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278 IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 19, NO. 2, JUNE 2004

Performance Optimization of Induction MotorsDuring Voltage-Controlled Soft Starting

Gürkan Zenginobuz , Student Member, IEEE , Isik Cadirci , Member, IEEE , Muammer Ermis , Member, IEEE , andCüneyt Barlak

Abstract— This paper deals with the performance optimizationof medium/high-power induction motors during soft starting byeliminating the supply frequency torque pulsations, and by keepingthe line current constant at the preset value. Starting torque pulsa-tions are eliminated by triggering back-to-back-connected thyris-tors at proper points on the first supply voltage cycle. Line cur-rent during starting can be kept constant at any preset value by asimple strategy composed of successive cosinusoidal and constantfunction segments of the triggering angle. These strategies are im-plemented by the use of an 8-b microcontroller. Transient perfor-mance analyses of the system are carried out by means of a hybridABC/dq machine model which takes into account the three-phase,

two-phase, and disconnected modes of operation in terms of actualstatorvariables. Experimental results obtained on a custom-designtest bed are found to be in good agreement with the theoreticalones.

Index Terms— Induction motor, soft starter, torque pulsations.

I. INTRODUCTION

D IRECT-ONLINE starting of large ac motors may presentdifficulties for the motor itself, and the loads supplied

from the common coupling point because of the voltage dipsin the supply during starting, especially in the case of a weakpower system [1]–[4]. An uncontrolled starting may cause a

trip in either the overload or the undervoltage relay, resulting instarting failure. This is troublesome for field engineers since themotor cannot be reenergized until it cools down to an allowabletemperature, in a long time period. Furthermore, the number of starts per day is limited to only a few attempts. Therefore, cur-rent and torque profiles of the motor during starting are to becarefully tailored according to the needs of the load [4], [5].

AC motor starters employing power semiconductors arebeing increasingly used to replace electromagnetic line starters,and conventional reduced voltage starters because of theircontrolled soft-starting capability with limited starting current.These may be classified in two groups:

Manuscript received January 6, 2003. This work was supported by the Elec-trical, Electronics, and Informatic Research Group of Scientific and TechnicalResearch Council of Turkey under Project EEEAG-65.

G. Zenginobuz is with the Information Technologies and ElectronicsResearch Institute, Scientific and Technical Research Council of Turkey(TÜBITAK-Bilten), Ankara, TR06531, Turkey (e-mail: [email protected]).

I. Cadirci and C. Barlak are with the Department ofElectrical and ElectronicsEngineering, Hacettepe University (HU), Ankara TR06532, Turkey (e-mail:[email protected]; [email protected]).

M. Ermis is with the Department of Electrical and Electronics Engineering,Middle East Technical University (METU), Ankara TR06531, Turkey (e-mail:[email protected]).

Digital Object Identifier 10.1109/TEC.2003.822292

i) back-to-back connected thyristor-based soft starters,which apply a reduced voltage to the motor;

ii) insulated gate bipolar transistor (IGBT) or gate turn-off thyristor (GTO)-based dc link converters, which give avariable-frequency output.

At the present time, pre-engineered versions of such acmotor starters are commercially available up to an operatingvoltage level of 15 kV l-l, and a power rating of 20 000 hp.With the present technology, the higher cost, complexity, andlarger volume occupied by the variable-frequency dc link

converter can be largely offset both by the high starting torquerequirement of “heavy to start” industrial loads, and the needfor speed control in a moderate range. Speed control in a widerrange makes necessary special or additional means of coolingfor the motor. Thyristorized soft starters, however, are cheap,simple, reliable, and occupy less volume, and hence, their useis a viable solution to the starting problem of medium-voltageand large ac motors, in applications where the starting torquerequirement of the load is not high.

Besides the developments and progress in commercial softstarter technology [5], [6], numerous attempts have been madeon the performance analyses and control techniques of three-phase induction motors fed from a thyristorized voltage-con-

troller [7]–[15]. Utilizing a dynamic function for the triggeringangle of thyristors in thevoltage controller proves to be a simpleandeffective way to improve the transientperformance [15].Byemploying a proper triggering function, the rate at which mainflux builds up has been decreased, and transient torque has beensmoothed.

An induction motor (IM) may produce severe pulsations inelectromechanical torque [16]–[21] depending upon the initialswitching instants of all three-phases to the supply, regardlessof the starting method: direct-online or soft starting. Minimiza-tion of the pulsating torque component is given in [16] and [21]for direct-online starting. The amount of electromagnetic torquepulsations reflected to the shaft depends on the parameters of the mechanical subsystem [18], [19]. These may cause shocksto the driven equipment, and damages in the mechanical systemcomponents, such as couplings and gears, in the long term.

In this paper, the performance of the IM during voltage-con-trolled soft starting has been optimized by eliminating thesupply frequency torque pulsations, and by keeping the linecurrent constant at the preset value, over the entire soft-startingperiod. The pulsating torque elimination strategy defined, andapplied in [16], [21] for direct-online starting of the IM hasbeen extended to cover all of the operating conditions of asoft starter, with back-to-back-connected thyristors on each of

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starting, a simple control strategy composed of successive cos-inusoidal and constant function segments of “ ” is employed,as described in Section IV. The soft starter developed in thiswork has some further features such as estimating initial firingangle value for the current setting, storing successful load andcurrent setting-dependent run-ups, etc. In the starting period,at any point on the acceleration profile, the IM will operate in

one of the following three operation states depending upon thepower factor of the motor, and starting current limit setting: 1)three-phase operation; 2) two-phase operation; and 3) no-phaseoperation (disconnected mode).

The analysis and design work of the system are carried out byusing hybrid models, in which the stator quantities are kept intheir actual form, whereas rotor quantities are transferred into areference frame fixed in the stator.

III. M ATHEMATICAL MODEL OF SYSTEM

The mathematicalmodelsused in the digital simulation of thesystemaregiven in this section. The numerical solution method,

and necessary transitions between various operation modes of the starter are described. At any time, the IM operates in one of the operation modes defined in Table I.

A. Three-Phase Model

Starting with the mathematical model of the IM in terms of ABC/abc axes quantities, all space angle, and hence, time varia-tions in inductances are eliminated by applying the well- knownthree-phase to two-phase, and commutator transformations (and ), only to the rotor side. This results in a hybrid math-ematical model in terms of ABC/dq0 axes quantities for whichthe rotor reference frame is fixed in the stator [ 21]. The hybridmodel is brought into the state space form for digital simulation,

as given in (1), shown at the bottomo of the page, and (3), shownon the next page.

B. Two-Phase Models

1) Mode 1: When only motor terminals A and B are con-nected to the supply, the two-phase model in (4) and (5), shownon the next page, is to be used. Since and , thebehavior of the motor is expressed in terms of the line-to-linevoltage .

TABLE IDEFINITION OF OPERATION MODES

2) Mode 2: This operation mode arises when motor termi-nals A and C are connected to the supply. By substitutingand into the mathematical model, and making thenecessary row operations, (6) and (7), shown on the next page,are obtained.

3) Mode 3: This operation mode arises when motor termi-nals B and C are connected to the supply. By substitutingand , the mathematical model in (8) and (9), shown onthe next page, is obtained.

C. No-Phase Operation (Disconnected Mode)

The mathematical model in (10) is used for the case wherenone of the motor terminals is connected to the supply. Since

, only the rotor voltage equations are to besolved in the transient-state

(10)

D. Numerical Solution Method To solve for the nonlinear first-order differential equation

sets given above, a fourth-order Runge –Kutta integrationalgorithm is used. To ensure the continuity of solutions when achange in the operation mode takes place, final values of theprevious mode are taken as initial values of the next mode.This first necessitates the detection of turn-off instant for theconducting thyristor in one of the three supply lines. This isachieved by the “Current Zero Detection Subroutine ” integrated

(1)

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ZENGINOBUZ et al. : OPTIMIZATION OF INDUCTION MOTORS DURING VOLTAGE-CONTROLLED SOFT STARTING 281

into the program. Forward or backward thyristor in one of thesupply lines turns off when its current ceases. “Current ZeroDetection Subroutine ” monitors continuously the current, andwhenever it takes a negative value at any numerical integrationstep “m,” the program terminates and returns to the previousintegration step “ ,” to detect the zero crossing point withan acceptable error, at a reduced time step. The next operation

mode can then be determined automatically by detecting theforward biased thyristors, which receive a firing pulse.

Possible transitions from old to new operation modes aregiven column wise in Tables II –VI. If the current of forward orbackward thyristor in any one of the three lines ceases beforethe antiparallel thyristor receives a firing pulse, possible opera-tion mode is either the two-phase or the no-phase operation. If a thyristor in one of the lines receives a firing pulse while theothers are not conducting, it cannot trigger into conduction untila reverse thyristor in one of the other two lines receives a firingpulse, resulting in no-phase operation. In Table II, transitionsfrom three-phase to two-phase operation are given. This occursif “ ” is greater than the power factor angle “ ” of the motor.However, in cases where , both the new and old operationmodes are three-phase operation. “ ” value depends on theoperating condition, especially on the shaft speed, and can bedetermined from the per-phase equivalent circuit at the steadystate. Transitions from two-phase operation modes (1 to 3) toany other are given in Tables III –V. In Tables II –VI, thyristors

TABLE IITRANSITIONS FROM MODE 0 TO OTHERS

TABLE IIITRANSITIONS FROM MODE 1 TO OTHERS

marked as active for no-phase operation are not in conductionstate, but they are only receiving firing pulses.

where

(2)

(3)

(4)

(5)

(6)

(7)

(8)

(9)

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TABLE IVTRANSITIONS FROM MODE 2 TO OTHERS

TABLE VTRANSITIONS FROM MODE 3 TO OTHERS

TABLE VITRANSITIONS FROM MODE 4 TO OTHERS

IV. R ESULTS

The firing angle control strategy, which yields a nearly per-fect starting current envelope, and the initial thyristor triggeringstrategy, which nearly eliminates the torque pulsations, will bedescribed in this section. These strategies are deduced from theresults of digital simulation studies, and are verified by experi-mental work. The proposed strategies are applicable with minormodifications, bothto the medium voltage squirrel-cage IM, andto the low-voltage ones.

A. Elimination of Torque Pulsations

The cause of pulsations in electromechanical torque at thesupply frequency is the uncontrolled switching of three motorphases to the supply on the first voltage cycle. Simultaneousswitching of the three motor phases always gives rise to a con-siderablepulsatingtorque component[ 16]. In solid-state startersfor medium-voltage IMs, the switching strategy reported in [ 21]nearly eliminates the torque pulsations. In direct-online starters,“ ” is zero; therefore, the mentioned switching strategy consti-tutes a special case of the torque pulsation elimination strategyapplicable to solid-state soft-starters. In Table VII, the initialswitching strategies that can be used in solid-state soft starters

TABLE VIITRIGGERING INSTANTS OF THYRISTORS ON THE FIRST CYCLE FOR

ELIMINATION OF TORQUE PULSATIONS

Fig. 2. Torque pulsation elimination strategy.

are given for both the continuous, and discontinuous line cur-rent cases. To eliminate the torque pulsations, the six thyris-tors in Fig. 1 should receive firing pulses in turn (in the num-bered sequence from 1 to 6), at prespecified points of the firstsupply voltage cycle, as illustrated in Fig. 2. Continuity of theline current waveform is largely dictated by the initial setting

of firing angle “ ,” which adjusts the starting current to thepreset value. For the discontinuous line current case, two dif-ferent initial switching strategies would arise, depending uponwhether “ ” is less or greater than the power factor angle of theIM at unity slip “ .” For most standard NEMA Class A, andB squirrel-cage IM under 1000 hp, typical data suggest the as-sumption of starting pf as being nearly equal to 0.2 [ 1].

The transition between continuous and discontinuous linecurrent waveforms occurs for an range smaller than, but closeto the pf angle at unity slip (i.e., ). At the criticalangle “ ,” some of the line currents become discontinuous inthe first cycle, and during a considerable portion of the accel-eration period, all line currents become continuous. Similar to

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the starting pf, only minor changes occur in “ ” for class A andB squirrel-cage IM. Indeed, “ ” is found to be 63 and 65 ,respectively, for the low-voltage and medium-voltage IMs.

In Table VII, firing angles of all thyristors are defined withrespect to the zero-crossing point on the ascending portion of phase R voltage. First, T1 receives a firing pulse at “ ,” but itdoes not trigger into conduction until T2 receives a firing pulse

at 120 for case 1, “ ” for case 2, and “ ” forcase 3. This initiates the two-phase operation. The next thyristorwhich is going to be triggered into conduction is T3, which re-ceives its firing pulse with a delay of 90 for cases 1 and 2, and“ ”forcase 3. In theeliminationof torquepulsations,onlythese two switching instants for T2 and T3 are important; theremaining thyristors T4 to T6 will receive firing pulses in thenormal sequence with delays of 180 , 240 , and 300 with re-spect to “ .” Expressions of “ ” and “ ” are givenin (11) and (12), respectively. The coefficients of these straightlines have been found within % accuracy from the resultsof simulation studies, by curve fitting techniques. For this pur-pose, several trials have been made in order to eliminate thetorque pulsations for different initial firing angle settings. Majorchanges in the coefficients of straight lines have not been ob-served from low-voltage IMs to medium voltage ones. There-fore, a common set of straight line equations for the optimumtriggering instants of T2 andT3 canbe used. These would nearlyeliminate, or minimize the pulsating torque component for ma-chinesof differentsizes. Initial values, which areconsiderablysmaller than the pf angle at unity slip, yield continuous line cur-rent waveforms (case 1 in Table VII). For a loaded motor, as thespeed increases during starting, motor pf increases too. How-ever, this will not be the case if the motor is going to be accel-erated at no load or light load. It is obvious that a firing anglecontrol range from zero to the critical firing angle “ ” at unityslip (case 1) results in no soft starting action over a considerableportion of the starting

(11)

(12)

period. The three line currents remain continuous, and the softstarter behaves just like a direct-online starter during acceler-ation. However, as the speed increases, gradually decreases,and may have a value lower than setting in the steady state.Therefore, this operating condition gives rise to discontinuousline current waveforms, and hence, to pulsations in the electro-

magnetic torque at six times the supply frequency in the steadystate. For case 2, some of the line current waveforms becomediscontinuous (during current holdoff angle ), and during aconsiderable portion of the starting period, all current wave-forms become continuous. A sample waveform set is given inFig. 3. When , the line currents become discontinuousover the entire starting period (case 3 in Table VII). A samplewaveform set is given in Fig. 4. It has been shown in [ 16] that thetransient supply frequency torque component at starting of theIM is almost entirely dependent on the transient asymmetricalmagnetizing current. In view of the very high value of ,with respect to the magnetizing current, it is necessary to con-nect each phase to the corresponding supply voltage at or near

Fig. 3. Line current waveforms for case 2 at = 7 2 [Time scale: 10 ms/div;(top), (middle) phase current during starting (100A/div); (bottom)phase currentat steady state (20 A/div)].

the instant of maximum value, to minimize this transient com-ponent. Starting from the expression of per unit magnitude of supply frequency torque component given in [ 16] for direct-on-line starting, the per unit value of the supply frequency torque

component produced by an IM during soft-starting can be ex-pressed approximately as given in (13)

(13)

where is the stator current holdoff angle in the first cycle, asindicated on Fig. 3. “ ” depends on , and remains almost thesame for all IMs with a starting pf in the range 0.2 –0.3.Note that

is minimized when both and. Here, since is

expressed with respect to the zero-crossing of phase R voltage,a phase-shift of rad. occurs with respect to . For rel-

atively high values of initial firing angles (i.e., ), asignificant sixth-order torque ripple superimposes on the supplyfrequency oscillations. This brings a further delay to the op-timum initial switching instants, which can only be determinedby means of numerical simulations.

For the elimination of supply-frequency torque pulsations,the soft starter selects one of the control strategies which aredefined as Case 1 to 3 in Table VII, by comparing “ ” withthe boundaries “ ,” and “ ,” (i.e., case 1: , case 2:

, case 3: ). Since minor changes occurin the values of “ ” and “ ” from one motor to another, torquepulsations will be nearly eliminated by using fixed values suchas and . Motor-dependent adjustment of “ ”

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Fig. 4. Line current waveforms for case 3 at = 8 0 . [Time scale: 10 ms/div;(top), (middle)phase current during starting (100A/div); (bottom)phase currentat steady state (20 A/div)].

and “ ” (e.g., and ), however, eliminatesentirely the torque pulsations. There remains the determinationof by the C for the given motor, and the set value of currentlimit. This can be achieved by using one of the methods given

below in the order of complexity:1) starting with a sufficiently large value (e.g.,

) and using closed-loop current control strategy;2) motor-dependent lookup table;3) calculating from the per-phase equivalent circuit by the

use of motor parameters;4) parameter estimation based on one cycle energization of

the motor.

B. Current Control Strategy

A simple current control strategy is proposed to keep thestarting current constant at the preset value. This closed-loop

current control strategy is implemented by the use of a C, andgives nearly level envelopes for the starting current waveforms.As illustrated in Fig. 5, the proposed current control strategy iscomposed of successive cosinusoidal and constant function seg-ments of “ ” in accordance with the feedback signal taken fromone of the line currents. Starting current limit can be set to anyvalue in the range from to by the operator,where and k are the peak value of rated motor current,and the index indicating the limit value of starting current, re-spectively. k varies in the range from 1 to 5 –8, depending on themotor. The controller acts with respect to a limit value whichis chosen to be only a few percentages smaller than .Here, the purpose is to keep fluctuations in the current around

Fig. 5. Illustration of proposed starting current limiting strategy.

. In Fig. 5, this lower limit value at which the controlleracts has been chosen as .

In Fig. 5, is the initial firing angle estimated by the con-troller, and gives the preset value of starting current just at theswitching instants. is theperiodof thequarter of cosinusoidal

variation, and0.95 is thelower limit of theband defined forthestarting current. , and the lower band limit constant determinethesuccess of thestartingcurrent pattern. Choosing valueun-necessarily large will cause an extended starting time duration;however, the starting current pattern will be perfectly smooth.Choosing too small will result in an increased number of

cosine and constant “ ” segments, during start-up. This wouldcause a distortedand undesirable startingcurrent pattern tendingto move outside the chosen band. , and the lower band limitconstant can be reprogrammed optimally by the application en-gineer using the field experience, and depending upon the sizeof the motor and load characteristics, if necessary. By means of sensitive programming, the possibility of high inrush currentsis eliminated, and the starting current waveform approximatesmore closely to a waveform with ideal envelopes. After ener-gizing the motor at , “ ” is varied by the controller cos-inusoidally until the current reaches, and then tends to exceedthe lower limit of the band: at . The cur-rent controller then keeps “ ”constant at untilthe current

returns to after making an overshoot.At time t2, the controller replaces constant “ ” by cosinu-soidal variation, resulting in an undershoot in starting currentenvelope in the time period from t2 to t3. The controller ac-tion repeats itself in this manner until the motor reaches thefull speed. It is worth noting that the variations in Fig. 5 arenot drawn to scale, and are exaggerated in order to illustrate theoperation principles of the current controller. At the final speed,“ ” reaches zero.

Line current and EM torque waveforms during starting of thesame medium-voltage IM are given in Fig. 6 for a current limitvalue of .

Following observations can be made on these waveforms:

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Fig. 6. Line current and electromagnetic torque versus time for I = 3 : 3 I .

1) starting current remains nearly constant at the presetvalue;

2) starting torque pulsations at the supply frequency are suc-cessfully eliminated;

3) since the line currents are discontinuous during starting,torque pulsations at six times the supply frequency withrelatively low amplitudes, are produced;

4) Since too many transitions would occur between cosineand constant function segments, and the control band

is narrow, versus t andversus t variations are not as marked as in Fig. 5.

C. Experimental Results

The performance of the developed soft starter equipped withtorque pulsation elimination, and starting current control strate-

gies is tested on a 4-kVA, 0.4-kV laboratory motor-generatorset equipped with a custom-design torque and speed measuringsystem in the dynamic state.

Fig. 7 shows the variations in the shaft torque and startingcurrent of the soft starter for current setting. As can beunderstood from Fig. 7(top), simultaneous connection of thethree motor terminals to the supply causes severe pulsations inthe shaft torque. If the thyristors of the soft starter were trig-gered at “ ,” “ ,” “ ' on the first ap-plied voltage cycle, the response of the system would be as inFig. 7(middle). This switching strategy yields less pulsationsin the shaft torque in comparison with simultaneous switching.The proposed switching strategy, however, eliminates entirely

Fig. 7. Starting performances of the system with the proposed current controlmethod (Ch4: Shaft torque, 50 Nm/div; Ch3: Line current, 60 A/div; Ch2:Speed, 1500 r/min/div). (top) Simultaneous switching, (middle) all phasesswitched at = 72 , (bottom) proposed pulsating torque elimination strategy.

the shaft torque pulsations, and reduces the starting time of themotor-load combination, as shown in Fig. 7(bottom).

A set of starting current versus time waveforms is presentedin Fig. 8 for a comparative assessment of the known open-looptriggering strategies (ramp and cosinusoidal triggering func-tions), and the proposed closed-loop current limiting strategy.In these case studies, the pulsating torque elimination strategyhas been applied, and the output of the driven machine has been

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Fig. 8. Comparison of startingperformances of various current control strategies withpulsating torque elimination. (topleft) Direct-online starting without torqueelimination. (middle left) Ramp triggering function. (bottom left) Cosine triggering function. (top right) Cosine-constant triggering function. (middle right) Ramptriggering with ramp time T = 1 5 s. (bottom right) Cosine triggering with period T = 7 s.

adjusted to the nominal value at steady-state. Direct-onlinestarting in Fig. 8(top left) is given only for comparison pur-poses. In Fig. 8(middle left ) and (bottom left), the firing angleof soft-starter thyristors is varied, respectively, as ramp andcosine functions during starting. The following observationsare made on these waveforms.

• The starting current does not remain constant, and es-pecially for ramp triggering function, may follow thedirect-online starting current envelope as the motor speedincreases.

• The periods of ramp and cosinusoidal triggering functionsin which “ ” is varied from “ ” to zero, aredependenton

the motor and load characteristics. For the load consideredinFig.8,longertimeperiodswouldapproximatethestartingcurrent moreclosely to a constantcurrent waveform.

However, for the current control strategy proposed in this work,the starting current can be successfully kept constant at a presetvalue, with nearly level envelopes, during the starting period[Fig. 8(iv)]. Since the control exercised on the starting currentis closed-loop control, the proposed control system inherentlytakes into account load characteristics. Note that the periods of ramp [Fig. 8(ii)], and cosinusoidal [Fig. 8(iii)] “ ” variationsare adjusted to the starting period of cosine-constant triggeringfunction [Fig. 8(iv)] which is 4.2 s.

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In order to approximate the starting current envelopes to idealones for the ramp and cosinusoidal triggering functions, sev-eral trials have been made. The optimum cases are as given inFigs. 8(v) and (vi). Corresponding variation periods are 15 and7 s, which are longer than the “ ” variation period of the pro-posed current control strategy. A good correlation is obtainedbetween the theoretical and experimental results.

V. CONCLUSION

Performance optimization of a voltage-controlled, thyristor-ized inductionmotor soft starter is carried outby theuseof somesimple control strategies implementable on a microcontroller atno additional cost. Nearlyperfect current and torqueprofilescanbe obtained during starting for better utilization of the availablestarting facilities. Using the proposed strategies, a good accel-eration profile can be tailored by smooth, pulsation-free torquesover the entire starting period. Since soft starting by voltagecontrol reduces the starting torque with the square of the ratio

of applied rms voltage to rated voltage, the applicability of thismethod depends on the stiffness of the grid, and torque require-ment of the load. On the other hand, for loads with high startingtorque requirement, and operating on very weak grids, the onlyalternatives are starting by frequency control or by conventionaltechniques employing wound rotor induction motors, for whichhigh torques can be produced at low currents.

APPENDIX

A. Machine Data

Medium voltage induction motor:Serial Number: 474437; Type: A3 12-42-8Y4;Manufacturer: Made in USSR;Three-phase, Y-connected, 50-Hz, eight-pole, 250-kWIM;Supply voltage: 6-kV l-l, Rated current: 31 A, .Parameters referred to stator side:

; ; ;

Low-voltage laboratory machine:Type/serial number: DUM 2550-2/4 B3 JP21;Manufacturer: SIEMENS;Three-phase, Y-connected, 50-Hz, 4-pole, 4-kVAsquirrel-cage IM;Supply voltage: 380-V l-l, Rated current: 8.1A, .Parameters referred to stator side

; ; ; .

B. Load Torque

.

REFERENCES

[1] A. J. Williams and M. S. Griffith, “Evaluating the effects of motorstarting on industrial and commercial power systems, ” IEEE Trans. Ind. Applicat. , vol. IA-14, pp. 292 –299, July/Aug. 1978.

[2] F. M. Bruce, R. J. Graefe, A. Lutz, and M. D. Panlener, “Reduced-voltage startingof squirrel-cage induction motors, ” IEEETrans. Ind.Ap- plicat. , vol. IA-20, pp. 46 –55, Jan./Feb. 1984.

[3] P. J. Colleran and W. E. Rogers, “Controlled starting of AC inductionmotors, ” IEEE Trans. Ind. Applicat. , vol. IA-19, pp. 1014 –1018,Nov./Dec. 1983.

[4] J.Nevelsteenand H.Aragon, “Startingof large motors-methods and eco-nomics, ” IEEE Trans. Ind. Applicat. , vol. 25, pp. 1012 –1018, Nov./Dec.1989.

[5] F. Blaabjerg, J. K. Pedersen, S. Rise, H. H. Hansen, and A. M. Trzynad-lowski, “Can soft-starters help save energy, ” IEEE Ind. Applicat. Mag. ,vol. 3, pp. 56 –66, Sept./Oct. 1997.

[6] J. Bowerfind and S. J. Campbell, “Application of solid-state AC motorstarters in the pulp and paper industry, ” IEEE Trans. Ind. Applicat. , vol.IA-22, pp. 109 –114, Jan./Feb. 1986.

[7] W. Shepherd, “On the analysis of the three phase induction motor withvoltage control by thyristor switching, ” IEEE Trans. Ind. General Ap- plicat. , vol. IGA-4, pp. 304 –311, May/June 1968.

[8] S. A. Hamed and B. J. Chalmers, “Analysis of variable-voltage thyristorcontrolled induction motors, ” Proc. Inst. Elect. Eng., B , vol. 137, no. 3,pp. 184 –193, May 1990.

[9] S. S. Murthy and G. J. Berg, “A new approach to dynamic modelingand transientanalysis of SCR controlled inductionmotors, ” IEEE Trans.Power App. Syst. , vol. PAS-101, pp. 219 –229, Sept. 1982.

[10] G. Nath and G. J. Berg, “Transient analysis of three-phase SCR con-trolled induction motors, ” IEEE Trans. Ind. Applicat. , vol. IA-17, pp.

133–142, Mar./Apr. 1981.[11] L.X. Leand G.J. Berg, “Steady-state performance analysis of SCR con-trolled induction motors: A closed form solution, ” IEEE Trans. Power App. Syst. , vol. PAS-103, pp. 601 –611, Mar. 1984.

[12] F. M. H. Khater and D. W. Novotny, “An equivalent circuit model forphase backvoltagecontrolof ACmachines, ” IEEETrans. Ind.Applicat. ,vol. IA-22, pp. 835 –841, Sept./Oct. 1986.

[13] T. A. Lipo, “The analysis of induction motors with symmetricallytriggered thyristors, ” IEEE Trans. Power App. Syst. , vol. PAS-90, pp.515–525, Mar./Apr. 1971.

[14] T. M. Rowanand T. A. Lipo, “A quantitative analysis of inductionmotorperformance improvement by SCR voltage control, ” IEEE Trans. Ind. Applicat. , vol. IA-19, pp. 545 –553, July/Aug. 1983.

[15] W. Deleroi, J. B. Woudstra, and A. A. Fahim, “Analysis and applicationof three-phase induction motor voltage controller with improved tran-sient performance, ” IEEE Trans. Ind. Applicat. , vol. 25, pp. 280 –286,Mar./Apr. 1989.

[16] W. S. Wood, F. Flynn, and A. Shanmugasundaram, “Transient torquesin induction motors due to switching of the supply, ” Proc. Inst. Elect. Eng. , vol. 112, no. 7, pp. 1348 –1354, July 1965.

[17] H. Rehaoulia and M. Poloujadoff, “Transient behavior of the resultantairgap field during run-up of an induction motor, ” IEEE Trans. EnergyConversion , vol. EC-1, pp. 92 –98, Dec. 1986.

[18] R. H. Daugherty, “Analysis of transient electrical torques and shafttorques in induction motors as a result of power supply disturbances, ” IEEE Power App. Syst. , vol. PAS-101, pp. 2826 –2836, Aug. 1982.

[19] A. A. Shaltout, “Analysis of torsional torques in starting of large squirrelcage induction motors, ” IEEE Trans. Energy Conversion , vol. 9, pp.135–141, Mar. 1994.

[20] J. Faiz, M. Ghaneei, and A. Keyhani, “Performance analysis of fast re-closing transients in induction motors, ” IEEETrans. Energy Conversion ,vol. 14, pp. 101 –107, Mar. 1999.

[21] I. Çadirci, M. Ermis, E. Nal çaci, B. Ertan, and M. Rahman, “A solidstate direct-on line starter for medium voltage induction motors with

minimized current and torque pulsations, ” IEEE Trans. Energy Conver-sion , vol. 14, pp. 402 –412, Sept. 1999.

Gürkan Zenginobuz (S’00) received the B.Sc. and M.Sc. degrees in electricaland electronics engineering from Middle East Technical University (METU),Ankara, Turkey, in 1997 and 2000, respectively. He is currently pursuing thePh.D. degree at METU.

Currently, he is a Research and Development Engineer in the InformationTechnologies and Electronics Research Institute at the Scientific and TechnicalResearch Council of Turkey, Ankara. His areas of research include inductionmotor soft starters at medium voltage and microprocessor control of motordrives.

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Isik Çadirci (M’98) received the B.Sc., M.Sc., and Ph.D. degrees in electricaland electronics engineering from Middle East Technical University (METU),Ankara, Turkey, in 1987, 1988, and 1994, respectively.

Currently, she is Associate Professor at Hacettepe University and is a Se-nior Researcher for Information Technologies and Electronics Research Insti-tute with the Scientific and Technical Research Council of Turkey. Her areas of interest include electric motor drives and switch-mode power supplies.

Muammer Ermis (M’98) received the B.Sc., M.Sc., and Ph.D. degrees in elec-trical engineering from Middle East Technical University (METU), Ankara,Turkey, in 1972, 1976, and 1982, respectively. He received the M.B.A. degree inproduction management from Ankara Academy of Commercial and EconomicSciences in 1974.

Currently, he is a Professor of Electrical Engineering at METU and Directorof the Power Electronics Group in Information Technologies and ElectronicsResearch Institute, Ankara. His current research interests are high power,medium voltage motor drives, and static reactive power compensation systems.

Cüneyt Barlak received the B.Sc. and M.Sc. degrees in electrical and elec-tronics engineering from Anadolu University, Eskisehir, Turkey, and MiddleEast Technical University, Ankara, Turkey, in 1994 and 1996, respectively.

His areas of interest include power electronic applications for electric motordrives.

Mr. Barlak is a member of the Institute of Electrical Engineers in Turkey.

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