William T. (Bill) Thomson is Directorand Consultant with EM Diagnostics Ltd.,in Alford, Aberdeenshire, Scotland. Hebegan as an apprentice electrician and has40 years’ experience covering the installa-tion, maintenance, design, performance,and condition monitoring of electricaldrives. He has worked as an electrician,R&D engineer with Hoover Ltd., consul-tant, and academic. In 1990, he wasappointed Professor at the Robert Gordon

University, in Scotland, in recognition of his research anddevelopment work on condition monitoring for electrical drives inthe 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 isa senior member of the American IEEE, a Fellow of the IEE, anda registered Chartered Engineer in the United Kingdom. He is alsothe visiting Professor in Electrical Engineering at the University ofAbertay, Scotland, and has published over 70 papers.

Philip (Phil) Orpin is Technical ServicesManager at Ingenco Ltd., in East Kilbride,Lanarkshire, Scotland. He is responsiblefor a team of 60 multidisciplinary en-gineers, including the electrical andmechanical condition monitoring supportand analysis group. He started his careeras an engineering apprentice in 1977 atMarconi Communications Ltd., working indesign of digital electronic circuits. Sincegraduating in 1981, he has been involved in

the design of data acquisition and measurement systems fortelecommunications, defense systems, and fiscal metering. In themid 1980s, he developed one of the first automatic vibrationmonitoring systems for large turbo-alternators through the SSEB,and continued this development with the CEGB North East regioninto the early 1990s with systems installed in nine major UnitedKingdom power generating plants. For the past 10 years, he hascontinued developing online plant condition monitoring equip-ment, concentrating on both electrical and mechanical conditionmonitoring techniques.

ABSTRACT

Induction motor drives are the most widely used electrical drivesystem and they typically consume 40 percent of an industrializednation’s total generating capacity. In the USA the total generatingcapacity is approximately 800,000 MW, consequently inductionmotor drives are major assets in the process and energy industries.The asset management of electrical drives requires reliablemaintenance strategies that include condition monitoring andonline diagnostics. Due to the complex electromagnetic andmechanical characteristics of an electrical drive system, a unifiedmonitoring strategy has distinct advantages over monitoring onlyone parameter (e.g., vibration) to diagnose problems. This paperfocuses on industrial case histories to demonstrate the applicationof current and vibration analysis to diagnose problems ininduction motor drives. The results show how the root cause of aproblem can be established when a combination of current andvibration monitoring is used in comparison to only analyzing onesignal.

INTRODUCTION

Many operators now use online condition-based maintenancestrategies in parallel with conventional planned maintenanceschemes. This has reduced unexpected failures, increased the timebetween planned shutdowns for standard maintenance, andreduced operational costs. However, it is still the operator who hasto make the final decision on whether to remove a machine fromservice or let it run based on information from conditionmonitoring systems. The root cause (RCA) of the fault has to beestablished and, ideally, the operator also requires a prognosis ofthe remaining run life; but the latter is complex and in most casesis an impossible task. Vibration monitoring and analysis are matureand effective techniques to diagnose mechanical problems. Currentmonitoring can detect problems such as broken rotor bars, airgapeccentricity, and more recently shorted turns in low voltage statorwindings.

However there still tends to be a historical culture thatmechanical 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 reportsare presented. With respect to an induction motor drive system, thisdoes not make sound engineering sense since an electrical drive isan interconnected, electromechanical system, and it is not sensibleto analyze signals in isolation. The fault may appear to be electricalbut the fundamental cause may be due to mechanical forces. For

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CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSISAND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES

byWilliam T. ThomsonDirector and Consultant

EM Diagnostics Ltd.

Alford, Aberdeenshire, Scotland

andPhilip Orpin

Technical Services Manager

Ingenco Ltd.

East Kilbride, Lanarkshire, Scotland

example, stator winding failures can be due to stator core or endwinding vibration. Likewise, abnormal levels of airgap eccentricityproduce high electromagnetic forces that can cause serious bearingwear and failure. Historically, condition monitoring has focused ondetecting problems, but there is now a need to provide reliableinformation to assist with the identification of the root cause of thefault—operators require solutions to prevent a recurrence of faults.An integrated monitoring strategy that analyzes both vibration andcurrent at the same time can provide a better assessment of thehealth of an electrical drive and meet the goal of identifying theroot cause of a fault.

SUMMARY OF PROBLEMSAND FAILURE MECHANISMS

Broken Rotor Bars or End Ringsin Squirrel-Cage Induction Motors

Broken rotor bars or end rings can be caused by the following asdiscussed by Bonnet and Soukup (1992) and Finley andHodowanec (2001):

• Direct online starting duty results in high thermal andmechanical stresses.

• Pulsating mechanical loads such as coal crushers can subject therotor cage to high mechanical stresses.

• Imperfections in the manufacturing process of the rotor cage

Although broken rotor bars do not initially cause an inductionmotor to fail, there can be serious secondary damage due to brokenparts of the bar hitting the stationary stator winding at highvelocity. The photos shown in Figures 1 and 2 show the case of abroken bar in a 450 hp/336 kW, four-pole, three-phase, inductionmotor-coal crusher drive and the consequential damage to thestator winding that resulted in a $50,000 complete rewind of the3.3 kV stator winding. These motors had not been previouslymonitored to detect broken bars via motor current signatureanalysis (MCSA). In hazardous environments, sparking at the faultsite (during the degradation process) can be a potential safetyhazard.

Figure 1. Broken Rotor Bar.

Airgap Eccentricity in Three-Phase Induction Motors

Airgap eccentricity is kept to a minimum and typical maximumlevels for a large induction motor are between 5 to 10 percent.There are two types of airgap eccentricity, namely static anddynamic. In the former, the minimum airgap is fixed in space and,in the latter, the minimum airgap revolves with the rotor. Inpractice there will always be an inherent level of both types due tomanufacturing tolerances, but abnormal levels can occur due to:

• Incorrect onsite installation of a large motor causing abnormalairgap eccentricity that results in vibration and bearing wear.

Figure 2. Stator Winding Failure as a Consequence of Rotor BarProblem.

• Following a repair or overhaul, the motor can be reinstalled withunacceptable levels of airgap eccentricity (e.g., an eccentricitylevel of 25 to 30 percent in a large motor is considered to besevere).

• In unusual cases thermal bowing of the shaft can cause dynamiceccentricity.

A catastrophic failure can occur if the airgap eccentricity is at alevel such that the resultant unbalanced magnetic pull causes arotor to stator rub.

Mechanical Problems in Induction Motor Drives

There is a very substantial knowledge base on vibrationmonitoring to detect mechanical problems, but what is not so wellrecognized (via condition monitoring) are the electrical problemsthat can be caused by mechanical phenomena. The classicalmechanical 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 arotor to stator rub, serious stator core damage, and stator windingfailure.

• High mechanical unbalance in the rotor—increased centrifugalforces on the rotor winding

• Looseness—decreased stiffness in the bearing pedestals canincrease the forces on the rotor winding.

• Shaft/coupling misalignment—results in forces on the bearingsand on rotor and stator windings.

• Problems in gearboxes—forces and vibration transmitted to theinduction motor

• Cavitation in pumps—cause disturbances to rotor alignmentreflected into stator current.

• Oil whirl/whip in plain bearings

• Critical speeds/shaft resonances—increased forces and vibrationon the rotor core and winding

Stator Winding Failures

Only a brief summary is given and further details are presentedby 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

PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 200262

Within these categories there are various subdivisions and combi-nations of these that enhance the probability of failure. Forexample, vibration is probably one of the major causes ofpremature degradation of a high-voltage stator winding. Thesources of vibration can come directly from electromagnetic forcesor via transmitted vibration from stator core and teeth vibration, orexternal vibration from mechanical problems.

CURRENT AND VIBRATION MONITORING—INDUSTRIAL CASE HISTORIES

Justification

An induction motor drive is a complex electrical and mechanicalsystem. For example, mechanical forces and consequentialvibration from a mechanical load such as a coal crusher or recip-rocating compressor are transmitted to the motor, andelectromagnetic forces from the motor act on the rotor system.Vibration signals therefore include a combination of componentsfrom mechanical and electromagnetic sources. A disturbance to theairgap flux waveform can result in additional flux componentsbeing produced, which in turn can induce current components inthe stator winding. These components can be detected via currentsignature analysis with the current being sensed via a currenttransformer around one of the supply cables. By analyzing thecurrent and vibration signals, the electromechanical health of thedrive system can be assessed and the root cause of a problem canbe established.

Case History One—Power Station Main Auxiliary Drive

FD fan motor drive in a coal fired power station: three-phase, 11kV, 1620/1150 hp (1208 kW/858 kW), 50 Hz, 495/425 rpm,78.5/62.5 A, two-speed pole amplitude (PAM), squirrel-cageinduction motor (SCIM). Directly coupled. Rotor bars = 112.Copper fabricated rotor winding with underslung (nose joint) barto end ring brazed joints.

The current and vibration were continuously monitored on all 11kV strategic drives. Figure 3 shows that twice slip frequencysidebands (±2sf1) are present in the current spectrum given byEquation (1), as discussed by Williamson and Smith (1982),Hargis, et al. (1982), Thomson and Rankin (1987), and Kliman andStein (1990).

(1)

where:s = Slip of the induction motorf1 = Supply frequency, Hz

Figure 3. Twice Slip Frequency Sidebands in Current Spectrum.

In the USA these sidebands are often referred to as the pole passfrequencies. The sidebands are 50 dB down on the supplyfrequency component, but with a rotor having 112 slots thisindicates the early signs of a broken rotor bar problem. The drivewas running in the high-speed mode (12 pole configuration), butthe motor was operating on a reduced load with an operationalcurrent of 57 A compared with a full-load current of 78.5 A. It hasbeen proven via industrial case histories that an estimate of broken

bar severity under full-load conditions can be given by thefollowing equation, as presented by Thomson and Rankin (1987)and Hargis, et al. (1982):

(2)

where:R = Number of rotor slotsN = dB difference between the sidebands and the supply

frequency componentp = 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 leastone bar was broken. It is important to note that if this had been atwo-pole motor with the sidebands at 50 dB down, there would notbe a broken bar problem. The number of bars would be consider-ably 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 showsthat there was one broken bar. The broken bar was repaired and themotor drive was reinstalled. Figure 5 shows the current spectrumafter the repair, and there are no sidebands present. This was asuccessful outcome for the diagnosis of the problem.

Figure 4. Faulty Rotor—One Broken Bar.

Figure 5. Current Spectrum after Repair of Broken Bar.

Unfortunately, following the reinstallation it was observed thatthe bearing vibration on the motor had changed from 0.05 in/secpeak (0.9 mm/sec rms) to 0.33 in/sec peak (5.9 mm/sec rms) at thedrive-end and from 0.03 in/sec peak (0.5 mm/sec rms) to 0.1 in/secpeak (1.9 mm/sec rms) at the nondrive-end. Clearly the motor hadbeen incorrectly reinstalled. The current was immediately analyzedand a comparison made between the spectra before and after thereinstallation. The most likely cause was that an increase in airgapeccentricity had been introduced, but current analysis wouldconfirm if that was indeed the case. Abnormal airgap eccentricity

CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSISAND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES

63

( )f f s Hzsb = ±1

( ) ( )n R pN≈ + ≈ × + =20 50 20/ /

can be diagnosed via current signature analysis. Equation (3) isused to detect the unique signature pattern as presented byCameron, et al. (1986):

(3)

where f1, R, s, and p are defined above and nws = integer values of1, 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 inducedby rotating flux waves due to rotor slotting, which are also afunction of airgap eccentricity. In Equation (3), the part given byf1(1�s)/p = fr = rotational speed frequency component of currentwas induced by a rotating flux wave—a function of dynamic airgapeccentricity and interaction (modulation) with static eccentricityflux 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 passingfrequencies, and it is this pattern that can be used to characterizeabnormal airgap eccentricity. There is a series of these slot passingfrequencies 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 Hzand the rotational speed frequency is at 8.27 Hz (496 rpm). Therotor slot passing frequency with the highest dB value was selected,and Figures 6 and 7 show the current spectrum before the outageand after reinstallation when the vibration had increased by nearlya factor of six at the drive-end bearing. Examination of these twospectra clearly shows there are no rotational speed frequencycomponents around 880 Hz before the outage, but that two distinctcomponents at ±8.27 Hz (± rotational speed frequency) are presentafter the reinstallation. These components are 32 and 36 dB downon the rotor slot passing frequency, and in this case is indicative ofan increase in airgap eccentricity. It was not possible to remove themotor from service due to strategic operational demands on theunit, but the bolts at the drive-end were tightened and the vibrationdropped 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 theincorrect size of shims had been installed. The motor is under closesurveillance and will be checked at the first available opportunity.Note that the motor is operating in the 12-pole mode and it is not ahigh-speed machine. This case history is a classic case of what canhappen at an outage and reinstallation. One problem was correctedbut another was introduced. However, by the use of current andvibration monitoring, the root cause of the high vibration wasestablished and, equally important, when and how it was introducedwas identified.

Figure 6. Current Spectrum—No ±Fr Components.

Figure 7. Current Spectrum—±Fr Components Exist.

Figure 8. High Bearing Vibration at the Drive-End afterReinstallation and Effect of Tightening Bolts.

Case History Two—Witness Tests

Bearing vibration measurements are carried out during witnesstesting as part of a manufacturer’s quality analysis and qualitycontrol procedures, and to provide the client with a record to provethat the levels are within the specified standards. By applyingcurrent signature analysis, the quality of the rotor winding and anyoperational airgap eccentricity problems can also be determined.The client is also provided with baseline signature patterns at thetime of manufacture.

Motor specification: three-phase, 6.6 kV, 60 Hz, 3.95 MW/5295hp, 404 A, 3540 rpm (two-pole), SCIM, efficiency = 95.4 percent,rotor slots = 46, stator slots = 60. Vibration velocity (rms over therange 10 Hz to 1 kHz) in the vertical, horizontal, and axialdirections 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-endbearing: 1.2, 1.1, 1.2 mm/sec in the vertical, horizontal, and axialdirections, respectively. Note: 1 mm/sec = 0.056 in/sec peak, 1.5mm/sec = 0.084 in/sec peak. These values were well within theupper limit and the rotor was running smoothly. The stator corestructure is shown in Figure 9.

Figure 9. Stator Core Assembly.

As a special exercise the vibration was measured on the core backat two different positions (at 12 o’clock and 3 o’clock as viewedfrom the drive-end): measured velocity of 0.29 in/sec peak (5.2mm/sec rms). The core vibration (velocity) was of the order of 4.5

PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 200264

( )( ) ( )f f R p s n f s pec ws= − ± ± −1 1

times greater than the bearing velocity levels. The fundamentalreason for this is due to the electromagnetic forces that act directlyon the stator core assembly. These forces are proportional to the fluxdensity squared, and the fundamental frequency of vibration due toelectromagnetic forces is therefore at twice the supply frequency,which in this case is 120 Hz. This is confirmed via the vibrationspectrum shown in Figure 10. Note also the classic vibrationcomponents due to rotor slotting as given by Equation (4) andpresented by Alger (1965), Yang (1981), and Cameron, et al. (1986):

(4)

where f1, R, s, and p were previously defined and nr = integervalues of 0, 2, 4, 6, . . .

Figure 10. Stator Core Vibration Spectrum.

There is a series of these components all spaced twice the supplyfrequency apart. The reason for highlighting these results is that thestator winding system is contained within the stator core, and interest-ingly, there are no standards for acceptable stator core vibration levels.

Current Analysis Results

Figure 11 shows the current spectrum for assessing the qualityof 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 speedwas 4 rpm higher than the nominal full-load value, and theoperational slip s = 0.0146 gives sidebands at 1.75 Hz. Thesidebands are 58.8 and 66 dB down on the supply component, andtaking the average dB difference gives N = 62.4 dB. From Equation(2), this gives a broken bar factor of n = 0.07, which means aperfectly healthy and high quality rotor cage winding.

Figure 12 shows the current spectrum for assessing theoperational airgap eccentricity. For this motor at an operational slipof 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 2777.5Hz at the time of measurement and is only 1.5 Hz different from theprediction (i.e., 0.05 percent difference). The components at ±fraround the selected rotor slot passing frequency are 50 dB down, andthis corresponds to a normal airgap eccentricity level. Contrast thiswith the fr components in Case History One where they were 32 dBdown when there was a problem and high vibration levels. Thismotor 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, 1.45MW (1944 hp), 11 kV, 103 A, 742 rpm, 50 Hz, eight-pole SCIM.Rotor bars = 62. Onsite personnel had measured the bearing

Figure 11. Current Spectrum—Healthy Rotor.

Figure 12. Current Spectrum—Normal Airgap Eccentricity.

vibration on one of the motors that was only running for 45 to 60minutes prior to the machine being tripped out at a bearingtemperature of 76°C (168.8°F). Drive-end bearing peak-to-peakdisplacement was 4.8 mils (122 µm) and 4.4 mils (111 µm), and atthe nondrive-end 2.7 mils (68 µm) and 2.4 mils (61 µm) in thevertical and horizontal positions, respectively. For this size andspeed of motor, the classification for a good running machine was1 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 measuredlevels of 4.8 mils (122 µm) and 4.4 mils (111 µm) were unaccept-ably high. Vibration spectrum analysis by onsite personnel resultedin uncertainty as to the root cause of the fault, although it wassuspected that high airgap eccentricity may be the problem.

Current spectrum analysis was used to determine the root causeof the problem. Figure 13 shows the sidebands at ±2sf1 are 64 dBdown on f1, and this corresponds to a perfectly healthy rotorwinding. Figure 14 is the current spectrum around one of the rotorslot 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 eccen-tricity in a large motor. The airgaps were checked at the drive andnondrive-ends and found to be 35 percent (a severe eccentricityproblem) and 20 percent, respectively. The gaps were set at the 12,3, 6, and 9 o’clock positions to within ±5 percent (manufacturer’sspecification) 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 tomonitor vibration. Misalignment in the rotor-coupling system can

CURRENT AND VIBRATION MONITORING FOR FAULT DIAGNOSISAND ROOT CAUSE ANALYSIS OF INDUCTION MOTOR DRIVES

65

( )( )f f R p s nrs r= − ±1

Figure 13. Current Spectrum—Healthy Rotor.

Figure 14. Current Spectrum—High Airgap Eccentricity.

cause vibration problems, bearing wear, and secondary damage toseals and connecting pipe work. Initial tests to detect misalignmentvia current signature analysis were carried out on a motorgenerator set. Figures 15 and 16 show the change in the currentspectrum due to a parallel misalignment of 0.25 mm (9.8 mils). Itis accepted that this is a high level, but it demonstrates the effect ofmisalignment on the current and vibration spectra. Thecomponents at ±fr around f1 have increased by nearly 16 dB (i.e.,the +fr component) and is due to forces acting on the rotor, whichsubsequently disturbs the airgap magnetic field in the motor.

Figure 15. Current Spectrum—Aligned.

Figure 16. Current Spectrum—9.8 mils (0.25 mm) Misalignment.

Figures 17 and 18 show the vibration spectrum and, as expectedwith parallel misalignment, the second harmonic of the rotationalspeed frequency component has increased, in this case by 16.8 dB.The rotational speed frequency component has also increased butby a lesser amount of 6.5 dB. The results confirm that both currentand vibration analysis can diagnose misalignment, and there ispotential for current signature analysis to detect mechanicalproblems when vibration sensors cannot be installed.

Figure 17. Vibration Spectrum—Aligned.

Figure 18. Vibration Spectrum—9.8 mils (0.25 mm) Misalignment.

Pump Wear in Electrical Submersible PumpsUsed for Artificial Lift to Extract Oil in Deep Wells

Due to the operational environment in deep wells, vibrationsensors cannot be installed. Motor specification: three-phase (twoin tandem), 2300 V, 210 hp/127 kW, SCIMs, 60 Hz, 40 A, two-pole, operating at 7500 ft (2286 m) in a well deviation of 45degrees. Figure 19 shows the change in the current spectrum overa four month period due to pump wear, and Figure 20 shows aphoto of the worn pump. There is obvious potential to detect anincrease in pump wear, which means the operators can reduce theload on the pump to obtain a longer life and increase the load onother pumps in the well. The removal of a failed electricalsubmersible pump (ESP) and its replacement can be very costly inan offshore oil installation ($750,000 to $1 million), hence anymeans to better manage their operation and to determine theproblems that can lead to improved designs is an advantage.

CONCLUSIONS

The case histories demonstrate the advantages of using anintegrated monitoring strategy via vibration and current analysis todiagnose faults and establish the root cause of problems. Inelectrical drive systems it is better to simultaneously analyzeelectrical and mechanical signals such as vibration and current, andin high-voltage motors (4.16 kV and above) partial dischargesshould also be monitored to ascertain the health of the statorwinding. In addition, temperature sensing on bearings and windingsshould be included and, where appropriate, the use of thermo-graphic surveys is a useful technique. Electrical and mechanical

PROCEEDINGS OF THE THIRTY-FIRST TURBOMACHINERY SYMPOSIUM • 200266

Figure 19. Difference in Current Spectra Due to Worn Pump.

Figure 20. Photo of Worn Pump Stage.

engineers must be encouraged to work together and cross thehistorical boundaries when applying condition monitoring andonline diagnostics for health care of electrical drives.

REFERENCES

Alger, P. L., 1965, Induction Machines, New York, New York:Gordon and Breach.

Bonnet, A. H. and Soukup, G. C., 1992, “Cause and Analysis ofStator and Rotor Failures in Three-Phase Squirrel-CageInduction Motors,” IEEE Transactions on Industry Appli-cations, 28, (4), pp. 921-937.

Cameron, J. R., Thomson, W. T., and Dow, A. B., May 1986,“Vibration and Current Monitoring for Detecting AirgapEccentricity 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 InductionMotors,” IEEE Transactions on Industry Applications, 37, (6),pp.1563-1573.

Hargis, C., Gaydon, B. G., and Kamish, K., 1982, “The Detectionof Rotor Defects in Induction Motors,” Proceedings IEEEMDA Conference, London, England, pp. 216-220.

Kliman, G. B. and Stein, J., 1990, “Induction Motor FaultDetection Via Passive Current Monitoring,” ProceedingsInternational Conference (ICEM’90), Massachusetts Instituteof Technology, Boston, Massachusetts, pp.13-17.

Thomson, W. T., 1999, “A Review of On-Line ConditionMonitoring Techniques for Three-Phase Squirrel-Cage Induc-tion Motors—Past, Present, and Future,” IEEE Symposium onDiagnostics for Electrical Machines, Power Electronics andDrives, Gijon, Spain, pp. 3-18 (opening keynote address).

Thomson, W. T., 2001, “On-Line MCSA to Diagnose ShortedTurns in Low Voltage Stator Windings of 3-Phase InductionMotors Prior to Failure,” Proceedings of IEEE Conference onElectrical Machines and Drives (IEMDC), MassachusettsInstitute of Technology, Boston, Massachusetts.

Thomson, W. T. and Barbour, A., December 1998, “On-lineCurrent Monitoring and Application of a Finite ElementMethod to Predict the Level of Airgap Eccentricity in 3-PhaseInduction 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 RotorWinding Fault Diagnosis in Induction Motors,” Proceedings2nd International Conference on Condition Monitoring,University College of Swansea, Wales, United Kingdom.

Williamson, S. and Smith, A. C., May 1982, “Steady StateAnalysis of 3-Phase Cage Motors with Rotor-Bar and End-Ring Faults,” Proceedings IEE, 129, Part B, (3), pp. 93-100.

Yang, S. J., 1981, Low Noise Electric Motors, Monographs inElectrical and Electronic Engineering, IEE, Savoy Place,London, England.

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