Transients Due to Switching of 400 kV Shunt Reactor

Ivo Uglešić,Sandra Hutter

University of Zagreb10000 Zagreb, Croatia

ivo.uglesic@fer.hr

Miroslav KrepelaSiemens

10000 Zagreb, Croatiamiroslav.krepela@siemens.hr

Božidar Filipović- GrčićCroatian National Electricity

10000 Zagreb, Croatiabozidar.filipovic-grcic@hep.hr

Franc JaklUniversity of Maribor

2000 Maribor, Sloveniafranc.jakl@uni-mb.si

Abstract - Transients phenomena originating fromswitching of high voltage shunt reactor with solidlyearthed neutral are considered. Two reactors (100MVAr and 150 MVAr) were compared on the basis ofresults obtained from calculations of the transients.The unsymmetrical phase inrush currents during ashunt reactor energizing and overvoltages provoked bycurrent chopping and reignition of electrical arc areanalyzed. Successful synchronous switching can reducethe mechanical and electromagnetic stresses enduredduring normal switching operations.

Keywords: Shunt reactor, Switching operations, Inrushcurrents, Chopping Overvoltages, Recovery voltage,Reignition, Controled switching.

I. INTRODUCTION

High voltage shunt reactors switching belongs to thenormal system operations that can be performed severaltimes a day.

Because of the shunt reactor technical characteristicsand their special purpose, its current is mainly inductiveand can be referred to as a small inductive current. It isconsiderably smaller (10 or 20 times) than nominalcurrents of today most commonly used SF6 circuitbreakers, and even up to 200 times smaller than theexpected short-circuit currents. The preliminary analysis of the switching operation withthe first 400 kV shunt reactor designed for the Croatiantransmission system is conducted. A high voltage reactor isrelatively frequently switched: during the periods of thesystem operations with low loads it is energized and withthe rise of load it is de-energized again. Shunt reactorsswitching operations result in electromagnetic transientsand some mechanical effects. At closing, high,unsymmetrical currents with long time constant can occur.During opening, interruption of small inductive currents(current chopping) will cause overvoltages. Reignitionbetween the circuit-breaker contact gap can lead to verysteep overvoltages that represents a special problem.

The above mentioned transient phenomena are themain concern of this paper.

II. ENERGIZING OF HV SHUNT REACTORS

Unsymmetrical phase inrush currents occuring at HVshunt reactor energizing depend on the time instant ofcircuit breaker pole operation with respect to the referencesignal. Switching operations at unfavorable instants cancause currents that may reach high magnitudes and have

long time constants. During HV shunt reactor energizing,mechanical vibrations and buzzing of transformers wereobserved. At shunt reactors with solidly grounded neutral,unsymmetrical currents cause zero-sequence current flowwhich can activate zero-sequence current relays. Current asymmetry becomes smaller if the circuitbreaker closes near the maximal value of the powerfrequency voltage. On the other hand this causes switchingovervoltages. Dielectric stress of reactor insulation causedby the very steep overvoltages originated from circuitbreaker contact gap prestrike immediately before closing issimilar to that caused by the reignition after reactor de-energizing with voltage across the circuit-breakerinterrupter of 1 p.u. However, it is impossible to obtainreduction of inrush currents and stresses originated fromthe overvoltages during switching operations at the sametime. It remains on the user to find the most acceptablesolution. A high voltage shunt reactor with the associatedequipment is planned to be installed in the 400 kVswitchyard, as shown in the simplified single-line diagram,Fig. 1., while in Fig. 2. the corresponding computer modelscheme is presented. The inrush currents at energizing of 100 MVAr and 150MVAr HV shunt reactors have been observed for twocases: switching at a random small value and near the peakvalue of the phase voltage. Parameters for the reactor with rated power 100 MVArwere L = 5.093 H/phase, C = 2 nF/phase and for thereactor with rated power 150 MVAr, L=3.395 H/phase, C= 2.6 nF/phase. The high voltage 400 kV circuit breakerconsists of two breaking chambers. Across each a potentialgrading capacitor of C = 500 pF is connected. Inrush currents can be expressed in the p.u., where 1 p.u.equals I Max, which obtains following values, depending onthe HV shunt reactor size:

- 100 MVAr, AII nMax 2052 == (1)

- 150 MVAr, AII nMax 3072 == (2)

Fig.1. Single-line diagram of the 400 kV switchyard.

Fig.2. Equivalent circuit for 100 MVAr reactor, circuitbreaker and network.

Maximal inrush currents obtained when energizing HVshunt reactor at a random small value of the phase voltagewere:- 100 MVAr, ..09.2 upIMax = p.u.0.975I0max = t = 1.3 s- 150 MVAr, ..05.2 upI Max = ..970.0max0 upI = t = 1.0 swhere t represents approximate duration of theunsymmetrical currents. The maximal inrush currents obtained when energizingHV shunt reactor near the peak value of the phase voltagewere:-100 MVAr, ..57.1 upI Max = ..527.0max0 upI = , t = 0.75 s-150 MVAr, ..49.1 upI Max = ..515.0max0 upI = , t = 0.6 s.

Fig.3. Initial inrush currents at energizing of 150 MVArshunt reactor.

Fig.4. Zero-sequence current at energizing of 150 MVArshunt reactor.

The closing time instants of the circuit-breaker contactsinfluence values of the switching currents and the durationof the unsymmetrical currents. Controlled closing of thecircuit breaker has a significant influence on the switchingcurrents magnitudes and on the current asymmetry.

III. DE-ENERGIZING OF HV SHUNT REACTOR

When interrupting small inductive currents, the mediumused for arc extinguishing will develop fast increase of theresidual column resistance, and abrupt current interruptionbefore its natural zero crossing occurs. Release of energystored in the reactor inductance will cause theelectromagnetic transients that lead to the switchingovervoltages. The switching overvoltage can be dangerous for theequipment if their peak value surpasses the rated switchingimpulse withstand voltage of the equipment. Beforemaking a decision about the choice of equipment or thetechnical solutions of the facilities, it is very important toknow the level of dielectric stress that can occur duringoperation in the system in order to avoid insulationfailures. The electric arc in the high voltage 400kV circuitbreaker was simulated with its conductance-dependentparameters. The method is based on the Cassie and Mayr-arc-equation:

( ) ))(

(1 2g

gPi

gdtdg s −

Θ= (3)

where g is the arc conductance, is is the switching current,P is the arc cooling power and Θ is the arc thermal timeconstant. Parameters P and Θ are conductance dependent.P is defined for two rates that correspond to the ignitionand burning phase:

- Ignition: ( )MWgPi748.023= (4)

- Burning: ( )MWgPb5.2387= (5)

In calculations a constant value of Θ=1µs was used.Parameters Pi, Pb and Θ are selected as recomended in [1]. Switching overvoltages that can occur during normalreactor switching operations place additional demands onthe dielectric strength of the equipment.

A. Chopping Overvoltages

If the current is chopped before its natural zero crossing,in the reactor remanent accumulated energy has to bereleased through the oscillations in the L-C circuit. Theovervoltage is caused by the release of magnetic energystored in the reactor inductance at the moment of currentchopping. Chopping overvoltages during de-energizing of reactorswith nominal powers of 100 MVAr and 150 MVAr havebeen analyzed. Field tests on the existing 400 kV switchyards withsimilar reactors and circuit-breakers [2], have shown that

the values of the chopping current will typically rangefrom 2 to 14 A. Fig.5. shows voltage oscillograms after currentchopping for the 100 MVAr reactor. Capacitive andinductive interphase coupling can be observed. In theforegoing interrupted phase, there will be an increase inoscillations after the second and third pole clearance.However they do not exceed the value of the firstmaximum. Voltage in the phase which was lastinterrupted, decreases with an exponential factor. Theoscillation frequency was approximately 1500 Hz for 100MVAr reactor and around 1600 Hz for the 150 MVArreactor. Transient recovery voltage will be established across thecircuit breaker contacts after current interruption. Itsmaximal expected value is equal, if damping is neglected,to the sum of the peak voltages of the source andovervoltage provoked by current chopping. For the 100 MVAr reactor the highest suppresion peakovervoltage factor was 2.05 p.u. but only when surgearrester was not connected. With connected surge arresterthis overvoltage factor never exceeded 1.5 p.u.Overvoltage factors in the cases with 150 MVAr reactorwere lower, as it would be expected. Without surgearrester the overvoltage factor was 1.67 p.u., and withconnected surge arrester 1.47 p.u. In all cases switching overvoltages were lower than thepeak value of the switching impulse withstand voltage ofthe isolation (1050 kV, wave shape 250/2500 µs). As aresult of the low oscillation frequency on the reactor side(around 1.5 kHz), the overvoltage slope is gentle and isequally distributed along the winding, what results in smallpotential differences between turns. It is suggested [3] that maximal switching overvoltagesbetween the circuit-breaker contacts generated by currentchopping should not exceed 80% of the peak value of theswitching impulse withstand voltage which is 3.63 p.u.(900 kV+345 kV).

With connected surge arrester computed ovevoltages inall cases were below recommended values. Maximalrecovery voltage on the circuit breaker contacts waskr=2.5 p.u (69% of the rated peak value).

Fig.5. Chopping overvoltages on the reactor 100 MVAr.

B. Reignition Overvoltages

Reignition overvoltages are generated by the reignitionfollowing the initial interruption and arc extinction.Reignitions are provoked when the recovery voltage acrossthe circuit breaker contact gap exceeds the dielectricwithstand of the residual column.

Reignition can be usually expected near the peak valueof the recovery voltage. In that moment the voltagedifference across the circuit breaker is around 2 p.u. Ifreignition occurs at that moment, the voltage across thereactor will very rapidly assume the value of instantaneousvoltage on the network side. The transient voltage acrossreactor can surpass 1 p.u., and because of its oscillatingnature can assume the opposite polarity with a valuebetween 2 p.u. and 3 p.u. For an analysis of dielectric stresses of the reactorinsulation, besides overvoltages to the ground, thedifferences in voltages between the two consecutive peakvalues of the opposite polarity after reignition need to beconsidered.

Several reignitions may occur before the final arc-extinction. This phenomenon is denoted as voltageescalation, and can occur when interrupting currents withvery steep slopes at the zero crossings. In praxis it wasobserved that in some cases reignition lead to circuitbreaker interrupter damage (nozzle, contacts). For thisreason, and since it may also affect shunt reactorinsulation, it is desirable to eliminate reignitionoccurrence.

The occurrence of reignition depends on the shuntreactor current to be switched. The lower the inductivecurrent, the higher possibility of reignitions.

Overvoltages caused by reignitions depend greatly onthe network configuration and on the circuit-breakercharacteristics. They expose the shunt reactor and otherdevice insulation to stresses similar to those of lightningovervoltages. In computer simulation elements of the substation aremodeled with their surge impedance and correspondinglength, similar to the study of lightning overvoltages.Calculations were conducted with and without connectedsurge arresters, to observe their protective influence. Reignition was observed for the 100 MVAr and 150MVAr reactors under the same conditions. In Fig.6. thevoltage oscillogram obtained during reignition after reactorde-energizing, without MO surge arresters is shown. Themaximal overvoltage across reactor evolves immediatelyafter reignition. Transients are strongly damped andoscillation frequency on around 400 kHz can be noticed.The voltage slope is very steep as the change between thetwo, consecutive peak values of the opposite polarity lastsless than 1µs. Calculation results are given in the Table 1. The highestovervoltage of 703 kV appears, as it could be expected,when the 100 MVAr reactor is de-energized withoutconnected surge arresters. With connected surge arrestersvoltage across reactor to the ground is lowered on 655 kV.The voltage rate of rise (kV/µs) between the positive andnegative peak value was approximately thesame in bothcases.

Fig.6 Reignition after 100 MVAr reactor de-energizing,without MO surge arresters connected at the reactor.

Table 1. Overvoltages caused by reignition

Reactor Surgearrester

UMAX ( )kV Sav ( )skV µ/

No 703.5 1322100MVAr Yes 654.8 1320

No 666.6 1159150MVAr Yes 632.7 1155

When reactors were protected with surge arrestersovervoltages to the ground were never critical. In all casesthe voltage rate of rise on the reactor was very steep. The protective influence of the surge arresters can beobserved in limiting overvoltages caused by currentchopping. In this way reignition is indirectly avoided.However, the surge arresters have practically no influenceon the voltage steepness. It can be seen that the surge arresters can limit peakvalues of overvoltages, caused by current chopping andreignitions to an acceptable level that will not causeinsulation damage. In all cases overvoltages caused byreignition had smaller peak value than the nominallightning impulse withstand voltage Uw =1425 kV, withwave shape 1.2/50 µs. For the reactor insulation, however, the voltage rate ofrise is more dangerous than its peak value. For transientscaused by reignition the period of the wave crest usuallylasts from one to couple of microseconds. Very steepovervoltages are distributed unevenly along the reactorwinding and the first few turns are exposed to the largestpotential differences. Overvoltage slopes duringreignitions, critical for insulation, are smaller for the 150MVAr reactor. At reignition, the electric arc resistance decays veryrapidly and in a few nanoseconds drops to a couple ofOhms. Basically, such a fast change has no influence onthe voltage slope, although it slightly reduces overvoltagevalue. The acceptable voltage rate of rise can be definedfrom the slope of the standard 1.2/50µs wave. The averagesteepness of the nominal lightning impulse withstandvoltage, with peak value of 1425 kV, is determined in theinterval between 30% and 90% of the crest value. Duringthat interval (t=0.72μs), wave can be approximated with astraight line.

Average steepness is:

skV

skVSav µµ

118772.0

1425)3.09.0( =⋅−= (8)

To avoid accelerated wearing of the insulation, safetyfactor of 0.65 is recommended [4], which corresponds tothe crest value of the standard wave limited to 65 %. Inthat case average rate of change becomes:

skV

avSav

Sµ

77265.0' =⋅= (9)

The initial part of the standard wave limited to 65% canbe described with the following function:

( )tkVtu ωcos1463)( −= (10)

with a half-cycle T/2 = 1.2 µs, frequency of 417 kHz andcrest value of 926 kV. The maximall instantenuous rate of change determinedfor the same case may be obtained from:

( )s

kVdt

tduSµ

1211max == (11)

It is assumed [4] that insulation can withstand such anovervoltage rate of rise applied twice a day. In the computer simulations conducted, the maximumovervoltage slopes were steeper or close to recommendedvalues and as such represent the danger for the reactorinsulation, since a high frequency transient is unevenlydistributed along the reactor winding, stressing theincoming turns with high interturn voltages. It has to beconsidered that the reactor switching operations can beperformed several times a day. Frequent exposure to thoseovervoltages causes aging and wearing of the insulationand the application of measures for their limitation isrecommendable. The internal reactor insulation is class A(paper in oil), that is not self-renewable. Even the testingvoltages that do not cause damages can influence theinsulation characteristics. Because of the testinglimitations, there are no statistical data about the realwithstanding voltage level for the equipment.

IV. SHUNT REACTOR CONTROLED SWITCHING

All circuit breakers exhibit a high probability ofreignition for arcing times shorter than Ta min (maximalarcing time at which reignition is still probable). Providedthe arcing time can be controlled such that it exceedsTa min, the probability of reignition is practically negligible.

Small inductive current interruption is similar to that ofno load and may be approximated by the cold gasdielectric recovery characteristic [5]. The increase ofpower frequency (50 Hz) withstand voltage as a functionof time and contact gap for cold SF6 gas can be, basedupon experiments on a 400 kV circuit breaker,approximated by equation [6]:

( ) 1.77.201010 23 ++×= − dddUd (12)

with d = v × t, where Ud is the circuit breaker contact gapwithstand voltage, d is the contact gap at a time instant t(ms) and v is the contact velocity at opening (9 m/s for thestudy case circuit breaker).

Dielectric recovery characteristic for the cold gas isshown in Fig.7.

0,00

500,00

1000,00

1500,00

2000,00

2500,00

3000,00

1 2 3 4 5 6 7 8 9 10 11 12 13 14t (ms)

U (kV)

Ud = f (ta)

Ur (100 M VAr)

Ur (150 M VAr)

Fig .7. Dielectric recovery charact. for the cold gas

In the same figure, recovery voltage (Ur) characteristicsas a function of arc duration, i.e. chopping number λ (inthe range 5 ms < Ta < 14 ms), for both study case shuntreactors (100 MVAr and 150 MVAr) are shown. It can beseen that the probability of reignition is inverselyproportional to the arcing time Ta and shunt reactor ratedpower.

For the study case circuit breaker, chopping number λ,based on the experimental records in accordance with IEC61233, can be approximated by the equation:

FATamean /10)922.0( 4×+=λ (13)

within the range of arcing times 5 ms < Ta < 14 ms, andwith the standard deviation s=0.8×104.

A principle of synchronous trip operation is shown inFig. 8.

Fig. 8. The principle of synchronous trip operation

The circuit breaker switching command is a randomevent with respect to the reference signal, at a time instanttcommand. As soon as the command is received, thesynchronous control device is blocked and a time delay is

initiated awaiting the first voltage zero crossing. Afterinitial checking and calculations, the controller triggers offcount-down of intentional synchronizing delay (Ty)determined by the opening time (Topen) and a targetedcontact separation instant with respect to current zerocrossing at which extinguishing occurs. This delay can be,in simple terms, expressed by the equation:

openarczeroy TTTNT −−×= (14)

where N is an integer number of half cycles, Tzero is halfcycle duration calculated immediately before synchronousdelay calculation, Tarc is an arcing time and Topen is amechanical opening time. The calculation is performed foreach phase separately, taking into account initial valuesmeasured at site at the circuit breaker commissioning andvalues measured immediately before delay calculation.

Similar considerations are valid for the shunt reactorclose operation.

In accordance with the study case circuit breakerexperimental record [2], the longest arcing time at whichreignition still occurred was Ta min ≈ 7.5 ms. Consequently,the synchronous switching controller shall be set in a wayto enable circuit breaker contact separation at the timeinstant tseparate = 8 ms prior to the natural current zerocrossing (Tarc = 8 ms), when reignition occurrence isimprobable.

Assuming the power frequency equals 50 Hz and initialparameters for the study case circuit breaker: Topen = 25ms; Tarc = 8 ms, a simplified calculation gives thesynchronous trip delay: Ty = 7 ms.

Chopping number for Tarc = 8 ms equals toλmax = 12.36×104 A/√F. The magnitude of the suppressionpeak overvoltage for worse case (shunt reactor 100 MVAr)will be approximately equal to:

kVupQ

Nka 5.538..57.12

312max

max==+=

ωλ (15)

and the magnitude of recovery voltage across the circuitbreaker:

kVupkk ar 5.881..57.21maxmax

==+= (16)

Both values, even if MO surge arresters were not takeninto consideration, are below the allowable values.

The reignition and voltage escalation will not occursince the synchronous switching controller allows thecontact separation at the time instant for which reignitionis improbable after natural current zero crossing.

Regarding circuit breaker close operation, contact gappre-strike will occur when the voltage magnitude reachesthe peak value, with arc duration up to 5 ms (1/4 cycle)prior to contact touch. With such closing, inrush current ismaximally reduced. Since it is impossible to achievesimultaneous reduction of inrush current and insulationstress due to steep voltage transient at prestrike, thecompromise solution is to be reached.

N x T z e r o

T y T o p e n

t c o m m a n d

T r ip in p u t

T r ip o u t p u t

B r e a k e r c o n t a c t sT a r c

Experiments have shown [7] that at synchronousswitching of circuit breaker with independent poles aphase inrush current can be limited to less than 1.5 p.u.with the zero-sequence current lower than 0.5 p.u. Thus,mechanical stresses of the shunt reactor and adjacenttransformers are reduced and false operation of zero-sequence current protection prevented.

V. CONCLUSION

Switching operations of shunt reactor are relativelyfrequent and primarily depend on power network loading.With regard to its inductive character, switching of shuntreactor rated current results, without exceptions, intransients that can jeopardize insulation of shunt reactoritself and other switchyard elements, and createmechanical stresses.

Conducted computer simulation, based on realswitchyard, shunt reactor and other apparatus technicaldata, has suggested the following conclusions:

− switching transients are inversly proportional to theshunt reactor rated power;

− chopping overvoltages at de-energizing, withoutreignition, never exceeded the insulation switchingimpulse withstand voltage (1050 kV, 250/2500 µs),since MO surge arresters provide low voltage level;

− dependant on contact separation instant with respect tothe natural current zero crossing, after currentchopping, the reignition voltages that developed hadlower peak value than insulation lightning impulsewithstand voltage (1425 kV, 1.2/50 µs);

− calculated reignition transients steepnesses werebigger or close to recommended values (eq. 9&11)and represent danger for reactor insulation;

− surge arresters have practically no influence onovervoltage steepness;

− shunt reactor energizing may produce eithermechanical stress (inrush current) or dielectric stress(prestrike);

− since frequent exposure of shunt reactor insulation totransients, especially steep reignition overvoltagesaccelerates aging and wearing, means for theirlimitation are recommended;

− shunt reactor controlled switching can influence(reduce) inrush current magnitude or prestrikeovervoltage at energising and decrease or completelyeliminate the probability of circuit breaker reignitionat de-energising.

VI. REFERENCES [1] R.Sun, “Transiente Überspannungen in SF6-isolirter

metallgekapselten Schaltanlagen” (“TransientOvervoltages in the SF6 -isolated Metal EnclosedSwitchgear”) Doctoral thesis, University of Stuttgart,Juni 1991.

[2] Bachiller J., Borras F., Degen W., Habedank U., Trapp N., “Switching of Shunt Reactors - Theoretical and Practical Determination of High-Voltage Circuit- Breaker Behaviour”, Proceedings of the CIGRÉ SC 13 Colloquium, paper 13-95, September 1995, Florianopolis (Brazil). [3] IEC1233 - Technical Report Type 2: "High-Voltage

Alternating Current Circuit-Breaker- Inductive LoadSwitching", First Edition, 1994.

[4] S.A.Morais, “Considerations on the Specification of Circuit-Breakers Intended to Interrupt Small Inductive Currents”, Electra, No. 147, April 1993, pp. 45-69.

[5] Z.Ma, C.A.Bliss, A.R.Penfold, A.W.Harris,S.B.Tennakoon, "An Investigation of TransientOvervoltage Generation when Switching High VoltageShunt Reactors by SF6 Circuit Breaker", IEEETransaction on Power Delivery, vol.13, No.2, April1998.

[6] A.Miliša, K.Meštrović, “Breaking capacity of HV SF6 circuit breaker at back to back switching”, Proceedings of the 1st Croatian CIGRÉ committee session, October 1993, Zagreb. [7] CIGRÉ WG 13.07, “Controlled switching of HVAC circuit breakers – guide for application lines, reactors, capacitors, transformers (1st and 2nd Part)”, Electra, No’s. 183, April 1999, pp. 43-73 and 185, August 1999, pp. 37-57.