induction motors versus permanent-magnet …people.tamu.edu/~panoskak/yliko/journals/induction...

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 61, NO. 8, AUGUST 2014 4315

Induction Motors Versus Permanent-MagnetActuators for Aerospace Applications

Panagiotis E. Kakosimos, Student Member, IEEE, Athanasios G. Sarigiannidis, Student Member, IEEE,Minos E. Beniakar, Student Member, IEEE, Antonios G. Kladas, Senior Member, IEEE, and

Chris Gerada, Member, IEEE

Abstract—This paper introduces a comparative study on thedesign of aerospace actuators concerning induction motor andpermanent-magnet motor technologies. In the analysis under-taken, the two candidate configurations are evaluated in terms ofboth their electromagnetic and thermal behaviors in a combinedmanner. On a first step, the basic dimensioning of the actuatorsand their fundamental operational characteristics are determinedvia a time-stepping finite-element analysis. The consideration ofthe thermal robustness of the proposed motor configurations isintegrated in the design procedure through the appropriate han-dling of their respective constraints. As a result, all comparisonsare carried out on a common thermal evacuation basis. On asecond step, a single objective optimization procedure is employed,considering several performance and efficiency indexes using ap-propriate weights. Manufacturing and construction-related costsfor both investigated topologies are considered by employing spe-cific penalty functions. The impact of the utilized materials isalso examined. The resultant motor designs have been validatedthrough manufactured prototypes, illustrating their suitability foraerospace actuation.

Index Terms—Actuators, aerospace engineering, finite-elementmethods (FEMs), geometry optimization, induction motors (IMs),machine design, permanent-magnet motors (PMMs).

NOMENCLATURE

v Magnetic reluctivity.A Magnetic vector potential.JS Source current density.σ Electrical conductivity.V Electric potential.HC Coercive force of the permanent magnet.R Phase resistance.L End-winding inductance.is Stator current.vs Supply voltage.l Active motor length.nc Number of conductors.

Manuscript received January 30, 2013; revised April 26, 2013 and June 28,2013; accepted July 6, 2013. Date of publication July 24, 2013; date ofcurrent version February 7, 2014. This work was supported by the EuropeanCommission in the framework of the “Clean Sky” Programme, Topic Nbr: JTI-CS-2009-1-SGO-02-010, under grant agreement 255811EMAS.

P. E. Kakosimos, A. G. Sarigiannidis, M. E. Beniakar, and A. G. Kladas arewith the Laboratory of Electrical Machines and Power Electronics, Departmentof Electrical and Computer Engineering, National Technical University ofAthens, 15780 Athens, Greece (e-mail: [email protected]).

C. Gerada is with the Department of Electrical and Electronic Engineering,The University of Nottingham, Nottingham, NG7 2RD, U.K.

Color versions of one or more of the figures in this paper are available onlineat http://ieeexplore.ieee.org.

Digital Object Identifier 10.1109/TIE.2013.2274425

S Total conductor surface.S+ Total surface of conductors with positive phase

current.S− Total surface of conductors with negative phase

current.Ft Tangential component of the magnetic force.Pt Mean tangential pressure.Sg Air-gap surface.T Electromagnetic torque.Rr Rotor radius.Bn Normal magnetic flux density component.Bt Tangential magnetic flux density component.θ Angular coordinate.σm Radial magnetic force density.Xk Design variable vector.Ltooth Length of the stator teeth.Wt1 Width of the stator teeth.Wt2 Width of the stator tooth tip.X0 Vector of initial values of the design variables.G1,...,5 Weight coefficients.Tm Mean electromagnetic torque.Tr Electromagnetic torque ripple.An Third and fifth normalized amplitude sum.Pc Copper losses.g Problem constraints.C1 Term associated with the winding fill factor.C2 Term associated with the slot shape.C Total technical cost.k Ratio of end-ring resistance to bar resistance.θm Mechanical angle.VDC DC bus voltage.Iabc Measured three-phase line currents.v∗q q-axis reference voltage.v∗d d-axis reference voltage.V ∗abc Three-phase reference voltages.

I. INTRODUCTION

THE LATEST general direction in the industry of electricmotors is biased toward the better utilization of rare-earth

materials. This tendency is focused on two main alternativestrategies, i.e., the complete substitution of existing and en-trenched technologies or their adaptation to the new topologytrends through their systematic refinement emphasizing ontheir operating efficiency increase. That technological transi-tion combined with the current advances in permanent-magnet

0278-0046 © 2013 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission.See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.

4316 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 61, NO. 8, AUGUST 2014

technologies renders the extensive use of rare-earth magnetsfeasible, even under severe thermal environmental conditions[1]. Permanent-magnet motors (PMMs) and PM-assisted mo-tors exhibit numerous advantages and constitute an attractivesolution for almost every particular application. However, aclear indication toward the potential limitation on the usage ofinduction motors (IMs) and their eventual replacement has notyet been established. This is more intensely evident in industrialapplications, such as pulp or paper/forest product industry,where IMs are traditionally the dominant technology [2]. Onthe contrary, in aerospace actuation applications, an increasingtendency toward the integration of PMM topologies is apparent,boosting their domination rates. This is basically attributed tothe fact that they satisfy the spatial limitations and constraints,dictated by the strict nature of such applications. Nevertheless,the choice of the more advantageous between the two candidatemotor configurations is application specific, as they both benefitfrom their particular operating characteristics [2]–[4].

More specifically, PMMs feature crucial inherent advantagesover the alternative motor topologies [5], [6]. Consequently,they have attracted a rapidly growing interest within the sci-entific community [7], especially for high-power-density actua-tion applications highlighting the necessity for their thoroughinvestigation [8]. In aerospace applications, along with thedemand for high performance, safety and reliability issues arise.The extremely strict nature of the specifications, both opera-tional and spatial, dictates the necessity for further investigationof their operation [9], [10]. The most important advantages ofPMMs lay on the fact that the PMs constitute an independentand strong excitation system. This feature enables substantialoverloading of the motor while providing higher torque densityvalues. The fact that no electromagnetic excitation system isemployed additionally enhances their transient behavior, whilesimplification and maintenance are also two significant ben-eficial factors. The aforementioned advantages have led PMmotors to be considered a viable and attractive solution [4],[11], [12]. On the other hand, their main disadvantages concerntheir operation under fault conditions. Their serious thermalconstraints under fault conditions, their limited fault tolerancecapabilities, the PM retention and demagnetization issues, andthe need for high braking torque are major bottlenecks againsttheir widespread usage [13].

The predominant motor technology in such applications formany years has been the cage IMs. Their superior dynamicbehavior in conjunction with their brushless nature, which en-ables their operation without the presence of commutator or sliprings, renders them suitable for high-performance controlledoperation in electric drive applications [14]–[16]. The dramaticadvance in the area of power electronics and automatic controltechnologies has significantly contributed to their establishmentas standard motors in electrical drives, when field-weakeningoperation in high speeds is required. However, the IM tech-nology exhibits numerous disadvantages as well, both con-structional and operational. For example, their relatively smallair-gap length and their inferiority against the synchronousmotors in terms of overall efficiency and power factor aremajor drawbacks. Great scientific effort has been made overthe past years in order to limit the impact of the aforementioned

Fig. 1. Three-dimensional actuator representation models. (a) 24-slot–28-polePM machine. (b) Eight-pole induction machine.

disadvantages. Advancements in the area of motor design andmotor drives have facilitated cost-efficient motor operation overa wide speed range without compromising their efficiency andreliability.

In this paper, the design process and the comparison in termsof operational characteristics of two motor topologies, a PMMand an IM, satisfying the specifications of a specific aerospaceapplication, are presented. Two distinct and extremely differentoperating conditions are considered for both motor topologiesaccounting also for emergency operation. In order to determinethe fundamental actuator operating characteristics and the pre-cise stator and rotor geometries, finite-element (FE) models areemployed, integrated in an advanced optimization routine. Thesevere environmental conditions in aerospace applications arealso taken into consideration during the design process. Asthe impact of the ambient temperature rise is of paramountimportance, extensive thermal analysis has been conducted inorder to satisfy the equal thermal evacuation basis criterion foreach configuration. The two candidate actuator configurationsare illustrated in Fig. 1. On a next step, manufacturing andmaterial-utilization-related issues are further investigated. Theanalysis is mainly focused on the formation of the end-ringsmostly concerning the IM topology. The implemented designmethodology has been validated through the manufacturing andexperimental testing of two prototypes, a PMM and an IM,respectively, while for the case of the IM, two different end-ring formations have been tested.

II. ACTUATOR SPECIFICATIONS

Certain geometrical and electromechanical specifications,dictated by the application nature, have to be satisfied by theactuator geometry. The main actuator specifications and dimen-sions are tabulated in Table I for the two different operationmodes. The actuator output torque capability under normaloperation is about five times less compared to the extrememode of operation. A respective increase in rotor speed is alsoevident. Such a duality in the motor specifications is quitetypical in aerospace applications. This is mainly attributed tosafety reasons to prevent undesirable effects of disturbanceson the actuator operation during certain maneuvers. It is worthpointing out that the speed value, as shown in Table I, is theactual machine mechanical speed without the usage of step-down gear. The main restrictions are now set by the controlsystem capacity, dc voltage level, and switching frequency.

In the process of determining the basic motor dimensions,a preliminary analysis, based on analytical formulas, has been

KAKOSIMOS et al.: IMs VERSUS PERMANENT-MAGNET ACTUATORS FOR AEROSPACE APPLICATIONS 4317

TABLE IMAIN ACTUATOR SPECIFICATIONS AND DIMENSIONS

carried out for the two actuator types. For the need of theanalysis, the fundamental design principles have been initiallyconsidered. However, each machine type required special at-tention during the design procedure regarding the satisfactionof the listed requirements and both the respective geometricaland spatial limitations.

III. MODELING AND DESIGN CONSIDERATIONS

The time-dependent magnetic field is governed by the well-known Maxwell’s equation in the form of the magnetic vectorpotential

∇× v∇×A = Js− σdA

dt− σ∇V +∇×Hc. (1)

The equation describing the interaction of the magnetic fieldwith the applied machine terminal voltage is of the followingform:

vs=R · is+L · disdt

+l · ncS

·

⎛⎝∫ ∫

S+

∂A

∂tdS −

∫ ∫S−

∂A

∂tdS

⎞⎠ .

(2)

On a first step, in order to specify the main motor designcharacteristics, a preliminary design analysis has been imple-mented. In particular, for the IM, conventional design principleshave been followed, mainly focusing on the efficiency and theoutput torque. The produced electromagnetic torque can becalculated from the mean tangential pressure of the magneticforces on the gap of the machine, which is expressed by meansof Maxwell stress tensor as follows:

Ft =Pt · Sg (3)

T =Rr · F → T = 2πR2r · l · Pt (4)

Pt =1

μ0

∮Bn ·Bt · dc. (5)

Different topologies and winding configurations suitable forthe specific application have been investigated and compared inorder to identify the one exhibiting the best performance withrespect to the considered specifications. The nonoverlappingwinding investigation has shown that, under such circumstancesand topology, a nonoverlapping winding is not favorable andeffective for the IM; thus, the classical winding layout hasbeen considered. In order for the IM to meet the requirements,the whole available space, maximum length and diameter, isrequired to be covered, taking into consideration the best perfor-

Fig. 2. Block scheme of the adopted optimization algorithm [22].

mance. In this case of inverter-fed motor, minimization of rotorresistance is desirable. This constraint has been considered atthe design stage of the proposed motor configuration. Finally, asensitivity analysis of the main motor parameters of the selectedconfiguration has been performed in order to define the detailedmotor design shown in Fig. 1.

Contrary to the IM case, as far as the PMM is concerned,a systematized analysis of the most popular alternative motorgeometries and of the most suitable winding configurations, forthis specific application, has been undertaken. The analysis wasmainly focused on the key points of performance, efficiency andreliability. The active machine length for the case of the PMmotor has been finally set to 40 mm, which is one third of thetotal available length. Moreover, both overlapping and nonover-lapping winding configurations have been considered, and theirpotential suitability has been evaluated. The nonoverlappingsingle-layer winding was considered the most favorable optiondue to its underlying advantages of lower copper losses andlower torque ripple [19]. Owing to the nonregularly distributedmagnetic forces along the air-gap, a unidirectional pulling forcethat rotates with time and generates noise and vibration inthe motor is present under such winding configurations [19].Radial magnetic forces are computed by Maxwell stress tensoras follows:

σm(θ, t) =1

2μo·(B2

n(θ, t)−B2t (θ, t)

). (6)

By the end of the preliminary design process and the se-lection of the final winding layout, a particular evolutionaryoptimization algorithm has been employed. The algorithmfacilitated the comparative approach on the stator geometryoptimization of surface PM machines involving fractional-slotconcentrated-winding configurations, considering the variationof key stator design variables. More specifically, a zero-orderdirect-search Rosenbrock-based optimization algorithm is in-troduced in order to minimize an application-specific penaltyfunction through a sequential unconstrained minimization tech-nique [17]. Fig. 2 illustrates the block diagram of the proposedoptimization procedure. The procedure is based on the interac-tion between a 2-D FEM parametric model and the optimizerblock.


Fig. 3. Torque profile versus slip frequency considering constant stator currentdensities. The dotted line corresponds to 28 and 20 A/mm2 for machine lengthsequal to 40 and 60 mm, respectively.

The penalty function at the beginning of the kth iterationshown in the block scheme is given in [17]

P k(Xk) =G1 ·Tm(Xk)

Tm(X0)+G2 ·

Tr(Xk)

Tr(X0)+G3 ·An(Xk)

+G4 ·Tm(Xk)√PCu(Xk)

·(

Tm(X0)√PCu(X0)

)−1

+G5 ·C(Xk)

C(X0)+Rk ·

∑1/ (gi(Xk)) (7)

where Xk is equal to [Lktooth,W

kt1,W

kt2] and gi(Xk) ≥ 0, i =

1, 2, 3. The total technical cost C is expressed as

Ck(Xk) =C1 ·max

(Lktooth

L,W k

t2

W

)

+ C2 ·max

(W k

t1

W kt2

,W k

t2

Lktooth

)

=

4∑i=1

3∑j=1

aij exp(− ((xi − bij)/cij)

2). (8)

IV. IM ACTUATOR CASE

The absence of excitation on the rotor of IMs is its mostimportant characteristic regarding the establishment of the ro-tor magnetic field. To achieve adequate torque production, itdemands higher current density to flow into the stator. As aresult, it is reasonable that, under these circumstances, the IMrequires larger envelope dimensions than a PM machine inorder to achieve the equal output torque under the same statorcurrent density levels. In Fig. 3, the output torque versus slipfrequency is illustrated for different stator current densities andactive motor lengths. For all cases, an actual slot fill factor equalto 0.5 and constant currents have been considered.

From Fig. 3, it can be derived that the eight-pole IMconfiguration meets the prescribed specification values at themaximum available machine length of 120 mm. It should benoted that a thermally acceptable current for the normal oper-

Fig. 4. Power factor versus speed under (a) normal and (b) extreme operatingconditions.

Fig. 5. Power loss density distribution under (left) normal and (right) extremeoperating conditions.

Fig. 6. Induced and terminal phase voltages under (left) normal and (right)extreme operating conditions.

ating conditions is about 6 A/mm2. Power factor versus speedunder normal and extreme operating conditions is illustrated inFig. 4. The motor nominal operating conditions according tothe design procedure involve low speed operating range shownin Fig. 4(a) presenting appropriate power factor values. Incounterparts, the low power factors encountered in high speedranges corresponding to extreme operating conditions are notcompromising the overall actuator behavior as such conditionsare rarely implemented in case of emergencies.

Fig. 5 shows the power loss density distribution for theIM, considering both modes of operation. The actual suppliedcurrent densities are 6 and 13 A/mm2, respectively. For thespecific current density values, the required torque levels areachieved, and the stator and rotor core materials are optimallyutilized, i.e., they operate at the knee of the lamination B–Hcurve, which is a design process criterion. The preservation ofsaturation at low levels leads to increased sinusoidality of theback EMF waveform. Fig. 6 illustrates the unsaturated core for


TABLE IIROTOR SLOT CHARACTERISTICS

the normal operation. For this case, the EMF is almost sinu-soidal. The EMF for the extreme operating conditions exhibitsinsignificant signs of harmonic distortion, also indicating theoperation at the knee of the magnetizing curve.

The reduction of parasitic torque, additional losses, noise,and vibration is crucial for the selection of the stator androtor numbers of slots. However, the construction of multipoleand small-sized machines significantly reduces the availableand feasible stator and rotor slot combinations. To evaluatethe most appropriate combination between the stator and rotornumbers of slots, several possible combinations have beeninvestigated. Due to manufacturing limitations, 24 stator slotswere considered. The cross sections of the copper bars and therotor bore have been kept constant in order not to compromisethe comparability of the results under the equal output torque.Table II presents the torque ripple under normal operation offive distinctive cases that were examined and indicates theimpact of the increase of the number of rotor slots on theproduced torque.

When the number of rotor slots increases significantly, therotor teeth are heavily saturated, resulting in an increase in thetorque ripple. Furthermore, due to the limited available space,a compromise between manufacturing feasibility and the theo-retical optimum approach should be made. From Table II, it isevident that the 30 slots in rotor benefit from the increased slotnumber constituting also an adequately satisfactory version.Another important factor that has to be mandatorily taken intoaccount is the rotor moment of inertia, as it directly improvesthe motor dynamic performance. The existence of copper inthe rotor slots deteriorates the machine behavior during speedvariations. The IM rotor involves a moment of inertia slightlylower than 5 · 10−4 kg · m2.

An important issue in achieving high efficiency in the IMdesign case concerns low rotor resistance enabled through arccopper welding formation of cage end-rings. However, in thereduced rotor geometries favored in aerospace applications, it isdifficult to implement such manufacturing procedures, resultingin usual end-ring resistance values on the order of 25% of therotor resistance. That is why silver joint end-ring formation hasto be considered, involving rather greater end-ring resistancevalues [14]. The use of die-cast rotor cage could reduce thehigh end-ring resistance values; however, it is quite demandingto be implemented in such space-limited applications. In orderto take into account this constraint, the variation of end-ringresistance impact on the motor performance has been investi-gated at the design stage. The respective torque–speed char-acteristics for ideal end-ring hypothesis (zero resistance) anddifferent end-ring resistance values have been investigated atthe design stage and are shown in Fig. 7. This figure illustrates

Fig. 7. Torque versus slip frequency for different estimated rotor resistances.

TABLE IIIWINDING LAYOUT CHARACTERISTICS

the importance of achieving low end-ring resistance values [18].The term k represents the ratio of end-ring resistance to barresistance. For the arc copper welding and silver joint end-ringformations, actual ratios of about k1 = 1.5 and k2 = 2.5 havebeen considered, respectively.

V. PM ACTUATOR CASE

Regarding the winding configuration for the PM actuator, thenonoverlapping single-layer winding was adopted as it has mostof the benefits that we can get from the investigated windingtopologies. Even though the alternate teeth wound winding isconsidered a more favorable choice, it presents some seriousdrawbacks. The number of poles and slots has to be chosenwith extreme care in order to maintain high values for thewinding factor and to minimize the values of the torque rippleand the rotor losses, respectively, which incur a significantincrease in machine noise [19]. In order to identify the favorablewinding layout and the numbers of rotor poles and stator slots,different winding topologies have been examined by using thetime-stepping FE analysis in conjunction with the optimizationprocedure.

Table III summarizes the main characteristics of some of theanalyzed winding layouts while considering identical embraceand thickness of the magnet configuration under the samevoltage supply.

The MMF harmonic component in PM machines with fewerstator slots than rotor poles that contributes in the mean torqueproduction is not the fundamental but a higher harmoniccomponent of the same order as the number of pole pairs.The 24-stator-slot and 28-rotor-pole single-layer fractional-slotmachine configuration has been selected for this particular


Fig. 8. Radial force density distribution for a rotating period.

Fig. 9. Power loss density distribution under (left) normal and (right) extremeoperating conditions.

Fig. 10. Induced and terminal phase voltages under (a) normal and(b) extreme conditions.

application and specifications, exhibiting low torque ripple withhigh winding factor and, subsequently, higher EMF capacity.

The radial force density distribution for the selected topologyunder nominal operating conditions, calculated by (6), versusthe rotor angle in the air-gap is illustrated in Fig. 8. It may benoted that the resulting radial magnetic force per one rotatingperiod is negligible and approaches 0.47 N. Such a low totalradial force is a result of the symmetry of the applied windinglayout. Fig. 9 shows the power loss density distribution of thePM under both operations considered.

The capability of the motor to operate under overloadingconditions is of high importance, enabling us to meet the re-quirements of the extreme loading condition. The torque angleunder normal operation is approximately equal to 30◦, andthe EMF is absolutely sinusoidal, which is mainly due to theunsaturated stator and rotor core materials. Fig. 10 depicts theinduced and terminal phase voltages considering a sinusoidalvoltage supply at the machine terminals. Under the extremeoperating condition, the maximum torque at maximum speedis significantly higher than the specified torque for the normaloperation.

The stator and rotor laminations are driven to high saturationlevels, presenting mean flux density in the stator tooth equal to2 T. Subsequently, the induced EMF is affected by the highly

Fig. 11. Fourier analysis of the (a) induced and (b) terminal phase voltages.

saturated laminations, presenting a third harmonic componentequal to 50.8% of the fundamental component. The inducedEMF and terminal voltage as well as the harmonic contentare illustrated in Figs. 10 and 11, respectively. In high-speedoperating conditions, PMM topologies are expected to presentreduced field-weakening capability [20]. The action of the field-weakening procedure is to lower the influence of the PM fluxlinkage on the resulting air-gap flux; thus, it is obvious thatthis constitutes a PMM drawback, especially when comparedto IMs which feature the absence of excitation. In this specificapplication, the highest speed foreseen is the nominal one; how-ever, the possibility of field weakening for inverter protectionunder small overspeed, which could be encountered due tomechanical load abrupt variation, is desirable.

VI. RESULTS AND DISCUSSION

IM drawbacks are derived primarily from the absence ofindependent electromagnetic excitation in the rotor. As it hasalready been mentioned, a higher current density in the stator isneeded in order to obtain equal output torque levels due to theestablishment of the rotor magnetic field. Therefore, overall IMsituation deteriorates compared to the PM machine.

From the preceded analysis, IM required the total availablelength, while the PM actuator demanded one third of the totallength in order to satisfy the specifications. However, the PMmachine, under the extreme operating condition, presents statorand rotor core heavy saturation in order to achieve the desirableoutput torque, resulting in distorted induced EMF and currents.Unlike the PM machine, the IM core saturation level is lowerdue to the fact that the magnetic field in the rotor has alreadybeen established, thus having the capability of being furtherloaded. It should be noted that geometry optimization has beencarried out for the normal operation for the two motor types. Forthe normal operation, the main PM and IM characteristics aretabulated in Table IV. It is worth mentioning that the appliedcurrent density of about 4 A/mm2 is too low for aerospaceactuators. In practice, a forced air or oil cooling is used, andthe current density is higher. However, the extreme-mode oper-ation necessitates higher current densities, setting the operationlimits and leading to lower densities under normal operation.Nevertheless, in the present study, natural air convection hasbeen considered.

Moreover, in aerospace actuation applications, the operationof more than one machine on the same shaft is a very common


TABLE IVPM AND INDUCTION ACTUATOR CHARACTERISTICS

Fig. 12. Instantaneous torque under three-phase short-circuit fault consideringconstant speed load. (a) PM machine. (b) IM.

practice, mainly for safety and reliability reasons, thus con-tributing to a fault-tolerant system. As a result, when a machinefault occurs, the remaining machines have to operate over-loaded in order to substitute the faulty machine operation andnot to be interrupted. The absence of rotor excitation favors theIMs, which exhibit zero braking torque under short circuit fault.Fig. 12 indicates the main disadvantage of the PMMs, whichaccounts for their thermal constraints under fault conditions.When a three-phase short circuit is imposed at 20 ms, the PMmachine continues to contribute with high braking torque equalto 0.5 N · m after 40 ms, while on the contrary, the IM presentsalmost zero braking torque.

The final choice of actuator type strongly depends on thespecific application demands and particularities. Both candidateactuator types exhibit major differences and correspondingadvantages. It should be noted that, for this specific aerospaceapplication, the equal thermal evacuation criterion has beenconsidered and has strongly influenced the results. As theextreme environmental conditions in aerospace applicationscontribute to the temperature rise, taking the thermal constraintsinto consideration is very important. Furthermore, another im-portant factor that should be mentioned is the fact that therotor temperature is much more crucial in the PMM than inthe IM due to the inherent retention issues of the PMs. Apotential temperature increase incurs a dramatic decrease in theremanent flux density of the PMs, which results in a significantdecrease in the air-gap flux density and consequently in thedeterioration of the motor’s overall performance. Fig. 13 showsthe temperature variation during normal operating conditionsfor the two configurations. In particular, the minimum andmaximum temperatures correspond to the rotor and the stator,respectively. The higher stator temperature values are caused by

Fig. 13. IM and PMM temperature variations under normal operatingconditions.

Fig. 14. (a) Power and control unit. (b) PMM. (c) IM.

the fact that it bears the main burden due to the winding estab-lishment. Maintaining approximately equal heat flux levels foreach configuration, the PMM topology seems to be favored asit involves an important reduction in size and weight with thesame thermal constraints.

VII. EXPERIMENTAL VALIDATION

In order to experimentally demonstrate the developedmethodology, the experimental setup shown in Fig. 14 has beenemployed. The experimental setup comprises the control andpower unit, and an electromechanical brake so as to applyvariable mechanical loads to both prototypes.

The constructed multipurpose control unit receives and pro-cesses signals of the voltage and current transducers, and out-puts the desirable pulses. Fig. 15 shows signal routing from andto the main component of the digital signal processor (DSP).The employed DSP presents an adequate computational capa-bility for this specific application, supplying the IM and PMMprototypes with a sinusoidal pulsewidth modulation (SPWM)voltage employing a carrier frequency of 10 kHz. In order tocapture the presented data in the following section, a TiePieoscilloscope and a National Instrument PXI (16 b and 4 MS/s)are used.


Fig. 15. Overall system configuration.

Fig. 16. (a) IM stator. End-ring formation by using (b) arc welding or(c) silver joint.

Fig. 17. Measured phase voltage and current of IM under no-load operationat 180 r/min.

Validating the design methodology, PMM and IM prototypeshave been manufactured, and in the case of the IM, two differentend-ring formations are also investigated. Fig. 16 shows the IMconfiguration, as well as both manufactured rotor prototypeswith end-ring formation by using either arc welding or silverjoint. The stator laminations employed to stator are ThyssenM 330-35 A/35JN230. For the PMM neodymium (Nd)-iron(Fe)-boron (B) NMX41-EH has been adopted as PM material.The measured phase voltage and current of IM under no-load operation at 180 r/min are depicted in Fig. 17. Althoughunder constant voltage supply the magnetizing current does notsignificantly depend on load, in this case of inverter supply, themagnetizing current of IM is low enough under no-load condi-tion. Thus, the machine presents low power losses associatedwith the core due to hysteresis and eddy currents.

Fig. 18. Measured phase voltage and respective harmonic content of IM at6000 r/min.

Fig. 19. Measured torque under constant current density as a function of slipfrequency for both end-ring formations. (a) Arc welding. (b) Silver joint.

In addition, the measured voltage and harmonic content ofthe IM at the highest speed of 6000 r/min are illustrated inFig. 18. Fig. 19 evidences the claims discussed in Section IVand the simulation results shown in Fig. 7 during the designprocess. Despite the use of silver end joint, end-ring formationinvolves greater end-ring resistance values; it is a preferablechoice in cases where machine dimensioning is defined by strictspace limitations. Torque versus slip frequency waveforms forboth rotor configurations are within the bounds set during thedesign phase, allowing safe selection of materials and manu-facturing process to be followed. It should be noted here thatthe upper and lower bounds correspond to the case where thesimulated end-ring resistance is comparable to bar resistance(k = 1.5) and to greater values (k > 1). Significant variationsin the produced torque for different end-ring formations andspeeds are observed in Fig. 19 even in lower slip frequencyvalues, when the motor is inverter driven.

Fig. 20 shows the stator and rotor of the PMM configuration.It is worth noticing here that the magnet retention issue is faced


Fig. 20. PMM configuration. (a) Rotor. (b) Stator.

Fig. 21. Comparison of measured and simulated phase voltages of PMMunder no-load operation.

Fig. 22. Comparison of simulated and measured torque–speed characteristics.

by employing glass fiber support of PMs. Behaving satisfacto-rily even under 6000 r/min, the voltage at the machine terminalspresents a linear behavior as illustrated in Fig. 21 since themachine is quite far from the knee of the lamination B–Hcurve. Contrarily, it is expected that, according to the simulationresults of under 6000 r/min and full-load condition, the machinelaminations are under heavy saturation. The simulated andmeasured torque–speed characteristics for the PMM are shownin Fig. 22, where the simulated results are in good agreementwith the measured ones, especially in the lower speed range.The measured phase voltage and current of PMM under no-load operation at 180 r/min are depicted in Fig. 23, where thecurrent is significantly higher than that of the IM.

Furthermore, the measured voltage and harmonic content ofthe PMM at the highest speed of 6000 r/min are presented inFig. 24. As previously mentioned, for the SPWM technique im-plemented by the inverter in PMM and IM, a carrier frequencyequal to 10 kHz is utilized. The maximum fundamental electricfrequency at the highest speed of 6000 r/min for the PMM is1.4 kHz, while for the IM, it is 400 Hz. Therefore, the selectedcarrier frequency of the SPWM is appropriate in driving bothmotors adequately for both operating conditions, introducinglow harmonic content in the stator current, due to the filteringaction of the winding inductance.

Fig. 23. Measured phase voltage and current of PMM under no-load operationat 180 r/min.

Fig. 24. Measured phase voltage and respective harmonic content of PMM at6000 r/min.

VIII. CONCLUSION

In this paper, a comparative analysis between PMM and IMaerospace actuators has been carried out. Different combina-tions of rotor–stator slots for the IM and different winding lay-outs for the PMM have been modeled, and their performanceshave been compared. The optimal IM configuration, featuredby the absence of rotor magnetic field, required larger dimen-sioning and weight than the respective PMM one for the samethermal evacuation ratio. However, the IM case presented lowerbraking torque and short-circuit currents with respect to thePMM case, leading the final choice of the actuator technologyto an overall system-consideration-dependent problem.

In the considered application, the two rival technologiespresented different drawbacks: the IM configuration, due tothe rotor dimensional constraints, has difficulty in achievingacceptably reduced end-ring resistance values compared tobar resistance, leading to motor efficiency compromise at lowspeed ranges. On the other hand, the PM configuration, involv-ing surface PM structure, presented increased losses in high-speed operating conditions due to a reduced field-weakeningcapability. Consequently, through the analysis developed inthis paper validated by experimental results, the proposed IMand PMM configurations present complementary advantages,the former concerning fault tolerance characteristics and thelater concerning power density issues. Thus, according to thispaper’s findings, on a strict motor basis, the PM topology is


TABLE VSTATOR LAMINATION LOSS PROFILE AND PM CHARACTERISTICS

favored in the considered aerospace application involving twodistinct operating speed ranges.

APPENDIX

The loss profile of the Thyssen M 330-35 A/35JN230 lami-nation and the basic characteristics of the NMX41-EH PMs aretabulated in Table V.

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Panagiotis E. Kakosimos (S’12) received theB.Eng. and M.Eng. degrees in electrical and com-puter engineering from Aristotle University ofThessaloniki, Thessaloniki, Greece, in 2009 and thePh.D. degree from the Department of Electrical andComputer Engineering, National Technical Univer-sity of Athens, Athens, Greece, in 2013.

His current research involves power gener-ation from renewable-energy sources, industrialdrives, and electric machine design optimization foraerospace and electric vehicle applications.

Dr. Kakosimos is a Registered Professional Engineer in Greece.

Athanasios G. Sarigiannidis (S’13) received theDiploma in electrical and computer engineering fromthe University of Patras, Patra, Greece, in 2010. Heis currently working toward the Ph.D. degree in theLaboratory of Electrical Machines and Power Elec-tronics, Department of Electrical and Computer En-gineering, National Technical University of Athens,Athens, Greece.

His current research involves permanent-magnetmotor modeling, design, and control for tractionapplications, with emphasis on field-weakening tech-

niques for electric motor optimal efficiency control.Mr. Sarigiannidis is a member of the Technical Chamber of Greece.

Minos E. Beniakar (S’08) received the B.Eng. andM.Eng. degrees in electrical and computer engi-neering from the National Technical University ofAthens, Athens, Greece, in 2008, where he is cur-rently working toward the Ph.D. degree in the Labo-ratory of Electrical Machines and Power Electronics,Department of Electrical and Computer Engineering.

His current research involves design optimizationof electric motors for aerospace and electric tractionapplications and industrial drives.

Mr. Beniakar is a Registered Professional Engi-neer in Greece.


Antonios G. Kladas (SM’10) was born in Greecein 1959. He received the Diploma in electricalengineering from Aristotle University of Thessa-loniki, Thessaloniki, Greece, in 1982 and the D.E.A.and Ph.D. degrees from the University of Pierreand Marie Curie (Paris 6), Paris, France, in 1983and 1987, respectively.

From 1984 to 1989, he was an Associate Assis-tant with the University of Pierre and Marie Curie.From 1991 to 1996, he was with the Public PowerCorporation of Greece. Since 1996, he has been with

the Department of Electrical and Computer Engineering, National TechnicalUniversity of Athens, Athens, Greece, where he is currently a Professor.His research interests include transformer and electric machine modeling anddesign, and analysis of generating units by renewable-energy sources andindustrial drives.

Dr. Kladas is member of the Technical Chamber of Greece.

Chris Gerada (M’05) received the Ph.D. degreein numerical modeling of electrical machines fromThe University of Nottingham, Nottingham, U.K.,in 2005.

He was a Researcher with The University ofNottingham, where he worked on high-performanceelectrical drives and on the design and modeling ofelectromagnetic actuators for aerospace applications.He is currently an Associate Professor of electricalmachines with the Power Electronics, Machines andControl Group, Department of Electrical and Elec-

tronic Engineering, The University of Nottingham. He is also a Project Managerwith the GE Aviation Strategic Partnership, Nottingham. His research interestsinclude high-performance electric drives and machines.

induction motors versus permanent-magnet …people.tamu.edu/~panoskak/yliko/journals/induction...

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