papers of hve
TRANSCRIPT
-
8/7/2019 Papers of Hve
1/14
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
-
8/7/2019 Papers of Hve
2/14
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
-
8/7/2019 Papers of Hve
3/14
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
-
8/7/2019 Papers of Hve
4/14
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
-
8/7/2019 Papers of Hve
5/14
[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
-
8/7/2019 Papers of Hve
6/14
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
-
8/7/2019 Papers of Hve
7/14
)(
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
-
8/7/2019 Papers of Hve
8/14
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.
-
8/7/2019 Papers of Hve
9/14
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.
-
8/7/2019 Papers of Hve
10/14
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.
-
8/7/2019 Papers of Hve
11/14
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] -
8/7/2019 Papers of Hve
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.
-
8/7/2019 Papers of Hve
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
-
8/7/2019 Papers of Hve
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.