Evaluation of the Detectability of Broken Rotor Bars for Double Squirrel Cage Rotor Induction Motors Jongbin Park, Byunghwan Kim, Jinkyu Yang, and Sang Bin Lee*
Ernesto J. Wiedenbrug, and Mike Teska
Seungoh Han
Department of Electrical Engineering, Korea University, Seoul, Korea *Email:
[email protected]
Baker Instrument CompanyAn SKF Group Company Fort Collins, CO, U.S.A.
Research Center for Convergence Technology, Hoseo University Asan, Chungnam, Korea
Abstract—Double squirrel cage rotor design is employed in induction motor applications that require high starting torque and high efficiency operation. The outer cage of double cage rotors is vulnerable to fatigue failure since it must withstand the large thermal/mechanical stresses experienced during a loaded startup due to the high starting current and long acceleration time. However, there are only a few publications that investigate broken bar detection for double cage induction motors. In this paper, the detectability of broken outer cage bars in double cage motors for the most commonly used rotor bar test methods is evaluated. A finite element and experimental study show that the sensitivity of on-line MCSA is significantly decreased, whereas that of off-line standstill tests is not influenced for broken outer cage bars. This suggests that one should be aware of the insensitivity of MCSA for double cage rotors, and there is a need for development of new on-line monitoring methods. Keywords—AC Machine; Broken Rotor Bar; Condition Monitoring; Diagnostics; Double Cage Induction Motor; Off-line Testing; Spectrum Analysis; Squirrel-Cage Rotor.
I. INTRODUCTION The double squirrel-cage rotor design is mainly employed to meet the high locked rotor torque and low rated slip requirements specified in NEMA design C (or “high torque” motor for > 500hp) or IEC design H induction motors [1]-[2]. Double cage motors can be characterized by the high starting torque and efficient steady state operation; therefore, they are used in applications that require a loaded startup and continuous operation such as conveyors, crushers, stirrers, compressors, loaded pumps, etc [1]-[3]. Although the manufacturing cost is increased with the double cage structure, superior starting characteristics can be obtained without sacrificing the steady state performance (efficiency) since the deep bar effect of single cage deep bar rotors is accentuated.
for the outer bar temperature to exceed 200oC, and the thermomechanical stresses combined with rotational forces, vibration, and residual stresses, make the outer cage of double cage rotors vulnerable to fatigue failure [8]. If broken outer bars are left undetected/unrepaired, they can vibrate in the slot, and the inter-bar current from the broken bars to the adjacent bars can damage the laminations due to arcing [5], [9]-[10]. Broken bars also result in redistribution and increase of the current (and thermal stress) in the adjacent bars at motor startup causing breakage of additional rotor bars. This degrades the starting performance (increase in startup time) of the motor until excessive heating or motor trip during startup (startup failure) becomes an unsustainable problem. A photograph of a double cage rotor with contiguous broken outer cage bars is shown in Fig. 2 [8]. This motor was used for driving a gangway conveyor for coal-mining, and the problem with the rotor was discovered because the motor failed to start. The arguments above show that it is important to detect and correct rotor failure in its incipient stage to prevent costly forced outages due to motor failure. Many on-line and off-line techniques for detecting broken bars have been developed and used in the field over a long period of time [5], and recently, there has been lot of research
It has been shown in many resources that rotor faults are caused by a combination of motor operating stresses and manufacturing defects, and they account for up to 10% of induction motor failures [3]-[6]. For the double cage design, the rotor current is essentially confined to the outer cage during the startup transient when s=1 due to the dominance of the high leakage inductance of the inner cage, as can be seen in Fig. 1. Since double cage motors are used for applications where the motor is loaded at startup, the outer cage must handle the large starting current for the long acceleration time with limited path for heat dissipation (Fig. 1) [7]. Therefore, it is not uncommon
978-1-4244-5287-3/10/$26.00 ©2010 IEEE
2493
Fig. 1. Electrical circuit representation of double squirrel cage rotor
Fig. 2. Photograph of a 300kW, 1000V double cage induction motor rotor that failed to start due to broken outer cage bars [8]
effort in academia on the analysis and development of advanced broken bar detection algorithms [11]-[12]. However, most of the work focus on the analysis and detection of single cage rotor failures, and only a few publications in the literature investigate broken bar detection for double cage rotors [9]-[10], [13]. In [9]-[10], it is shown that there are significant inter-bar currents in double cage induction motors with broken bars in either cage, which makes broken bar detection insensitive, especially for fabricated rotors. In [13], a steady state model for analyzing inner and outer bar failures is derived, and a method for discriminating between broken inner/outer bars is presented. The authors address the problem of detecting outer cage failures in steady state based on a theoretical analysis of their model, and suggest the negative sequence impedance as a means of detecting outer cage failures. However, the practicality is yet to be verified experimentally under actual motor operation. It can be seen that the outer cage of double cage rotors is vulnerable to failure; however, a literature search shows that there has not been much work performed on the analysis or detection of outer cage faults. In this paper, the detectability of rotor bar failures for double cage induction motors is evaluated for several types of fabricated copper bar double cage rotor designs using commercially available rotor bar test methods extensively used in the field. The rotor fault detection methods proposed in the literature but not commonly used in the field are excluded from this study. II.
design, the end ring impedance of Fig. 3 is included as a part of the corresponding inner and outer cage impedances. The resistance of the outer cage, Rr,o, is larger than that of the inner cage, Rr,i, due to the smaller cross sectional area (Fig. 1) and since it is usually made of brass or bronze (for fabricated rotors), which has higher resistivity than copper [3]-[4], [15]. The leakage inductance of the inner cage, Xlr,i, is larger than that of the outer cage, Xlr,i, since it is less tightly coupled to the stator, and because the inner bar leakage flux crosses the low reluctance path between the two bars (Fig. 1). During motor startup when the slip is high (s≈1), the large value of Xlr,i forces most of the rotor current, Ir, to be in the high-resistance outer cage. This enables double cage motors to have superior starting performance of high starting torque with low starting current. When operating in steady state at low slip (s≈0), the influence of Xlr,i and Xlr,o are negligible, and Ir stays mainly in the low resistance inner bar, as can be predicted from Figs. 1 and 3, resulting in high efficiency operation. The increase in Xlr results in lower pullout torque and power factor, and decreases the efficiency to some degree. In addition, the thermal load on the outer cage during startup and the increased number of rotor bars increases the chances of rotor failure. However, the large values of Rr,o and Xlr,i exaggerate the deep bar effect and results in high starting torque and steady state efficiency almost as high as that of a deep bar rotor. The parameters of a 440V-32A, 25hp, NEMA design C double cage induction motor with common end rings are shown
MODELING & ANALYSIS OF DOUBLE SQUIRREL CAGE INDUCTION MOTORS
In a double cage rotor, a single (common) end ring may be used to short both inner and outer cages, or two individual (separate) end rings may be used for each set of cage. All aluminum die-cast rotors are built with common end rings, but most fabricated copper bar rotors typically employ physically separated end rings [13]. Although the common end ring is advantageous for manufacturing cost, the separate end ring design has higher starting torque due to the larger value of Rr (higher resistivity and thinner end ring) [16]. This design is also less susceptible to bar breakage since the independent thermal expansion and contraction of the two cages and end rings applies less thermo-mechanical stress in the bar overhangs and bar-end ring joints of the outer cage [3], [15]. The electrical equivalent circuit of double cage induction motors with common end rings can be derived as shown in Fig. 3 [3], [7], [14]-[16]. In the circuit, subscripts i, o, and er represent the inner cage, outer cage, and end ring of the rotor, respectively, and Xlr,io represents the mutual leakage inductance between the inner and outer cages. Xlr,i, Xlr,o represent the leakage inductances of the inner and outer cage minus the mutual leakage inductance Xlr,io. The value of Xlr,o is very small and usually assumed to be zero since most of the outer bar leakage flux links the low reluctance path inner bar and contributes to Xlr,io, as can be seen in Fig 1. The end ring leakage Xlr,er is also very small and neglected in many cases, especially for fabricated rotors due to the physical separation between the end ring and rotor core. For the separate end ring
2494
Fig. 3. Electrical equivalent circuit representation of double squirrel cage rotor with common end ring Table I. Parameters of 25hp common end-ring double cage motor [14]
Fig 4. Rotor current in inner and outer bar of double cage rotor as a function of speed
in Table I [14]. The rotor current distribution in the inner and outer bars of this motor is shown in Fig. 4 as a function of slip. It can be clearly seen that the rotor current is mainly confined to in the outer bar under the startup transient at high slip, and transitions to the inner bar during steady state operation at low slip (inner/outer bar current 75.0A/33.9A at rated speed). The distribution between the inner and outer bar current depends on the rotor bar material and slot/end ring design. III. EVALUATION OF ON-LINE MONITORING TECHNIQUES FOR BROKEN OUTER CAGE BAR DETECTION A. Existing On-line Monitoring Techniques On-line monitoring of rotor faults relies on indirectly observing the influence of the asymmetry in the rotor flux pattern due to the re-distribution of the current caused by the broken bar. The frequency spectrum of the speed, vibration, torque, axial/radial flux, or current measurement or estimate can be used for broken bar detection [5], [11]-[12], [17]-[19]. Motor current signature analysis (MCSA) is popular since it provides monitoring based on a current sensor that is available in most systems. In MCSA, the frequency component shown in (1) can be observed under steady state operation for rotor fault diagnosis, where fs is the fundamental frequency, and s is the slip.
f brb = (1 ± 2s ) ⋅ f s
(1)
However, MCSA is known to be ineffective for applications where the load constantly changes, since the slip and fbrb also change. It is also unreliable under low slip operation when the motor is lightly loaded since the fbrb component is small and close to the input electrical frequency, fs. In addition to the extraction of small and varying fbrb signatures, other difficulties include the requirement for accurate speed information, interference with load or coupling problems, and heavy data processing capability. Recently, many different types of on-line MCSA-based broken bar detectors have become commercially available [18][19]. However, it is difficult to define the acceptance criteria for broken bars since the increase in fbrb depends on many factors such as the rotor design and operating environment. There are currently no international guidelines or standards for interpretation of test results on the fault severity of broken rotor bars. The default warning or alarm level of fbrb for broken bars is usually set between -45~-35dB with respect to the fundamental frequency component [18]-[19].
taken into account in the derivation of (2), it provides the approximate relationship between the magnitude of the fbrb component in the stator current and fault severity. It can be seen in (2) that the magnitude of the broken bar sideband current components is proportional to the magnitude of the rotor current for a given severity of fault. Since it is the asymmetry in the rotor current distribution due to broken bars that is observed, a strong signature for detection can be obtained if the rotor current is large. This applies to all on-line spectrum analysis based techniques whether it is the current, speed, vibration, torque, or flux measurement that is used. This can be observed in Fig. 10, which show the current spectrum for 0 to 3 broken bars as a function of load. It can be seen that the magnitude of the fbrb component increases with load (rotor current magnitude). B. Evaluation of the Detectability of Broken Outer Bars An equation similar to (2) can be derived for broken outer cage bars, as shown in (3), if it is assumed that the current in the broken outer cage bar is redirected to the adjacent outer cage bars (and not the inner cage bar).
I s ,brb ≈
nr I r ,o Nr
(3)
Ir,o represents the current in the rotor outer cage, as shown in Fig. 3. This equation shows that the broken bar sideband component observable in the stator current is small since the outer bar current is small for double cage motors operating under steady state. To observe the current distribution in the inner/outer cages in the presence of broken outer bars, a 2-dimensional finite element simulation was performed using the Finite Element Method Magnetics (FEMM) software. The 7.5hp double cage fabricated copper bar rotor used for experimental verification in V was used for the simulation (details presented in V.A). To simulate the rotor current distribution under rated load operating conditions, a steady state AC analysis was performed with the stator excited with a rotating magnetic field at the rated slip frequency of 2Hz (rated current), as shown in Fig. 5(a) (rotor reference frame). The simulated results of the magnitude of the slip frequency current component for each inner and outer cage bar for a faulty rotor with 3 broken outer bars (bars 21-23) are shown in Fig. 6(a) (percentage of the total bar current of healthy rotor is shown). It can be seen in this
In [11], [17], the relationship between the magnitude of the sideband components shown in (1) and the broken bar fault severity is derived for single cage rotors, as shown in (2). I s ,(1−2 s ) f + I s ,(1+ 2 s ) f I nr ≈ = s ,brb Nr Ir Ir
(2)
Nr and nr represent the number of rotor bars and contiguous broken bars, respectively, and subscripts (1±2s)f represent the frequency components of the stator current. Although the influence of the inter-bar current, magnetic asymmetry and saturation, end ring impedance, and bar reactance were not
2495
(a) (b) Fig 5. 2 dimensional AC steady state finite element simulation for broken outer cage bars for a common end ring fabricated Cu bar rotor under (a) rated slip (2hz) rotating field & (b) 60hz pulsating field excitation.
figure that the magnitude of the outer bar current is smaller by a factor of over 3, which will result in a small fbrb component in the stator current, as shown in (3). In addition to the magnitude of the outer bar current, the distribution of the current also influences the detectability of broken outer bars. For a single cage motor, the current in the bars adjacent to the broken bars will increase since there is no alternate path, and this distorts the rotor flux distribution [11]. If the outer cage bars break for a double cage motor, the current in the inner cage bars will also increase as it provides a low impedance path under low slip operation. It can be seen in Fig. 5 that most the broken bar current is diverted to the inner cage instead of the adjacent outer cage bars. This will make broken bar detection more insensitive since the increased inner bar current tends to smoothen the asymmetry in the rotor flux distribution caused by the broken outer bar. The presence of inter-bar currents, which are ignored in this simulation, will also make MCSA insensitive [10]. It can be predicted from Figs. 4, 6(a) and (3) that steady state spectrum analysis techniques are insensitive to broken outer cage bars in double cage motors when compared to broken bars in single cage motors. This is mainly due to the small current in the outer cage, and the smoothening effect of the increase in the inner bar current with a broken outer bar. Broken outer cage bar detection is expected to be more difficult for separate end ring rotors due to the higher end ring resistance. It is even more difficult if the outer bar made of high-resistance material, since the difference between the inner and outer bar resistances is larger. It should be noted that the brass outer cage separate end ring fabricated copper bar rotor, in which the broken outer bar is difficult to detect with MCSA, is most common in large double cage induction motors.
bar increases the equivalent rotor impedance, the current and power decrease at certain rotor positions (a broken bar is suspected if the variation in current magnitude exceeds 3%). This test is also very effective and does not require motor disassembly; however, manual rotor rotation is difficult when the rotor is placed in a hostile, inaccessible environment, when the motor is large (sleeve bearings) or of vertical type, or when the load must be removed. A common shortcoming of all off-line periodic rotor inspection techniques is the inability to perform testing frequently. There are also cases where bars disconnected or cracked during operation are shorted at standstill when the rotor is cold [5]. In some facilities, the off-line tests are performed after heating the rotor in an oven to test for broken bars under actual operating temperature.
B. Evaluation of the Detectability of Broken Outer Bars As described in IV.A, the high current excitation, growler, and single phase rotation tests are performed with the rotor at standstill under 60hz pulsating field excitation. To observe the current distribution in the inner/outer cages in the presence of broken outer bars at motor standstill, a steady state FE analysis was performed with the stator excited with a 60hz pulsating field (rated current). In this simulation, the relative position between the rotor and pulsating field was set so that maximum voltage is induced at the location of the broken bars (bars 2123), as shown in Fig. 5(b). The simulated results of the rotor current magnitude for each inner and outer cage bar for a faulty
IV. EVALUATION OF OFF-LINE TESTS FOR BROKEN OUTER CAGE BAR DETECTION
A. Existing Off-line Rotor Bar Tests Many off-line tests for inspecting the rotor for broken bars exist since they are essential for quality assurance in manufacturing facilities, regular inspection at end-user sites, and verification of repair at service shops [5]. The traditional off-line tests that require motor disassembly for rotor testing are visual inspection, fluorescent dye penetrant test (detection of hairline cracks), ohmmeter test, tap test, and high current excitation test. The growler test is a more commonly used test, wherein current is induced in the rotor bars with 60hz external excitation. The abnormal flux pattern produced by the induced rotor bar current is observed with magnetic imaging on the rotor surface or 120hz vibration of a hacksaw blade placed along the bar [5]. Although these off-line tests require the rotor to be removed from the motor, they are known to provide reliable assessment of rotor bar quality since they observe the influence of the broken bar directly at the rotor surface. In the single-phase rotation test, an alternating field is produced in the motor by exciting two phases with 1/8~1/4 rated AC voltage at 60hz with the motor assembled. The change in current magnitude or input power due to a broken bar is observed while slowly turning the rotor [5]. Since a broken
2496
(a)
(b) Fig 6. Distribution of rotor current in inner and outer bars with 3 broken outer cage bars (21-23) for a common end ring fabricated Cu bar rotor under (a) rated load operation; (b) pulsating field excitation (% of total bar current of healthy motor)
rotor with 3 broken outer bars are shown in Fig. 6(b) (% of total bar current of healthy rotor is shown). It can be seen that the majority of the rotor current is in the outer cage because the large leakage inductance of the inner cage has a dominant effect at 60hz standstill excitation. It can also be seen that the current of the broken bar is redirected to the adjacent outer bars and to the inner bar counterparts of the broken outer bars. As in Fig. 6(a), the increase in the inner bar current will tend to “smoothen” the flux distribution, but the effect is expected to be minor considering the relative magnitude of the inner bar current and its distance from the rotor surface. Since the outer cage current is dominant and the influence of the inner cage is not significant, the flux distribution at the rotor surface is similar to that of a broken bar for single cage motors. Therefore, it can be predicted from the results in Figs. 4, 6(b) that broken outer cage bars would be as clearly detectable as broken bars for single cage motors. For separate end ring rotors with high resistivity outer bar material, the sensitivity to broken outer cage bars is expected to decrease to some degree with off-line tests due to the relative decrease in the outer bar current.
To simulate broken bar conditions, the contact between the rotor bars and end rings were cut at the outer bar-end ring joint for fabricated rotors. For the aluminum die cast rotor, the broken bar condition was created by drilling a hole at the joint between the bar and end ring at the slot portion. On-line MCSA and off-line growler/single phase rotation tests were performed on the healthy and faulty rotor samples with 0~3 broken bars to evaluate the detectability of broken outer cage bars.
B. Experimental Results – On-line Monitoring Experimental tests for MCSA performed on samples 1-3 are shown in Figs. 10(a)-(c), respectively, when operating under 33-100% rated load for 0-3 broken bars. The motor load was controlled by adjusting the field voltage of a 30hp DC generator coupled to the motor. Commercial current sensors and a 16 bit data acquisition board were used to measure the current at 6khz sampling for 50secs under steady state conditions. The FFT of the fbrb left sideband component with
V. EXPERIMENTAL STUDY
A. Experimental Setup To evaluate the detectability of broken outer cage bars of double cage induction motors using standard test methods, an experimental study was performed on a 4P, 380V, 1740RPM, 7.5hp induction motor. The original rotor of this test motor is an aluminum die cast rotor with 44 rotor slots (sample 1). To test double cage motors, two 44 slot fabricated copper bar double cage rotors (samples 2 and 3) were designed and built to fit the stator of the original motor. The 3 rotor samples that were tested are summarized as follows, and shown in Fig. 7. • Sample 1: Single cage deep bar, Al die-cast rotor (Fig. 7(a))
(a) (b) (c) Fig. 7. Rotor test samples: (a) original Al die cast single cage rotor (sample 1); (b) fabricated Cu bar double cage rotor with common end ring (sample 2); (c) fabricated Cu bar double cage rotor with separate end ring (Cu inner bar & brass outer bar, sample 3)
• Sample 2: Double cage, fabricated Cu bar rotor with common end ring (Figs. 7(b), 8, 9(a)) • Sample 3: Double cage, fabricated Cu bar rotor with separate end ring (brass outer cage, Figs. 7(c), 8, 9(b)) Fabricated rotors with copper bars were built for the experimental evaluation because they are more representative of large motors, as motors rated above 500hp are typically fabricated rotor type. The design of the two double cage rotors was based on the typical design rules used for large motors. The laminations for samples 2 and 3 were punched to fit the stator and shaft of the original motor, as shown in Fig. 8. The inner and outer bars were inserted into the punched slots, and skewed by one stator slot pitch (identical to sample 1). After brazing the rotor bars to the end rings, the rotors were put through standard lathing (0.5mm airgap) and balancing operations. The bar and end ring design for the two double cage fabricated rotors with common (sample 2, Fig. 7(b)) and separate (sample 3, Fig. 7(c)) end ring design are shown in Fig. 9(a)-(b), respectively.
2497
Fig. 8. Lamination of double cage fabricated copper bar rotors (samples 2-3, Figs. 7(b)-(c))
(a) (b) Fig. 9. Bar and end ring design for (a) common end ring rotor (sample 2, Fig. 7(b)) and (b) separate end ring rotor (sample 3, Fig. 7(c))
respect to the fundamental frequency component are plotted in dB in Fig. 10 (The left sideband at (1-2s)fs was always larger than the right sideband at (1+2s)fs due to the large inertia of the 30hp DC machine [17]).
was calculated from Fig. 11 by extracting the component that repeats itself twice per 180 degree rotor rotation via FFT. The % change in current is plotted for the three samples as a function of the number of broken bars in Fig. 12.
It can be seen in Fig. 10 that the sensitivity of MCSA is lower for the broken outer bars of fabricated double cage rotors when compared to the broken single cage bar, as predicted in III.B. The fbrb component for sample 2 is lower by approximately 8-12dB compared to that of sample 1 under full load condition, which corresponds to a decrease in the fbrb component by a factor of 3~4. For the brass outer bar separate end ring rotor (sample 3), which is more common in large double cage machines, the fbrb components are even smaller due to larger difference between Rr,o and Rr,i in steady state. The fbrb component for sample 3 is lower than that of sample 1 by approximately 18-22dB under full load, which corresponds to a decrease by a factor of 8~12 times.
It can be observed clearly in Fig. 11 that there is a variation in the stator current pattern for all 3 samples with broken bars. The results in Fig. 12 show that there is not much difference in the sensitivity for completely broken single cage bars (sample 1, Fig. 11(a)) and broken outer cage bars of common end ring
There is a clear increase in fbrb in Fig. 10 with the number of broken bars for all three cases, which is observable with careful trending of the fbrb component over time. However, it should be noted that the tests were performed in a lab environment using high-end data acquisition equipment with the load controlled to be constant for 50secs. In the field environment, the background noise level will rise and the fbrb component could decrease due to variations in load. There could also be coupling or load related characteristics or problems that produce components in the vicinity of fbrb that may obscure fault detection. For this reason, commercial MCSA equipment for broken bar detection have the default alarm level set at -45~-35dB to avoid false positive alarms. It can be clearly seen in Fig. 10 that there is a significant difference in the sensitivity of MCSA for single and double cage induction motors. Due to the very small levels of fbrb in samples 2 and 3, MCSA-based fault detection can be ambiguous for double cage motors, especially for the separate end ring rotors that employ high resistivity outer bar material. Therefore, it is important for the inspection engineer to keep a record of the double cage motors since applying the fault threshold of single cage motors is meaningless. It is also important to be aware of the insensitivity of MCSA-based broken bar detection for double cage induction motors.
(a)
(b)
C. Experimental Results – Off-line Testing The growler test was performed based on the setup and test procedure described in [5] with 60hz pulsating field excitation. The 120hz vibration of a hacksaw blade placed axially on the bar surface was observed for fault detection. The results cannot be quantified or shown visually, but the broken bars were clearly observable for all samples with the growler test. The single phase rotation test was also performed with 1/8 of rated voltage applied between two phases. The measured stator current magnitudes as a function of rotor mechanical angle for samples 1-3 are shown in Fig. 11(a)-(c), respectively, for 0-3 broken bars with the rotor manually rotated in 5 degree intervals. It can be seen that the magnitude of the stator current fluctuates twice per 180 degrees with broken bars (4 pole motor), and the variation in current increases with the number of broken bars. The % change in the stator current magnitude
2498
(c) Fig. 10. Experimental results of on-line MCSA - fbrb measurement at 100%, 67% (or 75%), and 33% (or 50%) rated load for 0-3/44 broken bars from (a) sample 1; (b) sample 2; and (c) sample 3 (outer cage bars broken for samples 2 and 3)
(a) Fig. 12. Experimental results of off-line single phase rotation test - % change in stator current as a function of number of broken bars for samples 1 , 2, and 3 (outer cage bars broken for samples 2 and 3)
with an experimental study on fabricated copper bar double cage motors of common and separate end ring rotor design with 0~3 broken outer cage bars. The tests were also performed on a single cage aluminum die cast rotor with broken bars for comparison of the detectability. It was shown that the sensitivity of MCSA is significantly decreased for broken outer cage bar detection, especially for the separate end ring rotor design with high resistivity outer cage material. This is due to the small outer cage current as most of the rotor current is in the inner cage in steady state. The small outer bar current is diverted to the low impedance inner bar instead of the adjacent outer bars when it is broken, which tends to smoothen the rotor flux asymmetry making fault detection more insensitive. It was also shown that the sensitivity of off-line tests (growler/1φ rotation test) is not influenced significantly as they are performed at motor standstill when most of the current is in the outer cage. Although there is a decrease in sensitivity for the high resistance outer cage separate end ring rotor design, off-line standstill testing provides reliable detection since it is not influenced by operating conditions.
(b)
(c) Fig. 11. Experimental results of off-line single phase rotation test - stator current as a function of rotor mechanical angle for motor with 03/44 broken bars from (a) sample 1; (b) sample 2; and (c) sample 3 (outer cage bars broken for samples 2 and 3)
double cage rotors (sample 2, Fig. 11(b)). However, there is a decrease in the % stator current variation for the brass outer cage separate end ring rotor (sample 3, Fig. 11(c)) by a factor of 2-3 times. This can be attributed to the relatively smaller current in the outer bar due to the high resistance outer cage, as predicted in IV.B. Although the current variation is below the typical threshold of 3% for 1 broken bar, it can be seen in Fig. 11 that the presence of broken rotor bars can be detected if the resolution of the measurements is high enough. Since the single phase rotation test or other off-line tests are standstill tests, they can provide reliable assessment of rotor bars independent of variations in load or operating conditions unlike MCSA. VI.
CONCLUSION
The detectability of broken outer cage bars in double cage induction motors for the most commonly used on-line and offline field test methods was evaluated in this paper. A qualitative finite element analysis was provided, and supported
The results of this study provide important information for practicing motor inspection engineers on the limitations of MCSA on the detectability of broken outer cage bars for double cage motors. In industrial facilities, it is unknown in most cases whether an induction motor is of single or double cage construction unless it is obtained from the manufacturer or recorded when the motor is disassembled. The results of this study suggest that it is important to keep a record of double cage machines, since one should apply different standards on the level and increase in the fbrb component when applying MCSA. MCSA-based detection of double cage motor problems may be difficult depending on the motor design and operating environment, and can produce false negative indications. Since off-line standstill tests provide reliable and sensitive assessment of double cage rotor faults, they should be performed whenever possible. This paper also shows that there is a need for developing new on-line monitoring methods for double cage rotor fault detection. ACKNOWLEDGMENT The authors gratefully acknowledge Chaewoong Lim, Jongkyung Bae, and Shinyoung Park of Hansung Electric
2499
Industrial Company for their support on the design and fabrication of the double cage fabricated copper bar induction motors . REFERENCES [1] [2] [3] [4]
[5] [6]
[7] [8]
[9]
[10]
[11]
[12]
[13]
[14]
[15]
[16]
[17]
[18] [19]
Motors and generators, NEMA standards pub. MG 1-2006, 2006. Motors and generators, NEMA standards pub. MG 10-2001, 2001. H.A. Toliyat, G.B. Kliman, Handbook of electric motors, 2nd edition, Marcel Dekker, 2004. A.H. Bonnett, T. Albers, "Squirrel-cage rotor options for AC induction motors," IEEE Transactions on Industry Applications, vol. 37, no. 4, pp. 1197-1209, July-Aug. 2001. T. Bishop, “Squirrel cage rotor testing,” EASA Convention, 2003. I.D. Lomax, “Assessment of induction motor cage fatigue life,” Proceedings of the International Conference on Electrical Machines and Drives, pp.281-284, 1991. M.G. Say, The performance and design of alternating current machines, Sir Isaac Pitman and Sons LTD., London, 1955. J. Mroz, “Temperature field distribution of a double squirrel-cage motor during startup,” IEE Proceedings – Electric Power Applications, vol. 152, no. 6, pp. 1531-1538, Nov. 2005. I. Kerszenbaum, C.F. Landy, “The existence of inter-bar currents in three phase squirrel cage motors with rotor-bar and end-ring faults,” IEEE Transactions on Power Apparatus and Systems, vol. PAS-103, no. 7, pp. 1854-1862, July 1984. R.F. Walliser, and C.F. Landy, “Assessment of interbar currents in double cage induction motors with broken bars,” IEEE Transactions on Energy Conversion, vol. 9, no. 1, pp.159-164, March 1994. A. Bellini, F. Filippetti, C. Tassoni, & G.A. Capolino, “Advances in Diagnostics Techniques for Induction Machines,” IEEE Trans. on Ind. Elec., vol. 55, no. 12, pp. 4109-4126, Dec. 2008. S. Nandi, H.A. Toliyat, and X. Li, “Condition monitoring and fault diagnosis of electric motors – a review,” IEEE Transactions on Energy Conversion, vol. 20, no. 4, pp. 719-729, Dec. 2005. S. Williamson, and M.A.S. Abdel-Magied, “Steady state analysis of double cage induction motors with rotor cage faults,” IEE Proceedings – Electric Power Applications, vol. 134, no. 4, pp. 199-206, July 1987. C.H. Lee, “A design method for double squirrel-cage induction motors,” AIEE Transactions Part III - Power Apparatus and Systems, vol. 78, no.2, pp. 630-636, Jan. 1953. P.L. Alger, and J.H. Wray, “Double and triple squirrel cages for polyphase induction motors,” AIEE Transactions Part III - Power Apparatus and Systems, vol. 78, no.2, pp. 637-645, Jan. 1953. P. Jacobs, “Defining the equivalent circuit of the double-cage motor,” AIEE Transactions Part III - Power Apparatus and Systems, vol. 78, no.2, pp. 651-657, Jan. 1953. F. Filippetti, G. Franceschini, C. Tassoni, and P. Vas, “AI Tehcniques in Induction Machines Diagnosis Including the Speed Ripple Effect,” IEEE Transactions on Industry Applications, vol. 34, no. 1, pp. 98-108, Jan./Feb. 1998. G.C. Stone, and I. Culbert, “The testing tools,” IEEE Industry Applications Magazine, no. 14, vol. 6, pp. 48-53, Nov./Dec. 2008. EXP 4000 Users Manual, Baker Instrument Company – An SKF Group Company, Fort Collins, CO, Feb.2010.
2500
