Design of a synchronous reluctance drive

7/23/2019 Design of a synchronous reluctance drive

1/9

IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 27,

NO.

4,

JULYIAUGUST

1991

74 1

Design

of

a Synchronous Reluctance

Motor Drive

T. J . E. Miller, Senior Member,

ZEEE,

Alan Hutton, Calum

Cossar,

and David A. Staton

Abstract-A segmental-rotor synchronous reluctance motor is

used in a variable-speed drive with current-regulated

PWM

control. T he low-speed torque capability is compared with those

of an induction motor, a switched reluctance motor, and a

brushless dc

PM

hotor of identical size and copper weight. The

results suggest that many

of

the desirable properties of the

switched reluctance motor can be realized with the synchronous

reluctance motor but using standard ac motor and control

components. The torque capability is lower, but

so

is the noise

level.

I. INTRODUCTION

HE POLYPHASE synchronous reluctance motor was

T eveloped particularly in the 1960s as a line-start

(cage-type) synchronous ac motor

[

1

-

4] for applications

where several motors are operated synchronously from a

single voltage-source inverter. In some cases, it has been

replaced by cage-type ac permanent-magnet (PM) motors

that, although more expensive, have better performance and

permit more motors to run from the same inverter 171.

More recently, there has been increasing use of variable-

frequency ac induction motor drives with one motor per

inverter. At first, the six-step inverter was used, usually with

constant voltage/frequency ratio and often without speed

feedback. The development

of

pulse-width-modulated (PWM)

inverters *followedwith slip-control, and today, field-oriented

or vec tor control is the most advanced form of ac drive,

with performance characteristics that match those of the best

dc drives. Although the induction motor is the most common

in ac drives, synchronous PM motors are also used. With one

motor per inverter, there is no need for a rotor cage because

the motor does not have to start across the line.

It is perhaps surprising that there has been so little devel-

opment of the cageless synchronous reluctance motor instead

of the induction motor or PM ac motor for variable-frequency

operation [a-@],13], [14]. One reason is probably its

reputation for poor efficiency and low power factor, but the

removal of the rotor cage and the use of field-oriented control

Paper IPCSD 91-17, approved by the Electric M achines Committee

of

the

IEEE Industry Applications Society for presentation at the 1989 Industry

Applications Society Annual Meeting, San Diego, CA, October 1-5.

Manuscript released for publication February 5 1991. This work was

supported by the

UK

Science and Engineering Research Council, a grant

from the General Electric Compa ny, and the membe r companies of the

Scottish Power Electronics and Electrical Drives SPEE D) Consortium.

T. J . E. Mil ler, C. Cossar, and D. A . Staton are with the Department of

Electronics and Electrical Engineering, University of Glasgow, Glasgow,

Scotland.

A.

J .

Hutton is with Micro Marketing, Motorola Ltd., Glasgow, Scotland.

IEEE

og

Number 9100935.

provide the designer with two new degrees of design freedom

that do not appear to have been fully explored.

The main features of the synchronous reluctance motor are

as follows:

The rotor is potentially less expensive than the PM

rotor. Because it requires no cage winding, it is lighter

and possibly cheaper than an induction-motor rotor.

The torque per ampere is independent of rotor tempera-

ture, unlike that of the PM or induction motors.

The stator and the inverter power circuit are identical to

those of the induction motor or PM synchronous motor

drives.

The control is simpler than that of the field-oriented

induction motor drive, although shaft position feedback

is necessary.

Because

of

scaling effects, the poor efficiency and high

slip of small induction motors prevent

the

extension

of

ac drive technology down to low power levels. The PM

ac motor or brushless dc motor can be used instead, but

PM motors are more expensive. They are temperature

sensitive, susceptible to demagnetization, and may re-

quire additional inverter protection. The synchronous

reluctance motor offers an alternative means of obtaining

the advantages of a synchronous motor but at lower

cost.

The synchronous reluctance motors smoothly rotating field

distinguishes it from the

switched

reluctance motor [9],

[

11 .

It therefore fits in the family of ac drives, which is

represented by the motors along the diagonal of Fig. 1. This

family enjoys a high degree of uniformity of motor and

controller component parts while offering a wide range of

performance characteristics [9] obtained by changing

only

the

motor rotor and the control strategy.

The rotating field

permits smooth torque and good operation down to low

speeds, both of which are difficult to achieve in the switched

reluctance motor. Unlike the switched reluctance motor, the

synchronous motor is completely compatible with the stators

and controllers of other ac motor drives.

To provide a shorter name and to distinguish it from the

switched reluctance motor, the term SYNCHREL is used in this

paper [12]. The design of a small SYNCHREL motor drive is

described, and the performance is compared with those of a

switched reluctance drive, an induction motor drive, and a

brushless dc (BLDC) PM motor drive. The resplts presented

are confined to the preliminary experimental findings of a

study whose scope includes larger drives than the ones

0093-9994/91/0700-0741$01.00 0

991 IEEE


2/9

142

IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 21, NO. 4

JULYIAUGUST

1991

c c

wound-field

dc commutator

PM commutator

PM brushless dc

Fig.

2.

Phasor diagram

of

S Y N C H R E L motor.

6 6

ac synchronous

ac induction

the d and q axes of the rotor:

l + b IC

T

= - p I d I q

L q L d ) .

(1)

Here, is the number of phases,

p

is the number of

pole-pairs, and L , and L , are the direct- and quadrature-axis

synchronous inductances, respectively. Note that

x d

=

27rfLd and

X =

2 a f L , , where

f

is the frequency. The

torque is independent of speed, provided that the voltage is

boosted above the constant volts per Hertz level to compen-

sate for resistive voltage drop at low speed. The torque per

ampere is maximized if the phase current is oriented at 45 to

the q axis so that Id and I , are equal in magnitude. Since

L , < L , ,

I must be negative, and therefore, the current

leads the q axis in the phasor diagram (Fig. 2 ) . Note that the

convention adopted here, in which the

q

axis is the high-

inductance axis, is contrary to the convention used in the

literature on the line-start reluctance motor. This is because

ac PWreluctance hybrid

@

switched reluctance

Fig.

1 .

Family

of

motor types showing ac motors along the diagonal: the

S Y N C HR E L motor is the center motor with magnets removed [9].

described here, as well as the optimization of lamination

geometry and control parameters. It is plamed that those

results will be published later. Particular points of interest in

the present paper are the comparison of motors of different

types, all with essentially the same frame size and tested

under identical conditions.

11. BASIC HEORY

The inverter-fed

SYNCHREL

motor is freed from the old

constraints of the line-start version as follows:

the particular motors in this paper are related to the interior-

magnet hybrid motor in which the magnet axis is the low-

inductance axis; it is more consistent with classical syn-

chronous machine theory to make this the

d

axis.

Equation

(1)

is the starting point for designing the rotor

lamination. Evidently the

saliency

(i.e. , the difference L ,

L , or the ratio L , / L d ) must be maximized but within

constraints set by manufacturability It is interesting to con-

sider the theoretical limits to the saliency. For a four-pole

motor with sinusoidally distributed windings, if the rotor is

removed, the rotating magnetic field has the form of Fig.

3.

By suitable choice of time origin, the

q

axes can be aligned

with the reference axes of the flux

so

that all the flux is

q-axis flux, and the d-axis flux is zero. The saliency in this

flux pattern is therefore infinite. Now, the objective is to

design a rotor that can be introduced into this field without

disturbing its shape. Since the rotor must be ferromagnetic, it

must present infinite permeance to q-axis flux and zero

permeance to d-axis flux. The obvious way to achieve this is

to make an axially laminated rotor in the fashion described by

Cruickshank [ 2 ] in which the laminations

are shaped to

follow the flux lines in Fig. 3 and are separated by flux

barriers that inhibit the d-axis flux in such a way that if the

rotor were rotated 90 electrical degrees relative to the stator

mmf the flux would fall to zero. This construction (Fig. 3) is

perhaps the natural way to attempt to construct a reluctance

No starting cage is necessary. The rotor can therefore be

designed purely for synchronous performance.

Electronic control makes the motor autosynchronous.

Therefore, the torque angle can be set to maximize

torque per ampere at all loads and speeds without con-

cern for pullout.

There is no need for amortisseur currents to prevent

rotor oscillations. This makes it possible to design for

the highest possible ratio of the synchronous reactances

x nd x d without concern for stability.

~~

rotor with infinite saliency, but in practice, the flux-barriers

are not impermeable, and the saliency is finite.

Assume that the laminations and flux barriers are every-

where very thin, and let t be the average ratio of flux-barrier

thickness to the combined thickness of lamination and flux

barrier. Then,

1/(1 t )

s a measure of the flux concentra-

Because the

SYNCHREL

motor is a classical synchronous

machine, its electromagnetic torque is given by ( l ) ,where

Id

and

I

are components of the rms phase current I resolved

along the

d

and q axes of the phasor diagram; they corre-

spond to the space-vector components of stator mmf along


3/9

MILLER

et al.:

DESIGN

OF

A SYNCHRONOUS RELUCTANCE MOTOR DRIVE

743

Fig. 3 . Natural four-pole field of sine-distributed current sheet repre-

senting the stator winding, aligned with the axis. The shaded sections

represent flux guides interspersed with flux barriers whose surfaces follow

the natural flux lines of the field. This structure was used by Cruickshank

et

al.

[2]

in their line-start reluctance motor. Because of symmetry, only one

octant is shown.

tion that occurs in the laminations owing to the loss of cross

section to the flux barriers. For a peak airgap flux density of

0.8

T

and a saturation density of around 1.7 T, t must be

limited to the order of 0.5. Now, the synchronous reactance

X , is inversely proportional to the airgap length

g ,

and by

the methods of

[9],

it can be shown that

X d

is inversely

proportional to the sum of g and the combined thickness of

the flux barriers, which is very roughly equal to tR , where

R

is the rotor radius. Therefore, the saliency is given approx-

imately by

t R + g tR

=

1.

-

~

x d g g

With t

= 0.5

and R / g typically about 50 this indicates a

maximum saliency of about 25. Values achieved in practice

are usually much smaller (generally no more than

10-15),

partly because of leakage inductance, which effectively adds

a swamping term to both the numerator and denominator

of

2)

and partly because of saturation. Nevertheless, this

simple line of reasoning indicates some fundamental bounds

to the achievement of high saliency and illustrates some of

the factors that are important.

The axially laminated rotor is not easy to manufacture. A

transverse lamination with the pattern

of

flux barriers shown

in Fig. 3 would also be difficult to make by punching;

individual laminations would be flimsy and difficult to han-

dle. The geometry of Fig.

4

is a compromise. It can be

regarded as having just one flux barrier. If this is rectangular,

it can accommodate a permanent magnet, providing a simple

means for enhancing the performance of a small motor when

necessary. With the magnets, the motor is an

interior mag-

net

or

buried magnet

motor, which is sometimes also called

a hybrid PM/reluctance motor (PMH motor). The SYN

CHREL motors in this paper are all of this type.

In evaluating a series of rotor designs,

the

linear magnetic

theory developed in [9] was used to calculate values of

Ld

and

L

in terms

of

dimensions, turns, etc. The values were

checked against finite-element calculations and both ac and dc

measurements

[

121.

111. EVOLUTIONF T HE DESIGN

Three rotors have been built, and the cross sections of two

of these are shown in Fig. 4.The pole pieces are held by two

d-Axis

Web

Rlb

Rotor2

(b)

1 ; (b) rotor

2 .

Fig.

4.

Transversely laminated single-barrier

SYNCHREL

rotors: (a) Rotor

Fig

5 .

Components of hybrid synchronous reluctance/PM motor showing

optional permanent magnets.

thin ribs that attach to the q axis webs in the same way as in

the interior magnet motor described by Jahns, et al. [ 5 ] . Fig.

5shows the components of the disassembled motor.

To minimize Ld the ribs (Fig. 4) must saturate at a low

level of current. This requires them to be radially thin.

L

is

not sensitive to the airgap length because of the large reluc-

tance in the flux barrier. L , must be maximized; therefore,

saturation is undesirable in any part of the q-axis flux path.

Therefore, the pole piece needs to have adequate radial


4/9

744

TABLE

I

IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL.

27,

NO. 4, JULYIAUGUST 1991

Parameter Rotor 1 Rotor 2

Pole arc ( )

Rotor Diameter mm)

Airgap length

mm)

Rib width mm)

Web width (mm)

Flux-barrier thickness (mm)

Rotor m aterial

L, [measured]

mH)

L, [finite-element]

mH)

L, [measured]

mH)

L [finite-element]

mH)

R k o L, /Ld [measured]

68.0

40.5

0.45

0.5

1

o

5.4

Losil 800

10.8

10.2

28.3

21.1

2.6

62.3

41.1

0.15

0.5

2.5

5.4

Losil 800

10.3

11.3

41 O

50.3

4.0

depth, and the web needs to be sufficiently wide as well.

Rotor 1 (Fig. 4(a)) was designed to accommodate 5.4-mm-

thick magnets, and for operation as a SYNCHREL motor, the

webs are too narrow; therefore, they were widened in rotor

2 . With 16 slots, this ensures that the web does not saturate

when aligned with the axis of the phase winding. The param-

eters of the two rotors are summarized in Table I and [12].

The inductances quoted in Table I were measured and calcu-

lated at 3.0 A .

The synchronous inductance ratios quoted in Table I are all

well below the theoretical limit of

25

mentioned earlier. This

is because saturation decreases the q-axis inductance, whereas

leakage through the ribs increases the d-axis inductance. It is

the price paid for the convenience of a lamination that has a

simple punching geometry and the ability to accommodate

magnets when required. However, as stated earlier, an object

of the investigation is to determine whether acceptable per-

formance is obtainable while retaining these katures.

Fig. 6(a) and (b) show typical d- and q-axis finite-element

flux plots. The calculation of magnetization curves is a

straightforward exercise of the finite-element method [181

once the magnetization characteristics of the core steel are

accurately known. Fig.7 shows measured magnetization

curves for rotor 1, clearly showing the effect of saturation on

the inductance ratio. Fig.

8

shows the running torque of both

rotors as a function of rms phase current. The calculated

curves were obtained from equation

(1)

and L , and L , taken

from the appropriate magnetization curves at the appropriate

current level. This calculation is approximate, but it reflects

the general trend and underlines the superiority of rotor 2

with its higher inductance ratio. The torques in Fig.

8

were

measured at a low speed in order to minimize the effects of

windage and core losses and provide data for the comparison

described in Section V.

IV. ELECTRONICONTROL

The configuration of the electronic control for two-phase

motor is shown in Fig. 9. A 360-pulse magnetoresistive

encoder mounted on the motor shaft generates an indexed

pulse count representing the rotor position. This count is used

to address two EPROM's: one for the

d

axis and one for the

axis. The EPROM's contain sine and cosine values multi-

plied in MDAC's by the reference or command value of the

phase current. These analog signals are used as references for

Fig. 6.

(a) D-axis flux plot showing operation of flux barrier and leakage

through the ribs that hold the pole pieces in position. D-axis current

=

3.0

A; (b) Q-axis flux plot. Q-axis current

=

3.0 A .

ROTOR

2 :

F L U X L I N K A G E

v

CURRENT

F l u x Linkage t m k

x

I.Oe2

I .

80

lO.8O-l

2.50

5.00 7.50

Current ( A m p s )

x

1.0eC

Fig.

7 .

Phase flux linkage versus current over a range of rotor positions

between the d axis and the

q

axis; rotor 2.

two full H-bridge hysteresis-type current-regulators, one for

each phase. Power integrated circuits operating at

40

V are

used for the

H

bridges. In the simplest mode of operation,

the current phase angle is set at a fixed value of

45

electrical

degrees, and the motor is controlled entirely by its current

reference with torque approximately proportional to current


5/9

MILLER

et

al . : DESIGN OF

A

SYNCHRONOUS

RELUCTANCE

MOTOR

DRIVE

745

Torque

(mNm) x

l.Oe2

t 501

3.751

Rotor 2

3 00 y

\

1.00

2.00 3.00

Current Arms) x 1.0e0

Fig. 8 . Running torque versus phase current. The points are measured; the

lines are calculated (by

l),

with inductances read from Fig.

6).

Fig.

9.

Electronic controller block diagram.

squared (Fig.

8).

A speed loop can be added outside the

feedforward torque regulator.

A

more sophisticated strategy

is to control the orientation of the stator mmf vector accord-

ing to the operating requirements, and provision is made in

Fig. 9for the addition of a phase-shift to vary the orientation.

V. COMPARISONITH INDUCTION,M, AND SWITCHED

RELUCTANCEOTORS

Tests were performed to compare the SYNCHREL motor

with several other brushless motors of different types. The

dimensions and performance comparisons are summarized in

Table

II.

So many different types of brushless motors are possible in

this size range that it was not possible to test every one of the

different types. In particular, no tests were performed on the

classical brushless dc PM motor. However, it would be

unfortunate to omit this machine because of its commercial

importance, its simplicity, and its close theoretical relation-

ship to the dc commutator motor, and therefore, a column of

calculated results has been included in Table 11 for this

machine (labelled BLDC). In a sense, these figures are

purer than measured results taken on a particular model

because they are exactly defined and totally reproducible, and

since this motor conforms well with relatively simple design

calculations [9], [19], it is used here as the benchmark or

per-unit base to which the parameters of all the other motors

are normalized. The design equations used for this motor are

given in full in [9], and Appendix I contains details of the

design. Fig. 10shows the cross section of this motor. Note

that the slots are rectangular, whereas all the ac motors in

Table 11have round-bottom slots, as is shown in Fig. 4.

A . Description of Motors Tested

Column 1 contains the calculated base values for the

brushless dc PM motor BLDC, which is assumed to have

180 magnet arcs, 120 rectangular phase current wave-

forms, and a wye-connected three-phase stator (Appendix I).

Column 2 is the permanent-magnet hybrid (PMH-1) motor or

interior magnet motor based on the

SYNCHREL

lamination

(rotor 1) with NdFeB magnets of remanent flux density 1.1

T. This motor is labelled PMH-1. The dimensions of the

magnets are identical to those in Column 3 (PMH-2), which

is the PMH motor obtained by fitting

ceramic

magnets of

remanent flux density 0.4 T to rotor 1 of Fig. 4(a). The

BLDC motor in column 1 has the same magnet weight and

the same ceramic magnet material as PMH-2.

Columns 4-7 are induction motors with airgaps ranging

from 0.1 to 0.4 mm in steps of 0.1

mm

to show the

sensitivity of the performance to the airgap length, which is

an important parameter in the cost of manufacture. For

motors of this size and length/diameter ratio, 0 .2 mm is a

normal value for the airgap length.

Columns

8-

10 are switched reluctance motors with airgaps

ranging from 0.2 to 0.4 mm in steps of 0.1

mm

to show the

sensitivity of the performance to the airgap length. The

controller for these machines is based on an integrated-circuit

switched-reluctance drive control described in [20].

Column 11 is the SYNCHREL motor with rotor 1 (Fig. 4(a)).

This is not the best of the SYNCHREL rotors, but it shows the

effect of removing the magnets from the PMH-1 and PMH-2

interior-magnet motors in columns 2 and 3. Column 12 is the

best of the

SYNCHREL

motors described in this paper, with

rotor 2 (Fig. 4(b)). Both SYNCHREL motors, both PMH mo-

tors, and all the induction motors have exactly the same

stator and windings.

All of the motors have four rotor poles, but the induction

and SYNCHREL motors have two phases instead of three. This

does not affect the performance. The SYNCHREL,MH, and

induction motors were operated with the current-regulated

field-oriented PWM inverter described in Section IV. In the

case of the two synchronous machines, the torque angle was

adjusted experimentally to give maximum torque per ampere.

B . Test Conditions

Because of differences in voltage, speed range, and lami-

nation material between the motors, it was not considered

meaningful to compare efficiencies directly. Instead, the per-

formance parameter used for comparison was the torque at

low speed, under conditions of equal stator copper loss in all

the motors. The results have been normalized by calculation

to the same copper weight (0.29 kg) in the stator windings.

Results are summarized in Table 11. Comparisons are made at

a copper loss of about 14 W, which represents about two

thirds of the dissipation capability of the (nonventilated,

totally enclosed) frame for continuous rated operation with a

temperature rise by resistance of about 55C. Assuming that

the copper losses are of the order of 2/3 of the total losses at

maximum power, this also gives a rough idea of the perfor-

mance comparison at maximum power.

Even minor differences in frame size, length/diameter

~ _ _ _ _ _ _


6/9

IEEE

TRANSACTIONS ON INDUSTRY APPLICATIONS,

VOL.

2 1 , NO. 4, JULYIAUGUST

1991

46

Fig. 10.

Cross section of BLDC motor see also Appendix

I).

TABLE I1

MOTOR ERFORMANCEOMPARISON

PARAMETER

Phases

Poles

Stator OD

Rotor OD

Stack Lgth

O/A

Length

Airgap

Stator Cu

Magnet

Steel

Total wt.

Matl cost

Inertia

Ohms/ph

Current

P cu st)

P Cu ro)

Torque

Torque

T/Vr

T/Vs

T/Vtot

T/Weight

T/f Matl

T / J

1

Un BLDC

3

4

mm 77.0

mm 40.6

mm 50.0

mm

86.0

mm/ lO 4 .0

kg 0.29

kg 0.11

kg 0.80

kg 1.20

2.0

100

Ohm 2.9

Arms 1.3

W 14.1

W

mNm 329

100

100

100

100

100

100

100

2

PMH-1

2

4

77.8

40.6

50.0

90.0

4.5

0.29

0.16

1.28

1.73

13.3

134

1.9

1.9

13.6

702

213

213

209

200

159

147

32

3

PMH-2

2

4

77.8

40.6

50.0

90.0

4.5

0.29

0.11

0.98

1.38

2.1

119

1.9

1.9

13.7

295

90

90

88

84

75

78

85

4 5 6

IM-1 IM-2 IM-3

2 2 2

4 4 4

77.8 77.8

77.8

40.6 40.6 40.6

50.0 50.0 50.0

90.0 90.0 90.0

1.0 2.0 3.0

0.29 0.29

0.29

1.25 1.25 1.24

1.54 1.54 1.53

2.0 2.0 2.0

134 134 134

1.9

1.9 1.9

1.9 1.9 1.9

13.7 13.7 13.7

7.1

5.8 4.2

280 250 180

85 76 55

85 76 55

83 74 54

80 71 51

63 57 41

66 59 43

85 76

55

7

IM-4

2

4

77.8

40.6

50.0

90.0

4.0

0.29

1.24

1.53

2.0

134

1.9

1.9

13.7

4.1

150

46

46

45

43

34

36

46

8

SR-1

3

614

73.7

35.5

44.5

74.0

2.0

0.29

0.94

1.23

1.6

25

2.9

1.3

14.3

346

105

155

129

133

414

102

128

9

SR-2

3

614

73.7

35.3

44.5

74.0

3.0

0.29

0.94

1.23

1.6

25

2.0

1.3

14.3

256

78

116

95

99

307

76

94

10

SR-3

3

614

73.7

35.1

44.5

74.0

4.0

0.29

0.94

1.23

1.6

26

2.9

1.3

14.3

188

57

86

70

72

220

56

69

11

REL-1

2

4

77.8

40.5

50.0

90.0

4.5

0.29

1.28

1.57

1.8

109

1.9

1 .9

13.7

109

33

33

32

31

30

25

36

12

REL-2

2

4

77.8

40.0

50.0

90.0

2.0

0.29

1.28

1.57

1.8

109

1.9

1.9

13.7

250

76

78

74

71

70

58

83

Copper f/k g

>

4.0

Note: PMH-1 has NdIGT- 30H magnets interior) Steel f/k g f0.5

PMH-2 has Ceramic-8 magnets interior) Aluminum f/k g f4. 0

BLDC motor has Ceramic 8 magnets surface mounted)

f4 .0

eram. Mag. f/kg

NdFeB Mag. f/ kg E7

O

30-Oct-90

ratio, copper weight, airgap length, and other key parameters

can significantly affect performance. The test data in Table I1

is unusual in that all the motors are very close in physical

size and shape, and all the ac motors have exactly the same

stator. This eliminates the need for adjustments to take

account

of

differences in dimensions. The main adjustment

that had to be made was to the results for the switched

reluctance (SR) motors, which were tested as built with only

0.155 kg

of copper in the stator windings and a very low

slot-fill factor. Since the comparisons in Table I1 are con-

structed on the basis

of

equal stator copper losses, the

performance of these motors had to be recalculated with

almost double the amount of copper in the windings. The

adjustment was performed using the computer-program PC-

SRD [SI, and a careful study of the accuracy of these results

indicates that the calculated torque with 0.29 kg of winding

copper may be as much as

16

low. Moreover, no correc-

tion was made for the fact that the stack length

of

the

SR

motors was only 89 of that of the other motors. The

SR

motor torques in Table I1are therefore underestimated by as

much as 25 , but because these figures have not been

directly verified by test, the conservative adjusted values are

used in Table 11.

The results are valid only for small motors similar in size


7/9

MILLER et

al.:

DESIGN OF A S YNCHRONOUS RELUCTANCE MOTOR DRIVE

747

700 -

m i -

500

a

BLDC

Torque/MatedalCost [%]

[ loo%

=

166mNm/f l ]

SR-1

SR-1

1 2 3 4 5 6 7 8

9

10 11 12

Fig.

11.

Continuous low-speed torque of 78-mm-diameterbrushless motors

at the same stator copper loss with equal stator copp er weights.

1 2 3 4 5 6 7 8 9 1 0 1 1 1 2

Fig.

13.

Torque per unit of material cost of 78-mm-diameter brushless

motors at the same stator copper loss with equal stator copper weights,

expressed as a percentage of the torque per unit material cost of the BLDC

motor in Column

1.

300

250

XK

150

100

5

0

IM 2

1 2 3 4 5 6 7 8 9 1 0 1 1 1 2

Fig. 12. Torque/in ertia ratio of 78-mm-diam eter brushless motors at the

same stator copper

loss

with equal stator copper weights, expressed as a

percentage

of

the torque/ine rtia ratio of the BLDC motor in Column

1 .

to the ones tested, i.e ., in the 50-200-W range or

so.

Extrapolation up to larger powers is unlikely to be meaning-

ful.

C . Performance Comparison

A selection of results from Table I1 is plotted in Figs.

1 Torque: Fig. 11shows the continuous torque capability

of each type of motor relative to the BLDC motor torque.

The PMH-1 motor has the unfair advantage of high-en-

ergy magnets, which explains its high torque capability.

When fitted with the same volume of the same ceramic

magnet material, the interior-magnet motor (PMH-2) pro-

duces only about 90% of the BLDC motor torque and only

14 more than the SYNCHREL motor REL-2.

Even with an airgap of 0.1 mm, the induction motor

produces only 85 of the BLDC motor torque, and it also

has the penalty of 7.1 W of rotor losses that are not included

in the equal stator copper loss criterion. When the airgap

is increased to the more manufacturable value of 0.2 mm, the

11-13.

torque is reduced to 76 ,and if the total losses were reduced

to the same level as in the BLDC motor, the torque would be

derated to only about

50%.

This result is entirely expected:

Small induction motors suffer seriously from excessive mag-

netizing current and slip losses, and the results merely high-

light the known advantage of PM and synchronous motors.

As the airgap is increased still further, the torque falls off

rapidly. At 0.4 mm (the same airgap as used in the BLDC

motor), the induction motor torque is only 46 , even before

the derating due to rotor losses is applied.

The switched reluctance motor also suffers from this in-

verse relationship between torque (at constant copper loss)

and airgap length, but in this case, there is no slip loss, and

the degradation is due to increased excitation requirements as

the airgap increases. With constant stator copper loss, the

relative rate of degradation is very roughly the same as in the

induction motor. The SR motor produces some 30 more

torque than the induction motor with the same airgap but

without any significant rotor losses. Torque parity with the

BLDC motor is achieved only with a 0.2-mm airgap. The SR

motor cannot tolerate the large airgap (0.4 mm) of the BLDC

motor; with a 0.4-mm gap, its torque is only 57 . It was

observed in testing that the SR motor was the noisiest motor,

even at the preadjustment torque level (around 200 mNm for

SR-1 with 0.155 kg of copper). If this motor is actually

operated at 346 mNm, it is excessively noisy even though it

is still operating far below its theoretical electromagnetic

capability.

The

SYNCHREL

motor REL-1 in column 11 is included to

show the effect of removing the magnets from PMH-1 or

PMH-2; there is a marked reduction in torque to 33 , and

again, the large airgap of the BLDC motor cannot be toler-

ated. The motor REL-2 (column 12, using rotor 2 of Fig.

4(b)) has much better performance (76 ) and even outper-

forms the induction motor IM-2 with the same airgap length.

Moreover, it does this with negligible rotor losses. Although

the torque is lower than the 105 of SR-1 with the same

airgap length, it is comparable with the 78 of the SR-2


8/9

748

IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 27.

NO.

4, JULYIAUGUST 1991

motor, and if these motors had to be designed for the same

noise level, the SYNCHREL motor REL-2 would be competi-

tive. It has the further advantage of using a standard induc-

tion-motor stator.

2) Torquel lnert ia Rat io T/ J) : Fig. 12 shows the

torque/inertia ratio based on the continuous torque capabil-

ity. (Transient torque capability was not evaluated.) The high

value for the SR motor is mainly due to its low inertia, which

results from the small rotor diameter and the removal of

material between the rotor poles. The SYNCHREL motor REL-2

has about 20 better T/J than the induction motor IM-2 but

only

70

of that of the BLDC motor.

3

TorquelMaterial Cost Ratio: Fig. 13shows the torque

per unit of material cost based on the per-kilogram costs of

materials given in Table 11. These costs are approximate and

somewhat variable, but they do give an additional insight.

The PMH-1 motor is clearly penalized by the high cost of its

magnet material. Otherwise, the only motor to excel the

BLDC motor is the switched-reluctance SR-1 motor. The

SYNCHREL motor REL-2 is a few percent more economical

than the induction motor IM-2, but both are 15-20% less

cost effective than the BLDC motor.

VI. CONCLUSION

A simple synchronous reluctance (SYNCHREL)otor using

a 78-mm-diameter induction motor stator and fixed-phase-

angle variable-frequency control produces an efficient syn-

chronous motor drive with approximately 50% more torque

than the induction motor in the same stator based on equal

total motor losses. The low-speed torque is about 30 lower

than that of a switched reluctance motor having the same

framesize, airgap length, and copper weigh; however, in

other respects, the quieter SYNCHREL motor has many of the

attractive features of the switched reluctance motor-freedom

from permanent magnets, high-temperature and high-speed

capability, freedom from parameter variations due to temper-

ature, etc. It also uses standard ac motor parts and sinewave

control. Calculations show that the SYNCHREL motor as de-

scribed

in

this pape r

cannot equal the torque of the brush-

less dc motor with ferrite magnets in this size range, even on

the basis of torque per unit material cost; neither can it

tolerate the large airgap length of the BLDC motor. How-

ever, if manufacturing cost is taken into account, the SYN-

CHREL motor is quite competitive because of its simple rotor

and the common induction motor stator.

These results have been achieved with a single flux-barrier

design capable of accommodatipg permanent magnets. The

inductance ratio is much smaller than theoretically possible in

a pure SYNCHREL motor, and much better results would be

expected with an axially laminated construction or equiva-

lent.

The comparison of motor types underlines the superiority

of the PM brushless dc motor in raw torque production at

low speed and its ability to tolerate a large airgap length. The

comparison also highlights the weakness of the induction

motor in this small size range. The advantage of the SYN-

CHREL motor is that it uses a standard induction-motor stator

and provides a synchronous, efficient drive with parameters

independent of temperature and freedom from demagnetiza-

tion and other magnet-related problems. Although the perfor-

mance is exceeded by the switched reluctance SR) motor,

the SR motor has the disadvantage of a higher noise level and

higher torque ripple, and it cannot use a standard ac motor

stator.

These conclusions cannot be extrapolated to larger motors

because the effects of scale are too nonlinear. Future plans

include the extension to integral-horsepower motors

[

161 and

the investigation of small motors with much higher induc-

tance ratios.

APPENDIX

PARAMETERS

F

BLDC MOTORDESIGN

Stator/rotor diameters

Airgap length

Stack length

Magnet thickness/pole arc

Magnet/remanent flux-density

Coercive force

Poles/slots/phases

Tooth width

Slot area

Winding type

Coil throw (slot pitches)

Turns in series per phase

Self/mutual inductance

Airgap flux density

Torque constant

Speed at test point

77/50 mm

0.4 mm

50 mm

3 . 8

mm/180 elec

ceramic/0.329T

269 kA/m

4/24/3

2.5 mm

61.5 mm2

Lap, single layer

5

200/0.6 mm dia

12.7/1.8 mH

0.246T (open circuit)

0.2

/A

200 r/min

ACKNOWLEDGMENT

Thanks are due to R . S . Boughton (Sherman Electromech),

S. E. Wood (Brook Crompton Parkinson), P. Ibbotson (Wat-

son Marlow), A. J. Hutton (Motorola),

K.

Debebe,

I.

Young,

and J. Kelly of the University of Glasgow.

REFERENCES

[l ] P.

J .

Lawrenson and L. A. Agu, Theory and performance of

polyphase reluctance machines, Proc . Inst. Elec. Eng., vol. 111,

A.

J . 0

Cruickshank, A. F. Anderson, and R . M. Menzies, Axially

laminated anisotropic rotors for reluctance motors,

Proc .

Inst.

Elec. En g., vol. 113, pp. 2058-2060, 1966.

[3] W. Fong, and J .

S .

C. Htsui, New type of reluctance motor,

Proc . Inst. Elec. Eng., vol. 117, pp. 545-551, 1970.

[4] V. B. Honsinger, Steady-state performance of reluctance machines,

IEEE Trans. Power App. Syst. , vol. PAS-90, pp. 305-317, 1971.

[ ] T. M. Jahns,

G .

B . Kliman, and T.

W .

Neumann, Interior magnet

synchronous motors for adjustable-speed drives, IEEE Trans. In-

dustry Ap plications, vol. IA-22, pp. 738-747, 1986.

T. Fukao, Principles and output characteristics of super high-speed

reluctance generator system, IEEE Trans. Industry Applications,

A. Fratta and A. Vagati, A reluctance motor for high dynamic

performance applications, in

Conf.

Rec .

1987

IEEE Industry

Applications Soc. Ann. Mtg ., Part I

pp. 295-302, Paper ID-87-24.

A. Chiba and

T .

Fukao, A closed-loop control of

super

high-speed

reluctance motor for quick torque response, in

Conf.

Rec. 1987

IEEE Industry Applications Society Ann. Mtg., Part 1, pp.

289-294, Paper ID-87-23.

[9]

T.

J . E. Miller, Brushless Permanent-Magnet and Reluctance

Motor Drives.

T.

J . E.

Miller and

M . I.

McGilp, PC CAD for switched reluctance

pp. 1435-1445, 1964.

[2]

[6]

vol. IA-22, pp. 702-707, 1986.

[7]

81

Oxford: Oxford University Press, 1989.

[ lo ]


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drives, in IEE Conf. Publ . 282, pp. 360-366, 1987.

P.

J .

Lawrence et a l . , Variable-speed switched reluctance motors,

Proc. Inst. Elec. Eng., vol. 127, pt. B, pp. 253-265, 1980.

T . J. E . Miller and K. Debebe, Design of a synchronous reluctance

motor, in P CIM M O T O RCO N

Conf

roc. Munich), June 6-8,

T. Lip0 and L-Y Xu, A novel converter-fed reluctance motor with

high power density, in Symp. Electric Drives Cagliari, Italy), June

L.-Y. Xu and T. Lipo, Analysis of a variable speed singly-salient

reluctance motor utilizing only two transistor switches, in Conf .

Rec.

1988

IEEE Industry Applicat ions Soc. An n. M tg., Part

I,

pp.

T. J . E. Miller, P. G. Bower, R. C. Becerra, and M. Ehsani, Four-

quandrant brushless reluctance drive, in Proc. IEE Conf . Power

Electron. Variable-Speed Drives London), July 1988, pp. 273-276.

[16] M. R. Ham s and T.

J .

E. Miller, Comparison of design and

performance param eters in switched reluctance and induction motors,

in

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London), Sept. 1989,

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J .

E. Miller, C. Cossar, and D. Anderson, A new control IC for

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[13]

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[15]

pp. 303-307, CP310.

[17]

[18]

[19]

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[20]

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CP324.

T. J.

E.

Miller SM82) is a native of Wigan, UK.

He was educated at Atlantic College and the Uni-

versities of Glasgow and Leeds, U.K.

He spent 20 years in industrial research and

development, including eight years with General

Electric Corporate Research and Development,

Schenectady NY , before becoming Titular Profes-

sor in Power Electronics at Glasgow University.

He is author of two textbooks, numerous patents,

and IEEE and IEE publications.

Alan Hutton was born in Glasgow, Scotland, on

April 29, 1966. He received the B.Eng. Hons.)

degree in electrical and electronic engineering from

the University of Glasgow, Scotland, in 1987.

He has spent two years as a research assistant

under the S PEE D consortium at the University of

Glasgow. He was e ngaged in the design of switched

reluctance motors and drives and is about to submit

for a Masters degree based on this work. He

is

presently working for Motorola in technical mar-

keting.

Mr. Hutton is an associate member of the IEE.

Calum Cossar was born in Hamilton, Scotland, on

January 22, 1962. He received the B.Sc. Hons.)

degree in electrical and electronic engineering from

the University of Glasgow, Scotland, in 1983.

He spent five years as a design engineer with

Ferranti Defence Systems Ltd. and was involved in

high-speed DSP in radar systems. For the last two

years, he has been employed as a research technol-

ogist at the University of Glasgow. His field of

interest is digital techniques in motor control.

David A. Staton was born in Chesterfield, Eng-

land, on July 29, 1961. He received the B.Sc.

Hons.) degree in electrical and electronic engi-

neering from Trent P olytechnic, Nottingham, Eng-

land, in 1983 and the Ph.D. degree from the

University of Sheffield, England, in 1988.

From 1977 to 1984, he was employed by British

Coal, who sponsored him while he was undertal-

ing his B.Sc. degree. While at the University of

Sheffield, he developed CAD software for perma-

nent-magnet dc motors in collaboration with GEC

Electromotors Ltd. F rom 1988 to 198 9, he was with the Thorn EM1 Central

Research Laboratories and was engaged in the design of motors for the

Kenwood range of food processors. Over the last year, he has been

employed as a research assistant at the University of Glasgow and is

involved in optimizing the design of synchronous reluctance motors.

Dr. Staton is an associate member of the IEE.

Documents

Design of a synchronous reluctance drive