high-field self-starting permanent-magnet synchronous motor - 04643350 - 1981
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High-field self-starting perm anent-magnet
synchronous motor
K.J. Binns, B.Sc, D.Sc, C.Eng., A.F.I.M .A., M.I.E.E., and M.A. Jabbar, B.Sc.(Eng.), Ph.D.
Indexing terms: Magnetic
ields,
Permanent
magnets,
Synchronous motors
Abstract:
A
number
of
configurations
of
self-starting permanent-magnet synchronous motor have been
developed in recent years. This paper describes a form of motor having a high efficiency/power-factor pro-
duct. Use can be made of ferrite, Alnico or rare-earth magnets. The machine is particularly suitable for
variable-speed drives involving the combination of an inverter and one or more synchronous motors, where
the speed of rotation needs to be accurately determed. The characteristics of the m achine are discussed an d
:
results indicating the performance are presented.
1 Introduction
Over recent years, number of forms of self-starting permanent
magnet machine have been developed [ 1 , 2 ] . They mainly use
ferrite or rare-earth magnets and they start by induction
motor action using permanent magnets to provide the syn-
chronous torque. Some use a com bination of reluctance and
permanent-magnet/reluctance motors give good pull-in charac-
teristics with low starting current
but do not
give
the
highest
values of efficiency/power-factor pro duc t at synchronism. The
teristics with low starting current but do not give the highest
values of efficiency-power factor produ ct at synchronism. The
optimum combination depends on the number of motors
supplied from each inverter.
If
a single motor is used
for
each
inverter, the starting characteristics dominate the maximum
kVA demand. If many motors are used for one inverter and
one or two are started while the remainder are synchronised,
the dominant factor tends to be the efficiency/power-factor
product
of
the synchronised machines.
If
the starting current
of each machine is excessive, some inverter manufacturers are
prepared to start motors on a ramp function of frequency
against time. This must surely involve extra cost. For some
applications,
for
example the processing of man-made fibres, in-
verters control the speed
of
a number
of
self-starting synchron-
ous motors to a very high degree of precision. The inverter
cost tends to be considerably higher than that of the motors.
There is also an energy saving in using motors having a high
efficiency.
This paper describes
a
particular family
of
motors which
can make use of any of the well known permanent-magnet
materials, including rare-earth magnets. The choice depends
mainly on the initial cost of motor and inverter for a given
duty and to some extent on the running cost. If a rare-earth
samarium-cobalt magnet
is
used, power factors
in
excess
of
90
are achievable for 3-phase mach ines.
The demagnetisation characteristics
of
some permanent-
magnet materials are shown
in
Table 1.
In
making use
of
par-
Table : Properties of some commonly available permanent magnet
materials approxim ate values)
Material and type
H
c
BH),
High remanence Alnico
High coercivity Alnico
High remanence ferrite
High coercivity ferrite
MnAIC
Polymer-bonded rare earth
Sm-Co
5
R
2
Co
17
T
1.32
0.88
0.38
0.34
0.58
0.55
0.88
1.10
kA/m
56
120
135
23 0
1 90
4 0 0
6 4 0
537
kJ/r
4 8
4 0
26
22
4 8
5 5
150
2 40
Paper 1293B, first received 7th November 1980 and in revised form
24th February 1981
The authors are with the Department of Electrical Engineering, The
University, Southampton, Hants. SO9 5NH, England
ticular material, one has to balance the performance achieved
against the total cost of the m achine and its inverter supply,
if necessary. There is no overall optimum. An expensive
magnet
may
produce
a
high overall performance
but its use
must be economical for the particular application. Com-
puter programs have been developed which enable the assess-
ment of the relative cost of different designs for a particular
duty. In the long term, the price of high-energy magnets may
reduce
in
comparison with
the
cost
of the
remainder
of
the
machine and its control. The results presented in this paper
should serve to indicate what may be achieved with the
particular design to be discussed.
For any design, demagnetisation must not occur during
normal runnings or fault conditions. The cost of remagnetis-
ation, including
the
loss
of
use
of
the drive system
is
almost
always prohibitive.
2 Machine configuration
A sketch of the rotor lamination is shown (not to scale) in
Fig. 1 for a 4-pole design. The flux passing from each r otor
pole emerges from
the
sides
of two
magnets which have
nonradial axes; some flux
can
also pass under
the
magnets,
X / _
Fig. 1 Schematic diagram of geometry of typical rotor configuration
1^. direction of magnetisation
a Magnet
b Rotor iron
c Cage bars
d Conducting material
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giving an additional path for induction motor action at starting.
Some of the flux from the magnets can pass through this thin
magnetic bridge,
but
this bridge
is
normally
in
saturation.
The
pole-arc to pole-pitch ratio is considerably smaller than has
been used in some designs. The width of the iron bridge below
the magnet has to be carefully chosen; it also serves to give
increased mechanical strength to the rotor, thereby raising
the possible maximum running speed. A cage winding is in-
corporated
in
slots
in the
rotor poles
and the
region
to the
side of the pole also containing conducting material con-
nected
to the end
rings.
3 Flux distribution
in the
rotor
In order to achieve a good synchronous performance, it is
essential to be able to design on the basis of computation of
the flux distribution, at least under steady-state conditions,
for the range of possible rotor configurations. The field com-
putation under steady-state conditions can be achieved by
discretising only a pole pitch of the machine as shown by
broken lines in Fig. 1. Load conditions can be simulated by
invoking symmetry and periodicity conditions. In this par-
ticular geometry, the effective pole centre lies at an angle of
64 degrees from the line OX. Fig. 2 shows a flux distribution
for strontium-ferrite magnets
and Fig. 3
gives
a
flux distri-
bution for samarium-cobalt magnets. The flux passing be-
tween adjacent lines is the same in the two field maps. A
computer program prepared
at
Southampton
was
used
in
Fig.
2
Flux map of permanent-magnet machine under load, using
ferrite magnets (Feroba III
the computation of these flux distributions. Results shown
in Figs. 2 and 3 are at conditions when the machine is under
load. Because of the difference of the magnet excitation, the
airgap flux axis has moved from the pole axis by 30 degrees
in Fig. 2 and 19 degrees in Fig. 3.
From the results of such a program, it is possible to deter-
mine the output torque for any given set of dimensional
paramaters, the demagnetisation properties of the magnet and
th e B/H characteristics of the rotor iron. The airgap flux
distribution can be computed by a method described earlier
[3 ,
4] and is shown in Fig. 4 for three magnetic materials —
ferrite, Alnico and rare earth
—
for the no-load co ndition. The
operating characteristics are those for a high value of load
angle.
40 60 80 100 120
airgap positions degrees
140 160 180
Fig. 4 Airgap flux distribution at no-load condition
a Sm-Co
5
b Ferrite
c Alnico
4 Exp erimen tal results
4.1 erformance
at 5
Hz
Tests were carried out on a development prototype using a
standard induction motor stator supplied initially at 50 Hz.
The synchronous performance is indicated in Table 2 for a
nominal rated voltage. The rotor length is 8.2 cm and its
diameter is 9.3 cm. The magnets used were of samarium-
cobalt. It is seen that, at low loads, the power factor is of the
order of 0.96, falling to 0.91 at a load corresponding to 70
of the pull-out torque. The efficiency is of the order of 85
over a wide range of loading and the pull-out torque is 2.6 kW.
This machine is not an optimal design, but represents a par-
ticular practical design. The balance of output parameters can
be varied, depending on the requirements of particular appli-
cations. The choice of supply voltage level (or stator winding
turns) is clearly an important one for any particular application.
Table 3 shows the performance parameters for a voltage of
0.96 p.u.
and
Table
4
shows
the
same
for a
voltage
of 1.04
p.u.
Table 2: Synchronous performance at nominal rated volts at 5 Hz
Fig.
3 Flux map of per manent-m agnet mach ine under load using
rare earth magnets (Sm-CoJ
Current
A
0.84
1.10
1.48
1.82
2.39
2.89
3.36
3.91
4.72
6.00
7.84
10.20
/ „
=
0.6
Output
W
9 9
2 48
41 3
5 5 5
7 59
9 4 0
1121
1278
1514
1841
2206
2521
Power
factor
0.690
0.880
0.940
0.956
0.960
0.964
0.975
0.950
0.940
0.910
0.880
0.871
Efficiency
0.426
0.644
0.744
0.793
0.834
0.847
0.859
0.864
0.857
0.848
0.803
0.712
Pull-out =
Power facto r X
efficiency
0.294
0.567
0.699
0.758
0.801
0.817
0.838
0.821
0.806
0.772
0.707
0.620
2595 W
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Results show significant changes in efficiency, power factor,
no-load current and pull-out torque for relatively small changes
in supply voltage. It is not possible, of course, to control the
excitation level of the permanent magnets, and so the applied
voltage has a dominant effect on the operating power factor
as well as on the pull-out torque.
Table 3 : Synchronous performance at 0.96 p.u. rated volts at 5 0 Hz
Current
Output
Power
factor
Efficiency
Power factor X
efficiency
A
1.27
1.53
2.00
2.55
3.12
3.88
4.56
5.60
6.85
8.96
W
102
278
489
728
943
1212
1420
1710
1986
2316
0.512
0.712
0.820
0.911
0.934
0.954
0.952
0.951
0.935
0.887
0.433
0.670
0.782
0.823
0.850
0.859
0.858
0.843
0.814
0.765
/„ = 1.07 A
0.222
0.477
0.641
0.750
0.794
0.819
0.817
0.802
0.761
0.679
Pullout = 2500
W
Table 4: Synchronous performance at 1.04 p.u. rated volts at 5 0 Hz
Current
Output
Power
factor
Efficiency Power factor X
efficiency
A
1.01
1.34
1.93
2.56
3.28
4.10
4.76
6.00
7.82
9.87
/„ = 1.07 A
W
114
249
558
806
1058
1333
1533
1871
2245
2552
0.609
0.826
0.885
0.907
0.902
0.907
0.900
0.894
0.861
0.822
0.440
0.707
0.786
0.836
0.860
0.863
0.861
0.839
0.802
0.739
Pu
0.272
0.584
0.696
0.758
0.776
0.783
0.775
0.750
0.691
0.607
= 2784W
4.2 Operation over
a
range of frequencies
For variable-speed applications, it is important that a machine
performs well over the required range of frequencies. Tests
have been carried out over a range of frequencies, showing
good results and stable operation. Tables 5 to 10 show the
results of tests at different frequencies and m ain flux levels.
The results of test at 3 0 Hz for per-unit vo ltage of 0.94, 1.00
and 1.06 are shown in Tables 5 to 7. Comparative results at
75 Hz are shown in Tables 8 to
10.
It
is
apparent that the power
factor and efficiency of the machine remain high over this
frequency range. At 75 Hz, the supp ly voltage is increased
prop ortio nally . However, at 30 Hz the voltage is increased by
8.7 over such a value. This is necessary to keep the flux
level constant. Extensive experimentation has been done on
hybrid machines and an experimental relationship established
[4].
/„ is the noload current shown in Tables 2 to 8. In some
of the Tables it can be seen that /„ is larger than the first load
current. This happens when the motor is overexcited and
operating at a leading power factor.
Table 11 summarises the results at 70 of pull-out to rque
for three flux levels in each case. It is clear that the power
factor is critically dependent on the supply voltage. The
efficiency increases w ith frequency over this range, but has an
optimum related to flux level. Depending on the choice of
lamination, a supply frequency will be reached at which the
efficiency starts to fall owing to sta tor core loss.
5 Effects of cage win ding
on output
The design of the cage winding is important, since both the
synchronous and the asynchronous performance are affected
by the choice of the cage winding.
Table 5: Synchronous performance at 0.94 p .u. rated volts at 3 0 Hz
Current
Output
Power
factor
Efficiency Power factor X
efficiency
A
1.22
2.05
2.88
3.83
4.77
6.03
7.40
W
204
393
581
779
958
1128
1326
0.896
0.935
0.967
0.980
0.986
0.954
0.967
0.770
0.844
0.860
0.856
0.840
0.810
0.764
/„ = 0.59 A
0.690
0.789
0.832
0.839
0.828
0.773
0.739
Pull-out = 1525W
Table 6: Synchronous performance at nominal rated volts at 30 Hz
Current
Output
Power
factor
Efficiency Power factor X
efficiency
A
1.46
2.02
2.75
3.57
4.49
5.57
6.90
8.70
W
212
401
582
774
963
1157
1346
1534
0.778
0.934
0.966
0.965
0.969
0.967
0.948
0.938
0.717
0.819
0.847
0.865
0.853
0.826
0.791
0.724
/„ = 1.15 A
0.558
0.765
0.818
0.835
0.827
0.799
0.750
0.679
Pull-out = 1671 W
Table 7: Synchronous performance at 1.06 p.u. volts at 3 0 Hz
Current Output Power
factor
Efficiency Power factor X
efficiency
A
2.46
2.61
3.07
3.70
4.50
5.48
6.65
8.06
10.00
W
218
407
596
784
973
1162
1346
1534
1723
0.484
0.719
0.840
0.902
0.918
0.928
0.922
0.916
0.908
0.660
0.782
0.833
0.848
0.850
0.824
0.791
0.750
0.685
/ „ = 2.62 A
0.319
0.562
0.700
0.765
0.780
0.765
0.729
0.687
0.622
Pull-out = 1812
W
Table 8: Synchronous performance at 0.94
p.u.
rated volts at 75 Hz
Current
Output Power
factor
Efficiency Power factor X
efficiency
A
1.68
2.42
3.20
4.08
5.02
6.10
W
502
997
1469
1940
2436
2907
0.797
0.915
0.954
0.971
0.974
0.961
0.656
0.788
0.842
0.857
0.871
0.868
0.523
0.721
0.803
0.832
0.848
0.834
/„ - 0.98
A
Pull-out = 3 437 W
Table 9: Synchronous performance at nominal rated volts at 75 Hz
Current Output Power
factor
Efficiency Power factor X
efficiency
A
1.42
2.25
3.07
3.92
4.84
5.95
7.15
/„
=
0.66
A
W
800
1295
1795
2235
2790
3400
4030
0.929
0.949
0.964
0.941
0.951
0.943
0.930
0.642
0.770
0.812
0.850
0.865
0.855
0.833
Pull out =
0.596
0.731
0.783
0.800
0.823
0.806
0.775
3804 W
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Table 10: Synchronous performance at 1.06 p.u. rated volts at 75 Hz
Current
Output Power
factor
Efficiency Power factor X
efficiency
A
1.87
2.45
3.15
4.00
4.84
5.87
7.00
8.50
W
5 0 2
9 9 7
1445
1964
2412
2931
3379
3874
0.693
0.838
0.884
0.905
0.906
0.908
0.898
0.871
0.604
0.758
0.810
0.850
0.860
0.856
0.840
0.816
0.419
0.635
0.716
0.769
0.779
0.777
0.754
0.711
/„ = 1.50
Table 11
Frequency
Hz
30
50
75
Pul lout = 4323W
: Synchronous performance at 70 pull-out power over a
range of frequencies
Voltage
p.u.
0.94
1.00
1.06
0.96
1.00
1.04
0.94
1.00
1.06
Current Power
A factor
5.59
5.56
6.16
5.87
5.98
6.477
5.00
5.40
6.11
0.965
0.960
0.925
0.948
0.910
0.885
0.974
0.955
0.906
Efficiency
0.821
0.824
0.805
0.827
0.840
0.821
0.871
0.860
0.853
Power factor X
efficiency
0.792
0.790
0.745
0.784
0.764
0.727
0.848
0.821
0.772
Table 12 : Effects o f changes in the size of cage bars on pull-out torque
a function of pole arc and magnet material*
Pole arc
Size
of
cage bars
- 1 0
Standard
design
+ 10
Magnet
Type
- 1 0
Fe- l l l
5.95
5.82
5.75
Standard design
RE Fe -ll l RE
15.17 6.30 15.95
14.45 6.
14.15 6.
.23 15.70
.15 15.44
i + 10
Fe- l l l RE
6.15 16.40
6.10 16.18
6.03 15.96
*The standard design represents a reasonable balance between starting
and running conditions
Being a high-field machine, the accelerating torque due to
the cage is opposed by the generation of a speed-dependent
EMF due to the magnets [5]. The voltage level needs to be
sufficiently high to pull the machine into synchronism in
many applications, but if the voltage exceeds the back EMF
by a significant amount, the power factor starts to fall. The
output from a given frame size increases rapidly with supply
voltage, as does the maximum demand from the inverter.
The design of the cage winding on the rotor presents some
interesting problems. When rare-earth magnets are used, the
flux density in the rotor poles is such that the iron operates
in the nonlinear regime. As the area devoted to the cage in-
creases, the rotor flux is reduced and, for a design having an
effective cage winding, a balance between good ind uctio n
motor action and high synchronous torque has to be achieved.
Table 12 shows the variation in maximum torque at synchron-
ism as the area of the rotor slots is varied for both ferrite and
rare-earth magnets.
The torque is clearly dependent of the pole arc since the
level of saturation depends on the iron area for a given magnet
flux. It is apparent that, as the pole arc decreases, the cage de-
sign becomes more critical.
It is also important for good asynchronous torque that the
cage bars are positioned at an appropriate depth below the sur-
face.
In this particular design, the depth at which the con-
ducting bars are buried also affects the available space for the
synchronous flux and so changes the saturation level in that
part of the rotor.
6 Conclusions
The configuration of a high-field perm anent magnet synchro n-
ous motor is discussed. Performance characteristics for a
particular machine are presented and an assessment of the
problems encountered in designing such machines is given. The
balance between synchronous and asynchronous performance
characteristics is sensitive to fairly small changes in design
parameters. A computer program has been developed for the
computation of the synchronous performance. The efficiency/
power-factor product can be high for such machines compared
with that achieved using an induction motor of the same frame
size.
7 Acknowledgments
The authors wish to thank the UK Science Research Council
for financial support for this work, and also to acknowledge
the fruitful collabo ration with W alter Jones and Co. in the
development of permanent-magnet machines.
8 References
1 BINNS, K.J., BA RNARD, W .R., and JABBAR, M.A.: 'Hybrid
permanent-magnet synchronous motors', Proc. IEE, 1978, 125,
(3), pp. 203-2 08
2 SIEMENS, A.G.: 'An electric machine having permanent m agnets
mounted in the rotor between its pole segments', British Patent
1177247
3 BINNS, K.J., JABBA R, M.A., and BARNARD, W.R.: 'Comp utation
of the magnetic field of permanent magnets in iron cores', ibid.,
1975,
122, (12), pp. 1 377-138 1
4 JABBAR, M.A.: 'Analysis of the performance of a perman ent-
magnet a.c. machine'. Ph.D thesis, Southampton University, 1977,
p.
268
5 HONSINGER, V.B.: 'Permanent-magnet machines: asynchro nous
operation', IEEE Trans., 1980,
PAS-99,
pp. 1 503-1509
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