Wind-Turbine Fault Ride-Through Enhancement - IEEE Xplore

IEEE TRANSACTIONS ON POWER SYSTEMS, VOL. 20, NO. 4, NOVEMBER 2005

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Wind-Turbine Fault Ride-Through Enhancement Alan Mullane, Member, IEEE, Gordon Lightbody, and R. Yacamini

Abstract—Fault ride-through specifications listed in modern transmission and distribution grid codes specify that wind-turbine generators (WTGs) must remain connected to electricity networks at voltage levels well below nominal. Achieving reliable operation at greatly reduced voltage levels is proving problematic. A particular problem regarding power converter-based WTGs is that standard controllers designed for reliable operation around nominal voltage levels will not work as designed during low network voltages that can occur during a fault. A consequence of this is greatly increased converter currents, which may lead to converter failure. This paper presents a nonlinear controller design for a power converter-based WTG that ensures that current levels remain within design limits, even at greatly reduced voltage levels, thus enhancing the WTG’s fault ride-through capability. Index Terms—Current control, digital signal processors, power systems, variable-speed drives, wind energy, wind power generation.

I. NOMENCLATURE Capacitance (F). Current (A). , , , , , , Gains. Inductance (H). Reference frame angular velocity (rad/s). Resistance . , Voltage (V). Subscripts Superscripts and Operators Reference value. , Direct, quadrature axis component. Differential operator. II. INTRODUCTION

P

OWER system faults are often characterized by a momentary (i.e., 0.5–30 cycles) decrease in the RMS voltage magnitude. In the past wind-turbine generators (WTGs) were allowed to disconnect from the system during a fault. As wind turbines begin to displace conventional generation, there is an increasing requirement that they remain connected to the power Manuscript received February 15, 2005; revised June 23, 2005. This work was supported by Sustainable Energy Ireland (SEI) through the National Development Plan and also by Forbairt Ireland, Electricity Supply Board (ESB) Networks, ESB Power Generation, ESB National Grid, Commission for Energy Regulation, Cylon, Airtricity, and Enterprise Ireland. Paper no. TPWRS-000842005. A. Mullane is with the Electricity Research Centre, Department of Electronic and Electrical Engineering, University College Dublin, Dublin 4, Ireland (e-mail: [email protected]). G. Lightbody and R. Yacamini are with the Department of Electronic and Electrical Engineering, University College Cork, Cork, Ireland. Digital Object Identifier 10.1109/TPWRS.2005.857390

Fig. 1. Fault ride-through requirement of wind-farms; voltage versus time profile at point of connection.

system during faults. Due to this requirement, system operators in many countries have recently established transmission and distribution system grid codes that specify the range of voltage conditions for which WTGs) must remain connected to the power system. These are commonly referred to as the fault ride-through specifications. In July 2004, the wind-farm transmission grid code was released in the Republic of Ireland. The fault ride-through requirements for WTGs connected to the Irish transmission system are specified in WF1.4.1 of the Wind Farm Transmission Grid Code Provisions as follows [1]. “A wind farm shall remain connected to the transmission system for transmission system voltage dips on any or all phases, where the transmission system voltage measured at the HV terminals of the grid connected transformer remains above the heavy black line in Fig. 1.” Achieving this ride-through requirement is a significant technical issue on which turbine manufacturers are working [2]. Considering that there are various commercially available WTG designs, including squirrel-cage induction machine based, doubly-fed induction generator based, and full rated series converter-based designs, there are various problems that must be overcome in achieving the fault ride-through requirements and some solutions have been proposed. In [3], wind-turbine voltage ride-through capabilities were investigated for the squirrel-cage design using different reactive compensation techniques such as fixed capacitor and an SVC. This paper primarily addresses the support of network voltage during a fault. In [4], a ride-through solution is proposed for the doubly-fed induction generator wind turbine that limits the high rotor currents observed during grid faults. Little work has appeared detailing fault ride-through schemes for WTGs employing full rated series converters with a design being used by many manufacturers and outlined in Fig. 2, where back-to-back voltage source converters (VSCs) connect the

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Fig. 2.


Multipole synchronous-machine-based WTG.

generator to the electrical network. In Fig. 2, and represent the total resistance and inductance between the HV terminals of the grid-connected transformer and the WTG power converter. Some work has appeared detailing the effects of voltage sags on wind-turbine-based VSC controller performance. A PI-based dc-link voltage regulator using current feed-forward was examined in [5], where a variety of voltage sag conditions were considered. The dc-link voltage controller performed satisfactorily for a three-phase dip of 30% (70% remaining) during which the load current remained constant. While this work did examine converter controller performance during a voltage dip, it concentrated on the control of dc-link voltage level and did not examine in detail the impact of controller performance on converter currents; in addition, a voltage dip to the 15% level shown in Fig. 1 was not examined. Some fault ride-though schemes for power converters in non-wind-turbine applications have appeared in the literature but often involve hardware modification [6]–[8] and do not examine controller performance in detail. This paper will demonstrate that the inherent characteristics of the converters found in modern wind turbines can cause currents to exceed their limits during network faults due to the poor off operating-point behavior of standard PI-based controllers. The standard design will first be presented and exposed to increasingly severe voltage dips, reducing to the lowest (15% remaining) level specified by the grid code. It will be shown that currents well exceeding design specifications result in lower voltage levels. A nonlinear controller design will then be presented that will be shown to maintain converter currents within their design limits, even at the low voltage levels specified by the grid code. Finally, a series of plant-model mismatch tests will be performed to demonstrate the effectiveness of the proposed controller in nonideal operating scenarios. III. WTGs The back-to-back converter gives the WTG its variable-speed capability and allows it to operate effectively over a wide range of wind speeds. In order to produce the voltages necessary for the accurate speed control of the motor connected to the machine-side inverter, a constant dc-link voltage is essential. Maintaining a constant dc-link voltage requires careful balancing of the power at both the machine and network side of the converter. Similar to many motor-drive applications, the machineside converter of a WTG is configured to control the speed of the machine, while the purpose of the network-side converter is to regulate the dc-link voltage. A successful method

for controlling the dc-link voltage, used in many commercial variable-speed motor drives, is to vary the direct component of the network current. Using this philosophy, a cascade controller is constructed where the dc-link voltage controller provides the reference signal for an inner-loop current controller. The voltage and current controllers are usually designed using small signal analysis of the nonlinear equations describing the dc-link voltage level. The system resulting from the small signal analysis, however, is only valid about a specific operating point [9], [10]. In the real system, as the operating point moves, the obtained response will vary [9], resulting in nonoptimal behavior of the whole system [11]. In a WTG where the controllers are required to operate in accordance with grid codes at voltages well below nominal, the response designed at the nominal operating point can deviate largely during a voltage dip. Deviations in the responses of the current controllers in particular could cause converter currents, designed to remain within limits at nominal voltage, to exceed their limits at low network voltage. Solving the nonlinearity problem and increasing the fault ride-through performance for the VSC of a wind turbine can be achieved with an advanced control scheme such as feedback linearization. It is known that feedback linearization can provide a predictable system response over the complete operating range [12]. A few examples of feedback linearizing control of converter systems have appeared in the literature, where they have been shown to improve dc-link voltage regulation. Feedback linearization schemes have appeared in [9], [13], and [14], where the nonlinear control technique has been used to reduce the dc-link capacitor ripple current and hence reduce the need for a large capacitor. There are some minor differences in the designs that have appeared; for example, in [9] and [14], the network series resistance was neglected. and in [13], problematic differentiators were employed in the control loop. While these designs were shown to reduce the capacitor ripple current, the inherent property of a globally predictable response was largely unutilized and the use of a feedback linearizing controller to improve the fault ride-through capability of a wind turbine has yet to appear. IV. CONVENTIONAL CONVERTER CONTROL In order to produce the required voltages , , and necessary for accurate speed control of the machine connected to the machine-side inverter, a constant dc-link voltage is essential. To do this, the network-side converter is designed to enmatches closely the current (which sure that the current

MULLANE et al.: WIND-TURBINE FAULT RIDE-THROUGH ENHANCEMENT

Fig. 3.

i

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current controller setup.

varies with changes in wind speeds and machine torque). If accurate matching is achieved, then the dc-link voltage will remain constant. Mismatches in the currents will result in an increase or decrease of the dc-link voltage at an extent determined by the size of dc-link capacitor. As previously stated, a successful method for controlling the dc-link voltage, as used in many commercial variable-speed motor drives, is to vary the direct component of the network current. This method of link-voltage control results in reasonable control under most operating conditions. A. Current Control In order to vary the direct component of the network voltage, a current controller is typically developed from the system equations describing the operation of the network side of the ac/dc/ac converter. If balanced conditions are assumed, then these can be written in a rotating reference frame using the dq transformation [15] as

Fig. 4.

DC-link model.

the dc-link voltage to the current flowing to the machine-side converter and the direct current and voltage components at the network side of the converter: (5) As this is a nonlinear model, a linearization of the model about a chosen operating point is required for a linear dc-link voltage controller design. Equation (5) can be rewritten to yield the nonlinear first-order differential equation (6) (7) This can be linearized about the operating point

to yield

(1) The state equation for

can be rewritten as

(8) (2)

A controller for the component can be constructed as shown in Fig. 3. A similar controller may be used for the component. Consider for this design a 1-MW variable-speed wind turbine, where the transformer is rated at 1.2 MVA. The transformer windings are such that the wind-turbine voltage of 690 V is stepped up to the grid voltage of 110 kV. This transformer may be represented by a series resistor and inductor together with an ideal transformer. Using a value of 5% of base impedance and a results in values of resistance and inductance ratio of referred to the low side of the transformer of and , respectively. Using the root locus method and the and values from above, a current controller was designed with a rise time of approximately 2.5 ms

Filling in for the operating point as , , and yields the linear model

,

(9)

(4)

The linear model can be represented in block diagram form as shown in Fig. 4. The first term in this block diagram accounts for the current added to the capacitor due to changes in the component of the network voltage. The second term accounts for the current added to the capacitor due to the component of current flowing from the network to the converter. The third component accounts for the current removed from the capacitor due to current flowing from the dc-link to the machine-side converter. The last component accounts for changes in capacitor current due to dc voltage variation [16]. Assuming a 1-MW design, a controller for the dc-link voltage was designed using a dc-link operating [17] with a dc-link capacitor value voltage of of (100 per horsepower [18]). The operating point for the network voltage at the low side of the transformer corresponding to from (4). was was designed using the root locus The dc-link controller method as

With correct alignment of the reference frame, the term is zero, and hence, the following equation can be written to relate

(10)

(3) B. DC-Link Voltage Control losses due to the series resistance are neglected, If the then the power balance between the ac and dc sides of the converter is given by

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Fig. 5.


Conventional cascade controller block diagram.

1

01000 A change in link current i at nominal 01000 A change in link current i at

Fig. 6. (a) v response to network voltage. (b) i response to nominal network voltage.

Fig. 7. voltage.

Using this controller and current controllers designed earlier (3), the cascade controller was connected to the nonlinear converter model as shown in Fig. 5. Fig. 6(a) shows the dc-link voltage restep in dc-link current corresponding sponse to a to the maximum expected change in link current for this drive rating. It can be seen that the deviation in dc-link voltage caused by a 1000-A change in current flowing from the machine side inverter to the dc link is compensated for within two cycles of the fundamental. As can be seen in Fig. 6(b), this compensation is achieved by increasing the current flowing to the network. Fig. 6 shows satisfactory controller performances for changes in link current at nominal voltage levels. In order to evaluate the performance of the linear design under fault or sag conditions, a series of tests was performed. In each case, the transmission system voltage measured at the HV terin minals of the grid connected transformer (equivalent to Fig. 2) was set at differing levels within the 100%–15% range as specified by the grid code and Fig. 1. The network voltage

was maintained at the chosen level for the duration of the 90 ms test. With the network voltage set, a dc-link current disturbance was applied. As the network voltage was reduced for each test, the magnitude of the applied dc-link current disturbance was also reduced ensuring that the active power delivered to the network reduced with the retained network voltage as stipulated by the grid code WF1.4.2 a) [1]. “In addition to remaining connected to the transmission system, the wind farm shall have the technical capability to provide the following function: During the transmission system voltage dip the wind farm shall provide active power in proportion to retained voltage.” From Fig. 7, it can be seen that as the network voltage reduces from its nominal value, the network current response, to a change in load current, deteriorates significantly. Indeed it can be seen from Fig. 7 that while a change in load current at nominal voltage will result in a maximum current transient of approximately 1200 A, an equivalent change in load current at a network voltage of 15% of nominal results in a current

1

i

response to change in link current at reducing levels of network


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Fig. 9.

Fig. 8. Peak i magnitude reached for full-range dc-link current step i at various network voltage levels, plotted p.u. of peak i at nominal network voltage.

Feedback linearized system configuration.

where is the state vector, is the control input, is the output, and are smooth vector fields, and is a smooth scalar function. Using the notations of differential geometry, the process of repeated differentiation begins with (12)

transient of approximately 1400 A. Fig. 8 shows the percentage increase in peak current levels above nominal for network voltage in the range 0%–100% of nominal. The observed increase in network current is due to the fact that a linear controller is used to control a nonlinear system. The linear controller will perform well for small changes about the operating point, namely, the nominal network voltage. The grid code specifies, however, that the WTG must operate at voltages far below this nominal level; these reduced levels represent large deviations from the operating point and result in an increasingly underdamped network current response. As the network current response becomes less damped, the magnitude of the network current increases. Increased inverter current levels may result in damage to the power devices and may necessitate costly uprating of the converter to deal with off operating point dynamics. Alternative control schemes for the converter should be considered to ensure that current levels remain within their design limits, even at greatly reduced network voltage levels, thus enhancing the WTG’s fault ride-through capability.

where the Lie derivative is defined as (13) If as

(14) which results in a linear relationship between

The idea of feedback linearization is to transform the dynamics of a nonlinear system so that linear control techniques can be applied. The goal of the input-output feedback linearization technique is to obtain, using state feedback and transformation, a linear relationship between a new input defined as and the output of the plant . This is outlined in the block diagram in Fig. 9, where the measured disturbance is also cancelled. The basic method for achieving input-output linearization is to differentiate the output function until the input appears and then design the input to cancel the nonlinearity.

and , namely, (15)

If the inputs have not appeared corresponding to for all , differentiate again to obtain (16) If

V. NONLINEAR CONVERTER CONTROL

for all , the input transformation can be derived

remains zero for all , differentiate repeatedly (17)

until for some integer , where the number of differentiations required for the input to appear is called the relative degree of the system. It is possible, however, that when the input appears, its coefficient is zero at the operating point. In this case, the relative degree of the system is then undefined at the operating point and may not be able to cancel the nonlinearity. In general, when the relative degree is undefined, input-output linearization cannot be achieved. The input transform can be obtained as (18)

VI. INPUT-OUTPUT FEEDBACK LINEARISATION Consider a single-input single-output (SISO) system to give a linear relation between (11)

and (19)

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The above concepts can be extended to multi-input multi-output (MIMO) systems

These equations can be written in the following form: (27) .. .

(20) where is the state vector, is the control input is the vector of system outputs. Here and vector, and are smooth vector fields and is an matrix whose columns are smooth vector fields. Input-output linearization of MIMO systems is performed in a manner similar to the SISO case, using differentiation of the outputs until the inputs appear. Assume that is the smallest integer such that at least one , then of the inputs appears in

(28)

where (29) (30) It should be noted here that the function stated as is also . The term will be accounted for later, a function of however, as well as the states. Differentiating the outputs with respect to time

(21) (31) with for at least one . Performing the above procedure for each output yields

Here (32)

(22)

where the

matrix

This yields (33)

is defined as Now calculate

using the same method

(23) (34) If is invertible, then similarly to the SISO case, the input transformation

and , the control It can be seen that as inputs have yet to appear, so differentiate again. With in (21), this yields

(24)

(35)

yields

Noting also the presence of the disturbance in (34), the following is obtained, where the control input has now appeared:

equations of the simple form (25)

VII. FEEDBACK LINEARIZATION FOR LINK VOLTAGE CONTROL Consider the m-input, m-output nonlinear system described by (1) and (5)

(36) Now the equations may be rewritten in matrix form (37) where

(26)

(38)


Fig. 10.

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Controller, prefilter, transformation, and converter block diagram.

is the differential of the reference signal, which may where be obtained from the prefilter usually included as part of the reference signal section of the controller (see Fig. 10). The folfor lowing first-order error dynamics are required where stability: (39) (46) or more compactly as This is achieved by choosing

as

(40) The inverse of the

matrix is

(47) For the other input

, the choice (48)

(41) will yield second-order error dynamics for and stable with if

The controller is singular when

Since is nonsingular in the operating range of law is given as

(42)

(49)

, the control

The signal in (48) may also be obtained from the reference signal prefilter.

(43) and the following equations result: (44) Now by choosing the values of can be shaped. First examine

and

, which will be

, the error dynamics

(45)

VIII. SIMULATION RESULTS The feedback linearization controller described in Section VII was combined with the converter model and represented as shown in Fig. 10. The figure includes the reference signal prefilters that provide the signals , , , , and whose parameters are listed in Table I. The controller parameters , , and were tuned to ensure that the speed of response and the magnitude of network current at nominal network voltage level matched closely those of the cascade controller. In a similar fashion to the linear controller presented in Section IV, the feedback linearization-controlled

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TABLE I FEEDBACK-LINEARIZATION CONTROLLER PARAMETERS

Fig. 12. Peak i magnitude (plotted p.u. of peak i at nominal network at 15% nominal network voltage) for full-range dc-link current steps i voltage for L and R plant-model mismatches in range .

620%

Fig. 11. (a) i response to change in link current at reducing levels of network voltage. (b) Peak i magnitude reached for full-range dc-link current step i at various network voltage levels, plotted p.u. of peak i at nominal network voltage.

converter was then subjected to a series of tests. In each case, the operating network voltage was set at differing levels within the 100%–15% range and maintained at the chosen level for the duration of the 90 ms test. A dc-link current disturbance was then applied and the response of the WTG converter observed. From Fig. 11(a), it can be seen that as the network voltage deviates from its nominal value within the range specified by the grid code, the network current response of the WTG to a change in load current does not change by any significant level. Fig. 11(b) shows the percentage change in peak current compared with nominal peak levels at network voltages from 0%–100% of nominal. It can be seen that as the network voltage reduces from 100% to 15% of its nominal level, the peak current flowing through the network side inverter of the WTG changes by less than half a percent. As demonstrated in this test, the nonlinear controller ensures predictable network current levels over the entire range of network voltages defined by the grid code fault ride-through specifications. The primary reason for the improved performance of the nonlinear controlled WTG is that the transformation block in the feedback linearization controller inputs both the load current and the network voltage level. When a change in one of these parameters occurs, the linear relationship between the new inputs and the plant outputs is maintained. Therefore, provided that there is a good match between the WTG plant and model, the feedback linearization controller should be expected to achieve superior ride-through capability when compared with the conventional cascade controller. The problem of plant-model mismatch requires careful consideration when implementing on an operational WTG a modelbased controller such as that presented here. In order to examine

the effect of plant model mismatch, a series of simulations was performed, which included plant-model mismatches for and in the range . This range of mismatch was chosen to account for possible differences between assumed and actual and values measured between the WTG converter terminals and the HV terminals of the grid-connected transformer. While transmission or distribution network and values may change significantly during fault conditions, providing that the fault does not occur within the wind turbine itself, the and values noted here will not change significantly during a fault. range should represent a Therefore, mismatches in the reasonable range within which to examine the effect of discrepancies between model and nameplate data. It can be seen from Fig. 12 that plant model mismatches of up to 20% at a network voltage level of 15% of nominal will result in a maximum increase in peak current levels of less than 3% above nominal. It can also be seen from the figure that underestimation of inductance levels tends to result in increased peak currents, a consequence of a slower response with reduced damping, while an underestimation of resistance levels tends to result in lower peak currents, a consequence of a faster controller response with reduced oscillation. Despite a minor variation in current levels, this investigation illustrates that in an operational WTG, the nonlinear controller proposed here would be expected to outperform a cascade controller, even with considerable levels of plant-model mismatch. IX. CONCLUSION Modern fault ride-through requirements stipulated in transmission and distribution grid codes require WTGs to operate at voltage levels far lower than was previously required. As operating voltages reduce, increased consideration must be given to the design of wind-turbine controllers. This is particularly true for wind turbines that employ power converters in their design. Conventional controllers for power converters while operating satisfactorily at normal network voltage levels have operating characteristics that result in excessive currents at reduced voltage levels. This phenomena is due to the fact that a linear controller design is typically applied to control power


converters whose operation is nonlinear in nature. If a nonlinear controller design is applied, a considerable improvement to the fault ride-through capability of wind turbines results. The feedback linearization control technique has been used in a wind-turbine-based power converter and is shown to maintain current levels within their design limits, even at greatly reduced voltage levels. This improvement has also been illustrated during considerable levels of plant-model mismatch, thus demonstrating a robust enhancement to the WTGs fault ride-through performance when compared with the conventional controller design. REFERENCES [1] (2004) Wind Farm Power Station Grid Code Provisions. WFPS1. [Online]. Available: http://www.eirgrid.com. [2] P. Gardiner, H. Snodin, A. Higgins, and S. M. Goldrick, “The Impacts of Increased Levels of Wind Penetration on the Electricity Systems of the Republic of Ireland and Northern Ireland,” Garrad Hassan and Partners Limited, Glasgow, U.K., Tech. Rep. 3096/GR/04, 2003. [3] C. Chompoo-inwai, C. Yingvivatanapong, K. Methaprayoon, and W.-J. Lee, “Reactive compensation techniques to improve the ride-through capability of wind turbine during disturbance,” IEEE Trans. Ind. Appl., vol. 41, no. 3, pp. 666–672, May/Jun. 2005. [4] J. Morren and S. W. de Hann, “Ridethrough of wind trubines with doubly-fed induction generator during a voltage dip,” IEEE Trans. Energy Convers., vol. 20, no. 2, pp. 435–441, Jun. 2005. [5] G. Saccomando, J. Svensson, and A. Sannino, “Improving voltage disturbance rejection for variable-speed wind turbines,” IEEE Trans. Energy Convers., vol. 17, no. 3, pp. 422–428, Sep. 2002. [6] J. Duran-Gomez and P. Enjeti, “Effect of voltage sags on adjustable-speed drives: A critical evaluation and an approach to improve performance,” IEEE Trans. Ind. Appl., vol. 35, no. 6, pp. 1440–1449, Nov./Dec. 1999. [7] C. Klumpner and F. Blaabjerg, “Experimental evaluation of ride-through capabilities for a matrix converter under short power interruptions,” IEEE Trans. Ind. Electron., vol. 49, no. 2, pp. 315–324, Apr. 2002. [8] J. Holtz and W. Lotzkat, “Controlled AC drives with ride-through capability at power interruption,” IEEE Trans. Ind. Appl., vol. 30, no. 5, pp. 1275–1283, Sep./Oct. 1994. [9] D. C. Lee, “Advanced nonlinear control of three-phase PWM rectifiers,” Proc. Inst. Elect. Eng., Elect. Power Appl., vol. 147, no. 5, pp. 361–366, Sep. 2000. [10] H. Sugimoto, S. Moritomo, and M. Yano, “A high performance control method of a voltage type PWM converter,” in Proc. IEEE PESC Conf. Record, Kyoto, Japan, 1988, pp. 360–368. [11] R. Kennel and A. Linder, “Predictive control of inverter supplied electrical drives,” in Proc. Power Electronics Specialist Conf., vol. 3, 2000, pp. 716–766. [12] J. E. Slotine, Applied Nonlinear Control, 1st ed. Englewood Cliffs, NJ: Prentice-Hall, 1991.

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[13] D. C. Lee, “DC-bus voltage control of three-phase AC/DC PWM converters using feedback linearization,” IEEE Trans. Ind. Appl., vol. 36, no. 3, pp. 826–833, May/Jun. 2000. [14] J. Jung and S. Lim, “A feedback linearizing control scheme for a PWM converter-inverter having a very small DC-link capacitor,” IEEE Trans. Ind. Appl., vol. 35, no. 5, pp. 1124–1131, Sep./Oct. 1999. [15] P. Kundur, Power System Stability and Control. New York: EPRI, McGraw-Hill, 1994. [16] L. Malesani and L. Rossetto, “AC/DC/AC PWM converter with reduced energy storage in the DC link,” IEEE Trans. Ind. Appl., vol. 31, no. 2, pp. 287–292, Mar./Apr. 1995. [17] D. Schreiber, “State of the art of variable speed wind turbines,” in Proc. 11th Int. Symp. Power Electronics, Oct. 2001, pp. 1–4. [18] T. A. Belli, R. P. O’Leary, and E. H. Camm, “Evaluating capacitorswitching devices for preventing nuisance tripping of adjustable-speed drives due to voltage magnification,” IEEE Trans. Power Del., vol. 11, no. 3, pp. 1373–1378, Jul. 1996.

Alan Mullane (S’01–M’03) received the B.E. degree in electrical and electronic engineering in 1998 and the Ph.D. degree in electrical engineering in 2003, both from the Department of Electrical and Electronic Engineering, University College Cork, Cork, Ireland. In 2004, he joined the Electricity Research Centre, University College Dublin, Dublin, Ireland, as a Postdoctoral Research Fellow. His research interests include nonlinear modeling and control of dynamic systems, with particular interest in simulation and control of wind turbines and their integration into electrical networks.

Gordon Lightbody received the M.Eng. degree with distinction in 1989 and the Ph.D. degree in electrical and electronic engineering in 1993 from Queens University Belfast, Belfast, U.K. His research interests include nonparametric modeling, local model networks for process modeling and control, model-based predictive control, fuzzy/neural systems, and nonlinear control. This work focused on key application areas, including wind power, power system control, and harmonic analysis.

R. Yacamini, photograph and biography not available at the time of publication.