. to ground. Conductors extend from each ungrounded power conductor to an arrester and from the arrester to ground so that in resistance-grounded three-phase sys- tems, a minimum of three arresters is needed. For station- ary equipment on the surface, an arrester ground bed, such as a substation system ground bed, serves as the grounding medium (it must be low resistance); on portable equipment, the frame would be the ground. As the connec- tions exhibit inductance, they should be of No. 6 AWG solid-copper conductor or larger and as short as possible because the inductance of too long a conductor can render an arrester ineffective (19). Sharp bends should also be avoided, since a bend substantially increases the effective inductance. The inductance of arrester connections that adhere to these requirements is estimated at 0.4 /xH/ft, with the voltage drop produced during a surge estimated at 1.6 kV/ft, using a current wave front of 4,000 AJ/is (31). This voltage must be added to sparkover and IR discharge voltages to assess protective margins. As a general rule, all arresters should be located as close as possible to the equipment they are to protect. Ideally, they should be across the protected equipment terminals, the connections for three-phase systems being a wye configuration with the common arrester connection grounded. First of all, this location minimizes the possi- bility of a destructive surge entering the circuit between the protecting and protected devices (19). Second, close proximity also reduces the change of surge-voltage ampli- fication through refraction and reflection of a traveling wave. For instance, consider figure 11.29, where a tran- sient voltage surge is traveling along a power conductor toward surge-arrester-protected equipment (22). The ar- rester is located a distance, d, from the line end. As the wave passes the arrester location, the arrester sparks over at its protective level, but lets a traveling wave with a crest equal to its sparkover value through. The voltage wave reaching the equipment terminals is amplified by the addition of the incident and reflected wave, with the resultant magnitude depending upon the line-end surge impedance. Consequently, surge-arrester locations other than directly across the equipment terminals can lead to higher surge voltages at the protected apparatus than the arrester sparkover voltage. The terminal voltage rise will be aggravated by a greater separation distance, d. Note 30 kV Surge arrester Line termination Figure 11.29.— Surge approaching surge-arrester-protected equipment. 295 that with a wavefront rise time no greater than 0.5 fis, a maximum distance of 25 ft is perhaps allowable, but shorter distances always afford greater protection (28). Capacitors and System Capacitance Surge capacitors, also termed RM capacitors, are special units with low internal inductance that are used extensively to protect rotating machinery and dry- insulated transformers (28). This equipment is very sus- ceptible to line-end turn-to-turn failures caused by fast rise-time wavefronts, and the faster the rise time, the greater the probability for damage. Connected across equipment terminals in grounded wye, as shown in figure 11.30, the capacitors serve to limit the rate of rise of the transient voltage. Simply, the capacitor has to be charged before the overvoltage can be impressed on the system dielectric. Transient rise time is then largely determined by the charging rate. Coupled with the system inductance, the limiting criterion is that at least 10 (is is needed before the crest value of the protected-equipment nameplate voltage is reached (22). Low internal inductance of the capacitor is important because the presence of series inductance in the capacitor circuit deteriorates the wave- sloping action. Accordingly, the capacitors shown in table 11.3 have been standardized for this kind of protection (22). In combination with the recommended low-sparkover distribution-class surge arresters (table 11.2), the crest voltage of transients is considered restricted to harmless values for the utilization equipment (31). Table 1 1 .3.— Commonly used surge capacitors for limiting voltage rate of rise on rotating machinery and dry-insulated transformers Rated equipment Capacitance, voltage, V nF 650 or less 1.0 2,400 to 6,900 .5 11,500 or higher .25 Surge-Impedance Reduction Surge capacitors can also be used to control transient overvoltages by reducing the system's characteristic im- pedance. An increase of system capacitance, as exhibited by equation 11.8, can lower the surge impedance and therefore the possible peak transient voltage resulting from current chopping. A fine line exists here, because too much capacitance on the load side of switching apparatus can cause capacitive-switching or prestrike transients. For instance, prestrike events are dependent upon the capaci- tive inrush current during contact closure, and the mag- nitude of inrush current is controlled by the amount of load-side capacitance (30). With capacitive switching, the transient overvoltage is created by the ability of the load-side capacitance to hold a charge, thus causing a recovery voltage across the breaker so restriking occurs. System capacitance is, therefore, a critical factor in tran- sient protection. Hence, the problem of surge-impedance reduction mainly concerns how much capacitance is necessary to limit safely the transient overvoltages caused by current chopping. Actually, any magnitude below the minimum insulation withstand level or, if used, the surge arrester sparkover can be considered safe. Yet perhaps the most conservative approach would be to limit any chopping ^nmnmnnn^n Series inductance Standard arrester Low sparkover arrester Machine frame Figure 11.30.— Typical surge protection of rotating machinery and dry-insulated transformers. event to two times the peak system voltage. The most critical portion of the mine power system would be where capacitance is minimum, such as the case of a switchhouse connected to a power center, as illustrated in figure 11.14, or rotating machinery. It can be inferred that current chopping is the result of VCB operation, but other switching-apparatus types, including current-limiting fuses, could be chopping sources. Equation 11.8 can be used to select a value of capac- itance that will include the capacitance inherent in the distribution system, that is, the equivalent line-to-neutral capacitance of all devices on the load side of the switching apparatus. Considering a three-phase transformer as the load, the procedure can be as follows: 4 1. Determine the allowable peak voltage, V ; 2. Find the exciting current and assume interruption is at peak level, I m ; 3. Determine the transformer exciting inductance, L m ; 4. Assume all transient energy is absorbed by the capacitance, C; then 5. The necessary per-phase capacitance, in farads, referred to the transformer primary circuit is C = V 2 (11.21) However, a more useful form of this equation can be found if (24) 1. The allowable peak voltage is limited to two times the peak system voltage; 2. The inductance of the transformer and the power system is no greater than 20%; and 3. The peak exciting current is expressed in terms of the rated capacity and rms voltage of the transformer. With these parameters, C = 10 S (60) V 2 f (11.22) where C S V„ and f = = necessary capacitance, (i¥, = per-phase transformer capacity, VA = line-to-line voltage rating of transformer pri- mary, V (if it is desired to connect the capac- itance across the secondary, secondary line- to-line voltage is used), power-system frequency, Hz. 4 Personal communication from E. K. Stanek, West Virginia University, Aug. 1977. 296 The value resulting from equation 11.21 or 11.22 is total system capacitance per phase. If the level is above that supplied by the system, additional capacitance might need to be added. For three-phase systems, the common surge capacitor connection is shown here in a wye configuration with the center connected to ground. The typical location is directly across the protected equipment terminals, as in wave- sloping applications (fig. 11.30). Another popular location has been at the switching-apparatus load terminals. The philosophy here is that the interrupter sees the increased capacitance directly, and thus, chopping transients are limited more effectively; in other words, transients are best eliminated at their source. There is a general feeling that in this way the entire system downstream from the capacitor location would receive protection. However, surge capacitors have an extremely low internal series impedance; therefore, at the interrupter load terminals, they can be a significant source of capacitive inrush current as well as having the ability to exchange transient energy effectively to and from the system. Consequently, the best location for applying surge capacitors is at the protected-equipment terminals. Two advantages are gained through the ground con- nection: Transient energy is shunted to ground, and com- pared with delta connections, a lesser voltage rating is necessary. The capacitor working-voltage rating (WVDC) should be at least three times the exposed rms system voltage, usually line-to-neutral for wye grounded and line-to-line for delta connections. 5 However, for the same reasons given for surge arresters, it is perhaps best to rate surge capacitors in resistance-grounded systems to three times the line-to-line voltage. Buss (11) and Morley (30) have performed extensive tests on actual underground coal mining equipment to determine the severity of transients existing on these distribution systems. For the most part, these tests in- volved recording staged transients on unloaded and loaded systems by chopping (tripping the interrupter) and pre- strikes (engaging the interrupter). In every instance, the system segment was similar to figure 11.31 and consisted of a VCB-equipped switchhouse with various lengths of cable supplying a power center. Various switchhouses and power centers were used, all typical of actual mine instal- lations. Beyond the principal goal of uncovering the na- ture of any transients, the overall objective was to see if surge capacitors were necessary to limit chopping voltage transients and what effect they have on prestrike events. Buss (11) backed up the actual equipment testing with computer simulations but primarily addressed chopping events. Morley (30) extended the research to cover pre- strike activity. In terms of transient protection by surge capacitors, the results of both research programs were practically identical and apply directly to underground mine power systems or similar surface arrangements. Specifically, it was found that both grounded wye or ungrounded wye surge capacitors do significantly reduce chopping tran- sients when applied as given in table 11.3. With capacitors removed and practical cable lengths held to a minimum, the chopping transient voltages reached substantial levels (about 4.5 pu system nominal voltage for a common event) but were still below the sparkover level of correctly applied low-sparkover distribution-class surge arresters. Thus Mine distribution system u r^> Measurement point Feeder cable « «- SWITCHHOUSE Feedthrough Measurement point 4 rni — uzt iH Wye grounded 5 Personal communication from E. K. Stanek, West Virginia University, Aug. 1977. LOAD CENTER KEY C Surge capacitor -usually 0.25,6 9 ° 2 .0993 0.0850 0.0401 — g * 1 .1096 .0938 .0456 0.0345 < 250 .1598 .1368 .0776 0588 350 .1846 .1580 .0844 .0634 500 .2150 .1840 .0920 .0685 750 .2410 .2062 .0981 .0743 1,000.. .2740 .2345 .1118 .0789 NOTE.- — Dashes indicate cable is not made 0.040 .032 .024 .016 .008 ■P i^g 6C >°- "ZL§2°— — <2z: 1,200 1,800 Table 11.5.— Typical capacitances per phase of power-system components 100 200 300 400 500 MACHINE SIZE, hp Figure 11.32.— Capacitance for 2,300- V induction motors; for motors up to 7,200 V, value will not vary more than ± 15% of above. Power-system component Capacitance, /iF Nonshielded cable 0.02 to 0.05 per 1,000 ft; typically 0.03 per 1,000 ft. Nonshielded cable in conduit 0.02 to 0.06 per 1,000 ft; typically 0.04 per 1 ,000 ft. Overhead open-wire lines Negligible for line lengths used in typical mine distribution. Surge capacitors, by insulation class: 600 V 1.0. 2,000 V 1.0. 5 kV 0.5. 8 kV 0.25. 15 kV 0.25. 23 kV 0.25. Synchronous and induction motors, by insulation class: 600 V 0.032. 2,000 V to 23 kV See figures 1 1 .32 and 1 1 .33. at the high side of every distribution transformer. Such extensive use can pose problems in the high-resistance grounding system of a mine. As stated in chapter 7, the ground-fault current that is limited by the grounding resistor should not be less than the system capacitive charging current. This requirement restricts the amount of capacitance that can be placed on a system of a specific size. In other words, the size of the mine distribution system is limited by the total connected capacitance (11). Some relief is gained through power- transformer configurations, because the zero-sequence sys- tem is isolated in each voltage level. Nevertheless if there is excessive capacitance in distribution, for example, the capacitance can discharge during a line-to-neutral fault and feed the fault with capacitive ground current in excess of current limit. Tables 11.4 and 11.5 and figures 11.32 and 11.33 provide typical system capacitances to assist in estimating a per-phase system capacitance (38). The following can be used to compute system charging current per phase: o u o U Ql O u_ o rr < x o 0.020 .016 .012 .008 .004 p, jb0> q aH 9U 2800.. ^2g 100 200 300 400 MACHINE SIZE, hp 500 I =^ co X„ (11.23) Figure 11.33.— Capacitance for 2,300- V synchronous motors; for motors up to 7,200 V, value will not vary more than ±15% of above. where I co = per-phase system charging current, A, V ln = line-to-neutral system voltage, V, X co = per-phase capacitive reactance, Q, = l/27rfC , and C = lumped charging capacitance per phase, F. This current can be compared with ground-current limits to assess the effect of adding surge capacitors. The values in the table can also be compared with the results from equations 11.21 or 11.22 to obtain greater understanding of chopping events. 298 Other Suppression Devices The discussion thus far has been mainly concerned with surge suppression on high-voltage distribution. Transient-related failures can also be severe on mine systems below 1,000 V, an outstanding example being the destruction of solid-state elements in such equipment as ground-check monitors, communications apparatus, and power supplies. These transistors, integrated circuits, and thyristors are the devices most sensitive to transient overvoltages, and problems can occur even through induc- tion from transients occurring on power conductors to a neutral or communications circuit. Two suppressor types already presented also offer effective protection for low- voltage systems: valve-type surge arresters and surge capacitors (also termed snubbers). In this section, several other common protection devices designed principally for but not restricted to low-voltage applications will be discussed. These can generally be divided into two classes: transient suppressors and circuit-shorting devices. Transient Suppressors Transient suppressors, also called constant-voltage devices, are basically nonlinear resistances placed across the circuit to be protected (13, 18). They act directly to clamp or limit voltage rise much like the valve block of a surge arrester, but no series spark gap is used. Typical transient suppressors are power zeners and metal oxide varisters. Powers zeners are primarily for dc protection, working on the zener-regulator principle covered in chapter 5. They can be used in ac circuits when two devices are placed back to back or anode to anode to give bidirectional operations. Zeners have the capability to clamp transient voltage rise to a level largely independent of the impedance or voltage- current characteristic of the transient. The response is extremely fast, and the clamping action is very firm. Although they are effective for low-energy transients, many events common in industrial power systems can readily destroy all but the most expensive high-energy zeners, which are called avalanche diodes (18). Metal oxide varisters (MOV's) are ceramic suppressors that use zinc-oxide-based materials such as a zinc oxide and bismuth oxide ceramic body (18). Their construction provides a voltage-dependent, very nonlinear resistance with symmetrical conducting properties. The bidirectional breakdown allows their use on either ac or dc circuits. The response is similar to that of back-to-back zeners, but the clamping action is softer than with zeners, yet faster than with valve-type surge arresters (13). As with the silicon carbide valve block in surge arresters, transient energy is dissipated throughout the entire volume of material, mak- ing the MOV a very rugged suppressor. MOV's can be used for a wide range of applications from power-supply voltage regulation to power-system transient suppression, including high-voltage systems. Rated voltages extend from 22 V to the thousands (18). A very popular application in mining is to suppress transients across thyris- tors in solid-state motor starters, where an MOV is mounted across each thyristor. A projected employment at this time is for direct replacement of valve surge arresters in ac distri- bution and dc trolley-line applications. Circuit-Shorting Devices A circuit-shorting device or "crowbar" can be de- scribed as a device such as a spark gap, gas-discharge tube or thyristor that senses a high voltage and throws a short circuit or low resistance across the line (18). The low resistance is not removed until the current through the crowbar is brought to a low level; hence, these devices need to have the power removed momentarily before resetting can occur (13). To facilitate this requirement, crowbars are often used in conjunction with circuit breakers, where the device could be connected in series with a shunt-tripping element, and the combination located across the line on the load side of the breaker. Sensing an overvoltage, the current through the crowbar trips the circuit breaker; normal circuit operation is not resumed automatically. These quick-acting devices are available for ac or dc systems, typically 250 V and below. Faraday Shields Faraday shields are used to protect the low side of a transformer against surge voltages. The shield is a turn of nonmagnetic metal sheeting placed between the primary and secondary windings; it is insulated from all windings and connected solidly to ground (15). This location effec- tively destroys interwinding capacitance and substan- tially reduces the transfer of surge conditions. The shield has the further advantage of practically eliminating inter- winding faults. Other advantages of these shields are covered in chapters 12 and 14. Circuit Arrangements As induced voltages on low-voltage circuits from tran- sients on high-voltage or other power-system conductors can be a serious problem, there are four recommendations to reduce the possibility of induction. 6 1. Separate low- voltage circuits from high-voltage systems by a large distance. 2. Use shielded conductors on the low- voltage circuit or maintain shielding between circuits serving different purposes. 3. Twist low- voltage conductors (this effectively can- cels many induced voltages). 4. Place conductors on the high-voltage circuit close together. Protection of Overhead Lines Exposed overhead lines are very susceptible to direct contact from a lightning stroke, which obviously produces severe transient overvoltages. An excellent means of pro- tecting distribution lines from such occurrences involves the use of overhead ground wires or static wires (25, 37). One or two ground wires are strategically situated above and between the power conductors to provide a shielding effect, as illustrated in figure 11.34 (27). Here, line a bisects a line drawn from the ground wire to the outer 6 Personal communication from E. K. Stanek, West Virginia University, Aug. 1977. 299 power conductor; lightning strokes in area 1 are inter- cepted by the static wire. Line b is equidistant between the outer power conductor and the earth's surface, so that lightning streamers in area 2 will discharge to the earth. The position of line b in figure 11.34 varies with the relative height of the supporting structure, and line c is described by an arc whose radius depends upon the size and construction of the supporting structure. A stroke in area 4 will be borne by the supporting structure if it is metallic. Area 3 is the danger region (and may include area 4) where lightning flashes will strike the power conductor (27). Extensive tests and observations have been conducted to determine the optimum angle 9, which is measured from the vertical up to the line joining the power conductor and the static wire. The angle is dependent upon the height of the supporting structure, as shown in table 11.6 (27), but field results have shown that a good average for this angle is 30° (37). (Actually, 45° provides satisfactory performance if the line is situated on a level surface, but if on a hillside, the angle should be decreased from 45° by the slope angle of the hill.) Importantly, the design of static protection is practically independent of the system voltage. Figure 11.35 illustrates two approaches for shield- ing overhead lines by static wires when the support structure is wooden (37). Table 1 1 .6.— Protective angle versus structure height Tower height, ft Angle (Q), deg 50 45 100 25-30 150 10-12 In addition to intercepting direct strokes, the static wires perform another function: when static wires are struck by lightning, the resulting surge current is imme- diately halved since the impulse travels in both directions from the point of contact. The magnitude of the induced voltage on the static wire is thereby halved also, leading to a similar reduction in the induced potentials seen by the power conductors. Adequate grounding must be used in conjunction with the overhead static wires. The system generally consists of downleads extending from the static wire to grounding electrodes so that induced voltages are minimized by conducting stroke currents to earth as quickly as possible. This also helps to reduce the surge impedance of the overhead static-wire system. A grounding electrode may be provided by driving several ground rods at the base of each supporting structure and connecting them to the static wires via the downlead (23). An alternative method is based upon the use of a counterpoise, consisting of one or more buried horizontal conductors located at the base of each tower (5). The counterpoise may be a single continu- ous conductor buried directly beneath the power conduc- tors and running from tower to tower, or it may consist of several short conductors radiating outward from each tower base (9). The most common technique in mines is to wrap heavy wire around the base of each wooden pole before it is placed in the hole, the so-called butt wrap. Whichever technique is selected, the grounding electrode must always have as low a surge impedance as practical, and the downleads must adhere to the guidelines stated earlier for surge arrester connections. As a general rule- of-thumb, downleads and grounding electrodes should not KEY • Static wire o Line conductor 9 Protective angle '////// v//////, '///////, v////////. •'//////A Figure 11.34.— Overhead ground-wire shielding for low and high distribution towers. Wood braces. If steel braces, make allowances for loss of clearance. , Static wires -. Possible cross Insulator spacing Insulator spacing H-frame structure Single pole Figure 11.35.— Static-wire-protection designs of wooden support structures using 30° protective angle. be spaced apart more than 500 ft. Considering normal spans in mining, this translates to grounding the static wire no more than every other supporting structure. A few additional comments are needed on the appli- cation of static wires in mining. At the voltage levels used in conjunction with overhead lines to distribute mine power, the normal separation distance used for insulation between power conductors and static wires is perhaps insufficient to prevent a lightning stroke that hits a static wire from jumping directly to the power conductors (37). Often, the static wire is connected to earth only at the system ground bed at the substation. This can lead to a high-impulse impedance for the static-wire system, thereby reducing its effectiveness. An alternative method could be the use of protector tubes or surge arresters on the power conductors in conjunction with pole grounds (37). Additional points about static-wire applications in mining are presented in chapter 13. 300 Impulse Performance of Ground Beds The impedance of a ground bed is directly connected with its ability to help dissipate a transient. The general consensus about the performance of ground beds when subjected to impulse currents is that the surge resistance of an electrode (R s ) is almost invariably less than the normal resistance value (R n ) as measured by an earth tester (9, 14, 21). Examination of oscillograph records has shown that under surge conditions, the ground system exhibits a resistance equal to or slightly higher than its normal value for the first 1 or 2 us. The surge resistance then quickly decreases to a level usually from 20% to 80% of the normal value (36). Figure 11.36 illustrates that as the value of peak surge current increases, the ratio of impulse to 60-Hz resistance (Rg/R n ) decreases (7). Further- more, ground beds that exhibit a rather high normal resistance show a proportionally larger resistance de- crease when subjected to an impulse (36). Thus, two ground beds with dissimilar values of R n may exhibit R s values that are very close to each other (8). Experiments have shown that in many respects soil behaves as a dielectric material (17). When a certain critical potential gradient is reached, the soil breaks down or arcs internally. Thus, the ground-bed resistance de- crease under surge conditions is due to electric discharges within the soil, which spark across areas of high electric- field intensity. As shown in figure 11.37, the presence of moisture in the soil acts to increase its dielectric strength (17). Accordingly, very dry soil breaks down much more readily than soil containing a minor amount of water. Grounding systems can be designed in such a way that their performance under surge conditions is optimized. Figure 11.38 shows how increasing the pointedness of electrodes results in lower values of surge resistance (21). As a result, driven ground rods may be superior to meshes in lightning prone areas since the pointed electrodes are conducive to high stress concentration buildup and will therefore cause the soil to break down more readily under surge conditions. This chapter has described the electrical transient and its ramifications on mine power-system safety and integrity. An electrical transient has been defined as the outward manifestation of a sudden change in circuit conditions. The magnitude and fast rise time of the 1.0 uj y ^°2 .8 0_UJ J— _ 5U^ -6 — C/l U_< UJ „ oho; .4 Qui? u - IS) CREST IMPULSE CURRENT, kA overvoltage can cause damage to electrical components, particularly insulating materials. In addition to sound design practices, techniques and devices are available to reduce the damage done by surge voltages. The applica- tion of these in the design of mine power equipment is covered in the next two chapters. UJ < o > UJ CO CL < UJ a. 120 \ 4-in gap 90 - V ^^ " c ~ 1 ^^^ fin _ ?-in qnp **•*" — . 30 — - - 11% moisture i=i 0.38 % moisture i i i i 1 i i i i 1 i i i i 1 5 10 TIME TO BREAKDOWN, ^s 15 Figure 11.37.— Impulse breakdown of sand for two moisture conditions using spherical electrodes. 2,500 r A KEY No points 2,000 B C D E 2 -cm points 4-cm points 6-cm points 8-cm points 3 UJ u < H CO co UJ 1,500 1,000 500 \b \ ^ i i Figure 11.36.— Ratio of impulse to 60-Hz resistance as a function of peak impulse current, for driven rods. 20 40 60 PEAK IMPULSE VOLTAGE, kV Figure 11.38.— Impulse characteristics of spherical elec- trode, with seven attached pointed protrusions of various lengths. 301 REFERENCES 1. American National Standards Institute (New York). Stan- dard General Requirements for Distribution, Power, and Regulating Transformers. Stand. C57. 12.00-1973 (IEEE Stand. 462-1973). 2. . Standard for Surge Arresters for Alternating Cur- rent Power Circuits. Stand. C62. 1-1981 (IEEE Stand. 28-1981). 3. . USA Standard Guide for Application of Valve-Type Lightning Arresters for Alternating-Current Systems. Stand. C62.2-1981. 4. Stand. C84.1-1970. 5. Armstrong, H. R. Grounding Electrode Characteristics From Model Tests. Trans. Am. Inst. Electr. Eng., Part 3, v. 72, Dec. 1953. 6. Beehler, J. E. Capacitance Switching With Power Circuit Breakers. Pres. at IEEE Symp. on Capacitance Switching Capability, New York, Jan. 31, 1968; available from IEEE, New York. 7. Bellaschi, P. L. Impulse and 60 Cycle Characteristics of Driven Grounds. Trans. Am. Inst. Electr. Eng., v. 60, Mar. 1941. 8. Bellaschi, P. L., R. E. Armstrong, and A. E. Snowden. Im- pulse and 60 Cycle Characteristics of Driven Grounds-II. Trans. Am. Inst. Electr. Eng., v. 61, Dec. 1942. 9. Bewley, L. V. Theory and Tests of the Counterpoise. Trans. Am. Inst. Electr. Eng., v. 53, Aug. 1934. 10. Boehne, E. W. Energization Surges of Capacitive Circuits. IEEE Winter Power Meet., Conf. Paper 70CP235-PWR, 1970. 11. Buss, E. W., R. C. Dugan, and P. C. Lyons. Vacuum Circuit Breakers and Dry-Type Transformers Special Considerations for Mining Operations. Paper in Conference Record- IAS 12th Annual Meeting (Los Angeles, CA, Oct. 1977). IEEE, 1977. 12. Card, R. H. Earth Resistivity and Geological Structure. Trans. Am. Inst. Electr. Eng., v. 54, Nov. 1935. 13. Coyle, M. Effective Protection Schemes Smooth Transient Pains. Electron. Prod. Mag., v. 18, May 1976. 14. Davis, R., and J. E. M. Johnston. The Surge Characteristics of Tower and Tower-Footing Impedances. J. Inst. Electr. Eng. (London), v. 88, Oct. 1941. 15. Dornetto, L. D. The Importance of Grounding Systems in the Protection of Personnel and Equipment. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMinesIC 8694, 1975. 16. Eaton, J. R. Electrical Power Transmission Systems. Prentice-Hall, 1972. 17. Impulse Characteristics of Electrical Connections to the Earth. Gen. Electr. Rev., v. 47, Oct. 1944. 18. General Electric Co. (Syracuse, NY). Transient Voltage Suppression Manual. 1976. 19. Greenwood, A. N. Electrical Transients in Power Systems. Wiley, 1971. 20. Greenwood, A. N., D. R. Kurtz, and J. C. Sofianeko. A Guide to the Application of Vacuum Circuit Breakers. IEEE Trans. Power Appar. and Syst, v. 90, July/ Aug. 1971. 21. Hemstreet, J. G., W. W. Lewis, and C. M. Foust. Study of Driven Rods and Counterpoise Wires in High-Resistance Soil on Consumers Power Company 140-kV System. Trans. Am. Inst. Electr. Eng., v. 61, Sept. 1942. 22. Institute of Electrical and Electronics Engineers (New York). Recommended Practice for Electric Power Distribution for Industrial Plants. Stand. 141-1986. 23. Recommended Practice for Grounding of Industrial and Commercial Power Systems. Stand. 142-1972. 24. Kunjara, N. A., G. P. Russell, and E. K. Stanek. Prediction and Suppression of Electrical Transients in Mine Electrical Systems. Paper in Conference Record -IAS 12th Annual Meeting (Los Angeles, CA, Oct. 1977). IEEE, 1977. 25. Lewis, W. W. The Protection of Transmission Systems Against Lightning. Wiley, 1950. 26. Long, R. J. The Selection and Application of Vacuum Circuit Breakers for Open Pit Mining and Excavation. Paper in Con- ference Record- IAS 11th Annual Meeting (Chicago, IL, Oct. 1976). IEEE, 1976. 27. Marshall, J. L. Lightning Protection. Wiley, 1973. 28. McCann, G. D., and D. E. Morgan. Field Disturbances Pro- duced by Lightning. Trans. Am. Inst. Electr. Eng., v. 62, 1943. 29. McEachron, K. B. Multiple Lightning Strokes. Trans. Am. Inst. Electr. Eng., v. 53, 1934. 30. Morley, L. A., F. C. Trutt, and J. L. Kohler. Final Report -Evaluation of Coal-Mine Electrical-System Safety (grant 0155003, PA State Univ.). BuMines OFR 160-81, 1981; NTIS PB 82-139338. 31. Ohio Brass Co. (Mansfield, OH). How Does a Distribution Class Surge Arrester Work? Tech. Book. 2470-H, undated. 32. Pflantz, H. M. Analysis of Multiple Prestrike and Interrup- tion Phenomena in Capacitive Circuits. Pres. at IEEE Summer Meet, of Power Eng. Soc, San Francisco, CA, July 1972; available from IEEE, New York. 33. Rudenberg, R. Transient Performance of Electric Power Systems. MIT Press, 1967. 34. Smith, E. P. Lightning Arrester Applications for Mine Power Systems. Paper in Conference Record-IAS 11th Annual Meeting (Chicago, IL, Oct. 1976). IEEE, 1976. 35. Titus, C. H. Evaluation and Feasibility Study of Isolated Electrical Distribution Systems in Underground Coal Mines. Apr. 1972; NTIS PB 213 741. 36. Towne, H. M. Impulse Characteristics of Driven Ground. Gen. Electr. Rev., v. 31, Nov. 1928. 37. Westinghouse Electric Corp. (East Pittsburgh, PA). Elec- trical Transmission and Distribution Reference Book. 4th ed., 1964. 38. Westinghouse Electric Corp., Relay -Instrument Div. (Newark, NJ). System Neutral Grounding and Ground Fault Pro- tection. PRSC-4E, Ind. and Commer. Power Syst. Appl. Ser., Feb. 1986. 302 CHAPTER 12.— MINE POWER CENTERS The major power equipment in mining power centers, switchhouses and substations, was introduced in chapter 1, and an elementary overview of the protective circuitry involved was given in chapter 9. Other chapters have added numerous basic concepts and techniques whose thrust has actually provided the background for this chapter and the next, where the major power equipment used in mines is discussed in detail. In this chapter, the internal components and construction of typical mine power centers are covered. It should be noted that no formal standards have yet been developed for this equip- ment; hence, extensive use is made of information pro- vided by major manufacturers serving the industry and by several mine operators. Many circuit diagrams in this chapter apply directly to underground coal mining. With a few notable excep- tions, underground power equipment is the most complex electrical equipment found in mining. An understanding of this equipment should, therefore, lead to comprehension of that used in any other mining system. In the para- graphs that follow, some material contained in preceding chapters is reviewed, but usually the content is either changed or expanded. The main reason is to enhance many section presentations without requiring reference to other information. EQUIPMENT SPECIFICATIONS An important objective of this chapter and the next is to provide sufficient information for an individual to draw up the detailed specifications required when procuring a piece of power equipment to fit a mining need. It is often stated that a manufacturer supplies what the customer requests. Since there are no official standards nor recom- mended practices at present, and given the variability of most equipment, every piece of mine power apparatus must be specified individually. If these specifications are not drawn up in complete detail, down to each nut and bolt, the purchaser might not receive what he or she intended. Making detailed equipment specifications does not imply criticism of equipment manufacturers. It is a fact that many companies employ very capable applications- oriented engineers who have the total respect of the mining industry. Some can deliver precisely matched power systems just by knowing the mining equipment being used and the seam in which it is operating. Even in these cases, however, a complete specification from the buyer is still wise in terms of the protection it affords both the manufacturer and the customer. In situations where manufacturers receive specifications they believe are faulty or feel money can be saved by another approach, most manufacturers will contact the buyer for clarification rather than following the specifications rigidly and with- out question. A listing of the minimum number of items recom- mended for a specification is provided below. 1.0. 2.0. 3.0. 4.0. 5.0. 6.0. 7.0. 8.0. Time Schedule Work by Purchaser General 3.1. Intended use of equipment 3.2. Enclosure specifications 3.3. Special needs Internal Components 4.1. One-line diagram 4.2. Description of each component Drawing and Manual Requirements 5.1. With manufacturer's bid 5.2. Before construction starts 5.3. With equipment delivery Inspections 6.1. During construction and prior to delivery, by purchaser 6.2. Manufacturer compliance with local, State, and Federal regulations applicable to company 6.3. Equipment (being built) compliance with applicable local, State, and Federal regu- lations Guarantee 7.1. Minimum guarantee or warrantees de- manded by buyer 7.2. Request for proper storage requirements before service Delivery Dates 1 The author wishes to thank Thomas Novak, associate professor of mineral engineering, the University of Alabama, who prepared the origi- nal manuscript for this chapter while he was an instructor in mining engineering at The Pennsylvania State University. A document prepared to this format could be used as part of a purchase order directly to one manufacturer or as a solicitation for bids from several companies. Usually, such information is preceded by a general title, a phrase ex- pressing what the specification covers, and the number of units desired. The next paragraphs will present sugges- tions for the content within each item of the document. The individual making the specification is referred to as the buyer or purchaser. The time schedule alerts the manufacturer to when the buyer expects work on the equipment to begin. Fur- thermore, it serves notice that the manufacturer must have enough personnel, equipment, and material to com- plete work within the delivery time stated. (Obviously, if the number of sources for the equipment is restricted— and sometimes there could be only one, the buyer must accept the seller's delivery dates.) Item 2.0 should contain a statement that the purchaser will accept only equipment that meets the specifications. In other words, the units may be rejected if they do not comply, and the purchaser has no obligation to pay. A detailed account of the intended use for the equip- ment should be given under item 3.1 so the manufacturer knows exactly what the specification covers and what the buyer expects to do with the unit after receiving it. Item 303 3.2 primarily applies (but is not restricted) to unit- designed equipment. Content examples include: • Minimum material requirements. • Physical protection of internal parts. • Desired mounting (tire, skid, or rail). • Measures to prevent water from entering enclosure. • Maximum length, width, and height. Minimum materials should always be stated when the specification is going out for bid and when ruggedness for the mining environment is imperative. To prevent any inadequate bids, it is wise to provide minimum acceptable steel thicknesses for the base plate, frame, end plates, and covers. Specification of maximum dimensions is essential for underground operations. It may seem absurd, but there have been many instances where equipment has been purchased that is too large to fit down a shaft or slope. Special needs, item 3.3., refers to such things as engraved nameplates for all major components and labeling of internal wiring and terminal blocks. Item 4.0, internal components, is commonly the larg- est part of a specification. It should provide a complete listing of all components with one-line diagrams to show how they are to be connected. It is also helpful to state how each component is to be used and why each is included. Special needs such as transformer ratios, trip settings, and minimum insulation levels should be given where applicable. Item 5.0 is concerned with requirements for drawings and manuals. If the specification is being sent out for bids, it is good to request both one-line diagrams and schemat- ics of physical layout from prospective manufacturers. With unit-designed equipment, details for physical layout would include the basic frame, the mounting arrange- ment, and major component layouts. The manufacturer should be asked to submit all applicable outlines, arrange- ments, and schematics for approval prior to starting con- struction. The specification should also state the number of parts manuals and instruction manuals to be supplied on delivery of the equipment. Drawings, including some of reproducible quality, should also be requested where applicable. Item 6.0 should be included to protect the customer. The first entry must relate that the buyer has the right to inspect the equipment during manufacture and prior to delivery and that any discrepancies from the specification must be corrected. At times, orders can be construed as contracts; therefore, the buyer might be held (in part) legally responsible if the manufacturer violates any local, State or Federal ordinances, codes, acts, regulations, or laws. Thus, the specification must demand compliance with all applicable statutes; this should be extended to any subcontractors or suppliers that the manufacturer might use. The last entry for item 6.0 is a statement that the specified equipment shall comply with all Federal and State regulations and requirements that apply to the intended use in the State in which it will operate. Guarantees or warranties (item 7.0) cannot be ob- tained on some mine power equipment. Regardless, a request for a minimum guarantee is wise, for example, a 1-yr guarantee that begins after the equipment is placed in operation. The manufacturer should also be asked to supply any special instructions or precautions necessary for proper equipment storage prior to its being placed in service. Finally, the specification must relate the dates that the equipment is to be delivered to the mining operation. Establishing an adequate specification is not always an easy process. As will be seen, certain types of mine power equipment could require a document of numerous pages. However, a good specification might be the differ- ence between having a unit that performs superbly throughout the life of the mine or one that is a headache from the day it is placed in service. MINE POWER CENTERS The power or load center is one of the most essential power-system units for underground mines and, in a simpler form, for surface mines. Its primary function is to convert the distribution voltage to utilization voltage for operating equipment throughout the mine. It must incor- porate protective circuitry to ensure safe, efficient, and reliable operation. In effect, the power center could be considered a portable substation, but because of the ways in which it is used, its main component (and perhaps the enitre unit) might also be classified as a distribution transformer. The electrical components of the power center are usually metal clad, that is, housed in a heavy-duty steel enclosure that may be tire mounted, skid mounted, or track mounted. Illustrations of typical underground coal mining units are given in figure 12.1 (9, 12). 2 Towing lugs or pin-and-link couplers are commonly provided on each end of the enclosures to permit towing as mining advances or retreats. Bumpers or check plates are often installed to protect externally mounted components, such as couplers, from damage by mobile equipment. Similar enclosures can be used in surface mines, but here the simplest power center consists of outdoor components assembled on a flatbed trailer with a fence or gates to discourage unau- thorized entry. There is no such thing as a "standard" power center because of the variety of mining practices and the numer- ous types of mining equipment used. The power center may supply only 1 motor or as many as 20 pieces of machinery; it may be totally ac, dc, or a combination of ac and dc. The distribution voltage received by the power center can be 4.16, 7.2, 12.47, 13.2, or 13.8 kV. The outgoing ac utilization voltage may be 480, 600, 995, 1,040 V, or a combination of 995 or 1,040 V with one of the lower voltages. Dc can be at 300 or 600 V, but is almost always 300 V for face applications. As a result of this variety, manufacturers custom-build the units to meet the individ- ual needs and specifications of the customer. However, there is a general design philosophy central to all power- center types and this forms the basis of the following discussion of a typical ac power center and a combination ac-dc power center. 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. 304 Tire mounted The major electrical components of a typical mine power center are shown in the schematic of figure 12.2, and a possible component placement is provided in figure 12.3. The circled numbers in figure 12.2 will be used to indicate component location with respect to the overall system throughout the following descriptions of individual components. Skid mounted Track mounted Figure 12.1.— Typical power centers used in underground coal mines. HIGH-VOLTAGE CABLE COUPLER Incoming power is usually supplied to the center from the distribution cable through a high-voltage cable cou- pler (fig. 12.2, No. 1). The receptacle (typically with female contacts) is cable mounted, while the plug (male contacts) is gear mounted. Although the conductor pins are recessed in the coupler housing, dust caps should be provided for each side of the coupler pairs to protect the contacts when disengaged. When being disconnected, the coupler is designed so that the pilot contact is broken first, the line conductors second, and the grounding conductor last. As was dis- cussed in chapter 9, the pilot contact is interlocked with the upstream high-voltage circuit breaker, which protects the incoming line. If the incoming power is energized, the power will be tripped by the associated ground-check monitor prior to breaking the power contacts of the cou- pler. Having the grounding contact break last ensures that the frame is tied to earth ground whenever the power center is energized. A feedthrough receptacle may be provided as shown in figure 12.4; the contacts are typically female. This permits P o->o — n — n- f® HHHHHHHH h— j_ ff^T 5 KEY 1 High -voltage coupler 2 Interlock switches 3 Emergency-stop switch 4 Disconnect switch 5 Pilot-break monitor 6 High-voltage fuses 7 Surge arresters 8 Surge capacitors 9 Power transformer 10 Temperature device 1 1 Grounding resistor 12 Busway 13 Outgoing circuit breaker 14 Main circuit breaker 15 Voltage metering 16 Current metering 17 Outgoing cable coupler Figure 12.2.— Schematic illustrating major components in power center. 305 Outgoing cdbl LOW-VOLTAGE TRANSFORMER COMPARTMENT COMPARTMENT HIGH-VOLTAGE COMPARTMENT High-voltage coupler if 1 Feedthrough 1 J receptacle Figure 12.3.— Top view of mine power center showing place- ment of many internal components. To id-bn switch \ lo "X> load-break ■^ Qwil Figure 12.4.— Interconnections between input and feed- through receptacles. distribution power to be supplied (or continued) through the power center at other higb-voltage loads. A dust cap is again provided for use when the feedthrough receptacle is not in service. This cap also shorts the pilot contact to the grounding contact to provide a closed path for the ground- check monitor interlocking with the upstream circuit breaker. The conductors between the input and feed- through receptacles should be sized to the maximum capacity of the distribution system. For added safety, some manufacturers sectionalize their power centers into three separate compartments: high voltage, transformer, and low or medium voltage (fig. 12.3). If the equipment has no barriers, interlock switches should be placed on the low-voltage or medium-voltage exterior covers, with their contacts in series with the incoming pilot circuit. These additional switches are rec- ommended even with the compartment segregation, but then the interlocks are connected to trip circuit breakers associated with the transformer secondary, thus avoiding tripping the upstream switchhouse. An emergency-stop button (fig. 12.2, No. 3) should also be provided, and its function is similar to the interlock. It consists of a normally closed {NO set of contacts in series with the interlock switches. If the stop button is de- pressed, its contacts are opened, opening the incoming pilot circuit, and thus tripping the upstream circuit breaker. The switch should not automatically reset to the normally closed position after being depressed; manual resetting of the switch should be required. DISCONNECT SWITCH The disconnect or load-break switch (fig. 12.2, No. 4) is a mechanically operated air-type switch whose primary function is to allow a quick means of disconnecting the primary of the power-center transformer. A spring-loaded or torsion-bar mechanism provides the quick-make and quick-break operation, which is independent of the speed of the manually activated handle. Observation windows are provided in the power-center enclosure, and the switch can serve as a visual disconnect. Disconnect switches have no interrupting capability, but load-break switches do. Load-break switches are able to interrupt currents that are not in excess of their continuous-current rating. They are also rated for a max- imum interrupting capacity, but at this level they are designed for one-time interruption only. Load-break switches are more expensive than disconnects, yet many mine power engineers prefer them because of their extra ruggedness. Some States require their use. Continuous- current ratings of 400 and 600 A and a 15-kV voltage rating are common for mining applications. Table 12.1 illustrates typical ratings of a 400-A load-break switch. Table 12.1.— Typical current ratings of 400-A load-break switch INTERLOCK SWITCHES Electrical interlock switches (fig. 12.2, No. 2) are presently used in the high-voltage and transformer com- partments of mine power equipment. Strategically located around exterior top and side protective covers, they act to deenergize the internal power circuitry if the panels are removed, and thus help to prevent accidents caused by a worker's contacting energized components. The normally open (NO) contacts of the switches are connected in the pilot circuit of the incoming distribution cable as shown in figure 12.2. With covers in place, the interlock switches are depressed and their associated contacts are closed. This provides a closed path for the pilot circuit. When a cover is removed, the switch contacts are opened, thus causing the upstream circuit breaker to trip because of the loss of continuity in the pilot circuit. Full-load current Rated interrupting capacity Maximum interrupting capacity. Short-time current (1 s) Impulse current Close-and-latch current Rating, A 400 400 1,200 20,000 45,000 37,000 The only means of activating a load-break switch should be by the manually operated handle accessible from outside the load-center enclosure. Its operating mech- anism should not be tied into the power-center protective circuitry. More will be said about this recommendation soon. If a disconnect switch is allowed and employed, a pilot-break monitor (fig. 12.2, No. 5) is required to inter- lock the switch handle with the pilot circuit of the incom- ing distribution system. This allows the upstream circuit 306 breaker to interrupt the circuit prior to opening the switch contacts. It is also desirable to use pilot interlocking on load-break switches to extend the life of the contacts. Again, some States require both load-break switches and pilot -break monitoring in order to minimize any danger of operating the switch under load. The balance of this chapter will use or imply the term load-break for all disconnect-switch applications. HIGH-VOLTAGE FUSES A fuse has been defined as an overcurrent protective device with a circuit-opening fusible element that is heated and severed by the passage of current through it. Current-limiting fuses (fig. 12.2, No. 6) are the type used in mine power centers. Their main purpose is to protect the high-voltage section, particularly the transformer, during short circuits. High-voltage circuit breakers are not recommended here, because of their potential for creating transients. For resistance-grounded mine power systems, the fuse voltage rating should be based on the line-to-line voltage. The following two rules-of-thumb are recommended for determining the fuse current rating that will ensure that fuse action will not be triggered by the transformer inrush current. 1. The fuse should be able to withstand 12 times the rated current of the transformer for 0.1 s without element damage. 2. The element should be able to withstand 25 times the rated current of the transformer for 1/2 cycle. Based upon the above criteria, the continuous-current rating of the current-limiting fuses usually falls in the range of 1.5 to 2.5 times the rated transformer current. Transient overvoltages can be generated during the operation of the fuses. The magnitude of the overvoltage depends upon the point of the waveform at which initia- tion of the fault occurs and the size and design of the fuse element. For ribbon-element fuses, the generated peak arc voltage is primarily a function of the system voltage, as shown in figure 12.5 (10). The peak arc voltage should be compared with the minimum 60-Hz sparkover level of the surge arresters. The sparkover level of the arresters^ discussed in the next section, should be multiplied by V2 to obtain a peak voltage. If this value is greater than the peak arc voltage, arrester sparkover will not result from fuse interruption. The fuses are sometimes used in conjunction with an automatic (fused) load-break switch, which has the objec- tive of preventing the power center from operating under a single-phase condition if only one fuse blows. The fuses normally have actuators located at one end of their cas- ings. In the event that a fuse element is blown, the actuator extends outward from the fuse end and activates the trip mechanism of the switch. This causes the switch to open and eliminates the single-phase condition. How- ever, for a three-phase fault, the possibility exists that one fuse may blow before the others, which could result in the load-break switch attempting to clear a fault current higher than its designated rating. As a result, this type of automatic load-break switch is not recommended in power- center applications. Single-phasing protection is better accomplished at the secondary bus, using relays to monitor for a single- < b o > o ir < < UJ Q. X < uu 90 80 70 / / 60 50 40 30 ?0 10 5 10 15 20 25 30 CIRCUIT VOLTAGE, kV 35 Figure 12.5.— Graph illustrating transient crest voltage caused by ribbon-element current-limiting fuse operation. phase or phase-reversal condition. The relay contacts can be employed to activate the undervoltage release or shunt trip of the main breaker on the secondary. If a main breaker is not used, the contacts can be connected directly at the output of the control transformer secondary wind- ing, so all outgoing circuit breakers will trip through their undervoltage releases if the relay is actuated. SURGE ARRESTERS The surge arrester (fig. 12.2, No. 7) is a device de- signed to limit dangerous transient overvoltages to safe levels. Lightning, switching surges, and some faults result in transient overvoltages that can exceed the insulation levels of power-system equipment. Since it is not always economically possible to rate the equipment insulation above the surges, overvoltages must be clamped or sup- pressed to tolerable levels. The function of an arrester is • To discharge the energy associated with a transient overvoltage; • To limit and interrupt the 60-Hz current that fol- lows the transient current through the arrester; and • To return to an insulating state without interrupt- ing the supply of power to the load. Valve-surge arresters of the low-sparkover distribution class are used almost exclusively in mine power centers. A complete discussion of these devices as well as application tables was provided in chapter 11. The voltage rating of a surge arrester is the highest power-frequency voltage at which the arrester is designed to operate. As mine power systems utilize resistance grounding and fall into the category of noneffectively grounded systems, the surge-arrester voltage rating must be selected on a line-to-line basis. However, the rating should be 5% to 10% above the nominal distribution voltage to allow for voltage-regulation compensation. 307 The component protected by the surge arrester is the main power transformer. The key to protection is coordina- tion between the transformer insulation withstand, or BEL (basic impulse insulation level), and the characteristics of the transient voltage the arrester lets through. Dry transformers are used almost exclusively in present-day power centers, and their insulation strength does not increase significantly above their BIL as the duration of the applied pulse de- creases. Therefore, the voltage-time withstand characteristic can be plotted as a flat line with a value equal to the BEL, as shown in figure 12.6 (11). The arrester characteristic, includ- ing impulse sparkover and discharge voltage, is also illus- trated. The margin of protection is the difference between the transformer BEL and the arrester discharge characteris- tic at any given instant of time. This value should not be less than 20% if a 5,000-A surge discharge current is assumed. Note that the worst case discharge of 20,000 A need not be used in typical power-center locations. The physical location of surge arresters within the power center is important for effective operation. The two critical distances are the conductor lengths to the line conductor and to the power-center frame ground, and the distance between the arrester and the transformer. To maximize arrester performance, the arrester leads must be as short and straight as possible and made of No. 6 AWG solid copper or larger. Also, the surge arrester and arrester connections to the line conductors should be as close to the transformer primary terminals as possible. Preferably, the surge arrester leads should be connected right at these terminals. Surge capacitors (fig. 12.2, No. 8) can be found in numerous mine power centers. Their intended purpose is to lessen the severity of transient overvoltages caused by current chopping in vacuum circuit breakers. Their effec- tiveness is based upon the ability to limit the rate-of-rise of transient overvoltages and to reduce the system character- istic impedance. However, as was discussed in chapter 11, surge capacitor use in load centers is generally not re- quired, provided that the surge arresters are correctly applied, and a minimum practical cable length (100 ft) exists between the power center and the upstream circuit breaker. BIL (full-wave withstand) < o > Minimum margin of protection . Front-of-wave [ sparkover ■*- Discharge-voltage maximum characteristics of arrester Impulse voltage rising 100 kV/^s for each 12 kV of arrester rating 1 2 3 4 5 6 TIMERS Figure 12.6.— Comparison of transformer withstand characteristic and surge arrester withstand characteristic. TRANSFORMERS The main transformer (fig. 12.2, No. 9) can be consid- ered the heart of the power center, since its primary function is to convert the distribution voltage to utiliza- tion. Proper selection is therefore imperative from the standpoint of safety, efficiency, and reliability. Fortu- nately, the IEEE (2, 4) has established classifications and specifications for determining transformer characteristics as an aid in design and application, and these regulations are in general use in the mining industry. The IEEE uses three items for the general classifica- tion of transformers: distribution or power, insulation, and substation or unit substation. The capacity or kilovoltam- pere rating determines whether the transformer is classed as a distribution or power unit. Distribution transformers fall in the range from 3 to 500 kVA, with power transform- ers having capacities greater than 500 kVA. Power-center transformers can be of either classification, since they may range from 150 kVA for a rectifier or belt drive, up to 2,250 kVA for a longwall section. Capacities are rarely greater than 1,250 kVA. A transformer may be further classified by the insu- lation system as liquid or dry. Liquid insulation includes mineral oil or synthetic fluids, and dry transformers are ventilated or sealed gas-filled types. Ventilated dry units are used almost exclusively in mine power centers. With convection cooling and air insulation, dry transformers have the following advantages: • Toxic gases cannot be released. • The transformer cannot explode or catch fire. • There is no oil or other liquid to spill, leak, or dispose of (a critical problem with PCB, poly chlorinated biphenyls). • The transformer is virtually maintenance-free be- cause there are no valves, pumps, or gauges. The substation-transformer classification refers to direct or overhead-line termination facilities, while a unit- substation transformer has an integral connection to pri- mary or secondary switchgear. Primary refers to a voltage of 1,000 V or higher, while secondary refers to a rating less than 1,000 V. For practical purposes, 1,040 V is also considered 1,000 V. Therefore, a load-center transformer can be considered a unit-substation class. Specifications The IEEE transformer specifications (4) that apply to mine power centers include • Capacity or kilovoltampere rating, • Phases, • Frequency, • Voltage rating, • Voltage taps, • Winding connections, • Impedance, • BIL, • Temperature rise and insulation, • Cooling. Each will be discussed in general terms in the next para- graphs. There are a few exceptions to the IEEE standard. 308 There is no set formula for determining the kilovolt- ampere rating for a power-center transformer. For con- stant loads, it is a common rule-of-thumb to allow 1 kVA for each horsepower of connected load. However, the min- ing process does not produce a constant load (that is, all connected motors are not operating at the same time on a continuous basis); hence, using the rule-of-thumb will normally result in oversizing the transformer. Past expe- rience and demand factors established by manufacturers and operators, along with the horsepower of the connected load, are essential for determining the transformer capac- ity. For typical underground mining sections, the kilovolt- ampere rating may lie within the range of 50% to 80% of the connected horsepower. (Chapters 4 and 8 contain additional information on demand factors.) As an example of sizing, consider the following continuous-mining section: 1. 370-hp continuous miner (gathering head, pump, tram, and cutter motors), 2. Two 125-hp shuttle cars (traction, conveyors, and pump motors), 3. 80-hp roof bolter (traction and pump motors), and 4. 50 hp of auxiliary equipment (section fan, sump pump, hand tools, etc.). The total connected horsepower sums to 750 hp. If the demand factor has been determined to be 60%, the effec- tive load would be 450 kVA, and a 500-kVA transformer would be selected. However, flexibility must always be considered in mining applications, and as a result, it may be necessary to select the next higher kilovoltampere rating to accommodate anticipated additional loads. Over the years, power-center manufacturers have ar- rived at certain standard or typical transformer capacities, created mainly by repeat demand of the industry. These are- Ac output only (section may have outboard rectifier for dc load): 150, 225, 300, 500, 600, 750, 1,000, and 1,250 kVA. Dc only (often termed rectifiers; rating of unit com- monly in kilowatts): 100, 150, 200, 300, 500, and 750 kVA. Combination ac-dc (typical dc capacity; on larger units, 200 and 300 kW for dc also available): 300/150, 500/150, 600/150, 750/150, 1,000/150, and 1,250/150 kVA. Many manufacturers will supply capacities at any incre- ment of 50 kVA up to 2,250 kVA, the maximum allowed by Federal regulations for face applications in underground coal mines. Three-winding transformers are necessary in many power centers, for example, when ac face equipment has two rated voltages such as a 950-V continuous miner working with 550-V machines, or in the common applica- tion of mixed ac and dc machinery. In these cases, the capacity of each transformer winding (primary, secondary, and tertiary) must be individually rated. All mine power-center transformers are three phase, being either three single-phase units where each trans- former is rated at one-third of the total required capacity, or integral three-phase types with construction allowing field replacement of failed windings. The integral three- phase transformers are preferred because they have equal reliability, lower cost, and higher efficiency, take less space, and have fewer exposed interconnections than three single-phase units. Transformers should be rated at the standard com- mercial frequency in the United States, which is 60 Hz. If a transformer rated at 60 Hz is applied to a 50-Hz system, the voltage and the kilovoltampere rating must be reduced by 20%. Transformers rated at 50 Hz can be operated at 60 Hz, but the efficiency and regulation are reduced, since the reactance is directly proportional to the frequency. The primary voltage rating of the transformer is dictated by the distribution voltage (that is, 4.16, 7.2, or 12.47 kV, and so on). The voltage rating of the secondary winding must match the voltage of the utilization equip- ment. In practice, the transformer secondary should be rated at a higher voltage (10% is common) than shown on the nameplate of the utilization equipment, to allow for a voltage drop in the trailing cables. The secondary winding of power-center transformers is most commonly rated at 480 or 600 V, relating to 440- or 550-V equipment, respec- tively. Secondaries supplying 950-V machines are usually rated at 1,040 V, except for machines that cannot be remotely controlled or in States where face voltages over 1,000 V are not allowed. In these instances, 995 V is often specified. The transformer voltage for dc applications is dictated by the rectifiers and will be presented in a separate section. Power-center transformers should be provided with voltage taps on the primary windings to account for voltage fluctuations and line losses in the distribution system. Five fully-rated taps at 2.5% increments are usually available, with the middle tap being rated at the nominal distribution voltage. Figure 12.7 illustrates this arrangement for a 7.2-kV primary on a per-phase basis. Delta primary and wye secondary are the preferred connections for standard two-winding power transformers and are the connections commonly used in mine power centers. An external busing from the neutral of the wye secondary provides an easy means for resistance grounding, as required by Federal regulations. The delta-connected primary provides isolation of the distribution circuit from the utilization circuit with respect to zero-sequence currents resulting from exciting currents or secondary ground faults. The delta-wye connection also stabilizes the secondary neu- tral point and minimizes the production of third-harmonic voltages. Additional information concerning the reasoning behind using delta-wye transformers can be found in chap- ters 4 and 9, including the reasons wye-wye transformers are not recommended. In certain situations, a delta-connected secondary may be specified or required, and the primary may be delta or wye in this case. If neutral grounding is desired or required, a grounding transformer is needed to derive a neutral. The zig-zag grounding transformer, illustrated in figure 12.8, is sized according to the continuous-current SECONDARY Figure 12.7.— Typical primary winding taps on power cable transformer. 309 rating of the neutral-grounding resistor. Standard single- phase or integral three-phase transformers, arranged in a wye-delta configuration, can be used as grounding trans- formers in lieu of the special zig-zag type, as shown in figure 12.9. The transformer impedance is sometimes referred to an -impedance voltage because of the common means of measurement. Impedance voltage is the voltage required to circulate rated current through one of two windings of a transformer when the other winding is short-circuited, with the windings connected as for rated voltage operation (4). An example will help to illustrate this procedure. Figure 12.10 shows a three-phase delta- wye distribution transformer with the indicated ratings. The secondary of the transformer is short-circuited while a three-phase variable voltage source is connected to the primary. The voltage source is wye connected so the impedance voltage, Vj, will be a line-to-neutral voltage. The source voltage is increased until the ammeter reads the rated or full-load current of the transformer, I FL , which can be determined by W = s V3 V (12.1) where I FL = full-load current of transformer, A, S = rated transformer capacity, VA, and V = rated voltage at transformer winding termi- nals, V, or for the primary, Line leads Neutral leads Windings on coil Schematic Figure 12.8.— Zig-zig grounding transformer. To line conductors A A ILUU UuuJ Luuu Neutral grounding resistor I Figure 12.9.— Delta-wye connection for deriving a neutral. J-TTT — 500,000 V3 (7200) = 40.1 A. Consider the case where V ; = 250 V, when the ammeter reads 40.1 A. The impedance voltage is normally ex- pressed in terms of a per-unit value; therefore, 250 V nn „ V - = 4160V = 006 P u ' The line-to-neutral rating of the transformer (4,160 V) is used since the impedance is being measured as a line- to-neutral potential. For this example, the line-to-neutral impedance of the transformer, with all quantities reflected to the primary, would be LiTy V: 250 V I, 40.1A = 6.23 Q. 7, 200 /480 V 500 kVA Figure 12.10.— Technique for measuring transformer im- pedance. The transformer impedance can be measured in terms of a per-unit quantity by using the ratings of the transformer as base values, as follows: kVA b = 500 kVA, kV b = I b = 7,200 ~7T kVA = 4.16 kV (line to neutral), 500 V3kV b " V3 (7.2) " 4(U A ' _ V b _ 4,160 b I b "" 40.1 = 103.7 0. Thus, the per-unit value of the transformer impedance is Z„ 6.23 nn „ ^" ^ To3?Tfi = 006 pu - This yields the same results; therefore in terms of per-unit quantities, the terms impedance and impedance voltage are interchangeable. 310 The vector sum of the leakage reactance and the resistance of the windings comprise the impedance of a transformer. Figure 12.11 shows that the X/R ratio of a typical transformer increases as the size (megavoltampere rating) increases (13). The transformer impedance is ex- tremely important since it plays a major role in voltage regulation as well as dictating the amount of current that can be delivered to a fault on the secondary. Here, the maximum symmetrical secondary-fault current (three- phase fault) can quickly be estimated by L, u J Tpu (12.2) where I flmax) = maximum symmetrical fault current, A, I FLs = full-load current of transformer secondary, A, and Z.^ = transformer per-unit impedance, pu fl. For the previous example, 6 -. I— XIK UJ — ' <-J O UJ 40 30 20 10 ty. ) 1 2 5 10 20 50 100 200 500 1,000 SELF -COOLED TRANSFORMER RATING, MVA Figure 12.11.— Typical X/R ratio versus transformer capaci- ty \ 'A V3kV 500 V3 (0.48) = 601 and K 601 0.06 = 10,023 A. This technique produces a more pessimistic estimate than would actually occur, since it assumes that the trans- former primary is connected to an infinite bus. A more realistic value for fault current should take into account the system impedance upstream from the transformer as well as an equivalent impedance for the utility. However, this system impedance is constantly changing because of the dynamic nature of the mining process. By using the above method, it is possible to obtain a conservative determination of the required interrupting capacity for protective circuitry serving the utilization equipment. For transformers used in mine power centers, the per-unit values of impedance normally lie in the following range: • 0.04 to 0.05 pu for 150 to 450 kVA, and • 0.05 to 0.06 pu for 500 to 1,250 kVA. The 0.04-pu lower limit is usually needed to limit short- circuit currents to reasonably safe values. For large power transformers, the leakage reactance may be intentionally increased in the design of the transformer to further limit the available fault current. External air-core reactors in series with the secondary windings are sometimes used to restrict fault current from transformers that feed rectifi- ers. This will be discussed in the section on ac-dc combi- nation power centers. For transformers used in mine power centers, the following minimum insulation ratings should be used (the BIL's in parentheses are optional): 4.16-kV primary, 35-kV (60-kV) BIL; 7.2-kV, 60-kV (75-kV) BIL; 12.47-kV (13.2 and 13.8 kV), 95-kV BIL; and 23-kV, 150-kV BIL. Coordination with surge arrester characteristics must be maintained, as discussed earlier. (See chapter 11 for fur- ther information.) The kilovoltampere rating of a transformer is based upon continuous operation at rated voltage and frequency without exceeding the specified temperature rise or limit- ing temperature. The temperature rise depends upon the core and conductor losses. The core losses remain constant with load, while the conductor losses are determined by the I 2 R of the windings. Dry transformers used in mine power centers should have class 220 °C insulation. This classification terminol- ogy has replaced the old class H 150°C rise notation that was based on an average ambient temperature of 30°C, allowing a maximum 150°C temperature rise in the windings (2). This restricted the average insulation tem- perature to 180°C, but a 220°C hot-spot temperature was also allowed. The new classification simply places an absolute allowable maximum temperature on the trans- former when it is operating continuously under full capac- ity at rated voltage and current. Because of the use of trailing cables and mobile equip- ment, short-circuit conditions are frequently encountered in mining, and transformer windings can undergo considerable thermal and magnetic stress during these occurrences. Con- sequently, the transformer should be designed to withstand 25 times rated current in any winding for a period of 2 s. One method of curtailing expansion is to add an insulated brace across the windings, and some engineers request adjustable braces, which allow periodic readjustments to be made when expansion causes loosening. A temperature-sensing device (fig. 12.2, No. 10) can be placed in the transformer windings to prevent damage from overheating. The device controls a set of contacts located in the pilot circuit of the incoming distribution line. If the transformer temperature exceeds the specified limit, the contacts in the pilot circuit are opened, which results in tripping the upstream circuit breaker. As an alternative, the contacts can be used to activate the tripping element of a main secondary breaker or all outgoing breakers. Mine power centers are usually cooled by natural convection, and the side panels of the transformer com- partment normally have louvers to allow air circulation. The transformer windings are designed so that the heat generated by the I 2 R losses is exposed to an adequate amount of cooling to handle the expected loads. The effective cooling areas are the inside of the winding, the 311 outside of the winding, and the cooling ducts within the winding. The movement of air by convection carries away the heat generated by the winding (13). Transformer Construction The two main transformer components are the core and the windings. Figures 12.12 and 12.13 show a typical mine power-center transformer in the process of being constructed and in completed form. The transformer core provides a path of low reluctance to the flux produced by the primary windings. In effect, it comprises the magnetic circuit of the transformer. To reduce eddy currents, the core is constructed of laminated sheet steel. The eddy- current losses vary as the square of the thickness of laminations, and core laminations are usually from 0.01 to 0.02 in thick (13). The laminated material is cut from silicon-iron sheets. The silicon reduces the reluctance to hysteresis and prevents increased loss with age. The laminated material is specially annealed to obtain a high permeability and is also treated with a chemical coating to insulate the laminations from each other. The laminations are stacked one upon another to construct a closed mag- netic path. Alternate layers are staggered in an interlam- inar core-and-gap construction so all joints do not meet at the same place. This is done to reduce the effect of the air gap between the joints and make the entire structure function more like a solid piece of iron (13). The primary and secondary windings comprise the current circuit of the transformer. The windings are de- signed to get the required number of turns into the minimum of space through the core opening, while also providing enough room for insulation and cooling ducts. Transformer windings are made of copper or aluminum. Because of its lower conductivity, aluminum requires a larger cross section of conductor and thus a larger opening in the core. The conductors used in the windings may be round, square, or rectangular in cross section. Aluminum secondaries are often wound using sheet metal. All wind- ings may be insulated with enamel, mica paper, NOMEX paper, silicon glass tape, or a combination of these materi- als (8). Faraday Shields Transformer windings can be provided with a grounded Faraday (or electrostatic) shield to destroy inter- winding capacitance between the primary and secondary, thus reducing the danger of transferring distribution transients to utilization. Another important use of the shield is to prevent interwinding faults between layered windings. This is especially critical when the secondaries are isolated above ground, as the primary voltage can be impressed on the secondary without detection. Using the shield, a ground fault will occur in this situation, and the resulting ground current will be detected by upstream ground-fault relaying. The shield consists of a layer of nonmagnetic metal placed between the primary and secondary windings, insulated from all windings, and connected solidly to ground (1). Aluminum or copper can be used, but the shield should be made of the same material as the main windings. Physically, it may be a single turn of sheet metal or a closely wound single layer of wire. To obtain interwinding fault protection, the shield must have the same ampacity as the grounding conductors leading to the Figure 12.12.— Typical mine power-center transformer under construction. (Courtesy PEMCO Corp.) Figure 12.13.— Completed transformer prior to installation. (Courtesy PEMCO Corp.) power center, in order to carry maximum available ground-fault current (at least, equivalent to one-half the cross-sectional area of the primary-winding wire). GROUNDING RESISTOR Federal regulations require the maximum frame po- tential on medium-voltage and low-voltage circuits to be limited to 40 V. In an attempt to ensure this, the regula- tions further require the maximum ground-fault current of these circuits to be limited to 25 A. However, many modern power centers are designed to limit the ground fault current to 15 A for an additional factor of safety. Limiting the ground-fault current results in the following additional benefits (7): • Reduction in burning and melting of faulted elec- trical equipment, • Reduction in mechanical stress of faulted electrical equipment, • Ability to selectively clear the faulted circuit, and • Reduction in overvoltages that might cause insula- tion failure. 312 The grounding resistor (fig. 12.2, No. 11) is inserted between the neutral of the transformer and the power- center frame as a means of limiting the ground-fault current. The ohmic value of the grounding resistor is based on the maximum ground-fault current condition; that is, a ground fault at the secondary terminal of the transformer. For this situation and neglecting the trans- former impedance, the resistance is determined by R V >n (12.3) where R G = ohmic value of grounding resistor, U, V ln = line-to-neutral potential of transformer sec- ondary, V, and If = maximum ground-fault current, A. As an example, consider sizing a grounding resistor to limit the ground-fault current to 15 A for a 480-V system. The ohmic value of the resistor would be Rn = 480/V3 15 = 18.5 n. When sizing a grounding resistor, the time rating must also be considered. Under normal conditions, an insignificant amount of current flows through the resistor, but during the worst case situation, 15 A may flow through it. Thus the power (P) dissipated by the neutral- grounding resistor would be or P = I f 2 R G P = (15) 2 (18.5) = 4,162 W. (12.4) Normally the resistor would only be required to dissipate this power for the time it takes for a circuit breaker to trip, but the possibility of the circuit breaker's failing to trip must also be considered. Federal regulations require the resistor rating to be based on an extended time rating. This has been defined as 90 days of operation per year. The end of the resistor that is connected to the neutral of the transformer is also required to be insulated for line-to-line voltage. BUSWAY The power-center output requires numerous taps for feeding the utilization circuits. A busway (or busbars) provides a convenient and economical means of providing these taps (fig. 12.2, No. 12). The busway consists of flat, bare conductors supported within the power-center enclo- sure by means of insulators made of glass, polyester, or porcelain, as can be seen in figure 12.14. Busways are available with either copper or aluminum conductors. Aluminum has a lower electrical conductivity and lower mechanical strength and quickly forms an insulating film on its surface when exposed to the atmosphere. If alumi- num conductors are used, they should have electroplated contact surfaces (tin or silver), and bolting practices that accommodate aluminum mechanical properties should be used at electrical joints (5). The continuous-current rating of the busway is based on the cross-sectional area of the conductor and a maxi- mum temperature rise of 55° C from ambient. The short- Figure 12.14.— Typical bus work in power center under con- struction. (Courtesy PEMCO Corp.) circuit rating should exceed the maximum available fault current. The sizing of the busway is basically the same as that discussed for other conductors in chapter 8. OUTGOING CIRCUIT BREAKER Molded-case circuit breakers (fig. 12.2, No. 13) are usually employed to protect ac utilization equipment and associated cables. Some manufacturers have special mine- duty breakers available, which have greater ruggedness and reliability to meet the demands of mining practices. Dual-element fuses are rarely used because of single- phasing problems and the lack of auxiliary protective- relaying capabilities. A complete discussion of molded- case breaker and fuse applications, sizing, and problems is available in chapters 9 and 10. A presentation of molded- case devices with solid-state tripping elements is con- tained in chapter 14. The power-center circuit breaker compartment must be designed for easy access, but protection is necessary against accidental exposure to energized terminals and conductors. Dead-front panels are the preferred method of construction. Here the breaker is mounted on a recessed panel so that only the operating handle and adjustments are accessible from the outside. As an alternative, the breakers can be surface mounted, but here too a barrier should cover the power terminals. Doors should be used on the exterior frame to minimize exposure of the compart- ment to dust and other contaminants. The two kinds of protection that can be provided directly by molded-case circuit breakers are short circuit and overload. Short-circuit protection is required on all outgoing power (ungrounded) conductors, and maximum allowable instantaneous-trip settings are established by Federal regulations. Overload protection is mandated in some States. Magnetic elements give instantaneous-trip protection, and thermal elements afford overload protec- tion. Pickup settings for both are based on the smallest size of cable being protected. Undervoltage protection of each outgoing circuit is almost always required, the only exception being if the protected equipment has its own. An undervoltage release (UVR) is invariably used and is an auxiliary solenoid that trips the operating mechanism of the breaker whenever its 313 coil voltage drops below 40% to 60% of rated. Additional external protective relaying can be given to the outgoing circuit through the UVR. Note that even though under- voltage protection may be needed, each outgoing breaker in a mine power center should contain an UVR. Outgoing breakers are rarely provided with shunt-trip solenoids. Terminal connectors are available for molded-case breakers for either single- or multiple-conductor entry. Copper conductors require the use of copper terminals, and aluminum conductors require aluminum-compatible terminals. If multiple-conductor terminals are employed, the terminal must be constructed so that each conductor can be tightened without removing another conductor. Figure 12.15 is a representative view of conductor connec- tions to a molded-case breaker. The selection of molded-case circuit breakers is based on the voltage, frequency, interrupting capacity, continuous-current rating, and trip settings. Even though this topic is presented in chapters 9 and 10, additional coverage of tbe practical considerations for interrupting- capacity and continuous-current selection is warranted here. Molded-case circuit breakers often begin interrupting short-circuit currents during the first cycle after the fault, and so they must be selected on the basis of maximum first-cycle asymmetrical fault current. The breakers are usually rated on a symmetrical current basis, which eliminates applying dc offset multipliers when selecting the breaker. It is generally considered adequate to use calculated symmetrical short-circuit currents for load- center applications. As an example, consider a 750-kVA power center with a secondary utilization voltage of 600 V. If the transformer impedance is 6%, the maximum symmetrical fault cur- rent, If (max) , can be calculated as 750 1fLs " V3 (0.6) " 722 A ' 722 ■<"■"■■«> = O06- 12 ' 033A - This of course assumes a worst case condition, since transformer impedance is the only limiting factor for all fault current. In this case, even a 100-A standard mine- duty breaker could interrupt the fault; the typical inter- rupting capacity is 14,000 A symmetrical at 600 V. How- ever, if the power center contains a 1,000-kVA transformer with 6% impedance, fault current becomes *- = ^S = 962 a > 962 0.06 = 16,038 A. Under this situation, the standard 100-A molded-case breaker would have inadequate interrupting capacity. (A premium-duty 100-A unit would be safe; typical interrupt- ing capacity is 18,000 A symmetrical at 600 V. See chapter 9 for other typical ratings.) The simple example given here shows just one reason for the minimum transformer im- pedance stated earlier in this chapter. Figure 12.15.— Typical conductor connection to molded- case circuit breaker. (Courtesy PEMCO Corp.) The continuous-current rating for a molded-case cir- cuit breaker is actually the rating of the thermal-trip element, which can be less than the breaker frame size (a thermal-magnetic breaker). Magnetic-only devices have a continuous-current rating equal to the frame size. The rating is based upon 100% continuous current at 40°C, but sizing the breaker is normally related to 80% of the rating. This means that sizing is based upon 1.25 times the cable ampacity or the full-load circuit current. In other words, the breaker operates at 80% of the continuous-current rating at full-load current. (Note that some molded-case breakers are applied at 100% full-load current, particularly those with solid-state tripping elements.) For instance, consider sizing a breaker to protect a 4/0 cable for overload. A typical ampacity for 4/0 three- conductor unshielded trailing cable is 287 A. Thus, the breaker continuous-current rating would be (1.25) (287) or 359 A. The next highest standard size, or a 400-A rating, would then be selected. However, when sizing outgoing breakers for a mine power center, the method of selection is not so simple. The process can approach something of an art based on prior experience, especially when under- ground coal mining equipment is involved. An actual situation is used here to illustrate some difficulties that can be encountered. From a practical standpoint, 4/0 is the largest trailing-cable size that can be used in underground mining and is the most common conductor size for continuous miner applications. Assume that the continuous miner has the following motors; • Cutting motors, two at 90 hp; • Pump motor, 50 hp; • Gathering-head motor, 70 hp; and • Traction motors, two at 35 hp. Assuming a 480-V system, an average efficiency of 90%, and an average power factor of 90%, the maximum theo- retical full-load current for the machine (370 connected horsepower) would be from equation 8.3: I™ = (370) (0.746) V3 (0.48) (0.9) (0.9) = 410 A. This current exceeds the cable ampacity; thus, breaker sizing based on cable ampacity is not applicable. It should be realized that it would be a rare situation to have all 314 motors draw rated current. Each motor or motor pair has its own noncontinuous duty cycle, which is a function of many variables. It is for this reason that an undersized trailing cable can be used. Yet the circuit breaker must be selected to handle the load demands of the continuous miner. Therefore, the sizing of the continuous-current rating for a breaker does not follow a set procedure. Most power-center manufacturers are well aware of the machin- ery available and can usually provide the optimum circuit- breaker size for specifications. (Again, see chapter 9 for more discussion.) A main circuit breaker (fig. 12.2, No. 14) is often recommended if the number of outgoing circuits exceeds three. The breaker has the bus work as its primary protection zone, but also serves as a backup to the outgo- ing breakers. The rating can be based upon the full-load current of the transformer secondary or the ampacity of the bus work, whichever is lower. In general, the bus work is sized on the transformer full-load current. For example, the full-load current for the 600-V secondary of a 750-kVA power center is 750 V3 (0.6) = 722 A. Thus, the breaker continuous-current rating should be based on I FL = (1.25) (722) = 902 A. A 1,200- A frame with a 1,000- A continuous-current rating would be selected. The thermal elements would provide overload protection to the transformer and bus work; magnetic elements could be set to give short-circuit pro- tection to the bus work. Some main breakers are provided only with magnetic elements. Obviously, the main breaker must be coordinated with the outgoing breakers and the high-voltage fuses on the transformer primary. schemes work with a low-cost, low-burden CT. Since the CT is drastically underutilized, its secondary current cannot be predicted by knowing the turns-ratio of the CT. The pickup values must be determined by testing. The normally open (NO) contacts of the relay usually parallel the UVR of the associated circuit breaker. When the relay pickup value is exceeded, its associated contacts close and short out the UVR coil. This technique has been adopted to eliminate nuisance tripping due to bounce and vibration, which hamper circuits that have contacts in series. One problem with paralleled contacts as shown in figure 12.16 is that ground-fault protection is lost if the relay is removed from its socket. To prevent this, UVR power can be supplied through a jumper in the ground- fault relay case (fig. 12.17). Without the relay, the circuit breaker cannot be closed, except in some small molded- case units, such as 50-A units. Line conductors Molded-case circuit breaker To control transformer GTR(50G) Figure 12.16.— Zero-sequence relaying on outgoing circuit with control connections to breaker. GROUND-FAULT PROTECTION As mentioned earlier, a grounding resistor is placed between the transformer neutral and the power-center frame to limit ground-fault current to not more than 25 A. Ground-fault relaying must be used on each outgoing ac circuit to initiate circuit interruption during malfunc- tions. The common relay schemes for ac applications in load centers are zero-sequence, neutral (direct), and poten- tial. Zero-sequence relaying is the most common ground- fault protection and is utilized on practically all outgoing ac circuits. As shown in figure 12.16, the three line conductors are passed through a window-type current transformer (CT), and burden for the CT secondary is a ground-trip relay. Relay operation was described in chapter 9. The ground-trip relay must be set to pick up at 40% or less of the maximum ground-fault current. For this situa- tion, a low-ratio CT might be expected to be used, such as 25:5 or 50:5 (ampere-turns ratio). However, better sensitiv- ity has been realized by using a high-ratio CT, 350:5 or higher, with a voltage-sensitive relay (about 1.5-V pickup), instead of a current-sensitive relay for the tripping device. Some manufacturers employ a slight variation to this conventional scheme by rectifying the CT output and using a dc voltage-sensitive relay. Either way, these Line conductors ^- Molded-case circuit breaker To control transformer GTR(50G) Figure 12.17.— Zero-sequence relaying with jumper in relay case. 315 Neutral relaying is sometimes used in mine power centers when a main breaker is used. The neutral conduc- tor from the transformer is encircled by a CT, as shown in figure 12.18. The CT and the ground-trip relay are the same as discussed for zero-sequence relaying. However, a time delay should be introduced by the ground-trip relay so selective interruption of the faulted circuit occurs. If the circuit breaker of the faulted circuit fails to trip, the main circuit breaker will provide backup protection after the prescribed time delay. The time delay is on the order of 0.5 s and can be achieved by pneumatic or electronic means. The breaker trip device may be either a shunt trip or UVR, but the UVR is preferred. Potential relaying can also be used as backup protec- tion in conjunction with a main circuit breaker. Unlike zero-sequence and neutral relaying, potential relaying has the advantage of being able to detect a ground fault with the neutral grounding resistor open. Figure 12.19A shows a potential-relaying scheme using a PT for obtaining a voltage within the operating range of the relay whereas figure 12.19B shows a voltage divider to provide the same function. For both schemes, the maximum pickup voltage of the relay should correspond to the voltage developed across the relay at 40% of the maximum fault current. When using a voltage divider circuit, care must be taken in sizing the resistors so the total equivalent resistance of the parallel combination is not reduced significantly. The combined value (R x + R 2 ) must be significantly higher than the value of the neutral grounding resistor. The impedance of the relay should also be significantly higher than the resistance of R 2 for the same reason. The power (I 2 R) rating of the resistors should be capable of withstanding the currents expected through them under a maximum ground-fault condition. Figure 12.19C illustrates a popular alternative to backup relaying. Here the relay coil is replaced by a red warning light located in a rugged or explosion-proof hous- ing mounted on the outside of the power center. The light I r r r III UVR •— Wv *• To control transformer Main circuit breaker Figure 12.18.— Neutral relaying applied to grounding- resistor current as backup protection. bulb voltage is matched to the PT to give maximum brilliance under maximum ground-fault conditions but noticeable light output at 40% of current limit. Many engineers prefer this technique as it provides an obvious indication of ground fault and is especially useful when primary ground-fault relaying is inoperative either through tampering or malfunction. Test circuits for ensuring pickup at 40% of the maxi- mum ground-fault current can be incorporated into the power center: figure 12.20 is an example. A control trans- former with a secondary voltage of 12 V is used as a Neutral grounding resistor To circuit breaker trip A Potential transformer Neutral grounding resistor _L To circuit nr breaker trip B Resistive voltage divider Neutral grounding resistor -Ground-fault indicator C Warning light Figure 12.19.— Backup protection devices associated with mine power centers. Line conductors CT To control transformer 1 To circuit breaker trip 120/12 V 50 VA 6 turns GTR Figure 12.20.— Typical test circuit for zero-sequence relay- ing. 316 current source. The 12-12 resistor is inserted in the circuit to limit the secondary current to 1.0 A. The six turns through the CT at 1.0 A produce the same effect as 6.0 A of zero-sequence current for the power conductors. The current corresponds to 40% of a 15-A system so that depressing the test button simulates a ground fault at 40% of the maximum current. The associated circuit breaker should then trip. SINGLE-PHASE TRANSFORMERS ■« — -EZD ;> To 480 V K -* -n n — -> 240/480 V 50 VA 120 V To ground monitor Figure 12.21.— Simple control circuit incorporating one ground-fault relay (GTR) and one ground-check relay (GCR). Single-phase transformers are used in power centers to supply 120 V to the control circuit and 240/120 V to convenience outlets. The control circuit consists of ground- monitoring systems, undervoltage releases, and ground- fault circuitry for each associated machine circuit, as well as relay connections for other protection devices. Conve- nience outlets can be used for portable power tools, area lighting, or external test instrumentation. Control trans- former capacity normally ranges from 50 to 500 VA, whereas transformer capacity for 120/240-V outlets often falls in the range of 5 to 10 kVA. Larger capacities may be found, particularly in mine power centers with extensive control circuitry. A simple control circuit and a convenience-outlet circuit are shown in figures 12.21 and 12.22, respectively; in both circuits, fuses are used to protect the transformer primaries. Circuit breakers can be employed to protect convenience-transformer primaries, but fuses are recommended for control-power circuits be- cause breakers can be tripped by an unwary miner, thus deactivating the load center. Control-power fuses can be mounted in insulated dead-front holders or in a typical spring-clip arrangement that is only accessible to authorized personnel through a bolted cover. These two mounting types are illustrated in figure 12.23. The testing and changing of control-circuit fuses are among the more frequent electrical maintenance procedures for all types of equipment. Typical spring-clip fuse holders have uninsulated exposed metal clips, and rushed repair persons frequently remove the fuses from these holders for testing without deenergizing the circuits, thus placing their hands within close proximity of the energized clips. A slight inadvertent movement of the hand can result in electric shock. With dead-front fuse mounting, however, all energized components are enclosed in an insulated housing so that the fuses can be removed and replaced without exposing the electrician to the metallic clips and fuse ends. A variety of schemes exist for handling the control- and utility-circuit voltages. Each machine circuit can have its own individual control transformer, but it is more common for a single control transformer to supply all the control circuits. When all control circuits are supplied by a single source, the control-circuit transformer can be elim- inated, with the control voltage supplied by the secondary of the convenience-outlet transformer. Many mine power engineers dislike this combination because a failure or misuse of a convenience-outlet circuit could result in blowing the transformer fusing. This in turn would cause the loss of control power and all power to the mining machinery would stop. For this reason, some engineers prefer not to incorporate any 120/240-V outlets in their load centers, even with separate transformers. 240-V receptacle 120-V receptacle 480V-120/240-V 5kVA 15A Figure 12.22.— Simple convenience-outlet circuit for 120- or 240-V single phase. Dead-front panel X Insulated fuse housing Typical mounting Dead-front mounting Figure 12.23.— Fuse mountings. METERING CIRCUITS The mine power center should include metering cir- cuits to monitor line voltages and currents of all three phases (fig. 12.2, Nos. 15-16). The built-in instrumenta- tion is an invaluable aid to maintenance, as it allows a firsthand look at the electrical operation of the load center and gives a composite view of the loads it serves. The metering is usually for distribution or utilization voltages and current, but rarely for both. Figure 12.24 illustrates a common approach for me- tering voltage, where potential transformers (PT's) are used to reduce the line voltage to a value within the rating of the voltmeter potential coil. Two single-phase trans- formers are connected in open delta, and a four-position switch allows the three line-to-line voltages to be moni- tored in addition to the o/f position. As shown in figure 12.25, two CT's are sufficient for metering the current in all three lines and isolating the meter circuit from the line. The CT ratio should be as low 317 600 -V bus Voltmeter switch ^Hf V 0-750 V __^2_3_4 g-b X X b-c X X c-a X X 750:150 25 VA Figure 12.24.— Typical metering circuit for line-to-line voltages. ■H| — LZZr 1 2 3 4 - X X a X X - X X X b X X - X X c X X - X X X OFF X X 1,500:5 a b c "=■ Figure 12.25.— Typical metering circuit for line currents. as possible, without exceeding rated current in the second- ary winding. A common recommendation is that the ratio be such that normal CT secondary current is 0.5 to 0.75 of the full-scale rating of the meter (5). The CT turns ratio can cause dangerously high voltage on the secondary if it is opened, so the ammeter switch must be designed so the secondary is not open circuited during the transition from one switch position to another. These are known as short- ing switch contacts. OUTGOING CABLE COUPLERS The trailing cables for utilization equipment are al- most always connected to the power center by means of low-voltage or medium-voltage cable couplers (fig. 12.2, No. 17). The couplers are rated at 600 V for low- voltage operation and at 1,000 V for medium voltages. Standard current ratings are 100, 225, 400, and 800 A, but larger sizes are readily available. Additional information can be found in the coupler section of chapter 8. The outgoing cable couplers on the power center are female-contact receptacles, which connect with the male- contact plugs of the trailing cables. When disconnecting, the opening of contacts must follow a sequential order. The couplers serve as a visual disconnect for the low-voltage and medium-voltage circuits served by the mine power center, and for safety the couplers should always be locked out when not in use. Lockout is the process of deenergizing a circuit so that it cannot be energized without authority. Improper lockout or failure to lock out has been a leading cause of electrical accidents. For example, a repair person deenergizes a faulty circuit at the power center prior to performing repair work. While the repair person works on the faulty circuit, another worker who is unaware of the situation mistakes one cable for another, energizes the faulty circuit, and subjects the repair person to electric shock or electrocution. Effective lockout can be provided by routine use of locking dust covers or keyed couplers. With keyed couplers, the receptacle of each outgoing circuit is matched to fit only one cable-mounted plug so that mistakes cannot be made. On the power center, the dust covers are hinged to the receptacle or connected by a chain. The hinged covers are preferred from a safety standpoint since they are less likely to become disengaged and lost. Locking dust covers on cable-mounted plugs offer even better protection, as once the plug is locked it is impossible to connect it to any receptacle. The covers also prevent damage and dirt accu- mulation when the plugs are not in use. Although chain- connected locking dust covers are the only type presently available for plugs, hinged lids would again offer improved safety. Another lockout technique is to drill a hole in the spare pilot-circuit pin of the cable-mounted plug. A pad- lock can then be inserted in the hole by the electrician. This technique is extremely simple and inexpensive, but it provides excellent lockout protection by prohibiting plug- receptacle coupling. The installation of two receptacles for each outgoing machine circuit can also provide lockout protection. One receptacle is used for powering the outgoing circuit while the other is used strictly for locking out the circuit when necessary. Actually, the lockout receptacle is not a true receptacle since its housing need not contain internal contacts. The lockout receptacle can also protect the plug from damage while the associated circuit is not in use. GROUND-CHECK MONITORS Low-voltage and medium-voltage resistance-grounded systems are required to have a fail-safe ground-check circuit to continuously monitor the continuity of the grounding conductor. The monitor must cause its associ- ated circuit breaker to trip if the grounding conductor or a pilot wire is broken. An indicator lamp on the monitor should indicate a tripped condition. The monitors are usually enclosed in a dead-front package and mounted near the associated circuit breaker. As was seen in chapter 9, monitors in common use in mining can be divided into two general classifications: impedance types and continu- ity types. Impedance monitors require the trailing cable to have a pilot conductor. The monitor is calibrated to the imped- ance of the loop formed by the pilot and grounding conductors, and the device then monitors the change of impedance from the initial calibration. If the impedance of 318 the loop increases beyond a preset value, the monitor must trip its associated circuit breaker by opening a set of contacts in series with the undervoltage release. The maximum allowable increase in impedance is dependent upon the maximum ground-fault current per- mitted by the system. For a 15-A neutral-grounding resis- tor, the monitor should cause tripping if the impedance increases by 2.7 ft. This value is based upon the maximum allowable frame-to-ground potential of 40 V, as follows: 40 Z = -f* m ■ From the above, it is apparent that a ground-check moni- tor does not ensure that the frame potential of a piece of equipment will not rise above 40 V, since the device monitors only the change in impedance of the pilot and grounding-conductor loop and not the actual impedance of the grounding conductor. A schematic of a common impedance monitor is shown in figure 12.26. This particular monitor is powered from a 24- to 32-V source, but others use 120-V power. Out-of-phase induced currents during motor starting can result in cancelling out the monitoring current, which in turn can cause nuisance tripping. As a result, a polarity- reversal switch should be provided to change the phase relationship of the pilot current with respect to the in- duced current. Some manufacturers use impedance- matching transformers to amplify the change in imped- ance for easy detection. The monitor should also provide a test button. With the button depressed, the appropriately sized resistor is inserted into the pilot circuit, which should result in tripping the circuit breaker. A disadvan- tage of the impedance monitor is that it cannot detect a pilot-to-ground fault. This type of monitor is also suscep- tible to problems with parallel paths. Continuity monitors do not monitor impedance change, but only the grounding-conductor continuity. However, they must be adequately immune to parallel paths. Continuity motors are audio units and do not require a pilot conductor for operation. Figure 12.27 contains a block diagram of a common unit. Most makes can also be wired for pilot operation (fig. 12.28), but only the operation of the pilotless configuration will be discussed. The monitor generates an audio frequency, which is coupled to the grounding conductor by means of the transmitting coil. The pilot wire is eliminated by using the line conductors as a return path. Filters at the monitoring location and within the monitored machine are necessary for coupling and uncoupling the audio signal from the line conductors. If the grounding conductor is intact, the re- ceiver coil picks up the audio signal. If the grounding conductor is open, the receiver coil will pick up no signal, and the monitor will cause a set of contacts to open the undervoltage circuit of its associated circuit breaker. Chapter 9 contains a discussion of a continuity monitor that operates in a somewhat different fashion; the signal is r i/i/hmi. Frame ground ■g Large coupler I" Line Line 2 Transmitter coil Line 3 Frame ground -PP£ =>PPr Grounding conductor Small coupler "1 Frame ground I3J Receiver coil MONITORED MACHINE Ground monitor To breaker trip circuit To control | power(l20Vac)l r_j POWER-CENTER DISTRIBUTION BOX OR OTHER Figure 12.27.— Block diagram of continuity monitor con- nected in pilotless mode. 32 Vac Polarity switch Transformer Al LI Deenergized 3.0-11 resistor B 0@ Figure 12.26.— Typical impedance monitor circuit. r Transmitter coil rrn ~l Frame ground Pilot conductor Receiver coil Frame ground =PP>- Groundmg conductor Frame ground Ground monitor I To breaker trip circuit I I MONITORED MACHINE To control power (120 Vac) SUBSTATION, POWER-CENTER DISTRIBUTION BOX, OR OTHER Figure 12.28.— Block diagram of continuity monitor wired for pilot operation. 319 impressed on and removed from the line conductors, with the grounding conductor used for the return path. The major advantage of the continuity monitor is that it is immune to parallel paths and stray currents. It is also immune to pilot-to-ground faults if a pilot conductor is not used. However, the continuity monitor is expensive and complex when compared with an impedance monitor. A pilotless monitor must still be wired into the pilot and grounding contacts of the coupler in order to trip the circuit breakers when disconnecting the coupler. Another problem is that the grounding conductor must be isolated from the coupler shell, or the intended monitoring is bypassed. For metallic shells, a separate grounding con- ductor must be supplied through a spare coupler contact. POWER-FACTOR CORRECTION Some mining machines have notoriously poor power factors resulting from underutilization of induction mo- tors. Perhaps the most outstanding example is the contin- uous miner, which can have a power factor that averages 0.6 lagging during the operational cycle. Whether it is this machine or others that create excessive reactive power, the result is poor power-system efficiency and utilization. If the power factor is poor at the purchase points (under 0.80, for example), the utility company will attach a penalty to the power bill. There are many opinions on the best location at which to improve power factors. The most obvious place is at the utilization points. This strategy is common in surface mines, particularly with excavating machines. However, surface mining equipment has adequate space to incorpo- rate the correction equipment, an advantage not common in underground mining machines. If the improvement is made at a convenient location immediately upstream from offending motors, the entire system up to and including the substation can benefit through lower apparent power, reduced line currents, and better voltage regulation. This location is commonly on the utilization buswork in mine power centers. On the other hand, if correction is not attempted here, power factors will be increased by the inherent distribution-system capacitance (for example, that associated with shielded cables) or somewhat can- celled by the diversity of operation that always exists between mining machinery and equipment sections. As a result, some engineers would rather correct the overall power factor at the distribution side of the substation. Consequently, three general locations can be consid- ered if power factors need correction: the machine, in the power centers, or at the substation. Each individual mine power system must be analyzed to find the solution, and the final decision must consider both electrical operation and economics. One consideration should be the cost of load-center correction versus correction at the substation versus the penalty added to the power bill. If a decision is made to add correction, the common approach is to use a bank of capacitors, also known as a static bank. (Synchro- nous rotating machinery is not logistically practical for most cases.) Precautions must be taken when applying correction- capacitor banks, regardless of the location. They are a system load, and thus the bank must not be grounded on resistance-grounded systems. In the past, the common capacitor insulation was PCB (polychlorinated biphenyl), but because of the environmental hazards associated with this material, capacitors containing PCB must no longer be used. Owing to the dynamic operation of mining equipment, the correction should never be designed to obtain unity power factor on an average basis. If this were done, the resulting power factor could be leading (capaci- tive) for a significant portion of operation time, and the overall system operation could be worse than if no correc- tion had been attempted. The last precaution concerns resonance. There is a chance that undesirable resonance could be created if a resistance-capacitance-inductance (RCL) combination is formed by adding capacitance from the static bank to the system portion. This problem has been observed particularly with power centers (belt trans- formers) supplying solid-state belt starters. It manifests itself as voltage and current frequencies other than 60 Hz. Figure 12.29 shows the application of power-factor correction in a mine power center. The schematic is basically a reproduction of figure 12.2 with the addition of items 18 and 19, a static capacitor bank and its associated protection, respectively. Three capacitors connected in ungrounded wye are shown, but an integral three-phase capacitor with the same configuration can also be used. A molded-case circuit breaker affords short-circuit protection but also acts as a switch for removing the bank from the line. Sizing this breaker follows the same procedure re- lated earlier for other applications and can be based on normal capacitor current. The current can easily be calcu- lated from the per-phase capacity (kilovoltamperes reac- tive) and the system voltage. Some load centers are de- signed with more than one capacitor bank so that the magnitude of correction can be changed. The sizing of the capacitor is also straightforward and follows the basic procedures presented in chapters 3 and 4. For instance, consider a mining section that has been found to be consuming 172 kW at an average power factor of 0.6 lagging. The average apparent and reactive powers are then S = P pf 172 0.6 = 287 kVA, Q L = S[sin (cos -1 pf)] = 287 (0.8) = 230 kvar. With a 90-kvar capacitor bank, the resultant reactive power and new power factor would be Q T = Q L - Q c = 230 - 90 = 140 kvar inductive, Q T 140 Ian u-jy — -p = v7o> ^t = *-*" » pf = cos"^ = 0.78. Thus, the average power factor has been improved from 0.6 to 0.78. As a general recommendation, power-factor cor- rection in mine power centers should not try to correct an average power factor above 0.85. Figure 12.29 also serves as a summary of the material presented on ac mine power centers. The internal compo- nents of an ac mine power center can be grouped into those associated with the high-voltage side, the transformer, and the utilization side. 320 KEY High- voltage coupler Interlock switches Emergency-stop switch Disconnect switch Pilot -break monitor High-voltage fuses Surge arresters Surge capacitors Power transformer Temperature device Grounding resistor Busway Outgoing circuit breaker Main circuit breaker Voltage metering Current metering Outgoing cable coupler Power-factor correction capacitors Circuit breaker To utility transformer and control circuit Figure 12.29.— Application of power-factor correction in mine power center. DIRECT CURRENT UTILIZATION Although ac utilization now dominates the mining industry, some segments of the mining process still remain better suited to the use of dc motors. The classic example is the use of dc series-wound motors for traction. The speed-torque characteristics of these motors are particu- larly well suited to this application, especially for locomo- tives and shuttle cars. This is the reason why dc utilization still plays a significant role in the mining process. There are various ways of obtaining a dc voltage for powering the face equipment. If a mine uses rail haulage at a working section, the dc equipment can be powered directly from the trolley feeder line. However, if a working section does not have direct access to a trolley feeder line, a rectifier must be used to convert the three-phase ac voltage to dc voltage. The rectifier can be a separate piece of equipment housed in its own enclosure and powered by means of a feeder cable from the ac power center, or it can be incorporated into a single enclosure with the ac power center, with the total unit being referred to as an ac-dc combination power center. Some mining machinery man- ufacturers offer face equipment with on-board rectifiers so the benefits of dc motors can be obtained without the need for a section rectifier (see chapter 14). Even if the rectifier is supplying a trolley, the basic internal components of dc power equipment remain essen- tially the same, and so only the dc section of a combination power center will be discussed. A general arrangement of the dc components is illustrated in figure 12.30. This diagram will be used as a reference to indicate the placement of individual components with respect to the y £ To control circuit Figure 12.30.— General arrangement of dc components for combination power center. 321 overall system. The dc circuits that are frequently associ- ated with trolley feeder lines will also be mentioned. Components and circuits that have been discussed in the preceding sections will be presented only when they are required for clarity. RECTIFIER TRANSFORMER A three-winding transformer (fig. 12.30, No. 1) is commonly used in a combination power center and con- sists of primary, secondary, and tertiary windings. Nor- mally, the primary winding is delta connected with the secondary and tertiary windings wye connected, but the transformer may be designed such that the primary wind- ing can be connected in either a delta or a wye configura- tion so it can be applied at two different distribution voltages: for example, 4,160-V delta, 7,200-V wye. For the latter situation, either the secondary or the tertiary wind- ing must be delta connected. Thus, if a resistance- grounded system is to be used on the delta-connected winding, a zig-zag or grounding transformer must derive a neutral. The full-wave bridge (fig. 12.31) is the most popular rectifier configuration used in combination power centers or section rectifiers. Here the relationship of the ac rms input voltage to the dc output voltage is V ac = 0.74 V dc . The nominal 300-V system voltage is the one most com- monly used in mining applications. Although some 600-V systems are still in operation, they are usually used only for trolley supplies. For a nominal system voltage of 300 V, the line-to-line ac voltage feeding the input to the rectifier would be »V in V f .360-Hz ripple Figure 12.31.— Full-wave bridge rectifier. Air-core reactors Main circuit breaker ¥ ¥ ¥ To rectifier Figure 12.32.— Series reactance to reduce available short- circuit current. V ar = 0.74 (300) = 222 V. Hence, the voltage rating of the transformer winding feeding the rectifier would be 222 V. Some of the common ratings used in combination power center are listed in table 12.2. Table 12.2.— Typical ratings for combination power centers. dc, kW ac, kVA ac voltage, V dc voltage, V Input voltage, V 75 300 480Y/277 300 7,200 150 375 480Y/277 300 4,160 400 480Y/277 300 7,200 200 600 600/220 300 7,620/13,200 300 500 480Y/277 300 7,200 750 600/220 300 7,620/13,200 1,000 600/220 300 7,200/4,160 A dc is much more difficult to interrupt than an ac, and the amount of available short-circuit current must be limited to less than that of ac equipment with similar capacity. As a result, the impedance of transformer wind- ings feeding rectifier circuits is normally in the range of 0.06 to 0.08 pu. In order to obtain this high impedance, the leakage reactance of the associated winding is often delib- erately increased during transformer fabrication. If this method does not provide the adequate impedance, a small additional impedance can be obtained by using air-core reactors at the output of the transformer, as shown in figure 12.32. However, in some circumstances, it is neces- sary to use a separate transformer (fig. 12.33) to limit the Main circuit breaker (ac) To ac circuits Figure 12.33.— Separate transformer to increase impedance of dc circuit. 322 maximum fault current to a level that can be safely interrupted and a magnitude that will not damage the i-ectifier bridge. A main circuit breaker (fig. 12.30, No. 2) can be provided to protect the transformer winding that supplies the dc circuit. The thermal-trip setting of this breaker is usually based upon 125% of the rated current of the winding. As an example, a winding rated at 150 kVA would have a full-load current of 150 FL " V3 (0.222) = 390 A. The thermal-trip setting would be the next rated value above 1.25 (390), or 488 A. If the transformer impedance is 0.06 pu, the breaker should be capable of interrupting a fault current of If = 390 0.06 = 6,500 A. RECTIFIER As seen in figure 12.30, the output of the transformer tertiary winding feeds the rectifier (No. 3) via the main circuit breaker. Again, the rectifier configuration is almost always a three-phase full-wave bridge. The power or kilo- watt requirement of rectifiers used for mining is normally too high to permit the use of a single diode in each leg of the bridge. The current rating of the rectifier bridge, and thus the diodes, is selected on the expected bolted-fault current (faulted dc output) with the power-center input connected to an infinite bus. This current should also incorporate a multiplier to account for a worst case offset. The rectifier should be capable of handling this current for the time required for the circuit breakers to interrupt current flow. To achieve the adequate current-carrying capacity, diodes are paralleled in each rectifier leg to share the current. Figure 12.34 shows a full-wave bridge rectifier with two diodes paralleled per leg (No. 1). The number of parallel diodes is determined by the individual diode current rating (all must be rated equally) along with the maximum available fault current. Each diode must also be rated in terms of a peak inverse voltage (PIV), which is the maximum voltage that can be applied across the diode in a reverse-biased mode without causing breakdown. For mine rectifier applications, the diode PIV should not be less than 2.5 times the nominal dc system voltage if proper transient suppression is used. Thus for a 300-V system, a rating of 800 V is commonly employed. Matched diode characteristics should allow sufficient means for sharing current when used in a parallel config- uration. Replacement through failure of any diode, how- ever, would easily upset the balance. Therefore, although rectifiers are supplied with matched diodes, most manu- facturers use current-balancing reactors (fig. 12.34, No. 2) to force uniform conduction of all diodes in parallel. (See chapter 4 for an explanation of operation.) Each diode is accompanied by a circuit-protection fuse (fig. 12.34, No. 3). The purpose of the fuse is not to protect the diode but to prevent a catastrophic failure to a bridge leg (these diodes fail in a shorted mode). The fuse current rating must be such that it will not initiate interruption for a bolted fault at the dc bus until the clearing time for the downstream circuit-interrupting devices has elapsed (that is, the time-current characteristics of the two must be coordinated). Each fuse has a matching light (fig. 12.34, No. 4) to indicate when the fuse has blown and the diode needs to be replaced. Suppression devices should always be used to protect the rectifier from transient overvoltages occurring from either the ac or dc side. Owing to their successful history in mining applications, selenium voltage suppressors (fig. 12.34, No. 5) are presently the most popular devices used in rectifiers. They are available with rms ratings ranging from 30 V to 480 V in 30-V increments. The clamping voltage at rated discharge is approximately 2.5 times the rms voltage rating. For instance, the input ac voltage to a 300-V rectifier would be 222-V rms; therefore, a selenium suppressor rating of 240- V rms should be selected. The resulting clamping voltage would be 600 V, which is significantly below the 800-V PVI rating determined earlier. Although not shown in figure 12.34, the suppres- sor often has each connection (five in all) protected by a fuse, again with a paralleled light to indicate failure. For rectifiers serving face equipment (150 or 300 kW), one selenium suppressor is used, but two or more may be paralleled for trolley rectifiers (450 kW or more). An RC-snubber circuit, as shown in figure 12.35, is sometimes provided to reduce commutation transients. Figure 12.34.— Typical full-wave bridge rectifier with two diodes in parallel per leg. Diode Circuit- protecting I 2 t fuse Hvwv- -a- dv/dt snubber circuit Indicator light Figure 12.35— Diode with RC snubber protection. 323 The circuit serves to reduce the voltage rate-of-rise (dv/dt) that might be developed across the diode. Although snub- ber circuits are common in mine rectifiers, some manufac- turers have achieved reliable diode operation without their use. The application of this protection technique is discussed in chapter 14 as dv/dt protection is required on thyristors. DIRECT CURRENT GROUND-FAULT PROTECTION SCHEMES Five basic systems of dc ground-fault protection are presently being used in the United States: 1. Diode grounding, 2. Basic grounding conductors, 3. Relayed grounding conductors, 4. Neutral shift, and 5. Differential current. These are illustrated respectively in figures 12.36 through 12.40. The technique used dictates the complexity of the circuitry that will be associated with each outlet from the rectifier section. Because a complete discussion is pre- sented in chapter 8, the systems will only be summaried here. The first three systems are more commonly used when a trolley feeder is the dc source. The neutral-shift and differential-current systems can only be used in con- junction with rectifiers serving face equipment. It should be noted that rectifiers for trolley systems must have both positive and negative conductors isolated above frame ground; thus, the subject of grounding is not applicable in that instance. In the diode-grounded system (fig. 12.36) the machine frame is tied to the grounded conductor through a ground- ing diode. The basic grounding-conductor system (fig. 12.37) uses a separate conductor to provide grounding to the machine frame. In essence, adequate grounding-fault protection is not available from either system; for example, Machine frame Figure 12.36.— Diode-grounded system. (+) (-) -o o — ^ (+) (-) GND Machine frame ground faults in the trailing cable rely on an interrupting device for sensing and clearing. No additional circuitry, other than the grounded power conductor, is used in the power center. The relayed grounding-conductor system (fig. 12.38) can be sensitive to ground faults. A ground-fault relay coil is placed in series with the grounding conductor, and when a current threshold is reached, contact pickup can be used to trip a breaker UVR. However, parallel paths existing through the earth can cause fault current to partially bypass the relay coil, making for unreliable protection. The neutral-shift system is shown in figure 12.39, where the resistors R x and R 2 create a dc neutral point. Ground faults on either the positive or negative conductor will cause the neutral to shift. The voltage-sensing relays, M 1 and M 2 , detect the change. The result is sensitive ground-fault relaying, but the system is not selective with more than one outgoing circuit. For one type of differential-current system (fig. 12.40), a grounding resistor is placed between the transformer -o o — /- (+) =-o (-) iA GND s?=-o GTR Machine frame n To circuit breaker trip circuit Figure 12.38.— Relayed grounding-conductor system. To rectifier . M, ± To circuit breaker trip circuit (+) R 2 M?) ± (-) GND = — o To circuit breaker trip circuit Figure 12.39.— Neutral-shift system. To circuit breaker trip circuit Grounding resistor u Saturable reactor \ 1 — < . i > — < i i 1 — 1 ; 1 i 1 J™ • 1 i i i / i rO ► Ground- fault relay , (+) r , gndT, To ac control circuit Machine frame Figure 12.37.— Basic grounding-conductor system. Figure 12.40.— Differential current scheme. 324 neutral and load-center frame and both positive and negative conductors pass through a saturable reactor. The differential current created by a ground fault causes magnetic saturation of the reactor core, which in turn allows the ground-fault relay to pick up. The system is sensitive and selective and by using the technique for each outgoing dc circuit, practical ground-fault protection is available. As a result, this system is considered in more detail in the following discussion. DIRECT CURRENT CONTROL CIRCUITRY A control circuit using a differential-current tech- nique for ground-fault relaying is illustrated in figure 12.41. A main control transformer supplies 120 V to the entire circuit. Heat-sensing devices are mounted to the heat sink for each leg of the bridge rectifier. The overtem- perature relay (OTR) coil is connected in series with the heat-sensing devices and is energized by a filtered single- phase bridge rectifier. The relay contacts are in series with the UVR of the main breaker so when a rectifier overtem- perature condition occurs, a sensing device opens, the relay resets, and the UVR trips the breaker. This version of differential-current relaying is sup- plied from the control power through a stepdown trans- former with a 6-V secondary. With a ground-fault current from 4 to 6 A, voltage is impressed cross the ground-trip relay (GTR) in sufficient magnitude to cause contact pickup. This deenergizes the UVR on the machine circuit breaker, which in turn interrupts the offending circuit. DIRECT CURRENT INTERRUPTING DEVICES The problem mentioned earlier concerning interrup- tion of dc requires the interrupting device to force the current to zero with an arc voltage greater than the line voltage. The device must be capable of withstanding the energy dissipated during the time that the arc exists across its contacts. This energy is a function of the circuit inductance as well as the current magnitude (3, 14). There are two basic interrupters used to protect dc machine circuits: the dc contactor and the molded-case circuit breaker. For trolley systems, power circuit breakers are employed. A cross-sectional view of a dc contactor is shown in figure 12.42. This device consists primarily of stationary and movable contacts, which make and break upon ener- gizing and deenergizing the operating coil. When the operating coil is energized, the armature assembly is attracted to the pole piece, which completes the magnetic circuit and causes the contacts to close. The contactor is designed so that the contacts close with a self-cleaning wiping action to ensure a good electrical contact (6). Arc termination is similar to that of an air-magnetic circuit breaker (see chapter 9). A blowout coil is located between the front power terminal, Z, and the stationary contact. The purpose of the coil is to aid in extinguishing the arc and to minimize burning of the contact tips. To accomplish this, the coil produces a magnetic field perpen- dicular to the arc, which causes the arc to be lengthened and ruptured by deflection. Since the blowout coil is connected in series with the circuit to be interrupted, the strength of the magnetic field is proportional to the current. A flexible braided shunt completes the connection ■— o- '^mcid- Control transformer Lo^«- Hi OTR (49 R) ik il Rectifier (+) ( Ti * -• — •- l| ' V . , II Heat-sensing devices dc bus Figure 12.41.— Representative control circuit for rectifier. Top insulator Lo "f / Operating x insulator £ oi | a Top plate Blowout coil Arc horn Arc shield-* Stationary contact tip Stationary contact assembly Shunt Spring washer Armature hinge pin Lock screw Movable contact tip Armature assembly Moving contact assembly Figure 12.42.— Cross section of dc contactor. (Courtesy Joy Manufacturing Co.) 325 from the movable contact to the rear power terminal, Y. The arc horn protects the insulation of the blowout coil from being burnt by the arc, and an arc shield encloses the contacts to confine the arc and protect the adjacent parts (6). The contactors most commonly used as interrupting devices for dc machine circuits have a continuous-current rating of 250 or 500 A. Their interrupting capability is approximately 10 times their continuous-current rating, which is significantly less than that given for molded-case circuit breakers. One advantage of the dc contactor is its simple design and ease of maintenance. In some applica- tions, molded-case circuit breakers are used as backup protection for dc contactors. It should be noted that molded-case circuit breakers are not rated for 600-V appli- cations, but dc contactors can be applied at this voltage. Molded-case circuit breakers that are rated at 600 Vac are also normally rated at 300 Vdc. The dc interrupting- duty rating for mine-duty breakers is either 10,000 or 20,000 A, depending on the product. However, as with the dc contactor, the actual interrupting capability varies with the inductance of the circuit being interrupted. Labora- tory tests have indicated that for dc operation, the instan- taneous dc tripping current is about 1.3 times the magnetic-element trip setting for ac (14). Even though molded-case breakers have higher inter- rupting capacity, many engineers in mining operations have expressed concern over using them for dc applica- tions. The engineers relate that in practice molded-case devices cannot adequately interrupt dc faults, whereas low-voltage power circuit breakers appear to work satis- factorily. On the other hand, manufacturers report that they have had few problems with field applications of molded-case breakers on dc. In summary, the principal understanding that should be acquired from this power center presentation is that the equipment is assembled to match the mining operation and application. This is true whether the unit is a belt transformer, a trolley rectifier, or a combination ac-dc power center. Consequently, there cannot be a standard mine power center, but nevertheless, there are many practices recommended for mine power center construc- tion that must be considered in order to achieve an efficient operating system. REFERENCES 1. Dornetto, L. D. The Importance of Grounding Systems in the Protection of Personnel and Equipment. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 2. Dutton, J. C, and W. J. McNutt. Transformer Stan- dards-Status and Trends. IEEE Trans. Ind. Appl., v. 16, Jan./Feb. 1980. 3. Helfrich, W. J., P. M. Hall, and R. L. Reynolds. Time Con- stants-Direct Current Mine Power Systems. Paper in Proceedings of 5th WVU Conference on Coal Mine Electrotechnology (Morgan- town, WV, July 1980). WV Univ., 1980. 4. Institute of Electrical and Electronics Engineers (New York). General Requirements for Dry-Type Distribution and Power Transformers. ANSI/IEEE Stand. C57.12.01-1979. 5. Recommended Practice for Electric Power Distribu- tion for Industrial Plants. Stand. 141-1986. 6. Joy Manufacturing Co. (Franklin, PA). Direct Current Min- ing Machinery. 5th ed., 1971. 7. King, R. L. Development of an Electrical Engineering Course for Mining Engineers. M.S. Thesis, Univ. Pittsburgh, Pittsburgh, PA, 1977. 8. Kline, A. D. Design Consideration of Mining Transformers. Paper in Conference Record -IAS 15th Annual Meeting (Cincin- nati, OH, Sept.-Oct. 1980). IEEE, 1980. 9. Line Power Manufacturing Corp. (Bristol, VA). Electrical Power for the Mining Industry. Sales literature, undated. 10. McGraw-Edison Co. (Canonsburg, PA). Application catalog 240-60, undated. 11. McGraw-Edison Co., Power System Div. (Canonsburg, PA). Distribution-System Protection Manual. Bull. 71022, undated. 12. Ohio Brass Co., Ensign Electric Div. (Mansfield, OH). Mine Power Equipment. Sales literature, BHE37, undated. 13. RTE Corp. (Waukesha, WI). Transformers. Undated. 14. Shimp, A. B., and D. A. Paice. Application of Molded-Case Breakers on DC Electrical Systems in Coal Mines. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 326 CHAPTER 13.— SWITCHHOUSES AND SUBSTATIONS 1 Switchhouses, substations, and power centers comprise the major power equipment in mine power systems. Chapter 12 covered the design and construction of mine power cen- ters, and this chapter will follow a similar format to present switchhouse and substation arrangements. SWITCHHOUSES associated control circuitry for figure 13.1 is given in figure 13.2. The operation of this circuit will be discussed later. As in mine power centers, interlock switches are positioned around the side covers and top covers of the switchhouse and are wired into the incoming pilot circuit to trip the upstream breaker in the event that a cover is removed. An emergency stop button is also provided. Switchhouses are contained in metal-clad enclosures, similar in construction to those of power centers. This power equipment consists of visible disconnects and sectionalizing units, although the term switchhouse is normally equated with just the sectionalizing equipment. The principal func- tion of the disconnect switch is to remove power manually from downstream mine power equipment that is connected to distribution. As in power centers, it must be possible to determine the position of the disconnect (load break) switch visually through a window in the enclosure wall. In under- ground coal mines, the power removal function must be located within 500 ft of the point where power enters the mine, and additional disconnects are incorporated into all power equipment that receives power from high-voltage distribution. This disconnect-switch equipment can be de- scribed simultaneously with switchhouses. The high-voltage side of the power center that was shown in figure 12.2 contained the same components. The prime role of the switchhouse is to provide pro- tective relaying in the distribution system and to allow branching of the radial system. The principal component is an automatic circuit breaker. The equipment name is modified depending upon the number of circuit breakers and associated protected outgoing circuits: for example, a double switchhouse contains two breakers and circuits. Because of size limitations, the units in underground mine power systems are rarely larger than double switchhouses, but surface mines often incorporate four-breaker switch- houses, or switching skids as they are commonly called. The schematic diagram for surface or underground appli- cations is practically the same: the only basic difference is the repetition of internal components to correspond to the number of circuit breakers used. SWITCHHOUSE INTERNAL COMPONENTS A general arrangement for a single-breaker switch- house is illustrated in figure 13.1. Incoming high-voltage power enters the switchhouse through the input receptacle to the load-break switch and the feedthrough receptacle. Some circuit-interrupting devices have disconnect switches incorporated in their construction, as shown in this diagram; others have a load-break switch located on the line side of the breaker. In either case, the switch must have an external operating handle with a mechanism for locking the switch in the open position. The output (load side) of the interrupting device feeds the branch receptacle. The current-sensing de- vices for the protective-relaying circuits are usually situated between the breaker and the branch receptacle, while the control transformer is located on the incoming side. The 1 The author wishes to thank Thomas Novak who prepared original material for many sections of this chapter. Branch x receptacle _± To line ^ ground- overcurrent check relays and monitor ammeter Cover Emergency interlocks stop Figure 13.1.— Diagram for typical single switchhouse. I — ww- | ■ — WW- " — WW-' o -tpo ^/rm,^ 52 a 52 b i TgcsT GCS »To pilot — o — UV52 "UUUU Heater strips T 52a Q 52 trip coil Battery KEY I Instantaneous contact T Timed contact SI Seal- in contact K§h ISsT !-<§>. 'si X^ -W- X-%- GCS _i_ Trip button o o — 50/5H 50/51-2 50/51-3 51-G Figure 13.2.— Control circuitry for single switchhouse using battery tripping. 327 A typical arrangement of a double switchhouse is presented in figure 13.3, and its control circuitry is shown in figure 13.4. The figures again show breakers with incorporated switches, and the repetition of circuitry men- tioned earlier is obvious. In addition, the circuits in branch A work independently of branch B, and vice versa. Where the breakers do not have built-in switches, one load-break switch is usually located upstream from both breakers, and serves as a disconnect for all branch circuits, but not the feedthrough. The incoming circuitry to the circuit breakers, shown in figures 13.1 and 13.2, is nearly identical to the incom- ing circuitry to the power center transformer discussed in chapter 12 and shown in figure 12.2. Surge arresters and switch prebreak monitors are not incorporated in the switchhhouse schematics, but they are often found in practice and are recommended. A typical schematic for the disconnect switch (the equipment, not the component) would be similar to figure 13.1 if the circuit breaker and Input receptacle Cover Emergency interlocks stop ^rrrru Branch A receptacle rtt To ground- To phase trip relay overcurrent A relays-A JT TTl Branch B receptacle -, -rrm - To ground- To phase trip_relay overcurrent relays-B To control circuits AandB DC, Figure 13.3.— Diagram for typical double switchhouse. Control transformer I — VWV— | (i — v\M—-<> str 'P , , .... . , heaters 52 A/a Hi — 52A/b Closed 'M/W\rt -© ■• 50/51-3A 51-GA ■SI (SI) fa- — i¥- GCS-A KEY I Instantaneous contact T Timed contact SI Seal -in contact Trip button -A Trip button -B Figure 13.4.— Control circuitry for double switchhouse using capacitor tripping. 328 current transformers (CT's) were removed. This similarity is important in mine power-equipment design because when circuits serving identical functions have the same arrangement and construction, the maintenance crews can work on all equipment containing these circuits with- out surprise. This both simplifies mine training and increases the safety factor for maintenance personnel. SWITCHHOUSE PROTECTIVE RELAYING Induction-disk overcurrent relays are used for most protective-relaying applications in switchhouses. General relay operation and characteristics have already been discussed in chapter 9. For mining applications, inverse- time characteristics are usually employed for line- overcurrent relaying, while very-inverse-time relays are commonly used for ground-fault protection. These relays have a range of adjustments so they can be applied to a variety of situations: the operating time can be controlled by the time dial setting; pickup current is adjusted by the main coil taps. Since the resulting time- current characteristics are the same for each tap, curves are plotted in terms of multiples of the pickup value. A typical family of curves for an inverse-time relay is given in figure 13.5 (ll). 2 An instantaneous relay can be incor- porated with the induction-disk (timed) relay, which will respond to the same actuating quantity. The actuating currents that are delivered to overcurrent relays are ob- tained by CT's. The CT's provide insulation from the high-voltage circuit and supply the relays with currents that are proportional to the currents in the line conduc- tors, but sufficiently reduced. When applying a CT, critical consideration should be given to transformer burden and performance. Chapter 10 can be consulted for information on necessary calculations. The pickup of an overcurrent relay should be selected so it will operate for all short circuits within its primary zone of protection and provide backup protection in adjoin- ing zones. Figure 13.6 will be referred to as an example. Breaker 1 should provide primary protection for line 1. At the same time, its relays must be coordinated with the relays of breakers 2 and 3 to afford backup protection for lines 2 and 3. To ensure selectivity under all circum- stances, the pickup value of any given relay should be somewhat higher than the pickup of downstream relays. The time delay of overcurrent relays must also be adjusted to provide selectivity with the relays of the immediately adjoining zones. To provide proper selectivity, a bolted three-phase fault can be assumed as a basis for adjusting line- overcurrent relays, and a bolted line-to-ground fault would form the basis for the ground-relay setting. Selectivity under these conditions should ensure selectivity at lower currents. For the fault shown in figure 13.6, the relay at breaker 2 must close its contacts, and breaker 2 must trip and interrupt the flow of short-circuit current before the relay at breaker 1 closes its contacts. The calculation must also account for the overtravel of the relay at breaker 1. The time delay of breaker 1 that is needed to provide selectiv- ity with breaker 2 can be determined by i \ \ i i i | — i — i — i i i i i i i i i i i Ti = T 2 + B 2 + C-! + F, (13.1) 1 ''^L J. 1. '_ ' ' ' UU I I I 15 2 3 4 5 6 7 8 9 10 15 20 MULTIPLE OF PICKUP Figure 13.5.— Typical family of curves for inverse-time relay. Line I . — , . . Line 2 HZr¥ ^ 1 Fault Line 3 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. Figure 13.6.— Illustration of fault location for adjusting selectivity. where T 1 = operating time of relay at breaker 1, s, T 2 = operating time of relay at breaker 2, s, B 2 = short-circuit interrupting time of breaker 2, s, O x = overtravel time of relay at breaker 1, s, and F = factor of safety, s. As was shown in chapter 10, the sum of B 2 , 1? and F can normally be assumed to be 0.4 s. When coordinating the time delays of inverse-time relays, the process should start with the most downstream relays and work back toward the substation. The typical connections for the line-overcurrent relays and CT's used in switchhouses are illustrated in figures 13.2 and 13.4. The inverse-time units are labeled 51, while the instantaneous units are labeled 50. Figure 13.2 shows 329 three wye-connected CT's driving three wye-connected overcurrent relays, whereas figure 13.4 shows two open- delta-connected CT's driving two open-delta-connected overcurrent relays. Although protection for all three lines can be obtained with the open-delta connection, this approach is neither as accurate nor reliable as the wye connection. The best ground-fault protection is obtained through zero-sequence relaying, which is shown in both figure 13.2 and figure 13.4. The CT ratio is typically 25:5 or 50:5. Unlike the power-center applications in chapter 12, the secondary current from the CT, which has an induction- disk relay as burden, is proportional to the ground-fault current as dictated by the turns ratio of the CT. However, because of the necessarily large CT windows, burden matching and actual testing are essential to obtain reli- able sensitive relay pickup. Here the aim should be to obtain 50% of the current limit. Residual ground-fault relaying can be found in some switchhouses, but adequate sensitivity is usually not available with this system because of errors caused by CT saturation and unmatched characteristics. Some success in eliminating this problem has been achieved with the use of solid-state relays, as discussed in chapter 14. must have a close-and-latch rating, based on asymmetrical current, that is greater than 1.6 times the maximum symmetrical fault current. Sometimes an automatic load-break switch inter- locked with fuses (as discussed in chapter 12) replaces the power circuit breaker in the most downstream switch- house in the distribution system. A small delay is built into the switch tripping circuit to remove the problem of the switch's trying to clear fault currents. The time delay is coordinated with the maximum period that a second fuse takes to clear a line-to-line or three-phase fault. Manual interlocking with the fuse elements is therefore not applicable. The only automatic load-break switch suitable for this approach is the type where the fuses have elements that activate a contact set upon separation and can be used to power the tripping mechanisms in conjunc- tion with the control circuitry. For additional information, note that chapter 9 con- tains extensive information about different power circuit breakers and their advantages and disadvantages, and chapter 10 covers the selection of continuous-current, interrupting, and close-and-latch ratings for these devices. SWITCHHOUSE CONTROL CIRCUITS POWER CIRCUIT BREAKERS Three types of power circuit breakers can be found in mine switchhouses: live-tank oil (or minimum oil), dead- tank oil (OCB), and vacuum (VCB). Live-tank oil circuit breakers have been plagued with maintenance problems, the most outstanding being inadequate oil levels after five or fewer interruptions. This requires extremely frequent inspections, which are impractical for most mining opera- tions. Three-phase dead-tank OCB's in 250- to 500-MVA interrupting capacities are used exclusively by many min- ing companies with tremendous success. Despite this, their use has been substantially curtailed in recent years because of cost, availability problems, polychlorinated biphenyl (PCB) contamination, and their large size for thin coal seams. In fact, some States prohibit the use of OCB's in underground applications above 10 kV. VCB's have been criticized in terms of the transient overvoltages that their high efficiency can create. However, as demon- strated in chapter 11, these problems can usually be related to poor power-system design practices. When rea- sonable precautions are met, the VCB is perhaps the most effective high-voltage power circuit breaker compared with other types that have equal voltage and interrupting- capacity ratings. It is by far the most popular circuit breaker used in switchhouses. The voltage ratings of power circuit breakers corre- spond to the insulation class used in the mine distribution system, but manufacturers offer VCB's rated at 15 kV, which are applicable for 4.16-kV through 14.4-kV nominal voltage systems. Continuous-current ratings of 400 and 600 A are the most common in switchhouse applications. Since high-voltage circuit breakers begin interruption a few cycles after the first-cycle fault current peak, the interrupting rating of these breakers is based on symmet- rical fault current. (Many power breakers are also rated on an asymmetrical interrupting-current basis.) A 400-A VCB has a typical interrupting rating of 4, 000- A symmet- rical, with the 600-A model rated at 12,000 A. The breaker must also be able to withstand the physical stresses resulting from first-cycle fault currents. Thus, the breaker Most control-circuit schemes found in switchhouses have a similar operation. Thus, for convenience of discus- sion, the typical control shown in figure 13.4 is reproduced in figure 13.7 and the control shown previously as figure 13.2 is repeated as figure 13.8. Figure 13.7 illustrates the control circuit of a double-breaker switchhouse. Each circuit breaker has its own independent control circuit. Control circuit B is an exact duplicate of control circuit A; therefore, only circuit A will be discussed. The actuating currents for the line overcurrent relays (50/51-1A and 50/51-3A) are obtained by the open-delta- connected current transformers. The relays are also con- nected in an open-delta configuration through the amme- ter and ammeter switch. The ground-fault relay 51-GA is used in a zero-sequence ground-trip circuit. A single-phase control transformer steps down the distribution voltage to 120 V to supply both control circuits and the strip heaters. The strip heaters are resistive devices placed in the switchhouse enclosure to provide heat that minimizes moisture accumulation. The red lamp is connected in series with the normally open (NO) auxiliary contacts (52 A/a) of the circuit breaker, while the green lamp is in series with the normally closed (NO auxiliary contacts (52 A/b). The auxiliary contacts are mechanically controlled by the opening and closing mechanism of the breaker. When the breaker is closed, 52 A/a also closes and causes the red lamp to give a visual indication that the breaker is closed. At the same time, 52 A/b opens, which causes the green (open) lamp to be extinguished. The reverse procedure follows when the circuit breaker is opened. The ground-check system of branch A (GCS-A) also obtains its power from the control circuit as shown. The output of the GCS is connected to the pilot and grounding conductors of the outgoing cable of branch A to ensure continuity of the downstream grounding conductor. A capacitor-trip device is used to supply power to the circuit breaker trip circuit in the event of a fault or overcurrent condition. A diode is used as a half-wave rectifier to charge the capacitor (C). The capacitor stores energy to provide reliable tripping of the breaker and 330 minimize nuisance trips. The voltage-sensing relay (CR) operates when the capacitor is fully charged and causes the lamp to provide a visual indication of a full charge. Another NO set of auxiliary contacts (52 A/a) is placed in series with the trip coil of the circuit breaker to disconnect the tripping circuit from the control voltage when the breaker is open. The contacts of the line- overcurrent relays and the ground-fault relays are con- nected in parallel as shown. If either the instantaneous or the inverse-time unit of any of these relays is actuated, its associated contacts close, which completes the path and allows the capacitor to discharge through the trip coil. When the trip coil becomes energized, it causes the trip mechanism of the breaker to release and open the breaker. A push button and an NC set of ground-check contacts (GCS-A) are also parallel with the contacts of the overcur- rent relays. The GCS-A contacts should remain open, provided that the control circuit is energized and the integrity of the associated pilot and grounding conductor is maintained. If the continuity of the loop circuit is lost, the GCS-A contacts will close, which results in tripping the breaker. Depressing the push button provides a conve- nient means of tripping the breaker, as well as a means of testing the operation of the capacitor trip device. The control circuit of figure 13.8 is similar to the above control circuit, with some minor exceptions. Three CT's and three line-overcurrent relays are connected in a wye configuration to provide a more precise and reliable operation. A battery, which is charged by a full-wave bridge rectifier, is used as the energy storage device, rather than a capacitor. An undervoltage release (UVR) (UV 52) is also included, with an NO set of ground-check contacts in series with it. Control transformer nnm I — WW— i <| — w*—>> ii — WW—" Strip heaters 50/51-1A 50/51-3A 51-GA i SI (siV* ,SI CSD-* .SI GCS-A fa 52B/a 52B/b Closed Open 1 f GCS-B 2_f To pilot B ii WW- -@- o ^52B/b ^52B V/ trip coil I CR 50/51-1 B 50/51-3B 51/GB i SI (SI .SI .SI GCS-B R°fl ■PPF AMSW KEY I Instantaneous contact T Timed contact SI Seal -in contact Tripbutton-A Trip button-B Figure 13.7.— Typical control circuit for double switchhouse using capacitor tripping. 331 I — WW— i , — ww- ' — WW-' o ^p ^jrrr^ . 52 a 52 b o r.rc o GCS GCS »To pilot — o — UV52 "UUUJ Heater strips - 1 - 52 a Battery 5 52 trip coil KEY I Instantaneous contact T Timed contact SI Seal-in contact * 'si K^ x_^ l-@- -H^ i-^ GCS _i_ Trip button o o — - 50/51-1 50/51-2 50/51-3 51-G Figure 13.8.— Typical control circuit for single switchhouse using battery tripping. SWITCHHOUSE DESIGN At first glance the design of switchhouses is not nearly as complicated as that of mine power centers, but the simpler circuitry can be misleading. Incorrect sizing or adjustment of the internal components can have a disas- trous effect on the mining operation. Perhaps the most serious concern is coordination, because it is in switch- houses that most adjustments must be made to obtain sensitive and selective protective relaying. For this reason alone, load-flow and fault analysis of the entire mine power system must be available. Comparison of these data allows the judicious selection of the required transformer ratios, relay time-current characteristics, pickup ranges, and device ratings. Some additional comments about the process are in order. Because of the need for a uniform design, it is best for all switchhouses to have an identical assembly. In other words, all single switchhouses should have the same components, and one section of a multiple-breaker switch- house would be basically equivalent to a single switch- house. This may not be economically possible in some mines and not practical in others, yet there would be a specific advantage: switchhouses could be placed any- where in the mine and then adjusted to suit that location. This would benefit any major movement of equipment within the complex. To provide such uniformity, all power circuit breakers should be the same, with 1. The continuous-current rating sized to the maxi- mum continuous current through any switchhouse in the mine (demand factors should be applied), 2. The interrupting-current rating selected to stop the maximum short-circuit current at any location in the distribution system (opposed to the preceding statement, the required rating here might be greater than symmet- rical rms short-circuit current, depending upon the time to interruption), and 3. Close-and-latch rating sized to the maximum asym- metrical first-cycle rms fault current (convention is to use 1.6 times the interrupting-current rating). Load-break switches need this same close-and-latch rat- ing, rather than a multiple of their interrupting duty. The ampacity of all power conductors within the equipment would be sized and braced accordingly, but the size should never be less than 4/0. For overload and short-circuit protection, multiratio CT's might be necessary, but experience has shown that no more than one or two types are required for the majority of applications: for example, 150/300/600:5 A, 300/600:5 A, 75/150/300:5 A, and 150/300:5 A are common ampere-turns ratios. With precise burden matching and accuracy calcula- tions, common induction-disk relays can work reliably and selectively. Obviously the pickup range for the timed and instantaneous elements must correspond to the entire range needed for coordination. One range for each element might be inadequate for the mine's needs, but the relays are easily interchangeable, and one or two spare sets of relays can be maintained at the mine to meet all demands of line- overcurrent protection. Correct burden and accuracy must be verified for any relay-CT combination. Adequate ground-fault protection, however, can be a difficult goal to attain. As stated earlier, burden and accuracy matching become critical. Yet even with precise matching, reliable pickup at a desirable ground-fault current threshold sometimes cannot be achieved. The large window required on the zero-sequence CT introduces induction problems that can cause inadequate secondary current at rated burden. The main result is that both the timed and instantaneous elements of induction-disk relays cannot be used simultaneously for ground-fault protection with a 50:5 or even a 25:5 CT ratio. A solution to this ground-fault problem is to use only timed elements and apply the minimum time dial setting to the most downstream switchhouse. The radial distribu- tion arrangement can then be established such that no more than five relays exist between the substation trans- former secondary and the extreme end of distribution (see chapter 10 for discussion). This enables practical ground- fault coordination to be achieved from one tap-setting range and only one relay type. However, such problems as tolerances in off-shelf components can still remain. The best recourse is to apply testing that simulates a ground fault through the CT, then observe the response of the induction-disk relay. The optimum location for this testing is at a switchhouse in place within the mine power system. During initial equipment installation, the tests are not difficult, but subsequent maintenance checks may be impractical. As an alternative, it is possible to include a 332 test circuit, such as shown in figure 12.20, in the switch- house circuitry for each ground-fault relay. The resistor and wire turns about the CT would then be adjusted for the required relay pickup; for instance, 2 A and 6 turns would simulate a 12-A ground-fault current or 48% pickup on 25-A ground current limit. A final point must be made about switchhouse con- struction. Reference has been made to surge capacitors in chapters 11 and 12: in the past, some manufacturers and operators incorporated these devices to increase the sys- tem's characteristic impedance in an attempt to limit transient overvoltages created by VCB chopping. However, the use of discrete capacitors of any kind on the load side of a vacuum interrupter, including power-factor correction types, can set the stage for prestrike and capacitance- switching transients. Hence, such devices should never be specified within a switchhouse. (A surge capacitor used directly across transformer and motor windings to limit the rate of voltage rise is a feasible means of protection if needed.) For similar reasons, switchhouse and power- center components must never be combined into a single enclosure, for instance, by using a VCB in place of the high-voltage fuses that were shown in figure 12.2. Con- versely, surge arresters are desirable transient protection, and distribution-class types should be installed in all switchhouses. ever possible. A substation is used to receive the utility power before distribution to the mine and other facilities. This complex of equipment contains switching and protec- tive relaying, establishes the grounding system, and may or may not involve transformation to a distribution volt- age. The substation design should be based on personnel safety, reliability of operation, and ease of maintenance, and limit damage from fire, lightning, and equipment malfunction (2). The utility supply voltage for mines may range from 230 to 13.2 kV and less, depending upon utility company availability and the economics and location of the mine. The substation components are often mounted on concrete pads. Power-conductor terminal structures, com- monly made of galvanized steel but also of wood, support the conductors and cables that provide connections for transformers and circuit breakers. Insulators, usually made of glazed porcelain or Pyrex heat-resistant glass, insulate the overhead conductors from the supports. An overall view of a typical mine substation is given in figure 13.9. The main substation is usually a permanent instal- lation, but portable or unit substations are becoming very popular for small load requirements. BASIC SUBSTATION ARRANGEMENTS SUBSTATIONS It is common practice in the mining industry to purchase electrical power from a utility company when- The most popular distribution system used in mining is expanded radial A simplified sketch of this system is illus- trated in figure 13.10, with the substation noted. Radial techniques have the lowest initial cost since a single trans- former supplies all mine circuits and there is no duplication , ■ J :.,_ ...' * ■ M'4'-~ 4-" ~f\ . At wi V? * Figure 13.9.— Overall view of main substation serving mine. 333 Utility Utility Figure 13.10.— Radial distribution applied to underground mine and its surface facilities. of equipment. Operation and expansion is simple, but this simplicity leads to its major disadvantage: if a major compo- nent fails, the entire system can be effected. Nevertheless the logistics of mining, especially underground mining, often dictate the use of radial systems. Single-Ended Substations The substation design employed for radial systems is often termed single-ended. Figures 13.11 and 13.12 are one-line diagram examples, and the only difference be- tween the two is found on the primary side of the trans- former. The substation components, much like those con- tained in power centers, can be grouped into three general sections. The primary section provides connection with one or more incoming utility lines, switching and inter- rupting devices, and transformer protection. The trans- former portion may include one or more transformers, a neutral point to establish grounding, and automatic tap changing, if used. The secondary section or distribution side provides for the connection of one or more secondary feeders, each having switching and interrupting devices. Transformers (primary side) may be protected by high-voltage power fuses (fig. 13.11) or by a circuit breaker and associated overload and short-circuit relaying (fig. 13.12) or, less frequently, by both. Fuses or a circuit breaker is usually all that is necessary. Ground-fault protection could also be provided on the primary (not shown), but this relaying is restricted to circuit breaker applications and its inclusion depends to some extent on utility company requirements. The sizing of components is similar to that already discussed for power centers and switchhouses. The outgoing-circuit protective relaying covers ground faults, overloads, and short circuits, and establishes a primary protection zone as far as the first downstream switchhouse. There are two recommended backup ground- fault relays: neutral-current sensing between the trans- former bushing and the top of the grounding resistor, and potential relaying about the grounding resistor. The first ensures backup if the resistor shorts, while potential relaying provides protection against open-resistor hazards. Station y\ ground ^~ Surge arrester — o o — ? Gang-operated switch I Power fuse A uLu Power i— o 51-N 52 transformer -3 pt Circuit breaker 59-G 50/51 T Pilot . Disconnect / switch Surge arrester — o o — ^Station "V* ground Power conductors To borehole Grounding conductor Figure 13.1 1.— One-line diagram for single-ended substation with fuse-protected transformer. 59-G .v. Station \s* ground vy Station ^ ground * Power conductors Grounding conductor Figure 1 3.12.— One-line diagram for single-ended substation with circuit-breaker-protected transformer. 334 Both problems can occur in this surface installation be- cause of exposure. The ground-fault protection must be coordinated with that in the downstream switchhouses. Line-overcurrent relays must coordinate with downstream protection as well as with the transformer primary protec- tion and the utility. Circuit breaker, relay, and transformer sizing and selection generally follow the procedures given earlier, this time closely related to switchhouses. The substation has two separate ground beds: the station (or system) and the safety (or mine). The grounding resistor is tied to the safety bed, and a grounding conduc- tor extends into the mine. Ground-check monitoring of this conductor is required. The surge arresters, substation fencing, and other metallic parts are grounded to the station ground bed. All substation components should be located within a fence posted with danger signs; the internal area, includ- ing the fence, is called the substation area. Specific concerns about substation components will be given in the following sections. SUBSTATION TRANSFORMERS Transformers for permanent surface substations are almost always liquid (oil) immersed and built to IEEE standards (5). A typical substation transformer is shown in figure 13.14. Capacities commonly range from 5,000 to 30,000 kVA (sometimes less, but rarely more). In addition to the mine, the loads may include a preparation plant and general surface loads, such as pumps, ventilation fans, maintenance shops, and office and bathhouse facilities. When an underground mine is involved, it is recom- mended that a separate transformer be used for the underground power system. The selection of transformer capacity must be based on estimation of electrical load. Up to 2,000 or 3,000 kVA, a general rule of 1 kVA/hp of connected load is perhaps satisfactory. This provides more than adequate capacity Utility Utility Double-Ended Substations To enhance the reliability of the distribution circuit, some mines utilize a secondary-selective system. Two substation secondaries, as shown in figure 13.13, are connected with an NO tie circuit breaker. The combination is commonly but not always in the same substation area. Both substation halves are identical, and each is basically a single-ended unit but often with only one station and one safety ground bed. Under normal operation, each substa- tion independently feeds 50% of the load, and the distri- bution system on either side is actually expanded radial. If a primary feeder or a transformer fails, the main second- ary breaker on the inoperative circuit is opened, and the tie breaker is closed either manually or automatically. With automatic operation, relaying between the second- ary breakers and the tie breaker must be interlocked so that the tie breaker will not close on a faulted distribution system. For instance, if the main secondary breaker is tripped by its associated relaying, the same relays should prevent tie breaker closure. To permit continued operation with one transformer, consideration must be given to the following (8): • Oversizing both transformers (instead of 50%, 80% of load) so one can temporarily carry the total load (at 125% of its capacity), • Providing forced-air cooling to the transformer that remains in service during the emergency period, or • Oversizing both transformers so that one can carry the total load. These substations, also called double-ended, can be eco- nomical when the total substation capacity is above 5 MVA. Two separate incoming utility lines maximize the advantage of secondary-selective over radial systems. However, because of high transformer reliability, the eco- nomics of using a double-ended substation with one incom- ing line is considered questionable by some. Yet a definite advantage of this technique over single-ended types is that maintenance chores can be performed without power out- age, such as insulator cleaning, component adjustment and replacement, and no-load tap changing. I Mine loads Tie circuit breaker Mine loads Figure 13.13.— Simplified one-line diagram for double-ended substation. Figure 13.14.— Typical liquid-immersed transformer in substation. MB^MM 335 for the load and will allow minimal load growth. Note that a worst case demand factor of 0.75 to 0.85 is built into the selection. Such oversizing will also allow for an adverse power factor. When the requirements are greater than this, a more precise load estimate is required if optimum equipment costs are to be achieved. Considering the common capacity needs for the majority of mining operations, a load-flow analysis must be performed. The estimation may be made by using the demand factors discussed in chapters 4 and 8. In some cases, capacity might be determined on a per- connected-kilovoltampere basis (about 0.75 per connected horsepower) to allow for the growth in demand factor that always exists. Table 13.1 shows recommended impedance values on a percentage basis for a liquid-immersed transformer (5). These values cover a capacity range from 500 to 30,000 kVA and are based on primary voltage, whereas mine power-center values are based on capacity. To provide a reasonable limit to short-circuit currents, the reactance should never be less than 5.0% (0.05 pu). Past practice in the industry was to use three single- phase transformers for easy replacement of failed units, and a spare was often included to facilitate the changeover. However, improved manufacturing techniques have now resulted in transformers' being among the most reliable components; thus, the common practice today is to install three-phase integral units (13). The same advantages given for power centers hold for substations. The standard ratings for substation liquid-immersed transformers are based on an allowable average winding- temperature rise of 65°C. A rating related to a 55°C rise is also available but is no longer listed as a standard (5). A transformer rated with the 65 °C rise can be loaded to 112% of the 55°C-rise specification. The transformer ca- pacity always involves a self-cooled [over-air (OA)] rating but may also have a forced-air-cooled rating (FA). Standard capacities for OA-rated and FA-rated substation transform- ers are listed below (8): Secondary Secondary < 1,000 V 2,400 V and above 5.75 5.5 6.25 6.0 6.75 6.5 NAp 7.0 NAp 7.5 NAp 8.0 Table 13.1.— Standard impedance for liquid-immersed three-phase transformers, percent Primary voltage, V 2,400 to 22,000 26,400 to 34,500 43,800 67,000 115,000 138,000 NAp Not applicable. SUBSTATION SWITCHING APPARATUS Dead-tank OCB's, such as those illustrated in figure 13.15, are the most common interrupters used in substa- tions. In recent years, air-magnetic, vacuum, and sulfur hexafluoride (SF 6 ) circuit breakers have gained in popularity because of their reduced and simple maintenance, but for applications at utility voltages of 69 kV and higher, OCB's or SF 6 breakers are employed almost exclusively (13). SF 6 circuit breakers use a fluorocarbon gas for the arc- interruption medium, and their high cost usually restricts their use at 69 kV and above. On the other hand, the use of OCB's at lower substation voltages often results in lower costs and easier installation than the alternatives. Because of the transformer overload capacity and the lack of overload capacity in circuit breakers, main break- ers of the transformer secondary should have a continuous- current rating 25% greater than the anticipated top continuous-current rating of the transformer. Obviously, the voltage, interrupting-current, and close-and-latch rat- ings must be sufficient for the system to which the breaker is applied. These demands may be greater than for inter- rupters in switchhouses. Reclosers 55°C OA: 500 2,500 12,000 750 3,750 15,000 1,000 5,000 20,000 1,500 7,500 25,000 2,000 10,000 30,000 65 °C OA: 560 2,800 13,440 840 4,200 16,800 1,120 5,600 22,400 1,680 8,400 28,000 2,240 11,200 33,600 55°C FA: 575 3,125 16,000 862 4,687 20,000 1,150 6,250 26,667 1,725 9,375 33,333 2,300 12,500 40,000 65°C FA: 644 3,500 17,920 966 5,250 22,400 1,288 7,000 29,867 1,932 10,500 37,333 2,576 14,000 44,800 Oil circuit reclosers have proved to be extremely reliable as interrupters on the transformer secondary at distribution voltages to 15 kV (13). A circuit recloser is a circuit breaker If the transformer is so equipped, the unit could also have a forced-oil-and-air (FOA) rating. Figure 13.15.— Dead-tank OCB in substation. 336 with the necessary self-contained ability to detect line over- currents, to time and interrupt the overcurrents, to reclose automatically, and to reenergize the line. If the line overcur- rent is permanent, the recloser will lock open after a preset number of operations (usually three or four) and isolate the failure. Thus, the recloser can eliminate prolonged outages of the distribution system due to temporary faults or tran- sient overvoltage conditions (12). Reclosers can be hydraulically or electronically con- trolled (12). With the hydraulic control, an overcurrent is sensed by a trip coil in series with the line. When the minimum trip current of the recloser is exceeded, the trip coil actuates a plunger and causes the recloser contacts to open. The timing and sequencing are accomplished by the pumping of oil through separate hydraulic chambers. Electronically controlled reclosers provide a more easily adjusted, flexible, and accurate control then the hydrauli- cally controlled recloser. The electronic control gives a convenient means for changing time-current characteris- tics, trip-current level, and the sequence of the recloser operation without deenergizing the recloser. Auxiliary tripping devices are available to allow additional protec- tion, such as ground-fault and ground-check monitoring. Activation of the auxiliary tripping causes the recloser to lock open. The selection and application of a recloser is basically the same as for an interrupter. The necessary items to consider are system voltage, maximum available fault current at the recloser location, maximum load current, maximum fault current in the zone protected by the recloser, coordination with other protective devices up- stream and downstream, and ground-fault tripping or sensing. The common ratings are voltage, basic impulse insulation level (BID, continuous current, minimum trip current, and interrupting current. Disconnect Switches and Fuses To provide a visual disconnect for maintenance pur- poses, knife-blade load-break switches or fuse cutouts are located on the primary and secondary of the substation transformer. The switches are usually housed in metal- clad enclosures, while the cutouts are pole mounted. Gang-operated switches are recommended over hook-stick devices, since pulling one line at a time with the circuit under load can obviously single-phase the system. Fuse cutouts have an interrupting capacity instead of a short-time current rating. They should be used to isolate parts of a deenergized circuit, even if the cutouts used are designed with mountings that operate as a load-break switch. Cutouts alone cannot be relied upon to protect the transformer in all cases, especially when the available short-circuit current is above their interrupting rating. Power fuses should therefore be used as backup protection. High-voltage fuses may be employed as the main means of transformer protection. As with any protective device, their time-current characteristics must be coordi- nated with upstream and downstream devices, here the utility and the protective relaying on the secondary. Fuses alone cannot provide ground-fault protection of the trans- former primary. When fuses are used and ground-fault protection is needed, relaying could be used to trip the mechanism of an automatic load-break switch. A general protection rule, though, is to use a primary circuit breaker with protective relaying when the three-phase trans- former capacity is 5,000 kVA and above (13). For any capacity, the primary breaker prevents the single phasing that fuses allow and permits easier coordination with other relaying. PROTECTIVE RELAYING IN SUBSTATIONS The sizing of overload, short-circuit, and ground-fault relays is basically no different from that described already in this and previous chapters for other power equipment. Transformer protection is a protective-relaying problem in substations. With larger substation capacities, relay pickup for overloads and short circuits alone is normally too high to provide adequate protection. Sudden-pressure relays that sense the rate of pressure rise in the gas cushion of the transformer tank are sensitive to small arcs under oil, and their use is warranted (13). Further, percentage-differential relaying is also strongly suggested for internal fault protection of transformers rated at 5,000 kVA or higher, although it is not as yet a widespread practice. A standard percentage-differential relaying system is illustrated in figure 13.16. Figure 13.17 is a typical one-line diagram of a substation with this differential relaying added. The CT's on a wye-connected winding of a transformer should be connected in delta, while CT's on a delta winding are connected in wye. There are two basic requirements that percentage-differential relaying must satisfy (11). 1. The relays must not operate for load currents or external faults. 2. The relays must operate when internal faults are severe enough. Relay pickup is used to trip the breaker on the transformer primary. Operating coil Restraining coil Percentage - difference relay Figure 13.16.— Standard percentage-differential relaying system for transformer protection. H^M^H^ 337 Utility 59-G Operating coil Restraining coil — > To borehole Ground Power Pilot Figure 13.17.— One-line diagram of substation with percentage-differential relaying. Differential relays usually have coil taps to compen- sate when the CT's are not perfectly matched. When selecting a CT for differential relaying, the common prac- tice is to choose the highest CT ratio that will provide a secondary current as close as possible to the lowest-rated relay tap. This minimizes the effect of impedance in the wiring connections between the CT's and the relays. To assure that the relay will operate a maximum sensitivity, the current supplied to the relay under maximum load conditions should be as close as possible to the continuous- current rating of the tap (11). Percentage-differential relays usually have an adjust- ment to vary the percent slopes. The adjustment provides a means for preventing unreliable relay operation due to unbalances between CT's during external faults. Unbal- ances can occur from the following: • Tap changing of the power transformer, • Mismatch between CT secondary currents, and • The difference between CT errors on either side of the power transformer. If a power transformer is rated at 10,000 kVA or greater, a harmonic-restraint circuit is recommended in addition to the percentage-differential relays (13). This circuit causes a differential relay to be self-desensitizing during magnetizing inrush periods, but the relay is not desensitized if a short circuit occurs in the transformer during a magnetizing inrush period. Only the fundamen- tal component of the differential current is delivered to the operating coil. The harmonics are separated, rectified, and delivered to a restraining coil (11). The pressure relays mentioned earlier can play a valuable role as supplemental protection to differential relaying. In fact, with sensitive and reliable pressure relays, the sensitivity of the differential relays can be reduced to prevent undesirable operation due to inrush current (11, 13). Thermal-overload protection should also be provided as a third means of transformer protection. LIGHTNING AND SURGE PROTECTION IN SUBSTATIONS Since much of the equipment of the surface substation may be exposed, there must be protection against tran- sient overvoltages due to lightning as well as due to switching. Overhead static wires and shielding masts (see chapter 11) are commonly used to protect substation equipment from direct lightning strokes. Two static wires can be positioned above and between the line conductors to provide shielding for overhead distribution lines. In addi- tion, surge arresters are mandatory to limit the transient overvoltages to safe levels. The surge arresters on the incoming lines should be located as close as possible to the transformer terminals. Station-class valve arresters should be used. Again, ar- rester selection should be based on the primary voltage, the effectiveness of grounding, and insulation coordina- tion between the arrester and the transformer BIL. With dry-type transformers, the BIL is practically constant with the width of applied impulse (see chapter 12, figure 12.5). The margin of protection could then be less for the arrester sparkover voltage than for the IR-discharge voltage. It can be seen in figure 13.18 that in liquid-immersed transform- ers, the insulation withstand is not a linear function with the impulse width. Instead, the insulation-withstand level decreases from the front-of-wave to chopped-wave to the full-wave values (12). The full-wave value is the BIL rating. Standard values for oil-immersed transformers are listed in table 13.2 (9), and values of insulation classes below the standard are provided in chapter 11, table 11.1. As shown in figure 13.18, the BIL should be compared with the discharge voltage. Because the exposure avail- able in substations can lead to worst case surge conditions, Table 13.2.- -Standard BIL's for oil- transformers immersed power Primary winding phase-to-phase voltage, V Insulation class, kV BIL, kV 22,900... 25.0 25.0 34.5 34.5 46.0 46.0 69.0 69.0 92.0 115.0 138.0 161.0 150 23,000... 150 26,400... 200 34,500... 200 43,800... 250 46,000... 250 67,000... 350 69,000... 350 92,000... 450 1 1 5,000 . 550 138,000. 1 350 M50 650 161,000. 1 550 1 450 750 '650 1 550 1 Reduced BIL's may be applied if proper coordination is maintained with surge arresters. 338 > < o > Front-of-wave withstand Chopped-wave withstand BIL (full-wave withstand) Front-of-wave sparkover Minimum margin * of protection ^-Discharge-voltage maximum r characteristics of arrester Impulse voltage wave rising 100 kv///s for each 12 kV of arrester rating 1 3 4 TIME, //s Figure 13.18.— Insulation characteristic of liquid-immersed transformer compared with the characteristic of valve surge arrester. ii f (i 1> 1 ii < ,, 25 ft Figure 13.26.— Use of isolation transformer with utility substation. ADDITIONAL SUBSTATION DESIGN CONSIDERATIONS Over the years many additional design precautions have been developed through experience with permanent substations at mines. A listing of some of these follows: • Substation conductor insulators can be of post or cap-and-pin types, and either standard strength or heavy- duty design can be used (13). "Extra-creepage" insulators and bushings (or one voltage class higher than required) are recommended (7) as they extend the time between required cleanup and minimize flashover. • The high-voltage side should be located on the windward side of the substation (9-10). The prevailing winds will then help keep the substation insulators clean between regular cleanups. • The voltage drop on the distribution system should be maintained within + 5% to all load concentrations (7). This establishes the nominal voltage and necessary taps for the secondary of the substation, and corresponds to the taps stated earlier for transformer primaries in power centers. • Permanent substations should be located outside the influence of the mine. This is very critical for surface mines. • Buses and bus supports should be designed to withstand the stress from the average asymmetrical short- circuit current during the first 10 cycles after a fault (13). • Line tensions within the substation area should be assumed to occur under maximum wind-loading and ice- loading conditions (13). Structure design should also be based on these specific line tensions and wind loading. • Substation primary voltages above 34.5 kV cannot be used for distribution-class equipment (that is, cannot be transformed directly to low-voltage and medium-voltage mine usage) (13). This could create problems in supplying certain surface loads. • The utility will meter the substation power con- sumption on the primary or secondary, depending on the primary voltage. The substation should also contain the mine's own metering, of equivalent precision to that of the utility (7). This is invaluable for maintenance checks and also serves as a double-check on the utility. Metering could include kilowatt, ammeter, voltmeter, and power-factor 345 instrumentation for each outgoing feeder from the substa- tion. Demand meters, for either 15- or 30-min maximum demand, may also be needed. • The fence enclosing the substation area should be no less than 8 ft high (1). The fence should be provided with a door or gate, which should be locked except when authorized personnel are present. The fence should be grounded to the station ground bed. It should also extend to the ground so no one can crawl underneath. Danger signs should be posted on all sides of the substation area. • The substation area should be maintained free of weeds, trash, and combustible material that might create a hazard to personnel or the substation components (2). • All components containing liquid insulation should be mounted on pads to allow quick drainage and capture of any lost fluids. This is especially important when flamma- ble or environmentally hazardous liquids are involved. Chapters 12 and 13 have attempted to collate an enormous amount of information to provide a coherent presentation of typical construction practices used for mine power equipment. Some designs, like mine power centers, are so dependent upon the individual application that standard units cannot exist. At the same time, the assembly of others, such as switchhouses and substations, is rather uniform across the industry. Because the subject matter is so large, some generalization has been essential and the omission of some specialized material was inevi- table. Any engineer involved with mine equipment should of course work very closely with the manufacturer to determine the best solution to the specific needs of the mine. REFERENCES 1. American Standards Association. American Standard Safety Rules for Installing and Using Electrical Equipment in and About Coal Mines (M2.1). BuMines IC 8227, 1964. 2. Bifulco, J. M. How To Estimate Construction Costs of Elec- trical Power Substations. Construction Publ. Co., 3. Cooley, W. L., and R. L. King. Guide to Substation Ground- ing and Bonding in Mine Power Systems. BuMines IC 8835, 1980. 4. Dornetto, L. D. The Importance of Grounding Systems in the Protection of Personnel and Equipment. Paper in Mine Power Distribution. Proceedings: Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 5. Institute for Electrical and Electronics Engineers (New York). General Requirements for Liquid-Immersed Distribution, Power, and Regulating Transformers. Stand. C57.12.00-1979. 6. Guide for Safety in Substation Grounding. Stand. 80-1976. 7. Recommended Practice for Cement Plant Power Distribution. Stand. 277-1967. 8. Recommended Practice for Electric Power Distribu- tion for Industrial Plants. Stand. 141-1986. 9. Lordi, A. C. Electrification of Coal Cleaning Plants. Mechanization, v. 20, Oct. 1956. 10. Trends in Open-Pit Mine Power Distribution. Coal Age, v. 66, Jan. 1961. 11. Mason, C. R. The Art and Science of Protective Relaying. Wiley, 1956. 12. McGraw-Edison Co., Power Systems Div. (Canonsburg, PA). Distribution-System Protection Manual. Bull. 71022, undated. 13. Wade, E. C, and M. E. Kunsman. Practical Considerations for Selectiong and Coordinating Electrical Components in Outdoor Mining Substations. Paper in Conference Record -IAS 13th An- nual Meeting (Toronto, Ontario, Canada, Oct. 1978). IEEE, 1978. 346 CHAPTER 14.— SOLID-STATE CONTROL AND RELAYING 1 Except for a short introduction to electronics in chap- ter 5, discussion of solid-state devices has been avoided up to this point. The reason is that their use is rather new to the mining industry. However, their impact has been substantial, and if one area of electrical engineering can be singled out as having the greatest probable influence on future mining operations, it is electronics. This chapter will primarily discuss two solid-state applications that are already important: motor control and protective relaying. MOTOR CONTROL Almost all activity and interest in the solid-state control of industrial motors have centered around the use of the silicon-controlled rectifier (SCR) or thyristor. A model and circuit symbol for this four-layer, three-junction device are shown in figure 14.1. The outer two layers act as a p-n junction and the inner layers serve as an element to control that junction. The device has three external termi- nals: anode, cathode, and gate. If the anode is positive with respect to the cathode (forward biased) and if the gate is reversed biased in reference to the cathode, there exists a balance of electri- cal charges in the four layers, and current flow in inhib- ited. This process is termed forward blocking, and the SCR exhibits high resistance in both directions. When the gate is forward biased with respect to the cathode, the gate current upsets the electrical charge balance, and anode- to-cathode current can flow. Once this conduction starts, the gate loses all control; in other words, the gate can turn the thyristor on but not off. Gate control (or turning the thyristor off) can only be achieved by reverse-biasing the gate-cathode circuit once more and reducing the anode- cathode current essentially to zero. The typical characteristic curve for a thyristor illus- trated in figure 14.2 details the phenomenon just de- scribed. Voltages and currents are given with respect to the cathode. When the anode potential is positive and the gate is at cutoff (negative potential), the characteristics are similar to those when the anode is negative. At a specific positive potential, called the breakover voltage, the SCR will turn itself on. Here, anode current increases considerably, and the voltage drop is reduced substantially across the device. The breakover point can be altered by varying the gate current, and this is why the thyristor is so valuable in power control applications. The common technique for initiating conduction is to apply a current pulse at the gate (12). 2 The pulse alone neutralizes the thyristor blocking action at the desired breakover voltage. The simplest thyristor application is in the half-wave rectifier circuit shown in figure 14.3. Break- over is determined by the gate current, and thus the average value of dc through the load can be controlled by the gate control pulse (i G ). The process of starting the thyristor is often referred to as firing, and the angular 1 The author wishes to thank D. J. Tylavsky, assistant professor of electrical engineering, Arizona State University, who prepared the original material for the static protective relaying section of this chapter while he was a graduate student at The Pennsylvania State University. 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. position or timing of each pulse with respect to the source waveform is called the firing angle (angle a). In order to control ac, the load must receive both sides of the ac waveform. Two single-phase circuits that can be used as ac voltage controllers are shown in figure 14.4. The bidirec- tional arrangement uses two thyristors back to back and gives a symmetrical output with appropriate firing-angle pulses for the two gates. The unidirectional circuit with one thyristor and one diode presents an asymmetrical waveform to the load. The dc component thus created is considered a major disadvantage because the source wave- form must be ideal for the circuit to be of practical value (9). Voltage control for three-phase loads employs three single-phase controllers as shown in figure 14.5. Although a bidirectional circuit is illustrated, unidirectional circuits can also be applied. The dc components are mostly can- celled with the three-phase waveform, but the unidirec- tional circuits introduce a higher harmonic content to the line currents (9). Gate section , * , Anode P N P N Cathode A Gate terminal -oGate Anode o -o Cathode Figure 14.1.— Model and circuit symbol for thyristor. J Avalanche breakdown ON characteristics w Breakover voltage ~\~ V-±- OFF characteristics Figure 14.2.— Typical characteristic curve for thyristor. 347 'G VAK Q v=/2V sinort Vq = Vr CR ' G An J2V| 1.0 IT tx 1.0 o- - cat 2TT ojt ■IT 27T -•►aft « IT 2TT Figure 14.3.— Thyristor half-wave rectifier. «rt Qi -K 1 7*7^ Q. v=/2V sin art tt) Vo.'c R r\ + . T+oc 21T oc U^ art (2.) Bidirectional v=/2V sin art Unidirectional v o^R v o>'o i (/.) R n T \ oc V /2TT * 6Jt Figure 14.4.— Alternating current thyristor control. 348 Figure 14.5.— Three-phase control with bidirectional thyristor arrangement. Figure 14.6.— Full-wave thyristor bridge rectifier. Simple Motor Control One method of speed control in dc motors has already been covered, simply adjusting the dc voltage to the machine. This would apply to loaded series-wound motors, but shunt and compound dc motors can also be controlled by two direct means: armature voltage or field voltage. The field control is interesting because field excitation is usually considered constant, but in fact motor speed can be adjusted inversely with field current. The simple circuit in figure 14.3 could be used in any of these cases when the source is ac; however, the basic circuits of figures 14.6 and 14.7 are commonly applied for single-phase and three- phase sources, respectively The commutation diode shown in figure 14.7 serves a specific purpose. Commutation is the transfer of current from one device to another as voltage relations change (21 ). The commutation diode conducts armature current so that it will not be transferred to the thyristors and diodes of the bridge and the thyristors can turn off at zero voltage. Without this diode, thyristor current might not reduce to zero and turnoff would not be likely. Other commutation techniques are available (9). A chopper or dc-to-dc converter is another method of variable-speed dc motor control using thyristors; it is especially important because it operates between a dc source and a load. As shown in figure 14.8, the thyristors in the chopper circuitry switch the dc source on and off, supplying unidirectional voltage pulses. This reduces the effective voltage, V , to the load. Variable-speed control of a standard three-phase squirrel-cage motor is possible using thyristor voltage control (for example, figure 14.5). Motor speed varies because the motor allows a greater percentage of slip at reduced voltage, then the controller increases the terminal voltage, the allowable slip is reduced, and the motor speed increases. However, three specific problems arise in this application. Reduced-voltage operation causes more heat to be generated in the motor. Motor torque is also reduced at slow speeds with lower voltage. Finally, the waveform produced by the thyristors is rich in harmonics (frequen- cies other than line) and this creates even more motor heat. Nevertheless, if high-quality motors are used and reduced-voltage operation is restricted to a short time, the result can be effectively controlled starting of squirrel- cage induction motors. Commutating diode \ s> c\ 4 * - 6 dc motor armature Figure 14.7.— Three-phase thyristor-controlled rectifier. v oA V Vr 'ON tr 'ON •» *■ 'ON T — *U T— *l B Figure 14.8.— Simplified chopper control. 349 Control Systems lb this point, only thyristors and their application to variable voltage control have been discussed. Other than explaining how the device is fired, nothing has been stated about the control circuitry that supplies the gate signal. Perhaps the best way to visualize power control applica- tions with thyristors is through control-system principles. The two basic types of control systems are open loop and closed loop (12-13). In open-loop systems, the controlling element is unaware of the effect it produces in the con- trolled element, whereas in closed-loop or feedback sys- tems, information is delivered back to the controlling element so it can adjust control of the controlled element to produce a desired result. The feedback control systems of interest here are termed regulators whose purpose is to maintain the result constant. Feedback systems inher- ently provide more precise control than do open-loop systems. The operation of these systems can be described with the assistance of the basic block diagram in figure 14.9, where the lines in the diagram represent system variables. Each block represents a transfer function, the quantity that must be multiplied with the input to obtain the output. The comparison point in the block diagram is a position where two or more system variables are summed, and here the output is the algebraic sum of the entering variables. The variable to be controlled, C, may be voltage, speed, or any other quantity. The input or reference variable, R, is usually adjusted manually by an operator and determines the desired value of C. The feedback element, H, supplies information about the output to the comparison point, whose function is to compare the input and feedback signal. The difference or error, E, thus obtained drives the controller, G. In the simplest system, the controller might contain the device that needs to be controlled. (Note that G usually represents the transfer functions from input to output. H denotes transfer func- tions in a feedback path.) A feedback system contains all these elements and variables, whereas in open-loop sys- tems only the controller responds to the reference signal. Elaboration of these statements can be seen in the block diagram of figure 14.10, which illustrates the main parts of a motor controller (9). The power circuits, which primarily contain thyristors, provide variable direct or alternating voltage output to the controlled system. The controlled system can be a rotating machine or any other driven load. Appropriate feedback is supplied back to the controlling system and could, for example, represent cur- rent, voltage, temperature, speed, and so forth. The con- trolling system responds to the feedback and input infor- mation, and supplies control signals to the digital circuits. Finally, the digital circuits, acting as an additional trans- fer function, simply switch the thyristors in the power circuit off and on at appropriate times. Physical Characteristics of Thyristors The two common thyristor configurations used in industrial applications are illustrated in figure 14.11. Stud-mount types are intended for smaller loads and have average forward-current ratings up to around 150 A. Heat sinking is like that shown for diodes in chapter 5 and, because thyristors have p-n junctions, the same heat- dissipation theory applies to these devices also. The larger thyristor in the figure was introduced in the late 1960's and is termed a hockey puck, press pack, or disk (2). It has the ability to transfer junction heat on both sides and is clamped between two heat sinks as shown in figure 14.12. Comparison E t(f) Error or difference signal 6 Controller Input or reference -T Output or variable variable C R H Measurement and feedback Figure 14.9.— Basic control-system block diagram. Command Controlling ->- Digital circuits -»- f D ower Controlled Response system circuits system i Met Hit Po inf Figure 14.10.— Simplified block diagram of a motor con- troller. Anode --^S Anode Stud mount Disk type Figure 14.11.— Common thyristor configurations. Figure 14.12.— Heat sinking of disk-type thyristors. 350 The heat-sink cooling can be by either air or water. Single disk thyristors presently have average forward-current ratings up to 7,500 A. Their relatively low cost has made solid-state motor control competitive in areas where prior to 1970 only electromechanical arrangements were consid- ered applicable. DIRECT CURRENT APPLICATIONS There has been excellent success in applying solid- state or static control to dc equipment, especially equip- ment with less than 100 connected horsepower (4, 17, 35, 39). Examples include battery on-track and off-track vehi- cles as well as underground face equipment. The control has primarily used chopper circuitry, and two important advantages have been achieved: increased motor life by limiting armature current during acceleration, thereby reducing brush and commutator problems, and signifi- cantly reduced drive-train maintenance. General mechan- ical problems are reduced because motors have controlled acceleration. Even during plugging, which is the worst shock-loading instance, motors have controlled decelera- tion, reversal, then acceleration. For battery-powered ve- hicles, battery life is increased, since peak current demand is reduced and power is generally used more efficiently. (Applications in battery chargers are given in chapter 15.) Static control of dc motors is also applied in surface excavators, hoists, and conveyor belts, where there are variable-speed requirements and substantial demand fluc- tuation (39). The power supplied to surface excavators, for example, can swing from 200% demand to 120% regener- ation in 1 cycle. Currently, static control in surface excavators is lim- ited to machines of 20-yd 3 capacity or less. The form is similar to the conventional Ward-Leonard system (20, 39), but the motor-generator set is replaced by thyristors in a bridge configuration with feedback control. This change does improve operational performance, but it has a detri- mental effect on the ac distribution system to the machine (31). High-amplitude voltage transients and significant electrical noise commonly found on mine trolley systems have seriously hampered the adaptation of this technology to trolley locomotives (39). However, there has been some limited success with chopper control (4). It is interesting to note that problems have been encountered on practically all solid-state equipment receiving trolley power. An outstanding success of static control has been the ac-dc drive in underground ac face equipment. The design maximizes power-distribution efficiency, uses the traction advantages of dc series-wound motors, and has improved the performance available from thyristor control (35). Diagrams for systems used on shuttle cars and continuous miners are given in figures 14.13 and 14.14. In both, ac power is supplied through the trailing cable. After the main circuit breaker, ac is supplied through conventional contactors to the pump-conveyor (hydraulic) motor, the cutting motors, and an on-board three-phase transformer. The transformer secondary supplies ac for the dc systems, involving thyristor control of the traction motors. 3- phase transformer ac switching Rectifier dc switching Control timing and feedback ^ [ Traction \ ~^~V motor J ^ / Traction \ ^^ motor I ac power switching 3- phase power input Main circuit breaker Figure 14.13.— Block diagram of ac-dc shuttle car. 351 3- phase transformer ac switching Rectifier I dc switching Control timing and feedback ( Traction \ "*"\ m otor J _^J Traction \ ^^ motor J ac power switching Figure 14.14.— Block diagram of ac-dc continuous miner. ALTERNATING CURRENT APPLICATIONS Full variable speed cannot be accomplished with straight current or voltage control on ac induction motors. However, as motor speed is a function of the applied power frequency, speed control can be obtained by varying this frequency (12). A simple representation for one type of variable-frequency ac drive is shown in figure 14.15, and an elementary inverter circuit is given in figure 14.16. The incoming ac is rectified then filtered by capacitance to provide a high-quality dc. An inverter then converts the dc back to ac. By controlled switching of the thyristors, an alternating square-wave signal of the desired frequency can result. This can be used directly or filtered, as in the transformer in figure 14.16, to provide a sinusoid. Three- phase ac output can be obtained by employing three inverters and firing the inverters so that 120° timing is available. A direct adaptation of this technique has been used on production mining shovels as shown in the block diagram in figure 14.17 (20). This is reported to give better perfor- mance than Ward-Leonard dc motor control since machine characteristics are a function of the power electronics and not the motor. Another application of thyristor control of ac induction motors is in conveyor starters. Across-the-line starting of three-phase squirrel-cage induction motors on in-mine belt conveyors can be very detrimental. Belt conveyor installations can call for rather high horsepowers, and the resulting high starting current ac source -U Rectifier T Filter dc i I I i> Adjustable- period sinusoidal ac I Inverter V Filter ij Motor \T Rectified Adjustable-period Adjusted dc square wave speed output Figure 14.15.— Simple variable-frequency control. = v dc *- ■Hr Figure 14.16.— Elementary inverter circuit. 352 On-board transformer Dynamic-braking control VFD INV RECT DYN BRK PM HM CM SM KEY Variable-frequency drive Inverter Rectifier Dynamic brake Propel motor Hoist motor Crowd motor Swing motor Figure 14.17.— Use of variable-frequency drive on production mining shovel. (three to eight times full load) can produce protective- relaying problems and large voltage drops. The latter can cause a motor torque decrease, which in turn can hamper belt conveyor acceleration. In the past, wound-rotor motors with step starters were used extensively on belt conveyors to overcome these problems. The step starter is essentially a bank of resis- tors that is connected to the rotor winding (see chapter 6) and allows the motor to start with a high resistance- to-reactance ratio for limited starting current and high starting torque. The external resistance is then decreased in steps, each decrease resulting in an increase in motor speed. At full speed, all external resistance is shorted out, and the wound-rotor motor operates like a low resistance- to-reactance design with corresponding high efficiency. Step starters for belt conveyor applications require large- capacity switching and contacting equipment. Histori- cally, these electromechanical components have been maintenance problems, and the brushes and slip rings of the motor can be a continual source of difficulty. On the other hand, controlled acceleration and lim- ited starting current can be achieved using squirrel-cage induction motors with solid-state starters. Compared with wound-rotor motors, motor and starter maintenance is lower, and the following advantages are gained: Reduced belt and splice tears; Decreased stress on mechanical power-transmission components, resulting in increased life; Elimination of some mechanical components, such as extensive gear trains, clutches, and so on; Decreased belt slippage, thereby reducing belt burns and removing impulse stresses; Custom design for special application; and Increased motor life. All these features have been validated by experience. Control Systems All three-phase solid-state belt starters in common use today employ reduced-voltage motor starting. The technique is based on the principle that torque developed at the motor shaft is proportional to both the rotor current and the square of the terminal voltage. Therefore, if the terminal voltage is reduced to a given level to correspond with a desired torque value, then motor current is limited. Thyristors are used to control power turn-on to the ac motor, and voltage is reduced by not firing the thyristors until some angle past the source voltage zero. This effec- tively reduces the average voltage across the motor termi- nals to a value less than the source voltage. Firing control is by one of two means: open loop or closed loop (2, 5, 31). Without feedback, the open-loop systems are the sim- pler designs, and the terms voltage ramp and current ramp are often used to describe these systems. The controller element (G in figure 14.9) brings the voltage or current up to maximum in a predetermined time period independent of the motor conditions. The maximum value of current is not limited by the starter, but rather by the line voltage and motor characteristics. This limits the use of open-loop control to motors or power systems where the starting current does not create problems. In the closed-loop systems, some reference signal or feedback from the motor (H in figure 14.9) is compared with a reference to adjust the controller output. The result 353 is more control over the motor starting characteristics and allowances for loading or operating conditions. In the case of static belt-conveyor starters, and feedback signal typi- cally corresponds to either motor line current or motor speed. The line-current feedback control schemes can be divided into current-limit or current-regulated types. Motor-speed feedback techniques are also called linear acceleration or tachometer control. The basic current-limit scheme is shown in figure 14.18. Feedback control is used to compare an adjustable reference signal (the current-limit setting) with a signal that represents motor current. Motor current is monitored by current transformers, preferably in all three lines (2). The torque being developed at the motor shaft is then computed by the starter circuitry from the current being used. As a result, the current is limited to a preset value by the thyristors, and the resulting belt acceleration is nonlinear (almost logarithmic) and is a function of belt load. When the belt is empty, a full-voltage start can occur. In current-regulated control, the problem of a full- voltage start is removed by the addition of a second reference supplied to the control-system summing point. This is an adjustable ramped reference signal, which is usually a linearly increasing voltage with time. The starter circuitry now restricts motor current to rise over a preselected time period (the acceleration-time setting in figure 14.18) to the current-limit setting. Motor accelera- tion is smoother than current-limit control but is still somewhat nonlinear. In basic linear-acceleration starters, a tachometer generator is placed on the motor shaft, and its output provides either a feedback signal proportional to motor speed, which is compared with a ramped reference, or a feedback signal proportional to motor acceleration, which is compared with constant preset reference. As a result, motor voltage is adjusted to limit current, such that a specified rate of acceleration is obtained. Belt acceleration is linear and rather constant regardless of load, this is, whether the belt is empty or full. Some models combine the linear-acceleration feature with overriding current- limit and current-regulation control, as shown in figure 14.19. Thus, the starter tries to linearly accelerate the motor, but the current-regulation control keeps the rate of current rise within preset limits and the current-limit control keeps the maximum value of current below a certain level. Controlled deceleration is also available in some starters, but the deceleration time must be longer than that for the drive mechanical inertia. The start-stop circuitry is often relay controlled. (This is not shown in figures 14.18 and 14.19.) When off, relay contacts clamp the thyristor gates and the input of the control-system amplifier. Under a start command, the relay contacts sequentially unclamp the gates, allowing the control system to start the acceleration cycle (2). Because the thyristors do not physically disconnect the load, a circuit breaker is also provided on the incoming line for this purpose as well as for short-circuit protection. Control System Design Considerations When starting an induction motor on a belt drive, three major areas need to be considered: the power system, the motor, and the mechanical equipment (speed reducers, the belt, and so forth). From the standpoint of power- system protection and optimum operation, the motor should be started with as little current as practical to Current- limit setting (^ — (§H>- Firing control Acceleration- N time ' setting v ! Ramp j "generator' TTT L-i L 2 L 3 Figure 14.18.— Simplified diagram of current-regulated static belt starter. Motor Current- limit setting Acceleration- ,, Amplifier time ?\ » setting v_^ To other thyristors Jill coupling i Tachometer Q irmg control HT Tachometer signal processing L, L-, Figure 14.19.— Simplified diagram of linear-acceleration static belt starter. minimize the overcurrent and undervoltage effects on the power system. With motor protection and operation taken as the priority, the philosophy would be to bring the motor up to full speed as quickly as possible. The reason here is to prevent insulation breakdown caused by rotor overheat- ing. The optimum situation for protection of the mechan- ical equipment would be a smooth, easy acceleration of the belt. This reduces the wear and tear on the gears and the belt from excessive or uneven torque. Unfortunately, because the design considerations are mutually exclusive, all three areas cannot be completely protected. Protection of the power system and mechanical components calls for an extended low-level starting cur- rent, whereas protecting the motor from overheating re- quires a rapid-rise high-level starting current. The area that is given priority helps determine which control sys- tem should be used (31). When solid-state belt starters were first introduced, protection of the power system and conveyor components was the primary concern. Thus, early starters used a current-regulated or linear-acceleration scheme for smooth starting of the belt and an overriding current limit 354 to protect both the power system from overcurrent situa- tions and the conveyor components against overtorque problems. Either control system gave the same basic static-starter advantages given previously. Linear- acceleration controls had a special advantage in installa- tions where belt length was constantly changed, such as for panel belts, as little or no adjustment of the starter circuitry was necessary. With current-regulator control only, adjustment may be required for any conveyor change, but current-regulator controls do have the advantage of simplicity. Most manufacturers offered both options in their equipment. Many of these early static starters worked well, but some problems occurred because of initial design flaws. Foremost was the situation where either the current limit was set too low or the belt was overloaded. In either case, the motor would not reach full speed before the rotor overheated because of the somewhat limited acceleration imposed by the current limit, and either the thermal protection of the motor or a maximum current time limit would shut down the motor. Another common problem area was strictly associated with linear-acceleration mod- els: high maintenance requirements of the tachometer generator and its wiring. These problems can directly produce conveyor belt downtime and, in turn, a large loss-of-production cost for the mine. Thus, a recent trend with many manufacturers has been toward more reliability and a simpler control system, and the common design is an open-loop voltage ramp scheme without a current limit. Typically, the volt- age starts at roughly 60% of the line voltage then gradu- ally increases to 100% in a preset time limit. Because there is no current-limiting capability, thermal overload sensing in the motor is almost a necessity, and simple circuitry for tripping the main circuit breaker when the overcurrent situation has exceeded a specified time limit is usually recommended. This design is reported to have a lower breakdown rate than other more complicated sys- tems, and it also allows for full locked-rotor torque and full locked-rotor current when needed by the conveyor. How- ever, the simplicity does negate some of the advantages of soft starting, particularly limiting starting currents and voltage drops. Motor Designs There are four concerns when selecting a suitable motor and motor characteristics: providing sufficient starting torque, drawing symmetrical line current with minimum disturbance to other equipment, optimizing the thyristor characteristics, and providing an acceptable stress level on motor insulation under any condition (39). In terms of characteristics, NEMA design B appears to be a good match for linear-acceleration control because of its greater locked-rotor time, and NEMA C seems best suited for open-loop and current-regulated applications from its greater starting torque at a given current (31). However, the main consideration is thermal because the motor must accelerate to full speed for normal operation. Energy continues to be stored in the motor until the internal fan velocity can dissipate the heat faster than it is generated. Hence, NEMA C motors are often recommended for all control schemes because less energy output is needed to provide the same output starting torque. Some manufac- turers have even recommended the NEMA design D be- cause they feel the advantage of its very high starting torque outweighs the disadvantage of its inefficient full- speed characteristics. Some motor manufacturers are producing squirrel- cage machines specifically designed for solid-state start- ing. They have higher quality insulation, larger fans, larger rotors, and are built on a larger frame per given horsepower. These motors can have an energy-in-heat-out balance at 75% of full speed, whereas for the standard motor, it can be as high as 90% of full speed (39). Experience has shown that conventional thermal- overload relays do not provide adequate protection during the long starting times caused by static control (2, 30). Overtemperature detection devices installed in the motor are the best way to provide thermal protection. Of these, solid-state temperature sensors, installed internally in the motor windings, with a load-current detection backup, appear to provide the best technique (2). Thyristor Configuration Both thyristor configurations, unidirectional and bi- directional, are available for static belt conveyor starters, but there are difficulties with unidirectional control. Some problems are caused by the higher harmonic content that unidirectional control adds to line current when in the control mode. When applied to static induction motor starters, the harmonics tend to create excess heating but mainly during acceleration. If sufficient cooling time is allowed between starts and acceleration times are mini- mized, the heating problem is reduced considerably but not eliminated (5). However, the main difficulty with unidirectional con- trol is allowing dc to flow in the motor, and two situations illustrate the problem (5). During a motor ground fault, the diodes can rectify current, and dc can flow through the grounding resistor and grounding conductor, and not be detected by the zero-sequence ground-fault relaying. Thy- ristor failures are rare but are always in the shorted mode. If thyristor fusing is not available, there also exists a low-impedance path for dc through the motor windings and the other two diodes. Either case is a problem whether the starter is on or off. With the harmonic and dc difficulties presented by unidirectional control, it appears that bidirectional control is better despite its high cost. Some additional advantages are gained (39). Two thyristors are always in series from line to line, and hence, less stress is given per thyristor by transient and long-term overvoltages. Because of symme- try, smoother control is provided under varying conditions, and finally, motor windings and cable power conductors are near neutral potential when the thyristors are off. Firing Circuits When discussing the basics of thyristors, a single gate pulse was shown to fire the device (figs. 14.3, 14.20A). However, in practice, there are many instances where a single gate pulse would fail to fire the thyristor, for example, if there is low ambient temperature or low line voltage or when the devices themselves are old (39). Two types of firing systems overcome this problem: sustained pulse (fig. 14.20B) and multi-pulse (fig. 14.200. Generally, a pulse transformer is used in the multipulse system as isolation between the power circuit and the control circuit. Numerous pulses, each capable of turning the thyristor on, are applied to the gate to keep the thyristor on during 355 'g' I i . i oc IT 27T OC + 2TT 3ir Ir.A cut _L ■cut 'g* TT 2TT QC+2TT 3TT cut TT 27T CC+2TT Figure 14.20.— Types of thyristor firing pulses. 3TT its intended conduction period. The other technique, also termed the dc firing system, maintains a continuous gate current during the desired conduction period. This helps ensure that the thyristor will turn on, and turn on completely by continuous stimulation of the gate. Two parameters of the firing pulse are important: the level of the current and the current duration. If the pulse is only of minimal current and duration, it may cause conduction only in a limited area of the thyristor or no conduction at all. In particular, a small conduction area coupled with a high rate of change (di/dt) of anode- to-cathode current can result in concentrated heating and possible failure of the thyristor. To ensure conduction of the whole interface of the thyristor, a technique called hard firing is used in conjunction with either dc or multipulse firing. It consists of a high-level initial pulse with a steep wave front and enough duration to operate the thyristor at near the maximum input power level of the gate. The high initial gate current floods the conduc- tion region of the thyristor and turns it on completely. After the device is conducting, a sustained dc pulse or multipulse keeps the thyristor on even if the anode- to-cathode current approaches zero. This prevents thyris- tor turnoff during the critical end of the conduction cycle. Hard firing, thus, provides more consistent operation and reduced failures of thyristors, and allows the use of off- the-shelf devices (9, 39). Thyristor Ratings and Protection Thyristors are susceptible to damage from overvolt- ages, overcurrents, and rapid changes in voltage (dv/dt) or current (di/dt) with respect to time. Because these are rather common occurrences in mine power systems, thy- ristors should be protected from each of these phenomena if premature failure is to be prevented. The best way to protect the thyristors against the effects of high di/dt, which has already been mentioned, is proper design of the firing pulse. Two avenues are typically used to provide the other protection, thyristor ratings and protective circuitry, and a typical protection arrangement is shown in figure 14.21. I 2 t fuse 0.25AF 50 A MOV ^ MOV / Optional Q indicator /T^ lamp - t f Gate pulse to control (di/dt) Figure 14.21.— Thyristor protection for static belt starters. It is standard practice to select a thyristor voltage rating at 2.5 times the nominal line-to-line system voltage. For bidirectional control, this is quite adequate consider- ing the 5-pu utilization transients discussed in chapter 11. However, the thyristors themselves are a source of abnor- mal transient overvoltages, and some engineers prefer to use a device voltage rating no less than 3 or 3.5 times the system voltage (5, 39). Calculation of the required thyristor continuous- current rating from the load current should be easy if the thermal properties of the heat sink are known. Yet during starting, which is the worst stress case, the device is called upon to deliver much more current. The conventional approach for overload protection of the thyristor is a 30-s (overload) rating of 300% continuous current. Some engi- neers do not believe this is adequate, because the induc- tion motor starting current may be as high as six to eight times the running current; hence, certain manufacturers have selected thyristors with 300% of the continuous- current requirements and a 30-s overload rating at 500%. Others feel that it is better to have an overload rating 356 based on horsepower, 25 s at 600%, or 100 s at 300% of continuous current (39). Thus, the thyristor current-rating selection is not a straightforward matter, and special consideration is sometimes required for individual cases (5). Some short-circuit protection is afforded by oversizing (300%) the continuous-current rating. However, past prac- tice was to provide additional protection with semiconduc- tor protection or I 2 t (let-through energy) fuses with one fuse in series with each device. The philosophy was that, even though these fuses were expensive, the replacement cost for the thyristor was even more so. The fuse continuous-current rating was matched to that of the thyristor, but the I 2 t rating was less than the thyristor I 2 t capability. There were instances where the thyristor continuous-current rating was set at 150% of needed and then the I 2 t fuse was matched (39). However, this proxim- ity was found to lead to nuisance fuse activations, espe- cially during acceleration. Recent practice for overcurrent protection of the thy- ristors is to eliminate or oversize the I 2 t fuses. When the fuses are eliminated, the thyristors are sufficiently over- sized to withstand most long-duration situations, and thermal-overload sensors are used on the thyristor heat sinks. In cases where the fuses are used, the fuse current limit is set very near the failure point of the thyristor. Both of these schemes are intended to reduce starter downtime due to blown fuses, even to the point of sacrific- ing the main power thyristors. The general feeling among mine operators is that the production saved by eliminating false fuse tripping is worth an increased liability of the thyristors' failing. This situation has become justified because of increased production cost associated with belt downtime and the decreased cost of the thyristors. A high rate of forward-voltage increase can turn a thyristor on, even with a zero gate current (9). Inductive loads, such as induction-motor belt drives, can present such dv/dt problems (5). The common solution is RC snubber networks across each pair of devices as in figure 14.21. Typical values for R s and C s are 50 fi and 0.25 fiF, but values must be selected so that the dv/dt allowed is less than the minimum specified by the device (around 100 V//ts). To ensure that transient overvoltage does not de- stroy a thyristor, metal oxide varistors (MOV's) across each device are sometimes specified. The maximum crest al- lowed by the MOV is coordinated with the maximum voltage rating of the thyristor, allowing the usual 20% safety margin. The preceding paragraphs have not only introduced the basics of solid-state motor control but also indicated the special difficulties of applying static-starter concepts to mine belt conveyors. This is one firm example of the benefit of technology to industry. Even with the increased internal complexity, the end result has been an overall increase of belt-conveyor system reliability. STATIC PROTECTIVE RELAYING Most engineers define the term relay as an electrically controlled, usually two-state, device that opens and closes electrical contacts to effect the operation of other devices in the same or another electrical circuit (15). Historically, one important relay use has been the protection of people and electric circuits from electrical hazards. The operation of an electromagnetic protective relay was presented in chapter 9 but is repeated in figure 14.22 (36). This partic- Station bus CT u - b c — < 1 1 1 i — ( 1 nr JL nr nda 4_- i i i i i i i PT Circuit breaker C Trip C coil A ° [_ -o o- \ i £■ i eco s p Protected circuit Relay r y otential bus Figure 14.22.— Protective-relay connections. ular relay uses two actuating quantities (voltage and current) that directly affect the status of the relay con- tacts. Whenever current and/or voltage exceeds a prede- termined level, the current sensor and/or voltage sensor (that is, CT and/or PT) outputs cause the relay to close its contacts through electromagnetic attraction or electro- magnetic induction. The closed contacts permit current to flow through the trip coil, tripping the circuit breaker. Electromagnetic-attraction relays in common use are the solenoid, clapper, and polar types (38). The typical electromagnetic-induction relay is the induction disk. The basic concept in the design of solid-state relays (again also called static relays) is to replace the mechani- cal contact device with a solid-state device. The solid-state device is inserted in the trip coil circuit and controlled by the sensor circuit. When unactuated, the solid-state device acts as a very large resistance in the trip coil circuit, limiting the current through the trip coil to a very small value (known as leakage current), which is incapable of tripping the associated circuit breaker. When actuated, the solid-state device acts as a very small resistance, which allows ample trip coil current, thus tripping the circuit breaker. The solid-state devices commonly used are the transistor, the thyristor, and the triac. OPERATION OF SIMPLIFIED SOLID-STATE AND HYBRID RELAYS The electromagnetic relay of figure 14.22 is repre- sented schematically by figure 14.23, where the current and voltage input have been replaced with a manual push button, and the trip coil (load) circuit is supplied with a 120- V, 60-Hz power source. The static relay differs from this in that the contents of the dashed box of figure 14.23 are replaced by the semiconductor device shown schemat- ically in figure 14.24 (16). The load current now flows through the common terminal of the semiconductor de- vice, which is common to the trip (contact) and sensor (control) circuits. 357 Figure 14.23.— Simple electromechanical relay. Semi- conductor contact device \ Control current input Load Control Common Control current Output ( contact ) Figure 14.24.— Simple static relay. If the static device is an npn transistor, the circuit of figure 14.25A results. Since the transistor is a current- controlled device, if zero input (control) voltage is applied to the transistor base, no current will flow into the base. Because base current is required for current flow from collector to emitter, no current will pass through the load. Hence, the contact-circuit voltage supply will not be dropped across the load, and the voltage must therefore be across the collector-emitter terminals of the transistor. A positive input voltage applied to the transistor base causes a positive current to flow into the base. This base current is the controlling factor that permits a large collector- to-emitter current to flow through the transistor. When sufficient base control current is supplied, the collector current will increase until essentially all contact-circuit voltage is across the trip coil (load). Thus, the switching characteristics of figure 14.25B are obtained. A drawback to the circuit in figure 14.25A is the lack of electrical isolation between the control and the trip circuits. One isolation method is to use the photon-coupler circuit shown in figure 14.26 (16). Here, the transistor performs the same function as that in figure 14.25A except that it is light controlled, with a high-intensity light (from the LED) acting as a large base current. The light im- pinges on the phototransistor base region, allowing cur- rent to flow from collector to emitter. Without an input to the LED, no light is produced and consequently no collector-to-emitter current flows. Replacing the transistor with a thyristor results in the circuit of figure 14.27 (16). This device has been described earlier in the chapter; thus, its operation here should be clear. The triac, which is a contraction for triode ac semi- conductor, is a bidirectional solid-state device that acts 'load 'input Current- - Load 1 limiting resistor JC t Control V 1 1 Contact I circuit voltage (V) « ransistor 1 'load 'input A Circuit Control Control voltage voltage applied removed B Switching waveforms Figure 14.25.— Transistor used as relay. Current- limiting resistor I — VW\r— Figure 14.26.— Optical transistor as relay. V load ^WvV^ v input| Control Load V, scr rms Figure 14.27.— Thyristor used as relay. like two thyristors connected back-to-back as shown in figure 14.4A. The triac provides full-wave voltage control in one solid-state structure with only one gate control, as shown in figure 14.28A. The load-voltage characteristics, as shown in figure 14.28B, are very much like those exhibited by the thyristor. The single structure has heat- dissipation limitations, restricting the triac to small- current applications (9). 358 V. i i v load KAAAA Triac MT2 ^MArTlM ' input MT1 Load load 'triac V ( triac - LU (9 i q V vJ2-h < \ ^v^rate 1- \ _l O -\ \ N£ n - > \ \ \5h >^ _1 1 .8 — \ ^w — _l HI V.h \ 3h \ ^ o — i 1-h * rate 1 i *." Final volts 1.7 \ ,' — 1 • \,' - 1 6 1 . I.I.I , 1 i 4 6 8 TIME, h 10 12 Figure 15.2.— Voltage per cell of a typical lead-acid battery with varying continuous rates of discharge. UJ cr a: CJ o z o < X o 140 120 100 80 60 40 20 - 1 1 1 1 1 1 1 - Charging current^ — Cell voltage^ ^/\ - - 1 1 1 1 1 1 1 1.120 2.7 2.6 2.5 2 4 uj < 23 b £..0 Q > 2.2 j UJ o 2.1 2.0 120 too 80 o" ul 60 oc < X 40 lj cc u_ 20 -,100 °„ -1 80 lu Figure 15.3.— Typical charging process of cell from 18-cell, 725-Ah battery. Ambient temperature, 77°F. Specific gravity temperature adjusted. 370 quantities of current become available for electrolysis because of the higher state of charge in the cell. A certain amount of gassing is a necessary conse- quence of a good charge, which explains why water must periodically be added to batteries, but excessive gassing or overcharging causes damage to the plates, excessive water consumption, and excessive hydrogen emission. For this reason, the rate of charging current must be controlled as the battery charges. The large amounts of H 2 and 2 released during excessive gassing cannot be detected by mine personnel. However, the H 2 and 2 are sometimes accompanied by amounts of H 2 S0 4 released into the mine atmosphere, which is easily identified by smell. Undercharging a battery is also harmful, even if it is only practiced occasionally. Insufficient charging leads to a gradual sulfation of the negative plates, which eventu- ally causes a reduction in battery capacity and life (7). A similar situation can occur when batteries are left stand- ing in an uncharged state for long periods (18). BATTERY MAINTENANCE Proper battery maintenance has a very significant influence on battery life. Battery manufacturers normally include a recommended maintenance program with their batteries, based on specific gravity levels and equalizing schedules. Equalizing is the process by which all the cells in a battery are brought to the same voltage. In lead-acid batteries, the electrolyte specific gravity is a function of the state of battery charge. Consequently, a plot of electrolyte specific gravity versus discharge depth for a particular battery is important. All lead-acid batter- ies require periodic equalization, but excessive equaliza- tion can cause unnecessary battery deterioration. The following guidelines can be used to develop a good battery maintenance program, but see also references 3-4, 6, 8-9, and 14. There are three groups of activities in the program: those that should be performed daily or during each charge period, those that should be performed weekly, and those to be performed approximately once every 3 months. Accurate records should be kept of all maintenance activities for each battery as these are a convenient way to monitor individual battery perfor- mance. Deteriorating battery conditions can thus be de- tected before the battery becomes a safety hazard or the source of costly downtime. Daily battery maintenance activities should include the monitoring of one battery cell, which is called the pilot cell. Any battery cell can be used as the pilot. The following cell characteristics should be recorded during each charge: specific gravity before and after charging, electrolyte temperature before and after charging, and the water level in the cell. If any of the pilot parameters falls outside those specified as acceptable by the manufacturer, all cells should be checked and corrective action should be taken. Other daily maintenance activities should include checking the battery for physical defects such as cracked cell plugs and ensuring that the charger output voltage is correct. The weekly maintenance program includes checking all battery cells for proper water level (the water consump- tion of a good battery is generally equally distributed among the individual cells) and routine cleaning of the battery tops. Every 3 months, it is good practice to take a complete set of cell voltage and specific-gravity readings at the end of an equalizing charge to ensure that these parameters meet manufacturer specifications. CHARGERS The current supplied to a cell must be dc. It would be possible to obtain this current from a trolley-distribution system, but this source has a serious drawback: it is difficult to obtain the precise current requirements for charging. As a result, most mine batteries are charged from the ac distribution system, using a transformer- rectifier combination. Mercury arc rectifiers were used previously (10), but selenium or silicon rectifiers are now universally employed, and silicon diodes are considered to be the industry standard. Transformers generally have isolated secondaries and are most often three phase, although single phase can be used. Rectifiers are usually in a full-wave bridge. Several methods can be used to control the rate of charge. Most are designed to initiate the charge at a fairly high current or starting rate, commonly 20 to 25 A per 100-Ah capacity. The level is then tapered off to the finish rate as the battery charge is restored, for example, 4 to 5 A per 100 Ah, as shown in figure 15.3. Some chargers reduce the starting rate to the finish rate in one step when the charge is about 80% complete, but control devices that taper the charge rate are more common in mining. Taper chargers are either active or passive, and in both types the charge is usually stopped automatically when full charge is reached (7). The most popular passive system for taper charging employs some value of ballast resistance placed in series with the battery (32). This resistance limits the initial charge current and gives a relatively flat current-versus- time curve throughout the charging cycle. The advantage of this method (sometimes termed modified constant poten- tial) is its simplicity. Its disadvantages include the fact that the ballast resistor dissipates a rather large amount of energy and the charge rate does not coincide with that considered optimal for lead-acid batteries. A variation of the above system uses a timer to switch additional resis- tance in series with the battery, thereby reducing the rate of charge at some point in the charge cycle. A characteristic common to all active systems is control of the charging current by a feedback system that samples the battery voltage during charging. The simplest active system in taper chargers consists of a voltage- controlled relay that switches additional resistance in series with the battery at a cell voltage of 2.37 V. This and all active systems must have different voltage thresholds for batteries with different numbers of cells. Another active method, shown in simplified form in figure 15.4, utilizes a saturable reactor type of voltage transformer to feed the rectifier. The saturation level of the transformer core is controlled by a dc applied to a winding on the core so that the secondary voltage is regulated by voltage-controlled current feedback from the battery. Output regulation usually begins at a battery voltage of 2.37 V per cell and is varied to zero current at the end of the charge cycle (9). A relatively new method uses thyristors to replace half of the silicon diodes in the full-wave bridge, as shown in figure 15.5. The firing angle is determined by a feed- back circuit that senses battery voltage (17). The circuit given is for a single-phase charger. Three-phase chargers would have three SCR's in the full-wave bridge with three firing boards, but the control circuitry would be much like that for single phase. Although many manufacturers and users have had success with full-wave rectifier configura- tions, in some designs a phase reference for firing the 371 Transformer Core O-d Main contactor Saturable reactor Timer Figure 15.4.— Simplified schematic of saturable-reactor charger. I J Battery :=: being i charged m K h m r\ r\ r\ r\ Single-phase line voltage SCR bridge output voltage ■Battery voltage Firing board output pulse Current output Figure 15.5.— Simplified schematic of single-phase thyristor charger. SCR's cannot be obtained with a full wave because of negligible ripple. As a result, a half- wave must be used and this gives less efficient rectification and sometimes produces a chopped waveform, which is not entirely suit- able for battery charging. The last active method employs a ferroresonant trans- former and bucking coils (13). For the conventional trans- former, as modeled in figure 15.6, any change in primary voltage produces a corresponding change in secondary voltage, which is maintained until saturation commences, the knee shown in the transformer magnetization curve of figure 15.7. This is an advantage for most power applica- tions since transformer operation approaches ideal when used in the linear portion of the curve (to the left of the knee). However, there are many instances where fluctua- tions in the secondary voltage due to normal primary- voltage changes are unwanted, and here the secondary voltage must be regulated, or remain reasonably constant under a specific load condition. A more constant but distorted secondary voltage could be obtained by driving the transformer into the saturation region (to the right of the knee in figure 15.7). Here, any change in the primary- voltage magnitude would cause only a small change in the Figure 15.6.— Two-winding transformer model. x V V P1 P2 V V P3 P4 MAGNETIZING FORCE Figure 15.7.— Representative transformer magnetization curve. core flux and thus a small change in induction in the secondary winding, but this operation is unwise because primary current becomes excessive. Nevertheless, regula- tion can be obtained with normal primary current if the transformer is modified so the secondary winding is ex- posed to a saturated core portion while the primary operates under unsaturated conditions. This is the basic principle behind a ferroresonant transformer. A model of a basic ferroresonant transformer is shown in figure 15.8 (13). The leakage block of magnetic material provides a shunt path to bypass part of the flux produced by the primary winding, with the air gaps limiting the amount of flux bypass. The flux not bypassed causes induction in the resonant winding. The inductance and capacitance of this winding are selected so that additional flux is produced that is in phase with the primary flux when there is no load on the secondary winding. This increases the flux in the right core portion to about the knee of the magnetization curve. As a result, the right transformer portion will operate under saturated core conditions while the primary sees an unsaturated core. In the basic ferroresonant transformer, an increase in primary voltage will still produce a slight increase in secondary and resonant winding voltages. lb offset any increase in secondary voltage, a compensating or bucking coil, consisting of a few turns wound directly over the primary, can be connected in series with the secondary winding but with opposing polarity (fig. 15.9). Any in- crease in primary voltage will produce a proportional change in V b , which will offset any increase in secondary voltage and produce a rather constant output voltage over a specified range of primary voltages. 372 Figure 15.10 shows the adaptation of a ferroresonant transformer to a battery charger. Although a single-phase unit is illustrated, the technique can also be used in three-phase units. Some chargers do not use the bucking coils for reasons discussed later in the chapter. -^ Primary flux J2 Air gaps u; — i u^ i » Resonant <-rE ; r~l It winding ,PE \ =3 T Secondary -^^Ir' j' k Secondary ~* i ( ^&[ J C j) V c Leakage block Secondary flux Figure 15.8.— Ferroresonant transformer model. v f f Primary^ | coil 3 Ic_ ± > TV C Resonant 3 coil 5 C^ Bucking | v C^ coil : Secondary coil iut Load Figure 15.9.— Ferroresonant transformer. Active charging systems have an advantage over pas- sive systems in that feedback techniques can be used to give more accurate control of the charge rate. The most popular chargers for mine vehicle batteries use saturable reactors, thyristors, or ferroresonant transformers. Charge termination can be achieved by a timer or by monitoring the cell voltage or its rate of change to determine when the battery is fully charged. Although the success or failure of the battery-powered mine transportation system is largely a function of the operator's ability to get maximum life from the batteries, it is also a function of the safety factor involved in battery usage. The basics of battery safety will therefore be discussed in some detail in the following sections. CHARGING STATIONS Special charging stations are required in under- ground mines to charge vehicle batteries. These stations must be designed and constructed to meet specific venti- lation requirements, but because mining methods and plans vary widely, it is extremely difficult to define a rigorous set of guidelines. Hence, this section presents the principles that underly charging-station construction and explains the requirements that must be met. The first problem that must be addressed at battery charging stations is dissipation of the gas produced by the charging operation. It has already been stated that hydro- gen gas is liberated at the close of the charge cycle. Although modern chargers are designed to prevent exces- sive gassing by automatically dropping the charge rate to a very low value when a specified cell voltage is reached, it is impossible to charge a battery properly without To ac power supply ^ Line contactor ._ Timer T contacts s*\ Line ( ) contactor V-^ coil 6 Timer motor Primary (£_ coils ac fuse HI DZ_Th Line contactor ^-pv '-r - ^ ' Resonant coil Overvoltage suppressor Ammeter Figure 15.10.— Ferroresonant battery charger. 373 producing some gas. This gas is explosive and must be diluted to render it harmless in the mine atmosphere. The traditional method of achieving this is by forced ventila- tion of the charging room. The hydrogen concentration must be kept below its lower explosive level of 4%. Natu- rally, because of the catastrophic nature of underground explosions, hydrogen concentrations are closely controlled by Federal regulations. These limit the permissible con- centration of hydrogen in coal mine atmospheres to 0.8% by volume, which provides a safety factor of 5 (30 CFR 75.301.5). This limit is analogous to a maximum allowable concentration of one-fifth of the lower explosive level of methane. In order to maintain these requirements, it must be possible to monitor or estimate the evolution of hydrogen in the battery. By assuming a typical charge characteris- tic, manufacturers have made it possible to estimate hydrogen evolution from the number of cells and rated ampere-hour capacity. Table 15.1 lists such formulas from three different manufacturers. It can be seen that the H 2 evolution calculated from these equations is fairly consis- tent, ranging from 0.0024 to 0.0028 ft 3 per cell ampere- hour. These figures are applicable to the latter stages of the charge cycle at cell voltages of 2.37 V and greater. These estimated values have been confirmed by compari- son with actual concentrations and charging room airflow rates obtained by in-mine survey {19). _J Fresh-air , entry => Charger 2\ Fresh-air flow Corrugated metal siding */////////////////, Scoop tractor 140 ft 3 /min minimum i Roof support crib 1 •Blocking <^ Return-air entry Ventilation' opening L Y///////A /////////////////////////////////////// Figure 15.11.— Plan of underground charging station. Table 15.1.— Formulas to estimate hydrogen evolution Formula for ft 3 H 2 H 2 evolution, Manufacturer liberated in last ftVhper References 3 h of charge 1 cell Ah C & D (number of cells) (Ah) (0.0024) 0.0008 3-4 Exide (5) (Ah/100) (number of cells) .0008 6 (0.016) KW (number of cells) (Ah) (0.002948) .. .0007 17-18 For KW, last 4 h of charge. The second requirement for charging stations that is stipulated by Federal law is that all underground battery charging stations must be "housed in fireproof structures or areas." To render the station fireproof, it is common practice to line the charging area with corrugated metal siding so that all exposed coal (which has previously been rock-dusted) is covered. Concrete block is used as a lining in some instances but is more costly. The third Federal stipulation is that "air currents used to ventilate structures or areas enclosing electrical installations shall be coursed directly into the return." This implies that the charging station must have a sepa- rate split of fresh air. To comply with this ventilation requirement, stations are frequently located in unused crosscuts immediately adjacent to return-aircourse en- tries. A small opening is made in the ventilation blocking so that fresh air passes over the station and dumps immediately into the return. Figure 15.11 illustrates a charging station that meets Federal regulation. Many other configurations are possible. When each charging station is designed, it must be verified that the hydrogen concentration is below the maximum allowed. A simple but effective approach is to establish a worst case airflow quantity for each battery on charge, then substitute the safe condition by an actual airflow measurement of the charging station. This mea- surement can be made with an anemometer. From table 15.1 it can be seen that batteries in the final stages of charge evolve hydrogen to the following approximate level: (5) (Ah/100) (number of cells) (0.016) = ft 3 /h H 2 . Accordingly, dilution requirements can be calculated by: (5) (Ah/100) (number of cells) (0.016) Q(ff7min) = (0.008) (60) (5) (Ah/100) (number of cells) (0.016) 60 (15.4) where Q(ft 3 /min) is the minimum quantity of air necessary to dilute the H 2 produced to 0.8%, not allowing for dilution by room volume. To approach a worst case, it could be assumed that all batteries are 120-cell, 700 Ah, which is the present upper limit in battery size. Thus, each battery in tbe charging station requires about 140 ft 3 /min of air for good ventilation. This air quantity assumes that the battery lids re- main open or removed during charging. Although this is mandatory, it is a practice that is sometimes neglected, especially in low coal. For charging, vehicle storage bat- teries are placed in strong enclosures (trays) with remov- able lids. If these lids remain closed during charging, dangerous accumulations of hydrogen are likely to accu- mulate because of the small space above the battery top. A charging battery can be expected to liberate 60 ft 3 /h per 25,000 cell ampere-hours. The result is an explosive con- centration unless forced ventilation is applied, and this is considered impractical by many. Enlarging the venting slots significantly is not an answer to the problem since it would weaken the battery box structure. The only solution is to ensure that the lids are open or removed. 374 BATTERY BOX VENTILATION The ventilation of the traction battery enclosure is not only important while charging but also during operation. Following the charge cycle, the covers are closed and the vehicle is placed in service. From chemical reactions and entrapped gas within the cells, lead-acid batteries con- tinue to emit gas for several hours after receiving a charge, be they open-circuited or on discharge (22). Con- sidering the ignitability of hydrogen, it is possible that a dangerous air mixture will accumulate. Hence it is neces- sary to provide adequate ventilation for the evolved hydro- gen in the closed box while the battery is in operation. Prior to 1945, there was little mention in the litera- ture of possible gas emissions from lead-acid batteries during discharge. In fact, many early publications stated categorically that there were no emissions under normal conditions (33), and this statement was repeated in mining literature (15). However, the possibility of gas emission did receive much attention from authorities responsible for high-capacity battery installations in such confined spaces as submarines. Robinson (22) quoted a report revealing that lead-acid batteries in confined spaces are always liable to emit considerable quantities of hydrogen and oxygen for the first few hours after charge. Although the actual quantities varied greatly from battery to battery, an 80-Ah cell, standing idle at 80°F (26.7°C) with 1.26 specific-gravity electrolyte, would probably emit 5 to 20 mL/h H 2 for about 12 h after charge. The approximate rate was found to be • Directly proportional to the battery capacity, • Doubled for each 15°F (9.5°C) rise in temperature, and • Doubled for each 0.050 unit increase in electrolyte density. As a result of antimony contamination of the negative plates, the open-circuit emission also increased with cell age. Upon discharge, additional hydrogen was evolved from the negative plates, which was thought to be a release of gas entrapped during charging. Combining the open-circuit and discharge emissions on a worst case basis, it was postulated that a lead-acid battery cell could release up to 500 mL (about 0.2 ft 3 ) of hydrogen per hour at the end of its useful life. An advisory committee on coal mining, formed in the United Kingdom during the mid-1940's, investigated such problems related to underground battery vehicles (28). The committee recommended that traction battery enclosures be properly ventilated to prevent a hazardous accumula- tion, and accordingly, a regulation for UK. coal mines was promulgated in 1949 (27). In the United States, the Bureau of Mines subsequently incorporated a similar requirement in Schedule 2G, stating that "battery boxes shall be adequately ventilated" (31). Corresponding to the 1949 regulations, a test proce- dure was made statutory for the United Kingdom (22, 29) in which • Hydrogen generation was calculated at 3.0 ft 3 /h per 25,000 cell ampere-hours; • Tests were to be made in still air; • Maximum hydrogen concentration within the box under these conditions could not exceed 2.0% with tolerance. The emission calculations assumed that all factors con- tributing to the hydrogen-emission rate, such as cell temperature, acid specific gravity, and discharge rate, were simultaneously at worst case. Still air was selected to simulate a vehicle traveling at the same velocity as the mine ventilation so that the container's natural ventila- tion would receive no assistance. The maximum specified concentration provided a safety factor of 2.0 below the lower flammable level of hydrogen (4.0%). In his research, Robinson examined simple arrange- ments of container vents that would provide sufficient natural ventilation to meet these ventilation require- ments (22). He used the fact that hydrogen has the highest coefficient of diffusion into air of any gas; because of its extreme mobility, hydrogen is very difficult to retain within a leaky enclosure. The investigation used a simu- lated battery box containing a dummy 56-cell, 288-Ah lead-acid traction battery. Hydrogen liberation for dis- charge was calculated at 2.0 ft 3 /h by the foregoing rela- tionship. While hydrogen was pumped in at this steady rate, hydrogen concentrations were measured in the space between the battery top and cover (volume of 6.1 ft 3 ). Eleven different venting arrangements in the enclosure top were used, ranging from 40 to 140 in 2 , and sealed top vents were also investigated. End-plate venting was avail- able in all tests. Robinson's findings were as follows: 1. Equilibrium between the battery rate of hydrogen emission and rate of hydrogen escape through the vents was established within 1 h after each test commenced. Thereafter, the hydrogen concentration remained almost constant. 2. Maximum concentrations with natural top venting ranged from 1.3% to 2.8%, related inversely to vent area and following a near logarithmic curve. 3. When the top vents were sealed and the side vents open, hydrogen concentration rose sharply to 5.3%, achiev- ing 4.2% in 8.0 min. However, with 31 ft 3 /min of forced ventilation through the side vents, the 5.3% maximum dropped to 1.7% H 2 . It was suggested that normal haulage speeds and mine ventilation flow rates would create a high factor of safety. The overall conclusion was that a battery box could be readily vented to meet U.K. requirements and forced ventilation was not needed. In 1959, Titman (26) reported on gas emissions from lead-acid batteries for the 45-min period after charging, supplementing his earlier study for alkaline cells. It was then suspected that gas emission rates immediately after charge might be greater than those emitted by the battery after a standing period. Titman also investigated the parameters causing variations in emissions, such as acid strength, discharge rate, and increased cell temperature. The experiments employed a new 6-V, 309-Ah traction battery. The conclusions of the research were 1. Total gas emission for similar cells varied as much as 15%. 2. Hydrogen emission rate doubled for an increase of 0.050 in acid specific gravity (1.260 to 1.360). (Note: this increase could easily occur if the battery electrolyte was not topped up (19).) 3. The rate doubled for a 12.5°C increase in electro- lyte temperature. 375 4. Overall rates were similar to those for alkaline cells. 5. Immediately after charging, hydrogen rates up to 5.0 L/h per cell were observed (corresponding to 14.30 ft 3 /h per 25,000 cell ampere-hours). 6. After 45 min, the cells were found to liberate as much as 1.3 L/h per cell (3.72 ft 3 /h per 25,000 cell ampere-hours). 7. The minimum emission rate was always reached within 5.0 to 8.0 min with the battery standing after charging. As can be seen, the findings for specific gravity were similar to those reported by Robinson. However, hydrogen emission rates were considerably higher than the 3.0 ft 3 /h testing standard: 376% higher immediately after charging and 24% higher after a brief standing period. Using these emissions rates as a basis, Titman (25) further investigated the effect of high hydrogen emission rates on typical battery box ventilation. The enclosures investigated were basically the same as those used by Robinson (22). The findings included 1. A 2.0% H 2 concentration was not exceeded until the emission rate was 10 times the 3.0-ft 3 /h standard rate. 2. After x /i h, the hydrogen became uniformly distrib- uted, with no tendency to accumulate at a specific place. 3. For hydrogen concentrations of 2.0% or less, the enclosure concentration varied as the two-thirds power of the emission rate. (This was also verified theoretically.) 4. With the high emission rates, acid-drainage holes assisted enclosure ventilation. This research has been presented in detail to demon- strate the reasoning behind battery box venting require- ments and to emphasize its importance. A direct compar- ison of hydrogen emission rates versus the top venting area of the enclosure would be advantageous, yet no direct relationship has been found. In practice, however, it has been discovered that about 25 in 2 of top vent area will allow 1.0 ft 3 H 2 to escape per hour, while meeting the U.K. testing requirements (20). As an alternative, the use of catalyst battery caps appears to be an excellent solution for difficult venting situations. These battery caps are a fairly recent introduction and are not yet used widely in mining, although they have been in use for some time in equipment ranging from torpedo batteries to personnel carriers in salt mines. They have also had some limited use in metal-nonmetal mining, such as uranium mines. In all these applications they have proved to be reliable and effective. A catalyst battery cap converts any emitted hydrogen and oxygen back into water (23). The caps have the dual function of preventing the escape of any hydrogen from the cells and restraining the loss of cell water. The technique has considerable safety advantages. No additional venti- lation is required for the tray or for the charging station itself. Battery lids do not have to be opened for charging, which is a major advantage in low coal. The explosion hazard associated with batteries is practically eliminated. Since watering of the cells is not required, no electrolyte can be spilled on the battery top, and the possibility of surface leakage (see below) is minimized. Tray corrosion from spilled electrolyte is similarly reduced. One problem with catalyst battery caps, however, is that the palladium catalyst may be destroyed if the cap is turned over. Special care must be taken not to do this during maintenance. BATTERY SURFACE LEAKAGE AND FAULTS A battery is designed to be an electrically floating system, insulated from its tray, which is at machine frame or ground potential. Yet three situations can occur that could connect the battery with its tray. 1. Current may leak across the battery surface to the steel tray; this is referred to as surface leakage. 2. A poorly insulated or damaged cable, bushing, and so forth may contact the tray, causing a fault condition. 3. A rock fall or collision may force the battery box cover down onto the battery terminals, possibly shorting one or more to ground. Since lead-acid traction batteries are quite heavy, often weighing several hundred pounds, the only material presently available for tray construction is heavy-gauge steel. Steel is a reasonable conductor of electricity, a disadvantage when it is used to support an isolated system such as a battery. After a battery has been in operation for some time, dust mixed with spilled electrolyte can collect on the battery top. If these contaminants are permitted to accumulate, a number of low-resistance paths may form between the various cell terminals and the tray. Since the terminals are at different potentials, currents tend to circulate across the battery top and through the tray, and may cause the following problems: 1. Currents circulating in resistive loops cause heat- ing. This heating is proportional to I 2 R, where I is the leakage current and R is resistance of leakage path. When R is low enough to permit a substantial current to flow, a smoldering fire may ensue. This hazard is compounded by the often explosive hydrogen concentrations that occur in the battery above the electrolyte level. 2. The presence of paths from the cell terminals to the tray can be a shock hazard for mining personnel. A miner leaning against a battery box and touching an exposed terminal could be shocked. Terminal-to-tray conducting paths present an especially serious hazard when the battery is charging if the charger design or a fault within the charger permits dc to flow in the frame or ground. 3. Circulating currents waste battery power and may reduce the amount of time a battery can be used before recharging is necessary. They also tend to increase tray corrosion. Another potentially hazardous situation can be caused by a low-resistance ground fault, for example, produced by a damaged cable in contact with the battery tray. Since the battery is theoretically a floating system, a single such fault should cause no current flow, but the effect of surface leakage paths can result in a certain level of current, depending upon path resistance. Two simulta- neous low-resistance ground faults, however, would permit extremely high currents to flow. Depending on the contact resistance of the fault, currents on the order of 10,000 A may exist for a short duration because of the low internal resistance of lead-acid batteries. A very hazardous situa- tion could result because batteries cannot be deenergized; hence, a fire caused by faults would be difficult to extin- guish until the battery was discharged. The surface leakage problem has been recognized by battery manufacturers for some time, and they constantly expound the virtues of keeping battery tops clean. An 376 example is the following excerpt taken from a typical maintenance article (11): If the battery tops become wet and dirty, or if tray corrosion is visible, give the battery a soda wash. Mix a handful of bicarbonate of soda (baking soda) in a bucket of water. Pour this solution over the top of the battery, using one full bucket per tray. Be sure vent caps are in place. Water will dry, leaving some dry soda on the battery top. It is good practice to give batteries a soda wash once a month. If a battery is accidentally flooded (acid spilled on cell tops) due to overfilling cells, give soda wash as soon as possible. Indeed, if mine batteries were always kept clean and dry, the surface leakage problems would be greatly reduced. Unfortunately, cleaning is often neglected in the mining industry, especially in low coal where batteries are so difficult to access. It would therefore be advisable to find some alternative way of reducing this problem. Most batteries now in use underground have a thick coating of paint on their trays. This reduces leakage somewhat, but the paint is prone to deterioration by chipping, abrasion, and attack by battery acid. Recently, some manufacturers have been coating their battery boxes inside and out with a tough vinyl compound sold under the brand name Plastisol. This product is readily available and can be sprayed onto any properly prepared steel surface. It appears to reduce surface-leakage problems significantly. Battery safety could be increased further if the ex- posed intercell connectors were insulated in some manner. Unfortunately, this is not a straightforward matter since any insulating coating applied to the connectors reduces their heat-transfer capabilities and the intercell connec- tors remove about 80% of the battery heat. As an option on their mining batteries, leading bat- tery manufacturers now produce a dead-top battery with totally insulated intercell connectors. This system should greatly reduce the potential for surface leakage and also remove the possibility of dangerous arcing if a tool is dropped across the battery top. Since surface leakage and ground faults are potential hazards, a method is needed to detect and isolate them. Recognizing this, Statham and Littlewood (24) developed a fault-detection system designed to be fitted to the battery or battery charger. This system, shown in the circuit in figure 15.12, is a simple way to detect ground faults that might occur between the battery and its load. One problem with the circuit is that a large portion of the battery itself cannot be protected: for instance, a fault occurring at point A in figure 15.12 would produce no current flow through the current relay. Figure 15.13 is a plot of relay current versus fault position on the battery; it is assumed that the current-relay shunt is set at 1,000 fi and that a 200- V traction battery is used. In order to increase the sensitiv- ity of their circuit to battery surface leakage faults, Statham and Littlewood proposed installing switches be- tween point X and frame and between point Y and frame (fig. 15.12). These switches would be alternately opened and closed by a mechanical or solid-state device. The upper dashed lines in figure 15.13 show the new position- sensitivity curve resulting from the modified circuit. The entire battery could be protected in this manner, although sensitivity still varies with fault position for surface leakage faults. Figure 15.12.— Circuit for detecting faults in batteries. 70- >^ Relay characteristics 9 r -70 60- >. with switches between points , ♦ - -60 <-( v. X and frame or ,♦* E 50- >v Y and frame ,♦* - -50 h- 7* ^^ ♦ Ld 40- \^ ♦ - -40 or ^V ,* or =) o ^v ♦* 30- N. y >. - -30 V ^V. * ^^. < ^V ♦* ^« f _1 20 - ^V. * ^V. m 20 Ld or \^* Original circuit yf ♦* >^ characteristics^ \^ 10- - -10 C I T i 1 50 100 150 200 ASSUMED FAULT POSITION RELATIVE TO BATTERY VOLTAGE, V Figure 15.13.— Curve of relay current for various fault posi- tions on battery. Virr and Pearson (34) devised an electronic ac injec- tion system that would provide protection sensitivity in- dependent of fault location. They utilized a system of red and green lamps to indicate a not safe or safe condition. Although Virr and Pearson recommended mounting the device on each battery, this would not be necessary for trackless battery-powered vehicles since a single device on the charger would be sufficient to prevent an unsafe battery from being charged. Since the development of surface leakage presumably takes place gradually with the accumulation of conducting material on the battery top, continuous monitoring of each individual battery is probably unnecessary. There is an additional hazard when using lead-acid batteries in any environment: an explosive hydrogen-air mixture can be available inside the cell. This is true even 377 with catalytic caps. If ignition energy is sufficient within the cells or even close to the vent cap, the internal mixture can explode, possibly spraying acid and blowing bits of the cover toward personnel in the vicinity. Internal faults can produce this kind of explosion. Virr and Pearson related that a common source of such events also occurs when electrolyte leaks from a cracked cell, producing a spark when the acid level falls below the bottom of the plates. There is no known method of preventing or suppressing these internally initiated cell explosions. However, ground-fault protection will detect electrolyte leakage. Low-resistance faults to the tray or vehicle frame caused by damage to cables, and so on, represent a different problem. As previously mentioned, no provision is currently available for deenergizing a faulted battery, which would continue to discharge until its stored energy was dissipated. Some form of circuit breaker between the cells of the battery would provide a way to sectionalize the battery in the event of a fault. In conditions of excessive current flow, the intercell connectors might themselves act as protective fuses, melting down when their current- carrying capacity was exceeded. Proper isolation of the battery electrical system from the mine ground system is a prerequisite for safe battery use. The reliability of battery isolation can be greatly enhanced by following the maintenance and design sug- gestions given in this chapter. Total electrical safety for any battery installation is, however, a function of charger characteristics as well as battery isolation. BATTERY-CHARGING HAZARDS A brief review of four representative accidents sheds some light on the hazards associated with battery charging: 2 1. A scoop operator was electrocuted when his body came in contact with the frame of a scoop tractor that was being charged. During the subsequent investigation, it was found that a potential difference of 260 V existed between the tractor frame and mine floor when the charger was energized. The cause was a low-resistance surface-leakage fault between the battery and battery tray, which caused the tray and the tractor frame on which it was resting to become energized. Faulty insulation between the primary and secondary windings on one arm of the three-phase transformer permitted secondary cur- rent to flow in the ground. The glaring error here was the failure of mine personnel to ground the steel frame of the scoop tractor properly while it was being charged. 2. A utility worker received a fatal electric shock when his body contacted the frame of a battery charger. A fault occurred within the charger that caused 210 Vdc and 114 Vac to exist between the charger frame and earth. Despite the fact that the primary cause of the electrocu- tion was a worn bushing that failed to insulate the timer circuit from the charger frame, a proper frame ground could again have prevented the fatality. 3. An electrician was fatally injured when he came in contact with a bare conductor on the charging leads of a battery charger while connecting the charger to the vehi- cle. The electrician was standing in water while attempt- ing to connect the battery. Although the charger switch 2 MSHA reports of fatal coal mine electrical accidents. was in the off position, a primary-to-secondary fault in the charger transformer circumvented the switch (which in- terrupted only one primary conductor) and caused the charger leads to become energized. 4. A surveyor was electrocuted when he contacted the battery ground clamp that was attached to the charger frame. The charger frame was energized because of a fault within the charger. The investigation showed the accident to be due to inadequate safety grounds on the frames of all the electrical equipment in the mine. These accidents occurred using chargers containing the basic circuitry discussed earlier in this chapter, and with the possible exception of the enclosure, they were similar to chargers used by other industries. This implies that additional design concepts or components must be included in battery chargers in order to ensure safety in the mine environment. Poor grounding and component failure (most espe- cially the power transformer) have been the most notori- ous contributors to mine charger accidents and electrocu- tions. They have caused the ac source power to be impressed on the dc charging circuit. Consider an instance where a primary-to-secondary power transformer fault elevates the charging circuit by the primary potential. If this occurs simultaneously with a fault that energizes a battery tray, and the grounding is unsatisfactory, the battery tray will be a shock hazard. Of vital importance is adequate grounding because an intact grounding system ensures that frame potentials do not exceed reasonably safe levels, regardless of the contribution of other factors to the hazardous condition. A list of other problems causing accidents has been made through an analysis of battery charger accidents and input from operators, manufacturers, and State and Fed- eral regulatory agencies: • Bad charging-cable insulation; • Unconnected, exposed charging couplers (plugs) that are still energized; • Poorly designed charging couplers that have ex- posed ungrounded metal parts where the cables are at- tached and allow cable damage; • Charging couplers with poor contact as a result of frequent damage or inadequate maintenance; • Battery surface leakage and internal faults; • No overcurrent protection for the ac power input; • Faults causing the control circuitry and compo- nents to have an elevated potential; • Personnel connecting a charger to a battery of wrong polarity; • Repair work performed by unauthorized personnel; and • Charging stations in abnormally wet locations. In view of this poor safety record, it is advisable to consider the safety features that would reduce the potential for accidents and electrocutions when charging batteries: 1. The charger input cable should contain a monitored ground to ensure that the grounding conductors to the charger frame are intact. 2. The charger should be equipped with panel inter- locks that deenergize the charger at the outby source when access panels are removed. 378 3. The charger should have an emergency off switch located in a conspicuous place on the charger frame. This switch should not be spring-loaded, thus requiring reset- ting after use. 4. The power transformer should electrically isolate the battery being charged from the power source, with primary and secondary windings being so arranged to eliminate hazardous interwinding faults. 5. A separate grounding conductor and ground-check monitor circuit should be provided for each battery tray serviced by the charger. 6. The dc couplers should be of the type that inter- rupts the ground-check circuit before the charging circuit. 7. The dc connection between the charger and battery box should consist of a single cable with appropriate grounding and ground-check conductors. 8. The charger should contain battery surface leakage detection circuitry that prevents a leaky battery from being charged. 9. The power transformer secondary and all dc circuits and components should be isolated from the frame ground of the charger. 10. The charger should have a meter or similar device that indicates the state of battery charge. 11. The charger should have overload and short- circuit protection on both the input and output. 12. The charger should contain circuitry that prevents a battery of the wrong voltage from being charged; alter- natively, the charging connector must be keyed or sized such that no connection can be made with a battery with fewer cells than that for which the charger is designed. Figure 15.14 is a diagram of a typical solid-state controlled taper-rate charger. All the desired electrical components specified in the above list are signified by the letters in parentheses. These and other features will be discussed in the following paragraphs. It should be noted that in some instances there are alternative practical methods that could be employed to protect personnel. System Grounding Because of its ac-to-dc application, the charger illus- trated is considered to be portable electrical equipment and is fed by a resistance-grounded low-voltage or medium-voltage ac system. Under Federal regulations for coal mining, a grounding conductor and ground-check monitoring of this grounding conductor are required be- tween the power source (usually a power center) and the charger frame. The portion of the ground-check circuit within the charger is indicated at a. The panel interlock switches (o) and the emergency off switch (c) are shown in series with the ground-check circuit. Here, a pilot-type monitor is implied. If the circuit is opened in any manner, the circuit breaker at the power source is tripped. After tripping, the circuit breaker must be reset manually at the source and cannot be reset until the ground-check circuit is restored. Electrocutions involving contact with the charger frame have taken place when grounding practices were unsatisfactory and a fault existed between the frame and some charger internal component. One method for pre- venting this type of hazard would be to use an insulation coating to isolate the enclosure completely from possible contact with live circuits. There are, however, several shortcomings to this method of protection. Most battery chargers designed for underground use are mounted on skids for easy mobility. In rough mine use, any insulation would be quickly worn off, especially on the skids. Insula- tion coatings would be difficult to apply everywhere. Incoming cable check Ground < Toff -on (i/4", d = 0.004" Wi Cover i Shall be greater than Ve' a + b shall not be less than 3/4" Minimum engagement 1/2" Figure 16.3.— Typical step-flange joint; enclosure internal volume larger than 124 in 3 . Cover Hinge pin Total developed length to conform to 30 CFR 18.31 (a)(b)for class Ifit Figure 16.4.— Threaded joint. Figure 16.5.— Tongue-and-groove joint. 386 Bolting holes must not penetrate to the interior of the enclosure; otherwise, the omission of a bolt would provide entrance to the chamber. Through-holes must be blind or bottomed as shown in figure 16.6. The maximum spacing between fasteners for joints all in one plane is 6 in. The maximum spacing between fasteners for joints, portions of which are on different planes, is based on the size and configuration of the enclosure, the strength of the materi- als, and explosion test results. The bolts or other means of clamping the joints should be proportioned to minimize stripping of threads and to give adequate strength from the stress developed during an internal explosion Mini- mum bolt diameters are also given in table 16.1. The minimum distance from the enclosure interior to the edge of the bolt hole is important as it must provide an adequate flame path length. The minimum Federal re- quirements (considering just flange width) are (37) Distance Joint size, in in mm V* Vs 3.2 y 4 3 /i6 4.8 1 or over 7 /i6 11.1 Figure 16.3 shows how these measurements are applied. The Maximum Experimental Safe Gap To help understand the joint criteria in conjunction with the theory of explosion-proof enclosures, a presenta- tion of past research is in order. The external ignition- suppression properties of flange joints have been an active research area for several years, with extensive investiga- tions carried out in the United States, the United King- dom, and the Federal Republic of Germany. Most research Fuse clamp Fuse Switch-box wal Fuse and switch Machine screws holding switch block to casing wall. Hole is "blind" Terminal for positive lead conductor Figure 16.6.— Blind screw hole. work has resulted in similar conclusions, and the general international consensus is that design standards for explosion-proof enclosures are imperative. The maximum experimental safe gap (MESG) is a standard often used to determine the explosion- transmission properties of a flame path. The MESG is measured by igniting a flammable gas mixture inside a test system and observing if it ignites a surrounding gas mixture outside the enclosure (14). The MESG is the largest gap for a flame path length that does not permit ignition outside. Relatively speaking, a higher MESG implies a safer situation. Note that a luminous flame is permitted to pass through the flange, provided that it has insufficient energy to ignite the surrounding atmosphere. This does not coincide with the Federal requirements given earlier, since U.S. regulations do not allow passage of any flame (37). A gap that will not allow ignition is 7 to 12 times larger than a gap that quenches visible flame (14). The following general statements can be made about past research findings (14, 25-26); a parenthetical com- ment is made where these conflict with present Federal regulations. 1. The MESG increases as the enclosure volume decreases because thermal losses predominate in small enclosures. Accordingly, smaller flange widths with the same gap can be allowed for smaller enclosures. 2. It is not likely that gaskets or O-rings used to make explosion-proof containers weathertight have any signifi- cant effect on enclosure safety, as long as they are external from the flange surface. It can be noted that O-rings are not permitted under 30 CFR 18 (37). 3. There is no evidence that metallic materials are essential for the flange construction. (Federal regulations generally require metal-to-metal joints except for head- light and meter enclosures. Metal-to-metal flanges usu- ally provide a better heat-sinking effect.) 4. Surface finish is not considered critical by most authorities. (Federal regulations require that flat surfaces be planed to within one-half the maximum clearance allowed, and finished to not more than 250 ^in.) 5. The apparent safe gap between bolted flanges, the normal commercial construction, is considerably smaller than would be predicted from immovable or nonbolted flange tests. 6. Evaluating the effectiveness of explosion-proof en- closures or studying flange-gap quenching is still highly empirical. Certain researchers have dealt with other specific subjects pertaining to the MESG. Torry (33), for example, established that no effect can be ascertained in the MESG over the humidity range of 0% to 50%, but the MESG increases as humidity increases from 50% to 100%. Con- sequently, the MESG can be measured with confidence when humidity is under 50%. By varying the methane-air mixtures inside and outside, in addition to in the flange gap, Titman and Torry (32) found the most ignitable outside methane-air mixture was 6.5%, with a 9.5% methane-air mixture inside the enclosure. James (10) varied the temperature in several plane-flange experi- ments. Using a 1.0-in (25.4-mm) joint, he found the gap that would just permit propagation to the outside was 0.045 in (1.14 mm) at 80°F (25°C), and 0.035 in (0.89 mm) at 500°F (260°C). He stated that with a safety factor it is unwise to exceed a gap of 0.020 in (0.51 mm) for a 1.0-in flange width. Furthermore, a gap of 0.02 in or larger 387 cannot be applied to plane flanges without a consideration of enclosure material strength, since one part of the flange is a plate that could spring or bend during an internal explosion, thus increasing the gap. James also showed that explosion-proof enclosure safety is influenced more by slight changes in gaps than by changes in the flame path. Several researchers have discovered that doubling the flange width, with other conditions remaining constant, increases the MESG by about 1.3 (14, 25, 32). This statement is limited to flange widths 1.0 in and less because few data are available for wider flanges, which are uncommon. As the flange width nears 1.0 in, the relative effect of widening decreases. Phillips (25-26) determined quantitatively that the MESG should be smaller if the initial enclosure temper- ature is raised, and furthermore, that turbulence outside the enclosure increases the MESG. He substantiated empiri- cally the enclosure-protection theory presented earlier, and most of his work has been further substantiated by the experimental results of other researchers. The standards established in 30 CFR 18 (37) and by Underwriters' Laboratories, Inc. (36) do not permit pas- sage of flame and specify almost identical flange-gap and width values. This means that the allowable gaps found from MESG research are 8 to 10 times larger than the maximum allowable gap in the United States. Some foreign countries specify the gap values as a fraction of the MESG, and the question has been asked as to why the United States does not allow such larger gaps. The U.S. requirement can be readily justified by the need for a safety factor that will accommodate the problems encoun- tered with deterioration resulting from exposure to the mine environment or neglect. In fact, Short (28), in discus- sion with various foreign testing authorities, has found that none of them will accept a product with gaps as large as those indicated by MESG research. In practice, many countries permit gaps two to four times larger than those specified in the United States. Pressure vent has large effective open area for flow of gases from ignition but cools exiting gases and arrests flame Pressure vent Electrical enclosure walls Lines representing constant pressure Point of ignition Figure 16.7.— Pressure vent limiting pressure buildup during internal explosion. Vent body Swinging-door- vent cover Magnet holds ■ cover closed Bolt with / lock washer Enclosure cover Flame path meets MSHA J requirements -Steel retaining ring Vent retainer Flame-arresting vent material Foam stainless steel Bolt with lock washer Tapped blind hole as required by MSHA Figure 16.8.— Pressure vent assembly using metal-foam material. Pressure Vents Some manufacturers add special pressure-venting de- vices that supplement the conventional release of pressure through flange and shaft gaps. The concept is illustrated in figure 16.7 (6*). Properly sized pressure vents can reduce the internal pressure during explosion to 12 to 20 psi (6). The vents exhibit a large effective area for gases from an ignition but provide cooling and inhibiting to arrest the flame. A prototype vent developed under Bureau of Mines contract is shown in figure 16.8 (6). Here the escaping gas is filtered through flame-arresting material. Clamp fastened with bolts Locking screw Cable rO Radial clearance 0.003" maximum Internally threaded stuffing box Vs" minimum (1/4" maximum) with cable properly packed F~* Hose clamp ^O \ Hose conduit Figure 16.9.— Typical slip-fit straight stuffing box and packing-gland lead entrance. Cable Entrances The explosion-proof enclosure must also, obviously, have openings through which electrical connections can be made. These openings are particularly important because of the frequency with which trailing cables must be replaced on permissible mobile equipment. The most pop- ular type of cable entry incorporates a packing gland or stuffing box, which may have straight-through entry or angled entry. Figures 16.9 through 16.11 show cable entries with slip-fit stuffing boxes, where the box is a separate component with a cylindrical projection that fits into the enclosure wall (37). Specifications for enclosures exceeding 124 in 3 include a flame path length of 1.0 in and a radial clearance between the wall and box not to exceed Hose conduit Hose clamp Cable -*" Vs" minimum with cable properly packed Clamp fastened with bolts Radial clearance 0.003" maximum Externally threaded gland nut Figure 16.10.— Typical slip-fit angle stuffing box and packing-gland lead entrance with hose clamp. 388 Externally threaded gland nut Internally threaded Ve" minimum -^ ".' „ ' " ' ~}~7}„ stu ^ ' ng box - ' Clamp clearance ^iTr II I fastened with bolts) Radial clearance 0.003" maximum Figure 16.11.— Typical slip-fit angle stuffing box and packing-gland lead entrance. Plug shall be secured by spot welding or brazing. Weld may be on plug, clamp, or fastening bolt. 0.003 in. Figure 16.12 shows how a spare cable entrance hole must be plugged, and figure 16.13 illustrates a packing gland that is an integral part of an enclosure (37). A stuffing box is usually constructed so that when the packing-gland nut is threaded into the opening and tight- ened, it forces the stuffing material against the cable, making a very tight joint. The packing material is usually untreated asbestos or an MSHA-accepted asbestos substi- tute material. When compressed, there must be at least V£ in of packing along the length of the cable, and the clearance between the packing-gland nut and the stuffing box must be no less than Vs in (38). Repacking a stuffing box in the mine can be extremely awkward and time consuming. A recently introduced method that uses a tapered polyurethane flame-resistant grommet in place of the asbestos (fig. 16.14), greatly simplifies the assembly (6). The elastomeric grommet has high compressibility so that one size can accommodate a limited range of cable sizes. The requirements for leads that pass between explosion-proof compartments separated by a common wall are not as rigid as those for leads passing through an exterior wall. One type, an insulated stud entrance, is shown in figure 16.15. Here the conductor does not actu- ally pass through the wall but is connected to a finished brass casting stud that is isolated from the enclosure wall by an insulated tube and washers. Windows and Lenses Federal regulations state that MSHA may waive window and lens material testing except for headlight lenses (37). All window and lens material must be sealed in place or provided with correct flange joints and must be protected from mechanical damage, either by guarding or inherently through location and structural design. If the exposed material area exceeds 8.0 in 2 (51.6 cm 2 ), the window or lens (excluding headlight lenses) must be protected with guarding or the equivalent. Both thermal- shock and impact performance tests are outlined in the Federal regulations. Enclosure Mechanical Strength and Internal Pressures /■ Plug / s- Clamp (fastened with bolts) Radial clearance 0.003" maximum Figure 16.12.— Typical plug for spare lead-entrance hole. Stuffing box (integral) part-, of enclosure y Packing material I/2" minimum i/e" minimum, 1/4" maximum clearance Packing -gland nut / /gv* — Hose clamp Hose conduit Cable Metal tubing Locking screw Figure 16.13.— Typical threaded straight stuffing box and packing-gland lead entrance with provision for hose conduit. Retaining clip Cable entry body Tapered urethane grommet Figure 16.14.— Prototype trailing cable entry with polyurethane grommet. The maximum internal pressure developed during the ignition of an explosive air mixture in an enclosure is directly related to the amount of venting through flange gaps and any auxiliary pressure-relief devices. Tests of mechanical strength and internal pressures provide im- portant parameters used to design enclosures that will not transmit the explosion to a surrounding atmosphere. Mea- surements of explosion pressure are made to prove that the container strength is adequate. It is particularly impor- tant to ensure against flange-gap distortion. Insulated washers Insulated tube- Figure 16.15.— Insulated-stud lead entrance. 389 30 CFR 18.31(a) specifies that cast or welded enclo- sures must be designed to withstand an internal pressure of 150 psig (1.04 MPa) and that casting must be free from blowholes. Welded joints forming an enclosure must be continuous, gas-tight welds in accordance with American Welding Society standards. Minimum allowable thick- nesses for enclosure walls, flanges, and covers are also outlined. Again, a review of relevant research findings assists in understanding these regulations. Nagy (21) conducted extensive tests on methane-air and coal-dust explosions to determine pressures with vary- ing enclosure geometries. A range of 38 tests was per- formed on enclosures with internal volumes from 0.043 ft 3 (101.5 cm 3 ) to 905 ft 3 (25.63 m 3 ), and methane-air mix- tures from 6.0% to 14.1%. He found that with an ambient internal pressure of about 14.2 psia (98 kPa), the maxi- mum pressure developed for all vessels was 119.6 psia (0.825 MPa). When the ambient internal pressure was at 17.6 psia (121 kPa), the maximum was then 143.3 psia (0.988 MPa). Both maximums were obtained from exami- nation of a 1.0-ft 3 (2,359-cm 3 ) vessel with 9.4% methane- air mixtures. From 39 other experiments, using the same enclosure but a Pittsburgh coal-seam dust, a maximum of 119.3 psia (0.823 MPa) occurred with a concentration of 0.6 oz/ft 3 (601 g/m 3 ) in an 0.32-ft 3 (7.55-cm 3 ) vessel. Nagy concluded that • Vessel size does not affect the maximum pressure level if heat loss is neglected; • The rate of pressure rise decreases as the vessel size is increased; • The vessel shape does not affect the maximum pressure if heat loss is neglected; • A change in initial pressure produces a propor- tional change in the maximum pressure and the rate of pressure rise; • The changes in the initial temperature produce an inverse change in the maximum pressure and have little effect on maximum rate of pressure rise. Considering that the maximum pressure developed in an explosion is a function of venting, the closed-system enclo- sure tests carried out be Nagy could be taken as the worst case. The Federal regulations require enclosures to with- stand 164.7 psia (1.136 MPa) assuming standard atmos- pheric pressure, or two times the maximum pressure recorded during any tests without any pressure limit (37). The maximum permissible flange gap is rather small and therefore restricts venting. Consequently, the Federal specification does provide some margin of safety. Magison (14) suggests that one of the most significant general facts about the performance of explosion-proof containers is that higher maximum pressures and rates of pressure rise are associated with small flange gaps (for the same flange width with identical explosive gas mixtures). Referring to the work performed by the Safety in Mines Research Establishment (34), he states that the venting effect of the MESG in an 8-L sphere (488 in 3 ) reduces maximum pressures to a few pounds per square inch (ignition of 9.8% methane-air mixture). Magison indicates that, when considering heat loss, the actual maximum pressure achieved in an enclosure will depend upon the location of the ignition source, the container size and geometry, plus the arrangement of inside equipment. Theoretically, enclosure size has no explicit effect on the maximum pressure, but research has shown that size does have an effect for lean and rich mixtures (40). Here, the explanation given is that a longer reaction time occurs in larger vessels and allows for more heat loss. MSHA specifies twofold pressure tests to ascertain conformance with strength requirements. The first test is mainly to determine the explosion-proof characteristics, specifying that an explosive mixture within the enclosure will be ignited electrically and the subsequent explosive pressure will be recorded. The test must be repeated a minimum of 16 times to accommodate various ignition points, different atmospheric composition, and the operat- ing status of the motors. No part of the enclosure must rupture as a result of these tests, nor have permanent distortion exceeding 0.04 in per linear foot. If 125 psig (0.862 MPa) is exceeded during any test, the enclosure can be rejected unless it is reconstructed to reduce the pressure or, alternatively, is able to withstand a dynamic pressure twice the highest value recorded in the initial test. A second or static-pressure test is applied only when MSHA's Approval and Certification Center decides that visual inspection is inadequate to reveal casting or welding defects. The pressure applied in this test is 150 psig or 1.5 times the maximum value recorded during the explosion tests, whichever is greater. The external surface temperature of all permissible mechanical or electrical components for use on electrical face equipment must not exceed 150°C under normal operating conditions in U.S. mines (37). Underwriters' Laboratories, Inc., has a similar specification but it ap- plies only to dust-proof enclosures, which will be discussed later. The regulation is concerned with the possibility of causing an ignition through surface contact between the enclosure and any of the four potentially hazardous mix- tures: methane-air, coal-dust-air, combined methane and coal-dust mixtures in air, and collected coal dust on the enclosure surface. The lowest autoignition temperature must obviously form the basis of the regulation. Of the methane-air mixtures, a 2.9%-methane mix- ture had the lowest autoignition temperature of about 550°C, which can be eliminated immediately (41). Mea- surements to establish the minimum ignition tempera- tures of coal-dust clouds and dust layers have been per- formed by Nagy (20) on several U.S. coals. The dust cloud experiments on 22 different coals gave an average mini- mum ignition temperature of 617 °C. The lowest ignition temperature was 440°C for a Colorado high-volatile coal. Tests performed on layered dust of 16 U.S. coal samples showed an average minimum ignition at 222 °C, with the lowest temperature being 160°C for a high-volatile Illinois coal. Ten of the layer tests resulted in minimum ignition temperatures of 190°C or below. Hence, layered coal-dust ignition temperature is the limiting factor for the maxi- mum allowable surface temperature of explosion-proof enclosures and justifies the 150°C requirement of the Federal regulation. The requirement is particularly signif- icant for equipment in underground mines and prepara- tion plants where dust accumulates readily. Enclosure Hazards In recent years, some mining experts have questioned the inherent safety of explosion-proof enclosures. The concern is that when an explosion is triggered because of an arc or short circuit within the enclosure, gases can be generated that would defeat the containment properties of the enclosure. A number of failure modes are known to be possible when a high-voltage short circuit occurs within 390 the enclosure: these include burning through the enclo- sure wall, ignition transmission by hot gases or flame, particle ignition transmission, and enclosure deformation or bursting. The possibility exists that once an arc is established within an enclosure, it may jump across to the enclosure wall, and the intense heat and power might then cause the arc to burn through the wall. Killing and Tielke (12) have demonstrated this possibility quite convincingly. If organic insulation is used within the enclosure, combustible hydrocarbons may be present under fault conditions that could lead to defeat of an ignition- containment requirement that has been designed only for methane. Arc formation between two metal conductors (usually copper) within an enclosure can also cause ignition of the surrounding atmosphere by hot particles expelled through the flange gap. Heat from the arc causes metal particles to be separated from the conductors and subsequently forced through the gap by pressure buildup. These hot metal particles are capable of igniting a combustible gas-air mixture. Hence, the MESG must be small enough to prevent passage of such particles, further justifying the small gaps required by U.S. regulations. However, if enclosure gaps are made small enough to ensure that ignition of exterior gases is not caused by expulsion of hot gases, flame, or hot particles, the possi- bility then exists for high pressures to build up within the enclosure under fault conditions. Covers could then be blown off, with subsequent ignition of the surrounding atmosphere. Davidson and Lord (5) cited half a dozen such incidents in the Federal Republic of Germany, one in Canada, and two in the United Kingdom. The flameproof enclosures that burst in Canada and the United Kingdom had gap specifications similar to those required in the United States. In a typical instance, a steel cover was thrown several meters after shearing apart fourteen V^-in-diameter securing bolts. In all cases there was severe electrical fault damage within the enclosure, including charring of organic electrical insulating material. Ciok (4) reported that high-voltage short circuits within flameproof enclosures have become a matter of concern in Polish coal mines. Such short circuits have occurred several times, and the covers closing the equipment chamber have been blown off. These high pressures within the enclosures appear to be partly due to the evolution of gases from organic insulating materials incorporated in the enclosure con- struction. Simon (29) described experiments where organic insulating materials within enclosures were subjected to heating by an electric fault arc. He showed that in these tests, gases evolved that were capable of causing sufficient pressure to rupture the enclosure. The evolved gases (volatization products, consisting primarily of hydrogen, carbon monoxide, nitrogen, and methane) ignited sponta- neously upon contact with the external atmosphere. If the flange gaps or pressure vents would allow all the volati- zation products to escape, no enclosure failure would occur, although the nature of the escaping hot gases must be considered. Another possible contributing factor is an effect inves- tigated by Brown (3) and Bossert (2). It is known that hydrocarbon-air flames produce free ions that cause an electrical current to flow when the flame front bridges two points of opposite electrical polarity. This current may be sufficient to form an arc discharge; that is, an initial flame might produce ions that contribute to additional faults within the enclosure. The situation might be aggravated by the presence of organic insulating materials: the initial fault may produce hydrocarbon gases through thermal insulation breakdown, and the burning of these gases might induce additional arcing. This possibility would increase as voltage levels were raised, triggering a chain- reaction effect that could culminate in deformation or bursting of the enclosure. Materials that apparently contribute to overpressures are those that give off gases when subjected to heating. The greatest hazard is from materials that evolve combus- tible gases, here, insulating materials and accumulated water. When water is subjected to an electrical current, electrolysis takes place and hydrogen and oxygen are evolved. This constitutes a very undesirable situation; hence, water accumulation within enclosures should be avoided as much as possible, for not only does it have the potential of creating faults, but when a fault occurs it may contribute significantly to increased enclosure pressures. When organic insulation materials are subjected to electrical arcing, tracking or heating, decomposition takes place that yields various gases, both combustible and noncombustible. Over the past 30 yr, electrical equipment manufacturers have shown a tendency to replace tradi- tional insulants (cellulose, natural fabrics, asbestos, mica, porcelain, glass, etc.) with newly developed organic poly- mers. Yet at the same time, traditional insulants are still in common use. As a result it is possible to find almost any known insulation within an explosion-proof enclosure. The numbers are so great that it is extremely difficult to categorize the many types and variations of insulating and plastic materials that are likely to be found. However, a short summary of insulants that are known to evolve dangerous gaseous products follows (19). Cross-Linked Synthetic Polymers. This group contains the basic synthetic resins commonly utilized in the man- ufacture of plastic materials. Of these, the phenolic, epoxy, and silicone resins appear to offer the greatest hazard potential because of their high yield of combustible vola- tile products. The amino (melamine) polyester, polyure- thane, and isocyanate resins appear to have a lower yield of combustible volatile products. Materials made with these resins should therefore exhibit a lower, though still quite significant, hazard potential. Linear Synthetic Polymers and Elastomers. This group contains many of the well-known types of insulators. The worst potential hazards (similar in magnitude to those for the cross-linked group) appear to be associated with poly- propylene, polymethylene, polystyrene, polyethylene, neo- prene, and polymethylmethacrylate. Other materials such as nitrile butadiene (NBR) synthetic rubber, natural rubber, styrene butadiene (SBR) (GR S) synthetic rubber, Dacron polyester fiber, Teflon fluorocarbon polymer, poly- vinylchloride, poly ure thane, polyvinyl formal, and nylon show less though still very significant hazard potential. Other Possibly Dangerous Materials. Insulating oils should also be considered high hazard materials. Even though the gases formed are initially dissolved in the oil, upon saturation they may eventually evolve, causing an unsafe situation. A similar statement can be made about oil-impregnated paper. Other materials, such as cellulose and cotton, which do not appear dangerous at first glance, could also contribute to a hazardous condition because of the amounts of water and carbon monoxide evolved. Safe Materials. The electrical insulating materials that should be used whenever possible in explosion-proof enclosures are the electrical porcelains, ceramics, glasses, 391 asbestos, and mica. These materials are the most resistant to the production of gaseous products when exposed to heating or arcing. Note, however, that these recommenda- tions concerning materials consider only the possible contribution of evolved gases to increased enclosure pres- sures; toxicity is not considered. Although it does not appear feasible to eliminate all potentially hazardous materials from enclosures, the use of any materials that can evolve gaseous products under fault conditions should be avoided wherever possible. Of course, if a sustained arc occurs, even the conductors are vaporized, so all such statements are relative. Perhaps the best recommendation should be that the materials to be avoided are those that are readily volatilized, particularly those that evolve large amounts of combustible gases. PERMISSIBLE EQUIPMENT As defined at the beginning of the chapter, the term permissible equipment is applied to completely assembled electrical machines or components that have received official approval from MSHA. The term completely assem- bled means all equipment portions from the protection at the power source to all internal and external components of the machine, including the trailing cable. Permissibility requirements have been mentioned at various places in other chapters; details of grounding requirements were given in chapter 7, trailing cables and components in chapter 8, protective devices in chapters 9 and 10, battery equipment in chapter 15, and explosion-proof enclosures in this chapter. The aim here is to demonstrate how this information is tied together, by giving an overview of the procedures used by MSHA to investigate prospective per- missible equipment for safety. The overview is followed by a discussion of procedures recommended for checking equipment after it has been placed in service, and for maintaining explosion-proof enclosures and permissible equipment in proper condition. Permissible Equipment Schedule This is based on information in a 1954 Bureau of Mines Information Circular (9), as updated by MSHA's Approval and Certification Center. As already stated, the published regulations are contained in 30 CFR 18. Other pertinent regulations are found in 30 CFR 19 through 29 for electrical equipment and 30 through 36 for mechanical equipment. Each of these parts is often termed a schedule. Schedules are revised from time to time to conform to equipment development and to permit as much freedom as possible without lowering standards. Thus, some of the details in the following information may become outdated, but the general nature of the requirements will not change. Investigations are carried out by MSHA to determine the permissibility of such equipment as continuous min- ers, shuttle cars, battery-powered vehicles, pumps, distri- bution boxes and so on, and for certification of components such as explosion-proof enclosures, connectors, and battery assemblies. The investigations are divided into four major consecutive parts: review of drawings to verify that the design meets the requirements, detailed inspection of the equipment, tests of the equipment or internal components in explosive gas-air mixtures and/or adequacy tests where appropriate, then a final inspection of the tested accesso- ries in the completely assembled unit or machine for which approval is requested. Although investigations for the certification of components may follow the same pro- cess, only the first three steps are usually necessary. To initiate the procedure, a written application must be made to MSHA, accompanied by a set of detailed drawings, wiring diagrams, specifications, descriptions, and any related material. Any intrinsically safe compo- nents must be stated. When approval is being considered, only those compo- nents that have a bearing on permissibility are studied; only one motor, controller, protective device, or unit of a given design is required. The investigation starts with a check of the drawings and specifications in order to determine compliance with the applicable regulations, and then a detailed check of all parts against the draw- ings, to see that they coincide. Exact measurements are performed on the dimensions of joints, bearings, pressure vents, and other possible flame-arresting paths in enclo- sures. For explosion-proof enclosures, for example, the examination determines any unnecessary through-holes, the adequacy of design and construction of cable and lead entrances, the adequacy of electrical insulation and clear- ances between live parts and between live parts and the enclosure, any weaknesses in welds or flaws in casting, any distortion of enclosures before tests, and the adequacy of the fastenings, including their size, spacing, security, and possible bottoming. The quality of design, material, and workmanship receives careful scrutiny. Only equip- ment adhering to the following statements will be ac- cepted for further investigation (37): 1. "Electrically operated equipment intended for use in gassy mines shall be rugged in construction and shall be designed to facilitate inspection and maintenance." 2. "Only electrical equipment that is constructed of suitable materials, is of good quality workmanship, based on sound engineering principles, and is safe for its in- tended use" will be tested by MSHA. The testing phase emphasizes explosion-proof charac- teristics and component properties. The general nature of the explosion testing has been covered earlier. No less than 16 internal explosions are made at various ignition points, using a methane-air mixture for all, with bitumi- nous coal dust added for some tests. The actual internal electrical equipment, or dummies of equal dimensions, must be in place for a prescribed number of tests. Motor rotors are tested in both stationary and rotating modes. An enclosure can be rejected with the occurrence of any of the following conditions: • Discharge of flame from any joint or opening, • Ignition of an explosive mixture surrounding the enclosure, • Development of afterburning (gas drawn into the enclosure by the vacuum created by the explosion, then ignited within the enclosure), • Rupture of any part of the enclosure or any panel or divider within the enclosure, • Permanent distortion of the enclosure exceeding 0.04 in per linear foot. Other tests and examinations are made to determine the adequacy of components for the intended use. Some of these are performed at the discretion of MSHA investigators. 392 • Where the durability of a component is in doubt, mechanical tests will be performed to ascertain whether any points need to be strengthened. • Battery boxes are examined for ventilation, electri- cal clearances, insulation, drainage, and suitability for specific service. • Switches, circuit breakers, or contractors intended to function as switches are checked to see if they are capable of interrupting the maximum current permitted by the circuit's automatic protection device. • Cables, conveyor belting, and hoses are tested for flame resistance. At the end of the investigation, a final inspection is made of completely assembled new machines or of ma- chines that were previously approved but have since un- dergone substantial modification. The aim of the final inspection is to uncover any unsafe features, and the inspection includes such items as • Compliance with joint, lead-entrance, or other per- tinent requirements; • A check of wiring between components and the adequacy of cable clamping and mechanical protection for cables; survey of the positioning of cables, particularly those in proximity to hydraulic components; • Determining the adequacy of protection against damage to headlights, push buttons, and other vulnerable locations; • A check of the settings of overload and short-circuit protection; • Ensuring that there is a suitable means of connect- ing and protecting the trailing cable. Finally, MSHA has the option to have a staff engineer check the first machine produced, preferably at the factory where it is built. When approval is granted, MSHA issues a formal notice of approval, which is sent to the manufacturer. The notice is accompanied by a photograph of an approval- plate design. Plates reproducing the design and required information are mounted in a conspicuous place on the machine, serving to identify the machine or accessory as having met the applicable requirements of 30 CFR 18. Maintenance of Permissible Equipment For equipment to retain permissibility, it must be maintained in the same condition as that approved by MSHA. This is the advantage of familiarity with the schedule requirements: the essential details may be ex- tracted to establish a routine maintenance and inspection program that will be in full legal compliance. Safety can be compromised easily during mainte- nance procedures. One of the most common problems with explosion-proof enclosures is that small unapproved open- ings, which can constitute an immediate danger, can appear because of incorrect maintenance. Such openings are rarely caused by a manufacturing error and usually result from negligence or a lack of understanding on the part of the mine maintenance crew. One of the most common of such violations is the citation "open box— 0.005," which means that a plane flange has been found to exceed 0.004 in (16). Other typical problems that cause dangerous enclosure openings are bottomed bolt and screw holes that have been drilled through, holes drilled through when plates or components have been attached, loose or improperly assembled cable packing glands, or undue wear on bearings where shafts enter an explosion-proof enclosure (11). The following precautions and procedures are recom- mended in the maintenance of permissible equipment (11, 14, 16, 38), but reference should be made to the Federal regulations for precise compliance. Items 1 through 11 concern explosion-proof enclosures specifically, while items 12 through 19 relate to permissible equipment covered in more detail in other chapters. Items 20 through 23 list specific schedule requirements. 1. Before any apparatus is examined, the power source should be deenergized, locked out, and tagged. This proce- dure is particularly important before an explosion-proof enclosure is opened, although it should be obvious practice for any electrical equipment, permissible or not. 2. All joint clearances should be examined regularly with feeler gauges to determine that the maximum allow- able clearances given in table 16.1 are not exceeded. 3. All pressure vents should be examined for cleanli- ness. 4. Any missing bolts, lock washers, or ineffective fasteners must be replaced. All threaded inspection covers must be secured. 5. All cover flange and thread surfaces must be treated with great respect. They must not be handled roughly, and anything, including tools, that might possi- bly scratch or mar a joint or thread surface must not be allowed to come into contact with it. If a cover thread is damaged, it must be replaced. If a joint is scarred by any means, the equipment must be removed from service. 6. If an enclosure is opened, the cleanliness of joint and thread surfaces must be maintained so that foreign matter does not enlarge a gap. Hence, all joints and threads should be carefully cleaned before reassembly, and if necessary, a thin layer of lubricant should be applied. Care must be taken that the joint does not become recon- taminated. When reassembled, all joint clearances should be rechecked. 7. Frequent examination should be made to determine any corrosion on flanges, threads, shafts, bearings, and any other flame-arresting path. Corrosion inhibitors and lubricants may be used, but if any corrosion is found, the enclosure should be taken out of service and sent for repairs since corrosion products cannot be removed ade- quately from equipment in operation. 8. Enclosures should be examined regularly for burned holes. 9. All cable packing glands should be examined to see that the cable fits tightly, and the clearance between the gland nut and housing should be checked to see that it is adequate. 10. Headlights should be checked for loose or broken lenses, loose packing glands, missing or broken parts, and improper assembly. 11. As much as practical of the accumulated coal dust should be removed. 12. Portable or mobile equipment must be properly frame-grounded or provided with equivalent protection. Ground and ground-check conductors must be attached to separate studs that are not attached to a removable panel. If a separate grounding conductor is used on dc machines supplied by trolley power, the return (which is usually negative) and the grounding conductors must be attached to the rail or other grounded conductor with separate clamps. 393 13. All electric components must be solidly attached to the machine frame. Light fixtures must be grounded with a separate grounding conductor. 14. No splices are allowed in the external wiring on permissible equipment, except for intrinsically safe circuits. 15. All conduit hose must be flame resistant and MSHA accepted. The condition of mechanical cable pro- tection such as guards, conduit hose, and clamps should be examined regularly. Conduit hose should not be spliced. Worn or cut conduit hose may be repaired using MSHA approved flame-resistant cable-jacket repair material. 16. Trailing cables should adhere to Federal regula- tions in type, size, length and condition (see chapter 8). A temporary splice must not be made within 25 ft of the permissible machine, except in reeled applications. Short- circuit protection must be provided for all ungrounded power conductors, either by correctly adjusted circuit breakers or properly sized dual-element fuses (see chapters 9 and 10). Overload protection is recommended. 17. All trailing cables must be provided with effective strain relief at the entrance to equipment. 18. Machine cable reels and spooling devices must be insulated with flame-resistant material. Rollers, sheaves, and reel flanges must be maintained so as not to damage cables. Reels should maintain a positive tension on the trailing cable during reeling and unreeling. Reel collector rings should be examined for any deterioration that would cause a high-resistance contact. 19. All circuit breakers and overload protection on the machine must be maintained in working order. A main circuit breaker, contactor, or disconnect switch must be provided on the machine and be capable of deenergizing all power conductors on board the machine except the methane-monitor power supply. Headlights and flood- lights must have a separate two-pole switch to deenergize the power conductors. 20. All wheel-mounted equipment must be provided with brakes, unless the design of the driving mechanism prevents accidental movement when parked. 21. If a mobile transportation unit travels faster than 2.5 mph, headlights and red reflecting material are man- datory on both the front and rear of the vehicle. Such vehicles must have an audible warning device. 22. Guards and safe-off devices for push buttons must be maintained in working condition. 23. The approval plate must be attached to the equip- ment. Any unauthorized changes to the equipment not con- tained in the approval will render the equipment nonper- missible. Furthermore, no machine that has been changed from that approved may be placed in service until a field modification to the approval has been reviewed and ap- proved by MSHA. COAL DUST HAZARDS Coal dust can pose an ignition hazard of two types: a dust layer ignition, or dust cloud ignition. A dust cloud concentration above the lower explosive limit can be ignited by an arc, thus presenting a severe dust explosion hazard. The dust cloud is nonhomogeneous, and ignition is dependent upon the volatile content (the chemical compo- sition), the moisture content, and the ash content, as well as the size and shape of the cloud (14). When the dust is composed of material containing less than 8% volatile matter, as some anthracites and coke, there is almost no explosion hazard, but a fire hazard may still exist. Bitu- minous coals and lignite, which commonly have volatile contents ranging from 30% to 40%, present a serious dust explosion hazard. Bureau of Mines research (20) has determined the parameters required for the ignition of a bituminous coal dust that is minus 200 mesh dry (less than 5% moisture): Minimum energy mJ.. 5 Minimum autoignition temperature ... °C. 617 Lower explosive concentration g/L.. 0.05 Upper limit range g/L.. 2-5 It can be seen that the energy required for ignition is many times greater than that required for the ignition of methane-air mixtures. However, this energy level exists at a concentration of about 0.2 g/L and above. Concentrations can be measured quite readily, but the rule-of-thumb measure widely accepted among miners is that a dust cloud does not exceed an explosive concentration if a person can see his/her outstretched hand in front of his/her face (20). Layered dust that accumulates on equipment can also be a safety hazard. The parameter of most concern is the autoignition temperature, which depends mainly on the layer thickness and also the particle size related to the surface temperature of the equipment. Table 16.2 gives the minimum autoignition temperature for various thick- nesses of layered coal dusts (20, 24). The maximum surface temperature of 150° C specified in the Federal regulations for permissible equipment was based on this problem: A 150° C surface temperature would allow dust thicknesses up to 100 mm (about V& in) without incurring autoignition, but at that thickness there would be almost no safety factor. If a 1.25 safety factor is applied to a 150° C equipment surface temperature, table 16.2 indicates that no more than 20 mm of dust should be allowed to accumu- late. With a 1.50 safety factor there should be no more than 5 mm of dust. By comparison, the Federal Republic of Germany requires that the maximum surface temperature be 75° C below the autoignition temperature for a 5-mm dust layer (24). The moisture level is relatively unimpor- tant since moisture will quickly evaporate from any sur- face exceeding 100° C, whether or not it is covered with dust. Table 16.2.— Minimum autoignition temperatures (in degrees Celsius) versus layer thickness for bituminous coals. Thickness, mm Low-volatile coal High-volatile coal 5 300 240 20 250 190 50 ND 175 100... ND 160 ND No data. Classification of Dust Locations All areas where coal dust is a hazard are defined as class II locations. The decision flowchart in figure 16.16 can be used to assign the class II locations as either division 1 or division 2. It is based on information from the 394 None Does not exceed minimum depth Not a hazardous location Low Frequent or periodic Causes combustible dust cloud and electrical failure or causes electrical failure when cloud present Exceeds minimum depth Division 2 Division 1 Figure 16.16.— Decision flow chart of class II, division 1 and 2 hazardous locations. NEC (37) and the literature (17, 23). Division 1 locations occur with • A low-resistivity dust, • A high-resistivity dust in a frequent or periodic combustible cloud, or • A plant malfunction that allows formation of a combustible cloud and electrical equipment to spark or overheat. Division 2 locations are defined when • The thickness of layered dust exceeds the minimum depth to propagate flame, or • A high-resistivity dust infrequently forms a com- bustible cloud. An additional division 2 location should be noted: when the layered dust interferes with heat dissipation— but Magison (14) suggests that this condition is implicit in item 1 above for any combustible dust. Low resistivity has been defined as values less than 100 fi-cm; a conductive dust is one that breaks down under an electric field of 1,000 V/cm or less. A direct hazard exists with these dusts because they can provide conduc- tive paths between live parts and are susceptible to arc formation (14). Moodie (18), however, has found that coal dust has high resistance and is nonconductive, even when contaminated with mine drainage water. Hence, hazard- ous locations appear to exist when a cloud concentration exceeds 0.05 g/L or when the dust layered on electrical equipment exceeds a thickness of about 20 mm. These values have been extracted from the current literature and should be taken as a guide rather than a specific rule. Reducing Dust Hazards In class II hazardous locations, the types of enclosures in common use are dust-ignition-proof and dust-tight. Pressurized and intrinsically safe systems are used less frequently (14). The objective of these enclosures is to keep dust away from ignition sources and to prevent ignition of layered dust. Dust-ignition-proof enclosures meeting Un- derwriters' Laboratories, Inc., requirements conform to both objectives (36). The requirements are similar to those for explosion-proof containers but less severe. Joints, for example, must not be less than 3 /ie in wide, with maximum gaps of 0.0015 in. With wider joints, the maximum clear- ances are raised proportionately but must not exceed 0.008 in. Gasketing is allowed as an alternative between mated surfaces, but the surface and gasket width cannot be less than 3 /e in. Such gaskets must not deteriorate under normal use and cannot be glued to the surface. The goal is to provide an enclosure that will prevent hot particles from escaping and dust from entering (14). The maximum allowable temperature for the surface of such equipment is 150°C for equipment that can experience overload and 200°C for equipment that will not usually overload. Dust- ignition-proof motors are suitable for both division 1 and division 2 locations. Dust-tight enclosures are intended only for division 2 locations. The standards for their dust-tight construction are less restrictive than for equipment in division 1 locations, but they undergo the same tests by the Under- writers' Laboratories, Inc. (35). Hazardous Locations in Preparation Plants Most of this chapter has been projected at hazard reduction in underground coal mines. The reason should be obvious: all portions of underground coal mines inby the last open crosscut and in return airways are considered hazardous locations. Hazardous locations in surface mines and surface facilities of mines are not so easily defined. The following information on coal preparation plants is intended to provide an example of areas that can be hazardous. Coal preparation plants commonly experience three conditions that are classified as hazardous: • High levels of coal dust suspended in the plant atmosphere, • Significant accumulation of coal dust settled on electrical equipment and other surfaces, • Dangerous accumulations of readily ignitable gases, such as methane released from coal being processed. The extent of these hazards depends on the characteristics of the coal being handled, the preparation plant design, and the steps taken to modify or control the hazardous condition. The unit operations and plant locations susceptible to class II hazards include (1) • Transfer points in the materials-handling system, such as conveyor-to-conveyor and bin-to-feeder locations; • Coal-crushing and rotary-breaking systems, includ- ing operations that create new particles due to coal friability; • Coal wetting, sizing, and sorting operations; 395 • Plant cleanup systems; • Manual picking tables; • Thermal dewatering systems. Coal dust accumulates in these locations, particularly inside dust covers and covered conveyors and chutes where effective cleanup is difficult. Class I locations where methane can accumulate are mostly operations involved in particle size reduction and locations that have both limited ventilation and stored or slow-moving coal, such as silos, storage bins, pressure discharge chutes, and the tunnels and chutes leading to these locations (1, 15). The purpose of the chapter has been to introduce hazard-reduction techniques that are used in and about coal mines. Portions of coal-mine operations can be ren- dered dangerous without diligent adherence to the proce- dures and regulations presented in the foregoing para- graphs. Anyone desirous of gaining more knowledge in this subject should read reference 14. REFERENCES 1. Atallah, S., and P. Valence. Analysis of Coal Preparation Plants for Applicability of the National Electrical Code (contract J0166059, Arthur D. Little, Inc.). BuMines OFR 60-78, 1977; NTIS PB 283 411. 2. Bossert, J. A. Electric Aging During Flammable Gas Explo- sions. Can. Explos. Atmos. Lab. Rep. 342, 1974. 3. 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F. Preliminary Investigation Into the Causes of the Bursting of Group I Flameproof Enclosures, in Which the Elec- tric Current Caused Gases To Evolve From Plastic Material. SMRE, Sheffield, England, Transl. 6012; 14th Int. Conf. Mine Safety Research Establishments, Donetsk, U.S.S.R., 1971. 30. Stefanko, R., and L. A. Morley. Mine Electrical Systems Evaluation (grant G0133077, PA State Univ.). Explosion-Proofing of Mine Containers. BuMines OFR 76(2)-75, 1974; NTIS PB 245 928. 31. Thomas, V. M. Design of Intrinsically Safe Apparatus for Use in Coal Mines: A Review of Data and Techniques. Min. Electr. and Mech. Eng., v. 44, May and June 1964. 32. Titman, J., and R. Tony. Flameproof Enclosures for Mining Electrical Equipment; The Protection Afforded by Flanges of One- Half-Inch Radial Breadth for Mixtures of Methane and Air. SMRE, Sheffield, England, Res. Rep. 123, 1955. 33. Torry, R. Flameproof Enclosures for Mining Electrical Equipment -Influence of Atmospheric Moisture on Maximum Safe Gaps in Mixtures of Methane and Air. SMRE, Sheffield, England, Res. Rep. 202, 1962. 34. U.K. Ministry of Fuel and Power (now U.K. Dep. Energy). A Review of Electrical Research and Testing With Regard to Flame-Proof Enclosure and Intrinsic Safety of Electrical Ap- paratus and Circuits. 1943. 35. Underwriters' Laboratories, Inc. Electric Motor and Generators for Use in Hazardous Locations, Class II, Groups E, F, and G. Standards for Safety 674(A), 1973. (Updated periodically). 36. . Industrial Control Equipment for Use in Hazardous Locations. Standards for Safety 698, 1974. (Updated periodically.) 37. U.S. Code of Federal Regulations. Title 30 -Mineral Resources; Chapter I -Mine Safety and Health Administration, Department of Labor; Subchapter O-Coal Mine Health and Safe- ty; Part 18 -Electric Motor-Driven Equipment and Accessories; 1981. 38. U.S. Mine Safety and Health Administration. Coal Mine Safety Electrical Inspection Manual, Underground Coal Mines. Apr. 1979. 39. Widginton, D. W. Some Aspects of the Design of Intrinsical- ly Safe Circuits. SMRE, Sheffield, England, Res. Rep. 256, 1968. 40. Yao, C, J. deRis, S. N. Bajpai, and J. L. Buckley. Evaluation of Protection From Explosion Overpressure in AEC Gloveboxes (U.S. At. Energy Comm. contract AT(11-1)-1393). Dec. 1969. 41. Zabetakis, M. G. Flammability Characteristics of Combusti- ble Gases and Vapors. BuMines B 627, 1965. 396 CHAPTER 17.— MAINTENANCE 1 The ultimate goal of the power engineer is to maxi- mize system availability without compromising personnel safety. It has already been seen that protective circuitry with its relays, circuit breakers, and surge arrestors plays a vital role in achieving this goal. Yet system protection extends beyond protective circuitry: a good maintenance program can be equally important in providing a safe reliable mine system. Knowing the principles behind each type of maintenance and when and where to apply them can save both time and money as well as improve overall system safety. There are three types of maintenance: emergency, preventive, and predictive. While maintenance is a famil- iar topic for most engineers, this familiarity rarely ex- tends much beyond emergency repairs or routine mainte- nance. Typically, a component such as a motor fails, and in the heat of the production environment, maintenance crews make repairs as quickly as possible. Maintenance becomes an exercise in troubleshooting. Preventive maintenance (PM) is practiced by few mining operations, although its concept is as old as industrialization itself. PM is traditionally taken to be the periodic performance of various tasks that will extend the life of a component: an example is the regular lubrication of bearings. More generally, PM is taking measures that will prolong the useful life of a component, and then when failure is imminent, replacing it at a time when the minimum downtime and personnel hazard will be in- curred. This is the definition that will be used in this chapter. Techniques of predictive maintenance are rarely ap- plied in the mining industry, although the technology is routinely practiced in several other industries. Predictive maintenance is the prediction of component failure through measurement of key parameters and observation of any changes that occur over an extended period of time. An example is the measurement of insulation resistance and comparison with previously recorded values. When the recorded curve approaches a critical value, failure is impending and predictable. Advances in electronic tech- nology have opened the door for increasingly sophisticated prediction techniques, which, like the more common pre- dictive techniques, can be included as part of a compre- hensive maintenance program. In this chapter, various aspects of preventive and predictive maintenance programs are discussed, mainte- nance techniques are outlined, and finally, some of the elusive problems that can plague the mine electrical system are explored. But first, a few definitions are in order. The study of existing conditions in a variety of mine power systems indicates that a properly designed and main- tained power system is characterized by its safe, reliable, and economic operation; conversely, poorly designed or main- tained systems tend to be unsafe and experience low avail- ability (14). 2 Such conclusions are not surprising and are well known by power engineers working in surface transmission and distribution (13). What is significant is the close rela- 1 The author wishes to thank J. L. Kohler, assistant professor of mining engineering, The Pennsylvania State University, who prepared the original material for this chapter. 2 Italicized numbers in parentheses refer to items in the list of references at the end of this chapter. tionship between safety and availability. When the goal is to obtain the maximum availability of equipment, system safety is also maximized. Availability can be defined as the mean time of component operation, that is, the ratio of total operating time to the total time in which the component could have been operated (9). Mathematically, availability is a func- tion of reliability and maintainability, though it is impos- sible to define the precise balance of reliability to main- tainability that produces a specific availability. In order to understand and improve availability, it is necessary to examine the factors of reliability and maintainability more closely. Reliability can be defined as the mean time between failures. It is primarily affected by the component design, although poor reliability can be caused by abuse or im- proper application of equipment; for example, a portable cable rated for 8,000 V can be used on a 15-kV system, but the life and reliability of the cable will be considerably reduced. However, although unreliable components are sometimes found, and although there remain a small number of engineers who do not understand the principles of good system design, in general it can be stated that mine power-system reliability is essentially fixed. While manufacturers could try to improve the reliability of individual components, the results would probably not be commensurate with the effort and cost expended. Further- more, attempts to improve component performance often lead to increased complexity, which, paradoxically, can further degrade reliability. Based on this argument, it can be surmised that any significant improvement in the availability of a mine power system will have to come through increased maintainability. The U.S. Department of Defense has defined main- tainability as the quality of the combined features and characteristics of equipment design that ^permits mainte- nance to be accomplished under operating conditions by personnel of average skills (9). As implied, this factor is affected by the original design, but the skill and integrity of the maintenance personnel are also important factors in the maintainability of a component or of the full system. Another factor that is often overlooked is the ratio of PM to emergency maintenance. For instance, the maintainabil- ity of a trailing cable will be poorer if it is repaired only after a failure, as opposed to being repaired when signs of wear are discovered during a PM shift. In such a case, the reliability of the cable is also affected; this statement is substantiated by the fact that splices made on production shifts tend to be less reliable than those made on PM shifts. 3 To increase availability, the problem is then to im- prove maintainability. The initial responsibility for im- provement would appear to lie with the manufacturer, but once a system is operational, manufacturer changes rarely impact on total availability. Consequently, the burden lies on the maintenance effort at the mine. Emphasis must be switched from repair and emergency work to PM. This of course is easier said than done, and experience indicates that PM programs at the mine are generally haphazard and neglected (16). 3 Personal communication from R. H. King, Department of Mineral Engineering, The Pennsylvania State University, July 1980. 397 MINE MAINTENANCE PROGRAM At any mine or plant, a maintenance program must be justified before the company will release personnel, materi- als, and money. This is an unfortunate reality. The quantifi- cation of the costs and benefits of a PM program is always difficult, especially when maintenance and production de- partments have the misguided notion that they have oppos- ing goals. It is easy for a plant superintendent to visualize the program costs, but it is more difficult to see the long- range savings that will be reaped from a well-conceived PM program. Similarly, it is easy for a supervisor to compute the lost production costs when a machine is shut down for an hour of PM, but it is frustratingly difficult for the mainte- nance people to compute long-term savings due to increased machine availability. Unfortunately, the prevailing attitude among managers, from the faceboss to the superintendent, is that preventive maintenance is worthless unless proven otherwise. Economic Justification A good economic analysis is difficult to obtain if there are insufficient supportive data on component failure. At many operations, maintenance records of mine equipment are incomplete, at best. Consequently, it is necessary to make reasonable assumptions about the probable effects of a good PM program. Manufacturer data as well as data from related mine operations can be very helpful, but in the absence of this information there are some general guidelines that can be used to present a strong justifica- tion for a PM program. Three ratios are available for a first-order approxima- tion (25): cost of component repair or replacement cost of component PM cost of downtime cost of PM safety hazards cost of PM (17.2) (17.3) If the value of any one of these ratios exceeds 1.0, then PM can be justified. Although the application of these ratios should be obvious, an example of each is given. If a motor bearing fails, there is a good chance that the motor will be damaged and short out. The repair costs then become substantial because the motor has to be pulled and rebuilt. The PM costs in this case would be very small, consisting of monthly vibration measurements on the bearing. Ratio 17.1 is applicable to this case and, depending on the downtime involved, ratio 17.2 could also be used. The second ratio is more applicable to situations where the cost of the failed component is insignificant, but the failure causes a significant amount of downtime. For instance, if a contactor fails on the starter of the wound- rotor induction motor that drives the main conveyor belt, the contactor replacement costs will be minor compared with the cost for idling a major portion of the mine. Ratio 17.3 cannot be evaluated in dollars, at least not in good conscience. Although some industries, most nota- bly the automobile industry, have placed dollar values on human lives, this cannot be condoned. Instead, the goal should be to practice PM to the point of removing any reasonable chance of personnel injury. Other techniques, such as investment analysis, can be performed in the absence of extensive historical data. This method is probably best approached by considering the investment to be labor and overhead, and then conserva- tively estimating the effects of the PM program on the cost per unit of output. In the mine, the cost might be in dollars per raw ton, whereas in a preparation plant, it would be in dollars per clean ton. When such an estimation is carried out, the PM investment can be profit-adding rather than just profit-maintaining. However, it should also be noted that PM does follow the law of diminishing returns; carried to excess, it is not cost effective. Preventive Maintenance Program Implementation When the need for PM has been established, it is important that the top level of company management be involved in the decision to implement a suitable program. Since the returns from PM will initially lag behind its cost, such management support is crucial. Whenever com- pany profits fall and operating budgets must be slashed, the PM program is usually the first to suffer, but this is just the time when it is needed most. The backing of management can alleviate this industry-wide problem. The composition of the PM crew should be planned carefully. Experience in other industries indicates that at least half of the work crew should be seasoned mechanics who have a working knowledge of all the electromechan- ical system they will encounter. Circumstances may war- rant specialists. For example, as solid-state equipment becomes more prevalent in and out about the mine, an electronics technician would be a valuable addition to the PM crew. An established policy to have all PM personnel as salaried workers might be the most satisfactory. The maintenance superintendent typically assumes responsibility for the PM crew. However, this is not neces- sarily appropriate in the mining industry, where ideally the job should be filled by a separate individual very knowledgeable in electrical or mechanical engineering (preferably both), who has several years of mining experi- ence. The PM engineer should report directly to the plant or mine superintendent. Regardless of the crew makeup, a policy should be established that the PM personnel do not carry out normal repair maintenance. When implementing a PM program, an initial deci- sion must be made concerning the equipment to be in- cluded and the priorities to be followed. First priority should perhaps be given to new or expanding facilities. Maintainable equipment and installations do not just happen. From the planning stages through equipment purchasing to installation, a maintenance specialist should be involved in all the decisions. In fact, a recent study found that when this procedure was followed, the time required for a new preparation plant to reach capac- ity operation was reduced from an average of 3 yr to less than 18 months (23). When selecting the equipment for PM coverage in an existing operation, the components whose failure would have the greatest impact on production or safety should be con- sidered first. In an underground coal mine, these might be surface substations, ventilation fans and belt-drive motors, after which, rail-haulage motors and rectifiers, face machin- ery, and dewatering pump motors could be added. The 398 economic ratios presented earlier can be used to review the types of system components and assess different levels of coverage when designing the PM program. The next decision, and the one with the greatest impact on the success or failure of the program, concerns the organization of the data. The choice is basically between some type of card-filing system and a computer- ized system. Card files are rapidly becoming outdated; computers are becoming the normal operating method. Microcomputer systems now have extended capabilities and improved graphics and provide a convenient and flexible alternative that is adequate for handling the PM data of most companies. A wide variety of off-the-shelf software is available that can be readily tailored to indi- vidual company requirements. Such computerized card- filing systems have built-in analytical routines that sim- plify record keeping; they can print out summary reports of various types, print out daily calendars of work to be accomplished, and supplement data with graphs and charts. With minimal training, the PM crew can record data directly on to the system on a shift-by-shift, daily or weekly basis. TECHNIQUES OF PREVENTIVE MAINTENANCE The following sections discuss a variety of electrical and mechanical tests that will provide the data necessary for operating an effective PM program. lb these could be added a variety of other procedures; such simple tasks as measuring basic temperatures, clearances, or runout, for example, can provide the maintenance engineer with valuable information on the condition of equipment. In large transformers, spectroscopic analysis of the oil pro- vides sufficient data to detect many potential problems. Analysis of lubricating oils and greases can also be useful. The presence of a few parts per million of metal can indicate excessive wear or other problems. In oil-filled distribution transformers, for example, partial discharge on the windings is indicated by the presence of a few parts per million of copper (10). Many other techniques could be mentioned, but the objective of this chapter is to outline the measurements that would form the backbone of the PM program, and these are electrical and mechanical tests. In practice, it is important that the engineer not lose sight of the interrelationships between electrical and mechanical components in the overall system. It has been estimated that up to 75% of all electrical failures in motors can be caused by the initial failure of a nonelectrical component. Basic Electrical Measurements The main reason for taking measurements on the mine power system is to discover impending failures. Many potential problems can be detected with a simple voltmeter. Voltages on machines and at key points around the electrical system should be checked periodically to detect excessive voltage drops. Voltage drops of 20% to 30% under machine rated voltage are too common around mines, and the ramifications of such undervoltage can be serious, as stated in chapter 6. The voltmeter is also a valuable tool for troubleshooting the solid-state circuits that are becoming more common. Another useful voltmeter test is measurement of the voltage differences among phases at a given location, particularly at a transformer that supplies power to mo- tors. Motor damage can result if the voltage differences among phases are more than a few percent. In fact, one research group has found that the negative-sequence cur- rent caused by a 2% voltage difference can create enough heat to halve the life of a motor (25). The measurement of current is a simple task that can be performed for ac with the voltmeter and a CT or for dc with a resistance shunt. Current measurements taken periodically on a motor at no-load and full-load can provide valuable information on its condition. Although this type of measurement is better suited to stationary motors, as on belts or pumps, there is no reason why it cannot be used on mobile equipment. One specific application for current measurements is in the periodic checking of dual-motor belt drives. Here, it is normal for the load to be shared unequally by the two motors, for example, 52% to 48%, but it is not normal for the ratio to change by more than a few percent with time. If a significant change occurs, there may be an electrical or mechanical problem, such as a coupling misalignment. Resistance readings can also be invaluable. Although some of these readings must be made with a megohmmeter or bridge meter, many can be made with an inexpensive volt-ohm-milliammeter (VOM): resistor-bank values, open relay coils, and so on. In some cases, such as contact-resistance measurements, more accurate results can be obtained using a Kelvin bridge, although VOM measurements have some practical utility. Insulation Measurements Insulation is the only material used in electrical components that begins to deteriorate the moment it is manufactured. Strictly speaking, insulation should be called a dielectric, since it is simply one application of dielectric material (6). There are three primary stresses that can cause dielectric deterioration in the mine power system: mechanical, thermal, and electrical. Electrical stresses result from an overvoltage in excess of the rated withstand level of the dielectric. This may cause immedi- ate failure or gradual deterioration, depending on the level of the disturbance. Thermal and mechanical stresses are closely related. Most dielectrics do not have significant mechanical strength, so actual or circumferential stress may cause a failure, and they usually do not have much resistance to fatigue. Ther- mal stress, resulting from an operating temperature in excess of rated, causes a change in the physical properties of the insulator, which usually changes its electrical and me- chanical properties. The integrity of insulation is eroded as temperature increases, and several factors combine to cause a decrease in resistance or increase of current flow. The primary cause is the increased carrier mobility at elevated temperatures (6). Some types of insulation, especially some of the enamels used in motor wire, tend to become brittle, and as they expand and contract under the changing tem- perature conditions, they crack. The resulting leakage cur- rents will be larger than those in insulation with decreased resistance. Thermal stress can also manifest itself as a mechan- ical stress. This effect is more likely to occur where two or more dielectrics are sandwiched together, as in a portable cable. Since each insulator has its own particular coeffi- cient of thermal expansion, the expansion for each mate- rial will not be the same. This is particularly critical when an underlying material is being constrained, since me- chanical stresses will result. Strain may then occur, result- ing in voids or weakened areas. When failure eventually 399 takes place, it can appear to be mechanical or electrical in nature. Insulation is the material most likely to fail in a motor, and periodic measurements are a very valuable means of detecting incipient failure. The model of a dielectric shown in figure 17.1 indicates the parameters that can be measured to determine the integrity of the insulation (6). The two obvious ones are resistance and capacitance. Insulation resistance is measured directly: a perfect insulation would have R p = oo and R s = 0. Capacitance is not measured directly. Instead, 6, the phase angle between the voltage and the current, is measured, as shown in figure 17.2. The tangent of 5 (or the cotangent of 6) is known as the dissipation factor, and the cosine of 6 is called the unloaded power factor (6). At power angles close to 90°, the two factors are nearly equal. The voltage- current relationships for a sound dielectric will be almost purely capacitive, and the unloaded power factor and the dissipation factor should be very low or very close to zero. As insulation deteriorates, large leakage currents will tend to increase these factors significantly (24). The dissi- pation factor is used frequently in technical literature, whereas the unloaded power factor is used more fre- quently by manufacturers when specifying component tolerances (6). Hence, the latter might be more useful in preventive maintenance than the dissipation factor. Regardless of the factor selected, readings should be taken at periodic intervals, and records should be kept of each reading so that any trends can be detected. Typically, a manufacturer will specify the maximum acceptable unloaded power factor, usually 2% or less. If this value is exceeded, one of two problems is usually indicated: either 11 R, ic R« Figure 17.1.— Circuit modeling a dielectric. I Figure 17.2.— Current-voltage chracteristics in a dielectric. the insulation is severely deteriorated or it is water soaked. In either case, failure may be imminent, although allowing the insulation to dry would correct the latter problem. A suitable instrument for making these measure- ments is a capacitance and dissipation-factor bridge. Since the temperature affects the power-factor readings, a cor- rection factor must be applied. These factors are depen- dent upon the particular dielectric type and should there- fore be obtained from the manufacturer. The instrument connections to the apparatus being tested are identical to those of the megohmmeter tests, which are discussed in the next section. The shunt resistance of the leaky capacitor, R p , shown in the insulation model of figure 17.1, can be measured and used to predict insulation failure as well as to observe some measure of the insulation integrity. In practice, this measurement is usually performed with a megohmmeter. The current that flows because of the potential impressed on the dielectric is composed of two parts: a leakage component across the insulation surface and a current actually through the insulation. When the detection of deterioration is of interest, the superficial leakage current must be minimal, otherwise the insulation resistance reading will be artificially low. Consequently, it is neces- sary to examine those factors that can cause incorrect measurements. They are 1. Surface conditions: abrasion, foreign material, moisture; 2. Moisture; 3. Temperature; 4. Method of test: test-instrument potential, duration of test; 5. Residual charge. The surface condition of the insulation normally has the largest effect on the superficial current component; for instance, an abraded surface will collect airborne dust. If the dust is from a highly conductive material such as from carbon brushes, the possibility of establishing a conduc- tive path on the insulation surface is rather high. Even dust with a much lower conductivity (such as coal dust) can allow enough current to flow so a significant error is introduced. A nonabraded surface will have a tendency to collect some dust, but here the problem is not as severe. A conductive path can also be created if the dielectric surface is moist, and foreign material on the surface will aggravate the problem considerably. The surface moisture problem can usually be minimized by taking the measure- ments only when the temperature of the component is above the dewpoint. In the case of machines, surface moisture is of no significance at the normal operating temperature, but some researchers have expressed concern over the effect of high relative humidity during testing (4). While it is true that experiments conducted in a controlled environment do reveal a decreased insulation resistance with increased relative humidity, the changes are not troublesome if the data are correctly analyzed. More will be said about this later. The second factor is moisture other than surface moisture. Some insulating materials are hygroscopic; they absorb moisture. This effect is entirely different from that of surface moisture and presents a more serious problem because it compromises the dielectric integrity. Although a dielectric may exhibit a greater propensity for absorbing water as it ages, this characteristic is not a reliable 400 method of determining aging, at least not outside the laboratory. The hygroscopic moisture content can be reduced by bringing the component to normal operating temperature for a few hours before making any measurements. If a measurement is low and moisture absorption is suspected, caution should be observed when energizing the compo- nent, to prevent failure. In some cases, such as with large motors, it may be worthwhile passing a current equal to load current through the windings at a low voltage. The heating that occurs will dry out the windings. The third factor, insulation temperature, has a signif- icant effect on the measured resistance. If the insulation temperature is known, a normalized resistance can be calculated: R c = R t K t (17.4) where R c = insulation resistance corrected to 40° C, Q, R t = measured insulation resistance, fi, and K t = temperature coefficient. The temperature coefficient factor is obtained from a graph such as that shown in figure 17.3, which is valid for rotating machines (11). The measurement of insulation temperature is some- times difficult, but neglecting this factor can be risky. A suitable compromise can be reached in many cases; for example, taking the measurements consistently before or after the component is at operating temperature may stabilize the readings. The ambient temperature may vary over a small enough range to ignore its effect on the preoperating temperature of the component. Of course, steps such as these must be evaluated for each situation, and what might work well for a large motor could be meaningless in the case of a circuit breaker. The fourth factor, method of test, is important but presents no real application problems. The test potential of a megohmmeter is usually 500 to 1,000 V. Measure- ments on a given component should always be taken at the same test potential. The basic test format for a given component is fixed and will be suggested shortly. Beyond this, the test measures and procedures should always be duplicated precisely. Since the purpose of the test is to analyze a trend over time, small errors introduced by nuances of the test procedure will not pose a problem, provided that the methodology is consistent. Residual charge is the final factor. The leaky capaci- tor of the dielectric model must be discharged before a resistance measurement is taken. This is achieved by shorting to ground the outer skin of the dielectric and the enclosed conductor. The short should remain in place for a few minutes. When multiple readings are taken on the same insulation system, the dielectric should be shorted for at least 3 min, assuming the standard 3-min test (11). In specialized megohmmeter tests, where parameters such as the time and rate-of-rise of resistance are analyzed, this factor is critical. Megohmmeter Tests As was described in chapter 5, a megohmmeter is a portable instrument with an integral voltage source and a meter calibrated in megohms. The voltage source is either a handcrank generator or a battery pack. The type of instrument most suitable for routine insulation resistance measurements has only two probes, much like a voltmeter. Tests are conducted by placing the probes across the insulation system and measuring the resistance. Compo- nent manufacturers generally provide instructions for megohmmeter testing of their components. Three differ- ent tests are performed: spot reading, time resistance, and multi voltage. Spot Testing Readings for the spot test are taken consistently at 60s intervals, since the resistance in a good dry insulation will always increase with time, as shown in figure 17.4. -z. UJ o u_ u_ Ld o o UJ en cr UJ a. UJ o -z. < H CO CO UJ tr 0.05 h i i i i i i i i 20 40 60 80 WINDING TEMPERATURE, °C Figure 17.3.— Graph relating approximate insulation resistance variation with temperature for rotating machines. 4 6 8 10 TIME, min Figure 17.4.— Insulation resistance versus application time of test voltage. 401 Spot readings are taken periodically on system compo- nents and recorded so that any persistent downward trends that indicate potential failure can be detected. Conceptually, the most straightforward spot test is on cables, where one power conductor is checked at a time, with all other cable components shorted together and to ground. The connections for checking the line A conductor of a cable are shown in figure 17.5. After the connections are made, the test voltage is applied for 1 min and the final resistance value is recorded. Megohmmeter testing of motors is similar, but this test checks only the insulation system to ground and not the turn-to-turn insulation system. Figure 17.6 shows the test connections for line A testing of an ac motor, and figure 17.7 shows the dc motor connections. The minimum value for the spot resistance test should always be 1.0 MQ, plus 1.0 Mfi for each 1.0 kV of nominal voltage rating for the equipment. Consequently a low-voltage mine motor should have a minimum resistance of 2.0 MO. Typical spot resistance test records for motors are shown in figures 17.8 through 17.10. Disconnect other end of cable Any grounding conductor or shield "X Megohmmeter Figure 17.5.— Megohmmeter test connections for checking cable insulation in line A. Brush Megohmmeter Figure 17.7.— Megohmmeter test connections for dc motor. 100 c| 50 UJ o 11 CO UJ or 10 ~i ■ r 20 40 60 80 100 120 140 TIME, months Figure 17.8.— Spot resistance curve for normal motor. Megohmmeter 3 UJ u -z. < I- <£> CO UJ or t i i i I t~~ i r~ 1 I i i i i i_ Figure 17.6.— Megohmmeter test connections for ac motor. INCREMENTAL TIME — Figure 17.9.— Spot resistance curve showing effects of dust and moisture. 402 158 162 TIME, months Figure 17.10.— Spot resistance curve for defective motor. When spot-testing a transformer, the core iron must be grounded and the resistance connections or solidly grounded connections must be removed. All windings except the one under test must then be shorted together and grounded. The test connections are shown in figure 17.11. Similar testing procedures exist for starters, control boxes, relays, and circuit breakers, to name a few. Time-Resistance Tests With the same connections, time-resistance tests can be performed. Here successive readings are taken at specific time intervals to form time-resistance curves as shown in figure 17.12. Of particular interest is the point where the curve begins to level out, since a good insulation will have a continual increase in resistance with time (26). Figure 17.13 shows curves for an actual deteriorating motor. The polarization factor can be calculated by plotting the resistance values at 1 and 10 min: r _ R io Ri (17.5) The computed factor should have a minimum value of 2.0, as shown in figure 17.14. The polarization factor is inde- pendent of temperature and equipment size, which makes it very convenient for some applications. A polarization factor curve for a deteriorating motor is shown in figure 17.15. Multivoltage Tests Multivoltage tests are particularly valuable for as- sessing the efficiency of high-voltage components. Insula- tion resistance measurements are repeated at different Megohmmeter Grounding resistor Figure 17.11.— Megohmmeter test connections for transformer. TIME Figure 17.12.— Time-resistance curve. UJ o oo CO UJ Figure 17.13. — Three time-resistance curves for deteriorating motor. 403 100 120 Figure 17.14.— Time-resistance curves showing polarization for hypothetical motor. 77 80 83 86 89 92 95 98 101 104 107 110 TIME, months Figure 17.15.— Polarization factor curve for deteriorating motor. voltage levels, such as 100, 250, 500, and 1,000 V, and the resistance found at the higher voltage should always be equal to or greater than the resistance at the lower voltage. A large reduction in insulation resistance with measured voltage is indicative of insulation weakness, as shown in the curve for a deteriorating motor in figure 17.16 (26). Harmonic Analysis As insulation deteriorates, a small initial current that is rich in harmonics can be produced from electrical discharge across the voids (26). The mechanism producing the harmonics is probably partial discharge, which will be discussed later in this chapter. The presence of harmonic current provides a method of detecting insulation failure, 6 TIME, min Figure 17.16.— Multiple voltage curves for deteriorating motor. since researchers have shown that in new insulation the frequency component of a current is primarily that of the source, whereas in old insulation the current is composed of many frequencies (26). Harmonic analysis is carried out by a spectrum ana- lyzer, which performs a Fourier analysis on a signal and displays the amplitude of each frequency component. An instrumentation arrangement to carry out the analysis is shown in figure 17.17, where current is flowing from a 60-Hz source. The CRT display of the spectrum analyzer immediately indicates insulation of questionable integ- rity. The method has proved to be effective for cables, motors, and transformer-insulating oils. Power Factor Versus Voltage In this test the unloaded power factor is plotted against various voltage levels. For perfect insulators the resulting curve is a straight line, but in deteriorating dielectrics the curve rises with increased voltage to form a parabola, as shown in figure 17.18 (26). This is because voids in insulation contain entrapped air, which tends to ionize as voltage approaches the breakdown point, thus producing the extra loss that causes divergence of the plotted curve. A description of a simple power factor meter for in-mine measurements can be found in reference 26. Infrared Testing Although it is not an electrical test directly, the observation of infrared emissions with a portable instru- ment can detect abnormal hot spots in operating electrical 404 To unloaded system High-voltage transformer Figure 17.17.— Circuit for harmonic tests. 01 o h- o UJ o Q_ 1 Coil with / many voids / .. Coil with / few voids >^--- 1 Tip-up i i i i — '■ 1» VOLTAGE, kV Figure 17.18.— Power-factor versus voltage curves showing tip-up. equipment. At any temperature above absolute zero, all bodies radiate energy. The radiation in the infrared region is closely proportional to the body temperature. Excessive temperature can be caused by broken conductor strands or excessive insulation leakage in cables, defective coils in motors, or excessive leakage flux in transformers (26). When using routine infrared observations as part of the maintenance plan, any sudden temperature increase can indicate trouble. Real-Time Computer Analysis The two greatest problems concerning PM in the mining industry are probably the need to depend on personnel to make the necessary measurements and the inadequacy of existing techniques for predicting many of the failures that occur. The search for ways to alleviate these problems has been an ongoing process (18, 20). One recently developed method, real-time computer analysis, eliminates the human link in the data collection and analysis process. The method is also sensitive to deterio- ration processes that cannot be detected by more conven- tional tests. The technique is based upon the classification of a deterioration matrix composed of the values of elec- trical parameters, such as power-frequency harmonics and symmetrical components, that are collected continuously in a real-time environment. Economic analyses indicate that the cost of such a system could be amortized within 2 yr. However, practical implementation of the method is for the future. In the meantime, much can be gained through using the conventional tests described here. MECHANICAL MEASUREMENTS The failure of simple mechanical devices on a ma- chine, such as a bearing, can lead to a catastrophic electrical failure. Electrical problems can cause excessive vibration (as with open-end rings on rotor-bar circuits), which in turn can cause electrical deterioration as com- mutator damage, insulation cracking, and squirrel-cage rotor damage. In fact, a variety of electrical problems can mimic mechanical deterioration. The only way to pinpoint these things is to include mechanical measurements in the maintenance program. Indeed, in the case of rotating machines, it is the most applicable method. Vibration Machine vibration is an important symptom of elec- trical or mechanical problems. Electrical problems that can cause vibration include the presence of a subharmonic voltage component (15, 30, or 45 Hz) or shorted turns in one phase of a motor. Vibration can also be caused by such mechanical problems as unbalance, misalignment, or faulty parts. Unbalance can occur in a motor shaft or coupling as the result of improper installation, mishandling or defect. It may be static, occurring in a single plane, or dynamic, occurring in more than one plane. Whatever the cause, the vibration should be eliminated to prevent commutator damage, winding-lead fatigue or cracking of insulation. Some devices, such as fans, require balancing both before use and at periodic intervals. Coupler misalignment can also lead to vibration problems, especially when the motor coupling to a pump or air compressor is not in perfect alignment. Faulty parts can cause vibration. For example, when a bearing begins to fail, its vibration level will increase. If vibration levels are recorded routinely, for example, every 60 days, any significant increase will be detected, incipient failure can be predicted, and problems avoided. The measurement of vibration is not too difficult. The transducer used is typically a linear accelerometer, which for convenience is often attached to a magnet or probe, as shown in figure 17.19. The output signal is connected to an instrument that gives a readout of acceleration, veloc- ity (by integrating the acceleration signal), and displace- ment (the second integral of the acceleration signal). These instruments are also available with an oscilloscope display for examining the measurements in real time and in time or frequency domains. Although analysis of vibration data is beyond the scope of this chapter, a few comments are in order. First, taking readings involves engineering judgment. In the rather simple case of the motor and pump, readings would be taken along each axis at points A through D, as shown in figure 17.20. Careful analysis is required to avoid various pitfalls; here for example, an unwary reader might confuse cavitation with vibration. Table 17.1 gives a brief summary of some of the more common causes of vibration, and figure 17.21 provides some insight into the severity of different vibration levels (27). 405 Table 17.1.— Common causes of vibration Cause Amplitude Frequency Phase Remarks Unbalance Proportional to unbalance. Largest in radial direction. Misalignment of couplings Large in axial direction, or bearings and bent 50% or more of radial shaft. vibration. Bad bearings — antifriction Unsteady— use velocity Very high, several x type. measurement if possible, r/min . Eccentric journal Usually not large 1 x r/min Bad gears or gear noise... Low-use velocity measure Very high, gear teeth x if possible. r/min. Mechanical looseness No data 2 x r/min Bad drive belt Erratic or pulsing 1,2, 3, and 4 x r/min of belts. Electrical Disappears when power 1 x r/min or 1 or 2 x is turned off. synchronous frequency. 1 x r/min Single reference mark . Most common cause of vibration. Aerodynamic hydraulic No data 1 x r/min or number of forces. blades on fan or impeller x r/min. Reciprocating forces do 1 , 2 and higher orders x r/min. 1 x r/min usual; 2 and 3 Single, double, or triple.... Best found by appearance of large axial x r/min sometimes. vibration. Use dial indicators or other method for positive diagnosis. If sleeve bearing machine and no coupling alignment, balance the rotor. Erratic Bearing responsible, most likely the one nearest point of largest high-frequency vibration. Single mark If on gears, largest vibration in line with gear centers. If on motor or generator, vibration disappears when power is turned off. If on pump or blower, attempt to balance. Erratic No data. 2 reference marks. Usually accompanied by unbalance and/or Slightly erratic. misalignment. 1 or 2 depending on Strobe light best tool to freeze faulty belt. frequency. Usually unsteady. Single or rotating double If vibration amplitude drops off instantly mark. when power is turned off, cause is electrical. No data Rare as a cause of trouble except in cases of resonance. ..do. Inherent in reciprocating machines; can only be reduced by design changes or isolation. Magnetic probe BNC conne ction Magnet Magnetic probe Probe extension- Keeper (remove before placing magnet on machine) BNC connection Cable Figure 17.19.— Mounting techniques for two vibration transducers. Vertical Motor Shaft and Motor load , T j coupling / " Axia' Horizontal Figure 17.20.— Four typical vibration measurement points. E .E Z -O t_> 01 < 3 _1 u> Q_ O to 406 Acoustic Emission All rotating machines produce wide-band acoustic emissions that can provide considerable insight into the operating efficiency of a component. In fact, acoustic- emission analysis can predict most incipient mechanical failure, ranging from shaft defects and bearing problems to bad welds, many days in advance of the failure. It is a relatively new technology that has not yet been widely adopted in the mining industry. Portable instruments are available for measuring and displaying acoustic signals, using either contact or noncontact probes at specific loca- tions on the machine. Figure 17.22 illustrates the tech- nique (12). The distributions of greatest interest are the fre- quency and the amplitude, and both play a role in emis- sion analysis. A crack propagating in a bearing, for example, generates a narrow acoustic-emission pulse with a flat frequency distribution that is almost identical to friction noise in the bearing assembly. In this case, the frequency emission alerts the engineer to a potential problem. Subsequent analysis of the amplitude distribu- tion will identify the specific problem, since the emission pattern for a propagating crack and that caused by friction are readily distinguished. Acoustic emissions divide conveniently into low- frequency emissions, ranging from about 5 to 1,000 Hz, and high-frequency emissions that start around 65 kHz and rise into the megahertz range, but 1,000 kHz is considered the practical cutoff point. Although low- frequency signals can be used to diagnose many machine malfunctions, they have the disadvantage of containing many frequencies caused by normal phenomena, and these complicate interpretation of the readings (12). High- frequency emissions do not have this problem with back- ground noise, and furthermore, they are more defect oriented. The signals in the high-frequency range are rapidly attenuated, which makes it easier to locate the source of the emission. CONTINUOUS-MONITORING SYSTEMS Low-frequency vibration spectral data Envelope-detected high-frequency spectral data 1 — ■ — i — ■- Defective Good bearing 300 4000 100 FREQUENCY, Hz 200 300 400 Figure 17.22.— Comparison of acoustic-emission tech- niques for detecting failing roller bearing. Mine environment Sensors Communication link Figure 17.23.— Conceptual diagram of generalized mine monitoring and control system. Recent research in predictive techniques has focused on the development of reliable continuous-monitoring sys- tems that will automatically identify disturbances in patterns of normal operation and thus give warning of impending problems. The systems employ contact and noncontact sensors located at strategic locations on equip- ment and linked via a communication network to a central processing facility. Research designers favor distributed systems with outstations or local data processors that collect and display information and serve as a filter, passing only important data through to the central sta- tion. The outstations may have limited decision capability and may trigger actuators that are part of a system control network. These features are diagrammed in figure 17.23. Remote sensing and control systems have the potential to reduce downtime, thereby increasing productivity, and it is anticipated that the use of such systems will become routine in mines. Esoteric spectral analysis and remote sensing and control can seem far from the day-to-day routines of the electrical maintenance engineer who is concerned with basic repairs and troubleshooting defective equipment. Few tasks can be more daunting to the young, inexperi- enced engineer, armed only with theoretical knowledge, than being faced with a downed system while surrounded by an impatient production crew. Troubleshooting is a skill— some would claim an art— that is acquired through experience. Many of the techniques outlined earlier in this chapter are employed to diagnose equipment problems, and these together with troubleshooting tips and layouts provided by equipment manufacturers are usually suffi- cient to identify the cause of the trouble. There are, however, some instances where the problems are more puzzling and defy routine analysis; among these are corona and partial discharge. These topics will be dis- cussed in the remainder of this chapter. CORONA Corona is the name given to very small transient discharges that occur as part of the process of localized gaseous ionization associated with dielectric materials. Corona is most prevalent in high-voltage systems, and when the process reaches critical levels in regions of high electrical stress, the byproducts of ionization can cause degradation of insulation and lead to system failures. Awareness of the problem became widespread in the mining industry in the early 1970's when 12.47-kV distri- bution systems were adopted in several underground 407 mines. The increase in system integrity expected to result from the high voltage did not materialize; instead, the change was accompanied by anomalous failures in cou- plers, cables, and stress cones. Ensuing research added greatly to the understanding of partial-discharge phenom- ena and the practical implications for mine power systems. Corona is now recognized as a concomitant feature of high voltages. In this section the conditions for the inception of corona, its subsequent behavior, and the effects on compo- nents in the power system are outlined. Figure 17.24 shows a general graph of conduction effects in a gas such as air. These effects can be explained in terms of the ionic theory of conduction (28, 30). Up to point A in the graph, Ohm's law is valid; at point A, saturation occurs because of the space-charge density. After the potential is increased to point B, the gas begins to ionize and the field changes from a subdischarge field to an ionizing field. With any further increase in potential there will be complete breakdown; in other words, a discharge field is attained. The proportionality between current and voltage up to point A occurs because the electric field is so weak that it ionizes very few gas molecules. Hence, the number of ions and free electrons is very small compared with the number of gas molecules. Multiplication effects are minimal, so in general the current depends only on the mean speed of the ions and their relative numbers. The formation of ions and free electrons is approximately equal to the number of recombinations, and the current will depend on the rate of this progress. Saturation between points A and B occurs when the electron density reaches such a level as to decrease the field intensity; the flux of ions and electrons, thus current, remains constant. A further increase in potential causes the ions and free electrons to acquire sufficient velocity that, when they collide with neutral molecules, enough energy is imparted to split the mole- cules into ions and free electrons. This process increases exponentially; thus, the current also increases exponen- tially. This ionization by collision is the most important mechanism of conductivity in gases. The most complete breakdown in the process of ion- ization by collision gives rise to a spark. Essentially, a spark consists of a quantum of electrical energy traveling through the gas, with the associated current limited only by the source. It is very unstable because the passage of current lowers the voltage. Below complete breakdown, there are three other types of discharge that are known collectively and rather loosely as corona. These are the discharges of interest here. In descending order of magni- tude they are: • Brush discharge: These discharges are often consid- ered an anomaly of corona and consist of a very small number of sparks ending in air. • Corona discharge: The onset of this discharge is signaled by a glow or halo ("corona" is the Latin word meaning "crown"). • Partial discharge: This discharge cannot be de- tected visually; chemical changes occur in the gas. VOLTAGE Figure 17.24.— Conduction in gas. Ld CD < O > Ionization by collision C0r0na RriKh discharge J£SL Partial , discharge discharge ., 1 Spark Arc CURRENT Figure 17.25.— Discharge sequence in an ionizing field. for completeness, but this discharge is not a product of ionization by collision; it is an unstable condition that occurs when the electrodes are hot enough to supply electrons for the current, thus vaporize the anode, and create positive particles that heat the cathode by impact (see chapter 9). Any system where corona may be initiated is known as a corona source. The specific area of discharge, the corona site, is the source of transient currents that pulsate for a few nanoseconds with a current too small to be measured directly. When the corona pulses occur at regu- lar intervals for several minutes, the phenomenon is known as continuous corona; where the periods succeed each other at increasing intervals, it is known as intermit- tent corona. The frequency may vary from 1 pulse per minute for dc, to 100,000 pulses per second for 60-Hz ac (8). The shape of the corona pulse varies widely, being affected by the corona current at the site. Discharges are a response to electrical stress. Mathe- matically, the conditions for the occurrence and mainte- nance of the discharge are given by Townsend's continuity theorem (28): In this chapter the term corona will be used to describe the phenomena of partial, corona, and brush discharges. Figure 17.25 is a plot of the ionizing field for gases showing the different types of discharges. In any given situation, the sequence of discharges will always progress in the order shown, but the discharges can begin at any point in the sequence (28). The arc in the figure is included [ ' a[exp ( X (j8 - a)dx]dx = 1, (17.6) where a = ionizing coefficient of positive ions, j8 = ionizing coefficient of negative ions, and I = path of ionization (e.g., wire). 408 A very important part of Townsend's theory shows that discharges occur only in nonuniform electric fields; these are also the points of greatest electrical stress concentra- tion. Common geometric configurations conducive to high stresses and thus to the formation of corona are shown in figure 17.26. Corona Behavior The inception and behavior of corona from an ac potential can be illustrated by considering a single bare conductor that is parallel to a ground plane. In this case, the dielectric or insulator is air. A high-voltage test set is connected to the conductor, and as the voltage of the conductor is increased, the gradient between the conduc- tor and the ground plane rises. When the gradient reaches a critical value, the air molecules at the conductor inter- face are ionized. The voltage at this point is referred to as the corona inception level. The ionization rate is a function of the air tempera- ture, pressure, and the potential gradient. Higher gradi- ents result in higher ion and electron velocities, hence, greater energies. When these particles strike other mole- cules, the energy transfer is sufficient to cause ionization, and there will be an exponential increase in ionized gas. The ionization process will continue outward until the ionized particles no longer have sufficient energy to split any other molecules. During this process, the conductor will have the characteristic violet halo of corona discharge, extending outward for two or three times the conductor diameter. Closer examination of the conductor will reveal a group of bright beads, evenly spaced and superimposed on the otherwise uniform halo. The beads are negative co- rona, which occur on the negative side of the sign wave; the uniform glow is positive corona, which occurs on the positive half. The inception level for negative corona is higher than that for positive corona and negative corona is believed to be more destructive (22). Sensitive detection equipment has been used to determine the macroscopic ionization frequency, which varies from 1 to 10,000 Hz (8). If the line voltage is reduced to the corona inception level once more, the corona will be significantly reduced but it will not disappear. In the case of the bare conductor, a surface discharge, the corona extinction level is slightly below the inception level. For a discharge in a void, the extinction level is 15% to 20% less (24). In the literature, particularly the early literature, researchers tended to be indiscriminate in their use of corona terminology. Two terms are commonplace: thresh- old corona level, meaning the point at which it occurs, and visual corona level, meaning the point at which it can be seen. Since corona is now detected and quantified using sophisticated instrumentation that includes ultraviolet light imperceptible to the human eye, the interpretation of visual corona level is subject to confusion, and great care must be exercised when comparing data given in the literature. The characteristics of dc corona are basically the same as those of ac corona, for their respective polarities. However, dc corona usually occurs intermittently and at a slightly higher potential. Since the dc conductor will not change polarity, unlike the ac conductor, a surface charge is deposited by the initial discharge, which must leak away before another discharge can occur. This decreases the frequency of discharges. The ac corona is often consid- A Corner B Point Conductor C Interface Insulation D Abrupt change in electric E Irregular symmetry field Figure 17.26.— High-stress geometries. ered to be the mean of positive dc and negative dc coronas (15). Harmonics due to corona can range into several thou- sand, but the third harmonic is the chief development (21 ). There is a dual relationship between harmonics and corona: harmonics are caused by corona, and there are also corona losses associated with the harmonics. In transmis- sion systems, the latter are the most important. In 60-Hz systems, corona occurs during part of each half cycle, and the corona discharge is pulsated at twice the line frequency. This causes a cyclic change in line admittance, which results in a modified or distorted wave- form (6). The sinusoid has considerable harmonic content, particularly the third harmonic; the fifth, seventh, and ninth harmonics are often present. Fourier analysis of the symmetrical components of the current shows that the third harmonic must be composed only of zero-sequence currents or voltage; therefore, if the lines are connected in a grounded-wye configuration, the triple-frequency (third-harmonic) currents caused by corona will flow through the lines and into the ground loop. Zero-sequence currents do not flow in delta-connected systems; instead, a triple-frequency voltage pulsates between lines. The ap- pearance of other harmonics as currents or voltages is easily determined by knowing the sequence of the har- monic in question (18). Ionization during corona causes radiation of radio frequency noise. Below the visual corona level, radio frequency noise is negligible, on the order of 10 /xV. The radio interference voltage increases rapidly up to 100 or 200 /xV with the occurrence of visual corona (17). When air is ionized during a surface discharge, ozone and nitrogen oxides are the principal byproducts. If mois- ture is present, the nitrogen oxides combine with the water to form nitric acid, which causes gradual deteriora- tion of the insulation. Ozone is very unstable and changes quite rapidly to harmless diatomic oxygen, but a thin layer of ozone persists in the vicinity of the insulation. This can cause the insulation to harden, to become brittle; and if the site is under flexure stress, it will develop large cracks. The effects of corona in small or microscopic voids within insulation can cause more serious damage because the site is not visible to external inspection. Virtually all commercially available insulated conductors contain air within small random dielectric voids. Such voids usually 409 occur as bubbles with diameters ranging from 0.1 to 0.01 mm. Despite excellent quality control, present economics and technology prohibit the manufacture of insulated conductors that do not contain some voids. An air film may also exist between the conductor-semiconductor interface, or between layers of insulation, and occluded gases have also been observed in the interstices of the dielectric, as shown in figure 17.27 {24). Voids can also result from careless handling or poor splicing practices in the mine. Little is known about the air spaces within a given insulation. The size, location, distribution, pressure, and gas content of the voids are all unknown, making corona evaluation difficult. The boundary between the dielectric void and the insulation can be represented as two materi- als having dielectric constants E x and E 2 with a potential across them (fig. 17.260 (24). This series combination of the two materials results in an effective dielectric con- stant, E k , which is less than either E x or E 2 . From electromagnetic theory, it is known that this effect is due to the increased electric-filled intensity at the interface, caused by the discrete change of the dielectric property (7). The breakdown potential of the contained air is lower than that of surrounding air; thus in a void, corona can occur at a voltage that is below the rated corona extinction level. Degradation of insulation by corona can be explained by two separate mechanisms that result in the chemical decomposition of the dielectric. Electron bombardment is believed to be the primary mechanism; chemical degrada- tion is a secondary mechanism. As explained earlier, ionization by collision depends upon the continuation of energy transferal during collision. Depending on the po- tential gradient, the ionization extends outward for some finite distance. However, in a dielectric void, the contained gas will be ionized, and then the ions and free electrons strike the molecular structure of the dielectric. This en- ergy transfer is usually sufficient to break the weak bonds and produce volatile products with a lower molecular weight (5). As the void begins to deteriorate, the electrical stresses increase, causing a snowball effect, until the surface of the insulation is pierced. Depending on the location of the punchthrough and the specific cable appli- cation, a line-to-line or a line-to-ground fault may occur. The corona site may not result in an immediate failure but nevertheless remain active, continuing to deteriorate the dielectric material. During surface ionization, the chemical products of the process are solely responsible for dielectric damage, but in voids chemical degradation can be considered a secondary mechanism. Volatile byproducts of ionization, which include some very caustic acids, enhance the dete- rioration process during ion bombardment. Corona Detection Numerous tests have been devised to determine the presence of corona, but the tests have been designed primarily for checking corona in commercially produced cables and, hence, are of limited use in the mine. Never- theless, the principles underlying these devices could be applied with slight modification for mine use. Since the current of the corona pulse is too small to be measured, its effect on a traveling wave is monitored. The change in the traveling wave can be related to the apparent charge of the corona pulse, which is proportional to the damage caused by the corona. An apparent charge greater than 4 pC is usually considered harmful. A basic system for detecting corona discharge (fig. 17.28) consists of a partial discharge detector and display, power-separation filter, power supply and voltmeter, and high-voltage transformer. The power supply is used to deliver a voltage high enough to initiate corona. The 60-Hz signal and its harmonics are filtered out using the power- separation filter, leaving only the high frequency due to corona discharge. The detector contains the electronics necessary to integrate the pulses and determine the peak pulse values. Various circuits are used to perform these operations. Practically all detectors work on this principle, though variations are sometimes incorporated to obtain results under special or adverse conditions. The technol- ogy in this area is still changing rapidly. A fundamentally different technique that uses ultra- sonics has possible application to mine power systems. This method is based on the fact that corona breakdowns cause both audible and ultrasonic pressure waves at the A Missing semiconductor tape B Void created by conductor damage C Metal particle with associated air pocket D Bubble E Blister in splice jacket Figure 17.27.— Typical dielectric voids in cables. Partial-discharge detector and display -f- Power supply and voltmeter High-voltage transformer Figure 17.28.— Block diagram for corona-detection system. 410 corona source. If the medium containing the corona source is in free-moving air, these pressure waves can be detected by an ultrasonic transducer. Successful tests using a barium titanate transducer have been reported (2). The advantages of this method are the portability of the equipment and the relative simplicity with which the presence of discharges is determined. Disadvantages in- clude the need to add complex equipment to obtain quan- titative data. When traveling through a solid, the pressure wave will follow the path of least resistance, thereby increasing the time of travel and making it difficult to determine the exact location of the corona site. An addi- tional problem is that noise in transformer cores, known as magnetostriction noise, renders the ultrasonic detector useless around transformers. Partial-Discharge Problems in Mining In mine power systems, some corona destruction goes undetected until failure occurs. Subsequent analysis of failed cables, couplers, or stress cones frequently identifies the culprit as partial discharge, the lowest level of dis- charges associated with corona. In other cases, partial discharge is the suspected cause but a lack of corroborative data precludes a direct correlation. Instead, such failures can be attributed to anything from "bad cable" to "tran- sients," which may not be entirely accurate. In high-voltage systems that require a 15-kV insula- tion class, problems with partial discharge are common in cables and couplers. Most transients do not have sufficient energy to cause failure in good cable, but when a cable has been weakened by partial-discharge degradation, it be- comes susceptible to failure or at least to a higher rate of deterioration because of the increased stress. Ultimate failure can come from a single large energy transient or from a number of smaller transients over a period of time; hence, it is important to minimize transients in the mine power system. Partial discharge can be initiated by improper cable handling, particularly bending the cable sharply or pass- ing cable directly over metal parts or through insulation. Any abrupt change in the electric field along the cable length causes sufficient stress to initiate partial dis- charge. The subsequent chemical degradation can eventu- ally reach the conductor surface and terminate with a fault. Cable applications should therefore be investigated carefully before they are installed in power centers, distri- bution transformers, or other units. A specific installation can be partial-discharge resistant in one case, but suscep- tible in another. The severe discontinuity in the electric field that can occur at the termination of a high-voltage cable is a prime site for partial discharge, and various stress-relief systems are employed to prevent its inception. Some of these methods are shown in figure 17.29 and were discussed in detail in chapter 8. Similar problems are found in the confined spaces of high-voltage couplers, where stress relief is provided by preformed filler moldings or stress- control tape. With both methods, it is extremely important to eliminate air voids, since partial-discharge site can be created by any manufacturing or mounting defect. Couplers can be subjected to other damage caused by partial-discharge byproducts as the result of careless han- dling in the mine. A coupler with a folded insulator tube, for example, can create a restricted area with lower surface resistivity at the fold. Ozone and nitric acid can Copper braid Semiconductive tape Semiconductive tape Molded cone used with semiconductive tape and copper braid Cone molded with stress-control semi- conductive tape Stress-control tape Heat-shrink stress-control sleeve Stress-control tape Figure 17.29.— High-voltage cable terminations Stress-control heat- shrink form at this site and cause rapid deterioration of insula- tion inside the coupler housing. Recent research into discharge phenomena within high-voltage couplers identified typical sites for the forma- tion of partial discharge and demonstrated the importance of adopting impeccable standards of workmanship when repairing or reassembling couplers (19). It was shown that the smallest void, cavity, or discontinuity can become a partial-discharge site; examples are • Small nicks or cuts in cable insulation. • A void at the termination of extruded semiconduct- ing tape. • A service loop in the grounding strap that is too long and is located too close to the uninsulated power conductors when assembled within the shell. • An uninsulated void between the pin and the orig- inal cable insulation (fig. 17.30). • Nuts on conductor pins that, when turning, have caused conductor strands to untwist, forming voids. • Voids in insulation caused when pinned conductors are inserted into insulator tubes and the middle section of the coupler shell. (A well terminated coupler end might have discharge levels below 3 pC but may increase to 30 pC or more when inserted into the insulator tubes.) • Voids at the bottom of insulator tubes (fig. 17.31). • Voids formed when potting compounds cure and shrink away from the coupler shell. The use of a detector such as that diagrammed in figure 17.28 was found to be essential when attempting to isolate partial-discharge problems in failed couplers. The extinction level of partial discharge is an impor- tant parameter in all parts of the mine power system. The critical value for initiating or extinguishing partial dis- charge is the potential to the ground plane. Since the neutral in high-voltage mine distribution systems estab- lishes the ground plane, the main concern is line-to- neutral voltage. Obviously, the extinction level must be safely above the nominal line-to-neutral system voltage; 411 Figure 17.30.— Major insulation void sometimes found in high-voltage coupler terminations. Shell Figure 17.31.— Possible stress site in high-voltage coupler insulators. otherwise, partial discharge initiated perhaps by a tran- sient overvoltage would continue. A safety factor of 25% above the maximum line-to-neutral voltage is consistent with power-engineering practice (3). This converts to an 11-kV extinction level for the 15-kV couplers used on mine distribution systems. In general, the recommendations for reducing partial discharge in the mine system are • Minimize high-voltage transients. • When installing cables inside power centers, belt transformers, and so on, take care not to bend or stretch the cables unless they have been designed for that purpose. • Pay special attention to eliminating insulation voids and possible partial-discharge sites in high-voltage systems. INTERMACHINE ARCING Intermachine arcing, another maintenance problem, refers to electric arcing between the frames of under- ground electrical face equipment that is of sufficient magnitude to ignite explosive methane-air mixtures. Wolf (29) has described the sources and corrections for ac intermachine arcing on mobile face equipment, and this section summarizes his work. When the 1969 Coal Mine Health and Safety Act specified that all low-voltage and medium-voltage trailing cables for ac mobile equipment must contain an insulated conductor for the ground-monitoring circuit, mine opera- tors could not comply immediately because available ca- bles did not meet the specified requirements. The demand prompted conversion from three-conductor type G round and flat cables to a newly developed type G-GC. In this cable, one of the grounding conductors is insulated to serve as a pilot conductor, and the size of the remaining ground- ing conductors is increased. The result is a cable with asymmetrical cross section. In 1971, mining companies began replacing their mining equipment trailing cables with this new type, but soon after, sparks were observed arcing between equipment frames. Following exhaustive tests by the Mine Enforcement and Safety Administration (MESA, now MSHA), it was concluded that this sparking was caused by an induced voltage in the grounding con- ductors of continuous miner trailing cables that originated from the cable asymmetry. Subsequent investigations con- firmed that the energy released by the arc was incendive. A review of the conditions associated with this phe- nomenon indicates that the problem increases with trailing-cable length, the power demand of the machine, and the asymmetrical geometry of the cable. The presence of the insulated ground-check conductor in the cable causes a physical asymmetry that aggravates the problem. Such is true whether the cable is shielded or nonshielded. There are four basic methods for solving the induced- voltage problem and the subsequent arcs, which will be discussed in turn: 1. Transposition of the phase conductors, 2. Use of symmetrical-cable types, 3. Use of diodes to suppress arc currents, and 4. Use of saturable reactors to suppress arc currents. Transposition means to change the power-conductor position with respect to the individual neutral conductors at specific locations along the cable length. Induced volt- ages in the grounding conductors can be cancelled by dividing the cable precisely into thirds, and cutting and resplicing it as shown in figure 17.32. This method is based on the fact that if a grounding conductor is located the same distance from the three line conductors, the voltages induced by the line currents sum vectorially to zero. Equal transposition of the power conductors along the cable length has the same effect. Although this method is both simple and effective, its disadvantages inhibit its use in mining. It shortens cable life and physical strength, and in actual practice equal transposition lengths are difficult to maintain, and per- haps most important, correct transposition can be easily lost during subsequent splicing. The use of symmetrical cables is an alternative solu- tion. In theory, symmetrical cables exhibit no induced- voltage or arcing-current phenomena. The common sym- metrical cables are the type G round and type G + GC round (see chapter 8). However, a symmetrical cable that has been spooled or one that has been physically abused exhibits considerable distortion and is not truly symmet- rical. Further, any unbalanced power-conductor load or conductor defect will alter the cable symmetry, and the resulting induced voltage will be impressed on the ma- chine frame. Despite these problems, the use of symmet- rical cables appears to be the best long-term solution to the problem of intermachine arcing. It should be noted 412 that induced grounding-conductor voltages can be negated only if the continuity of each individual neutral conductor is assured through monitoring. This is obviously impossi- ble with SH-D cable types. Another approach to solving the problem is to insert a nonlinear impedance in the grounding circuit. The device must have a high effective impedance at low induced voltages, but a low effective impedance to the higher voltages and currents that are available during a ground fault. Any such device must have a continuous-current rating of at least 25 A and have a short-circuit capability equivalent to the cable grounding conductors. Diodes connected in a bridge arrangement, as shown in figure 17.33, apply this concept. The center diodes conduct at all times while the other diodes conduct every alternate half-cycle. The devices that are usually used conduct about 1.0 A with a voltage drop of 0.6 V. Since an incendive arc occurs with about 1.0 A, the danger of explosion can be eliminated with a sufficient number of diodes. For instance, with an induced voltage of 6.0 V, 10 diodes in series are needed, which means that a total of 12 diodes are required in the bridge arrangement. Because simultaneous line-to-neutral faults on different lines at different system points are possible, if the diodes fail, it is essential that they fail short, not open. A saturable reactor is a nonlinear element, which can limit arcing current without affecting fault current. The saturable reactor has several advantages over diodes, mainly, that there is no junction of semiconducting mate- rial subject to damage. The device is simply an iron-core inductor that has a high impedance to the point where the core is saturated. The inductance of the reactor is much less after saturation than before. This means that the device will limit induced voltages and arcing currents to a safe value, while operating as a linear device. For fault currents and voltages, the saturable reactor operates in both linear and nonlinear regions; in other words, the currents are not sinusoidal, but the effective impedance will be low because the greater part of operation is in the nonlinear region. Thus, the grounding system is still effective. The selection of a suitable saturable-reactor charac- teristic is somewhat similar to that for diodes. The main input is the amount of induced voltage present. Once this is known, a reactor can be chosen that will operate in a linear region for the induced voltage, limiting arc current to 1.0 A or less. The reactor should saturate for a slightly higher voltage and thus be effective in limiting machine frame potentials during faults. An acceptable saturable- reactor characteristic is shown in figure 17.34. Saturable reactors can be damaged by high separate- line, simultaneous-fault currents, and the winding conduc- tor can be fused by the large current if its capacity is not great enough. As with diodes, the device must fail short in these cases. Another problem is that saturable reactors can store energy, which lowers the atual incendive cur- rent. Such energy storage can be restricted by selecting a low quality factor (X/R ratio). Saturable reactors and diodes are equally effective if they are installed between the machine frame and the grounding conductor or in the power center between the grounding conductor and frame ground. But when they are installed in power centers, the grounding pin in metal-enclosed cable couplers must be isolated from ground; otherwise, the device will be shorted out. (This is the same precaution as that recommended in chapter 9 for wireless ground-check monitors.) '/3- point 2 /3- point I Red I i ! Red ! Red Ground i \ 71 !\ ■J' White ! \ ! White ! K ! White Ground ! V : iV Black \/ y \ i Black |/ > \ Black Ground Splice Splice Figure 17.32.— Power-conductor transposition on three- conductor type G cable. ■ Power center Figure 17.33.— Application of diode-suppression bridges in power center. 'rms 10 5 •rms 5 10 Figure 17.34.— Typical saturable-reactor characteristic. GROUND DIRECT CURRENT OFFSETS The dc offsets in grounding can be a serious problem wherever mixed ac-dc systems are used. Intermachine arcing can result from stray dc as well as from ac induced voltage on grounding conductors. The presence of dc ground currents on the ac grounding system can also cause 413 protective-relaying malfunctions with ground-fault relays and ground-check monitors. This section discusses the methods available to eliminate these offset currents, while information on the sources of the currents can be found in chapter 7. A discussion of corrective methods cannot be re- stricted to dc offset currents alone since they are only one anomaly caused by dc superimposed on the ac ground in a bipower ac-dc system. All major problems, including ground-bed deterioration by electrolysis, must be consid- ered. The analysis can be conveniently divided into three parts: » 1. Feasible power-system modifications, 2. Grounding-system construction, and 3. Water-distribution systems. Power-System Modifications The dc finds its way onto the ac ground by direct contact between equipment frames and the mine floor and also through two-wire dc shuttle car supplies where the negative lead is tied directly to the frame of the shuttle car and rectifier. In the first case, stray current from the mine's trolley system flows through the mine floor, roof or ribs, then moves through an ac equipment frame and into the ac safety-ground circuit. Stray current tends to flow in the mine floor because here it finds a lower resistance path back to the rectifier through the track. The resistance is usually higher in the track because it is difficult to bond track systems satisfactorily underground. Some bonding procedures are hard to follow and lead to improper bonds, while other methods are easier to use but result in a bond that is physically unsound and easily damaged. Even if a good bond is achieved, it can be destroyed by heavy rail usage or even the type of abuse caused when moving continuous miners into the section over the tracks. A parallel feeder, bonded at each rectifier and also bonded to the track at 100-ft intervals, should greatly attenuate the amplitude of stray dc flowing in the mine floor. Stray dc from trolley sources should present little problem when this arrangement is used in conjunction with careful ac equipment and rectifier placement, for example, positioning equipment a minimum of 25 ft from the track or, if this is impossible, placing ac equipment on insulating mats. The use of isolated or floating dc feed for shuttle cars would help to alleviate the problem with shuttle car loads. A two-wire negative-ground system is cheaper than a three-wire arrangement, but this represents a serious electrical compromise, as does diode grounding. A two- conductor type G shuttle car trailing cable is advocated to prevent offset problems. Although it is not in practice, ground-check monitoring of the dc cables is worthwhile. Grounding-System Construction The dc offset problems are minimized with correctly constructed grounding systems. Construction of a ground mat or grid is governed by three needs: to achieve a low value of earth resistance, to control potential gradients, and to prevent corrosion. Low earth resistance is achieved by placing a sufficient length or surface area of metal in intimate contact with the soil. If the soil has a high conductivity, less metal is required than if soil resistivity is high. For safe grounding, a very low resistance is required between the soil and the buried metallic grid. Potential gradients are controlled by the depth of burial beneath the surface and the placement of the grounding conductors in relation to one another. The manner in which a grounding grid responds to the flow of current through it depends on the magnitude and duration of the loading. Generally, all metal structures should be tied together to eliminate potential-gradient hazards. How- ever, corrosion protection necessitates the isolation of underground metallic structures from the corrosive effects of the soil. Detailed information on these subjects is provided in chapter 7. Water-Distribution Systems There is no doubt that underground water-pipe sys- tems can serve as excellent sinks for any stray currents moving through the mine floor. Water pipes should be isolated from current sources as much as possible. Where stray currents are present, nonconductive pipe sections could be inserted in the waterline at intervals, which would serve to limit the amount of current carried by the pipe. However, a problem remains because of the water inside the pipe. Although distilled water is an insulator, the presence of dissolved ions renders it conductive, and most mine water is rich in impurities. Hence, the best protection is to maintain isolation (as much as practical) of the water-distribution system from possible stray ground- current sources. The suggestions given here on dc offset ground cur- rents must be general in nature because no two mine systems have identical problems. Hence, guidelines are more effective than specific rulings in this instance. When this material is combined with that in chapter 7, a thorough list of correction methods for reducing the effects of ground dc offset currents can be assembled. SUMMARY The foregoing sections in this chapter have covered many aspects of maintenance, including justification, measurements, and planning, as well as some specific problems related to the subject. Many other areas fall under this general title and its supervision in the mining industry. These aspects are major portions of chapters 6 through 16, where they are presented in detail. Using this knowledge as background, the main objective of chapter 17 has been to integrate additional information into a concept for a feasible PM program. The importance of adequate PM cannot be overempha- sized. Through its employment, equipment availability can be maximized while personnel hazards are minimized. On the other hand, poorly done PM is worse than worth- less, and managers who have experienced this are very opposed to any approach to PM in their mines. Chapter 17 brings to a conclusion this publication. The objective has been to assemble a comprehensive engineering reference on mine power systems. Because not all aspects of mine electrical systems could be included and because this area of mining is rapidly changing and growing, the aim has been to collect as much significant information as possible that will provide the basic tools needed to continue a knowl- edgeable involvement in mine electrical applications. Such an involvement is important today, but in the future it will have even greater significance as the use of sophisticated electrical systems expands. 414 REFERENCES 1. American Society for Testing and Materials (Philadelphia, PA). 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Anaconda Co., Wire and Cable Div., Marion, IN, undated. 416 APPENDIX.— ABBREVIATIONS AND SYMBOLS UNIT OF MEASURE ABBREVIATIONS 1 A ampere Ah ampere-hour A/m ampere per meter A/m 2 ampere per square meter A//is ampere per microsecond As ampere-second A/s ampere per second A/V ampere per volt AAVb ampere per weber C coulomb °C degree Celsius cm centimeter cm 3 cubic centimeter C/m 2 coulomb per square meter cmil circular mil C/s couloumb per second C/V coulomb per volt °C/W degree Celsius per watt deg degree F farad °F degree Fahrenheit F/m farad per meter ft foot ft 3 cubic foot ft?/h cubic foot per hour ft- lb foot pound ft«lb/V foot pound per volt ft 3 /min cubic foot per minute ft//ts foot per microsecond g gram g/L gram per liter g/m 3 gram per cubic meter H henry h hour H" 1 recriprocal henry H/m henry per meter hp horsepower Hz hertz in inch in 2 square inch in 3 cubic inch in/ft inch per foot in/s inch per second J joule J/K joule per kelvin J/s joule per second K kelvin kA kiloampere kg kilogram kg/m 3 kilogram per cubic meter kn kilohm kn/v kilohm per volt kPa kilopascal kV kilovolt kVA kilovoltampere kVA/hp kilovoltampere per horsepower kvar kilovar 1 Standard IEEE format (rather than Bureau of Mines format) is used for unit of measure abbreviations in this publication. kV/ft kilovolt per foot kV/ M s kilovolt per microsecond kW kilowatt kWh kilowatthour kWh/°C-kg kilowatthour per degree Celsius L liter lb pound lb/ft pound per foot lb/Mft pound per thousand feet lb/mi pound per mile lb/yd pound per yard L/h liter per hour m meter m 2 square meter mA milliampere mA microampere m f microfarad mH millihenry pH/ft microhenry per foot min minute jtin microinch mJ millijoule mL/h milliliter per hour mm millmeter mm 2 square millimeter MQ megohm mU milliohm nil microhm MPa megapascal mi/h mile per hour ms millisecond m/s meter per second m/s 2 meter per second squared MS microsecond MV megavolt mV millivolt *»V microvolt MVA megavoltampere MV//1S megavolt per microsecond N newton N-m newton meter n ohm n/°c ohm per degree Celsius Q-cm ohm-centimeter fl/cm 2 ohm per square centimeter Q-cmil ohm-circular-mil fi-cmil-ft ohm-circular-mil-foot O-cmil/ft ohm-circular-mil per foot a/cmil-ft ohm per circular-mil-foot fi-ft ohm-foot fi-in ohm-inch Q-m ohm-meter O/Mft ohm per thousand feet fi/mi ohm per mile Q/V ohm per volt oz/ft 3 ounce per cubic foot pC picocoulomb pF picofarad psi pound per square inch psia pound per square inch, absolute psig pound per square inch, gauge rad radian rad/s radian per second 417 r/min revolution per minute r/s revolution per second S Siemens s second S/m Siemens per meter T tesla ton/ft 2 ton per square foot V volt VA voltampere VIA volt per ampere Vac volt, alternating current var voltampere reactive V/cm volt per centimeter Vdc volt, direct current V/m volt per meter V/Mft volt per thousand feet V/mi volt per mile V/jis volt per microsecond Vs volt-second W watt W/A watt per ampere Wb weber WbA weberampere Wb/A weber per ampere Wb/m 2 weber per square meter Wh watthour W/(m-°C) watt per meter degree Celsius wt% weight percent yd 3 cubic yard yr year OTHER ABBREVIATIONS AND ACRONYMS ac alternating current ACSR aluminum conductor steel reinforced AMSW ammeter switch AWG American Wire Gauge BIL basic impulse insulation level BS breaking strength cemf counterelectromotive force CFR Code of Federal Regulations (U.S.) ckt bkr circuit breaker CRT cathode-ray tube CSP chlorosulfonated polyethylene CT current transformer dc direct current DS disconnect switch emf electromotive force EPR ethylene propylene rubber FA forced air FET field-effect transistor FOA forced oil and air GCR ground-check relay GCS ground-check system GND ground GTR ground-trip relay h.s. high-strength (guy grade) IC integrated circuit ICEA Insulated Cable Engineers Association IEEE Institute of Electrical and Electronics Engineers LCD liquid-crystal display LED light-emitting diode MCC motor control center MCM thousand circular mils (wire gauge) MESA Mine Enforcement and Safety Administration (U.S.) MESG maximum experimental safe gap m-g motor-generator MOS metal oxide semiconductor MOV metal oxide varistor MSHA Mine Safety and Health Administration (U.S.) NBR nitrile butadiene rubber NC normally closed NEC National Electrical Code NEMA National Electrical Manufacturers Association NESC National Electrical Safety Code NO normally open NTIS National Technical Information Service, U.S. Department of Commerce OA over air (self-cooled) OCB oil circuit breaker OL overload relay, overhead line OTR overtemperature relay PCB polychlorinated biphenyl pf power factor PIV peak inverse voltage PM preventive maintenance PT potential transformer pu per unit PVC polyvinyl chloride RC remote control RFI radio frequency interference RM rotating machinery rms root-mean-square SA surge arrester SBR styrene butadiene rubber SCR silicon-controlled rectifier SEL sensitive earth-leakage (system) SF safety factor SI International System of Units s.m. Siemens-Martin (guy grade) TBS two-breaker skid TDR time-domain reflectometer tpdt three-pole double throw (switch) UV undervoltage UVR undervoltage relay VCB vacuum circuit breaker VOM volt-ohm-milliammeter VR voltage regulator VTVM vacuum-tube voltmeter WVDC working volts, direct current XLP crosslinked polyethylene 418 ELECTRICAL AND ELECTRONICS SYMBOLS A a D d F f GC H h I I J K k L I LF anode area turns ratio of transformer (chapter 3) unit vector eJ 120 (chapter 4) radius of rod, spacing between electrodes (chapter 7) susceptance, magnetic field flux density area (chapter 7) base (chapter 8) capacitance collector (chapter 5) capacity (chapter 8) controlled variable (chapter 14) electric flux density (chapter 5) diameter, in feet or meters (chapter 8) diode (chapter 9) conductor diameter; coil diameter, in inches (chapter 2) distance (chapter 11) electric field strength emitter (chapter 5) potential (chapter 7) difference or error (chapter 14) dielectric constant (chapter 17) magnetomotive force factor of safety (chapter 13) force, frequency (chapter 6) safe bending stress (chapter 8) conductance cable code: grounding conductor gate, galvanometer (chapter 5) ground-fault device (chapter 9) cable code: ground-check cable magnetic field strength feedback element (chapter 14) thickness total operation time for motor (chapter 6) current current phasor, total circuit current instantaneous current current density at electrode surface (chapter 7) current through circuit breaker, system current (chapter 11) conjugate of complex current (chapter 3) imaginary operator, V^T (chapter 2) Boltzmann constant (1.38 x 10- 23 J/K) (chapter 5) torque constant, proportionally constant (chapter 6) reflectance factor (chapter 7) winding series capacitance (chapter 11) cathode (chapter 14) relay (chapter 15) coefficient of coupling (chapter 3) inductance total length, distance, in feet or meters (chapters 7-8) length, in inches (chapter 2) path of ionization (chapter 17) actual average power consumed divided by rated average power (chapter 8) M MP MPF MP-GC N Q q R U y v v W X x mechanical moment mutual inductance (chapter 3) machine contractor (chapter 9) voltage-sensing relay (chapter 12) cable code: mine power cable code: mine power feeder cable code: mine power feeder with ground-check conductor north number turns of the coil (chapter 2) number semiconductor with excess negative charge (chapter 5) armature speed (chapter 6) power, permeance instantaneous power (chapter 2) semiconductor with excess positive charge (chapter 5) number of magnetic poles presented by stator (chapter 6) charge, reactive power electrical charge, stored electrical charge in capacitor resistance, reluctance reference variable (chapter 14) radial distance or moment arm (chapter 6) equivalent radius (chapter 7) apparent power south distance (chapter 8) transformer capacity (chapter 11) displacement (chapter 2) motor slip (chapter 6) s complex power SH cable code: shielded SH-C cable code; see tables 8.2, 8.3 SHC-GC cable code; see tables 8.2, 8.3 SH-D cable code; see tables 8.2, 8.3 SHD-GC, cable codes; see tables 8.2, 8.3 SHD+GC SS substrate T circuit configuration (chapter 2) junction temperature (chapter 5) torque (chapter 6) time (chapter 7) tensile strength, tension (chapter 8) t temperature time taper (chapter 8) Ta ambient temperature (chapter 5) velocity of propagation (chapter 11) voltage, electromotive force, potential difference voltage phasor instantaneous voltage energy, work weight wind (chapter 8) cable code; see tables 8.2, 8.3 work done (chapter 2) reactance distance (chapter 11) 419 Y admittance e y wye circuit configuration z impedance |Z| magnitude of impedance z burial depth a temperature coefficient (chapter 2) X \ no-load position, angular position (chapter 6) n firing angle (chapter 14) Mi ionizing coefficient of ions (chapter 17) P |3 a divided by (1 - a) a 7 conductivity A delta circuit configuration $ 5 soil density (chapter 7) e permittivity >7 efficiency (chapter 3) phase angle maximum allowable soil temperature rise (chapter 7) power-factor angle (chapters 4-5) protective angle (chapter 11) thermal resistance relative permittivity soil thermal conductivity (chapter 7) permeability (absolute) relative permeability resistivity conductivity soil specific heat (chapter 7) potential difference, magnetic flux flux direction (chapter 6) phase angle (chapter 2) flux (chapter 15) electric flux angular frequency 420 INDEX -A- Page Abnormal transient 282, 355 Accuracy 75, 118-120, 174, 208, 246-247, 270, 272-274, 331, 361, 362 Active power 63, 137, 319 Admittance 55, 64, 72, 268 Air circuit breaker 87, 227 Alternating current 2, 45, 48, 50, 129, 244, 246-247, 301, 347, 351 Alternating current mine power centers . . 280, 302-307, 310, 313, 315-317, 319-320, 325, 331-345 couplers 9, 13, 15-16, 191-195, 206, 208, 222, 224, 252, 254, 303-304, 317, 319, 360, 377-381 protective circuitry 13, 75, 98, 103, 159, 165, 179-180, 186, 224, 247 Alternating current reclosing breaker (see Recloser) Alternating current time-overcurrent relay . . 90, 166, 243-245, 247, 249-250, 256, 270- 271, 328-330, 334, 358, 360-362, 364-366 Aluminum 122, 130, 174, 178, 184, 194, 196, 203, 211, 216-217, 242, 244, 312, 339 bus 5, 7, 9, 10-11, 13, 108-109, 201, 224, 236, 255, 262, 264, 266, 268, 306, 308, 310, 312, 314, 322, 344 cables 184 connectors 5, 191-192, 208-210, 228-231, 313, 376-378 Faraday shields 298, 311, 379 transformer windings ... 68, 70-71, 74, 86, 95, 109, 180, 290, 308-311, 321-322, 341-383 Ambient temperatures, correction factor 195-196, 271, 399 American Wire Gage (AWG) 161, 184 Page Ammeters 51, 72, 89-90, 115-117, 119-120, 123-124, 127- 128, 273-274, 309, 317, 329, 344, 380, 398 Ampacity 195-199, 201, 255, 271, 276, 311, 313-314, 331 Ampere 20, 86, 93, 95-96, 103, 119, 154, 160-162, 171, 196-197, 229, 235, 240, 260, 262, 271 Ampere-turn 245, 272-273, 314, 331 Amplifiers 22, 109-114, 127, 353, 362 Apparent power 59-60, 62-63, 66, 68, 74, 81, 83, 93, 103, 118, 246, 319 Approval 186-187, 194, 207, 303, 382-383, 391-393 Arc . . 3, 97-98, 125, 150, 159-161, 166, 181, 192, 211, 215, 217, 224- 229, 232-234, 237-238, 255, 261, 275, 281-282, 285-287, 290, 292, 299-300, 306, 324-325, 335-336, 338, 360, 362, 367, 370, 376, 380-381, 384, 389-391, 393-395 Arc, incendive 362-384 Arc chutes 227-229 Arc fault 98, 261, 384, 390 Arc interruption 225, 227, 229, 232-234, 335 Arc quenching 98, 227, 229 Arc tracking 192, 390 Arcing, intermachines 166, 255, 384 Armature 1, 77, 130-132, 136, 146-154, 229, 241-242, 324, 348, 350 Armature reaction 147-148 Armature windings 130-132, 136, 146, 148-149 Arresters, surge 22, 292-296, 298-299, 301, 306-307, 310, 327, 332, 334, 337, 343, 396 Asymmetrical current 229, 262-263, 268, 313, 329, 344 Atoms 20, 104 Autotransformers 74, 114, 141, 159 Availability 177, 217, 220, 225, 233, 329, 332, 396-397 Avalanche diodes 298 -B- Backup relaying 254-256, 315, 360 Base line 9-11, 93, 219 Basic impulse insulation level (BIL) 290, 307, 336 Basic power circuit 76, 84 Battery 2, 42, 87, 117, 124, 192, 198, 330-331, 350, 359, 366-381, 391-392, 400 Battery boxes 374, 376, 392 Battery charging 367, 371-373, 377, 379-381 Battery tripping 331 Belt-conveyor starters 319, 353-355, 360, 365 Bias 104-106, 109-114, 199, 257, 346 Bidirection thyristor control 348 Bimetal 229-232 Blowout coils 324-325 Bolted fault 98, 102, 261, 263, 270, 275, 322 Bonds 164-166, 185, 207-208, 210-211, 215, 229, 252, 290, 339, 345, 367, 379 rails 216 welded 216 Borehole cable installation 203 Boric acid fuse 238-239, 276 Braking, dynamic 150-158 Branch 8, 35-39, 4M2, 45, 56-57, 65, 80, 220, 247, 274, 326-327, 329 Brazed connection 179 Breakers (see Circuit breakers) Breakover voltage 144, 346 Bridge circuits 34-37 Kelvin double 123, 398 Wheatstone 122 Bridge rectifiers 106, 143, 321-322, 324, 330, 348, 363 Broken-delta relaying 251 Brushes (see Motor brushes) Bucket-wheel excavators 8 Burden 8, 118-120, 246-247, 272-274, 314, 328-329, 331, 361-362, 365, 396 Bus 5, 9-12, 108, 201, 236, 255, 262, 264, 266, 268, 306, 310, 312, 322, 344 protection 7, 13, 314 Bus bar 109, 224 Bus bar connections 244 Butt-wrap grounding 299 421 -c- Page Page Cable-fault location 98, 102, 159, 207, 244, 263- 264, 266, 268, 328, 376 Cable-reel locomotives 367 Cables 1, 3, 14, 29, 81, 85, 87, 91, 101, 109, 162, 165-166, 216-220, 228, 230, 271, 275-276, 278-279, 283- 288, 297, 305, 307-310, 312, 319, 323, 329, 332, 354, 362, 377-380, 382, 384, 387-388 alternating current mine 2, 46, 48, 50, 129, 244, 246-247, 301, 351, 366 aluminum 184 ampacity 195-198, 201, 271, 313 battery 375 borehole 13, 186, 191, 203 capacitance 30, 268, 282, 285 charger 198, 350, 367, 370-373, 375-381 clamps, in mine shafts 193, 203-205 copper 184, 213 corona in 185, 192 couplers 5, 191, 222, 254, 304, 317, 379 direct current mine 30, 129, 131, 135, 158, 252, 256, 320, 324, 350 drag 183, 190, 205-207 entrance 193, 391 fault current 97-98, 102, 159-161, 171, 180, 218, 224-226, 229, 236, 243, 246, 254-255, 257, 260-264, 266-268, 270-277, 287, 292, 306, 310-315, 322-323, 329, 336, 339, 363, 375 feeder 9, 16, 76, 96, 103, 124, 154, 182, 184-185, 195, 197, 204, 211, 213-214, 260, 270, 289, 320 flameproof 395 flat 187-188, 193, 197, 210 high-voltage 13, 193 installation 193-194, 202, 204, 296 insulation 183, 185-187, 190, 193, 196, 199, 205, 207, 209-211, 398, 401, 403 interlocked armor 191, 203-204 jackets 183, 185-187, 193-195, 199, 203, 205-206, 208, 210, 393 locating faults 207 maintenance 202 mine power feeder 191 portable 2, 182, 184, 187, 193, 197, 201, 204, 206, 209, 222, 290, 396, 398 rating 159, 185, 195, 199 reactance 265 resistance 123, 164, 207 round 187, 188, 190, 193, 197-198, 204 shielded 98, 186, 190-191, 210 shuttle-car 183, 367 size selection 200, 202 splicing 183, 207, 210, 222, 258, 290 stresses 183, 191 symmetrical . . 82-83, 97-103, 117, 179, 186, 190, 225, 229, 232, 235-237, 239, 246, 250, 261-263, 266, 268, 270, 273, 298, 310, 313, 329, 331, 346, 354, 404 trailing (see Trailing cables) types 182, 186, 187, 190, 204, 257 voltage-drop criteria 195, 202 Capacitance 21, 28-31, 33, 45^8, 50-51, 53, 56-57, 59, 61-64, 81-82, 107, 115, 122, 160, 167, 170, 179, 186, 207, 260, 268, 276, 280, 282-288, 290-291, 295-298, 311, 319, 351, 359, 360, 363, 371, 383-384, 399 Capacitance switching 282-284, 286, 301, 332, 359 Capacitive circuit 301-302, 384 Capacitive reactance 54, 297 Capacitive susceptance 47 Capacitor motor 157 Capacitor-trip device 329 Capacitors 28, 30, 47-48, 54, 106, 111-113, 115, 124, 147, 154, 156-157, 207, 237, 248, 253, 280, 284, 292, 298, 328-331, 359, 399^00 power-factor improvements 283, 332 protection 296, 319 rating 2% surge 87, 260, 268, 282, 295-297, 307, 332 Capacity, interrupting 18, 224, 228-229, 232-234, 238, 268, 305, 310, 313, 325, 329, 336 Catalyst battery caps 375, 380 Cathode ray oscilloscope 126 Cathode spot 226-227, 233 Cathodic protection 178, 181 Certification 382, 389, 391 Chain conveyor, longwall 12 Characteristic impedance 284, 286, 288, 295, 307, 332 Charge 20-21, 28-29, 47, 104-105, 115, 126, 167, 179, 207, 225, 260, 280, 282, 284, 290, 295, 329- 330, 346, 368-370, 372-374, 376-377, 380 Charge cycle 369-370, 372-373, 380 Chargers 350, 367, 370-373, 375, 377-381 Charging, battery 367-368, 373, 377 Charging current 179, 260, 268, 284, 296, 297, 370, 380 Charging station 372-373, 375, 377, 380 Chemical treatment, ground beds 177, 180 Chopper motor control 349-350 Chopping transient 285, 296 Circuit breaker 7, 9-10, 13, 75, 86, 97-98, 179-180, 184, 224, 239-240, 244, 247-248, 250-256, 259-260, 263, 267-279, 282-283, 286, 290-292, 298, 301, 304-307, 310, 312-319, 322, 326-327, 329-336, 342, 344, 353, 356, 360-367, 377-378, 400, 402 air 87, 227 air magnetic 227-228, 232, 324 alternating current 232 contactor 90, 114, 150-151, 159, 257, 287, 324- 325, 350, 359, 379-380, 393, 397 direct current 232 frame size 228-230, 232 high-voltage 226-227, 232-233 low-voltage 226-228, 231 magnetic 227, 229-230 medium-voltage 226-228, 231 mine duty 229 minimum oil 232-234 molded case 226-232 oil, dead tank 233, 329, 335 oil, live tank 233 power 226-228, 332 ratings 227-228, 232-233 high-voltage 232-233 low-voltage 227 medium-voltage 227 molded case 228 oil 232-233 422 Page Page power 232 thermal-magnetic 229, 230-231 vacuum 5, 232, 234 Circuit reduction 31, 34, 36, 39, 41, 44, 56 Circuits ... 2, 5-9, 11, 13, 19-50, 53-59, 61-73, 75-79, 82-86, 89, 91, 93-95, 97-98, 100-117, 119-120, 122-125, 127-129, 136, 144-146, 150-152, 154, 159-161, 163, 179-182, 200, 215, 224-232, 234-242, 244, 246-263, 266-268, 270-272, 274, 278, 280, 282-292, 294-296, 298, 300-308, 310-319, 321-337, 339, 346, 348-365, 367-368, 370, 376- 380, 383-384, 389-390, 392, 393-396, 398, 404 alternating current 53, 57-58 bridge 34-37, 123 common-base 110-111 common-collector 110, 112 common-emitter 110-112 comparator 363-364 control 75, 141, 151, 154, 316, 324, 327, 329-330, 331, 349, 354, 370 direct current 56 distribution 237-238, 260, 308, 334, 343 integrated 114, 298, 364 magnetic 274, 311, 324 parallel 24-25, 44 protection 184, 216 rectifier 105-109, 117, 321, 346 regulator 349 series 22-25, 54, 76, 94 solid-state 360, 362 thyristor 113-114, 237, 287, 298, 323, 346-360, 370-372 transistor 109-114, 298, 356-359 triac 356-360 tripping 184, 216, 255, 305, 310, 318, 329- 330, 344, 356, 359, 367, 378 Circular mil 22, 184 Clapper relay 241-243, 356 Clear 19, 73, 90, 161, 220-221, 224, 228, 232, 234, 285-286, 306, 311, 329, 357 Close-and-latch current 225, 229, 267, 268, 305 Coal dust ignition 12, 17, 367, 382, 389, 391-395, 399 Coal preparation plants 394-395 Code, National Electrical (NEC) 195, 382, 395 Code of Federal Regulations (CFR) 19, 182, 207, 222, 259, 278, 395 Coefficient of coupling 66, 69 Coefficient of diffusion 374 Coefficient of grounding 293 Coils 29, 66, 69-70, 82, 115-120, 124, 133, 136, 138-139, 145, 150, 152, 231, 237, 241, 247-249, 251-253, 256-257, 271-272, 287, 290, 313-316, 323, 328, 330, 336-337, 356-361, 372-373, 398, 404 blowout 227, 324-325 long 26-27 polarity 22, 40, 65, 90, 98, 104-106, 109, 131-132, 161, 242, 244, 246, 258, 280, 318, 377, 380, 390 toroidal 28 Coincident demand 103 Commutating diodes 348 fields 148 poles 148 Commutation 148, 322, 348 Commutator 4, 131-132, 136, 147-148, 154, 350, 404 Compensating winding 148 Complex algebra 48-51, 53, 55 Complex power 59-62, 68, 73-74, 81 Complex quantity 52-55, 59 Components, symmetrical 98-103, 179, 250, 261- 262, 266, 268, 404 Compound generator 132 Compound motor 148-149, 151 Computers, use of 114-115, 125-126, 200, 202, 260, 263, 268, 278, 2%, 364, 398, 404 Concentration 9, 13, 15, 29, 104-105, 127, 133, 175, 287, 300, 344, 373-375, 382, 389, 393-394 Conductance 22, 25, 32-33, 55, 64, 67, 292 Conductivity 104, 165, 171, 174-175, 177-178, 184, 201, 211-212, 251, 311-312, 399 Conductors . . 2, 5, 8, 13-14, 20-21, 26-29, 65, 68, 73-75, 77, 79, 85- 86, 91, 97-98, 100-104, 113, 118, 120, 124, 127, 129, 132, 133-134, 136, 137, 141-142, 146-149, 154, 156, 160-164, 166, 167, 171, 172, 174, 178, 180-183, 192, 199, 200, 208, 211, 213, 216, 224, 227, 229, 231, 233, 236-237, 244-250, 252-253, 256, 259, 261, 273, 276, 280-281, 287- 288, 290, 292, 299, 304-305, 307, 310- 312, 314-318, 324, 328-331, 334, 338-342, 354, 375, 377-380, 383- 384, 388, 390-393, 400, 404 aluminum 184, 194, 196, 312, 313 ampacity 251, 275-276 cable 9, 81, 165, 184-185, 187, 194-195, 210, 217, 220, 257-258, 262, 268, 275, 329, 332, 401 capacitance 167, 170, 268 grounding 159, 165-166, 186-187, 191, 193, 218, 255, 257-258, 319, 323 impedance 101, 179 inductance 170, 176, 273, 294 insulation 71, 130, 138-139, 183, 185, 186-187, 194, 202, 204, 206-207, 209-210, 271, 344 reactance 139 resistance 22, 31, 70, 123, 139-140, 159, 195 shields 186-187, 190, 298, 337 size 3, 184, 195-197, 199, 202, 217, 230, 264, 275 spacing, overhead lines 217 span, overhead lines 217-221 stranded 185, 201 strength 184, 202-203, 209 Conduit 219, 297, 388, 393 Conjugate 49, 60, 97 Connected load 5, 84, 100-101, 103, 151, 270, 285, 308, 334 Connections 3, 9, 78-79, 82, 84, 86, 92, 95, 116, 118-120, 124, 128, 132, 136, 148-149, 151, 157, 159-160, 181, 187, 192-193, 207-209, 211, 224, 231, 244, 246- 249, 250-251, 261, 272, 289, 294, 2%, 299, 307-308, 313-314, 316, 328, 332, 337, 339, 341, 343, 356, 379-380, 387, 399, 401-402 Connectors 5, 191-192, 208-209, 228, 231, 313, 376-377, 391 Contact ... 98, 129, 132, 141, 143-145, 150-152, 159-162, 164, 166, 169, 183, 186-187, 190-194, 204-205, 209-210, 215- 216, 218-222, 225, 226-229, 231, 233-235, 237, 240-243, 244-246, 248, 251-254, 256-259, 263, 268, 270-271, 278, 280, 282, 284-287, 295, 298-299, 302, 304-306, 310, 312, 314, 317-319, 323-325, 328-330, 336, 339, 353, 356-360, 363, 375, 377-379, 381-385, 389-390, 392-393, 398 423 Page Page bounce 286, 360 breaker 248, 254, 259, 285-286, 323-324, 328-329 relay 240, 242, 253, 314, 330, 358-359 switch 253, 305 whiskers 286 Contactor 90, 114, 150-151, 159, 257, 287, 324- 325, 350, 359, 379-380, 393, 397 Continuity of service 6, 276 Continuous of current . . . 195-197, 225-226, 228-234, 237-238, 245, 260, 274, 276, 305-306, 308, 312-314, 325, 329, 331, 335-337, 355-356, 362 Continuous miner ... 3, 12-13, 15, 81, 103, 153-154, 190, 198-199, 201, 209, 308, 313-314, 319, 350-351, 391 Continuous rating 153, 228, 231, 236, 239 Control systems 4, 349, 352-354, 365 Control transformer 306, 315-316, 324, 326, 329, 343 Control wiring 224, 226, 248 Controllers (see Starters) Convenience outlets 247, 316 Conventional mining 3, 11-13, 182, 367 Conveyors 3, 8, 12-13, 16, 113, 140-141, 143, 153, 182, 308, 342, 350-354, 356, 360, 367, 392, 394-395, 397 Cooling 7, 106, 134, 155, 183, 191, 197, 227-228, 233, 238, 307, 310-311, 334, 350, 354, 384-385, 387 transformers 307, 311, 334 Coordination, protective relaying 254, 336 Coordination-curve plots 276-278 Copper 22, 70, 136, 139, 161, 177-179, 184-186, 190, 193, 195-196, 200-203, 211-213, 215-216, 226, 231, 236, 307, 311-313, 339, 390, 398 cables 184, 213 Faraday shields 186, 201, 311 Core loss 69-70, 72-73, 137, 310 Cores 27, 65-66, 68-70, 72-75, 82, 109, 119, 124, 127, 133, 136-140, 149, 177, 208, 246, 250, 258-259, 272, 285, 310-311, 324, 361, 370-371, 379, 402 rotor 130 stator 132, 137-139 transformer 75, 87 Corona 185, 192 Corrosion 180, 184, 192, 194, 203, 375-376, 380 conductor 178, 339 electrode 177-178, 181 explosion-proof enclosures 392 ground-bed conductors 177, 180 Coulomb 20 Counterelectromotive force (cemf) 149 Counterpoise 299, 301 Couplers, cable (see Cable couplers) Coupling 9, 64-67, 69, 119, 160, 163, 192, 194, 291, 317-318, 338, 360, 398 magnetic 27, 65-67, 105, 145 motor 145 Crest voltage 47, 51, 285-286, 292-293, 295-296, 306 Cross bonding 166, 216 Cross field, motor 156-157 Crosslinked polyethylene insulation 185 Crowbars 298 Cumulative compound motor 151 Current . . 2-3, 18, 20-48, 50-57, 59-76, 78-95, 97-127, 129, 131-133, 135-142, 145-157, 159, 160-167, 170-176, 178, 179-181, 183-186, 190-192, 194-199, 201-202, 206-207, 211, 215-216, 218, 220-222, 224-232, 234-268, 270- 290, 292-301, 305-309, 310-326, 328-329, 331-332, 335-341, 343-344, 346, 348- 365, 367-373, 375-380, 384, 390, 392, 394-395, 398^00, 403 asymmetrical 229, 263, 267-268, 329, 344 balance 25, 79, 249, 261 carrying capacities 192, 196, 211, 216, 222 bus 322 charging 179, 260, 268, 284, 296-297, 370, 380 chopping 284 continuous 195-197, 225-226, 228-234, 237-238, 245, 260, 274, 276, 305-306, 308, 312-314, 325, 329, 331, 335-337, 355-356, 362 definition 20 density 26, 119, 129, 133, 137, 140, 148, 167, 171 earth 166, 323 eddy 70, 136-137, 311 exciting 69-73, 94-95, 274, 295, 308 fault 97-98, 102, 159-161, 171, 180, 218, 224-226, 229, 236, 243, 246, 254-255, 257, 260-264, 266-268, 270-273, 275-277, 287, 292, 306, 310-313, 315, 322- 323, 329, 331, 336, 339, 363, 375 inrush, motor 277, 283 inrush, transformer 150, 244, 270, 276, 285- 286, 295-296, 306, 337 interrupting 225, 229, 232-234, 238, 268, 305, 310, 313, 325, 329, 336 leakage 190, 206-207, 356, 360, 375, 399 line 79-80, 83-85, 92-93, 97, 100-102, 118, 120, 137, 141, 146, 197, 225, 245, 249-250, 254, 260-261, 263, 346, 353-354, 362 locked-rotor 141 loop 36-38, 57, 129 magnetizing 69-70, 72, 284-285 momentary 225, 360 motor-starting 125, 143, 145, 151, 157, 200, 236, 238- 239, 277, 283, 298, 318, 363-365 overload 274, 275 phase 79-80, 83-84, 92, 101, 108, 120, 250 short-circuit . . 44, 73, 202, 225, 229, 232, 236-237, 260-264, 270, 275, 277, 310, 313, 321, 328, 331, 336, 344 symmetrical three-phase 263, 313 transient 283 unbalanced 102, 247 Current-balancing transformers 109 Current-limiting capacity 18, 87, 224 Current-limiting fuses 235-237, 239, 247, 276, 306, 338 Current-limiting resistors 165 Current transformers 83, 108 accuracy 118, 246, 272-273 burden 118, 246, 272-274, 328-329 errors 272-275 model 272 ratios 244-246, 272-274 relaying 89, 118, 244-246, 249, 272-274, 328-329 saturation 272-273 Cutting machine 12-13, 183, 187, 198 Cylindrical-rotor motor 144-145 424 -D- Page Page Damper winding 145 Damping 242-243, 283, 285 D'Arsonval meter 115-117, 119-120, 124-125 Data ... 19, 86, 126, 180-181, 197-199, 202, 211, 222, 266, 273-274, 277, 331, 364-365, 381, 387, 393, 395, 397-399, 404 cable 213, 271 Dead lock 14 Dead front 312, 316-317 Decrement factor 267 Deep-bar rotor (see Rotor) Definite-time relay 151 Delta connections 35, 79-80, 82, 84-85, 100, 108, 138, 141, 258, 296, 308, 321 Delta-delta transformers 83, 92, 261, 308, 343-344 Delta-wye transformations 34, 83-84, 92 Delta-wye transformers 90, 92-93, 261, 308-309, 344 Demand 24, 17, 68, 103, 122, 153-155, 197, 199, 201, 255, 271, 275-276, 302-303, 308, 312, 314, 331, 335, 345, 350, 364 Demand factor 103, 308, 335 Demand meter 90, 122, 127, 345 Diagrams, one-line 86, 91-92, 95-96, 163, 201, 255, 262, 264, 276-277, 302-303, 333-334, 336, 344 Dielectric 28, 186, 191-192, 210, 221, 228, 234, 280, 285- 286, 290, 292, 295, 300, 398400, 403 Different-current relaying (direct current) 36, 323-324, 336, 337, 367 Differential compound motor 151 Differential protective relaying 90 Diode 3, 22, 90, 104-105, 108-112, 114, 125, 165, 181, 237, 257-258, 275, 287, 298, 322-323, 329, 346, 348-349, 354, 363, 370 Diode grounding 164, 187, 194, 256-258, 275, 323 Direct current cables 202, 256 Direct current circuits 21-22, 28, 30-31, 45, 50, 53, 55-56, 58, 107, 117, 166, 227, 232, 242, 247, 253, 258-259, 286, 298, 321- 322, 324, 378 Direct current circuit breaker 90, 228 Direct current generators 2, 90, 133, 147-148, 150, 152 commutator 131-132 excitation 132 Direct current ground-fault relaying 259, 323 Direct current motors 3, 147, 156, 198, 287, 351, 380, 401 chopper-driven 349-350 compound 148-149, 151, 153, 348 separately excited 4, 132, 148, 152 series 2, 4, 135, 148, 151-154, 320, 348 shunt 148-151 Direct current offset current 166 Direct current overcurrent relay 90, 247, 256, 361 Direct relaying 246-250, 255 Directional relay 88, 90, 240, 242, 244 Discharge, partial 185-186, 192, 393, 403 Disconnect switch 5, 9, 13, 220, 226, 237, 254, 256, 305-306, 326-327, 336, 338, 393 Discontinuity 207, 283-289 Dissipation factor 399 Distribution . . . 2-11, 13-15, 17-19, 22, 25, 30-31, 35, 55, 75, 78-79, 82, 96, 103, 107, 124, 128, 131, 147, 153-154, 158- 159, 161-162, 166, 174, 179, 181-183, 185-186, 190-191, 197, 199-202, 211, 216-219, 222, 224, 228, 234, 237-238, 244, 254-255, 259-260, 262-263, 266, 272, 277-281, 284, 286-287, 291-299, 301, 303, 304-311, 316, 319, 321, 325-326, 329, 331-338, 340, 342-345, 350, 361-362, 365, 370, 381-382, 391, 396, 398 Distribution system ... 3, 7, 13, 124, 147, 153, 159, 166, 182, 190- 191, 200-201, 211, 234, 272, 278-279, 281, 292, 295-297, 305, 308, 319, 325, 329, 331-332, 336, 344, 365, 370 direct current 2, 164-165, 187-188, 226-227, 229, 232, 247, 275, 286, 287, 298, 322, 350 open pit mine 8, 11, 19, 185, 216, 219, 255, 301, 345 preparation plant 8, 17-19, 129, 334, 340, 342, 382, 384, 394-395, 397 primary-selective 9 radial 5-6, 10, 326, 333-334 secondary-selective 6, 334 strip mine 4-5, 9-11, 185, 216, 219-220, 232 Distribution transformers 5, 7, 13, 75, 262, 297, 303, 309, 398 Double-cage rotor 140-141 Double-ended substation 7, 334 Double switchhouse 5, 9, 326-327, 330 Draglines 4-5, 8, 166, 182-183, 190 Driving potential 30, 102 Drum controller 150 Dry-type transformer 293, 301, 337, 379 Dual-element fuse 235-236, 239, 276, 312, 393 Duality 64 Dust-cap, receptacle 304-305 Dust-ignition-proof enclosures 134, 382-383, 394 Dust-tight enclosures 394 Duty-cycle operation 153-156, 197-199, 314 Dv/dt protection 323 Dynamic breaking 150-152 Dynamometer 115, 117-120, 122 -E- Earth resistance 167-168, 174, 178, 180-181 Earth resistivity 178, 180-181, 301, 341 Eddy currents 70, 136-137, 311 Efficiency 3, 12, 59, 71, 73, 77, 81, 84, 106, 137, 139- 141, 154, 158, 186, 197-198, 232, 239, 284, 292, 307-308, 313, 319, 329 motors 138, 270 transformers 146 Electric charge 20, 28 Electric field 28-29, 104-105, 167, 181, 186, 225, 281, 286, 290, 394 Electrochemical cell 178, 367 Electrocution 208, 217, 219, 221, 317, 377-378 Electrodes, grounding 172, 177, 181, 299, 301 425 Page Page Electromagnetic attraction 240-243, 356-358 Electromagnetic induction 240, 242, 244, 356, 358 Electromagnetic torque 133 Electromechanical devices 240, 362 Electromotive force (emf) 130, 290 Electron theory 20 Electrostatic force 20 Electrostatic shielding 311 Emergency-stop switch 254 Enclosure 130, 159, 178, 220, 232, 236, 238, 240, 302-305, 320, 326, 329, 332, 378-380 battery 373-374, 377 dust-tight 389, 394 explosion-proof 382-387, 389-392, 395 joints 385 metal-clad 303, 326, 336 motors 134 Endosmosis 171 Equations 22-23, 25-26, 28, 31, 34, 39, 42, 48-55, 59-66, 69, 74-80, 84-85, 92-96, 99-101, 105-107, 110-111, 120, 123, 137, 146-149, 155-156, 162, 170- 171, 198, 263, 271, 282-283, 285, 288- 289, 291, 295-297, 313, 368, 373 circuit 30-36 loop 36-40, 43, 57, 64, 67, 102 node 38, 40, 43 Equipment 1-19, 62-65, 75-76, 81-82, 85, 93-98, 103-108, 114- 117, 120-125, 129, 135, 147, 153-154, 159-166, 174, 178-187, 190-191, 194, 197-201, 204-206, 218-222, 224, 228-236, 239-241, 246-248, 251, 263-264, 265, 268-270, 274-281, 287- 313, 316-328, 331-340, 343-345, 350- 354, 360-362, 367, 375, 377-379, 382-383, 387, 389-398, 401-404 mine power 4 mining 2 Equivalence 25, 31, 33-34, 42, 80 Equivalent circuits 40, 57, 68, 70-71, 73, 76, 85, 94-95, 113- 114, 260-261, 266, 284-285, 291 (EPR) insulation 185 Error, current transformer (CT) 272-275 Ethylene propylene rubber 185, 222 Euler's theorem 49, 52 Excavators, surface 129, 147, 152, 252, 278, 350 Excitation 72, 90, 132, 143-146, 150-151, 260-261, 272, 348 Exciter 90, 143, 147, 152 Explosion proof 134, 191-192, 315, 382-392, 394-395 Explosion-proof enclosures 382-387, 389-392, 395 Extended-time rating 179, 312 Extinction, corona 185, 192 -F- Factors 3, 19, 52, 63-64, 81, 83-85, 90-91, 99, 105, 107, 110, 117-119, 123, 127-128, 146-147, 153-154, 161, 166- 167, 169-170, 174, 180, 184-185, 191, 195, 199- 200, 201-203, 206, 208, 217, 219, 221, 225, 234, 246, 261-264, 268, 270-272, 275, 278, 280, 290, 293, 308, 311, 313, 319, 328, 331, 335, 339, 342, 357, 363, 365, 367, 372-374, 377, 380, 383, 386-387, 389- 390, 392-393, 396, 398, 400, 403 ambient temperature 196 decrement 267 demand 103 dissipation 399 diversity 103 elevation 155 load 61-62, 80, 103, 138-140, 197-198 polarization 402 power 59, 62, 114, 120, 138, 201, 399 reflection 139, 295 service 134 Fall-of-potential measurements 172, 179 Farad 295 Faraday shield 298, 311, 379 Faraday's law 129, 290 Fault 88, 97-98, 101-102, 119, 159-163, 165-166, 171, 179-181, 184, 186, 192, 207, 218, 220, 222, 224-226, 229, 235-236, 243-244, 246-247, 248-251, 253-255, 257-268, 270-273, 275-278, 280, 287, 290, 292, 293, 297-298, 301-302, 306, 308, 310- 319, 322-325, 328-329, 331, 333-334, 336-337, 339, 344, 354, 360-364, 367, 375- 380, 383-384, 390-391 Fault, battery 375-377 Fault, cable, locating 207 Fault calculations 101, 260-263, 265, 267, 268, 383 Fault current 97-98, 102, 159-161, 171, 180, 218, 224-226, 229, 236, 243, 246, 254-255, 257, 260-264, 266-268, 270-273, 275-277, 287, 292, 306, 310-313, 315, 322-323, 329, 331, 336, 339, 363, 375 Fault-current sources 260 Fault-point impedance 268 Fault-through stress 224 Feed-through receptacle 305, 326-327 Feedback control 349-350, 353 Feeders (electrical) . . 5-7, 9-10, 13, 15-16, 72, 76, 81, 96, 103, 124, 154, 182-185, 187, 190-191, 195, 197-198, 201, 204, 211, 213-215, 260, 270, 287, 289, 320-321, 323, 333- 334, 342, 345, 366-367 Fencing 334-338 Ferroresonance 287 Ferroresonant transformers 371-372, 381 Field-effect transistor (FET) 112 Field emission 226, 234 Field excitation 150, 348 synchronous motors 4, 89, 135, 137, 143-148, 152, 260-261, 268, 297 Field strength, electric 29, 286 Field windings 130-132, 143-146, 148-149, 151, 153 Filter, rectifier 351 Flameproof enclosure 383, 390, 395 Flat-compounded motor 148-149, 151 Flux, magnetic 26-27, 66, 69, 77, 109, 129, 131-133, 148-149, 151, 280 426 Page Page Flux linkage 66, 69 Forced response 48 Forward bias 104-106, 109-111, 257, 346 Forward blocking 346 Frame, motor 133, 148 Frame size, circuit breaker 228-230, 232 Full-load currents 75, 140, 197-199, 247, 270-271, 273, 276, 287, 309-310, 313-314, 322, 364 Full-wave rectifier 106, 108, 117, 132, 364, 370 Fuses 87, 97-98, 180, 224, 226, 244, 248, 254- 257, 259, 266, 277-278, 292, 295, 312, 316, 322, 329, 336, 338, 377, 380, 393 boric acid 238-239 class 235-237 current-limiting 235-238, 284, 356 dual-element 235-236 expulsion 237 high-voltage 235, 237-238, 276 I 2 t 236 low-voltage 235, 237-238, 263, 271, 276 power 237 ratings 235, 239, 260, 266, 270, 306, 322 semiconductor 237 transformer 247, 314, 333, 342 Fusible element 98, 235-238, 306 Galvanometers 123, 125 Gap, explosion-proof enclosure 387 Gaps, spark 292 Gassing, battery 380 Gate, thyristor 113-114, 346, 360 General-purpose fuses 237, 239 General-purpose motor 140 Generator (see Direct current generators, synchronous machines) Ground 9, 13, 15, 18, 38, 79, 87-90, 106, 139, 153, 159-187, 190-194, 201-204, 206, 208-221, 224-226, 233, 236- 237, 247-261, 263, 268, 270-272, 275-276,278- 281, 284, 286-287, 290-301, 304-309, 311- 312, 314-319, 321, 323-325, 328-345, 354, 361-363, 366-367, 375-380, 383, 391-393, 400-402 Ground bed, mesh or electrode . . 124, 160, 162-164, 166, 169-172, 174-175, 177-181, 294, 299- 300, 334, 338-341, 343 Ground-bed measurement 172, 181 Ground-check monitor 164 Ground fault 159, 162, 165, 178-179, 181 Ground-fault current 160-161, 180 Ground insulation 120, 139 Ground protective relay 90, 165, 179, 248-249 Ground resistance 166, 169-170, 172-173 Grounded conductor 159, 220, 257, 323, 393 Grounded system 160, 251, 261, 271, 272, 276, 306 effective 293 noneffective 293 resistance 163, 293, 319, 402 solidly 293 ungrounded 164, 284, 286 Grounding 5, 18, 22, 82, 99, 119-120, 160, 162, 167, 170, 172, 176-177, 181-182, 185, 208, 293-294, 299, 305- 306, 325, 337, 339, 363, 380, 383, 391 conductor 159, 161, 163-166, 171, 179-180, 184, 187, 190- 191, 194, 206, 209-210, 217, 224, 247-253, 255-256, 258, 268, 275-276, 281, 290, 304, 311, 317-319, 323, 329, 330, 334, 338, 340-342, 377-379, 392-393 diode 257, 323 instruments 180 resistor 90, 161, 163-165, 179-180, 224, 247, 249-250, 253, 255, 268, 272, 278, 287, 312, 314-315, 323, 333-334, 340-341, 343-344, 354, 361 system 1-2, 9, 13, 98, 159-161, 165-166, 169, 171- 173, 178-179, 181, 251, 254, 271-272, 281, 300-301, 332, 340-342, 345 transformer 79, 179-180, 308-309, 321, 340, 343, 344 Group motor control 17 Guy wires 218 -H- Half-cell 178 Half-wave rectifier 106, 108, 329, 346-347 Hall effect 126-127 Harmonics 70, 286-287, 337, 346, 348, 354, 403-404 Hazard reduction 383, 394-395 Hazardous atmosphere 382-384 Hazardous locations 382-383, 394-395 Headlight lens, permissible 388 Heat sinks 106-107, 110, 349-350, 355-356, 359, 386 Heaters, strip 229, 329 Heating of soils 175 Henry 26-27 Hertz 46, 124 High field emission 225 High voltage 3, 6, 8, 13, 15-17, 63, 71, 75, 120, 128, 135, 139, 159-160, 165, 179, 181, 185-187, 190-195, 210- 211, 217, 220-222, 226-227, 229, 232-235, 237-239, 248, 250-251, 253-256, 260- 263, 266, 267-268, 271-273, 276-277, 280, 283, 285-287, 290, 293, 298, 304-306, 308, 314, 317, 319, 326, 328-329, 332-333, 336, 339- 340, 342, 344, 362, 365, 368, 390, 395, 402403 High-voltage circuit breakers 13, 232, 239, 263, 268, 273, 306, 329 High-voltage couplers 191, 193-194, 304 High-voltage distribution 217, 234, 250, 256, 266, 272, 286, 298, 326, 365 427 Page Hipot 206-207 History, mine electrical 2 Hoists 129, 142, 152-153, 158, 218-219, 221, 278, 350 Horsepower 2^t, 8, 84, 103, 134-135, 138, 140-142, 149, 152-154, 197-198, 241, 260-261, 264-265, 270, 308, 313, 335, 350-351, 354, 356 Page Hot-spot temperatures 310 Hybrid relay 356, 358, 360 Hysteresis loss 70, 285, 311 -I- I 2 t fuses 234 Ideal sources 42-43, 57, 102 Ideal transformers 66-71, 73-74, 91, 108, 363 Idealization 29 Imaginary numbers 48-49 Imaginary power 59-60 Impedance ... 20, 53-57, 61, 63, 64, 66-68, 71-74, 76, 78-80, 83-86, 91-98, 101-102, 107, 109-113, 115, 118-120, 122- 123, 125, 141, 156-157, 166, 179-180, 186, 200- 202, 224, 241, 246, 252-253, 256, 258, 260- 266, 268, 272-275, 276, 281, 284-289, 294-296, 298-301, 307, 309-310, 312- 313, 315, 317-319, 321-322, 332, 335, 337, 342, 361-363 Impedance angle 67, 85 Impedance diagram 97, 266 Impedance voltage 272, 309, 363 Incendive arcing 362, 384 Inductance ... 20, 26-31, 33, 45^18, 50-51, 53, 63-67, 69-73, 75, 118, 122, 145, 156-157, 170, 174, 176, 230, 246, 280, 282- 288, 290-291, 294-295, 324-325, 371, 383-384 Induction-disk relays 136, 242-243, 249, 270-273, 278, 328-329, 331, 361-362 Induction motors 60, 62, 84, 89, 122, 135, 137, 140-141, 144- 145, 152, 158, 242, 244, 260-261, 263, 265, 267-268, 270-271, 297, 319, 348, 351, 353-356, 363, 365 single-phase 156-157 three-phase 91, 136, 138-139 wound-rotor 142-143, 283, 352, 397 Inductive reactance 47, 54, 63, 139, 146 Infrared 207, 40S404 Instantaneous trips 229-230, 275-276, 312, 362 Institute of Electrical and Electronics Engineers (IEEE) . . 18, 181, 199, 206, 222, 226, 245- 246, 259, 261, 271, 273, 278- 280, 301, 307-308, 325, 334, 346, 365-366, 379, 381, 382, 395 Instruments 75, 89, 114-120, 122-128, 173- 174, 176, 270, 399, 400, 403 Insulated Cable Engineers Association (ICEA) [Insulated Power Cable Engineers Association (IPCEA)] 185-186, 194-201, 204, 206-207 Insulation 71, 119-120, 123, 135, 139, 155, 159, 179, 183, 184-187, 190-196, 200-202, 204, 206-211, 220, 246, 260, 271, 283, 290-295, 297, 299, 303, 306-307, 310-312, 319, 325, 328-329, 336-339, 345, 353-354, 377-379, 382-383, 390, 391- 392, 396, 398-404 Insulation, cable 183, 185-186, 192, 195, 204, 206-207, 209-210, 290, 404 Insulation, conductor 139, 187, 195, 206, 208, 299 Insulation, motors 135, 139 Insulation, transformer 159 Insulation class 135, 195, 246, 297, 310, 329, 337, 382 Insulation coatings 378-379 Insulation coordination 293, 337 Insulation failure 138, 312 Insulation shield 186-187, 190 Insulation systems 135, 139, 186, 192, 290, 294, 306, 400, 401 Insulator 20, 113, 166, 174, 179, 185-186, 192, 194, 216, 218, 221, 233, 284, 312, 332, 334, 344, 367, 390, 398, 403 Integrated circuit 114, 298, 364 Interlocks 226, 254, 305, 360, 377, 379 Intermachine arcing 166, 255, 384 Intermittent duty 153, 198-199, 222 Interpoles, direct current motor 148 Interrupting capacity 18, 224-229, 232-234, 238, 268, 305, 310, 313, 325, 329, 336 Interrupting current 225, 229, 232-234, 238, 268, 305, 310, 313, 325, 329, 336 Interruptors 284 Interwinding faults 298, 311, 378 Intrinsic safety 382, 383, 391, 393, 394-395 Inverse-time characteristic 230, 235, 243-244, 276, 328 Inverters 351-352 Ionization by collision 227, 233-234 IR-discharge voltage 293, 337 Isolation transformers 342, 344 -J- Jackets, cable 183, 185-187, 193-195, 199, 203, 205-206, 208, 210, 393 Jogging 90, 141 Junction 23-24, 36, 38-39, 43, 45, 78, 104-107, 109- 114, 164, 194, 226, 289, 346, 349, 359 Junctions, semiconductor 112 428 -K- Page Kelvin double bridge 123, 398 Kinds of protection 248, 256, 312, 362 Kirchhoffs current law KirchhofFs voltage law Page 23-25, 32, 35, 38-40, 79, 100, 110 22-24, 36-38, 42, 47, 57, 67, 76, 146, 149 -L- Lagging phase angle 47, 146 Laminations, core rotor 130 stator 132, 137-139 transformer 311 Lead-acid batteries 368-370, 374-376, 381 Lead entrances, enclosures 387-389, 391-392 Leading phase angle 47 Leakage current 190, 206-207, 356, 360, 375, 399 Lenz's law 27, 136 Let-go current 161-162, 181 Let-through energy 236 Lightning 159, 163, 166, 175, 256, 280-281, 290, 292, 296, 298-301, 306, 332, 337, 339-342, 344, 379 Lightning protection 166, 218, 301, 337 Line-end failures 295 Liquid-immersed transformers 262, 335, 337-338 Liquids, insulating 307, 337, 345 Load-break switch 226, 235, 239, 256, 305-306, 326- 327, 329, 331, 336, 342-343 Load centers (see Power centers) Load factor 103, 154, 197-199 Loading machine 11-12 Locked-rotor torque 139-140, 354 Locomotives, trolley 164, 287, 367 Long-time delay 230-232, 275 Longwall mining 3, 11-13, 229 Loop 5, 14, 22-23, 36-40, 43, 55, 57, 64-65, 67, 70, 79, 102, 117, 129, 132, 137, 151, 227, 251- 253, 317-318, 330, 349, 352, 375, 379 Loop equations 36-38, 40, 43, 57, 64, 67, 102 Losses 2-3, 6, 8, 66, 79, 108, 140, 149, 206, 284, 289, 308 core 69-70, 72-73, 137, 310 eddy-current 70, 136-137, 311 friction 137, 146 hysteresis 70, 285, 311 I^R 137 transformer 68, 285 windage 146 Low-resistance groundbed 177-178, 339, 341 Low voltage ... 3, 5, 9-10, 15, 135, 154, 165-166, 185, 190-192, 194, 207, 216, 226-228, 232, 235-239, 244, 254-256, 259, 261, 263, 266, 274-275, 278, 287, 298, 305, 311, 317, 325, 342, 344, 348, 362, 378-379, 400, 403 Low-voltage systems 154, 181, 261, 276, 362 -M- Magnetic blowout 227 Magnetic circuit breaker 231 Magnetic circuits 227, 274, 311, 324 Magnetic fields 26-27, 66, 70, 77, 115, 126-127, 129- 130, 133, 136-139, 146, 148, 156- 157, 227-228, 231, 258, 324 Magnetic-flux density 26-27, 66, 69, 77, 109, 129, 131-133, 148-149, 151, 280 Magnetic poles 26, 137-138, 143 Magnetic starters (see Starters) Magnetically coupled circuits 27, 66, 105 Magnetization curves 115, 119, 371 Magnetizing current 69-70, 72, 284-285 Maintenance 2-3, 6, 18, 79, 128, 135, 143, 153, 158, 182, 202, 207, 221-222, 224, 228, 232-234, 261, 285, 307, 316, 325, 328-329, 331-332, 334-336, 344, 350, 352, 354, 358, 361, 366, 368, 370, 375-377, 379-383, 391-392, 396-399, 404 Manual starters, direct current 150 Margin of protection 307, 337-338 Maximum experimental safe gap 386 Maximum power transfer 40, 42-43 Maximum trip setting 270, 274-275, 313, 325, 363 Medium voltage 3, 16, 135, 179, 185, 190, 194, 226-227, 231- 232, 234, 251-252, 254, 256, 261, 268, 271-272, 274, 305, 311, 317, 344, 378 Megohmeter 123-124, Mesh, ground-bed 159, 166 Messenger wire Metal-clad enclosure Metal oxide varister Meter 20, 103, 115-125, 133, 136, 176, 290, 316, 366, 378, 386, ammeter 51, 72, 89-90, 115-117, 119-120, 273-274, 309, 317, 329, 344, d'Arsonval 115-117, 119-120 demand 90, 122, electrostatic megohm moving iron multimeter ohmmeter 22, power-factor var 90, 118, voltmeter 51, 90, 115-117, 119- 127-128, 316, 344, volt-ohm-milliammeter (VOM) watt 72, 90, 115, 117-120, 122. watt hour 90, 122, 127 Methane ignition Microelectronics Mine explosions 181, 206 , 169-171 203-204 326,336 298,356 207, 272, 390,403 123-124, 380,398 ', 124-125 127, 345 .. 115 .. 181 117, 119 117, 125 117, 206 .. 118 , 127-128 120, 125, 398,400 .. 398 , 127-128 128, 136 .. 384 .. 114 .. 367 429 Page Page Mine gases 3, 234, 367, 395 Mine motors 135, 153, 155-156, 222, 401 Mine power centers 1, 5, 200, 279, 302-308, 310-311, 313, 315-317, 319, 320, 325, 331, 335, 345 Mine Safety and Health Administration (MSHA) 395 Mine ventilation 12, 17, 147, 334, 342, 373, 374, 375, 380 Mines shaft 13, 15 strip 4-5, 9-11, 185, 216, 219-220, 232 underground 3-5, 11-13, 15, 19-20, 76, 81, 103, 164, 166, 182, 187, 190, 192, 198-202, 204, 207, 210-211, 223, 228, 234, 238, 254-257, 259-261, 264, 275, 279, 281, 285, 289, 296, 301-303, 308, 313, 319, 326, 333-334, 340, 343, 365-367, 372, 379, 382-383, 389, 394-395 Minimum-oil circuit breaker 233-234 Mining cycle 12 Mining methods, coal area, strip 8-9 contour, strip 8-9, 220 longwall 3, 11, 13, 229 open pit 8, 11, 19, 185, 216, 219, 255, 301, 345 room-and-pillar 11-12 shortwall 13 strip 4-5, 9-11, 185, 216, 219-220, 232 Model, circuit 53, 72 Moisture, soil 171 Molded-case circuit breaker 226-232, 240, 247, 256, 270, 274, 275, 278, 312-313, 319, 324-325, 342, 362, 365 Momentary-current rating 225, 360 Motor 1-4, 9, 13-14, 16-17, 19, 26, 30, 76-79, 85, 92, 103- 104, 113-114, 125, 133-158, 160, 176, 197-198, 200, 222, 226, 231, 233-234, 236, 238, 261, 275-279, 283, 285, 287, 290-291, 298, 303, 308, 313-314, 318, 332, 346, 348-356, 360, 380, 383, 385, 389, 391, 394-404 brushes . . 90, 129, 132, 135, 142-143, 148-150, 350, 352, 366, 399 capacitance 268 commutating-pole 148 compound 148-149, 151, 153, 348 construction 135, 148 enclosures 134 induction 60, 62, 84, 89, 91, 122, 135-143, 152, 156, 158, 242, 244, 260-261, 263, 265, 267-268, 270-271, 297, 319, 348, 351-356, 363, 365, 397 insulation 135-139 mine 135, 153, 155-156, 222, 401 series 2, 4, 135, 148, 151-154, 320, 348 shunt 148-151 single-phase 156-157 squirrel-cage induction 135-136, 139-140, 143, 154, 348, 351-352 synchronous ... 4, 89, 135, 137, 143-148, 152, 260-261, 268, 297 three-phase 3, 84, 135-139, 154, 283 wound-rotor induction 142-143, 283, 352, 397 Motor control (also see Speed Control and Starters) Motor-generator, sets 4, 129, 147, 350 Moving-coil instrument 115-118 Multiple restrike 283-284 Mutual inductance 27, 64-67, 69, 70, 75, 170, 173 Mutual-resistance effect 168-169, 172 -N- National Electrical Code (NEC) 195, 271, 278, 382, 395 National Electrical Manufacturers Association (NEMA) 134, 382, 395 Natural frequency 260, 282-283, 285-286, 287 Negative sequence 98-102, 180, 224, 255, 261, 263, 361, 398 Negative-sequence components 100, 361 Network . . 2, 5, 23, 40-41, 43-44, 48, 55, 101, 166, 267-268, 356, 358 equivalent 33, 34 four-terminal 176 three-terminal 33, 34 two-terminal 33 Network protector 7 Neutral . . 20, 77-79, 82-83, 86, 89-90, 93, 95, 98-102, 104, 108, 125, 147-148, 159-161, 163-166, 171, 177, 179-181, 216, 224- 226, 236, 249-250, 255-256, 258-259, 261, 264, 268, 272, 276, 278, 287, 290-293, 298, 301, 308-309, 312, 314-315, 318, 321, 323-324, 333, 340, 343-344, 346, 354, 362-363 Neutral-shift protection 256, 258-259, 323 Nip 14, 256-257 Node 36, 38^»2, 55, 64-65 Node equations 38, 40 Normal transient 282 Norton's theorem 40, 43, 44, 57 N-p-n transistor 109-113 Nuisance tripping 141, 166, 184, 216, 251, 253, 270-271, 290, 314, 318, 330, 360, 363-364, 367, 380 Null 122-123 -o- Ohm 21-22, 42, 53, 93, 116, 123, 162, 174, 361 Ohmmeter 22, 117, 206 Ohm's law 21, 23-25, 30-31, 35-36, 38-40, 46, 53, 261-262 Oil circuit breaker (OCB) 5, 232, 329 Oil-immersed transformers 294, 337 One-line diagrams 86, 91-92, 95-96, 163, 201, 255, 262, 264, 276-277, 302-303, 333-334, 336, 344 Open circuit 30, 33, 43, 72, 97, 196, 206-207, 215, 253, 289, 317, 368, 374 Open delta 82, 120, 249, 251, 316, 329 Open pit mining 8, 11, 19, 185, 216, 219, 255, 301, 345 Operating mechanism, breaker 90, 228-229, 231, 233-235, 239, 274, 286, 305, 312 430 Page Oscillograph 125, 300 Oscilloscope 125-126 Outby 182, 205, 377, 379 Overburden 4, 8, 11 Overcurrent protection 141, 232, 250, 275, 278, 356, 377, 380 Overcurrent relay, alternating current time 90 Overcurrent relay, direct current 90 Overhead-line distribution 217, 219 Overload 6, 97-98, 107, 119, 134, 141, 154, 171, 224, 226, 229- 237, 248-249, 254-256, 258, 270-271, 273-276, 278, 312-314, 331, 333, 335-336, 340, 354-356, 360-362, 364, 365, 378, 380, 392-394 Page Overload protection 107, 255-256, 271, 273-276, 278, 312, 314, 331, 340, 355, 362, 364- 365, 378, 380, 392, 393 Overtravel, relay 278, 328 Overvoltage ... 7, 64, 88, 90, 119, 159-161, 224, 240, 242, 246-247, 260, 278-280, 283-287, 290, 292, 295, 298, 300, 306-307, 311, 322, 329, 332, 336-337, 354-356, 358, 361, 364, 379, 398 Oxide film, aluminum 174 Ozone resistance 185 -P- Packing gland 192-193, 205, 387-388, 392 Packing-gland lead entrance 387-388 Parallel circuits 24-25 Parallel-ground path 252, 379 Parallel resonance 64 Parallel-series circuits 143, 368 Paralleling reactors 109 Partial-discharge 185-186, 192, 398, 403 Peak inverse voltage 105, 322 Peak load 4, 103, 249, 270 Peak voltage 132, 282, 286, 295-2%, 306 Percent quantities 93 Percent ratio error 242, 247, 272-274 Permanent splice 207 Permissible equipment 382, 387, 391-393 approval 382, 391 explosion-proof 382-383, 387, 391 intrinsic safety 382, 393 mobile equipment 387, 392 Per-phase reduction 84-85, 90 Per-unit quantities 93, 309 Phase angle 46-47, 51, 52, 54, 59-61, 90, 92, 146, 157, 244, 247, 249, 363, 399 balance 83, 88 current 79-80, 83-84, 92, 101, 108, 120, 250 protection 248, 256 sequence 90, 120, 122, 124, 137 Phase-sensitive short-circuit protection 362-365 Phase-sequence indicator 122-124 Phasor 51-54, 56, 61, 77, 91, 97-98, 100, 146, 198, 201, 287 Pilot wire 88, 90, 318 interlocks 306 monitoring 187, 251-254, 305-306, 317, 319, 378-379 Plants, power . . 8, 17-19, 129, 334, 340, 342, 382, 389, 394-395, 397 Plugging 141, 191, 350 Plugs and receptacles (see Connectors) P-n junctions 104-107, 110, 112-114, 346, 349 P-n-p transistor 109-110, 112-114 Polar relay 242 Polarity of windings 65 Polarizing diode 257-258 Poles, motors nonsalient 130 salient 130, 136, 138, 143, 244 Polychlorinated biphenyls 307, 329 Polyvinylchloride (PVQ insulation 390 Portable cables (see Cables) Portable substation 3, 5, 19, 166, 303, 332, 342 Positive sequence 77, 98-100, 101-102, 255-284, 261, 263 Positive-sequence components 100 Potential difference 20, 23, 29, 115, 161, 171, 280, 290, 377 Potential gradient 163, 166, 169-172, 175, 178, 287, 300, 339 Potential relaying . . . 246-247, 249-250, 255, 278, 315, 333, 340, 344 Potential transformers . . 75, 86, 118, 120, 126, 237, 243, 246, 248 Pothead 87, 191-192 Power apparent 59-60, 62-63, 66, 68, 74, 81, 83, 93, 103, 118, 246, 319 average 21, 50, 59-60, 63, 72-73, 80, 81, 83 complex 59, 60-62, 68, 73-74, 80-81, 83, 84, 103, 117, 119-122, 125, 197, 319 imaginary 59-60 reactive 59-60, 62-63, 81-82, 97, 118, 319 real 59-60, 62 three-phase ... 59, 76, 79-80, 81-82, 98, 103, 144, 154, 157, 281 Power centers 1, 3-5, 8, 11-13, 16, 19-20, 33-34, 36, 44, 48, 65, 76, 81, 94-95, 193, 199-201, 228- 229, 279, 295-296, 310-317, 335, 344 alternating current 255, 280, 302-307, 310, 313, 315- 317, 319-320, 325, 331, 345 alternating current-direct current 320, 325 breakers 257 bus 312, 319 couplers 5, 191, 254, 304, 317, 379 disconnect switch 5, 13, 226, 254, 256, 305- 306, 326-327, 336, 338 fuses 254-257, 278, 295, 306, 312, 314, 316, 322, 329, 332-333, 336, 377-380 grounding ... 1, 5, 13, 164-165, 190, 252, 255-257, 308, 311-312, 314-315, 317, 319, 321, 325, 329, 333, 344, 379 instruments 11, 75, 173-174, 176, 270 surge arresters 295-296, 306-307, 310, 332, 334, 337-338 transformers 200, 308, 311 Power circuit breaker 75, 226-228, 232, 274-275, 329, 331 Power factor ... 59, 61-63, 80-82, 84-85, 90-91, 117-120, 127-128, 138-139, 146-147, 154, 197-198, 201-202, 237, 246, 270, 282, 313, 319-320, 363, 399, 403 correction 283, 332 Power rectifier 90, 93, 107 Power systems (also see Distribution) Power transformers 66, 70-71, 73, 75, 82, 86, 179-180, 238, 276, 297, 307, 308, 310, 325, 337, 340-341, 344, 377, 378-379 431 Page Page Power zeners 298 Prefault voltage 262, 264 Preparation plants 8, 17-19, 129, 334, 340, 342, 382, 389, 394-395, 397 Pressure relay 337 Pressure-relief devices 388 Prestrike transient 286, 295-296 Primaiy-selective distribution 9 Propagation velocity 288 Protective relaying 75, 104, 118, 125, 240, 244-246, 248-249, 254-256, 259, 273, 278, 312-313, 326, 328, 331-333, 336, 340, 343, 345- 346, 352, 356, 359, 362, 365-366 Protective relays 90, 119, 160, 248, 259, 272, 344, 356, 365-366 Pull-in torque 145 Pull-out torque 146 -Q- Quality factor, a 64 Quick-break, quick-make mechanism 226, 229 -R- Radial system 5-6, 10, 326, 333-334 Rail 2-5, 12-14, 164-166, 204, 211, 215-216, 220, 252, 256-257, 303, 320, 367, 392, 397 Rail bond 165, 211, 215-216, 252 Ratchetting, relay 360 Rating 82, 86, 102-103, 105, 107, 118-119, 134, 140-141, 149, 153-154, 179-180, 187, 207, 222, 239, 241, 254, 260-266, 272-279, 290, 292, 305-310, 312-317, 321-322, 325, 331, 337, 349-350, 362 cable 159, 185, 195, 199 circuit breaker 226-229, 231-233, 270, 292, 313-314, 329 grounding resistor . . 179, 224, 247, 255, 272, 278, 312, 314-315 motor 135, 142, 261 switching apparatus 8, 224-226, 240, 255, 257, 264, 267-268, 284, 295-296, 335 transformer (see Transformer rating) voltage 71, 135, 144, 185, 200, 232, 236, 239, 246, 262, 292-293, 295-296, 305-307, 321-322, 329, 355-356, 379, 401 Ratio error 246, 272-274 Reactance 47, 54-57, 59, 63-64, 68, 71, 75, 93-94, 96, 118, 140, 145-146, 160, 200-201, 261-265, 308, 310, 321, 335 capacitive 54, 297 inductive 267, 271 leakage 69 of cables 265 of conductors 139 of motors 266, 267, 270-271 Reactive power 59-60, 62-63, 81-82, 97, 118, 139 Reactors . . 87, 109, 114, 252, 258-259, 310, 321-322, 324, 370, 372 Real numbers 4&49 Real power 59-60, 62 Receiver 76, 90, 318 Reciprocity 4fr42 Recloser 335-336 Recorders 124-126 Recovery voltage 282, 283, 295 Rectifiers 13, 90, 104-111, 117, 128, 164-166, 211, 228, 232, 256, 258, 290, 307, 310, 320-325, 364, 397 battery charging 367, 371-373, 377, 379, 380-381 control circuitry 75, 141, 151, 154, 316, 324, 326- 327, 329, 330, 331, 349, 370 full-wave 370 bridge 143, 370 three-phase 108, 348 half-wave 106, 108, 329, 346-347 mercury arc 370 mine 3, 323 overloads and faults 97-98, 224 ratings 308 silicon 108-109, 147, 370 silicon-controlled 113, 346 three-phase 107-108 thyristor 237, 346, 348 transformers 105-108, 321 Reduced-voltage starters 141, 352, 365 Reed relays 358 Reels, cable 9, 187, 204, 222, 257, 367, 393 Reference node 38-40 Reference phasor 51, 54, 56, 60-61, 68, 73, 97 Reflected wave 288-289, 294 Refracted wave 288-289 Regulation, voltage 1, 3, 7, 18, 73-74, 76, 85, 197, 199, 200, 298, 310, 319 Relays 14, 75, 86, 89, 104, 118-119, 125, 143-145, 150- 151, 157, 159-161, 180-181, 224, 226, 230- 231, 263, 274-279, 287, 290, 301, 306, 311, 313-316, 323-324, 326, 334, 336- 337, 340-341, 343-346, 353-354, 366, 370, 376, 379-380, 396, 398, 402 alternating current time overcurrent ... 90, 166, 243, 245, 247, 249-250, 256, 270-271, 328-330, 334, 358, 360-362, 364-366 clapper 241-243, 356 contacts 358, 363 cylinder 242, 245 direct current ground-fault 259, 323 direct current overcurrent 90, 247, 256, 361 differential 244-247 directional 88, 90, 240, 242, 244 electromechanical 240-241 ground protective 90, 165, 179, 248-249 hybrid 356, 358, 360 induction disk 136, 242-243, 249, 270-273, 278, 328-329, 331, 361-362 432 Page Page instantaneous 241, 243, 255, 328, 364, 368 overcurrent 240, 243, 248 overtravel 278, 328 residual 249-251 solid-state 240, 356, 358-360, 361-362, 364-366 tap settings 243-244, 249, 271-273, 331, 361 terminology 240 thermal 240-241 time-delay 248 undervoltage 240-248 voltage 240, 246 Reliability 5-6, 12-13, 18-19, 135, 153, 221, 251, 307- 308, 312, 332, 334, 354, 356, 358, 361-362, 365, 368, 377, 396 Reluctance 69, 311 Residual relaying 249-251 Resins 135, 139, 221, 361, 390 Resistance 2, 20-21, 23, 24, 26, 32-36, 3945, 50-57, 59, 61, 64, 66, 83, 93, 98, 101-102, 105-108, 111-112, 116, 132, 138, 141-143, 146, 157, 160-162, 166, 167-169, 176, 178-181, 190, 192, 200, 211, 213, 215-216, 224, 241, 246-249, 252, 254-256, 260, 262, 264-270, 273- 276, 283-298, 306-308, 310-315, 317- 321, 338-346, 352, 356, 359-365, 368-370, 375, 378-379, 382, 394 alternating current 46-47 cable 123, 164, 207 conductor 22, 31, 70, 123, 139-140, 159, 195 direct current 184 ground-bed 124, 163, 170, 172, 174, 177, 300, 339 insulation 123, 185, 207, 396, 399, 400, 402-403 parallel 25, 30, 43, 117 series 29-31, 117, 123, 125 winding 68-73, 148-153 Resistivity 166-171, 174-187, 190, 339, 394 earth 178, 180-181, 301, 341 material 22 Resistor 22, 24-25, 29-30, 33, 3741, 46-48, 54, 89, 112, 114-118, 123, 126, 141-142, 145, 150, 160, 171, 246, 253-256, 258-260, 292, 309, 311-318, 323, 332-334, 338, 352, 358, 363, 370, 398 current-limiting 165 grounding 90, 161, 163-165, 179-180, 224, 247, 249- 250, 253, 255, 268, 272, 278, 287, 297, 312, 314-315, 323, 333-334, 340- 341, 343-344, 354, 361 starting 151 Resonance 287, 319 parallel 64 series 64 Response, frequency 361 Restrike 98, 226, 282-285, 287 Retentivity 70 Reverse bias 105-106, 110, 112, 114, 322, 346 Reversing controls, motor 150 Rheostats 90 Ripple voltage 108-109, 132 Rod, ground bed 168 Roof bolter (drill) 81, 153, 190 Room-and-pillar mining 11-12 Root-mean-square (rms) 50-51, 59 Rotating lines 51 Rotating magnetic field 136-137, 146, 157 Rotor 130-133, 135-138, 141, 142, 156-157, 244, 260, 353-354, 383, 391 bars 136, 139-140, 142, 404 construction 140 cylindrical 143-145 wound 142-144, 283, 352, 397 Rules, ground bed 178 -s- Sacrificial anode 178 Safety factor, borehole cables 198, 203, 261-262, 268, 275, 278, 328, 367, 373-374, 380, 386-387, 393 Safety grounding 160, 163-164, 166, 170, 178- 181, 334, 338-341, 343, 377 Sag, overhead line 217 Salient pole 130, 136, 138, 143, 244 Saturable reactor 252, 258-259, 324, 370-372 Saturable transformer 258 Saturation current 104-105, 110 Saturation curves 273 Schedule 142, 194, 207, 302, 367, 370, 381, 391-392 Schedule 2G 259, 374 Secondary-selective system 6, 334 Secondary-spot network 5, 7 Sectionalizing unit 326 Segment 3, 13-14, 17, 20, 76, 93-94, 113, 132, 147-148, 150, 153-154, 201, 211, 254, 276, 284, 288, 296, 320 Selective relaying 159, 224 Selective -system operation 18, 224 Selector device 90 Selenium suppressor 322 Self-inductance 27, 65-66 Semiconductor devices 104, 114, 356 Semiconductor fuses 237 Semiconductor shields 210 Sensitive earth-linkage system 362 Sensors 207, 221, 354, 356-357 Series circuits 22-25, 54, 64, 76, 94, 124 Series motors 4, 151 Service factor 134 Settings, maximum instantaneous 275 Shaft mines 13, 15 Shell, coupler 319 Shield 13, 98, 183, 189, 194-196, 200-201, 206, 210, 224, 231, 281, 292, 299, 325, 342, 363, 383 conductor 186-187 electrostatic 311 Faraday 298, 311, 379 insulation 186-187, 190 nonmetallic 186 overhead lines 298 433 Page Page transformer winding 67, 74-75, 95, 119, 246, 308-310, 321, 336, 379 Shielded cable 9, 11, 186-187, 189-191, 199, 204, 210, 219, 260, 319 Shielding failure 281 Shock, electric 160, 161-162, 180-182, 186, 222, 316- 317, 339, 366, 367, 377, 380 Short circuit . . 23, 33, 40, 42-44, 57, 73, 90, 95, 97-98, 107, 146, 180, 184-185, 197, 207, 224, 228-229, 232-236, 239, 248-249, 254-256, 261, 265, 270, 274-278, 286, 289, 298, 306, 309-312, 328, 333, 336-339, 361-364, 367, 389-390, 395 Short-circuit currents 44, 73, 202, 225, 229, 232, 236-237, 260-264, 270, 275, 277, 310, 313, 321, 328, 331, 336, 344 calculations 278 Short-circuit protection 231, 236-237, 249, 256, 270-271, 273- 276, 279, 312, 314, 319, 331, 340, 353, 356, 362-363, 365-366, 378, 380, 392 Short-time ratings 336 Shorting switch 317 Shovels 4-5, 8, 10, 85, 147, 158, 166, 182, 190, 270-271, 281, 351-352 Shunt 4, 42, 90, 106, 116-118, 125-126, 132, 148-151, 231- 234, 240, 242, 247-248, 256, 292, 296, 298, 306, 313, 315, 324, 348, 362, 371, 376, 398-399 Shunt field 132, 150-151 Shunt motor 148-151 Shunt trip 231-234, 248, 298, 306, 313, 315, 362 Shuttle car 2-3, 12, 15, 29-30, 81, 96-97, 151, 153, 165, 183, 185, 187, 190, 196, 198, 204-206, 208-209, 222, 308, 320, 350-351, 367, 391 Silicon-controlled rectifier (SCR) (thyristor) 113, 346 Silicon diode 3, 108, 165, 287, 370 Silver-sand fuse 238 Single-cage rotor 140 Single-ended substation 333 Single-line diagram 86, 256 Single-phase analysis 84 Single-phase motors 156-157 Slip . . . 129, 136-140, 144, 153-154, 156, 209-210, 348, 352, 387-388 Slip rings 90, 129, 131, 135-136, 142-143, 257, 352 Snubber 298, 323, 356 Soil heating 170-171, 180 Soil resistivity 166-169, 171, 174-175 treatment 177 Solenoids 27, 141, 231, 234, 241-242, 312-313, 356 Solid-material fuse 238 Solid-state belt starters 319, 352-353 Solid-state device . . 104, 107, 125, 232, 237, 346, 356-360, 376, 379 Solid-state motor control 356 Solid-state trip elements 312-313, 365 Solidly grounded system 276 Source transformation 40, 42-45, 57 Source transformer 249, 278, 343-344, 363 Sources 19, 23-24, 39-43, 45, 48, 55, 69, 76-79, 84, 97, 100, 110, 159-160, 185, 225, 242, 247, 260- 263, 270, 280, 295, 302, 343, 348, 383, 394 Spacing, overhead lines 217 Span, overhead lines 217-221 Spark gap 98, 292, 298, 363 Sparkover voltage 292-296, 306-307, 337-338 Specifications 2, 93, 135, 158, 185-186, 194-195, 197-198, 200, 204, 208, 211-212, 229, 246, 280, 302-303, 307, 314, 335, 370, 381, 383, 387-391 Speed control, motor chopper 349-350 compound-motor 148-149, 151, 153, 348 induction motor 60, 62, 84, 89, 91, 122, 135-143, 152, 156, 158, 242, 244, 260-261, 263, 265, 267-268, 270-271, 297, 319, 348, 351-356, 363, 365, 397 series-motor 2, 4, 135, 148, 151-154, 320, 348 shunt-motor 148-151 solid-state 356 Ward-Leonard system 152-153, 350-351 wound-rotor motor 142-143, 283, 352 Speed-torque curves 133 Splices, cable 186, 195, 207-210 Squirrel-cage motor 139-140, 154 single-phase 135, 348 three-phase 135-136, 143, 351 Squirrel-cage rotor 145, 156, 404 Standard burdens, current transformers 119 Starters, alternating current 16, 113, 125, 145, 150, 238-239, 298, 351, 356, 366, 402 across-the-line 141, 149 autotransformer 74, 114, 141, 159 belt conveyor 319, 353-355, 360, 365 magnetic 87, 142 reduced voltage 141, 352, 365 solid-state 319, 352-353 synchronous motor 4, 89, 135, 137, 143-148, 152, 260-261, 268, 297 wound rotor 142-143, 283, 352, 397 Starting torque 77, 133, 135, 140, 144-145, 149, 151, 157, 354 Static device 358-362 Static lines 341-342 Station arrester 293-294 Stator 130-132, 135-146, 156-157, 244, 260 Steady-state response 48 Storage battery 42, 367-368, 373, 381 Stranding conductor 185, 404 Stray ground current 367 Stress cone 191, 194 Strip-chart recorders 124-126 Strip heaters 329 Strip mining 4-5, 9-11, 185, 216, 219-220, 232 Stuffing box 387-388 Substation ... 1, 3, 11, 13, 15, 17, 19, 76, 83, 91-92, 96, 179, 191, 199-201, 217, 220, 233, 237-238, 256, 265, 281, 293-294, 299, 302-303, 319, 326, 337, 343 area 339-341, 344-345 bus 255 capacity 7-8, 334 circuit breakers 10, 333 connections 328 design of 332, 342 double-ended 334 economics 8 fuses 333 grounding 9, 166, 180-181, 249, 338, 342, 345 loads 335, 342 location 8, 344 portable 5 relaying 249, 336 switches 5, 336 transformers 128, 163, 307, 331, 334-336, 342 434 Page Page unit 307 Substitution 34, 40, 43 Subtransient reactance 261-268, 270-271 Subtransmission 8-9, 13, 342 Superposition 40-41 Suppressor, transient 298, 322 Surface leakage, battery 375-379 Surface mining 2, 4-5, 7-11, 13, 19, 104, 159, 166, 182-183, 185, 190-191, 194, 199, 204, 206-207, 210- 211, 216, 218-220, 237, 255, 260-261, 270, 275, 279, 281, 289, 296, 303, 319, 326, 343-344, 382, 394 Surface oxide film, aluminum 192 Surge arrester 22, 87, 163, 292-2%, 298-299, 301, 306- 307, 310, 327, 332, 334, 337-338, 3% Surge capacitor 260, 268, 282, 295-297, 307, 332 Surge impedance 166, 285, 288-289, 295, 299 Susceptance 47, 55, 64 Switch 9-10, 87, 89-90, 105, 116, 120, 121, 144, 147, 151, 157, 191, 215-216, 224, 231, 235, 239-240, 253, 257, 280-281, 285, 291, 316-319, 329, 331- 332, 342-344, 348-349, 370, 376-380 disconnect 5, 13, 220, 226, 237, 254, 256, 305-306, 326-327, 336, 338 interlock 150, 305-306, 378-379 interrupter 226 load-break 226 push-button 330, 356, 360, 392-393 ratings 260 Switches, fusible 235, 237, 254, 306 Switchgear 14, 159, 261, 267, 307 Switchhouses ... 1, 5, 10, 13, 15, 128, 166, 182-183, 193, 201, 219, 224, 255-256, 263, 296, 302, 305, 335, 345 breakers 9, 326, 329 connections 285, 295 couplers 191 design of 331 relaying 278, 326, 328, 334 Switching apparatus 8, 224-226, 240, 255, 257, 264, 267-268, 284, 295-296, 335 Switching skid 326 Switching transients 234, 281-282, 286-287, 295 Switchyard 8-9 Symbols 21-24, 27-28, 30, 38, 47-48, 55, 66, 74, 82, 86-88, 98-99, 104-105, 109, 112-115, 240, 343, 346 Symmetrical components . . 98-103, 179, 250, 261-262, 266, 268, 404 Symmetrical currents 263 Synchronous generators 131, 260 Synchronous motors 4, 89, 135, 137, 143-148, 152, 260-261, 268, 297 Synchronous reactance 261, 265 Synchronous speed 137, 139-140, 145-146, 153, 156-157 System 1-22, 30, 33-34, 36, 48, 50, 62-63, 65, 71-73, 75-86, 90-103, 105, 118-120, 129, 131, 135-136, 139-140, 147, 152-154, 157-173, 175, 178-182, 186-188, 190-192, 194-195, 197, 199-201, 203-206, 210- 211, 216-266, 275-290, 292-302, 304-308, 310-313, 316-318, 321-326, 329, 331-345, 355-359, 362-367, 370, 372, 375-377, 379-383, 386, 394-398, 400-401, 404 alternating current-direct current 164 control 4, 349, 352-353, 354 diode-grounded 164, 187, 257-258 expanded radial 5 grounding 160, 165, 378 high-voltage 160, 165, 239, 251, 253, 268, 271, 293, 298 load 98, 215, 319 low-voltage 154, 181, 261, 276, 362 medium-voltage 179, 227, 231, 268, 271-272 power (see Power systems) primary-loop 6 primary-selective 6, 9 reliability 5, 356 resistance-grounded ... 179, 276, 296, 317, 321, 340, 362, 365 secondary-selective 6, 334 simple radial 9 solidly grounded 276 trolley 5, 14, 154, 164-166, 182, 211, 215, 228, 256-258, 270, 323-324, 350 ungrounded 160, 272, 276, 287, 293 -T- Tachometer 353-354 Tap changers 333-334, 337 Tap settings, relay 243-244, 249, 271-273, 331, 361 Tape recorders 125-126 Temperature 68, 90, 105-107, 110-112, 134-135, 153, 155-156, 174, 177, 180, 183-184, 194, 196-199, 201, 206- 208, 217, 222, 226, 230, 240-241, 246, 271, 274, 307, 310, 312, 349, 354, 358-359, 369, 370, 374-375, 383-384, 386-387, 393-395, 398-400, 402, 404 coefficient of resistance 22, 175 effect on ampacity 195 ground bed 171 limits 235 permissible enclosure 389 ratings 185, 195 Temperature rise 68, 134-135, 153, 155-156, 171, 195, 222, 230, 235, 271, 307, 310, 312, 383 Temporary splice 207, 393 Testing 123-124, 222, 246, 314, 329-331, 360-362, 375, 387-388, 391, 395, 399-403 cable 206-207 system 316 transformer 316 Thermal conductivity 171 Thermal device 90, 241 Thermal-magnetic breaker 229, 313 Thermal overload 271, 337, 354, 356 Thermionic emission 226, 234 Thevenin's theorem 43-44, 57 435 Page Page Three-phase 3, 57, 59, 76-85, 90-93, 97-103, 107-108, 120-121, 124-125, 131, 135-141, 143-144, 154, 156-157, 164, 179, 216, 225, 229, 239, 246, 248, 250, 255-256, 260-261, 263-264, 266, 268, 270, 274-275, 278, 281, 283-284, 294-296, 306, 308-309, 316, 319-320, 322, 328-329, 335-336, 346, 348, 350-352, 363, 370, 372, 377 Three-phase generators 77, 131, 135 Three-phase rectifier 107-108 Three-phase systems 76-81, 83, 85, 99, 131, 136, 239, 246, 294, 296 Thumper 207 Thyristor 113-114, 237, 287, 298, 323, 346-360, 370-372 Time-current curves 229, 236, 271, 360 Time-delay relay 242, 248-249, 255-256, 263, 266, 278 Time dial setting 244, 270, 278 Time-domain reflectometer (TDR) 207 Time rating 140, 179, 312 Tinned conductors 200 Tolerance 194, 197, 199, 331, 360, 374, 399 breaker settings 270, 274-275, 278-279 voltage 227-228, 232, 248 Torque 77, 115, 117-118, 122, 124, 133-135, 137-138, 139-157, 242, 244, 348, 352-354, 361 Torque-speed characteristics direct current motors 147 inductor motors ... 60, 62, 84, 89, 91, 122, 135-145, 152, 156, 158, 242, 244, 260-261, 263-265, 267- 268, 270-271, 297, 319, 348, 351-356, 363, 365, 397 synchronous motors 4, 89, 135, 137, 143-148, 152, 260-261, 268, 297 wound-rotor motors 142-144, 283, 352, 397 Touch-and-step potentials 171, 338-339 Track ... 8, 154-155, 164, 166, 211, 215-216, 220, 303, 367, 376, 390 grounding 165 resistance 192 Track bonds 165, 216, 367 Traction locomotives 154, 155, 156 Trailing cables (see Cables, portable) Transducers 115, 119, 126, 224, 240 Transfer functions 349 Transformation . . 3, 13, 33-34, 40, 42^15, 55, 57, 83, 92, 164, 332 Transformers ... 27, 59, 64, 78, 81-83, 85-87, 89-90, 94-96, 99, 103, 111, 129, 139, 158-159, 163-164, 166, 183, 200, 224, 228, 237-238, 240-241, 243-251, 254-256, 259, 264, 265, 271-273, 276-278, 293-298, 312-316, 318-328, 331-338, 340-342, 350- 354, 361, 370-372, 377-381, 402-404 auto 74, 141, 159 air-cooled 134 basic impulse insulation level (BIL) 290, 366 burden 119-120, 328 capacitance 296 capacity 68, 81, 103, 295, 306, 316, 334, 356 class 307 construction 130, 311 control 306, 315-316, 324, 326, 329, 343 cooling 334 core 70 core loss 69-70, 72, 73, 137 current . . 83, 89, 108, 118, 243-244, 246, 249, 272-273, 328-329 delta-delta 83, 92, 261, 308, 343-344 delta-wye 83, 92, 261, 308, 343-344 design 69, 173, 310 distribution 5, 7, 13, 75, 262, 297, 303, 309, 398 dry-type 2, 93, 301, 337, 379 eddy-current loss 70, 136-137, 311 efficiency 146 electrostatic shielding 311 equivalent circuits ... 40, 57, 68, 70-71, 73, 76, 85, 94-95, 113- 114, 260-261, 266, 284-285, 291 exciting current 69-73, 94-95, 274, 295, 308 Faraday shielding 298, 311, 379 fault current 310 ferroresonant 371-372, 381 forced-cooled 334 frequency 72-73 grounding 79, 179-180, 308-309, 321, 340, 343-344 high-voltage 305, 319, 334 ideal 66-71, 73-74, 91, 108, 363 impedance 94 input impedance 66-67, 110, 112, 125 inrush current . . . 150, 244, 270, 276, 285-286, 295-296, 306, 337 instrument 118-119, 270 insulation 159 interwinding faults 298, 311, 378 isolation 342, 344 kilovoltampere rating 261, 307-308, 310 leakage reactance 68-69, 71, 75, 310 liquid-immersed 262, 335, 337-338 magnetizing current 69-70, 72, 284-285 oil 294, 337 open-delta 82, 120, 249, 251, 316, 329 performance 68 potential 75, 86, 118, 120, 126, 237, 243, 246, 248 power 200, 308, 311 protection 118, 271, 301, 334, 336-337, 344 pulse 354 ratings 246, 308-309 rectifier 105-108, 321 regulating 301, 345 resistance loss 118 saturable 258 saturation 108 single-phase 67, 82-83, 180, 316, 335 source 249, 278, 343-344, 363 substation 128, 163, 307, 331, 334-336, 342 taps 308 temperature 310 three-phase 82, 92, 179, 295, 308-309, 336, 350, 377 three-winding 95 transient voltages 280, 282, 283-287, 291-292, 295-296, 298, 301 turns ratio 66-68, 70, 72, 75, 83, 91-92, 95, 196, 262, 270, 272-273, 291, 314, 317, 329 unit substation 307 utilization factor 107 voltage 74, 91-92, 247, 251, 262, 295, 305, 308-309, 313, 316, 321, 363, 370 windings 67, 74-75, 95, 119, 2%, 308-310, 321, 336, 379 wye-delta 85, 90, 92, 179 wye-wye 90-92, 308 zig-zag 79, 89, 179-180, 308-309, 321 436 Page Page Transient 20, 48, 125, 159-160, 186, 192, 197, 224, 234, 252, 261, 265, 278, 298, 300-301, 306-307, 311, 322, 329, 336-337, 350, 354-356, 360, 379, 395 characteristics 282 protection 292, 295-296, 332 sources 280 Transient reactance 261 Transient response 48 Transient suppressors 298 Transistors 109-114, 298, 356-359 Transmission 2-3, 7-8, 65, 76, 127, 162, 181, 216, 218, 222- 223, 259, 282, 287, 293, 301, 384, 390, 395 Traveling waves 286-288, 289-290, 294, 296, 342 Triac 356-360 Trip-free relay 90, 229 Trip settings 270, 274-275, 303, 313, 325, 363 Tripping, circuit breaker ... 75, 184, 229, 231, 235, 240, 248, 251, 253-254, 255, 298, 305, 315, 318- 319, 323-324, 330, 336, 354, 356, 359, 367 Tripping, element . . . 228-231, 240, 248, 274, 310, 312-313, 362, 365 Trolley rectifiers 232, 322, 325 Trolley systems 5, 14, 154, 164-166, 182, 211, 215, 228, 256-258, 270, 323-324, 350 Trolley wire 2, 13, 164, 211, 214-215, 256, 270, 278 insulators 166, 367 Two-layer earth structures 176, 181 -u- Unbalanced loads 10, 79-80, 82, 97 Unbalanced system 100-102 Underground coal mining ... 4-5, 15, 19-20, 76, 81, 103, 164, 166, 187, 190, 192, 198-202, 204, 207, 210-211, 223, 228, 234, 238, 254-257, 259, 261, 264, 275, 279, 281, 285, 289, 296, 301-302, 313, 319, 326, 333-334, 340, 343, 365- 366, 372, 379, 382-383, 389, 394, 395, 397 advance versus retreat 182, 303 continuous 3, 12-13, 16, 201, 308 conventional 11-12, 367 longwall 3, 11-13, 229 room-and-pillar 11-13 shortwall 13 Undervoltage release 231, 256, 306, 312, 316, 318, 330, 342 Underwriter's Laboratories 228, 259, 382, 387, 389, 394-395 Undergrounded systems 160, 272, 276, 287, 293 Unidirectional thyristor control 348, 354 Uniform design 331 Unit substation 5, 10-11, 200, 249, 307, 332, 341-342 Units 3, 5, 14, 17, 20-22, 27, 46, 50, 55, 60, 82, 103, 107, 114, 116, 133, 166, 182, 218, 220-221, 226, 228-230, 232- 233, 236-237, 241, 243, 246-249, 256, 274-275, 292, 295, 302-303, 307-308, 314, 318, 326, 328, 335 Unloaded power factor 399, 403 Utility, electric system 7 Utilization 4-6, 9-10, 14, 96, 107, 153-154, 158, 180, 182, 185, 197, 199, 224, 263, 277, 291, 293, 295, 303, 307- 308, 310-313, 316-317, 319-320, 355, 362, 366 -V- Vacuum circuit breaker (VCB) Var 232, 234-235, 284-285, 295- 296, 329, 332, 343 90 Varmeter 90, 118, 127-128 Velocity propagation 288 Ventilation 129, 381-382, 395, 397 battery boxes 373-374, 375-376, 378, 392 charging stations 373, 375, 377, 380 mine 12, 17, 147, 334, 342, 373-375, 380 Visible disconnects 326 Volt 5, 20, 23, 25-26, 36, 40, 42, 73, 90, 93, 95, 279 Voltage . . . 2-7, 9-11, 13, 17-48, 50-57, 59-61, 63-68, 70-86, 90-102, 104-112, 114-122, 124-132, 134-137, 139, 141, 145- 146, 149, 150, 152-154, 156, 159-162, 165-166, 171, 175, 178-181, 185-187, 190-192, 194-195, 197, 199-202, 207, 210-211, 216-217, 220- 222, 224-229, 231-232, 234-236, 238-241, 243-244, 246-248, 250-258, 260-265, 268, 270, 272-275, 277, 280-301, 303, 305- 317, 319-325, 329-330, 332, 335-339, 342-344, 346, 348-361, 363-365, 368- 373, 378-380, 395, 398-401, 403404 arc 227, 286, 306, 324, 338 cable ratings 159, 185, 195, 199 control 346, 349 drop calculation 199 drop maximum 199 gradients 175, 178, 180 rating 71, 135, 185, 194, 200, 232, 236, 239, 246, 262, 292-293, 295-2%, 305-307, 321, 322, 329, 355, 356, 379, 401 regulation 3, 7, 18, 19, 73, 74, 76, 88, 197, 199, 200, 298, 310, 319 relay 240, 246 ripple 108-109, 132 standard 216, 246 three-phase 77 transformer ratings 246, 265, 308-309 Voltage classes 216 Voltage gradient 175, 178, 180, 280, 339 Voltampere 60, 66, 73-74, 93, 107, 119, 247, 274 Voltmeter 51, 90, 115-117, 119-120, 125, 127, 128, 316, 344, 398, 400 Vulcanization 184-185, 203, 207, 210 437 -w- Page Page Ward-Leonard system 152-153, 350-351 Watt 60, 106, 118, 137 Watthourmeter 90, 122, 127-128, 136 Wattmeter 72, 90, 115, 117-120, 122, 127-128 Wave sloping 295-296 Waves, traveling 286-290, 294, 296, 342 Wenner array 176, 178, 181 Wheatstone bridge 122 Windage 137, 146 Windings 26-27, 86, 94, 108-109, 118, 152-153, 156- 157, 159-160, 180, 240-241, 249-250, 258, 271-272, 298, 317, 322, 335-337, 341, 352, 354, 363, 370-371, 377- 378, 383, 398, 400, 402 armature 130-132, 136, 146, 148-149, 151 compensating 148 response to transients 290-291 transformer ... 67, 74-75, 95, 119, 246, 308-310, 321, 336, 379 Windows and lenses 388 Wire (see Conductor) Wooden poles 216-218, 281 Wound-rotor motor 142-143, 283, 352, 397 Wye connections ... 34, 77, 79-87, 93, 100, 138, 141, 179, 249, 258, 261, 296, 308-309, 321, 329, 336, 343, 363 Wye-delta transformations . . . 33-36, 55, 77-87, 90, 92-93, 100, 108, 120, 138, 141, 179-180, 249, 261, 308-309, 321, 336, 344, 363 -Y- YBUS load-flow analysis 268 -Z- ZBUS fault analysis 268 Zener diodes 105, 298 Zero-sequence relaying 249-251, 255-257, 272, 314-315, 329, 340, 343, 362, 365 Zig-zag transformers 79, 89, 179-180, 308-309, 321 Zones of protection 254, 259 * U.S. GOVERNMENT PRINTING OFFICE : 1991 - 287-709