A Strategy for Real Power Control in a Direct-Drive PMSG Based WECS

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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 28, NO. 3, JULY 2013 1297

A Strategy for Real Power Control in a Direct-DrivePMSG-Based Wind Energy Conversion System

Omid Alizadeh , Student Member, IEEE , and Amirnaser Yazdani , Senior Member, IEEE

Abstract— As the penetration of wind energy into the powersystem continues to grow, wind energy conversion systems(WECSs) are increasingly expected to be able to control theiroutput real power, while retaining their maximum power-point

tracking (MPPT) capability. This paper proposes a simple realpower control strategy, which augments the MPPT feature of modern WECSs, and is based on rapid torque control as opposedto the traditional pitch-angle control. This paper presents the

implementation of the proposed control strategy for a direct-driveWECS that employs the permanent-magnet synchronous gener-

ator, even though the proposed method can also be extended toother classes of electronically interfaced WECSs. The paper alsopresents a parameter-tuning procedure for the proposed control

strategy. The effectiveness of the proposed control strategy isdemonstrated through mathematical analysis and time-domainsimulation studies.

Index Terms— Control, damping, direct drive, eigenvalue anal-ysis, permanent-magnet synchronous generator (PMSG), windenergy.

I. I NTRODUCTION

T HE anticipated large-scale integration of wind energy

conversion systems (WECSs) into the electric power

system indicates that system operators should be able to con-

trol the output real and reactive powers of the WECSs, to

more effectively take part in the control of the power system

and to ride through faults and other contingencies [1]. While

the reactive-power controllability of electronically interfaced

WECSs is widely recognized, their real power controllability

has received insignificant attention and, thus far, been merely

utilized for maximum power-point tracking (MPPT).

This paper proposes a strategy for controlling the output real

power of a direct-drive WECS that employs a high-pole per-

manent-magnet synchronous generator (PMSG) [2]. The choice

is based on the expectation that PMSG-based WECSs will be

widely deployed in the future, due to their low-loss generators,

low maintenance requirements, and quiet drive-trains [3]. The proposed real power control strategy is based on rapid torque

Manuscript received June 29, 2011; revised November 04, 2012; acceptedApril 04, 2013. Date of publication May 03, 2013; date of current version June20, 2013. This work was supported in part by the Natural Sciences and Engi-neering Research Council (NSERC) of Canada, andin part by theFaculty of En-gineering, the University of Western Ontario. Paper no. TPWRD-00558-2011.

O. Alizadeh is with theUniversity of Western Ontario, London, ON N6G5B9Canada (e-mail: [email protected]).

A. Yazdani is with Ryerson University, Toronto, ON M5B 2K3, Canada(e-mail: [email protected]).

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

Digital Object Identifier 10.1109/TPWRD.2013.2258177

control, rather than the (slow) pitch-angle control presented in

[4]–[6]. Thus, the pitch-angle control is not exercised for output

real power control, but it is employed, exclusively and, as per

the common practice, for limiting the mechanical power if the

turbine overspeeds.

Rapid torque control, however, can excite drive-train tor-

sional modes, especially since the drive-train stiffness is, in

general, inversely proportional to the number of generator

poles [7] and is therefore low in a WECS with a high-pole

PMSG. Moreover, a high-pole PMSG possesses no inherent

damping [8]. Drive-train oscillations, if not damped, impact the

operation and may even lead to instabilities. Thus, a supple-mentary active damping scheme is designed for the proposed

power-control strategy, based on a detailed mathematical model

and eigenvalue analysis of the WECS. The proposed control

strategy and its supplementary active damping scheme enable

the control of the WECS output real power, from a low value

up to the maximum power corresponding to the prevailing wind

speed. The active damping strategy, however, is not unique and

may be achieved through other reported techniques [9]–[11].

II. STRUCTURE OF THE DIRECT-DRIVE WECS

Fig. 1 illustrates a simplified schematic diagram of a direct-

drive PMSG-based WECS. The WECS is composed of a windturbine, a high-pole PMSG, and a power-electronic ac-dc-ac

converter, which interfaces the PMSG to the host utility grid.

The ac-dc-ac converter, in turn, consists of two back-to-back

voltage-sourced converters, VSC1 and VSC2. The converter

VSC1 controls the PMSG torque and, thus, the power that the

PMSG extracts from the wind turbine, whereas VSC2 regu-

lates the dc-link voltage by controlling the real power that it ex-

changes with the grid. The converter VSC2 can also exchange

reactive power with the grid, to support the grid or enhance

voltage stability [12]. In Fig. 1, the composition of the wind tur-

bine, PMSG, VSC1, and the scheme that controls VSC1 is la-

beled as the energy capture subsystem, whereas the compositionof the dc-link capacitor, VSC2, the tie reactor , and the con-

trol scheme for the regulation of the dc-link voltage and reactive

power is identified as the controlled dc-voltage power port [13].

This paper exclusively studies the dynamics of the energy cap-

ture subsystem; this is possible since the arrangement shown in

Fig. 1 effectively decouples thedynamics of the utility grid from

those of the wind turbine, drive-train, and PMSG.

III. MATHEMATICAL MODEL AND CONTROL SCHEMES

This section presents a mathematical model for the energy

capture subsystem of the WECS of Fig. 1. For the sake of com-

pactness, hereafter, the same notation will be adopted for a vari-

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1298 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 28, NO. 3, JULY 2013

Fig. 1. Simplified schematic diagram of the direct-drive WECS.

able and its Laplace transform. In addition, the superscript sig-

nifies the reference value (setpoint) for a variable.

A. Mechanical Torque and Drive-Train

The mechanical torque of a wind turbine is given by:

(1)

where is the turbine torque (in Nm), is the turbine

power (in watts), is the turbine angular speed (in radians per

second), is the turbine radius(in meters), is the air massden-

sity (in ), is the wind speed (in meters per second),

is the turbine tip-speed ratio (unitless), is the

pitch angle (in degrees), and (unitless) is the so-called tur-

bine power ef ficiency [14].

Drive-train dynamics may be represented by models of dif-

ferent levels of complexity. However, as discussed in refer-

ences [15] and [16], a two-mass model is adequate for cap-

turing the dynamics that affect stability; higher-order models

are commonly employed for studying the mechanical fatigue of

the turbine drive-train. Thus, ignoring the mechanical losses, the

drive-train is represented by the following two-mass model:

(2)

(3)

(4)

where is the PMSG rotor speed (in rad/s); and re-

spectively signify the turbine and PMSG moments of inertia (in

); is the drive-train stiffness (in Nm/rad); thevariable

represents the torsional displacement of the drive-train (in rad);and denotes the PMSG torque (inNm). Equations (1) through

(4) constitute a state-space drive-train model for the energy cap-

ture subsystem.

B. Permanent-Magnet Synchronous Generator (PMSG)

The PMSG torque control is performed in a rotating frame

whose direct axis is aligned with the PMSG rotor flux vector,

as discussed in [13]. The control can be tuned such that

responds to its setpoint based on the following first-order

transfer function:

(5)

for which the time constant is a design parameter [13].

Fig. 2. Block diagram of the PMSG power control scheme.

If the PMSG torque is controlled, the PMSG power can be

regulated. Fig. 2 shows the block diagram of a control scheme

whose main objective is to force the PMSG power to track

the power setpoint . The setpoint is determined based on

the mode of operation, as will be explained in the next section.

Fig. 2 also illustrates the pitch-angle control process whosefunction is to ensure that the turbine and PMSG speeds do not

exceed the maximum permissible value, . Thus, if ex-

ceeds , a PI compensator increases to decrease the tur-

bine power and regulate at ; if is smaller than ,

the compensator output is saturated at its lower limit, , to

maximize the turbine power. Fig. 2 further shows that the PI

compensator output passes through a rate limiter which repre-

sents the limited speed at which the pitch angle can be changed

in practice.

IV. PROPOSED CONTROL AND MODES OF OPERATION

The objective of the proposed control is to enable the WECSto exercise power-flow control in addition to the MPPT. Thus,

two modes of operation are defined for the WECS of Fig. 1: (1)

the MPPT mode, and (2) the controlled-power (CP) mode. The

operating mode is determined by the way that the setpoint is

stipulated, as explained next.

A. MPPT Mode of Operation

In the MPPT mode, the objective is to maximize the power

that the turbine extracts. This can be achieved if is maxi-

mized. To maximize , must be kept constant at its optimum

value , regardless of the wind speed. The objective is ful-

filled if the PMSG power setpoint is determined based on the

following law [17]:

(6)

in which the constant is

(7)

It then follows from assuming a fast control that .

Thus, (6) can be rewritten as

(8)

Fig. 3 illustrates the characteristic curve of a wind turbine

(heavy solid line), for a wind speed. The figure also plots the



ALIZADEH AND YAZDANI: STRATEGY FOR REAL POWER CONTROL 1299

Fig. 3. Characteristic curves of a wind turbine for a wind speed and two dif-ferent values of pitch angle, that is, (heavy solid line) and (lightsolid line).

PMSG power versus rotor speed, based on (8) (dashed line). It is

observed that if the WECS is in the MPPT mode, the operating point settles at point A (intersection of the two curves) which

corresponds to the maximum turbine power at the given

wind speed.

B. Controlled-Power (CP) Mode of Operation

In the CP mode, the objective is to regulate the WECS output

power at the command value , regardless of the wind

speed. Thus, is given the value of . Let us assume

that, initially, the WECS is in the MPPT mode, ,

, and ; then the value of

(and therefore ) is rapidly changed from to (i.e.,subsequent to a switching from the MPPT mode to the CP

mode). As Fig. 3 indicates, this causes the PMSG power to

drop below the turbine power and results in an increase in

towards a new value, . Depending on the wind speed,

can be larger than the maximum permissible rotor speed,

, as for the example illustrated in Fig. 3. The situation is

circumvented by the pitch-angle control mechanism; thus, once

exceeds , the pitch-angle control scheme increases

and consequently alters the power-speed characteristic of the

wind turbine, to the one shown by light solid line in Fig. 3,

such that drops to and the rotor speed settles at

(corresponding to point C in Fig. 3). To ensure that the PMSGand turbine power-speed curves have at least one crossing

point [see Fig. 3], in the CP mode is limited to the value

. Therefore, if is so large that cannot overtake

it at the given wind speed, then will be limited to

and, effectively, the system continues to operate in the MPPT

mode until either there will be a rise in the wind speed (thus

increasing the corresponding ) or the system operator steps

down the command .

Fig. 4 illustrates a mechanism for selecting between the

MPPT and CP modes of operation. As Fig. 4 shows, the set-

point is obtained from the output of a hard limiter whose

input and upper saturation limit are and , respec-

tively (the lower saturation limit is zero). Thus, is equal to

and the CP mode is exercised, as long as is smaller

Fig. 4. Block diagram illustrating the generation of the power setpoint .

than ; otherwise, is equal to and the energycapture subsystem operates in the MPPT mode. Therefore, to

permanently leave the system in the MPPT mode, it is suf ficient

to assign an adequately large value (e.g., larger than the

value of that corresponds to the rated wind speed).

V. EIGENVALUE A NALYSIS

An eigenvalue analysis is performed to reveal the dynamic

properties of the energy capture subsystem and to tune the pa-

rameters of the proposed power control strategy. To that end,

a linearized model is developed and analyzed. The numerical

examples presented hereafter are based on an example WECS

whose parameters are reported in Appendix A. The same ex-ample WECS has also been simulated in time domain for pro-

ducing the results reported in Section VII.

As Fig. 2 indicates, the compensator of the power control

scheme can be described by

(9)

Replacing with in (9) and expressing the resultant in

time domain, one finds

(10)

As Fig. 2 indicates, if saturation is ignored, in (5) is equal

to the compensator output and one finds

(11)

Eliminating between (10) and (3), and then sub-

stituting for from (11) in the resulting equation, one

deduces

(12)

Equations (2) through (4), (11), and (12), along with the alge-

braic (1), constitute the following nonlinear state-space model

for the energy capture subsystem:

(13)

for which is the vector of state

variables (superscript T denotes matrix transposition), and

are the (inter-related) control inputs, and is the

disturbance input. is a vector of nonlinear functions of

the state variables and inputs. It should be noted that (13)




Fig. 5. Absolute value of the real part of the dominant eigenvalue of as afunction of and .

assumes a constant pitch angle (i.e., it ignores the dynamicsof the pitch-angle control mechanism) to keep the mathemat-

ical model tractable. The approximation is plausible since

the pitch-angle is normally settled at its minimum value, and

varies only if the turbine speed exceeds its maximum value.

Moreover, the ignored dynamics are, by design and by nature,

remarkably slower than those of the state variables in .

However, the simulation model used for the assessment of the

proposed control strategy includes the pitch-angle control loop

(Section VII).

In the MPPT mode, is determined based on (6). Taking

derivatives with respect to time from both sides of (6), and elim-

inating between the resultingequationand(3), one finds

(14)

Substituting in (13) for and , respectively, from

(6) and (14), and linearizing the resulting set of equations, one

deduces

(15)

where and are matrices whose elements are functions of

the steady-state operating point of the system, and “~ ” denotes

the small-signal perturbation of a variable; the matrices are in-troduced in Appendix B.

For the example WECS, Fig. 5 plots the absolute value of

, that is, the real part of the dominant eigenvalue of , as a

function of and ; the dominant eigenvalue is defined as

the eigenvalue with the smallest real part(in absolute value), and

has been calculated for the operating point that corresponds to

9.0 m/s. As Fig. 5 illustrates, is maximized if 1.0

and 2.4; these values are adopted for subsequent analyses.

In the CP mode, and, thus, the linearized version

of (13) takes the form

(16)

TABLE IEIGENVALUES OF THE E NERGY CAPTURE SUBSYSTEM; ,

FOR THE CP MODE

Fig. 6. Block diagram illustrating the implementation of the active dampingcontrol.

Fig. 7. Control block diagram of the active damping scheme.

where and are matrices whose elements are functions of

the system steady-state operating point; the matrices are intro-

duced in Appendix B.

Table I reports the eigenvalues of and . It is observed

that in the MPPT mode, the system has two pairs of complex-

conjugate eigenvalues that correspond to two stable but poorly

damped eigenmodes. The situation is even worse for the CP

mode; in the CP mode, the energy capture subsystem has an

unstable oscillatory mode. Thus, both modes of operation callfor an active damping mechanism, which will be described in

the next section.

VI. ACTIVE DAMPING STRATEGY

A. Structure

The active damping scheme designed hereafter augments the

PMSG torque setpoint with a high-pass filtered measure of the

rotor speed, through the scheme illustrated in Fig. 6. As Fig. 6

shows, first the ac component of the rotor speed is extracted

by passing through a high-pass filter, . Then, a compen-

sator processes the error between zero and thefilter output, and determines the supplementary component for the

PMSG torque setpoint . As illustrated in Fig. 7, the active

damping scheme of Fig. 6 results in a control loop whose ob-

jective is to (rapidly) force the ac component of to zero.

Let be a second-order high-pass filter of the form

(17)

for which and are the corner frequency and quality factor,

respectively. For the example WECS, Fig. 8 shows a family of

curves that plot the imaginary part (frequency) of the unstable

complex-conjugate eigenvalues in the CP mode (see Table I),

for a corresponding set of wind speeds, as a function of the ratio




Fig. 8. Variation of the damped natural frequency as a function of the normal-ized power command in the CP mode.

. As the figure indicates, the frequency of the un-stable mode varies over a fairly narrow range about 4.05 rad/s.

Considering this frequency band, the choices of and

rad/s result in a small difference between and ,

in terms of magnitude and phase angle. As will be discussed in

the next subsection, the choice of a pure gain for results

in stable operation.

B. Eigenvalue Analysis

To assess the effectiveness of the presented active damping

scheme, and to find the optimum value for the gain ,

the state-space model of Section V is modified.

If , as Fig. 6 indicates, the supplementary compo-

nent of the PMSG torque setpoint is

(18)

It then follows from replacing by in (5) [see

Fig. 6], expressing the result in the small-signal time-domain

form, and substituting for from (18) in the final form that

(19)

Eliminating between (10) and (3), linearizing the re-

sult, and substituting for from (19) in the final form, one

deduces

(20)

Differentiating (3) with respect to time, eliminating

between the result and (4), and expressing the result in the small-

signal form, one deduces

(21)

Substituting for from (19) in (21), one finds

(22)

Eliminating between (22) and the expression of

(17) in the small-signal time-domain form, one concludes that

(23)

Equations (19), (20), (23), and the small-signal versions of

(2)–(4) constitute a state-space model for the energy capture

subsystems augmented with the presented active damping

scheme. Replacing and in the state-space model

with their expressions corresponding to each mode of opera-

tion, the following linear state-space models are obtained:

for the MPPT mode (24)

for the CP mode (25)

where ; the matrices

, , , and are introduced in Appendix B.

For the example WECS, Fig. 9 plots the migration of the (ini-

tially) unstable complex-conjugate eigenvalues in the CP mode,

when the active damping mechanism is enabled and is varied

from zero to . The migration plot is sketched for theoperating point that corresponds to 9.0 m/s and

1.5 MW. It is observed that the eigenvalues, which

are unstable for , migrate toward the left-half plane as

is increased, but move back toward the imaginary axis once

surpasses a certain value. This behavior indicates the exis-

tence of an optimum value for . The optimum value is com-

puted such that the complex-conjugate eigenvalues possesses

the maximum damping ratio. This, for the example WECS, cor-

responds to the choice of , which results in the

smallest angle between the real axis and the tangent to the mi-

gration plot.

Table II reports the eigenvalues of the energy capture sub-

system under the MPPT and CP modes of operation. A com-

parison of the results reported in Table I confirms the improved




Fig. 9. Migration plot of the (initially) unstable eigenvalues in the CP modefor different values of .

TABLE IIEIGENVALUES OF THE E NERGY CAPTURE SUBSYSTEM WITH ACTIVE DAMPING

CONTROL; 9 m/s, , FOR THE CP MODE 1.5 MW

Fig. 10. Migration plot of the dominant eigenvalues for , whenthe wind speed va ries from 6 to 12 m/s. (a) a nd (b)

.

damping of the complex-conjugate eigenvalues with higher un-

damped natural frequency, under the presented active damping

scheme. More important, the active damping scheme has stabi-

lized the system in the CP mode, as all of the eigenvalues lie inthe left-half plane.

For the example WECS, Fig. 10 plots the migration of the

dominant eigenvalues for a constant and two different values

of the ratio , when the wind speed varies from 6

to 12 m/s. It is observed that some eigenvalues approach the

imaginary axis as the wind speed increases. Nonetheless, the

system remains stable over the entire wind speed range, and the

eigenmodes are well damped.

To verify the accuracy of the developed mathematical model,

the response of to an abrupt switching from the MPPT mode

to the CP mode is depicted in Fig. 11. The response is obtained

from a detailed switched model of the example WECS, devel-

oped in the PSCAD/EMTDC environment [18], for the gains

, , and . As Fig. 11

Fig. 11. Response to the operation mode change from MPPT to CP.

shows, for , the response is oscillatory and unstable,

while results in sustained oscillations. How-

ever, as expected, results in a damped re-

sponse. Fig. 11 further indicates that the frequency of oscil-

lations closely match those predicted by the eigenvalue anal-

ysis. For example, Fig. 11(b) indicates that the period of oscil-lations for is about 1.56 ( 4.69/3) s, which

corresponds to an angular frequency of about 4.03 rad/s. This

frequency is very close to that indicated by Fig. 9 for

.

VII. SIMULATION R ESULTS

To further demonstrate the effectiveness of the proposed

power control strategy, the detailed switched model of the

example WECS has been subjected to various operating con-

ditions. In the graphs to follow, the angular velocities are

expressed in rad/s, the torques are expressed in MNm, the

powers are expressed in MW, the dc-link voltage is expressedin kV, the wind speed is expressed in m/s, and the pitch angle

is expressed in degrees.

Figs. 12 and 13 illustrate the WECS response to changes in

the operating mode and wind speed. Before s,

MW and the wind speed is assumed to be 9.0 m/s. In this case,

the WECS is operating in the MPPT mode and the turbine yields

the maximum power. The example WECS then experiences the

following sequence of events: (1) at s, is stepped

from 6.0 MW down to 1.5 MW and, since the maximum power

corresponding to m/s is about 2.1 MW, the operation

mode is changed from the MPPT mode to the CP mode; (2) at

s, the wind speed assumes a step change from 9.0 to 12

m/s and thus the WECS remains in the CP mode; (3) at

s, is stepped further down to 0.5 MW and thus the WECS




Fig. 12. Response to changes in the operation mode and wind speed.

continues to operate in the CP mode; (4) at s, the wind

speed changes stepwise from 12 to 9.0 m/s. However, since the

maximum power corresponding to m/s is larger than

MW, the WECS retains its CP operating mode; (5)

at s, is stepped up to 3.5 MW. This command is

larger than the maximum power for m/s, that is, 2.1

MW. Therefore, the WECS experiences a change from the CPmode to the MPPT mode and, as such, its output power settles

at 2.1 MW; and (6) at s, the wind speed again rises

stepwise from 9.0 to 12 m/s and, since the maximum power

corresponding to m/s (about 5.0 MW) is larger than

3.5 MW, the operating mode reverts back to the CP

mode and the output power settles at 3.5 MW.

Fig. 12 shows that the output power rapidly tracks

during the periods when the WECS operates in the CP mode,

that is, from to 150 s, and from s onwards.

The figure also indicates that the rises in the wind speed, at

s and 110 s, when the system is in the CP mode, result

in transient excursions in the output power, but have no effects

on the steady-state command following. It is interesting to note

that at s when is changed from 0.5 to 3.5 MW,

Fig. 13. Response to changes in the operation mode and wind speed (cont.).

TABLE IIIWIND TURBINE PARAMETERS

TABLE IVPMSG PARAMETERS

TABLE VCOMPENSATORS AND OTHER PARAMETERS

the output power transiently overshoots, but reverts to its steady

state value of 2.1 MW (that is, the maximum power for

the wind speed of 9.0 m/s). The reason for the overshoot is the

stored mechanical energy of the drive-train inertia, which is

momentarily released.

Fig. 13 shows the waveform of the supplementary compo-

nent of the PMSG torque setpoint [Fig. 13(a)], and the pitch

angle waveform [Fig. 13(b)]. It is observed that tran-

siently responds to each disturbance, but settles down at zero.

By contrast, the pitch angle only responds to those disturbances

that cause the drive-train speed to exceed (and to be in need of

regulation at) the maximum permissible value of 1.35 rad/s [see

Fig. 12(a) and (b).




(26)

(27)

VIII. CONCLUSION

This paper presented a simple real power control strategy

based on rapid control of the generator torque. The implementa-

tion of the proposed control was demonstrated for a direct-drive

WECS that employs a PMSG. It was shown that the proposed

strategy enables rapid control of the WECS output power, from

small values up to the maximum power that corresponds to

the prevailing wind conditions, but results in the instability of

the drive-train torsional modes. Therefore, the paper also pre-sented a supplementary active damping scheme and a procedure

for parameter tuning. The effectiveness of the proposed con-

trol strategy was demonstrated by mathematical analyses and

time-domain simulation studies.

APPENDIX A

PARAMETERS OF THE EXAMPLE WECS

The parameters of the example WECS are reported in

Table III and Table IV. The other parameters are introduced

in Table V. The saturation limits on the PMSG torque setpoint

(see Figs. 2 and 6) are set to 4.8 MNm.

APPENDIX B

A NALYTICAL FORMS OF MATRICES

The a nalytical forms o f the m atrices , , , , , a nd

are presented in (26) through (31). Matrices and can

be obtained by eliminating 6th and 7th rows and columns from

matrices and , respectively. The entries of these matrices

are functions of the steady-state values of the variables (denoted

by the overline). See (27) at the top of the page.

(28)

(29)

(30)

(31)

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“Comparative study on transient stability analysis of wind turbine gen-erator system using different drive train models,” IET Renew. Power

Gen., vol. 1, no. 2, pp. 131–141, Jun. 2007.[16] G. Ramtharan, N. Jenkins, O. Anaya-Lara, and E. Bossanyi, “Influence

of rotor structural dynamics representations on the electrical transient performance of FSIG and DFIG wind turbines,” Wind Energy, vol. 10, pp. 293–301, Jul./Aug. 2007.

[17] S. Morimoto, H. Nakayama, M. Sanada, and Y. Takeda, “Sensorlessoutput maximization control for variable-speed wind generationsystem using IPMSG,” IEEE Trans. Ind. Appl., vol. 41, no. 1, pp.60–67, Jan./Feb. 2005.

[18] PSCAD/EMTDC v. 4.2, Manitoba. Winnipeg, MB, Canada: HVDCResearch Centre.

Omid Alizadeh (S’12) received the B.Sc. degreein electrical engineering from Khajeh-Nasir ToosiUniversity, Tehran, Iran, in 2005, the M.Sc. degree inelectrical engineering from the University of Tehran,Tehran, Iran, in 2008, and is currently pursuingthe Ph.D. degree in electrical engineering at the

University of Western Ontario (UWO), London,ON, Canada.

His research interests include design, modeling,and control of wind energy conversion systems.

Amirnaser Yazdani (M’05–SM’09) received thePh.D. degree in electrical engineering from theUniversity of Toronto, Toronto, ON, Canada, in2005.

He was with the University of Western Ontario(UWO), London, ON, Canada. Currently, he isan Associate Professor with Ryerson University,Toronto, ON, Canada. He is a co-author of the book Voltage-Sourced Converters in Power Systems

(IEEE/Wiley, 2010). His research interests includemodeling and control of electronic power converters,

renewable electric power systems, distributed generation and storage, andmicrogrids.

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A Strategy for Real Power Control in a Direct-Drive PMSG Based WECS

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