Vector Control of Front-End Converters for Variable ... - IEEE Xplore

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 53, NO. 4, AUGUST 2006

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Vector Control of Front-End Converters for Variable-Speed Wind–Diesel Systems Roberto Cárdenas, Member, IEEE, Rubén Peña, Member, IEEE, Marcelo Pérez, Jon Clare, Senior Member, IEEE, Greg Asher, Senior Member, IEEE, and Fernando Vargas, Student Member, IEEE

Abstract—This paper presents a novel power-balance control method for a wind–diesel generation feeding an isolated grid. The system is based on a variable-speed wind energy conversion system (WECS) connected to an ac load using a power converter. An energy storage system (ESS), connected to the ac load using an additional converter, is used to balance the power generated by the WECS with the load. In this paper, the vector control systems for both interfacing power converters are discussed; the control uses the WECS converter to regulate the ac load voltage and the ESS converter to regulate the power flow to achieve a power balance. A small signal model is used to design the control systems. Finally, the proposed control is implemented in a 2-kW experimental prototype and the experimental results are fully analyzed and discussed in the paper. Index Terms—Diesel engines, pulsewidth-modulated (PWM) power converters, wind-power generation.

I. I NTRODUCTION

V

ARIABLE-SPEED operation of wind turbines has many advantages that are well documented in the literature [1]–[3]. Torque peaks in the gearbox and shafts are reduced, the wind turbine can operate with maximum aerodynamic efficiency and power fluctuations can be absorbed as inertial energy in the blades. In some applications, the wind turbine is augmented by an additional source, usually a diesel generator [4]–[7]. These generation schemes are called wind–diesel systems. In wind–diesel systems, wind speed variations may produce not only power fluctuations but also frequent start/stop cycles of the diesel engine in response to periods of unacceptably low wind speed. Simulation results presented in [5] illustrated than only 2 min of storage (i.e., to be able to supply the load for 2 min without diesel generation) can reduce the number of diesel starts from 30 per hour to two per hour, with a consequent reduction in overall fuel consumption. Therefore, in wind–diesel systems, an energy buffer is very important in order to avoid unnecessary deterioration of the diesel engine.

Manuscript received December 29, 2004; revised May 16, 2005. Abstract published on the Internet May 18, 2006. This work was supported by the Chilean Government through Fondecyt Grant 1020721. R. Cárdenas, R. Peña, M. Pérez, and F. Vargas are with the Electrical and Electronics Engineering Department, University of Magallanes, Punta Arenas, Chile (e-mail: [email protected]). J. Clare and G. Asher are with the School of Electrical and Electronic Engineering, University of Nottingham, NG7 2RD Nottingham, U.K. (e-mail: [email protected]). Digital Object Identifier 10.1109/TIE.2006.878321

Power smoothing for wind–diesel system in which the wind energy conversion system (WECS) and energy storage system (ESS) shared a common dc bus was considered in [8] and [9]. This paper addresses a system in which the WECS and ESS are connected to an isolated ac grid as depicted in Fig. 1. A variable-speed WECS is used to supply electrical energy into a stand-alone load. A power converter is used in the WECS side, to control the electrical torque of the generator, driving the wind turbine to the curve of maximum power capture for a given wind speed. The control of the WECS electrical generator and its associated power converter can be found in [1], [2], and [10] and will not be repeated here. Previous publications related to wind–diesel systems for stand-alone or isolated ac grid systems mostly consider a fixed speed wind turbine for the WECS [4]–[7] although the operation of variable-speed systems have been reported in [11]–[14]. In [11] and [12], simulation results of a wind–diesel system based on a variable-speed wind turbine are presented in which continuous operation of the diesel engine regulates the voltage and frequency at the grid, and power balancing is achieved by regulating the energy captured by the WECS using pitch control. However, continuous operation of the diesel is not desirable because of the high fuel consumption, while power balancing through controlling the pitch angle of the blades has relatively slow dynamics. In [13], a simulation study investigates a wind turbine and the diesel engine, both operating at variable speed for maximum wind energy capture and optimal diesel fuel consumption; however, the paper does not consider the control for power balancing, grid voltage, and frequency regulation. Another simulation study [14] investigates a variable-speed wind turbine connected to an isolated ac grid with a power smoothing system. Scalar control is used in the WECS front-end converter. This paper extends [14] through a rigorous presentation of the dual vector control schemes and its verification through experiment. In this paper, the control of a variable-speed WECS and an energy buffer are considered (see Fig. 1) for the operation in a wind diesel system. Energy is stored and released from the energy buffer, via an ESS and interfacing converter, to match the power absorbed from the wind with the load power. A flywheel or batteries are used for energy storing [15]. For sustained periods of low wind power, the flywheel speed or battery charge will drop below a threshold and the diesel generator is started, synchronized and connected to the load. When the flywheel speed or battery charge is above an upper threshold, the surplus energy has to be dissipated using a resistive load or the energy capture has to be reduced using

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Fig. 1. Wind–diesel system with a variable-speed wind turbine.

pitch control of the blades. It is assumed throughout the paper that the diesel generator is disconnected, since the main goal of the ESS is to reduce to a minimum the need to operate the diesel. It is assumed however that the diesel-generator set can be synchronized to the grid voltage and switched on when the wind energy, supplemented transiently by the ESS, becomes too low to supply the load. When the diesel generation is connected the grid voltage and frequency is then controlled by conventional exciter and governor control. This paper addresses the control systems for the two interfacing or “front-end” converters, shown in Fig. 1. The WECS front-end converter is used to regulate the voltage and frequency at the load. The electrical frequency is regulated using vector control of the WECS front-end converter orientated in a rotating axes system whose angular velocity is ωs . The ESS front-end converter is used to balance the power by regulating the dc-link voltage EG of the WECS converter to a constant value (maintaining IG ≈ IW in Fig. 1). Both, the WECS and the ESS, front-end converters are vector controlled to achieve fast and decoupled control of the active and reactive power supplied to the grid/load. The ESS machine-side converter (or the equivalent power electronics associated to a battery bank) is controlled to regulate the dc-link voltage EC (maintaining IC ≈ If in Fig. 1). The regulation of the dc-link voltage EC , using the ESS machineside converter, has been discussed in [8] and [9], and details will not be given here. This paper considers a balanced linear R–L grid load for verifying the basic idea of the control strategy. Under unbalanced phase loading, negative sequence and zero sequence distortion may occur creating unbalanced phase voltages with unequal phase shifts, together with 2ωs oscillations in the current and voltage d−q components. For sourcing power to unbalanced loads, four-leg WECS and/or ESS converters may be used in which the fourth leg connects to the load neutral [16], [17]. Nonlinear loads are normally handled through filtering. The application of the proposed control system to unbalanced and/or nonlinear loads, is considered outside the scope of this paper. Section II discusses the vector control system used in this paper. Section III discusses the small signal model, and in Section IV, the experimental results, obtained from a 2-kW experimental prototype, are presented and discussed.

Fig. 2.

Control system for the WECS front-end converter.

II. C ONTROL S YSTEMS FOR THE F RONT -E ND C ONVERTERS A. Control of the WECS Front-End Converter In this paper, the WECS front-end converter is used to regulate the frequency and voltage at the load. The vector control is based on a rotating axis system whose angular velocity ωs is set in the controller and defines the electrical frequency at the load. Fig. 2 shows the vector control structure of the WECS front-end converter. The load is represented as a series resistor RL and inductor LL . A second-order LC filter is used to reduce the load voltage harmonics; to avoid resonance problems, the resonance frequency of the filter is set to more than ten times the line frequency and less than one half of the switching frequency [18]. The design of the voltage control loop has to consider that the load is varying over a wide range, and that the ESS frontend converter is connected to the load. The small-signal model used to design the voltage control system, and the interaction between both front-end converters, are discussed in Section III.

CÁRDENAS et al.: VECTOR CONTROL OF FRONT-END CONVERTERS FOR WIND–DIESEL SYSTEMS

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III. S MALL -S IGNAL M ODEL A. Small-Signal Model for a Resistive Load Assuming that the power factor of the RL −LL load is close to unity, the reactive power supplied by the front-end converter is negligible, i.e., the quadrature components iqG , iqC of the WECS and ESS front-end converters are both zero. Considering a constant RL , the d−q equations for the ac side of Fig. 1 are idG + idC = idL

(2)

vd idG EG

(3)

vd = idL RL

(4)

IW ≈

where idG is the direct component of the current supplied by the WECS front-end converter, idC is the direct component of the current supplied by the ESS front-end converter, and vd , idL are the direct component of the load voltage and current, respectively. The dc-link voltage EG is obtained as Fig. 3.

Control of the ESS front-end converter.

EG =

The current control loops are standard PI [19] and have been designed for a closed-loop natural frequency of ≈ 70 Hz. B. Control of the ESS Front-End Converter The power captured by the wind turbine may be represented by IG in Fig. 1, while the power supplied by the WECS to the rest of the system may be represented by IW . If the generation and the load are balanced, then IG ≈ IW and the ESS does nothing. When the generation and the load are not balanced, the ESS supplies (or absorbs) energy to maintain the balance in IG and IW . The ESS front-end converter is vector controlled using a d−q frame orientated along the load voltage vd , as shown in Fig. 3. The relevant equations are didC − iqC ωe Lf + vd dt diqC + idC ωe Lf = iqC Rf + Lf dt

vqp

(1)

where vdp , vqp are the d−q voltages at the ESS front-end converter, idC , iqC are the d−q components of the “compensating” current iC to/from the ESS (see Fig. 1), vd is the load voltage, and Rf , Lf are the equivalent resistance and inductance between the ESS front-end converter and the load. Power balance is achieved by regulating the WECS converter’s dc-link voltage EG . When the dc-link voltage decreases, the idC component of the ESS front-end converter (which represents real power flow) is controlled to supply energy into the system. Conversely, if EG increases, idC is controlled to absorb energy from the system. The design of the dc-link voltage (EG ) controller is discussed in Section III. For the ESS front-end converter, the current control loops have been designed for a natural frequency of about 70 Hz.

(5)

where CG is the WECS converter’s dc-link capacitance. Using (1)–(5), the control diagram of Fig. 4(a) is obtained. In Fig. 4, GcE (s) is the dc-link voltage controller shown in Fig. 3 and GcV (s) is the controller of the ac load voltage. The symbols π and ÷ represent multiplication and division, respectively. For simplicity, the current control loop dynamics have been neglected in Fig. 4. The control systems of Fig. 4 are nonlinear with coupling between the dc-link voltage control loop and the load voltage control loop. Linearizing around a quiescent point (EG0 , vd0 , idL0 , IG0 , IW 0 ) yields ∆idG = − ∆idC ∆IW

vdp = idC Rf + Lf

IG − IW sCG

∆vd idG0 + ∆idG vd0 vd0 idG0 = − ∆EG 2 EG0 EG0

∆EG =

∆IG − ∆IW . sCG

(6) (7) (8)

During normal operation, the small signal model is simplified by considering that the variation in ∆EG , ∆vd are small compared to the variation in the currents. Therefore vd0 idG0 vd0 idG0 ∆idG ∆vd − ∆EG . 2 EG0 EG0 EG0

(9)

Using (6)–(9), the small-signal model shown in Fig. 4(b) is obtained. In the control system of Fig. 4(b), there is still some coupling between the dc-link voltage control loop and the load voltage control loop. However, because of the high inertia of wind turbines [1], [2], [20], the current idC will vary slowly compared with the natural frequency of the load voltage control loop. Therefore, for the dc-link voltage control system the load voltage is considered almost constant and, using (6), the open

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B. Control System Operation Considering a Low-Power-Factor Load The transfer function (12) is obtained considering a resistive load. For the general series RL −LL case, the transfer function of the load and filter capacitance is obtained as vd idG (s + RL /LL ) ≈ . 2 vq Cf (s + (RL /LL )s + 1/(LL Cf )) iqG

(13)

In (13), the cross coupling terms between the d and q axes have been neglected and vd , vq are the d−q components of the voltage. If the load has a near-unity power factor, a PI controller on each axis can be used to control the load voltage. However, if the power factor is low, the zero of (13) may be too close to the PI controller integrator and a slow dynamic response will result. Therefore, a PI controller may be not appropriate to regulate the voltage for loads with a very low power factor. The approach used in this paper is to design the load voltage controllers of Fig. 4(b) using a nominal load RL0 and then to test the response over the desired load range. Using this approach, we have designed and experimentally tested the control system for loads between 2% and 100% of the nominal value and power factors > 0.85. For more demanding conditions, a more sophisticated approach may be needed. However, such consideration is beyond the scope of this paper.

IV. E XPERIMENTAL R ESULTS Fig. 4. Block diagram for the proposed control system. (a) Control system proposed. (b) Small signal model of the control system proposed.

loop transfer function ∆EG /∆idC is obtained as ∆EG vd0 ≈ ∗ . ∆idC sCG EG

(10)

Using (10), a PI controller [for GcE (s)] can be designed for a bandwidth of about 4 Hz. The dc-link voltage controller can be compensated against variation in the term vd /EG using a variable gain (see Fig. 3). The open loop transfer function for the load voltage control loop can also be obtained from Fig. 4(b) as ∆vd ≈ RL . ∆idG

(11)

Considering the filter capacitance, and neglecting the cross coupling between the d and q axes, the open loop transfer function ∆vd /∆idG is ∆vd RL ≈ ∆idG 1 + sRL Cf

(12)

where Cf is the filter capacitance. Using (12), a PI controller [for GcV (s)] can be designed for the operation of the load voltage control loop.

Fig. 5 shows the experimental system implementing that of Fig. 1. The diesel generator has been excluded, since the control philosophy of the paper is not used when the diesel generator is switched in. The control structures were implemented on a DSP board based on a TMS320C31 processor. Two vector-controlled 2-kW pulsewidth-modulation (PWM) inverters with a switching frequency of 1 kHz are used. A 1-kHz chopper-based system is used to emulate the variable-speed wind turbine, supplying a current profile IG to the WECS dc link capacitors; ∗ is sent from the host PC this profile for the demand current IG to the DSP board. Four voltage transducers are used to measure the load voltage, and the dc-link voltages EG and EC . Four current transducers are used to measure the currents supplied by the front-end converters. An additional current transducer is used in the chopper-based IG control system. The frontend converters are connected to the load using 12-mH 0.3-Ω filter inductances. The capacitor bank is selected for a resonant frequency of approximately 500 Hz. In a real application, a flywheel system or a battery bank is connected to the ESS dc link capacitors. In the experimental system considered here, the battery bank is replaced by an additional chopper-based system. This is used to regulate the voltage in the ESS dc link capacitors, allowing bidirectional power flow in the ESS front-end converter. More information about the parameters of the experimental system is presented in the Appendix. Unless otherwise stated, all experimental tests are carried out with the demand current i∗qC = 0; all reactive power is supplied from the WECS front-end converter.


Fig. 5.

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Experimental system.

Fig. 6. Control system performance for a wind gust. (a) Current profile corresponding to a wind gust. (b) Current profile at the input of the system. (c) Current supplied by the WECS front-end converter and corresponding load voltage. (d) Current supplied by the ESS front-end converter and corresponding dc-link voltage.

A current profile corresponding to a typical wind gust is shown in Fig. 6(a). A succession of these is supplied to the WECS dc-link capacitors [as shown in Fig. 6(b)]. With a ∗ of 500 V, and RL chosen for a load power demand voltage EG of ≈ 720 W, Fig. 6(c) shows the direct component of the load voltage vd and the current idG . As expected, the shape of the current waveform is similar to that of the IG current. Fig. 6(d) shows idC supplied by the ESS front-end converter. The ESS supplies a current between ≈ 1 and 5 A. Note that idG + idC is approximately 5 A (the load current) for the entire duration

of the test corresponding to Fig. 6. The regulation of the load voltage vd and dc-link voltage EG is good. The dc-link voltage variations are below ±5 V, (about 1% of the reference voltage). For the load voltage, the voltage variations are also below ±5 V (about 4.5% of the reference load voltage). The system of Fig. 1 is also tested using a current profile obtained from a real wind profile lasting about 60 s. Fig. 7 a shows the wind profile and its corresponding current IG , the latter obtained by simulating a variable-speed wind turbine with an artificially zero inertia and applying the wind profile of

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Fig. 7. Control system performance considering a wind profile. (a) Wind profile and corresponding current profile. (b) Current supplied by the WECS front-end converter and load voltage corresponding to Fig. 7(a). (c) Current supplied by the ESS front-end converter and dc-link voltage corresponding to Fig. 7(a).

Fig. 7(a). Because wind turbines normally have large inertia, they behave like a low-pass filter [9], [20], [21], filtering out much of the wind turbulence. If the wind turbine has no inertia, the wind speed fluctuations are reflected on the generator output power. This represents a worst case condition for power smoothing since the high frequency content of IG is maximized and the natural smoothing normally associated with the turbine inertia is minimized. The current profile of Fig. 7(a) has a peak current of 3.5 A with an average value I¯G of 1.9 A and a dispersion coefficient σIG of 0.76 A. The mean turbulence intensity value, defined as σIG /I¯G , is about 40%. This is a very high value of turbulence for a variable-speed wind turbines [20], [21] and represents a worst case for testing purposes. Fig. 7(b) shows idG of the WECS front-end converter and the load voltage vd . The load is ≈ 720 W (about 35% of nominal) for t < 10 s. At t = 10 s, the load is stepped to ≈ 1900 W (95% nominal). After 50 s, the load is stepped again to 720 V. The shape of the idG is identical to that of the current IG , with a small disturbance at the sudden load variations. The dip and the overshoot on the vd voltage are ≈ 17 V. The response of the voltage controller is quite good considering the magnitude of the load step, and also considering that the control loops are not decoupled for fast load changes (see Fig. 4). When the load is fixed, there is small variation in the voltage vd produced by the IG current fluctuations. Fig. 7(c) shows the dc-link voltage EG and the direct component idC of the current supplied by the ESS front-end converter. Before the load step, the ESS front-end converter is absorbing energy from the system (idC < 0). After the load step, the additional load power is supplied from the ESS converter and the current idC changes from 0 to ≈ 3 A. When the load is disconnected, the current idC is again reduced to a low value.

Notice that at the end of the current profile, because of the low load, the ESS front-end converter is again absorbing energy from the system. Fig. 7(c) shows also the dc-link voltage EG , when the load is connected and disconnected. The dip and the overshoot are less than 15 V (3% of the nominal voltage). During application of the current profile, there is also a small variation of ±3 V in the dc-link voltage EG . Given the high turbulence of the IG current profile used, this variation is very small and shows the high performance of the proposed control system. In order to test the robustness of the control method, its performance is tested using a load with a low pf of 0.53 (20 Ω in series with 100 mH—both per phase). The shape of the current profile used in the experimental tests is identical to that of Fig. 7(a), but with an average value I¯G of approximately 1.3 A. Fig. 8(a) shows the d−q currents supplied by the ESS and WECS converters. For this test, the ESS front-end is absorbing much of the energy supplied from the WECS. This is because the load has a low-power factor and there is a large voltage drop in the load inductance. Fig. 8(b) shows the dc-link voltage EG and the load voltage vd . The performance of the control system is good even for this low-power-factor load. There is some increase in the switching noise and the oscillation magnitude has increased slightly respect to the performance of Fig. 7. Nevertheless, the experimental test of Fig. 8 shows the robustness of the proposed control system. Even when the power factor is relatively low, the performance obtained from the control system is still quite good. Fig. 9 shows the load voltage and line currents from both front-end converters. Fig. 9(a) shows the currents and voltage for a 1-kW load, with constant IG ≈ 1.5 A. Because the power supplied from the WECS is not sufficient for the load, the ESS


Fig. 8. Control system performance considering a load with low power factor. (a) d−q components of the front-end converters for a low-power-factor load. (b) Load voltage and dc-link voltage corresponding to Fig. 8(a).

Fig. 9. Voltage and currents for steady-state operation. (a) Load voltage and current for a 1-kW load. (b) Load voltage and currents for a 100-W load.

is supplying energy to the load. Therefore, iaG and iaC are in phase with the phase voltage va . Fig. 9(b) shows the voltage and currents (for identical input current) when the system load is reduced to ≈ 100 W. Because of the reduced load, most of the WECS energy is absorbed by the ESS front-end converter. Therefore, iaC in Fig. 9(b) is 180◦ out of phase respect to va . The experimental results

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Fig. 10. Control system operation for a step change in the input. (a) Load voltage and WECS currents for a step change in IG . (b) DC-link voltage and ESS current for a step change in IG .

of Fig. 9 also show that, even for wide variations in the load power, the harmonics content in the voltage and current are acceptable. The robustness of the proposed control system has also been tested by applying step changes in the input current IG . For a load power of ≈ 700 W, the input current is varied between ≈ 0 and 3.5 A, equivalent to an input power step of 1750 W. The experimental results are shown in Fig. 10. Fig. 10(a) shows the load voltage and idG , iqG of the WECS. As expected there is a sudden change in idG from ≈ 0 A to 5 A (there is also a small change in the iqG current) when the input power step is applied. The current idG changes again from ≈ 5 A to 0 A when the input power step is disconnected. For the experimental test of Fig. 10(a), the dip and the overshoot in the load voltage are ≈ 25 V and 35 V, respectively, which is acceptable considering the magnitude of the power step applied to the system. Fig. 10(b) shows the idC component of the ESS converter. Before the input power step, the ESS front-end converter is supplying most of the load power. When IG changes to 3.5 A, the ESS front-end converter absorbs the surplus energy from the system. Notice in Fig. 10(b) that the dip and the overshoot in the dc-link voltage are relatively small considering the magnitude of the power step applied to the system. The experimental test of Fig. 10 shows the robustness of the proposed control system. A power step is not realistic in wind energy applications. However, even for this drastic test, the performance of the proposed control system is still quite good. In order to test the performance of the control system, input disturbance current profiles IG are injected to the WECS dclink capacitors with a variable period T . Varying the period can test the control scheme for a number of conditions found in practice: wind gusts, high-turbulence wind speed variations,

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Fig. 11. Control system considering a current profile of variable frequency. (a) Current profile with variable period T . (b) DC-link-voltage variation corresponding to a fundamental frequency of 0.1 Hz. (c) DC-link-voltage variation corresponding to a fundamental frequency of ≈ 5 Hz.

tower shadow effects [21] and IG ripple content arising from the WECS generator converter. The current profile IG representing a power variation between 0 and ≈ 1 kW (0% to 50% of nominal power) was applied at two frequencies: ≈ 0.1 Hz representing a wind gust test [see Fig. 11(a)] and 5 Hz, representing, e.g., turbulence or tower shadow. As a worst case scenario, the WECS side capacitance is set at a relatively low value of CG ≈ 1270 µF. Fig. 11(b) and (c) shows the dc-linkvoltage control response to the 0.1 and 5 Hz disturbances, respectively. For 0.1 Hz, EG has ±5 V (1%) variation, with the variation in the direct component of the load voltage vd at 3 V. For 5 Hz, which is outside the controller bandwidth, the EG variation is 486–523 V (−2.8% to 4.6%), which is acceptable given the large power variation. At 20 Hz (not shown), an EG variation of 492–512 V (−1.6% to 2.4%) was observed with a vd load voltage variation of 5 V. The experimental results show that the maximum voltage oscillations occur for an IG disturbance frequency of ≈ 5 Hz. However, even in this case the regulation achieved by the proposed control system is good. V. C ONCLUSION This paper has presented a new control strategy for power balancing in a variable-speed wind generation or wind–diesel system feeding an isolated grid. The WECS comprises a variable-speed wind turbine and back-to-back power converters. The ESS system comprises a front-end power converter interfacing an energy storage medium (flywheel or batteries). The WECS is connected to the load using a vector controlled front-end converter, which regulates the load voltage and frequency. Power balancing is achieved by regulating the dc-link voltage of the WECS converter using the direct component of the current supplied by the ESS front-end converter. A complete small-signal model has been derived for the dynamics of the

ESS front-end converter, the WECS-side dc-link capacitors and the WECS front-end converter. A control system, based on the model, was developed. Experimental results, using real wind profiles, wind gusts, and power steps, have been presented and the results demonstrate excellent performance. For a typical wind profile the voltage regulation in the ac load is almost perfect. Even with a relatively large load step of 1.2 kW (for a 2-kW system), the dip and the overshoot of the load voltage are acceptably small. The control system has also been tested for operation with low-power-factor loads, step changes in the input power and disturbance current profiles with variable frequency content; the performance obtained is also good. The experimental and simulation results obtained in this paper are very promising and illustrate the advantages and improvements that can be expected when modern control techniques and power conversion are applied to wind–diesel systems. Although this paper has concentrated on a standalone system, the techniques presented are equally applicable to systems, which have a weak connection to a larger grid system. Further work will address the problem of unbalanced and nonlinear loads, which may be a particular problem for isolated three-phase grids. The paper considers a balanced linear R–L grid load for verifying the basic idea of the control strategy. A PPENDIX Natural frequency of the load voltage controller natural Frequency of the dc-link-voltage controller natural frequency of the WECS d−q current control systems natural frequency of the ESS d−q current control systems switching frequency of the PWM front-end converters

7 Hz; 4 Hz; 70 Hz; 70 Hz; 1 kHz;


frequency of the stand-alone grid resonant frequency of the second-order filter ∗ demand dc-link voltage EG ∗ demand load voltage vd dc-link capacitance dc-link capacitance for the experimental results shown in Fig. 11

50 kHz; ≈ 500 Hz; 500 V; 110 V; 2400 µF; 1270 µF.

ACKNOWLEDGMENT The authors would like to thank the British Council, through their academic links program, which has made the collaboration between the Universidad de Magallanes and the University of Nottingham possible. R EFERENCES [1] A. Miller, E. Muljadi, and D. Zinger, “A variable speed wind turbine power control,” IEEE Trans. Energy Convers., vol. 12, no. 2, pp. 181– 186, Jun. 1997. [2] E. Muljadi and C. Butterfield, “Pitch-controlled variable-speed wind turbine generation,” IEEE Trans. Ind. Appl., vol. 37, no. 1, pp. 240–246, Jan./Feb. 2001. [3] M. Steinbuch, “Optimal multivariable control of a wind turbine with variable speed,” Wind Eng., vol. 11, no. 3, pp. 153–163, 1987. [4] A. J. Rudell, J. A. M. Bleij, and L. Freris, “A wind diesel system with variable speed flywheel storage,” Wind Eng., vol. 17, no. 3, pp. 129–145, 1993. [5] R. Dettmer, “Revolutionary energy; A wind/diesel generator with flywheel storage,” IEE Rev., vol. 36, no. 4, pp. 149–151, Apr. 1990. [6] F. Hardan, J. Bleij, R. Jones, P. Bromley, and A. J. Rudell, “Application of a power-controlled flywheel drive for wind power conditioning in a wind/diesel power system,” in Proc. IEE 9th Int. Conf. Elect. Mach. and Drives, 1999, pp. 65–70. [7] I. J. Iglesias, L. Garcia, A. Agudo, I. Cruz, and L. Arribas, “Design and simulation of a stand-alone wind diesel generator with a flywheel energy storage system to supply the required active and reactive power,” in Proc. IEEE PESC, 2000, pp. 1381–1386. [8] R. Cárdenas, R. Peña, G. Asher, and J. Clare, “Power smoothing in generation systems using a sensorless vector controlled induction machine driving a flywheel,” IEEE Trans. Energy Convers., vol. 19, no. 1, pp. 206– 216, Mar. 2004. [9] ——, “Control strategies for enhanced power smoothing in wind energy systems using a flywheel driven by a vector controlled induction machine,” IEEE Trans. Ind. Electron., vol. 48, no. 3, pp. 625–635, Jun. 2001. [10] R. Cárdenas and R. Peña, “Sensorless vector control of induction machines for variable speed wind energy applications,” IEEE Trans. Energy Convers., vol. 19, no. 1, pp. 196–205, Mar. 2004. [11] S. Hurtado, G. Gostales, A. de Lara, N. Moreno, J. M. Carrasco, E. Galvan, J. A. Sanchez, and L. G. Franquelo, “A new power stabilization control system based on making use of mechanical inertia of a variablespeed wind-turbine for stand-alone wind–diesel applications,” in Proc. IEEE IECON, Nov. 5–8, 2002, vol. 4, pp. 3326–3331. [12] J. A. Sanchez, N. Moreno, S. Vazquez, J. M. Carrasco, E. Galvan, C. Batista, S. Hurtado, and G. Costales, “A 800 kW wind–diesel test bench based on the MADE AE-52 variable speed wind turbine,” in Proc. IEEE IECON, Nov. 2–6, 2003, vol. 2, pp. 1314–1319. [13] Z. Chen and Y. Hu, “A hybrid generation system using variable speed wind turbines and diesel units,” in Proc. IEEE IECON, Nov. 2–6, 2003, vol. 3, pp. 2729–2734. [14] R. Cárdenas, R. Peña, J. Clare, and G. Asher, “Power Smoothing in a variable speed wind–diesel system,” in Proc. IEEE PESC, Acapulco, Mexico, Jun. 2003, vol. 2, pp. 754–759. [15] R. Hebner and A. Walls, “Flywheel batteries come around again,” IEEE Spectr., vol. 39, no. 4, pp. 46–51, Apr. 2002. [16] R. A. Gannet, “Control strategies for high power four-leg voltage source inverters,” M.Sc. thesis, Virginia Polytechnic Inst. and State Univ., Blacksburg, VA, Jul. 30, 2001. [17] O. Ojo and P. M. Kshirsagar, “Concise modulation strategies for four-leg voltage source inverters,” IEEE Trans. Power Electron., vol. 19, no. 1, pp. 46–53, Jan. 2004.

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[18] T. Wang, Z. Zhilong, G. Sinha, and X. Yuan, “Output filter design for a grid-interconnected three-phase inverter,” in Proc. IEEE PESC, Acapulco, México, Jun. 2003, pp. 779–784. [19] R. Peña, R. Cárdenas, J. Clare, and G. Asher, “Control strategies for voltage control of a boost type PWM converter,” in Proc. PESC, Vancouver, BC, Canada, Jun. 2001, vol. 2, pp. 730–735. [20] J. A. M. Bleij, A. W. K. Chung, and J. A. Rudell, “Power smoothing and performance improvement of wind turbines with variable speed,” in Proc. 17th BWEA, Warwick, U.K., 1995, pp. 353–358. [21] W. E. Leithead, “Dependence of performance of variable speed wind turbines on the turbulence, dynamics and control,” Proc. Inst. Elect. Eng., vol. 137, pt. C, no. 6, pp. 403–413, Nov. 1990.

Roberto Cárdenas (S’95–M’97) was born in Punta Arenas, Chile. He received the Electrical Engineering degree from the University of Magallanes, Punta Arenas, Chile, in 1988, and the M.sc. and Ph.D. degrees from the University of Nottingham, Nottingham, U.K., in 1992 and 1996, respectively. From 1989 to 1991, he was a Lecturer at the University of Magallanes. He is currently with the Electrical Engineering Department, University of Magallanes. His main interests are in control of electrical machines, variable-speed drives, and renewable energy systems. Dr. Cárdenas is the principal author of the paper that received the Best Paper Award from the IEEE Industrial Electronics Society, for the best paper published in the IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS during 2004.

Rubén Peña (S’94–M’96) was born in Coronel, Chile. He received the degree in electrical engineering from the University of Concepcion, Concepcion, Chile, and the M.Sc. and Ph.D. degrees from the University of Nottingham, Nottingham, U.K., in 1984, 1992, and 1996, respectively. From 1985 to 1991, he was a Lecturer with the University of Magallanes, Punta Arenas, Chile. He is currently with the Electrical Engineering Department, University of Magallanes. His main interests include control of power electronics converters, ac drives, and renewable energy systems.

Marcelo Pérez was born in Punta Arenas, Chile. He received the degree in electrical engineering from the University of Magallanes, Punta Arenas, Chile, in 2003. He is currently a Research Assistant with the Electrical Engineering Department, University of Magallanes. His main interests include power electronics and digital control using DSPs.

Jon Clare (M’90–SM’04) was born in Bristol, U.K. He received the B.Sc. and Ph.D. degrees in electrical engineering from the University of Bristol, Bristol, U.K. From 1984 to 1990, he was a Research Assistant and Lecturer at the University of Bristol, involved in teaching and research in power electronic systems. Since 1990, he has been with the Power Electronics, Machines and Control Group at the University of Nottingham, Nottingham, U.K., and is currently Professor in Power Electronics and Head of Research Group. His research interests are power electronic converters and modulation strategies, variable-speed drive systems, and electromagnetic compatibility. Prof. Clare is a member of the Institution of Electrical Engineers, U.K., and is an Associate Editor of the IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS.

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Greg Asher (M’98–SM’04) graduated in electrical and electronic engineering and received the Ph.D. degree in bond graph structures and general dynamic systems from Bath University, Bath, U.K., in 1976 and 1979, respectively. In 1984, he was appointed Lecturer in Control in the School of Electrical and Electronic Engineering at the University of Nottingham, Nottingham, U.K., where he developed an interest in motor drive systems, particularly the control of ac machines. He was appointed Professor of Electrical Drives in 2000 and is currently Head of the School of Electrical and Electronic Engineering at the University of Nottingham. He has published over 180 research papers, and has received over $5M in research contracts. Prof. Asher was a member of the Executive Committee of the European Power Electronics (EPE) Association until 2003. He is an Associate Editor of the IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS and is currently Chair of the Power Electronics Technical Committee for the IEEE Industrial Electronics Society.

Fernando Vargas (S’03) was born in Punta Arenas, Chile. He received the Electrical Engineering degree from the University of Magallanes, Punta Arenas, Chile, in 2004. He is currently working toward the M.Sc. degree in the Electrical Engineering Department, University of Magallanes. His main interests are power electronics, control of electrical drives, and digital control using DSPs.