Design of 400 Hz saturistor motors with improved

starting performanceB.J. Chalmers, B.Sc, Ph.D., D.Sc, C.Eng., F.I.E.E., L.G. Atkinson, C.Eng., M.I.E.E., J.W.R. Cox, B.Sc, Ph.D.,

and M.M.K. El-Attar, M.Sc, Ph.D.

Indexing terms: Computer applications, Induction motors

Abstract: A method is presented for predicting, from physical design data, the performance of saturistorinduction motors having idle bars of permanent magnet material in the rotor slots. It is shown how theanalysis may be used to synthesise the physical design of a motor in order to achieve a specified starting per-formance. The procedure is applied to the design of a 400 Hz aircraft motor to obtain reduced startingcurrent compared with an original cage motor, without significant detriment to its normal running perfor-mance. Measured B-H loop data is presented for Alcomax III, Hycomax III and Hycomax IV materials. Testresults demonstrate the validity of the treatment and the success of an experimental design incorporating barsof Hycomax III.

List of symbols

A = area of B-H loopDs = saturistor energy densityd = depth of saturistor/ = supply frequencyHa = peak magnetising force in slot airgapHs = peak magnetising force ii: saturistorlb = bar current1ST = starting currentL = length of saturistorPL = hysteresis lossRs = effective saturistor resistanceRT = total effective resistance of equivalent circuits = slipTc = torque component due to copper cageTs = torque component due to saturistor materialTST = starting torqueVm = volume of magnetic materialwa = width of slot airgapvvs = width of saturistor barXs = effective saturistor reactanceZ = impedance of equivalent circuitZs = effective saturistor impedanceS2 — number of rotor slots<!> = peak flux through saturistor barcos = angular synchronous speedPrime indicates rotor parameter referred to the statorphase

1 Introduction

For some years, aircraft constructors have been concernedabout the high starting current of aircraft 400 Hz inductionmotors,1 which, in some modern aircraft may account forsome 30% of the supply capacity. The starting current ofsuch motors has a significant influence on the capacity ofthe power source, the cable weight and the protectionarrangements, and its reduction would have a beneficial

Paper 942B, first received 4th March and in final form 22nd July1980Dr. Chalmers is, and Dr. El-Attar was formerly, with the Universityof Manchester Institute of Science and Technology, Sackville Street,Manchester M60 1QD, England. Dr. El-Attar is now with theFaculty of Engineering, Alexandria University, Alexandria, Egypt.Mr. Atkinson and Dr. Cox are with the Engineering Physics Depart-ment, Royal Aircraft Establishment, Farnborough, Hampshire,England

IEE PROC., Vol. 127, Pt. B, No. 6, NOVEMBER 1980

effect on the aircraft power-supply system as a whole. Onemethod of reducing the starting current without significantdetriment to the performance in its normal working regimeis by the use of saturistors in the form of idle bars of hardmagnetic materials in the rotor slots. The original concept ofthe saturistor was described by Alger et ai? in 1963 andwas patented in 1966.3 This patent describes various formsof the saturistor including the idle-bar cage-rotor type, thephysical arrangement of which is shown in Fig 1. The typeof saturistor comprising a separate assembly connected toa wound rotor has been extensively investigated.2'4"7 Cagesaturistor rotors are considered to possess useful technicaladvantages in 50 and 60 Hz applications, but a quantitativeprocedure for their design has not been described pre-viously, nor has their full potential for successful use with400 Hz supply been demonstrated in practice. Some experi-mental results have recently been reported.8

This paper relates to the cage type of rotor and describesa method by which the idle-bar material and dimensionsmay be determined to achieve specified values of startingcurrent and torque, together with adequate running per-formance. The validity of this design technique is con-firmed by comparison with tests on a commercial 3-phase,6-pole, 400 Hz aircraft motor rated at 5 h.p. (3-73 kW), andthe achievement of improved starting performance meetinga specified target is demonstrated by results for an experi-mental saturistor motor.

7nonmagneticmnt<»rinl(airgap)

- —

? a | | - Ws

Ha " Hs

Slot

saturistormaterial

copperbar

width

Fig. 1 Arrangement of saturistor in rotor slot

341

0143-7038/80/060341 + 08 $01-50/0

2 Motor model

2.1 Equivalent circuit

The saturistor is modelled simply by a series impedance inthe rotor circuit, comprising a resistive part, representingthe energy that it dissipates in hysteresis loss, and a reactivepart produced by the increased flux linkage that resultsfrom the presence of the saturistor in the magnetic circuitexcited by the rotor-bar current. Thus the per-phase equiva-lent circuit of the saturistor machine referred to the statoris that of a conventional machine modified to the formshown in Fig 2. It is necessary to derive the equivalent cir-cuit parameters of the saturistor in order that the perfor-mance of the modified machine may be calculated.

The approach adopted assumes that the presence of thesaturistor material, and any changes to the rotor slot dimen-sions, do not significantly alter the main flux distribution,i.e. the stator leakage reactance, the magnetising reactanceand the iron losses are assumed to be unaffected by suchchanges.

The derivation of equivalent circuit parameters for thesaturistor, for use in an equivalent circuit with sinusoidalvoltages and currents, presents a conceptual difficultyowing to the nonlinearity of the B-H properties of thematerials used. The approximate method used here for thedetermination of the saturistor equivalent circuit para-meters concentrates on the fundamental torque-producingeffects and is as follows.

2.2 Determination of equivalent circuit parameters

The saturistor impedance per slot in rotor terms Zs =Rs + j'Xs is calculated from the four-quadrant B-H loops ofthe appropriate hard magnetic material, the area of whichrepresents the hysteresis loss per unit volume per cycle ( i.e.the energy density of the material) for a given peak magnet-ising force.

If, for a given peak alternating field H8 across the satu-ristor, a peak flux density B8 is produced within it and thearea of the resulting B-H loop is A, then the power loss perslot in the material of volume Vm is given by

= VmAf 0)

If an r.m.s. bar current Ib is necessary to produce such afield, then the effective resistance of the saturistor idle baris

Rs =~n (2)

The peak flux 4> that passes through the saturistor andhence around the bar is

= BAL

M A ,

(3)

I'r

Fig. 2 Equivalent circuit of saturistor motor

342

This generates in the bar an e.m.f. which, on theassumption of a sinusoidal variation of flux, produces anr.m.s. voltage of

(4)

The r.m.s. value of the bar current is related to its peakvalue by

h = (5)

where K was assumed to have a value of unity in recog-nition of the rectangular nature of the B-H loops of thesaturistor materials used. Ib peak is related to the slot-saturistor configuration by Ampere's law, which statesthat

<§Hdl = I

where H is the magnetic field vector and / is the totalcurrent within the path around which the integral is per-formed. Applying this to the arrangement shown in Fig. 1,

'bpeak (6)

The effective impedance per slot of the saturistor Zs is thendefined as

7 *rms

which, on substitution of the above expressions forKr.m.fc and/b , gives

Z~t e

y/2 nf&K (7)

Hsws + Hawa

Thus the bar current determines the field and thereforeP and <&, and uniquely defines Rs and Zs, whence theeffective saturistor reactance is

X. = (8)

The above procedure was tested against published measure-ments2'4'5 on static saturistors and in most cases the calcu-lated values generally showed good agreement with themeasured values for Alcomax III, using the assumptionsthat the applied voltage was sinusoidal and the resultingcurrent was of square waveform. The effect of eddycurrents in the saturistor material are not considered to besignificant in the present work because, assuming a meanpermeability of four, the skin depth is of the order of 10mm at 400 Hz,9 which is substantially larger than the depthof the saturistor bars used in the motors described herein(typically 1 -5 to 2 mm).

For the purpose of computation, a family of 20 sets ofmagnetic data was compiled from measured characteristicsfor each saturistor material. This data related peak H topeak B and hysteresis loss density. The remaining para-meters of the equivalent circuit (Fig. 2) were evaluated asfollows. R\ was measured directly. Xm and Rm wereobtained from the no-load test on the original cage motor.R'2 and X>2 were calculated from rotor-slot and bar dimen-sions for each rotor design, neglecting skin effect becausethe bar depths in the motors under consideration are muchless than the penetration depth at 400 Hz. It was foundnecessary to cater for saturation of leakage reactance, asillustrated by the substantial variation of Xi in Fig 13 later,and to achieve this in a convenient manner all saturation

IEEPROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980

was assigned to X\. This was considered justifiable, as it is-well known that induction-motor performance is notgreatly influenced by variation in the subdivision of a giventotal leakage reactance into stator and rotor components.Total leakage reactance was measured by a locked-rotortest on the original cage motor and, deducting the calcu-lated rotor impedance of that rotor, X\ was evaluated as afunction of load current.

2.3 Performance analysis procedure

The saturistor impedance is a function of both frequencyand rotor current. In the equivalent circuit, the rotor impe-dance is referred to the stator, using the usual ratios, andcan then be considered to operate at the supply frequency.It is then necessary to establish the relationships betweenslip, rotor current and input current, and the method des-cribed by Dubey and De6 was adopted.

Referring toFig 2, the primary impedance is replaced byits TheVenin equivalent shown in Fig 3, where

\E'\ =

Table 1: Design data

(9)

R =(RlRm-X1Xmf +(RlXm+RmXl +RmXmy

(10)

X =RmX\X

From Fig 3,

l'r

or

s = ' 2

X'2+Xs)2

(11)

(12)

(13)

Rs and X8 are functions of l'r and, if a value of l'r is assumed,R8 and X8 can be calculated as described in Section 2.2.These, when transformed to their equivalent primaryvalues, are then used to calculate s directly, avoiding themore complicated procedure required if values of slip wereassumed initially.

Fig. 3 Thevenin equivalent of Fig. 2

IEEPROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980

Original 400 Hz, 6-pole motor

StatorNumber of slotsConductors/phaseResistance/phase at 20° CBore diameterStack o.d.

RotorNumber of slotsBar diameterBar cross-sectional areaEffective bar lengthSkewBar resistance at 75° CEnd ring (per bar) resistance at 75°CSlot detail

Experimental saturistor rotor (Details aswhere indicated below)

Saturistor materialSaturistor volume/slotSlot detail

2790

0O8f t57-38 mm

100-1 mm

362-23 5 mm3-923 mm2

102mm1 slot pitch5-439 X 10 " a0O84X 10"" ftFig. 4a

original cage rotor except

Hycomax III574 mm3

Fig. Ab

l'r is first adjusted iteratively to determine its value atwhich s = 1. The motor performance is then calculated forthis value of l'r and for a series of progressively lower valuesof l'r to establish the complete characteristics from s = 1 to

(14)

The torque and input current are calculated from theexpressions

3/;2

Is

and

R-.

Input power and power factor follow directly from theequivalent circuit.

A computer program was prepared to solve theseequations and to calculate the performance characteristics.Xi was varied with l'r, using data obtained as described inSection 2.2.

2.4 Verification of the program

In the initial phase10 of the study, the computational pro-cedure was verified in comparison with experimental resultsusing

(a) the original cage motor, whose data are given inTable 1 and Fig. 4a

(ft) a preliminary rotor coreplate modification havingdeeper slots, similar cage-bar size to (a), but no idle bars

(c) as rotor (b) but including idle bars of Alcomax III.The original stator was used throughout this study

To obtain full performance data over the full range ofspeed including the unstable region, a test installation andprocedure was developed1!, in which all quantities wererecorded during dynamic starting with a large inertial loadcoupled to the motor via a torque transducer. It was con-firmed that runup was always slow relative to the electrical

343

time constants, and so could be regarded as giving pseudo-steady-state performance. Rotor heating during runup had asignificant effect and, to standardise the procedure, all suchtests were made with the motor initially at room tempera-ture.

Following development of the model as described inSections 2.2 and 2.3, calculated performance characteristicsgenerally showed acceptable agreement with measuredtorques and currents, but input power showed some diver-gence from the measurements. Discrepancies may beattributed to power-measurement errors associated withlow power factors, changes of temperature and any residualimperfections in the model including its neglect of parasiticphenomena.

In illustration of the correlation obtained, Fig. 5 givesmeasured and computed performance of the original cagemotor. To show the influence of temperature variations,computed results are presented for rotor-cage tempera-tures of 20°C and 200°C, representative of actual rotortemperatures at the start and finish of a typical test runup.It is seen that the computations at 20° C give best agree-ment at start, in respect of current and torque, while thoseat 200°C agree well at the end of the acceleration.

1 020 mm

1 -60 mm

2 39 mm did.

0-38 mm

2 08 mm

1-40 mm

2 79 mm

Fig. 4 Rotor slot dimensions

a Original cage rotorb Experimental saturistor rotor

Steady-state running tests at normal temperatures, suchas those summarised in Table 2, also gave results whichsupported the validity of the computations and in generalterms, the agreement achieved is considered to be fully upto normal standards for the size of motor under consider-ation.

3 Design procedure for saturistor rotors

3.1 Design considerations

The synthesis of a design that is optimised to the user'srequirement is essentially an iterative process, carried outby the designer changing the input parameters in a system-atic way at each iteration. This procedure calls for a degreeof skill in the interpretation of the output from theprogram, and in making appropriate alterations to the inputdata to produce performance figures suited to the designer'sneeds.

Eqn. 6 shows how the magnetising field of a given barcurrent is utilised to magnetise the saturistor material. Tofully exploit the material properties, the magnet should, asfar as manufacturing tolerances permit, occupy the wholeof the available space in the slot and be excited, at start, tothe value of Hpeak at which the most satisfactory value ofZs is obtained. A change of slot dimensions will alter theproportions and/or volume of the magnet material, andhence Zs, and also the slot permeance. This change in rotorimpedance will clearly modify the rotor current, as them.m.f. across the magnetic material is a function of bothslot dimensions and current. The relationships are suf-ficiently complex to prevent a direct calculation of the slotand saturistor dimensions to produce the desired target per-formance. An iterative approach is therefore adopted,whereby the starting performance is computed for each of anumber of slot designs until an arrangement is found whichsatisfies these requirements. Some computer time is savedby making an initial estimate of the volume of magneticmaterial required, as follows.

Given the rated current / and the rated torque T of aparticular conventional induction motor, a target startingcurrent IST of Kjl and target starting torque TST of KTTare specified. The approximate torque contribution of thecopper cage at start (s = 1), ignoring the fact that fST

includes the magnetising current, is

Tr. =3ISTR2 (16)

The required torque contribution of the saturistor materialis

- TST Tc (17)

From the data for the appropriate saturistor material, thesaturistor energy density Ds is determined at 400 Hz and atthe excitation level at which the saturistor resistance ismaximum. Whence the volume of saturistor material perslot is

(18)DsS2f

Having allotted a value to the slot width, the program isthen used to calculate the starting performance of themachine. The dimension wa is then progressive increased,decreasing the volume of saturistor material, to explore themachine's performance over a range of values. Effects of

344 IEEPROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980

Table 2: Comparative performance data

Machine type

Original cage rotor

(1) (2)

Experimentalsaturistor rotor

(1)

Revised saturistordesign

(3)

Output

h.p.456

456

456

kW2-983-734 4 8

2-983-734-48

2-983-734 4 8

Slip%4-15-87-8

4-86-59O

3-34-15-1

p.f,

0 4 90-570-66

0 4 40-510-58

0-560-620-66

Efficiency%899089

868685

878888

TSTNm

9-3

7-5

8-6

lSTA

102

71

84-8

Q

1

1-66

1-34

'peakComputedNm

14-9

11-9

13-7

MeasuredNm

15

10-7

(1) Measured(2) Has harmonic torque dip, which may introduce differences between measured and calculated values(3) Computed results for rotor temp, of 20°C, slot width 3 05mm, saturistor depth = 1 -27 mm, vol/slot - 384 mm3, wa = 0-076 mm

variation of slot depth and width are similarly examineduntil the most suitable combination of dimensions is deter-mined. Maximum slot width is limited by tooth saturation,manufacturing and mechanical strength considerations.

3.2 Design development

The objective -was to achieve a saturistor design withstarting and running performance similar to the originalcage machine, but with reduced starting current.

The starting data for the original cage machine were, fora start from ambient temperature, TST = 9-3 Nm and/ S T = 1 0 2 A . The rated output was taken to be 5 h.p.

(3-73 kW), 4-7 Nm, at 7500rev/min with an input currentof 22 A. Thus, the original cage motor had KST = 1 -98 andKj = 4-64.

A starting torque of 1 -6 times the rated torque and astarting current of 3-2 times the rated current were set astargets, and a design was developed on this basis. Thus, thetargets were 7'S T=7-5Nm and / S T = 7 0 - 4 A . UsingAlcomax III idle bars and with maximum slot width deter-mined simply on the basis of tooth saturation, it was cal-culated that the best starting performance achievable wouldbe 7-02 Nm and 71 A. From a manufacturing viewpoint itwould be necessary to make the tooth width appreciablywider than this minimum and, as a consequence, the

\U-\

12-

10-

E 8"

u-

2-

0J

1000 2000 3000 4000speed , rev/ min

5000 6000 7000 8000

Fig. 5 Performance characteristics of original cage motor

torque, measuredcurrent, measured

— - - — input power, measuredComputed characteristics• torque, 200°C A current, 200°C • power, 200°Co torque, 20°C ^ current, 20°C a power, 20°C

IEEPROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980 345

volume of Alcomax would have to be reduced. It was there-fore clear that a material of higher energy density would berequired if the target starting performance were to beachieved in a practical design.

Energy density data were not available from magnetmaterial manufacturers and the simple BHmax product is oflittle relevance since materials having nominally the sameBHmax value may have widely-differing maximum energydensities, as illustrated by Figs. 6 and 7. Materials may beselected, in the first instance, for the suitability of theirvalues of remanence Br and coercive force Hc. Hc should beless than the available slot magnetising field and Br shouldrepresent a good compromise between excessive saturistorreactance and adequate energy density. Four-quadrant B-Hloops were measured12 for various materials, by means of apermeameter using electronic integration of search-coil vol-tages. Figs. 8, 9 and 10 present these characteristics forAlcomax III, sintered Hycomax III and sintered HycomaxIV, respectively.

For the present application, Hycomax III with a highervalue of Hc, wellmatched to the available slot m.m.f., alower value of Br and a higher energy density, appeared tohave distinct advantages over Alcomax III. Thus to achieve

flu>

constant BH xcontour

\ i<

. - . • • " ' . / / .

3 2 / • 1

/ // /

Hycomax TS j '

y-'j£'\ )^ ^ Hycomax ID

density, T

r•

0-4

02

0 2

04

y

10

^y

XL

/

•10

•0-6

0

j

• /L•\L

1. /1/

/ /

j 1!

/ 2 3magnetising for

AT/m x 1 0 " ^

^,0-8Alcomax ID

the maximum value of Rs using Alcomax III, it would benecessary to allow a substantial airgap wa, whereas withHycomax III this could be achieved with a virtually fullslot. The computed starting performance of a range of

flux density, T

Fig. 8 B-H characteristics of Alcomax III

Fig. 6 Four-quadrant curves for some permanent magnetmaterials

32

180r

!15O

o120

^. 90

- 6 0

Q- 30

Fig. 9 B-H characteristics of sintered Hycomax III

Hycomax DI >•' y ^ '

Alcomax ID

7// / Hycomax ffi

/ /

Fig. 7 Power density curves for some permanent magnetmaterials

346

Fig. 10 B-H characteristics of sintered Hycomax IV

IEEPROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980

designs, using Hycomax III with a slot width of 2-79 mm,are shown in a convenient form in Fig. 11. In a practicalmachine it is clearly necessary to assume a small clearancebetween the idle bar and the slot; for this reason, designshaving an airgap wa of 0076 mm (0-003 in) were examined.

The set of dimensions which most closely met thetarget starting requirements are indicated by the encircledpoint in Fig. 11. An experimental saturistor motor was con-structed with these dimensions, which are recorded at thefoot of Table 1 and in Fig. 4b. The measured characteristicsof this motor, together with the computed performance forcage temperatures of 20°C and 200°C, are shown in Fig. 12.The agreement is generally as good as that obtained for theoriginal cage motor, although the peak torque is some 10%lower than that predicted and peak input power is alsolower than that computed. As summarised in Table 2, theexperimental saturistor motor was fully successful inmeeting the starting performance targets of 7-5 Nm and70-4 A, for a start from room temperature. The measured

steady-state performance at normal operating loads and

saturistor volumeper slot, mm3

90r262

557

saturistor depth, mm

Fig. 11 Saturistor rotor design data for locked-rotor condition

Hycomax IIISlot width = 2-79 mm

0 1000 2000 3000 4000 5000 6000 7000 8000speed, rev/min

Fig. 12 Performance characteristics of experimental saturistormotor

torque, measuredcurrent, measured

_ . . —input power, measuredComputed characteristics• torque, 200°C * current, 200°C • power, 200°Co torque, 20°C A current, 20°C o power, 20°C

temperatures is also given in Table 2, showing little depar-ture from that of the original cage motor.

The calculated variations of impedance components areshown in Fig. 13. This illustrates the pick-up of the satu-ristor impedance at just above rated slip, so that it does notinfluence running performance at normal loads, and its risetowards a maximum near standstill. At standstill, the satu-ristor impedance is mainly reactive and has the desiredcurrent reducing effect.

To illustrate the design flexibility offered by saturistormotors, a revised design of saturistor was computed, whoseperformance more closely approached that of the originalcage motor, but which had a useful reduction in startingcurrent. These performance data are included in Table 2.The ratio TSTjl\T is proposed as a useful figure of meritfor comparative purposes and is shown in Table 2 as thefactor Q, expressed relative to the value for the originalcage motor.

4 Conclusions

A method of design analysis has been presented whichsatisfactorily predicts the performance of saturistor-typecage induction motors. Its use in a procedure for designsynthesis to meet a specified starting performance has beendemonstrated in practice.

The study has established that use of idle bars of suitablemagnetic material in 400 Hz aircraft induction motors caneffect substantial improvement in the TST/IST ratio, with-out significant loss of performance in the normal workingregion. In the experimental motor incorporating HycomaxIII, starting current was reduced by 30% and starting torquewas reduced by 17%, compared with the original com-mercial induction motor. A recent experimental study8

using Alcomax III reported a 10% reduction in startingcurrent accompanied by 25% reduction in starting torque.Relative to their respective original cage motors, the TST/1ST r a t i ° w a s improved by 66% in the present study andreduced by 9% in Reference 8.

It is clear that, apart from the high starting current,the best performance can be obtained from an inductionmotor of conventional design. In the case of the saturistormachine, the TSTII%T ratio improves with increase of satu-ristor material, but this aspect of performance is accom-panied by an increasing loss of starting torque, peak torqueand a more modest degradation of normal running per-formance. However, the introduction of saturistor barsdoes allow the designer a much better compromise betweenreduced starting current and running performance than isobtainable in conventional squirrel-cage machines.

0 1000 2000 3000 4000 5000 6000 7000 8000speed, rev/min

Fig. 13 Impedances of experimental saturistor motor

IEEPROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980 347

5 Acknowledgment

The authors are grateful to the Ministry of Defence (PE)for the sponsorship of the work done at UMIST. Theyalso wish to thank UMIST for the facilities provided andM.H. Salama for his involvement in the early stages ofthis work.

6 References

1 RATCLIFFE, J.P.: 'A constructor's viewpoint'. IEE andR. Aeronaut, Soc. Colloquium on 'Is constant frequency powergeneration worthwhile', 1966, pp. 15—17

2 ALGER, P.L., A*NGST, G., and Schweder, W.M.: 'Saturistorsand low starting current induction motors', AIEE Trans.,1963, PAS-82,pp. 291-298

3 ALGER, P.L.: 'Saturistors comprising hard magnetic materialsenergised by alternating currents'. US Patent 3293468,Dec. 1966.

4 ALGER, P.L.: 'The nature of induction machines', (Gordonand Breach, 1965), pp. 28-33

5 GUNN, C.E.: 'Improved starting performance of wound rotormotors using saturistors', AIEE Trans, 1963, PAS-82, pp. 298-302

6 DUBEY, G.K., and DE, G.C.: 'Saturistor control of inductionmotor', Proc. IEE., 1973, 120, (4), pp. 491-496

7 SABIR, S.A.Y., SIMMONS, S., and SHEPHERD, W.: Designof saturistors for induction motor speed control', IEEE Trans.,1975, IA-11, pp. 489-493

8 ALGER, P.L., and SHEPHERD, W.: 'A 400 Hz saturistor motorfor aircraft pump duty', IEE J. Electr. Power Appl, 1979, 2,(4), pp. 145-148

9 BRAILSFORD, F.: 'Physical principles of magnetism' (VanNostrand, 1966)

10 ATKINSON, L.G., and COX, J.W.R.: 'An investigation of amethod of reducing the starting current of 400 Hz squirrel cageinduction motors using saturistors'. RAE Technical Report78048, Royal Aircraft Establishment, Farnborough, England,1978

11 ATKINSON, L.G., and COX, J.W.R.: 'An induction motor testfacility', ibid., Report 78152,1978

1.2 COX, J.W.R.: 'Measurement of the magnetic properties ofpermanent magnet materials', ibid., Report 78047,1978

348 IEE PROC, Vol. 127, Pt. B, No. 6, NOVEMBER 1980