Ieee 141

IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 39, NO. 2, MARCH/APRIL 2003 313

Self-Excitation of Induction Motors Compensatedby Permanently Connected Capacitors andRecommendations for IEEE Std 141-1993

Muammer Ermis, Member, IEEE, Zafer Çakir, Isik Çadirci, Member, IEEE,Gurkan Zenginobuz, Student Member, IEEE, and Hakki Tezcan

Abstract—Self-excitation of induction motors compensated bypermanently connected capacitors is investigated in this paper.Theoretical analyses of self-excitation phenomenon are carriedout by using some simplified equivalent circuits, and a hybridmathematical model in axes, respectively, in steadystate and transient state. An unusual operating condition aboutwater pumping stations is reported, in which water within thepipeline may drive the motor in the reverse direction at speedshigher than synchronous, when a supply interruption coincideswith a check-valve failure. In order to prevent the motor fromdangerous overvoltages due to self-excitation, it is recommendedto connect a simple static protection circuit consisting of a resistorin series with a thyristor switch between any two lines of the motor.Critical resistance boundaries, which will lead to loss of excitation,and demagnetization of the rotor core are determined separatelyas a function of operating speed. A suitable resistance valuechosen in the Safe Design Area constitutes a reliable protectionmechanism against self-excitation.

Index Terms—Compensation, induction motors (IMs), self-exci-tation.

I. INTRODUCTION

SELF-EXCITATION phenomenon in induction motors hasbeen well-known since the 1930s [1], [2]. When an induc-

tion machine is disconnected from the supply, and driven bya mechanical source, terminal voltage builds up if its laggingvar demand is supplied externally, and sufficient residual mag-netism is present in the rotor core. This is known as self-ex-citation phenomenon in the literature. Shunt compensation ca-pacitors are the most common var supplies for the self excita-

Paper ICPSD 00–02, presented at the 2000 Industry Applications Society An-nual Meeting, Rome, Italy, October 8–12, and approved for publication in theIEEE TRANSACTIONS ONINDUSTRY APPLICATIONSby the Power Systems En-gineering Committee of the IEEE Industry Applications Society. Manuscriptsubmitted for review October 15, 2000 and released for publication December11, 2002.

M. Ermis, I. Çadirci, and G. Zenginobuz are with the TÜBITAK-METUInformation Technologies and Electronics Research Institute, TR06531Ankara, Turkey, and also with the Electrical and Electronics EngineeringDepartment, Middle East Technical University, TR 06531 Ankara, Turkey(email: [email protected]; [email protected]; [email protected]).

Z. Çakir is with the Electrical and Electronics Engineering Department,Middle East Technical University, TR 06531 Ankara, Turkey (email: [email protected]).

H. Tezcan was with the Iskenderun Iron and Steel Plant, TR31319 Hatay,Turkey. He is currently with Tezcan Elektrik Mühendislik, Karaköy/Istanbul,Turkey.

Digital Object Identifier 10.1109/TIA.2003.808978

tion of induction motors. The use of an induction machine as anautonomous generator due to self-excitation phenomenon hasbeen extensively investigated by several researchers, especiallyfor wind power generation [3]–[8].

Recommended practices for power-factor improvement of in-duction motors supplied from the utility grid are given in var-ious standards and handbooks in detail [9], [10]. Since the motorreactive power does not change too much from no load to fullload, fixed shunt capacitors can be installed in various mannersfor power-factor improvement. Among these, the use of capac-itors directly connected to motor terminals is the cheapest andthe simplest solution to this problem. Maximum capacitor ratingshould not be in excess of 100% and 90% of no-load reactivepower consumption of the motor as recommended, respectively,in [9] and [10]. A capacitance higher than the recommendedvalues leads to overvoltages at the motor terminals owing toself-excitation, when the motor is disconnected from the utilitygrid. Even though the capacitor ratings are chosen according torecommended practice, they may lead to self-excitation of themotor in some cases. These cases are clearly specified in theNever do List[9] as “Never connect the capacitors directly tothe motor when…”.

This paper investigates the effects of capacitors directly con-nected to the motor on self-excitation phenomenon during asupply failure. It also reports an unusual operating conditionabout water pumping stations which is not included in theNeverdo List. If a check-valve failure coincides with supply interrup-tion, then the gravitational force acting on water may drive thepump motor in the reverse direction at a speed higher than syn-chronous. Also, in wind energy conversion systems and smallhydroelectric plants containing constant-voltage and constant-frequency induction generators connected to the grid, and com-pensated by permanently connected capacitors, sudden discon-nection of the generator from the grid inexorably results in over-speed. Directly connected capacitors may then cause self-exci-tation of pump motor, or the generator. A completely static pro-tection circuit is proposed in this paper in order to avoid harmfuleffects of generated overvoltages on the machine and capacitorfor such cases. Guidelines for the design of a shunt compensa-tion system for water pumping stations are presented. The re-sults of steady-state and transient-state analyses are verified byexperiments conducted on a universal machine set in the labo-ratory.

0093-9994/03$17.00 © 2003 IEEE

314 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 39, NO. 2, MARCH/APRIL 2003

(a) (b) (c)

Fig. 1. Common reactive power compensation techniques for multi-motor applications. (a) Individual compensation capacitors directly connected to motorterminals. (b) Individual compensation by contactor switched shunt capacitors. (c) Group compensation.

II. REACTIVE POWER COMPENSATION OFGRID-CONNECTED

INDUCTION MOTORS

A. Common Compensation Techniques

A comparative evaluation of power factor (PF) improvementtechniques for squirrel-cage induction motors (IMs) has beengiven in [11]. Among these techniques, the use of shunt ca-pacitor banks is a common solution to the var compensationproblem of grid-connected IM. Common installation techniquesof shunt capacitors for multi-motor applications are as illus-trated in Fig. 1. Direct connection of a fixed capacitor bank tothe motor terminals is the simplest and cheapest solution to thevar compensation problem, if each IM has long hours of use.The main advantage is that this technique [Fig. 1(a)] also uti-lizes the motor contactor (MC) to switch the capacitors and themotor together as a unit so that the capacitors are on the systemonly when required [9]. The scheme in Fig. 1(b) is a more costlysolution since each fixed capacitor bank for each motor needs itsown contactor (CC). On the other hand, the group compensationtechnique in Fig. 1(c) is more suitable especially if duty factorsof most of the motors are considerably low. High cost of con-trol elements, which make capacitive var generation adjustable,is offset to a certain extent by low installed var capacity. ThePF of a squirrel-cage motor at full-load is usually between 80%and 90%, depending upon the rated speed, type, and size of themotor [9].

Even though the PF of an IM varies significantly from noload to full load, the motor reactive power does not change verymuch. With a properly selected capacitor, the operating PF be-comes excellent over the entire load range of the motor. Fig. 2shows the variations in kVA, kvar, kW, and PF at the input ter-minals of a three-phase four-pole 50-Hz 380-V 8.1-A IM whichis part of a universal machine set (see the Appendix). The exper-imental work presented in this paper has been conducted on itand, therefore, it will be called the test machine in the text. Thelevel to which the PF should be improved depends on the eco-nomic payback in terms of utility PF penalty requirements, andsystem energy saved due to lower losses. It is generally in excessof 95% at full load, and higher at partial loads. The suggestedmaximum capacitor ratings for different motor types are givenin [9], in tabulated form. Higher values can cause self-excitationas the motor runs down and, in consequence, a substantial over-

Fig. 2. Measured input and output data of the test machine.

voltage at the motor terminals. Contributions of capacitors ratedat 100%, 80%, and 72% of no-load reactive power consumptionon the input PF of the test machine are also given in Fig. 2. 19.3

F per-phase delta yields 100% compensation capacity. If themotor has long hours of use at partial loads, an 80% or lowercapacity cannot raise the PF to a satisfactory level on a monthlybasis. For large- and/or medium-voltage IMs, contributions ofthe above compensation capacities would be much better.

B. Problem Definition in the Case of Water Pumping Stations

The cooling water requirement of various processes in theIskenderun Iron and Steel Plant, Hatay, Turkey, is partly sup-plied from the Mersin River by means of six centrifugal pumpsdriven by medium-voltage squirrel-cage IMs (Fig. 3). The wateris pumped to the plant at a rate of 5000 m/h through 80-cm-di-ameter and 2500-m-long pipelines (Fig. 4). The water pumpingstation is at sea level, however, the plant is located at an altitudeof 15 m. The water is highly alcaline as can be understood from

ERMIS et al.: SELF-EXCITATION OF INDUCTION MOTORS AND RECOMMENDATIONS FOR IEEE STD 141-1993 315

Fig. 3. General view of Mersin River water pumping station.

Fig. 4. Schematic diagram of the hydraulic system.

the test report given in the Appendix. The nameplate data of themotors manufactured by two different companies are as given inthe Appendix. Their PFs at full load are lower than 0.90. In viewof utility PF penalty requirements, average system power factorshould be improved to 0.92 lagging in terms of kilowatthourand kilovar-hour values appearing in bills. Since each motor haslong hours of use nearly at full load, it has been decided to con-nect sufficient amount of capacitors directly to the terminals ofeach motor as a cost-effective solution to the var compensationproblem.

In the design phase of the project, operational data haveshown that the mean rate of utility grid interruptions at thepumping station caused by faults and/or equipment failure in

harsh climatic conditions is around ten interruptions per siteper month. During interruptions with long duration, putting thebackup supplies into service takes a considerable time. Amongfive of these supply interruptions on the average per annum,gravitational force drives water within the pipeline back tothe water pumping station. Water flow in the reverse directionfirst brings the centrifugal pump and the motor to a quick stop,and then accelerates them in a short time period in the reversedirection.

Final running speeds of two different motors are measured tobe 1650 and 1100 r/min, respectively. These values are nearly10% higher than the corresponding synchronous speeds. Thereason for such an unusual event is that, sometimes one of thecheck valves does not function properly to prevent water flowin the reverse direction because of substantial amount of saltdeposits and iron rust on the inner surface of hydraulic systemelements. The amount of solid materials which were depositedon the inner surface of check valves are given in the third columnof the analysis report in the Appendix. The solid material hasthe appearance of iron rust. The performance of the cathodicprotection system seems to be inadequate.

The operational experience on the Mersin River Pumping Sta-tion reveals that special attention should be paid to the designof var compensation system for pump motors in similar appli-cations, especially if permanently connected fixed capacitorsare going to be used. Otherwise, the supply failure may leadto self-excitation in pump motors, and generation of dangerousovervoltages at the motor terminals which may have harmful ef-fects on motor insulation and human life.

C. Static Protection Circuit Against Self-Excitation

For the unusual operating conditions, a static protection cir-cuit as given in Fig. 5 can be used to prevent the motor from dan-gerous overvoltages during self-excitation. It is based on con-nection of a damping resistor between any two motor lineswhenever the motor self-excites. Fast connection of the dampingresistor is achieved by using a pair of back-to-back-connectedthyristors. This is a cheap solution because the protection circuitemploys only one resistor bank instead of a three-phase bankswitched by three static switches, one for each phase.

The idea behind this choice is to cause loss of excitation,which is the rapid decay of terminal voltage to zero and, at thesame time, to destroy residual magnetism so that it cannot berestored until the motor is reconnected to the supply. In orderto guarantee that motor iron is brought to a throughly demagne-tized state, the damping resistor is kept connected to the motorterminals for a predetermined time period. This is achieved bythe use of a timing circuit as shown in Fig. 5. However, in a prac-tical application, it is better to apply firing pulses to the gatesuntil recovery of the supply.

The success of the latter method is based on the fact thatonce a load resistance leads to loss of excitation for a certaincompensation capacitance, and shaft speed, the machine doesnot self-excite for the same conditions if the load resistance re-mains connected to it. The control circuit continuously monitorsthe terminal voltage and status of the MC. Whenever the ter-minal voltage builds up and tends to exceed a preset thresholdvalue, the controller sends a signal to the firing circuit to apply


Fig. 5. Proposed static protection circuit against self-excitation.

firing pulses to the gates of back-to-back-connected thyristors.The threshold value of the voltage can be set to a value in therange from 50% to 120% of the motor rated voltage dependingupon the application, as will be discussed in Section IV.

Forward and backward thyristors will receive firing pulseson the zero-crossing points of the positive and negative halfcycles of generated voltage, respectively. In order to discrim-inate whether the overvoltage at the motor terminals is owingto self-excitation or switching action of the MC, a signal canbe taken from the auxiliary contact of the MC. Since the ter-minal voltage, and hence the resistor current decays in a fewtens of cycles, depending upon the protection circuit design, andmotor type, size, and parameters, the thyristor switch and thedamping resistor can be implemented as simple and cheap nat-urally cooled units.

III. A NALYSIS OF SELF-EXCITATION PHENOMENON

Self-excitation of induction motors compensated by fixed ca-pacitors directly connected to the stator terminals will be inves-tigated theoretically in this section in both steady state and tran-sient state.

A. Steady-State Analysis

Although self-excitation is initiated by the residual magneticfield of the rotor, the frequency and magnitude of inducedvoltages appearing across the motor terminals are dictated byshaft speed and shunt compensation capacitance. If the motor isdriven by its load in either direction, i.e., forward or reverse at aspeed of in revolutions per minute, then the slip,becomesnegative with a value very close to zero, e.g.,0.2% for thetest machine resulting in a frequency of 49.9 Hz for inducedemfs. This operating condition assumes that, self-excitationoccurs when the motor–capacitor combination is disconnectedfrom the utility grid. The synchronous speed in revolutionsper minute is, therefore, less than, but very close to it. Onecan assume that . As an engineering approximation,the frequency of induced voltages,, can then be computedfrom (1)

(1)

where is the number of poles of the IM.For the test machine, the magnetization characteristic (

versus ) in Fig. 6 is obtained experimentally by driving the

Fig. 6. Magnetization characteristic of test machine.

IM by a dc motor, while the stator terminals are connected toa three-phase 380-V 50-Hz utility grid via an autotransformer.Stator voltage is adjusted by the use of an autotransformer insteps, up to a voltage level of 15% higher than the nominal value.Higher voltage values are obtained by capacitive excitation. Foreach voltage value, active power transferred from the supply tothe stator is made zero by adjusting the shaft speed of the dcmotor. These operating conditions ensure that IM draws only themagnetising current from the 50-Hz supply, and by subtractingstator leakage impedance voltage drop from statorterminal voltage at a phase angle of 90lag, the variations in

can be computed point by point as a function of.Equivalent circuit parameters of the IM are found by car-

rying out locked-rotor and no-load tests at rated voltage andfrequency, and dc resistance measurements (see the Appendix)[12]. The magnetization characteristic at any speed

can be deduced point by point from the one at rated fre-quency Hz by equating ratio toratio for the same , where is the rated value ofsynchronous speed, and 1500 r/min for the test machine.

The variations in terminal voltage against shaft speed and ca-pacitance for a self-excited IM can be determined from one of


(a) (b)

(c)

Fig. 7. Equivalent circuit representations of IM for self-excitation. (a) Exactequivalent circuit. (b) Approximate circuit. (c) The most approximate circuit.

the equivalent circuits in Fig. 7. In these equivalent circuits, thecapacitive reactance , stator and referred rotor leakage reac-tances and , and magnetizing reactance are all com-puted at rated frequency. The per-unit frequency, and speed

are defined as in (2) and (3)

(2)

(3)

, , , and are voltage and current phasors at anyshaft speed and, hence, corresponding frequency. Amongthese, the equivalent circuit in Fig. 7(a) is the most accurateone, and is used by several researchers in the analyses of self-excited induction generators [6], [13], [14]. In the utilizationof these equivalent circuits, circuit equations should be solvedsimultaneously with nonlinear magnetization characteristic at.

During self-excitation, capacitor losses, copper losses, andcore losses are supplied by the driving force of the mechanicalload through electromechanical energy conversion process rep-resented by negative slip-dependent rotor resistance

in Fig. 7(a). That is the main reason whyis much lowerthan full-load slip and . The approximate equivalentcircuits in Fig. 7(b) and (c) can, therefore, be obtained from theexact equivalent circuit in Fig. 7(a), by using the fundamentalassumption given in (1), i.e., by substituting . This is per-missible because of self-excited induction generator operationat no load.

The two approximate equivalent circuits are valuable andtime-saving tools for the design of compensation system formotors. The variations in terminal voltage against capacitanceat constant speed are given in several papers [2], [3], [8], [13],especially for self-excited induction generators operating at noload. However, more drastic variations will occur in terminalvoltage as the shaft speed increased for a fixed capacitance aswill be discussed in Section IV. Therefore, the variation of crit-ical capacitance with shaft speed has a crucial importancein the design of shunt compensation system directly connected

Fig. 8. Critical capacitance versus speed.

to the motor. A capacitance larger than may lead to self-ex-citation of motor when stator terminals are disconnected fromthe supply. can be computed from any one of the circuits inFig. 7 by using unsaturated value of . Unsaturated at

is the slope of the air-gap line (a.g.l.) which is tangent tothe linear portion of experimental magnetization characteristicin Fig. 6. The approximate value at any and can bedirectly computed from (4)

(4)

where for approximate equivalent circuit inFig. 7(b), and for the most approximate equivalentcircuit in Fig. 7(c). Stator leakage inductance , and unsatu-rated value of mutual inductance are to be calculated from

and unsaturated at .Theoretical values against are given in Fig. 8 in com-

parison with experimental results. The experimental values arekept slightly higher than actual values in order to maintainself-excitation at a stable equilibrium point near the knee point.The most optimistic results have been obtained for the exactand the most approximate equivalent circuits in Fig. 7(a) and(c), which are very close to actual values. The most pessimisticresults have been obtained for the approximate equivalent cir-cuit in Fig. 7(b), because of the neglect of all losses. Anotherinteresting result is that the most approximate equivalent circuitgives very accurate and much better results than the approximatecircuit for the calculation of . This is because the neglectof in Fig. 7(c) largely compensates for the errors owing tothe neglect of machine losses. Therefore, the most approximateequivalent circuit in Fig. 7(c) can be safely used in the designof permanently connected compensation system even for large-and medium-voltage IMs. Furthermore, it requires minimum in-formation about the motor, i.e., only the air-gap portion of ex-perimental magnetization characteristic at .

For large- and/or medium-voltage IMs, test certificates in-cluding no-load and blocked-rotor test results and resistancemeasurement are usually present. Equivalent circuit parameters


can be computed and magnetization characteristic at rated fre-quency can be deduced from the given test data [12].valuesfor different rotor speeds can therefore be calculated easily from(5), where is the critical capacitance at rated frequencyobtained from the test data by using the most approximate equiv-alent circuit in Fig. 7(c). It is seen from (5) that the critical ca-pacitance is inversely proportional to the square of speed ratio

. Therefore, the maximum shaft speed that the machine mayattain when a supply failure occurs should be taken as the mostimportant factor affecting the design of compensation system.At high speeds, even a capacitance value considerably smallerthan the recommended values in standards may cause genera-tion of dangerous voltages at the machine terminals.

(5)

B. Transient Analyses

The static circuit in Fig. 5 has been proposed for protectionagainst self-excitation of the motor. Whenever an overvoltageappears across the machine terminals owing to self-excitation,

is immediately connected between any two lines of the motor.Since demands active power from the machine in addition tosystem losses, this action may bring the operating point to anunstable point on the unsaturated region of plane, re-sulting in loss of excitation. On the other hand, if a resistancehigher than the critical value is connected to the machineterminals, generated voltage reduces, but the operating point re-mains at a stable equilibrium point on the nonlinear portion ofmagnetization characteristic, resulting in no protection action.Therefore, the variation of with for the specified valueof has an utmost importance in the design of static protectioncircuit. This is determined by the use of hybrid mathematicalmodel for IM in terms of axes quantities (Fig. 9 and(6), as shown at the bottom of the page).

The mathematical model in (6) has been obtained from the ac-tual one in terms of axes quantities by applying thewell-known phase and commutator transformationsand[15], only to the rotor side. The equations of three-phase capac-itor bank connected to the motor terminals, and damping resistorconnected across two lines are not included in (6). Although the

Fig. 9. Schematic representation of IM inABC=dq axes.

speed is assumed to be constant for each digital simulation trial,the model remains nonlinear because of the saturated portion ofthe magnetization characteristic. It is essential in the analysis ofself-excitation phenomenon to preserve magnetic nonlinearityin order to be able to find a stable equilibrium point in the satu-rated region. Numerical integration of the mathematical modelhas been carried out by the use of a fourth-order Runge–Kuttaalgorithm which can integrate overdiscontinuities. In the math-ematical model, all machine parameters except the mutual in-ductance between stator and rotor (its stator referred value isdenoted by ) are assumed to be constant. Potential drop on

is the resultant electromotive force (EMF) on the stator side, and is the cause of terminal voltage appearing across the

machine terminals during self- excitation. is produced by thecombined effect of stator and rotor magnetic fields, i.e., by theresultant magnetomotive force (MMF) . For the self-excitedIM, is nearly equal to stator MMF . At each numericalintegration step, correct value of is to be substituted intothe model. This makes necessary the determination of resultantMMF and, hence, the corresponding saturation level. In the nu-merical simulation, this is achieved by determining the degreeof saturation at theth integration step, updating value and,then, using it in the next integration step . For that purpose,the magnetization characteristic in Fig. 6 is normalized with re-spect to angular frequency , and expressed in termsof peak resultant mmf per stator turns . It is subdividedinto three regions (Fig. 6). In region 1, a small value representing

where

and (6)


residual voltage has been introduced. In regions 2 and 3, respec-tively, the air-gap line and the saturated portion of the magne-tization characteristic are to be used. The nonlinear portion issatisfactorily approximated byas given in [16]. The coefficients of exponential expression havebeen found by curve fitting technique. At any integration step,

can be computed from (7) and (8)

(7)

(8)

where is the actual number of stator turns, the effectivenumber of stator turns, and and i are the andaxes stator current components. The three-phase to two-phasetransformation has been applied to the stator side only forthe calculation of . For the calculated value of ,the saturation factor can be determined from Fig. 6 by using(9)

(9)

where and are peak volt-seconds read on the sat-urated portion and air-gap line of the normalized magnetizationcharacteristic, respectively. The saturated value that willbe used in the th integration step can, therefore, be deter-mined from (10), by using saturation factordetermined at the

th integration step [17]

(10)

where is the unsaturated mutual inductance coefficientdetermined from the slope of the air-gap line.

The transient model has been run on a computer several timesfor different and values. Two sample outputs are givenin Fig. 10. In Fig. 10(a), since is greater than , loss ofexcitation does not occur, resulting in a reduction of terminalvoltage only. However, as sufficient damping is introduced intothe self-excited IM, terminal voltage collapses, resulting in lossof excitation as can be understood from Fig. 10(b).

IV. EXPERIMENTAL RESULTS

A series of tests have been carried out in the laboratory on theuniversal machine set to make clear the self-excitation phenom-enon, and to verify the theoretical results given in Section III. Inthe experiments, in order to keep the shaft speed constant at anyset value, the dc machine is supplied from a three-phase thyris-torized ac/dc converter system.

First, several trials have been made in order to find outwhether the operating conditions in the water pumping stationin the Iskenderun Iron and Steel Plant may lead to self-exci-tation of the IM or not. It is observed that terminal voltagebuilds up in all cases with a compensation capacity of 100% ofthe no-load reactive power demand ( F per-phasedelta). The minimum speed which leads to self-excitation isfound to be 1450 r/min. It is less than the synchronous valueof 1500 r/min for 50-Hz excitation. As pointed out in [5], it is

(a)

(b)

Fig. 10. Simulated line-to-line voltage waveform with the static protectioncircuit (n = 1590 r/min). (a)R = 160 . (b)R = 60 .

more difficult to self-excite large machines, because the ratioof their magnetising inductance in linear operating range to theone at very small currents is higher than those of small IMs.This fact does not eliminate the risk of self-excitation of largeIMs entirely, because the machine may attain a high speedas in the case of the water pumping station with sufficientcapacitance directly connected to its terminals for satisfactorycompensation.

For a fixed capacitance, a drastic increase in terminal voltageagainst shaft speed has been obtained, as given in Table I. Inorder to compute the terminal voltage for the given compen-sation capacitor by the use of equivalent circuit in Fig. 7(c),terminal characteristic of the capacitor should be plotted onversus or versus the curve of the IM (Fig. 6), andthe intersection point between them is to be found. In the casewhere the graphical solution will be based on versus the

curve, (11) is to be used as the equation of the load line,where is the capacitor voltage. Although the capacitance re-mains the same at higher operating shaft speeds and, hence, fre-quencies, the slope of the load line decreases proportionately


TABLE IVARIATIONS OF TERMINAL VOLTAGE WITH SHAFT SPEED

Fig. 11. The variations in terminal voltage (Ch 2) and resistor current (Ch3) during loss of excitation (n = 1530 r/min, R = 60 , and decay timeT � 1:5 s).

with the square of speed ratio (or square of frequency ratio), re-sulting in the drastic increase in terminal voltage given in Table I

(11)

The variations in terminal voltage and damping resistor cur-rent against time are obtained during self-excitation in order totest the performance of the static protection circuit proposedin the paper. A sample record of these variations is shown inFig. 11. In this case, thyristors receive firing pulses for a periodof s, and terminal voltage and resistor current decay tozero in a shorter time period of s. During loss of exci-tation, electrical energy stored in capacitive- and inductive-typeenergy storage elements of the system are extracted almost en-tirely. Although time constants and are very ef-fective on decay time , loss of excitation occurs in a time pe-riod considerably bigger than those estimated from. This isbecause, there is considerable electrical power input to the rotorcircuit with decaying amplitude, from the mechanical systemby the negative slip-dependent rotor resistance in Fig. 7(a). Inorder to be able to define aSafe Design Area(SDA) for the pro-tection circuit on damping resistance vs speed plane, the resultsof the above tests are evaluated and the boundaries defining crit-ical resistances as a function of shaft-speed and thyristor firing

Fig. 12. Critical resistance versus speed for a compensation capacitance ratedat no-load reactive power consumption. Region 1: no loss of excitation, terminalvoltage drops only. Region 2: loss of excitation occurs, thus destroying residualmagnetism. Region 3: Loss of excitation occurs, but residual magnetism is notdestroyed.

period (timer setting) are given in Fig. 12. These results havebeen obtained for F per-phase delta and kept con-stant throughout the tests. Two boundaries marked asand in Fig. 12 have been obtained defining three oper-ating regions. boundary is unique for a given machineand compensation capacitance, as expected.

A large value of in Region 1 brings the operating pointto a new equilibrium point on the saturated portion of themagnetization characteristic resulting in no loss of excitationand protection action. The self-excited IM starts to generatea lower voltage to balance active power demanded by,machine losses, and generated power by slip-dependent rotorresistance. A sample waveform set which shows the variationsin terminal voltage and damping resistor current is given inFig. 13(a). However, for Region 2, an value suitable for thegiven speed will lead to both loss of excitation, and a successfuldemagnetization of the rotor core, in such a way that afterremoval of firing pulses, terminal voltage does not build upbecause of the lack of sufficient residual magnetism in the rotorcore. A sample waveform set is given in Fig. 13(b), for which

is adjusted to 2 s. Region 3 arises from the adjustment ofto very low values in comparison with those of Regions 1 and2. In this region, terminal voltage decays very rapidly, but in afinite time period, after removing the firing pulses, the machinereexcites. This is attributed to the fact that loss of excitationoccurs more rapidly than demagnetization process.

Complete demagnetization of a core can only be obtained byvarying the exciting current in either direction very slowly [18].A sample waveform set for operation in Region 3 is given inFig. 13(c) for which s when is adjusted to 2 s. Itis seen from these waveforms that terminal voltage builds up in

5 s time after removal of the firing pulses. A solid short-circuit


(a) (b)

(c)

Fig. 13. Variations in terminal voltage (Ch1) and resistance current (Ch2) against time for different values ofR with a timer setting ofT = 2 s (n = 1560 r/min,C = 19:3 �F). (a) Case study illustrating operation in Region 1,R = 160 . (b) Case study illustrating operation in Region 2,R = 60 . (c) Case studyillustrating operation in Region 3,R = 4 .

applied to the motor terminals does not lead to a considerabledemagnetization effect in the rotor core as pointed out in [7].Keeping thyristors triggered for longer time periods enlargesRegion 3, resulting in a family of boundaries insteadof a unique one. Only two boundaries of Region 3 for sand 2 s are plotted in Fig. 12. An example SDA for shaft speedsin excess of 1600 r/min may be the shaded area in Fig. 12.

For sizing the elements of protection circuit and thermal man-agement of the system, Regions 1 and 3 in Fig. 12 are developedon versus speed plane as given in Fig. 14,where and are, respectively, rated rms current forthe motor ( A for the test machine), and the maximumrms value of the resistor current just after triggering the thyris-tors. From Fig. 14, can be chosen in the range from75% to 200% of for terminal voltage at maximum attain-able shaft speed. For the test machine, ifis chosen to be 60,

and will be, respectively, 7.3 A rms and 2 s, resultingin 3.2-kW peak power dissipation and2-kJ energy absorptioncapability for the damping resistor. For this purpose, carbon ce-ramic resistors capable of sustaining high transient currents andFig. 14. Current versus speed.


high surge energies [19] or capacitor-discharge-type resistorsmay be used. A surge resistor suitable for printed circuit board(PCB) mounting with a 3-kJ surge and 9-W average power dis-sipation capability is suitable for the above example. Its dimen-sions in millimeters are , , and .Small thyristors ( V, A) are used inthe tests without any heatsink. To be on the safe side, a heat sinkhaving nearly 4-kW thermal resistance with natural air coolingcould be used.

One important point of interest in the design is the thresholdvalue of induced voltage at which the thyristor switch will betriggered into conduction. For protection against self-excitation,for rotation only in the reverse direction, as in the case of waterpumping stations, the design constraints of the thyristor switchand damping resistor can be alleviated by applying a differenttriggering strategy. Whenever the terminal voltage rises, andtends to exceed a low value, for example, 50% of the rated value,the thyristor switch can be triggered into conduction if MC isalready open. This assumes the IM does not self-excite in theforward direction of rotation during its rapid deceleration.

The precharged capacitor will discharge rapidly through thestator after supply failure during deceleration period. Sincethis switching strategy makes the protection circuit smallerand cheaper, its contribution is more marked for large and/ormedium voltage IMs than that of a small machine. However,for applications in which the IM is driven by the high inertiaload at overspeeds in the forward direction after a supplyinterruption, the above triggering strategy will obviously makeno contribution to the alleviation of design constraints. Anotherpoint of interest in the design is the timer setting. The exper-imental results have shown thatdirectly determinesboundary and, hence, the success of the proposed protectioncircuit. Since it is impractical to carry out tests in the field ashas been conducted in the laboratory on a small machine,setting of the protection circuit should be made independent ofsize and ratings of IM, and type of the ferromagnetic materialused in the rotor core. This can be achieved by applying firingpulses to the gates of thyristors until the recovery of the supply.This will automatically avoid the uncertainty ofboundary and the design of the protection circuit can then bebased only on the boundary.

V. DESIGN OFCOMPENSATIONSYSTEM FORPUMP MOTORS

First, problems which may arise from the presence of powersystem harmonics have been investigated theoretically, and har-monic spectra of both voltages and currents have been measuredin the field. The results show that pump motors can be success-fully compensated by permanently connected fixed capacitorswithout a need to tuned filter reactors. This is because, the pumpmotors are supplied from a 34.5-kV busbar via two 34.5/6-kVstep-down transformers in two separate groups. After installa-tion, these findings have been verified by harmonic measure-ments at the pumping station.

Magnetization characteristic has been deduced, and equiva-lent circuit parameters have been calculated from the results ofno-load test, blocked-rotor test, and resistance measurementsgiven in the test certificates of the two types of pump motors.

Fig. 15. Magnetization characteristic of Pump Motor Type I.

Fig. 16. Determination of maximum permissible capacitance for Pump MotorType I.

Fig. 15 shows the magnetization characteristic of motor type1 (see the Appendix) for 50-Hz excitation. The variations incritical capacitance with shaft speed are then obtained from theair-gap line in Fig. 15, (5), and the approximate equivalent cir-cuit in Fig. 7(b). It is given in Fig. 16 for motor type 1. Themaximum shaft speed in the reverse direction that will be at-tained by the motor in the case of a check valve failure can bedetermined from a detailed model of the overall system, whichconsists of pipelines, centrifugal pump, and motor mechanicalsubsystem. Instead of this theoretical approach, the maximumshaft speed value for each motor type is measured separatelyby establishing the same operating conditions in the field artifi-cially. This critical speed r/min for motor type 1 ismarked in Fig. 16 which intersects the critical capacitance curveat 11.5 F per-phase wye. In order to eliminate self-excitation


risk, the shunt capacitance that can be connected directly to themotor terminals should be less than this value.

On the other hand, shunt capacitors rated at 100%, 90%, and80% of no-load reactive power demand of IMs at rated voltagecan be calculated from no-load test data for motor type 1 , whichare found to be 16.2, 14.6, and 13.0F per-phase wye, respec-tively. It can be concluded from Fig. 16 that the compensationsystem designed according to 80% of no-load reactive powermay lead to self-excitation at maximum shaft speed of 1650r/min in the case of a supply failure. Therefore, a smaller com-pensation capacity (11.3F per-phase wye) has been assignedfor motor type 1, which increases the PF to 0.92 lag at ratedvoltage and full load. It can be calculated simply from the name-plate data by taking into account the utility penalty limit of 0.90.The recommended capacitance may lead to self-excitation atspeeds in excess of 1650 r/min.

Since the most expected value of terminal voltage is 6.3 kV,the contribution of 11.3 F per-phase wye to the improvementof power factor will be less due to the character of the saturatedportion of the magnetization curve. Operational experience fora few months after installation has shown that the realized meanpower factor is close to 0.92.

The penalty for PF for many utilities is however 0.95. Ahigher capacitance (13.0F/phase wye) would be needed formotor type 1 in such cases instead of the design value of 11.3Fper-phase wye. This capacitance would obviously lead to self-excitation, even at speeds near to synchronous value. Therefore,in such cases, either the shunt compensation system that will beconnected directly to the motor terminals should be equippedwith the protection circuit proposed in this paper, or anothercompensation technique such as shunt capacitors switched byCC should be adopted.

VI. CONCLUSIONS

Recommendations for IEEE Std. 141-1993

1) Never connect the capacitors directly to the motor inconventional manner using the capacitor kilovar valuesrecommended by motor manufacturers when wateris pumped through pipelines. If a supply interruptioncoincides with check-valve failure as an unusual con-dition, then water within the pipeline may drive themotor in the reverse directions at speeds higher thanthe synchronous speed. This may lead to generation ofdangerous over-voltages at the motor terminals owing toself-excitation phenomenon.

2) Capacitors can be connected directly to the motor termi-nals in a water pumping station only if:

a) var compensation system is equipped with a static pro-tection circuit against self-excitation, or

b) a var capacity smaller than recommended practicemeets the utility PF penalty requirements, and doesnot lead to self-excitation at maximum speed in thereverse direction. This makes necessary the avail-ability of both motor test data and maximum shaftspeed value.

TABLE IIMERSIN RIVER WATER ANALYSIS (METHOD: Fe, Ca,AND Mg

DETERMINED BY FLAME ATOMIC ABSORPTIONTECHNIQUE USING PERKIN

ELMER MODEL 305BAAS)

TABLE IIINAMEPLATE DATA OF PUMP MOTORS

APPENDIX

A. Machine Parameters and Nameplate Data of Test Machine

Nameplate data: three-phase, wye-connected, 50-Hz, four-pole; stator: 220 V/phase, 14 A; rotor: 140 V/phase, 17 A. Pa-rameters: ; ; ; ;

at rated operating voltage.

B. Mersin River Water Analysis Report

See Table II.

C. Nameplate Data of Pump Motors

See Table III.

REFERENCES

[1] E. D. Basset and F. M. Potter, “Capacitive excitation for induction gen-erators,”Elect. Eng. Trans., pp. 540–545, 1934.

[2] C. F. Wagner, “Self excitation of induction motors,”Elect. Eng. Trans.,vol. 58, pp. 47–51, 1938.

[3] N. Mohan and M. Riaz, “Wind-driven capacitor excited induction gener-ators for residential electric heating,” presented at theIEEE PES WinterMeeting, New York, 1978, Paper A 78 050-7.

[4] D. W. Novotny and G. H. Studtmann, “Self excitation in inverter driveninduction machines,”IEEE Trans. Power App. Syst., vol. PAS-96, pp.1117–1125, July/Aug. 1977.

[5] J. M. Elder, J. T. Boys, and Prof.J. L. Woodward, “The process of self-excitation in induction generators,”Proc. Inst. Elect. Eng., pt. B, vol.130, no. 2, pp. 103–108, Mar. 1983.

[6] G. Raina and O. P. Malik, “Wind energy conversion using a self excitedinduction generator,”IEEE Trans. Power App. Syst., vol. PAS-102, pp.3933–3936, Dec. 1983.

[7] J. M. Elder and J. T. Boys, “Self-excited induction machine as a smalllow-cost generator,”Proc. Inst. Elect. Eng., vol. 131, no. 2, pp. 33–41,Mar. 1984.

[8] S. P. Singh, B. Singh, and M. P. Jain, “Performance characteristics andoptimum utilization of a cage machine as capacitor excited inductiongenerator,”IEEE Trans. Energy Conversion, vol. 5, pp. 679–685, Dec.1990.


[9] IEEE Recommended Practice for Electric Power Distribution for Indus-trial Plants, IEEE Std. 141 (Red Book), 1993.

[10] SIEMENS Electrical Installations Handbook Part 2, 2nd revised andenlarged ed. New York: Wiley, 1987.

[11] R. Spee and A. K. Wallace, “Comparative evaluation of power factor im-provement techniques for squirrel cage induction motors,”IEEE Trans.Ind. Applicat., vol. 28, pp. 382–386, Mar./Apr. 1992.

[12] A. E. Fitzgerald and C. Kingsley,Electric Machinery, 2nd ed. NewYork: McGraw-Hill, 1961.

[13] L. Shridhar, B. Singh, C. S. Jha, B. P. Singh, and S. S. Murthy, “Selectionof capacitors for the self regulated short shunt self excited inductiongenerator,”IEEE Trans. Energy Conversion, vol. 10, pp. 10–17, Mar.1995.

[14] T. F. Chan, “Analysis of self-excited induction generators using an iter-ative method,”IEEE Trans. Energy Conversion, vol. 10, pp. 502–507,Sept. 1995.

[15] C. V. Jones,Unified Theory of Electrical Machines. London, U.K.:Butterworth, 1967.

[16] M. Ermis, “Representation of the magnetization characteristic of self-excited induction generators for computer use,” inProc. AMSE Conf.,vol. 2C, July 1988, pp. 139–150.

[17] , “Modeling and analysis of a wind turbine driven self excited in-duction generator,” Ph.D. dissertation, METU, Ankara, Turkey, Apr.1982.

[18] H. C. Roters,Electromagnetic Devices. New York: Wiley, 1941.[19] K. C. E. Smart, “Low Profile Power Resistor,” inIEE Pulse Power

Colloq., 1997, pp. 26/1–26/3.

Muammer Ermis (M’99) received the B.Sc.degree in electrical engineering from Middle EastTechnical University, Ankara, Turkey, in 1972, theM.B.A. degree in production management from theAcademy of Commercial and Economic Sciences,Ankara, Turkey, in 1974, and the M.Sc. and Ph.D.degrees in electrical engineering from Middle EastTechnical University in 1976 and 1982, respectively.

He is currently a Professor of Electrical Engi-neering at Middle East Technical University, and isalso the Director of the Power Electronics Group,

TUBITAK Information Technologies and Electronics Research Institute,Scientific and Technical Research Council of Turkey. His current researchinterests are static reactive power compensation systems and high-powermedium-voltage motor drives.

Zafer Çakir received the B.Sc. and M.Sc. degreesin electrical and electronics engineering from MiddleEast Technical University, Ankara, Turkey, in 1997and 2000, respectively.

Between 1997–2000, he was a Research Engineer,working in the field of ac motor drives for auxiliarypower supplies of electric locomotives. Presently,he is fulfilling his military obligation in Turkey. Hisresearch interests include modeling and simulationof induction motor drives, and high-frequency powerconverters.

Isik Çadirci (M’98) received the B.Sc., M.Sc.,and Ph.D. degrees in electrical and electronicsengineering from Middle East Technical University,Ankara, Turkey, in 1987, 1988, and 1994, respec-tively.

She is currently an Associate Professor in theElectrical and Electronics Engineering Department,Middle East Technical University, and also a SeniorResearcher for the TUBITAK Information Technolo-gies and Electronics Research Institute, Scientificand Technical Research Council of Turkey. Her

areas of interest include electric motor drives, switch-mode power supplies,and static reactive power compensation systems.

Gurkan Zenginobuz (S’00) received the B.Sc.and M.Sc. degrees in electrical and electronicsengineering in 1997 and 2000, respectively, fromMiddle East Technical University, Ankara, Turkey,where he is currently working toward the Ph.D.degree.

He is also currently a Research and DevelopmentEngineer with the TUBITAK Information Technolo-gies and Electronics Research Institute, Scientificand Technical Research Council of Turkey. Hisareas of research are induction motor soft starters at

medium voltage and microprocessor control of motor drives.

Hakki Tezcan received the B.Sc. degree from YildizTechnical University, Istanbul, Turkey, in 1975.

From 1975 to 1980, he was an Electrical Engineerwith the Iskenderun Iron and Steel Plant (ISDEMIR).Between 1980–1985, he was a Chief Engineer withFoster Wheeler Intercontinental Corporation. From1985 to 2000, he was Vice Director of the ElectricalMaintenance and Repair Department of ISDEMIR.He presently directs Tezcan Elektrik Mühendislik, Is-tanbul, Turkey, a private company specializing in re-active power compensation systems.

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