real time measurements of flux distribution in the air gaps of the DC ... is driven by DC machines, for example, electric rope shovels and draglines used in open ...
Performance Evaluation of DC Motors for Electric Rope Shovels Based on Air Gap Flux Measurement G. Mirzaeva, R.E. Betz and T.J. Summers School of Electrical Engineering and Computer Science University of Newcastle, Australia, 2308
A BSTRACT
Two industrial DC motors of the same size connected back-to-back, one being the test motor, the other working as a programmable load (these roles can be swapped); • Two sets (one per each motor) of a thyristor converter, digital drive, reactive power compensator and other devices that are normally used with the rope shovel DC motors; • A large number of sensors installed internally and externally to the motors to collect information about the machine performance during the tests; • PC-based data acquisition system to record, process and visualise the experimental data. The rest of the papers is organised as follows. It first explains the major points of the test methodology. It then provides information about the sensor calibration, test results for a laboratory motor, and the full size motor instrumentation. It speculates about the expected flux distributions and how these are related to commutation quality. It then presents experimental results for two industrial DC motors and their physical interpretation, followed by the conclusions. •
DC motors remain an attractive solution for mining applications, such as rope shovels and draglines, due to their robustness, good dynamic performance and vast experience accumulated within the industry. To achieve optimal performance of the DC machines under increasingly demanding operation conditions, it is important to thoroughly understand their performance limitations under conditions dictated by mining environment. For this purpose, a unique test facility has been developed in Hunter Valley (NSW, Australia) to be able to complete extensive experiments, under controlled conditions, on the full size DC motors used by Australian mining industry. This paper will explain about the structure of the facility, the test methodology and the instrumentation. The main focus of the experimental program is to obtain real time measurements of flux distribution in the air gaps of the DC machines, together with other important parameters. Physically measured flux distribution under different conditions will allow to accurately determine a variety of important characteristics, including electromagnetic conditions of commutation. To the best of the authors’ knowledge, only theoretical or indirectly obtained results on flux distribution in the air gap are currently available. The paper includes theoretical predictions as well as experimentally measured flux distributions for two commonly used types of DC motors, physical interpretation of these results and their relation to the motors’ performance.
II. T EST METHODOLOGY The dynamometer arrangement consists of two DC machines, one being considered as the load machine and the other as the test machine. Because of the symmetry of the system, these functions are interchangeable between the machines. Both machines are similarly instrumented with a large number of sensors installed both internally and externally, such as: • main pole and interpole flux sensors; • temperature sensors of the machine core; • armature and field voltage and current; • high frequency component of the armature current (that contains signatures of sparking events); • torque and speed measured on the shaft connecting the two machines; • ambient conditions of operation. To make a better use of the sampling bandwidths, a separate 96-channel data acquision card Adlink-2208e and a separate PC are used for each machine, the measured data being synchronised at the data processing level. The test methodology is based on direct measurement, under different loading conditions, of the air gap flux distribution in space and in time, and relating these to the simultaneously
I. I NTRODUCTION Up to the present, a big share of electrical mining equipment is driven by DC machines, for example, electric rope shovels and draglines used in open cut mining. Though the market share of DC machines will continue to decrease, it is reasonable to expect that they will be still in service for the next few decades. Their performance requirements will continue to grow due to increasingly difficult digging conditions and production pressure. A project has been set up between the University of Newcastle and industry partners that includes development of a full size DC motor testing facility and conducting of an extensive experimental program with the DC motors used in electric rope shovels. The facility allows for both static and dynamic testing, and has the following features: 1
B(I) Characteristic
Step test measurements 4
1.20
3.5
1.00
y = 0.0589x + 0.008
Secondary emf
2
Secondary current
1.5 1 0.5 0 -0.1 -0.5
0.80
0.60
0.40
Experimental data 0.20
Linear approximation
0.00
-0.05
0
0.05
0.1
0.15
0.2
0.25
-0.20 -2
0
2
4
6
8
10
12
14
16
18
Primary current (A)
Time (sec)
(a) Calibration rig set up
Flux Density (Tesla)
Current (A); Voltage (V)
3 2.5
(b) Measured voltage and current
(c) Estimated flux density characteristic
Fig. 1: Step test of the calibration rig.
recorded parameters of the operating conditions. Such conditions are varied in a controlled manner and are determined by the magnitude and dynamics of the load (i.e. the armature current), rotation speed and ambient conditions. In the first series of experiments we only vary one of the major operating parameters. For example, when running the motor at a constant speed we apply 0%, 25%, 50%, 75%, 100% and 125% of the rated load. Or, when applying a constant load, we command to the motor to run at 0, ±1/3, ±2/3 and full rated speed. After mechanisms of the influence of each specific factor are understood, combinations of two or more varied factors will be applied. More details on the requirements to the test facility, its structure and the experimental methodology can be found in [1].
To ensure that the flux sensors would work correctly inside a machine, a laboratory testing on a small DC machine was carried out first. The test rig was a small dynamometer setup with two DC motors connected back-to-back so that the test machine could be loaded to allow a reasonable armature current to flow. The machine was instrumented with two flux sensors – one on the main flux pole of the machine, and the other one on an interpole. Fig. 2a shows the flux sensors mounted on the main pole and on the interpole faces. Fig. 2b and Fig. 2c show details of the main pole and interpole flux densities recorded during the test. On both plots one can see high frequency flux oscillations due to a slot effect: as an armature slot passes the flux sensor, the air gap increases and the flux density drops. Variation of the main flux density with a period of one revolution seen in both plots were attributed to the rotor eccentricity. From Fig. 2b and Fig. 2c one can also appreciate that the recorded signals were reasonably clean, and that the electrical noise was effectively kept out of the measurement system by shielded cables being used.
III. S ENSOR CALIBRATION AND LABORATORY TESTING The direct flux measurement in the machine air gap is the vital part of the experimental methodology, hence a thorough into the available flux sensors was performed. As a result, the Melexis MLX90251 Programmable Linear Hall Effect Sensor was chosen. It satisfied our requirements on measurement range, accuracy and bandwidth. It is also supplied with the programmed thermal compensation (see [2] for the details) but would require calibration of its main characteristics “with the application”. To get the sensors ready for installation inside the motor we had to construct a calibration rig capable to provide known flux density values in the range from zero to almost 2 Tesla. Fig. 1a shows a picture of the flux sensor calibration test rig. A step test was performed on the calibration rig, which was found to be the most adequate way to estimate its magnetic characteristics (see, for example, [3]). A DC voltage step from zero to 18 volts was applied to the primary winding; the voltage and the current induced in the secondary winding were measured (shown in Fig. 1b). The flux linkage of the secondary winding and the flux density in the air gap were calculated according to Faraday Law, and the flux density characteristic relative to the primary current resulted, as shown in Fig. 1c. Based on the known B(I) characteristic of the calibration rig, the flux sensors were calibrated using a two point method. More details on the test rig and its calibration procedure can be found in [1].
IV. F ULL SIZE MOTOR INSTRUMENTATION As implied by the test methodology, the flux sensors were installed inside two industrial motors of similar (around 2000 HP) ratings. Fig. 3a and Fig. 3b show how the sensors were positioned on the main pole and interpole faces of the two motors respectively. In both cases circumferential as well as axial flux distributions were captured, the circumferential flux distibution being of a particular interest as it relates to the motor magnetic compensation. Hence large number of sensors installed circumferentially on the selected pole set of each motor. Two other pole sets of the same polarity (not shown here) were instrumented in each machine by a reduced number of sensors, so that effects related to equilisation between the poles (see, for example, [4]) could also be observed. The total number of 62 flux sensors per each machine was used, which ensured a good redundancy level given the assembling challenges. V. T HEORY AND EXPECTED RESULTS Theoretical models of the flux distribution in the air gap of DC machines can be tracked back to 1920s-1940s when classical textbooks on DC machines were first published by 2
Main pole flux measurement
Interpole flux measurement
2.1
2.35
2.3
1.9
Sensor output voltage (V)
Sensor output voltage (V)
2
1.8 1.7 1.6 1.5 1.4
2.25
2.2
2.15
1.3 0
(a) Laboratory DC motor with flux sensors
0.01
0.02
0.03
0.04
0.05 0.06 Time (sec)
0.07
0.08
0.09
2.1 0
0.1
(b) Main pole flux density measurement
0.01
0.02
0.03
0.04
0.05 0.06 Time (sec)
0.07
0.08
0.09
0.1
(c) Interpole flux density measurement
Fig. 2: Flux density recordings on a laboratory DC machine – speed 1500rpm. Main pole Armature reaction Interpole M+A+I M+A+I (saturated)
Main pole Armature reaction Comp.winding Interpole M+A+C+I Field strength
Field strength
Field strength
Main pole Armature reaction M+A M+A (saturated)
Distance (circum.)
Distance (circum.)
(a) No compensation for armature reaction
(b) Compensation using interpoles
Distance (circum.)
(c) Compensation comp.windings
using
interpoles
and
Fig. 4: Theoretical flux distribution in the DC motor air gap.
are shown in Fig. 4 for three cases: (a) non-compensated DC machine - in Fig. 4a; (b) DC machine with interpoles - in Fig. 4b; and DC machine with interpoles and compensation windings - in Fig. 4c, respectively. On the plots of Fig. 4 field strength in the air gap is plotted versus distance, from a certain fixed point, along the circuference of the armature. Presented waveforms cover the span of two adjacent pole sets: main south pole (positive field) corresponds to the left half, and main north pole (negative field) - to the right half of each plot. Polarity of the armature reaction field corresponds to a motor rotating in the counterclockwise direction or a generator rotating in the clockwise direction. For a DC machine with no interpoles, the armature reaction waveforms may have a saddle-shape feature (as in Fig. 4a) due to magnetic reluctance drop in an increased air gap between the poles. Armature reaction field when superposed with the main pole field results in two major effects: magnetic saturation and undesirable field in the commutation zone. To mitigate the latter effect, higher power DC machines are designed with interpoles, which fields are to compensate the armature reaction fields in the commutation zone and to provide some extra field of the opposite polarity to support commutation. This can be seen from Fig. 4b, where the field in the mid-point between the main poles is fully compensated. To account for the main field saturation, additional compensation windings are laid in slots on the pole shoe of each main pole. Fields created by the compensation windings compensate for the armature reaction within the span of each main pole, as can be seen from Fig. 4c. Both motors of interest are designed with interpoles and
(a) Flux sensor locations in Motor 1
(b) Flux sensor locations in Motor 2
Fig. 3: Flux sensor locations inside two different DC motors.
such authors as A.E.Clayton (UK, [4]) and M.P.Kostenko (USSR, [5]). Recently, some interesting work has been done on commutation modeling ( [6]) and experimental studies ( [7]). Qualitative plots of flux distributions for DC machines 3
Flux density under the pole set 3 at speed = 227 RPM
2
Flux density under the pole set 3 at Ia = 500A
2
1.5
At At At At At At
1.5
speed speed speed speed speed speed
= = = = = =
-227 RPM 227 RPM -490 RPM 490 RPM -625 RPM 625 RPM
+ Average Bip = 0.060 T
Ia (ref) = 0 A, Ia (act) = 100 A
1
Ia (ref) = -0 A, Ia (act) = 42 A
Flux density (Tesla)
Flux density (Tesla)
Average Bip = -0.066 T
1
Ia (ref) = 200 A, Ia (act) = 318 A 0.5
Ia (ref) = -200 A, Ia (act) = -149 A Ia (ref) = 500 A, Ia (act) = 632 A Ia (ref) = -500 A, Ia (act) = -436 A
0
Ia (ref) = 1000 A, Ia (act) = 1162 A
0.5
0
Ia (ref) = -1000 A, Ia (act) = -917 A Ia (ref) = 1500 A, Ia (act) = 1695 A -0.5
Ia (ref) = -1500 A, Ia (act) = -1389 A -0.5
Ia (ref) = 2000 A, Ia (act) = 1988 A -1 0
Ia (ref) = -2000 A, Ia (act) = -1869 A 50
100
150
200
Circumferential position (mm)
250
300 -1
Fig. 7: Motor1 - Flux density in the air gap 1.5
0
50
100
150
200
Circumferential position (mm)
250
300
Fig. 9: Motor1 - Interpole flux estimation
Flux density in the middle of the pole set 1 at speed = 227 RPM Flux density under the pole set 3 (mid.point) at Ia = 500A
1.5
1 Ia (ref) = 0 A, Ia (act) = 55 A 1
Ia (ref) = -200 A, Ia (act) = 274 A Ia (ref) = 500 A, Ia (act) = -425 A Ia (ref) = -500 A, Ia (act) = 543 A
Flux density (Tesla)
Flux density (Tesla)
Ia (ref) = 200 A, Ia (act) = -137 A 0.5
Ia (ref) = 1000 A, Ia (act) = -910 A Ia (ref) = -1000 A, Ia (act) = 999 A
0
Ia (ref) = 1500 A, Ia (act) = -1391 A Ia (ref) = -1500 A, Ia (act) = 1462 A Ia (ref) = 2000 A, Ia (act) = -1871 A
-0.5
0.5
0
Ia (ref) = -2000 A, Ia (act) = 1944 A Ia (ref) = 2500 A, Ia (act) = -2353 A
At speed = -227 RPM At speed = 227 RPM At speed = -490 RPM At speed = 490 RPM At speed = -625 RPM At speed = 625 RPM Average B+ = 0.066 T
Ia (ref) = -2500 A, Ia (act) = 2419 A -0.5
-1 0
50
100
150
200
Circumferential position (mm)
250
300
ip
-1
Average B-ip = -0.086 T 0
50
100
150
200
Circumferential position (mm)
250
300
Fig. 8: Motor2 - Flux density in the air gap Fig. 10: Motor2 - Interpole flux estimation compensation windings. At the time when the motors were designed, only approximate or indirect methods were available to verify that the magnetic compensation was sufficient. Such tests included (see, for example, [5]): steady state test at different loads; black band test; brush pencil test. Besides limited accuracy of the above methods, they can only be applied under the steady state conditions. The operation experience of the DC motors in the dynamic environment such as mining clearly indicates that the dynamic character of the load significally affects the commutation quality. Our test methodology will allow to thoroughly investigate the dynamic as well as steady state performance of the motors under conditions similar to those in mining operation.
flux and the interpole flax data for both motors, for Motor 1 more than for Motor 2. Note that Fig. 5 and Fig. 6 only show signals from one main pole flux and one interpole flux sensors for each motor, while the experimental results were obtained, in a synchronised manner, for all flux sensors mounted in the air gaps. The other set of experimental data, not included in this paper for the sake of space, was obtained when both motors were loaded with 500A and the commanded speed was changed, in equal time intervals, from zero to +227 rpm, −227 rpm, then +490 rpm, −490 rpm and so on. B. Flux densities in the air gap and discussion
VI. E XPERIMENTAL RESULTS
The flux density distributions under one pole set are shown in Fig. 7 for Motor 1 and in Fig. 8 for Motor 2. The flux density in each point was found as a time average from the corresponding signal for the steady state condition at each of the applied load values. The “sawtooth” shape of the main flux profile for Motor 1 (see Fig. 7) is attributed to a discretisation effect of the compensation winding. The “sawtooth” component of the main pole flux created by the compensation winding follows, in magnitude and direction, the armature current, so does the interpole flux. Flux densities are slightly higher at the edges of the main pole, possibly because the main pole edge provides a magnetic path with smaller reluctance than the adjacent interpole, therefore attracting additional magnetic lines of the armature reaction. Apart from this small non-linearity, the flux
A. Raw experimental results Experimental results obtained for motors of the two different types are shown in Fig. 5 and Fig. 6. The measured data is shown in the left column, and the processed data is shown in the right column of both figures. The processing of data included filtering out the 300Hz ripple in the armature current, armature voltage and interpole flux, as well as some low pass filtering of the main pole flux signals. The source of the 300Hz ripple is a 6-pulse silicon controlled rectifier (SCR). The oscillations seen in the main pole flux signal are due to the armature slot effect, similar to that observed in the laboratory testing. Given the very small sizes of the air gaps, the above mentioned slot effect is strongly pronounced in the main pole 4
Fig. 5: Motor1 - Application of variable load under constant speed
Fig. 6: Motor2 - Application of variable load under constant speed
5
density under the main pole is flat and symmetrical, which indicates appropriate design of the compensation winding. The arc shape of the main pole flux profile for Motor 2 (see Fig. 8) is most likely related to slightly different curvatures of the inner surface of the main pole and the outer surface of the armature. On its right edge the main flux is slightly irregular, possibly because the main pole links a part of flux of the adjacent interpole. The discretisation effect of the compensation winding is not present at all, as for Motor 2, due to the main pole design constraints, the flux sensors were installed in the middle of each main pole bar, i.e. exactly between the compensation winding conductors (see Fig. 3b). Overall, the main pole flux profile is also symmetrical and even with respect to the zero load line, hence the compensation winding works properly. For both motors the experimentally obtained plots of the main pole flux densities are reasonably close to the ideal theoretical flat shape shown in Fig. 4c. This means that the saturation effect of the armature reaction on the main pole flux density is appropriately compensated for. The flux densities across the interpoles in both Fig. 7 and Fig. 8 also look symmetrical with respect to the zero load line. The flux values under the interpoles are more accurately estimated based on the constant load - variable speed test (for which the raw experimental data is not included in this paper). The results of these estimation, averaged for all test speeds, are shown in Fig. 9 for Motor 1 and in Fig. 10 for Motor 2. Note that the reduced main pole flux densities at higher speeds are due to weakening of the main field but no change is observed in the flux densities under the interpoles as they depend only on the armature current. As earlier explained in chapter V, the flux densities measured under the interpoles are the result of superposition of several fields, therefore we will use the term “resulting interpole flux densities” to distinguish from “interpole flux densities” created solely by the interpoles. As follows from Fig. 9, the resulting interpole flux densities measured for positive and negative direction of rotation of Motor 1 are almost equal in magnitude (0.060 Tesla vs −0.066 Tesla), hence the brush is set almost perfectly to its magnetic neutral position. For Motor 2, as seen from Fig. 10, the resulting interpole flux densities are not equal (0.066 Tesla vs −0.086 Tesla), therefore the brush is slightly shifted from the magnetic neutral axis. When making a comparison to the ideal theoretical plot from Fig. 4c, one can see that the resulting flux densities in the middle of the interpole are non-zero for both motors, particularly at higher loads. They have the correct sign (opposite to that of the armature reaction field) to assist commutation process. The magnitude of the resulting field is a very important for commutation. In the case of undercompensation, commutation is of a retarded nature which leads to higher current densities and, if strongly pronounced, to sparking under the trailing edge of the brush. In the case of overcompensation, commutation is accelerated with similar consequences but for the leading edge of the brush. An adequate magnetic compensation largely depends on
the motor operation cycles, hence further investigation will be performed including experimental cyclic load testing as close as possible to the actual motor operation in the digging application. VII. C ONCLUSIONS This paper has presented and explained: • The structure of the dynamomenter facility for dynamic testing of the industrial DC motors; • Major points of the chosen test methodology; • Some details on the selected sensors and their calibration technique; • Test results obtained on the small laboratory DC motor; • Details of full size motor instrumentation; • Theoretical basis and the expected waveforms of the air gap field to be measured; • The experimental data obtained for two industrial motors; • The plots of experimentally obtained flux densities in the air gap for the two motors and their comparison to the theoretically expected plot; • Discussion of the magnetic compensation of the two motors and possible consequences for their commutation performance. The major contribution of this paper is seen in that it presented experimentally obtained waveforms of the air gap fields in two big industrial DC machines and compared these waveforms to those known from the literature. The authors are not aware of other published results of such experimental measurements. Their view is that such a direct measurement has only become possible with recent advances in sensoring and data acquisition technologies. R EFERENCES [1] G. Mirzaeva, R. Betz, T. Summers, and I. Marxsen, “Development of a unique dc motor test facility in the hunter valley,” in Proc. of the Australian Mining Technology Conference: Smart Technologies for Sustaining the Minerals Boom. 16-18 September 2008, Twin Waters, QLD, Australia, 2008, pp. 95–106. [2] MLX90251 Programmable Linear Hall Effect Sensor Data Sheet. [3] A. Mariscotti, “Evaluation and testing of off-the-shelf hall sensors for compliant magnetic field measurement,” Instrumentation and Measurement Technology Conference, 2006. IMTC 2006. Proceedings of the IEEE, pp. 72–75, 24-27 April 2006. [4] A. E. Clayton, The performance and design of direct current machines, 3rd ed. London: Sir Isaac Pitman and Sons Ltd, 1969. [5] M. Kostenko and L. Piotrovsky, Electrical Machines. Part1. Peace Publishers, 1961, translated from Russian by A.E.Tchernukhin, El.Eng. [6] A. Vauquelin, J.-P. Vilain, S. Vivier, N. Labbe, and B. Dupeux, “A new modelling of dc machines taking into account commutation effects,” in 18th International Conference on Electrical Machines, 2008. ICEM 2008 Conference Record of the, 6-9 September 2008. [7] Z. Qingliang, T. Ueno, and N. Morita, “Basic studies for accurate commutation analysis which enables commutation spark energy estimation,” in 18th International Conference on Electrical Machines and Systems, 2008. ICEMS 2008 Conference Record of the, 17-20 October 2008.
6
