IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 46, NO. 4, JULY/AUGUST 2010
1341
A Low-Intrusion Load and Efficiency Evaluation Method for In-Service Motors Using Vibration Tests With an Accelerometer He Zhang, Pericle Zanchetta, Member, IEEE, Keith J. Bradley, Associate Member, IEEE, and Chris Gerada, Member, IEEE
Abstract—This paper presents a novel low-intrusion load and efficiency evaluation method for in-service induction motors based on vibration measurements. This method enhances the traditional vibration analysis by providing motor load and efficiency information in addition to the mechanical health information. The application is in multimotor plants, where individual motor monitoring is too expensive to implement yet where motor operating conditions need to be known to work at improved plant efficiency. The vibration signature of the machine, measured by an accelerometer and processed by fast Fourier transform (FFT) is used to extract frequencies defining shaft speed and the supply frequency. This data, in conjunction with the basic motor performance data, enables the determination of the actual load and indirectly, the efficiency which the motor is operating at. For motors supplied form a variable-speed drive (VSD), the loss segregation method is used to yield the motor losses indirectly and thus, the efficiency. Index Terms—Energy and efficiency measurement, induction machines, industrial power system, vibration.
I. I NTRODUCTION
E
LECTRICAL motor-driven systems represent more than 70% of the industrial-related electrical power consumption [1] and [2]. The majority of the electrical motors in the field are induction motors both for fixed- and variable-speed applications. On average, in-service induction motors operate at no more than 60% of their rated load partly due to the variable loading requirements and partly due to the oversized installations [3]. Although most motors are provided with basic manufacturer’s data such as rated efficiency, the nominal efficiency alone is not sufficient to realistically estimate the energy con-
Manuscript received April 21, 2009; revised July 1, 2009 and November 2, 2009; accepted November 10, 2009. Date of publication May 6, 2010; date of current version July 21, 2010. Paper 2009-EMC-082.R2, presented at the 2008 Industry Applications Society Annual Meeting, Edmonton, AB, Canada, October 5–9, and approved for publication in the IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS by the Electric Machines Committee of the IEEE Industry Applications Society. H. Zhang is with the University of Nottingham, Nottingham, NG7 2RD, U.K., and also with WRc plc, Swindon, SN5 8YF, U.K. (e-mail: eexhz2@ nottingham.ac.uk). P. Zanchetta, K. J. Bradley, and C. Gerada are with the Power Electronics, Machines, and Control Group, University of Nottingham, Nottingham, NG7 2RD, U.K. (e-mail:
[email protected]; Keith.Bradley@ nottingham.ac.uk;
[email protected]). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TIA.2010.2049550
sumption in field conditions due to the variable supply frequency and variable load of the motor. This is particularly important for multimotor systems, where individual motor monitoring is too expensive and thus, motor efficiency cannot be determined. It is sometimes necessary to assess each individual motor’s energy consumption in real time to be able to actively manage the motors in the plant and improve the overall efficiency. Monitoring the power consumed by each individual motor in a cost-effective way makes it possible to understand the motors’ efficiency conditions and enables system efficiency improvement. Energy saving can be achievable, for example, by actively shedding motors, sharing a common load if they are operated at a nonoptimal operating point which can be more efficiently delivered with a combination of fewer motors operating closer to their rated conditions. This paper introduces a relatively low-cost, low-intrusion load efficiency evaluation which is applicable to in-service motors based on the vibration measurements. The traditional vibration analysis is still the dominating best preventive maintenance practice used in most industries. Therefore, this method provides a great added value to the traditional vibration analysis. Traditional efficiency-determining methods defined in IEEE Standard 112 and IEC 60034-2(2007) cannot be used for inservice motors due to the need not to interrupt the industrial processes. The Standard method to determine efficiency is to measure the input and output power at several load points with a power meter and dynamometer or torque transducer, such as IEEE 112-2004 Method B [4], which is quite accurate. This requires a variable voltage and load for the no-load and the load test, respectively. An alternative method requires the following: 1) a no-load test; 2) a load test; 3) a rotor removed; and 4) a reverse rotation test, such as IEEE 112-2004 Method E [4]. None of these tests are suitable for field evaluation due to their intrusive nature. There are many methods relevant to field efficiency evaluation as discussed in [5]–[9]. Most of them suffer from at least one of the following problems: 1) high intrusiveness; 2) poor accuracy; or 3) cost effectiveness. As reported in [10], the nameplate-based methods, the basic slip method and the current measurement method have relatively low intrusiveness but lack accuracy; Equivalent Circuit methods, loss segregation methods, and torque measurement methods can achieve higher accuracy but at the cost of high intrusiveness.
0093-9994/$26.00 © 2010 IEEE
1342
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 46, NO. 4, JULY/AUGUST 2010
Fig. 1. Torque–speed curves for variable-frequency supply. Fig. 2.
Estimating the output power based on the slip frequency is one simple method based on the torque–speed curve for an induction motor as shown in Fig. 1, which shows traditional characteristics for an induction motor controlled using a constant voltage/frequency (V /f controller) where the supply frequency to the motor is modified to achieve changes in speed while the voltage magnitude is controlled in direct proportion to the applied frequency to assure the motor operates at rated flux. Most controllers generally apply a voltage boost at lower speeds to compensate for the significant stator voltage drop. Within the normal motor operation range, the torque–speed curve is almost linear. The approximate output power of the motor can be estimated simply by Pout = 2π
s n fs × × × Trated 60 srated frated
(1)
where Pout is the output power, n and s are the actual speed and slip, srated and Trated are the rated slip and torque, respectively, fs and frated are the supply frequency and rated frequency, respectively. The slip method requires the measurements of speed and supply frequency to yield the actual slip. Direct measurement of the rotor speed needs a shaft-mounted speed encoder or optical speed sensor. But in many cases, it is not possible to access the motor shaft in existing installations. Therefore, it is difficult to attach or install the transducers. Alternative methods have been developed to detect the induction motor shaft speed without speed sensor. For instance, a noninvasive speed estimation method based on current harmonics has been developed by [3] and [11]. This speed detection algorithm is complicated and the accuracy needs further improvements. It can achieve an accuracy of within 5 r/min at high speed though it is not reliable with some motor slot combinations. Furthermore, monitoring the current waveform can be an intrusive operation, requiring the presence of an electrician in the field. In order to address the energy consumption in a plant, it is essential that the load and efficiency of induction motors can be estimated quickly with minimum disturbance to the service and installation.
Typical vibration spectrum of a four-pole induction motor.
II. L OW-I NTRUSION L OAD AND E FFICIENCY E STIMATION M ETHOD This paper proposes a novel low-intrusion load and efficiency method (LILEM) for in-service induction motor efficiency estimation based on the vibration measurements using an accelerometer. In conjunction with the manufacturer’s data or nameplate data, the output power can be evaluated by the motor slip method and the equivalent circuit (EC) method. A. Vibration Signature The vibration signature of the motor offers the possibility of determining the rotor speed and supply frequency without interrupting the electrical supply to the motor or the drive system. When induction motors are operated under mains supply, the typical vibration spectrum of a four-pole motor is shown in Fig. 2. In a general, in a rotating system, such as an induction motor, there are numerous sources of vibration. Among them, the largest low-frequency vibration harmonic is normally due to the mechanical imbalance of the rotating parts [12]. As illustrated in Fig. 2, the largest low-frequency harmonics of the vibration spectrum is associated with the rotational frequency. This can be distinguished easily from the low frequency range of the spectrum. The mechanical vibration due to the mechanical imbalance is a once per revolution force, thus, the rotation velocity is 60 × fm r/min [13]. Induction motors vibrate at twice the supply frequency due to their electrical excitation. The air-gap field, which rotates synchronously with the supply excitation, produces a magnetic attraction force between the rotor and stator to pull them together. The attraction force can be expressed as F =k×
B2 A μo
(2)
where k is a constant, B is the alternating air-gap flux density in the motors air gap, and A is the pole area. Thus, the attraction force remains in the same direction irrespective of the direction
ZHANG et al.: A LOW-INTRUSION LOAD AND EFFICIENCY EVALUATION METHOD FOR MOTORS USING VIBRATION TESTS
1343
of the flux. This results in the magnetic attraction force at any point on the stator frame pulsating at twice the supply frequency. The attraction force acts on the stator and induces the motor frame to vibrate, which leads to a significant harmonic peak at twice the supply frequency in the vibration spectrum as shown in Fig. 2. The rotor speed is between the maximum synchronous speed and synchronous speed minus maximum slip. Therefore, there is only a narrow frequency range where to locate the twice supply frequency component for which bounds can be determined by fs_ min = 2 × fm × P/2 fs_ max = 2 × (fm + smax ∗ ffundamental /60) × P/2
(3)
where P is the pole number, fm is the mechanical rotational frequency, fs_ min and fs_ max are twice of the minimum and maximum supply frequency, respectively, smax is the maximum slip at the maximum load, and ffundamental is the fundamental supply frequency, which is very close to fs_ min /2. B. Vibration Measurement The vibration generated by induction motors can be detected by using transducers for displacement, velocity or acceleration measurements. Piezoelectric accelerometers nowadays are universally used for vibration measurements converting vibratory motion into an electrical charge signal. These devices are compact, exhibit more stable and reliable characteristics than any other type of vibration transducers [14] and are therefore used for this application. The vibration measuring system includes the following instrumentation: 1) the piezoelectric transducer; 2) signal conditioning equipment; and 3) a data acquisition and analysis system. The vibration frequencies are in the same basic coordinate directions, i.e., radial [12]. Therefore, a single-axis accelerometer ADXL103CE is utilized in this research. The ADXL103CE single-axis accelerometer used has high precision and a 0.001-g resolution at 60 Hz with a full-scale range of ±1.7 g. The requirement for the motor rotational speed and supply frequency vibration detection does not demand a wide frequency range. Normal two-pole three-phase induction motors run at speeds of up to 3600 r/min at up to 60-Hz supply frequency, producing a fundamental rotational speed vibration at 60 Hz or below. Frame vibration due to the magnetic attraction force is at twice the supply frequency at 120 Hz or below. A frequency range from 2 Hz to 200 Hz is specified, allowing the fundamental frequency and one or two harmonics to be captured. The data sampling frequency is set to 500 Hz. A resolution of 1 r/min in speed and 1/60 Hz in supply frequency is obtained with a 60-s sample. Greater resolutions can be obtained with a longer duration data set. The sensitivity required was determined by experimenting with a range of induction motors and it has been found that the vibration intensity between 0.01 and 1 g is expected. An accelerometer attached to a motor in the laboratory is shown in Fig. 3. The accelerometer has a signal-conditioned voltage output proportional to the acceleration and uses a
Fig. 3. Motor with an accelerometer installed in the laboratory.
single, cheap, monolithic integrated circuit. The motor speed measurement is much less sensitive to the means of mounting the transducer than applications such as machine condition monitoring. This is due to the relatively low frequencies to be measured. It was therefore possible to have a very simple arrangement of mounting the PCB with accelerometer in a plastic housing and holding it against a motor body with an elastic cord. The end of the plastic housing was machined into a blunt screwdriver end to press against the motor casing between the air cooling fins as shown in Fig. 3. C. Estimation of Motor Output Power Based on the previous analysis, it is normally straightforward to locate the mechanical rotational frequency by monitoring the vibration spectrum and finding the most significant peak in the rotational frequency range expected. Defining the expected range is simple and depends on the maximum and minimum supply frequencies and motor pole number. This information helps to filter out unwanted vibration noise. After locating the largest peak harmonic representing the mechanical speed and subsequently the supply frequency, the output power of the induction motor can then be estimated based on the assumption of linear torque–speed relationship near the rated speed fs (P/2) × 60 − fm × 60 60 × fm × × Pr (4) Pout = nr fr (P/2) × 60 − nr where Pout , fm , fs , Pr , fr and nr are the output power, the rotational frequency, the supply frequency, the rated power, the frequency and speed, respectively. D. Estimation of Motor Efficiency 1) Motor Fed From Fixed Frequency Mains Supply: The efficiency of an induction motor fed from a fixed mains supply can be evaluated either from a known performance curve of efficiency versus load or from a typical efficiency versus load curve of similar power rated and efficiency class motors. There are several factors that make the actual input in-service power different from the standard test results, such as the following: 1) temperature rise; 2) frequency variation; 3) supply voltage; and 4) phase imbalance. Therefore, there will be a difference between the actual input power and the data derived by
1344
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 46, NO. 4, JULY/AUGUST 2010
Fig. 4. Efficiency–load curve for three 200-kW motors.
Fig. 5.
applying the efficiency–load curve method. However, for mains-supplied motors, the efficiency is relatively high and small uncertainties in the efficiency curve make little impact on the input power or efficiency evaluation results. So there will be minimum difference between the actual efficiency values and those derived by applying this efficiency curve method. The uncertainties of the induction motor efficiency determination with different international standards are discussed in [15], and the largest uncertainty is 0.67% for IEC standard 34-2. Using the manufacturer’s data on efficiency against load curves to estimate the input power and efficiency from the known load is an effective and simple way to evaluate motor efficiency. If there is no data available from the motor manufacturer, an alternative method is to use typical values of efficiency for similar power rating motors of the same efficiency category, including the following: 1) EFF1; 2) EFF2; and 3) EFF3 (European targets). To address this problem, three 200-kW EFF1 induction motors made by different manufacturers are tested carefully by IEEE 112 standard method B. The efficiency curves of these machines together with an average efficiency curve are presented in Fig. 4. The maximum efficiency discrepancy from the average curve is 0.7%. 2) Motor Fed From Variable Speed Drives: The estimation of the input power for a motor fed from a VSD is more complicated since there is limited data provided for motors fed from such variable frequency supplies (only drives with constant V /f control are considered in this analysis). The input power evaluation can be obtained by estimating the power losses of the induction motor under variable frequency as will be described in this section. The motor losses can be estimated using the loss segregation method indirectly. The five components of the motor loss can be divided into two categories according to their load dependencies. Windage and friction losses Pwf and core loss Pcore are assumed to be load-independent, while stator copper loss Ps_copper , rotor copper loss Pr_copper , and stray load loss PLL are load-dependent. Stator copper loss Ps_copper is statorcurrent-dependent and rotor copper loss Pr_copper and stray load loss PLL are rotor-current-dependent. The total loss Ploss is
In conjunction with the motor input power, the efficiency can be expressed Vs Is cos φ − Is2 Rs + Ir2 Rr + kIr2 + (Pcore + Pwf ) η= Vs Is cos φ (6)
Ploss = Ps_copper + Pr_copper + PLL + Pcore + Pwf = Is2 Rs + Ir2 Rr + kIr2 + (Pcore + Pwf ). (5)
Variations of motor losses with load torque at 50 Hz.
where Vs , Is , and φ denote the supply voltage, the stator current, and the phase angle between Vs and Is . To separate load-independent and load-dependent losses, some assumptions have to be made. The stator current Is is the vector sum of the rotor current Ir and the magnetizing current Im . The maximum efficiency lies in the approximate range of 60%–80% rated load for conversional motor design [2]. For constant V /f controlled loaded motors, Im is assumed to be constant and relatively smaller than Is , resulting in Ir being close to Is in magnitude. With this approximation, the motor efficiency η can be expressed as a function of stator current Is Vs Is cos θ − Is2 (Rs + Rr + k) + (Pcore + Pwf ) . (7) η= Vs Is cos θ Equation (7) is continuous and monotonic; thereby, the maximum efficiency ηmax occurs when ∂η = 0. ∂Is
(8)
Ps_copper + Pr_copper + PLL = Pcore + Pwf .
(9)
Consequently,
As a result, the maximum efficiency occurs when the loaddependent loss components are equal to the load-independent losses. From the efficiency curve, the maximum efficiency point can be located, as shown in Fig. 5. Identifying the maximum efficiency point on the efficiency versus load curve in Fig. 5 allows the two components of losses to be separated and defined by (9), then, load-dependent loss curve and load-independent loss curve can then be derived Ploss_indep = Ploss_dep =
Pmax × (1/ηmax − 1) 2
(10)
where Pmax and ηmax are the output power and efficiency at maximum efficiency point, Ploss_indep is the load-independent loss at 50 Hz, and Ploss_dep is the load-dependent loss at 50 Hz.
ZHANG et al.: A LOW-INTRUSION LOAD AND EFFICIENCY EVALUATION METHOD FOR MOTORS USING VIBRATION TESTS
The load-independent loss includes the following: 1) core loss; 2) windage; and 3) friction losses. The core loss includes: 1) hysteresis; and 2) eddy current loss. Hysteresis loss is normally assumed to vary linearly with frequency and eddy current loss varies approximately with supply frequency squared. Windage loss varies with speed cubed, and therefore frequency cubed, and friction loss is assumed to vary linearly with speed. A suitable approximate assumption is that the constant losses vary with speed squared, and thus, with supply frequency squared. The load-dependent losses including the following: 1) stator loss; 2) rotor loss; and 3) stray loss, remain essentially the same and dependent only on load. Then, the estimation of power loss against load curves for any other operating frequency other than 50 Hz mains supply can be constructed by 2 f + Ploss_dep (11) Ploss = Ploss_indep × 50 where f is supply frequency, Ploss is power loss at frequency f , and Ploss_dep can be determined from (10) and Fig. 5. Many VSD systems are used to control pumps and fans with the aim of improving efficiency by removing flow control valves and dampers. The operation frequency range is normally quite low from about 50%–100% speed due to the cubed power law for speed power variation for such drives. Thus, the assumptions taken in the loss estimation introduce only minor errors. III. I MPROVEMENT T ECHNIQUES The LILEM technique based on the basic slip method can be affected by a number of inaccuracies. The main error relies on the accuracy of the manufacturer’s data, but also on the resolution of the frequency spectrum and the motor operating conditions. Several improvement techniques can be adopted to mitigate these errors, including the following: 1) rated speed correction; 2) temperature and voltage compensation; or 3) the use of a more sophisticated approach based on the motor equivalent circuit. Inevitably, these techniques require more information on the system. A. Rated Speed Correction The nameplate rated speed is allowed a deviation of as much as 20% by standard NEMA MG1 [16] and IEC 34-2-1 [17]; therefore, the rated speed on the nameplate potentially introduces the largest error for the LILEM. This significant influence can be mitigated by actually determining the fullload speed, instead of using the one quoted from the nameplate. This can be done low intrusively by monitoring the power supply from one feeder supplying a group of motors. There will be a step change in total input power when a particular motor is started or when there is a significant load change due to a process instigated by the plant control. Associating the step change of input power with the predicted motor power change enables an individual motor power consumption to be separated from the group of motors supplied from the same
1345
feeder. Assigning the step input power change for a particular motor from a value of Pinput1 to a value of Pinput2 , with a total variation of input power equal to ΔPinput , and using the vibration measurement of synchronous speed nsyn and actual motor speed n under the two load situations, the actual full load speed can be approximately derived. The output power delivered by the induction motor under the two load conditions can be expressed as Pinput1 − Ploss1 = Trated ×
nsyn1 − n1 2 · π · n1 × ∗ nsyn50 − nrated 60
(12)
Pinput2 − Ploss2 = Trated ×
nsyn2 − n2 2 · π · n2 × nsyn50 − n∗rated 60
(13)
where nsyn50 is the motor synchronous speed under 50-Hz frequency mains supply and n∗rated is the corrected rated speed. Subtracting the above equations (12) and (13), rearranging and substituting for ΔP yields the following expression for the corrected rated speed: 2 · π · Trated ∗ nrated = nsyn50 − 60 (n2 (nsyn2 − n2 )) − (n1 (nsyn1 − n1 )) · . (14) ΔP − (Ploss2 − Ploss1 ) The change in power loss is much smaller in comparison to the change in motor output power, particularly for large motors. For high-efficiency motors, this assumption, i.e., Ploss1 − Ploss2 = 0, only introduces a small error, much less than the potential 20% error. For further improvement, the assumption that the efficiency is constant and equal to rated efficiency, i.e., Ploss2 − Ploss1 = Δp(1 − ηrated ) and (14) can be expressed as 2 · π · Trated ∗ nrated = nsyn50 − 60 (n2 (nsyn2 − n2 )) − (n1 (nsyn1 − n1 )) . (15) · ΔP × ηrated The rated speed correction is suitable in the situation in which the motor input power can be measured or a single-motor load change can be identified in a group of motors and the input power fed to this group of motors is somehow monitored. B. Temperature Correction Another major source of error in the prediction of output power is the rotor bar temperature, which effectively changes the rotor resistance resulting in a considerably different torque–speed curve from the manufacturer’s specified one. Motor slip may typically vary from cold to full-load hot conditions by about 30%–50% depending on size and design. The LILEM technique is more accurate close to full load but does not show appreciable error unless the motor is very lightly loaded. It should be noted that all standard efficiency values are quoted at full-load temperature rise and are therefore subject
1346
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 46, NO. 4, JULY/AUGUST 2010
the two different voltages and corresponding full-load speeds are related by ∗ nrated V2 n∗ Prated = r 1 − rated Rr nsyn nsyn ∗∗ 2 ∗∗ nrated V n = m 1 − rated . (17) Rr nsyn nsyn Rearranging it gives the slip ratio ∗∗ Vr2 n∗∗ rated (nsyn − nrated ) . = Vm2 n∗rated (nsyn − n∗rated )
(18)
Equation (4) then can be compensated by Fig. 6. Temperature rise of winding, case, and ambient.
Pout =
to exactly the same error for continuous part load operation. Rotor temperature can be predicted by measuring the ambient temperature and motor case temperature. A simple temperature measurement arrangement can be achieved by attaching two thermocouples at both ends of the accelerometer, one against the motor case and the other at the cable end for the environment temperature, Fig. 3. The rotor resistance and the machine model can then be compensated accordingly. To verify the temperature compensation, three thermocouples are used to monitor the following: 1) winding; 2) motor case; and 3) ambient temperature during the motor heat run, as presented in Fig. 6. There is a relatively constant temperature difference between winding and case temperature except for the starting point. The rotor temperature is assumed to be the same as the winding temperature. Then, based on such a generic temperature rise curve, the rotor temperature rise can be estimated according to the following: 1) load condition; 2) running time; and 3) motor case temperature rise. C. Voltage Compensation Another correction for actual full-load speed can be made by monitoring the voltage on the feeder to a group of machines whose power consumption is to be determined. The supply voltage may be obtained indirectly from a power socket, sharing the same incoming supply for induction motors. When induction motors are running under loaded condition, the voltage drop across the stator resistor can be neglected; therefore, the motor torque Pout will be approximately dependent upon the supply voltage, the rotor resistance Rr and the slips as presented in Pout ≈ 3 ×
V 2 × s × (1 − s) . Rr
(16)
This enables a new full-load speed n∗∗ rated to be derived which is related to n∗rated on the assumption that n∗rated is evaluated for operation at the rated terminal voltage of the motor. Then, the new full-load speed n∗∗ rated is referred to the new terminal voltage Vmeasured and n∗rated is the evaluated full-load speed of motor supplied from rated voltage Vrated . Then, according to (16), when the motor delivers the same full-load torque at two different full-load speeds under different voltage supply,
Vm2 n∗∗ rated sm Pr 2 Vr n∗rated sr
(19)
which can be approximated to Pout =
Vm2 sm · Pr Vr2 sr
(20)
where Vm and Vr are the measured supply voltage and rated voltage, respectively. D. Equivalent Circuit Method The assumptions for the slip calculation method are based on the linearity of the torque–speed curve in normal operation range when fed by variable voltage and frequency. However, the curves are not actually parallel with each other and not linear particularly under high load. The LILEM technique is therefore more accurate when the load is close to full load but less accurate when overloaded. An improved accurate method for high-load points can be introduced by using the equivalent circuit. The values of the EC parameters are critical when determining the performance of the motor. They can be estimated by the following: 1) no-load; 2) part-load; and 3) locked-rotor tests. Doing these tests is not always a viable option. Some large machines are provided with all or some of these parameters from their respective manufacturers. More likely than the parameters themselves, information such as the following: 1) stator resistance; 2) no-load current; and 3) starting current are provided in the manufacturer’s catalogue, from which the parameters can be approximately derived. The parameters can also be estimated from standard test results, such as IEEE 112-B and IEC 60034-2(2007). Alternatively, estimation of the machine parameters can be made by using average values for same power rating machines when there is no more information available. IV. E XPERIMENTAL R ESULTS The proposed LILEM has been verified by testing a 55-kW motor on a laboratory test rig. The motor was mechanically loaded by a dc machine which provides a controlled stable torque to the test motor. A vibration test was carried out after a heat run under rated full load for both fixed frequency and
ZHANG et al.: A LOW-INTRUSION LOAD AND EFFICIENCY EVALUATION METHOD FOR MOTORS USING VIBRATION TESTS
1347
Fig. 8. Speed and frequency comparisons.
Fig. 7.
Signal output and FFT spectrum.
variable frequency conditions. The vibration of the motor is detected by an accelerometer, which is interfaced with a carefully designed signal conditioning system, which includes the following: 1) constant current supply; 2) band pass filter; and 3) an amplifier. The signal output is processed with FFT to extract shaft speed and supply frequency. In these laboratory tests, the voltage compensation and equivalent circuit improvement techniques listed in the previous section were included in the procedure. Fig. 9. Comparison of output power of IEEE 112-B, LILEM, and EC.
A. Motor Fed From Fixed Frequency Mains Supply The output signal of the accelerometer and its FFT spectrum, in the case of a 55-kW motor, are presented in Fig. 7. The spectrum is generated using a 500-Hz sample frequency and a time window of 60 s. The shaft speed frequency and twice supply frequency peaks of the spectrum are clear and uncluttered with noise. The search for the most significant vibration peak in the frequency range, which is below 30 Hz for a four-pole motor, is straightforward. The twice supply frequency range can be fixed by using (3). Thus, searching and locating the most significant vibration peak gives the magnetic attraction force frequency, which is twice the supply frequency. For evaluating the accuracy of the results obtained with LILEM, the motors were tested according to the IEEE Standard 112-B. A shaft-mounted, strain-gauge-type torque transducer which is carefully calibrated is used to measure the motor mechanical torque with accuracy of 0.2%. The speed of the motor is also monitored by an optical speed sensor with an accuracy higher than 1 r/min. Using 60 s will result in 1 r/min and 1/60 Hz for speed and supply frequency. Greater resolutions can be obtained with a longer duration data set. The test results show an excellent agreement between these measurements and the LILEM method as presented in Fig. 8. The resolution of the accelerometer is determined by the duration of the time window. With these measurements, the output power of the motor can be evaluated according to the procedure described in Section II. The results obtained from the LILEM and low voltage supply at
Fig. 10. Efficiency comparison of IEEE 112-B, LILEM, and EC.
390 V are compared with results from the IEEE Standard 112-B test, as shown in Fig. 9 for output power and efficiency. The result shows that evaluating the output power and efficiency under normal operating conditions has quite good agreement with the results from IEEE Standard 112-B. The estimation for load higher than the rated load overestimates output power due to the nonlinearity of the torque–speed curve as shown in Fig. 1. If a more accurate induction motor model based on the equivalent circuit is used instead of the basic slip method, as described in Section III, improved result can be obtained as illustrated in Figs. 9 and 10. Voltage compensation technique is evaluated by testing the motor under 380 V and 420 V for comparison. The Fig. 11
1348
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 46, NO. 4, JULY/AUGUST 2010
Fig. 11. Comparison of output power of IEEE 112-B, low voltage, high voltage, and voltage compensation.
Fig. 12. Comparison of output power of IEEE 112-B, LILEM, and rated speed correction.
Fig. 13. Comparison of output power of IEEE 112-B, LILEM with and without temperature compensation.
presents the results under different voltage supplies and the improvements by using the voltage compensation technique. The relatively good results in terms of accuracy for this case are partially due to the accurate rated speed stated on the nameplate, which presents only a 3.4% deviation from the test results and also due to the rotor temperature remaining almost constant between tests. Therefore, rated speed correction and temperature correction are not included in the comparisons above. As previously stated, all standard efficiency values are quoted at full-load temperature rise and are therefore subject to errors when operated continuously at part load.
Fig. 14.
Comparison of output power from IEEE 112-B and LILEM.
Fig. 15. Efficiency comparison of IEEE 112-B, LILEM, and equivalent circuit.
The rated speed correction can be evaluated by the input power measurements. Assume the rated speed is unavailable in the nameplate, the rated speed can be estimated from (15). The output power derived by corrected rated speed and IEEE-B test results are presented in Fig. 12. In this case, the input power of this specific motor are carefully monitored, therefore, the rated speed can be corrected with high accuracy. There is only less than 1-r/min error compared to the test results. It will be less accurate to employ this method in the field because it is difficult to monitor a single-motor input power variation. The temperature compensation is applied for the winding temperature from 30 ◦ C to 90 ◦ C. The output power results with and without temperature compensation are compared with the actual test result in Fig. 13. The temperature compensation can significantly improve the accuracy at the motor cold conditions. B. Motor Fed From VSD The vibration spectrum for a VSD-fed motor is also clear and uncluttered with other noise, so it is easy to find out the rotational frequency and supply frequency. Together with the presumption that speed–torque curves are almost linear and parallel with each other within the normal operation range as shown in Fig. 1, the output power for the VSD-fed motor can be estimated by (4) or (20). The results in Figs. 14 and 15 show that the output power and efficiency estimation of the motor fed from VSD are less accurate than motors fed from the mains supply. This is mainly because the actual speed–torque curves
ZHANG et al.: A LOW-INTRUSION LOAD AND EFFICIENCY EVALUATION METHOD FOR MOTORS USING VIBRATION TESTS
are not quite parallel with each other in the normal operation range. Improvements for this method of operation are obtained using the equivalent circuit method as shown in Figs. 14 and 15. V. C ONCLUSION This paper has proposed a novel low-intrusion method for in-service induction motor load and efficiency estimation based on vibration test and motor nameplate, with the possibility of using additional sensors and measurements to improve the method accuracy. This method adds a great value to the traditional vibration analysis often carried out for motor mechanical health check. Analyzing the vibration spectrum from an accelerometer placed on the motor frame provides the rotational speed and supply frequency. Basic information, such as: 1) rated speed; 2) current; 3) voltage; 4) efficiency; and 5) rated supply frequency, obtainable from the machine nameplate, are also required for output mechanical power evaluation. In conjunction with the specific information about the motor efficiency curve (or a typical efficiency curve for similar power rating motors of the same efficiency category), the efficiency of in-service motors fed from mains supply or constant V /f VSDs can be accurately evaluated. Further improvements have also been discussed and presented, which includes the following: 1) rated speed correction; 2) temperature compensation; 3) voltage compensation; and 4) equivalent circuit method. The LILEM method has been demonstrated to be suitable for induction motors for both fixed- and variablespeed applications. R EFERENCES [1] A. Boglietti, A. Cavagnino, M. Lazzari, and M. Pastorelli, “International standards for the induction motor efficiency evaluation: A critical analysis of the stray-load loss determination,” IEEE Trans. Ind. Appl., vol. 40, no. 5, pp. 1294–1301, Sep./Oct. 2004. [2] H. Auinger, “Efficiency of electric motors under practical conditions,” Power Eng. J., vol. 15, pp. 163–167, Jun. 2001. [3] B. Lu, T. G. Habetler, and R. G. Harley, “A nonintrusive and in-service motor efficiency estimation method using air-gap torque with considerations of condition monitoring,” IEEE Trans. Ind. Appl., vol. 44, no. 6, pp. 1666–1674, Nov./Dec. 2008. [4] IEEE Standard Test Procedure for Polyphase Induction Motors and Generators, IEEE Std 112-2004, 2004. [5] J. S. Hsu and P. L. Sorenson, “Field assessment of induction motor efficiency through air-gap torque,” IEEE Trans. Energy Convers., vol. 11, no. 3, pp. 489–494, Sep. 1996. [6] Y. El-Ibiray, “An accurate low-cost method for determining electric motors’ efficiency for the purpose of plant energy management,” IEEE Trans. Ind. Appl., vol. 39, no. 4, pp. 1205–1210, Jul./Aug. 2003. [7] T. Phumiphak and C. Chat-uthai, “An economical method for induction motor field efficiency estimation for use in on-site energy audit and management,” in Proc. POWERCON, Singapore, Nov. 21–24, 2004, pp. 1250–1254. [8] E. B. Agamloh, A. K. Wallace, A. von Jouanne, K. J. Anderson, and J. A. Rooks, “Assessment of nonintrusive motor efficiency estimators,” IEEE Trans. Ind. Appl., vol. 41, no. 1, pp. 127–133, Jan./Feb. 2005. [9] J. R. Holmquist, J. A. Rooks, and M. E. Richter, “Practical approach for determining motor efficiency in the field using calculated and measured values,” IEEE Trans. Ind. Appl., vol. 40, no. 1, pp. 242–248, Jan./Feb. 2004. [10] B. Lu, T. G. Habetler, and R. G. Harley, “A survey of efficiency-estimation methods of in-service induction motors,” IEEE Trans. Ind. Appl., vol. 42, no. 4, pp. 924–933, Jul./Aug. 2006. [11] A. Ferrah, K. J. Bradley, and G. M. Asher, “An FFT-based novel approach to noninvasive speed measurement in induction motor drives,” IEEE Trans. Instrum. Meas., vol. 41, no. 6, pp. 797–802, Dec. 1992.
1349
[12] P. L. Alger, Induction Machines: Their Behavior and Uses. New York: Taylor & Francis, 1995. [13] M. L. Adamis, Rotating Machinery Vibration: From Analysis to Troubleshooting. Boca Raton, FL: CRC Press, 2001. [14] W. Boyes, Instrumentation Reference Book, 3rd ed. Oxford, U.K.: Butterworth–Heinemann, 2002. [15] W. Cao, K. J. Bradley, H. Zhang, and I. French, “Experimental uncertainty in estimation of the losses and efficiency of induction motors,” in Conf. Rec. IEEE IAS Annu. Meeting, Oct. 2006, vol. 1, pp. 441–447. [16] Motors and Generators, NEMA Standards Pub. MG1, 1993. [17] Rotating Electrical Machines—Part 1: Rating and Performance, IEC60034-2-1, 2004.
He Zhang received the Ph.D. degree in electrical machines and drives from the University of Nottingham, Nottingham, U.K., in January 2009. He was a Research Follow in the Power Electrionics, Machines and Control Group, University of Nottingham, until September 2009. He is currently a KTP Associate at the University of Nottingham and WRc plc, Swindon, U.K. His recent work has focused on energy efficiency and design of electrical machines.
Pericle Zanchetta (M’00) received the Laurea degree in electronic engineering and the Ph.D. degree in electrical engineering from the Technical University of Bari, Bari, Italy, in 1993 and 1997, respectively. In 1998, he became an Assistant Professor of power electronics and control at the Technical University of Bari. Since 2001, he has been a Lecturer in control of power electronics systems in the Power Electrionics, Machines and Control Research Division, University of Nottingham, Nottingham, U.K. His main research interests are in the field of power quality and harmonics, active power filters, power systems impedance estimation, advanced control of power converters, control design, and system identification using genetic algorithms. He has published over 110 papers in international journals and conference proceedings.
Keith J. Bradley (A’93) received the Ph.D. degree from the University of Sheffield, Sheffield, U.K., in 1974. For a period of one year, he was with YARD, Limited, where he worked on low-vibration induction motors for nuclear submarines. He is currently with the Power Electrionics, Machines and Control Group at the University of Nottingham, Nottingham, U.K. His current research interests include tailoring machine design to optimize variable-speed drive performance and efficiency.
Chris Gerada (M’05) received the B.Eng. and M.Sc. degrees in electrical and electronic engineering from the University of Malta, Msida, in 2000 and 2002, respectively, and the Ph.D. degree from the University of Nottingham, Nottingham, U.K., in 2005. He is currently a Lecturer at the University of Nottingham, where he works on electrical machines and drives. His current research interests include aircraft actuation and numerical modeling and design of electrical machines.
