Studies on impulse behaviour of a transformer winding with simulated faults by analogue model I ing

C.K. Roy J.R. Biswas

Indexing t e r m : Analogue modelling, SimulatedJiiultr. Serge performance. Transformer windings. Winding /nuits, Winding responses

Abstract: The paper highlights some critical observations made in course of investigations on surge performance of a transformer winding by analogue modelling. The studies cover winding responses to both full and chopped standard lightning impulses where time-to-chop was varied over a wide range to find the maximum stresses across winding discs. Three specific zones along the winding were selected for studies. These were the line-end and earth-end sections each spanning about 6"4 of the entire winding length, while the midwinding disc was chosen to represent the midwinding section. Two types of faults were simulated: series faults both permanent and voltage-dependent, and parallel faults through resistance with preset faulting levels. Winding responses to faults under impressed impulses were observed and scann'ed using both neutral- current (Hagenguth) and tank-current methods of fault detection. Important observations are cor- roborated by oscillograms of winding responses. The results suggest and confirm appropriate methods of fault detection on fault occurrence in the specified sections along the winding.

L i s t of symbols

A,

A, B,,, C C,, C,, C,, = capacitance between mid-limb winding and

C, = total ground capacitance C,,

C,,

C,,

C,,

= second disc coil from line terminal representing

= fifth disc coil representing line-end section = 40th disc coil representing mid-winding section = capacitance to earth per disc of winding = 76th disc coil representing earth-end section = 79th disc coil representing earth-end section

line-end section

tank

= capacitance between side-limb HV winding and

= capacitance between impulsed HV winding and

= capacitance between impulsed HV winding and

= capacitance between nonimpulsed side-limb HV

tank

LV winding on the same limb

tank

winding and core

( ' , IEE. 1994 Paper 1218C (P7). received 14th June 1993 The aulhors are with the Electrical Engineering Department, Iadavpur University, Calcutta-700 032, India

1k:L Proc.-Grner. Tronsm. Divrrih., Vol . 141. N u . 5. Sepiemhrr I994

K k = coupling coeffjcient L L, 1.;. L> = self inductance of coils kf = mutual inductance M , M,, = self inductance N t ,

1 Introduction

A power transformer, being a vital and expensive piece of equipment in a power system, needs critical attention from the standpoint of its insulation design and per- formance under both steady-state and transient stresses.

It is highly desirable to predict the insulation per- formance against impulse voltage even at the design stage. to guarantee an amenable impulse-voltage and stress distribution along the winding. Impulse behaviour has therefore been considered to he a major criterion in designing HV power transformers.

Insulation with such transformers is costly, occupies space, impedes removal of heat from the core and con- ductor and increases reactance. This consequently neces- sitates use of a greater amount of copper and iron in the transformer than may he required for the electromagnetic design. Especially in high-capacity, high-voltage power transformers, an optimised design is to he aimed at, with appropriate insulation design aided if necessary by stress controlling techniques.

In the present paper extensive ohservational studies with standard full-wave and chopped impulses have been made on an electromagnetic analogue model of one HV winding limb of a 3311 1 kV, 3 MVA, Dy l l . O N 50 Hz power transformer, assuming that the response of a well constructed model is fairly in good agreement with that of the original [ I 1, 141.

1.1 Previous work Great progress has been made since 1915 in understand- ing the physical phenomena which govern the transient response in transformers. As a result, practical computa- tional techniques have been evolved.

The fundamental work on transients in a magnetic winding was done by Wagner [ I ] in 1915. Wagner divided the entire phenomenon into three time domains: a period of initial voltage distribution, a period of free oscillations and the pseudofinal distribution. He, however, did not consider the effect of mutual induct- ances hetween turns. This effect was. however, considered

40 1

= series capacitance per unit length of disc

= inductance per unit length = total inductance of two coils

= inductance of one disc coil

= number of turns per disc = time to chop in impulse-voltage wave

by Blume and Bayajian [Z] in their investigation. Further investigations were undertaken by Heller, Veverka et al. [3] and by Abetti [4]. 1411 these reports were based on the standing-wave concept.

Rudenberg [5] was the foremost amongst those who investigated the phenomena from the viewpoint of travelling-wave propagation along the surge-stressed winding and left out the effect of mutual inductances from their considerations.

In 1978 Miki, Hosay;~ and Okuyama [ 6 ] confirmed and explained that the iron core had hardly any role to play in the response of an impulse-stressed winding.

The best acceptable theories based on standing waves and travelling waves, for computation of surge distribu- tions of an impulse-stressed winding, assumed uniform winding. Obviously, the presence of one or more wind- ings of other phases on tlhe same leg and nonuniformities in the winding will introduce modifications in the response which can hardly be belittled in sophisticated designs. As expected. more complications are introduced in such cases which make the computational work extremely lengthy.

1.2 Scope of the present work Shortcomings of the computational method of transient-

nalysis may be overcome by analogue studies on an electromagnetic model which, if constructed with care. represents [ I 11 the actual1 winding very accurately.

I n computer-aided computation. scanning has to be carried out by discretising the period of sustenance of the voltages by the winding amd i t may be quite likely to miss some important features, of the oscillation pattern and reach incorrect conclusions. In studies on the model. however, though there is discretisation of parameters and the high resistance of the model winding effectively damps the oscillation amplitudes, the pattern of oscil- lation may be understood and selective studies may easily be undertaken with the help of a sophisticated versatile transient recording CRO.

In the present work, the authors employed a highly sophisticated version of recurring surge generator (RSGI with variable repetition frequency and widely controlla- ble time-to-chopping.

The studies cover winding responses to both full and chopped standard impulses. The time-to-chop was varied over a wide range to identify the most critical chopping instant with regard to thle winding. Specific winding sec- tions and coils were picked up t o represent properly the winding sections, namely the line-end section, mid- winding and earth-end sections, and performance of these coils were studied when ithe whole winding was subjected to impulses from the RSG. Two types of fault, the long- itudinal (disc-to-disc) fault and the transverse (disc-to- earth) fault were simulated by suitable faulting circuits, establishing both voltage-sensitive and resistive faults.

Winding responses to faults were investigated by both neutral-current (Hagenguth) [~9] and tank-current methods 'of fault detection.

The paper highlights the pattern of fault response on the basis of the nature of fault and location of the faulted section of winding, and thereby attempts to suggest the most appropriate method of fault detection to be adopted selectively with regard to its nature and zone of occurrence.

2

The transformer winding is represented by a network with distributed paramelers 1:4, 61 as shown in Fig. I . C,

402

Analogue model and the parameters

K , L and M represent the capacitance to ground, series capacitance, self inductance and mutual inductance per unit winding length, respectively. Magnetic losses have been ignored.

J

The model is based on the following basic assumptions: (i) The modelling is carried out by considering it as an

isolated winding. This is because of winding connections for impulse test resulting in mutual cancellation of the main fluxes in the core due to hit limb with that gener- ated by the short-circuited L V winding. The magnetic coupling by leakage flux between the impulsed HV winding and short-circuited LV winding is disregarded.

(ii) In the model, the iron core may be replaced by an air core owing to the above main flux cancellation [ I t , 121. In the iron core, only the leakage flux contributes to the impulse response of the winding under test. Removal of the iron core may amount to disregarding the slower electromagnetic component [ I O ] of transient stress dis- tribution. This omission is insignificant 16, 10, 121 in impulse studies within the time domain up to 200 ps . The cl'fect of the core may, however, be significant in switching-surge and transferred-surge studies [b, lo].

The model constructed is based on the design data of a practical 3 MVA, 33/1 I kV. 3-phase 50 Hz, ON, DY 11

I E E P r o ~ . . - ( ; e w r Transm. Di.urih.. P o l 141. ho 5, S q w m h e r 1YY4

transformer whose HV winding is considered to be impulse tested as per IEC Publication 76-3, 1980 [ 131.

2.1 Lumped parameters The above transformer HV-winding dimensions are as follows :

Outer diameter = 524 mm Inner diameter = 424 mm Winding depth (radial') = 5 0 m m Axial height of a disc = 6.6 mm Number of discs = xo Average number of turns per disc = 19

Self inductance of one disc is evaluated at M O = 0.324 mH by appropriate methods [6, 161.

The capacitances between windings are calculated by treating the windings as cylindrical electrodes [7. 8, I I , 121. Capacitances between side-limb windings and the tank (C+) are found by treating the winding and the rele- vant portion of the tank as a coaxial cylindrical electrode system (with certain approximations). The midlimb- winding-to-tank capacitances C , , are evaluated by the image method. The ground capacitance of the impulsed HV winding comprises the capacitance (Chi) of impulsed HV winding to LV winding on the same limb (which is short-circuited and earthed during impulse testing) and the capacitance (CJ between the impulsed HV winding and the tank. For a midlimb winding under test, a third component of ground capacitance exists owing to the physical proximity of the other side-limb cores along with the yoke. This capacitance (C,,') of nonimpulsed side-limb windings is generally insignificant compared with the other components of ground capacitance. By conventional connection for impulse testing, the LV winding, at earth potential. screens the core. The capa- citances of LV winding to core is thus shorted and madc ineffective. The numerical values of winding capacitances are presented in Table 1 . With 80 coils in the winding of 21 side limb under test, the parameters pcr disc (coil) arc as follows:

Self inductance (MO) = 0.324 mH Capacitance to earth (C,) = 1722/80 = 21.52 pF Series capacitance ( K ) = 11.7 x 80 = 936 pF

In the present model, lumped values for capacitances as distributed parameters, are taken to be

C', = 22 pF K = 1000 pF (per disc coil)

The above values chosen to approximate to the calcu- lated values while taking account of commercial avail- ability in the market.

2.2 Construction of the model Bakelite tube of diameter 15.5 cm and height 1 50 cm was taken as the winding former and the iron core was replaced by air. Each disc or coil in the model was con- structed to give the design value of self inductance of i~

real winding disc.

Table 1 : Winding capacitances

The necessary number of turns N in a disc was found to be 28. A three-layer disc was constructed with the average number of turns per layer 9 + (I) and with one tapping from mid-coil. The ends are marked 'start' and 'finish' to indicate magnetic polarity. Superenamelled copper conductor of SWG 23 was used to give the desired resistance. Varnished cotton tape (0.34 mm thick) was used for disc insulation.

Each disc measured 2 mm radius and 6.44 mm thick

The resistance of each coil is found to be 0.91 2 Q making the entire winding resistance 72.96 Q. This high coil resistance in the analogue model was choscn to take account of dielectric and eddy losses a s well as skin eft'ect in the conductors at high frequencies when the real trans- former winding is under impulse tcsl. I t may be men- tioned that a per-phase winding resistance of i R in actual transformer is represented by ;I model winding using 23 SWG copper wire to takc into account and partly accommodate the increased dielectric losses in the actual transformer at the specified test voltage. and partly to optimise the features of a hand! but reliable model that will ensure easy fiorkability.

The self capacitance o f each coil (three-layer. nine con- ductors per layer) is computed by finding the capacitance between adjacent conductors of circular cross-section and with the minimum distance of twicc the thickness of enamel insulation. This was found analytic+l) case of two parallel cylindrical electrodes with composite dielectric o f air and enamel insulation.

To find the self capacitance of the disc. the equivalent capacitance of the network (Fig. 2h) was determined. The numerical value becomes 3 p F and is negligible com- pared with the 1000 p l lumped capacitance connected between the coils.

The mutual inductance M between coils is determined from oscillation frequencies of an Lc' series circuit (Fig. 3a) excited by the RSG. L,. is the total inductance between two successive coils additively, connected so that L,. = 1: + L: + 2 M , where M = k , v:(L'iI,;) =: k C . ( k = coupling coefficient. 1,; = I*> = L).

The equivalent electrical nelwork of the model winding in the experimental arrangement is shown in Fig. 4.

In this investigation, the coupling factor k in the model has becn fixed for an interdisc spacing of 6 mni (edge-to- cdgc along the axis) maintained throughout. Variations of mutual inductance M and couple factor k with inter- coil spacing are shown in Fig. 3h and ('. Thcie plots have been obtained by the dynamic method of total- inductance determination from the change in the ringing frequency, as explained above.

2.3 Simulation of faults I n the present investigation three basic types o f fault have been recognised. Thesc are

( i ) simple short circuit of a disc or coil of the winding ( i i ) the above fault made voltage sensitive so that the

Winding limb Components of Total Series ground capacitance ground capacitance

capacitance c,. Ch. c,. C.

PF PF PF PF PF Side-limb winding 1561.8 160 negligible 1721.8 ( C v 2 ) 11.7 Mid-limbwinding 1561.8 246.6 161 6 1970.0 (Cg,) 11.7

I E E Proc.-Genrr. 7 ' r m . s m Disrrih.. Vof. 141. Nt,. 5, S c p t r m h w IYY4 403

fault is established a t a desired level of voltage across the coil, and

(iii) voltage-sensitive resistive transverse fault (coil to ground-shunt faults) being established at preset voltages.

I I axls of the col1

. -1. a

Fig. 2 a Cross-section of coil b Equivalent circuit representing a l capacitance of each coil

Coil and its equioalent circuit

These faults have been simulated as follows: (i) A short-circuited coil fault is simulated by shorting

the coils under study by placing a shorting link tempor- arily across the disc or coil.

(ii) Voltage-sensitive short-circuits of coils are achieved by a nonlinear bilateral circuit which is used to trigger the fault at the desired voltage, called the threshold fault voltage. This is shown in Fig. Sa. The circuit is made up of two similar elements, one being the inverse of the other for bipolar sensitivity. In each element, the thyristor is triggered by an appropriate gate voltage. The device is designed to create a polarity-symmetric low-resistance fault path, when the applied voltage across the faulting device exceeds about 1.6 V of either polarity. This is to simulate a fault condition of equivalent magnitude 2 kV across the faulted points in actual impulse testing of a 33 kV winding at 170,kW peak level. This is computed on the proportion of 1.6 V to the applied impulse of 136 V peak from the RSG. This helps us study the fault which may be initiated across a disc or a few turns of the winding. However, this initiation-voltage magnxtude can be easily increased as desired with more diodes D (IN4007) in series. The thy- ristor provides a low-resistance conduction, when trig- gered, simulating discharge in the real fault.

(iii) Voltage-sensitive resistive transverse faults of coils to ground are established by a nonlinear, bilateral circuit which is used to initiate the fault at preset threshold potentials. The arrangement shown in Fig. 5b is basically almost same as that in (ii). It consists of two elements, one sensitive to positive polarity and the other to negat- ive polarity of the voltage oscillations at fault. In the

404

- 0 08- -

E’

0 04-

0- 0

r

1-

0

20 40 60 80 separation distance, mm

b

0 20 40 €0 80 separotion distance,mm

C

01

Fig. 3 Mutual inductance a Circuit for determining dynamic mutual inductana with RSG b Variation of mutual inductance with intercoil spacing c Variation of coefficient of coupling with intercoil spacing

I E E Proc.-Gener. Transm. Distrib., Vol. 141, No. 5, September 1994

positive element an appropriate Zener diodes in series gives a preset trigger potential for the thyristor to turn on and establish the fault over a resistance R set at 25% of the impulse impedance of the coil. The other element

this probe is to be connected io the point where the voltage IS to be measured - - - - - - - - -

impulse generator

Fig. 4 Block d. I ram of experimental arrangement

" to coil ends

to earth

b

Fig. 5 Y Shon-circmt faulting b Shunt faultmg D = IN4007 Thvnstor = s3506

Voltage-sensitive sho~t-circuit-raaulting and shunt-fault devices

R = Z0/4 Z, = impulsz impedanoc of transformer winding Z = EC403

responds to negative polarity and includes a diode in the trigger circuit of its thyristor. The fault-path resistance is omitted in this element for a quick response, even at low amplitudes of negative swing.

The authors have noticed during their long association with impulse testing of transformers that, in most cases, failure by shunt faults occurs on full voltage application to transformers. At reduced level which is kept to 50-75% of full voltage level, fault occurrences are few. With this is view in the present investigation, the shunt- fault-simulating circuit is designed to trigger faults only when the terminal voltage across the circuit exceeds 60% of the full voltage of any given polarity.

IEE Proc.-Gener. Transm. Distrib., Vol. 141, No. 5, September 1994

With reference to the full impulse-wave output 136 V peak (positive polarity with respect to earth) from the RSG, the faulting level at 60% is found to be 81.5 V, approximately. This is ensured with the help of series- connected Zener diodes whose total operating voltage is set at the above predetermined value.

Addition of further Zener diodes in series may increase the operating level as desired. In this case, two series- connected Zener diodes (EC 403) and one IN4007 diode were used. The circuit, once activated on any polarity, will create a low-resistance fault path in either direction. In this case the circuit is set to trigger a fault at a positive polarity of applied impulse wave.

3

The surge performance of the winding under various fault conditions has been observed on a CRO. Instead of attempting analytical exposition, these observations are mostly presented by oscillographic records and meticu- lous observations. Findings are systematically based on these factual studies and inferences are carefully drawn.

Out of 80 similar coils of which the entire winding is made up, two coils, the second and fifth from the line end, were selected for investigation of the surge per- formance of the line-end section (A) of the winding. These were marked and designated as coils A2 and A,, respect- ively.

The mid-winding part (B) was represented by the 40th disc from the lineend and designated as B,, . The earth- end section (C) was studied from the response of the coils similarly marked and designated as C,, and C,, .

3.1 Occurrence of maximum potential across discs against time-to-chop of fhe applied chopped impulse

The RSG was adjusted to give a continuously variable time-to-chop t, over a wide range with an output impulse peak magnitude of 136 V, (positive polarity). The maximum amplitudes with polarity of voltage oscillation across the selected coils were observed on a CRO. The time of their occurrence was also noted to understanding of their severity. It may be remembered that a time of occurrence of more than 2 0 p s has little importance owing to attenuation by HF losses in the conductor and dielectrics. The results are reproduced in Table 2 and pre- sented by the oscillographic records (Figs. 6-10).

3. I . 1 Oscillographic records Oscillogra-hic records of the above observations are pre- sented in Figs. 6-10. It is observed that the most critical time-to-chop t , is about 3.1 ps for a line-end coil. In a midwinding coil, the stress situation arising due to

Table 2: Maximum stress across disc against time-to-chop

Coil Amplitude and Time-to-chop Time of polarity occurrence

V P P

+ 20 0.8

+14.5 2

+7.0 56

+6.5 60

+7.0 150

Investigational observations, records and their analyses

A, -23 3.1 3.7

A, -16.5 3.2 5

8, -10.5 6.0 8.0

C,, -12.5 7.3 12

c,, -12.0 7.0 11

405

chopped-impulse application with the most critical t , of 6 ps, is of much less concern. For an earth end the most critical t, is around 7.2 p s , giving rise to fairly moderate stress.

It has also been observed by a probing study that, when maximum stresses are occurring in the earth-end coil, line-end coils are also simultaneously stressed to an appreciably high order of about 85% of the maximum stress envisaged. Thus. a line-end coil is not relieved of

Fio. 7 Stre.q< oscillation in A.

- the stress even when earth-end stress concentration becomes high.

A careful scrutiny of the oscillograms highlights the following points worth emphasising.

The pattern of voltage oscillation across coil A,, and hence the stress oscillation in the coil, shows a sharp decay of amplitude spanning the first 100 11s. Occurrences of peaks are not systematic and regular.

Stress oscillations in coil A, exhibit an almost expo- nential decay of peak amplitudes, occurring more regu- larly than in A,.

Coil B,, shows a gradual decay of stress amplitudes after 50 ps . In the later period the swings are more close to sinusoidal variation.

The patterns of stress oscillation in coils C , , and C,, are almost identical to those coils A, and A,. respect- ively. This may be due to positional symmetry of the coils with reference to the midwinding coil B,, .

In the midwinding coils. the voltage/time oscillation pattern indicates that a series of reflections along the winding is likely to be over in about ? O o p s and there-

Fig. 9 Stress oscillation in C , , =. - ~ , - - ~~ ~ 3 . .

Scale5 I O V.divirion

Fig. 8 Scales 10 V.d,\,\,<," 5 0 ps;division

406

Stress oscillatinn in fit.,,, Fig. 10 scaler: i n v!dlvalon TO ps,dwision

Stress oscillation in C , ,

I E E P r o c C e n r r . 7runsm IXwih. . Vol. 111. .WO. 5. Sepremher 19Y4

after a standing wave with small amplitude begins to dominate in this section of the winding.

For the coils in the line-end and earth-end sections, the series of reflections continues for some time and the coils do not completely settle to standing waves within the time frame scanned by the CRO.

3.2 Short-circuit coil- fault responses and detection In this paper, fault investigations are mostly limited to permanent-type series faults established by short- circuiting representative coils along the winding, one at a time. Responses of each to both full and chopped impulses, in both neutral-current and tank-current methods of detection, are observed and scrutinised.

3.2.1 Short-circuit coil-fault detection The winding under test is first connected (Fig. l l a ) as usual, as per IEC Publication 76-3, 1980, for detection by

shunt

insulated lank

winding

a b

Fig. 11 and tank-current mrthods

a Neutral-current method h Tank-current method

neutral current. Faults are created as explained above. The oscillographic records of current growth with faults are set against the current-wave structure without a fault, for detection and marking of the response characteristics and deviations therein on a time spread of 200 ps .

Figs. 12 and 13 depict the responses of the same faults against standard full-wave and chopped impulses ( t , = 4 p s ) having identical wave shapes and magnitudes.

It is observed that deviations due to a Fault in a line- end coil against full impulse, may be significant only in

Connection forfault re.qponse and detection hy neutral-current

Fig. 12 Fault responses to full-wave impulse uyainst unfuultrd current

Traces of neutral-currcnt struc~ure: from the top down: unlaulted-current, ldulls in A,. A, B,,, . C?+, , C T 9 Scales: 0.2 Vldivwon. 20 psidivision

the time domain exceeding 100 p s . The pattern of current growth is easily distinguishable for a midwinding fault but for the earth-end faults deviations, although present, are not markedly noticeable with reference to the unfaulted-current structure.

a

b

Fig. 13 by neutral-current method

U t -mm the top down: unfaulted current. faults in A,, A, h From the tap down;: faulty in B40. C,, , C,, Scales: 0.1 V/division. 20 fis/dnision

Fault response to chopped impulse against unfaulted current

With chopped-impulse application (t, = 4 p s ) , the current structure becomes highly oscillatory. Deviations are noticeable in the midwinding-fault response in the time domains 40--50ps and 120-160ps. It is observed that fault responses for the line-end coil show small devi- ations with some shifts of oscillation amplitude in the time phase. For the earth-end faults, however, small but detectable deviations are observed in the time domain exceeding 100 ps. Shifts in the time phase are also present in this time domain.

In the tank-current method (Fig. l l h ) of fault detec- tion the deviations (Figs. 14 and 15) in the fault response in a mid-winding coil are clearly traceable in the entire time spread scanned. For the line-end coils, the devi- ations are markedly present only in the initial spread spanning upto 40 ps. Here the fault responses are distin- guished by HF oscillation, the current beginning to build up. The earth-end coils are not responsive to faults when using the tank-current method of detection. This is estab- lished by fault response against both full and chopped waves, as shown in Figs. 14 and 15, respectively.

It may therefore be concluded that the response to a mid-winding fault may show recognisable deviations, whereas a fault in the extreme winding sections, specially

407 I E E Proc.-Gener 7’ran.sm. D!.strih., Vol . 141, N o . 5. Srpiemher IYY4

at the line-end, may not give sufficient indication of failure when using the neutral-current method of detec- tion. Detection will be easy by the tank-current method for faults occurring in the midspan and line-end. It may also be possible to specify the fault zone by proper and careful scanning of the oscillographic records of current response.

3.2.2 Voltage-sensitive short-circuit faults in the coils In this study, the faulting device shown in Fig. 50 is set to trigger short-circuiting of the coils at the desired voltage levels of fault initiation.

These simulated faults are established in the selected coils A, , A , , B,, , C,, and C,9, representing line-end, midwinding and earth-end sections, respectively.

Fig. 14 Lure by tank-current method From thetopdown: unlaultedcurrent,faullsin A,, A,, Bao,C,6 ,C,9 Scales: 0.2 Vldivision, 20 &division

Fault responses to $111 wave against unJaulted-current .sLruc- Fig. 16 srruclure by neutral-i urrenr method From the lop down: unlaulled current, fnulla in A,. A, . E,,, , C.*. t::." Scales: 0.2 Vtdivision. 20 p:division

Fault responses tofull-wave impulse against unfaulted-current

a

Responses to all these faults against both prescribed full-wave standard impulse and chopped wave, are observed in Figs. 16-19 and deviations are pointed out both for neutral-current (Figs. 16 and 17) and tank- current (Figs. 18 and 19) methods of fault detection.

Faults established in the line-end coils A, and A, are prominently responsive in the time domain beyond 1OOps although there are minor deviations within the domain of 100 ps. The major deviations are marked by damping and flattening of oscillation amplitudes in above time spread. There are occasional advances of oscillations in the time phase, as found by comparison with the unfaulted current structure by the neutral-current method.

Faults in the midwinding coil are indicated by promin- ent deviations in the entire time frame scanned. The smaller amolitudes in current growth for the unfaulted

b tures in the neutral current. However, fiults in the line- end coil and distinguishable by noticeable H F osci[lation at the beginning, spanning only the first 20 p s . Similar oscillations are noticeable in the corresponding current responses (Fig. 19) to chopped impulses. Corresponding

Fig. 15 by tank-currr,nr method U Prom the top down unfaulted current. iaults in A,. A , h bramlhctopdown faultsmB,, ,C,, .C,,

F'oult responses to chopped impulse ayainsl unfaultrd currmt

Scales: 0.1 Vidivision. 20 &division

4on I E E Proc..-Gmcr. Trunsm. Disirih., Vol. 141~ N o . 5 . .S<,pr<vnber IY94

fault-current responses to chopped impulses in the neutral current as well as tank-current methods differ appreciably (Figs. 17 and 19) but it can not he said with certainty that one is more responsive to faults than the

a

b

Fig. 17 .$tructure by neutral-current method c i From the top dawn: unfaulted current, faults in A, , A, I , From the top down: faults m B,,, , C,, , C,, Scales: 0.1 Vldivision, 20 psjdivision

Fault responses to chopped impulse ayainsl unfaulted current

Fig, 18 .structure by tank-current method From the top down: faults in A,, A,. Baa. C , 6 , C,, Scales: 0 2 V/division. 20 psidivision

Fault responses to Jull-wave impulse against unJaulted-current

other, except the initial point of HF oscillations in the tank-current response indicating a line-end fault.

For voltage sensitive faults, in summary, a midwinding fault is highly responsive to both the neutral-current and

I E E Proi..-Cener. Transm. Distrih., V o l . 141. N o . .7, Srptcmher IYY4

tank-current methods of detection. Line-end faults may easily be detected and identified by initial H F oscillations in the tank-current responses.

a

b

Fig. 19 hy tank-current method t i From the top down: unfaulted current, faults m A,, A, h From thetopdown:laultsin B,,,C,,.C,, Scales: 0.1 Vldivision. 20 psjdivision

Fault responses to chopped impulse against unfaulted current

These faults are further confirmed by marked devi- ations in structure and occasional phase advances of oscillation in the later (above 80 p) part of current growth.

Earthed end faults have responses similar to those in their line-end counterparts. The former is to be identified and distinguished from the latter by the tank-current method of detection, as already stated.

3.2.2.1 ’Hold’ voltage across a coil faulted by a voltage-sensitive short-circuit: The nature of ‘hold (residual) voltage on fault set against voltage oscillations without fault across the representative coils are shown in Figs. 20-22 featuring the line-end coils ( A Z r As), mid- winding (B40) and earthend coils (C,6 and C,,), respect- ively.

The magnitude of residual voltage on fault does not exceed 4 V with a positive polarity swing. For negative polarity swings, this is limited to 1.6 V, as indicated e a r I i e r .

Note that voltage oscillations across unfaulted coils, caused by a standard full-wave impulse struck across the model winding from RSG, maintain some similarity to the corresponding line-end and earth-end coils. However, for the midwinding coil, the oscillation pattern is quite different, indicating subsidence of reflected waves from

409

the extremities in about 200 p s and leaving this winding have a low probability of occurrence. Notwithstanding, section to standing waves. the most vulnerable sectors for such faults are tapping

leads of ratio switch to earth, line end to tank by dis- 3.2.3 Voltage -sensiitive shunt faults charge in trapped gas pockets, and discharges along the In a well designed and carefully assembled transformer, lower part of the winding limb to earth in a heavily transverse faults between winding elements and earth sludging transformer which has been in use without

proper maintenance for a long period. These shunt faults are mostly voltage sensitive, having

I threshold potential of fault initiation and nonlinear impedance character. In this study. these simulated faults have been made resistive and the series resistance i.; referrcd to 25"% of the winding impulse impedance.

Fig. 23 shows voltage-sensitive resistive shunt faults under full-wave impulse established in A,, A,, B,, . C-,,

lopper plot) und .A. (lower plot) s..,,iV, NI \' J t ~ r t o n , 20 M > diviwn

Fig. 21 S.J~C\ Io V J ~ L M ~ ? \ I , , * J t r w ! n

l r n . rllurion with )ut Juult trnd bi11d t,oltuW on / d u l l In I ] , ,

Fig. 22 (upper plot) and C , , (lower plot) Scales: I O V:dwismn, 20 &division

410

Free ascillarion withoutfuult and hold voltage onfault in C7h

b

Fig. 23 U From the top down: unfaulted current, laulls in A,, A, h Framthetopdown:laultsinB,,,C,,.C,, Scales: 0.2 V/division, 20 ps/division

Fault response to full impulse by neutral-current method

and C,, and set against unfaulted current. Deviations of fault current compared with unfaulted current wave are more pronounced in midwinding and earth-end coils than in the line-end. However, fault-current structures for coils beyond line end show noticeable reductions in amplitude.

The fault response against chopped impulses (f, = 4 ps ) starts to be oscillatory for A,. The pattern of oscil- lations is almost the same as for an unfaulted current. For a fault at A,, the pattern is retained with some dampening of current amplitudes after 100 ps . The struc- ture of unfaulted current and fault currents for A, and A, are hardly distinguishable from these presented in Fig. 13a, and hence these are not presented here. Significant deviations by pronounced dampening in midwinding

I E E f'roc.-Gener. 7ransm. Distrib.. Vol. 141. No. 5. September 1994

fault and reappearance of large oscillations within 150 p in earth-end fault currents become perceptible. These are, therefore, presented in Fig. 24.

Figs. 25 and 26 depict fault responses by the tank- current method. Initial fault-current amplitudes are obvi- ously quite high for line-end coils. The response of the

midwinding and rear coils shows dampening of oscil- lations after 40 ps . Fault responses in the earth-end coil can hardly be differentiated from the unfaulted-current structure.

Responses under chopped impulses are represented in Fig. 27. The pattern of current growth in line-end faults is

Fig. 24 From the top down: B,,, C,*, C,, Scales: 0 I V/division, 20 psldivision

Fault response to chopped impulse by neutral current

Fig. 27 method (1 From the top down: faults in A , . A , , unfaulted currenl Scales: 0.5 Vldivision. 20 ps/dwtsmn h From lhe top down: unfaulled ~ u r r e n l . faulu in B,,,. C 3 $ . C',, Scales: 0.2 V/division. 20 pi,'division

Fault responses to chopped impulse uoltage hg tank-currmr

almost retained, with a very sharp rise at the beginning. The fault-current structure is found to vary significantly

1 Fig. 25 From the lop down: faults in A,, A, , unfaulted current Scales' 0.2 Vidiviaion. 20 &division

Fault response to full-wave impulse b y tank-current method

afterward. with shift? of high-frequenc! oscillations in time phase. kor thc rcar end o f the winding, the responses are hardly distinguishable from thc unfaulted current. A small variation. though not remarkable. is. howcvcr. noticed it1 thc response of the midn'inding coil

4 Concluding comments

I- rom various observations made during thc investigation and from an in-depth meticulous scrutiny 0 1 the oscillo- graphic records. certain featurcs of uinding performance t o full-wave and choppcd impulses have been exposed and manifestcd. Thew are highlighted and :inal>sed to lead t o thc following inkrenccs and conclusions:

(I) A n arbitrarily chosen time-to-chop in chopped- impulw applicution enforces difkrent stress \ituations in dilreent coils along thc winding. 1 he critical time-to- chop-lor a coil incrc;ises from line-end coil t o earth-end

Fig. 26 From the top down: unfaulted current, laulln in Baa, [,,,, C., (ii) When maximum stresses are envisaged for earth- Scales: 0.2 V/divaion, 20 &division end coils, the line-end coils are also simultaneously

I E E Pro(.-Gener. Trunsm. Distrih., Vol. 141, No. 5, Seplemher 1994 411

Fault response to full-wave impulse by rank-current method COilr

stressed to about 85% of the maximum stresses expected in them.

(iii) The patterns of stress oscillation in corresponding line-end and earth-end coils are almost alike. In the mid- winding coil, a standing wave sets in after settling of reflected waves in about 200 ps. Coils at the extremities take longer to settle to standing waves.

(iv) In detection of short-circuited coil faults under full-wave impulse by the neutral-current (Hagenguth) method, pronounced deviations in current wave may be expected in the time domain not earlier than 100ps. Hence it may be advisable to scan the growth of current till 150ps. Deviations in current growth in earth-end faults are not significantly noticeable.

Fault responses to chopped impulses are comparat- ively more significant for midwinding coils than in end coils. Appreciable deviations may be noticed in time domain of 40-160 ps. Midwinding faults may be advant- ageously confirmed by the tank-current method of detec- tion. Earth-end coils are not responsive to faults using the tank-current method. Both the methods of detection may be suitably employed to specify the fault zone in the winding.

(v) Responses to voltage-sensitive short-circuit faults in end coils are almost alike, showing noticeable deviations in the time domain exceeding 100 1s. Responses to mid- winding faults may also be prominent in the entire domain 0-250 ps to be scanned.

(vi) Line-end faults are indicated by initial H F oscil- lation in the current structure yielded by the tank-current method of detection, which makes it confirmatory.

(vii) Responses to voltage-sensitive shunt faults against both full and chopped impulses, are marked by high amplitude of initial fault current in the tank-current study for the line-end coils. Such faults in midwinding and earth-end sections are also significantly responsive. In the neutral-current study shunt faults in line-end coils are less responsive than those occurring elsewhere.

(viii) It may be often judicious to employ both the Hagenguth method and the tank-current method to

detect, confirm and approximately locate the fault zone in the winding. It should, however, be remembered that each method should be applied separately and not simul- taneously.

5 References

1 WAGNER, K.W.: ‘Das Eindrigen einer elektromagnetischen welle in &ne spule mit wind ungska padtal’, Elektrotech. Masch. bau, 1915, p. 89

2 BLUME, L.K., and BOYAJIAN, A.: ‘Abnormal voltages within transformers’, Trans. Amer. Inst. Electr. Eng., 1919, p. 577

3 HELLER, B., HLAVKA, J., and VEVERKA, A.: ‘Narazove zjevy v transformatorah - Parts I and II’, Elektrotechn. Obz.. 1948, pp. 93, 100

4 ABETTI, P.A.: ‘Transformer models for the determination of tran- sient voltages’, Trans. Amer. Inst. Electr. Eng., 1953, p. 468

5 RUDENBERG, R.: ‘Performance of travelling waves in coils and windings’, Trans. Amer. Inst. Electr. Eng., 1940, p. 1031

6 MIKI, A., HOSOYA, T., and OKUYAMA, K.: ‘A calculation method for impulse voltage distribution and transferred voltage in transformer’, IEEE Trans., 1978, PAS-97, p. 930

7 FERGESTAD, P.L. and HENRIKSEN, T.: ‘Transient oscillation in multi winding transformer’, IEEE Trans., 1974, PAS-93, p. 510

8 MUNSHI, S.: ‘Digital computation of potential distribution in a transformer winding due to lightening impulse’. MEE thesis, Jadav- pur University, 1985

9 HAGENGUTH. J.H., and MEADOR, J.R.: ‘Impulse testing of power transformers’, Trans. Amer. Inst. Electr. Eng., 1952, 71, pp. 697-704

IO ABETTI, P.A., and DAVIS, H.F.: ‘Surge transfer 3-winding trans- formers’, Trans. Amer. Inst. Electr. Eng., 1954, 73, pp. 1395-1407

I I MUNSHI, S.. ROY, C.K., and BISWAS, J.R.: ‘Computer aided analysis of surge performance of transformer winding with one end earthed‘, .I. Inst. Eng. (India), Electr. Eng. Diu., 1989, 70, (ELl), pp. 9-23

12 MUNSHI, S., ROY, C.K., and BISWAS, J.R.: ‘Computer studies of performance of transformer windings against chopped impulse volt- ages’, IEE Proc. C, 1992,139, (3), pp. 286-294

13 ‘Power transformer - insulation levels and dielectric tests’, IEC Publication 76-3, 1980

14 HELLER, B., and VEVERKAR, A.: ‘Surge phenomena in electrical machines’ (ILIFFE Books Ltd., London, 1968)

15 STEIN, G.M.: ‘A study of initial surge distribution in concentric transformer winding’, IEEE Trans., 1964, PAS-83, pp. 877-893

16 GOLDING, E.W., and WIDDIS, F.C.: ‘Electrical measurements and measuring instruments’ (First Indian Edition, 1979)

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