CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSIS
AND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES
by
William T. Thomson
Director and Consultant
EM Diagnostics Ltd.
Alford, Aberdeenshire, Scotland
and
Philip Orpin
Technical Services Manager
Ingenco Ltd.
East Kilbride, Lanarkshire, Scotland
ABSTRACT
William T. (Bill) Thomson is Director
and Consultant with EM Diagnostics Ltd.,
in Alford, Aberdeenshire, Scotland. He
began as an apprentice electrician and has
40 years’ experience covering the installation, maintenance, design, performance,
and condition monitoring of electrical
drives. He has worked as an electrician,
R&D engineer with Hoover Ltd., consultant, and academic. In 1990, he was
appointed Professor at the Robert Gordon
University, in Scotland, in recognition of his research and
development work on condition monitoring for electrical drives in
the offshore oil industry and power utilities.
Professor Thomson has a BSc (Hons, Electrical Engineering,
1973) and an MSc (1977) from the University of Strathclyde. He is
a senior member of the American IEEE, a Fellow of the IEE, and
a registered Chartered Engineer in the United Kingdom. He is also
the visiting Professor in Electrical Engineering at the University of
Abertay, Scotland, and has published over 70 papers.
Induction motor drives are the most widely used electrical drive
system and they typically consume 40 percent of an industrialized
nation’s total generating capacity. In the USA the total generating
capacity is approximately 800,000 MW, consequently induction
motor drives are major assets in the process and energy industries.
The asset management of electrical drives requires reliable
maintenance strategies that include condition monitoring and
online diagnostics. Due to the complex electromagnetic and
mechanical characteristics of an electrical drive system, a unified
monitoring strategy has distinct advantages over monitoring only
one parameter (e.g., vibration) to diagnose problems. This paper
focuses on industrial case histories to demonstrate the application
of current and vibration analysis to diagnose problems in
induction motor drives. The results show how the root cause of a
problem can be established when a combination of current and
vibration monitoring is used in comparison to only analyzing one
signal.
INTRODUCTION
Many operators now use online condition-based maintenance
strategies in parallel with conventional planned maintenance
schemes. This has reduced unexpected failures, increased the time
between planned shutdowns for standard maintenance, and
reduced operational costs. However, it is still the operator who has
to make the final decision on whether to remove a machine from
service or let it run based on information from condition
monitoring systems. The root cause (RCA) of the fault has to be
established and, ideally, the operator also requires a prognosis of
the remaining run life; but the latter is complex and in most cases
is an impossible task. Vibration monitoring and analysis are mature
and effective techniques to diagnose mechanical problems. Current
monitoring can detect problems such as broken rotor bars, airgap
eccentricity, and more recently shorted turns in low voltage stator
windings.
However there still tends to be a historical culture that
mechanical faults are the sole domain of the mechanical engineer,
and likewise electrical problems belong to the electrical engineer.
Signals are often analyzed at different times and separate reports
are presented. With respect to an induction motor drive system, this
does not make sound engineering sense since an electrical drive is
an interconnected, electromechanical system, and it is not sensible
to analyze signals in isolation. The fault may appear to be electrical
but the fundamental cause may be due to mechanical forces. For
Philip (Phil) Orpin is Technical Services
Manager at Ingenco Ltd., in East Kilbride,
Lanarkshire, Scotland. He is responsible
for a team of 60 multidisciplinary engineers, including the electrical and
mechanical condition monitoring support
and analysis group. He started his career
as an engineering apprentice in 1977 at
Marconi Communications Ltd., working in
design of digital electronic circuits. Since
graduating in 1981, he has been involved in
the design of data acquisition and measurement systems for
telecommunications, defense systems, and fiscal metering. In the
mid 1980s, he developed one of the first automatic vibration
monitoring systems for large turbo-alternators through the SSEB,
and continued this development with the CEGB North East region
into the early 1990s with systems installed in nine major United
Kingdom power generating plants. For the past 10 years, he has
continued developing online plant condition monitoring equipment, concentrating on both electrical and mechanical condition
monitoring techniques.
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PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 2002
example, stator winding failures can be due to stator core or end
winding vibration. Likewise, abnormal levels of airgap eccentricity
produce high electromagnetic forces that can cause serious bearing
wear and failure. Historically, condition monitoring has focused on
detecting problems, but there is now a need to provide reliable
information to assist with the identification of the root cause of the
fault—operators require solutions to prevent a recurrence of faults.
An integrated monitoring strategy that analyzes both vibration and
current at the same time can provide a better assessment of the
health of an electrical drive and meet the goal of identifying the
root cause of a fault.
SUMMARY OF PROBLEMS
AND FAILURE MECHANISMS
Broken Rotor Bars or End Rings
in Squirrel-Cage Induction Motors
Broken rotor bars or end rings can be caused by the following as
discussed by Bonnet and Soukup (1992) and Finley and
Hodowanec (2001):
• Direct
online starting duty results in high thermal and
mechanical stresses.
• Pulsating mechanical loads such as coal crushers can subject the
rotor cage to high mechanical stresses.
• Imperfections in the manufacturing process of the rotor cage
Although broken rotor bars do not initially cause an induction
motor to fail, there can be serious secondary damage due to broken
parts of the bar hitting the stationary stator winding at high
velocity. The photos shown in Figures 1 and 2 show the case of a
broken bar in a 450 hp/336 kW, four-pole, three-phase, induction
motor-coal crusher drive and the consequential damage to the
stator winding that resulted in a $50,000 complete rewind of the
3.3 kV stator winding. These motors had not been previously
monitored to detect broken bars via motor current signature
analysis (MCSA). In hazardous environments, sparking at the fault
site (during the degradation process) can be a potential safety
hazard.
Figure 2. Stator Winding Failure as a Consequence of Rotor Bar
Problem.
• Following a repair or overhaul, the motor can be reinstalled with
unacceptable levels of airgap eccentricity (e.g., an eccentricity
level of 25 to 30 percent in a large motor is considered to be
severe).
• In unusual cases thermal bowing of the shaft can cause dynamic
eccentricity.
A catastrophic failure can occur if the airgap eccentricity is at a
level such that the resultant unbalanced magnetic pull causes a
rotor to stator rub.
Mechanical Problems in Induction Motor Drives
There is a very substantial knowledge base on vibration
monitoring to detect mechanical problems, but what is not so well
recognized (via condition monitoring) are the electrical problems
that can be caused by mechanical phenomena. The classical
mechanical problems in induction motor drives are as follows:
• Bearing
wear and failure. As a by-product of bearing wear,
airgap eccentricity can increase, and there is the potential for a
rotor to stator rub, serious stator core damage, and stator winding
failure.
• High mechanical unbalance in the rotor—increased centrifugal
forces on the rotor winding
• Looseness—decreased
stiffness in the bearing pedestals can
increase the forces on the rotor winding.
• Shaft/coupling misalignment—results in forces on the bearings
and on rotor and stator windings.
• Problems in gearboxes—forces and vibration transmitted to the
induction motor
• Cavitation
in pumps—cause disturbances to rotor alignment
reflected into stator current.
Figure 1. Broken Rotor Bar.
• Oil whirl/whip in plain bearings
• Critical speeds/shaft resonances—increased forces and vibration
on the rotor core and winding
Airgap Eccentricity in Three-Phase Induction Motors
Airgap eccentricity is kept to a minimum and typical maximum
levels for a large induction motor are between 5 to 10 percent.
There are two types of airgap eccentricity, namely static and
dynamic. In the former, the minimum airgap is fixed in space and,
in the latter, the minimum airgap revolves with the rotor. In
practice there will always be an inherent level of both types due to
manufacturing tolerances, but abnormal levels can occur due to:
• Incorrect onsite installation of a large motor causing abnormal
airgap eccentricity that results in vibration and bearing wear.
Stator Winding Failures
Only a brief summary is given and further details are presented
by Bonnet and Soukop (1992), Thomson (1999), and Thomson
(2001). There are four main root causes of failure:
• Normal ageing and thermal stresses
• Mechanical forces and stresses
• Electrical overloads and switching transients
• Environmental pollution
CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSIS
AND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES
Within these categories there are various subdivisions and combinations of these that enhance the probability of failure. For
example, vibration is probably one of the major causes of
premature degradation of a high-voltage stator winding. The
sources of vibration can come directly from electromagnetic forces
or via transmitted vibration from stator core and teeth vibration, or
external vibration from mechanical problems.
CURRENT AND VIBRATION MONITORING—
INDUSTRIAL CASE HISTORIES
Justification
An induction motor drive is a complex electrical and mechanical
system. For example, mechanical forces and consequential
vibration from a mechanical load such as a coal crusher or reciprocating compressor are transmitted to the motor, and
electromagnetic forces from the motor act on the rotor system.
Vibration signals therefore include a combination of components
from mechanical and electromagnetic sources. A disturbance to the
airgap flux waveform can result in additional flux components
being produced, which in turn can induce current components in
the stator winding. These components can be detected via current
signature analysis with the current being sensed via a current
transformer around one of the supply cables. By analyzing the
current and vibration signals, the electromechanical health of the
drive system can be assessed and the root cause of a problem can
be established.
63
bar severity under full-load conditions can be given by the
following equation, as presented by Thomson and Rankin (1987)
and Hargis, et al. (1982):
n≈ R
(
N / 20
+ p) ≈ ×
(
50 / 20
+
)=
(2)
where:
R = Number of rotor slots
N = dB difference between the sidebands and the supply
frequency component
p = Pole pairs
Taking account of the fact that the motor was on reduced load,
hence the rotor current was less and the estimate was that at least
one bar was broken. It is important to note that if this had been a
two-pole motor with the sidebands at 50 dB down, there would not
be a broken bar problem. The number of bars would be considerably less (e.g., 46 is a typical value) and the broken factor n = 0.29,
which corresponds to a healthy unit. The photo in Figure 4 shows
that there was one broken bar. The broken bar was repaired and the
motor drive was reinstalled. Figure 5 shows the current spectrum
after the repair, and there are no sidebands present. This was a
successful outcome for the diagnosis of the problem.
Case History One—
Power Station Main Auxiliary Drive
FD fan motor drive in a coal fired power station: three-phase, 11
kV, 1620/1150 hp (1208 kW/858 kW), 50 Hz, 495/425 rpm,
78.5/62.5 A, two-speed pole amplitude (PAM), squirrel-cage
induction motor (SCIM). Directly coupled. Rotor bars = 112.
Copper fabricated rotor winding with underslung (nose joint) bar
to end ring brazed joints.
The current and vibration were continuously monitored on all 11
kV strategic drives. Figure 3 shows that twice slip frequency
sidebands (±2sf1) are present in the current spectrum given by
Equation (1), as discussed by Williamson and Smith (1982),
Hargis, et al. (1982), Thomson and Rankin (1987), and Kliman and
Stein (1990).
f sb = f1 ( ± s) Hz
(1)
where:
s = Slip of the induction motor
f1 = Supply frequency, Hz
Figure 4. Faulty Rotor—One Broken Bar.
Figure 5. Current Spectrum after Repair of Broken Bar.
Figure 3. Twice Slip Frequency Sidebands in Current Spectrum.
In the USA these sidebands are often referred to as the pole pass
frequencies. The sidebands are 50 dB down on the supply
frequency component, but with a rotor having 112 slots this
indicates the early signs of a broken rotor bar problem. The drive
was running in the high-speed mode (12 pole configuration), but
the motor was operating on a reduced load with an operational
current of 57 A compared with a full-load current of 78.5 A. It has
been proven via industrial case histories that an estimate of broken
Unfortunately, following the reinstallation it was observed that
the bearing vibration on the motor had changed from 0.05 in/sec
peak (0.9 mm/sec rms) to 0.33 in/sec peak (5.9 mm/sec rms) at the
drive-end and from 0.03 in/sec peak (0.5 mm/sec rms) to 0.1 in/sec
peak (1.9 mm/sec rms) at the nondrive-end. Clearly the motor had
been incorrectly reinstalled. The current was immediately analyzed
and a comparison made between the spectra before and after the
reinstallation. The most likely cause was that an increase in airgap
eccentricity had been introduced, but current analysis would
confirm if that was indeed the case. Abnormal airgap eccentricity
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PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 2002
can be diagnosed via current signature analysis. Equation (3) is
used to detect the unique signature pattern as presented by
Cameron, et al. (1986):
(
)
f ec = f1 R p ( − s) ± nws ± f1 ( − s) p
(3)
where f1, R, s, and p are defined above and nws = integer values of
1, 3, 5, 7, . . .
In Equation (3), the part given by f1(R/p(1⫺s) ± nws) = frs =
rotor slot passing frequencies in the current spectrum were induced
by rotating flux waves due to rotor slotting, which are also a
function of airgap eccentricity. In Equation (3), the part given by
f1(1⫺s)/p = fr = rotational speed frequency component of current
was induced by a rotating flux wave—a function of dynamic airgap
eccentricity and interaction (modulation) with static eccentricity
flux waves as proved by Thomson and Barbour (1998).
Equation (3) therefore gives the rotor slot passing frequencies
(frs) with ± rotational speed frequencies (fr) around the slot passing
frequencies, and it is this pattern that can be used to characterize
abnormal airgap eccentricity. There is a series of these slot passing
frequencies and they are spaced at twice the supply frequency apart.
For this motor, operating on reduced load and at this particular slip,
the rotor slot passing frequencies occur at 780, 880, 980, 1080 Hz
and the rotational speed frequency is at 8.27 Hz (496 rpm). The
rotor slot passing frequency with the highest dB value was selected,
and Figures 6 and 7 show the current spectrum before the outage
and after reinstallation when the vibration had increased by nearly
a factor of six at the drive-end bearing. Examination of these two
spectra clearly shows there are no rotational speed frequency
components around 880 Hz before the outage, but that two distinct
components at ±8.27 Hz (± rotational speed frequency) are present
after the reinstallation. These components are 32 and 36 dB down
on the rotor slot passing frequency, and in this case is indicative of
an increase in airgap eccentricity. It was not possible to remove the
motor from service due to strategic operational demands on the
unit, but the bolts at the drive-end were tightened and the vibration
dropped from 0.33 in/sec peak (5.9 mm/sec rms) to 0.27 in/sec peak
(4.8 mm/sec rms) as shown in Figure 8. It is therefore likely that the
incorrect size of shims had been installed. The motor is under close
surveillance and will be checked at the first available opportunity.
Note that the motor is operating in the 12-pole mode and it is not a
high-speed machine. This case history is a classic case of what can
happen at an outage and reinstallation. One problem was corrected
but another was introduced. However, by the use of current and
vibration monitoring, the root cause of the high vibration was
established and, equally important, when and how it was introduced
was identified.
Figure 8. High Bearing Vibration at the Drive-End after
Reinstallation and Effect of Tightening Bolts.
Case History Two—
Witness Tests
Bearing vibration measurements are carried out during witness
testing as part of a manufacturer’s quality analysis and quality
control procedures, and to provide the client with a record to prove
that the levels are within the specified standards. By applying
current signature analysis, the quality of the rotor winding and any
operational airgap eccentricity problems can also be determined.
The client is also provided with baseline signature patterns at the
time of manufacture.
Motor specification: three-phase, 6.6 kV, 60 Hz, 3.95 MW/5295
hp, 404 A, 3540 rpm (two-pole), SCIM, efficiency = 95.4 percent,
rotor slots = 46, stator slots = 60. Vibration velocity (rms over the
range 10 Hz to 1 kHz) in the vertical, horizontal, and axial
directions to not exceed 2.67 mm/sec (0.15 in/sec peak).
Vibration Results
Drive-end bearing: 1.1, 0.5, 1.5 mm/sec in the vertical,
horizontal, and axial directions, respectively. Nondrive-end
bearing: 1.2, 1.1, 1.2 mm/sec in the vertical, horizontal, and axial
directions, respectively. Note: 1 mm/sec = 0.056 in/sec peak, 1.5
mm/sec = 0.084 in/sec peak. These values were well within the
upper limit and the rotor was running smoothly. The stator core
structure is shown in Figure 9.
Figure 6. Current Spectrum—No ±Fr Components.
Figure 9. Stator Core Assembly.
Figure 7. Current Spectrum—±Fr Components Exist.
As a special exercise the vibration was measured on the core back
at two different positions (at 12 o’clock and 3 o’clock as viewed
from the drive-end): measured velocity of 0.29 in/sec peak (5.2
mm/sec rms). The core vibration (velocity) was of the order of 4.5
CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSIS
AND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES
65
times greater than the bearing velocity levels. The fundamental
reason for this is due to the electromagnetic forces that act directly
on the stator core assembly. These forces are proportional to the flux
density squared, and the fundamental frequency of vibration due to
electromagnetic forces is therefore at twice the supply frequency,
which in this case is 120 Hz. This is confirmed via the vibration
spectrum shown in Figure 10. Note also the classic vibration
components due to rotor slotting as given by Equation (4) and
presented by Alger (1965), Yang (1981), and Cameron, et al. (1986):
(
f rs = f1 R p ( − s) ± nr
)
(4)
where f1, R, s, and p were previously defined and nr = integer
values of 0, 2, 4, 6, . . .
Figure 11. Current Spectrum—Healthy Rotor.
Figure 10. Stator Core Vibration Spectrum.
There is a series of these components all spaced twice the supply
frequency apart. The reason for highlighting these results is that the
stator winding system is contained within the stator core, and interestingly, there are no standards for acceptable stator core vibration levels.
Current Analysis Results
Figure 11 shows the current spectrum for assessing the quality
of the rotor winding. For a nominal full-load speed of 3540 rpm,
the slip s = 0.0167 and any twice slip frequency sidebands (±2sf1)
due to rotor asymmetry will be at ±2 Hz. The actual rotor speed
was 4 rpm higher than the nominal full-load value, and the
operational slip s = 0.0146 gives sidebands at 1.75 Hz. The
sidebands are 58.8 and 66 dB down on the supply component, and
taking the average dB difference gives N = 62.4 dB. From Equation
(2), this gives a broken bar factor of n = 0.07, which means a
perfectly healthy and high quality rotor cage winding.
Figure 12 shows the current spectrum for assessing the
operational airgap eccentricity. For this motor at an operational slip
of s = 0.0146, the rotor slot passing frequencies (frs) from Equation
(3) are: 2539, 2659, 2779, 2899 Hz, etc.—all spaced 120 Hz apart,
and the rotational speed frequency fr = 59.12 Hz.
The principal (nws = ⫹1) slot passing frequency was in fact 2777.5
Hz at the time of measurement and is only 1.5 Hz different from the
prediction (i.e., 0.05 percent difference). The components at ±fr
around the selected rotor slot passing frequency are 50 dB down, and
this corresponds to a normal airgap eccentricity level. Contrast this
with the fr components in Case History One where they were 32 dB
down when there was a problem and high vibration levels. This
motor was clearly a very healthy unit at the time of manufacture.
Case History Three—
Oil Exporting Pump at Oil Tank Farm
Motor specification: four identical motors, three-phase, 1.45
MW (1944 hp), 11 kV, 103 A, 742 rpm, 50 Hz, eight-pole SCIM.
Rotor bars = 62. Onsite personnel had measured the bearing
Figure 12. Current Spectrum—Normal Airgap Eccentricity.
vibration on one of the motors that was only running for 45 to 60
minutes prior to the machine being tripped out at a bearing
temperature of 76°C (168.8°F). Drive-end bearing peak-to-peak
displacement was 4.8 mils (122 µm) and 4.4 mils (111 µm), and at
the nondrive-end 2.7 mils (68 µm) and 2.4 mils (61 µm) in the
vertical and horizontal positions, respectively. For this size and
speed of motor, the classification for a good running machine was
1 mil (25 µm) peak-to-peak, but an upper limit of 2 mils (50 µm)
peak-to-peak is considered to be just acceptable. The measured
levels of 4.8 mils (122 µm) and 4.4 mils (111 µm) were unacceptably high. Vibration spectrum analysis by onsite personnel resulted
in uncertainty as to the root cause of the fault, although it was
suspected that high airgap eccentricity may be the problem.
Current spectrum analysis was used to determine the root cause
of the problem. Figure 13 shows the sidebands at ±2sf1 are 64 dB
down on f1, and this corresponds to a perfectly healthy rotor
winding. Figure 14 is the current spectrum around one of the rotor
slot passing frequencies (frs) given by Equation (3), and there are
±fr components around frs. Since they are only 13 dB down on frs,
this corresponds to an unacceptably high level of airgap eccentricity in a large motor. The airgaps were checked at the drive and
nondrive-ends and found to be 35 percent (a severe eccentricity
problem) and 20 percent, respectively. The gaps were set at the 12,
3, 6, and 9 o’clock positions to within ±5 percent (manufacturer’s
specification) of the nominal airgap length of 100 mils (2.5 mm),
and the vibration levels returned to their normal levels.
Case History Four—
Misalignment and Pump Wear in Electrical Submersible Pumps
Misalignment
In electrical submersible pumps it is not always possible to
monitor vibration. Misalignment in the rotor-coupling system can
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PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 2002
Figures 17 and 18 show the vibration spectrum and, as expected
with parallel misalignment, the second harmonic of the rotational
speed frequency component has increased, in this case by 16.8 dB.
The rotational speed frequency component has also increased but
by a lesser amount of 6.5 dB. The results confirm that both current
and vibration analysis can diagnose misalignment, and there is
potential for current signature analysis to detect mechanical
problems when vibration sensors cannot be installed.
Figure 13. Current Spectrum—Healthy Rotor.
Figure 17. Vibration Spectrum—Aligned.
Figure 14. Current Spectrum—High Airgap Eccentricity.
cause vibration problems, bearing wear, and secondary damage to
seals and connecting pipe work. Initial tests to detect misalignment
via current signature analysis were carried out on a motor
generator set. Figures 15 and 16 show the change in the current
spectrum due to a parallel misalignment of 0.25 mm (9.8 mils). It
is accepted that this is a high level, but it demonstrates the effect of
misalignment on the current and vibration spectra. The
components at ±fr around f1 have increased by nearly 16 dB (i.e.,
the +fr component) and is due to forces acting on the rotor, which
subsequently disturbs the airgap magnetic field in the motor.
Figure 18. Vibration Spectrum—9.8 mils (0.25 mm) Misalignment.
Pump Wear in Electrical Submersible Pumps
Used for Artificial Lift to Extract Oil in Deep Wells
Figure 15. Current Spectrum—Aligned.
Due to the operational environment in deep wells, vibration
sensors cannot be installed. Motor specification: three-phase (two
in tandem), 2300 V, 210 hp/127 kW, SCIMs, 60 Hz, 40 A, twopole, operating at 7500 ft (2286 m) in a well deviation of 45
degrees. Figure 19 shows the change in the current spectrum over
a four month period due to pump wear, and Figure 20 shows a
photo of the worn pump. There is obvious potential to detect an
increase in pump wear, which means the operators can reduce the
load on the pump to obtain a longer life and increase the load on
other pumps in the well. The removal of a failed electrical
submersible pump (ESP) and its replacement can be very costly in
an offshore oil installation ($750,000 to $1 million), hence any
means to better manage their operation and to determine the
problems that can lead to improved designs is an advantage.
CONCLUSIONS
Figure 16. Current Spectrum—9.8 mils (0.25 mm) Misalignment.
The case histories demonstrate the advantages of using an
integrated monitoring strategy via vibration and current analysis to
diagnose faults and establish the root cause of problems. In
electrical drive systems it is better to simultaneously analyze
electrical and mechanical signals such as vibration and current, and
in high-voltage motors (4.16 kV and above) partial discharges
should also be monitored to ascertain the health of the stator
winding. In addition, temperature sensing on bearings and windings
should be included and, where appropriate, the use of thermographic surveys is a useful technique. Electrical and mechanical
CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSIS
AND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES
67
Bonnet, A. H. and Soukup, G. C., 1992, “Cause and Analysis of
Stator and Rotor Failures in Three-Phase Squirrel-Cage
Induction Motors,” IEEE Transactions on Industry Applications, 28, (4), pp. 921-937.
Cameron, J. R., Thomson, W. T., and Dow, A. B., May 1986,
“Vibration and Current Monitoring for Detecting Airgap
Eccentricity in Large Induction Motors,” IEE Proceedings,
133, Part B. (3).
Finley, W. R. and Hodowanec, M. M., November/December 2001,
“Selection of Copper Versus Aluminium Rotors for Induction
Motors,” IEEE Transactions on Industry Applications, 37, (6),
pp.1563-1573.
Hargis, C., Gaydon, B. G., and Kamish, K., 1982, “The Detection
of Rotor Defects in Induction Motors,” Proceedings IEE
EMDA Conference, London, England, pp. 216-220.
Kliman, G. B. and Stein, J., 1990, “Induction Motor Fault
Detection Via Passive Current Monitoring,” Proceedings
International Conference (ICEM’90), Massachusetts Institute
of Technology, Boston, Massachusetts, pp.13-17.
Figure 19. Difference in Current Spectra Due to Worn Pump.
Thomson, W. T., 1999, “A Review of On-Line Condition
Monitoring Techniques for Three-Phase Squirrel-Cage Induction Motors—Past, Present, and Future,” IEEE Symposium on
Diagnostics for Electrical Machines, Power Electronics and
Drives, Gijon, Spain, pp. 3-18 (opening keynote address).
Thomson, W. T., 2001, “On-Line MCSA to Diagnose Shorted
Turns in Low Voltage Stator Windings of 3-Phase Induction
Motors Prior to Failure,” Proceedings of IEEE Conference on
Electrical Machines and Drives (IEMDC), Massachusetts
Institute of Technology, Boston, Massachusetts.
Thomson, W. T. and Barbour, A., December 1998, “On-line
Current Monitoring and Application of a Finite Element
Method to Predict the Level of Airgap Eccentricity in 3-Phase
Induction Motors,” IEEE Transactions on Energy Conversion,
13, (4), pp. 347-357 (includes discussion and closure).
Thomson, W. T.and Rankin, D., 1987, “Case Histories of Rotor
Winding Fault Diagnosis in Induction Motors,” Proceedings
2nd International Conference on Condition Monitoring,
University College of Swansea, Wales, United Kingdom.
Figure 20. Photo of Worn Pump Stage.
engineers must be encouraged to work together and cross the
historical boundaries when applying condition monitoring and
online diagnostics for health care of electrical drives.
REFERENCES
Alger, P. L., 1965, Induction Machines, New York, New York:
Gordon and Breach.
Williamson, S. and Smith, A. C., May 1982, “Steady State
Analysis of 3-Phase Cage Motors with Rotor-Bar and EndRing Faults,” Proceedings IEE, 129, Part B, (3), pp. 93-100.
Yang, S. J., 1981, Low Noise Electric Motors, Monographs in
Electrical and Electronic Engineering, IEE, Savoy Place,
London, England.
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PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 2002