Integrating Hybrid Power Source Into an Islanded MV Microgrid Using ...

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Integrating Hybrid Power Source Into an Islanded MV Microgrid Using CHB Multilevel Inverter Under Unbalanced and Nonlinear Load Conditions Mohsen Hamzeh, Student Member, IEEE, Amin Ghazanfari, Student Member, IEEE, Hossein Mokhtari, Member, IEEE, and Houshang Karimi, Senior Member, IEEE

Abstract—This paper presents a control strategy for an islanded medium voltage microgrid to coordinate hybrid power source (HPS) units and to control interfaced multilevel inverters under unbalanced and nonlinear load conditions. The proposed HPS systems are connected to the loads through a cascaded H-bridge (CHB) multilevel inverter. The CHB multilevel inverters increase the output voltage level and enhance power quality. The HPS employs fuel cell (FC) and photovoltaic sources as the main and supercapacitors as the complementary power sources. Fast transient response, high performance, high power density, and low FC fuel consumption are the main advantages of the proposed HPS system. The proposed control strategy consists of a power management unit for the HPS system and a voltage controller for the CHB multilevel inverter. Each distributed generation unit employs a multiproportional resonant controller to regulate the buses voltages even when the loads are unbalanced and/or nonlinear. Digital time-domain simulation studies are carried out in the PSCAD/EMTDC environment to verify the performance of the overall proposed control system. Index Terms—Cascaded H-bridge (CHB) multilevel inverter, fuel cell (FC), hybrid power source (HPS), multiproportional resonant (multi-PR), photovoltaic (PV), supercapacitor (SC).

I. INTRODUCTION ICROGRIDS aim to provide a solution to reform the conventional power system toward a new concept for future energy distribution systems. A microgrid plays a key role for renewable energy integration and energy management capability improvement. The increasing use of medium voltage (MV) microgrid demands more reliable components and advanced control strategies [1]–[3]. A microgrid may inherently be subjected to significant degrees of unbalanced conditions due to the presence of singlephase loads and/or distributed generation (DG) units. Moreover, the inclusion of nonlinear loads in an islanded microgrid leads to

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Manuscript received November 3, 2012; revised February 7, 2013; accepted June 1, 2013. Paper no. TEC-00578-2012. M. Hamzeh, A. Ghazanfari, and H. Mokhtari are with the Center of Excellence in Power System Management and Control, Sharif University of Technology, Tehran, Iran (e-mail: [email protected]; [email protected]; [email protected]). ´ ´ H. Karimi is with the D´epartement de G´enie Electrique, Ecole Polytechnique de Montr´eal, Montr´eal, QC H3T 1J4, Canada (e-mail: [email protected]). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TEC.2013.2267171

various power quality problems, e.g., distortion of voltage and current. Nevertheless, a microgrid should be able to operate under unbalanced and nonlinear load conditions while complying with the existing standards, e.g., IEEE Std 141 and 519 [4], [5]. Based on the IEEE standards [4], [5], the voltage unbalance factor (VUF) and the voltage total harmonic distortion (THD) should be maintained within 2% and 5%, respectively, in a distribution network. Microgrids usually consist of multiple DG units interfaced to the system via power electronics inverters [6]. In order to improve power quality, utilization of multilevel inverters in an MV distribution network has attracted growing interests in the recent years. Due to modularity and ability to operate at higher voltage levels with negligible distortion, the cascaded H-bridge (CHB) multilevel inverters are preferred for high-power applications among the other topologies [7]. A hybrid power source (HPS) offers superb scalability and more flexibility for power management capability. Due to the clean and environment-friendly specifications, the fuel cell (FC) and photovoltaic (PV) systems are widely used as the main power sources. The slow dynamic response of the FC stack, intermittent nature of the PV, and quick load changes necessitate the use of supercapacitor (SC) as a storage system with high power density [8]. Therefore, parallel hybrid FC/PV/SC power source and CHB multilevel inverter are used by each DG unit. In comparison with the previous works, e.g., [8]–[11], the proposed HPS simultaneously ensures the enhancement of system modularity, transient response, and power quality of the MV microgrid, particularly, in the presence of nonlinear and unbalanced loads. Converter-based DG units may introduce harmonics into a microgrid and result in power quality issues. However, welldesigned and well-controlled converters are able to improve the power quality and efficiency of the microgrids. Besides the primary purpose of the DG units for power generation, many services can also be provided, e.g., voltage support, power factor correction, flicker mitigation, and harmonic and unbalance voltage compensation [12]–[18]. In this paper, a voltage controller for a CHB multilevel inverter is proposed to enhance dynamic response and power quality of microgrid in the presence of unbalanced and nonlinear loads. The multiproportional resonant (multi-PR) controller is used to regulate the load voltage. When the load is nonlinear, the use of a multi-PR controller is more advantageous as compared to the conventional PR controllers [17], [18].

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(a)

(b)

Fig. 2.

Fig. 1. (a) One-phase structure of a CHB multilevel inverter and (b) circuit diagram of a three-phase, three-wire DG unit.

The multi-PR controller compensates harmonic and negativesequence currents of nonlinear and unbalanced loads, respectively, and provides a set of sinusoidal balanced voltages at the microgrid buses. In addition, the proposed HPS configuration and the power management scheme can actively distribute the power demand among the main and auxiliary power sources of the DG unit even when unbalanced load switchings are imposed. The effectiveness of the proposed control strategy is verified through simulation case studies conducted in the PSCAD/EMTDC environment. II. DG UNIT STRUCTURE One-phase structure of a CHB multilevel inverter is shown in Fig. 1(a), where n multilevel inverter modules are connected in series to obtain an output voltage waveform of 2n + 1 steps. The inverter output phase voltage is the sum of the output voltages of n H-bridge cells with isolated dc sources. Fig. 1(b) shows the circuit diagram of a three-wire DG subsystem whose detailed mathematical model is described in [8]. To detect a phase-toground fault, the multilevel inverter modules are star connected with a grounded neutral [19]. The structure of the hybrid FC/PV/SC power source is shown in Fig. 2. Each HPS is connected to the dc link of a specified cell of the CHB multilevel inverter. The proposed HPS consists of proton exchange membrane FC stacks and PV arrays which together provide the main power, and the SC modules which accommodate the fast transients in power demands. The SC modules also guarantee the power quality and reliability of the HPS, and ensure proper functionality of the corresponding microsource. To maximize the fuel savings of the FC stacks, the

Proposed structure of the hybrid FC/PV/SC power source.

PV arrays must share the maximum possible portion of power demands despite their intermittent nature. In order to achieve high efficiency and galvanic isolation, the power sources of each HPS are connected to the dc link of multilevel inverter modules through full-bridge dc/dc converters. The unidirectional and bidirectional full-bridge converters of the FC and SC units make the output current smooth and regulate the output voltages of the units at the desired values. Moreover, the bidirectional power flow of the SC module increases power management flexibility. The full-bridge converter of the PV unit is controlled such that the maximum power point tracking (MPPT) is achieved. III. OPERATION PRINCIPLES OF THE PROPOSED CONTROL STRATEGY The proposed control strategy comprises 1) a power management for the HPS system, and 2) a voltage control for the CHB multilevel inverter. To manage the power and regulate the dc-link voltage of the HPS unit, two independent controllers are designed. Furthermore, a voltage control loop is proposed to provide a set of balanced sinusoidal voltages at the terminals of CHB multilevel inverter in the presence of nonlinear and unbalanced loads. A. Control Strategy of the HPS The proposed control strategy of the hybrid FC/PV/SC power source is shown in Fig. 3. The HPS uses the FC and PV units as the main power sources and the SC as the complementary power source. The PV unit enables the FC to obtain an appropriate operating point at which the hydrogen consumption is minimized. The SC modules support the FC and PV to achieve good transient response and meet the grid power demand. The utilization of three separate full-bridge converters in parallel facilitates the power management capability and increases the overall performance and flexibility of the HPS. The HPS controller is designed such that the SC converter regulates the dc-link voltage, and the FC and PV converters fulfill the dc-link power demand.

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Fig. 3.

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Proposed control strategy of hybrid FC/PV/SC power source.

The unidirectional power flow of the FC and PV converters results in decoupled dynamics for FC, PV, and SC systems. Therefore, the control design for each converter is carried out individually. Parameters of the SC and FC controllers are determined by the appropriate selections of bandwidth and phase margin using MATLAB Control Toolbox. According to the proposed control strategy, the dc current of the SC module must accurately follow its reference to zero. A PI controller determines the duty cycle of the FC converter. The reference signal generated by the controller is limited not to exceed the FC capability in injecting the current. The corresponding limitation for the current demand is calculated according to the typical range of utilization factor, which ensures the desired operation of FC stack. Furthermore, the FC current slope is limited to avoid the fuel starvation phenomena and to guarantee the safe operation of the FC stack. For control design purposes, the dc/dc converters are modeled using the state-space averaging technique. Based on the average model of the full-bridge converter, the FC control-to-input current transfer function is obtained as [20]: iˆL = 1+ dˆ

2I L RFC CFC S) (1−D ) (1 + 2 n2 LFC n2 LFC CFC R F C (1−D ) 2 S + (1−D ) 2

S2

(a)

(b)

(1)

where IL is the inductor current, D is the nominal duty cycle, n is the transformer winding ratio, and LFC , CFC , and RFC are, respectively, the inductor, capacitor, and equivalent output resistor of the FC converter. The bode diagrams of the closedloop transfer functions are shown in Fig. 4. As it is observed, the closed-loop systems show good robust stability margins. To attenuate the current ripple of the downstream inverter and to accommodate the slow dynamics of the FC stack, parameters of the controller are designed such that the current-loop bandwidth is more than 628.3 rad/s [21]. Therefore, as seen from Fig. 4(a), the current-loop bandwidth of the FC converter is set to 4360 rad/s to obtain a phase margin of 88.4◦ . The FC processor plays a vital role in regulating hydrogen flow according to the output

Fig. 4. Bode diagrams of closed-loop transfer functions. (a) FC current-loop response. (b) SC voltage-loop response.

power from the FC stack. The detailed mathematical model of the FC processor is described in [22]. The SC voltage control loop regulates the dc-link voltage using a PI controller. When the SC is charging (discharging), the duty cycle of the SC converter will decrease (increase) to maintain the dc-link voltage regulation. Based on the average model of the full-bridge converter, the SC control-to-output voltage

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TABLE I PARAMETERS OF CONTROLLERS

phase-shifted PWM strategy balances the dc capacitors voltages and mitigates the input current harmonics of the CHB multilevel inverter. According to the internal model principle, a reference (disturbance) can be asymptotically tracked (rejected) if the controller contains the Laplace transform of the reference signal in its transfer function. The output currents (Ioα and Ioβ ), which can be considered as disturbances in the control system, contain fundamental and higher order harmonics when the load is nonlinear. Notice that since the loads are connected to the microgrid buses via Y /Δ transformers, neither zero-sequence nor third-order harmonic currents exists in the inverter side of the DG units. To achieve zero steady-state error in the presence of harmonic currents, a multi-PR controller is proposed as follows:   (3) K(s) = C(s) G1 (s) + G5 (s) + G7 (s) where C, G1 , G5 , and G7 are

Fig. 5.

C(s) = k1

Block diagram of the proposed multi-PR controller.

s + k2 s + k3

G1 (s) = k4

transfer function is obtained as [20]

s2 + k5 s + k6 s2 + ωc s + ω02

2

Vc n LS C Vˆc (1−D ) (1 − R S C (1−D ) 2 S) = 2L n2 LS C CS C 2 SC 1 + R SnC (1−D dˆ ) 2 S + (1−D ) 2 S

(2)

where Vc is the output voltage of capacitor. The frequency response of the compensated SC system is shown in Fig. 4(b). The parameters of the designed controllers are listed in Table I. B. Control Strategy of the Inverter The circuit diagram representation of the CHB-based DG subsystem is given in Fig. 1(b). The dynamic model of a threewire inverter in the stationary reference frame (αβ-frame) is given in [16]. The ac-side controller should robustly regulate the load voltages in the presence of load uncertainties and the external disturbances. It should be noted that since the reference signals are sinusoidal in the αβ-frame, the use of a PR controller is more advantageous as compared to a PI controller [16], [23]. Fig. 5 shows the block diagram of the proposed voltage controller in the αβ-frame. The magnitude and frequency of the reference voltage (Vα∗ and Vβ∗ ) are determined by the droop controller [16]. To protect the inverter against over current and to increase the internal stability of the voltage control loop, an inner current loop is also incorporated. The current controller is a simple gain, kc , whose value is calculated such that the damping factor of the dominant poles of the inner loop system becomes 0.7. To eliminate the impact of load dynamics, the output currents, i.e., Ioα and Ioβ , are feedforward to the output of the voltage control loop. The resultant signals are then applied to the current controllers to generate the control signals Uα and Uβ . Finally, the control signals in αβ-frame are transformed to the abc-frame and then applied to the modulation unit. The phase-shifted pulse width modulation (PWM) reference required is used as the modulation strategy since it provides an even power distribution among the units [24]. Moreover, the

G5 (s) = k7

s2

G7 (s) = k10

s2 + k8 s + k9 + 5ωc s + (5ω0 )2

s2 + k11 s + k12 . s2 + 7ωc s + (7ω0 )2

(4)

G5 (s) and G7 (s) are harmonic compensators, and C(s) is a lead compensator which is employed to guarantee the robust stability of the closed-loop voltage control system. According to the bandwidth of the voltage control system (400 Hz), only the fifth- and seventh-order harmonics can be compensated. In (4), the coefficients ki , i = 1, . . . , 12, are the design parameters of the multi-PR controller. To obtain the coefficients ki s, the following performance characteristics are to be met. 1) The closed-loop system achieves good stability margins. 2) The bandwidth of the open-loop system should be less than 10% of the switching frequency. 3) The reference should be tracked within two cycles with zero steady-state error. 4) The disturbance (harmonic currents) should be rejected. Considering the aforementioned performance indices and using MATLAB SISO tools, the coefficients of the controller are designed and listed in Table I. The frequency response of the open-loop controlled system considering is shown in Fig. 6. It is clear that the desired phase margin, gain margin, and bandwidth for the system are achieved. The phase margin is almost 30◦ , gain margin is infinite, and the bandwidth of the system is close to 400 Hz. In addition, the gain of the system at 50 Hz and the other harmonic frequencies is high which results in significant disturbance rejection. IV. SIMULATION RESULTS To investigate the effectiveness of the proposed control strategy, the microgrid system of Fig. 7 is simulated in the

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TABLE II PARAMETERS OF THE MICROGRID’S SUBSYSTEMS

(b)

Fig. 6.

Bode plots of loop transfer function of controlled system.

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Fig. 7.

Single-line diagram of MV microgrid consisting of two DG units.

PSCAD/EMTDC environment. The microgird system is composed of a three-feeder distribution system and two DG units. Each DG unit is connected to the corresponding feeder using a CHB multilevel inverter. For the sake of simplicity, each DG unit employs a two-cell CHB multilevel inverter. The loads are connected to the feeders via Y /Δ transformers. It is assumed that the microgrid system operates in the islanded mode. Each CHB multilevel inverter is equipped with the proposed multiPR controller, HPS power management, and a droop control strategy. The slope of FC current is limited to ±0.0625 p.u.s−1 to prevent the fuel starvation phenomenon. The PV system is equipped with an MPPT control strategy. Maxwell Technologies Boostcap BMOD0165-type SC is used as the energy storage. The dc-link voltage of each HPS system is regulated at 1 kV. The microgrid parameters are given in Table II. Initially, the microgrid is operating under balanced and linear load conditions. At t = 4 s, a six-pulse diode rectifier with 570 kVA and PF = 0.95 is connected to the LV side of feeder F1 . Subsequent to the first load change, phase-c of the LV side of feeder F1 is disconnected from the microgrid at t = 10 s. The instantaneous real and reactive powers of the feeders with respect to the load changes are shown in Fig. 8. Due to the presence

Fig. 8. Microgrid response to unbalanced and nonlinear load changes in feeder F1 . (a) and (b) Instantaneous real and reactive powers of feeders.

of unbalanced and nonlinear loads, the double- (100 Hz) and high-frequency ripples are introduced in the power components of feeder F1 . Fig. 9 shows the positive- and negative-sequence components of the load currents, and the harmonic currents of one phase at the LV side of the feeders. The positive-sequence component and the harmonic current of feeder F1 increase due to the nonlinear load inclusion at t = 4 s. At t = 10 s, when phase-c of the LV side of feeder F1 is disconnected, the positive- and negativesequence components of the current are, respectively, decreased and increased. Fig. 10 shows the instantaneous real and reactive powers of the DG units due to the load changes. The average of the real and reactive powers of the loads is shared between the DG units according to their droop characteristics. The positive- and negative-sequence currents and harmonic currents of the DG units are depicted in Fig. 11. The

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Fig. 9. Microgrid response to the unbalanced and nonlinear load changes applied to feeder F1 ; positive-sequence, negative-sequence, and harmonic components of loads currents at (a) feeder F1 and (b) feeder F2 .

(a)

Fig. 11. Microgrid response to the unbalanced and nonlinear load changes applied to feeder F1 ; positive-sequence, negative-sequence, and harmonic currents of (a) DG 1 and (b) DG 2 .

(a)

(b) (b)

(c)

Fig. 10. Dynamic response of DG units to unbalanced and nonlinear load changes applied to feeder F1 . (a) and (b) Real and reactive power components of DG units.

positive-sequence currents of DG units are in accordance with the average apparent power, shown in Fig. 10. The negativesequence currents of the unbalanced loads are shared between the DG units based on the equivalent negative-sequence impedance between the DG units and the loads of feeder F1 . Moreover, the harmonic currents of nonlinear loads are shared between the DG units based on the equivalent harmonic impedance of the lines and transformers. Since, the distance of feeder F1 from DG1 is shorter than its distance from DG2 , the main portion of the negative-sequence and harmonic currents are supplied by DG1 . Fig. 12(a) shows the instantaneous currents of DG1 prior and subsequent to the nonlinear load connection. The cascade connection of the two cells results in a five-level output voltage as shown in Fig. 12(b). The proposed multi-PR controller provides

Fig. 12. (a) Instantaneous current waveforms, (b) five-level-inverter output voltage, and (c) voltage waveforms of each phase of DG 1 ’s CHB inverter due to the nonlinear load connection to feeder F1 .

a set of sinusoidal voltages at the DG unit terminals as illustrated in Fig. 12(c). Fig. 13(a) shows the instantaneous currents of DG1 prior and subsequent to the single-phase load disconnection. The cascade connection of the two cells leads to a five-level output voltage as shown in Fig. 13(b). The proposed multi-PR controller provides a set of balanced sinusoidal voltages at the DG unit terminals as shown in Fig. 13(c). Fig. 14(a) shows the voltage THD for DG1 during the load switchings. After nonlinear load connection, the voltage THD is increased to 2.2% which is within the permissible ranges [5].

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Fig. 15. Fig. 13. (a) Instantaneous current waveforms, (b) five-level-inverter output voltage, and (c) voltage waveforms of each phase of DG 1 ’s CHB inverter due to the single-phase load disconnection from feeder F1 .

Voltages of dc links for DG 1 ’s units.

(a)

(a)

(b) (b)

(c) Fig. 14.

(a) Voltage THD and (b) VUF at DG 1 ’s terminal.

As shown in Fig. 14(b), the VUF of DG1 is below 2%, while injecting 85 A (0.29 p.u.) negative-sequence current to feeder F1 . The results show that the proposed voltage control strategy effectively compensates the harmonic and negative-sequence currents. Therefore, the acceptable values for the THD and VUF for all feeders are ensured. Fig. 15 shows the dc-link voltages of the HPSs of DG1 . The SC modules controllers regulate the dc-links voltages of the HPSs within 5% of the rated values even under the sudden load changes. Moreover, the power oscillations of the CHB multilevel inverters impose the double-frequency (100 Hz) ripples at the dc-link voltages. The dc-link voltages of each module of CHB multilevel inverter are separately balanced by the proposed HPS control strategy. The proposed control strategy of each inverter

Fig. 16. Dynamic response of DG 1 to load changes; currents of FC stacks and PV units for each HPS. (a) Phase a, (b) phase b, and (c) phase c.

balances the power sharing between the modules. Fig. 16 shows the currents of the FC stacks and PV units of the HPSs of DG1 . Each PV unit at phases a and b generates 34 kW, and each PV unit at phase c generates 20 kW. Fig. 17 shows the average (dc) currents of the SC modules for HPS units of DG1 . The power management strategy of the HPS system supports the different of power generation of PV units at each cell of

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FC/PV/SC power source and a voltage control strategy for the CHB multilevel inverter. The main features of the proposed HPS include high performance, high power density, and fast transient response. Furthermore, a multi-PR controller is presented to regulate the voltage of the CHB multilevel inverter in the presence of unbalanced and nonlinear loads. The performance of the proposed control strategy is investigated using PSCAD/EMTDC software. The results show that the proposed strategy: 1) regulates the voltage of the microgrid under unbalanced and nonlinear load conditions, 2) reduces THD and improves power quality by using CHB multilevel inverters, 3) enhances the dynamic response of the microgrid under fast transient conditions, 4) accurately balances the dc-link voltage of multilevel inverter modules, and 5) effectively manages the powers among the power sources in the HPS system.

REFERENCES

Fig. 17. Dynamic response of DG 1 to load changes; average current of SC module of each HPS. (a) Phase a, (b) phase b, and (c) phase c.

CHB multilevel inverter. Therefore, the FC stacks in phase c generate 14 kW more than the other FC stacks. At t = 4 s, the FC stacks of all phases of the CHB inverter increase their output powers to track the reference powers which are determined by the HPS controllers. The SC module compensates the power shortage of the FC stack, while the FC stack increases its output power at a limited response rate. The SC modules will be in standby state when the power capacity of the FC stacks is equal to the load demand. After unbalanced load switching at t = 10 s, the output currents of phases a and c of the CHB inverter are decreased. However, the current of phase b remains unchanged. As shown in Fig. 16, the FC stacks currents of phases a and c are decreased to meet the load demand in these phases. The power difference between the FC stacks and the load demand charges the SC modules in phases a and c of the CHB inverter, as shown in Fig. 17. At t = 15 s, the radiation intensity drops and the power of PV units decreases to 80% in about 2 s, as depicted in Fig. 16. As shown in Figs. 16 and 17, the SC module compensates the power shortage of the FC stack, while the FC stack increases its output power at a limited response rate. V. CONCLUSION This paper presents an effective control strategy for an islanded microgrid including the HPS and CHB multilevel inverter under unbalanced and nonlinear load conditions. The proposed strategy includes power management of the hybrid

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Mohsen Hamzeh (S’09) received the B.Sc. and M.Sc. degrees from the University of Tehran, Tehran, Iran, in 2006 and 2008, respectively, and the Ph.D. degree from the Sharif University of Technology, Tehran, in 2012, all in electrical engineering. Since 2010, he has been the Senior Research Engineer with SGP Company, Tehran. His research interests include distributed generation, microgrid control, and applications of power electronics in power distribution systems.

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Amin Ghazanfari (S’11) received the B.Sc degree in electrical engineering from the K. N. Toosi University of Technology, Tehran, Iran, in 2008, and the M.Sc. degree in power electronics from the Sharif University of Technology, Tehran, in 2011. Since 2010, he has been the Senior Research Engineer with the MGRayaneh Company and Niroo Research Institute, Tehran. His research interests include microgrid, renewable energy systems, power quality, and application of power electronics in power distribution systems.

Hossein Mokhtari (M’03) received the B.Sc. degree in electrical engineering from the University of Tehran, Tehran, Iran, in 1989, the M.Sc. degree in power electronics from the University of New Brunswick, Fredericton, NB, Canada, in 1994, and the Ph.D. degree in electrical engineering from the University of Toronto, Toronto, ON, Canada, in 1999. Since 2000, he has been with the Department of Electrical Engineering, Sharif University of Technology, Tehran, where he is currently a Professor. His research interests include power quality, power electronics, and the application of power electronics in power systems.

Houshang Karimi (S’03–M’07–SM’12) received the B.Sc. and M.Sc. degrees from the Isfahan University of Technology, Isfahan, Iran, in 1994 and 2000, respectively, and the Ph.D. degree from the University of Toronto, Toronto, ON, Canada, in 2007, all in electrical engineering. He was a Visiting Researcher and a Postdoctoral Fellow in the Department of Electrical and Computer Engineering, University of Toronto, from 2001 to 2003 and from 2007 to 2008, respectively. He was with the Department of Electrical Engineering, Sharif University of Technology, Tehran, Iran, from 2009 to 2012. From June 2012 to January 2013, he was a Visiting Researcher in the ePower Lab of the Department of Electrical and Computer Engineering, Queen’s University, Kingston, ON. He ´ ´ joined the D´epartement de G´enie Electrique, Ecole Polytechnique de Montr´eal, QC, Montr´eal, Canada, in 2013, where he is currently an Assistant Professor. His research interests include control systems, distributed generations, and microgrid control.