IEEE Power Engineering Society

Transmission and DistributionCommittee

IEEE Transactions on Power Apparatus and Systems, Vol. PAS-104, No. 6, June 1985

SINGLE PHAS SWITCHING TESTS CN 765 kV AND 750 kV TRANSMISSICN LINES

H. N. Schnerer, Jr., Fellow, NEB. R. Stperling, Senior NimberJ. W. Chadwick, Jr., Fellow,,

USA

Abstract - Single.. phase switchingcompensation schemes for transposed anduntransposed lines were tested during stagedfault tests on the 243 km, 765 kVKammer-Marysville line (USA) and the 417 km,750 kV Vinnitsa-Dnieper line (USSR).Conventional four-legged reactors were usedon both lines, and in addition a switchedfour-legged reactor bank was utilized on'theuntransposed Kammer-Marysville line. Duringthe tests, a large -range of single phaseswitching parameters was simulated, whichintentionally varied the secondary arccurrent and the recovery voltage on theopened phase. The power frequency componentof the secondary arc current was varied from3 A r.ms to 80 A rms. The upper limit of therecovery voltage rate of rise reached about.15 kV/ms.. The tests definitely proved theapplicability of high-speed .,single-phasereclosing on shunt-reactor-compensated EHVlines.

INTRODUCT ION

Single-phase-to-ground faults account for70% to 95% of all faults on EHV powertransmission lines, some 70% to 90% of thesefaults are arcing 'faults that can be clearedby proper single phase switching (SPS)techniques without interrupting the remainingtwo phases. This makes SPS an efficientmethod of enhancing the reliability of powersupplied over EHV lines. In addition, theuse of SPS prevents certain kinds ofovervoltages and reduces others, whencompared to the overvoltages produced by athree-phase opening and a three-phasereclosing.

The secondary arc extinction time, andconsequently the required SPS dead time,increases with the secondary arc current andwith the recovery voltage on the open phasefollowing arc interruption. Higher systemvoltages and longer EHV transmission linesresult in increased secondary arc currents.In most cases, the addition of shunt reactorsto a given line raises the recovery voltagebut may reduce its rate of rise.

To reduce the secondary arc current onEHV transmission lines, the use offour-legged shunt reactors was suggested by

84 SM 711-8 A paper recommended and approvedby the IEEE Transmission and Distribution Committeeof the IEEE Power Engineering Society for presenta-

tion at the IEEE/PES 1984 Summer Meeting, Seattle,Washington, July 15 - 20, 1984. Manuscript submit-

ted August 30, 1983; made available for printingMay 15, 1984.

N. N. BelyakovV. S. Rashkes

K. V. Khoetsian

USSR

N. KnudsenEl] and E. Kimbark[2) andimplemented for the first time by TVA on a500 kV line[3). A similar four-iinductancearrangement was initially suggested: b"y W.Petersen[4] to suppress fault current insystems with an ungrounded neutral.

In the USSR, 750 kV lines are transposedand therefore the electrical parameters oftheir individual phases are approximately-equal. This makes the direct use offour-legged reactors applicable. In the USA,765 kV lines are untransposed and thereforehave considerably dif'ferent interphaseparameters. To effectively limit thesecondary arc current on these lines, specialswitching schemes were developed tosupplement the conventional four-leggedreactor[5,6].

At present, the secondary arc extinctiontime is determined mainly on the basis ofexperimental data. Because utilities andresearch organizations are interested inminimizing the SPS dead time, joint USA-USSRSPS tests were carried out in 1979 on EHVtransmission lines equipped with four-leggedreactors. The American Electric PowerCompany conducted SPS tests on the 765 kV,243 km Kammer-Marysville transmission line.The Soviet Electric Power Research Institutecarried out SPS simulated fault tests on the750 kV, 417 km Vinnitsa-Dnieper transmissionline.

TEST DESCRIPTION

Simplified diagrams of the test lines Prepresented. in Fig. 1. The 60 Hz self cindmutual impedances (Z) and admittances (Y) ofthe untransposed Kammer-Marysville (K-M) lineare given below:

Zaa=Zcc=0.113+jO.561 Ohm/kmZbb=0.115+jO.557 Ohm/kmZab=Zbc=0 .098+jO.238 Ohm/kmZac=0 .096+jO .187 Ohm/kmYaa=Ycc=j4. 39.10-6 Siemens/kmYbb=j4.51 10-6 Siemens/kmYab=Ybc=jO .72 .10-6 Siemens/kmYac=jO .l9 10-6 Siemens/km

The 50 Hz equivalent positive and zerosequence impedances and admittances of thetransposed Vinnitsa-Dnieper (V-D) line are as

follows:

Zl=0.02+jO.29 Ohm/kmZO=0.17+jO.62 Ohm/kmYl=j4.15-10-6 Siemens/kmYO=j3.15 10-6 Siemens/km

Main parameters of the conventionalapparatus used in the SPS tests are given in

Fig. 1. The parameters of the neutralreactors are presented in more detail inTable 1.

0018-9510/85/0006-1537$01.00 1985 IEEE

1537

Table 1 Neutral Reactor Parameters

(b)

Fig. 1 Diagrams of the test lines

Kammer-Marysville Line (a)

Rl, R2RN1RN2RSANS1-NS4SAl-SA4

- Shunt reactors, 300 MVAr- Neutral reactor, 300 Ohms- Neutral reactor, 800 Ohms- Reactor Switch, 765 kV- Vacuum neutral switches- Zinc-oxide surge arresters

Vinnitsa-Dnieper line (b)

Rl-R3

RN1,RN2 -

DSSG1,SG2 -

Shunt reactors, 300 MVArNeutral reactors, 310 OhmsDisconnect switch, 750 kVStabilized sparkover gaps

*Substations where fault applicationstructures were installed

The necessary insulation levels of theneutral reactors in both networks weredetermined by temporary overvoltages acrossthese reactors during the SPS cycle and inaccordance with the insulation ofcorresponding shunt reactor neutrals.Switching overvoltages across the neutralreactors in the conventional four-leggedreactor banks during line breaker operationsand in the switched four-legged reactorsduring vacuum switch openings, were limi'tedwith the help of zinc-oxide arresters (K-Mline) and stabilized sparkover air gaps orvalve-type arresters (V-D line).

Single phase switching parameters,namely, the secondary arc current and theopen phase recovery voltage after arcextinction, were varied from test to test bychanging the reactances of neutral reactorsandi also, the number of shunt reactor banksconnected to a test line. To a lesserextent, the secondary arc current and the

Reactors KalmerIMarZsville Vinnitsa, Dnieper

Type Air core Steel core withan air gap

Continuous 20 15currentA rms

Short term 250* 200**currentA rms

Withstand 130*** 65**** 85***voltagekV rms

Impedance 800 300 310Ohm__ _ _ _

Taps 200 150 220Ohm 400 235

600 255275

* 2 sec rating** 3 sec rating*** 1 min rating**** 30 sec rating

recovery voltage varied with changes in powerflow as well as with fault location. In theK-M tests the calculated 60 Hz value of thesecondary arc current, Is , ranged from 8 Arms to 80 A rms and the calculated value ofthe recovery voltage, Vr , on the openedphase varied from 0.08 p.u. to about 1.5p.u. The respective calculated parametersfor the V-D tests ranged from 3 A rms to- 49 Arms and from 0.07 p.u. to about 1.3 p.u. Thearc resistance in the steady-stAtecalculations was assumed constant and equalto 10 Ohms since its influence on Is wasfound negligible for resistances smaller than200-500 Ohms, depending on the scheme. Thevalues of Vr larger than 1.1 p.u. werecalculated taking into account the effect ofcorona losses.

TEST PROCEDURES AND MEASUREMENT TECHNIQUES

Ini order to minimize system distrbancesand facilitate test procedures, the majorityof tests in both the K-M and V-D series, wereconducted with real or simulated faultsapplied to a line open at one end. Three K-Mtests were performed with the line carryingup to 950 MW. With zero power flow along aline, the secondary arc current consistsessentially of an electrostatic componentonly. Absence of the electromagnetic currentcomponent in a given scheme results in thereduction of the secondary arc current whencompared to the current under load transferconditions. This reduction can be partiallycompensated by an increased electrostaticcomponent corresponding to a lowercompensation of interphase capacitances. Forexample, on the Vinnitsa-Dnieper line withtwo four-legged shunt reactors (neutralreactors of 250 Ohms each) the calculated 50Hz secondary arc current was 22 A rms for a

1538

MARYSVILLE.* KAMMER

15391500 MW power flow prior to the fault. If movement was started by a spring loadedone end of the line is open and one neutral mechanism. The rod was used to swing areactor is shorted out, the secondary arc stainless steel wire attached to the groundcurrent reaches 26 A rms. At the same time bus. In spite of the relatively slow, 2 m/s,it is necessary to keep in mind that the linear speed of the moving end of the rod,secondary arc current is not the only the power fault initiation angles varied fromparameter which determines the arc extinction 35 to 105 degrees, which in turn produced atime. relatively large range of dc offset in the

fault and secondary arc currents.The test procedures for both the K-M and

V-D test series were quite different. andtherefore need some explanation. The K-Mtests were performed using a common singlephase switching sequence of operationsconsisting of the establishment of a phase toground power fault, the single phase faultclearance by the circuit breaker poles in 2.5cycles, and the single phase reclosing in 5.1m

approximately 30 cycles after faultinitiation. In case of an unsuccessful

m

reclosing, a three phase breaker operationcleared the line. Opening and closing of theneutral switches (Fig.la) followed 4.8mcorresponding single pole circuit breakeroperations with time delays less than one-ycle for switch openings and less than twocycles for switch closings. The 30 cycles or0.5 s used in the K-M tests is the high speed 6 'nreclostng time now used on the AEP 765 kV 4 5.5m.system. l-4

During V-D tests. the secondary arc wasestablished with no fault on the system.This method resulted in a different sequenceof breaker operations. A line, phase wasopened and isolated a few cycles prior to theclosing of a breaker pole in series with analuminum wire. At the time of fault currentinitiation through the fuse wire, the openphase voltage increased slightly above thesystem voltage because of a low frequencybeating process. To avoid line-to-groundfaults on the system in case the arc did notextinguish during a chosen SPS dead time,reclosing in the V-D tests took place onlyafter the fault current initiating breaker-was opened. The Soviet authors believe that-the secondary arc current in simulated faulttests is similar to the one in real faulttests. Analysis of the transient secondaryarc current in the V-D tests shows that itconsists of three components, namely, theoscillatory component depended on thedischarge process of line capacitance with20 - 50 ms attenuation time, the dc componentwith 30 - 300 ms attenuation time and initialvalue depended on the shunt reactor currentsat the fault initiation moment, and the 50 Hzsteady-state component. The maximum value oftransient fault current in V-D tests variedfrom 1.0 to 3.4 kA while the initial dccomponent value didn't exceed 1.0 kA.According to the K-M test analysis, however.the equivalent dc component time constant isequal to 20 - 30 ms.

The structures for fault initiation areshown schematically in Fig. 2. In the V-Dtests the fault was initiated through a poleof a 750 kV breaker. The utilization of aconventional circuit breaker for this purposeallowed a wide range of fault initiationangles to be attained, resulting in a varietyof initial dc component values. During theK-M tests the power faults were initiatedusing a fiberglass insulating rod whose

(a)

(3_-4v750hV BUS

( .._(i) 6Jm

(b)

Fig. 2 Fault application structures.

Marysville substation (a)

1 - wooden poles, 2 - insulators,3 - ground mesh, 4 - insulatorsupport, 5 - 765 kV tube bus (150mmin diameter), 6 - fiberglass rod,7 - fault steel wire (0.8 mm indiameter), 8 - current transformers

Vinnitsa substation (b)

1 - 750 kV tower truss, 2 - 750 kVinsulator string, 3 - insulator,4 - grading rings, 5 - faultaluminum wire (0.5 mm in diameter),6 - current shunt, 7 - circuitbreaker pole

1540

The symmetrical portion of the powerfault current in the K-M tests was 2.7 kA rmswhen the line was opened at Marysville and8.6 kA rms in other cases. The peak valuesof the fault current, depending on the faultinitiation angle, ranged from 4.1 kA to 6.6kA for the line open at Marysville and from15 kA to 17.5 kA when the line was energizedfrom both ends.

Pre-test voltage at the substation busesvaried within 2% of 765 kV in the K-M testsand from -1% to 5% of 750 kV in the V-Dtests. The wind velocity did not exceed 5m/s in most of the tests but reached 10 m/sin a few tests.

In both the K-M and V-D series of testsvoltage measurements were performed usingcapacitance taps on EHV current transformersor on shunt reactor bushings. Specialcurrent transformers (K-M tests) andlow-inductance current-measuring shunts (V-Dtests) were utilized for currentmeasurements. A high fidelity opticallyisolated measurement system[71 was employedin the K-M tests. Magnetic tapes were usedto record most of the monitoring parameters.In the V-D tests voltages and shunt measuredcurrents were recorded by electronicoscillographs while current transformermeasured currents were recorded byelectromagnetic oscillographs. Thecoordination of the oscillographs with thefault initiation was executed by a countdownprocedure in the K-M tests. A specialelectronic device was utilized during the V-Dtest.s to coordinate the timing of thebreakers and oscillographs at both ends ofthe test line. Power line carriers were usedfor signal transfer in both test series.

DESCRIPTION OF TYPICAL TEST OSCILLOGRAMS

A total of 27 SPS tests was performed inthe K-M test series and 15 tests wereexecuted on the V-D line. The value of IS onthe K-M line varied from 8 A rms to 80 A rmsby changing the neutral reactances as well asthe faulted phase and the line loading. Onthe V-D line, five configurations withdifferent numbers of shunt reactor banks andfour-legged shunt reactors were tested.These configurations correspond to 3 A rms,26 A rms and 49 A rms, 50 Hz calculatedvalues of Is. All performed tests are listedin the Appendix.

Typical oscillograms of the V-D tests areshown in Figs. 3-6. The oscillograms in Fig.3 illustrate the extinction of the secondaryarc in a scheme with only two shunt reactorbanks connected to the line (neutral reactorswere disconnected). The calculated value ofIs for this test was 49 A rms while themeasured crest secondary arc current duringthe quasi steady-state condition reached 110A. The oscillograms in Fig. 4 demonstratethe arc extinction in a scheme with one shuntreactor bank and one four-legged shuntreactor connected to the line. The value ofIs in this test was 26 A rms or about onehalf that of the previous test because ofpartial compensation of the interphasecapacitances. The oscillograms in Fig. 5illustrate a similar case which differs from

r A R ~~~~~~AR 240kv

Vi

t1 t2 lIOA AR

2kA ARRIA

t2 Vinnitsa Dniener

4u lV

(T)

Is

V2

Fig. 3 Secondary arc current (Is) andrecovery voltage (VIC) in a testwith two shunt reactor banks on theV-D line (T). Calculated 50 Hzsecondary arc current is 49 A rms.

A - arc interruptionR - arc restrikeE - arc extinction

Ilt I\ ( 0\ / 01A R E

VIA\J /\;/120kV 01

640kVVI I

t t2 A R E

Is-05A

Is1.75kA-

t:2130kV 610A

VRN2

tl t2

(T)

Vinnitsa

t1 VIA--,

Dnieper

Fig. 4 Secondary arc current (Is) andrecovery voltage (VlA), and neutralreactor voltage (VRN2) and current(IRN2) in a test with a four-leggedand a shunt reactor banks on the V-Dline (T). Calculated 50 Hz secondaryarc current is 26 A rms

A - arc interruptionR - arc restrikeE - arc extinction

the preceding test by an additional shuntreactor bank connected to the line. This did-not change the calculated value of thesecondary arc current but did result in ahigher recovery voltage of a resonantnature. In this test the sparkover voltageof the protective gap across the neutralreactor at Vinnitsa substation was reduced to100 kV and led to a sparkover with a.follow-current through the gap for 2.38 s,until the opened phase was reclosed. Also,the excitation of the second harmonic voltagetook place in this test after reclosing of

1541

61 5kv

V2C

VRN2

IRN2

0,&V V V aOnOOkV v >,4/ __OA -H

1MAA c'

1 965kV t2 soS'1 '1 1

'vJ2Ar\cJJJJV\\\\\\r'0<n q \r\F'AfVVVMv0 WVM\VVRN2-, 5rH 615k'- %'!e- I 8 ~~~~~~~~~~83OkV- 11

IRN2: A 420A *n -

ISG21k A JA ..... ..._IV V9t+l.8s t4 SGE

Vinnitsa

IT)

F ) t eI2Is

t2

Fig. 5

Dnnieper

V2

G T I

VRN2

ISG2 j2 ;IRN2

Secondary arc current (Is), recovery voltage (V2C),neutral reactor voltage (VRN2) and current (IRN2), andspark gap current (ISG2) in a test with a four-leggedand two shunt reactor banks on the V-D line (T).Calculated 50 Hz secondary arc current is 26 A rms.

E - arc extinctionSGS - spark gap flashovexSGE - spark gap arc extir

the opened phase. The oscillograms in Fig. 6show the arc extinction in a scheme where thesecond four-legged reactor was connected tothe line and the value of Is was reduced to 3A rms. In this test, a valve-type arresterwith a 130 kV sparkover voltage was usedinstead of the protective gap. This arresterflashed over from the opening of the linephase until the fault was initiated.

Typical oscillograms from the K-M testsare presented in Figs. 7 and 8. In the testillustrated by Fig. 7, the neutral reactanceswere 200 Ohms at Kammer and 150 Ohms atMarysville and the calculated 60 Hz value ofthe secondary arc current was 44 A rms. Thefault was initiated on an outer phase withthe line open at Marysville. The power faultinitiation moment corresponded to an 85degree angle of the line voltage. The faultCurrent (3800 A) was interrupted at tl = 41.4ms by the circuit breaker at Kammer. Thevacuum switches NS1 and NS3 opened at t2 = 48ims bringing the compensation scheme intoeffect. The first interruption of thesecondary arc current occurred at t3 =' 65.4ms. Subsequently, the voltage across the arcpath reached 75 kV in an additional 4.1 msand the arc restruck at t4 = 69.5 mis. Thesecond current interruption took place at t5= 99.1 ms after which the voltage startedbuilding up in a beating process, reaching apeak of 80 kV in another 4.8 ms after the arcinterruption and achieving 425 kV at t6 = 150,ms. At this time another restrike occurredand an intermittent secondary arc remained inthe gap until the reclosing of the circuitbreaker resulted in the transition of thesecondary arc into the power arc which wascleared by a three-phase opening in 2.5

t 2

125kV\ 570A

VRN2 2t \-/\tl t2

Vinnitsa

I 1v1(T) I

IARI7 IRN1o s

S I

Dn ie per

V2

IR

4II~RN2Fig. 6 Secondary arc current (Is), recovery

voltage (VlB), neutral reactorvoltages (VRN1, VRN2) and currents(IRN1, IRN2), and the sum of IRN1and neutral arrester current (IARl)in a test with two four-leggedreactor banks on the V-D line (T).Calculated 50 Hz secondary arccurrent is 3 A rms.

E - arc extinctionAS - neutral arrester sparkoverAE - neutral arrester current

extinction

1542

VA

VB / \I\V\/vVc /V BN V /

VN V/\Q IX 9i/ / I-

5375A-) < U7VV---

Isf d , : ilb A JIiI|AVr + j |4 ? \70 ~~~195A }XI6

to. ti t2 t3 t4 t5 t6

cycles. In the test presented by Fig. 8 theneutral reactance at Kammer was increased to400 Ohms. The calculated value of Is was 37A rms for the middle phase fault and the lineopen at Marysville. The fault was initiatedat a 45 degree angle of the line voltage.This resulted in a dc offset which increasedthe peak value of the fault current from 3800A to 6200 A. The primary fault current wasinterrupted at tl = 44 ms by the breaker atKammer. Full compensation went into effectat t2 = 61 ms when the neutral switches NS1and NS4 opened. The secondary arc currentwas maintained for another 52 ms. After thesecondary arc extinction at t3 = 113 ms, the.recovery voltage across the arc path built upin a beating process with a modulatingfrequency of 5.5 Hz. It reached a first peakof 45 kV in 4.4 ms and a maximum value of 455kV in 85 ms. The restrikes did not occur inthis test and a single phase reclosing wassuccessful.

Fig.7 Secondary arc (Is) and pc(If) currents, recoveryphase (VA, VB and VC) volneutral reactor voltageunsuccessful SPS testfour-legged shunt reactorthe K-M line. Calculatsecondary arc current is 44

VA

Vc

-80kV

VN-

If

6

Is

Vr

I II

to tI t2 t3

Fig. 8 Secondary arc (Is) and p(If) currents, recoveryphase (VA, VB and VC) volneutral reactor voltagesuccessful SPS testfour-legged shun-t reactorthe K-M line. Calculatsecondary arc current is 37

ower faul t(Vr) and

.tages, andISECONDARY ARC CURRENT AND RECOVERY VOLTAGE

VN) in an The power frequency component of awith two measured secondary arc current can bebanks on determined only when the arc lasts at least a

ed 60 .Hz few cycles and reaches quasi steady-stateA rms. -conditions as in the tests illustrated by

Figs. 3 and. 7. The nonlinear characteristicof the arc resistance results in the presence

Af\v of harmonics in the secondary arc current.especially the third harmonic. The contentof the third harmonic does not varysignificantly from one scheme to another andequals about 40 percent of the powerfrequency component. At the same time themeasured power frequency component of thesecondary arc current agrees closely with thecorresponding calculated current values. Adetailed analysis of the harmonic content ofthe secondary arc current under different-boundary conditions is given in [6].

After arc extinction, the open phaserecovery voltage builds up to a steady-statevalue in a beating process. This process ischaracterized by large time constants whichdepend mainly on the line and reactorparameters. The recovery voltage in amajority of the tests did not reach asteady-state condition during 0.5 s in theK-M tests and an even longer time in the V-Dtests. Nevertheless, a comparison of themeasured steady-state recovery voltages andthe corresponding calculated values showedgood correlation. The measured steady-statevoltages were determined, where applicable,as average values of the adjacent maxitum andminimum values of the modulating recoveryvoltage.

The steady-state voltage magnitudes on adisconnected. phase in the V-D tests reachedquite high values. In a configuration withtwo shunt reactor banks and one four-legged

ower fault reactor the measured steady state voltage was(Vr) and 1.12-1.18 p.u. and in a scheme with three

tages, and shunt reactor banks this voltage reached(VN) in a 1.18-1.24 p.u. Calculations not accountingwith two for corona losses resulted in 1.45 p.u.banks on voltage for the first scheme and above 3.0

ed 60 Hz p.u. for the second one. It is interesting7 A rms. to mention that published earlier analyses[8]

1543regarding the influence of the corona losseson resonant overvoitages gave practically thesame results, namely, 1.15-1.20 p.u. voltagefor a 750 kV line with similar parameters.This underlines once more the importance ofcorona loss representation when analyzingresonant overvoltages. The transient part ofthe recovery voltage in these two schemes wasaffected by corona losses as well. Thus, asillustrated by Fig. 5, the corona lossessignificantly reduced the time constants ofthe transient recovery process resulting in asmooth beat-free buildup of the open phasevoltage after arc extinction.

EXTINCTION OF THE SECONDARY ARC CURRENT

Secondary arc extinction attempts usuallystart appearing after the end of the firstpart of the transient secondary arc current(Figs. 3 and 7). These extinction attemptstake place at zero crossings of the secondaryarc current. The secondary arc current isaffected by the initial dc offset which inits turn depends on the fault initiationmoment. Therefore, the time span between thefault clearing moment and the first zerocrossing of the secondary arc current alsodepends on the initial dc offset and may varyby a few cycles from test to test. Incompensation schemes corresponding torelatively low steady-state secondary arccurrents, the arc extinction occurs usuallyat the very first secondary arc current zerocrossing (Figs. 5, 6, and 8). When the faultinitiation angle is close to the crest of thephase voltage, the dc offset is practicallyabsent, but an immediate extinction of thesecondary arc may be slightly hindered by thearc path thermal inertia which results in theremaining high conductivity of the arcchannel. In cases corresponding to largersteady-state values of the secondary arccurrent a final arc extinction, if it tookplace, as a rule was preceded by a fewunsuccessful interruption a,ttempts (Figs. 3,4, and 7).

The arc extinction process, as shown byfilm recordings of the previous tests[9), ischaracterized by the appearance of growingdark portions in the arc path while the mainpart of the arc channel maintains itsintensity -and, therefore, high conductivity.Thus, the arc channel is no longerhomogenious after the current interruptingattempts occurred. Sparkovers alongnon-conducting portions of the arc channeldue to the buildup of the arc recoveryvoltage may result in the arc reappearance inthe dark portions of the initial path.Evidently, after each interruption attemptthe arc becomes more and more unstable, thedark spots become greater in size and the airgap withstand voltage increases steadily.Test oscillograms have confirmed thedescribed arc extinction mechanism. Thismechanism results in a substantially shorterarc burning time than that associated withthe arc elongation to a critical length.

The secondary arc extinction timesregistered in both series of tests are givenin Fig. 9. The extinction time for the K-Mtests is defined as the time from the momentof the last neutral or high side reactorswitch opening which constitutes the

.50

C.)w 0.

w

z

0O.

C.).

z

w

cr 0.

A 00A~~~~~

~ -t

-~~-j~---~ ~ I

800 20 40 60CALCULATED SECONDARYARC CURRENT, A rms

Fig. 9 Secondary arc extinction time versuscalculated power frequency secondaryarc current.

A - Vinnitsa-Dnieper tests0 - -Kammer-Marysville tests

Arrows ihdicate tests in which thearc did not extinguish in a giventime and the single phase reclosingswere unsuccessful.

initiation of the compensation scheme, to themoment of the final arc extinction. Becauseof different testing techniques, theextinction time for the V-D tests is alsodetermined differently, namely, as the timefrom the moment of the secondary arc

initiation to the arc extinction moment.During the arc extinction time, as followsfrom the definitions, the compensation schemeparameters do not change.

Arc extinction time for the secondary arc

current less than 40 A rms did not exceed0.15 s in the K-M tests and 0.30 s in the V-D

experiments. The later tests arecharacterized by noticeably greater

dispersion of the arc extinction times

(Fig. 9).

In four tests on the K-M line with 60 Hzvalues of the secondary arc current, Is

larger than 45 A rms, the arc did notextinguish in 30 cycles and single phasereclosings were unsuccessful. In seven tests

performed on the V-D line the 50 Hz value ofIs was 49 A rms. Extinction time in six of

these tests varied from 0.33 s to 0.68 s.One test, however, resulted in the arcburning time of 1.42 s. Particularly largedeviation of the extinction time in thosetests is a function, also, of relatively highrate of rise of the recovery voltage whichwas close to 8 kV/ms in two tests and about15 kV/ms in five others. In addition, the

text a

I7rad

1544arc extinguishing mechanism for longer arcingtimes becomes more complicated with arcelongation. Wind velocity in these casesalso is a factor in the arc extinguishingprocess as is illustrated by da.ta in theAppendix. In eight tests on the V-D linewith Is equal or less than-26 A rms the rateof rise of the recovery voltage did notexceed 5 kV/ms. This contributedsignificantly to the -fast extinction of thesecondary arc in these tests.

Secondary -arc restrikes which took placeduring the tests on. K-M line gave an*opportunity to determine the rate of rise ofthe withstand voltage for a 4.2 m air gap.ionized by the power fault current. Theaverage withstand voltage for this air gap'can be approximately expressed as Vg = 10t,at least for the first 50 ms after arcextinction. In this expression the time tshould be in ms and the air gap withstandvoltage Vg in kV. The withstand voltagedev'iation from test to test in the K-M serieswas relatively small. Comparing thewithstand voltage of the 4.2 m air gap 'withthe rate of rise of recovery voltage in theV-D tests performed with a 6.2 m gap along aninsulator string, one can conclude that 15kV/ms rate of rise observed in the tests withIs = 49 A rms, is close to the withstandcapability of this gap. This could alsocontribute to the longer extinction times inthe V-D tests. Also, it should be said thataccording to previous tests on Soviet 750 kVlinesEl0], the withstand voltage rate of riseof a 6.2 m air gap along an insulator stringwas about 7.5 kV/ms after the arcextinction. Voltage withstand capability of2.0 p.u. was reached in 0.15 - 0.20 s inthose tests.

OVERVOLTAGE PROTECTION OF NEUTRAL REACTORS

Protective apparatus was used in bothseries of tests to limit overvoltages acrossthe neutral reactors to the levels determinedby Table 1.

On the K-M line the insulation levelswere identical on the neutral reactors andthe neutrals of the shunt reactors. Theneutrals were protected by specially designedzinc-oxide surge arresters (Fig. la). Themiddle neutral arrester, SA2, connected inparallel with the 800 Ohm reactor at Kammerhas a protective level of 185 kV and consistsof two identical zinc-oxide disk stacks. Twooutside neutral arresters, SA3 and SA4, wereused at Kammer to limit the voltage at theshunt reactor neutrals, in case of neutralswitch-malfunctions resulting, in simultaneousopenings of two adjacent switches. Each ofthese arresters is designed to withstand fullshunt reactor current and to limit thevoltage at the corresponding reactor neutralsto the same 185 kV level. These arrestersconsist of six zinc-oxide, stacks in paralleland are capable of withstanding 375 A forfive consecutive cycles without change ofparameters. The Marysville arrester, SA1,connected in parallel with the 300 Ohmneutral reactor has a protective level of 85kV and consists of four disk stacks capableof conducting 130 A for 12 cycles.

Shunt reactor neutrals as wel'l as neutralreactors in four-legged reactor banks on theV-D line were protected either by valve-typearresters or by stabilized spark gaps. Aconventional valve--type arrester connected inparallel with the neutral reactor could notassure fin,al current interruption after anarrester sparkover, because of the relativelyhigh, close to 2 kHz, frequency of thevoltage buildup process across the neutralreactor. Although in some tests thearresters could cope with the currentextinction (Fig.6), their frequent operationsled to two arrester failures. After thesefailures occured, stabilized spark gaps wereused for neutral reactor protection. Thestabilized gaps with a sparkover voltage ofabout 130 kV operated in three tests withcurrent values reaching 630 A. The maximumspark gap current conducting time in thesetests was 2.38 s (Fig. 5), although underactual operating conditions current durationthrough the spark gaps should be considerablyshorter. It is necessary to notice, also,that an operation of a spark gap shorts outthe corresponding neutral reactor, changesthe compensation scheme parameters and,therefore, may influence the secondary arcextinction process. The neutrals offour-legged reactor schemes introduced onSoviet 750' kV lines later were protected byzinc-oxide arresters. The protective levelof these arresters is 110 kV. The arrestersare designed to withstand 60 kV rms voltagefor 3.5 s and 1.2 kA current of 1.2 x 2.5 mswave shape.

PERFORMANCE OF REACTOR SWITCHES

No special measures were employed toprevent prestrikes and reignitions in thevacuum switches of the switched four-leggedreactor at Kammer. Reignitions in the vacuumswitches occured in 7 and prestrikes in 11operations out of 21 SPS tests when theneutral switches were used. A typicaloscillogram shown by Fig. 10a illustrates aseries of reignitions during. the openings ofswitches NS1 (12 reignitions) and S3(1 reignition) after a fault on phase A. Therate of rise of the voltage across a switchand, therefore, across the neutral reactorafter each interruption was determined mainlyby the reactor inductance and -the capacitanceof a divider used to measure the neutralvoltage. Duration of a series of reignitionsin any given test did not exceed 2.4 ms withthe number of singular reignitions reaching25. For comparison, 'an oscillogram from atest without vacuum switch reignitions ispresented in Fig. 10b.

The prestrikes in the vacuum switchesoccured in about a half of the tests. Atypical oscillogram with switch prestrikesis given in Fig. 11. The number of singularprestrikes was less than the number ofreignitions and did not exceed 6. At thesame time, due to different initialconditions, the intensity of the transientprocess during reignitions was slightlygreater than during prestrikes. Reignitionsand restrikes, however, did not affect theperformance of the compensation schemeconsidering the presence.of neutral arresters.

I I620A

I i.it,

*-325A

*I8

1 1

1 I.S62S6tmsec(a)

15.625 msec

( b)

1545opening and to 4 ms or 8 ms, depending on theswitch, for closing operations. Switchoperations during the tests did not produceany noticeable overvoltages.

K

Fig.10 Neutral voltage (VN) and vacuum

switch currents (INSl, INS3) in theswitched reactor bank at Kammer.

a - switch opening with reignitions

b - switch opening without reignitions

Fig. 11 Neutral voltage (VN) and vacuum

switch current (INS4) during aneutral switch closing withprestrikes in the switched reactorbank at Kammer.

Resistors of 7000 Ohms were employed inthe high side reactor switches (Fig. la) forclosing and opening operations. Timeinterval between the main and auxiliarycontact systems was equal to about 16 ms for

CONCLUSIONS

1. Staged fault tests on the 765kV, 243km Kammer-Marysville line and the secondaryarc extinction tests on the 750 kV, 417 km:Vinnitsa-Dnieper line have clearlydemonstrated the applicability' of high-speedsingle-phase reclosing on theshunt-reactor-compensated EHV lines. Therequirea reduction 'of the secondary arccurrent was achieved by using conventional aswell as switched four-legged shunt reactorbanks.

2. The secondary arc extinction time,when the calculated secondary arc current wasless than 40 A rms, did not exceed 0.15 s onthe K-M line and 0.3 s on the V-D line. Therate of rise of recovery voltage after arcinterruptions was less than 8 kV/ms in thesetests.

3. The secondary arc extinction time inthe V-D tests varied from 0.33 s to 1.42 s inthe schemes corresponding to the calculatedsecondary arc current, Is , of 49 A rms. Infour tests on the K-M line with Is largerthan 45 A rms the secondary' arc did notextinguish in 0.5 s.

4. The secondary arc extinction time inschemes with the calculated secondary arccurrent less than 30 A rms, is determinedmainly by the first zero' crossing of thesecondary arc current. A larger initial dccomponent in the secondary arc currentresults, as a rule, in a' longer extinctiontime.

5. Gapless zinc- oxide arresters weresuccessfully used for protection of theneutral reactors against the switching surqesand temporary overvoltages during SPS'operations. Stabilized spark gaps can alsobe used for this purpose.

6. Resonant overvoltages on an open phase ofthe V-D line were limited to 1.25 p.u. due tocorona losses.

ACKNOWLEDGEMENTS

The authors are grateful to numerouscolleagues within the American Electric PowerService Corporation and the All-UnionElectric Power Research Institute 'whocontributed to this project. The field testswere performed with the permission andsupport of the Ohio Power Company and theUSSR Power System.

REFERENCES

1. Knudsen, N. - "Single PhaseSwitching on Transmission Lines UsingReactors for Extinction of the SecondaryT\rc", CIGRE, Peport 310, 1962

2. Kimbark, E.W. - "Suppression ofGround-Fault Arcs on Single Pole 'Switched EHVLines by Shunt Reactors", IEEE Transactions,Vol. PAS-83, March 1964, pp. 285-290

INSI

INS3

VN

155kV -

VN

1546

3. Edwards, L., Chadwick, Jr., J.W.,Reich, H.A., and Smith, L.E. -"Single PoleSwitching on TVA's Paradise-Davidson 500 kVTJine. Design Concepts and Staged Fault Test

Results", IEEE Transactions, Vol. PAS-90NTo.6, 1971, pp. 2436-2450

4. Petersen, W. -"Protection ofAlternating Electric Current Systems", U.S.

Patent Office, No. 1537371, 1925

5. Shperling, B.R., Fakheri A.,Ware, B.J. - "Compensation Scheme forSingle-Pole Switching on UntransposedTransmission Lines", IEEE Transactions, Vol.PAS-97, July/August 1978, pp. 1421-1429

6. Shperling, B.R., Fakheri, A.J.,Shih, C.H., Ware, B.J. - "Analysis of SinglePhase Switching Field Tests on the AEP 765 kVSystem", IEEE Transactions, Vol. PAS-100,No. 4, 1981, pp. 1729-1735

7. Malewski, R., Nourse, G."Transient Measurement TechniquesSystems", IEEE Transactions, Vol.May/June 1978, pp. 893-902

R.

in EHVPAS-97,

8. Rozavskaja, S.N., Shperling, B.R. -

"Resonant Overvoltages on EHV TransmissionLines", Power Transmission by AC and DCSystems, Vol. 19, 1973, pp. 135-148

9. Beliakov, N.N., Komarov, A.N.,Rashkes, V.S. - "Results of InternalOvervoltages and Electrical EquipmentCharacteristics Measurements in the Soviet750 kV Network", CIGRE, Report 33-08, 1978

10. Beliakov, N.N.,Rashkes, V.S., Khoetzian,autoreclosing on 750 kVreactors", Elektrichestvo,6-12.

Bourgsdorf, V.V.,K.V. "Single-phaselines with shuntNo. 7, 1981, pp.

APPENDIX

MAIN RESULTS OF SINGLE PHASE SWITCHING TESTS

TABLE Al. Kammer-Marysville Tests (USA)

Calculated 60 Hz Recovery SecondaryTest Faulted Neutral Secondary Recovery Fault Current Voltage Arc WindNo Phase Reactances Arc Current Voltage Initial Crest Maximum Extinction Velocity

Ohm Value Beat Value Time

RN1 RN2 Is, A rms Vr, p.u.* If, kA Vr, p.u.* t, s m/s

1a A 300 800 0.08 4.8 0.15 0.018 2-52a A 300 800 8 0.08 6.0 0.12 0.042 2-52a A 300 800 8 0.08 5.2 0.10 0 2-54a A 300 800 8 0.08 4.2 0.13 0.016 2-55 A 150 800 12 0.14 5.5 0.16 0.034 56b A 150 800 12 0.14 16.5 0.20 0.018 57C A 150 800 14 0.15 15.0 0.22 0.012 58 B 150 800 23 0.24 4.5 0.38 0.024 5gc B 150 800 24 0.25 17.5 0.43 0.019 510 C 300 400 24 0.30 4.5 0.39 0.011 511 C 300 400 24 0.30 5.6 0.45 0.34 512 C 300 400 24 0.30 4.1 0.40 0.016 213d C 300 400 24 0.30 4.1 0.42 0.013 514d C 300 400 24 0.30 4.2 0.44 0.023 515e A 300 - 25-30 0.3-0.5 5.1 0.50 0.026 2-516f A 0 800 25 0.33 4.8 0.64 0 217 B 300 400 28 0.30 5.0 0.45 0.022 918d C 150 400 30 0.42 5.5 0.55 0.040 5lgd C 150 400 30 0.42 4.8 0.57 0.038 5

20 B 150 400 37 0.45 6.3 0.73 0.051 921d B 150 400 37 0,45 6.6 0.71 0.160 f 522d B 150 400 37 0.45 6.3 0.70 0.063 523 C 0 400 40 0.69 6.0 0.95 0.041 7249 A 150 200 45 0.77 5.4 - - 2259 B 0 400 51 0.74 5.6 - - 7269 B 0 400 51 0.74 5.7 _ _ 9279 B 0 0 80 1.50 6.5 _ _ 5

* lp.u.=625kVa. Optimal compensation schemeb. 200 MW power flowc. 950 MW power flowc. 950 MW power flowd. High side reactor switch operatione. RN2 was partially shorted out due to its failuref. Artificial delay for the last switch openingg. Arc did not extinguish in 0.5 s. Single phase switching was unsuccessful

1547

TABLE A2. Vinnitsa - Dnieper Tests (USSR)

Calculated 50 Hz Recovery SecondaryTest Number Neutral Secondary Recovery Fault Current Voltage Ar-c WindNo of Shunt Reactances Arc Current Voltage Initial Crest Maximum Extinction Velocity

Reactors Ohm Value Beat Value Time

RN1 RN2 Is, A rms Vr .u.* If, kA Vr, .u.* t, s m/s

1 2 235 235 3 0.07 1.4 0.07 0.16 22 2 235 235 3 0.07 1.45 0.08 0.14 23 2 0 235 26 0.14 1.0 0.19 0.24 2-54 2 0 235 26 0.14 1.75 0.18 0.22 2-55 2 0 235 26 0.14 2.0 0.23 0.24 2-56a 3 0 255 26 1.15** 2.6 1.18 0.09 27a 3 0 235 26 1.15** 2.1 1.16 0.22 2-58a 3 0 235 26 1.15** 2.1 1.12 0.30 2-59 2 0 0 49 0.19 2.0 0.37 0.36 5-1010 2 0 0 49 0.19 1.35 0.43 0.67 2-5llb 3 0 0 49 1.19** - 1.24 0.68 212b 3 0 0 49 1.19** 1.5 1.19 1.42 2

13b 3 0 0 49 1.19** 1.5 1.22 0.51 2-514b 3 0 0 49 1.19** 1.9 1.24 0.33 5-1015b |3 0 0 49 1.19** 3.4 1.18 0.53 2-5

* 1 p.u.=643 kV** Corona losses are taken into accounta. Beats are absent. Recovery voltage reaches the steady state condition in 0.28 s.b. Beats are absent. Recovery voltage reaches the steady state condition in 0.15 s.

DiscussionS.R. Lambert (Douglas G. Peterson & Associates, Inc.): The authorshave presented some interesting information on single phase switchingand secondary arc extinction time, and it is most welcome. Coupled withthe work of Makopar [11, Haubrich [2] and Kappenman [3], there is nowa wealth of solid data supporting single phase switching applications.The authors report longer extinction time (often no extinction) especial-

ly when neutral reactors are not present. This is often the result of highersecondary arc current magnitudes, but I also suspect that it is partiallydue to the result of a change in the recovery voltage magnitude andespecially time to crest caused by changes in the effective resonant cir-cuit. Perhaps the authors also came to this conclusion?

I am also surprised that the withstand of ionized air gaps appears asconsistent as reported in the section entitled "'Extinciton of the Secon-dary Arc Current." Considering the vagaries of wind velocity and thestatistical nature of air insulation, I would have expected a significantstandard deviation. Would the authors comment further?

Regarding their comments on overvoltage protection of the neutralpoint, I strongly support their concern for the insulation. Addition ofa neutral reactor substantially changes the zero sequence circuit associatedwith the line/reactor combination. Neutral point shifts increases becauseof the now higher XO/Xl ratios, and sustained, as well as switching surgeovervoltage magnitudes can dramatically increase over the case wherethe neutral reactor is not present [4]. These are by no means insurmoun-table problems, but they do require careful consideration.

REFERENCES[1]. A.S. Maikopar, "Minimum Time of Automatic Reclosing," Elec-

tric Technology, U.S.S.R., pp. 302-315, 1960.[21. H.J. Haubrich, G. Hosemann and R. Thomas, CIGRE 1974,

pp. 31-39[3]. J.G. Kappenman, G.A. Sweezy, V. Koschik, and K.K. Mustaphi,

"Staged Fault Tests with Single Phase Reclosing on the Winnipeg-Twin Cities 500 kV Interconnection," IEEE Summer Meeting 1981,81 SM 366-4.

[4]. S.R. Lambert, V. Koschik, C.E. Wood, G. Worner and R.G.Rocamora, "Long Line Single-Phase Switching Transients and TheirEffect on Station Equipment," IEEE Summer Meeting 1977, F 77687-7.

Manuscript received August 6, 1984.

H.N. Scherer, Jr., B.R. Shperling, J.W. Chadwick, Jr., N.N. Belyakov,V.S. Rashkes, and K.V. Khoetsian: We thank Mr. Lambert for his in-terest in the papers. We agree that the existing experimental data fromdifferent staged fault tests, including the tests analyzed in the paper aswell as from simulated fault tests, permit the formulation of technicalrequirements for single phase switching schemes applicable to a varietyof transmission line configurations.

It is also interesting to note how much progress has been made duringthe last 10-15 years into the understanding of the secondary arc behaviorduring single phase switching. For example, Maikopar's and Haubrich's,et al papers analyzed single phase switching applications on uncompen-sated lines. For these lines, the steady-state values of the recovery voltageare proportional to the ratio of interphase and phase-to-groundcapacitances and are therefore relatively low. The rate of rise of therecovery voltage on uncompensated lines is, however, quite high andresults in multiple restrikes during the arc extinction process. In bothpapers, the authors suggested that shunt reactors would not substantiallyimprove the prospects of successful single phase reclosing and that thehigh recovery voltage resulting from the open phase resonance condi-tion might even prevent final arc extinction. In other words, the rateof rise of the recovery voltage was not considered to be an importantparameter.The results of our studies, as well as studies by others, have proven

that shunt reactors and especially conventional and switched four-leggedshunt reactor banks, noticeably influence the arc extinction process. Forexample, if one compares the arc extinction times for uncompensatedlines, as summarized by Maikopar and Haubrich, et al, with the testresults described in our paper, it becomes quite clear that the additionof shunt reactors significantly decreases the arc extinction time.At the same time, the calculated power-frequency components of the

secondary arc current and the recovery voltage as well as the rate of riseof the recovery voltage, are all closely interrelated for transmission linesutilizing four-legged reactor compensation. The neutral reactors, selectedto reduce the secondary arc current to a very small value by equatingthe phase-to-phase inductive impedance to the phase-to-phase capacitorimpedance produce a very large equivalent phase-to-phase impedance.Once the secondary arc is extinguished, the open-phase-to-ground voltageis relatively small because of the large ratio of the equivalent phase-to-phase impedance to the equivalent phase-to-ground impedance.

Standard deviation values for the air gap withstand voltage after arcextinction were not presented in the paper because the tests did not pro-vide a sufficient number of actual secondary arc reignitions for a thorough

1548

statistical analysis. Nevertheless, a general analysis of test results showsthat the deviations of measured reignition voltages are similar to theflashover voltages for long air gaps (Fig. 12). Thus, we would expectthe standard deviation of the reignition voltages to be approximately 5percent. We agree that the gap withstand capability may vary for theextreme wind velocities. However, for the test conditions, the windvelocities (less that 10 m/s) had negligible effect.

During the papers's review process, as well as at the meeting we receivedcomments indicating that insufficient emphasis had been placed on thedifferent test sequences used in the two test programs and the extent towhich these differences affected the results. Thus, the switching sequencesused to obtain the secondary arc currents are depicted in Fig. 13. It shouldbe noted that the K-M test sequence initiated a primary fault to establishthe arc path for the secondary arc current. The V-D tests did not placea primary fault on the system; instead, a fault was placed on a discon-nected phase and coupled energy was used to establish the arc path forthe secondary arc current.

Table 2Test Parameters

A tabulation of test parameters, both calculated and measured, is givenin Table 2. The differences in switching sequence should be borne in mindwhen comparing the results. The preconditioning of the arc path resultsin a more stable, shorter secondary arc with substantially different timeconstants, initial offset, and extinction time as was illustrated, for ex-ample, by Fig. 9. The rate of the air gap withstand voltage capabilitywas also different in both test series (Fig. 12). Detailed analysis verifiesthese substantial differences.

Finally, we would like to add that Figs. 3, 4, 5 and 6 in the paper cor-

respond to tests 9, 4, 6 and 2 of Table A2 and Figs. 7 and 8 correspondto tests 22 and 24 of Table Al.

Manuscript received December 18, 1984.

40O

30

20

10

O 10 20 30 40 ms

Fig. 12. Air gap withstand voltage capability after arc interruption.

2 3

CE--

-0---

K-M TEST SEQUENCE 1 - 2 - 3

t 1,2 = (40-44)ms

t2,39 0.58

V-D TEST SEQUENCE 2 - I - 4

t2,1 (40-200)ms

ti,4 > 3s

Fig. 13. Sequence of operations in the single phase switching tests:

I - Fault application

2 - Opening of line breakerson the faulted phase

3 - Closing of line breakerson the faulted phase

4 - Line breaker opening onthe unfaulted phases

V* FLASHOVER VOLTAGE

o WITHSTAND VOLTAGE

0~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~011 04 -IK-M TESTS - I

A-,-V-D TESTS

.-p~ ~ ~ -

~o _ _

0

-I~~~~~~~~~~~~~~

PARAMETERS K-M TESTS V-D TESTS

Short Circuit Current 6-12KA RMS

Fault Current Initial 4.2-17.5Crest Value, KA

Simulated Fault 1.0-3.4Current Initial CrestValue, KA

Calculated Secondary 8-80 3-49Arc Current, A RMS

Calculated Faulted 35-660 30-1300Phase Recovery Voltage,KV RMS (Ignoring CoronaLosses)

Calculated Faulted 35-660 30-515Phase Recovery Voltage,KV RMS (with CoronaLosses)

Air Gap Clearance, m 4.2 6.1

Rate of Rise of the 10 7.5Air Gap WithstandVoltage, Kv/ms

i