Saturation Curve

Curves

of

Delta-Connected Transformers from Measurements

Washington

L. A.

Neve;

Student Member, IEEE

Fellow, IEEE

Hermann W. Dommel

Department of Electrical Engineering

University of British Columbia Vancouver, B.C.,

Canada

V6T 124

Saturation

An algorithm is described

Abstract

for the computation characteristics of three-~hase deltacoMected

of the saturation transformers from measured

V-

-

Z-

curves and no load losses at rated frequency. For these transfor", positive sequence excitation tests are usually carried out with the delta connection closed. While triplen harmonic currents are present in the deltaamected windmgs, they do not appear in the line connections to the three-phase source. This paper presents an algorithm which accounts for the missing triplen har- monics on the line side in the derivation of the saturation chancteris- tics.

Key words:

Transformer, delta connection, saturation, harmonics.

1

Introduction

insulation coordination studies are needed to specify the insulation levels of power apparatus and systems. These studies are usually done with digital computer programs, such as the EMTP (Electromagnetic Transients Program).

They

require mathematical models for the various power system components.

The type of model depends on the type of study. For inrush current and ferroresonance studies, models for transformers and reactors must represent saturation enects reasonably well. For this, the instanta- neous saturation characteristic, which gives flux llnkage

R

as a function of current i , is neded. Numerical methods have been used for some time to produce these peak fluxcurrent curves from measured

Vms

-

I-

curves[lJ].

A simple circuit consisting of a nonlinear resistance in parallel with a nonlinear inductance represents the transformer

core

reasonably we11[3,4]. An algorithm to obtain the nonlinear resistance (piecewise linear v

-

i, curve) and the nonlinear inductance (piecewise linear

L

-

il curve) from the measured

V-

-

Z-

curve and no-load losses was presented in [SI. This algorithm assumed all cdd harmonic current components to be present in the measured values.

On leave from Universidade Federal da Paraiba, Campina Grande

-

PB

-

Brazil.

For three-phase transformers, the standard excitation test data available are the positive sequence V-

-

Z-

curves and no-load losses. Figure 1 shows a symmetrical three-phase voltage source supplying a no-load delta-connected transformer. The

delta

branches consist of nonlinear elements. In general, excitation tests are carried out with a closed delta. In that case, ammeters, placed in series with the line, will not contain the triplen harmonic currents because these circulate in the delta connection.

This

paper presents a method for genera- the piecewise linear saturation curves (nonlinear resistance and nonlinear inductance), which accounts for the

fact

that triplen harmonics circulate in the closed delta, but do not appear in the measured line currents.

Figure 1: D ella-con n e c l e d transform er- Posit i v e seq u en c e e: c i t at i on test.

2

Basic Considerations

In the circuit of Figure 1, the three branch elements of the delta

connection

are assumed to

be nonlinear and identical.

The branch

currents can be Written as a Fourier time

series

containing

odd

harmonic components only. Then:

The triplen harmonic currents (Z3. Zg ...) are in phase (zero sequence harmonics).

The nns current in each branch is 94 SM 459-8 PWRD A paper recommended and approved

by the IEEE Transmission and Distribution Committee of the IEEE Power Engineering Society f o r presentat- ion at the IF,EE/PES 1994 Summer Meeting, San Francisco, CA, July 24

-

28,

1994.

Manuscript submitted December 30, 1992; made available for printing April 20, 1994. Thelinecurrentsare: i a ( r ) = id(r)-ic,,(r) i b ( r ) = i k ( r ) - i d ( t ) i c ( r ) = ic,,(r)-t&)

.

(4) 0885-8977/95/$04.00 0 1994 IEEE

From equations (1) to (8) one can make the following observations: 0 triplen harmonic currents, although present in each branch, are not present in the line currents.

if triplen harmonic currents in each delta branch are removed from their nns values (equation (4)) and scaled by

4 ,

the nns

line currents (equation (8)) ax obtained. This information is the basis of the algorithm developed next.

3

Saiuration Curves

Each delta branch in Figure 1 is represented by a nonlinear inductance in parallel with a nonlinear resistance (Figure 2). Their nonlinear characteristics are computed with the following assumptions' :

the v

-

i r A and A

-

i[A c w e s (Figures 2b and 2c) are symmetric with respect to the ongin (Rk and & are the slopes of segment k of the v

-

ij-A and 2

-

i / A curves, respectively);

the transformer winding resistances and leakage inductances are ignored.

The algorithm works then as follows:

1. For the construction of the v

-

irA curve (Section 3.1):

compute the peak values of the branch current irAl, i r A p . . point by point from the n*load losses.

2. For the " c t i o n of the A

-

i/A curve (Section 3.2):

0 from the v

-

irA c w e , compute the nns values ZrA-m, remove

the triplen harmonic

currents and

obtain the resistive

h e

current Ir-nns;

obtain the nns values I/-- of the line current due to the nodinear inductance fkom I,---, the total line current It--,

end

the applied voltage V ;

compute the peak values of the inductive current i/Al, i/&,

...

point by point iteratively.

3.1

Computali'on

of the

v

-

irA curve

Let us assume that the three-phase ndoad losses Pi,

P2,

...,

P m are available as a function of the branch voltages VmSl,

Vm2,

...,

V-,,,

(Figure 3).

If we

assume

that the applied voltage

is

sinusoidal, the conversion of rms voltages to peak values (vertical axis of Figure 2b) is simply

vk =

f i

( 9 )

for k = L Z . * * , m .

a

Figure 2: ( a ) N on lin ear elem en Is; (b) y - i,, curve; (c) 1- i,,curve.

I

Due to symmetry reasons, voltage and current waveforms need only be evaluated over 114 of a cycle. For a sinusoidal voltage v(6) =

vk

sin

e,

pk can be written in the form2:

f

f l

For the

f i

linear segment in the v-rd curve, the current is sinusoidal. The computation of the frrst peak current frAl is therefore straightforward.

Since

P 1 = 3 v m ~ , I m , in the linear case,

From the second segment onwards (t 2 2), equation (10) is evaluated at each segment A , with only i r k being unknown, as explained in more detail in Section 2.1 of (51. The computation of the peak current i r b is done segment by segment,

starting

with i.4 and endug with the last point i r b . Whenever a point i r 4 is found, its mu value is calculated as well. A Fourier program (Appendix A) is used to cam- pute the triplen harmonics (13. 19, ...). They are removed from I r A m s to obtain Ir-ms, which is needed later for the construction of the peak flux

-

peak current

II

-

ilA curve.

3.2

Computation of the

curve

The conversion of the peak branch voltages vk to flux

4

is again a re- scaling procedure. Hence, for each linear segment in the

II

-

i/A m e ,

A&+ (12)

Let us now compute the peak values of the currents il4 through the nonlinear inductance. First, the mu values of the line currents II+,,,~ are evaluated with

where the line current It+,,,, is available f h m the measurements, and

where

For the first linear segment, the computation of iiA, is straightforward since there are no harmonics yet. Therefore,

has already been computed h m the previous section.

From the second segment onwards

(k 2

2) the algorithm works iteratively as follows (see R

-

ilA curye in Figure 4):

1. guess ig;

2. with A((B)=Aksine, find 1/4 of a cycle of the

distorted

current d*cally,

3. compute the nnr inductive branch current whose peak is ig ; 4. use a Fourier program to find the triplen harmonic inductive currents in the delta branch (Appendix A);

5. remove the triplen harmonics from the estimated nns branch current. Scale the estimated result Ilest by

Js

and compare it to Il-- in equation (1 3);

6. if the absolute value of the difference id = I 1 - w

-

is less

than

the specified tolerance, convergence is achieved. Otherwise, the residue id is added to ig and the iterative process i s repeated fitnu step 2 onwards.

In general, convergence is achieved in less

than

20 iteration steps with a tolerance of ilA1 and initial guess

for every

k

2 2. In some cases as many as 40 iteration steps may be n - = = Y . c o r r e c t c u r v e

‘T

.

x 8

T

81 ZT 2 8 F i g u r e 4 : G e n e r a t i n g c u r r e n t w a v e f o r m from s i n u s o i d a l f l u x .

4 Casestudy

Consider a 50

Hz

three-phase five-legged core type transformer. The following infomation is known[6]:

2. rated voltages

-

420

kV/

27

kV

(line to line values);

3. wye connection on 420

kV

side, delta connection

on

27

kV

side. Table I

shows

the data

from

the positive sequence excitation tests measured at the closed delta 27

kV

side.

1. rated power

-

750 MVA (three-phase);

Table I: Three-phase transformer test data

I

v m

I

It-m

I

P

orv)

I

(A)

I

orw

22.76 8.20 206.21 24.29 25.64 27.00 27.50 28.47 29.10 32.50 11.35 15.50 21.16 24.68 31.63 38.30 80.97 240.26 270.13 311.00 323.03 355.48 385.41 560.00

Figure 5 shows that the correct curve goes deeper into saturation. For the flu value of 1.2 P.u., there is

a difference

of approximately 14% between the peak currents of the incorrect curve

and

the wrmt one. In

facf

for transient studies, it is often neceSSafy to know peak flux- peak current curves beyond that point. The usual way is to extend the curve up to a value necessary for the study (this extension is sometimes done with

a

straight line passing

through

the previous to the last and the last point in the peak flux-peak current curve).

This

may lead to larger errors for the peak values of the cun-ent since the curves diverge as the flux goes up towards deep saturation. The transformer magnetizing current would always be underedmated if triplen harmonics inside the delta windings were not taken into

accoullt Errors can also affect the air core inductance value. A parametric study was done considering typical values for the air core reactances (0.2 p.u to 0.5 P.u.). For the study, a straight line segment was connected to the last point of the correct curve of Figure 5 . Errors on the slope of the saturation characteristics, for this case, are between 18% and 25% when the magnetizing current reaches the transformer rated current. These differences may be important in ferroresonance or inrush current studies.

The proper way to represent saturation effects is to look at the transformer magnetic circuit. Saturation is related to fluxes in the wre and tank. There tire several types of core construction. To model the core reasonably well, one should know what kind of geometry the transformer has. One should h o w if the core is three-legged, five legged or shell-type. The knowledge of the zero sequence magnetizing impedance is also required. It can be estimated for each type of core coIlstruction

[A.

The saturation curve must be obtained for each leg of the transformer and placed across the lowest voltage

terminal.

In the iteration scheme of Section 3.2, harmonics up to the 99* order

were included. An average of 23.86 iteration steps was necessary (the "um number of iterations was 34). In order to check the numerical accuracy of the method, the m line currents were were found to be very small (less

than

0.001%).

recomputed back from v

-

irb, and 1

-

jib, Curves. errom

Vms is the m line to line excitation voltage, Ir-ms is the nns

excitation current (three-phase average) and P are the n d o a d losses (three-phase values).

Figure 5 shows two computed 2

-

jib, c w e s (points connected by straight line segments). One of them assumes that all odd harmonic components of the current are present in the measured values and it is therefore incorrect. The other curve is the correct one; it has

been

pro- duced with the algorithm of Section 3.2. Figure 6 shows the piecewise linear v

-

irA curve.

1 . 2

-

1.0 - 0.8

-

o ' 2

1

0.002 0.004 0.006 0.008 0.010 0 . 0 1 2 C u r r e n t

(

p. U

.)

Figure 5:

A -

i i a c u r v c . 0.0005

C u r r e n t (

P.u.) Figure 6 : V - i f A c u r v e . I 0.0010

5

Uncertainties

To

model

transformers exactly is very complicated, if not impossible. The ferromagnetic properties of steel laminations are not well understood yet. There is no theory available today that could predict losses accurately as a function of frequency, even for low kequencies [8,9,10]. There is a big discrepancy between calculated and measured losses due to the very complicated domain structure of these materials, There is also the

further

discrepancy #at transformer core loss pa kilogram is always greater

than

the nominal loss of the steel as measured in standard testers. The ratio between the transformer

per

unit loss and the nominal or standard unit loss is the so-called building factor of the core. Building factors usually range from 1.1 to 2.0. The extra loss is due to factors such as [ l 11:

0 Non-uniform flux distribution due to difference

in

path lengths

0

Distortion

of

flux

waveform due to magnetic saturation;

0 Circulating flux due to magnetic anisotropy of m a w , 0 Flux directed out of the rolling direction;

0 Transverse flux between layers due to joints. among magnetic circuits;

n

e

flux distribution

in transforma laminations

isn0tl"eVtn at low

fkqueames.

For

a Sinusoidal applied flux, the flux m each lamhation is, in genaal,notsinwidal although the fluxcomponents addup to produce the sinusoidal total flux [12. 131. These at- Mcertain tiesmironm modelling. Amther major problem is the availability of data

Flux Linkage(V.s) 0.000000 0.040327 0.059046 0.075777 0.095659 0.114790 0.151553 0.169560 0.189066 0.207823 The core nonlinear elements, in this paper, are obtained from one

single frequency. According to this model. transformer no-load losses would then be underestimated for frequencies below the rated frequency and overestimated for frequencies above the rated frequency. However, ferroresonance and inrush current studies show that the

power

system is much more sensitive to variations of the saturation curves

than

to variations of core losses.

Figure 7 is the measured positive sequence excitation characteristic of a three-phase five-legged 75 kVA distribution transformer (secondary side

-

120 V). The x-axis is the ratio of the no-load exciting current

I , (measured in each phase) to the no-load current at rated voltage

I-,a#d (average value of rated current in all phases) for the same transformer. The solid line represents the average curve.

1.4,-

- -

-

---,-

-

7-7 Cment(A) 0.0000000 0.426281 0.747657 0.905318 1.1933 16 1.240726 0.815392 0.396883 0.123851 0.162665 2 4 6 8 1 0 lexcdexc- rated

Figure 7: V,,

-

1,characteristic for three-phase transformers There are discrepancies below the rated voltage, but the curves are very close as the transformer goes into deep satmition.

The V ,

-

I , curve shown in Figure 8 was obtained from a recently manufactured 10 kVA single-phase distribution transformer

.

The nns current at 90% of the rated voltage is smaller

than

the m s current at 50% of the applied voltage. Here, the exciting current in the unsaturated region is very much

affected

by stray capacitances. As the transformer goes into deep saturation, stray capacitances aEkct very little the computation of saturation curves.

rms C u n m t (A) 1

(a) (b)

Figure 8 a)V,

-

L

curve for a recently manufactured transformer. b) Corrupted fluxcurrent curve (not to scale).

A crude estimation of the saturation curves and the open-circuit capacitance

can

be found iteratively. First, the algorithm generates the peak fluxcurrent characteristic "corrupted" by stray capacitance effects as shown in Figure 8b. The peak current is decreasing in the region between points A and B. The inductance is very high in this region. It can be assumed that it is

infinite

between turning points A and B. The peak cucrent 1. through the open circuit capacitance C, is given by the difference between the current at turning point A and

the current at any point in between A and B. For instance at point B, the open-circuited capacitance can be obtained by:

Copon =

+.

m AB

Table

II

shows *corrupted* fluxcurrent points obtained using Figure Sa as input data.

The following information is needed: peak current at turning point A = 1.240726 A; peak current at tUming point B = 0.123851 A; peak flux at turning point B = 0.189066 Vs.

Then, the open circuit capacitance referred to the 120 V side is

Copen = ( 2 z x 60)2x0.189066

1.240726- 0.123851 ~ ~ 4 1 . 5 6 ~ . If

.

,

C

is referred to the primary side (14400 V),

Copen =41.56~(120/14400~~2.89~F.

Table

II:

Corrupted fluxcurrent curve.

To represent the transformer exactly, one should represent saturation, hysteresis and eddy currents in the core as well as eddy currents and stray capacitances effects in the windings.

Modern

power transformer cores have very low losses and saturation is the predominant effect. In this paper the main focus is on estimation of the saturation curves of transformers.

4

Conclusions

An approach for the computation of instantaneous saturation curves of deltaannected transformers has

been

presented. It uses positive sequence excitation test data as input, and is suitable for situations in which the tests are performed

with

a closed delta.

Once the v

-

irA and R

-

ilA curves have been obtained, they can then be used to model the excitation branch of transformers in transient and harmonic studies.

7

Acknowledgments

The authors would

like

to

thank

the r e v i m for their valuable suggestions. Also. the authors would like to

thank

Powatech Labs Inc. for providmg the

t

"

m

test data of Section 5.

The financial support of Mr. Washington Nevw from Univmidade

Federal

da Paraiba, Campina Grande

-

PB

-

Brazil, and

from

The University of British Columbia is gratefully acknowledged.

1437

8

References

1.

S.

N. Talukdar, J. K. Dickson. R. C. Dugan, M. J. SprinZen, C. J. Lenda, On Modeling T m f o n n e r and Reactor Satumtion Chamcteristics for Digital and Analog Studies, IEEE Trans. on PAS, vol. PAS-94.1975, pp. 612621.

2. S. Prusty d M. V. S. R m , A Dimxtp'ecewise Linean.zedAppd

toGmveli musalumtionchomcrenstr

.

'c to Instantaneous Wmtim

CUM, IEEE Trans. Mag., vol. Mag-16, NO. 1, J ~ ~ M I Y 1980, p ~ . 1 5 6 160.

3. L. 0. Chua and K. A. StrO"oe, Lumped Circuit Models for Nonlinear Inductors Exhibiting Hysteresis Loops,

IEEE

Trans. on 4. J. G. Santesnases, J. Ayah, S . H Cachero, Analytical Approximation

of Dynamic Hysteresis Loops and its Application to a Series Ferroresonance Circuit, Proc.

IEE

117, No. 1, J a n 1970, pp. 234-240. 5. W.

L

A N e v e H. W. Donrmel, On Modelling Imn Core

Nmlinamities.

IEEE

T d o n s on Power Systems, vol. PWRS-8. h4ay 1993, pp. 417425.

6. ATP Rule Book, Leuven

EMTP

Center (LEC), Section XIXG, Revision July, 1987.

7. H. W. Dommel, Electromagnetic Tmnsients Progmm Reference Manual, Section 6, Department of Electrical Engineering-The University of British Columbia, Vancouver, 1986.

8. G. E. Fish, So) Magnetic Materials, F'mxdqs of the

IEEE,

Vol. 78, No. 6, June 1990, pp. 947-972.

9. T. H. O'Dell, Fe"amtoC+"ics, Chapters 1 and 5, Mcmillan Circuit Theory, vol. CT-17, NO. 4, NOV. 1970, pp. 564-574.

For segment R = 2 ,

1 .

Substituting

r

-

-!-= !!h and

r,

=

-

mto the above equation, the fundamentalcurrent(n=l) anditsharmonics(n23)are:

' - 4

4

4

4 4 =

r24

+

$rI -

r2

x+z

+ 4 C O 4 ) 1 10. 11. 12. 13. PressLtd,LonQns 198L and

G. Hener and H. Hilzinger, Recent Developnents in Sop A h p d c

4 - 1

Matetials, physlat

Scripta,

Vol. T24,1988, pp. 22-28. b , , = L ? { - , - r l ~ ~ ~ n , e f - l ~ ~ k + - C ~ n ~ j - l

T. Nakata, Numerical Analysis of Flux and Loss Distribution in 1-2

Electrical Machinety, (Invited Paper),

IEEE

Trans. on Magnetics, Vol. Mag

-

20, NO. 5 , September, 1984.

Ponw

Larw

in Elecbiml Sheet Steel. Proc. IEE Vol. 112,

NO.

4,

where

r,

=

i

'm, =

$of

-

i s i n 20,)

F. J. Willcins and A E. Drake, Meann"t and Intmpretation of Lk '

April, 1%5, pp. 71-785. and

Behaviour In a

IOOkVA

Disbibutim

Tmnr/onner

Core, IEEE Trans.

sin[(n-lP,I

-

sin[(n+l)811 A.

Basak

and

AAA

Qader,

FtI"enta1

and Hanncnic

Flux

gm(n.w =

qn-1)

2(n+l) onh4ag.,Vol. Mag-l9,NoS,Septanber 1983,pp. 2100-2102.

Appendix A

-

Computation

of HmmOniC

Componenis

Although only the computation of triplen harmonics of the current is needed, it is appropriate to show the derivation of a l l odd harmonic components that may be produced by saturation curves during excitation tests. The equations below are developed for nonlinear inductances. For nonlinear resistances, one needs to replace 1 by V and L by R accordingly.

Consider the piecewise nonlinear inductance of Figure 2c. For a sinusoidal flux A(@) = sin(@, the current can be written in a Fourier series form containing odd harmonics only. So,

i,A (@=

2

b,,sinne (A. 1 )

R I

for n=l,3,--.,

with

- -

Washington

L

.

A. Neves was bom in Brazil on March I, 1957. He re- ceived the B. Sc. and M. Sc. degrees in Electrical

Engineaing

h m Universidade Federal da Paraiba in 1979 and 1982 respectively. From

1982 to 1985 he was with the Departmat of Electrical Engineerkg of Faculdade de Engenharia de Joinville, Santa Catarina, Brazil. Since November 1985, he is with the Department of Electrical hguleering of Universidade Federal da Paraiba, Campina Grande-PB

-

Brazil. He is currently a Ph. D candidate at the University of British Columbia, Vancouver, Canada.

Hennunn K Dommel was born in Germany in 1933. He received the Dip1.-Ing. and Dr.-Ing. degrees

in

Electrical

hgu~eerhg from Technical University, Munich, Germany in 1959 and 1962 respectively. From 1959 to 1966 he was with the Technical University Munich, and from 1966 to 1973 with Bonneville Power Administration, Portland, Oregon. Since July 1973 he has been with the University of British Columbia in Vancouver, Canada.

Dr.

Dommel is a Fellow of IEEE and a registered professional engineer in British Columbia, Canada.