Industrial Power Systems Handbook [First ed.] 9780070043015

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Industrial Power Systems Handbook D O N A L D BEEMAN, Editor Manager, Industriaf P w e r Engineering Industrial Engineering Seclwn General Electric Company, Schenectady, New Yorlc

FIRST EDITION

McGRAW-HILL BOOK COMPANY, INC.

1955

New York

Toronto

London

Ch.UPh?r 1

by Donald Beeman, Alan Graeme Darling, and

R. H. Kaufmann

Short-circuit-current Calculating Procedures FUNDAMENTALS OF A-C SHORT-CIRCUIT CURRENTS The determination of short-circuit currents in power distribution systems is just as basic and important as the determination of load currents for the purpose of applying circuit breakers, fuses, and motor starters. The magnitude of the shoncircuit current is often easier to determine than the magnitude of the load current. Calculating procedures have been so greatly simplified compared with the very complicated procedures previously used that now only simple arithmetic is required to determine the short-circuit currents in even the most complicated power systems. SHORT-CIRCUIT CURRENTS AND THEIR EFFECTS

If adequate protection is to he provided for a plant electric system, the size of the electric power system must also be considered to determine how much short-circuit current i t will deliver. This is done so that circuit breakers or fuses may he selected with adequate interrupting capacity (IC). This interrupting capacity should be high enough to open safely the maximum short-circuit current which the power system can cause to flow through a circuit breaker if a short circuit occurs in the feeder or equipment which it protects. The magnitude of the load current is determined by the amount Of work that is being done and hears little relation to the size of the system supplying the load. However, the magnitude of the short-circuit current is somewhat independent of the load and is directly related to the size or I

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

2

capacity of t,he power source. The larger the apparatus which supplies electric power t o the system, the greater the short-circuit current will be. Take a simple case: A 440-volt three-phase lo-lip motor draws about 13 amp of current a t full load and will draw only this amount whether supplied by a 25-kva or a 2500-kva transformer bank. So, if only thc load currcnts arc considered when selecting motor branch circuit breakers, a 15- or 20-amp circnit, breaker wnuld he specified. However, the size of t,he power system back of the circuit breaker has a real bearing on the amount of the short,-circuit,current. which can flow as a result of a short circuit on the load side of the circuit breaker. Hence, a much larger circuit breaker would be required to handle the short-circuit current from a 2500-kva bank than from a 25-kva bank of transformers. A simple mathematical example is shown in Fig. 1.1. These numbers MUST BE CAPABLE OF

El

INTERRUPTING

MOTOR IOOV 100 A

APPARENT IMPEDANCE 20 OHMS

CIRCUIT CURRENT =

E ZT

MUST

LOAD

CURRENT 5 AMP

~ ~ 1 0O.HM 1S

SHORT

1000 AMPERES

:

I00 = 1000- AMPERES 0.1

BE CAPABLE OF INTERRUPTING 10,000 AMPERES

w

I000 A 2 1 = 0.01 OHMS

FIG. 1.1

MOTOR LOAD CURRENT 5 AMP

Illustrotion showing that copocity of power source has more effect on rhortcircuit-current magnitude than load.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

3

have been chosen for easy calculation rather than a representation of actual system conditions. The impedance, limiting the flow of load current, consists mainly of the 20 ohms apparent impedance of the motor. If a short circuit occurs at F , the only impedance t o limit the flow of short-circuit current is the transformer impedance (0.1 ohm compared with 20 ohms for the motor); therefore, the short-circuit current is 1000 amp, or 200 times as great as the load current. Unless circuit breaker A can open 1000 amp, the short-circuit current will continue to flow, doing great damage. Suppose the plant grows and a larger transformer, one rated a t 1000 amp, is substituted for the 100-amp unit. A short circuit a t F , (bottom in Fig. 1.1) will now be limited by only 0.01 ohm, the impedance of the larger transformer. Although the load current is still 5 amp, the shortcircuit current will now he 10,000 amp, and circuit breaker A must be able t o open that amount. Consequently it is necessary to coiisider the size of the system supplying the plant as well as the load current, to be sure that circuit breakers or fuses are selected which have adequate interrupting rating for stopping the flow of the short-circuit current. Short-circuit and load currents are analogous t o the flow of xvater in a hydroelectric plant, shoivn in Fig. 1.2. The amount of water that flows under normal conditions is determined by the load on the turbines. Within limits, it makes little difference whether the reservoir behiiid the dam is large or small. This flow of water is comparable to the flow of load current in the distribution system in a factory. On the other hand, if the dam breaks, the amount of water that will flow will depend upon the capacity of the reservoir and will bear little relation to the load on the turbines. Whether the reservoir is large or small will make a great difference in this case. This flow of water is comparable t o the flow of current through a short circuit in the distribution system. The load currents do useful work, like the water that flows down the penstock through the turbine water wheel. The short-circuit currents produce unwanted effects, like the torrent that rushes madly downstream when the dam breaks. SOURCES OF SHORT-CIRCUIT CURRENTS

When determining the magnitude of short-circuit currents, it is extremely important that all sources of short-circuit current he considered and that the reactance characteristics of these sources be known. There are three basic sources of short-circuit current: 1. Generators 2. Synchronous motors and synchronous condensers 3. Induction motors

4

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

All these can feed shorecircuit current into a short circuit (Fig. 1.3). Generators are driven by turbines, diesel engines, water wheels, or other types of prime movers. When a short circuit occurs on the circuit fed by a generatar, the generator continues t o produce voltage because the field excitation is maintained and the prime mover drives the generator at substantially normal speed. The generated voltage produces a shortcircuit current of a large magnitude which flows from the generator (or generators) to the short circuit. This flow of short-circuit current is limited only by the impedance of the generator and of the circuit between the generator and the short circuit. For a short circuit a t the terminals of the generator, the current from the generator is limited only by its own impedance.

FIG. 1.2

Normal load and short-circuit currents are analogous to the conditions shown in

the hydroelectric plant.

SHORT-CIRCUIT-CURRENT ULCULATlNG PROCEDURES

5

METAL CLAD SWITCHGEAR

SHORT CIRCUIT

CURRENT FROM INDUCTION MOTOR

FIG. 1.3

Generators, synchronous motors, and induction motors all produce short-circuit

current.

HOW SYNCHRONOUS MOTORS PRODUCE SHORT-CIRCUIT CURRENT

Synchronous motors are constructed substantially like generators; i.e., they have a field excited by direct current and a stator winding in which alternating current flows. Normally, synchronous motors draw a-c power from the line and convert electric energy to mechanical energy. However, the design of a synchronous motor is so much like that of a generator that electric energy can be produced just as in a generator, by driving the synchronous motor with a prime mover. Actually, during a system short circuit the synchronous motor acts like a generator and delivers shortcircuit current to the system instead of drawing load current from it (Fig. 1.4). As soon as a short circuit is established, the voltage on the system is reduced to a very low value. Consequently, the motor stops delivering energy to the mechanical load and starts slowing down. However, the inertia of the load and motor rotor tends to prevent the motor from slowing down. In other words, the rotating energy of the load and rotor drives the synchronous motor just as the prime mover drives a generator.

6

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

The synchronous motor then becomes a generator and delivers shortcircuit current for many cycles after the short circuit occurs on the system. Figure 1.5 shows an oscillogram of the current delivered by a synchronous motor during a system short circuit. The amount of current depends upon the horsepower, voltage rating, and reactance of the synchronous motor and the reactance of the system to the point of short circuit.

LOAD CURRENT

FIG. 1.4 UlILITY SYSTEM

SYNCHRONOUS MOTOR

,-

\

Normally motors draw load current from the source or utility system but produce rhortcircuit current when a short cirw i t occurs in the d a d .

-€t SHORT CIRCUIT CURRENT FROM MOTOR

. .-.. . SYSTEM

SYNCMOYOUS

'

Yoroll

'.

SHORT CIRCUIT

-

I

-.

__

FIG 5 IBmlowl c. e f. 0s. . 1.._ ,. ..,. l.r o. .o . . . cillogrclm of short-circuit current produced by a synchronous motor

SHORT CIRCUIT CURRENT DELIVERED BY A SYNCHRONOUS MOTOR.

SHORT.CIRCUIT-CURRENT CALCULATING PROCEDURES

7

HOW INDUCTION MOTORS PRODUCE SHORT-CIRCUIT CURRENT

The inertia of the load and rotor of an induction motor has exactly the same effect on an induction motor as on a synchronous motor; i.e., it drives the motor after the system short circuit occurs. There is one major difference. The induction motor has no d-c field winding, but there is a flux in the induction motor during normal operation. This flux acts like flux produced by the d-c field winding in the synchronous motor. The field of the induction motor is produced by induction from the stator rather than from the d-c winding. The rotor flux remains normal as long as voltage is applied to the stator from an external source. However, if the external source of voltage is removed suddenly, as it is when a short circuit occurs on the system, the flux in the rotor cannot change instantly. Since the rotor flux cannot decay instantly and the inertia drives the induction motor, a voltage is generated in the stator winding causing a short-circuit current to flow to the short circuit until the rotor flux decays to zero. To illustrate the short-circuit current from an induction motor in a practical case, oscillograms were taken on a woundrotor induction motor rated 150 hp, 440 volts, 60 cycles, three phase, ten poles, 720 rpm. The external rotor resistance was short-circuited in each case, in order that the effect might he similar to that which would he obtained with a low-resistance squirrel-cage induction motor. Figure 1.6 shows the primary current when the machine is initially running light and a solid three-phase short circuit is applied a t a point in the circuit close to its input (stator) terminals a t time TI. The current shown is measured on the motor side of the short circuit; so the shortcircuit current contribution from the source of power does not appear, but only that contributed by the motor. Similar tests made with the machine initially running a t full load show that the short-circuit current produced

T.

FIG. 1.6

Tracer of oxillograms of short-circuit currents produced running a t light load. ,

by an induction motor

8

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

by the motor when short-circuited is substantially the same, regardless of initial loading on the motor. Note that the maximum current occurs in the lowest trace on the oscillogram and is about ten times rated full-load current. The current vanishes almost completely in four cycles, since there is no sustained field current in the rotor to provide flux, as in the case of a synchronous machine. The flux does last long enough to prodnce enough short-circuit current to affect the momentary duty on circuit breakers and the interrupting duty on devices which open within one or two cycles after a short circuit. Hence, the short-circuit current produced by induction motors must he considered in certain calculations. The magnitude of short-circuit current produced by the induction motor depends upon the horsepower, voltage rating, reactance of the motor, and the reactance of the system to the point of short c. "cuit. The machine impedance, effective a t the time of short circuit, cmesponds closely with the impedance a t standstill. Consequently, the i iitial symmetrical value of Short-circuit current is approximately equnl to the full-voltage starting current of the motor. TRANSFORMERS

Transformers are often spoken of as a source of short-circuit current. Strictly speaking, this is not correct, for the transformer merely delivers the short-circuit current generated by generators or motors ahead of the transformer. Transformers merely change the system voltage and mag; nitude of current but generate neither. The short-circuit current delivered by a transformer is determined by its secondary voltage rating and reactance, the reactance of the generators and system to the terminals of the transformer, and the reactance of the circuit from the transformer to the short circuit. ROTATING-MACHINE REACTANCE

The reactance of a rotating machine is not one simple value as it is for a transformer or a piece of cable, but is complex and variable with time. For example, if a short circuit is applied to the terminals of a generator, the short-circuit current behaves as shown i n Fig. 1.7. The current starts out a t a high value and decays to a steady state after some time has elapsed from the inception of the short cirroit. Since the field excitation voltage and speed have remained snbstantially constant within the short interval of time considered, a change of apparent react,ance of the machine may he assumed, to explain the change in the magnitude of short-circuit current with time. The expression of such variable reactance at any instant after the

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

9

occurrence of any short circuit requires a complicated formula involving time as one of the variables. For the sake of simplification in short-circuit calculating procedures for circuit-breaker and relay applications, three values of reactance are assigned to generators and motors, viz., subtransient reactance, transient reactance, and synrhronous reactance. The three reactances can be briefly described as follows: 1. Subtransient reactance X y is the apparent reactance of the stator winding at the instant short circuit occurs, and it determines the current Row during the first few cycles of a short circuit. 2. Transient reactance X i is the apparent initial reactance of the stator winding, if the effect of all amortisseur windings is ignored and only the field winding considered. This reactance determines the current following the period when subtransient reactance is the controlling value. Transient reactance is effective up to 45 see or longer, depending upon the design of the machine. 3. Synchronous reactance X d is the apparent reactance that determines the current flow when a steady-state condition is reached. It is not effective until several seconds after the short circuit occurs; consequently, it has no value in short-circuit calculations for the application of circuit breakers, fuses, and contactors but is useful for relay-setting studies. Figure 1.8 shows the variation of current with time and associates the various reactances mentioned above with the time and current scale. Previous loading has an effect on the total magnitude of short-circuit

CURRENT DETERMINED BY SYNCHRONOUS

OCCURS AT THIS TIME.

OF TOTAL OSCILLOGRAM

ONLY TWO ENDS SHOWN HERE. THIS REPRESENTS THE BREAK BETWEEN THE TWO PARTS.

FIG. 1.7 Trace of orcillograrn of hart-circuit current produced by a generator.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

10

/-

MAX, SUBTRANSIENT

CURRENT- USE SUBTRANSIENT REACTANCE X"d

TIME(8)

FIG 1.8

Variation of generotor short-circuit current wilh time.

current delivered by a generator. The value of X i or X y generally given by the machine designer is the lowest value obtainable. Hence, its use will show maximum short-circuit current. Certain characteristics of short-circuit currents must he understood before a system analysis can he made. SYMMETRICAL AND ASYMMETRICAL SHORT-CIRCUIT CURRENTS

These terms are used to describe the symmetry of the a-c waves about the zero axis. If the envelopes of the peaks of the current waves are symmetrical about the zero axis, the current is called symmetrical current (Figs. 1.9 and 1.10). If the envelopes of the peaks of the current waves are not symmetrical about the zero axis, the current is called asymmetrical ENVEWPES OF PEAKS OF SINE WAVE ARE SYMMETRIGAL ABOUT THE ZERO AXIS. ZERO

AXIS

FIG. 1.9 Symmelrical a-c wove.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

11

THE ENVELOPES OF PEAKS ARE SVHHETRICAL ABOUT

ZERO AXIS

FIG, 1.10

Symmetrical

d t e r n a t i n g current f r o m a short-circuited generotor.

ENVELOPES OF PEAKS ARE NOT SYMMETRICAL ABOUT ZERO AXIS

AX1 S TOTALLY 0 F F SET PARTIALLY O F F S E l

FIG. 1.11 Asymmetrical (I-c waver. The conditions shown here ore theoreticol a n d ore for the purpose of illustration only. D-C component will r a p i d l y d e c a y to zero i n a c t u a l circuits.

FIG. 1.12

Trace of o r c i l l o g r a m of a t y p i c a l short-circuit current

12

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

current (Fig. 1.11). The envelope is a line drawn through the peaks of the waves, as shown in Figs. 1.9 to 1.12. For the sake of explanation, many of the illustrations, such as Figs. 1.11, 1.15 to 1.19, show sine waves of current uniformly offset for several cycles. It should be noted that in practical circuits the amount of asymmetry decreases rapidly after the occurrence of the short circuit in the system. This decrease of asymmetry is shown qualitatively in illustrations such as Figs. 1.12, 1.20, 1.23, and 1.24. Oscillograms show that short-circuit currents are nearly always asymmetrical during the first few cycles after the short circuit occurs. They also show that the asymmetry is maximum at the instant the short circuit occurs and that the current gradually becomes symmetrical a few cycles after the occurrence of the short circuit. The trace of an oscillogram of a typical short-circuit current is shown in Fig. 1.12. WHY SHORT-CIRCUIT CURRENTS ARE ASYMMETRICAL

In the usual industrial power systems the applied or generated voltages are of sine-wave form. When a short circuit occurs, substantially s i n e wave short-circuit currents result. For simplicity, the following discussion assumes sine-wave voltages and currents. In ordinary power circuits the resistance of the circuit is negligible compared with the reactance of the circuit. The short-circuit-current power factor is determined by the ratio of resistance and reactance of the circuit only (not of the load). Therefore the short-circuit current in most power circuits lags the internal generator voltage by approximately 90" (see Fig. 1.13). The internal generator voltage is the voltage generated in the stator coils by the field flux. If in a circuit mainly containing reactance a short circuit occurs at the peak of the voltage wave, the short-circuit current would start at zero and trace a sine wave which would be symmetrical ahout the zero axis (Fig. 1.14). This is known as a symmetrical short-circuit current. If in the same circuit (i.e., one containing a large ratio of reactance to resistance) a short circuit occurs at the zero point of the voltage wave, the current will start a t zero but cannot follow a sine wave symmetrically about the zero axis because such a current would be in phase with the voltage. The wave shape must be the same as that of voltage hut 90' behind. That can occur only if the current is displaced from the zero axis, as shown in Fig. 1.15. In this illustration the current is a sine wave and is displaced 90' from the voltage wave and also is displaced from the zero axis. The two cases shown in Figs. 1.14 and 1.15 are extremes. One shows a symmetrical current and the other a completely asymmetricd current.

WORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

13

GENERATOR TRANSFORMER INTERNAL VOLTAGE OF GENERATOR APPLIED HERE

ONE LINE IMPEDANCE

ioxazx

7x 0.m

REACTANCE, X = 19% RESISTANCE. R = 1.4%

I

RESISTANCE I S LESS THAN OF THE REACTANCE BE NEGLECTED WITHOUT AN APPRECIABLE ERROR

HENCE MAY

INTERNAL VOLTAGE OF GENERATOR

-

NEARLY 90'

SHORT CIRCUIT CURRENT

DIAGRAM SHOWING SINE WAVES CORRESPONDING TO VECTOR DIAGRAM FOR ABOVE CIRCUIT

FIG. 1.13

Diagrams Illustrating the phase relations of voltage and short-circuit current.

14

SHORT-CIRCUll-CURRENT CALCULATING PROCEDURES

GENERATED VOLTAGE SHORT CIRCUIT CURRENT

ZERO AXIS

SHORT CIRCUIT OCCURRED AT THIS POINT

FIG. 1.14

Symmetric01 short-circuit current and generoted voltage for zero-power-factor

cirwit.

-SHORT CIRCUIT CURRENT

FIG. 1.15 circuit.

Asymmetrical short-circuit current and generated voltage in zero-power-factor Condition i s theoretical and is shown for illustration purposes only.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

IS

If,in a circuit containing only reactance, the short circuit occurs a t any point except a t the peak of the voltage wave, there will be some offset of the current (Fig. 1.16). The amount of offset depends upon the point on the voltage wave at which the short circuit occurs. It may vary from zero (shown in Fig. 1.14) to a maximum (shown in Fig. 1.15). I n circuits containing both reactance and resistance, the s~,?&&,R&!~~ amount of offset of the shortCURRENT circuit current may vary between the same limits as for circuits containing only reactance. However, the point on the voltage wave a t which the short circuit must occur to produce maximum asymmetry dependsupon the ratioof reactance to resistance of the circuit. Maximum asymmetry is obtained when the short circuit occurs a t a time angle equal to 90" 0 (measured forward in degrees from the zero point of the voltage wave) where tangent 0 equals thereASYMMETRICAL actance-to-resistance ratio of FIG. 1.16 Short-circuit current and generated the circuit' The short-circuit voltage in zero-Dower-factor circuit. Short circurrent will be symmetrical cuit occurred between the when the fault occurs 90"from point and peak of the generated voltctge wove. that point onthe voltage wave. This condition i s theoretical and for illustration an example, assumeacir- purporer only. The short-circuit current will gradually become symmetrical in practical cuit that has equal resistance CiTCUit., and reactance, i.e., the reactance-to-resistance ratio is 1. The tangent of 45" is I ; hence, maximum offset is obtained when the short circuit occurs a t 135' from the zero point of the voltage wave (Fig. 1.17).

+

D-C COMPONENT OF ASYMMETRICAL SHORT-CIRCUIT CURRENTS

Asymmetrical alternating currents when treatedas a single current wave are difficult to interpret for circuit-breaker application and relay-setting purposes. Complicated formulas are also required to calculate their magnitude unless resolved into components. The asymmetrical alternating currents are, for circuit-breaker applications and relay-setting

16

SHORT-CIRCUIT-CURRENT CALCUUTING PROCEDURES

MAXIMUM OFFSET

FIG. 1.17 Short-circuit current and generated voltage in circuit with equal reactance and resistance. This condition i s theoretical and is shown for illustration purposes only. The short-circuit current will gradually become symmetrical in practical circuits.

purposes, arbitrarily divided into simple components, which makes it easy to calculate the short-circuit magnitude a t certain significant times after the short circuit occurs. The asymmetrical alternating current behaves exactly as if there were two component currents flowing simultaneously. One is a symmetrical a-c component and the other a d-c component. The sum of those two components a t any instant is equal t o the magnitude of the total asymmetrical a-c wave a t the same instant. The d-c component referred to here is generated within the a-c system with no external source of direct current being considered. I n some cases, particularly in the neighborhood of the d-c railways, direct current from the railways flows through neighboring a-c systems. This type of d-c current is not considered in this discussion or in the calculating procedures which follow. As an example of the resolution of asymmetrical alternating currents into components, refer to Fig. 1.15 which shows an asymmetrical shortcircuit current which is resolved into a symmetrical a-c and a d-c component in Fig. 1.18. If the instantaneous values of the two components (dashed lines) are added a t any instant, the resultant will be that of the asymmetrical current wave.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

17

F I N S T A N T AT WHICH SHORT CIRCUIT OCCURS

ASYMMETRICAL

AC COMPONENT

FIG. 1.18 current.

Theoretical Ihort-circuit-cvrrent wove illustrating components of asymmetrical In practical circuits, d-c component would decay to zero in o few cycler.

INSTANT

OF SHORT CIRCUIT

TOTAL CURRENT

DC COMPONENT AC COMPONENT

ZERO A X I S

a = b = D C COMPONENT FIG. 1.19 Components of asymmetrical short-circuit current in which short circuit occurred at some point between the zero point and p e a k of the generated voltage wave. This is a lhsoretical condition similar to that shown in Fig. 1.18.

I8

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

As mentioned previously, the examples shown in Figs. 1.13 and 1.18 are for purposes of illustration only. In practical circuits the d-c component decays very rapidly, as shown in Fig. 1.20. INITIAL M A G N I T U D E OF D-C C O M P O N E N T

The magnitude of the d-c component depends upon the iustant, the short circuit occurs and may vary from zero, as in Fig. 1.14, to a maximum initial value equal to the peak of the a-c symmetrical compoiieiit, as i n Figs. 1.15 and 1.18. When the short circuit occurs at any other point, such as shown in Fig. 1.19, the initial magnitude of the d-c componciit is equal to the value of the a-c symmct,riral component a t thc instant of short circuit. The above limit,s hold true for the initial magiiitudc of d-c eomporient in a system regardless of the reactance and resistance. Ilowever, the d-c componeut does not continue to flo~va t a constant value, as shown i n Figs. 1.18 and 1.19, unless there is zero resistauce i i i the circuit. DECREMENT

There is uo d-c voltage in the system t o sustaiu the flax of direct current; therefore the energy represeuted by the dirert. component of current will be dissipated as ZZR loss from the direct current flowiug through the resistance of the circuit. If the circuit had zero resistance, the direct current would flow at a constant value (Figs. 1.18 and 1.19) TOTAL ASYMMETRICAL CURRENT C

COMPONENT AC

COMPONENT

FIG. 1.20 Trace of orcillogrom showing decay of d-c component and how orymmetricd short-circuit currenl gradually becomes symmetrical when d-c component diroppearr.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

19

until the circuit was interrupted. However, all practical circuits have some resistance; so the d-c romponent decays as shown in Fig. 1.20. The combination of the decaying of d-c and symmetriral a-(*components gives an asymmetrical wave that changes to a symmetriral wave whcti the d-c component has disappeared. The rate of decay of the currents is called the decrement.

X/R

RATIO

The X / R ratio is the ratio of the reactance to the resistance of the circuit. The decrement or rate of decay of the d-c component is proportional to the ratio of reactance to resistance of the complete circuit from generator to short circuit. The theory is the same as opening the circuit of a battery and an inductive coil. If the ratio of reactance to resistance is infinite (i.e., zero resistance), the d-c component never decays, as shown in Figs. 1.18 and 1.19. On the other hand, if the ratio is zero (all resistance, no reartance), it decays instantly. FOFany ratio of reactarice to resistance in between these limits, the d-c component takes a definite time to decrease to substantially zero, as shown in Fig. 1.20. ! I n generators the ratio of subtransient reactance to resistance may be as ?much as 7 0 : l ; so it takes several cycles for the d-c component to disappear. In circuits remote from generators, the ratio of reactance to resistance is lower, and the d-c component decays more rapidly. The higher the resistance in proportion to the reactance, the more IaRloss from the d-c c.omponent, and the energy of the direct current is dissipated sooner. D-C TIME CONSTANT

Often it is said that generators, motors, or circuits have a certain d-c time constant. This refers again to the rate of decay of the d-c compoO C COMPONENT

a

C-

= 37Y. OF b (APPROX

TIME

CONSTANT I N SECONDS FIG. 1.21

OF D C COMPONENT

Graphic illustration of time constant.

)

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

20

nent. The d-c time constant is the time, in seconds, required by the d-c component to reduce to about 37 per cent of its original value a t the instant of short circuit. I t is the ratio of the inductance in henrys to the resistance in ohms of the machine or circuit. This is merely a guide to how fast the d-c component decays. Stated in other terms, it is the time in seconds for the d-c component to reach zero if it continued t o decay a t the same rate it does initially (Fig. 1.21). RMS VALUE INCLUDING D-C COMPONENT

The rms values of a-c waves are significant since circuit breakers, fuses, and motor starters are rated in terms of rrns current or equivalent kva. The maximum rrns value of short-circuit current occurs at a time of about one cycle after short circuit, as shown in Fig. 1.20. If there were no decay in the d-c component, as in Fig. 1.18, the rrns value of the first cycle of current would be j.732 times the rrns value of the a-c component. I n practical circuits there is always some d-c decay during the first cycle. An approximate rrns value of one cycle of an offset wave whether it is partially or totally offset is expressed by the equation

where C

=

a b

= =

rrns value of offset or asymmetrical current wave over one cycle rrns value of a-c component value of d-c component at one-half cycle

MULTIPLYING FACTOR

Calculation of the precise rrns value of an asymmetrical current a t any time after the inception of a short circuit may be very involved. Accurate decrement factors to account for the d-c component a t any time are required, as well as accurate factors for the rate of change of the apparent reactance of the generators. This precise method may he used if desired, but simplified methods have been evolved whereby the d-c component is accounted for by simple multiplying factors. The multiplying factor converts the rrns value of the symmetrical a-c wave into rms amperes of the asymmetrical wave including a d-c component. The magnitude of the d-c component depends upon the point on the voltage wave a t which the short circuit occurs. For protective-device application, only the maximum d-c component is considered, since the circuit breaker must be applied to handle the maximum short-circuit current that can occur in a system.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

21

In the general case for circuits rated above 600 volts, the multiplying factor to account for d-c component is 1.6 times the rms value of the a-c symmetrical component at the first half cycle. For circuits rated 5000 volts or less where there is no local generation, that is, where the supply t,o the bus is through transformers or long lines, the multiplying factor to ralculate the total current at the first half cycle may be reduced to 1.5. For circuits 600 volts and less, t,he multiplying factor to calculate the total current at the first half cycle is 1.25 when the circuit breaker is applied on the average current in three phases. Where single-phase conditions must be considered in circuits GOO volts and less, then to account for the d-c component in one phase of a three-phase circuit a multiplying factor to calculate the total current at the first half cycle of 1.5 is used. For some calculations, rms current evaluations a t longer time intervals than the first half cycle, such as three to eight cycles corresponding to the interrupting time of circuit breakers, are required. Multiplying factors for this purpose may be taken from the curve in Fig. 1.22. Table 1.2 gives the multiplying factors commonly used for applying

e

FIG. 1.22 Charts showing multiplying factors to account for decoy of d-c component for various X / R ratio of circuits.

22

SHORT-CIRCUIT-CURREM CALCULATING PROCEDURES

short-circuit protective devices. These factors range from 1 t o 1.6, depending upon whether the short-circuit calculation is being made t o determine the interrupting or momentary duty on the short-circuit protective device. SHORT-CIRCUIT RATIO OF GENERATORS

This term is referred t o frequently in short-circuit discussions. With present AIEE procedures of short-rircuit ralrulations, it has become a n accessory with no practical significance from this standpoint. For the sake of completeness, a definition is given here. Short-circuit ratio field current t o produce rated voltage a t no load -~ field current t o produce rated current at sustained short circuit

No further mention will he made of short-circuit ratio. TOTAL SHORT-CIRCUIT CURRENT

The total symmetrical short-rirruit current is made up of currents from several sourves, Fig. 1.23. At the top of the figure is shown the shortcircuit current from the utility. This act,ually comes from ut,ility generators, but generally the industrial system is small and remote electrically from the utility generators so that the Symmetrical short-rircuit current is substant,ially constant,. If there are generators in the indust,rial plant, then they cont,ribute a symmet,rical short-circuit rurreiit which for all practical purposes is constant over the first few cycles. There is, however, a slight decrement, as indicated in Fig. 1.23. The other sources are synchronous motors which act something like plant generators, except that t,hey have a higher rate of decay of the symmetriral component, and induction motors whirh have a very rapid rate of dccay of the symmetrical component of current. When all these currents are added, the total symmetrical short-circuit rurrent is typical of that shown a t the bottom of Fig. 1.23. The magnitude of the first few cycles of the t,otal symmetrical shortcircuit, current is further increased by the presence of a d-c compouent, Fig. 1.24. The d-c component, offsets the a-c ware and, therefore, makes it asymmetrical. The d-c component decays t o zero within a few cycles in most indust,rial power systems. It is this total rms asymmetrical short-circuit current, as shown in Fig. 1.24, that must he determilied for short-circuit protective-derice appliration. The problem of doing this has been simplified by standardized procedures to a poiut xhere t o determine the rms asymmetriral current one need only divide t,he line-to-neutral roltage by the proper reactance

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

RG. 1.23 Tracer of orcillogramr of rymmetrical short-circuit currents from utility, panerator, synchronous motors, and induclion motors. The shape of the total combined currents is illurtmted by the bottom hace.

23

FIG. 1.24 Arymmelrical short-circuit current from dl sources illustrated in Fig. 1.23 plus d-c component.

24

SHORT.CIRCUIT-CURRENT U L C U U l l N G PROCEDURES

or impedance and then multiply by the proper multiplying factor from Table 1.2. BASIS OF RATING A-C SHORT-CIRCUIT PROTECTIVE DEVICES

The background of the circuit-breaker rating structure as well as the basic characteristics of short-circuit currents must be understood to enable the engineer to select the proper rotating-machine reactances and multiplying factors for the d-c component to determine the sbort-circuitcurrent magnitude for checking the duty on a particular circuit breaker, such as momentary duty or interrupting duty. The rating structure of circuit breakers, fuses, and motor starters is designed to tell the application engineer how circuit breakers, fuses, or motor starters will perform under conditions where the short-circuit current varies with time. In discussing these rating bases, and for the sake of clarity, they will be arbitrarily divided into two sections, i.e., the rating basis of high-voltage short-circuit protective devices above 600 volts and the rating basis of low-voltage Short-circuit protective devices 600 volts and below. HIGH-VOLTAGE SHORT-CIRCUIT PROTECTIVE DEVICES (ABOVE 600 VOLTS)

Power-circuit-breaker Rating Basis. The standard indoor oilless power circuit breakers as used in metal-clad switchgear will be used here t o explain power circuit-breaker ratings. The same fundamental principles apply to all other high-voltage power circuit breakers. The circuit-breaker rating structure is complicated because of the time of operation of the circuit breakers after a short circuit occurs. The few cycles needed for the power circuit breaker to open the circuit and stop the flow of short-circuit current consist of the time required for (1) the protective relays to close their contacts, (2) the circuit-breaker trip coil to move its plunger to release the breaker operating mechanism, (3) the circuit-breaker contacts to part, and (4)the circuit breaker to interrupt the short-circuit current in its arc chamber. During this time, the short-circuit current produces high mechanical stresses in the circuit breaker and in other parts of the circuit. These stresses are produced almost instantaneously in phase with the current and vary as the square of the current. Therefore, they are greatest when maximum current is flowing. The foregoing discussion showed that t,he short-circuit current is maximum during the first cycle or loop, because of the presence of the d-c component and because the motors contribute the most short-circuit current a t that time. Thus, the short-circuit stresses on the circuit breakers and other parts of the circuit are maximum during the first loop of short-circuit current. During the time from the inception of the short circuit until the circuitbreaker contacts part, the current decreases in magnitude because of the

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

25

decay of the d-c component and the change in motor reactance, as explained previously. Consequently, the current that the circuit breaker must interrupt, four or five cycles after the inception of t.he short circuit, is generally of less magnitude than the maximum value of the first loop. The fact that the current changes in magnitude with time has led to the establishment of two bases of short-circuit-current ratings on power circuit breakers: (1) the momentary rating or its ability to withstand mechanical stresses due to high short-circuit current and (2) the interrupting rating or its ability t,o interrupt the flow of short-circuit current within its interrupting element. What Comprises the Circuit-breaker-rating Structure. Circuitbreaker-rating structures are revised and changed from time to time. It is suggested that where specific problems require the latest information on circuit-breaker ratings the applicahlc American Standards Association (ASA), National Electrical Manufacturers Association (XEMA), or American Instituteof Elect,rical Engineers (AIEE) standards he referred to. To illustrate the various factors that comprise the circuit-breakerrating structure, an oilless power circuit breaker for metal-clad switchgear rated 4.16 kv 250 mva* has been chosen. The complete rating is shown on line 5, Table 1.1. The following will explain the meaning of the several columns of Table 1.1, starting at the left. The rircuit-breaker-type designation, column 1, varies among manufacturers. For the sake of completeness the General Electric Company nomenclature is used in this column. The remainder of the items are uniform throughout the industry. 1. Type of Circuit Breaker (AM-4.16-250) AM = magne-blast circuit breaker 4.16 = for 4.16-kv class of circuits (not applicable to 4800- and 4800volt circuits) 250 = interrupting rating in mva a t 4.16 kv

2-4. Voltage Rating 2. Rated kv (4.16): the nominal voltage class or classes in which the circuit breaker is rated. 3. Maximum design kv (4.76): the maximum voltage a t which the circuit breaker is designed to operate. The 4.16-kv circuit breakers, for example, are suitable for a 1330-volt system plus 10 per cent for voltage regulation or 4.76 kv. (Note: 4330 is 4% X 2500.) Some utility syst.ems operate a t 1330 volts near the substation. 4. Minimum operating kv a t rated mva (3.85) : the minimum voltage a t which the circuit breaker will interrupt its rated mva or in this case it is 3.85 kv. At any voltages below this value, the circuit breaker

* blegavalt-amperes t

i.

(see Appendix).

16

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

I !

I (

I

t

a /

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

27

is not designed to interrupt the rated mva but will interrupt some value less than rated mva. This is very significant in the rating of power circuit breakers for, as poiuted out later, the circuit hreaker will interrupt a maximum of only so many amperes regardless of voltage. At any voltage less than the minimum operating voltage the product of the maximum kiloampere interrupting rating times the kv times the square root of 3 is less than the mva interrupting rating of the circuit breaker. 5-6. Insulation Level (Withstand Test) 5 . Low-frequency rrns kv (19): the 60-cycle high-potential test. 6. Impulse crest kv (60) : a measure of its ability to withstand lightning and other surges. This is applied with an impulse generator as a design test.

7-9. Current Ratings in Amperes 7. Continuous 60 cycles (1200 or 2000): the amount of load current which the circuit breaker will carry continuously without exceeding the allowable temperature rise. 8-9. Short-time Rating 8. Momentary amperes (60,000) : the maximum rms asymmetrical current that a circuit breaker will withstand including short-circuit cnrrents from all sources and motors (induction and synchronous) and the d-c component. This rating is independent of operating voltage for a given circuit breaker. This is just as significant a limitation as mva interrupting rating. It defines the ability of the circuit breaker to withstand the mechanical stresses produced by the very large offset first cycle of the shortcircuit current. This rating is nnusually significant because the mechanical stresses in the circuit hreaker vary as the square of the current. It is the only rating that is affected by the square law, and therefore is one of the most critical in the application of the circuit breakers. The rating schedules of power circuit breakers are so proportioned that the momentary rating is about 1.6 times the maximum interrupting rating amperes. 9. Four-second (37,500): the maximum current that the circuit breaker will withstand in the closed position for a period of 4 sec to allow for relaying operating time. This value is the same as the maximum interrupting rating amperes.

10-13. Interrupting Ratings 10. Three-phase rated mva (250): the three-phase mva which the circuit breaker will interrupt over a range of voltages from the maximum design kv down t o the minimum operating kv. In this case the

28

SHORT-CIRCUIT-CURREM CALCULATING PROCEDURES

interrupting rating is 250 rnva between 4.76 and 3.85 kv. The mva to be interrupted is obtained by multiplying the kv a t which the circuit breaker operates times the symmetrical current in kiloamperes to be interrupted times the square root of 3. The product of these must not exceed the rnva interrupting rating a t any operating voltage. 11. Amperes a t rated voltage (35,000): the maximum total rms amperes which the circuit breaker will interrupt a t rated voltage, i.e., in the case of the example used above 35,000 at 4.16 kv (4.16 X 35.000 x fi = 250 mva). These figures are rounded. This figure is given for information only and does not have a limiting significance of particular interest to the application engineer. 12. Maximum amperes interrupting rating (37,500) : the maximum total rms amperes that the circuit breaker will interrupt regardless of how low the voltage is. In this example, this current is 37,500 amp. At minimum operating voltage, 3.85 kv, this corresponds to 250 mva, and, for example, a t a voltage of 2.3 kv this corresponds to 150mva. The circuit breaker will not interrupt this much current a t all voltages, i.e., i t will not interrupt this much current if the product of current, voltage, and the square root of 3 is greater than the mva interrupting rating. This current limit determines the minimum kv a t which the circuit breaker will interrupt rated mva (column 4). At any voltage lower than that given in column 4, this maximum rms total interrupting current determines how much the circuit breaker will interrupt in mva. Therefore, when the voltage goes below the limit of column 4, the mva which the circuit breaker will interrupt is lower than the rnva rating given in column 10 by an amount proportional to the reduction in operating voltage below the value of column 4. 13. Rated interrupting time (8 cycles on 60-cycle basis): the maximum total time of operation from the instant the trip coil is energized until the circuit breaker has cleared the short circuit. What limits the Application of Power Circuit Breakers an on interrupting-and Momentary-duty Basis? In so far as applying power cir-

cuit breakers on an interrupting-duty basis is concerned i t can be seen from the foregoing that there are four limits, none of which should be exceeded. These must all be checked for any application. 1. Operating voltage should never at any time exceed the limit of column 3, Table 1.1, i.e., the maximum design kv. 2. Interrupting rnva should never be exceeded a t any voltage. This limit is sig’nificant only when the operating voltage is between the limits of columns 3 and 4, Table 1.1. It is not significant when the operating voltage is below the limit of column 4, Table 1.1, because maximum interrupting amperes limit the mva to values less than the rnva rating. 3. Maximum interrupting rating amperes should never be exceeded

SHORT-CIRCUIT.CURRENT CALCUUTING PROCEDURES

29

even though the product of this current times the voltages times the square root of 3 is less than the interrupting rating in mva. This figure is the controlling one in so far as interrupting duty is involved when the voltage is below that of column 4, Table 1.1 (minimum operating voltage a t rated mva). 4. Momentary current should never be exceeded a t any operating voltage. Modern power circuit breakers generally have a momeutary rating in rms amperes of 1.6 times the maximum interrupting rating in rms amperes. As a result, where there is no short-circuit-current contribution from motors, a check of the interrupting duty only is necessary. If this is within the circuit-breaker interrupting rating then the maximum Short-circuit current, including the d-c component, mill be within the momentary rating of the circuit breaker. Where there is short-circuit contribution from motors, the momentary rating of the circuit breaker may be exceeded, before the interrupting rating is exceeded in a given cirruit. Whenever there are motors to be considered in the short-circuit calculations, the momentary duty and the interrupting duty should both be checked. How to Check Momentary Duty on Power Circuit Breakers. Siuce the short-circuit current is maximum a t the first half cycle, the short-circuit current must be determined a t the first half cycle to determine the maximum momentary duty on a circuit breaker. To determine the short-circuit current a t the first half cycle, it is necessary to consider all sources of short-circuit current, that is, the generators, synchronous motors, induction motors, and utility connections. The subtransient reactances of generators, synchronous motors, and inductiou motors are employed in the reactance diagram. Since the d-r component is present a t this time, it is necessary to account for it by the use of a multiplying factor. This multiplying factor is either 1.5 or l.G, as outlined in Table 1.2. Typical circuits where the 1.5 multiplying factor can be used are shown in Fig. 1.25. The procedure is the same, regardless of the type of power circuit breaker involved. How to Check Interrupting Duty on Power Circuit Breakers. To check the interrupting duty on a power circuit breaker, the short-circuit current should be determined a t the time that the circuit-breaker contacts part. The time required for the circuit-breaker contacts to part will vary over a considerable range, because of variation in relay time and in circuitbreaker operating speed. The fewer cycles required for the circuitbreaker contacts to part, the greater will be the curreut to interrupt. Therefore, the maximum interrupting duty is imposed upon the circuit breaker when the tripping relays operate instantaneously. In all shortcircuit calculations, for the purpose of determining interrupting duties, the relays are assumed to operate instantaneously. To account for

SEPES-DIVEN SEN-RIO-EIELI', tCA

1

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

30

HIGH VOLTAGE INCOMING LINE

2400 4160 4800 VOLT INCOMING L I N E FROM UTILITY

(0)

$

T O P L A N T LOAD NO GENERATION IN THE P L A N T

o,:4600 A6,0 V BUS

TO P LANT L O AD NO GENERATION IN THE P L A N T

(b)

13.6 KV U

u.-L

U

USE 1.6 MULTIPLYING FACTOR NO GENERATION ON THIS BUS NO GENERATION

2400, 4160 OR

(C)

TO LOAD

FIG. 1.25 One-line diogrom of carer where the multiplying factor 1.5 may be used on circuits rated less than 5 h.

c

,,

,..

.:

.. . .

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

..

,

.

31

variation in the circuit-breaker operating speed, power circuit breakers have been grouped into classes, such as eight-cycle, five-cycle, three-cycle circuit breakers, etc. It is assumed for short-circuit-calculation purposes that circuit breakers of all manufacturers, in any one speed grouping, operate substantially the same with regard to contact parting time. Instead of specifying a time a t which the short-circuit current is to he calculated, it is determined by the simpler approach of specifying the generator and motor reactances and using multiplying factors. These factors are listed in Table 1.2. In industrial plants, eight-cycle circuit breakers are generally used. Normally, the induction-motor contribution has disappeared, and that of the synchronous motors has changed from the subtransient to the transient condition before the contacts of these circuit breakers part. Therefore, in calculating the interrupting duty on commonly used power circuit breakers, generator subtransient reactance and synchronous-motor transient reactance are used and induction motors are neglected. The elapsed time is so long that usually all the d-c component has disappeared. What d-c component is left is more than offset by the reduction in a-c component due to the increase in reactance of the generators. Hence, a multiplying factor of one (1) is used. In very large power systems, when symmetrical short-circuit interrupting duty is 500 mva or greater, there is an exception to this rule. In such large power systems, the ratio of reactance to resistance is usually so high that there may be considerable d-c component left when the contacts of the standard eight-cycle circuit breaker part. To account for this, the multiplying factor of 1.1is used in determining the total rms short-circuit mva that a circuit breaker may have to interrupt in these large systems. The multiplying factor of 1.1 is not applied until the symmetrical shortcircuit value reaches 500 mva. High-voltage Fuses. High-voltage fuses are either of the currentlimiting type, Fig. 1.26, which open the circuit before the first current peak, or of the non-current-limiting type, which open the circuit within one or two cycles after the inception of the short circuit. For the sake of standardization, all fuse-interrupting ratings are on the basis of maximum rms current that will flow in the first cycle after the short circuit occurs. This is the current that will flow if the fuse did not open the circuit previously, i.e., fuses are rated in terms of “available short-circuit current.” To determine the available short-circuit current a t the first cycle for the application of high-voltage fuses, use the subtransient reactances of all generators, induction motors, synchronous motors, and utility sources and allow for the maximum d-c component. The multiplying factor for allowing for d-c component is 1.6, the same as for allowing for d-c compo-

u w

TABLE 1.2

Condensed Table of Multiplying Factors and Rotating-machine Reactances

To Be Used for CaLdatina Swt-dreuit Cunanh for Circuit-breaker, Fuse, and Motor.rtartor Applicdons

1 Generators. 1

I

I

1

I

I

frequency changers

0

w

C

a

2

Interrupting duty Eight cycle or slower (general case). Rva cycle..

.......... Above 600 wlh

..............................

Any ploee where symmetricmi short-circuit kva i s loss than 500 mva

Above 600 volt,

I .O

Subtransient Subtransient

1.1

ii

s

Momentary duty

........................... ..........................

Generol GOSO.. Lar than 5 k..

Above 600 volt) 601 to 5000 volh

Near generoting station Remote from generating dolion (X/R rotio l e u thon I01

s z

Subtransient Subtransient

1.6 1.5

High-voltaqe Fuses

5 Three-phose I n o interrupting duly

All typos, including dl wrront-limiting fuses.

All types, including dl current-limiting fuses.. Non-current-limiting lypes only..

.... Above 600 wih ... Above 600 volt'

............. 601 to 15,000 wlh

Anywhere in system

I

1

I .O

Subhqndent

1

Transient

1

Neglect

Maximum rms ampere interrupting duty Anywhere in system Remote from generoting %to. tion ( X / R mtio leu lhm 41

1.6 1 .?

i

i

i

Subtronsient Svbtronrient Svbwmrient Subwoniiont Subhmrient Subtransient

All h e p o w e r ratings..

....................

2400 and 4i60Y

Anywhere in system

Wlh

All horsepower rotingr..

....................

1.0

2400 and 4160Y

Anywhere in system

I .6

Yolh

CIrmit breaker w conladm l y p e . .

Cirwit b r w b r or contocto~lype. Clrcvit b r e e b r or contartor type..

...........

601 10 5000 volts

............ 601 to MOO volts ........... 601 lo 5000 volts

0 bywhere in system temote from gener.ting 11.lion lX/R ratio leis than 101

1.6 1.5

Subtransient Subtrmdent Subtransient Subtrmdent

Subtransient Subtransient

R 0 m

z

Apparatus. 600 Volts and Below Interrupting or momentary duty Air circuit breakers or breaker-contactor combino. lion motor stoners.. Low-voltacp furas or fused combination motor

.................... Slarte" ...............................

8

600 volts and below Anywhere in system

I .25

Subtransient Subtianrient

600 volt* and below Anywhere in system

1 .25

Subtransient Subtransient Svbtraniient

Svbtronrienl

34

SHORT-CIRCUIT.CURRENT CALCULATING PROCEDURES

nent when determining the momentary duty on a power circuit breaker (see Table 1.2). The interrupting rating of fuses in amperes is exactly parallel, in so far as short-circuit+urent calculations are concerned, to the momentary rating of power circuit breakers. The ampere interrupting rating of high-voltage fuses is the only rating that has any physical significance. For the sake of simplicity of application in systems with power circuit breakers, some fuses are given interrupting ratings in three-phase mva. The three-phase mva interrupting rating has no physical significance, because fuses are single-phase devices, each fuse functioning only on the current which passes through it. WAVE OF AVAILABLE

THE FUSE ELEMENTS MELT BEFORE PEAK VALUE OF AVAILABLE SHORT CIRCUIT CURRENT I S REACHED

1 FIG. 1.26 Grophic sxplonotion of the current-limiting action of current-limiting fuses. See Fig. 1.27 for method of determining available short-circuit current.

SHORT-CIRCUIT-CURRENT CAKULATING PROCEDURES

35

These three-phase mva ratings have been selected so they will line u p with power-circuit-breaker ratings. For example, a high-voltage fuse rated 150 mva and a power circuit breaker rated 150 mva can he applied on the basis of the same short-circuit-current calculations. Of course, the application voltage must he factored in each case. High-voltoge M o t o r Starters. High-voltage motor starters generally employ for short-circuit protection either current-limiting fuses or power circuit breakers. The short-circuit-current calculations for applying these motor starters are the same as those for high-voltage fuses and power circuit breakers, respectively. LOW-VOLTAGE CIRCUIT PROTECTIVE EQUIPMENT (600 VOLTS A N D BELOW)

low-voltage Air Circuit Breokers. The present designs of low-voltage air circuit breakers differ from those of high-voltage power circuit breakers because they are substantially instantaneous in operation a t currents near their interrupting rating. The contacts often begin to part during the first cycle of current. Therefore, low-voltage air circuit breakers are subject to interrupting the current a t the first cycle after short circuit and withstanding the mechanical forces of that rurrent. It is necessary to calculate the current a t only one time for the application of low-voltage air circuit breakers. The current determined should be that of the first halt cycle and should be determined on exactly the same hasis as for checking the momentary duty of high-voltage power circuit breakers, except for a change in the multiplying factor as discussed in the next paragraph. The suhtransient reactances of generators, induction motors, and synrhronous motors are used, and the d-c component is considered (see Table 1.2). The multiplying factor for the d-c component is not so high in lowvoltage circuits as in some high-voltage circuits. This is due to the generally lower level of reactance-to-resistance ( X I R ) ratio in low-voltage circLits, which causes the d-c component to decay faster than in some high-voltage circuits. In rating low-voltage air circuit breakers, the average d-c component of the three phases is used, which is somewhat lower than that for the maximum phase. The generally lower ( X / R ) ratio and the use of an average d-c component for the three phases result in a considerably lower multiplying factor in low-voltage circuits. The multiplying factor has been standardized at 1.25 for the average for the three phases. This is equivalent t o a multiplier of about 1.5 to account for the d-c component in the maximum phase. Application of High-voltage Oil Circuit Breokers to 600-volt Systems. In the 192Os, 5-kv oil circuit breakers were used extensively on 600-volt

36

SHORT-CIRCUIT-CURRENT CALCULAnNG PROCEDURES

systems. The procedure for determining short-circuit currents in systems of 600 volts and below is slightly modified for checking duty on oil breakers of the 5-kv class as compared with low-voltage air circuit breakers. Both the momentary duty and interrupting duty must be checked for the oil-circuit-breaker application. To check the momentary duty, use the same procedure as for low-voltage air circuit breakers, i.e., generators, utility sources, induction motors, and synchronous motors (subtransient reactance). However, a multiplying factor of 1.5 is used instead of 1.25 as for low-voltage air circuit breakers. Oil-circuit-breaker momentary ratings are based on the maximum current through any one pole, not on the average current in the three phases which is employed in the rating of low-voltage circuit breakers. To determine the interrupting duty, use the generator subtransient reactance and utility-source reactance plus the synchronous-motor transient reactance and a multiplying factor of 1.0. Low-voltage Fuses. Several low-voltage fuses with published a-c interrupting ratings are appearing on the market. There are no industry standards to follow, but most of these seem to be following air-circuitbreaker standards, i.e., using the same rating base and same method of determining short-circuit duty as is used for low-voltage air circuit breakers. Hence, the procedure will not be repeated here except to point out that the 1.25 multiplying factor is used (see Table 1.2). So-called National Electrical Code (NEC) plug and cartridge fuses have no established a-c interrupting ratings. Many tests have been made to determine their a-c interrupting ability, but to date the industry has not applied a-c interrupting ratings. Low-voltage M o t o r Starters. Low-voltage motor starters are of two types: those using fuses and those using air circuit breakers for shortcircuit protection. Those using air circuit breakers for short-circuit protection are applied 04 exactly the same basis as low-voltage air circuit breakers in so far as short-circuit currents are concerned. Motor starters using fuses for short-circuit protection are applied on exactly the same basisas fuses in so far as short-circuit current is concerned. AVAILABLE SHORT-CIRCUIT CURRENT

In determining the short-circuit current, the impedance of the circuit protective device connected in the faulty feeder is neglected. The shortcircuit current is determined by’ assuming that the protective device is shorted out by a bar of zero impedance (Fig. 1.27). The short-circuit /

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

37

current which flows in such a circuit is commonly called available shortcircuit cumat. The procedure for determining the available short-circuit current is based on setting up impedance or reactance diagrams. The impedance of the short-circuit protective device that is nearest the short circuit (electrically) is omitted from the impedance diagram. Practically all protective devices are so rated and tested for shortcircuit interrupting ability; hence this procedure may be followed in short-circuit calculations. This greatly simplifies the calculations and removes the effect of impedance variations between different types and makes of devices having the same interrupting rating. I t means that one set of short-circuit-current calculations for a given set of conditions is all that is needed for applying any type of protective device, regardless of the impedance of the devices themselves.

0

GENERATOR

TRANSFORMER

MOTORS CABLE

SHORT ClRCUlTED 8 1 J UMPER OF Z E R O IMPEDANCE

CABLE SHORT

CIRCUIT

FIG. 1.27 Connections

for determining available short-circuit current for testing rhort-

circuit protective devices.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

38

HOW TO MAKE A SHORT-CIRCUIT STUDY FOR DETERMINING SHORT-CIRCUIT CURRENT FORMULAS FOR SHORT-CIRCUIT STUDY'

1. Changing ohms to per cent ohms, etc.:

Per cent (%) ohms reactance

=

(90 ohms reactance

=

Per-unit

(ohms reactance) (kva.base) (1.1) (kvt)*(lO) (ohms reactance)(kva base) (kv)*(1000) (1.2)

[see Eq. (1.34)] Ohms reactance

=

( % reactance)(kv)2(10)

Per-unit ohms reactance

=

kva base per cent ohms reactance 100

(1.3) (1.4)

2. Changing per cent or per-unit ohms reactance from one kva base to another:

% ohms reactance on kva base 2 base 2 - kva X (% ohms reactance on base 1) (1.5) kva base 1 9f reactance on kva base 2 - kva base 2 X (% ohms reactance on kva base 1) (1.36) kva base 1 3. Converting utility-system reactance to per cent or per-unit ohms reactance on kva base being used in study: a. If given in per cent ohms reactance on a kva base different than that used in the study, convert according to Eq. (1.5). b. If given in short-circuit kva, convert to per-unit ohms thus:

kva base used in reactance diagram (1.6) short-circuit kva of utility system c. If given in short-circuit amperes (rms symmetrical), convert t o perunit ohms thus:

9i reactance

Yi reactance =

=

kva base used in reactance diagram (short-circuit current) ( d $ ) ( k v rating of system)

(1.7)

d. If only the kva interrupting rating of the incoming line breaker is known, * See pp. 54 to 57 for more prr-unit formulas

1 kv

= line-to-line kilovolts.

SHORTT-CIRCUIT.CURRENTCALCULATING PROCEDURES

39

9f ohms reactance -

kva base used in reactance diagram kva interrupting rating of incoming line breaker

(1.8)

4. Determining kva base of motors:

The exact kva base of a motor

=

EI 4 3

(1.9)

where E = name-plate voltage rating I = name-plate full-load current rating When motor full-load currents are not known, use the following kva bases: Induction motors: kva base = horsepower rating (1.10) 0.8-power factor synchronous motor: (1.11) kva base = 1.0 (horsepower rating) 1.0-power factor synchronous motors: (1.12) kva base = 0.80 (horsepower rating) 5. Changing voltage base when ohms are used: Ohms on basis of voltage 1 -

')* X (ohms on basis of voltage 2) (voltage 2)2

(1.13)

In Eqs. (1.1) to (1.4), ohms impedance or ohms resistance may be substituted for ohms reactance. The final product is then per-unit or per cent ohms impedance or resistance, respectively. 6 . Determining the symmetrical short-circuit kva: Symmetrical short-circuit kva

=

~

(kva base)

(1.14)

'& (kva base)

(1.15)

% X*

- y? -~

(line-to-neutral voltage)2 ohms reactance X 1000 kv2 X lo00 ohms reactance 7. Determining the symmetrical short-circuit current: (100) (kva base) Symmetrical short-circuit current = (% X*)(v%(kvt) kva base (% X*)(&)(kvt) k v t X lo00 ( d ) ( o h m s reactance) * X = reactance or impedanoe. t kv = line-&line kilovolts. = 3

(1.16)

(1.16a)

(1.17) (1.18) (1.19)

.

TABLE 1.3 Factor ( K ) to Convert Ohms to Per Cent or Per-unit Ohms for Three-phase Circuits*

0 L

Base kvo

loot Pr.

216Y/125 240 480

600 2,400 4.1 60

c*nt

1 50 Per-""it

'14 73 43.4

2.14 1.73 0.434

27.7 1.73 0.56

0.277 0.0173 0.00576

Per cent

321.5 260.4 65.21

200 Per-""it

Per cant

300

500

- __

Por-un1t

Per cent

Per-""it

Per cent

Per-""it

-0.712

3.215 2.604 0.6521

128 147 86.8

4.28 3.47 0.868

t43

4.166 2.604 0.808

0.4166 0.02604 0.00808

55.5 3.47 1.15

0.555 0.0347 0.0115

0.868 0.42 0.386

0.00868 0.0042 0.00386

1.302 0.63 0.579

0.01302 0.0063 0.00579

1.05 0.965

1.38 0.0868 0.0288 0.0217 0.0105 0.00965

30.2

6.43 5.21 1.302

071 868 217

83.3 5.21 1.72

0.833 0.0511 0.0172

I38 8.68 2.88

a1

8.68 2.17

4,800 6.900 7,200

0.435 0.210 0.193

0.00435 0.0021 0.001 93

0.651 0.289

0.00651 0.0031 5 0.00289

l1,OOO 11.500 12,000

0.0825 0.0755 0.0695

0.000825 0.000755 0.000695

0.123 0.113 0.104

0.00123 0.00113 0.00104

0.165 0.151 0.138

0.00165 0.00151 0.00138

0.247 0.226 0.208

0.00247 0.00226 0.00208

0.413 0.377 0.347

0.00413 0.00377 0.00347

12,500 13.200 13,800

0.064 0.0574 0.0525

0.00064 0.000574 0.000525

0.096 0.086 0.0787

0.00096 0.00086 0.000787

0.127 0.114 0.105

0.00127 0.00114 0.00105

0.192 0.172 0.157

0.001 92

0,00172 0.001 57

0.32 0.286 0.262

0.0032 0.00286 0.00262

23,000 37.4M) 46,000

0.0187 0.00711 0.00471

0,000187 0.000071 I 0.0000471

0.0283 0.0107 0.00708

0.000283 0.000106 0.0000708

0.0378 0.0142 0.00945

0.000378 0.000142 0.0000945

0.0567 0.0213 0.0141

0.000547 0.00021 3 0.0001 41

0.045 0.0355 0.0236

0.00045 0.000355 0.000236

0.0031 5

0.000031 5

0.0042

0.000042

0.0063

0.000063

0.0105

0.000105

69,OCU

0.0021 2 0.0000212 -

* For per-unit, K

=

0.315

-

kva base , kva base For per cent, K = kv' X 1wO kv' X 10

2.17

kv = line-to-line kilovolts

t To determine multiplying factors far any other base use figures under 100-kvs base columns multiplied by new base in kva, 100

v)

I

2 KE 2

B

-I

n

2 5 f

0

6c

R v,

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

41

8. Determining the asymmetrical short-circuit current: Asymmetrical short-circuit current = (symmetrical current) (multiplying factor) (1.20)

Asymmetrical short-circuit kva = (symmetrical kva) (multiplying factor) DIAGRAMS

One-line Diagram. The first step in making a short-circuit study is to prepare a one-line diagram showing all sources of short-circuit current, i.e., utility ties, generators, synchronous motors, induction motors, synchronous condensers, rotary converters, etc., and all significant circuit elements, such as transformers, cables, circuit breakers, etc. (Fig. 1.28). M a k e an Impedance or Reactance Diagram. The second step is to make an impedance or reactance diagram showing all significant reactances and resistances (Pig. 1.29). In the following pages this will be

I

C

GENERATOR

UTILITY SYSTEM TRANS

D

GENERATOR

CABLE E SHORT CIRCUIT LARGE MOTOR

CABLE J

480 VOLT MOTORS e diagram c

FIG. 1.28

, typical large industrial power system.

H

-SHORT

FIG. 1.29

INFINITE BUSES

CIRCUIT CURRENT GOES THROUGH HERE

Reactonce diagram of system shown in Fig. 1.28.

42

SHORT-ClRCUIT.CURRENT CALCULAltNG PROCEDURES

referred to as an impedance diagram, recognizing of course that only reactances will be used in many diagrams. The circuit element,s and machines considered in the impedance diagram depend upon many factors, i.e., circuit voltage, whether momentary or interrupting duty are to be checked, etc. The foregoing discussion and Table 1.2 explain when motors are to be considered and what motor reactances are to he used for checking the dut,y on a given circuit breaker or fuses of a given voltage class. There are other problems, i.e., (1) selecting the type and location of the short circuit in the system, (2) determining the specific reactance of a given circuit element or machine, and (3) deciding whether or not circuit resistance should be convidered. SELECTION OF TYPE AND LOCATION OF SHORT CIRCUIT

Three-phase Short Circuits Generally Considered. I n most industrial systems, the maximum short-circuit current is obtained when a three-phase short circuit occurs. Short-rircuit-current magnitudes are generally less for line-to-neutral or line-to-line short circuits than for the three-phase short circuits. Thus, the simple three-phase short-circuitcurrent calculations will suffice for application of short-circuit protective devices in most industrial systems. Unbalanced Short Circuits in Large Power Systems. In some very large systems where the high-voltage-system neutral is solidly grounded, maximum short-circuit current flows for a single phase-to-ground short rircuit. Such a system might be served from a large delta-Y transformer bank or directly from the plant generators. Hence the only time that single-phase short-circuit-current calculations need be made is on large high-voltage systems (2400 volts and above) with solidly grounded generator neutrals or where main transformers that supply a plant from a utility are ronnected in delta on the highvoltage side (incoming line) and in Y with solidly grounded neutrals on the low-voltage (load) side. The calculations of unbalanced short-circuit currents in large power systems can best be done by symmetrical components, see Chap. 2. Normally, generator and large delta-Y transformer secondaries are grounded through a reactor or resistor to limit the short-circuit current for a single line-to-ground short circuit on the system to letis than the value of short-circuit current for a three-phase short circuit. Bolted Short Circuits Only Are Considered. Several tests have been made to evaluate the effect of arc drop at the point of short circuit in reducing the short-circuit-current magnitude. It was felt by some engineers that the current-limiting effect of the arc was pronounced. These tests showed, however, that for circuit voltages as low as 300 volts

43

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

there may be no substantial difference in the current that flows for a bolted short circuit and when there is an arc of several inches of length. These test,s also confirmed modern calculating procedure as an accurate method of estimating the short-circuit-current magnitude in systems of 600 volts and less. .4rcs cannot be counted on to limit the flow of short-circuit currents even in louvoltage circuits; so short-circuit-current calculations for all circuit voltages are made on the basis of zero impedance at the point of short circuit, or, in other words, a bolted short circuit. This materially simplifies calculation because all other circuit impedances are linear in magnitude, whereas arcs have a nonlinear impedance characteristic. At What Point in the System Should the Short Circuit Be Considered to Occur? The maximum short-circuit current will flow through a cir-

cuit breaker, fuse, or motor starter when the short circuit occurs at the

4160V.

I

I

I

$?

$-

MAX.SHORT CIRCUIT DUTY ON BREAKERS ON THIS BUS $EW:RS FOR SHORT CIRCUIT

1 T ?;

A&?? Y T T - 3 &

r

+

y

* +

r-x

MAX. DUTY FOR THESE BREAKERS OCCURS FOR SHORT CIRCUIT HERE

FIG. 1.30

Location of faults for maximum Short-circuit duty on circuit breakers.

44

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

terminals of the circuit breaker, etc. (Fig. 1.30). These devices, if properly applied, should be capable of opening the maximum shortcircuit current that can flow through them. Therefore, only one shortcircuit location (at the terminal of the device) need be considered for checking the duty on a given circuit breaker, fuse, or motor starter. DETERMINING REACTANCES AND RESISTANCES OF CIRCUITS AND MACHINES

Typical reactances of circuit elements and machines are given at the end of this chapter. Resistances are included for certain items. These tables may be used as a basis for assigning values to the various elements of the impedance diagram. The reactances and resistances are all lineto-neutral values for one phase of a three-phase circuit. Where the reactances of a specific motor, generator, or transformer are known, these values should he used in lieu of the typical reactances in this chapter. The following is a guide to general practice in selecting and representing reactances. U s e R e a c t a n c e s of All S i g n i f i c a n t Circuit E l e m e n t s . Whether or not the reactance of a certain circuit element of a system is significant depends upon the voltage rating of the system where the short circuit occurs. In all cases, generator, motor, and transformer reactances are used. In systems rated above 600 volts, the reactances of short bus runs, current transformers, disconnecting switches, circuit breakers, and other circuit elements of only a few feet in length are so low that they may be neglected without significant error. In circuits rated 600 volts or less, the reactances of low-voltage current transformers, air circuit breakers, disconnecting switches, low-voltage bus runs, etc., may have a significant hearing on the magnitude of total shortcircuit current. As a general guide, the reactance of the low-voltage secondary-switchgear section in load-center unit substations with closely coupled transformers and secondary switchgear is not significant for all voltages of 600 volts and below. However, where there are several transformers or generators paralleled on one bus, or connections several feet long between a single transformer and its switchgear, reactances of the bus connections will generally be significant and should be considered in short-circuit calculations. I n systems of more than about 1000 kva on one bus a t 208Y/120 or 240 volts, reactance of all circuit components such as short bus runs, current transformers, circuit breakers, etc., should be included in the short-circuit study. I n systems of more than about 3000 kva on one bus a t 480 volts or 600 volts, reactances of all components such as current transformers, circuit breakers, short bus runs, etc., should be considered. It should be remembered that the lower the voltage, the more effective

SHORT-CIRCUIT-CURRENl CALCUUTING PROCEDURES

45

a small impedance is in limiting the short-circuit-current magnitude. That is why extreme care should he used to include all circuit elements in the impedance diagram, particularly for large ZORY/lZO-volt or 240-volt systems. If care is not used, the calculations will result in a value of current far higher than will actually be realized in practice. See the example outlined in Figs. 1.46 and 1.47. This often results in the adoption of low-voltage switchgear of higher interrupting rating and higher cost than are actually required. If care is used in including all reactances, the calculated reiults will be close to the short-circuit currents obtained in practice. Short-circuit calculations are of most value if they reflect accurate answers. When Is Resistance Considered? The resistance of all generators, transformers, reactors, motors, and high-capacity buses (above about 1000-amp rating) is so low, compared with their reactance, that their resistance is not considered, regardless of their voltage rating. The resistance of all other circuit elements of the high-voltage system (above 600 volts) is usually neglected, because the resistance of these parts has no significant bearing on the total magnitude of short-circuit currents. In systems of 600 volts and less the error of omitting resistances of all parts of the circuit except cables and small ampere rating buses is usually less than 5 per cent. However, the resistance of cable circuits is often the predominant part of the total impedance of a cable. When appreciable lengths of cable are involved in the circuit through which short-circuit current flows in a system of GOO volts or less, the resistance as well as the reactance of the cable circuits should be included in the GENERATOR

OF-THESE CIRCUIT ELEMENTS. IN GENERAL USF REACTANCE AND RESISTANCE OF THESE

SHORT CIRCUIT CURRENT CONSIDERING REACTANCE ONLY :20800 AMPERES

___

-. . . -. 1100 FT. 101

---(20 FT

SHORT CIRCUIT CURRENT CONSIDERING REACTANCE OF A LL PARTS PLUS RESISTANCE OF COW VOLTAGE CABLE = 11500 4MPERES.

FIG. 1.31 One-line diagram showing effect of resistance in cable circuits.

46

SHORT-CIRCUIT.CURRENT CALCULATING PROCEDURES

impedance diagram. The example of Fig. 1.31 shows the error that might result in neglecting cable resistance. I n secondary network systems of 600 volts and less, the resistance as well as the reactance of the tie-cable circuits between substation buses should be included in the impedance diagram. The example of Fig. 1.32 shows the effect of cable resistance in reducing short-circuit current in a typical industrial network.

n n SHORT CIRCUIT CURRENT USING REACTANCE ONLY = 51000 AMPERES, SHORT CIRCUIT CURRENT USING REACTANCE PLUS RESISTANCE OF T I E CIRCUIT= 41000 AMPERES.

T I E CIRCUITS 208 Y / l Z O V O L T S .

200 FT 2- 250 M,CM 3 CONO. CABLES ~~~~~T I N PARALLEL

200 F T

FIG. 1.32

One-line diogrtlm of low-voltage secondary network system showing effect of resistance of cable tie circuits.

Where to Use Exact Multiplying Factors. I n low-voltage systems having considerable lengths of cahle, the X / R ratio may be so low that the 1.25 multiplying factor would be considerably in error. Hence in these systems where resistance is considered, determine the correct X / R ratio and then use minimum multiplying factor. GUIDE FOR REPRESENTING THE REACTANCE O F A GROUP O F MOTORS

A group of motors fed from one substation or from one generating station bus may range in rating from fractional to several thousand horsepower per motor. All motors that are running at the time a short circuit occurs in the power system contribute short-circuit current and therefore should be taken into consideration. Motors Roted 600 Volts and Below. I n that portion of the power system operating at 600 volts or less, there are generally numerous small

A?

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

motors, i.e., under about 50 hp. I t becomes impractical to represent each small motor in the impedance diagram. These motors are constantly being turned off and on; so it is practically impossible to predict which ones will be on the line when a short circuit occurs. Furthermore, it would be impractical to obtain the characteristics of each small motor and to account for the effect of the impedance of their leads. Where more accurate data are not available, the following procedure may be used with satisfactory results for representing the combined reactance of a group of miscellaneous motors operating a t 600 volts or less. 1. In systems rated 240, 480, or 600 volts a t each generator and/or transformer bus, assume that the maximum horsepower of motors runniug a t any one time is equal to the combined kva rating of the stepdown transformer and/or generators supplying that one bus (see Figs. 1.33 and 1.34). 2. 10 systems rated 208Y/120 volts, a substantial portion of the load usually consists of lights and a lesser proportion of motor load than in 240-, 480-, or 000-volt systems. Hence in 208Y/120-volt systems where more accurate data are not available, assume a t each generator and/or transformer bus that the maximum horsepower of motors running a t

TO UTILITY SYSTEM

REbCTbNCE OF UTILITY SYSTEM REbCTbNCE OF 7 5 0 K V b TRbNSF.

QOW, OR5.,s

5.5%

0.25% OR

25 % REbCTbNCE OF EQUIVALENT MOTOR

IMPEObNCE O I b G R b M 750 K V b BASE SHORT CIRCUIT

EQUIVALENT MOTOR 750 KVb

240, 480, 600 VOLT SYSTEMS

El hKVA TO UTILITY SYSTEM

REbCTbNCE OF UTILITY SYSTEM

0.50% OR

50 % REACTbNCE OF EQUIVALENT MOTOR

SHORT CIRCUIT

FIG. 1.33

REbCTbNCE OF 7 5 0 KVb TRbNSF. EQUIVILENT MOTOR 375 K V b IMPEObNCE OIbGRbM 750 K V b BASE 2 0 8 Y / 1 2 0 VOLT SYSTEMS

Oiagromr illustrating how to include motors in low-voltage radial systems.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

40

any one time is equal t,o 50 per cent of the combined rating of all stepdown trausformers and/or generators supplying power to that one bus, Fig. 1.33. For large commercial buildings the 50 per cent figure may be too low. Check carefully the mot,or load on all large 208Y/120-volt systems. I n the generalized rases referred t o in paragraphs 1 and 2 , no specific ratio of induction t o synchronous motors or no specific number of motors which prcduce unusually high short-circuit current,s has been set fort,h. T o account for these variables, a n average motor reitctance ihcluding leads is assumed t o be 25 per cent for the purpose of preparing application tables like Table 1.5 and in making short-circuit st,udies where no more accurat,e data are available. It will he noted that the average motor reactance of 25 per cent is based on the transformer or supply-generator kva rating. This figure is between the values of 28 per cent for induction mot,ors and 21 per cent for synchronous motors given in Table 1.14. Where the division between synchronous and iuduction motors is known, then more accurate calculations can be made by using the assumed motor reactances of Table 1.14. T h e reactances given in Table 1.14 are based on motor kva ratings and not supply transformer or generator ratings.

750 KVA

A 500 KVA

T

750 KVA

-480 VOLTS 500 KVA v

EQUIVALENT MOTORS WOULD BE 250 KVA AND 375 K VA FOR 280Y/120 VOLT SECONDARY SYSTEM

FIG. 1.34 rvrternr.

Diagram illustrating how lo include motors in lowvoltage secondary network

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

49

Although a portion of the load connected to a bus rated GOO voks or less may be heaters, lights, a-c welders, solderitig irons, appliances, arid other devices which produce no short-circuit curreiit, the total installed horsepower of motors connected t,o such a bus is geiierally much greater than the kva rating of the supply transformers and generators. Hovever, allowing for diversity, generally the total comhitied horsepower rat,ing of all mot,ors running a t one time ix-ould trot produce short-circuit currents in excess of the values obtained when using the ahore assumptions. I n systems of 000 volts or Icss, the large motors (i,e., mot,ors 011 t,he order of several hundred horsepomerj are usually few i n number and represent only a small portion of the tot,al connected horsepower; therefore, these larger motors are generally lumped in with the smaller motors and the complete group is represented as one equivalent motor i t i the impedance diagram. Synchrouous and induction motors need not be segregated when combining the motors in these low-voltage systems, because lorn-voltage air circuit breakers operr so fast that only the current flow duritig the first half cycle is considered; i.e., only suhtraiisient reactances ( X y ) of marhiiies are considered. Motors Rated above 600 Volts. High-voltage motors (rated 2200 volts and ahove) are generally larger in horsepower rating thau motors on systems operating under 600 volts. These largcr motors may have a much more significant hearing on short-circuit-current magnitudes than smaller motors, and, therefore, more exact determinatiou of the reactances of the larger motors is in order. Therefore, it is often foutid convetiient t o represent each large high-voltage motor individually in the impedance diagram. However, in large plants like steel mills, paper mills, etc., where there are numerous motors of several huridred horsepower each, it is often found desirable t o group these larger motors iii one group arid represent them by one reartaiire in the impedance diagram. Individual motors of several thousand horsepoitrer should be coiisidered individually and their reactances accurately determined hefore starting the short-circuit study. Whether considering motors individually or in groups, regardless of voltage rating of the motors, it is necessary t o obtain an equivalent kva rating of the individual or group of motors. This can be done precisely for large motors by Eq. (1.9) or can be approximated hy Eq. (l.lO), ( l , l l ) , or (1.12), when the full-load current is not known. The latter equations are used when considering a single reactance t o represent a group of miscellaneous motors.

50

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

I n high-voltage systems, complete motor data may not be available. Lacking these data, the connected horsepower is assumed to he equal t o the generator and/or transformer capacity supplying a given highvoltage bus. If the reactance of the leads between the transformer and/or generator bus and the motors is significant, the reactanre of these leads should be included. MAKING THE IMPEDANCE DIAGRAM

After it has been decided what elements of the one-line diagram are to be considered in the impedance diagram, the mechanirs of making the impedance diagram and of determining the short-circuit-current magnitude are as follows.

7

GENERATOR OR MOTOR OF ZERO IMPEDANCE

Treatment of Sources of Short-circuit Current. The generators and motors

are treated as if they comprised a generator of zero reactance plus an external reactor to represent the reactance of the EXTERNAL TO machine windings, Fig. 1.35. The first REPRESENT IMPEDANCE OF step in making an impedance diagram GENERATOR OR MOTOR. is torepresent every generator and motor or groups of motors and utility supply FIG. 1.35 One-line representation by a reactance connected to a zero imof generator or motor in impedance pedance bus or so-called “infinite bus,” diogmm. Fig. 1.36. This bus represents the internal voltage of the generators and motors. Completing the Impedance Diagram. The second step is to add the reactance of cables, buses, transformers, current transformers, circuit

flG. 1.36 Representation of reactances of generators, motors, and utility supply of system shown in one-line diagram form in Fig. 1.28.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

51

breakers, switches, etc., in their proper location to complete the impedance diagram, top of Fig. 1.37. Choice of Ohms, Per Cent Ohms, or Per-unit Ohms Method. The next step is to decide whether to use ohms, per cent ohms, or per-unit ohms to represent the various circuit impedances in the impedance diagram. INFINITE BUS

SHORT CIRCUIT

INFINITE BUS

6.04V

STEP NO i COMBINE SERIES REACTANCES H

C+D=0.04+0.15~0.19% = 2 . O t ~0.0+0.10~ 12.10%

+1+J

STEP NO.:!

COMBINE PARALLEL REACTANCES

+

J) F,G AND I H + I = _' + L + I XI F G H + I + J

'

I

I

- _ 3 + p-j=j 2 . 5 + 0 . 2 + 0 . 0 8 3 -o,'40t

I

-=

XI

2.783 X =0.3698

STEP N 0 . 3 COMBINE SERIES REACTANCES X,,AND

X t = XI

STEP NO. 4

E

+ E = 0.36+0.04'0.40%

COMBINE PARALLEL REACTANCES X o , A . B . AND IC+D) 1 l +i

- 1 + 1 + L+-

I XR

XI A

B

C+D

'0.40 0 2 5

+I I+ 2.0

0.19

2 . 5 + 4 + 0 . 5 +5.3=12.3 X,

RESULTANT SINGLE REACTANCE

X

I ~

0.0805 % O ~ z

FIG. 1.37 Complete reaclomce diagram for system shown in Fig. 1.28. bining reactances into o single resultant value.

Steps for com-

52

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

Ohms are generally not used because of the difficulty of converting ohms from one voltage base to another without error and because of the very small numbers, which make accurate and easy calculation more difficult than the per cent or per-unit system. In many of the examples in this book, the assumed or given impedance or reactance data are listed in per cent, hut in the reactance dia,-rams these are converted to per-unit. N o notation will he made when that is done as it will be obvious. Equations (1.1) to (1.4) show how to convert ohms to per cent ohms, ohms to per-unit ohms. The Per-unit System for Electrical Calculations.* A per-unit system is a means of expressing numbers for ease in comparing them. A per-unit value is a ratio: a number (1.21) Per-unit = base number ~~

The base number is also called unit value since in the per-unit system it has a value of 1, or unity. Thus, base voltage is also called unit voltage. Any convenient number may be selected for the base number. For example, for the columns below, a base of 560 is used: Number 93 125 560 2053

Per-unit Volue with 560 as a Base

0.17 0.22 1 .oo 3.65

Each number in the second column is a per-unit part of the base number. In the first column, to compare the numbers, first mentally determine the ratio of one to the other. In the second column this is already accomplished. The comparison can be aided by selection of the base number which will illustrate the comparison best. In the foregoing example, if it is desired to show how much larger each uumber is when compared with the smallest number, the number 93 might have been selected as the base. This would then be obtained as follows: Per-""it Valve Number

with 93

( I ,

(I

93

I .oo

125 560 2053

I.35 6.00 22.20

Base

The value of a per-unit system is particularly useful when comparing

* From material originally Company.

prepared by H. J. Finison. iormrrly of General Ekctrir

SHORT-CIRCUIT.CURRENT CALCULATING PROCEDURES

53

numbers that are similarly related to two different base numbers. example : Core A 2300

Norm01 "0th Volts during motor starting

2020

For

Cole B 460 420

The above figures in themselves have little significance until they are compared each with its normal condition as follows: Vollr during starting per-unit of normal

0.91

0.88

Per Cent. Obviously per cent and per-unit systems are similar. The per cent system is obtained by multiplying the per-unit value arbitrarily by 100 to keep many frequently used per-unit values expressed as whole integers. By definition,

Per cent =

a number base number

x

100

(1.22)

Thus to change per cent to per-unit, divide by 100. For example, a transformer which has an impedance of 6 per cent has an impedance of 0.06 per-unit. The per cent system is somewhat more difficult to work with and more subject to possible error since it must always be remembered that the numbers have been arbitrarily multiplied by 100. For a simple example, money may draw interest a t the rate of 4 per cent per year. Early in arithmetic one learns to determine the interest by multiplying the principal by 0.04. It is thus necessary to remember to convert to the per-unit value before using the figure. In a complex calculation, this repeated conversion may invite errors. In effect it is safer and more convenient to say that interest is a t the rate of 0.04 per-unit. Impedances of electric apparatus are usually given in per cent. I t is usually convenient to convert these figures immediately to per-unit by dividing by 100 and thereafter do all calculating in terms of per-unit rather than attempt to remember always during the calculations whether a number should or should not be multiplied or divided by 100 to obtain the true value. Symbol. Just as the per cent system has a symbol (%) to designate that a given number is expressed in terms of per cent (as 6%) so also does the per-unit system have a symbol. The symbol for per-unit is (%). Thus 0.06 per-unit is written as 0.06 91. Selection of Base Number. In a per-unit system as used for expressing electrical quantities of voltage, current, and impedance, it is necessary to select numbers arbitrarily for the following: Base volts Base amperes

54

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

Do not then in addition arbitrarily select base ohms since it has already been fixed by the first two selections because of Ohm’s law. z = -E

I

base volts (1.23) base a m p z s Using the selected base values, all parts of an electric circuit or system may be expressed in per-unit terms as follows: volts Per-unit volts = (1.24) base volts amperes Per-unit amperes = (1.25) base amperes ohms Per-unit ohms = (1.26) base ohms In practice it is more convenient to select: Base volts Base kva The base values of other quant.ities are thus automatically fixed. Hence, for a single-phase system, base kva X 1000 Base amperes = (1.27) base volts base kva Base amperes = (1.28) base kv base volts Base ohms = (1.23) base amperes where base kva is single-phase kva and base volts is single-phase volts. For a three-phase system: base kva X 1000 Bme amperes = (1.29) X base voks base kva Base amperes = (1.30) 4 X base kv hase volts Base ohms = (1.31) X base amperes where base kva is three-phase kva, base volts is line-to-line, and hase ohms is per phase. Per-unit Ohms. In practice i t is desirable to convert directly from ohms to per-unit ohms, without first determining base ohms. By Ohm’s law, base volts Base ohms = (1.23) base amperes Base ohms

=

4

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

55

Substitute Eq. (1.27) (which gives the base amperes) into Eq. (1.23), to obtain base volts Base ohms = (base kva X 1000)/base volts (base volts)P Base ohms = bsse (1.32) kva x 1000 By definition: ohms Per-unit ohms = (1.26) base ohms Substitute Eq. (1.32) into Eq. (1.26) to obtain ohms Per-unit ohms = (base volts)e/(base kva X 1000) ohms X base kva X 1000 Per-unit ohms = (1.33) (base voltd2 ohms X base kva Per-unit ohms = (1.34) (base kv)2 X 1000 where base kva is single-phase kva and base kv is single-phase kv. When dealing with a three-phase system, i t is usual to select three-phase kva and line-to-line volts for the base values. Convert the above expressions to these bases to obtain ohms X base kva X 1000 X 3 Per-unit ohms = (base volts X d .3 ,)z ohms'X base kva X 1000 Per-unit ohms = (base volts)2 ohms X base kva Per-unit ohms = (1.35) (base kv)* X 1000 where ohms are per phase, kva is three-phase kva, and kv is line-to-line voltage. Usual Base Numbers for System Studies. If per cent or per-unit ohms reactance is used, the next step is to choose a kva base. In system studies it is usually desirable to select as the base voltage the nominal-system voltage or the voltage rating of the generators and supply transformers. Base kva will usually be selected as the kva rating of one of the machines or transformers in the system, or a convenient round number such as 1000, 10,000, or 100,OOO kva. After choosing the kva base, convert ohmic reactance of cables, wires, current transformers, etc., to per cent or per-unit ohms reactance on the chosen base, using Eq. (1.1) or (1.2) or Table 1.3. If ohms reactance is used, convert all per cent reactances to ohms by Eq. (1.3). Where two systems of differing voltage are interconnected through a

56

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

transformer, select a common kva base for both systems and the rated voltage of each system as its own base voltage. (These base voltages must have the same ratio t o each other as the turn ratio of the transformer connecting the two systems.) Base ohms and base amperes for the two systems will thus he correspondingly different. Figure 1.38 shows a typical example. Once the system values are expressed as per-unit values, the two interconnected systems may be treated as a single system and any calculations necessary carried out. Only in reconverting the per-unit values of the results to actual voltage and current values is i t necessary t o remember t h a t two different voltages actually existed in the system. Change of Base Number. Frequently the impedance of a circuit element may be expressed in terms of a particuiar base kva, and it may be desirable t o express it in terms of a different base kva. For example, the reactance of devices like transformers, generators, and motors is given in per cent on their own kva rating, and their reactances must be converted to the common base, chosen for the study by means of Eq. (1.5) or (1.36). Per-unit ohms on kva base 2 - base kva base kva 1

x

(per-unit ohms on kva base 1) (1.36)

Similarly, a machine rated a t one voltage may actually be used i n a circuit a t a different voltage. Its per-unit impedance must thus be changed to a new base voltage. GENERATOR 1000 KVA

MOTOR o(lOOO KVA)

I0;YKVA

13800 VOLTS

2300 VOLTS PRIMARY RATING

SECONDARY RATING

13200 VOLTS

2400 VOLTS

TRANSFORMER RATIO= 13 200/2400=5.5 (A)HIGH VOLTAGE SYSTEM

13 800

FIG. 1.38

BASE VOLTS

2500

BASE KVA

(A1 (El 5.5

RATIO

-

1000

I .o

41.6

EASE AMPS

233

115 5

190

BASE OHMS

6.2 5

I000

onother.

( 8 )LOW VOLTAGE SYSTEM

(5.5?

Method of converting bore volts, kva, amperes, and ohms from one value to

n Reference to Eq. (1.35) shows that per-unit ohms is inversely proportional to the square of base volts. Thus: Per-unit ohms on new base volts - (old base volt.s)* (1.37) Per-unit ohms on old base volts (new base volts)* and Per-unit ohms on new base volts = per-unit ohms on old base volts (old base volts)2 (1.38) (new base volts)2 Equations (1.37) and (1.38) may be used for per cent ohms as well as perunit ohms. Converting Ohms to a Common Voltage Base. When using ohms instead of per cent or per-unit in the impedance diagram, it is important to convert the ohmic values to a common voltage base by Eq. (1.13). For example, if the short-circuit current is being calculated in a 480-volt system (supplied by transformers rated 480-volt secondary) fed through a cable and a transformer from a 2400-volt system, the ohms impedance of the cable in the 2400-volt circuit must be multiplied by 48O2/24OO2to convert it to ohms on a 480-volt base. The transformer ratings, i.e., 480, 240, etc., and not system ratings, if different from transformer rating, are used as the voltage base for short-circuit-current calculations. Representing the Utility Supply System. The utility system must be represented by a reactance in the impedance diagram. Sometimes this utility-system reactance is available in per cent on a certain base. If so, it is merely necessary to convert this value to the common base used in the impedance diagram. To do this, use Eq. (1.5). In some cases the utility engineers will give the short-circuit kva or current that the utility system will deliver a t the plant site. In otker cases, only the interrupting capacity of the incoming-line circuit breaker is known. In these cases to convert short-circuit kva, current, or incoming-line breaker interrupting rating to per cent reactance on the kva base used in the reactance diagram, proceed as follows: If given short-circuit kva, convert to per cent by using Eq. (1.6). If per-unit is desired, use also Eq. (1.4). If given short-circuit amperes (rms symmetrical), convert to per cent by Eq. (1.7) and to per-unit by Eqs. (1.7) and (1.4). If only the kva interrupting rating of the incoming line circuit breaker is known, convert to per cent by Eq. (1.8) and to per-unit by Eqs. (1.8) and (1.4). SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

DETERMINING THE EQUIVALENT SYSTEM IMPEDANCE OR REACTANCE

After completing the impedance diagram and inserting the values of reactance or impedance for each part of the diagram, it is necessary to reduce this network to one equivalent value. This can be done either by

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

58

longhand calculation or with the aid of a calculating board. Since so few engineers have access to calculating hoards and must use longhand methods, this method will be covered in sufficient detail to enable solving the short-circuit problems commonly encountered. Use of Calculating Boards. A d-c calculating board will permit accurate solution of all short-circuit problems where reactance only is considered. In most cases where resistance is a significant factor and must be considered, the d-c calculating board cannot be used readily. However, in some problems involving resistance, certain approximations can be made to obtain reasonably accurate answers on d-c calculating boards. For exact calculating-board solutions of problems factoring resistance and reactance, the a-c calculating board may he employed. A-c calculating boards have boxes to represent both the resistance and reactance of a circuit. The procedure for using calculating boards is beyond the scope of this book. Longhand Method of Combining Reactances. Longhand methods of combining reactances vary in some respects. To illustrate the principles involved, refer to Figs. 1.37 and 1.39. Arbitrary values of reactance have been assigned to the various branches. Combining the various branches of the diagram is merely a question of reducing two or more series reactances to one value and reducing two or more parallel reactances to one value until one single equivalent value is obtained. The following shows how to combine reactances and resistances. 1. Combining reactance and resistance to determine impedance, z

=

m

(1.39)

z=r+jz

wherej = 47 2. Adding series reactance of circuits where resistance is neglected add reactances arithmetically, i.e.,

x,

+ x2 + xa = x.

= equivalent reactance z,,z2, and x 3 = reactances of circuit components zs= equivalent reactance

3. Combining parallel reactances, zo = equivalent reactance

For two reactances only x, and z2 XI =

(d(z2) 22

21

+

For combining several parallel reactances 1 1 1 1 1 -=_ 2.

2,

+ - + - +X -I + E 2, 2 2

(1.40) 1

(1.41)

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

59

INFINITE

c. REACTANCE DIAGRAM OF CIRCUIT SHOWN IN ONE LINE DIAGRAM TO THE LEFT.

ONE-LINE DIAGRAW

P~T&

$*,

T(

* T . Pm

EQUIVALENT Y CONVERT P I T I , PITI e c , TO EQUIVALENT Y. STEP x z

c. STEP# I COMBINE SERIES REACTANCES PI~TI,RBT~,ETC.

I

a"

--&&Pa.

L

3+c*

Ct

c4

cs

DRAW NEW DIAGRAM STEP-* 3

+

COMBINE 2 C t , 3 + C+ AND THEN REPEAT STEPS 2.3 e 4 UNTIL ONE EOUIVALENT REACTANCE IS OBTAINED. STEP t t 4

FIG. 1.39 Example of the method of combining remtmces of a network-type system into a single resultant value.

MI

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

Some systems are such that they cannot he reduced by merely combining series and parallel rgactances. For example, take the one-line diagram of a circuit as show in upper left-hand corner of Fig. 1.39. The reactance diagram is shown the ypper righehand corner of Fig. 1.39. In addition to combining serieszind parallel reactances, it is necessary to TI, PzrT , and C1to an equivaconvert a triangle of reactances such as PI, lent Y of reactances by the formulas of Fig. 1.40. By these conversions,

\

I

B=

c=

ob

+ a c + be

a=-

b

a b + a c + bc

b:

A = ob+oc+bc

C:

a

0c A+B+C

"

A+B+C A8 A+B+C

FIG. 1.40 Formula for converting a triangle or delta of three impedances to a Y of three equivalent impedances, and vice verso.

any commonly encountered system reactance diagram can be reduced to one equivalent reactance. Combining Impedances. Sometimes i t is desirable to consider the resistance and reactance of a circuit. This involves combining impedances. The procedure for combining impedances is outlined here. The combining of parallel impedances necessitates multiplication and division of impedances (complex quantities) and is outlined here. Adding Series Impedances. When two or more impedances are in series, the resistance and reactance components are added separately to combine the series into one equivalent value. Refer to Fig. 1.41. The three series impedances are

+ + ja

z1 = TI jzl za = 72 ijxa

zz = Tp

61

SHORT-ClRCUIT+CURRENT CALCULATING PROCEDURES

EQUIVALENT IMPEDANCE

3 SERIES IMPEDANCES

FIG. 1.41

Example illustrating the combining

The equivalent impedance

+ + +

rl VZ 73 j(z1 Using the numerical values of Fig. 1.41, 2 %=

2,

1+ j 2

22

2 +j3 0.5 + j l

= = = 21 =

(1

of series impedances.

+ zz+ 4

(1.42)

+ 2 + 0.5) + j ( 2 + 3 + 1) = 3.5 + j G

The above is applicable when impedances are expressed in ohms, perunit or per cent. Combining Parallel Impedances. Parallel impedances may be reduced to one equivalent impedance as follows (see Fig. 1.42):

TWO PARALLEL IMPEDANCES

FIG. 1.42

Example illustrating L e combining

EQUIVALENT IMPEDANCE

of parallel impedances.

61

WORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

(1) Reduce the per cent values of resistance and reactance in each of the given parallel circuits to a per-unit basis by dividing per cent figures by 100 or convert the per cent values to ohms. Per cent values can be used in the following if the multiplier 100 is applied properly, e.g., X

T

(Branch 1) 0.05 (Branch 2) 0.008

0.15 0.108

(2) Calculate the impedance squared z2 of each circuit 2% =

(Branch 1) rlz (Branch 2) r 2

r'

+

2 '

1

+ = ZI', e.g., 0 .052+ 0.1547-0>25 + zz2= zz2, e.g., 0.008z+ 0.108* = 0.0117 21'

(3) Obtain the ratios of r/z' of each circuit Tl 0.05 (Branch 1) -', e.g., -= 2.0 21 0.025 rz (Branch 2) -, e.g., 0.0°8 - 0.683 z'2 0.0117 ~

(4) Add the foregoing r / z z = Ga = 2.683

(5) Obtain the ratios of x/z* for each circuit 21 0.15 (Branch 1) -2 e.g., -= 6 21 0.025 XP 0.108 (Branch 2) e.g., -= 9.2 22 0.0117 (6) Add the foregoing 7 j

X/L'

(7) Ya2 = 02

(8) ra

=

=

Ba

=

15.2

+ Ba2,e.g., = 2.683' + 15.24 = 238.2

Ga e.g.,

-9

Y3'

BJ e.g.,

;2

=

~

- 0.0112

15" = 0.0642 __ 238.2 The foregoing may be tabulated for convenience in solving a number of parallel pairs of circuits: r z z4 = r' z2 r/z' 2/22 (Branch 1) 00 0 0 0 (Branch 2) 00 0 0 0 (Branch 3, etc.) ( ) ( ) ( ) ()() By addition- Go( )Bo( )

(9) xa

=

-2

Ya2

=

+

The combination of the circuits results in

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

63

1

Any number of parallel circuits may be accommodated by additional horizontal columns as fo branch 1 and branch 2, etc., their resultant (r/z2)’s and (x/z2)’s heling added to obtain G O and Bo. Multiplying and Dividing Impedances. Two impedmces may be multiplied as per the following equations: (21) 21

(22)

=

=

23

TI +jXl

ZP = T S

+j x , + +

= r8 jxa 2 3 = (TI jXl)(Tt = (TIT2 - 2 1 2 2 ) 13 = (nrz - XIXZ) j a = j(r1zz rczJ 23

+

+ +

jZ2) j(TIX2

1

+

TBZL)

(1.44)

Two impedances may be divided according to the following equations:

TI +j x , =-x-

r2

+j x 2

TZ TZ

- jxt

- jxt (1.45)

DETERMINING THE SHORT-CIRCUIT-CURRENT MAGNITUDE

After the reactance diagram has been reduced to a single value, the value of symmetrical short-circuit kva can be determined by Eq. (1.14), (1.15), or (1.16). To determine the symmetrical short-circuit current, use Eq. (1.17), ( l . l S ) , or (1.19). Equations (1.14) to (1.19) do not allow for any d-c component. Table 1.4 gives figures for converting kva to amperes. Apply Proper Multiplying Factor. The final step is to apply the proper multiplying factor from Table 1.2. To determine the total rms short-circuit current or kva, use Eq. (1.20).

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

64

TABLE 1.4 Three phore line-to-line, volh

Amperes per Kva Amperes rer phase per kvo

Amperes ier phase per kva

"0 wire

V d h

c or d-l

Amperes er kro or d-c kw

110 115 120

5.25 5.02 4.81

13.200 13,800 14,400

0.0437 0.0419 0.0401

24 48 110

180 I99 208

3.21 2.90 2.78

22,000 23,000 24,000

0.0263 0.0251 0.0241

115 120 125

8.70 8.33 8.00

220 230 240

2.63 2.51 2.41

33,000 34,500 36,000

0.0175 0.0167 0.0160

220 230 240

4.55 4.35 4.17

440 460 480

1.31 I .25

1.20

44,000 46,000 48,000

0.0131 0.0125 0.0120

250 275 300

4.00 3.64 3.33

550 575 600

1.05 1 .oo 0.962

66,000 69,000 72,000

0.00875 0.00838 0.00803

440 460 480

2.27 2.17 2.08

1,100 1,150 1,200

0.525 0.502 0.481

I I0.000 120,000

0.00525 0.00502 0.00481

550 575 600

I .82

1 I5.000

2,200 2,300 2,400

0.263 0.251 0.241

132,000 138,000 144,000

0.00437 0.0041 9 0.00401

650 750 1,200

1.54 1.33 0.833

3.300 3.450 3,600

0.175 0.167 0.160

154,000 161,000 168,000

0.00375 0.00359 0.00344

1,500 2,200 2,300

0.666 0.455 0.435

3,800 4,000 4.160

0.152 0.144 0.138

220,000 230,000 240,000

0.00263 0.00251 0.00241

2,400 3,000

0.417 0.333

6,600 6.900 7.200

0.0875 0.0838 0.0803

330,000 345,000 360,000

0.00175 0.00167 0.00160

11,000 11,500 12,000

0.0525 0.0502 0.0481

-

41.7 20.8 9.10

I .74 1.67

~

/

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

65

EQUIVALENT CIRCUITS

The redurtion of impedance diagrams to a single value of impedance can he greatly simplified by using equivalent circuits for duplex reactors and three-winding transformers. Equivalent Circuit for Duplex Reactors. The duplex reactor consists of two sections of winding per phase on the same core, with a t a p brought out from the junction point. The current ratings and reactances of the two sections arc generally equal. Aside from the midtap connections, whirh necessitate a total of nine leads, the construction is similar to that of the series reactor. If 1, and l2 are the self-inductances ( X , and X , are the corresponding reactances) of the individual sections, and f c is the “coupling factor” of the mutual inductance betmeen sections, then the simplified equivalent

LEX REACTOR

O N E LINE DlAGRPlM I

J

k

GENERATOR

-XI

FIG. 1.43

fc

One-line diagram and equivalent circuit for duplex reactor.

66

SHORT-CIRCUIT-CURRENT CALCUUllNG PROCEDURES

circuit for the duplex reactor is as shown in Fig. 1.43. For preliminary calculations, an average figure off. = 0.5 should give results of sufficient accuracy. Equivalent Circuit of Three-winding Transformer. When making short-circuit calculations of power systems which include three-winding transformers, there is a question on how to use the designer's reactance values. Designers give reactance values between pairs of windings. Figure 1.44A shows a three-winding transformer, and Fig. 1.44B shows its equivalent circuit. The following equations are easily derived and are the proper ones to use in short-circuit studies:

x. = x,. + 2 xs = + X2e c x,= + XAC

XIB

XBC

XdC

XBC

XAC

(1.46)

XdB

2

All reactance6 must be on same kva base. NOTE:The equivalent circuit and equations for a four-winding transformer are more complicated and will not he evident by simple analogy from Eq. (1.46).

,

(A1

FIG. 1.44 (A1

A

mi

One-line diagram and (61 equivalent circuit diagram of three-winding

transformer.

EXAMPLES OF SHORT-CIRCUIT-CURRENT CALCULATIONS'

The following examples are indicative of methods of applying the shortcircuit-current calculating procedures outlined in the foregoing. Systems 600 Volts and Below. The system shown in Fig. 1.45 involves one source of supply through a transformer from a primary system. The kva base for the short-circuit calculations is taken as the kva *NOTE:Numbers in parentheses in Figs. 1.45 and 1.47 to 1.50 refer to numbers of formulas used.

67

SHORT-CIRCUIT-CURRENT CALCUVITING PROCEDURES

INCOMING LINE

A

A

0.25 Yt

SOURCE

MOTORS

TRANSFORMER

I

750 KVA 5.5 x x (0.055%)

? T ? ?

480 VOLTS

REACTANCE DIAGRAM USE 750 KVA BASE FOR CALCULATIONS

M$ (0)

SOURCE REACTANCE ON 750 KVA BASE

0.0625 1

2

5

(1.61

1x=--XIXI% +x2-0.0625t025 0.0625XC125-0,05% T

v

750 - yj;xo,4& I

loo,ooo 750 - 0.0075%

:

5

%

(d)

18,000 AMPERES SYMMETRICAL [ 1.18)

X

o,050 18,000 X 1.25"22.500 AMPERES ASYMMETRICAL (1.201

(el FIG.1.45 Illustration of procedure for calculation of short-circuit currents in radial loadcenter system.

68

SHORT-CIRCUIT-CURREHI CALCULATING PROCEDURES

rating of the transformer. The kva of the connected motors is assumed to be 750 with an equivalent reactance of 25 per cent. Only reactances are used in these calculations. This problem is the type on which Table 1.5 is based. Large 208Y/120-volt Systems. Problems, particularly those involving secondary-network systems in the downtown area of the large cities or in large buildings, require the determination of the short-circuit current on a 208Y/120-volt basis. In these systems it is particularly important that the reactance of all circuit elements, however small, be taken into account, as they have a much more significant effect in reducing the short-circuit current a t 208Y/120 volts than a t 480 or 600 volts.

FEEDERS BREAKERS

PLAN

CHANNEL B U $ - 4 0 0 0 A

150'

n

w

I

y

Z

NETWORK TRANSFORMER

1 3 2 0 0 - 2 1 6 ~ / I 2 5 VOLTS KvA

INCOMING LINE 5 O YVA SC

II

nus o' NETWORK PROTECTOR Z500 A

L-

CIRCUIT BREAKER

.""... CD"Yl

rnY"l

4000

ELEVATION

FIG. 1.46 Arrmgement of equipment for large 208Y/120-volt spot network system.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

69

The equipment for this example is arrauged as shoirn iir Fig. 1.46. The one-line diagram is shown in Fig. 1.47.4 which iurludes the hayir reartanre data on the circuit elemenk. The impedauce diagram is shown i n Fig. 1.47B. Figure 1.47C shows the condensed diagram to illustrate t,he relative distribution of reactance in the system. It will be noted t,hat the overhead bus R has 70 per cent as much impedance as the romhinatiotr of all the transformers an8,huses ahead of it,. Elimiiiatiug this item would cause a serious error in t h magnitude of short-circuit, curretit. The intermediate steps etween Figs. 1.47H and 1.47C can be worked out by followiug t h e fa oing text. The short circuit is located just ahead of the maiii 4000-amp circuit breaker as this determiires the available short-circuit, curreut, which this circuit breaker must interrupt. As pointed out previously, air circuit breakers are applied 011 the basis of availahle rurreiit, and therefore \\.heir calculat,ing the short,-rirruit duty oil them, t,he impedalire of t,he rirciiit breaker is not included. Large High-voltage Power System. T h e examplc shown in Fig. 1.48 is typical of what might, be eucouritered i n a steel mill. The kva base chosen is 100,000 kva. Precise data are available 011 large motors and are used in the short,-circuit, st,udy. Since the large mot,ors roiistitute only part of the motor load, the remaining motor load is estimated. For short circuits on the 22-kv system t,he motor load is assumed to be equal to the capacity supplying each 22-kv bus, or 62,500 k r a aiid 20,000 kva. Should more precise data be available regarding ronnevted mot,or load, these data should be used for simulating motor ront,ribution for faults on the 22-kv system. In t,his example, the connected horsepower 011 the 6.0-kv bus mas known t,o be as shown in t,he diagram. To check the momentary dut,y at F , 011the KY-kv bus, the primary system should be represented by its equivalrut, subt,raiisieiit reartaure nf 12.2 per cent. For interrupting d u t y on the 6.9-kv bus, t,he primary syst,em should be represented by a reartanre equivalent t o the iirterruptirig duty on t,he 22-kv system, or 17.5 per cent. These large complicated syst,ems should he set up 011 a calculating board to enable accurate ausivers t,o he obtained easily.

J

SHORT CIRCUITS IN SINGLE-PHASE LIGHTING A N D WELDING POWER SYSTEMS (600 VOLTS A N D LESS)

A common p r a h c e is t o use single-phase trausformers roiiuected to three-phase primary systems t,o supply single-phase loiv-voltage power for welders and for lightirrg rircuits in some of the older syst,ems. When determining the short-circuit current a t the serondaries of these transformers, it, is necessary t o use the proper impedance t o represerrt the primary system. I n three-phase short-circuit calculations, the reactance

70

SHORT-CIRCUIT-CURRENT CALCUATING PROCEDURES

FIG. 1.47

One-line diogram, reactance diagram,

SHORT-CIRCUIT-CURREM CALCULATING PROCEDURES

71

and short-circuit-current calculation procedure for spot network system show in Fig. 1.46.

72

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

c:

74

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

of a conductor is the reactance from the center of the condurtor to the theoretical neutral. Assume that for eaeh phase the rurrent leaves on the phase conductor and returus through the neutral. In a three-phase short circuit, the three currents balance; so there is no rurrent flowing in the neutral. With single-phase line-to-line short cirruits, the eurreut leaves on one phase conductor and returns ou the other. Therefore this rurrent sees the reactance of two condurtors as beiug in series. Heure, for siuglephase tramformers conuected line-to-hie on the primary, twire the primary system impedance must be used to represent it in a true relation to the rest of the circuit. The remaining calculatious are essentially the same as for three-phase circuits using the transformer and loiv-voltagecircuit reactances. Single-phase tramformers used for supplying 120/240-volt single-phase lighting circuits usually have the midtap available for ronnerting to threemire neutral and ground by the user and are usually relatively low iu kva. These small transformers have a relatively high resistatire-t~reactance ratio compared with three-phase trausformers of a higher seroridaryvoltage rating and of larger kva rating.

7

100,000 KVA 3 PHASE SHORT CIRCUIT OUTY

BASE 500 KVA

4tL PRIMARY SYSTEM REACTANCE ON 3-PHASE BASIS.

PRIMARY SYSTEM REACTANCE ON SINGLE PWSE BASIS = 0.005X 2 * 0.01Va

i

PRIMARY SYSTEM X

:0.01%

TRANSFORMER X =O.O3Ym

TOTAL X ~ 0 . 0 4 %

1%'

o , 0 4 ~ ~ , , e o :%:26000 AMP SYMMETRICAL 0.0192 11.18 MODIFIED)

1.25 X 26000

FIG. 1.49 system.

= 32500

A M P ASYMMETRICIL K2Ol

Short-circuit-current calculating procedure for single-phase two-wire 480-volt

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

75 .

The most severe short-circuit condition in this case is a line-to-neutral short circuit because it involves a much higher primary-to-secondary turn ratio than does a line-to-line short circuit. Hence, this is the basis on which protective equipment should be selected. Since the reactance and resistance of the transformers are given on the basis of a full winding, it is necessary to convert to the proper values when only one-half the secondary winding is involved as is the case when a line-to-line neutral short circuit occurs. The reactance is increased by a factor of 1.2 and the resistance by a factor of 1.41. Therefore, the published reactances and resistances of these transformers are multiplied by those figures. Figure 1.49 shows a typical example where reactance only is used, as would be the case for a relatively large 480-volt transformer supplying a welder circuit. I n these calculations it is necessary to use twire the lineto-neutral reactance of the primary system. In the example of Fig. 1.50 use twice the line-to-neutral reactance of the primary. Use the proper

F+

100 000 KV4 3 PH4SE SH& ClRCUlT DUTY

1

120,240-V )IR X =:3 1.2% X ON ,. FULL ,. WINDING 84SIs

B4SE 5 0 KV4

-

PRIM4RY SYSTEM RE4CT4NCE ON 3 P H 4 K 84%

0.00198 PRlM4R"X

:0

0005~~

II 61

TR4NS X 0036% PRIM4RI SYSTEM RE4Cl4NGE ON 4 SINGLE PH45E B451S~00a)5X2iOO019~ H4LF WlNDlNG RE4CT4NCE OF TRMIYORMER42 XO0310036X TWW R 00172% 'I RESIST4NCE " .' ~144X0012~001720/1

1.25 X 10300

FIG. 1 .SO

i

I2900 4MPS ASYMHETRICbL 11.201

Short-circuit-current colculating procedure for single-phase three-wire 120/24Q.volt system.

76

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

reactance and resistance for line-to-neutral short circuit a t the secondary of the transformer. In both cases there is assumed to be no motor feedback.

TABLES AND CURVES FOR ESTIMATING SHORT-CIRCUIT CURRENTS To make short-circuit protective equipment application easier, particularly in circuits of 60 volts or less, many charts, tables, and curves have been prepared to eliminate the necessity for detailed calculations. Some of the more usef 1 ones are presented here. UNIT SUBSTATIONS

:o.

Standard low-voltage unit substations so widely used have standard transformer section impedance and voltage ratings. Hence, the secondary short-circuit currents available can be easily tabulated, as shown in Tables 1.5 and 1.6. The available short-circuit duty may be read directly from the table as a function of transformer kva, secondary voltage, and available primary short-circuit kva. Example of Use of Table 1.5. Assume a lonn-kva unit substation for 480-volt power service having an available SHORT CIRCUIT primary short-circuit capacity of 150,000 kva. "2 x, See 480-volt application table. Follow FIG. 1.51 0 n e - k diagram the vertical column under the 1000-kva suhshowing location of short circuit station rating down to the 150,000-kvaavailfor determinotion of short-circuit able primary three-phase short-circuit kva currents shown in Table 1.5. line in thetable. The availableshort-circuit current a t the 480-volt bus is indicated as 30,400 amp.

% ?]-'"

REDUCTION OF SHORT-CIRCUIT CURRENT DUE TO FEEDER IMPEDANCE

The unit substation application Tables 1.5 and 1.6 make it easy to determine the short-circuit current a t the main unit substation bus. By the use of the simple estimating curves the short-circuit, current at the end of the secondary feeders can he easily determined too. Henre these tables and the curves shown in Figs. 1.52 and 1.53 make it easy quickly to estimate the short-circuit current a t any point in a secondary system 600 volts and less fed by standard load-center unit substations. The curves are for 60-cycle operation. Figure 1.52 is for cable cirruits and Fig. 1.53 for bus feeders. The results are in terms of the three-phase average asymmetrical rm value during the first cycle corresponding with the basis of rating for low-

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

77

voltage air circuit breakers. The effect of circuit resistance both in increasing the impedanre and speeding the decay of the d-c component 'has been included. The range of operat,ing conditions encompassed is as follows: System operating voltage (nominal) : ZOSY/lZO volts, three phase, four wire; or 208 volts, three phase, three wire 480 volts, three phase, three wire; or 480Y/27' volts, three phase, four wire 600 volts, three phase, three wire Short-circuit-current magnitudes: 10,000 t o 100,000 amp Feeder-circuit construction : Three-conductor cable, No. 4 Awg to 500 MCM Busway, plug-in bus of representative designs in current ratings from 225 to 800 amp. Interlared loiv-reactance feeder bus (LVD) rated 2,000 amp, t,hrec phase (four bars per phase). y/

3

.

CABLE FLLOER LENCTM- FEET

FIG. 1.52

Chart for determining short-circuit current a t end of cable circuit consisting of three-conductor cable in conduit or interlocked-armor cable (60cycler).

NOTE:

12.2

12.3 12.4 12.5 12.6

11.9

10.0

417

150

15.9 16.5 16.7 16.9 17.1 17.2

625

500

Normol current, amp

300

750

1000

20.7 21.7 22.1 22.4 22.6 22.9

834

2080

2780

4170

I500

270

361

1 542

32.4 35.0 36.0 36.8 37.5 38.1

~

42.3 46.8 48.5 50.0 51.3 52.5

53.3 60.4 63.3 65.9 67.9 70.2

~

48.7 61.3 74.5 80.0 85.5 90.0

I

9.4 9.6 9.7 9.7 9.8 9.8

11.2 11.5 11.6 11.7 11.8 11.8 15.1 15,6 15.8 16.0 16.1 16.2

Total low-voltoge short-circuit Curlenh, thousands of amperes

1388

1 1 1 1 1 1

225

1 1 I I I 1 1

10.3 10.4 10.4 10.5 10.5

313

112.5

Substation kva rating

19.7 20.6 21.0 21.2 21.5 21.7

722

I

I

1

31.1 33.3 34.2 34.9 35.5 36.1

1203

rmal current, en

-

41.3 45.1 46.6 48.0 49.0 50. I

-

-

1804

-

Substotion kra rclting

52.2 58.3 60.8 63.0 64.8 66.7

71.2 82.5 87.5 92.0 95.9 100.0

%

impedance,

former

4.0

4.5

5.0

5.0

5.0 5.5

5.5

5.5

4.0

4.5

5.0

5.0

5.0

5.5

5.5

5.5

or different voltoge bare, multiply short-circuit current values in table by NOTE: 3. For differed wltmge hose. I tipiy 9 208 240 the ratio values in toble by the ralio naw voltoge n o r *olt.*e NOTE: 2. Motor short-circuit current contribution is 2.5 times the transformer normal NOTE: 4. Motor short-circuit current-contribution is 5.0 t i m n lhe t m n r current for 50% connected motors. former norm01 current for 100% connected moton.

~

Unlimited

50.000 100.000 150.000 250,000 500.000

.

Fi.C"it kw

short-

Available Primary threephase

SECONDARY RATING: 240 VOLTS, THREE PHASE

Available Short- circuit C u r r e n t f r o m ' t o n d a r d T h r e e - p h a s e Unit S u b s t o t i o n s

SECONDARY RATING: 2 0 8 Y / l 2 0 VOLTS, THREE PHASE

TABLE 1.5

a

$

2

5

2

'$

2

80

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

FIG. 1.53 Chart for determining short-circuit current (it end of feeder bur. designations refer to General Electric Company bus I60 cycles).

The type

Required Data. The basic data needed to enable the use of Figs. 1.52 and 1.53 are the following: 1. System operating voltage 2. Available short-circuit current at the source bus (average asymmetrical) 3. Length and construction of the feeder circuit 4. Connected motor load at the feeder terminal Procedure for Use of Figs. 1.52 and 1.53. The evaluation of feeder terminal short-circuit current involves only four simple steps (see Fig. 1.54): 1. Locate the magnitude of source-end short-circuit current on the proper left-hand operating voltage scale. 2. From this starting point move along to the right following along a curve or an interpolation between adjacent curves until the desired length of specific feeder construction (horizontal scales) is reached. 3. Project the latter point horizontally to the left and read the shortcircuit current contributed by the feeder on the same scale as used in 1.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

81

4. Add the feeder terminal connected motor-current contribution (five

times the sum total of the motor full-load current). MODIFICATIONS FOR SPECIAL CONDITIONS

Parallel Circuit Feeders. A feeder circuit composed of two or more identical circuits in parallel can be readily treated by making a correction in the apparent length. The impedance presented by a feeder consisting of two circuits in parallel will be identical to that of a sing16 circuit of half the length; that of three circuits in parallel will be identical to that of a single circuit of one-third the length; etc. In the case of parallel circuit feeders, divide the true feeder length by the number of circuits in parallel and proceed on the basis of single-circuit data. .

I

Y

t

H k--

I .

1 i I ,/CP)OICI

I

I

l

I_

FEEDER L m m "

FIG. 1.54

I

Example rhowing how to use the charts of Fig. 1.52 and 1.53.

l

850

I

25 137.5

I I 1 50

75

100

150

1 1 I 1 333

250

200

500

Available primary Normal eurrenl. ornperes a1 240 volts

lhree-phose hoil-circull kro

104

I 1 I 1 I 1 156

208

313

417

625

833

1042

I

1388 12083

Tolo1 lox-vollage shw-circuil c ~ ~ r e nlhousandr l, of rms omperes for m e 120-volt winding short-circuited, lhe olhei opon-circuiled

- 25,000 50,000 100,000

I

6.5 9.6 12.6 6.7 10.0 13.3 6.8 10.2 13.6 6.8 10.2 13.7 6.9 10.3 13.8 6.9 10.3 13.9 6.9 10.4 14.0 - -

150.000 250,000

500,000

15.9 16.9 17.5 17.8 17.9 18.0 18.1

20.4 22.1 25.1

23.5 23.7 24.0

I 28.4 31.8 33.9 34.7

I

I

I

32.3

37.3

5.0

5.5

48.9 40.8 36.9 43.5 60.2 44.2 39.7 47.4 68.1 45.5 41.0 49.0 71.1 35.3 46.6 41.7 50.1 73.9 1 3 5 . 8 1 4 7 . 5 1 4 2 . 3 1 5 1 . 2 176.0

35.2

__

Unlimited Transformei full-winding impadançe: Per cent R . . Per cent 2..

... ....

1.4

1.4

1.2

1.2

1.21

3.0

3.0

3.0

3.5

3.5

1.21 3.5

1.21 3.5

1.0 5.5

- -

A short circuit invalving one of the secondrtry half windings (terminals Xi to X 2 ai terminals X, to X , ) , Fig. 1.51, allows eansiderahly more short-çireuit current to flow than a short circuit involving the full seeondary minding (terminals X i to X d . Consequently, the circuit-hreaker seleetions are based on the half-winding value of shortcircuit current. The eonditions on whieh the tables are hased are summsrizcd below: 1. A salid half-winding short cireuit at the tcrminals (scc Fig. 1.51). 2. Primary three-phase short-eircuit capacities vsrying from 25,000 kva to unlimited kva. For the worst case, the single-phasr short-cireuit capaeity is me-half the threephase primsry short-circuit capacity, and this value has bem used in thc celculations. This worst csse involves the assumption t h a t the primary of the transformer is connected line-to-line on the high-voltage system, not line-to-neutral. 3. The full-winding per cent impedance and per cent resistances m e given in Table 1.6. 4. The half-winding reactance was taken as 1.2 times the full-winding reactance, while the half-winding resistance was taken as 1.44 times the full-winding resistance, on full kva base. 5. The d-e offset multiplier for the first half eycle was taken as 1.25. 6. It is sssumed that the 120/240-volt units will supply lighting loads only, i.e., no motor feedbaek. 7. The only source of power connected to the secondary bus is one transformer of the capaeity indicated.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

83

Feeders Consisting of Different Circuit Construction in Series. Make an independent evaluation of each common circuit construction starting at the source end. 1. Evaluate the short-circuit current a t the end of the first section of common feeder construction in the standard manner. 2. Using the answer derived from 1 as the source short-circuit-current value for section 2, proceed in the standard manner t o evaluate the shortcircuit current a t the end of the second section. 3. Using the answer derived from 2 as the source short-circuit-current value for the third section, proceed in the standard manner to evaluate the short-circuit current a t the end of the third section. Results obtained for sections beyond the first will be somewhat on the conservative side (higher than the true short-circuit-current value). This follows from the fact that the basic analysis assumes an X / R ratio of 12 a t the source end of the feeder. The true X / R ratio at the source terminals of any feeder section beyond the first will necessarily be less than 12 since no feeder construction exhibits an X / R ratio as high as 12. Interpolation for Intermediate Cable Conductor Sizes. Specific cable feeder length scales have been inscribed for conductor sizes of 500 MCM, 250 MCM, No. 2/0 Awg and No. 4 Awg. For intermediate valuesof cable size locate the horizontal scale points for the desired length of adjacent cable sizes which are charted, and interpolate between these values. For example, a No. 3/O-Awg conductor is about midway between a No. 2/0Awg and a 250-MCM. To evaluate the effect of a 100- f t run of No. 3/0Awg cable based on Fig. 1.52, locate the 100-ft point on the No. 2/0-scale and on the 250-MCM scale. A point midway between these two points will closely represent 100 ft of No. 3/O-Awg conductor. Three Single-conductor Cables in Conduit. Results obtained from the estimating curves without correction can be safely used to select protective interrupters. If desired, a closer approximation of the actual value can be obtained by increasing the apparent feeder length to account for the higher impedance of single-conductor feeder circuits. Conductor Sirs 500 M C M . . 250 M C M . . No. 2 / 0 A r g . . No. 4 Awg

Use an Appored Lenglh of

........ 130% of lhe acluol feeder Imglh ........ 120% of the o c h d feeder lenglh ..... 110% of lhe amal feeder lmglh ......... No correction

Both the 60-cycle resistance and reactance of a three-single-conductor cable feeder in conduit are greater than those of a three-conductor cable feeder in conduit or steel armor in the ratios reflected in the accompanying table:

84

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

Residence,

Conductor sire

I

..............

500MCM No. 2/0 A w g . . No.4Awg

I

......... .............

7%

Reactance.

Yo

I i 25

106 102

I

1-50 150

150

N OTE: Spaced open-wire circuits should be treated by conventional calculation procedures; a suitable one is given under Circuit AnalysisGeneral Case. Single-phase Circuits. Results obtained from the curves, Figs. 1.52 and 1.53, may be used with safety for the selection of protective interrupters. The true short-circuit-current value for a two-wire single-phase circuit operating at line-to-line voltage will be about 87 per cent of the t h r e e phase evaluation. Frequency. The curves, Figs. 1.52 and 1.53, are restricted t o 60-cycle operation. For operating frequencies other than 60 cycles, conventional calculations should be used, such as outlined under Circuit AnalysisGeneral Case. Note that feeder circuit resistance is not appreciably affected by frequency, while reactance varies directly with frequency.

UhIN SOURCE BUS 48O"OLTS ,.P"**E 6 0 C I C L E S SHORT CIRCUIT C W l R E N T i

4CCOOAYP

2 5 0 YCY 3IC INTERLOCKED ARMOR CABLES IN PARALLEL

FIG. 1.55 System diagram used as on example to illustrate the determination of short-circuit currenk a t the end of feeder circuits.

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

Example of Application-Fig. 1.55.

85

Short-circuit current at bus A ?

Source short-cirruit current = 40,000 amp Equivalent single cable feeder length = 1595 = 75 ft From curve Fig. 1.52 (4GO-volt short-circuit current scale; 250-MCM feeder Irngt,h scale) : Contribut,ion via feeder cable = 23,000 amp Motor contribution, bus A = 5 X 310 = 1,550 24,550 amp 315 Motor contribution, bus R = 5 X 03 = Short-circuit current bus A = 24,8G5 amp ~

~

Short-circuit current a t bus B? Source short-circuit current for section 2 = 24,550 amp (say 25,000) Feeder lengt,h = 75 f t From curve (4GO-volt short-circuit current scale) interpolate between the 7 5 f t point on ;To. 210 and KO.4 feeder length scales-Ko. 2 about onethird of the way from Xo. 4 to No. 2/0. Contribution via feeder cable = 11,000 amp Motor ront,ribution, bus R = 5 X G3 = 315 Short-circuit current bus R = 11,315 amp ~

CIRCUIT ANALYSIS- GENERAL CASE

The circuit, problem involved in resolving short-circuit-current magnitudes in low-voltage feeder systems is outlined in Fig. 1.56. I n general, low-voltage short-circuit current,s are expressed in terms of three-phase average asymmetrical rms amperes during the first cycle of currcnt flwv. Since main low-voltage source systems exhibit a n X / R rat,io of about, 10, it, is standard convention t o multiply the symmet,rical short,-rirruit, current, by 1.25 t o obtain the short-circuit current a t the main buses (this corresponds with a n X / R ratio of 12) (see Table 1.2). Therefore, at the main bus 1.25 E Short-circuit current = 1.25 X I symm = - X -

v5 z*

z , = 1.25 -& x

E short-circuit current

Considering the source system X / R ratio 1.25

z . = -4 x

E short-circuit current

(A +

=

12

jl) = R.

+jX.

86

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

+ + + + + _-

2, (obtained from reference tables) = R, j X , 2, (impedance to end of feeder) = R, R, j ( X . X , / R , ratio a t end of feeder = x. XI - x, R. Izi Rt

+XI)

M is the factor to account for d-c offset and is a direct function of the X , / R , ratio

......... ...................

XdRt ratio.. K

I 1 I I I 1 1;;s

I:*

s1:

!l

Ii6

2 1.02

I, is the local motor contribution, and the three-phase average assymmetrical rms value may be taken as five times the motor full-load rated amperes. Available short-circuit current at X = I, (three-phase avg assymmetrical rms) I,

+

61 2 -SOURCE SVSTEY IMPEDANCE Rg+ j X s OHMS/PHASE

'1

I

MAIN LOW-VOLTaGE BUS

4

FEE0ER:Zf:Rf

:'I

\J

LOAD

tjxf

OHMSIPHASE

IFROH TABLES)

VAIL4ELE SHORT ClRCUlT CURRENT DESIRED HERE IS'CURRENT CONTRIBUTION FROM SOURCE *"STEM ly*CURRENT CONTRlBUTlON FROM LCCAL YOTORS

FIG. 1.56 One-line diagram for rhortcircuit-current calculation ot the end of feeder circuits-genernl core.

07

SHORT-CIRCUIT-CURRENT CALCULATING PROCEDURES

TABLES SHOWING EFFECl OF CABLE LENGTH

Another useful way of showing the effect of length of cable in reducing short-circuit currents is given in the Tables 1.7 to 1.10. These show how much cable length is required to reduce the short-circuit current from one protective-device rating level to another for circuits GOO volts and less. Standard protective-device rating levels are: 100,000 amp rms asymmetrical 75,000 amp rms asymmetrical 50,000 amp rms asymmetrical 25,000 amp rms asymmetrical 15,000 amp rms asymmetrical 5000 amp rms asymmetrical The tables show how long a cable with a given confiom a short,-rircuit standpoint, it makes no difference whcthcr the rircirit breakers are of t,he nil or oilless type. Ratings Available. High-voltage power circuit breakers are availahle in ratings from 2.4 kv up to over 300 kv and in interruptirig ratings from 15 mva up to 25,000 mva. Complete listings of power circuit breakers can he found iii the latest copy of S E R l A Standards SG&l954. T h e circuit, breakers most comtnonly used in industrial plants are the oilless or air type, sho\rn i n Fig. 3.14. The available ratings of this type of circuit breaker are given in Table 1.1 (Chap. I).

FIG. 3.14 Typical ille err (air) power circuit breaker ar wed in metal-clad switchgear for c i t w i t s rated 2.4 to 13.8 kv.

172

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

FIG. 3.15 Outdoor frome-type oil circuit breaker This circuit breoker i s rated 34.5 kv.

01

used in circuits rated above 15 kv.

Application. Power circuit breakers are applicable anywhere in the syst,cms rated 2.4 kv up t,o the highest a-c voltages in use today. They combine all the essential characteristics for circuit switching and protect,ion and therefore may be used at main buses supplied by generators or transformers or i n connection with unit substations. They are also applirable at, loral switching points and for protection of primary branch circuits (see Fig. 3.16). Motor Starting or Other Repetitive Duty. Certain of the power circuit breakers, particularly the oilless type, are suitable for motor-st,arting duty within the limitations outlined by the manufacturer. It should be noted that compared with contact,ors the principal limitation of power circuit breakers for motor-starting duty is the degree of repetitive duty that can be withstood. Contactors are designed for more operations and longer life under severe operating duty cycles than are power circuit breakers.

A-C SHORT-CIPCUIT PROTECTIVE DEVICES ANO CIRCUIT EQUIPMENT

173

Selection of Interrupting Ratings. The selection of interrupting ratings of power circuit breakers for industrial applications is out,lined in Chap. 1. A detailed description of the various faetors to consider in applying oilless eircuit breakers as used in metal-clad switchgear is given there.

I

Q

GENERATOR

I

Q P

69 KV

OUTDOOR POWER C I R C U I T BREAKERS

T TYWI

GENERATOR CIRCUIT BREAKER

TRANSFORMER SECONDARY CIRCUIT BREAKER

!

' MAlN FEEDER CIRCUIT BREAKER

A AHEAD O F L I N E OF L I M I T A M P MOTOR STARTERS

LARGE OU HIGH VOLTAGE MOTORS

FIG. 3.16 One-line diogrorn rhowing where oilless power circuit breakerr in metal-clad rwitchgeclr and outdoor power cirwit brecikerr may be applied in industrial power dirtribution ryrtemr.

174

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

Standards. Poiver circuit breakers are eovered by NEMA Standards SG1-19% POWER FUSES A N D OIL-FUSE CUTOUTS

There are many types of power fuses available for circuits rated 2.4 kv and above. These t,ypes of fuses, generally speaking, divide t,hemselves into three categories. The first is the power fuse, typical examples of which are shown in Fig. 3.17 which are for high-rapacity power circuits. The second type that is slightly differeni, i n construct,ion i s the oil-fuse cutout, which i s really a combination of a cntout and a fuse immersed in a container of oil, Fig. 3.18. The third type of fuse is used mainly in distribntion cutouts for overhead opcir-wire outdoor distriliutioii systems of utilit.ics in urban and suburban areas, Fig. 3.119.

FIG. 3.17 Typical high-voltage (above 600 volts1 power furer: Ifeft) current-limiting nonenpulrion silver-rand type, (right] "on-current-limiting expulsion outdoor type.

176

A-C SHORT-CIRCUIT PROTECTIVE DNICES AND CIRCUIT EQUIPMENT

The last type of fuse mentioned is applicable toindustrial power systems for outdoor installations only where the interrupting rating is less than the duty on the system. This fuse is not metal-enclosed and is not for indoor installation. I n general, power fuses divide thcmselves into two classes, i.e., currentlimiting and non-current-limiting. Typical of the current-limiting category are the silver-sand fuses, Fig. 3.17(left). Typical of the non-current-limiting type are the oil-fuse cutout, Fig. 3.18, the expulsion fuses, Fig. 3.17(right), as well as the “boric acid” fuses and “liquid” power fuses. A further classification is that some are expulsion type, i.e., expel hot gases when they operate. These are not suitable for indoor application because of the hazard of the expelled hot gases. Such fuses are the expulsion fuse, Fig. 3.17(right), and the “boric acid” fuse without a condenser and the “liquid fuse”. Typical of the nonexpulsion type are the silver-sand fuse, Fig. 3.17(left), and the “boric acid’’ fuse with condenser. Application- General. All types of power fuses operate faster than power circuit breakers a t or near their interrupting ratings. Because of the fast operating time of the fuses, they are generally employed as the last circuit protective device in each voltage level in a primary power system, as shown in Fig. 3.20. Typical applications are in motor starters and ahead of primaries of transformers stepping down to a lower voltage. The silver-sand fuse, Fig. 3.17(left), is often the preferred type of fuse for power circuits because of its fast operating time and currentlimiting ability. However, in some cases where coordination is required, it may be necessary to use non-current-limiting types of fuses which have longer time delay. However, when the longer time delay is obtained, the benefits of reduction of damage to the circuit through which shortcircuit current passes is lost to a large degree. Interrupter Switches and Fuses. Nonexpulsiori-type power fuses suitable for indoor use are often applied in a metal enclosure with an interrupter switch to form a switch-and-fuse cornbination for high-voltage circuits. Interrupter switches are desirable for this application because they have interrupting ratings usually in the range of 100 to 400 amp. Plain disconnecting switches are generally not satisfactory for this service because they have no interrupting ability, and therefore the combination of the plain switch and fuse cannot be used as a load-switching device. The oil-fused cutouts combine in one unit the fuse and the interrupter switching element. Interrupter slyitches and fuses and oil-fused cutouts find wide application in industrial plants as the primary swit,ching and protecting section of a load-center unit substation (see Chap. 11). Application of Fuses in O p e n Switching Structures. Open-structure switches or disconnect,ing mountings without current-interrupting ability are often used with power fuses. These can be considered for isolation purposes only. Hazards in operations are materially increased in this

A-C SHORT-CIRCUIT PROTECTIVE DNICES AND CIRCUIT EQUIPMENT

177

type of appliration. That is the reason that such applications should be limited t o outdoor structures Ivhere the operator is a considerable distance from the disconnecting switch when operating it. The use of such isolatiug switches i n series with fuses in indoor metal-enclosed structures is not coilsidered safe practice bemuse of thc proximity of the operator to t,he sivitrh and the possibility of the operator inadvertently operating the switch under roiiditions i u which the switch will hare to interrupt or close in on currents ronsiderably beyond its ability. Failure may result eveti though there is a fuse in series with such switches. 33 K V

P

UTDOOR TYPE FUSE SMALL POWER

IyTy\ TRANSFORMER

I LIMITING

AHEAD OF SMALL LOAD CENTER UNIT SUBSTATIONSUSE INTERRUPTER SWITCH AND POWER FUSE OR FUSED OIL CUTOUTS. FIG. 3.20 One-line diagram rhowing where high-voltage (above 600 may be applied in industrial power distribution systems.

VOllS)

Power

178

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

Selecting Fuse-interrupting Rating. Fuses are generally rated in amperes interrupting ahility. CMculate the short-circuit duty in rms amperes asymmetrical at the first half cycle as outliiied in Chap. 1, and select a fuse whose interrupting rating is greater than the duty imposed. Equivalent three-phase iiiterrupt,ing ratings may also be considered. Since the ratings of fuses are not too well st,andardized, refer t o t h e fuse manufacturer for complete data before applying fuses. Standards. Power fuses are covered by S E M A Standards, Cutouts, I’orer Fuses, and Current-limit,ing Resistors, Publication S(2-1954, and AIEE Standards S o . 25. MOTOR STARTERS

There are in general three kinds of motor starters: 1 . The contactor 2. The combination motor starter 3 . The circuit breaker Contactors are in general of two types, the most common variety being t,hose which have a n interrupting rating of only ten times normal rated current. These are completely inadequate for short-circuit protection and must have addit,ional short-circuit protection provided b y either fuses or circuit breakers. When a short,-circuit protective device such as a fuse or circuit breaker

FIG. 3.21 Typical lowvoltage 1600 volts and below) combination motor starter with current-limiting silver-rand furer for short-circuit protection.

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

179

is used in comhinatiori with contactors, it forms what is commonly called a combination motor starter. Circuits 600 Volts and Less (Fig. 3.21). In systems 600 volts and less there are t w i types of cornhination motor starters, both employing the same type of cont,act,or. The first employs a fuse disconnecting sm.itch alirad of t,he conta,ctor, and the other a circuit, breaker, usually a moldedcase-type circuit breaker, ahead of t,he cout,actor. The select,ion between t,hese two is based mainly on the fuudamerital differelice betveen fuses and circuit, breakers as short-circuit protective devices. The fused combinatirin motor starters have an over-all interrupting ahility so that the combination motorst,arter can successfully irit,errupt an available short-circuit current equal t,o 50,000 amp rms asymmetrical wheri equipped wit,h high-interrupt,ing-capacity currentlimiting fiises. This is for a short circuit outside ” the case of t,he mot,or starter and using type E.14 fuses. The molded-case circuit-breaker comhiiiat,ion mot,or starters are limited to a maximum duty of 15,000 or 25,000 amp rms asymmet,rical. Circuits above 600 Volts (Fig. 3.22). For circuits of 2.4 kv aud up t o 5 kv, the combination motor starter commonly used consists of current-limiting silver-sand fuses and contactors with the fuses mount,ed in disconnecting-type supports and placed in a metal enclosure s o interlocked that the fuses cannot be disconneeted unless the coritactor is in the open position. In this way the disconnecting fuse mounting has no current to interrupt. Since the FIG.3.22 Typicalhighfuses are for short-circuit protection only, suit,able (2,4to 4,8 kv) running overload relays should he provided in the bin tio motor motor st.art,er. These motor starters have inter- starter using current-limrupting ratings of 150 mva at 2.4 kv and 250 mva iting rilver-rand power a t 4.16 kv. From a short-circuit standpoint they furel for short-circuit protection. may be appIied up to their momentary and int,errupting rat,iiig. Since these devices contain fuses as the short-circuit protective element, they are naturally best suited t o application as the last protective device in the circuit. When used as motor starters, they are t,he last protective device, and therefore the fast operating time of the fuse is a very dist,inct advantage in limiting damage in the motors when a failure occurs. The fast operating time of the fuse also permits low settings on other relays further back in t.he system.

180

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

SELECTION OF CONDUCTORS AND OTHER CIRCUIT COMPONENTS FROM A SHORT-CIRCUIT STANDPOINT The floiv of short-circuit current in an electric system imposes mechanical and thermal st,resses (heating) on all component,sof the system through which such currents flow. This includes cables, bus bars, current transformers, disconnecting switches, as \veil as circuit breakers, fuses, and motor starters. The following is intended t,o aid in the selection of circuit component,s, ot,her than circuit breakers, fuses, and motor starters, from a short-circuit,-current standpoint. POWER-CABLE SELECTION FROM A SHORT-CIRCUIT STANDPOINT

Multiple-conductor power cables possess high mechanical strength because of the compact conductor lay and the continuous concentric binding arsist,ed many times by armor or lead sheath. KOlimit on mechanical stresses in such cables has been assigned. This is not true with regard to thermal effects. In common with ot,her current-carrying parts of the electric system during short-circuit-rurrent flow, the abrupt elevation in conductor temperature will be limited only by the ability of the conductor metal to absorb the heat developed. The magnitude of the temperature increase is greater (1) as the current magnitude becomes greater (as the square of the current), (2) as the conductor cross section becomes smaller, and (3) as the duration of current flow becomes greater. Temperature limits. Power-system short-circuit-current magnitudes, feeder-conductor cross section, and short-circuit protective device interrupting time should be coordinated to avoid severe permanent damage to cable insulation during an interval of short-circuit-current flow in the system. The effect should be limited to a moderate reduction in useful cable life (possibly 1 per cent of normal life). Reasonable maximum-peak transient temperatures for various cable insulations and operating potentials have been designated and in general are approximately 150 C (see Table 3.6). At a slightly higher temperature (approximately 175 C), destructive disint,egration of organic materials may occur, accompanied by smoke and combustible vapors. At somewhat higher temperatures large quantities of combustible vapors are expelled which increases the risk of explosion and fire. It is important to note that the abnormal temperature persists much longer than t,he duration of short-circuit-current flow. For example, the flow of 20,000 amp in a KO. 4/0-Awg copper conductor will elevate the

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

181

copper temperature from an initial temperature of 75 C to 150 C i i i ahout 34 see. With the current then redured to zero, about 1000 see d l be required for the copper temperature to return to 75 C in a 30 C ambient. TABLE 3.6 Conductor Rated Maximum Continuous Operating Temperature and Peak Transient (Momentary) Temperature for Various Types and Operating Voltages

-

MOX

lronrienl

lollogl Cable type

d.**,

copper temp.

kv

C

Vc type V or VL, single conductor or three conductor..

Impregnated paper (slid), single conductor..

lmpregnalod paper (did), three mnductor.

Type R*.

T i p s RH

Coronol

.......

..............

...............

............................................ ............................................

.............................................

* Applies to new

I 5 8 I5

85 85 84 77

I50 145 135 120

1 5 8 I5

85 85 85 81

I50 I45 140 135

1 5 8 I5

85 85 85 81

I40 135 I30 125

I

60

5

60

8

60

15

60

I40 135 130 125

1 5 8

75 75 75

I50 145 140

1 5 8 15

80 80 80 80

I50 145 I40 I30

-

-

type R (1947 code specification). t Actual operating temperature may be lompr because of consprvative application or a favorable ambient temperature.

Conductor H e a t i n g . On the basis that all heat produced by shortcircuit-current flow is initially absorbed by the rondurtor metal (wbirh

A-C SHORT-CIRCUIT PROTECTIVE DNICES AND CIRCUIT EQUIPMENT

I82

TABLE 3.7

Quick Estimating Table o f Minimum Conductor Sire' A. Low-voltoge Air-circuit-breaker Protection

Short-circuit current. ~

omp

11.25

X rymmetricall

~~

1.5 to 2 cycles

>g s*c

linrt. trip) No. 8 Awg No. 4 Awg

5,000

I0,OOO 15,000

No. 2 Awg No. I Awg

25,000

No. 1/0 Awg No. 3 / 0 A w g

35,000

5o.m

300 M C M 350 M C M

75,000

100.000

Interrupting kvo at

Short-circuit

No. No. No. No.

4 Awg

No. 2 Awg No. 1/0 A w g No. 3/0 Awg 300 M C M

1 Awg 2/0 Awg 4/0 AWQ

300 M C M 400 M C M 600 M C M 800 M C M

400 M C M 600 MCM 800 M C M I000 M C M

Duration of hort-sircuit current

current, amp

(1.0

x

symmetricoll

3,000-3.500 3,500-4.000 4.000-4.500 4.500-5.000

5,000-6.000 6.000-7.000 7,000-8,000

8,000-9.000 9,000-1 0,000 10,000- 12.500 12.500-15,000 15.000-20.000 20,000-25.000 25,000-30.000 30.000-35.000 35.000-40.000

...... 25mvo ....... 75 m w N o . 6 A w g ........................... No. 4 Awg ............. 50 mva ....... No. 4 Awg ........................... No. 4 Awg ........................... No. 2 A w g 25 mva 50 m w ....... I50 mva No. 2 Awg ........................... No. 2 Awg ............. 100 mva ....... No. 1 Awg ........................... No. 1 Awg 50 mvo .............. 250 mvo No. 1/0 Awg ...... 100 mva 150 m w ....... No. 2/0 Awr ........................... No. 3 / 0 A w r 00 m

w

150 m w 250 m w 500 m w No. 4/0 AWI 250 M C M 300 M C M 350MCM

........................... ...... 250 mva ....... 750 mvo 5 0 m r o ....... 500 mvm .......

No. No. No. No.

2 Awg 2 Awg

No. No. No. No.

1 Awg

2 Awg 2 Awg

1 Awg I / O Awg 2/0 Awg

No. 2 Awg No. 1 Awg No. 1 Awg No. 1/0 Awg No. 2/0 No. 2/0 No. 3 / 0 No. 3/0

Awg

Awg Awg Awg

No. 2/0 Awg No. 4/0 Awg 250 M C M Na 3/0 Awg No. 4/0 Awg 300 M C M 300 MCM 400 M C M

350 MCM 400 MCM 500 MCM 600 MCM

500 M C M

600 M C M 750 M C M 750 M C M

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQVIPMENT

183

has been proved to be valid for canductor sizes of No. 8 Awg or larger*), the conductor heating is governed by the following: For Copper:

t I

duration of current flow, see rms amperes during entire interval of current flow em conductor cross sect,ion, cir mils T I= initial copper temperature, C T 2= final copper temperature, C To simplify a n application, these relationships are presented graphically in the large chart in Fig. 3.23. The permissible time for various ternperature ranges can be quickly evaluated with the aid of the auxiliary curve B , shown in Fig. 3.23. For quick estimating purposes, minimum safe conductor size is given in Table 3.7, subject to application conditions as shown. For Aluminum. The problem of joining and terminating aluminum conductors without creating local “hot spots” deserves very careful attention. There are available, however, materials and methods which laboratory tests and experience have proved to be satisfactory. In the absence of abnormal local heating, a rough approximation of permissible current duration may he made on the basis of the same limiting temperatures as for copper. (For a particular current and conductor cross section, the permissible duration of short-circuit-current flow will he 45 per cent of that for copper.) It may be more convenient to make an artificial correction in current. Consider the current to be 150 per cent of the actual value, and proceed on the chart (Fig. 3.23) as if the conductor were copper. Rms Current. Rms current as used here is defined as the root-rneansquare value for the total interval of short-circuit-current flow. The temporary d-c component encountered in a-c circuits increases the rms current, but to a lesser extent as the interval of current flow becomes longer. The appropriate factor K , by which the symmetrical current value shall be multiplied to determine the true rms current is given in chart A , Fig. 3.23, for several typical ratios of circuit 60-cycle reactance * B. W.Jones and J. A. Seott, Short-time Current Ratings for Aircraft Wire and = = =

Cable. AIEE Technical I’sper, 1946.

184

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

"I.

18.5

-8.

FIG. 3.23

Short-time bhort-circuit) heating limits of copper cables and correlation of current and time to elevate the copper temperature from 75 to 150 C (dlheat is oirumed to be stored in the copped.

to resistance (distribution circuits will generally fall in thc region of X / R = 10 or less). Circuit X / R ratio is generally not known and requires numerous circuit constants for an evaluation. Conservative factors ( K , ) for the more common application conditions are

K , = 1.25 Low-voltage circuit breakers tripped instantaneously Power circuit breakers, eight cycle, instantaneously tripped K I = 1 . 1 Any industrial power-distribution problem with current duraK , = 1.0 tion of 35 sec or more Short-circuit Protective-device Interrupting Time. Circuit Breakers. The minimum time duration of short-circuit-current flow in a rircuit protected by a circuit breaker tripped by an instantaneous element will vary with the type of circuit breaker used. Typical values are shown in the lower left-hand corner of the large chart in Fig. 3.23. When interruption is purposely delayed by time-delay relays or timedelay trip coils, the duration of current flow will be governed by the timedelay relay or trip coil plus the inherent delay in the circuit breaker. Fuses (Current-limiting) , Current-limiting fuses (silver-sand and National Electrical Code low voltage) tend progressively to limit the time interval of current flow to lesser values as the magnitude of current increases. As the current magnitude increases toward the maximum interrupting ability of the fuse, the magnitude of Z't approaches a fixed value (approximately) for a particular fuse ampere rating. This is equivalent to a fixed temperature rise in a particular size of conductor. Data accumulated indicate that a fuse (of the types mentioned in this paragraph) whose ampere rating is not greater than 1.5 times the conductor continuous-current rating will protect against dangerous conductor

106

A-C SHORT-CIRCUIT PROTECTIVE DNICES AND CIRCUIT EQUIPMENT

temperatures for severe overcurrents up to the maximum interrupting rating of the fuse. Table 3.8 shows the wire sizes which will have less than 75 C conductor TABLE 3.8

Silver-sand Fuse Protection at High Overcurrents Based on Copper Conductor Fuse roting,

S m a l l ~ twire normally vied,

amp

RH insulation

30 60 I00 IS0 200

No. 10 Awg No. 6 Awg No. 3 Awg No. 1/0 Awg No. 3/0 Awg

Sm.lle.t

wire

protected

No. 14 No. I2 No. 10 No. 0 No. 6

Awg Awg Awg Awg Awg

temperature rise because of the flow of short-circuit current when protected by silver-sand fuses. Fuses (Nou-current-limiting). Non-current-limiting fuses accomplish current interruption at a normal current zero, and thus the current conduction time cannot be reduced below that of the first current loop of short-circuit-(.urrent flow which may be as much as about one cycle of the power frequenry. Applications should thus recognize one cycle as the minimum time of short-circuit-current flow. Application Procedure. 1. Evaluate the symmetrical short-circuit current or currents that may be critiral. 2. Define the short-circuit protertive device clearing time at this or these current magnitudes. 3. Apply the rms correction factor to allom for the d-c component for each time interval involved. 4. Make an initial check on the current-time chart for the smallest conductor size being considered (permissible time should exceed shortcircuit protective-device interrupting time). 5 . If critical, it is advisable to rorreet for the exart temperature range (see Table 3.6 and temperature-range correction curve). F. If an oversize ronduetor is considered, but the continuous-load rursent is to remain fixed, advantage can be taken of the lower initial ronductor temperature. EXAMPLES

Example 1. A transformer feeder cable is being selected to accommodate a 1000-kva 2.4-kv transformer. The rated current of the t,ransformer (240 amp) indirates a rahle conductor of 250 MCM. The trans-

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

I87

former iri question is good for full short-circuit current (sixteen times normal) for 5 sec. It is desired that the feeder cable have the same ability. Solution: Rms symmetrical amperes = rated current X 16 = 240 x 16 = 3900 amp. The duration of this current as defined by the conditions of the problem is 5 sec. Assume X / R ratio = 10 or less From chart A of Fig. 3.23, K 1 = 1; ( X / R ratio of 10 and time of 5 sec) Henre, the total rms amperes affecting cable heating = K , X 3900 = 1.0 X 3900 = 3900 amp On the large rhart of Fig. 3.23, locate the intersection of the horizontal 3900-amp line and the 250-MCM conductor diagonal line. The permissible time (read on the bottom scale) is indicated to be 12 sec (75 to 150 C hasis). The 250-MCM cable will adequately meet the 5-sec requirement. Example 2. Feeder circuits are t o be run from a 480-volt 60-cycle load-center unit substation at which point the short-circuit duty is 25,000 amp (20,000 symmetrical rms amperes). What is the smallest reasonable feeder conductor size based on the use of a 25,000-amp interrupting rating air circuit breaker which trips instantarieously (1.5 cycles) a t currents in excess of fifteen times the normal rating? solulion: Symmetrical current = 20,000 amp Time duration = 1.5 cycles Rms amperes = 20,000 X 1.25 = 25,000 See preceding text for explanation of 1.25 factor K , On the large rhart of Fig. 3.23, locate the intersection of the horizontal 25,000-amp line and the vertical 1.5-cycle line. The minimum size conductor (75 to 150 C basis) whose curve is above the intersection is a KO.1 Awg. Example 3. A 4-kv feeder is t o be run from a substation at which the symmetrical short-circuit current is 25,000 amp. A continuous load caparit,y of 1000 kva is desired (113 amp), and a KO.2/0-Awg coronol cable run is being considered. Line relaying is to consist of standard time-overcurrent relays on the & tap and S o . 5 time-lever setting v i t h 250/5-amp rurrent transformers. Instantaneous attachments are not planned, but could be used if set at 3000-amp line current. Solution: Symmetriral short-circuit current = 25,000 amp Case 1. No instantaneous attachment on relay Rms symmetrical short-circuit current = 25,000 amp Relay operating time = 50 cycles; (From published time-current curves) Circuit-breaker operating time = 8 cycles

188

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

+

Total time = 50 8 = 58 cycles Assume X / R ratio = 15 or less From chart A of Fig. 3.23, K 1 = 1 Hence, total rms amperes affecting cable heating = KI X 25,000 = 1.0 X 25,000 = 25,000 amp On the large chart of Fig. 3.23, locate the intersection of the 25,000-amp horizontal line and t,he 58-cycle vertical line. The smallest conductor whose curve lies above this intersection is a 500 MCM. Therefore, a Xo. 2/0-Amg conductor is inadequate. Case 2. Instantaneous attachment, on relay set to operate at and above 3000 line amperes. Two point,s must, he checked: (1) a current of 3000 amp and time delay of overcurrent relay (just below the operating current of the instantaneous element) and (2) the maximum current with instantaneous relay operation. 1. From published relay data, the relay time a t 3000 line amperes is 66 cycles, circuit-breaker time is 8 cycles, making a total time of 66 8 = 74 cycles. From chart A of Fig. 3.23 for X / R ratio of 10 and time of 7 1 cycles, K I = 1. Total rms amperes affecting cable heating = KI X 3000 = 1.0 X 3000 = 3000 amp. The intersection of 3000 amp and 71 cycles on the large chart of Fig. 3.23 s h o m that a Xo, 2/0-Awg conductor is amply large to carry the 3000 amp for 74 cycles. 2. At the maximum current, instantaneous relay operation will be obtained. The total current duration will be the relay time ?,g cycle plus circuitbreaker time 8 cycles, or 836 cycles. For 8>i-cycle time interval, K , = 1.1. Total rms current affecting cable heating = K , X 25,000 = 1.1 X 25,000 = 27,500 amp. The intersection of the 27,500-amp horizontal line and the 84g-cycle vertical line on the large chart of Fig. 3.23 indicates a No. 4/0-Awg conductor (75 to 150 C basis) and shows that point 2 cont,rols the cable size. However, a No. 4/0-Awg conduct,or mould operate at less than rated temperature. A specific check may show that a KO. 3/0-Awg conductor is adequate. Rated conductor temperature coronol cable = 80 C (see Table 3.6), ambient temperature = 40 C. Xormal temperature rise produced by rated current = 80 - 40 = 40 C. Rated continuous current for No. 3/0-Awg coronol cable = 185 amp. The temperature rise will be roughly proportional to the square of the current.

+

r

I

A-C SHORTT-CIRCUITPROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

T i m e - seconds

189

190

A-C SHORT-CIRCUIT PROTECTIVE DEVICES AND CIRCUIT EQUIPMENT

Hence, the normal conductor temperature of a No. 3/0-Awg conductor operating a t 143 amp would he expected to be

(g)’

(full-load rise)

+ ambient

=

=

63.8 C, or 64 C

The maximum momentary temperature for coronol at 5 kv is 145 C (see Table 3.6). From detail chart B , Fig. 3.23, the correction factor K for an initial conductor temperature of 64 C and final of 145 C is K = 1.13. From the large chart of Fig. 3.23, the permissible time for 27,500 amp in No. 3/0-Awg conductor (75 to 150 C basis) is 6.7 cycles. The permissible time corrected t o a 64 to 145 C basis is K X 6.7 = 1.13 X 6.7 = 7.6 cycles. Therefore, a No. 4/O-Awg conductor is the correct selection since a No. 3/O-Awg conductor would fail t o meet the 8.5-cycle requirement. FUSING CURRENT TIME FOR COPPER CONDUCTORS

The fusing current time curves for copper conductors are shown in Fig. 3.21. The curves are based on the folloiving assumptions: 1. Radiatiou may be neglected because of the short time involved. 2. Resistance of 1 cu cm of copper at 0 C is 1.589 microhms. 3. Temperature-resistance coefficient of copper a t 0 C is 1/234. 4. Melting point of copper is 1083 C. 5. Ambient temperature is 40 C. Data are an adaptation from the eight,h edition of “Standard Handbook for Elect,rical Engineers.”* * A . E. Knowlton (editor-in-chief), “Standard Handhook for Electrical Engineers,” 8th ed., Chap. 4, McGraw-Hill Book Company, h e . , S e w York, 1949.

Chapter 4

by W. R. Crites and Maynord N. Halberg*

Voltage-Standard Ratings, A llowable Variations, Reduction of Variations, Calculation of Drops The purpose of any industrial power system is to maintain voltage a t the terminals of power-using equipment. This voltage should bewithin acceptable limits-equal to the rated voltage of this equipment. The standard voltage ratings for utilization equipment are discussed in this chapter, along with the standard voltage ratings for power generation and distribution equipment. KOpractical power system can maintain voltage a t rated value a t the utilization equipment a t all times. The voltage variations allowable and the methods which can be used in the design of a power system to keep the variations within acceptable limits are discussed. I t is necessary to calculate the voltage drop in the power system for steady-state conditions and during the starting of the larger motors to determine whether or not the voltage mill remain within acceptable limits. Methods of calculating these voltage drops are presented in this chapter.

* The following men, formerly in Industrial Pawcr Enginwring. General Electric Company, made substantial contributions to the material in this chapter: W. K. Boice, General Electric Company, l e w Haven. Conn.; D. F. Capehart. General Electric Company, Cincinnati, Ohio; J. R. Eliason, General Electric Company, Sehenectady, N.Y. 191

192

V O L T A G F S T A N D A R D RATINGS, VARIATIONS, CALCULATION OF DROPS

VOLTAGE DESIGNATIONS *

It is necessary t o have a n understanding of the voltage names of systems and t,he voltage rat,ings of various pieces of apparatus used in the system before start,ing a discussion on system-voltage problems so t h a t the proper voltage identification can be used throughout. It is also necessary t,o know v h y the voltage designat,ions are applied t o help in understairding the system-voltage disussion in the following sections. The volt,age-identification structure is summarized iu Table 4.1. For each of the nominal syst,em voltages listed, t,he table gives voltage ratings of generators, transformers, motors, and (in some cases) lamps. T o illustrate the use of Table 4.1, consider a 13,800-volt system. The generators would be rated 13,800 volts. Transformers stepping power down from transmission voltage would have secondary windings (I?, Fig. 4.1) rated 13,800 volts. Transformers steppiug power down t o utilization vokage in load-center substations would have primary mindings (C, Fig. 4.1) rated 13,800 volts. Motors connerted directly to the 13,800-volt bus would lie rated 13,200 volts. From the foregoing summary and Table 4.1 it is evident that care must tie exercised in using the proper voltage ident,ifiration for each piece of equipmelit as well as for the system. Some fundamental rules are as follo\vs : 1. When speaking of equipment, the rated voltage is used, aud it is the voltage to which the operating characteristics are referred. 2. When speaking of systems, rat.ed voltage is not an applicable term because various piwes of equipment in a given system often have different voltage ratings. Therefore, t,he term n o m i n a l s y s t e m vollage is used for convenient designation of systems and circuits t o define the voltage class. The problem of proper identification would be easier if all apparatus of a given voltage class had the same vokage rating. Then, of course, tem voltage could have that same value. Possibly if the industry were starting over again, vokage ident,iticatioii mould be made that simple. But, as syst,ems grew, voltages were ini,hed up t o compensate for t,he voltage drop between source arid load. As a result, of t,hese changes that have taken pla(.e over a period of years, transformer arid generator voltage rat,ings are generally higher than utilization-eiiuipment vnltagc rat,ings. There is logic in this in that the voltage rating of transformers, for example, is t,heir no-load rating. Since most plants are supplied by transformers, the concept has beeri acceptcd that, supply equipment will have a higher voltage rating than utilization equipment,. This means that in a 480-volt system, for cxam* For a iiirthrr rrpansion of t h i s srihjpet F W l < I ~ ~ I - X 1 5 MKPport., A l’refrrrrd Voltage I h t i n g s of :\(: Systrrris and Equipmmt, N I X l’uhliration lo. R-6. S E M A I’ulilirstion l o . 117, \lay, 1‘JIU.

VOLTAGE-STANDARD

RATINGS. VARIATIONS, CALCULATION OF DROPS

193.

ple, transforniers or geiierators supplying motors ivoiild have a ratiiig of 480 volts whereas t,he motors irould have a ratiiig of 440 volts. Part of this ditrereiicc: is compeiisated for by voltage drop iii the traiisformer aiid in the distributioii system betiveeii the traiisformers aiid the motors. Therefore, in general, the voltage at the motors is reasoiiably iiear thc iiame-platc ratiiig iii the average system. I n older types of distrihiitioii systems it i m s commoii prartire to use step-doivii trmçformers irith a Iower primary voltage ratiiig thaii thc transformers which ivould siipply that systcm. For example, the ti'aiisformer steppiiig dowi from the iitility voltage ofteii hnil a ratiiig of 2400 volts oii the secoiidary, aiid the traiisformer steppiiig doi\-ii to the utilizatioii voltage of 480 or 240 volts had a ratiiig of 2300 volts oii tlie primar?.. Becausc of the desigii of preseiit-day systems n-itli smaller drgi'ers of volt,age drop, aiid judirioiis m e of taps i i i traiisformers, the prartirc is, as INCOMING

4

MASTER U N I T SUBSTATION ( P R I M A R Y SUBSTATIONI

\

1

( A I P R I M A R I WINDING

u

(IF USEDI

m l SECONDARY

WINDING

X P L A N T P R I M A R Y D I S T R I B U T I D N VDLTAGE

LOAD C E N T E R U N I T SUBSTATION (SECONDARY SUBSTATION IN FACTORYI PRIMARY WINDING WINDING

FIG. 4.1

Typicol industrial plont power ryrtern

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

194

evident from Table 4.1, t o use the same voltage rating for all traiisformer windings connected t o a given system voltage. This is true whether the transformers are stepping down to this system or steppiug down from this system. TABLE 4.1 No min0 I

Genordor

system

rated

*olt.ge

voltage

Boric Pattern of Voltage Identification Transformer secondory rated voltage

Transformer primary rated voltage

120 or 120/240 I20 or 120/240 I20 or 120/240 240 or 120/240 240 or 120/240 240 or 120/240 208Y/120 208Y/120 208Y/120

120 240 I20

Motor and control rated rottoget

230 115

Three-phase Systems

208Y/120' 240 480* 600 2,400' 4.160' 4,800 6,900* 12,000 13,200 13.800' 23,000 34,500 46,000 69,000 1 1 5,000

20sY/l20 240 480 600 2.400 4,160 4,800 6,900 12.500 13.800 13.800

........ ........ ........

208Y/120 240 480 600 2,400 4,160 4.800 6,900 12,000 13,200 13,800

........ ........ .... .... .......

208 or 120 240 480

600 2,400 4,160 4.800 6,900 12,000 13.200 13.800 22,900 34,400 43,800 67,000

L.mp rated YoltDge

118or120

I

220 or 208 208.118. or 120 236 220 440 165 2,300 4,000 4,600 6,600 11,000 13,200 13.200

I I0,OOO

* In ~ P I Vinstallations, or W ~ P ~ P Ya srlwtion P ~ oi voltngr can l i p ~ n a d rthrsr . i ~ r cprcferrrd s y s t m valtagrs. t Specifying t h e w valiirs for motor voltsgcs is itnportarrt: For instnnw. motors to opprste on -IltiO-. GWC-, or 18,800-volt systrins should Iw rntcil 4000. (i(iO0. or 1:1,200 volts, resp2ctively. The one-line diagram (Fig. 4.1) shows a t y p i i d method of distributing power in industrial plants and will be used as referenre to identify some portions of the systems and equipment referred to. RATED VOLTAGES OF TRANSFORMERS

Transformer voltage ratings are hased on the no-load values, and the ratio of primary to secondary rated wltages is equal t o the turn ratio. The transformers have a voltage rating for each xindiiig. These are

VOLTAGSSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

195

the voltages a t which characteristics are measured. What then are standard transformer voltage ratings for industrial plants? First, consider primary or master unit substations and transformers which step down from some voltage above 15 kv to plant primary distribution voltage, which is generally below 15 kv (see Fig. 4.1, top substation). The standard primary-winding ( A , Fig. 4.1) voltage ratings of this class of substation and transformers are 110, 67, 43.8, 34.4, and 22.9 kv. These are the actual transformer-minding ratings. They are derived from the old rating structure based on secondary ratings in multiples of 115 volts. When secondary ratings were boosted to multiples of 120 volts, the high side rating was raised to maintain the same turn ratio. For instance, 33,000-2300 volts was once a standard rating. Thc corresponding present-day transformer would be rated 34,400-2400. The familiar designations 115, 69, 46, 34.5, and 23 kv refer to the classes of insulation used with these transformers. Secondary-winding ( B , Fig. 4.1) voltage ratings of this class of industrial substat,ions and transformers are 13.8, 13.2, 12, 6.9, 4.8, 4.16, and 2.4 k v . S e x t consider transformers in load-center unit substations (see Fig. 4.1, bottom substation) used in t,he industrial plants for stepping down from plant primary distribution voltage to utilization voltage. As stat,ed above, the plant, primary voltage is usually less than 15 kv. Therefore, the list belox includes only voltages below 15 kv. The primary-winding (C, Fig. 4.1) voltage ratings of load-center unit substations are 13.8, 13.2, 12, 6.9, 4.8, 4.16, and 2.4 kv. . Note that the primary voltage rating of this class of transformers (bottom, Fig. 4.1) i s the same as the secondary voltage rat,ing of the primary substation transformers (top, Fig. 4.1). The voltage ratings of secondary substations in the plant which supply motors and other utilization equipment are divided into two classesthose for serving utilization equipment above 600 volts and those for serving utilization equipment below BOO volts. Standard rat,ings are listed in Table 4.2. TABLE 4.2

Transformer Secondary Voltage Ratings ( I ) , Fig. 4.1) Supplying Utilizoti0n Equipment Roted

Above 600 Volts, Kv 6.9 4.8 4.16 2.4

Supplying Utilization Equipment Roled 600 Volt, or 0e1ox. Volt. 600 IY or delta1 400 IY or delta1

240 208Y/l20

All standard unit substation transformers have taps in the primery winding to allow compensation for voltages that vary from the transformer rating. The most common are four 255 per cent taps, two above

196

VOLTAGkSTANDARD RATINGS, VARIATIONS. CALCULATION OF DROPS

aiid two below normal, giving a total adjustment of plus or minus 5 per cent,. With these t,aps in the primary winding, a transformer actually has five different ratios. I t vould he very cumbersome to refer to all five of these ratios in all discussions; therefore, when in the following discussion a transformer is referred t o as having, for example, a rating of 2400-480 volt,s, the discussion will apply equally well whether the transformer is operated 00 the cenher t a p or other taps. Regardless of the tap used, the t,raiisformerwill still be referred to as a 2100-480-volt transformer, Comhined light arid power systems are frequently used where motors are supplied a t 180 volts, for example, and lights are supplied at 120 volts from the same 480-volt system, using dry-type transformers. The standard primary volt,age ratings for t,hese light,ing transformers are 600 volts, 480 volts, arid 240 volts, aiid the standard secondary vohage ratings are 208Y/120 volts and 120/240 volts. Two rated kva 5 per cent below normal t,aps are provided in these transformers t,o allow for operation of 120-volt lamps near t,heir rated voltage when the voltage on the 480-volt system is below 480 volts as it normally vill be.

TRANSFORMER VOLTAGE REPRESENTATIONS

Transformer voltage designations become rather complex. For illstance, windings may have series-parallel connections. Or they may be designed for connection line-to-neutral on higher rated volt,age systems, such as 3400-volt transformers which are suitable for line-toneutral operation OIL 4160-volt systems. These and other complex arrangements make exact identification desirable. These variables in t,ratisformer voltage ratings have long been expressed by various symbolic met,hods. Such methods are essential because t o fully describe the \\-indings of transformers often would require a fairly lengthy paragraph. However, t o bc of any value a transformer rating so expressed should meao the same to everyone. To further a consistent use of symbols, hot,h KERIA and ASh standards have been established t,o rci~ommenda standard transformer “shorthand.” Four symbols are used: the dash (-), t,he slant (/), the X, and the Y. In general terms, their uses are as follows: Dash (-). Used to separate the voltage ratings of separate windings in a specific transformer. Slant ( I ) . Used t o separate voltages t o be applied to or obtained from the same windiug. X. Used to designate separate vokagcs obtainable by reconnection of the coils of a winding in series or multiple combinations. Y. t!sed t o designat,e a winding t,hat is Y-connected. The absence of

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

I97

this symbol in a three-phase transformer rating indicates that the winding is delta-connected. The use of the dash, slant, and Y can he easily illustrated by the voltage rating of the transformer for a typical load-center suhstation. 4160-480Y/277: Note that this meaus the 4160-volt high-voltage winding is delta-corinected while the 480-volt winding is Y-couiiected with t,he neutral brought out. A three-winding t,ransformer might have this voltage rating: 13,800-2400-480Y/277. In three-phase transformers the slaut is ofteri used to indicate wiiidiiigs connectable either in delta or Y. For iiistauce, a 2400/4lCiOY windiug can be couiiected either for 2400 volts deka or 4160 volts Y. Xote that the delta voltage is expressed first. When a Y-connected winding has the neutral brought out it is siguified like this: 2OSY/lZO. Here the line voltage is expressed first, fol1oir.d by the line-to-neutral voltage. If the neutral is brought out with reduced insulation, that fact is shoivu by 208 Grd Y/120. Another use of the slant is to indicate taps, especially 011 single-phase transformers. For instance, a 240-volt wiuditig with a midtap is expressed 240/120. When a single-phase t,ransformer with a series-multiple winding is vound to be suitable for three-wire service on the series conoectioii, it is designat,ed 120/240. When a winding has several taps close to the rated volt,age, it is cust,omary to specify them as illustrated in t,his specific case: four 255 per ceut rated kva taps, t x o above and tI5-o below rated voltage. The X symbol is used to separate t,he volt,ages obtainable in a seriesmukiple minding not, suitable for three-wire operation. For example, a minding rated 120 X 240 can be connected with t,he coils in parallel to obt,ain 120 volts or Tr.it,h the coils iu series for 240 rolt,s. RATED VOLTAGES OF GENERATORS

Siiice the generator is a source of elect,ric poir-er aud is ofteu i u parallel wit,h primary substation transformers (see Fig. 4,1), its voltage aud ('oiiscquently its rat,itig is in practically all cases the same as the transformer in a giveu voltage class. Listed in Table 4.3 are the three-phase generator ratings that, are recommended by the latest EEI-SE5I.i report. TABLE 4.3 208Y/120 "Olt. 240 volts 480 volts 600 volts

* Ratings

Generator Voltage Ratings* 2,400 volts 4,160 volts 4.800 volts 6,900 volts

13.800 "011. 14,400 volts

of 11,500 and 12,500 volts are n s ~ dfor genrrators on smnr rstablislird hut are, not rrrommmdrd for nmv systim~s. Thc corrcsponrling trnnsfornii,r rating is 12,000 w i t s and transformcr taps sllon for paralirl oprration. systems

198

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

The 14,400-volt rating has been adopted largely in large generating stations where the input is transformed up to higher voltage in a unit transformer generator arrangement (see Fig. 4.2).

2

mTwI

4 I

FIG. 4.2 HIGH VOLTAGE BUS

Unit transformer generator arronge-

merit,

RATED VOLTAGES OF MOTORS

A t the other end of the system are the motors, and their rat,ings reflect the fact that voltage at utilizatioii equipment is somewhat loirer t,haii a t the sources of power because of voltage drop. Single-phase motors are usually rated at 115 or 230 volts. The standard voltage rat,ings of polyphase motors are given in Table 4.1. TABLE 4.4

M o t o r Voltage Ratings

110 "0111

550 "011.

6,600 Volt.

208 volt. 220 wit.

2,300 ~011s 4,000 ~ o l t i 4,600 volts

I1.000 volt,

440 rolls

13,200 volts

hlot,or-cotit,rol equipment has the same voltage rating as the associated motor. RATED VOLTAGES OF LAMPS

Inrandescent lamps are standardized at 120 volts. Higher voltages have not in general heeo found sat,isfactory. Fluorescent lamps offer a wider range of operation and are commotily rat,rd a t 118, 208, 230, and 265 volt,s (for line-t,o-neut,ral on 480-volt systems). OTHER APPARATUS

Some other types of equipment such as capacitors and industrial heating equipment have compromised between the extremes of generator

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

I99

rating and motor rating in a given voltage class. For instance, industrial heating devices are rated at 115,230, 4G0, and 575 volts. Capacitors are rated at 230, 460, 575, 2400, 4800, 7200, 12,470, and 13,800 volts. NOMNAL SYSTEM VOLTAGES

The choice of the numerical value t o represent nominal system voltage is purely arhitrary and does not attempt to indicate an average system voltage. It is merely a name. However, it is very desirable that a consistent practire in designating nominal voltages be followed. When used properly, the nominal voltage should give a good picture of the voltage struct,ure of a system with a minimum of misunderstandings. The standard values for nominal system voltage correspond t,o the ratings of source equipment. TABLE 4.5

Standard Nominal System Voltages Singlo Phase

120 120/240 240 Three Phore

208Y/l20

240 480 600 2,400 4,160

4,800 6,900 12,000

13,200

34,500

46,COO 69.000 115,000

13,800 23,000

Table 4.5 is not complete but is representative of industrial practice. To repeat, it is extremely important to identify properly the voltage rating of each piece of apparatus in a system as well as to identify the nominal system voltage. The voltage ratings of the various pieces of apparatus, as ran he seen from the foregoing, may be different even though the apparatus is for use on the same given voltage class system. Therefore, correct identification of each piece is of paramount importance. For example, if one is buying equipment to supply a 180-volt system, the secondaries of t,he transformers should he specified as -180-volt rating. The motors and control should be specified as 440-volt rating. The system nominal voltage is referred to as 480 volts. Other apparatus on this system may have different voltage ratings. For example, capacitors would be rated 460 volts; heating equipment would be rated 460 volts. It is also important to remember that transformer and generator voltage

200

VOLTAGE- STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

ratings are always higher than utilization-device ratings. This is logical because the transformer voltage ratings are the no-load voltage ratings, and as load is applied to the system the voltage drops to near the nameplate rating of the lower rated utilization apparatus.

VOLTAGE SPREAD AND FLICKER REQUIREMENTS* STEADY-STATE VOLTAGE REQUIREMENTS

An ideal electric power system is one which will supply constant frequency and volt,age at rated name-plate value t o every piece of apparatus in the system. I n modern power systems, frequency is a minor problem. It is impractical, however, t o design a power system which will deliver absolutely constant rated name-plate voltage to every piece of apparatus. Since this cannot he attained, what are the proper limits of voltage variation in a n industrial plant? These should be determined by the characteristics of the utilization apparatus. First, certain definitions are essential to underst,arid clearly the discussion of this problem. Voltage Spread. Voltage spread is the difference between the maximum and minimum voltages which appear at any location in a system under riormal operating conditions. Voltage spread is not intended to cover momentary voltage changes uf a transitory nat,ure such as those due t o switching surges, motor starting, welders, etc. The first part of this discussion is primarily concerned with voltage spread a t utiliaatiori equipment. This is the diKercnce between the maximum and minimum voltages a t the terminals of the utilization equipment under normal system operating conditions (Fig. 4.3). Maximum values usually appear during light load and minimum values a t full load on the electric system. Another important type of voltage spread is primary or supply voltage spread which is the difference between the maximum and the niinimum voltage a t the service entrance or plant primary bus of a particular plant under normal operating conditions. Voltage Zone. Voltage zone is the envelope of all voltage spreads for a particular voltage class of system. For any specific voltage class designated by a nominal system voltage there inherently exists an appreciable range of operat,ing voltages between the systems having the highest and lowest voltages for this class. Countrywide, this zoue is larger thaii the voltage spread at, ariy one location because of recognized differences in practices of different companies.

* The data in this sretion arc l a r ~ c l yadapted from an AIEE Industrial Power System Coinmittre 1Lpurt. Industrid Voltag- Ilrquirpmeats, Elec. Eng., vol. 6 i , 1948, pp. 358-374.

3.3 7.

z

5 ,

PRIMARY SYSTEM

LONGEST SECONDARY FEEDER

_ _ _ _ _ _ _ _ _ _ _ _NO_ LOAD _ _ _VOLTAGE _~ _ - - - - - - - - - - - - - - 480 2500-PRIMARY VOLTAGE SPREAD. NO LOAD TO F U L L LOAO AT PLANT SERVICE ENTRANCE

:z:

s

r

2400-

-

0

0 Y

Y

> E

2 N 0

1

SPREAD IN SECONDARY SYSTEM

v) Y

k9 2

E >

TRANSFORMER VOLTAGE DROP

Q

2300-

(L

P

FEEDER VOLTAGE DROP

T. Y < 0

NO LOAD VOLTAGE

5

> 0 PRIMARY VOLTAGE SPREIO, NO LOAD TO F U L L LOAD AT PLANT SERVICE LNTRANCE

> 2200-

*

I

I

E

L

I

0

MINIMUM FULL LOAD VOLTAGE

TRANSFORMER VOLTAEE DROP

C U R V E A - T R A N S F O R M E R OPERATING ON HIGHEST TAPRATIO 2 5 2 0 - 4 8 0 VOLTS AT NO LOAD. CURVE 8

TRANSFORMER OPERATING O N LOWEST TAPRATIO 2260-480 VOLTS AT NO LOAD.

FIG. 4.3

Examples of voltage zone, spread, and drop.

VOLTS

VOLTAGFSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

202

difference in voltage in various parts of the power system. The other cause is primary voltage spread a t the service entrance of the plant.

EFFECT O F VOLTAGE DROP

To show the effect of voltage drop in a plant it will be assumed that the primary voltage is maintained a t a constant value regardless of plant load. The simple circuit shown in Fig. 4.4 will be used as an illustration. The primary voltage is assumed to be of such magnitude that the secondary voltage on the transformer is 480 volts a t no load. Referring to Fig. 4.5, a t extremely light load there is essentially no voltage drop through the transformer or in any of the secondary circuits connected to the transformer. Consequently, the voltage is substantially the same throughout the plant, and any lights or other incidental load connected a t this time is subject to practically the no-load voltage. It is particularly significant a t this point to recognize that transformer voltage ratings are the no-load SECONDARY BUS

TRANSFORMER CIRCUIT

FIG. 4.4

480

.-

Typical industrial plant power circuit,

400 VOLTS

ZERO VOLTAGE DROP A

2

2

470-

:rp 4 6 0 ~

_ _ _ _ _ ~. TRANS FA NO

y)

3

9

]----____________________

----

460

TOTAL VOLTAGE

450

FIG. 4.6 spread.

l,z

LOAD VOLTAGE-480 VOLTS VOLTAGE DROP VOLTAGE DROP IN THRU 15 VOLTS TRANSFORMERSECONDARV FEEDER-IOVOLTS

sE~!~48oro~?p~"2~Ts G?!E?

IN BRANCH DROP CIRCUIT-

--- -___

_________________-__

Full-load voltage conditions for circuit shown in Fig. 4.4.

5 VOLTS

_

A

No primary voltage

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

203

ratios. For example, a transformer rated 4160-450 volts will produce 480 volts a t no load with 4160 volts applied to the primary. When load is connected to the transformer, current flows, and this causes a voltage drop in the secondary circuits as shown in Fig. 4.6. At t,he secondary bus the voltage drop caused by the current flowing through the transformer is assumed to be 15 volts. With constant primary voltage the secondary bus voltage varies from 450 volts a t no load to 465 voks at full load--the voltage spread a t this point is 15 volts. There are assumed additional drops of 10 volts in the secondary feeder and 5 volts in the branch circuit, making a total drop to load A of 30 volts. If the lowest voltage in the plant exists a t load A , then the maximum voltage spread is 30 volts (450 a t no load to 450 volts a t full load, or 30 volts). In designing an industrial power system the voltage spread should be kept t o a minimum consistent with reasonable first cost. If the spread is too great,, the voltage may be too high a t light load, causing equipment operating during that period to burn out, or voltage may he too low a t full load a t much of the utilization apparatus, impairing the performance and reducing the production obtained from the equipment, The second cause of voltage spread is the primary voltage spread a t the plant service connection. This may be caused by voltage drop in the primary system, or it may be due to regulation of the primary system by voltage regulators. To show the effect of primary voltage variation, assume that the primary voltage drops as load comes on in the plant. The transformer taps have been selected so that the no-load voltage is 450 volts as in Fig. 4.5. When load comes on the power syst,em,the same voltage drop occurs as in Fig. 4.6, but in addition, the primary system voltage is assumed t,o drop sufficiently to cause an additional 10-volt drop in the vokage at the secondary of the transformer. This primary voltage spread adds to the total voltage spread in the plant, making the spread 480 to 440 volts or a total of 40 volts as is shown in Fig. 4.7 instead of only 30 volts as shown in Fig. 4.8 where there was no primary voltage variation. The primary voltage spread may not always be in the direction shown in Fig. 4.7. The primary voltage may rise when the load comes on because of voltage regulators in the primary feeder circuit or because of other voltage regulators in the primary power system. This voltage rise of the primary reduces the voltage spread in the plant, as shown in Fig. 4.5. Very weak primary systems with a high drop or regulated primary systems whose load cycle does not coincide with the load cycle of the plant may cause excessive voltage spread in the plant-beyond the limits shown in Table 4.9. This is illustrated in Fig. 4.9. Automatic voltage regulation is required in such cases to bring the voltage spread within the limits shown in Table 4.9. Changing transformer taps to increase the vo1t:ige a t full load will not solve the problem because that will increase the no-load voltage beyond 450 volts.

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

204

480-

----

470-

VOLTAGE DROP THRU TRANSFORMER 15 VOLTS

y1

9 450

VOLTAGE DROP IN SEOWDARI FEEDER VOLTAGE DROP IN

-

TOTAL VOLTAGE SPREAD 480 TO 440 VOLTS 140 VOLTS1

_________________________________

440

I

-______

---

FIG. 4.7 Full-load voltage conditions for circuit shown in Fig, 4.4 with 10 volts (on 480volt baris) primary voltage spread. Primary voltage varies from maximum at no load to minimum a t full load.

VOLTAGE DROP I N SECONDARY FEEDER10 VOLTS

FIG. 4.8 Full-load voltoge condition3 for circuit shown in Fig. 4.4 with 10 volt. (on 480volt basis) primary voltage spread. Primory voltage varier from minimum at no load to maximum at full load.

_____

470

;1

LOAD VOLTAGE

4SO

PRIMARY VOLTAGE SPREAD

440

___---

430

VOLTAGE DROP THRU TRANSFORMER

420

-

480 VOLTS

~

460;

G

_________NO

2s__vw3

. 410 . J

TOTAL VOLTAGE SPREAD 4 8 0 TO 410 VOLTS 170 VOLTS)

- 40

VOLTS

VOLTAGE DROP IN SECONDARY FEEDER lo

VOLTAGE DROP

V O L T A G b S T A N D A R D RATINGS, VARIATIONS, CALCULATION OF DROPS

205

EFFECT OF VOLTAGE SPREAD O N UTlLlZATlON EQUIPMEN?

G e n e r a l Effects. Whenever the voltage a t the terminals of a utilization device varies from name-plate rating of the de\.ice, something is sacrificed either in life or performanre of t,he equipment. The effert,may be minor or serious, depending upon the chararteristirs of the device, how the device is applied, and the amount the voltage deviates from the device rating. KESIA Standards provide for rert,ain tolerances whirh may he taken advantage of without seriously affertiiig the performanre of the apparatus. However, with usbge of electrir pover for precise operations, there is often a major sacrifire in produrtion for volt,age variations of considerably less than given in t,he NERlA Standards. So that the plant engineer can better judge the effect 11f vokage variation on t,he electric equipment in his plant, the rharacteristics of many commonly used derires are given here. I t is these rhararteristirs rvhirh have been used as a st,arting point for establishing the desired voltage spread of Tables 4.8 and 4.9. Effect on Induction Motors. Induction motors are the most rommoir utilization derires in industrial plants. Thr variatioii i n rharactrristiw as a function of voltage for the widely used inductiotr motors is shoivn i n Table 4.G. The material in this section deals only n-ith the cffert 011 motor chararterist,ies of rhaiiges in voltage magnitude. The effect, of unbalanced voltages is also very importatit and shonld he rotrsiderrd. The rurrent may hecomc esressive for only a small voltage iuihalanre. The XEBIA St,andards should be consulted for detailed information on this subject,. Principal Effects of l o w Voltage on Induction Motors. The most significant effects of too lox voltage are reduction in starting torque a t i d increased full-load t,emperature rise. The redurtion of st,arting torque may be significant i n mot,or applications driving high-inertia rqnipmeirt. The lower torqne i d 1 result, in longer armleration periods. Torque mot,ors are also very materially affected hy redured voltage as thi. torque decreases as the square of the voltage; thus a t 10 per reut helow normal voltage, the torque is redured 19 per cent. The increased heating at low voltage aiid full load rediirrs thr lifr of the insulat,ion. Principal Effects of High Voltage on Induction Motors. The most, significant efferts of too high voltage are inrreased tnr(lue, inr,rrasrd starting rurrent, and decreased p o r e r factor. The increased torque may muse rouplings to shear off or damage t o driven equipment. Increased starting curretit raiiscs greater voltage drop in the power system, henre increases light, flirker. Uecreased p o ~ v z r factor is particularly disadvantageous where power-fartor peualty rlanses

206

VOLTAGE-STANDARD RATiNGS, VARIATIONS, CALCULATION OF DROPS

TABLE 4.6

General Effect of Voltage Variation on Induction-motor Characteristics

I

Voltage Variotion

90% voltage

Starting and maximum running torque... Synchronous speed.. Per cent d i p . . Full-load speed.

................. Decrease 19%

Functionof voltage

(Voltage)’

.......... No change Cons1.nt ............... Increase 23% 1 (voltagel~ .............. Decreore 136% ISyn. speod--.llpl Efficiency: Full load.. ................ Decrease 2 points .............. 9% load. .................. Proclicolly no change .............. )i l o a d . . ................. Increase 1 to 2 point$ .............. Power faclor; Full land.. Increase 1 point load.. Increase 2 lo 3 point! load. Incrcoie 4 lo 5 points Full-load ~ u r r e n t . Increase I1 Yo Starting w r r e n l . . Decrease 10 to 12% Voitoge Temperature rise, full load. Increose 6 to 7 C Maximum torque capocity.. IV0ltogeJ~ Decrease 19% Magnetic n0ire.m load in parliculor.. Decrease slightly

I( 36

................

................. .................. ............. ............. ..... .... .....................

110% voltage

InCreOle 21 N o change

Decrease 17%

Increase 1 % Small increo*e Procticdiy no change Decrease 1 to 2 points

.............. Decrease 3 points .............. Decrease 4 points .............. Decrease 5 lo 6 points .............. Decrease7% Inc,eo.e 10 to 12% .............. Decrease I lo 2 C

..............

Increa3e 21 %

Increase slightly

This table s h o w gencral effcets, which will vary somewhat for specific ratings.

are applied by the utilities. The higher the motor voltage rises, the lower the power fartor mill become. This may result in a greater penalty and hence a higher power bill. While the temperature rise at full load on standard motors decreases slightly for moderate overvoltages, the temperature rise may increase on certain types of sperial motors a t even very small overvoltages. Overvoltages of 10 t o 1.5 per cent have caused numerous burnouts on special four-speed grinder motors. Motors rated for intermittent load are also materially affected by overvoltagcs. While marry drive applications are not seriously affected by voltage deviations as much as plus or minus 10 per cent from rated voltage, there are import,ant applications that are. Effect on Synchronous Motors. The effect of voltage variation on the performance of synchronous motors is similar t o that on induction motors. However, while t,he starting torque varies as the square of the voltage, the maximum or pull-out torque varies directly with the voltage. From the above discussions it will be noted that, in general, voltages slightly in excess of motor name-plate rating have less detrimental effect

V O L T A G k S T A N D A R D RATINGS, VARIATIONS, CALCULATION OF DROPS

207

on motor performance than voltage helow the name-plate rating. This is one of the bases on which the voltage spreads in Table 4.9 mere determined. A s a n example, the figures show a recommended spread of 420 t o 180 volts for the 480-volt nominal system voltage, which is approximately 4 per cent below and 9 per cent above the 440-volt motor rating. Effect on Incandescent lamps. The light output and life of incandescent filament lamps are critically affected by the impressed voltage. I n Table 4.7 is shown the relationship of lamp life arid output t o voltage for a vokage range from 80 t o 120 per cent of rated voltage. I n general it may be said that for incandescent filament lamps a 1 per cent deviation from rated voltage causes a change of 3 t o 335 per cent in light output. It can be seen from Table 4.7 that a 10 per cent reduction in lamp voltage results in a 30 per cent reduction in light output. In other words, when the voltage is 10 per cent low, the investment in the lighting system is working at only 70 per cent efficiency-thus, 30 per

! i! 3 a

2

0 c

3

PER CENT NORMAL VOLTS

9 a FIG. 4.10 Characteristics of large gar-filled incandescent type C lampr. average of many lampr.

There are the

208

V O L T A G b S T A N D A R D RATINGS, VARIATIONS, CALCULATION OF DROPS

cent of the investment is lost. With an overvoltage of 10 per cent the lamp-life is reduced t o less than oue-third~-t,hus lamp-replacement costs are three times as great as a t normal voltage. Other dat,a arc shown in Fig. 4.10, from which it, should be noted that the lumens per watt., or lamp efficieilcy, rises sharply at voltages above 100 per cent. I n some cases, operating eronnmies result from hurriing lamps at higher efficiency and short life, or vice versa. TABLE 4.7

Effect of Voltage Variations on Gar-filled Incandescent-lamp Choracteristics

Socket voltage

Per cent rated voltage

Per cent rated lighl output

Per a\suming 4"(i volts on bus H with same load,

450 5-kv cahlr amperes = 193 X 4%

v =

4 3

X 204 X O.GO2

Hns B voltage

=

204

=

212 volts 480 (3900 - 212) 4160 =

Secondary load voltage, assuming q20 volts a t load, 250 0.420 X Cable resistance = 0.0072 ('able reactance = 0.0090 Load amperes

v =

ros B

=

0.7,

sill B =

c = 4 3 = = =

=

=

4

344

4x r ( x cos B + X sin 8 )

0.714

x

344(0.007"

x

0.7

+ O . O O ~ Ox 0.714)

4X 344(0.00504 + 0.00643)

4 3 X 344 X 0.01147 G.9 volts

Load voltage = 425 - G.9 = 418.1 volts Since the most i.ritii.al feeders n-ith respect, t o voltage drop have been selerkd, the ralrulated load voltages a t hus A , bus B , arid at the secondary-load trrminals provide sufkieirt information t o analyze the system from the standpoint of voltage drop. Xct,ually, the 480-418 voltage spread at the serondary-load terminals iiidicates that the system is on the horder line and should he stiffened, possibly'hy using a larger 5-kv feeder cable. Howevw, this is beyond the scope of this problem, which i s mcrcly iriteiidcd to out,liiic t h e method of det,ermitriiig voltage drop. CALCULATION OF VOLTAGE DROPS DUE TO MOTOR STARTING INTRODUCTION

I t is rharactrristic of most a-c motors that the riirrent, which they draw startiirg is mu(.h higher t,han t,heir rrormal running ( w r e n t . Syni~hronousand sqnirrel-rape iudi~rtionmotors started 011 full voltage may draw a c u r ~ w i tas high as sevt!ii or eight t,imes their fnll-load running rurrcnt. This sriddeir increase in the (.usrent, drawn from the power system may r c s i i l t iii csressive drop i n volt,age unless it is considered in oii

VOLTAGbSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

249

the design of the system. Folloii-ing are methods for ralculatiug the voltage drop which results from startiug of three-phase induction aud synrhronous motors. M O T O R - S T A R T I N G METHODS

The motor-startiug kra, imposed on the power-supply system, and t,he available motor torque are greatly aKected by the method of starting used. l'ahlc 4.13 gives a conlparisoii of several common methods. Full-voltage Starting. This method usually provides the most torque hut muses the greatest load t o be applied to the system. The load applied equals (at motor rated voltage) the full-voltage starting kva of hhe mot,or. Frill xwlt,age is the least espeosive method of startiug. The full-voltage starting kva of syurhroiious and squirrel-cage indurtion motors ruuges from 230 to 800 per cent of their full-load h a input. The latter is approsimately cqual to t,hc horsepower rating of induction and 0.8-pomr-factor syirrhrorious motors and is approximately 80 per eelit of the horsepower rating of 1.0-poiver-factor syrichronous motors. If the starting curreut in ampercs is kno\vu, the startiug kva (of threephase motors) may he ralrulated from the formula Kva = 1.73 X amperes X

line-to-line volts

looo

Reactor Starting. With t,his method, a reactor is connect,ed in series with the motor aud is shorted out when the motor approaches full speed. 4 reactor starter redures the line current in proportion t o the tap used. For example, with a 50 per cent tap, the current is cut in half. The torque is reduced hy the square of the tap used. Hence, the torque is reduced more rapidly than the line current. Reactor st,arting is commouly used for large motor-generator sets. Resistor Starting. Resistor starting is similar t o reactor starting except that a resistor is used in series with the motor, instead of a reactor. The torque available for a given reductioti in startiug current is the same as with a reactor. The hie-voltage drop may be somewhat less because of the better power factor of a resistance-st,arting load. Resistor starting seldom offers a cost advantage, except wheu several steps are required, t o meet limitations established for the maximum kva applied at any oue step. Power companies sometimes establish such limitations. Use of several steps may permit a generat,or voltage regulator to restore voltage between steps. It also tends to make light flicker less noticeahle, even if most of the drop is in the distribution system and cannot be reduced by regulators.

250

VOLTAGE- STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

Autotransformer Starting. If an autotransformer starter is used, the line current is reduced approximately as the square of the tap setting. For example, if an autotransformer at a 50 per cent tap is used, a motor starting load of 100 per cent of the rating of a generator will be redured to about 25 per cent. Table 4.13 shows 30 per cent because it allows for autotransformer magnetizing current. Autotransformer starting may cost more than reactor starting, but may be needed to provide adequate torque. The tap selected should always be high enough to accelerate the motor to a speed a t which the current will not be excessive after transfer t o the running connection. If the load torque is high a t the time of transfer to the line, a high transient inrush for a few cycles may occur at this time even if the speed is high. This is seldom sustained long enough to cause troublesome voltage dip, but may cause tripping of instantaneous overcurrent protection for the motor circuit.

TABLE 4.13

Comparison of Motor-starting Methods Line voltage = motor-rated voltage

Type of

starter* Motor voltage

line voltage

........................ ........................ ........................ ........................

Full.roltage stmrter. Autotransformer: 80 Per Cent t o p . . 65 per cent tap.. 50 per cent tap.. Resistor storm, single step [adjusted for motor voltage to be 80 per cent of line voltogel Reoctor; 50 per cent tap.. 45 per

5 50 Y

u 40 IL

f 30 20

ov

10

0

2

3

4

5

6

TIME- SECONDS

MOTOR-STARTING XVd*IDDPfR

A B N

CENT OF

DENEMTOR RATING

-

NO INITIAL LOeiD ON GENERATOR

-

5 0 PER CENT INITIAL LOAD ON GENERATOR

- NO REGULATOR

FIG. 4.38

Typical generator voltage behavior.

151

V O L T A G S S T A N D A R D RATINGS, VARIATIONS, CALCULATION OF DROPS

motor is essentially similar, up to the time of pull in. I n the case illustrated, a full-voltage starter is used, and the full-voltage starting kva is ahout 100 per rent of t,he generator rating. I t is assumed that the generator is provided with an automatic voltage regulator. Curves .A and R show the performance, with the regulator operating, for init,ial loads on the generator of zero and 50 per cent, respectively. The minimum voltage is about 75 per cent and is not affected much by the iriitial load. This is typical with most initial loads which consist of a combination of lighting loads and partially loaded iuduction motors. The voltage regulator restores the voltage ton-ard normal in about 2 see. At, this time the motor is usually st,ill at low speed and drawing a high current. The initial load on the generator has an important effect on the value t o which the voltage is restored by regulator action. This is illustrated by curve B , for whirh the voltage is restored by the regulator to only about 85 per cent of normal. This restored voltage is the voltage available for breaking away and accelerating the motor. When the motor comes up to speed, its current becomes much less, so that t,he regulator then restores the generator voltage to 100 per rent. The reason the regulator usually cannot restore the voltage to 100 per cent when a large motor is started on a heavily loaded generator is that the exciter maximum (ceiling) voltage limits the available generator excitation. Sometimes it is only necessary to calculate the minimum voltage. In other cases it is also necessary to calculate the restored voltage available for break away and accelerations. Methods of estimating each of these voltages are included. Minimum voltage is needed to determine whether undervoltage devices and contactors connected to the system mill drop out, or running motors stall, during the disturbance. The minimum voltage is also a determining factor in light flicker. The restored voltage is necessary to estimate the torque available for starting the motor. Usually it is sufficient to determine the minimum voltage and the restored voltage based upon the current drawn by the motor at standstill, i.e., upon the locked-rotor current. It is sometimes necessary, however, to determine the restored voltage throughout the acceleration of the motor. Although the current drawn by a motor decreases as it comes u p to speed, resulting in an increasing generator voltage, the load torque may also increase with speed so that a higher voltage is necessary to ensure acceleration. In the case of a synchronous motor i t may be necessary to check the restored voltage at the speed at which field excitation is applied (95 per cent of synchronous speed or higher) to make sure that the motor will pull into step. The pull-in torque of a synchronous motor varies approxi-

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

253

mately as the square of the voltage at the motor terminals just before application of field. Distribution-system Voltage. Frequently there are transformers, lines, or cables between the motor starter and the generator or generators supplying the power for starting. The drop in the transformers, lines, or cables will be additional to the generator drop. Often practically all the drop is in this distribution equipment. The drop in this equipment is not reduced by the action of voltage regulators. Consequently, when practically all the drop is in transformers, lines, and cables, the voltage falls immediately and docs not rerover till the motor approaches full speed. ESTIMATING GENERATOR VOLTAGE DROP

Minimum Voltage. The curves of Fig. 4.39 may he used for estimating the minimum voltage occurring at the terminals of a generator supplying power to a synchronous or squirrel-cage induction motor which is being

254

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

started. The initial load on the generator, if any, is assumed to be of the constant-current type. The three sets of curves shown are for three ranges of generator speed. The generator reactances assumed to apply for each speed group are also given in Fig. 4.39. The curves show the minimum voltage, in per cent of the initial generator voltage, plotted against the "motor-starting kva" in per cent of generator rated kva. The "motor-starting kva" is the kva which would be drawn by the motor being started if the generator voltage were maintained at rated value. Since there is a drop in generator voltage, the actual kva drawn by the motor will generally be less than the value defined above, but the effect of this is taken into account by the curves. The several curves in each speed group-except those marked N a n d Eapply for various values of a factor K. This factor is the exciter response in volts per second divided by the exciter voltage for rat,ed generator voltage at rated load and multiplied hy the generator open circuit field time constant in seconds. Approximate values of K are given in Fig. 4.40. The values of Ii in Fig. 4.40 are based on the use of a self-excited excit,er controlled by a direct-acting rheostatic voltage reguhtor (such as the

GENEIIbTOR e I T E O K"& ~

W

T DIRECT-CONNECTED " EXCITER

.....~ WIT" ~ BELTED ..

EXCITER NUMBERS ON CURYES ARE R P N

NUMBERS I" BRACKETS &RE EXClTER R P H

*

FOR "IMIAITION OF EXClTER RESPONSE WlT" GENERATOR IN1TIAL L o l o lNlTlbL LOAD (PER C E N T , UULTlPL" I( B" (00 ,70 75 I55 50 I"5 25 I25 HULTlPLlERS TO *ILLOW

0

FIG. 4.40 Typical valuer of performance factor K for (I-c

,oo generators.

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

255

A-C GENERATOR

EXCITER FIELD

GENERATOR VOLTAGE REGULATOR

FIG. 4.41

1

u

''lL

Excitotion system for a-c generator.

General Electric Company Type GDA) as shown in Fig. 4.41. In this system the response of the exciter depends not only upon its design hut also on the setting of the exciter field rheostat. The latter is determined by regulator requirements. The rurves of Fig. 4.40 are based on a setting of the exciter field rheostat which makes available a maximum generator field current of 120per cent of its rated value. The K fartors given by the curves are typical only, and in an individual case K may vary considerably from the value shown. The curves of Figs. 4.39 and 4.40 allow an estimate to be made of the generator miuimum voltage directly from the generator kva rating, the generator speed, the exciter speed, and the motor starting kva. If guarantees of performance are required, a study based on romplete data should be made considering the characteristics and adjustments of generator, exciter, regulator, exciter rheostat, initial load, and the motor being started. Restored Voltage. The curves of Fig. 4.42 may be used for estimating the restored voltage of a generator, that is, the voltage attained after the regulator has acted to apply maximum excitation current to the generator (or has restored the voltage to its initial value) following the starting of a squirrel-cage induction or synchronous motor. The curves show the restored voltage in per cent of rated generator voltage plotted against the kva which would be drawn by the motor being started if rated generator voltage were maintained. The several curves apply for various values of initial load which is assumed to be a constantcurrent load of 0.8 lagging power factor. The excitation system is assumed to be such that a maximum excitation current of 120 per cent of rated generator field current can he obtained. If guarantees of performance are required, a study based on complete data should be made considering the characteristics and adjustments of generator, exciter regulator, exciter rheostat, initial load, and the motor being started.

156

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION

MOTOR S T I R T l N G

I("&

OF

DROPS

I N P E R C E N T O F G E N E R l T O R K Y A A T R A T E O O E N E R ~ T O AVOLThQE

( B A S E D O N Y l x l Y u Y EXCITATION-IZOPER C E N T O F R A T E D G E N E R A T O R F l E L O C U R R E N T I NOTE: RESTORED VOLTAGE E W I L S VALUE READ FROM CURVE OR THE INITIaL YOLTlGL (REGULATOR SETTING1 WHICHEVER I S LOWER

FIG. 4.42

Restored generator vollage.

Advontages of Voltage Regulators. Figures 4.38, 4.39, and 4.42 show dashed curves, marked N , which indicate the results t o be expected if no regulator is used. It is apparent that regulators are very beneficial. They practically always justify their cost whenever the starting of large motors is involved. For example, consider a 480-volt 125-kva 1200-rpm generator. From Fig. 4.40 this may have a performanre factor K of about 1.7 with a regulator. From Fig. 2.39, 110 per cent motor-starting load or 138 kva will cause a 28 per cent voltage dip. This load would correspond t o starting a 25-hp motor at full voltage. T o obtain the same motor-starting performance without a regulator would require a 438-kva generator, because the curve N shows that about 32 per cent motor-starting load will cause a 28 per cent voltage drop if no regulator is used. (138 kva is about 32 per cent of 438 kva.) The 438-kva generator would cost over twice as much as a 125-kva machine. The best and least expensive arrangement mould be t o provide a regulator adding less than 15 per cent to the cost of the 125-kva generator. This mould permit successful starting of the 25-hp motor even against full-load torque and would improve normal generator performance. I n Fig. 4.39 are curves, marked E , which show the performance available when using an electronic exciter or some other very high-response excitation system. It shows there is a definite limit to the improvement

VOLTAGE-STANDARD RATINGS, VARIATIONS. CALCUUTION OF DROPS

157

which can be obtained by greatly inrreasing response; that is, the generator voltage will dip a t least a certain amount before the excitation system can do anything about it. Effect of Initial Voltage. Often the voltage rating of the generator supplying a motor is higher than that of the motor. A 440-volt motor might he supplied by R 480-volt generator and a 2200-volt motor by a 2400-volt generator. In such cases, the motor-start,ing kva should be adjusted t o take this into account,. The kva drawn hy a motor increases as the square of the line voltage. If t,hr startiiig inrush of a 410-volt motor is 1000 kva a t 440 volts, it will be 1190 kva at 480 volts because (480/440)* = 1.19. This is the value which should be used to determine the generator minimum voltage (from Fig. 4.39) regardless of the actual initial voltage. For example, assume that, with an initial voltage of 480 volts, the starting of the 440-volt motor (drawing 1190 kva at 480 volts) causes the voltage t o drop t o 75 per rent of the initial value, or 3G0 volts. If the voltage regulator is set t o hold a voltage of 440 volts, starting of the same motor will produre approximately the same voltage drop in per cent of the initial voltage, i.e., the voltage will drop t o approximately 75 per cent of 440 volts, or 330 volts. This shows that, from the standpoint of the minimum voltage, the regulator should be set t o maintain rated voltage on the generator even though the motor voltage is lower. As far as the restored voltage is concerned (Fig. 4.42), this is not affected by the initial voltage except that the voltage mill not recover t o a value higher than the initial voltage since this represents the setting of the voltage regulator. For example, if the initial voltage (setting of voltage regulator) is 90 per cent of rated generator voltage, the recovery voltage in per cent of rated generator voltage will be as shown by the curves of Fig. 4.42, except that all curves will become horizontal lines at 90 per cent voltage. Effect of Initial load. The voltage curves of Figs. 4.39 and 4.42 were prepared on the basis that the initial load on the generator draws constant current duririg the voltage disturbance. This sort of load characteristic is representative of many systems and results from the use of induction motors, all of which are not fully loaded. An induction motor at no load will draw a current approximately proportional t o the applied voltage, because the current is principally magnetizing current. A fully loaded induction motor will tend t o have constant kva input since its speed and power factor do not change much with variations in line voltage. Consequently, a fully loaded induction motor will draw more current if the voltage is lower, t o maintain the power constant, A system load consisting of both heavily loaded and

258

VOLTAGbSTANDARD RATINGS. VARIATIONS, UL6UUTION OF DROPS

lightly loaded motors will therefore tend to draw nearly constant current since a lowering of the voltage causes a reduction in the current to some motors and an increase in the current to others. A constant-current type of load will have very little influence on the minimum voltage during motor starting. It will, however, have an important effect on the value of the restored voltage of generators, as previously described. Lighting loads usually have little effect upon voltage disturbances due to motor starting. This is true because lighting loads usually constitute a small proportion of the total load on a generator, and also because of their high power factor. If the system load consists primarily of lightly loaded induction motors, the per cent minimum voltage and recovered voltage will both tend to be higher than indicated by the curves. If the initial load consists entirely of heavily loaded induction motors, the voltage disturbance from motor starting will be more severe than indicated by these figures. Initially connected synchronous motors are beneficial in reducing the disturbance due to motor starting. They are most beneficial when lightly loaded. Therefore, it is helpful to start synchronous motors first in a plant so that they will be on the line to help in the starting of large induction motors later. Synchronous motors will not be helpful, however, if the voltage disturbance is so great as to cause them to pull out of step. Although the curves in this section are based on initial loads of the constant-current type, they may be used for cases involving other types. This is done by adjusting the motor-starting kva by an amount corresponding to the change in current to the initial load, caused by the drop in voltage. The increase or decrease in motor-starting kva is such as to change the motor-starting current, a t the minimum voltage, by the same amount as the change in the lagging wattless component of the initial load. That is, the effect of the initial load is primarily due to a change in the wattless component, and this can be simulated by a change in the motor-starting kva. Since the change in current and the minimum voltage are dependent upon each other, a trial-and-error procedure is involved. The first trial is often sufficient,if the change in current is determined a t the voltage corresponding to the case of a constant-current initial load. For example, consider a generator whose voltage would dip to 75 per cent if a 100 per cent motor-starting load were applied when a 50 per cent constant-current initial load is being carried. If, instead, the initial load consisted of fully loaded induction motors a t 0.8 power factor, the dip would be more severe, because a t 75 per cent voltage the lagging wattless current to the running motors would be increased from 30 per cent of the

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

259

generator rating t o about 40 per cent. This increase could be approximately simulated by an increase of the motor-starting kva from 100 per cent to 113 per cent. This is true because a motor-starting load which would draw 13 per cent of generator rated kva a t full voltage will draw 10 per cent current a t 75 per cent voltage. Figure 4.43 shows the amount by which motor-starting kva should be increased to allow for the effect of an initial load consisting of fully loaded induction motors. Effect of Starting Power Factor. The power factor of most motorstarting loads lies between 10 and 40 per cent. Variations within this range do not materially influence voltage drop of generators. Wound-rotor motors have a starting power factor of about 80 per cent lagging. At this power factor the resulting voltage drop (initial voltage minus the minimum voltage) will not generally exceed 75 per cent of the drop caused by the same kva a t low power factor. Resistor starters

__

PF -R C F N T VOLTAGE

DROP

18 17 U

> Y

I !

g

10 09 08

20

13

15

12

'

07

2

03

LL

06

a w

04

1

25

14

z0 2

F

30

16 15

10 5

05

02

3

01

=

o

0

0'2

04

06

08

RATIO OF

10

12

14

16

18

20

I N I T I A L LOAD KVA MOTOR STARTING KVA

INCREASE MOTOR STARTING KVA BY MULTIPLIER SHOWN BEFORE USING CURVES OF FIG.4.39AND FIG.4.42 ( ! N I T I A L LOAD MAY THEN BE CONSIDERED AS CONSTANT CURRENT T Y P E )

FIG. 4.43

Approximate effect of initial lood consisting of fully loaded induction motors.

260

VOLTAGE-STANDARD

RATINGS, VARIATIONS,

CALCULATION OF DROPS

seldom cause the starting power factor to he high enough to reduce voltage drop greatly, except for the first steps when several are used. Effect of Drop in Generator Speed. Since the power factor of motorstarting kva is low, the amount of kw load applied to a generator is seldom large. Furthermore, the voltage drop, by reducing the electrical output, also reduces the new load applied. For example, a motor-starting load of 100 per cent of generator-rated kva at 0.3 power factor will involve a suddenly applied km load less than 30 per cent of rated kva, or 37.5 per cent of rated kw for an 0.8 power-fartor generator. The speed drop is not likely to be excessive if good governing means are employed. For most motor-starting problems, it may safely be neglected. As speed dips, a corresponding dip appears in the voltage, which is in addition t o the voltage drops considered in this section. For cases where speed dip may be sufficiently great to be important, this should be considered, but calculation of speed drop is beyond the scope of this book. ESTIMATING DISTRIBUTION-SYSTEM VOLTAGE DROP

The voltage drops in lines, cables, and transformers are often as important as generator voltage drop. In fact, they are frequently more important. For example, if the total kva of connected generators in the power system is more than 100 times the horsepower rating of the motor being started, then the generator voltage dip will be less than 1 per cent, and it will be quickly eliminated by regulators. In such a case, however, the motor will probably be supplied through a transformer bank. If the transformer-bank kva rating is only slightly larger than the motor rating, the voltage drop may be quite severe. Voltage Drop of Transformers. The curves of Fig. 4.44 may he used for estimating the voltage drop through typical transformers when starting a synchronous or squirrel-cage induction motor connected to the secondary of the transformer. The secondary voltage on starting of the motor, in per cent of the initial secondary voltage, is plotted against the motor starting kva. The latter is expressed in per cent of the transformer-hank kva rating and is the kva which wouldhedrawnhythemotor being started if rated transformer secondary voltage were maintained. The curves of Fig. 4.41 neglect the effect of primary-voltage drops caused by motor starting. Methods of taking these into account will he explained later. Note that the secondary voltage is plotted in per cent of its initial value. This initial secondary voltage is determined by the initial primary voltage, the t a p setting, and the initial load. It may he determined by measurement or by suitable calculations. It is usually slightly less than the rated secondary voltage.

VOLTACbSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

261

MOTOR STARTING KVA

1% OF B & N I K V A aiT RATED TRANSFORMER SECoNOAR” VOLTAGE1

FIG. 4 44 Transformer secondary voltage

The curves of Fig. 4.44 were prepared on the basis that the initial load, if any, draws constant current during the voltage disturbance. This is typical of a system consist,ing of both lightly loaded and heavily loaded inductiou motors. If the initial load consist,s largely of fully loaded induction motors, the curves of Fig. 4.44 may still be used provided that the motor-starting kva is first multiplied by the fartor shown in Fig. 4.43. The curves of Fig. 4.44 apply for motor-starting power factors in the usual range of 10 t,o 40 per cent. For wound-rotor motors which have a starting power factor of about 80 per cent, the drop in voltage will be about 70 per cent of that shown. Voltage Drop of Cables and Overhead Lines. The curves of Figs. -1.45 and 4.4G may be used for estimating the voltage drop through cables and overhead lilies n-hcn start,iiig synchronous and squirrel-cage induction motors supplied through these circuits. I n using these figures, it is first necessary t o determine the length of the circuit in feet, the initial voltage at the load end of the circuit, and the motor-starting kva a t the iuitial voltage. These quantities are combined to obtain the loading factor .If as follows: motor-starting kva at the initial voltage x (% ci:ri ne):,t M = (initial voltage)2

)

262

VOLTAGkSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

For example, if the motor-starting load were 1000 kva, the circuit 1000 ft long, and the initial voltage 2400 volts, the loading factor M would be

1000 x 1000 = o,1,3 (2400)' Figure 4.45 shows that for this case the voltage drop at the load end of a typical three-conductor cable is 1.5 per cent. This illustration gives data for three circuits: a three-conductor cable, a single-conductor cable, and an overhead line. It will be noted that the voltage drop in an overhead line is greater than that for a cable. If two circuits are in parallel, the drop is equal to that for a single circuit of one-half the actual length of each circuit. The voltage drop in a line or cable depends upon the conductor size and spacing. Consequently, for different cases than those illustrated in Fig. 4.45, the voltage drop may be somewhat different. This is illustrated by Fig. 4.46 showing the voltage drop for a range of circuit configurations. The points corresponding to the circled cases in Fig. 4.45 are circled in Fig. 4.46. Figure 4.46 applies for the condition hf = 1.0. It may be noted, however, that the curves of Fig. 4.45 are nearly straight lines. Hence, the voltage drop for other values of M may be estimated by multiplying the values of Fig. 4.46 by M . This provides a simple method of estimating the voltage drop for motor-starting loads. The power factor of the motor-starting load is assumed to be 0.3 power factor. For conductor sizes above No. 0 Awg, variations over the usual range from 0.2 to 0.4 power factor will not have an important effect on voltage drop. Figures 4.45 and 4.46 are based on a frequency of 60 cycles per sec. Lines and cables for systems operating a t lower frequencies mill have less voltage drop. The voltage drop will be reduced approximately in proportion to the frequency for all couductor sizes above KO. 0 Awg. For smaller sizes, the reduction will he less. Voltage Drop of Reactors. The voltage drop in a current-limiting reactor on starting a squirrel-cage induction or synchronous motor may be estimated from the transformer curves of Fig. 4.44. Current-limiting reactors are usually described as having a certain per cent reactance on a specified system-kva and syst,em-voltage base. The motor-starting kva of Fig. 4.44 should be that drawti at the specified system voltage expressed in per cent of the specified system kva. If the per cent reactauce of a reactor does not lie between 5 and 8 per cent, multiply the motor-starting kva by the ratio X / 5 , where X is the actual per cent reactance of the reactor, and read the voltage corresponding to this equivalent motor-starting kva on the 5 per cent reactance curve.

V O L T A G E - S T A N D A R D RATINGS, VARIATIONS, CALCULATION OF DROPS

263

0

100

" Y

90

10

80

20

70

30

0

0 0.1 02 0.3 0.4 05 0.6 0.7 0.8 0.9 1.0

LOADING FACTOR, M =

~

~

1.1

Q

1.2 1.3 1.4

( L E N G T H IN FT.)

~

(A) 3- CONDUCTOR C A B L E - (NO. 4 / 0 - A W G - I 5 K V )

(6) I - CONDUCTOR C A B L E - ( N O . 4 / O - A W G - 6

IN. SPACING)

(C) O V E R H E A D L I N E - N 0 . 4 / 0 - A W G - 1 5 K V CIRCLED P O I N T S APPEAR O N FIG. N O 4 46

FIG. 4.45

Variation of voltage drop with looding factor M for typical liner and cables.

c:

CONDUCTOR DIAMETER (INCHES) FOR MOTOR-STARTING LOADS OF 0 3 POWER FACTOR LO~DING FACTOR M:

(MOTOR-STARTING KVAI (LENGTH IN FEET1 (INITIAL VOLTAGE)^

*FOR FL4T SPACING, EQUIVALENT TRIANGULbR SPACING'; ADJACENT PHASES

FIG. 4.46

I0

I 2 6 TIMES SPACING BETWEEN

Voltage drop in lines and cables with loading factor M of unity.

264

V O L T A G b S T A N D A R D RATINGS, VARIATIONS, CALCULATION

OF DROPS

Effect of Series Capacitors. Sometimes it is advantageous t o include series capacitors in the distrihut,ion system t o neutralize the reactance of lines, cahles, or t,rausformers. Series capacitors redure voltage drop. The amount of redurt,ion depends upon the raparitor rating. For further informat,ion on series capacitors, refer to Chap. 8. Voltage Drop of Power Systems. Motors are frequently supplied from power systems cotisistirig of complicated uetworks of lines and cables for which a calculation of the voltage drop ~vouldhe difficult. The voltage drop may be est,imated, however, if t,he short-circuit current is known at the point of power delivery. The short-circuit rurrent is usually expressed in kva. When motor-starting kva is drawl from a system, the voltage drop in per cent of the initial voltage is approximately equal to 100 times the motor-startiiig kva divided by the sum of this kva and the short-circuit kva. The motor-starting kva used should be that drawn by the motor if the initial system \&age were maintained. For example, if a 1000-hp motor has a startirig kva of 5000 if initial system voltage were maintained and the system short-cirruit kva is 50,000, the voltage drop will be approximately

5000/(5000

+ 50,000) X 100

=

9 per cent of the initial voltage

In many systems the short-circuit kva varies over a wide range, depending upon the number of parallel h e s that are in service, system interconnections, etc. In such cases the highest short-circuit kva is the one usually determined since it must he the one used in selection of equipment which is t o carry or iritcrrupt the short-circuit current. For calculating voltage drop, oil the other hand, the minimum short-circuit kva should be used since the corresponding operating condition will give the highest voltage drop. The short-circuit kva of power systems varies over a wide range, as shown in Table 4.14. A corresponding variation occurs in the voltage drop produced by a certaiu motor-starting kva. TABLE 4.14

Power-system Short-circuit Kva Usual Range of

System Voltage 2,400 4,160 6,900

Short-circuit Kvo 15.000-1 50,000

25.000-250.000

50.000-500.000

13.800

100.000-1,000,000

23,000 34,500 69,000

150.000-1,500,000 150,000-I,500,000

I 15.000

250.000-2.500.000

I50.000-1,500,000

The method of calrulating voltage drop given above is not applicable at system locations where the short-circuit kva would be appreciably

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION DF DROPS

265

affected by reactance of generators. I t should be used only when the impedances of transmission lines, transformers, reactors, and cables largely determine the short-circuit current. COMBINED VOLTAGE DROP

Series Circuits. Often a motor is supplied through cables, transformers, overhead lines, and generators, all in series. In such cases, the total voltage drop may be roughly estimated as the sum of the voltage drops given by the foregoing illustrations for each of the different parts of the system. However, the simple addition of voltage drops is not quite accurate because addition of impedance in series tends to diminish the current supplied to the motor. For more accurate work, the following procedure is suggested: 1. Determine the voltage drop in the circuit element nearest the motor, neglecting the other elements. For example, for a motor supplied from a generator, transformer, and cable in series, determine the drop in the cable first. 2. Multiply the motor-starting kva by the ratio of the load-end voltage to the initial voltage of the cable just determined. 3. Using this new value of motor-starting kva determine the voltage drop in the next circuit element. In the example selected, this is the transformer drop. 4. Now multiply the motor-starting kva by the product of the ratio of the load-end voltage to the initial voltage of the cable and the ratio of the secondary voltage to the initial secondary voltage of the transformer. 5. Using this new value of motor-starting kva determine the voltage drop in the next circuit element. In the example selected this is the generator voltage drop. 6. Continue the process until all elements in series have been considered. 7. Multiply the initial voltage a t the motor by the product of the final to initial voltage ratios of all the circuit elements. This result is the final voltage a t the load. An example a t the end of this chapter illustrates the procedure described. Parallel Circuits. If several sources are in parallel, the voltage drop is less than if the motor-starting load is supplied through any one of them. To determine the combined voltage drop, it is suggested that groups of similar generators may be treated as a single generator having the same total kva rating and the same performance factor as the individual machines. Transformer banks may also be grouped if they are supplied from the same primary bus and have the same per cent reactance and the same tap settings. To find the combined voltage drop for several parallel sources of different characteristics, it is suggested that the motor-starting load first

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION

266

OF DROPS

be divided equally and the corresponding voltage drops determined. Then a new trial division of load can be made so as to increase the proportion of load carried by the sources with the least voltage drop. Usually oiily one or two trials are required to obtain a sufficiently accurate result. For example, consider the case of a motor which has a startiug kva of 1000 and is furnished with power by a 500-kva generator and a 300-kva transformer bank. The first trial division of load will he 500 kva each. Let us assume that this results in a minimum voltage of 75 per rent on the generator and 90 per cent on the transformer secondary. This means that the generator will actually accept less thari half the load. The drop in the generator is 2.5 times as great as in the transformer. Then assume that the transformer accepts 2.5 times as much load as the generator. This results in 285 kva being accepted by the generator and the remainder, 715, being imposed on the transformer (715 is 2.5 X 285). The voltage drop in the transformer for 715 motor-starting kva will be found to be practically the same as for 285 motor-starting kva applied to the generator. The drop obtained is the combined voltage drop. For the case illustrated, this voltage drop is about 14 per cent. A final check of the amount of voltage drop through each source is advisable, because the drop in a generator does not always vary directly with the amount of motor-starting load applied to it. This is especially true of the restored voltage obtained through the action of voltage regulators. FORMULAS FOR CALCULATING VOLTAGE DROP

The various curves and other data that have been presented allow estimates of the voltage drop due to motor starting to be made quirkly with minimum iuformation on the motor and circuit elements involved. For cases not adequately covered by these data, the formulas given below may he used. Static Circuit Elements Only. First assuming that all the voltage drop occurs in static circuit elements such as transmission lines, cables, transformers, and reactors, the voltage at the motor starter mill he equal to

+

Z.W

d(ttMRd2 where Z,

+ ( X , + Xd* X initial voltage at motor starter

(4.9)

impedance of motor being started (ratio of applied voltage to current drawn) R, = z, cos a, X , = Z, sin eM =

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

cos B.,

267

power factor of current drawn by motor being started total resistance of circuit, between motor and point in system where voltage remains constant, i.e., is not affected by start.ing of motor X s = hotal reartance of circuit between motor and point in system where voltage remains constant The impedance, resistances, and reactances in the above formula should all he expressed in ohms or all in per rent (or per-unit) on any convenient kva and voltage base. The mot,or impedance Z.,, expressed in ohms is

Rs

=

=

4

Voltage rating of motor in volts X starting current in amperes a t rated motor voltage

(4.10)

If a reduced voltage type starter is used, the starting current is that drawn from the line with rated motor voltage on the line side of the starter. Similarly, cos Ox is the poiver factor of the current drawn from t,he line. The voltage at the starter must, be multiplied by the motor voltage-line voltage ratio of the starter (see Table 4.13) to obtain the voltage at the motor t,erminals. The resistance and reactance of a transformer hank ran he expressed in ohms by multiplying its per cent resistance and per cent reactance, respectively, by (Secondary voltage rating in kv)2 X 10 Kva rating of bank

(4.11)

Circuit elements separated from the motor by a transformer should have their actual resistance and reactance values in ohms multiplied hy the square of the no-load voltage transformation ratio, that is, by (4.12) before adding to the ohmic resistances and reactances of the motor and other circuit elements on the serondary of the transformer. If two or more transformers are in series between the circuit element and the motor, the actual resistance and reactance in ohms should be multiplied by the square of the product of the various no-load voltage transformation ratios. For transformers equipped with taps 011 either primary or secondary winding, the voltage ratings used in the above formulas should correspond to the t a p setting. Using the per-unit system, it is generally convenient to select as base kva the kva drawn by the motor at rated motor voltage, which is X starting current in amperes X rated motor volts

1000

(4.13)

268

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION

OF DROPS

and select rated motor voltage as the base voltage. Iii this case Z , = 1. The per cent resistanre and reactance of a transformer, with the motor connected t o its secondary, should be multiplied by

f Motor-starting kva at rated motor voltage \ G t i n g of transformer

\

)

Isecondary voltage ratiug\' of transformer rated motor voltage

\

1 (4.14)

A second transformer in series would have its per cent resistance and reactance multiplied b j the above expression and also bj' the square of the no-load volhage transformation ratio (secondary voltage divided by primary voltage) of t,he first transformer. The resistaiice and reactaiice of circuit elements that are expressed iii ohms should be multiplied hy

Motor starting kva at rated motor voltage ." (Rated "r volts ' x 1000

)

(4.15)

except where the circuit element is separated from the motor by a transformer, in vhich case the multiplier is Motor-starting kva at rated mot,or voltage Prjmary voltage r a h g of transformer rated motor volts -x Secondary voltage rating of transformer 1000

(~~__

) x 1000 (4.16)

If t v o or more transformers are in series bet,ween the circuit element and the motor, the transformer no-load voltage ratio which appears in the above espression should be replaced by the product of the no-load vokage transformation ratios of the various traiisformers. Where voltage taps are provided on a t,ransformer, the voltage ratirigs used in the above formulas should correspoiid t o the t a p sett,ing. The resistance and reactance of circuit elements connected in series can be added directly. For circuit elemeots connected in parallel, equivalent wlnes of resistance and react,ance can be det,ermined hy the method given in Chap. 1. If current to other loads is flowing in one or more of the circuit elements between the motor and the const,ant voltage point mhen the motor is started, the above formula for voltage a t the motor mil1 still apply, assuming that these other loads are of the constant-current type, i.e., the current drawn does not change ivhen the voltage drops. Such load currents must, of course, be considered io determining the initial voltage at the motor starter. A method for taking into account loads whose current varies u i t h voltage will be given later. Often it is desirable t o know the effect of motor starting on the voltage

V O L T A G b S T A N D A R D RATINGS, VARIATIONS. C A L C U U T I O N OF DROPS

269

a t various points in the system as well as a t the motor. The voltage a t the motor starter divided hy rated motor voltage and multiplied hy the current drawn at rated motor voltage gives the actual current drawn from the line. This current can be comhined with any load current flowing through the various circuit elements and the voltage a t any point calculated hy the methods given earlier in this chapter. For the case where motor starting current only flows in the circuit elements between the motor and a point in the system, the voltage a t this point will he equal to

d(Rw d(RX

+ + Rs)*+ (x.w (X, + + R i ) 2+

where R I Xi

x1)2 X XS)2

initial voltage a t motor starter

(4.17)

resistance of circuit betweeo motor starter and specified point reactance of circuit between motor starter and specified point R a , X,, Rs, and X s are as previously defined If any load drawing current through the circuit elements in series with the motor is not of the constant-current type, the voltage a t the motor starter can still be calculated hy the formula given provided that the initial voltage a t the motor starter is calculated using the current drawn by the various loads aft,er the motor is started. Since these currents will depend upon the voltage drop occurring when the motor is started, a trial-and-error solution is necessary. Thns the voltage a t the various loads eaii first be estimated from calculations based on ali loads drawing a constant current. The current drawn by each load a t the estimated voltage is used to calculate a new value of initial voltage a t the motor starter from which the voltage a t the motor starter and a t the various loads can be recalculated. If the load voltages do not agree closely with those estimated, nem estimates can be made and the process repeated. In many cases the voltage drop can he caleulated with little error, considering only Lhe reactance of the circuit elements in series with the motor and using the formula =

=

Voltage a t motor starter

where X s

=

zx z.w + x s x

initiai voltage a t motor starter (4.18)

total reactance of circuit betmeen motor and point in system where voltage remains constant Z, = impedance of motor heing started When the reactance-to-resistance ratio of the eircuit elements (X,/Rs) is 2 or greater, this formula gives a voltage drop which is generally within 10 per eent of the correct value. Transformers rated 100 kva or larger usually have a reactince-to-resistance ratio greater than 2. =

270

VOLTAGF-STANDARD

RATINGS, VARIATIONS. CALCULATION OF DROPS

Effect of Generators. h’ext consider the case ivhere generator voltage drop as well as the voltage drop through static circuit elemeiits must be considered. If there is no initial current flowing through the circuit elemeiits mhen the motor is started, the voltage a t the gerierator terminais may be determined from the curves of Figs. 4.39 aiid 4.42 using a value of “motor-startiiig kva at rated geiierator voltage” equal t o

Starting kva drawn by motor if voltage at motor starter mere maintained at the initial value X

Z I,

d(fi,, + n,)z +

rated geiierator voltage

(x,“+ xsjz

)

x(.initialgeiieratorvoltage ..

(4.19)

where Z,,,, R , , and X , are as previously defined Ra = resistance of circuit betweeii motor starter and geiierator terminals X, = reactaiice of circuit betmeeii motor starter aiid geiierator termina Is The pomer factor of the current drawn from the generator will equal (4.20)

Haviiig dctermiiied the voltage a t the generator termiiials, the voltage at the motor starter cari be calculated as it xill equal

zw

d(&, + K s ) ? + (XI, +

XS)*

voltage a t generator _ terminals _ X initial ~ motor voltage initiai generator voltage

(4.21)

If currents t o other loads (of constant-current type) are floniiig through the circuit elements mhen the motor is started, the voltage drop may be determined by trial aiid error. The formula gireri above, Eq. (4.19j, for motor-starting kva a t rated generator voltage may be used for the first estimate and the correspoiiding value of generator voltage determiiied. From this the voltage at the motor starter may I i c calculated. It is equal t o /initial voltage at motor\ starter which mould apz >< (4.22) pear if generator voltage drop had already occurred Having the voltage a t the motor starter, the kva drawn by thc motor caii be calculated. The equivalent motor-starting kra a t rated generator voltage wili equal the actual kva drawn by the motor multiplied by

V O L T A G L S T A N D A R O RATINGS. VARIATIONS, CALCULATION OF DROPS

271

(Rated generator voltage)z Actual voltage at motor starter x actual voltage at, gcrterator

(4.23)

~~

~

~

If there is a transformer between the generator arid the motor, the vollagc a t the mot,or starter should be multiplied by the no-load volt,agc trausformation ratio (primary voltage ratiiig divided by secondary voltage rating) of the transformer before suhstitutiiig it in the above formula. With t,wo or more transformers in series, use as a multiplier the product of their no-load voltage transformation ratios. If the calculated mot,orst,arting kva a t rated generator vokage differs appreriably from the first estimate, a serond estimate based on the calculated value can be made and the calculatioiis repeated until a close rherk is obtained. Motor-starting Power Foctor. Use of the preceding formulas requires a knoivledge of the motor-start,ing power fartor ((WS 8.,,). The starting power factor of squirrel-cage induction and synchronous motors varies over a rather wide range, depending upon the rating and desigii characteristics. Approximat,e values of starting power factor for typiral squirrel-rage induction motors are given in Fig. 4.47. Low-speed (450 rpm aiid below) synchronous motors for reriprovatirrg compressor drive usually have a start,itig p o m r fartor bet,ween 0.20 aiid 0.40. Synchronous motors for rrntrifugal pump drive, on the other hand, have starting power fartors generally between 0.15 and 0.35. Where motor-start,ing power factor must be kuo\vn more acrurately, a value should be ohtailled from the motor manufacturer. With reduced voltage starting, the p o m r factor of the rurreut drawl from the line may be somewhat different from the motor-starting power factor. An autotransformer starter has oiily a small effect on the porver fact.or, but the magnetizing current of the autotransformer makes the power factor of the current drawn from the line slightly less t,han the motor-starting p o w r factor. With a reactor st,arter, the power factor

-"

0.70

50.60 0

-=

0.50

A

0.40

0.M

=w 0.20

B

O.I0 0.001 5

FIG. 4.47

I

K)

I

I5

I

20

I l l

I

I

30 40 50 75 100 HORSEPOWER RATING

1

1

150 M o

I

300

I

500

I

I

700 1000

Approximate 3tor:ing power factor of typical squirrel-cage induction motors.

272

VOLTAGE-STANDARD

RATINGS. VARIATIONS. CALCULATION OF DROPS

of thc. riirri.iit dmwi from t h r liiie \ri11 eqiial the motor-startiiig power f w t o r miiltiplird Iiy thi: volt,age ratio (motor volt,age divided hy liiie voltage) of t h r startcr. .i rc:sist,»r starter, oii the other haiid, results i i i a power fartor for t h r riirreiit drawii from the liiie equal t o

To illiistrate, assume t h a t a motor Iiaviiig a startiiig poirer factor of 0.30 is providrd with a resistor starter dcsigtied t o reduce the voltage applied t o the mot,or t o íi3 pcr cciit of ratcd motor voltagr. Thc p o w r factor of t h r (wrrriit drawii hy tliis motor-start,cr combiiiatioii \vil1 iie ~~

-\/I-

~(0.íi5)2X [1--(0.30)*]

=

0.785

REDUCED-FREQUENCY STARTING

Ociasioiially i i i ordcr to start a largc motor, t,he system frequeiicy is rcduccd to a Ioiv valiic i i i ordcr to iiirrcasc t,hc ratio of tlie motor torquc to thc motix-startiiig ciirreiit. At rcduced frcqiiciicy the applied volt,age is l o w r , liiit i i i thc iisual applicatioii of tlic schcmc, thc applied voltage is rrdiii,rd oiily to t h r samr rxtriit as the frrqiieiivy; that is, t,hc geiicrator exvitatioii is maiiitaiiied at tlie same valiie as heforr. Motor torqiie aiid wrrriit varg irith rediiiiiig frequeiivy i i i t,he samp iray as t,hey do with iiirrrasiiig spwd. sitiw i i i either rase t,he rotor frequeiicy is redured. C:«iiseqiieiitly, at 10 pcr i c i i t frcqiiciicy, the torqiie delivered and the wrreiii d r a w i \vil1 iic approsimatcly thc same as at 90 prr r r n t speed. IItari: tlir torqiie is griierally highcr aiid the ciirreiit loi\-er thaii at standstill. h t ttirse lon- freqiiriiries the effertive liiir resistarire is grcatly iiirrrast:d so that RII this iortliic is iiot rcalizcd. Severtheless, t,he scheme will eITrctively iii(.rcase t,lic toripie a\-nilat>lefor startiiig aiid aweleratiiig tlie motor. H o w v e r , thcre are scvcral disadvaiitages which usually makr it impractical: i . T o ohtaiii miich improvemriit thí! frequciicy must be redured t o a \-cry Iow valiie, iisiially M o \ \ - 50 per w i i t frequeiii.y, ivhich is difficult for some typcs of geiirrator ilrives. 2. i i i i iiidcpciideiit drive for the exeiter must he proridcd as direetcoiiiiectcd (ir I>rltrd exriters uill iiot provide suficieiit excit,atioii at 1ow geiicrator speeds. 3 . Loivcriiig tho system freqiieiiry may adversely affwt other equipmeiit coiiiiected to tlie systrm. Coriscquciitly, t h r svhrme is usually applicahle only for a generator supplyiiig a siiiglr motor ivheir excitatioii is supplicd by ai1 excitcr driven by a sepaiate steam tiirhiiic or aii eqiially iritiepeiidciit excitation source. 111 siicli cases, t h e schemc may be quite advaiitagrous.

V O L T A G ~ S T A N D A R DRATINGS, VARIATIONS, CALCULATION OF DROPS EXAMPLE OF CALCULATION

OF

VOLTAGE DROP DUE T O M O T O R STARTING '

Data (see Fig. 4.48) Generators: Two identical turbine-driven generators, 3600 rpm Total output rating = 10,000 kva Voltage rating = 6900 volts Voltage-regulator setting = 6700 volts Overhead line: 3-ft equivalent delta spacing Length = 5000 f t Conductor size = KO, 4/0 Awg Transformer hank: Output rating = 2000 kva (three-phase) Transformer voltage rating = 6600-2400 volts Motor starter: Autotransformer type Tap = 65 per cent Motor: Synchronous motor Output rating = 1000 hp Full load input = 1000 kva, 0.8 power factor Voltage rating = 2200 volts Full-voltage starting kva = 500 per cent Full-voltage starting torque = 65 per cent Initial conditions: Initial voltages At generator bus

=

6700 volts (regulator setting)

I rrT"

FIG. 4.48 Circuit diagrcm of power supply to motor.

LINE TRANYORMER BANK

MOTOR

273

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

274

At transformer primary = 6700 volts At transformer secondary = 2440 volts Initial loads At generator bus = 5000 kva (50 per cent of generator rating) of constant-current type No initial load on overhead line or on transformer Requirements : Minimum allowable voltage a t generator bus = 90 per cent of initial voltage Motor starting torque must be at least 25 per cent Voltage calculations: Starting kva drawn with rated motor voltage a t autotransformer starter = full-voltage starting kva X multiplier from Table 4.13 = 5 X 1000 X 0.46 = 2300 kva Kva applied to transformer a t rated secondary voltage =

starting kva a t 2200 volts X

=

2300 X

2400

(2200)

=

rated secondary voltage 2200

2735 kva = 137 per cent of bank rating

Transformer secondary voltage (neglecting primary voltage drop) is obtained from Fig. 4.44. For banks rated 15 kv and below and a starting kva of 137 per cent, it is 93 per cent of the initial secondary voltage. Kva applied to transmission line a t initial voltage secondary voltage of transformer = starting kva a t initial voltage X initial secondary voltage =

2300 X

( )' . 4

X 0.93 = 2620 kva

~~,

Loading factor = kva applied a t initial voltage length in feet - = 2620 X 5000 - 0.292 (6700)* (initial volts)* ~

From Fig. 4.46, for M = 1, 4/0 line, 3-ft spacing, voltage drop is 11.5 per cent. Since M = 0.292, drop in line is 0.292 X 11.5 = 3.36 per cent. Voltage a t end of line (neglecting generator voltage drop) is 100 - 3.36 = 96.64 per cent of initial voltage (6700 volts). Kva applied to generator a t rated generator voltage voltage a t end of line = starting kva a t rated generator voltage X initial line voltage

VOLTAGLSTANDARD RATINGS, VARIATIONS, CALCULATION OF DROPS

( itransformer n i t i a l secondary secondary voltage voltage X 0.9664 X 0.93

=

=

)

=

2300 X

(

275

i

mo

6900 X 22002400)

2690 kva 26.9 per cent of generator rating

From Fig. 4.40, performance factor K for a 5000-kva 3600-rpm generator a t 50 per cent initial load is 1.9. From Fig. 4.39, minimum generator voltage, for 26.9 per cent starting kva and K = 1.9, is 92.5 per cent of the initial voltage (6700 volts), or 6200 volts. From Fig. 4.42, restored generator voltage for 26.9 per cent motorstarting kva and 50 per cent initial load is equal to the initial voltage or 6700 volts. The minimum voltage a t the motor starter is equal to the initial voltage a t the motor starter multiplied by secondary voltage of transformer Minimum generator volts Initial generator volts initial secondary voltage voltage a t end of transmission line initial voltage a t end of line = 2440 X 0.925 X 0.93 X 0.966 = 2030 volts

) (

(

The restored voltage a t the motor starter is equal t o the initial voltage at the motor starter multiplied hy secondary voltage of transformer Restored generator volts initial secondary voltage Initial generator volts voltage a t end of transmission line initial voltage a t end of line = 2440 X 1.00 X 0.93 X 0.961 = 2200 volts

) (

(

Since the restored voltage is equal to rated motor voltage, the starting torque on the 65 per cent autotransformer tap = 65 X (0.65)' = 27.5 per cent The minimum voltage a t the generator bus (92.5 per cent of initial value) and the motor starting torque (27.5 per cent) both meet the requirements. Next the formulas for calculating voltage drop will be used to solve this problem. I t will be assumed that Motor-starting power factor Transformer resistance Transformer reactance Transmission-line resistance

= = = = =

30 per cent 0.7 per cent 5 per cent 0.0573 ohm per 1000 f t 0.287 ohm total

276

VOLTAGE-STANDARD RATINGS, VARIATIONS, CALCULATION O F DROPS

Transmission-line reactance

=

=

0.121 ohm per 1000 f t 0.605 ohm total

The per-unit system will be used n.ith base kva equal to the motorstarting kva a t rated motor voltage (2300 kva) and base voltage equal to rated motor voltage (2200 volts). On this basis, the motor constants are

z, = 1 cos On/ R.u

B M = 72.5'

=

0.3

=

Z . w cos BM = 0.3

X M = Z M sin Bar

=

0.954

The resistance and reactance of the transformer vill equal the per cent values multiplied by secondary voltage rating of transformer rated motor voltage

Motor-st,arting kva a t rated motor voltage Kva rating of transformer

Transformer resistance = 0.7 X 0.0137 = 0.0096 Transformer reactance = 5 X 0.0137 = 0.0685 The resist,ance and reactance of the transmission line will equal the ohmic values multiplied by Motor-starting kva a t rated motor voltage rated motor Primary voltage rating of transformer volts x 1000 1000 Secondary voltage rating of transformer 2300 -

=

0.06275

(4.16)

1000 Line resistance

= =

0.287 X 0.06275 0.605 X 0.06275

= =

0.0180 0.0380

The total resistance and reactance between the motor starter and the generator will be Rs = 0.0096 0.0180 = 0.0276 X s = 0.0685 0.0380 = 0.1065

+

+

The equivalent motor-starting kva at rated generator voltage = starting kva drawn by motor if voltage at motor starter were maintained at the initial value

RATINGS, VARIATIONS, CALCULATION OF DROPS

VOLTAGE.-STANDARD

X

d ( R , ,+ &)2 =

2300 X

z

277

(rated generator voltage initial generator voltage 1 X d ( 0 . 3 0.0276)* (0.954 0.1065)'

Y

+ (X,, + Xs)* ~~

(-)2440

+

2200

+

x (o)*

+

= 2700 kva

(4.19)

This is substantially the same as previously determined; so the generator voltage drop will h r essentially the same, that is, the minimum voltage will be 6200 volts a i d the restored voltage, 6700 volts. The voltage at the motor starter will equal the voltage at the geuerator multiplied by initial motor voltage zw x ( .initial . . generator voltage d(~,, + n,)z + ( x ,+~xS)2

zo0.328 + 0.276)' + (0.954 + 0.1065)* X '6700 1

-

~

d(0.3

=

Thus t,he minimum voltage at the motor starter will he 6200 X 0:328 and t,he restored voltage

i d

=

2030 volts

=

2200 volt,s

he

F700 X 0.328

(4.21)

Chafiter 5

by R. H. Kaufmann and Maynord N. Halberg

Sys tern Overvoltages-Causes and Protective Measures Electric insulation in energized systems is continuously under stress. To make the most economical use of insulation, operating overvoltages should he curbed in so far as is reasonably possible. The application of additional insulation to accept higher overvoltage levels entails several rather obvious disadvantages: (1) increased cost, ( 2 ) increased size and weight, (3) increased resistance to the flow of heat from the currentcarrying conductors. In the case of a-c systems, the electric potential is varying substantially as a sine wave. The crest potential will be 41 per cent greater than the rms value. Under ideal conditions the line-to-ground voltage stress mill he less than the line-to-line operating voltage. In the case of direct current or single-phase alternating current, this ideal line-to-ground voltage would he E L L / 2 ,or 50 per cent of the line-to-line value. In the case of three-phase a-c systems, this ideal line-to-ground voltage would be E L L / f i , or 58 per cent of the line-to-line value. Throughout this section, overvoltages will he expressed as multiples of the ideal balanced voltage stress in three-phase systems. Electric systems are subject to disturbances of many types which unavoidably produce overvoltages. However, the application engineer has at his command many system design principles which will greatly curb the magnitude of overvoltages. It is important to note that a-c systems are subject to many types of overvoltages not to be found in d-c systems; hence a-c systems deserve more careful consideration of the overvoltage problem. Electric insulation exhibits the effect of fatigue. Insulation will fail upon repeated or prolonged application of a given voltage stress which is 278

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

279

far below the single-impulse \vit,hstaiid abilit,y. One may conclude that a reduct,ion in either the magnitude or duration of overvoltage stress will in general result in longer useful life. OVERVOLTAGE SOURCES

There are many varied sources of overvoltages of sufficient magnitude to be damaging t o the insulation of a-c industrial power distributioii systems. 111t,liis chapter the mechanism by which the more prominent overvolt,ages are created v i l l be described and preventative measures suggested. ‘Treatment of t,he following will be included: I . Static 2 . Physical contact nith a higher voltage system 3. I1esouani.e effects ill series inductive-capacitive circuits 4. Repetitive int,ermittent short circuits 5 . SIT-itrhing surges (i. Forced-current zero-current interruptiou 7. Autotransformer connections 8. Lightiiitig Of these, most are the result of effevtsdirectly within the electric system itself. I n contrast, lightning (a vicious source of overvokage) is communicated to the electrical system from nature’s powerhouse in the heavens above. STATIC

Wind-blown sand or dust can become highly charged and impart relatively high voltage to exposed overhead electric conductors. Moving belts rutiiiing on iioiimet,allic pulleys can also develop high voltages by st,at,icmeans which may in turn be communicated t o electric system conductors if electric enclosing frames arc improperly grounded. The rate a t wtrirh electric i,harge is communicated t o electric system conductors by stat,ir means is extremely low. Even a rather high-resistance ground i~iiincrtionon the electric system n d l discharge these stat,ic currents t o ground as fast as they are rereived with negligible overvoltages. I n addition to grounding the elect,ric service system, it is important that electric machiue frames arid all metallic enclosures which contain electric circuit conductors be effectively grounded (see Chap. 7). PHYSICAL CONTACT WITH A HIGHER VOLTAGE SYSTEM

If the conductors of a high-voltage electrical circuit come in contact with those of a lower voltage circuit, then the same potential will exist on

280

SYSTEM OVERVOLTAGES-CAUSES A N D PROTECTIVE MEASIJRES

both circuits at the point of contact. If Lhe low-voltage circuit does not have its neutral grounded, its potential will be increased t o t h a t of the high-voltage system or flashover mil1 occur. If Lhe low-voltage system is anchored close t o ground potential as hy Lhe use of a solidly grounded neutral, high values of current may flow from the high-voltage system, b u t a much lower voltage will appear than with an isolated neutral system. Accidental cootacts hetmeen primary and secondary voltages on industrial systems are guarded against by the use of metal enelosures and metal barriers which separate conductor systems of different operating potentials. In some cases overhead circuits have both primary and secondary on the same pole, but substantial clearances reduce Lhe danger of accidental contact t o a minimum. Occasional cross-ups have occurred between primary and secondary on overhead circuits, and a few cases are known where failure has occurred between primary and secondary inside a transformcr. UNINTENTIONAL CONNECTION

7 PHYSICbL CONNECTIONS

P,

I I

'..

/

2 I 3LI 1

N O R M b L POSITION O F 4 8 0 V VOLTAGE TRIbNGLE

C xq I

I

Eb= 2 4 0 0 V

\:ol,

\,\

L,--' I

e0

b RESULTING VECTOR VOLTbGE DIbGRbM

FIG. 5.1 Overvollage on 480-volt ungrounded ryrtem rerulting from contcxt with a higher roltoge ryrtem.

Figure 5.1 illustrates this type of fault connection. It can be responsible for dangerous overvoltages on ungrounded low-voltage systems. The most effective protection against that type of overvoltage is grounding of Lhe lowvoltage system mith the grounding impedance made low enough t o accept Lhe maxirnum line-to-ground fault current of the high-voltage system without biasing the neutral of the low-voltage system by a dangerous amount.

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

281

RESONANT EFFECTS I N SERIES INDUCTIVE-CAPACITIVE CIRCUITS (LIMITED TO A-C SYSTEMS)

Ungrounded-neutral a-o systems are most commonly subject t o overvoltages originatiiig from this cause. It is import,aot t o recognize that ungrouiided-iieutral systems are actually capavitively roupled t o ground rather than truly divorred from ground. They are ungrounded in the sense that no int,er(.oiinection with ground has purposely been made, but every element of the electric system incorporates some capacit,aiice t o ground which constitutes an inherent caparitive impedance interconnection tietween the elertrir system conductors and ground. Every ungrounded elertric system contains the essential elements presented i n the upper diagram of Fig. 5.2. The electrical behavior of any one phase conductor relative t o ground rail be determined by a much simpler equivalent rirruit, as indicated in the lower sketrh of Fig. 5 . 2 . A

.

'S

A PHASE

I

GENERATOR OR TRANSFORMER

3- PHASE ESSENTIAL ELEMENTS

xs

A PHASE

"A"*"

Eg

-E'.c %

EQUIVALENT CIRCUIT REFERRED TO A PHASE CONDUCTOR

FIG. 5.2

Elemental composition of an ungrounded system

In terms of this simpler equivalent circuit it will be possible t o understand readily the effect of connecting different types of impedance hetween line and ground as shown in Fig. 5.3. I t becomes evident that the connection of any value of either resistance or raparitanrc tietween one line and ground produres no dairgerous overvoltages. The potential on the phase to which the impedance is connected progressively diminishes from normal value t o zero. The potential t o ground on the remaining two phase conductors will be increased t o full line-to-line value at the time the first

282

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

6

5

I

I

.

" 4

Y

J

z

LL

4 3 "Y 7

Ec ' 2

. i

FIG. 5.3 Overvoltages on ungrounded systems result

I

0

0

2 3 4 R4TlO OF ZF TO X C O , ~

S

6

from a high-inductive-reactonce connection between line and ground.

phase conductor has been reduced to zero potential. This represents an overvoltage of only 73 per cent, which is not dangerously high and will normally produce no ill effect unless continued for a long time. The connection of an inductive reactance between line and ground, on the other band, can be responsible for the production of serious overvoltages to ground. It is the ratio of the inductive reactance of the lineto-ground circuit to the total capacitive reactance of the system to ground which controls the degree of overvoltage. The highest overvoltage will occur when these two reactances are equal, and a t this point they may be as much as ten times normal. It is significant to note, however, that over a two-to-one range of reactance, overvoltages of three times normal or more would be produced. The unintentional connection of an inductive reactance between a phase conductor and ground can occur in a number of ways, some of which are illustrated in Fig. 5.4. The operating magnetic coil of a motorstarter contactor may be inadvertently connected between phase and ground by a ground short circuit in the control wire to the push-hutton station or the slip of a maintenance man's screwdriver. Any time that the inductive-reactance value, which becomes connected from phase to ground, falls in the danger region indicated on Fig. 5.3, dangerous overvoltages to ground will be produced which are communirated over the entire metallic conductor system of that operating voltage. Overvoltages originating from this canse can be completely suppressed by a relatively light-resistance ground on the electric system neutral. A grounding resistor of about the same ohmic value as the total charging

C

B

"--

C4SE I

AN INDUCTIVE WINDING CONNECTED BETWEEN 01 GROUND

FIG. 5.4

-

f

I

CbSE 3

Examples of unintentional high-reactance connections between line and ground.

ONE BROKEN OVERHEbO LINEGROUNDED ON T H E L O A D SIDE OF T H E B R E A K CONNECTS T H E REACTANCE OF TRANSFORMERS 12 AND 13 I N PARbLLEL B E T W E E N L I N E b N D GROUND [NOTE II

BROKEN L I N E GROUNDED

b GROUNO FAULT bT A F U S E PROTECTED T R I N S F O R H E R C I N BLOW ONE FUSE LEbVING THE REACTbNCE OF TRANSFORMERS 12 4 N D T 3 I N P b R A L L E L B E T W E E N L I N E AND GROUND 1 NOTE I I

CbSE 2

_

Y U N G R O U N D E D T R I N S F O R M E R CONNECTIONS WOULD PRODUCE T H E SAME EFFECT

NOTE I

:CIDENTbLLY PHASE h N D

II

284

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

capacitive reactance t o ground is sufficient t o eliminate overvoltages almost completely. It will be evident t,hat there is good reason t o adopt electric system neutral grounding with a much lower value of grounding resistance for ot,hcr reasons (see Chap. 6). Figure 5.3 has been computed on the basis t,hat the inductive reactance is linear. If this reartance incorporates a n iron core which during the mode of operation being considered should encounter magnetic saturation, the performance will be somewhat different. Under such conditions the effective reactance of t,he inductive circuit can become much lower than the unsaturated reactance, and the voltage will tend t o oscillate automatically betveen vokage limits which cause the effective inductive reactance to match the capacitive-reactance value. This character of operation has been named ferroresonance. The maximum voltage so developed may not be so high as would be produced by a linear reactor but, may still be in excess of two or three times normal. Substantial overvoltages may result by ferroresonance when the unsaturated reactance is many t,imes the capacitive reactance to ground. The application of grounded-Y potential transformers on ungrounded systems with a Y or broken-delta secondary connection can be responsible for damaging overvoltages as a result of resonant or ferroresonant action since the magnetizing reartance of the pot,ential transformers becomes connected from phase conGROUNDED WYE- BROKEN DELTA POTENTIAL TRANSFORMEIS FOR GROUND ductors t o ground. A comINDICATOR OR ZERO SEQUENCE VOLTAGE plete descriptionof thisphenomenon need not be taken UNGROUNDED NEUTRAL SYSTEM u p here as it has been adequately treated in an AIEE technical paper (see reference3). Thesesystemvoltage oscillations will not occur if the electric system U neutral is grounded. Freedom from this particular type of voltage oscillation TO INSTRUMENTS can be obtained even with R ungrounded-neutral operation byusingpotential transTO INSURE FREEDOM FROM UNWANTED LINE-TOformers with a line-to-line GROUND VOLTAGE OSCILLATIONS :

r Tz. T3, W I T H THE L I N E - T O - L I N E RATED VOLTAGE 2 APPLY A SECONDARY LOADING RESISTOR WITH A RESISTANCE NOT GREATER THAN 4 0 PERCENT OF THE TRANSFORMER MAGNETIZING REACTANCE. NOTE- THE LOADING RESISTANCE CAN BE APPLIED TO EACH SECONDARY BUT WILL THEN CDNSUME POWER AND LIBERATE HEAT CONTlNUOUSLY I SELECT P T s TI.

FIG. 5.5 Grounded-Y brokendelta potential transformers for ground indicator or zero-sequence voltage detector.

SYSTEM OVERVOLTAGES-CAUSES

AND PROTECTIVE MEASURES

285

voltage ratiiig and the applirat,iori of shuiiting resistors on the secondary windings as is outliried iii Fig. 5.5. Series-capacitor melders are occasionally applied, particularly in the case of large-sim machiiies because of their ability t o reduce the kva demaiid aiid improre the operating poiver fartor t o substantially unity. However, the series-raparitor welder preseiits a definite voltage hazard to aii uiigrouiided-iieutral a-c supply system. Duriiig welder operation the voltage arross hoth t,he series raparitor and the weldiiig transformer primary v i l 1 he severa1 t,imes the rated line-to-liiie voltage. The physical electrir roiiiieitioiis aiid the associated vector voltage relationships are iiidicated iii Fig. 5.íi. 48OV 3-PH 6 0 C Y

o

s i i o m m ~ uTO~GROUND -

?--.

PHVSlCbL CONNECTDNS

NORMAL POSITION OF IP UOLTME TRIANGLE;

-,,, /' b

RESULTING VECTOR VOLTbGE OIAGRPH

FIG. 5.6

Overvoltager on ungrounded syrtemr os a rerult ot o ground contact on a ieriei capocitor welding mochine.

Should a fault t o grouiid occur at the juiiction hetveeii the series capacitor aiid the weldiiig transformer (poiiit, P ) , the lorat,ioii of ground poteiitial will teiid t o become t,hat of this juiictioii poiiit iiistead of the center of the a-r system voltage triangle. The t,otal system eapacitiye impedaiice t,o grouiid would geiierally be expeeted to be high, relatire to that of the welder series rapacitor, aiid thus offers practically no opposition t o this shift in the loration of ground potential. Iii the case illustrated iii Fig. 5.G, it will be evideiit that the poteiitial of the A-phase roiiductor may be elerated to ahout 2000 volts to grouiid, which is about seveii times iiormal. As iii the other cases, this overvoltage is commuiiicated to a11 equipmeiit metallically iiiterroniierted a t this commoii operatiiig voltage. AI1 these resonant inductive-caparitive overvoltage hazards can be elimiiiated by electric system neutra1 groundiiig.

186

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

INTERMITTENT GROUND FAULTS

Substantial overvoltages can he developed in ungrounded a-c industrial systems by sputtering or intermittent ground-faulting connertions. The intermittent character of the short-circuit path may be the result of vibration which causes an electrical conductor to make contact intermittently with ground, the result of scattering particles of molten conductor metal which intermittently establishes a conducting path to ground, or as a result of successive breakdown and seal off of the separating space between conductor and ground. In the last case involving a fixed separation between conductor and ground, a progressively increasing breakdown voltage across this gap is an essential element in the build-up of severe overvoltages. Intermittent ground-fault conditions on lom-voltage ungroundedneutral systems have been observed to create overvoltages of five or six times normal quite commonly. An unusual case involved a 480-volt ungrounded system. Line-to-ground potentials in excess of 1200 volts were measured on a test voltmeter. The source of trouble mas finally traced to an intermittent ground fault in a motor-starting autotransformer. About two hours elapsed while the source was being located, during 13 hich time between 40 and 50 motors broke down. Electric systems which are grounded through reactanre of too high an ohmic value ( X a more than ten times XI) are also subject to overvoltage by this same mechanism acting in a little different form. An understanding of the manner in which a discontinuous electric connection can he responsible for the generation of overvoltages can he most easily acquired by examining the case of a sputtering or intermittent line-to-ground fault on an ungrounded-neutral system. I n Fig. 5.7 at A is shown the vector voltage pattern of a three-phase a-c system as it would operate under normal balanced conditions. The voltage vectors E., Eb, and E, rotate about the neutral at synchronous speed. The electric neutral is a point of central symmetry and remains constant at ground potential if the individual phase voltages are pure fundamental-frequency sine waves. Should the A-phase conductor become grounded, the system voltage triangle mould become displaced as illustrated in B . At the phase position illustrated in B , the A-phase voltage is at its maximum value at which instant the charging current to ground (90' ahead of the voltage) is passing through zero. In case the short circuit contains a small gap or an arc, the arc current would become extinguished at this point. Note that the trapped charge on the line-to-ground capacitance will tend to maintain the voltage triangle in the same displaced position. I n other words, the potential of the neutral (relative to ground) would tend to

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

287

remain at a d-c potential equal to the crest value of the a-c voltage wave. All this merely says that there will be little tendency for any voltage to reappear across the gap in the short circuit immediately following the current zero which occurs at B . During the next half cycle, however, the a-c generat,ed voltages will reverse their polarities (vectors rotate 180°),which would cause the threephase vector voltage pattern to assume the position shown in the upper part of C . Kote that during this one-half cycle time interval, the potential of the A phase has progressively inrreased from zero value to about twice the normal line-to-neutral crest voltage relative to ground potential. This value of line-to-ground potential of the A phase may he sufficient to break down the gap in the ground-fault circuit arid reestablish the corinection between the A phase and ground. If so, the A-phase potential will tend to be suddenly yanked to ground potential. Iuevitably there will be some system reactance in the A-phase conductor to the ground shortcircuit point which would result in an oscillation of the A-phase-conductor potential between plus 2 and minus 2 at a frequency probably 20 to 100 times normal. If the short circuit consisted of a solid metallic connection, this oscillation would decay to zero, leaving the A-phase conductor at ground potential. Xote that associated with this high-frequency transiEi"'

E;

Y--

+ $

CYCLE

--tC

$

CYCLE

4

€0

NORMAL

A

FIG. 5.7

B

C

D

Overvoltages on ungrounded systems due to repetitive momentary contact between one line and ground.

288

SYSTEM 0VERVOLTAGES.-CAUSES

AND PROTECTIVE MEASURES

tory oscillation will he a corresponding trarisitory charging current t o ground. This transitory charging current t o ground, or restrike current, will again reach zero value when the system voltage swing is at the maximum excursion i n the negative direction, as showri in the lower part of C . Thus, an opportunity is afforded for the gap in the ground short circuit t o rerlear. If reclearing does occur, a charge is again trapped on the system rapacitance t o ground which would tend t o maintain a constant d-c potential t o ground on the system neutral. In the course of the next following half cycle, the voltage vector system will again rotate 180°, causing the potential of the A-phase conductor to ground t o he elevated from minus 2 to minus 4 as indicated by the transition from the lower part of C to the lower part of D. This increased voltage across the short-circuit gap may again result in restrike, in which case the voltage triangle would tend t o be thrown in the positive direction in the form of a high-frequency oscillation between poteutial limits of minus 4 and plus 4, which in the presence of a solid metallic connection would gradually decay t o zero. I n this explanation of the mechanism, it will be noted that all conditions have been most favorable to the creation of the highest possible restrike voltages in the shortest possible time. The restrike has been assumed t o occur at the time the maximum recovery voltage was reached but not before. Likewise it has been assumed that a reclear occurs at the first current zero after restrike. Under these conditions a line-to-ground potential of five times normal has been developed in less than two cycles. I n practical cases, t,he restrike may occur before the maximum recovery voltage has been reached, and several cycles of the transitory oscillation may take place before the short cirruit reclears. While in theory it might he possible progressively t o increase the line-to-ground voltage by successive restrike without limit if the dielectric strength progressively increases, voltage measurements on actual systems indicate that voltage levels of five t o six times normal are rarely exceeded. There is reason t o believe that damaging overvoltages of repetitive restrike origin are far more common on ungrounded-neutral systems than mould a t first he suspected. The case which was mentioned in an earlier paragraph is unusual in t,hat the obnoxious restriking conditions persisted for a long interval of time while t.he source was being located. A farmore common occurrenre is one in which several pieces of electric equipment on the system suffer electrical breakdown apparently simultaneously and one or more of the fault conditions were known or believed to involve ground. These multiple failures are commonly associated with ungrounded-neutral system operation. It is also known that a solid metallic ground connectioo on one phase may exist for subshntial intervals of time without producing multiple breakdowns in equipment,

SYSTEM GV'ERVOLTAGES-CAUSES

AND PROTECTIVE MEASURES

289

although it does produce 73 per cent overvoltage on two of the phase conductors. It therefore seems reasonable to assume that the multiple failures result from the appearance of overvoltages considerably in excess of 173 per cent normal. Distribution-system ox,ervoltages of repetitive-restrike or intermittentground origin can be entirely eliminated by effective system neutral grounding (see Chap. 6). Resistance grounding with a resistance ground fault of any value upward of the line-to-ground charging current mill be effective. For various other reasons it mill he evident that higher values of available ground-fault current will he desirable. If reactance grounding is contemplated (it rarely finds application in industrial systems), it is important to keep the reactance of the grounding circuit sufficiently low so that the ratio of X o is no more than ten times X , . If this grounding reactanre value is exceeded, opportunity is given for another type of repetitive restrike action which can result in overvoltages t o ground. The ground-fault neutralizer (Petersen coil) represents one special case of high-reactance grounding which is free of overvoltages by repetitive restrike action. This is due t o the fact that the reactance value is carefully selected so that the oscillating circuit formed hetmeen it and the system-to-ground capacitance will oscillate a t normal line frequency. Following a ground-fault cnrrent shutoff point as at B in Fig. 5.7, the potential of the electric system neutral with respect t o ground would oscillate between plus and minus 1 at fundamental frequency as controlled by the tuned grounding reactor and system capacitance t o ground. Thus as the potential of the n-phase conductor with respect t o the neutral due to the generat,ed voltage in the supply system alternates from minus 1 t o plus 1, the free oscillation of the zero-sequence circuit remains in step with it, with the net result that the potential of the A-phase conductor tends t o remain at ground potential. Voltage of normal frequency gradually reappears as the free oscillation in the zero-sequence circuit decays. I n general, some 15 or 20 cycles will elapse before the potential of the previously shorted phase increases t o three-quarters of normal value. Thus, the freedom from restrike is due t o the long-delayed reappearance of voltage across the line-to-ground circuit. SWITCHING SURGES

Circuit switching operations constitute abrupt changes in circuit parameters and can be responsible for the creation of overvoltages although generally of short duration and not in excess of two to three times normal. It will be important t o recognize that normal a-c switching interrupters offer very little opposition t o the flow of circuit current during the course of current flow but do act t o build up dielectric strength rapidly during a

290

SYSTEM OVERVOLTAGES- CAUSES AND PROTECTIVE MEASURES

normal current zero and prevent reestablishing current flow during the following half cycle. As a result of this action it is unnecessary that the stored magnetic energy in the inductance of the circuit be disposed of during interruption. Interruption takes place at a normal current zero, at which time the stored magnetic energy is zero. A quaiitative understanding of the mechanism whereby such overvoltages are generated will be useful. Of first consideration is the amount of voltage change which would tend to appear across the switching contacts if they were switched open. For example, in Fig. 5.8, a line-to-line short-circuit condition between phases A and B is illustrated. With the circuit breaker still closed, the potential of a' and b' must be common and will lie midway between potentials e, and ea, as indicated in the vector diagram. With the vector relationships shown in the figure, the current in the faulted circuit will be going through zero, which affords an opportunity for the circuit breaker to make an interruption if the contacts have parted. If current flow is interrupted at this current zero, the potential of a' tends to return to e. while that of point b' tends to return to eb. OVERVOLTAGE IN CLEARING A LINE- TO- LINE CIRCUIT FAULT

SHORT CIRCUIT

;-,

- ___-

-\,

, , , "

'

"

VOLTAGE RELATIONSHIP WITH SHORTCIRCUITON AT THE T I M E OF A CURRENT ZERO IN THE SHORT CIRCUIT CIRCUIT e.' = eb' (VOLTAGE mob AT MAX VALUE1

-

IF CURRENT INTERRUPTION OCCURS AT THIS CURRENT ZER? THE POTENTIAL OF POINTS 0 AND b WILL TEND m SNAP BACK TO ea AND Ob RESPECTIVELY BUT DUE TO PRESENCE OF L AND C I T WlLL TAYE THE FORM OF h TRINSITORI OSCILLATION W l C H WlLL OVERSHOOT END

POINT eb'' 113 PERCENT OR

MAX e.'DR

73 PERCENT OVERMLTAGE

FIG. 5.8

Overvoltages due to interruption of

(I

line-to-line short circuit at current zero.

SYSTEM O V E R V O L T A G E S - C A U S E S A N D PROTECTIVE MEASURES

291

There will inevitably be inductive capacitive constants which cause this return to take the form of an oscillation of relatively high frequency; this causes the potential of points a' and b' to overshoot their final value by about a n equal amount. In this illustrative example, the potential of point b' would transitorily swing t o a value of 1.73 times normal crest voltage in the positive direction while that of the point a' would make a corresponding swing to 1.73 times normal crest value in the negative direction. Circuit breakers which introduce substantial resistance drop during current flow tend to reduce the magnitude of switching transient voltages. As a result of the higher power factor of the short circuit, the point at which a current zero is reached will approach more rlosely to the point at which a voltage zero would also he reached, which thus lessens the magnitude of voltage that tends to appear across the contacts immediately following current zero. Another form of switching transient which develops overvoltage primarily on the utilization machine on contact closing is illustrated in Fig. 5.9. Here illustrated is a n open-cycle autoPOSSIBLE SWITCHING OVERVOLTAGES ON CLOSING L I N E BREAKER WITH OPEN CYCLE AUTOTRANSFORMER START

ASSUME0 VOLTAGE RELATlONSHlP JUST PRIOR TO CLOSING L I N E BREAKER IAUTOTRANSFORMER STARTI CBPOLENOI ISTHE FIRST TO CLOSE

4e, c,

MOTOR TERMINAL B W I L L TEND TO ABRUPTLY JUMP TO e. BUT OUE TO

MAX TRANSLTORI VOLTAGES-MOTOR TERMINALS TO GROUND TERMINAL B - 2 5 0 PERCENT 1150 PERCENT OVERYOLTAGEI TERMINALS A 8 C - 325 PERCENT'1225 PERCENT OVERVOLTAGE1

FIG. 5.9 Possible switching overvoltage when motor running breaker closes lopen-cycle autotransformer start).

292

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

transformer starting arrangement. It has been assumed that 65 per cent voltage has been applied on the starting connection and the machine rotor brought up to near synchronous speed. The motor was then disconnected from the starting tap preparatory to reconnection across full-line voltage. During this interval it is possible that the internal generated voltage within the motor has dropped to 50 per cent of rating and has slipped back in angle so as to he 180' out of phase with respect to the supply system. At this point, the potential difference appearing across each of the three line-switching contacts is one and a half times normal line-to-neutral voltage as indicated by the vector relationships. Suppose that the line-switching unit is now closed and that pole 1 is the first to make contact. The potential of motor terminal R would tend to abruptly assume the potential e,, but the inevitable transitory overshoot would carry it on up to 250 per cent normal with respect to ground. The potential of the motor terminals A and C would tend to be carried along and suffer a transitory excursion to 325 per cent voltage with respect to ground unless contacts 2 and 3 close at almost the same instant as contact 1. Closed-cycle starting arrangements such as reactor starting or Korndoerfer autotransformer starting minimize the overvoltages which may be developed in this manner. One of the most severe sources of switching overvoltages is associated with the separation of two system sections which have become unsynchronized and are switched apart when the generated voltages in the two sections are nearly 180' out of phase. The elements of this case are illustrated in Fig. 5.10, which shows a synchronous motor t h a t has pulled out of step and the internal generated voltage of which is 180' out of phase with respect to the system. The main supply system on the left is considered to be operating with grounded neutral and contains a much smaller reactance than the motor circuit shown on the right. All three poles of the switching interrupter have been maintained in a closed position up to the time indicated by the vector diagram. It has been assumed that, in the course of pull-out operation, the demagnetizing reactive current which has been flaming in the motor stator windings has caused the internal generated voltage ahead of transient reactance in the motor to he depressed to 50 per cent of normal value. With the vector system in the position shown, the current in the A phase is going through zero, which affords an opportunity for interruption if the contacts have parted. If the current in the A-phase conductor is interrupted at this point, the potential of the motor A-phase terminal (point a2) will tend to jump to the right to its new steady-state position E,. The inevitable transitory overshoot will cause its potential to swing about an equal distance the other side of point E,, as shown hy the dotted line. At the

SYSTEM OVERVOLTAGES-CAUSES A N D PROTECTIVE MEASURES

293

maximum of this transitory excursion, the potential of point a2 mould reach about 3% times normal crest t o ground in the positive direction. I n coutrast t o the examples just cited, the more usual switching operation which is involved in separating a normally operating rotating machine or composite system of rotating machines involves very little switching surge voltage. The systems on both sides of the switching interrupters contain internal sourres of generated voltage which are of almost the same magrrihde and very close t o the same phase position. Very little change in potential tends to occur on either side of the switching device at the time interruption takes place. Arc-furnare circuits can be sources of rather severe overvoltage if switched off while an arc is in progress within the furnace. As the priS

e

.

d

.

5

\i, c2

e,.... i

.. . .... .. A L L C B POLES S T I L L CLOSE0 ASSUME C B TRIPPED AND POLE I (PHASE A 1 IS THE FIRST TO INTERRUPT AT T H I S CURRENT ZERO 0 1W I L L T E N 0 TO JUMP TO e0,AND

Q p T O T H E NEW EA WITH .TRANSITORY EXCURSIONS SHOWN BY DOTTED LINES

FIG, 5.10

Possible overvoltager when interrupting o synchronous motor during out-of-

step conditions.

294

SYSTEM OVERVOLTAGES-CAUSES

AND PROTECTIVE MEASURES

mary circuit-breaker contacts part, current at the breaker contacts can be forced to zero while current still continues to flow in the furnace arc. Thus the circuit breaker accomplishes a n interruption of line current with current flow still Continuing in the secoudary circuit. As the furnace internal current diminishes, the potential across the furnace arc increases in accordance with the normal inverse volt-ampere characteristic of an arc. The arc voltage progressively increases as the current dimiuishes and can result in a substantial voltage drop as the arc snaps out. While this voltage may not be high as referred to the arc-furnace anode, it still may be many times normal operating voltage and will be reflected to the primary side of the transformer by the turn ratio. The voltage developed at the transformer high-tension terminals may be dangerously high and sufficient to produce flashover. Special consideration is given to arcfurnace transformers, and preventative measures take the form of shuntcapacitor applications at the transformer terminals on older uuits or internal Thyrite* shunting resistors across sections of the winding 011 iiew units.

FORCED-CURRENT-ZERO INTERRUPTION

The discussion of switching overvoltages so far has considered interruption only a t a normal current zero. The term forced current zero or interruption of of current zero is used to describe an interrupting mechanism (be it a fuse, switch, section of small wire conductor, etc.) that has the property of developing a large countervoltage in opposition to rurrent flow which can force current to zero value at a time quite different from the normal inherent current zero of the rircuit. Should any element in an electric circuit have the ability to develop a high potential drop during current flow, the potent,ial so developed would appear on connected circuit conductors. The overvoltages so developed would persist until the stored energy in the inductive elements of the circuits has been dissipated (a current zero has been forced). A high rurrent short circuit created through a length of small wire conductors can be responsible for developing dangerous overvoltages in this manner. As current builds up in such a circuit, stored magnetic energy is heing accumulated in all inductive elements of the circuit. When the fusing point of the conductor is reached, the conductor copper tends t o separate into a loiig string of tiny globules of molten copper with a small arc between adjacent globules. The total voltage drop across the entire section of conductor may be several times the normal operating voltage of the circuit. During this interval of overvoltage, the magnitude of current is being diminished;

* Registwed

tradr-mark of Grncral Elrrtrir Cornpang-.

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

295

however the overvoltage will persist until the magnitude of current has been returned t o zero value. Because of the overvoltage problems, the vacuum contact switch finds little applicatioii. The vacuum switch tends t o shut off current completely the instaiit that the contacts part. Unless suitable overvoltage suppressors are associated with such an interrupter, high voltages will be developed if applied in inductive circuits. The overvoltages so produced may he sufficient t o sparkover the outside of the vacuum switch unless some other portion of the circuit breaks dowu a t a lower voltage. Current-limiting fuses constitute an example of a forced current interrupter. They possess the property of being able t o reduce the rurrent t o zero value ahead of a normal current zero. Overvoltages are developed during the operation of such an interrupter. As supplied by reputable manufacturers, the design of the internal elements contains special features mhirh rontrol the magnitude of such overvoltages, and full-srale tests are applied to prove the resulting performance t o ensure that overvoltages so developed d l be within the safe withstand value of the electric insulation of the voltage class t o which it is t o be applied. Because of the overvoltage problem, current-limiting fuse interrupters of a particular voltage rating should not be applied t.o electric systems of lower operating voltage. I n other words, a 7500-volt rated currentlimiting fuse should not he applied on a 2400-volt operating system because overvoltages developed iu its operation will be dangerous t o a 2400-volt insulation level. AUTOTRANSFORMER CONNECTIONS

Autotransformers for interconnecting two electric systems of different insulation level should be avoided in industrial systems unless both are solidly neutral grounded. The common metallic interconnection between t,he two systems which is formed by the autotransformer windings tends t o subject the lower voltage system t,o nearly the same transitory voltages as would be expected on the higher voltage system. There are some exceptions, and a specific example mill serve t o illustrate the nature. Should a system be planned which is to operate initially a t 2400 volts and later be converted t o 4160 volts with all equipment therein contaiuing insulation levels commensurate with 4100-volt operating potential, i t would be sitisfactory t o employ a suitable autotransformer for interconnecting this 2400-volt system with another 4160-volt syst,em. An unusual var'ation of autotransformer action which has been responsible for system overvoltages in a number of instances is represenled by a transformer with extended windings operating on an ungrounded-neutral system such as illustrated in Fig. 5.11. Applications of this sort are most

296

SYSTEM OVERVOLTAGES-CAUSES A N 0 PROTECTIVE MEASURES

often found in test areas or developmental areas which contain multipurpose transformers with a multiplicity of taps to permit a wide variety of output voltages to be obtained. If operated with system line voltage impressed across a fraction of the total winding, the vector voltage at the end of the winding extension will be as illustrated in Fig. 5.11 because the volts per turn developed in the winding extension will be exactly the same as the volts per turn in the excited winding. Should the end of the winding extension be inadvertently connected t o ground or develop a short circuit to ground, the point of ground potential would tend to move away from the center of the voltage triangle to the potential of the extreme end of the winding extension resisted only by the high system-to-ground capacitance coupling. It will be evident that, as a result of this action, the presence of any extended winding would cause the potential of one phase conductor to be elevated to more than 173 per cent of normal operating potential. The degree of overvoltage may be much more severe if greater amounts of winding extension are present. It is important to realize that these overvoltages would be carried to all apparatus connected to the same metallic system. Thus, a ground short circuit on a winding extension of a transformer in a small test area at one corner of a building might impose overvoltages on all equipment fed from the same load-center substation which might include half the productive machinery in that building. As has been true so many times before, grounding of the electric supply system neutral will cure this type of potential overvoltage also. A system grounding equipment which makes available a ground-fault current which is equal t o or greater than the short-circuit current resulting from short circuit of the extended winding portion of the offending transformer will keep the system line-to-ground potentials within safe bounds. It is quite generally true that transformers of this a

480V W Q H 60 CY

I

"

I

PHYSICAL CONNECTIONS

\ \ \

RESULTING (IOLTAGE

vEcmR

I

i

I

'"' b DIAGRAM

FIG. 1 1 1

Overvoltage on ungrounded systems due to a ground connection on the winding former.

of an autotrans-

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

1P7

character to be found in test areas are of relatively small physical size and do not impose restrictive requirements on the necessary system grounding equipment. As a matter of fact, on all low-voltage-system equipment (GOO volts and less) it is the standard practice to ground the neutral solidly. The application of three-phase transformers or three-phase banks of single-phase transformers, mhich do not incorporate a closed-delta winding in their make-up, should in general be avoided or quite rarefully examined to ensure that the resulting operation will be free of damaging overvoltages. This would be equally true of Y-connected autotransformers (see reference 4). Berause of the nonlinear shape of transformer magnetizing curves, the required transformer magnetizing current to produce a fundamental frequency sine wave of voltage will contain rather prominent amounts of harmonic currents. In a Y-connerted transformer system energized from a three-phase supply in the absence of a deltaconnected winding, the transformers are unable to obtain a sourre of third-harmonic current or multiples thereof because these are of zero sequence. As the result of the inability to obtain a third-harmonic exciting current, there will appear a third-harmonic voltage whirh may be as much as 50 per cent of the normal operating potential. Should the neutral of such a transformer system become grounded intentionally or accidentally and the supply system be ungrounded or high-resistance grounded, this third-harmonic voltage will be imparted to and appear on the system phase conductors and represent a sustained source of overvoltage. Even though the transformer system neutral is ungrounded, some fraction of the third-harmonic voltage will appear on the phase conductors, depending on the ratio of capacitance to ground within the transformer structure to the distributed capacitance to ground of the rest of the system. Core-type three-phase transformers present a fairly low zero-sequence magnetizing reactance which would hold the zero-sequence voltage to much lower levels than shell-type three-phase transformers or banks of three single-phase transformers and are thus much less susceptible to overvoltage difficulties. If operated with grounded neutral on an ungrounded-neutral system, a careful check should be made to ensure freedom from neutral instability, as treated in reference 3. While grounding the electric system neutral may not solve all the troubles of the Y-Y transformer connections, it will eliminate appearance of overvoltage on the phase conductors of a system to which such a bank of transformers might be connected. Overvoltage Example. A great many specific cases of system overvoltages have been analyzed, identified, and catalogued. All types are well represented. Space will not allow a lengthy treatment of these

298

A

SYSTEM OVERVOLTAGES-CAUSES A N D PROTECTIVE MEASURES

DISTRIBUTION BUS (UNGROUNDED SYSTEM)

1

B C FUSE CUTOUTS

-

I

Q Q

PT2

PHYSICAL CIRCUIT CONNECTIONS

(A1

POWER SYSTEM EOUIVALENT CIRCUIT

PROTECT1V E EOUIPMENT CIRCUIT

A-PHASE FUSE OPEN

EOUIVALENT CIRCUIT FORMED BY OPENINGOF T H E A- PHASE FUSE ( FJI

FIG. 5.1 2

Circuit conditions responsible for an orenoltoge experience on an ungrounded power system.

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

299

specific overvoltage cases. However, it will be interesting to review one case. The one here described has been selected because it discloses how obscure may be the basic overvoltage cause. Note that the series resonant circuit created by the opening of one fuse might very easily fail to be identified, leaving the overvoltage source to remain a mystery. A metal-products plant in the North Central section of the country had made application of a set of rotating-machine protective capacitors and arresters a t the main bus of a medium-voltage distribution system through a set of fuse cutouts. To monitor the fuses, two potential transformers and voltmeters had been applied on the load side of the fuses, as illustrat,ed in Fig. 5.12A. As a result of opening of the fuse unit in the A phase it was observed that voltmeter V , went off scale, potential transformer 1 overheated and melted out the compound, the gap shunting resistor 011 the A-phase arrester was destroyed, and phase-to-ground overvoltages appeared on the phase conductors of the service system. Not until the resulting circuit is redrawn as in Fig. 5.12B is it apparent that the overvoltages result from series resonance (probably of ferroresonance character). System-neutral grounding is to be adopted to ensure freedom from overvoltages on the distribution system conductors. (Additional corrective measures are needed to ensure freedom from overvoltage trouble in the local protective equipment circuit-potential transformer and capacitor shunting arrester.) PROTECTION OF POWER SYSTEMS AGAINST THE OVERVOLTAGES CAUSED BY LIGHTNING The highest overvoltages to which industrial power systems are subjected are those caused by lightning. Limiting these overvoltages by suitable protective measures is essential if costly equipment failures and service interruptions are to be avoided. NATURE

OF THE OVERVOLTAGES

A lightning stroke to earth represents the sparkover of a highly charged condenser, a cloud forming one plate, the earth the other, and the air between the dielectric. The initial charge has been estimated to be as high as 1 billion volts, and stroke currents as high as 200,000 amp have been measured. Although lightning may strike directly a t the terminals of outdoor electrical equipment, this can generally be avoided by proper shielding. Thus, the overvoltages usually reach the equipment (both indoor and

300

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

outdoor) through exposed overhead lines which often bring power t o the plant or, in some cases, distribute power withiu the plant. Direct Strokes and Induced Surges. Lightning may produce an overvoltage on a transmission line either by a direct stroke to the line or by electrostatic induction from a stroke t o earth iri the vicinity of the line. The probable maximum voltage appearing ori a liiie by a direct stroke is 15 million volts and for an induced surge, 500,000 volts. These voltages appear between conductor and ground. Wave Shopes. Although the voltage surges produced hy lightning have high magnitudes, their duration is very short. I t is measured in microseconds (millionths of a second). Typically, the voltage rises very rapidly (in 1 t o 10 psec) t o the maximum or “crest,” value and theu decays more slowly, reaching 50 per cent of the crest value in 20 t o 150 psec. As illustrated in Fig. 5.13, the shape of a voltage or current, surge produced hy lightning (and those produced artificially for test purposes) is customarily expressed by two numherç. The first, is the time from the “virtual zero” of t,he wave front t o the time the wave reaches crest value, while the second numher is the time from the virtual zero t o the time the voltage or current has decreased t o 50 per cent of the crest value. The

-WAVE-FRONT

-

t

WAVE - TA1 L -CREST

VALUE

f

. I I

ZERO TIME O F CURRENT WAVE ZERO TIME O F VOLTAGE WAVE

I

b

&tut

3

I I

4

T i a N MICROSEMXIDS

-

I

1.

_ I

WAVE- SHAPE OF VOLTAGE WAVE ti X 12 CURRENT WAVE t 3 X t e

FIG. 5.13

Termr ured to dercribe voltage cind current waves.

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

301

virtual zero of a wave front is the intersection with the zero axis of a straight line drawn through the points on the front of the wave which are 30 per cent and 90 per cent of the crest value for a voltage wave and 10 per cent and 90 per cent of crest value for a current wave. Both times are usually expressed in microseconds. To illustrate, a 95-kv lf.5 X 40-psec wave is one that has a crest value of 95 kv, rises to crest value in 134 pser from the time of virtual zero, and decays to 50 per cent of crest value (47.5 kv) in 40 psec from the time of virtual zero. Traveling Waves. The voltage surge produced on a transmission line by lightning does not appear simultaneously at all points on the line; instead, it appears at successively later intervals of time as the distance from the point of the st,roke increases. Furthermore, the magnitude and shape (voltage vs. time) of the surge remain approximately the same at all points of a uniform line, but are simply displaced in time phase. In effect then t,he surge which appeared as a voltage-time wave on the line where the stroke occurred becomes two identical voltage-distance waves on the line which travel at uniform velocity in oppvsite directions from the point of origin. Keglecting all resistances, it can be shown that 1. The voltage waves travel along the conductor without change in magnitude or shape with a velocity equal to l / d T C fps, where L is the inductance in henrys per foot of line and C i s the capacitance in farads per foot of line. 2. A current wave accompanies the voltage wave and is of exactly the same shape, that is, a t any instant at any point on the line, the current flowing in the conductor is directly proportional to the voltage from conductor to ground. 3. The ronstant of proportionality between the current and voltage is called t,he surge impedance Z and is equal to 4 r C ohms, where I, i s the inductance in henrys for any unit length of the line and C is the capacitance in farads for the same unit length. The current in amperes is equal to the voltage in volts divided by the surge impedance in ohms. The inductanre and caparitance of an overhead line are such that the velocity of a current or volt,age wave (called velocity of propagation) is equal to the velocity of light in free space, which is 984 ft per psec. In most ralrulations the round number 1000 is used. The propagation velority in a cable varies with its construction, but a typical value is 600 f t per psec. The surge impedanre of an overhead line varies with the size of the ronductor and its height aboveground, but is usually between 400 and 500 ohms. A typical value for a cable is 30 ohms. Reflection of Traveling Waves. A change occurs in a traveling wave when it reaches the junction between two conductors of different surge

302

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

impedance, for example, an overhead line and rahle. The original wave, called the inrident wave, gives rise to two waves at the t,ransition point, namely, a “refracted” wave whirh rontinues on through the second conductor and a “reflected” wave which starts traveling hack over the first conductor. If, at any instant, E is the voltage of the incident wave at the junction, then E X (Z, - Z,)/(Z, ZJ is the voltage of the reflected wave, where Z, is the surge impedance of the first rouductor (over whirh the surge arrived) arid Z , is the surge impedaure of the second ronductor. The voltage of the refracted wave at the junrtiorr is the sum of the voltages of the incident and reflected waves, that is, it equals E X (222)/(Z2 Zi). Reflected and refracted current waves accompany the corresponding voltage waves, the constant, of proportionality being t,he surge impedanre ZIor Z2 of the conductor the wave is traveling oil. A reversal of dirert,ion of a voltage wave, without change i n polarity, reverses t,he direction of flow of current. As indirated by t,he equations, if Z 2 is greater than Z,, a voltage wave reflects positively at, the junctioo and the voltage a t the junrtion (equal to the voltage of the refracted wave) is greater than the vokage of the incident wave. In the limiting rase if 2%is infinite (the line is open), the voltage at t,he junction is double the voltage of the inrident wave. On the other hand, if Z,is less than Z , , the wave reflerts negatively and the refracted wave is less than the incident wave. For the limiting rase of Z2 equal t o zero (the line is shorted t o ground), the volt,age a t the junrtion is, of course, equal t o zero. The current t o ground will equal twire the current of the incident wave. Although neglecting all resistances represents an idealized condition, the simplified relations this makes possihle are useful in many practical situations.

+

+

INSULATION CHARACTERISTICS

It is characteristic of most insulations that t,he maximum voltage which they can successfully ivithstatid varies inversely with the duration of the voltage. Since power systems are subject t o various types of overvoltage, some of long and some of short duration, power distribut,ion equivment is usually required t o withstand at least tivo different types of dielect,ric tests. The first are the so called “lorn-frequency” (00-cyrle) tests, usually of 1-min duration, that cstahlish the ahility of the insulation t o withstand moderate overvoltage of relatively long durat,ion. The others are the “impulse” tests which prove that, the insulation will not break down on vokage surges of high magnitude but short duration. Since the overvoltages produced by lightning are surges of high magnit,ude and

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

303

short duration, it is the impulse tests that are important as far as protection against these overvoltages is concerned. Basic Impulse Insulation levels. The impulse test which is most commonly used consists of the application of a 155 X 4O-psec full-wave voltage surge of a specified crest value to the insulation of the equipment involved. The crest value of the wave is called the basic impulse insulation leuel (abbreviated BIL) of the equipment. T o simplify the design and appliration of elertrical equipment, the Joint Committee on Coordination of Insulation of the American Institute of Electrical Engineers (AIEE), the Edison Electric Institute (EEI), and the Xational Electrical Manufacturers Association (KEMA) have established a series of Standard Basic Impulse Insulation Levels. These are listed in Table 5.1. It was the intent that the impulse level assigned t o any equipment should he taken from the standard series. This has generally been done, but in some cases the value adopted for a given insulation class is that shown in Table 5.1 for a different reference class. TABLE 5.1

Standard Basic ImDulse Insulation Levels Boric

Reference

6 X 40 wave which will cause sparkover on 50 per cent of the applications of this wave. Sparkover occurs on the tail of the wave. The other is the average voltage at which front of wave sparkover occurs with the voltage wave rising at the rate specified in the AIEE standards for arrester tests, namely, 100 kv per psec for each 12 kv of arrester rating. This sparkover voltage is generally higher-as much as 50 per cent higher for some arresters-than the crit,ical sparkover voltage for a I f 5 X 40-psec wave. Arrester discharge voltages usually published are the average crest values of the voltage appearing across the arrester terminals when discharging a 10 X 20-psec current wave having various crest values such as 1500, 3000, 5000, 10,000 and 20,000 amp.

300

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

From the average protective rharacteristirs of lightning arresters xyhivh are puhlished, the masimum values can be determined h y means of iirdustry rerogniaed toleraiires. As shown in Table 5.3, these give the amount hy whirh the masimum sparkover and disrharge voltages of a n arrester may be eupeited to exceed the average values. The various types of arresters listed in Table 5.3 are defined under the heading (:lassification of High-voltage Arresters which follovs. TABLE 5.3

Tolerances in Performance of Valve-type Lightning Arresters

Type of Arrester

1

ayeroge "(IiYe, per cent

I Sparkover voltage .......... ................. .............

Distribution. Line Stotion.

25

Discharge voltage

20

20 15

I5

10

Effect of Altitude. Since the sparkover voltage of a gap varies with the atmospheric pressure, the protective characteristics of arresters are afferted by the altitude a t which they are installed. This is true even if the arrester has a sealed gap since the seals employed are not expected t o maintain a pressure different from the surrounding atmosphere for any extended period. Standard arresters are considered suitable for altitudes up to GOO0 ft. Special arresters are available for altitudes of 6001 t o 12,000 ft and for altitudes of 12,001 t o 18,000 f t . Classification of High-voltage Arresters. Arresters in ratings of 1000 voks and higher are classified in accordance with their principal charact,eristirs and field of application as follows: 1. Distribution-type arresters 2. Line-type arresters 3. Station-type arresters Distribution-type arresters are available in voltage ratings of 1, 3, 6, 9, 12, 15, arid 18 kv. Though designed primarily for the protection of dist,ribut,ion transformers, they are also used to protect other equipment such as metering and switching devices, voltage regulators, distribution rapacitors, and cable. The arresters are small, lightweight units t h a t are readily mounted on poles or crossarms, have reasonably good protective rharacteristics, and are very low in cost. Line-type arresters are available in voltage ratings of 20, 25, 30,37, 40, 50, GO, and 73 kv. They are relatively small and lightweight, are moderate in cost, and have good protective characteristics. They are used for the protection of the smaller transformers and substations in the mediumvoltage range.

309

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

From the average protective rharacteristirs of lightning arresters xyhivh are puhlished, the masimum values can be determined h y means of iirdustry rerogniaed toleraiires. As shown in Table 5.3, these give the amount hy whirh the masimum sparkover and disrharge voltages of an arrester may be eupeited to exceed the average values. The various types of arresters listed in Table 5.3 are defined under the heading (:lassification of High-voltage Arresters which follovs. TABLE 5.3

Tolerances in Performance of Valve-type Lightning Arresters

Type of Arrester

1

ayeroge "(IiYe, per cent

I Sparkover voltage .......... ................. .............

Distribution. Line Stotion.

25

Discharge voltage

20

20 15

I5

10

Effect of Altitude. Since the sparkover voltage of a gap varies with the atmospheric pressure, the protective characteristics of arresters are afferted by the altitude a t which they are installed. This is true even if the arrester has a sealed gap since the seals employed are not expected to maintain a pressure different from the surrounding atmosphere for any extended period. Standard arresters are considered suitable for altitudes up to GOO0 ft. Special arresters are available for altitudes of 6001 to 12,000 ft and for altitudes of 12,001 t o 18,000 f t . Classification of High-voltage Arresters. Arresters in ratings of 1000 voks and higher are classified in accordance with their principal charact,eristirs and field of application as follows: 1. Distribution-type arresters 2. Line-type arresters 3. Station-type arresters Distribution-type arresters are available in voltage ratings of 1, 3, 6, 9, 12, 15, arid 18 kv. Though designed primarily for the protection of dist,ribut,iontransformers, they are also used to protect other equipment such as metering and switching devices, voltage regulators, distribution rapacitors, and cable. The arresters are small, lightweight units that are readily mounted on poles or crossarms, have reasonably good protective rharacteristics, and are very low in cost. Line-type arresters are available in voltage ratings of 20, 25, 30,37, 40, 50, GO, and 73 kv. They are relatively small and lightweight, are moderate in cost, and have good protective characteristics. They are used for the protection of the smaller transformers and substations in the mediumvoltage range.

310

SYSTEM OVERVOLTAGES-CAUSES

AND PROTECTIVE MEASURES

t,he arresters must withstand is 100,000amp for thestation typeand 65,000 amp for the distribution and line types. TABLE 5.4

Industry Average Protective Characteristics of Valve-type Lightning Arresters A v e r a g e impulse rporkover voltage on AlEE test wove, kv

V0ltog

rating, kr

Average discharge oltage with 10,00O-~mp 10 x zo-psec CUrlent wave, kv

__ Distribution OrreSler.

~

~

3 6 9 I2 I5 18

I8 34 48 61 71 84

13 23 35 43

53

...

Line

30 37 40 50 60 73 97 I09 121 145 169 195 242

75 93 110 136 147 183 220 267

... ... ... ... ... ... ...

44 55 69 78

11 22 33 44 54

Line

Or,e.te,S

20 25

I5 30

O,,&e,*

72 89

92 Ill

I06

I35

131 136 178 214 261 345 388 430

I64

51s 602 691 860

I77 222 271 328

... ... ... ... ... ... ...

72 90 108 132 144 179 217 262 349 394 438 523 610 698 872

Arresters and Capacitors for Rotating-machine Protection. A variant of the station-type arrester designed particularly for rotatingmachine protection is offered by some manufacturers. One version (see Fig. 5.16) has characteristics similar t o that of standard station-type arresters but differs mechanically in that it has a porcelain top with the line-terminal connection brought out through the center. This allows placing the three arresters of a three-phase installation close t o each other, thus reducing space requirements to a minimum. The arresters are available in voltage ratings of 3 t o 27 kv with the 3-, 4 . 5 , 6-, 7.5-, 9-,

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

31 I

12-, aiid 15-kv rat,ings of particular interest for industrial npplicat,ions. The 4.5- and 7.5-kv voltage ratiirgs are not, available in t h e standard station-type arresters. They are iiicliidrd in this line to give bet,ter protection to 1.16- and 6.0-kv machines tliaii rim be provided by t h c standard 6- aiid 9-kv arrest,ers. Tlic latter ~vouldotherwise be required where the paver syst,cnis are riot, sufficiently well grounded to permit t,lie USC of 3- and ti-kv arresters on 4.16- and G.!bkv rnachiiics ( Arrester T’oit,age Ratings). The coristructioii fcatorcs niid additioiial voltage ratiiigs available make these arresters dcsirahle for iit,her app1ii.atioiis such as t.he protection of switchgear. Surge protective capacitors are also available for rotatiiig-mii~hiiie protection. They are used to reduce tlic stcepriess of the wdve front of lightning surges aiid arc available in ratirigs of OM50 volts with 1 .O pi per pole, 2.1, 1.16, 1.8, arid 6.9 kv with 0.5 pi pf’r pole, and 11.5 and 13.8 k v with 0.2.5 pf p t pole. ~ l’liese capacit,ors differ from thc staridad porver-fact,or impr(iviiig capacitors i i i that they are designed t o withstaiid higher test, voltages and have low interrid inductance. A typiciil unit is shovii in Fig. 5.17. Low-voltage Arresters. For thc prntectiou of etluilimixit on circuits whose line-to-ground voltage is iri the 110- to 125-volt range, a 175-volt

FIG. 5.16 Rotating-machine form of station-type lightning arrester rated 6 kv.

FIG. 5.17 Surge protective capacitor rated 6900 volts, 25 to 60 cycler, 0.5 ilf.

312

SYSTEM OVERVOLTAGES--CAUSES AND PROTECTIVE MEASURES

liglituing arrester is avsilal-lie. This is built in a two-pole lorm; so a single unit will provide protect,ioii to the common 1 15i230-volt siriglephase tliree-wire grounded-iieutral circuit,. A t y p i d iiistallation is showii in Fig. 5.18. For a two-wire circuit, grounded o r 1 oiie side, the two poles of the wrcster arc generally roniiecied in parallel between the uiigrouiided h i e arid gruund. For three-phase circuits such as those supplied from a208Yjl20-volt grounded-imitral system, t x o arresters arc required. For the protection of equipmerit on higher voltage circuits-up t o 600 volts-~-~twu forms of arresters areavailahle, both rat,cd 650 volts. One has a port:elaiii housing (see Fig. 5.19), is for oiitdoor service oiily, arid is availablein a singlepoleaiidatno-poleform. T h e other has t i niet,al enclosure (see Fig. 5.20), is suitable for either indoor or outdoor service, axid is availsblc in one., two-, arid t,lirce-polc forms. This unit also has better prok c l i v e characteristi(,s and so FIG, 5.18 lnrtallotion of two-pole 175-volt is t,hc oIic usually selected for lightning orrester on o 115,'230-volt single-phore protection of indust,rial plaiit three-wiro circuit. equipmeti t. Arresters for D-C Systems. .krrcst,ers designed for use on a-c power systoms are iiot getierally suitable for service on d-c employed t o interrupt follow r'urrrirt is not cffectiv diics not periodir:ttlly go through zero. Arresters, hinrevcr, arc availrtiilo for d-? scrvicc. The moderir forms arc simply capacitors having iiot less tiinti 4 pf of capnt,itaricc. Tiicy are coiiriected from line to ground arid limit, tlit. tw:st, - i d t i e of a volt,age surge by absorbing the current as a charge o i i ttic capwitor. fleiice llie effectiveiiess of the arrester in limit,irig the \-oltnge of ail iticwmitig surge depends upon the duration as i\-ell as the magriitude of the surge. lrmvever, it, also ser t o slope t,he froiit of tiit? ivavc a r i d tliiis reduce the turii-to-turn voltage sircss on the d-c rotating mnr.liities. 'I'hc arresters are available i i i three voltage classes, iiamely, Obi30 volts (illustrated iii Fig. 5 . 2 ! ) , 751-2000 v o h , and 2001:?&00\-o1ts. (I

SYSTEM OVERVOLTAGES-CAUSES

FIG. 5.19

AND PROTECTIVE MEASURES

313

Single-pole lightning arrester

FIG. 5.20 Three-pole lightning or-

with porcelain housing roted 650 volts, for outdoor service.

rester in metol core roted 650 volt>. for indoor or outdoor service,

FIG. 5.21 cirwiti.

Capacitor-type lightning orrerter rated 0 to 750 volts, 4 rrf for w e on d-c lnruloting cop and sleeve removed ot one end to show terminol.

314

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

APPLICATION PROCEDURE

Every exposed overhead line distributing power within or supplying power t o an industrial plant represents a possihle sourre of destruitire overvoltages. Lightning arresters should be so applied that a voltage surge from any of these sources will be reduced to a ralue ~ v e l lheIoi\- the impulse strength of all apparatus involved. The application procedure consists of (1) selecting t,he voltage rating of the arresters t o he used, (2) choosing t,he types of arrest,ers needed, and (3) determining where the arresters should be located to ensure adequate yet economical protection. Selection of Arrester Voltage Ratings. The protective characteristics of an arrester are hetter and, in general, its cost, is lower, the lower its voltage rating. On the other hand, if the line-to-ground system voltage after sparkover of a n arrester should exceed its voltage rating, the arrester may not interrupt follow current and then iI-ill fail very quickly. This makes it important t o determine the maximum lilie-to-groutid system voltage at the point at which the arrester is applied. 111 so doing it is necessary t o consider all abnormal conditions which ran exist, particularly those conditions which are likely t o exist when the arrest,er sparks over. Under normal balanced operatirig conditions, the voltage from each line t o ground on a three-phase system is the syst,em line-to-line voltage divided by the square root of 3. This applies vhethcr the system neutral is grounded or ungrounded. There are, however, many abnormal rotiditions which can occur that result in higher t,hari normal line-to-ground voltages. Hut the one that is most likely t o exist a t the time of arrester sparkover is a line-to-ground fault. For example, if a lightning stroke causes flashover and hence a fault on one phase of a transmission line, the voltage indured on the sound phases is apt, t o cause sparkover of the arresters connerted t o these phases. These arresters must then interrupt, follow curreut, with a line-to-ground fault on the system. The voltage ratings of arresters are, therefore, generally selected 011 the hasis of the system voltage t o which they arc subjected under line-to-ground fault condi t,ions. The voltage from sound conductors t o grouud with a line-to-ground fault 011 a system depends upon how the system neutral is grounded. For the usual ungrounded or resistailre-grounded system, t,his vokage will be essentially equal to the system line-to-line voltage, and the lightning arresters used must be selected 011 this basis. Thcse are siimetimes referred to as ‘‘ 100 per cent arresters.” However, for solidly grounded or reactance-grounded systems the sooiid-rotidurtor-to-ground voltage with one line grounded may be as low as the system line-to-neutral volt-

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

315

age. It depends upon the relation between the zero- and positivesequence impedances of t,he syst,em. For example, if the ratio of zerosequence reactance X Ot o the positive-sequence reactance X I is positive and less than 3 and the rat,io of the zero-sequence resistance R , t o the positive-sequence reactance X I is less thau 1, the voltage from sound conductors to ground will not exceed 140 per cent of the system liue-toneutral voltage or about 80 per cent of t,he system line-to-line volt,age. Such a system is said to he “effei.t,ivelygrounded,” and t,he arresters used are referred to as “80 per rent arrest,ers.” Some syst,ems are grouudcd so that arresters of even lower voltage rating can he used as far as the orervoltage caused by line-to-ground f a u h is concerued. This, however, should he done only after a careful check of the possible overvoltages from all sources t,o make sure that v o h g e s in excess of t,he arrester rating are not likely to occur at the time of sparkover. Table 5.5 lists the voltage ratings of arresters usually selected for (1) ungrounded or resistauce-grouuded systems and (2) “effectively grounded” systems. Selections are show1 for all system voltages likely to he encountered in industrial plants. As shown in Table 5.5, 3-kv arresters are often used on 2.4/4.1C,Y-kv grounded-oeutral systems and 9-kv arresters on 7.2/12.47-kv grouridedneut,ral syst,ems, akhough in t,hcse cases the arrester rating is only 125 per cent of t,he nominal system line-to-neutral voltage. Before using these lmi-er rat,ed arresters, the maximum operating voltage and the rise iu soulid-conduct,or-to-ground rokages with a linn-t,o-grouud fault, should be determitied t o make sure that under these conditions the voltage applied to the arresters will not exceed their rating. I n geueral they should not be used on industrial pmver systems unless (1) the ratio of zero-sequenre reart,ance X o to the positive-sequence reactance X I is less thau 1.5 and (2) the ratio of the zero-sequence resistance Ro t,o t,he positive-sequence reactance X I is less thau 0.5. Even though a system meets the qualifications of an eflectively grounded system at the power source, it may not a t other points in the system because of the impedance of intervening lines. Furthermore, the system may be “effectively grounded” under uormal operating conditions, but certain faults or other emergencies may result in the opening of switches which leaves a portion of the system ungrounded but still energized either from generators or from mot,ors whirh can temporarily act as generators. Such possibilities should he considered before selecting the voltage rating of arresters to he applied on what appears t o be an effectively grounded system. Choice of Arrester Type. Where the arrester voltage ratiug required is 3 t o 15 kv, a choice must be made between the distribution-type and the station-type arrester. Similarly, if the rating required is hetween 20

316

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

TABLE 5.5

Voltage Ratings of Arresters Usually Selected for Three-phase Systems Voltage roting of arrester, kv

Nominal system voltage, kv

0.120/0.208Y 0.240 0.480 0.600 2.4 2.4/4. I6Y 4.16 4.8 6.9 12 7.2112.47Y 13.2 (or 13.81

23 34.5 46 69 115

I38

Sy*tom "e"Ir.1 ungrounded or 'eiirtonce groundeq

0.65 0.65 0.65 0.65

System neutral effectively grounded

0.175 0.65 0.65 0.65

3

t 4.5,;

4.5. or 6 4 . 9 or 6 6

3

or 6 4.5. or 6 4.5* or 6

6

7.5*or9

12

I5

9 t o r I2

15 15

12

25 37 50

20 30 40

73 121 145

60 97 121

* The 4.5- and 7.5-kv arresters are available only

;he station type. 1 less than that necessary to make the system "effectively grounded" (see accompanying text)

t The use of these arresters requires an X o / X ,

I

and 73 kv, either the line-type or the station-type arrester must he selected. The value of the equipment protected and the importance of uninterrupted service in an industrial plant generally warrants the use of stationtype arresters throughout their voltage range. However, for the smaller (liquid-filled) transformers and substations, say 1000 kva and less, distribution- or line-type arresters are frequently used. Similarly, for the protection of short lengths of cable joining overhead lines and apparatus, these lower cost arresters are generally chosen. They are also used to protect small breakers, disconnecting switches, and similar outdoor switching equipment. Finally, distribution-type arresters are often used in the protection of rotating machines, thereby supplementing the protec-

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

317

tioii provided h y statioii-type arresters (sec l'rotcrtioii of h-C Iliitatiiig hlarhiiirsj. Location of Arresters. The ideal location of lightiiiiig arrestcrs, from the staiidpoi~rt,:if the prutrrtioii whirh they provide, is directly at the terminals of t h e apparatus heiiig protwted. .kt this location, aiid with the arrester groutid leads i.oiinerted direi.tly to the tank, framc, or other metallii, strnctiire h i r h supports the iiisrilated parts, thc surge voltage applied to the itisrtlatioti will he limited to the sparkover vultage aiid the discharge voltage of the arresters. Iii some cases, howe\-er, it might I)P quite costly or aivk\\-ard t o muiiiit the arresters at tlie apparatus tcrmiiials. Furthermore, i i i somc iiistatlations, if the arresters are mo\-cd away from the trrmitials 11f the protected equipme~it.a single set of arresters caii he lorated \\-here they will intercept all lightiiiiig surges to two or more pieres of apparatus. H o ~ e v e rsuch separation hetwecti lightning arresters alid thc eqriipme~it,does mean some itiiwase i n the magiiitude of the voltage surge 11-hivh is applied t o t h r eiluipmmt. First, the equipmelit protevted will ofteir have a highrr surge impedtilice than that iif the h i e or mhle over \\-hich the lightiiiiig srirge arrives. This means that thc voltage wave will refle1.t positix-ely nt the equipmetit termiiials aiid the 1-oltage rearhed at this poiiit n-ill al\\-ays he lriglrcr than the sparkover v d t a g e of the arrester. T h e amoriiit of the itirvmse will depend upoii ( I ) the steepiirss of the froiit of the srirgc viiltagr, (2) the relative surge impedance of the eqnipmeiit aiid the circuit hetiweti the arrester and the protected equipmeiit. (3) the sparkowr \-iiltnge of the arrester, and (4) the length of the rirt,nit hrtivreli the arrester and the protwtrd eiluipmeiit. The greatest i i i i ~ r a s riii voltage wciiss if the cirruit is iipeir at the protected eiluipmetit (iiititiite surge i m p d a t i w j . 111 this rase tlie voltagr will IK dinible the arrester sparko\-er voltagr if the sepitration distairre is such that parko over ownrs before tlic voltage wave reflected from the eiliiipmriit arrives hack at the arrestrr, U'ith less separatiiiii the voltage will iiot iiirreasr a s miidi. This is showi h y the iwrves of Fig. 3.22. Citrve ;I applies if the overhead liiie. over \\-hirIi the surge arrives. estends past the arrester to the priitri,tid eqiiipmetrt, i\-hile curre B applies i f a i,ahle of typical chnravteristiis forms the cirruit het\\-eeti the arrester aiid t h r proterted equipment. The voltage whirh appears arross an arrester after spnrkovrr, i.e., its disrharge voltage. is also magnified by separation atid priidrwrs ii Iiighrr voltage at the protevtcd equipmeiit. Fnrthermiirr. if thew is ail? appreriable lciigth of lead hetween t h e h i e rolidrii~tiir atid the arrester or het\\-een the arrester atid griiulid, the voltage drop wross surh a lead adds to the discharge voltage of the arrester aiid is also itiiwased by separation betxi-eeii the arrester and the protected eqiiipmrtit. Finally, if a11 arrester located away from the protected equipment has a11isolated co~iiicctioiito

318

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

ground, the additional voltage drop resulting from discharge current flowing through the ground resistance also adds to the line-to-ground voltage a t the arrester and a magnified addition appears a t the protected equiqment. Certain installation practices help to reduce the difference between arrester discharge voltage and the corresponding voltage a t the pro-

FIG. 5.22 Effect of reparotion between a lightning arrester and the protected equipment on the rotio of the maximum voltage a t the equipment to the sparkover voltage of the arrester (doer not include any effect of the voltage at the arrester following its sparkover).

SYSTEM OVERVOLTAGES-CAUSES

AND PROTECTIVE MEASURES

319

tected equipment. For example, where an arrester is connected between an overhead line and ground, the leirgth of the line and ground leads can both he reduced to a minimum by use of the V connection. The arrest,er is placed a t ground level, and the line coriductor is brought down to the arrester and then back up, forming a V. The angle hetween the two sides of the V should not be less than 30" to minimize their mutual inductance. The effect of high ground resistance a t the arrester ran be minimized by interconnection of the arrester ground terminal with the tank or enclosure of the protected equipment, the station steelwork, and the ground mat. Finally, where the circuit hetween an arrester and the protected equipment consists of cable having a contirruous metallic sheath, the arrester ground terminal should he connected directly to the cable sheath and the sheath connected to the equipment tank or enrlosure. In this may arrester lead lengths can he kept to a minimum and the effect of ground resistance eliminated. More specific recommendations covering the application of arrest,ers for the protection of various types of equipment,, including suggested maximum separation distances, are given in the remainder of this chapter. PROTECTION OF TRANSFORMERS

Transformers generally constitute one of the must important elements of any industrial power system. Furthermore they are frequently connected directly to exposed overheadlines and so are suhject to destructive overvoltages unless properly protected by lightning arresters. A liquid-filled (oil or askarel) transformer having arresters mounted a t its terminals is well protected against the overvoltages produred hy lightning, with the possible exception of those result,ing from severe direct strokes to the transformer terminals or to the conneitcd lines close t,o the transformer. Furthermore, the possibility of such direct strokes can he essentially eliminated by proper shielding. Often, howerer, in order t o protect (with the same set of arresters) switching and other equipment located between the transformer and the exposed lines, or to protect two or more t,ransformers connected to the same line, it may appear desirable to mount the arresters some distance away from the transformer terminals. The maximum permissible separation distanres depend, among other things, upon the magnitudes and rates of rise of the voltage surges which can he expected to reach the arresters. Until more statistical data on these surges are available, no determination of permissible separation distances can be considered final. Hou,ever, making That appears to he reasonable assumptions, a Working Group of the AIEE Suhcommittee on Lightning Protective Devices (of the AIEE Committee on Protective Devices) proposed the maximum separation distances shown in Table 5.6.* The installation conditions on which these distances are

* Ser AIEE Misccllaneous Paper 51-285.

320

SYSTEM OVERVOLTAGESS-CAUSES AND PROTECTIVE MEASURES

based arc that ( 1 ) the transformer is fully insulated (liquid-filled), (2) statioii-type arresters are used, (3) arrester lead lengths are zero (V r~~nncctioir or eqnivalent), (4) ground resistance is negligible, and (5) the transformer is a t the elid of a single overhead line (the worst condition) with the arresters located on the line directly in the path of incoming sr1rgcs. TABLE 5.6

Separation Distance Permissible between Station-type Arresters and Transformer Bushings Separation diitonce, ft

Tiomformer i"lYl.ti0" CI.I.,

kv

Botic impvke inrvlotion

IWel, kv

System neutral

System neutrd

ungrounded or esistance grounded 1100% arrosten)

effectively grounded

180% arresters)

-___ 25 34.5 46 69 92 115 138

I50 200 250 350 450 550 650

25 25 25 ~. 25

30 35 30

70 70 70 70 75 85 95

For transformcrs of lower volt,age ratings (15-kv class and below) which are not covered i t t Table 5 . 6 , permissible separation distances have not becti estahlished. Severtheless it appears that for these ratings any apprecialile scparatiott should be avoided, that is, the arresters should he mounted 011 t,he transformer itself or closely adjacent to it. In ratings of 15 kv and helow, transformers are often connected t o exposed overhead lines through a length of cable. I n this case fully insulated liquid-filled power transformers ronnerted t o the overhead line through a cable having a continuous metallic sheath will be adequately protcvtrd by statioii-t,vpe arrestcrs located at the junction of the cable and the overhead l i t e. Thc arrester ground terminals must he connected directly t o thc catilc sheath, and at the transformer the cable sheath must lie rotinerted t o the transformer tank. If the transformers are of the distribution rathcr thaii the po\ver rlass or if distrihution-type rather than station-type arresters are provided at the junction of the cable and overhead line, it may he necessary to add a set of arresters at the t,ransforme,r terminals to eiisurc adequate protection. Dry-type transformers, \\-hose impulse level is ahout half that of the

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

321

liquid-filled type, are not generally recommended where connection t o exposed overhead lines is required. If used they should definitely be protected by station-type arresters located at the transformer terminals regardless of whether the connection t o the exposed overhead h i e is direct or through a cable. If a liquid-filled transformer is connected t o an exposed overhead line only through another transformer which is adequately protected by lightning arresters, no additional protection is required. In the same situation a dry-type transformer should, preferably, have station-type arresters mounted a t its terminals since analysis iiidicates that the surges that come through the other transformer can have magnitudes greater than the recognized impulse level of the dry-type units.

PROTECTION

OF

METAL-CLAD SWITCHGEAR*

Metal-clad switchgear (used on 2.4 t o 13.8-kv circiits) is often connected t o an exposed overhead line either directly from roof bushings or through a moderate length of cable. In either case it is esseiitial that adequate lightning protection be provided. If the switchgear is connected directly t o the overhead line from roof bushings, lightning arresters should always be provided a t the gear. Although the arresters are sometimes mounted on the first structure away from the gear which supports the overhead h e , the resulting separation between the arresters and the protected equipment substantially reduces the effectiveness of the protection. Heiice locating the arresters at the gear is definitely recommended. They may be mounted on the roof of the switchgear enclosure adjacent t o the bushings or inside the enrlosure but on the line side of the breaker. Since the former arrangement generally requires an extra ground bus, the inside mounting is usually selected. The arresters should preferably be of the station type (rotating-machine form), but space limitations may sometimes make i t necessary t o use the distribution type. The voltage rating selected should he the lowest that is consistent with the system voltage and method of grounding. If the metal-clad switchgear is connected by cable t o the exposed overhead line, the first requirement is that arresters be provided a t the junction of the cable and the overhead line in order t o protect the cable. Then if the cable does not have a continnous metallic sheath, a second set of arresters should be provided at the switchgear. I n this case dis-

* Adapted from Dillow, Gittings. Halherg. Hoffman. Howard. and Hontrr, Lightning Protection of Mptalclad Saitchgear and Unit Substations Connected to Overhead Lines, Gen. Eke. lieu., March, 1949.

322

SYSTEM OVERVOtTAGE5,-CAUSES

AND PROTECTIVE MEASURES

tribution-type arrcsters are usually used a t the junction, but those at the switchgear should preferably be of the station type (see Fig. 5.23). If the cable eoiiiiectiiig niet,al-clad gear t o an exposed overhead line does have ii coritiiiiious metallie sheath, the set:orrd set, of arresters at the gear may or may not be rec.]uired. It depends upon (1) tho iiisulation level of the gear, ( Z j the type and ... . . . .. voltage rating of the arresters provided at the junction, aiid (3) the length of the cable. An analysis of this ease was made on t,he basis of the followiiig assumptions : 1. The arrestors at the jiinctioii maintain a voltage at, this point which does not exceed the sparlrover voltage of the arrester as given in Table 5.8. 2. The maximum voltage at the switchgear must be limited to 80 per cent, of its BIL. 3. The volt,age waves which appear on the overhead liue arid reach the cable junction have a const,ant rate of rise which does not exceed 1000 kv per psec. 4. The surge impedance of the overhead line is 500 ohms, and t h a t of the cable is 30 ohms. FIG. 5.23 Stofion-type lightning orresterr 5 . The velocity of propagat,ion (rotating-mochine form) mounted in metalof the surge iii the cable is GOO f t clad switchgear. per $see. The r e s u k of the analysis are shown in Table 5.7. I n all cases the grouiid terminal of the junction arrester should be coriiiected to t,lre cable sheath as me11 as t o ground, aiid at the switchgear the cable sheat,h should he eonri d to the ground bus (see Fig. 5.24A). This is essential if no arresters are provided in the gear aiid is desirable in any case. Where large single-coiiductor cables are used, it may not be desirable t o ground bot,b ends of the sheath because of excessive sheath curreiit. I n this case the lightning-arrester ground terminal should still be coririected directly t o the cable sheath arid the sheath grounded at the switchgear, hut the connection t o ground at the arrester should be made through aii isolatiiig gap, as shown in Fig. 5.24R.

SYSTEM OVERVOLTAGES-CAUSES AND PROTECTIVE MEASURES

TABLE 5.7

Protection of Metal-clad Switchgear Connected to Overhead lines through Continuous Metallic Sheath Cable Arresters in witchgeor (required or not requiredl

System voltage

Voltage roting m d

811 of switchgear. kv

4.16

Neutral ungrounded

effectively grounded, kv

or

resistance grounded, kv

2.4or4.16

13.8

4.8 4.8 or 6.P 11.5 13.8

’The use of

Voltoge rating of .2,T&*,S

With dirtribution-typo

iundion, k?

arresters 01

the iunction

4.5 6

....

Not required Not required

Required

(55 h)f

t

Not required Not required Not required Not required (30 ftlt Required*

4.5

4.8

6

6.9

7.5

(75 tilt

t

P

6.P 11.5 13.8

12 15

-

With station-Wpe arresterr rrt the junction

t

Not required

3.

2.4 4.16 4.16

160 811)

I95 8111

323

Required Required Required

arresters on a 4.16-kv system requires an X d X , ratio IPSS than that necessary t o make the system “effectively grounded” (see Selection of Arrester Voltage Rating). t The 4.5- and 7.5-kv arresters are available only in the station type. t Arresters required in snitchgesr if length of cable exceeds this value. TABLE 5.8

Y

Sparkover Voltage of Arresters Used in Analysis of Protection Rewired for Metal-clad Switchgear Sparkover voltogs. k r

V0lt.go rating of arraters, kv

-I

Distribution-type .r,der*

Stofion-type .lr,e*t*r*

I

3

22

15

6

42.5 60 74 81.5

25 37 52 64

P 12

15

324

SYSTEM OVERVOLTAGLS-CAUSES AND PROTECTIVE MEASURES

type arresters and a set of protective capacitors (as used for rotatingmacliitie protection) at the junction of the rable and overhead line. The ground terrniual of both the arresters aiid the caparitors should be connected to the rable sheath as \yell as t o ground (directly or through a n isolating gap), aud the (.able sheath should he eoiniected t o the ground bus a t the switchgear.

EXPOSED OVERHEAD LINE

11,

a l l - P ~ ~ I?

CABLESHEATH

THIS ARRESTER MAY NOT BE REQUIRED

-

4

q&2 ’

-T

. 1

-i rneonr of neutrd bur and switchgear.

374

SYSTEM GROUNDING

to assure safety to personnel. If disconnecting switches are used (as with some outdoor installations), they should be elevated or metal-enclosed and interlocked in such a manner as to prevent their operation except when the transformer primary and secondary switches or generator line and field breakers are open. As shown in Fig. 6.21, it is necessary to provide only two neutral breakers, and only one of these is closed, although all generators may he in operation. This will eliminate any circulating harmonic zero-sequence currents. When the generator whose neutral is grounded is to be shut down, the second generator is grounded by means of its neutral breaker before the line and neutral breakers of the first one are opened. This procedure will assure that the system is grounded a t all times. I n the case of multiple transformers, all neutral breakers may he normally closed because the presence of delta-connected windings (which are nearly always present on at least one side of each transformer) minimizes circulation of harmonic currents between transformers. Selection of Arrangement. When total ground-fault currents with several individual resistors would exceed about 4000 amp, it is suggested that neutral switchgear and a single resistor be considered for reuistancegrounded systems. When only one source is involved hut others may be added to the station, it is suggested that space be allowed for neutral switchgear to be added if this will be necessary later. For similar generators with reasonably equal load division, circulating currents are negligible, and it is often found practical to operate with neutral breakers of two or more generators closed. This simplifies operating procedure and increases assurance that the system will be grounded at all times.

CALCULATION OF GROUND-FAULT CURRENT

The magnitude of current which will flow in the event of a ground fault on a solidly grounded system is usually determined hy the impedance of the grounded apparatus, plus the impedance of the lines or cables leading to the fault and the impedance of the ground return path. For interconnected systems, calculation of the rurrent may he rather complicated. For simpler cases, an approximation of the available fault current may be obtained from Table 6.4. This table applies only for faults near the transformer terminals when power is supplied by a single transformer hank wit,h its neutral directly connected to earth and with the primary connected to a system of relatively large short-circuit capacity.

S Y S T W GROUNDING

375

RESISTANCE GROUNDING

When a single line-to-ground fault occurs on a resistance-grounded system, a voltage appears across the resistor (or resistors), nearly equal to the normal line-to-neutral voltage of the system. The resistor current is equal t o the current in the fault. Thus the current is practically equal to line-to-neutral voltage divided by the number of ohms of resistance used. For example, consider a 13,800-volt three-phase system grounded by a 4-ohm resistor. Normal line-toneutral voltage for this system is 13,800/-\/3, or 8000 volts. The ground current is, therefore, very nearly equal to 8000/4, or 2000 amp. If two such resistors were used on the system, the ground current would be approximately 4000 amp. Resistors have a voltage rating equal to line-to-neutral voltage and an ampere rating equal to the current which flows when this voltage is applied t o the resistor. Thus, for example, a maximum ground-fault current of approximately 2000 amp will he obtained on a system when using a 2000-amp resistor. This very simple method of calculating the ground-fault current is not suitable except when the ground-fault current is small compared with the three-phase fault current for a fault a t the same location. Horucver, it is usually suitable for systems grounded by resistance of ohmir values normally used. The method just outlined applies to faults on lines or buses, or at the terminals of machines or transformers. If the fault is internal to a rotating machine or transformer, the ground-fault current will be less. The reduction in current is primarily due to the internal voltage of the apparatus. I n the case of Y-connected equipment, this internal voltage is a t full value a t the terminals and is zero a t the neutral. If the fault occurs a t the neutral of any apparatus, no voltage will appear across the system grounding resistor; so the fault current will be zero. At intermediate points in the winding between the neutral and a terminal, the fault current will he intermediate between zero and the current to a terminal fault, as shown in Fig. 6.22. For example, at a point 10 per cent of the winding length from neutral, the ground-fault current will bc approximately 10 per cent of the value for a terminal fault. For a fault anywhere between this point and a terminal, the current will be more than 10 per cent of the amount for a terminal fault. In the case of delta-connected machines the internal voltage to neutral may he considered to he 100 per cent a t the terminals and 50 per cent a t the mid-point of the windings, as shown in Fig. 6.22(c). The mid-points have the lowest potential with respect to the electric neutral of any d h e r

SYSTEM GROUNDING

376

UNGROUNDED GENERATOR NO.1 FAULT OCCURS ÇOMEWHERE IN WINDING E E TWEEN NEUTR AND L I N E TERMINAL

GROUNDED - N E U T R A L GENERATOR NO.2

GROUNDING RESISTOR _íGROUND)

( o ) FLOW OF GROUND-FAULT CURRENT FOR I N T E R N A L FAULT IN WYE-CONNECTED GENERATOR b

N0.I

b

-

N0.2

-

(GROUND P O T E N T I A L )

-' " " ' ~ " ~ ' ~ ""'~ ~ ' ~ t

t

PHASE

IF

VI Vr I F

R

I N T E R N A L VOLTAGE T E R M I N A L VOLTAGE FAULT C U R R E N T GROUNDI NG RESIÇTANCE

IF =

(0)

O F EUS

VT -(V,

-VI]

= -V I

R

R

( b ) MAGNITUDE OF CURRENT FOR INTERNAL GROUND FAULT I S PROPORTIONAL TO I N T E R N A L VOLTAGE

INTERNAL VOLTAGE

100%

@

VOLTAGE FROM ELECTRICAL NEUTRAL TO ANY POINT ON WINDING I S EETWEEN 50% AND 100%

INTERNAL VOLTAGE 50% (c)MINIMUM GROUND-FAULTCURRENT FOR DELTA CONNECTED APPARATUÇ IS 50% OF MAXIMUM GROUND-FAULT C U R R E N T

-

FIG. 6.22 Magnitude of currenl for interna1 g m m d faulb in maichinei connected to ryrtem having a rerirtance-grounded neutial.

(I

part of the windings. Therefore, a ground fault at any point in the winding w i l l produce a ground-fault current of 50 per cent or more of the resistor current rating. REACTANCE GROUNDING

In a rractance-grounded system with a single line-to-ground fault, the ground-fault current may he compiited from the formula

SYSTEM GROUNDING

377

(6.1)

(resistance may usually be neglected) where X I = system positive-sequence reactance, ohms per phase X , = system negative-sequence reactance, ohms per phase X o = system zero-sequence reactance, ohms per phase X , = reactance of neutral grounding reactor, ohms E = line-to-neutral voltage, volts I , = ground-fault current, amp An illustration of the method of calculating the ground-fault current in a reactor-grounded system is given under Selection of Reactor Rating (see page 381 of this chapter). SOLID GROUNDING

In a solidly g r o u n d 4 system with a single line-to-ground fault, the ground-fault current may be computed from the formula

RATING OF GROUNDING EQUIPMENT Grounding resistors, reactors, and transformers are normally rated to carry current for a limited time only. The standard time-interval rating usually most applicable for industrial systems, with relays arranged to protect the grounding equipment, is 10 sec. The voltage rating of a grounding resistor should be the line-to-neutral voltage rating of the system. The insulation class of a reactor is determined by the circuit line-toneutral voltage. The voltage rating may be less than line-to-neutral voltage, it being cakulated by multiplying the rated current by the impedance of the reactor. The voltage rating of a grounding transformer should be system line-toline voltage. Grounding resistors are rated in terms of the initial current which will flow through the resistor with rated resistor voltage applied. Conventional cast-grid or corrosion-resistant steel resistors will average approximately 7 per cent increase in resistance for each 100 C rise in temperature. The rated current of a grounding reactor is the thermal current rating. I t is the rms neutral current in amperes which the reactor will carry for its rated time without exceeding standard temperature limitations. The rating establishes an rms current which is assumed to be constant during

378

SYSTEM GROUNDING

rated time for purposes of design, ralrulation, and test. In service it is expected that the current may be greater than rated value during the initial cycles of the fault. If a grounding transformer neutral is solidly connected t o ground, the current which will flow during a ground fault is primarily determined by the reartance of the grounding transformer. When a resistor is used between neutral and ground, the current rating of the grounding transformer is based on the resistor rated current. I n either case the transformer israted t o carry the required current for rated time nithout exceeding its rated temperature limits. Ratings of neutral grounding equipment are summarized in Table 6.8. TABLE 6.0

Ratings of Neutral Grounding Equipment

Equipment

......................... ......................... ............

Reridor.. Reador.. Grounding transformer..

Time, sect

Reactonce

I

.............

I 10

* Insulation rlass is drtrrrnintd hy circuit line-to-neutral v a l t a g ~ . t Tcn sxonds is ntlrquate ior the conventional system. Standard ratings oi 1 mi", 10 mi", and continnous are svailablc.

SELECTION OF RATING OF GROUNDING EQUIPMENT RESISTOR RATING

The determination of the resistor ohmic value, thus the magnitude of ground-fault cnrrent, is based on (1) providing suflicient current for satisfactory performanre of the system relaying scheme and ( 2 ) limiting ground-fault current t o a value which will produce minimum damage at the point of fault. 111most cases, the ground-fault current may be limited hy the iieut,ral resistor t o a value from 5 t o 20 per cent of that which would flow for a three-phase fault. T o determine the minimum ground-fault rurrcnt required, a diagram of the system must be available giving ratings of current transformers and types of relays for each circuit. This diagram should include Consideration of future changes. The magnitude of ground-fault current innst, he sufficient for operation of all relays. In general, if the current is high enough t o operate the relays on the larger circuits, it will he adequate for the smaller circuits. The ground currents required for satisfactory operation of various types

379

SYSTEM GROUNDING

TABLE 6.9

Selection of Grounding Resistor

(Values given ore minimum recommended sround-fault current in per

cent

OF rrrred current of current

tronrformcr.)'

Type of relay Equipmsnt protected

per Cent

Y-connected generators. motors, ond transformer.. Delta-connected gcncraton, motors. ond t r m s ,

........................... former,... ......................... Foeden and tie liner.. .................. B",e* ................................

40 100

...

40

...

Pilot wire-100% Current b d m c e - l 0 0 %

501

* For further discussion and analysis of ground-fault rdsying. ser Chap.

0. ground differential is a d d d to the generator, the ground-fault current may be lirnitzd to lower values (if othw systzrn requircmmts permit). $ Based on current differential. If voltage differential is providcd, the groundfault current may be limited to lower values.

t If

of relays, expressed in terms of current-transformer rating, are given in Table 6.9. Note that the ground-fault current under all system operating conditions must equal or exceed the minimum required for relaying each circuit connected to the system. This value is established by selecting the highest of those currents which meet the requirements of the several conditions set forth in Table 6.9. An example of the proper use of Table 6.9 for the system shown in Fig. 6.23 follows: Determine from Table 6.9 the ground-fault current each generator must produce when it is the only pover source. The larger machine must produce a ground current of at least 1200 amp (100 per cent of the rating of the current transformers for differential overcurrent protection in the larger generator circuit). This ground current of 1200 amp is higher than is required by Table 6.9 for any other circuit, in the system. With the larger generator disconnected, the smaller machine must provide a ground current of only 800 amp for its own relaying requirements (again 100 per differential overcent of the rating of the current transformers for its 01~11 current relays). The 1200-amp circuit need not be considered under this operating condition, and the 800 amp needed in the smaller generator circuit is found from Table 6.9 to be adequate for relaying requirements in all the other circuits. If all sources are grounded, it can be show1 that

SYSTEM GROUNDING

380

there will always be sufficient ground current for relaying requirements as long as each source produces what is needed when it is the only supply source. For Fig. 6.23 it is correctly concluded that the larger machine needs a 1200-amp resistor and the smaller oue only ao 800-amp unit. If ground differential relaying were added to the two generators, a

P$

1200 AMPERE GROUNDING RESISTOR

1200/5 C T

FEEOERS T i € CIRCUIT

8 0 0 / 5 CT

I

$

T

400/5C.T.

FEEOER

FIG. 6.23 Selectim of grounding rerirtor bcired on cvrient-tronrforrner rotingr

SYSTEM GROUNDING

381

further analysis might be made. I n this case other system requirements may determine the rating of the resistor. The largest feeder circuit demands a t least, 400 amp, but 800 amp is required t o satisfy pilot-wire relaying on the tie circuit, if present. This would establish a minimum rating for both resistors. If this value also satisfies the requirements of adequate gcnerator differential protection of the larzer generator, the rating of both resistors may be 800 amp. REACTOR RATING

The reactance of a neutral grounding reactor should be chosen to limit theground-fault current and the current in the faulted phase t o the desired value. As previously stated, in order t o minimize transient overvoltages the ground-fault current must not he less than 25 per cent of three-phase fault current. This corresponds to a ratio of X o / X 1 equal to 10. For reartance grounding of generators the current in any winding must not exceed the three-phase fault current. This corresponds t o a ratio of X , / X , equal t o 1 . This establishes the criteria for maximum and minimum values of neutral reactance. It can be shown that under the condition of X , / X , equal t o 1 for any given generator on the system the current contribution in one phase winding of this generator t o a line-to-ground fault any place on the system (external t o the generator) cannot exceed the three-phase fault current of the machine. However, the neutral current may exceed this value, as shown later. The calculations concerning momentary duty (which is of interest for mechanical strength and transient overvoltages) are made using suhtransient values of machine reactance. The calculations concerning the thermal rating of apparatus are made using transient values of machine reactance. I n calculating the reactance of a neutral reactor, the positive-sequence reactance XI is taken t o equal the machine subtransient reactance. The calculation for determining the required reactance in the neutral t o limit the current in the machine winding to three-phase fault current becomes a very simple procedure, as illustrated below: lo =

3E

XI

+ x2 + xo + 3x8

I (three-phase)

=

E

x1

~

where I, = ground-fault current, amp (for a single generator this also equals the fault current in the machine winding) E' = line-to-neutral voltage, volts

SYSTEM GROUNDING

382

X , = positive-sequence reactance of generator, ohms per phase Xp = negative-sequence reactance of generator, ohms per phase Xo= zero-sequeiiw reactance of generator, ohms per phase 5,= reartance «f neutra1 reaitor, ohms If I,,

=

I (three-phase) and X ,

=

X,,

a

+ xo + 3x.v _- -xi 3E

ZX, 2x,

(6.5)

+ x o+ 3X." = 3x1 3xx=

x,-x.

The rated riirreiit of a iieutral groundiiig reactor is the thermal current rat,iiig. It is the rms iieiit,ral curreiit iii amperes which the reactor will rarry uiider staiidard conditioris for its rated time without exceeding staiidard ternperature lirnitatioiis. T h e rating establishes an rms current xhich is assumed to be eoiistant duriiig the rated time, for purposes of dcsigii, calculatioti, aiid test. I n service it is expected t h a t the current may he grcatcr thaii the rated value duriiig the initial cycles of the fault. The ixrreiit ratiiig of a iieutral groundiiig reactor is equal t o the rms symmctrical vurreiit (deulated i)y usiiig the t,ransient reactaiice t o represent syiii'hroiious ma(.hiiie positire-sequeiice reactaiice and the proper negative- aiid zero-sequenre reactanie values of the system. The mrreiit whirh will floiv throiigh a generator iieutral reactor is iiot iiidepeiidrnt of systrm coiistaiits, hut mil1 vary mith the number and siae

N0.I

N0.2

A

KVA

x'd =

x'd = 26% x'h = I3 % xo = 7 %

13%

x0 = 7%

& I000

KVA

-

480 VOLTS-60 CYCLE

-

TOTAL CONNECTEO SYNCHRONOUS MOTOR LOAD 1000 KVA

1) FIG. 6.24 volt ryrtem.

Xh = 31% Xd = 2 5 %

Reectance-grounded generotorr and rolidly grounded tranrformer on 480-

SYSTEM GROUNDING

383

of power sources. Thus the current rating of a neutral reactor is determined by the number and characteristics of system sources and whether they are grounded or ungrounded. The following example illustrates the calculation of ratings for generator neutral-grounding reactors to limit the fault current in generator windings t o three-phase fault current. Assume a system as show1 in Fig. 8.24. To determine the reactance of each grounding reactor from Eq. (6.6), XI is taken as the subtransient reactance X y of the related generator and Xuas the zero-sequence reactance of the related generator

X (ohms)

=

x(% __ kvz base kva

lo

(ohms per phase)

(6.7)

No. 1:

No. 2:

xu =

0'482 lo = 0.0129 ohm 1250 0.0239 - 0.0129 = o,oo37 ohm XN = 3 To calculate the current rating of each reactor, it is first necessary t o calculate the total ground-fault current le from Eq. (6.3). The positivesequence reactance of the system X , is calculated using the transient reactance X : of synchronous machines and the negative-sequence reactance of the system X 1 is calculated using the subtransieut reactance of synchronous machines.

No. 1:

SYSTEM GROUNDING

304

No. 2:

2'0

0'482 1250

3

lo = 0.0129 ohm

Transformer:

x,= x,= x o= 5'5

0'482 1000

lo = 0,0127 ohm

Connected load:

x2

=a

25

x

0.4s2 x 10 = 0.0575 ohm 1000

An equivalent circuit with values indicated is illustrated in Fig. 6.25. From Eq. (6.3)

3 (480/d%)

I'

0.00803

+ 0.0063 + 0.00705

*830 = 39,000 amp

0.0214

N0.I

XI

x2

XO

3% FIG. 6.25

N0.2

TRANSF

CONN. LO40

,0920

,0479

,0127

,0713

.00803

.0479

,0239

,0127

,0575

,0063

.02 5 8

,0129

,0127

:022I

.0111

.00705

Connection of positive-, negative-, and zero-sequence impedance networks for calculating ground-fault currents for system shown in Fig. 6.24.

SYSTEM GROUNDING

385

From inspection of the equivalent circuit it is evident that this total ground-fault current will divide through the paths to ground in inverse proportion to the impedance in the path.

No. 1:

I,,

= 0'00705 l o = ~

0.0479

0.147 X 39,000 = 5900

No. 2:

I,,

=

0.007051 o - 0.294 X 39,000 = 11,500 0.024

~~

To complete the picture, the ground-fault current a t the transformer will be 0.00705 ~ = 0.555I X 39,000 o 0.0127

IDT = ~

=

21,600

The reactor for generator No. 1 must be rated for a t least 5900 amp, and for No. 2 a t least 11,500 amp. This serves to indicate the method of determining the reactor current rating and proves that this rating is determined by system characteristics. The rating may be considerably greater than the three-phase short-circuit current of the related generator, as shown shove. GROUNDING TRANSFORMERS

The electrical specifications of a grounding transformer are as follows: 1. Voltage. The line-to-line voltage of the system. 2. Current. The maximum neutral curyent. I n a resistance-grounded system, this current is determined by the neutral resistor. I n a solidly grounded system, the current is determined by the grounding transformer impedance and the system impedance. 3 . Time. Usually designed to carry rated current for a short time, such as 10 see or GO sec. 4. Reactance. This quantity is a function of the initial symmetrical system three-phase short-circuit kva (use Fig. 6.26). The theory behind the determination of grounding transformer reactance is discussed in the following. When the grounding transformer is resistance grounded, the criteria for limiting transient overvoltages is either Xo/X, equal to or less than 10, or R o / X oequal to or greater than 2. It should he noted that Ro as it appears in this relationship is equal to 3 times the resistance of the neutral kesistor. When the grounding transformer is solidly grounded, the criterion for limiting transient overvoltages is X , / X , equal to or less than 10. The criterion for using groundedneutral-type lightning arresters is that X , / X , should be equal to or less

SYSTEM GROUNDING

386

than 3, and R o / X , should be equal to or less than 1 (see Chap. 5 ) . A summary of criteria for selecting neutral reactance is shown in Table 6.10. In a system having a grounding transformer, its reactance is the principal part of X, in the above criterion. Also, the positive-sequence reactance XI is equal to the reactance of the system to initial symmetrical rms three-phase short-circuit currents. Thus, the grounding-transformer reactance is a function of the initial symmetrical system three-phase short-circuit kva. On a system otherwise ungrounded, the groundingtransformer reactance required to provide any specified X o / X 1ratio is given by the following formula: 100

%

50

4

I

a W a

30 20-lf

10

\ 5

,

I

.5

,110

31 00

MAXIMUM SYSTEM SHORT CIRCUIT MVA CALCULATED USING SUBTRANSIENT REACTANCES OF ROTATING MACHINES FIG.6.26 Maximum allowable reactances of grounding transformers lo limit ground-fault current to 25 per cent of three-phase fault current.

SYSTEM GROUNDING

387

(-Y,,/.YJ x kv' x 1000 x,, = system symmetrical three-phase short-rircuit kva (masiniom ohms prr ]>haw)

((i.8)

Taking the specific case X d X , = 10, the desired grounding-transformer reactance may lie idriilated Ily thr fiirmula 10,000 X kv' xo, = system initial symmetrical three-phase sh;rt-rircuit

kva ~

(lj.!))

Curves shoiving typical values of groniidiiig-transformer rrartaiire for this condition are shown in Fig. 6.2(i. For example, it is desired t o apply a groundirig transformer i n thr folloning system: 2400-volt 50,000 itiitial s y m m e t r i d short-ririwit Iivn. The grounding transformer reactatice should be 10,000 x 2.4% - = 1.15 ohms per phase (mas) 50,000

XCT = -~

Grounding Transformer Grounded Solidly. The gri)rtiidiiig-traitsformer voltage, reactatwe, and time are determined as outlined al>ove. When grounded-neutr&type lightning arresters arc t o IIC applied, ttie grouiiding-transformer reactatice may tie determitied by 90, =

_ 3000 _ X _ kv? _ _~ system initial symmetrical three-phase short-rircuit k v s

(li.10)

When grounding transformers are solidly grounded, care should lie taketi that the reactanre is selerted at a value 1011- enough to provide sufficient fault current for tripping relays, Tuses, and circnit tjreakers. Grounding Transformer Resistance Grounded. I u this CBSP it is not necessary t o provide less groiiiiditig-tratisformer reactanre than t,hc values giveu in Fig. 6.26 siiice groiuidcd-tteotral-type lightning arrrsters are not applicable in resistailre-grounded systems. The values of reactatire given in Fig. (j.Z(i are equal t o ten times the system reavtatm t o threephase initial symmetrical rnis short-circuit current. This is cqnivalent to the ratio Xo/XI equal t o 10. Where the ratio of Ro/Xo is equal to or greater than 2 , the ratio of X , / X , may be greater than 10 without the dsirger of severe transient overvoltages. However, I?, must be low enough to permit sufficient current for good relaying. On systems of 600 volts or lo\\-er it is usually desirable to permit currents of magnitude considerahly greater than 25 per rent of initial symmetriral rms three-phase short-circnit current in order to assure positive tripping of protertive devices. 111 such systems the grouuding transformer is connected solidly t,o ground. The minimum current required for tripping is determitied by esamination of the system aud the ground-

SYSTEM GROUNDING

388

ing-transformer reactance selected t o permit at. least that much current t o flow in the event of a ground fault. TABLE 6.10

Summary of Criteria for Selection of Neutral Reoctavce

For limiting transient OVerYOltoge

XdX,

Ro/Xo

For application of grounded-neutral lightning arresterr

XdX,

_____

........................ ......... ......

R e o d m C e grounding.. Grounding transformer solidly grounded.. Grounding transformer resistance grounded'.

* Either criterion

10 or leis I 0 or leis

........ ........

10 or leis

2 or more

30, 11 .. 3 or less

is mtisfactory.

OTHER METHODS OF GROUNDING LINE GROUNDING

I n lorn-voltage systems (600 volts and below) which in the past have almost universally been connected in delta, it was sometimes advocated that one line be grounded, as illustrated in Fig. 6.27. This was done i i r order t o obtain some of the advantages of grounding at minimum expense. Because of its limitations and disadvantages it is strictly a compromise method and is rarcly encountered in modern industrial systems. Staudard load-center unit substations are now readily availahle with Y-connected secondaries at 480 and 600 volts i n all standard kva ratings. For

( A ) N E U T R A L GROUNDING

-( 8 ) CORNER- OF- THE- DELTA GROUNDING

FIG. 6.27

Two melhodr of grounding a low-voltage power system.

SYSTEM GROUNDHG

389

existing 480-volt delta systems dry-type zigzag grounding transformers provide a relatively inexpensive method of establishing a neutral. One of the outstariding disadvantages of corner-of-the-delta grounding is the necessity for positive identification of the grounded phase throughout the entire system. Instruments, meters, and overload relays should not be connected in the grounded phase. MID-PHASE GROUNDING

Where existing systems at 600 volts and below are supplied by three single-phase transformers with midtap available, it is possible to gain some of the advantages of neutral grounding by grounding the midtap of one phase. This method is illustratrd in Fig. 6.28.

FIG. 6.28

One phore of grounded ot the mid-point.

(I

delta system

THE INFLUENCE OF GROUNDING METHOD ON CONTROL-CIRCUIT SAFETY IN SYSTEMS 600 VOLTS AND BELOW Frequently the safety of a control rirruit is offered as a reason for a particular method of grounding. In all cases where motor-starter control eircnits are set up without control transformers, it becomes evident that there are problems with regard t o circuit arrangement which must be considered in order t o minimize operating difficulties and persolinel hazards. Accidental motor starting due t o faulted control circuits may be associated with ungrounded systems as well as most types of grounded systems. During such times as accidental motor starting may constitute a hazard, it should be standard practice to open the discomiiecting means whether the system is grounded or not, and regardless of the method of grounding. Analysis of the fault performance of motor control circuits from the standpoint of safety reveals that hazards may exist with all types of ungrounded and grouuded systems. Three methods quite commonly used are described. A similar analysis should be made with any other contemplated arrangement. Figure 6.29 shows a direct,ly connected control circuit on a n ungrounded system. A ground fault on any phase will remain unnoticed, and protective devices will not trip. Assume that a ground exists on either phase

390

SYSTEM G-ROUNDING

a L.

44 FIG, 6.29

Control circuit on ungrounded system without control power transformer.

-

.

A

L 6

L

B

0.L .

A

C

SYSTEM GROUNDING

391

2 or 3. A subsequent ground fault,at point R will impress full line-to-line voltage across the coil arid close thc contactor. A ground fault a t point C will pick up the coiitact,or, and the stop button will not stop the motor. Figure 6.30 shows a system wit,h solidly grounded neutral. A ground fault on any phase x i l l cause circuit tripping, and the fault mill be isolated. A ground fault a t point R or C will impress liiie-to-neutral voltage (58 per cent) amass thc coiitactor roil. This will usually not pirk up the contactor, but it will prohably burn out the (.oil. If the “start” button is closed during t,his period, full fault curreiit Xi-ill flow until interrupted by a protert,ire devirc. .Iground fault a t C ii-hilc t,hc motor is running ivill prevent stoppitig the motor from the stop button if the contactor fails t o drop out on 58 per rent voltage. Furthermore, the stop button may be called upon t o interrupt a fault, current in excess of its capability. Figure 6.31 shows one method of connecting a control circuit on a line grounded system. Here a ground fault on any phase except 1 d l cause circuit tripping. A ground fault a t R or C ivill not pick up the contactor and remain unnoticed. Closing the start hutton under this condition will cause full fault current t o flow through the start button. SPECIAL PROBLEMS AUTOTRANSFORMERS

Poiver autotransformers are quite frequently used in public-utility poiyer transmission and distrihutiori systems; however, their use in industrial power systems as a part of the power distribution system is relatively infrequent. Autotransformers are quite common, however, in control and utiliaatioii equipment. Systems using autot,rausformers may be subject t o dangerous fundamental frequency overvoltage during system faults or from high-frequency or steep wave-front transient overvoltages on the lines, originating from lightning or switching surges. Since the magnitude of these overvoltages depends in part upon the method of grounding the system and autotransformer, the nature of these overvoltages will be explained. Consider the case of a n ungrounded system using an autotransformer as represent,ed in Fig. 6.32. Lines a , b, c represent the loiv-voltage system normally operating at line-to-line voltage and points d , e , f represent the terminals of a step-up autotrarisformer normally operating a t line-to-line voltage E2. 111the event of a line-to-ground fault on the line connected t o terminal d, thc loiv-voltage phases b and c are elevated aboveground by the amount

392

SYSTEM GROUNDING

8:dc=

1

Ed,

= -dE,2

4 3

+ E,2 + h,,E*

(6.11)

For example, in the case of an autotransformer rated 13.8/34.4 kv operating ungrounded on an ungrounded system, a line-to-ground fault on one of the high-voltage lines will impress a voltage t o ground on two of the loii.-voltage lines of

0.58

m+ 34.4' + (13.8)(34.4) = 25 kv

Obviously, this is an undesirable situation and cannot be tolerated. Solid grounding of the autotransformer neutral eliminates this type of overvoltage. Another type of overvoltage called transient inversion can occur in a n autotransformer, as illustrated in Fig. 6.33. Steep wave-front transient overvoltages produced by lightning or slyitching surges coming in over lined and arriving a t point, .J are impressed across a portion of the aut,otransformer winding .IK, point K remaining a t it,s normal frequency value until C , can he rharged. The result is that the port,ion of the ivinding J K has impressed upon it practically the entire voltage disturhance. Since the port,ion of the winding K N is closely coupled t o J K , the voltage ivill be stepped up in K N by t,he turn rat,io of K.V t,o K J . Since the initial disturbance may he several times normal voltage, and since this may be stepped up tivo t o twenty times or more by inversion (depending upon the winding ratios), it, is evident that a serious overvoltage may be experienced. The hazard due t o transient inversion is greatest for autotransformers in which the high- t o lowvoltage ratio approaches unity. This type of overvoltage can be eliminated by solidly grounding the neutral. I n cases where this is not feasible, a lightning arrester or Thyrite* resistor connected between the neutral and ground can he used t o minimize this voltage. The presence of a tertiary delta on the autotransformer also tends t o minimize transient overvoltages of this hature. Another system autotransformer connection which is subject to both normal frequency inversion and transient inversion is operation with the supply system neutral grounded and the autotransformer neutral isolated, as shown in Fig. 6.34. A line-to-ground fault on the high-voltage line 2onnected t o terminal h forces the voltage of point h t o that of N , . This inverts the phase of winding hd by impressing voltage N , , from point h t o d . The hd portion of the winding induces in the d N , portion of the winding a voltage of corresponding phase and of a magnitude depending upon the turn ratio

* Registwed

trademark of General Electric Company.

SYSTEM GROUNDING

393

T O HIGH VC LTAG E UNGROUNDED SYSTEM

a

FIG. 6.32

Ground fault on ryrtem with autotransformer connecting ungrounded systems.

(1

;ca

TO UNGROUNDED HIGH VOLTAGE SYSTEM

I

FIG. 6.34 Autotransformer neutral isolated, supply-system neutral grounded.

of the two parts of the winding. This results in a shift of point Nz,as shown in Fig. 6.35. Note that phase voltages N P jand N 2 k are far above normal for the case

SYSTEM GROUNDING

394

illustrated, where the step-up ratio was 2 : l . If the step-up ratio had heen 1 . 1 : 1, that is, the autotransformer normally boosting the low voltage hy 10 per rent, the faulted phase would be overexcited by ten times normal, resulting in a much more severe shift of N 2 and overexcitation of the other phases. That is, the closer the autotransformer ratio is t o unity, the more severe is the overvoltage from this type of fault. Overvoltages from this cause can be prevented by solidly grounding the neutral of the autotransformer. The resultiirg voltage magriitudes are given by the following relations:

For example, in the case of a n autotransformer stepping up 10 per cent = E.V%h=

E Z = 1 . 1 Der unit 1.1 = 6.35 per unit d T ( l . 1 - 1)

1

= 7.32 per unit

S o t e also in this rase that only the high-voltage lines and connected apparatus are subject t o overvoltage. The lorn-voltage lines are not subjected to any ahnormal voltages in this case. The foregoiiig examples illustrate the nature of the overvoltages which ran be obt,ained with autotransformers. I n general, solidly grounding the neutral of t,he autotransformer is a satisfactory means of eliminating

h

t

EZ b

t El

1 -

FIG. 6.35 Vector diagram illustrating normal frequency inversion of clutotmnrformer.

SYSTEM GROUNDING

395

overvoltages. The disadvantage of solid neutral grounding is that thirdharmonic currents aiid telephone interfereiice may heromc excessive iii rertaiii cases. These harmonic problems ran usually he eliminated h?, use of a tertiary delta 011 the ant,otransformer. See referetire 5 for a more romplet,e discussion of this s n b j w t . SYSTEMS WITH PUBLIC-UTILITY SUPPLY

Some iiidustrial systems are directly roiinerted at t,heir operat,iiig voltage t o public-utility systems. The scheme of grounding the industrial system should be properly coordinated \\-it,ht.hat, for the utility system. If two systems are interconnected by means of a transformer bank, at least one \\-inding of t,he bank will normally be roiinevted i t 1 delta, and t,his delta-rotiiierted i~iiidingwill make grorindilig of earh systrm itidepetident of grounding of the other. GENERATOR-TRANSFORMER UNIT INSTALLATIONS

Figure (i.36 shorn an arrangemelit using a distrihut,ioti-t,ype t,raosformer, loading resistor, and relay in the gciierat,or neutral. This scheme may he provided x i t h a 5- or 10-miti ratiiig t o permit time for traiisferriiig load off the atrected mavhiiie before it is takeii out of service. The distribut.ioii transformer will usiially have a rating of 25 t o 50 h a , aiid the relay may be connerted t o operate on resistor current, or volt,age, depeiidiiig 011 the particular illstallation. This system is used bei,aose sometimes the rost of the resistor and distribntioii-type transformer is less than the vost of + high-voltage Ionrurreiit resistor roiiiiertrd dirwtlv hetween the neutral and ground. DISTRIBUTION TRAN 5 F O R M E R

RELAY

Y

-I-

Grounding the neutral of distribution transformer.

FIG. 6.36

(I

generator-tr~nrformerunit with resistance-loaded

396

SYSTEM GROUNDING

This results in a n effective high-resistance ground which, because of the limited system and the absence of switching devices, is satisfartory from the standpoint of transient overvoltages, and since no problem of relay coordination is involved, the relaying problem is simple. THREE-PHASE FOUR-WIRE SYSTEMS

I n these systems, single-phase loads are connerted between phase cotiductors and the neutral conductor. The neutral conductor is insulated over its entire length except where it is grounded at its source of supply. The neutrals of such systems should be grounded so solidly that during a ground fault the voltage between any phase conductor and ground does not appreciably exceed normal line-to-ground voltage; otherwise, abnormally high voltage t o ground mill be impressed on the unfaulted circuits. T o be adequately grounded, therefore, four-wire systems must use solid or reactance grounding with ground-fault currents approximately equal to three-phase fault currents. This is usually accomplished by direct connection of transformer bank neutrals t o ground. FAULT DUTY MAY BE INCREASED BY SOLID GROUNDING

Solid grounding of the service transformer neutral can be responsible for fault currents exceeding the three-phase values. This may i n some cases necessitate larger circuit breakers than would be dictated by threephase faults. Here is another advantage of limiting the ground-fault magnitude. A specific example (Fig. 6.37) incorporating a representative arrarigement will serve to explain what factors contribute t o a greater line-toground duty. The positive-, negative-, and zero-sequelire impedance diagrams for the system in Fig. 6.37 are shown in Fig. G.38. Three-phase fault-momentary duty :

I,,

=

x- (IJ(l.5) El

1.0 (1040)(1.5) 0.0832

=

~

18,750 amp asymmetrical

=

Interrupting duty:

I,,,

=

$ (Id(1.0)

=

~

0.0985

(1040)(1.0)

=

10,570 amp symmetrical

Line-to-ground fault-momentary duty : 3E,

I",,

=

Xi'

+ x,+ xo

(IB)(l.5)=

3

0.0832

+ 0.0832 + O.OF (1040)( 1.5) =

20,650 amp asymmetrical

397

SYSTEM GROUNDING

Interrupting duty: lint

=

X;

3 +3XEI, + Xu (zB)(l'o) = 0.0985 + 0.0832 + 0.06 (1040)(1.O) =

12,900 amp symmetrical

Ratio of line-to-ground t o three-phase fault duty: Momentary:

Interrupting :

The key t o this problem is the fact that three-phase fault current is coutrolled by the factor l / Z 1 while line-to-ground fault is controlled hy thc factor 3/(Z, Z, Zo). If Z,, Z?, and ZU\yere all equal, the two fault currents would be equal. Any system condition uhich acts to reduce

+ +

A

IND MOTORS 3000 K V A X- 25 PERCENT

X" = 6 2 . 5 PERCENT ON 7500 KVA

7 5 0 0 UVA X = 6 PERCENT

SYN MOTORS 2000 KVA X " / X ' = 2 0 / 2 5 PERCENT

X ' ? X ' = 7 5 / 9 3 . 7 5 PERCENT ON 7 M 0 K V A

FIG. 6.37 Typical system where ground-fault current may be greater than three-phore fault current.

SYSTEM GROUNDING

398

-

, 1 6 1040 ~ bMP BASE CURRENT(IB)= 4 7500

POSITIVE SEQUENCE X"

if

N

I /O.Il 9.09 110.75 1.33 1 / 0 . 6 2 5 = ~ 12.02

1/12.02= 0 . 0 8 3 2

0.625

X" EQUIVALENT

*

0.0832 N

IF

.eF

x ' ( FOR

INTERRUPTING DUTY )

I/O.lI 1/0.9375= '9.09

10.16 0.9375

I/I0.16=0.0985

X ' EQUIVALENT

El

N

-

F

0

NEGATIVE SEQUENCE X 2 (SAME AS POSITIVE SEQUENCE X" EXCEPT OMlTTlhO E l ) X,

EQUIVALENT N

2NLh/L.

IF

0.0832 ZERO SEQUENCE XO N

-2AN\r

IF

0.06

FIG. 6.38

Sequence impedoncer expressed in per-unit on 7500-kva 4160-volt threephase bare, for circuit shown in Fig. 6.37.

SYSTEM GROUNDING

399

Z, or Z, or any condition which tends to increase Z, mill make the line-toground fault current greater than the three-phase value. In the example the utility service line coutaiiis a fair amouut of reactance ii-hic,h becomes iiicluded i n the positive-sequence uct\vork but, riot in the zero-sequence netn-ork. Thus, referred to the 4I6O-volt bus Xo is smaller than X , and XI. Had the iiwomiug h i e shortkiri.uit dut,y been 500 mva instead of I50 mva (lower X , slid X2),there wor~ldbe scarcely auy difference hetween line-to-ground and three-phase fault-currrnt values. 111 the case of load-renter suhst,ations for inataiii,e, the highvoltage supply system reactanre \ d l he very small compared il-ith the transformer rcactauce; thus solid ueutral groundiug i n prarhically all cases results i i i 110 iircreased short-riruuit duty. In passing it is iuterest,ing t o note that grountliug any other ueutrals of 4160-volt equipment i n the esample would redwe %,, and (.&useline-togrouud fault rurrciit to be elevated.

EXAMPLES

OF PRACTICE

Example 1. Consider the syst,em of Fig. 6.39. ;\pplirat,ion proredure is as follows: I . All necessary data are giacit oii the diagram. 2. Select groundiug method. From Table (i.5, voiiditioti d , resist,anre grounding is suggested. 3. Select all three generator urutrals as grounding point,s, to assurr that the system \\-ill alij-ays he grounded. 4. Review system rclayiiig. a. Ground-fault curretit required for relaying is as follo\vs, from Table 6.9: Generators.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 300 amp Feeder irsiug (iOO;5 current tmisformers.. . . . . . . GOO amp Smaller freders.. . . . . . . . . . . . . . . . . . . . I.ess than (i00 amp

h. Ground-fault protection is available 011 every circuit esvcpt the bus, which can be protected by neutral back-up rdays. 5 . Select neutral circuit arrangement. Resistors should be rated at least GO0 amp. Three iudividual rcsistors would provide a maximum total ground-fault curretit of only 1800 amp. Therefore, individual resistors are selei.ted, as suggested uuder Seiitral (:iri,uit Arrangements. 6. Select neutral grnuilditlg equipment. Since gtwerator breakers may be used for back-up protwtion, a IO-sei. timc interval rating may hc used. Resistors for iudoor mounting are suggested. 7. The oiily additional devices needed arc three iirwtral current transformers and three neutral overr.urrent relays.

SYSTEM GROUNDING

51

PHASE OVERCURRENT RELAYS(TH0SE ON GENERATORS HAVE VOLTAGE RESTRAINT)

@) 07

@

GROUND OVERCURRENT RELAYS GENERATOR PHASE DIFFERENTIAL RELAYS(NEUTRAL CT'S NOT SHOWN) GROUND OVERCURRENT BACK-UP RELAYS FIG. 6.39

Example 2.

Circuit diagram tor Example 1.

Consider the system of Fig. 6.40. Procedure is as follows: 1. All necessary data are found on Fig. 6.40. 2. Select grounding method. From Table 6.5, condition R. 3. Select location of grounding point, as the neutral of the main bank. 4. Select neutral circuit arrangement, as a direct connection to earth (from Table 6.5). 5 . Review system relaying. We shall consider in turn each breaker used for system protection. a. Transformer primary breaker. This breaker will operate on trans-

SYSTEM GROUNDING

401

former secondary faults. For such faults the ground-return path will be so short that its resistance will be negligible, if the transformer case is properly bonded to the system neutral. From Table 6.4,a ground-fault current of about 17,300 amp may flow to a terminal fault. This is over ten times the circuit rating and hence is sufficient for operating phase overcurrent relays in the primary circuit. 21. T r a n s f o r m e r secondary breaker. This breaker is primarily for bus faults. Since the maximum ground-fault current is only about ten times the circuit rating, fast tripping may not be 1600 AMP .) I 480 V provided by the breaker, but the primary breaker will give satisfactory protection, as discussed for transformer faults. G. Feeder breakers. Since the FIG. 6.40 Circuit diagram for Example 2. current maximum . mound-fault , (17,300 amp) is a t least twenty times the rating of the largest feeder breaker, these provide adequate ground-fault protection. 6. No neutral grounding equipment is required. The transformer neutral must, of course, be available for grounding. OPERATING EXPERIENCE Case 1. The following is quoted from theexperience of an engineer of a large glass-manufacturing company as related to an AIEE group recently. “ A few years ago in one of our large plate glass plants two feeder grounds occurred on two different phases about 2000 feet apart in two departments. When the fireworks and excitement had subsided the two departments involved were shut down for several hours until repairs could be made. Within a few hours after this trouble happened a number of motors, seven altogether, were brought into the electric shop with burnt out or grounded windings. This, we believe, is an important effect of the system surges that occur during very bad fault conditions such as this., The loss of production and damage on that occasion amounted to several thousand dollars. “With continuous process operations the hunting of ground faults is very difficult, and two grounds on the same phase but on two different feeders are exceedingly difficult to trace. This is because all the feeders must be opened a t once and closed one at a time to find the trouble. Our

402

SYSTEM GROUNDING

experience is that the first ground remains on the system because we cannot open the feeder breakers to hunt it. The result is that the system operates with two phases a t line-to-line voltage to ground, and the operating electrician hopes that no other grounds occur before he has an opportunity t o find the first one. “It was because of our experiences, such as I have mentioned, and the need in our operations for the highest possible service continuity, that we began to seriously consider the use of grounded neutral low voltage distribution systems. “The cost of a grounded neutral low voltage system is slightly higher than an ungrounded system. The additional transformer neutral bushing and connections, the neutral bus and wiring are items that add t o the cost. These are first costs that do not add more than one per cent t o the total cost of a unit substation. “Two of our plate glass plants are now operating 100 per cent with 600 volt grounded neutral systems, and t.wo other works are about 50 per cent cut over. Two window glass plants, operating a t 460 volts, are completely changed over t o grounded neutral and a third is in the process of being changed. Several new plants, one paint plant and two fabricating plants were built with 460 volt grounded neutral systems. Our total transformer capacity operating a t 600 volts or 460 volts grounded neutral is now 40,000 kva, consisting of 30 units. “Our experience with these systems has been very satisfactory. There is no question that the service reliability has greatly improved. A majority of the faults occur on branch feeders and are cleared by the local branch protection devices such as fuses. Troubles are localized and promptly repaired. As the electricians become used t o the new systems they are more enthusiastic and quickly learn, for instance, that a single blown fuse probably indicates a ground. None of them has expressed any desire t o return to nongrounded systems.” Case 2. An engineer from a large steel company reported as follows on experiences with a grounded-neutral 6900-volt system which was placed in operation in 1947: “The operating record of the system since the grounded neutral was installed is most gratifying. The ground faults experienced show a marked reduction in number and severity. For instance, during the year 1944, the number of ground indications recorded totaled 34. Of these 34 indications, 19 resulted in equipment failures such as grounded motor coils or flashed-over bushings. During the year 1951, there were two ground relay operations resulting in one equipment failure, and the first fifty weeks of 1952 show a similar record. Particular attention has been paid t o the severity of the damage caused by these ground faults. I n each instance i t appears that the relaying has been fast enough t o clear the fault before any destructive burning resulted.”

SYSTEM GROUNDING

403

REFERENCES 1. Concordia., C.., and H. A. Peterson. Arcine Faults in Power Svstems. Trans. A I E E . vol. 60, pp. 340-346, 1941. 2. Concordia, C., and W. F. Skeats, Effect of Restriking on Recowry Voltage, Trans. AIEE, vol. 58, 1939. I

3. AIEE Standards No. 32, “Neutral Grounding Devices.” 4. National Electrical Code. 5. Blume. L. F.. “Tranaformer Eneineerins.” John Wilry & Sons. Inc., 1951. 6. Allen, J. E., and S. K. W a l d o r f , k n g Ground Tests i n a Normally Ungroundcd 13-Kv, 3-Phase Bus, Tmns. AIEE, 1946, p. 298. 7. Shott. H. S., and H. A. Peterson, Criteria for Neutral Stability of Y-Grounded Primary, Broken Delta Secondary Transformer Circuits, Tiann. AIEE, 1941. p. 997. 8. Clarke, Edith, “Circuit Analysis of A-C Power Systems (Symmetrical Components),” John Wiley &Eons, Inc., New York, vol. I, 1943, vol. 11, 1950. 9. Application Guide for the Grounding of Synehronoua Generator Systems, AIEE Committee Report, Power Apparatus and Systems KO. 6, June, 1Y53, pp. 517-526. 10. Quinn, R. F., Should the High Voltage Neutral of a Wye-delta. Stepdown Transformer Bank Be Grounded? Gen. Elm. Rev., June, 1945.

chapter 7

by L. J. Carpenter, Shelby C. Cooke, Jr.,

R. H. Kaufmann, and David Stoetzel

Equipment Grounding STATIONARY EQUIPMENT, BUILDINGS, AND STRUCTURES OBJECTIVES OF EQUIPMENT GROUNDING

Equipment grounding consists of the connecting to ground of noncurrent-carrying metal parts of the wiring system or apparatus connected to the system. This includes all metal conduits, metal raceway, metal armor of cables, outlet boxes, cabinets, switch boxes, motor frames, transformer cases, switchgear enclosures, metal enclosures of motor controllers and other frames, and metallic enclosing cases of all electric equipment and electrically operated equipment. One objective of equipment grounding is to limit the potential between non-current-carrying parts of the plant and between these parts and earth to a safe value under all conditions of normal and abnormal system operations. To accomplish this objective, a plant grounding system is required. The purpose of this is t o seek to achieve a uniform potential in all parts of the structure and apparatus, as well as to provide that operators and attendants shall also be a t the same potential a t all times. By achieving more nearly a uniform potential throughout the grounding system, the chances of large differences of potential within reasonable reaching distance of a person, great enough to shock or injure an attendant when short circuit or other abnormal occurrences take place, are reduced. A grounding system is very likely called upon to function very infrequently, and inadequacy may become evident only a t that time. It is like the gun that nobody thought was loaded until someone pulled the trigger. When a ground fault occurs on an electric system, lives may depend on an adequate equipment-grounding installation. 404

EQUIPMENT GROUNDING

405

A second objective of equipment grounding is a low-impedance return path for ground-fault current. The hazard to personnel exists a t the time a ground fault occurs. Forcing the current to flow through a highimpedance grounding connection may create a dangerous potential difference. Also, high impedance a t joints and connections or insufficient cross section in grounding circuits may cause arcing and heating of sufficient magnitude to ignite nearby combustible material or explosive gases.

IMPORTANCE OF EQUIPMENT GROUNDING

The published data of the Division of Industrial Safety, Department of Industrial Relations, State of California, states that in the year 1952 there were 909 recorded electrical work injuries, of which 40 were fatal. For comparison, similar figures for several previous years are listed in the accompanying table. Iniuries Year

1946 1947 1948 1949 1950

Rdio Total

Fatal

305 572 755 693 690

28 57 48 42 33

10.9 10.1 15.7 16.5 20.9

I I I ;;:;

:;:: :;

Of the 909 recorded injuries, 153 could be related directly to contact with frame case or non-current-carrying metal parts. It was found that in these 153 recorded injuries either no grounding or inadequate grounding could have been responsible for the injury. Typical injury descriptions are as follows. “Refrigerator Repairman. Electric drill shorted out-severe shock; employee knocked out for about fifteen minutes.” “Carpenter. Operating portable electric hand-saw on wet groundreceived shock and dropped saw. Laceration, severe, dorsal surface a t base of distal phalanx, third finger left hand.” “Cabinetmaker. Ground wire broken off in drill. Hot wire grounded to frame, took hold of another grounded drill-unconscious about one minute.”

406

EQUIPMENT GROUNDING

Inasmuch as adequate equipment grounding tends to keep the potei tial difference between equipment frame and ground within safe limits, it can he safely said that these 153 accidents, or approximately 17 per cent of the total, were attributable to inadequate equipment grounding. The National Fire Protection Association consistently reports that about 10 per cent of all fires, representing about 10 per cent of losses from fires, are specifically attributed to “Electrical, fixed services, fires due to misuse or faulty wiring and equipment.” They also report that another 10 per cent, representing about 30 per cent of losses from fires, is of unknown origin. I t is inevitable that many of these fires were caused by inadequate equipment grounding and insufficient attention given to return paths for ground-fault currents. The increased use of system-neutral grounding has focused attention on the necessity for good equipment grounding systems to obtain lowimpedance return paths for ground-fault current. For safety to personnel, it is generally recognized that equipment grounding is required but is often provided as an afterthought and, consequently, may or may not he adequate-for the purpose intended. With a little careful consideration, it becomes apparent that a well-planned equipment grounding system must be provided whether the system neutral is grounded or not. Ungrounded neutral systems often operate for extended periods with a single-phase faulted to ground. During such periods, a contact between another phase conductor and a metallic enclosure raises the enclosure to full-line potential aboveground. Failure to provide a suitable connection between enclosure and ground presents a serious hazard to personnel. Also, the flow of fault current through a high-impedance connection during a double line-to-ground fault may create differences of potential of dangerous proportions. COMPONENTS OF AN EQUIPMENT GROUNDING SYSTEM

Definitions. For the purpose of further explanation of the grounding system the following definitions are established (see Figs. 7.1 and 7.2). Grounding electrode is a conductor embedded in the earth, used for maintaining ground potential on conductors connected to it and for dissipating into the earth currents conducted to it. Ground bus is a protective ground network used to establish a uniform potential in and about the structure. I t is tied solidly to the grounding electrodes. h typical ground network is illustrated in Fig. 7.3. Grounding conductor is a conductor used to connect equipment frames or wiring-system enclosures to the ground bus. The distinction between a neutral conductor (white) and a grounding conductor (green) is illustrated in Fig. 7.4.

FIG. 7.1

Typical grounding system

for on outdoor substation.

c Y 0

EQUIPMENT GROUNDING

408

FIG. 7.2

pical grounding system for a building and heavy electric opparatvr in the

building.

GROUNDING

ELECTRODES

CONNECTlON TO WATER PlPiNG

JOINT

CONNECTIONS TO BUlLDlNG STEEL E A C H SlDE OF EXPANSlON .IOlNT

FIG. 7.3

Typical ground bur.

410

EQUIPMENT GROUNDING

Types of Grounding Electrodes. A continuous underground waterniping system provides a very satisfactory grounding electrode (Fig. 7.5). Consideration should be given t o the size of pipe arid the extent of the system if this is t o be the sole means of connection t o earth. Table 7.1 tahulates the size of wat,er pipe in terms of equivalent grounding conductor or bus. Artiiicial grounding electrodes should also be used. Such electrodes may be rods, pipes, plates, or conductors embedded in the earth. They should be of noncorrosive metal, such as copper or copper-hearing steel. They are embedded in the earth by bring driven or by burial. The Ground Bus. The importance of a continuous metallic circuit of low FIG. 7.5 ~ ~ method ~ of grounding i ~ i~npedarice ~ l in the returo path for to o large w o t e i pipe. ground-fault currents is illustrated in Fig. 7.G. Fig. 7.Fa shows a 120/240-volt single-phase system with transformer neutral connccted to ground through a grounding electrode ivhich measures 10 ohms resistauoc to earth. The conduit is connected t o earth through a separate grounding electrode which measures 20 ohms t o earth. h fault oc(?urs lictween conductor B and the conduit.

Ground-fault current =

20

120 in

+

=

4 amp

Voltage drop from conduit to earth equals 4 X 20, or 80 volts. Figure 7.iib shows the same system with both transformer neutral and conduit connected to a common ground network which is connected t o earth through a single grounding electrode which measures 25 ohms resistance. A fault occurs between coilduotor B and t h e conduit. A high fault current will flow through the low-resistance ground-return path causing fault interrupting devices t o operate. Little, if any, current flows through the 25-ohm resistance, and therefore the conduit will remain very close t o earth potential. It should not be inferred from the above t h a t 80 volts potential is necessarily fatal. T h e example used simply ,illustrates t h a t appreciable resistance in the ground-return path results in a difference in potential during a ground fault which may be great enough t o be fatal t o a person

EQUIPMENT GROUNDING

41 I

stepping or reaching from one point t o another. A continuous path of low impedance i5 effeoted by means of a properly designed ground hus. The size of the ground bus is determined by the magnitude of current and the time of flow, based on the maximum allowable temperature rise. For bolted joints the temperature rise should be limited t o 250 C , and for brazed joints t o 450 C:. While ground buses and connections must be adequately braced t o withstand the mechanical stresses due t o the initial asymmetrical line-to-ground fault current, the heating effect of such current can generally be disregarded because of its short duration. The following equations may be used in determining the size of ground bus when copper is used for conductors. For bolted joiuts with initial temperature of 26 C and temperature rise of 250 C,

For brazed joints with initial temperature of 26 C and temperature rise of 450 C,

1

0 VOLTS

Illustrating the importance of a continuous metallic ground-return path of low imc-edance.

FIG. 7.6

412

EQUIPMENT GROUNDING

where A = cross section, circular mils I = ground fault current,, amp S = time of floy, see For ungrounded, impedance-grounded, arid solidly grounded systems it is usually easy t o determine the magnitude of fault current t h a t could flow in the ground hus. For uiigrouiided arid impedaiice-grourided systems this will approximately equal line-t,o-line fault current,, a i d for solidly grounded systems i t will approximately equal three-phase fault current. For the average grounded or ungrounded power distribution system wit,h adequate (auk protective devices, a time of flow of current of 10 sec is conservative atid may be used in thc above calculation. Aside from the theoretical considerations t,here are practical limits which may finally determine the maximum or minimum size of ground bus. For mcclianical st,rength the ground bus should not be smaller t h a n S o . 2/0-.4~vgconductor. It is not usually necessary t o exceed 500 MCM or equivalent for large generating stations and substations, or S o . 4/0 Awg for small stations and industrial plants. However, it may he desirable t o exceed these values where exceptional precaihon is required or where extremely high ground-fault currents are expected. A ground bus of adequate size for the installation should be run completely around the periphery of the building (see Fig. 7 . 3 ) . Grounding conductor material should he soft-drawn or medium-hard-draxn copper wire or copper bar. For steel-frame buildings the ground bus should be conneoted t o each outside building column (Fig. 7.7). In large buildings a network should he provided t o include internal buildiiig columns. The around bus should he connected t o electrodes a t int,ervals of 200 f t or less. If the building consists of more than one floor, each floor should have its own ground bus, these floor ground hnses in turn should be connected by a number of condiictors t o the main ground bus 011 the first floor. In buildings having no steel frameivork, a grouiid network equiualent t o the above should be provided. Where no steel framework is available, all grounding conductors must he taken directly t o the ground bus. Better aocessihility is obtained if an exposed bus is provided in the upper structure, and often it is inore economical t o install it in this FIG. 7.7 connection of ground bur to building column. manner.

EQUIPMENT GROUNDING

413

T h e gronnd hos may be installed in the form of a loopoiilsidetlie boundaries of the buildings, biiried in the backfill for the columii footings arid foundation wall. The loop ground bus should be installed a minimum of 18 in. orit,side &hebuildiiig wall and 18 in. below the finished grade. Where exposed to mechanical injury the conductor should be suitably prot.ectcd by pipes or other Siibstaritial guards. If guards are iron pipe or other magnetic matcrial, the conduct,or should be electrically coiiiiected t o both erirls of the guard t,o prevent, inductive choke ef- S,G,7,8 Codwe,d-type buried groundfcct. ( h i d a c t o r s laid iiiidergroii~id ing" connection, should, urilcss other\\ protected, he laid slack t o prevent their being readily broken. 811 hiiritd grounding corinectioiis should be made by brazing or (:nd\\-eld-t,ype joiiit (Fig. 7.8). All ot,her grouridirig comicctions may be ma& by brazing, Cad\reld, or with approved pressure terminals. Steel-to-copper coririectioiis should he made abovegrouiid wherever possible. The ground bus should he coiiiie , at least at two points. to a continuous uiidcrgroiirid vater-piping ciri or t o suitable grouiiding e k trodes (Fig. 7.9).

FIG. 7.9

Cadweld connection to water pipe.

414

EQUIPMENT GROUNDING

Grounding Conductors. Grounding conductors should he large enough to carry the ground fault current safely. Where the grounding conductor is insulated (green wire), it should be the same size as phase conductors. Where the conductor is bare, the temperature rise is limited to the same as the ground bus except that where exposed or adjacent to inflammable materials the total temperature is limited to 100 C. Although neutral conductors may be grounded a t the source, they should not be used for equipment grounding. Separate conductors should be used for neutral conductors and for equipment grounding. Equipment-grounding conductors must be identified with green color code. This is distinguished from neutral conductors which should be white color code. An illustration of this requirement is shown in Fig. 7.4. limiting Values of Resistance from Ground Bus to Earth. In large stations the resistance of the ground bus to earth should not exceed 1 ohm and should be made as much lower as can be realized economically. In small substations a resistance from ground bus to earth of higher values than that in large stations is generally permissible because the ground-fault currents are relatively smaller and they are in general only accessible to qualified personnel. Preferably, however, it should not exceed 5 ohms and should be as much lower as can be realized economically. For residential customers it is common practice to ground one side or the neutral of electric services on the premises. In cities this connection is ordinarily made to a water pipe, which usually provides a very lowresistance grounding connection. I n rural locations water systems may not be accessible, requiring that a driven pipe or rod must be installed for the grounding connection. The National Electrical Code requires that such grounding connections shall have a resistance not to exceed 25 ohms. Methods of measuring resistance 60 ground are discussed later.

TABLE 7.1

Minimum Size Water-pipe Electrodes

Size of Grounding Conductor

Sire of water Pipe,

or Bus, Awg

Inches I.P.S.

8 6

$5

A

I 1 >i 1% 2

2 1 /o 2/0 4lQ

$/a

2%

EFFECT OF IMPEDANCE IN EQUIPMENT-GROUNDING CIRCUITS

Of course, no one would contest the fact that reactance as well as resistance influences the return path taken by ground-fault currents. Recent tests, however, indicate that reactance his a much more marked effect than has been previously appreciated. This is particularly true

EQUIPMENT GROUNDING

415

where circuit conductors are enclosed in magnetic materials such as steel conduit or husways. I t was found that, when the enclosing conduit and a return coiiductor (of about equal resistance) outside the conduit were paralleled, the current divided approximately 20 parts in t.he conduit to 1 in the conductor at low current,s and 10 parts in the conduit t o 1 in the conductor at high rurreiits. About the same ratio also held true when the rurrent was allowed to divide hetween the conduit and a very low-resistauce steel-frame huildiirg structure. When the path by may of the eonduit was opened, a substantial voltage appeared between conduit and ground. With this same ret,rirrr conductor iriside the conduit., the current tended to divide about equally hetween conductor and corrduit. This leads to the conr~lusionthat the 60-cycle reactance of any groundreturn circuit remote from the outgoirig circuit conductor will likely be high compared ivith its resistance and limit the magnitude of groundreturn current which i t will carry. It may also he concluded that t,he conduit or enrlosing metallic structure will tend to carry an appreciable port,ion of thc fault rurrent and that failure to provide a continuous path will result i n arcing and heating, which may cause fires in combustible materials whirh may he rrear. This may account for the many fires that are reported tiy insurance statistics as caused hy faulty electrical circuits or of unkiioirii origin. POWER PLANT AND DISTRIBUTION EQUIPMENT

The frames of stationary or permanently located rotatirig electric equipment and the frames and enclosures of static equipment such as transformer t,anks and associated equipment permanently lorated should he grounded by dirert roniieotion t o the building grouiid b u s through a grouudiug condurtor equal in size t o the largest conductor in the line connected t o the equipment hut not less than No. 6 Amg nor greater than S o . 4/0 Awg (Figs. 7.10, 7.11, and 7.12). Driven ground electrodes should be employed a t earh outdoor substation. To provide a vorrvenient method of grounding switchgear, a ground hus should be provided as part of the equipment for structures or panels contaiiring such primary apparatus as current transformers, potential transformers, pon-er circuit breakers, and disconnecting switches and such other apparatus as relays, instruments, and meters which require grounding. Each of these metal structures, metal panels, or metal supports should lie individually connected to the switchgear ground hns, which must not he smaller in current-carrying capacity than 25 per cent of the highest rontinuous-current rating of any piece of primary apparatus t o which it is connected. Usually a 2- by >&in. bar is used.

416

EQUIPMENT GROUNDING

FIG. 7.10 Grounding connection-utdoor circuit breaker

FIG. 7.12

FIG. 7.1 1

Grounding connection-tronr-

former tank.

Grounding connection-motor

frame to building column

EQUIPMENT GROUNDING

417

This switchgear ground hus should, in turn, be connected t o the common station ground bus by suitable conductors having a current-carrying capacity equal t o t h a t of the switchgear ground bus (Pig. 7-13). I n many cases ( f o r example, i n metal-clad switchgear or other metal strue,ture) apparatus may be considered adequately grounded through their mountiiig on the structure. In some suhstatioii installations, where all connect,ioris are underground and there is no possibility of energizing t,he enclosing FIG. 7.13 Switchgeor ground bur. fence by falling overhead wires, it i s desirable t o keep ly reduced as it is much more subject t,o failure because of forrign objects coming in contact with the conductors. Because of the lack of such inst,allatioris in present-day plants, no further reference will he made t o this type of circuit. The metal-enclosed bus is gaining wide favor in industrial plant installations because of its safety arid reliability as contrasted t o that offered by the open-wire system. There are two principal types of bus which are used. These are the plug-in type and the feeder type. The feeder h u s is often called a low-reactance hus hecause of the interleaved phasc-coriduct,or arrangement (see Fig. 12.9). This type of bus is gerierally used for major circuits carrying large currents to specific pieces of equipment. For example, this bus is often used t o feed large welding equipment. Also, i t is sometimes used to supply power to the plug-in type oE bus.

FIG. 12.9

Cross-sectional view of low-reactance feeder bur rhowing interleoved bur-bar

arrangement.

698

SECONDARY DISTRIBUTION SYSTEMS

The plug-in bus coiisists of the necessary conductors enclosed in a suitable housing designed to permit the insertion of tap-off circuit breakers or fused-switch types of plug connectors. The tap-off points are located at close intervals along the bus run to provide a high degree of flexibility. This hits offers a safe reliable system for power distribution but still retains the advarit,age of the open-wire installation with respect t o the easr of tapping off for branch circuits. Figure 12.10 shows a typical plug-in hus iustallation. Cable in conduit i s wed widely for supplyiug all types of loads in the secoudary dist,ribution area. I n past years braided-type cables with either rubber or varnished-cambric insulation have been used for such inst,allations. I n recent years, however, the development of synthetic insulations has rrsulted in cables with improved characteristics. For example, synthetic-ruhher cahle with a neoprene protective jacket is commonly used today in conduit systems. Figure 12. I1 shows a typical cable of this type. The neoprene jacket provides a very tough and

FIG. 12.10 Typicol installation of plug-in-type bur with flexible bur drop cables.

700

SECONDARY DISTRIBUTION SYSTEMS

FIG. 12.13

Interlocked-ormor power cable instolled in overhead rocks

of the sa,me ' s i x in steel conduit. In addit,ion the cable installation is extremely flexible, being out in the open for visual inspect,iori and easy repair, and can be relocated in case of system changes. A typical iiiterlocked-armor-cable installation in overhead racks is shown in Fig. 12.13. The prececliiig types of secondary-feeder cirmits will bc discussed in the next section with regard t o their application t o the design of the secondary feeder and branch circuits.

SECONDARY-FEEDER AND BRANCH-CIRCUIT DESIGN

The secondary-feeder and hranch-circuit design of a n industrial plant vary t o a cert,ain degree, depending upon the type of plant involved. Basically, there are two types of maehirie loads in industrial plants. There are machines wiiich are not pemanent,ly located and machiiies which are permariciitly located, Examples of plants with the first type of load are mctal-fabricating plants, automohile plant,s, etc., which consist principally of machine tools which are very suhjcct t o changes in location. The second type of load is found in a process type of plant in which all parts are more or less permanently installed. Typical examples

SECONDARY DISTRIBUTION SYSTEMS

701

of this type of plant would be steel-mill motor rooms, paper mills, oil refineries, etc. There are many industrial plants which cannot be classified either as the metal-fabricating type or the process type but are sonic conihiiiation of the two. The power distribution syst,cm for any plant, consisting of such a combination can easily be designed if the fundamentals for t,he two extreme types are understood. Metal-fabricating Type of Plant. The plug-in h u m a y type of installation offers the most flexible system possible for distributing power in the areas where the machine tools are Iocated. As production models and designs change, these machine tools are often relocated, and it is essential that the power system supplying them be sufficiently flexible t o permit these changes with minimum disturbance. The plug-in busway is generally installed overhead, running completcly across the area at not more than 40- t o 50-ft intervals. This means that the maximum horizontal run t o reach a position over any machine tool will not exceed 25 ft. The power feed dropping down t o the machine tool from' the bus can he either by means of a flexible bus drop cable or by means of wire in rigid conduit. The flexible types of bus drop ofkr a material advaritage in flexibility in utilization-equipment movements. The most economical may t o supply power to the plug-in bus in most plaqts is hy means of interlocked-armor cahlc run overhead in racks or trays. This cable begins a t the secondary-feeder breaker at the loadcenter unit substation and terminates in the cable entrance box on the run of plug-in busway. Oftentimes cranes prohibit the use of an overhead bus system because the busway and the bus drops would he in the may of the crane hook. The flexibility of a plug-in bus installation can still he ohtaiiicd in these cases by installing the bus in a different manner. The bus can he run horizontally along each line of columns with the drop cables running down the columns and, if necessary, under the floor t o the machines. Another method would be t o run the cable feeder along the row of columns with short lengths of plug-in busnay mounted vertirally on certain columns, as shown in Fig. 12.14. When deciding on a plug-in bus arrangement, it should be rcmemhered that there is a very definite advantage to having the cable drop from overhead. I n this manner the cables are free from the oil and grease on the floor, and it is not necessary t o dig up the floor t o install or move any of the machine tools. Process Type of Plant. The other extreme in plant t,ype is that where the utilization equipment consists of machinery which will be permanent,ly located. For loads such as this, the type of syst,cmsmentioned under t,he metal-fabricating plant can also be used even though the machines never

702

SECONDARY DISTRIBUTION SYSTEMS

R U G IN BUSWP;

FIG. 12.14

Suggested arrangement of plug-in burway where overhead crones ore

prerent.

need to be moved. The use of plug-in busway fed by interlocked-armor cable is a very economical method of supplying a number of relatively small loads from a load-center unit substation. In many cases, however, the loads are rather large spot loads and can best be served by an individual feeder. This feeder can either consist of cable or of a metalenclosed feeder type of busway. The selection of the type of feeder would be solely one of economics. In many plants there are spot loads of relatively large size, which require some voltage other than 480 volts. For example, there are many ovens rated 240 volts used in industrial plants. For such loads i t is suggested that they be connected to the 480-volt system through individual transformers stepping the voltage down to the 240-volt level required. Depending upon the size of these loads, they might either be connected into the overhead plug-in busway or perhaps would warrant their own individual feed from the load-center unit substation. It is common in a process type of plant to find areas with large numbers of motors permanently located. I n such cases it is common to install grouped motor-control equipment with a number of motor starters in a common housing. Such motor-control centers are generally supplied by means of interlocked-armor cable from their own load-center unit substation secondary feeder breakers. From the motor-control center to the

703

SECONDARY DISTRIBUTION SYSTEMS

individual motors, cable circuits are commonly used either in conduit or in interlocked armor. It is apparent from the foregoing discussion that there are many variations possible to supply the common types of loads in the secondary distribution system. Figure 12.15 illustrates severa1 of the methods which have been discussed. It is not intended that what has been said should

I

a

L

1-r480y’277 3 \1

CAELE

INTERLOCKED ARMOR CABLE

PLUG-IN BUSWAY

4 CONDUCTOR INTERLOCKED ‘ARMOR

FOR OVEN

:&

LICHTING FEEDER

CONTROL

LARGE OVEN DR MACHINE TO WAD FIG. 12.15 Severai methodr of supplying powei to loodr in the recondory dirtribution ryrtem.

704

SECONDARY DISTRIBUTION SYSTEMS

be a complete discussion of this suhject, but merely t o point out how modern metal-enclosed bus or interlocked-armor rahle can be used in various manners t o feed the loads safely and reliably in a modern industrial plant. Lighting Feeder Circuits. The foregoing discussion of secondary distribution feeder circuits has not included any mention of the method of supplying the lighting circuits. When high-voltage fluerescent lighting is used as discussed in Chap. 10, a four-conductor interlocked-armor feeder cable is commonly used to feed a numher of the small combination motor sbarters which control the lighting circuits. T h e lighting circuits are switched in balanced three-phase banks, which is more advantageous from a system standpoint. When small dry-t,ype transformers are used t o step down t o 120 volts for the lighting load, thrcc-conductor cable should he used t o feed several such transformers. The transformer secondary will then connect t o a lighting panel hoard. Panel hoards, which normally have a n interrupting rating of 5000 amp, are suitable for this service since the small step-down transformer sufficiently limits the shortcircuit current. The lighting load either can be on separate feeder circuits or can be taken from the power feeder circuits. Most engineers look more favorably upon having the lighting circuit completely separate so that it is not suhjert t o the outages and possible voltage fluctuations of the load circuits. Also, when using the fourth conductor in the plug-in busway as a ground conductor, it becomes impractical. t o snpply line-to-neutral lighting loads from the busway. Circuit Arrangements. The circuits associated with the secondary distribution system are practically all radial circuits. A radial circuit i.s one in which there is only one path through which the power can flow t o reach the utilization equipment. It is the simplest form of circuit and because of its simplicity is commonly used in these systems. Such an arrangement is normally adequate for this type of service and offers a sufficient degree of reliahility. Since there is a great deal of exposure t o oil, grease, dirt, mechanical damage, etc., in the jvorkirig areas, the secondary circuits undoubtedly experience more fault,st,han do other parts of the electrical distribution system. However, the amount of load affected by the outage of a secondary feeder is relatively small, and therefore additional paths for power t o flow t o the utilizat,ion equipment cannot normally he justified. Therefore, it is extremely rare t o find anything except radial circuits in the secondary distribution area. Although practically all circuits are radial, there occasionally arises a condit,ion where extra reliability is required because of the nature of certain process equipment. T o build reliability into the distrihution system for such loads, it is necessary t o consider not only the secondary distrihu-

SECONDARY DISTRIBUTION SYSTEMS

705

tion system but the over-all poirer dist,ributioIi arrangement from the source of power down t,o the load. Chapters 11 and 13 discuss various circuit arrangements for the load-center system and the primary system in order to assure higher degrees of reliability for such essential processes. When an adequate primary and load-center system have been designed t o ensure addkional reliability, then steps can be t,aken in t,he secondary distribution system to carry this extra reliability down t o the utilization equipment.

Chapter 13

by Norman L. Hadley

Primary Distribution Systems The primary distribution system of an industrial plant is generally the higher voltage portion of the system, starting with the purchased-power service and including generators, switching equipment, circuits, and all transformers with secondary voltages higher than 600. Figure 11.1 is a simplified one-line diagram that illustrates the dividing line between the primary distribution and the other portions of the system described in Chaps. 11 and 12. The material in this chapter is presented under several headings relating t o different parts of the primary distribution system. While this is a convenient method of presentation, i t may fail to emphasize the importance of planning the parts together. Coordinated planning guided by over-all system characteristics is the only way desired objectives can be achieved. Such system characteristics as cost, safety, reliability, flexibility, and simplicity should be viewed together because they will be interrelated in varying ways. System arrangements tending to favor a particular desirable characteristic will most often tend to produce compromises in one or more other desirabl: characteristics. System cost is the characteristic that receives the most planning attention. I n comparing alternative layouts it is very helpful t o hear in mind that lower cost is by no means synonymous with better value. On the other hand, a determination of values relies heavily on judgment because it is very difficult to appraise recognized differences in such characteristics as safety, reliability, and flexibility. Service reliability is considered t o be improved when the arrangements are modified in ways that promise to reduce outage time during maintenance operations or in the event of trouble. The general attack is to provide more than one power channel around system components that need maintenance or might fail. Increased investment for such provisions may be money wasted unless the system is well planned in some 706

PRIMARY DISTRIBUTION SYSTEMS

707

other respects. The primary requirements for good service reliability are that good-quality adequate equipment will be selected, that it will be rarefully installed, and that it will be well maintained. Industrial systems should be planned with good Hexibility as a characteristic so that possible needed changes can be made readily. It is particularly important t o he able t o expand the power system, because industrial plants historically develop more and more load, even without Hoar-space expansion. There are numerous ways in which growth possibilities may be restricted. A common mistake has been the application of circuit breakers with little or no margin in interrupting rating above initial requirements. A good working rule is to anticipate the most logical future steps in a possible expansion program for the purpose of checking probable future interrupting duties. Conscious planning for flexibility will avoid much future waste and many unnecessary barriers to expansion, sometimes with little or no increase in initial investment. PURCHASED-POWER ARRANGEMENTS

Whether an industrial plant generates power or not, i t is typical to have a power connection t o a utility system. Certain characteristics and requirements of the utility service must be understood. It is therefore important t o investigate the purchased-power arrangement as early as possible because the power company must first study the request with relation to its own problems. A number of matters should be settled with the power company before even preliminary engineering of the primary distribution system can he undertaken. Whether there will be a consumer-owned main substation or not is an important matter. The service voltage that can be made available must be known, and there will sometimes be a choice. It is also desirable to know the maximum and minimum no-load supply voltages, or the voltage spread. Another characteristic always needed is the symmetrical rontribution from the utility system to a three-phase short circuit. The system engineer is better informed if he knows the actual present maximum and minimum duties for checking voltage conditions, as well as the anticipated maximum duty for selecting suitable switching equipment. Maximum and minimum line-to-ground short-circuit-current contributions should be known, but this information along with metering and line-protection requirements may not be essential for preliminary system planning. Particularly in plants with large purchased-power demands, hut also in some relatively small plants that need extra power-supply reliability, two power-company lines, and sometimes three, will be made available. It is not enough t o know simply that there will be two lines; the various

708

PRIMARY DlSTRlBUTlON SYSTEMS

possible differences in two-line characteristics can require radically different main-bus or substation arrangements. I t is necessary to be able t o classify the service as one of the following kinds: 1. The two lines can be used one a t a time only-in some cases as alternates, in other cases on a preferred-emergency basis. 2. The two lines can be used simultaneously and may be operated in parallel. 3. The two lines can be used simultaneously but must not be operated in parallel. 4. The two lines constitute a loop service with plant-owned normally closed sectionalizing switching to permit a net power transfer past the industrial substation. If there are two supply lines, the power-company contribution to substation short circuits should be inquired about with particular care. If the symmetrical contribution to a three-phase short cirnuit, is known for each line separately, nothing further is needed for nonparalleled lines. But if the lines are to be connected in parallel by the user’s equipment, then the total duty with the two or more lines simultaneously contributing to the same three-phase short circuit should be known in addition. From data in this form a suitable equivalent circuit can be derived for determining short-circuit duties a t various points and for checking voltage drops. Chaptcr 16 covers some further matters of mutual interest to the power company and the industrial plant, if it is desired to operate plant generators in parallel with the utility system. PRIMARY-SYSTEM BREAKERS

Three classes of circuit breakers are used in the primary distribution system. These are outdoor oil circuit breakers in outdoor stations, air circuit breakers in statioh-type cubicles, and oilless circuit breakers in metal-clad switchgear. Outdoor oil circuit breakers are typically used in outdoor stations where there are overhead lines operating above the 14.4-kv level. Figure 13.1 shows a 115-kv circuit breaker a t the right in an outdoor substation used by a large steel mill. Circuit breakers in this class have always become available as larger power systems develop a need for them. Listed circuit breakers include sizes up t o 330 kv with an interrupting rating of 25,000 mva. Metal-clad switchgear is the most widely used class in the primary distribution system. Figure 13.2 shows an indoor installation of metal-clad switchgear with one of the circuit-breaker units removed. This line of switchgear includes circuit breakers up to the 13.8-kv 2000-amp size with a maximum interrupting rating of 500 mva. Metal-clad switchgear is

PRlMARY DISTRIBUTION SYSiEMS

709

.. FIG. 13.1

Main substation rhowing 115-kv outdoor circuit breaker a t the right and power tronrformerr throot-connected lo 13.8-kv metol-clod witchsear in the center.

FIG, 13.2

Indoor metol-clod rwitchgeor with one circuit-breaker unit removed.

710

PRIMARY DISTRIBUTION SYSTEMS

well suited to haiidle the main-hus arid distribution switching problems in the 2.4- t o 13.8-kv range for a11 hut a few unusually large plants. It is technically and economically sound t o employ statioii-type cubicle snitchgear, as illustrated by Fig. 13.3, in some parts of very large industrial systems. This s\ritchgear is available at volt,age ratings of 14.4 and 31.5 kv. The higher voltage level is seldom needed, hut the 14.4-kv eqnipineiit is risefill for main buses where the sources are large, especially if they are higher than 2000 amp, and the short-circuit duty cannot be limited to 500 niva. Even in such large systems, the major part of the primary-syst,em switchgear w i l l still be of the metal-clad construction out in the plant and on snhhuses. It, becomes t,oo costly to use station-type s\i-iti,hgear for the many smaller circuits, and the problem may be solved as shoii-ri in Fig. 10.25 of Chap. 10 hy the iise of ciirrerit-liinit,ing feeder reactors. As mentioired in Chap. 10, the most p~ofit,aiileiise of statioiitype sxvitchgrar is at the 1500-mva iiiterrupt,ing rating rat,her than 1000 mva a n d at 13.8 kv rather than 6.9 kv. At 14.4 kv, this class of smitchgear includes a 5000-amp k i w i t breaker wit,li a11 interrupting rating of 2.500 mva. The circuit-breaker equipment in m y primary distribution syst,em has an important ct'Eect, on over-all system cost as \re11 as on the performance. In particular, t,he availability and costs of circuit breakers help t o tfeterten1 voltage, the most practical sizes of soui the liest bns and feeder arrangements.

FiG.

i3.d

inrtoiiviion o i i4.4-iv rtotion-type cubicle switchgear.

PRIMARY DISTRIBUllON SYSTEMS

71 1

A cost sense for circuit-breaker application is a little hard to develop. because estimating data are not readily put in the familiar dollars-per-kva form used for many apparatus items. Circuit voltage and interrupting duty typically call for circuit breakers where only a small-but still widely variable-fraction of the continuous-current ability will be needed. For example, at 115 kv a single circuit breaker costs about the same as a 5000-kva transformer. However, the same circuit breaker with a current rating of 1200 amp is capable of handling a three-phase load of 239,000 kva. Some useful cost data on circuit breakers are included in Chaps. 10 and 17, and further reference to manufacturers’ schedules may be needed in making cost comparisons. TRANSFORMER CONNECTIONS*

The transformers used in main substations and a t other points in industrial primary distribution systems are typically three-phase twowinding units with delta-connected primary windings and Y-connected secondary windings. Two objectives are answered by the delta-Y connection. First, the Y connection provides a neutral connection point for system grounding purposes according t o the preferred practices covered in Chap. 6. The transformer secondary winding is a power source for the system it supplies, and the alternative of using a grounding transformer in connection with a delta source would be more expensive and less dependable. Second, the delta-connected primary winding is the simplest and surest way of stabilizing the Y-secondary neutral. Occasionally a Y-connected primary winding will be found desirable. For these a Y-Y two-winding transformer should be generally avoided in favor of adding a delta-connected tertiary winding. Among the reasons given for wanting a Y-connected primary winding is one based on a very common misunderstanding. When a four-wire service is available for supplying a substation transformer, an assumption is often made that it is compulsory t o use the neutral fourth wire, thus requiring a Y-connected transformer winding. This is not true; in fact, it is generally preferable not to use the fourth wire even though a Y-connected transformer primary winding is used for some other reason. A Y-connected autotransformer can be an unsatisfactory and dangerous system component. This connection should not be employed unless the possibilit,ies of trouble are thoroughly understood and will be guarded against in the application. Three-winding transformers may be used to interconnect parts of a * See Blume, Boyajian, Carnilli, Lennox, Minneei, and Mantsinger, “Transformer Engineering,” John Wilcy Q Sons, h e . , New York, 1951.

712

PRIMARY DlSTRlWTlON SYSTEMS

system a t two or three different voltages. These unit,s have the windings on a single core mounted in a single tank. They are t,echnically sound and will permit, a saving for an occasional applicat,ion, but they are not used so oft,en as might be expected. The reactances hetweeo each pair of windings and the loading relations are fixed by design, and t,here will consequently be less opportunity for later system modifications than if separate transformers are used. MAIN SUBSTATIONS

Xot all plants will o m and operate a main substation for supplying the primary dist,ribution system. A plant main bus serves the Same purpose if the purrhased-power voltage is suitable vit,hout transformation for the plant primary system. The principal functions of a main substation are indicated in Fig. 13.4A, which is a simple arrangement answering the requirements of a great many Emaller plant,s. More complivated substation arrangements result when there arc two or more incoming lines, two or more power transformers, or one of a number of other bus arrangements: Also in plants with power generation, the snbstation output may not supply a plant main bus but may be connected t o a synchronizing bus, as described later in this chapter. The substations in a few vary large plants with heavy loads in widely separat,ed areas may require transmission-voltage feeders connected t o the incoming-line bus, as mentioned in Chap. 10. Figure 13.4B differs from Fig. 13.4A in using power fuses instead of a circuit breaker in the incoming line. Circuit breakers are generally preferable, but fnses will be useful in satisfying over-all objectives in some of the smaller and simpler substations. When substat,iori primary fuses aTe used, it is better t o employ solid neutral grounding of the transformer secondary than t o limit the ground-fault current in the primary distribution system. The remaining substation examples all show two supply lines. I n these stations it will often be necessary t o accept some functional compromises in the high-voltage switching equipment for cost reasons. Inasmuch as smaller plants must sometimes be served from higher voltage systems, the main substation high-voltage circuit-breaker equipment can be disproportionately expensive among the other substation components. Stated in another way, a given high-voltage breaker arrangement for a given supply system will cost just about the same regardless of the substation size. The discussion is intended simply t o indicate what the several arrangements offer. Figure 13.4C shows a two-line single-transformer substation using two high-voltage circuit breakers. This arrangement might be used whether

713

PRIMARY DlSTRlWTlON SYSTEMS

the two lines are alternate, paralleled, or part of a loop. For a loop supply, the substitution of a circuit breaker for the transformer horn-gap switch would avoid opening thc loop hy the transformer protection scheme. T h e use of either two or three circuit breakers might he hard t o justify in particular cases. For alternate-line or preferred-emergmry

0 IBI I

I

(El

FIG. 13.4

Some typical main substation orrangementr used by industrial plants.

714

PRIMARY DISTRIBUTION SYSTEMS

supply, the single circuit breaker of Fig. 13.40 with interlocked incomingline switches has a minor deficiency in not permitting an aut,omatictransfer between lines. Either Fig. 13.4C or Fig. 13.40 permits expansion by adding one or more transformers t,o the high-voltage bus. Figure 13.4E is simply an extension of Fig. 13.4C for a two-transformer substation where the two incoming lines are alternate, paralleled, or part of a loop. As illustrated with four high-voltage breakers, this substation arrangement can provide a n unusually high degree of service reliability except for a high-voltage bus fault. For the special case of two incoming lines that may be operated in parallel hut are not a loop supply, the arrangement of Fig. 13.4F is often a good solution. By omitting the high-voltage bus, and paralleling on the low-voltage side of the transformers, a saving in high-voltage breakers and structure is accomplished. The arrangement reduces the availability of the total transformer capacity because each unit has a transmission line in series. However, the station-cost reduction may be so significant for smaller substations that load curtailment during an outage becomes an acceptable risk. I t is moreover possible t o reinvest part of the circuitbreaker saving in additional size of transformer units to achieve service cont,innity for all t.he load or to reduce the amount of load curtailment during half-capacity operation. Referring again to Fig. 13.4E in connection with loop supplies only, the high-voltage part of the substation employs almost all the circuit breakers that can be fitted into a single-bus arrangement. However, a fifth circuit breaker could be added in the bus. With appropriate relaying, it would ensure continuity of service through one transformer under the condition of a high-voltage bus fault. I t is perhaps more profitable t o observe how reliability and flexibility are modified by removing circuit breakers one a t a time, as illustrated in Figs. 13.W t o 13.45. I n the three-breaker echeme of Fig. 13.4G the main functional compromise is that transformer protection requires opening the loop supply. A utility would not ordinarily consider this as a serious shortcoming, but it could Ee avoided in the alternative three-breaker scheme of Fig. 13.4H. Either of these arrangements provides service continuity through one transformer for any single fault, including a high-voltage hus fault. I n the two-breaker scheme of Fig. 13.41, operation of the protective relaying of either transformer not only opens the loop but drops the whole subst,ation load. The loop can be reclosed and plant service can be reestablished t,hrough the unfaulted transformer circuit by manual switching. A permanent high-voltage bus fault must, of course, be repaired before either circuit breaker can be reclosed. I n attempting to use a single breaker as shown in Fig. 13.4J, a problem

PRIMARY DISTRIBUTION SYSTEM

715

is encountered. Any high-voltage fault down to the transformers will be cleared by the single circuit breaker and the utility as a single-line short circuit, leaving uninterrupted plant service through one transformer, However, there will he a level of transformer fault current below which the utility cannot trip, and the faulty unit cannot be automatically disconnerted from the system a t such a level of overcurrent except by transferred tripping of a power-company circuit breaker using carrier or a pilot wire. BUS ARRANGEMENTS A bus is a junction of three or more incoming and outgoing circuits. The most common plant bus arrangement consists of one source or supply circuit and two or more feeder circuits. The numerous other arrangements and variations are mainly intended to improve the service reliability through the bus t o all or part of the load during expected maintenance, or in the event of equipment failure or source outage. Some very complicated bus arrangements have been used in trying to improve service reliability or continuity. Some of these arrangements are technically unsound and will not provide actual benefits. Other arrangements that do qualify from an engineering viewpoint are useful in meeting the rather typical requirements in the heavy industries that handle large amounts of power through main and subdistribution buses. These same bus arrangements will seldom prove acceptable for cost reasons in medium-size and small systems even when service continuity is considered to he unusually important. The highest quality of service reliability can often be obtained more economically for smaller plants, particularly for those with load-center systems, by over-all system arrangements that employ simpler and less costly bus arrangements. The double-bus arrangement shown in Fig. 13.5A is an example of the more complicated arrangements that is technically sound if good-quality equipment is used, but it is very costly for the usual sizes of feeder circuits. The arrangement is suitable for outdoor circuit breakers, station-type cubicles, or metal-clad construction. I n metal-clad equipments, some requirements can be met a t lower cost by employing two positions and one circuit breaker per circuit plus one spare removable circuit-breaker element, as illustrated for one of the several circuits. This variation still allows transferring any circuit or maintaining any circuit breaker without a feeder interruption. Figure 13.5.4 was intended t o indicate a preferred physical arrangement with companion circuit-breaker compartments in separate standard equipments facing each other across an operating aisle. A cable cannertion would usually join the circuit breakers. Occasionally a special

716

PRIMARY DISTRIBUTION SYSTEMS

metal-clad equipment has been built with both buses running the lemgth of the equipment, placing companion circuit breakers in adjacent compartments. This is an inferior construrtiori because i t provides a possibility of involving both buses if there should he a serious switchgear failure. Most of the more complicated arrangements have in common the gen-

A

9

T 7 5 e

I

ICI

$Y

$Y

8Y

(Dl FIG. 13.5

Some typical bur arrangements used by industrial plants.

PRIMARY DISTRIBUTION SYSTEMS

717

era1 charact,erist,icthat, individual lines can he connected to either of tivo buses (often without service interruption) with good maintertarice aci'ess t o most of the apparatus. Intermediat,e flexihility and reliability can be more economically obtairied for multiple-source bus arrangcment,s hy sectionalizing straight single buses. Figure 13.58 illustrates a typical two-source sectionalized-bus arrangement with a single cirriiit breaker per line. Figure 13.5R or some variation places lines and hreakers on the same basis of availability. Where metal-dad switchgear is used in the primary system, a feeder outage for circuit-breaker maiut,enance can he reduced to a matter of minutes with a spare removable circuit hreaker on hand. I n extending reliahility from a main bus t o a subhus in an important load area, parallel feeders may he used. In t,he load-center system as described in Chap. 11, each load-renter transformer has the same availahility as its primary feeder and supply breaker. Improvement in service reliability is secured by intercoiinect,ion at secondary voltage. When three or more sources are available at a main bus, Fig. 13.5f,"is a natural extension of Fig. 13.5R. However, Fig. 13.5D is more flexible and is usually preferred eveit when another circuit breaker is needed. This arrangement may be referred t o as a star bus, but it also is somet,imes called a synchronizing bus arrangement whether any of thc sources i s a generator or not. Particularly if reactors are needed t o parallel the sources, Fig. 13.5D will he preferable t o a straight bus (or a riug bus) with the current-limiting reactors installed hetween each tie circuit breaker and the common hus. The need for tie circuit breakers nil1 be obvious in some straight buses, but there will be other cases where the value may be in doubt. Experience shows they are too often omitted where a choire rail he made in t,he planning stage. The following remarks are intended t,o summarize thc various ways in which hus-tie circuit breakers may be useful initially and later. When two sources are used simultaneously hut must not he opcrat,ed in parallel, a normally open bus-tie circuit hreaker interlocked with the source rirruit breakers permits serving hot,h hus sections from one of the sources mheri the other is not available. Reasons for not paralleling t h e sources might he that they are not synchronized or h a r e a vokage phase difference. .4nother reason could be t o reduce the bus short-cirruit duty either initially or in t,he future if the duty might he increased beyond desired limits through additions t o the source capacity. For alternate (or preferred-emergency) or normally paralleled sources, a single straight bus may bc used. It is preferable to use a normally closed bus-tie circuit breaker so that one bus section can he kept availatile

718

PRIMARY DISTRIBUTION SYSTEMS

when the other is out for maintenance or repair or t o permit additions during a plant expansion. For paralleled sources, relaying of the tie circuit breaker may be employed to split the system so that service continuity is retained on one bus if the other bus should feil or i t became necessary to back up a feeder circuit breaker on that bus. FEEDER ARRANGEMENTS

A feeder carries energy to a substation or bus or to several loads. The several feeder arrangements discussed below are illustrated together in Fig. 13.6. The primary feeders that supply load-center unit substations, as described in Chap. 11, are not included as further arrangements because they are fundamentally radial feeders. The main function of a tie feeder is t o connect two sources. It may connect two substation buses in parallel t o provide stiffness or service continuity for the load supplied from each bus. If either source has plant generation, then the tie feeder maintains the two parts of the system in synchronism and provides a circuit for transferring normal power and kilovars in either direction between the sources. A loop feeder also has its extremities connected to a source (usually a single source), but its main function is to supply two or more load points between. Each load point can be supplied from either direction; so it is possible to remove any section of the loop from service without causing an outage at any load point. A radial feeder connects between a source and a load point, and it may supply one or more additional load points between. If the connection to an intermediate load point is an “in-and-out’’ or “loop” Connection instead of a tap, the feeder does not, of course, assume loop characteristics inasmuch as each load point can be supplied from one direction only.

FIG. 13.6

Four primary-system feeder arrangements.

PRIMARY DISTRIBUTION SYSTWS

719

Radial feeders are the most widely used because they are simple, easy t o protect, and low in cost. They are simple because there is only one path for current to any given load point. They are easily protected by simple overcurrent relays a t the supply circuit breaker. The cost is low because there is no duplication of equipment. These comments apply to “single” radial feeders and not to “parallel” feeders which have characteristics resembling those of loop feeders. Parallel feeders consist of two or more feeders bused together both at the sending aud receiving ends. I t is sometimes more economical to design a heavy ctlhle circuit with two or more cables in parallel, but these cables cannot he considered as parallel feeders when single-circuit switching equipment is used. Parallel feeders as illustrated by Fig. 13.6 provide a high degree of service reliability, or continuity, if one of several methods of protecting parallel lines is employed. This protectiou is more complicated and expensive than the simple overcurrent protection ordinarily installed in single radial feeders. Separated circuits are desirable, and each will typically have the ability to handle the normal load with the other circuit out of service. Additional circuit breakers are needed; in general, four circuit breakers are required for parallel feeders supplying one load point, while only one circuit breaker is needed for a single radial feeder. There are obviously some heavy cost penalties against the parallel-feeder arrangement; so its excellent characterist,ias can he justified only for serving large loads such as suhdistribution buses or smaller loads with unusual service requirements. Several variations intended to improve upon the performance of single radial feeders have costs below those for parallel feeders. They all sacrifice service continuity as a characteristic. For particular requirements, one or another of these variations may be a preferred solution; hut t,hey teud in general to have questionable value. One rariatiou is t,o m e a conventional parallel-feeder scheme, but to use overcurrent protection only. This does not represent an important saving, arid a short, circuit in either line interrupts service. Moreover, the fault location is not indicated, and corisiderable time may be spent in finding which lioe is in trouble unless it is considered acceptable to take the chalice of closing again on t,he same fault. .I second variation is like the one above, except that one of the two circuits is operated normally open and is held as an alternate or reserve feeder. The equipment is the same, and the only advantage is that the fault location is indieated when service is interrupted. A third variation makes a further saving by using a total of four metalclad circuit-breaker positions but only two circuit breakers. The performance is similar to that of the second variation evcept that the outage

720

PRIMARY DISTRIBUTION SYSTEMS

FIG. 13.7 Feeder-equipment compclriron for three alternative arrangements each rupplying the same three loads.

will he longer t o allow extra time t o move the circuit-breaker elements. Circuit-breaker maintenance rvit.hout service interruption requires that two spare circuit-breaker elemerits be available. Figure 13.7 is included in combination with Tahle 13.1 to lead into a discussion of loop feeders. Three load point,s are arbitrarily selected for compariug the number of cables and circuit breakers necessary for each of the three feeder arrangements illustrated. Tahle 13.1 then indicates how large the cost differencesmight he without considering any differences in protection costs. TABLE 13.1

Circuit Breakers and Typical Amounts of Cable Required for Supplying N Load Points by the Three Feeder Arrangements Shown in Fig. 13.7

......, . . . ......... .... ......, .....

Number of Circuit breakers.. Cable current rating.. . . Cable footage.. .

4N

2N

+2

Loop feeders perform much the same as parallel feeders, and the operating features can usually he ohtairied a t less cost for the two or more load points. But this is still a high-cost feeder system that will prove economically sound in about the same k i d s of situations where parallel feeders can be justified. Loop feeders look so attractive to many engineers that variations with lower costs are often considered and sometimes adopted. Most of these arrangements only resemble good loop systems and can he aualyzed into positions of lower value than radial arrangements of still lower costs.

72 1

PRIMARY DISTRIBUTION SYSTEMS

Before esamiriirig any of these arrangements, there should be an understanding of the kind of equipment needed in a loop feeder t o produce the characteristics associated with it. h loop feeder is intended t,o provide service cont,inuity. It must therefore be operated normally closed and must have two sectiorializirig circuit, breakers at each load point. These breakers must be adequate for interrupting system short circuits, aud they must be a part of a complete protective system, including the source circuit breakers, so that any faulty section of t,he loop can be automatically removed from service mithout dropping any load. The preferred loop relaying is wire-pilot ditferent,ial. Directional overcurrent is slower but may be attempted t o reduce costs if there are only two or three load points. Figure 13.7 shows the switching equipment needed for the above performance in a loop feeder. One variation is shown in Fig. 13.8.4 where half the sectionalizing circuit breakers are omitted. Any feeder short circuit causes an outage at one load point, and an acceptable protection system is sometimes hard t o design, particularly for more than two or three load points v i t h any considerable distance between them. Figure 13.8B shows a further variation with no sectionalizing points. Scarcely any of the loop-feeder benefits remain. The service reliability is about the same as is provided by a single radial feeder, except that a feeder circuit breaker can be maintained without interrupting any load. Another variation of a good loop-feeder arrangement is the one illustrated by Fig. 13.8C. As a-ar true for Fig. 13.8B, a single fault drops all load. Service continuity for half the load can be obtained by operating the loop normally with a central sectionalizing switch open. Following

IEI

(A1 FIG. 13.8 Variations from the loop-feeder practice.

best

YY-Y (CI

722

PRIMARY DISTRIBUTION SYSTEMS

a short circuit, service t o all the load can be reestablished by isolating the faulted section. The arrangement can serve the same complement of substations that can be served by a single radial feeder without needing individual protection of the substation transformers. Without further analysis this arrangement appears t o have attained fair service reliability with a moderate cost increase. However, most operating engineers will reject this system for cable circuits using metal-enclosed disconnects at the sectionalizing points because there are serious hazards involved in the switchingoperationslikely t o bc followed after a short circuit has occurred. Whether t,he loop was open or closed when the fault occurred, the problem remains t o find the section in trouble so that service can be restored to all loadpoints. While it is not the best procedure, a common routine involves energizing thc feeder from one end by repeated trial, after adding a section at a time, until the faulty section is located by immediate tripping of the source circuit breaker. These operations present repeated opportuuities t o make swit,ching mistakes. Even if the operator appreciates the hazard, there is still the temptation t o energize the sections one at a time by means of the manual disconnects. These switches are also hazardous to operate under normal conditions because every section of a closed loop usually carries current. A loop feeder carried on open overhead lines with structure or polemounted disconnects for sectionalizing is much safer. One difference is t h a t short circuits on overhead lines are more often temporary than permanent. h more importaut difference is that any switch failure is considerably less likely t o injure a n operator because of the remote operation. Several precautions should he observed in loop and parallel feeder lay-. outs. It is well t o avoid compound loops-those with more than two source circuit breakers or with more than one path from one of two source circuit breakers back t o the other. Loop systems such as these may develop in an attempt t o relieve overloaded circuit conductors. This is seldom an efficient solution compared t o starting a new single loop. The load division is difficult t o calculate, requiring close estimates of circuit impedance and assumed iixed loads at each load point. Normal load variations and load growth can cause unexpectedly wide changes in the load division, and compound loops therefore tend t o give only temporary relief at best. Other compound-loop systems simply grow without sound planning. They tend t o be associated with poor load division, poor protection, and compromised service. Particularly where disconnects are used instead of circuit breakers, operation is more complicated and therefore more hazardous. It was implied earlier that tie feeders do not supply load points between the two sources. This arrangement, combining the usual functions of

PRIMARY DISTRIBUTION SYSTEMS

723

tie and loop feeders, is sometimes advantageous. At the same time, ccrtain complications and compromises of such a dual-purpose circuit should be recognized and weighed before finally deciding in favor of it. Listed beloiv are several possible shortcomings: 1. The metering of load transferred between sources is more complicated. 2. Additional protection at the load points may he needed to avoid opening the tie circuit on load-tap fault. 3. If synchronism is maintairied by t,he tie circuit, suitable precautious must he taken to prevent closing the tie by any sectionalizing circuit breaker because it will not he feasible to install synchronizing provisions at the several sectionalizing points. Another precaution applies equally to parallel and loop feeders. These arrangements may complete a closed circuit around a bus-tie circuit breaker or a current-limiting reactor when the source circuit breakers are connected t o different buses. Such connections are oftcu overlooked in system design work. I t is obviously a mistake to short-circuit currentlimiting reactors by a feeder arrangement, and it is at least questionable to have a bus-tie circuit breaker short-circuited by independent primary switching operations a t one or more remote points in the system. A review of the several reasons for using bus-tie circuit hreakers, as covered in another part of this chapter, will indicate how seriously the over-all system characteristics may be altered by au oversight of this kind. FEEDER INSTALLATION

Power lines are installed in a number of ways to suit a variety of local conditions. Even under similar conditions, practices vary according to the experience arid preferences of users. Outdoor Feeders. Outdoor circuit,s may he run overhead on poles with open-mire construction. This is a low-cost method, and it cau be credited with some further minor advantages for particular circumstances. However, it is being used less and less in industrial plants hecause of its disadvantages. The main objections to open-wire overhead construction are the hazard and outage possibilities from weather, pole breakage, and accidental contact with lines by mobile machinery such as cranes and shovels. There have also been many cases of direct contact by maintcnance personnel working on building roofs or at other places close to exposed lines. Open lines present further problems in contaminated atmospheres. For example, in some steel mills the dust problems may require overinsulatiou and insulator cleaning as often as once or twice a year to prevent flashovers. There are other types of plants where conducting dust or chemicals will similarly affect line insulation levels. Overhead lines on poles may use aerial cable, and this alternative

724

PRIMARY DISTRIBUTION SYSTEMS

method is finding increasing use. Aerial cable employs insulated conductors compactly bound t o a mcssenger Ti-ith one or more circuits on the same poles. This construction cost,s more, but it offers better value for many installations by modifying the disadvantages of open-wire circuits. (1) The construction is compact and does not require the largc mechanical and electrical clearances of open wiring. The chance of arcidental contact is therefore reduced. (2) Thc hazard arid outage possibility from accidental contact are greatly reduced. (3) There is no insulator cleaning problem in adverse atmospheric conditions. I t is t,rue that insulation failure of aerial cable will cause a longer outage time, but this disadvantagc is usually an acceptable risk. A second altcrnative for outdoor circuits in industrial plants is underground cable. The excavation, duct vork, and perhaps manholes bring the cost of this method above t,hat of aerial cable. It is certainly desirable t o he able t,o eliminate overhead conductors and the supporting structures from congested industrial areas, and underground cable work has been the most popular between buildings and even t o outlying areas where the distances are relatively short. Direct burial is the less eupefisive (:onstruction even if it is slabbed for mechanical protection, and it is used t o a small extcih for single circuits where extended outage in the event of insulation failure can he tolerated. Outdoor underground construction provides somewhat less protection than might be assumed. One of the most common causes of failure is mechanical, from construction or maintenance work. It is also true that some soil condit,ions may at,tack a cable sheath and that water, oil, or chemicals in ducts or manholes may reach and deteriorate the insulation. Indoor Feeders. There is a definite trend away from underground conduit and duct work for the cable circuits inside buildings. Increasing preference is being given t o the use of interlocked-armor cable installed overhead for these reasons: 1 . The cost, is lower than for underfloor circuits. 2. Overhead circuits are accessible and can be more easily changed. 3. Their location will not interfere with the fnture installation of machine foundations. 4. Overhead installation protects the cable from possible damage by water, oil, or chemicals that may collect in underfloor conduits. I n some places it Tvill be desirable for appearance reasons t o conceal the primary-system wiring. Another objection t o the exposed circuits was t h a t they might he mistaken for secondary circuits or that, conduit might be mistaken for piping. Possible confusion of this kind is hardly a valid argument against a generally desirable installation method, and it seems t o have disappeared from the thinking of most plant engineers.

PRIMARY DISTRIBUTION SYSTEMS

725

POWER GENERATING STATIONS

Chapter 16 should be consulted for general information about industrial power generation. The following material deals mairily with the generating-station bus arrangement and some problems that may be cncoui1tered in meeting the general requirement of integrat,ing the whole primary electric system, including any normally used tie with a utility. Generators are integral parts of turbine-generator sets; so generator details tend to be decided while dealing with the problems of prodnoing heat energy at a planning stage when no thought may have been given to the effect these characteristics can have on electrical problems. Coordinated planning will produce the best over-all solutions and may prevent unsatisfactory or costly compromises. Generator voltage rating and winding conriection are examples of characteristics that should be chosen entirely on the basis of what is best in the electric system. Generator kva rating is closely related t o the prime-mover size; so the selection of turbine-generator set size should recognize both heat and electrical objectives. The important effects of generator kva and voltage on the primary-syst,em smitchgcar mill be indicated in the following discussion. There mill he many eases vhere direct parallel connection of two or more sources on a generator hus will answer all the requirements for the initial inst,allation, but available switchgear will set a limit on the future size of the system. Even the initial station may be too large to permit direct parallel connection at generator voltage. The usual solution for industrial systems is a synchronizing bus, as illustrated in Fig. 13.9, which also s h o w the generally preferable method of connecting to purchased power through the synchronizing bus. Some curve data applying t o an elementary synchronizing bus pattern will help to make clear a few points t.hat are worth remembering. Figure 13.10 shows several identical generator-bus sections interconnected hy a synchronizing bus and identical reactors. The reactance values are

FIG. 13.9

Typicol generator synchronizing-bus arrangement with purchased-power tie.

716

P R M R Y DISTRIBUTION SYSTEMS

FIG. 13.10

Typical reactance values for

CI

generotor synchronizing-bus ponern.

reasonable and may be considered to be either those suitable for momentary- or interrupting-duty calculations. All reactanczs are per-unit values on the kva base of a single generator. Figure 13.11 shows how the generator-bus duty is limited by the synchronizing reactors. The total short-circuit duty includes a direct contribution of fifteen times normal from one generator and the motors supplied by one bus. The remaining cont,ribut,ion comes through one synchronizing reactor and varies with the numher of generators. The maximum contribution through a 10 per cent reactor is ten times normal for an infinite number of machines, makiug the total maximum load-bus duty equal to twenty-five times normal. The dashed line shows the direct variation of load-bus duty with machines directly paralleled and emphasizes how very effective 10 per cent synchronizing reactors are in limiting the duty. A further point of interest is the duty on the synchronizing bus. Even with reactance as high as 10 per cent, the synchronizing-bus duty cxceeds the load-bus duty for four or more machines. It is plain that the synchronizing bus may have to he braced for highcr short-circuit currents than thc load buses. It, is also apparent that the synchronizing bus should not be regarded as a load bus because the short-circuit duty may later exceed the ability of available or practical circuit-breaker equipment. The advisability of installing eit,her circuit breakers or disconnects t o isolate a faulty reactor from the synchronizing bus should also he questioned. The most important matter illustrated by Fig. 13.11 is that an early limit in size of system or numher of generators can be reached if the initial machine sizes are too large. Using the reactance values of Fig. 13.10, the load-bus duty for a single generator supplying an equal kva of motor load is fift,een times normal and the maximum generator rat,ing for 250mva switchgear would then be 250,000 divided by 15, or 16,667 kva. For an infinitely large system, the 250-mva ability mould be exceeded if the generator ratings were higher than 250,000 divided by 25, or 10,000 kva. In a similar way, the use of 5 per cent instead of 10 per cent synchronizing reactors would require limiting the machinc size to 250,000 divided by 35, or 7,140 kva, to he conserwhive in planning a large system using 250-mva

PRIMARY DISTRIBUTION SYSTEMS

727

switchgear. I t is riot recommended that actual systems he planned using these assomed reartatices; the closest estimates that can be made for the particular apperet,us should be used instead. The proper valuc of synchronizing reactance is sometimes a perplexing question. The least reactance that will meet short-circuit objectives is desirable. The value of 10 per cent mentioned several times is actually

z

I-

6or

2 50 W

z c)

40

0 -

I

3

4

NO. O F GENERATORS ( O R LOAD BUSES)

__

FIG. 13.1 I Effect of synchronizing reoctori on short-circuit duties for the rynchroniringbur arrangement of Fig. 13.10.

728

PRIMARY DISTRIBUTION SYSTEMS

fairly high, as \ \ i l l be shiiivti. First it should he remcmf,cred that the ohnrir value of r r w h t l r e rontrols short-circuit crtrretlt \\-bile the per rent reaibtwe 011 the amorlnt of load transferred is at1 iildex of voltage drop for that load. I t should also he rioted that synchronizing reactors are oftctl designed for handling only a f r a h o n of t,he generator kva and v i l l thcti oftrii be referred t o it1 per cent oil their own base. Figure 13.12 hclps to visiidizc the effect of varyitrg the amount, of synchroirizitig rrnrtatlw f o r the same generator and motor feartatlccs risrd beforc. .\I1 rra~~tatlres are again per-unit, values h s e d on the k v a rat,ing

FIG. 13.12

Decreoring effect of higher values of reactance in t h e synchronizing-bur

orrongement of Fig. 13.10.

PRIMARY DISTRIBUTION SYSTEMS

729

of a single generator. It is evident that the first few per cent of synchronizing reactance are the most eRective in reducing short-circuit levels and that relatively little further henefit results from increasing the reactance from, say, 5 t o 10 per cent or higher. There are some minor objections t o larger reactors, such as higher costs, higher losses, and finally system instahilit,y. However, the most practical objection is the voltage drop accompanying load transfer, and this drop varies directly with the amount of reactance. For example, a 10 per cent reactor on a given base rauses twice the amount of voltage change caused by a 5 per cent reactor on the same base when transferring the same load. Figure 13.13 illustrates this point for three reactor values. Figure 13.13 illustrates another matter of interest. Reactors typically have relatively small values of resistance so that voltage drop a t a load power factor of unity is very low, but it increases sharply as the power factor departs from unity in the lagging direction. (The curves of Fig. 13.13 assume a reactor X / R of 50.) It is therefore important t o recognize probable operating power factor in contemplating the use of larger values of synchronizing reactance. Even where load power factor will be high, it should he remembered that low-power-factor demands such as those caused by starting motors may produce objectionably high temporary voltage drops. The curves of Fig. 13.13 show voltage drops from a point of constant voltage in the system. This point will he at active generator buses for

LOAD P F ( L A G G I N G ) FIG, 13.13

Voltage drop ot load bur when transferred through synchronizing reactors.

CI

load equal to generator rated h a is

730

PRIMARY DISTRIBUTION SYSTEMS

actual systems and a t the synchronizing bus as well only if an infinite number of generators supply the synchronizing bus. Estimates of actual voltage drops therefore must use an “effective” value of synchronizing reactance. For example, if normal generator kva is obtained through a 10 per cent reactor from two active generators of a three-generator scheme, the total voltage drop includes the drop through one 10 per cent reactor to the synchronizing bus plus the drop through two more reactors effectively in parallel, and the total effective reactance is 15 per cent. Two bus arrangements with interconnection reactors produce operating results similar t o those from a synchronizing-bus scheme, but neither is preferred from an over-all characteristic standpoint. Both arrangements employ tie reactors between otherwise isolated generator buses, one in the farm of a ring bus and the other in the form of a straight bus with one less reactor. Both arrangements use more circuit breakers than a synchronizing-bus scheme and are inferior with respect to voltage drop for several machines. A minor advantage is that neither arrangement has a synchronizing bus where a failure could force isolated generator operation. A bus-tie reactor between two generator buses is basically the same as a synchronizing-bus scheme. If only two generators are installed initially, a single reactor should be used. At the time of installing the third machine, two additional reactors would be installed in the familiar synchronizing-bus configuration. Several other arrangements have been used in generating stations either t o limit short-circuit duty or to reduce the continuous-current requirement of generator switching equipment. Generator series reactors have sometimes been used to assist in controlling short-circuit duty in modernizing or expanding existing systems, but they can be regarded as generally undesirable. Duplex reactors offer a more efficient control of short-circuit level and also permit a reduction in continuous-current rating of the two breakers per generator. Double-winding generators produce similar results without needing external reactance. All these arrangements are somewhat specialized solutions. A very satisfactory arrangement is the unit system of directly transforming the generator output in a step-up transformer to eliminate all switching equipment a t generator voltage. An example of such an arrangement is shown in Fig. 10.26. This scheme has not often been used in industrial systems, which seldom have a transmission problem where the higher transformer output voltage is more suitable than the generator voltage level.

Chpter 14

by Donald S. Brereton

Power Systems for Commercial Buildinqs c

The subject of power systems for commercial buildings involves both the basic power-supply system consisting of the main switchboard or substations and principal secondary feeders and panel boards. I t also involves the many details of installation of branch circuits, out,lets, wall switches, lighting fixtures, and many other related types of equipment. This chapter will deal primarily with the discussion of the basic powersupply system and not with the details of installation of branch circuits and associated equipment. The range of commercial-building loads is from a few kva t o over 20,000-kva demand for a single building. Commercial buildings are considered to include office buildings, hospitals, schools, and practically any type of building except those occupied by manufacturing plants. They also include the officebuildings in industrial plants. Small buildings having less than a few kva demand and located outside the utility secondary network areas are almost invariably supplied secondary power by the utility, generally a t single phase, 120/240 volts. The power systems in these buildings involve service-entrance equipment and primary branch circuit equipment for their own use. Larger buildings, particularly those located in secondary network areas of the utilities, are generally supplied power a t three phase, 208Y/120 volts. The utilization equipment for larger buildings outside the network area in which the power component is relatively low and the main load is 120 volts may be supplied by 208Y/120-volt unit substations. A typical one-line diagram of a building supplied 208Y/123 volts by the utility is shown in Fig. 14.1, and a building supplied a t high voltage but having its own 208Y/120-volt substation may have a typical one-line diagram like that shown in Fig. 14.2. 731

732

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

As buildings become larger, say 500 kva or more, there is an increasing tendency t o use higher voltages than 208Y/120 volts. These types of buildings have, in general, three kinds of distribution systems. The first is the building with considerable power load which may have 480 volts supplied either by the utility or by its r-,---UTILITY I 1 own local substations for the power II I either by the local utility or by its own II ?( 0: 0: I substations for the lighting load. A ? I typical one-line diagram of such a building when power is supplied by its own substation is shown in Fig. 14.3, and one when power is supplied by theutility network systems is shown in Fig. 14.4. I n other cases the entire building supply is 480 volts, either from a spot network of the utility or from the building's own substations. (This is discussed in detail later in this chapter FIG. 14.1 Typical one-line diagram of and is shown in Fig. 14.18.) The 480 CI commercial building supplied with volts is used directly in power-utiliza2 0 8 Y / 1 2 0 volts from the network system tion equipment and then transformed of the utility. from 480 to 120 volts for local floorutilization. Thistypeof distributionsysr----------1 tem is particularly applicable in old II -r -----_----_ UTILITY I I buildings where riser conduit sizes, etc., I $ I are such that increased capacity can be obtained economically only by using higher voltage risers, i.e., 480-volt instead of 208-volt risers through the building. It is also applicable t o larger buildings where thedistancesand loads are such that the savings in switchgear and circuit conductors are enough to more than offset the cost of the 480120-volt transformers for supplying the 120-volt load. Such a building might he a large department store where there is considerable incandescent lighting and where the outlet FIG. 14.2 Typical one-line diagram of load and showcase lighting load are commercial buildine. rupplied With .~ large medium voltage b y the utility.

____--__

L ,A ..;

,""V,

~

L * ~ A ,

I"

1

---

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

733

r----------1 UTILITY

I

-__

-

I I

FIG. 14.3

Typical one-line diagrom of o commercial building when power is supplied by its own substation.

FIG. 14.4 Typical one-line diagram of a commercial building when power i s supplied by the utility network.

734

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

Figure 14.5 illustrates a large modern commercial building receiving power from the utility a t 13.8 kv and utilizing 480 and 208 volts for lighting, 2400 volts for air-conditioning motor loads, and 480 volts for auxiliaries and elevators. The reason 208Y/120-volt services are shown for use in the lower floors of the building is to provide the services necessary for the large quantity of showcase lighting on these floors. It should be noted that additional 120-volt loads are required on other floors but that $UTILITY SERVICE

7TH FLOOR

FIG. 14.5 A typical one-line diagram of (I large commercial building receiving mediumvoltoge power from the utility and tranrforming to various utilization voltages.

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

735

it will prove more economical to sci'vc those by a dry-type t.ransformer stepping down from the 480-volt, system. The newest system is applicable to large buildings of several hundred kva or more demand xvhere there is not individual metering arid xhere t,he general-area lighting is by means of fluorescent lamps. The 480Yj277volt higher voltage combined power and lighting system is finding wider anti wider appliratiori lx!cause of its lower over-all cost for these t,ypes of installat,ioiis. I t is with this system that the rest of this chapter is particularly concerned. 480Yj277-VOLT COMBINED LIGHT AND POWER SYSTEMS FOR LARGE COMMERCIAL BUILDINGS This type of system is receiving increased reiognition because of the increased power load in commercial buildings and because the need for adequat,e lighting in commercial buildings has inarcased the light,ing load. Further, the trend has been almost entirely toward the use of fluorescent lighting for gcrieral-area light,irig, which can he readily adapted t o a higher volt,age supply. This is shoxn ii Pig. 14.6. Cntil the a d r c n t of the fluoresirnt, lamp, most distribution of p o w r for lighting loads was at 120 volts for t,he most e f f i c i e n t operation of incan-

FIG. 14.6

General-oreo fluorescent lighting

in a modern office building

736

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

descent lamps. Higher voltage incandescent lamps are fragile and generally not satisfactory. Modern lighting practices, however, include more arid more the use of fluorescent lamps either alone or in combination with other types. Since the fluorescent lamp has a ballast in series with it, the 120-volt limitation no longer applies. The required voltage at instant starting is always above 120 volts and is supplied by a transformer built into the ballast. Thus the use of fluoresccnt lamps has opened new possibilities for the use of higher voltage circuits to reduce the cost of the power system. Combined light and power systems have been used for a number of years in industrial p l a n k Combined light and power systems using 480Y/277 volts have been in service over 10 years (see references 2 and 3). The fluorescent-lamp ballasts are connected line-to-neutral and the motors connected line-to-line in the 4XOY/277-volt system. This system has resulted in a substantial reduction in cost of the power system for supplying these plants compared with separate power and lighting distribution systems. I n the industrial plants the control of these lamps has been accomplished by the use of fused combination motor starters to serve large blocks (up to 15 kva each) of fluorescent lamps, but this type of control is not suitable for office buildings. Examples of new large modern commercial huildiugs using the 480-volt system for both air-conditioning and lighting loads have been described in the tcchnical press (see references 10 t.o 16). Examples of large arid small huildirigs are given in A I E E papers (see references 4 to 9). The economies of a 4x0-volt system for supplying fluorescent ceiling lights and integral horsepower motors can be obtained in office buildings. This is brought about by the development of a small control relay which can be mounted in the higher voltage fluorescent lighting circuits for controlling the 277-volt distribution circuit in the lighting fixture. This relay is operated by coils whose nominal voltage rating is 24 volts. Since only 24 volts has to be brought t o the wall switches, this obviates the necessity for using 277 volts on standard 250-volt wall switches. Special 277-volt wall switches are available. HOW COMPONENTS OF THE HIGHER VOLTAGE SYSTEM ARE APPLIED

A typical building is shown in Fig. 14.7 in a cutaway view. The various components of the 480Y/277-volt system are shown. I n the basement, the unit substation (or main 480-volt switchboard in cases where secondary power is purchased) (1) steps the supply voltage down to the required 480Y/277 volts for distribution within the building. The busway risers (2) are used to distribute the 480Y/277-volt power to each floor in the building. Fusible plugs at each floor are used to connect from

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

737

the husway to the panel boards (3). The panel boards supply 277 volts line-to-neutral for the lights (4). The 24-volt remote-control switches (5) control the lights through relays (6) and a transformer (7). A drytype transformer (8) rated 480:120 volts, single phase, supplies the panel hoard (9) for the floor circuits (10). Elevator, fire-pump, and air-conditioning motors (11) operate on 480 volts, line-to-line. The application of these componcnts is also shown in the one-line diagram, Fig. 14.8.

ECONOMIC COMPARISON OF THE 480Y/277- AND THE 208Y/120-V0LT SYSTEMS Because of the 120-volt limitation imposed by incandescent lamps, the 208Y/120-volt three-phase four-wire system has heen commonly used in the past. Motors are supplied with the line-to-line voltage of 208 volts. The lighting load is distributed on the three phases, but connected from one line to neutral a t 120 volts. Similarly, the 480Y/277-volt system is a three-phase system with neutral readily availeb!e, 480 volts line-to-line and 277 volts line-to-neutral.

POWER SYSTEMS FOR COMMERCIAL BUILDINGS

738

MOTOR FEEDER STbNDARD 3 CONDUCTOR CIRCUITS

~

''
$ l l 5 0.7 1.6 1.4120 1 . 2 2.2 2.5 95 0.6 1.4 1.5134 1 . 1 2.2 1.6 I l > 6 2 130 0.9 1.9 1.5140 1.4 2.8 2.0110 0.7 1.8 1.6154 1.2 2.3 1.5 0 2 2 I 5 0 1.0 2.4 1.6155 1.6 3.0 1.9125 0.9 2.3 1.8179 1.4 2.6 1.4 2 175 I . ? 2.6 1.5185 1.9 3 . 3 1.8145 1.0 2.4 1.7205 1.6 2.8 1.3 00 2 235 200 1.4 2.9 1.4210 2.2 4.3 2.0165 1.2 2.6 1.6237 1.8 3.0 1.3 000 2 1.8 3.8 1.7235 2.6 1.7 2.0195 1.5 3.6 1.8271 2.1 3.4 1.3 0000 2)62)$230 255 2.3 4.4 1.7270 2.9 5.7 2.1215 1.8 3.9 1.8300 2.3 3.7 1.2 250 2'13 3 310 3.0 5.8 1.9325 3.6 6.4 2.0260 2.4 5.1 2.0371 3.0 4.4 1.2 350 3 3'6380 4.0 6.8 1.8405 4.7 8.3 2.1320 3.3 6.1 1.9462 3.9 5.3 1.1 500 3 475 5.9 9.5 2.0500 6.511.0 2.2 587 5.7 7.3 1.2 750 3 3 5 4 2.2535 7.914.7 2.7 1000 4 5 545 7.411.9

_...... _._. ....

14

U

2 0

n 6

0

K

0

2 z x

w 10

n

0

9J

K

u)

P

W

K

3 18 g S

n

0

g

K

22

28

500

1000

Synchronous motors, 500 to 5oM) hp.

2000 2500 3000 MOTOR HORSEPOWER

FIG. 17.15

1500

3500

4000

4500

5oGQ

COMPARATIVE SYSTEM STUDIES DATA : MOTORS, 500 TO 5000 HP SYNCHRONOUS MOTOR l.OPF.IZ00 RPM

LOAD- AND COST-ESTIMATING DATA

922

tt 0 0

M3MOd3SMOH

I 0 n

I

::

M3d S M V l l O O 31VRIXOMddW

FIG. 17.17

Wound-rotor induction motors and control, 20 to 500 hp.

u

9 0

-~

FIG. 17.18 Synchronous motors and conlrol, x ) to 500 hp

Appendix

Compiled by D. B. Armstrong

CONVERSION FACTORS Acres Acres Acre-feet Amperes per square centimeter Amperes per square inch Ampere-turns Ampere-turns per centimeter Ampere-turns per inch Atmospheres Atmospheres Atmospheres Atmospheres

BY 43,560 1.562 X 10-5 43,560 6.452 0.1550 1.257 2.540 0.3937 76.0 29.92 33.90 14.70

To Oblain Square feet Square miles Cuhic feet Amperes per square inch Amperes per square centimetei Gilberts Ampere-turns per inch Ampere-turns per centimeter Centimeters of mercury Inches of mercury Feet of water Pounds per square inch

British British British Britiah

0.2520 778.2 3.931 X lo-' 2.930 X lo-'

Kilogram-calories Foot-pounds Horsepower-hours Kilowatthours

Ccntimcters Circular mils Circular mils Circular mils

0.3937 5.067 X 10-6 7.853 X 10-' 0.7854

Inches Square centimeters Square inches Square mils

Degrees (angle) Dynes

0.01745 2.248 X

Radians Pounds

Ergs

7.378

Farads Foot-pounds Foot-pounds Foot-pounds Foot-pounds

108 1.285 x 10-3 5.050 X lo-' 0.1383 3.766 X lo-'

Gallons per minute Gallons of water a t 62 F Gallons of water per minute

2.228 8.345 500.7

Mulliply

thermal thermal thermal thermal

units units units units

x

x

925

10-8

10-8

Foot-pounds Microfarads British thermal units Horsepower-hours Kilogram-meters Kilowatthours Cubic feet per second Pounds Pounds per hour

APPENDIX

926

Gausses Gilherts

6.452 0.7958

To Obtain Lines per square inch Ampere-turns

Henrys Horsepower Horsepower Horsepower Horsepower Horsepower Horsepower (boiler) Horsepower-hours Horsepower-hours

108 42.44 33,000 1.014 10.68 0.7457 33,475 2544 1.98 x 106

Millihenrys Btu per minute Foot-pounds per minute Horsepower (metric) Kilogram-calories per minute Kilowatts Btu per hour British thermal units Foot-pounds

Inches Inches of mercury Inches of mercury Inches of water Inches of water Inches of water

2.540 1.133 0.4912 25.40 0.5781 0.0361

Centimeters Feet of water Pounds per square inch Kilograms per square meter Ounces per square inch Pounds per square inch

Joules Joules

9.480 x 10-4 0.7377

British thermal units Foot-pounds

Kilograms Kilogram-calories Kilolines Kilometers Kilowatts Kilowatts Kilowatts Kilowatthours Kilowatthours

2.205 3.968 108 3,281 0.6214 56.88 737.8 1.341 3,413 2.656 X lo6

Pounds British thermal units Maxwells Feet Miles Btu per minute Foot-pounds per second Horsepower British thermal units Foot-pounds

1og.N or In N Log,,N Lumens per square foot

0.4343 2.303 1

l0gi.N 1og.N or In N Foot-candles

Maxwells Megalines MPgohms Meters Meter-kilograms Microfarads Microhms Microhms per centimeter cube Microhms per centimeter cube Miles Miles

10-2 10' 108 39.37 7.233 10-6 10-6 0.3937 6.015 5,280 1.609

Kilolines Maxwells Ohms Inches Pound-feet Farads Ohms Microhms per inch cube Ohms per mil foot Feet Kilometers

Multiply

yonreters

BY

APPENDIX

Multiply

P17

To Obtain

BY

lo-'

Ohms Ohms Ohms per mil foot Ohms per mil loot

106 0. I662 0.06524

Megohms Microhms Microhms per centimeter cube Microhms per inch cube

Radians

57.30

Degrees

Square centimetors Square feet Square fcet Square inches Square inches Square kilometen Square meters Square miles Square miles Square millimeters Square mils

1.973 X 10' 2.296 X 1 0 F 0.09290 1.273 X lo6 6.452 0.3861 10.76 640 2.590 1.973 x 108 1.273

Circular mils AWW Square meters Circular mils Square centimeters Square milcs Squarz fcet Acres Square kilometers Circular mils Circular mils

1 1.8 1 5/9 2240 2205 2000

Ahsolute temperature C Temperature F Ahsolute temperature F Temperature C Pounds Pounds Pounds

+ 273%) + 17.8) + 460)

Temperature (C Tcmperature (C Temperature (F Temperature (F Tons (long) Tons (metric) Tons (short)

32)

EQUIVALENT VALUES OF ELECTRICAL AND HEAT UNITS Eyuioalrnt Value in Othei l i i t s 1.341 horsepower-hour 1 kilowatthour 3413 British thermal units 0.7457 kilowatt-hour 1 horsepower-hour 2541.1 British thermal units 11.341 horsepower 44,254 footIponnds per minute 1 kilowatt 737.56 foot-pounds per second 3413 British thermal units D W hour 0.7457 kilowatt 33,000 foot-pounds per minut? 1 horsepower 550 loot-pounds per second 2544.1 British thermal units Der houi 1 j0"k = 1 watt-second 778.26 foot-pounds 1 British thermal unit' = . 1054.8 j o u l ~ sor wattbeconds of t h e quantity of heat rpquired * T h e British thermal unit (Btu) is defined as to raise t h e temperature of 1 Ib of water from 32 F t o 212 F at 1 atmosphere (14.696 psi).

limt

={

=I {

APPENDIX

928

Unit Degrees absolute (Kelvin) Dcgrees absolute (Rankine) Degrees centigrade Degrees Fahrenheit

Equivalent Value i n Other Units 273 460

= Degrees centigrade = Degrees Fahrenheit = 56 (deg F - 32) = 1.8 (deg C ) 32

+ +

+

NOMENCLATURE The following tables, compiled from ASA Standsrds, represent the letter symbols and abbreviations most commonly used in power system engineering. For a complete listing of letter symbols and abbreviations, refer to ASA Standards 210.5-1949, 210.4-1943, and 210.1-1941.

The Greek Alphabet Quantities Cmnmonly Designated by the Small Gmek Letter Name Resistance- temperature coefficient alpha beta Phase constant Conductivity gamma Increment * delta Dielectric constant epsilon zeta Efficiency eta Angular phase displacement theta iota Magnetic susceptibility kappa. Wavelength lambda Permeability mu Reluctivity nu xi omicron 3.1416 pi Resistivity rho Summation. Sigma Time constant tau upsilon Magnetic flux’ phi Angular phase displacement phi chi psi Flux of displacement* omega Angular velocity commonly designated by the large Greek letter.

Large Letter A

Smell Letter

B

B

r A

E 2

H

n Y

6

r 7l

e

e

I K

x

A

x

M N

I”

I

.j

€

0

0

n

r

P

P

z

m

T T

7

” 9 9

X Y

n

x. .

+ w

* These items are

letter Symbols Used in Electrical Formulas Quantily Admittance Angular phase displacement

Symbol (8,

Y 9) (theta, phi)

Zllvstratiue Unit

Mho Radian

APPENDIX

Quantity Capacitance Charge, or quantity of electricity Conductance Current Dielectric constant Efficiency Electromotive force Energy Flux density Frequency Frequency, angular velocity Frequency, rotational Impedance Inductance Magnetic flux Magnetizing force Magnetomotive force Number of conductors or turns Number of poles Permeability Permeanee Power ReactancQ,capacitive Reactance, inductive Reactance, subtransient Reactance, transient Reluctance Resistance Resistsnee-temperature coefficient Resistivity Rotative operator, 90’ Rotative operator, 120’ Temperature

929

Illusttotiue Unit

Symbol

c

Q G I

Farad Coulomb Mho Ampere Farad per inch Volt Joule Lines per square inch Cycle per second Radian per second Revolution per second Ohm Henry Line Ampere-turn per inch Ampere-turn

Line per ampere-turn per inch Line per ampere-turn Watt Ohm Ohm

Ohms

(a)(alpha)

(P) (rho)

Ohms Ampere-turn per line Ohm Decimal parts per degree C Ohm-in

i a

Degree C Degree K (Kelvin) Dearee Time Second Wavelength Meter Work W Joule NOTE: When distinctions between maximum, instantaneous, and r m s values are necessary, Emand I, are recommended for maximum values, e and i for instantaneous values, and E and I for rms values. * In the “Industrial Power Systems Handbook,” X without the subscript L is used to denote inductive reactance. t

T

Letter Symbols Used in H e a t and Thermodynamic Formulas Quonlity Efficiency Enthalpy Entropy Volume flow

Symbol

(eta) Hi h 8, 8 , 9

(7)

Q

lllvslralive Unit Per cent Btu per pound Btu per degree F per pound Cubic feet per second

APPEN~JIX

930

Quantity Heat flow Theoretical steam rate Actual steam rate Temperature, Fahrmheit Superheat Terminal difference Velocity Specific volume Pressure, gauge Pressure, ahsolute Gas constant Specific heat Energy

Illustrative Unit Btu per hour Pounds per kilowatthour Pounds per kilowatthour Degree F Degrcc F Degree F Feet per second Cubic Feet per pound Pounds per square inch gauge Pounds per square inch absolute

Symbol

Q TSR

SR 1 , FTT

FS td

V

"

psig psia R c

U

Btu per pound

Abbreviations Used in Electrical Formulas pf = power factor hp = horsepower r = 3.1416 va = volt-ampere kva = kilovolt-ampere mva = megavolt-ampere

kv var kvar w kw mw

line-to-line voltage in kilovolts reactive volt-ampere = kilovars, reactive kilovolt-ampere = watt = kilowatt = megawatt

=

=

DEVICE NUMBERS AND FUNCTIONS FOR SWITCHGEAR The devices used in switching equipment are referred to by standarddevicenumbem. These numbers are used on diagrams, in instruction books, and in specifications to permit ready understanding of the function and operation of switching equipment. The following list of device numbers and functions &retaken from American Standard C37.2-1945. ~~

,, '

Detrice No. Function and Definition 1 Master element .. 2 Time-delay starting or closing relay 3 (Reserved for future application) 4 Master cantactor or relay 5 Stopping device 6 Starting circuit breaker, contactor, or switch 7 Anode circuit breaker 8 Control power switch 9 Reversing device 10 Unit sequence switch 11 Control power transformer 12 Overspeed device Synehronous-speed device 13 14 Underspeed device 15 Speed-regulating device 16 Battery-charging control device Series-field shunting circuit breaker or contactor 17

APPENDIX

Deuiee No. 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50* 51 * 52

53 54

55 56 57 58 59 60 61 62 63' 64 65 66

Function and Definition

Accelcrating or decelerating eontactor, circuit breaker, or relay Starting to running transition eontactor or relay Electrically operated valve Impedance relay Equalizer circuit breaker or contactor Tcmpeetture-rPgulating device Bus-tic circuit breaker, contactor, or switch Synchronizing or paralleling device Apparatus thermal device A-C undervoltage relay Resistor thermal device Isolating circuit breaker, contactor, or switch Annunciator relay Separate excitation device D-C reverse-power relay Position switch Motor-operated sequence switch Brush-operating or slip-ring short-circuiting device Polarity device Undercurrent or underpower relay Eearing thermal device Field reducing contactor Field relay Field circuit breaker, contactor, or switch Running circuit breaker, contactor, or snitch Transfer device Unit sequence starting contactor or relay D-C overvoltage rday Reverse-phase, phase-balance current, or power-rectifier misfire- relay Single- or reverse-phase voltage relay Incomplete-sequence relay A-C thermal relay A-C instantaneous overcurrent relay A-C time-delay avercurrent relay A-C circuit breaker or contactor Exciter or generator relay High-speed circuit breaker Power-factor relay Field application relay or device (Reserved for future application) (Reserved for future application) A-C overvoltage relay Voltage balance relay Current balance relay Time-delay stopping or opening relay Fluid pressure, level, or flow relay Ground protection relay Governor Notching relay

Shows change from, or addition to, ASA-C37.21945.

931

APPENDIX

932

Function and Definition Device N o A-C power directional or a-c directional overcurrent relay 67. D-C thermal relay 68 Permissive control device 69 Electrically operated rheostat 70 D-C line emergency circuit breaker or contactor 71 D-C line circuit hresker or contactor 72 Load resistor circuit breaker or contactor 73 Alarm relay 74 Positian-changing mechanism 75 D-C overcurrent relay 76 Impulse transmitter 77 Phase-angle measuring relay 78 A-C rcclosing relay 79 D-C undervoltage relay 80 Frequency device 81 D-C reclosing relay 82 Selective control, or transfer, contactor or relay 83 Operating mechanism 84* Carrier or pilot-wire receiver relay 85 Locking-out relay 86 DiEerential current relay 87 Auxiliary motor or motor generator 88 Line switch 89 Rpgulating device 90 D-C voltage directional relay 91 D-C voltage and current directional relay 92 Field-changing contactor or relay 93 Tripping or trip-free relay or contactor 94

(Reserved for special application)

98 99

Shows change from, or addition to, ASA-Cj7.2-1945.

GRAPHICAL SYMBOLS FOR POWER-SYSTEM ONE-LINE DIAGRAMS One-line diagrams are very useful in showing, by means of graphical symbols and conventional nomenclature, an over-all power system arrangement. These symbols, when used consistently and in conformance with general practice, provide a valuable

tool. The following table, compiled from ASA Standards, represents the graphical symbols most commonly used in power-system one-line diagrams. For a complete list of graphical symbols for one-line diagrams, refer to ASA Standards Y32.1.1-1951and 232.3-1946.

933

APPENDIX

Equipment

Variations Definition

Arresters lightning

Arrester plus ground (mrge arrester)

Valve-type arrester Cable terminations

Single-conductor tern naticin Three+onductor termination

Capacitor

Capacitor plus ground (surge capacitor)

Capacitor hushing

Capacitor hushing patential device

Coupling capacitor potential device

Circuit breakers air

Breaker with drawout feature

Breaker with drawout feature, and operating coil

power

Breaker with drawout fcature

Breaker with diseonnecting switches

APPENDIX

934

Variations Equipment

Basic symbol Symbol

Contact

I T

Definition

L T

Vormally open (NO)

\Tormally closed (NC)

Coil operating

k. T

:ontact with blowout coil

Drswout mounting

Fuse

High-voltage primary fuse cutout, dry type; or fuse disconnecting switch lamc as above

High-voltage primary fuse cutout, oil type

Gap protective

-DQ+

:ap plus ground (surge gap)

Synchronous generato,

Generator

3epamtcly excited d-c generator' I

D-C shunt two-wire generator' * T h e broken a part of the sy

- - - indicates where line connection

symbol is made snd is not

935

APPENDIX

Equipmeni

Basic symbol

*

Variations

Symbol

Definition

I

Generator (Cr

D-C compound two-wire generator*

I

D-C three-wire shunt generator'

I -

Ground

Mechanical connection

____ dashes c necting equipme

ihort

Meters and instruments

0 (etteror letters shall he placed within the circle t o indicate the type of instrument:

A D F GD MA

Ammeter Demand meter Frequency meter Ground detector Milliammeter PF Power-factor meter R D Recording demand meter R E D Recording

* T h e broke a part of the s

R H Varhour meter S Synehroseope T Temperature V Voltmeter VA Voleammeter VAR Varmeter \I' Wattmeter WH Watthour meter

e - - - indicates where line connection to symbol is made and is not 01.

APPENDIX

936

Equipment

triations

Basic symbol

Definition

Symbol

Induction

Motor

Synchronous

Reactor (nonmagnetic core

+ ac

pawer

Relay

Q 0 'be relay device function number ihould be placed within the circle

APPENDIX

Equipment

P37

Variations

Basic symbol Symbol

Definition

Relay protective functions The following symbols arm used to indicate protective func r be placed adja tions and cent to th, .sic relay symbol

Over Under Directional (directional over)

Balance

Differential

-

Pilot wire

cc

Carrier current

Actuating quantity: The actuating quantity is indicated by the following letters, placed either on or above the relay protective function symbol shown above: C Current' Z Distance V Voltage W Power F Frequency

+

Phase T Temperature GP Gas pressure S Synchronism

*Generally accepted practice ia to omit any designation for currentactuated devices.

APPENDIX

938 ~

riations Equipment

Relsy protectivi functions (Cont'd)

Basic symbol

Definition

Ivereurrent

bvervoltage Xreetional overcurrent )irectional residual overcurrent

Jndervoltage 'ower directional

3slanee current

Differential current

Distance Directional distance

Over frequency

Under frequency

Ovcr temperature

Phase balance

Phase rotation

APPENDIX

Equipment

Basic symbol

P39

Variations Definition

Relay proteeti, functions (Cont'd)

Pilot wire, differential curent Pilot wire, directional eomparison

Carrier pilot Positive phase sequence undervoltage Negative phase sequence overcurrent Gas-pressure relay

Resistor

Switch air break

Thermal elemen

-

o u t of step

Grounded

2Double throw

-

Switch with horn gap

APPEHDLX

940

Variations Equipment

Transformer

Basic symbol

Definition

Two-winding transformer with taps Adjustable mutual inductor. constant-current transformer Threewinding transformer

Autotransformer

Potential transformer

Current transformer

Bushing-type current transformer

step-voltage regulator or load-ratio-control autotransformer

Load-ratio-contml transformer with taps

941

APPNDIX

Variations Equipment

Basic symbol Symbol

Trsnsformer (Cont’d)

@

Definition

Single-phase induction voltage regulator

Three-phase induction voltage regulator

Transformer winding connections rhe following symbols are used to indicate transformer winding connections and may symbol:

A

Three-phw three-wire delta

P -

Three-phase three-wire delta grounded

f -

Three-phase four-wire delta grounded

A

Three-phase Y

$

Three-phase Y grounded neutral

APPENDIX

942

riatibns Equipment

Basic symbol Definition

Transformer winding connections (Cont’d)

rhree-phase zigzag

rhree-phase zigzag grounded six-phase star (or diametrical)

Six-phase Star with grounded neutral

Three-phase open delta Three-phase open delta grounded at common point

1.06

1.12

% .440

241

130

210 252 300 371

l.w

1.05

1

1.07 1.00

185

180

203

172 194 213 253 301 367

161 181 199 236

0.85 0.80

0.90

0.95

1.04 1.w

, t Temperature

281 a42

r

0.88 0.83

1.04

.... ....

378

__

0.87

1.04

213 251 297 355

159

221 261 313

136 154 174

141

125 141

151

107 121

86

111

83 I34

79

83

125

101 115

114

66

~

51 67

...

Synthetic rubber. temp 80 c. 8.W15,WO Volts ahieldedt

shielded3 49 64

ihialded

-

Synthetic rubber. tsmp 85 c. 0-8000 volt.

108 119

77 99 111

60

68

87

46 60

46

rubber. tamp 75 c. )-Bw "Oilof

Synthetic

52

Correction Factor8 foi Varioua Am1

408

::&

.... 0.93 0.84 0.92 ,... 122 0.71 0.85 0.88 0.83 National Electrical Code ratiom (eorreoted for 40 C). t R a t i n 0 not $Overed by induatry stand*rds: d u e s liated are t h w e used by the General Electrio Company. Inavlated Power Cable Eoglnears ALYIODIB~~OO (IPCEA) ratings.

1.00

480

328 373

750 1OW

%

176

i ! 418

213

218

I)

rubger or "81nished-cambrio temp 77 C. 15.0W volts

I 1 I

118

46 58 76

shzgdt Shielded?

& ::$ :$ ;

Butyi synthetic rubber or varnished cambric.

224

114

101

40 57 75

Synthetic rubber, temp 75 c. 0-600volta*

262

160

135

103 119

33 45 57 78 90

Rubber 01 thermoplasti temp 60 C. 0-6wvolts'

500

01

I N CONDUIT Three single-oonductor cables or one three-conduotor cable per Conduit, 104 F (40 C) ambient air temperature

250 350

0 00 000 0000

1

6 4 2

8

Siie. Awa MCM

SIMPLIFIED SELECTION O F CONDUCTOR SIZE

Current-carrying Capsoitis in Amperes for Cable in Conduit. in Underarovnd Ducta, and in Intarlooked Armor

0.84

1.05

1.11

1.25

I85 213 244

161

138

52 68 81 121

f "

x

3P

%

v44

APPENDIX

NATURAL TRIGONOMETRIC TABLES

-

-

And