IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 17, NO. 3, SEPTEMBER 2002
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A New Synthetic Loading for Large Induction
Machines With No Feedback Into the Power System
Jafar Soltani, Member, IEEE, Barna Szabados, Senior Member, IEEE, and Gerry Hoolboom, Member, IEEE
Index Terms—Measurement of losses, phase modulation, synthetic loading.
I. INTRODUCTION
H
to the shaft. The two-frequency method [1] has major advantages over the other methods. It is applicable to both wound rotor
and squirrel cage machines as opposed to Romeira’s method
[2] which can be applied to wound rotor machines only. Furthermore, it does not require that all six leads of the rotor be
brought out as in Fong’s method [3]. In recent work [7] our
group has established that the two-frequency method did not
lead to good results because of the very high fluctuations of
stator voltage applied and, therefore, the degree of acceptance
of the method is dubious. Moreover, a geared generator and a
multiphase transformer is usually required both rated at the test
machine rating, and the full rated power oscillates at a low frequency into the power system. The investigation of three synthetic loading methods at full-load temperature evaluation using
calorimetric methods has shown the major drawbacks of the
two-frequency method. Results have been compared with the
conventional direct loading method [10]. Unfortunately, these
results are limited to low capacity induction motors due to the
power swing in the power system.
In this paper, a new synthetic loading is proposed. The
assumption made is that motor manufacturers prefer to build
rotating machines used for the test rig rather than buying
special purpose large transformers. The method is based on a
bang–bang phase modulation technique to generate a variable
frequency voltage generated by a synchronous generator. The
use of two systems in parallel, one system with the test motor
and one system with a “recovery” unit, limits the power swing
with the power system to the total losses of the five machines
used.
The method has been proven first by simulation whereby the
constraints of the method have been identified, and it has been
experimentally verified on a test set up in the laboratory.
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Abstract—Full-load testing of large induction machines is
constrained by the limitations in the power-supply and loading
equipment of the manufacturer’s facilities, resulting in costly set
up time. A new synthetic loading method is proposed based on
a bang-bang phase control strategy. The rated power oscillation
created is routed to an auxiliary system and the source hydro has
to provide only the total losses of the system, without seeing the
excessive power swings observed in other synthetic loading techniques. In this technique, only induction machines are used which
would enable motor manufacturers to build the test rig in-house.
The control stage is very simple to implement and requires only
unregulated dc supplies for the excitation windings. The method
is suitable for any induction machine and does not requires any
set up time. It is possible to strictly maintain constant DEFINE
RMS voltage and current at rated values for the duration of the
heat runs.
EAT RUN tests performed on electric machines are
extremely important to both manufacturers and users.
The test results verify the predictive performance calculation
methods used, therefore lowering business risks during the tendering process. On the other hand, the test results demonstrate
to the user that the contractually agreed upon performance has
been met. Specifically, it can be verified that at full load the
machine does not exceed the temperature insulation class limit.
Established business patterns have demonstrated that the
number of motors produced decreases as the machine power
output increases, and that the provision of motor test plant
follows the opposite trend [9]. Direct methods suffer from high
costs due to the complex coupling and loading mechanisms, as
well as the sophisticated sensor and measurement instrumentation required. Indirect methods suffer from large inaccuracies
and mainly cause unacceptable power swings on the power
system. There is obviously a need for alternate methods to
produce full-load heat runs economically.
A number of proposals have been devised [1]–[3], [5], [9]
which attempt to produce accurate performance verification of
induction machines without having to attach mechanical loads
Manuscript received July 25, 2000; revised February 7, 2002. This work was
supported in part by the U.S. Department of Commerce under Grant BS123.
J. Soltani was with McMaster University, Hamilton, ON, Canada, L8S
4K1. He is now with Isfahan University of Technology, Isfahan, Iran (e-mail:
j1234sm@cc.iut.ac.ir).
B. Szabados and G. Hoolboom are with the Power Research Laboratory, McMaster University, Hamilton, ON, L8S 4K1 Canada (e-mail: szabados@mcmaster.ca).
Publisher Item Identifier 10.1109/TEC.2002.801728.
II. PRINCIPLE OF EQUIVALENT LOADING
The internal air-gap voltage of an unloaded induction machine is very close in magnitude and phase angle to the apand the no load current
is small
plied armature voltage
[Fig. 1(a)]. As the machine becomes loaded, the load angle increases and a larger armature current is produced [Fig. 1(b)].
The principle of equivalent loading is to increase the internal
load angle without connecting any mechanical load onto the
shaft. The two-frequency method [1], [2], [7] essentially uses
one voltage at 60 Hz in series with another voltage at 50 Hz that
is applied to the armature of the machine under test. This voltage
will have a modulated phase angle, causing the rotor speed of
the machine under test to try to follow. The inertia of the rotor
acts as an energy storing device.
0885-8969/02$17.00 © 2002 IEEE
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IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 17, NO. 3, SEPTEMBER 2002
Fig. 2. Dual-field winding control. (a) Excitation windings. (b) Resultant flux
control in quadrature.
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Fig. 1. Principle of load angle control.
It has been shown [7] that the two frequency method sees
variations of terminal voltage between 140% and 60% of rated
value. This range drives the machine into saturation as well as
undervoltage, introducing major inaccuracies in the resulting
measurement. Furthermore, the full rated power swings at the
modulation frequency between the power grid and the test
motor.
III. NEW PHASE MODULATION
A. Principle
The principle of the new phase modulation method is to generate a voltage using a three-phase synchronous machine. The
synchronous machine needs to have a rotor with two windings
with a spatial phase shift, as shown in Fig. 2(a). By appropriately exciting each winding, any phase angle within the range
and
shown in Fig. 2(b) can be achieved for the excitation
field. The modulation in amplitude and phase can be achieved
by appropriately modulating the excitation currents and of
the field windings. A sinusoidal modulation would requires two
precisely controlled dc sources, which is very difficult to implement especially since large inductive values are present in the
field windings.
B. Practical Implementation
A synchronous generator using a wound rotor induction machine is created as shown in Fig. 3. Winding (A) of the rotor is
excited by a controlled source DC1. Windings (B) and (C) are
connected in series and excited with another fixed supply DC2.
Fig. 4 shows that the resultant dc flux created by the rotor windings, represents the direct axis in the synchronous frame of the
machine. By appropriately controlling the current in winding
(A), one can modulate the phase w.r.t. the utility source phasors.
Practically, the source DC1 is switched in a bang-bang mode,
Fig. 3. Creating a controlled dc field in the rotor of an induction machine.
Fig. 4. Creating a modulated synchronous field in the rotor of an induction
machine.
and due to the reactance of the field winding, the excitation current will vary exponentially. The vector diagram of Fig. 4 also
illustrates the range of the phase shift when a bang–bang modulation is impressed on winding (A).
The phase modulating synchronous generator implemented
with this scheme is driven by a “driver” motor fed from the
utility. The output of this synchronous generator feeds the test
motor which, in this single rig implementation, would see the
total power swing go through to the utility connection as described in a previous disclosure [11].
We propose here to use two systems similar to the one
described in [11]. Both systems share the “driver” motor
(D0) on the same shaft as shown in Fig. 5. Generator G1 of
of the
System(1) is rated at the maximum power rating
test rig (highest motor rating to be tested), and feeds the motor
M1 under test. The test motor could be either a wound rotor
machine or a squirrel cage machine. Both the generator G2 and
.
the “recovery” machine M2 of System(2) are rated at
If the field modulation of each generator is in opposition of
phase, the power generated by each system is also in opposition
of phase and, therefore, when one system absorbs power, the
other generates it and vice-versa. By adjusting the magnitude
of the excitation swing of generator G2, one can adjust the
SOLTANI et al.: A NEW SYNTHETIC LOADING FOR LARGE INDUCTION MACHINES WITH NO FEEDBACK INTO THE POWER SYSTEM
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power exchange with System(2) to exactly match the power
swing of System(1). When this equilibrium is reached, the
driver motor (D0) needs to provide only the losses in all five
machines. Driver D0 is preferably a synchronous motor, but
can be a dc motor or an induction motor with low slip. In the
latter case, the motor would have to be slightly over rated in
order to perform the test at close to synchronous speed.
There is a set of performance curves describing the magnitude of the modulation as well as the depth of the modulation
(frequency swing). It has been shown [11] that the optimum performance to reach full rated load is achieved with a 10-Hz modulation depth (55 to 65 Hz).
C. System Simulation
Fig. 5. Test set up.
The system shown in Fig. 5 has been simulated with G1 and
G2 being identical three phase, 2–kW wound rotor induction
machines, and M1, M2 and D0 are 2–kW squirrel cage motors,
with the corresponding parameters shown.
kg-m
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Motors squirrel cage) Induction Generators
poles KW
Hz
V
Nm-sec/rad (windage
iron)
kg-m
[with for resistance,
for leakage reactance,
for magnetizing reactance and for moment of inertia; subscripts
and
for stator and rotor].
The simulation uses the standard Park’s equations to convert
three phase into the conventional DQ axis model [8]. The Appendix shows the relevant equivalent circuits used for each one
of the interconnected systems (1) and (2) as well as the main
equations used.
The set of flux equations is solved using a static Runge–Kutta
fourth-order method. Fig. 6 shows the results of the simulation, providing the modulated phase voltage applied to motor
M1 and the corresponding phase current. One can note that although there is a definite phase modulation between current and
voltage, the magnitude of the voltage is nearly invariant and
at near rated value, while the magnitude of the current varies
considerably with the modulating frequency. The instantaneous
power is computed as the instantaneous products of current by
voltage. The active power is computed as the average of the instantaneous power over one pseudo cycle of the 60–Hz-modulated signal. Finally, the total losses in the system are computed
as the average of the active power over the modulation interval
[11]. Fig. 7 shows the power flow at the terminal of machine M2
and illustrates the opposition of phase with M1 compared with
Fig. 6.
Fig. 8 clearly shows that the power flow through driver D0 is
constant. Losses in M1 and M2 are found to be 0.22 PU each
(base power 2 kW), while the total losses in the 5 machines
show 1.2 PU, which is commensurate with the name-plate data
provided by the manufacturer.
Fig. 6. Power flow in machine M1 (test motor).
D. Experimental Results
Fig. 9 shows the experimental values of phase voltages and
phase currents in machines M1 and M2. It clearly shows that
both terminal voltages are modulated with the same frequency,
although slightly out of phase are nearly constant in magnitude.
However, the currents in the two machines are of different amplitudes, and also show definite phase differences at each instant
in time, reflecting the power oscillation between the two machines. The computed instantaneous power and the computed
average power or losses in the test machine M1 are shown in
4
Power flow in machine M2.
a phase shift exists between the voltages V1 and V2 and the
instantaneous power generated by one machine is not matched
by an opposing absorbed power in the other machine. We have
proven this in the simulation by including a phase difference
between the direct axis on machines G1 and G2, and the result
showed clearly that a power oscillation occurs and its magnitude
depends upon the phase angle. In our experimental set up, the
two machines were connected through a keyed mechanical coupling with a cogged rubber ring. The cogs allowed only an axis
adjustment within steps of 40 . We verified that power oscillations vary with realigning the two shafts, but due to the physical key of the coupler constraint we could not arrive at better
alignment than about 15 . In order to achieve perfect alignment,
we would need a slanted keyed coupler which, of course, could
easily be constructed on a final test rig.
control
Previous studies [7] advocate that a constant
strategy should provide the best loss measurement together with
a sinusoidal modulation. In this method, we do not use these
constraints. The experimental results showed clearly that the
use of the phase-modulation technique draws advantageously
on the armature reaction of the generators which tend to maintain the voltage of the generator output as constant as possible.
In fact, the phase modulation means that we actually modulate
the angle between the internal voltage of the generator and the
motor effectively, realizing a quasi current source. The results of
the simulation show that there is no power swing between the
utility supply and the test rig when System(2) is adjusted exactly to match the power swing of System(1). Although the experimental set up prevented us from creating this perfect match,
the power oscillation was minimized, and certainly would be acceptable in a practical platform.
We fully realize that this technique uses five machines. However, according to machine manufacturers we are working with,
they much prefer constructing extra machines, rather than purchasing bulky transformers and gear boxes. Furthermore, one
really needs only the driver D0 and the two generators G1 and
G2 as fixed installation with carefully aligned shafts. Nothing
prevents the use of two motors M1 and M2 to be tested at the
time. There is always one extra motor in the plant to act as a recovery unit, and since there is no need for mechanical coupling,
this is an easy installation. The extra two machines required (instead of a multiwinding transformer and a gear box) is considered a small investment versus achieving no power swings in the
power system.
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Fig. 7.
IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 17, NO. 3, SEPTEMBER 2002
Fig. 8. Power flow in driver D0 and total losses.
Fig. 10. Finally , Fig. 11 shows the instantaneous power as well
as the average active power flowing through driver D0, representing the total losses in the five machines. This is found to
be slightly different than the simulated results. One has to note
that the value of 1.5-PU power flow through the driver motor is
misleading. This is true for this small system where the losses
in each machine are approximately 20%. In industrial applications, this would be only around 5%.
The speed of the test motor was measured with a tachogenerator. The signal obtained showed tooth ripple and brush commutation noise well in excess of the expected small variation of
speed.
IV. DISCUSSION
Fig. 11 shows clearly a power oscillation through driver D0
while the average power still represents the total losses of the
system. This oscillation has been traced to two main effects. The
first cause comes from the phase modulation itself. The excitation currents of G1 and G2 have to be exactly in opposition of
phase. While it is possible to apply phase opposition voltages to
the windings, it is difficult to guarantee that both time constants
are identical. Hence, the two excitation fluxes will not necessarily establish their final values symmetrically. We have tried
to minimize the power ripple due to this cause by adjusting the
relative excitation voltage on G2.
The second main cause for the power ripple comes from the
misalignment of the electrical axis between machines G1 and
G2. If the two machines do not have their direct axis aligned,
V. CONCLUSIONS
We have demonstrated the viability of a new phase shifting
method to generate a modulated three-phase feed to a test machine. This method can be easily implemented using a conventional three-phase wound rotor induction machine excited by
two uncontrolled dc sources. Using a simple bang–bang control
on one of the dc sources provides the modulation to the output
voltage.
Furthermore, if two similar systems run back to back, a
rated power swing between the two systems can effectively be
achieved. Only the total losses of the five machines have to be
provided by the source.
SOLTANI et al.: A NEW SYNTHETIC LOADING FOR LARGE INDUCTION MACHINES WITH NO FEEDBACK INTO THE POWER SYSTEM
Fig. 11.
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Instantaneous power in D0 and total losses.
Fig. 9.
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when the test motor rating is much lower than the rating for
the recovery system, since one has to match the instantaneous
power in both machines. Even if the direct axis of both generators was perfectly aligned with the open-circuit voltage
test, during loading, the internal angles will be different in the
machines because of the different parameters of the attached
motors. Therefore, some power oscillation cannot be prevented
in practical applications without a sophisticated controller.
However, as shown in our experimental verification, this power
swing through driver D0 will be small compared to the rated
power of the machines For practical implementations it would
still be acceptable.
Voltages and currents in machines M1 and M2.
APPENDIX
SYSTEM SIMULATION
Fig. 10.
Computed power and losses in test machines M1.
Using simulation and experimental verification, we have
constant requirement is not a necessity. It
shown that the
is possible to maintain constant RMS voltage and current in
the test motor by adjusting the two field winding currents of
the generators. The simulation and experimental verification
proves that the main criterion of the test is to maintain the RMS
quantities of current and voltage to the rated values computed
over the modulation cycle.
The main problem identified during the experimental part
was the need for a closed-loop control to maintain the constant
ratings on the test motor and the matching power exchanged
with the recovery system. This problem becomes quite involved
Fig. 12 shows the usual equivalent circuits with the conventional naming of parameters used for each one of the generators
and motors in the interconnected systems (1) or (2). The left side
of the equivalent circuit contains the lumped stationary compoand
and the right side the rotor quantities of
nents of
referred to the stator. Indices
and
, respectively,
refer to direct quadrature axis and zero sequence, and indices
and
refer to stator and rotor quantities, while
refers
refers specifically to the
to the excitation field. The subscript
excitation of the induction machine working as a synchronous
in the parameters.
generator, generally indicated with indice
Furthermore, is the base angular frequency of the shaft while
and
are the rotor angular velocities of the driver motor
, respectively. The voltages
are calculated
and motors
from the equivalent circuit models of the stator of the generators. Fig. 13 shows the equivalent circuits for the rotors of the
working as synchronous generators.
machines
The fluxes shown on the models are
(A1)
and
(A2)
The induction generator flux equations follow as
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IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 17, NO. 3, SEPTEMBER 2002
(A3)
The field voltages are defined as
(A4)
(a)
leading to
(A5)
(b)
G
M ). (a)
The torque developed for each machine is
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Fig. 12. Equivalent circuit for stationary parts and rotor for ( ) and (
Direct axis. (b) Quadrature axis.
(A6)
REFERENCES
(a)
(b)
[1] A. Meyer and H. W. Lorenzen, “Two-frequency heat run—A method of
examination for three-phase induction motors,” IEEE Trans. Power App.
Syst., vol. PAS-98, pp. 2338–2347, Nov./Dec. 1979.
[2] M. P. Romeira, “The superimposed frequency for induction motors,”
Proc. IEEE, vol. 36, pp. 952–953, 1948.
[3] W. Fong, “New temperature test for polyphase induction motors by
phantom loading,” Proc. IEEE, vol. 60, pp. 883–887, July 1972.
[4] Std. Jpn. Electrotechn. Comm., JEC-37-1979, 1979.
[5] H. R. Schwenk, “Equivalent loading of induction machines for temperature test,” IEEE Trans. Power App. Syst., vol. PAS-96, pp. 1126–1131,
July/Aug. 1977.
[6] C. Grantham, E. D. Spooner, and M. Sheng, “Synthetic loading of machines using power electronics,” Elect. Eng., vol. 67, no. 8, pp. 60–68,
Aug. 1990.
[7] A. Mihalcea, “Equivalent loading methods for determining total power
loss in induction motors,” Master’s thesis, AU: At what University was
this thesis written, CITY?, COUNTRY, Sept. 1999.
[8] P. Krause, Analysis of Electric Machinery. New York: McGraw-Hill,
1986.
[9] D. H. Plevin, C. N. Glew, and J. H. Dymond, “Equivallent load test
for induction machines—The forward short circuit test,” IEEE Trans.
Energy Conversion, vol. 14, pp. AU: Page numbers?–, Sept. 1999.
[10] B. Szabados, J. Soltani, and G. Hoolboom, “A new synthetic loading of
induction machines based on phase modulation,” in Proc. INDUSCON
Conf., Porto-Alegre, Brazil, Nov. 2000, pp. 7–11.
[11] A. Mihalcea, B. Szabados, and J. Hoolboom, “Determining total losses
and temperature rise in induction motors using equivalent loading
methods,” IEEE Trans. Energy Conversion, vol. 16, pp. 214–219, Sept.
2001.
(c)
Fig. 13. Equivalent circuit for the rotor of the generators. (a) Zero sequence.
(b) Direct axis. (c) Quadrature axis.
Jafar Soltani received the B.Sc. degree from the University of Tabriz, Tabriz,
Iran and the Master’s and Ph.D. degrees from University of Manchester Institute
of Technology (UMIST), Manchester, U.K.
Currently, he is Associate Professor at Isfahan University of Technology,
Teheran, Iran, and was on leave at McMaster University, Hamilton, ON, Canada.
His main area of research is electrical machines and drives.
SOLTANI et al.: A NEW SYNTHETIC LOADING FOR LARGE INDUCTION MACHINES WITH NO FEEDBACK INTO THE POWER SYSTEM
Jerry Hoolboom (M’60) received the B.Sc. and M.Sc. from the University of
Delft, Delft, The Netherlands.
His expertise in electrical machine design was achieved while he was with
McMaster University, Hamilton, ON, Canada, as Assistant Professor, and Westinghouse Canada, AU: Location of Westinghouse Canada?, where he held the
position of Director of Technology.
He is a registered power engineer in the province of Ontario.
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Barna Szabados (SM’87) received the B.Sc. degree from Grenoble University,
Grenoble, France, the Master’s and Ph.D. degrees from McMaster University,
Hamilton, ON, Canada.
He is currently Professor of electrical and computer engineering at McMaster
University and is the Director of the Power Research Laboratory. His main interests are power electronics and power apparatus in the field of control, measurement, and modeling of machines and transformers.
Dr. Szabados is a registered P.Eng. in the Province of Ontario.
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