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8/7/2019 Papers of Hve

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LIST OF PAPERS

1. Electro-Thermal Analysis of an Induction motor

Abstract- The purpose of this paper is to presentthe temperature distribution of an induction

motor during no load and blocked rotor

conditions, where the phenomenon of

electromagnetically induced currents and heat

transfer are coupled under balanced and

unbalanced conditions. In this paper

electromagnetic and thermal fields of an

induction motor of rating 3 phase, 400v ,1hp is

verified using Finite Element method(ANSYS)

and experimental setup

Keywords squirrel cage induction motor,

electromagnetic model, thermal model, ANSYS.

I.INTRODUCTION

In all electric machines the electric, magnetic and

thermal processes are intrinsically coupled each

other. The temperature distribution is dictated by

power loss, which is in turn affected by the

temperature dependence of the properties of the

conducting and magnetic materials. Therefore

electromagnetic and thermal behaviors are

interdependent. The performance of the electric

machines depend on temperature distribution inside

the machine

Changes in the electrical conductivity lead

to a different slip in induction motors and

changes joule loss distribution[2].

The lifetime of machine firmly depends on

the hot spots in the insulation.

The prediction of these temperatures is important

and requires the solution of a coupled set of partial

differential equations representing electromagnetic

and thermal diffusion. This paper presents the

general purpose finite element method (ANSYS) for

the coupled set of electromagnetic and thermal. TheANSYS thermal-electric analysis only accounted for

the Joule heat as a coupling mechanism between the

thermal and electric fields.

. II. ELECTROMAGNETIC FIELD MODELING

In this paper, the two-dimensional model of an

induction motor is shown in Fig. 1. Since the

electromagnetic end-effect in radial field squirrel

cage induction motor is not very significant. So, the

two-dimensional FEM is preferred over three

dimensional FEM for electromagnetic analysis.

Considering the cross section of the induction motoras the X-Y plane, the partial differential equation(1)

describing the two dimensional domain[3,4,5]

Fig.1. Model of an induction motor

The Maxwells equation leads to

ezz

ez

e jAjsy

Av

yx

Av

x=

+

)()(

(1)

Where zA is complex vector magnetic

potential, ev is reluctivity, is conductivity S/m,je is applied current density A/m

2, angular slip

frequency.

The corresponding energy functional is given by

F(A) = d xAjAjy

Av

x

Av ee }

2

1])()([

2

1{ 222 +

+

(2)

The boundary conditions are [3]

4

21

3

0

TT

TT

AA

AA

=

==

Where 1T

& 2T

= outer and inner surfaces of

3T& 4T=left and right surfaces of

III.THERMAL FIELD MODELLING

120


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The axial heat flows from the end winding to

the winding in the core is so negligible that the two

dimensional thermal analysis is well accepted .In

this paper, a 2-D finite element model of a thermal

field over the whole cross section of a 1HP squirrel

cage induction motor is developed in order tosimplify the analysis. Some assumptions are

made[6]

Non-axial thermal heat flow

No heat flow from rotor core to the shaft

At steady state, concerning a homogenous and

isotropic medium with constant thermal conductivity

, the 2-D heat diffusion equation(3) in Cartesian

coordinates is given by[5,8,9]

vqy

Tx

T =+

2

2

2

2

(3)

Where T is the temperature in K,is the

thermal conductivity of the medium W/m.K and

vq are the heat sources per unit volume W/m3.

In the outer surface, the boundary condition (4) is

given by

0)(. =+ fTTnT

(4)

Where n is the unit is normal to the outer surface,

is the convection factorW/m k.2 and fT is the

ambient temperature.

In the inner surface of the rotor core, there

is absence of tangential heat flow, then the boundary

condition (5) is given by

0. = nT (5)

Where n is the unit vector normal to the boundary

The corresponding functional equation is presented

as formula (6)[5]

+

=D D

v dsTqdsy

T

x

TTJ

2

1)(

22

( )TdLTTtt

f32 ,

2

=min (6)

Where D is the solving region

IV. COUPLED ELECTROMAGNETIC-

THERMAL MODEL OF AN INDUCTION

MOTOR

The electromagnetic and the thermal areweakly coupled because the time constants for these

two aspects are dissimilar. The variation of the

magnetic quantities is much faster than that of the

thermal ones. Hence, at every time step a magnetic

solution is performed taking into account the rotor

slip. Rotor slip is decreases related to the time

step[1].

The electromagnetic motor model for the

induction motor is constructed in the stator reference

frame. Its inputs are the torque reference, the voltage

and the frequency, and the outputs are the currents

and the electromagnetic torque of the motor. Inputsfor the thermal model are the different losses that are

calculated according to the output currents of the

electromagnetic model.

Resistance parameters of the EM model are updated

according to calculated temperatures from the

thermal model. Also the resistances in the loss

calculation block are updated, as the resistive losses

strongly vary with the temperature. The block

diagram of the entire model, where different blocks

and signals can be shown in Fig.2 [1,7]

121


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Fig.2.Block Diagram of coupled electromagnetic

and thermal model

A. EFFECT TEMPERATURE ONPERFORMANCE:

The temperature rise in electrical machines leads to

[7]

Change in motor resistance parameter

which makes the characteristics of the

motor to be also temperature depended.

Increase of the stator resistance will

decrease the torque production capability,

Increase in the rotor resistance effects

mainly on the slip and rotor losses

In order to ensure high dynamic performance, torque

production capability should be high and also the

slip low.

Induction motors are dynamically high-performance

machines, the common method to increase the

dynamic performance is to use high flux densities.

High flux density causes

Saturation in induction machines

Decrease in magnetizing inductance

Increase stator copper losses

Decreased power factor and

Decreased torque-to-current ratio.

The rise of the temperature of the rotor bars by

1 C will increase the slip approximately by 0.4 % at

the rated point. A rise of 100C of the rotor

temperature therefore means that the slip will

increase by 40%[7].

The high slip causes

Increased rotor losses

Lower air gap flux

Complicates the control

These drawbacks can be avoided by properly

adjusting the flux level, depending on the operation

point

The slip of an induction motor is

determined by the rotor resistance The increase of

the slip due to temperature can be compensated by

increasing the air gap flux of the motor, as the slip is

approximately proportional to the inverse of the

square of the air gap flux. As the increased rotor

resistance causes the motor to run at higher slip, a

bigger proportion of the air gap power is transferred

into the heat at the rotor bars, which furth

increases the rotor resistance

B.. EFFECT OF THE FLUX ON THETEMPERATURE:

As the torque of an electric motor i

proportional to the Square of flux density , it is

beneficial to use high flux densities in high torque

machines[7].

The flux level has a significant effect on

the motor current, especially at partial loads,

because the share of the magnetizing current is

higher at smaller loads. But even at the rated load, a

decrease of 20 % of the air gap flux decreases the

current by nearly 4 A [7]. As the stator iron losses

are proportional to the flux density squared and the

copper losses to the current squared, optimizing theflux level can have a huge effect on the losses and

consequently on the heating of the machine.

.

V. ELECTRO-MAGNETIC RESULTS

The squirrel cage induction motor , which is

considered for simulation purpose under healthy

operation and unbalance condition.

A. Distribution of magnetic field:

Under the rated load condition, the distribution of

magnetic field for the case of healthy condition issymmetrical shown in Fig.3, but it is distorted in the

case of unbalance condition shown in Fig.4. and a

higher degree of magnetic saturation can be

observed around the broken bars as a results of the

lack of local demagnetization slip frequency induced

currents in these rotor slots, which might result in a

degradation mechanical performance of the

induction motor[10].

Fig.3.Distribution of Magnetic Field

Fig.4.Distribution of MagneticField for One Broken Bar

122


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B. Iron core loss density distribution on rotor:

The iron core losses distribution on rotor tooth

adjacent to broken bars is computed in each element.

We can observe that the regions in the vicinity of thebroken bars in the rotor have a much higher core

loss density as compared to the other regions of the

rotor.

Vi.THERMAL FIELD ANALYSIS AND

RESULTS

The temperature distribution of the motor operating

at the rated speed is estimated using Finite Element

Method (ANSYS)

Fig.5. Thermal Field Distribution

It can seen that the temperature of the rotor

is highest and that because of the large thermal

conductivity of the rotor core and the bars. The

temperature distribution throughout the radial cross

section of the motor is given as Fig.5

Fig.6.Radial Temperature Distribution

Vii.RESULTSThe induction motor has the following ratings:

3 phase, 1440 rpm,1 hp , 50 Hz and a squirrel cage

rotor is studied under No load in Fig.7 and Blocked

rotor conditions inFig.8

V. CONCLUSIONS

Fig.7.No Load Temperature Distribution

Fig.8. Blocked Rotor Temperature Distribution

Fig.9.Radial Temperature Distribution

Viii.CONCLUSION

In this work, the coupled analysis of an induction

motor is verified using ANSYS and the temperature

distribution of the motor operating at No load and

Blocked rotor conditions are shown. Furthermore,

estimate the magnetic field distribution at healthy

and broken bar conditions. A higher degree of

magnetic saturation can be observed around the

broken bars as a results of the lack of l

demagnetization slip frequency induced currents in

these rotor slots, which might result in a degradation

mechanical performance of the induction motor. .Itcan be seen that the temperature at rotor is highest

because of the larger thermal conductivity of the

rotor and the bars, temperature difference between

the rotor core and bars is rather low.

REFERENCES[1] Smail Mezani, N. Takorabet, and B. Laporte, ACombined Electromagnetic and Thermal Analysis of Induction

Motors A Combined Electromagnetic and Thermal,IEEETrans. Magnetics, vol. 41, no. 5, pp.1572-1575 ,may 2005

[2] Johan Driesen, Ronnie Belmans and Kay Hameyer,Coupled Magneto-Thermal Simulation of Anisotropic

Machines. IEEE Trans Magnetics ,pp 469-471, jul 1999.

123

TEM PERATURE DISTRIBUT

20

40

60

mpind

eg


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[3] C.C. Chan, Tietong Yan, hang Chen, Qezhong Wang,X.T. chau, Analysis of Electromagneitc and thermal

fields for Induction Motorsduring starting ,IEEE

Trans. Energy Conversion, Vol. 9, No. 1, Mar1994.

[4] V. K. Garg and J. Raymond, Magneto-Thermalcoupled analysis of Canned Induction Motor,IEEE

Trans. Energy Conversion, Vol. 5, No. 1,pp.110-113,

Mar 1990.

[5] Xie Ying, Li Weili,,Li Shoufa, ElectromagneticField and Thermal Field on Asymmtrical peration used

in Electric Vehiclec, IEEE2006.

[6] P. K. Vong and D. Rodger,Coupled

ElectromagneticThermal Modeling of Electrical

Machines,IEEE Trans. Magnetics, vol. 39, no. 3,may 2003.

[7] J. Puranen and J. Pyrhonen,Optimization of the

loadability of an induction servomotor with a coupled

Electromagnetic Thermal model, IEEE Trans.Magnetics, vol. 52, no. 3, july 2006.

[8] Ying Huai ,Roderick V.N. Melnik, Paul B.

Thogersen,Computational analysis of temperature rise

phenomena in electric Induction Motors,Applied ThermalEngineering23 (2003),pp. 779795.

[9] J.Sakellaris, j.Xypteras and T.Tsiboukis, A Coupled

Magnetic Thermal Model for Transients of Asynchronous

Machines, IEEE Trans. Journal of heat transfer.

[10] Behrooz Mirafzal, Nabeel. A. 0. Demerdash.

"Induction motor broken bar fault diagnosis using the rotor

magnetic field space-vector orientation," IEEE Trans. Ind

Apple., vol. 40, pp. 534 -542, Feb 2004

124


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2. STUDY ON TRANSIENT BEHAVIOR OF CCVTP. Arunkumar, V.Gowri sree.

College of Engineering, Anna University, Guindy.

Email:[email protected]

Abstract- In this work, a coupling capacitor

voltage transformer (CCVT) model to be used in

connection with the EMTP (Electromagnetic

Transients Program) is presented. Coupling

Capacitor Voltage Transformer (CCVT) is

widely used in power system for high voltage

measurements and also as an input device to

protective relays. The power system is subjected

to transient voltages of external and internal

nature and these overvoltages affect the

reliability of the power system. In addition the

performance of the measuring and protective

device is affected by these overvoltages. Thisnecessitates an indepth study of the transient

performance of measuring devices for stability of

power system. In this study the transient

behaviour of the CCVT is performed by an

Electromagnetic Transients Program (EMTP)

simulation considering the ferroresonance in

particular. A Ferroresonance Suppression

Circuit (FSC) is designed to suppress the Ferro

oscillations and the transient analysis of CCVT

was carried out.

Keywords Coupling Capacitor Voltage

Transformer, Electro Magnetic Transient

Program, Ferro resonance.

I. INTRODUCTION

The Coupling Capacitor Voltage Transformer

(CCVT) is one of the most widely used equipment

in power system for measurement of high voltage

(above 230 kV) and also as input sources to

protective relays. The steady-state performance of

the CCVT is well known.

More investigations are necessary when these

equipment are submitted to transient over voltage.

During normal switching conditions some

unexpected over voltages are producing in several

230 and 500 kV CCVT units, affecting the reliability

of the power system and even causing failures in

some CCVT units . The transient overvoltages

produced in CCVT are due to mainly ferroresonance

oscillations [5]. Ferroresonance oscillations may

take place if the circuit capacitances resonate with

the iron core nonlinear inductance. These

oscillations cause undesired information transferred

to the relays and measuring instruments.

Many works including field measurements,

laboratory tests and digital simulations have been

conducted to study the performance of the CCVT

[10]. However, there are many problems in

obtaining accurate models, especially due to the

need of laboratory tests. In this work Coupling

Capacitor Voltage Transformer (CCVT) is modeled

and simulated by using EMTP (Electromagnetic

Transients Program).The potential transformer

magnetic core and the silicon carbide surge arrester

nonlinear characteristic are included in the CCVT

model in order to improve the transient response to

overvoltages [1].

The obtained CCVT model was used to predict its

transient response. It was observed that the

ferroresonance suppression circuit and the protection

circuit are very effective in damping out transient

overvoltages produced inside the CCVT when a

short circuit is cleared at the CCVT secondary side.

.

II. BASIC PRINCIPLES

The basic diagram of a CCVT is shown in Fig. 1.

The primary side consists of two capacitive elementsC1 and C2 connected in series. The PT primary is

connected to C2 and provides a secondary voltagevo for protective relays and measuring instruments.

The inductance Lc is chosen to avoid phase shifts

between input voltage vi and output voltage vo at

power frequency.

Fig.1. Basic diagram of CCVT

The condition to avoid phase shift between vi and vo

is


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)(

1)(

21 CCLL TC

+=+

(1)

Where Lc represents inductance of compensator

and LT represents equivalent inductance of the

transformer referred to h.v side.

A. Effects of CVT on power system:

The CVT is well used in power system for

measuring and relaying purpose. In normal

conditions CVT will operate accurately but it will

affect the power system when subjected to

transients. The transients produced in CVT are

mainly due to Ferro resonance [2, 5].

Ferro resonance oscillations may take place if the

circuit capacitances resonate with the iron core

nonlinear inductance. These oscillations causeundesired information transferred to the relays and

measuring instruments. Therefore, a ferroresonance

suppression circuit (FSC) is normally included in

secondary of the CCVT windings.

Ferro resonance in a CVT can result in any of the

following:

High sustained over voltages, both phase

to phase and phase to ground

High sustained over currents

High sustained levels of distortion to the

current and voltage waveforms

Electrical equipment damage (thermal ordue to insulation breakdown)

Mis-operation of protective devices.

III. CCVT MODELLING

The basic diagram of CCVT as shown in Fig. 1 is

valid only near power frequency (50 Hz). A model

to be applicable for frequencies up to a few kilohertz

needs to take at least the potential transformer

primary winding and compensating inductor stray

capacitances effects into account [3, 4, 7].

In this work, the circuit shown in Fig. 2 was used

to represent the CCVT. It comprises of a capacitive

column (C1, C2), a compensating inductor (Rc, Lc,

Cc), and a potential transformer (Rp , Lp , Cp , Lm ,Rm) [2, 7].

Fig.2 CCVT model for high frequencies

The iron core nonlinear characteristics and surge

arrester non linear characteristics are included in the

model to give more accurate results for th

simulated CCVT transient response.

A.Surge arrester nonlinear characteristics:

The CCVT protection circuit comprises a silicon

carbide (SiC) surge arrester connected in parallel

with the capacitance C2. . The surge arrester is a non

linear resistor in series with spark gap, its spark over

voltage was measured at power frequency. The

surge arrester effect is also included in simulation of

CVT [3]. The nonlinear characteristics i.e. surge

arresterV-Icurve is estimated from laboratorymeasurements manufacturer. The surge arrester

nonlinear characteristic is shown in Table I.

Table I Silicon carbide surge arrester nonlinearcharacteristic

Current (A) Voltage (kV)

100 20.8

200 27.9

500 39.0

1000 42.9

2000 45.5

B.PT magnetic core nonlinear characteristics:

The magnetic core non linear characteristics is

defined by the equation

( )if= (2)The PT nonlinear peak fluxcurrent (I)

characteristic was obtained from measured rms

voltagecurrent (VI) data in laboratory test [8].

Peak voltages are converted to peak fluxes and the

rms values of the current through the nonlinear

inductance are converted to peak values. The

conversion of peak values of V to flux is again a

re-scaling procedure. Hence, for each linear segment

in the - Icurve


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w

Vk

= (3)

Where w is the angular frequency .The PT

magnetic core nonlinear peak fluxcurrent (I)

characteristic is shown in table II.

Table II PT iron core nonlinear characteristic

Peak Current (A) Peak Flux (V-sec)

0. 076368 0.025772

0.720881 0.189066

1.429369 0.396889

2.511675 0.748388

3.662012 0.863553

4.587227 0.903317

5.712037 0.942706

55.527018 1.556415

5552.7018 1.562242

IV. SIMULATION RESULTS

To perform the simulations on CCVT, EMTP

package (MICROTRAN) was used [9]. The CCVT

model is shown in Fig.3. The magnetizing

inductance Lm was replaced by a nonlinear

inductance connected across the CCVT secondary

terminals whose- Idata points are shown in

Table II. The potential transformer was represented

by the three winding single-phase transformermodel. The protection circuit composed by a silicon

carbide surge arrester was included as well. Its v - i

nonlinear characteristic is shown in Table I. At point

A the system was represented by its Thvenin

equivalent. With this valid CCVT modelsimulations were performed to predict theCCVT behavior under transient overvoltages.

In order to analyze the transient response of

CCVT the simulation consist of a close-open

operation of a switch SW connected across the

CCVT secondary terminals, as shown in Fig. 3. Theswitch closes at t= 100 ms and remains closedduring 5 cycles, when the short circuit is cleared.

Fig.3 CCVT model for Electromagnetic transient

studies

Fig.4 Secondary voltage of CCVT

Fig.4 shows the CCVT secondary voltage

waveform. The oscillations remain up to 500 ms,

when the steady-state is reached. The oscillations are

due to ferroresonance and will cause undesired

operation of relays. In order to suppress th

oscillations a parallel resonance circuit known as

ferroresonance suppression circuit (FSC) is

connected across CCVT secondary terminals.

A. Design of FSC:

To suppress the oscillations FSC was designed.

FSC in an active operational mode consists of

capacitor and iron core inductor connected in

parallel and tuned to the fundamental frequency.

They are permanently connected on the secondary

side and, will improve the CCVT transient response.


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Fig.5 Ferroresonance suppression circuit

The FSC which consists Capacitor Cf is

connected in parallel with an iron core inductor Lf

tuned to the fundamental frequency. Resistor Rfis a

damping resistor designed to damp the

ferroresonance oscillations [Fig.5]. The circuit is

tuned with a high Q- factor in order to attenuate

ferroresonance oscillations at any frequency except

the fundamental.

The performance of CCVT with the FSC was

analyzed. The oscillations are damped in a time

smaller than 100 ms with FSC when compared to

CCVT without FSC where the oscillations were

present upto 300 ms. This is evident from the Fig. 6.


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Fig.6 Secondary voltage of CCVT with FSC

Comparison of CCVT secondary voltage with

and without FSC reveals the importance of FSC in

improving the transient response which is as shown

in Fig. 7. Curve 1 shows the voltage variation

without FSC and curve 2 shows voltage variationwith FSC.

Fig.7 Secondary voltage variation with and

without FSC

V. CONCLUSIONS

In this work, a CCVT model for electromagnetic

transient studies is presented. The used CCVT

model was validated for frequencies in the range

from 10 Hz to 10 kHz. The model includes the PTmagnetic core and SiC surge arrester nonlinearcharacteristics in order to improve the transient

response to overvoltages and is simulated by using

EMTP. The result shows the importance of the FSC

in damping out transient voltages when a short

circuit is cleared at the CCVT secondary terminals.

VI. REFERENCES

[1] D. Fernandes Jr., W. L. A. Neves, and J. C. A.Vasconcelos, "A Coupling Capacitor Voltage

Transformer Representation for Electromagnetic

Transient Studies," inProc. 2003 InternationalConference on Power Systems Transients, New

Orleans, USA.

[2] M. Kezunovic, Lj. Kojovic, V. Skendzic, C. W.Fromen, D. R. Sevcik, and S. L. Nilsson, "Digital

Models of Coupling Capacitor Voltage Transformersfor Protective Relay Transient Studies," IEEE Trans.

Power Delivery, vol. 7, no. 4, pp. 1927-1935, Oct.

1992.

[3] M. R. Iravani, X. Wang, I. Polishchuk, J. Ribeiro,and A. Sarshar, "Digital Time-Domain Investigation

of Transient Behaviour of Coupling Capacitor

Voltage Transformer," IEEE Trans. Power Delivery,vol. 13, no. 2, pp. 622-629, Apr. 1998.

[4] D. A. Tziouvaras, P. McLaren, G. Alexander, D.Dawson, J. Ezstergalyos, C. Fromen, M. Glinkowski,

I. Hasenwinkle, M. Kezunovic, Lj. Kojovic, B.Kotheimer, R. Kuffel, J. Nordstrom, and S. Zocholl,

"Mathematical Models for Current, Voltage and

Coupling Capacitor Voltage Transformers,"IEEETrans. Power Delivery, vol. 15, no. 1, pp. 62-72, Jan.

2000.

[5] H. M. Moraes and J. C. A. Vasconcelos, "Overvoltages in CCVT During Switching Operations," (InPortuguese), in Proc. 1999 National Seminar on

Production and Transmission of Electrical Energy,

Foz do Iguau, Brazil.

[6] J. R. Lucas, P. G. McLaren, W. W. L. Keerthipalaand R. P.Jayasinghe, Improved Simulation Models

for Current and Voltage Transformers in RelayStudies, IEEE Trans. on Power Delivery, Vol. 7,

No. 1, pp. 152-159, January 1992.

[7] Lj. Kojovic, M. Kezunovic, V. Skendzic, C. W.

Fromen and D. R. Sevcik, A New Method for the

CCVT Performance Analysis Using Field

Measurements, Signal Processing and EMTP

Modeling, IEEE Trans. on Power Delivery, Vol. 9,No. 4, pp. 1907- 1915, October 1994.

[8] W.L.A. Neves and H.W. Dommel, On modeling

iron core nonlinearities, IEEE Trans. On Power

Syst., Vol. 8, No. 2, pp. 417-425. May 1993.

[9] Microtran Power System Analysis Corporation,

Electromagnetic Transients Analysis Program,Vancouver, 1999.

[10] D. Fernandes Jr., W.L.A. Neves, J.C.A.Vasconcelos, M.V. Godoy, Couplingcapacitor voltage transformer: laboratorytests and digital simulations, in: 2005International Conference on PowerSystems Transients, IPST05, Montreal,Canada, June 1923, 2005.


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3. Breakdown Characteristics of Transformer under Non-

Standard Impulse VoltagesDhayalan.J, Dr K.Udaya kumar,

Department of EEE, College of Engineering,Guindy, [email protected], [email protected],

Abstract The optimal and efficient design of

any high voltage apparatus depends on reliable

design of its insulation, which is tested with the

standard lightning impulse voltages of wave

shape 1.2/50s.during the testing of large power

transformers, it is difficult to adjust the impulse

generator to get the standard 1.2/50s wave

shape, and also part of the winding can get

stressed with voltages of non standard wave

shapes. A fundamental study on breakdown indielectric media like gases, liquid, solid and

composite materials for different electrode

configurations from uniform to highly non

uniform field configuration is essential. The test

voltages are of different types viz. power

frequency, lightning impulse and switching

impulse. As the insulation strength is not the

same for all the waveshapes, a detailed study on

the behaviour of insulation under various

voltages is essential to make an optimal design.

As most of the transformer failures are due to

these small insulations, it is must to asses the

breakdown characteristics of air, transformeroil and OIP in small gaps. This is necessary to

represent the actual conditions in transformer,

which would be helpful in the reliable design of

transformer.

Index Termsimpulse generator, test kit, sphere

gap, PSPICE.

1. INTRODUCTION

The equipments and materials used in all powersystem installation are designed and constructed

such that they are capable of withstanding electricstresses due to lightning impulse voltages [1]. The

ability of the equipment to withstand the dielectric

stress is checked with impulse voltage withstand

test. But in actual conditions the stress occurring

varies widely from that applied during standardlightning impulse test. A brief explanation has been

given on the various sources of non standard

impulse voltages and their effects on transformer

insulation. The practical study of transformer

insulation under oscillatory impulse voltage

requires the experimental generation of such

oscillatory waves. The simulation of the modified

impulse generator circuit for generation of

oscillatory impulse voltage and the practical

generation of unidirectional and bidirectional

oscillatory impulse voltage at different frequencies

are also explained. The behavior of air, oil and oil

impregnated paper insulation under bidirectionaloscillatory impulse wave shape of frequencies 3.5

KHz to 125 KHz using modified Marx circuit [2].

The thickness of air, oil and oil impregnated paper

ranges from 1mm to 5mm. The oil and

impregnated paper insulating medium are mainly

used as inter turn and inter disc winding [3]. The

test voltages are of different types viz. power

frequency, lightning impulse and switchingimpulse. In case of tests with lightning impulse

voltage, standard waveshape of 1.2/50s is used to

test the transformer. Even when the transformer is

tested with standard waveshape, due to part

winding resonance, the winding insulation isstressed with non standard waves, which are

oscillatory.The types of insulation within atransformer can be broadly classified as majorinsulation, end insulation and winding insulation. It

is reported in the literature that more than 50% of

the failures in power transformers are due t

insulation failure in the windings. Though there are

methods reported in the literature to evaluate the

breakdown strength, a detailed study on various

systems of insulation with varying thickness under

extreme field conditions representing the actual

conditions in transformer winding is mandatory.

2. DEFINITION OF STANDARD AND NON-

STANDARD IMPULSE VOLTAGE

3. Standard Impulse Voltage

As per IEC 60060, a standard lightning impulse

is defined to have a front time of 1.2s with

30%tolerance and a tail time of 50s with 20%

tolerance and with a peak overshoot of 5%.
mailto:[email protected],%[email protected]:[email protected],%[email protected]


12/14

V=Vo [exp (-t)-exp(-t)]

(1)

where

=0.0146, =2.467, Vo=1.04

The parameters and controlsthe front time and tail time of the impulse wave

respectively.

4. Non-Standard Impulse Voltages

Non oscillatory

Impulse wave shapes having different time

to front and time to tail beyond the specifiedtolerance limit and without any super imposed

oscillations are grouped as non oscillatory non

standard impulse voltages.

OscillatoryImpulse wave shapes having oscillations super imposed in

the wave front or the wave tail and with the peak overshoot greater than 5% and peak oscillations above are

groped as oscillatory non standard impulse voltages.

III. MODIFI

ED CIRCUIT FOR IMPULSE GENERATION

The modified Marx circuit is used to generate

bi-directional impulse of frequency range from

29KHz to 128KHz. The value of inductance

range from 20H to 500 H.

Fig. 1Modified Marx circuit for bi-directional

impulse

Time

0s 10us 20us 30us 40us 50us 60us 70us 80us 90us 100us

V(R3:2,0)

-100V

-50V

0V

50V

100V

Fig.2 Non-standard bi-directional impulse of 65

KHz

Ti me

0 s 1 0 us 2 0 u s 3 0 u s 4 0 u s 5 0 u s 6 0u s 7 0 u s 8 0 u s 90 u s 1 0 0 us

V (R 3: 2,0 )

-4 0V

-2 0V

0V

2 0V

4 0V

6 0V

8 0V

FIG .3NON-STANDARDBI-DIRECTIONALIMPULSEOF 128

KHZ

Table 1

S.no Inductance (H) Frequency kHz)1. 20 65

2. 30 843. 60 1054. 100 128

5. 200 466. 400 337. 500 29

Fig .4 Variation of Frequency W.R.T Inductance

IV.EXPERIMENTAL SETUP

A 120KV, 250 KJ impulse voltage generator of

MWB make is used for generating standard (1.2/50

s) impulse voltages. Figure 1.2 shows the circuit

of the non standard impulse voltage generator. Asthe gap distances chosen for the analyses are very

small, fine voltage control is effected. Voltage ismeasured using capacitive divider connected to an

8-bit digital storage oscilloscope (Tektronix, TDS

420). The oscilloscope is interfaced to the computer

via IEEE 488.2 interface through which the

digitized waves are acquired.


13/14

Fig. 5 Modified Marx circuit for uni-directional

impulse

REF1

-15

-10

-5

0

5

10

-1.50E-

05

-1.00E-

05

-5.00E-

06

0.00E+

00

5.00E-

06

1.00E-

05

1.50E-

05

time(s)

Voltage

REF1

FIG .6 Uni-Directional Impulse

The long breakdown time lags for fast frontedwaves and the breakdown on the front of the slowfronted waves are observed for small air gaps atdifferent configurations. The oil impregnatedbushings fail due to deterioration in the insulationcaused by the high frequency surges generatedswitching[7]. The analysis of oilpaper insulation

under steep fronted impulse shows that breakdownstrength is lower for steep fronted impulses andhence oil-paper insulted equipment subjected tosteep front transients may fail below the lightningimpulse design level (BIL)[5]. At a constantdamping factor when the oscillation frequency isvaried the breakdown voltage is found to increasewith the frequency of oscillation. The SF6 gas has atendency to have higher withstand capability undernon standard LI voltage than under standard LI.Using the circuit simulation software (Orcad-PSpice) the modified circuit is simulated and therange of inductance is estimated. It is found fromthe analysis that inductance values from 10H to30H generate bidirectional waves with required

range of frequency (3 KHz to 125 KHz) ofoscillations. The inductance values are in the rangeof 10H to 15mH of 66Kv rating are designed andfabricated to generate bi-directional waves in thelaboratory.

Fig. 7 Modified Marx circuit for uni-directional

impulse

V1- 230V/140KV, Cs-Charging Capacitor=100nF,D1, D2-Diodes, S-Sphere gap,R-Charging Resistor=2.5M Cd-CapacitiveDivider=1200pF R1=245,R2=1200.

REF1

-8

-6

-4

-2

0

2

4

6

8

10

-2.00E-

05

0.00E+0

0

2.00E-

05

4.00E-

05

6.00E-

05

8.00E-

05

1.00E-

04

REF1

Fig. 8 NON-STANDARDBI-DIRECTIONALIMPULSEOF 65

KHZ

REF3

-10

-5

0

5

10

15

-2.00E-

05

0.00E+0

0

2.00E-

05

4.00E-

05

6.00E-

05

8.00E-

05

1.00E-

04

REF3

Fig. 9 Non-standard bi-directional impulse of 128

KHz.

IV. CONCLUSION

During testing large powertransformers due to low inductance values it is very

difficult to maintain the standard impulse waveshape. Different factors viz., electrode material,

electrode geometry and type of voltage ( AC, DC or

impulse), which influence the breakdown in small

insulation gaps are explained. Plane-plane, plane-

cone, cone-cone and plain-needle electrodes are

chosen for air & oil insulating media. From the


14/14

measurement of break down voltages of air it can

be observed that the difference between AC, DC &

impulse voltages increase with gap distance. For oil

insulating media the analysis shows that the

difference between AC, DC & impulse voltages arealmost the same for all the gap differences. For OIP

insulation, there is a maximum of increase inimpulse breakdown voltage when compare to AC

breakdown voltage. The voltage time

characteristics explaining about the instant of break

down on standard lightning impulse play an

important role in the insulation

co-ordination of the entire power system. Its foundthat the voltage time characteristics dispersion is

more for plane-cone configuration than for plane-

plane configuration for air and oil.REFERENCES

[1] S. Usa, K. Udayakumar and V. Jayashankar,

"Modified disruptive effect method as ameasure of insulation strength for non-standard

lightning waveforms", IEEE Transactions on

Power Delivery, Vol.17, No.2, April2002.

[2] S. Venkatesan, S. Usa and K. Udayakumar,

Unconditionally sequence approach to

calculate the impulse strength of air for non

standard impulse voltages", IEEE Transactions

on Conference Proceedings, 2002. [5]

[3] Shiemetsu Okabe, Mmasonari Kotou et al."Dielectric characteristics of oil filled

transformer insulation models under non-

standard lightning impulse voltages, VIIIInternational Conference on High Voltage

Engineering, August 2003.

[4] G. Danikas, "Breakdown of transformer oil",IEEE Electrical Insulation Magazine, Vol.6,

No.5, Sept./Oct. 1990.

[5] K.D. Srivatsa and J.B. Neilson, Electricalbreakdown characteristics oil filled paper

insulation under steep fronted impulse

voltages", IEEE Transactions on Power

Delivery, Vol.9, No.4, October 1994.

[6] J.J. O'Dwyer, "Breakdown in solid dielectrics",IEEE Transaction on Power Delivery ET 17,

No.6, December 1982.

[7]Y. Kamata, S. Fufukawa and K. Endoh,"Dielectric strength of oil immersedtransformer insulation with super imposed c

and lightning impulse volatge, IEEE

Transaction on Electrical Insulation, Aug.

1990.

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