A Comprehensive DC Short-Circuit Fault Ride Through ... - IEEE Xplore

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IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 31, NO. 11, NOVEMBER 2016

A Comprehensive DC Short-Circuit Fault Ride Through Strategy of Hybrid Modular Multilevel Converters (MMCs) for Overhead Line Transmission Shenghui Cui, Student Member, IEEE, and Seung-Ki Sul, Fellow, IEEE

Abstract—The modular multilevel converter (MMC) is a promising candidate for voltage-sourced-converter-based high-voltage direct current (VSC-HVDC) transmission. The dc short-circuit fault management is a crucial issue especially for overhead line transmission, where the nonpermanent dc short-circuit faults occasionally occur. In this paper, a comprehensive dc short-circuit fault ride through (FRT) strategy is proposed for a hybrid MMC which combines half-bridge submodules and full-bridge submodules. By the proposed method, the hybrid MMC-based VSC-HVDC system is able to ride through a pole-to-pole short-circuit fault supporting the ac grid as a static synchronous compensator. Moreover, the transmission system can still transmit one-half of the rated active power in case of a pole-to-ground short-circuit fault without increasing power semiconductor stresses or requiring higher insulation levels of the electric equipment. The proposed method ensures successful regulation of arm currents and submodule capacitor voltages during the FRT process. Validity of the proposed strategy is verified by both computer simulation and experiment results. Index Terms—DC short-circuit fault, fault ride through, highvoltage direct current, modular multilevel converter, overhead line.

I. INTRODUCTION OLTAGE-SOURCED-CONVERTER-BASED highvoltage direct current (VSC-HVDC) transmission is a promising solution for future smart grids which would integrate a great amount of renewable energy sources to existing grids. For VSC-HVDC, compared to the conventional two-level or three-level converters, a modular multilevel converter (MMC) is a competitive candidate and is attracting worldwide attention [1]–[3]. The MMC presents many advantages such as very low harmonics, low dv/dt, modularity and simple scaling, no necessity of series connection of power semiconductors, and dc-bus capacitor elimination [4]–[6], etc. Fig. 1 shows the simplified schematic of an MMC station in the HVDC application. In the HVDC application, hundreds of half-bridge submodules (HBSMs) would be contained in each arm. Recently, the MMC converter with 401 submodules per arm has been reported for a ±320-kV, 1000-MW transmission project [7]. Typically, the

V

Manuscript received August 16, 2015; revised November 12, 2015; accepted December 21, 2015. Date of publication December 30, 2015; date of current version June 24, 2016. Recommended for publication by Associate Editor M. Hagiwara. S. Cui is with the Institute for Power Generation and Storage Systems, RWTH Aachen University, Aachen 52062, Germany (e-mail: scui@eonerc. rwth-aachen.de). S.-K. Sul is with the Department of Electrical and Computer Engineering, Seoul National University, Seoul 151744, Korea (e-mail: [email protected]. ac.kr). 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/TPEL.2015.2513426

Fig. 1.

Conceptual structure of the MMC station in the HVDC application.

MMC is grounded by a star-point reactor grounding device in the transformer’s secondary side [8], and smoothing dc reactors are usually installed in the dc side not only to reduce the transmission line current ripple but also to limit the increasing rate of the fault current in case of dc short-circuit faults. Dealing with the dc short-circuit faults is one of main concerns of VSC-HVDC transmission system, especially in the case that the overhead line is used for dc transmission. Protection schemes against the dc short-circuit faults are required to ensure safety of the MMCs. Generally, the dc short-circuit faults are caused by lightning strikes, pollution, and the broken tree branches in overhead line transmission. The dc short-circuit faults can be divided into two types, respectively, pole-to-pole short-circuit faults and pole-to-ground short-circuit faults. While a pole-to-pole short-circuit fault occurs, a large inrush current would be induced in the transmission line, and the power semiconductors of the MMC would be destroyed if no proper protection scheme is employed. Several protection schemes have been proposed to deal with the pole-to-pole shortcircuit fault [9]. The first approach is a bypass thyristor protection scheme as shown in Fig. 2(a) [4]. While a fault occurs, the IGBTs are switched OFF and the MMC is blocked within a few microseconds. Since the free-wheeling diode cannot withstand high surge-current, the bypass thyristor would be triggered to protect the diode. For this method, since the short-circuit fault cannot be extinguished naturally due to diode freewheeling effect, the circuit breaker in the ac side of the MMC should be opened and it may take several hundreds of milliseconds [10] or even longer [11] to restart the system. So, such protection scheme would be suitable for cable application where the poleto-pole short-circuit fault can be regarded as a permanent fault. However, for overhead line transmission, the system is required to clear the fault and restart in no time to minimize the loss

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CUI AND SUL: A COMPREHENSIVE DC SHORT-CIRCUIT FRT STRATEGY OF HYBRID MMCS FOR OVERHEAD LINE TRANSMISSION

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Fig. 2. Variable submodule circuits to deal with dc short-circuit fault. (a) HBSM with bypass thyristor protection scheme. (b) HBSM with double thyristor protection scheme. (c) CDSM. (d) FBSM.

of the transmission power [12]. An improved scheme has been proposed to deal with nonpermanent pole-to-pole short-circuit faults by employing the so-called double thyristor switches as shown in Fig. 2(b) [11]. By using the double thyristor protection scheme, the nonpermanent pole-to-pole short-circuit faults can be extinguished without opening the ac-side circuit breaker and it enables fast restart. A new concept of MMC employing clamp diode submodules (CDSMs) as shown in Fig. 2(c) has been proposed to handle the pole-to-pole short-circuit faults [13]. In case of the nonpermanent fault, the CDSM-based MMC is able to extinguish the dc arc inherently by blocking the IGBTs, while the penalty is a few additions of power semiconductor devices and around 50% higher semiconductor conduction losses compared to the conventional HBSM-based MMC. In some applications, the MMC is required to provide the reactive power as a static synchronous compensator (STATCOM) to support the ac grid even while a dc pole-to-pole short-circuit fault occurs [12]. Significance of supporting reactive power under the nonpermanent fault to enhance transient stability has been discussed in [10]. For such case, the MMC should be with fault ride through (FRT) capability for continuous operation without blocking it. However, neither the double thyristor protection scheme nor the CDSM-based topology is able to ride through the fault without blocking the converter. The full-bridge submodule (FBSM) solution that is shown in Fig. 2(d) has been proposed by both Siemens Co. and Alstom Co. to deal with this issue since the FBSM is able to output bipolar voltage regardless of the direction of the arm current [14], [15]. The main drawback of this solution is that both the number of power semiconductors and the semiconductor conduction losses are double of those of the conventional HBSM-based MMC. Besides the efforts on various submodule circuits, an alternate arm converter (AAC) which combines the two-level converter and the MMC has been proposed by Alstom Co. [16], as shown in Fig. 3. The AAC is able to operate as a STATCOM to ride through the pole-to-pole short-circuit fault without blocking the converter, and investigation on it shows that the conduction losses are even lower than those of the conventional HBSMbased MMC due to its alternate arm operation characteristic. Drawbacks of the AAC are as follows: 1) high-voltage capacitors are required in the dc link, while they are not necessary for the MMC, 2) it calls for series connection of IGBTs for direct

Fig. 3.

Conceptual schematic of the AAC.

switches, 3) a considerable sixth-order harmonic is included in the dc current, and 4) it is only able to balance arm capacitor voltages during an overlap period of arm current commutation. Currently, the symmetric monopole schematic is preferred for MMCs in HVDC applications. While a pole-to-ground short-circuit fault occurs, the pole-to-ground voltage of the healthy pole would be increased to as high as twice the rated value. Moreover, the transformer secondary-side phase-toground voltages would get offset by one-half of the rated dc-bus voltage leading to a serious offset voltage between the transformer neutral point and the ground [17]. Thus, the MMC-based VSC-HVDC system is unable to continue operating during a pole-to-ground short-circuit fault since it might lead to breakdown of the arresters, and insulation of the transformer and the healthy pole. Recently, a generic control scheme of FBSM-based MMC-HVDC systems was proposed for cable applications to ensure continuous operation in the case of permanent pole-to-ground short-circuit fault by reducing the HVDC transmission line voltage to one-half of its rated value [18]. This method can successfully avoid the overvoltage problem of the healthy pole, however, there are still several drawbacks behind this control scheme. At first, the transformer secondary-side phase-to-ground voltage would still get offset by one-half of the simultaneous dc-bus voltage corresponding to 25% of the rated HVDC line voltage. This is usually not allowed for the transformers used in the symmetric monopole VSC-HVDC systems. Moreover, if the HVDC system is grounded in the ac side (e.g., by the star point reactor), 25% of the rated HVDC line voltage would be subjected to the grounding device. It means that in overhead line applications, the fault current would keep flowing through the grounding device and the arc at the fault point cannot be extinguished. Recently, a hybrid MMC topology employing both the HBSMs and the FBSMs has been proposed to provide dc pole-to-pole short-circuit FRT capability as shown in Fig. 4 [19]–[22]. In each arm of the hybrid MMC, one-half of the submodules are HBSMs and the other half of the submodules are FBSMs. Compared to the FBSM solution, the hybrid one saves 25% of the power semiconductor devices and reduces

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Fig. 4.


Conceptual schematic of the hybrid MMC.

conduction losses by 25%. However, its dynamic control and arm capacitor voltage balancing for FRT have not been fully investigated. In this paper, a comprehensive FRT strategy is proposed for dc short-circuit faults including the pole-to-pole short-circuit fault and the pole-to-ground short-circuit fault. By the proposed method, the hybrid MMC-based VSC-HVDC system is able to ride through pole-to-pole short-circuit faults and the hybrid MMC supports the ac grid as a STATCOM. Also, an active-damping method is proposed to reduce the fault current extinguishing time conspicuously still keeping the submodule capacitor voltages at the rated voltage in the meantime. Moreover, the transmission system can still transmit one-half of the rated active power in case of a pole-to-ground short-circuit fault without increasing any power semiconductor stresses or causing additional voltage stresses on arresters or insulation of the transformer and the healthy pole. The FRT capability of transmitting active power can significantly improve the transient stability in terms of angle stability [10]. The proposed method can also enable the transmission system to operate continuously in case of a permanent fault of one pole of the transmission line. The proposed method ensures successful regulation of arm currents and submodule capacitor voltages during the FRT process, and it allows fast dynamic response of FRT and recovery. The validity of the proposed method is verified by extensive simulation and experiment results. Following parts of this paper are organized as follows. In Section II, the control method of the hybrid converter in normal operation is described. The MMC control and arm capacitor energy-balancing method conducted by authors in previous research [23] for normal operation has been briefly reviewed, which is the basis for the arm capacitor energy-balancing strategies proposed in this paper for dc short-circuit FRT. In Section III, the proposed FRT strategy of the hybrid converter in case of the pole-topole short-circuit fault is presented, including two novel voltage injection and current injection methods to balance leg capacitor energies in the null dc-bus voltage situation. In Section IV, the proposed FRT strategy to deal with the pole-to-ground fault is presented. In Sections V and VI, simulation and experiment

Fig. 5. Normal operation of the hybrid MMC. (a) Operation scheme of the upper arm. (b) Operation scheme of the lower arm. (c) Principle of operation.

results are provided to support the validity of the proposed strategy. II. NORMAL OPERATION In normal operation, a bipolar output voltage of the FBSM is not required. Then, the switches Su 6 , Sl6 of the FBSMs of both upper and lower arms should be normally turned ON, while the switches Su 5 , Sl5 should be normally turned OFF as shown in Fig. 5(a) and (b). In such case, the circuit of the hybrid MMC is topologically equivalent to that of the conventional HBSM-based MMC. The principle of the normal operation is Σ Σ and vc,xl represent abstractly shown in Fig. 5(c), where vc,xu sums of submodule capacitor voltages of upper and lower arms of “x” phase, respectively. The upper arm output voltage, the lower arm output voltage, and the ac grid-side current shown in Fig. 1 are denoted as vxu , vxl , and ixs respectively, and Vdc,rated denotes the dc transmission line rated voltage. A closed-loop indirect modulation [24]–[27] based control strategy of the MMC conducted by the authors in previous research is employed for normal operation [23]. The highlight of the control method in [23] is that it can fully decouple the ac grid current control, the dc-bus voltage and current control, and the arm capacitor energy balancing among six arms of the MMC. The twice line frequency circulating current can be inherently prevented with this control method [24], [25]. A. Modeling of the MMC Modeling of the MMC while the indirect modulation is employed is shown in Fig. 6 [23], [24]. The leg current ixo of “x” phase is defined as the average value of upper and lower arm currents as given in (1). A circulating current in “x” phase, ixo,cir , is defined as the difference between the corresponding phase leg current ixo and the average current of the dc-bus current that equally flows into each phase, as given in (2). Since the sum of leg currents of three phases is always identical to the dc-bus current, idc , (3) is valid which means that the circulating current


Fig. 6.

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Modeling of the MMC under indirect modulation.

is a current component that flows only inside the MMC ixo = (ixu + ixl ) /2. ixo,cir = ixo − iuo,cir + iv o,cir

idc . 3 + iwo,cir = 0.

(1) (2) (3)

In [23], [28], and [29], a universal MMC model under generalized nature of the dc bus was reported as shown in Fig. 7. In the universal model, the circuit of the MMC under indirect modulation in Fig. 6 has been split into three equivalent circuits to describe, respectively, the ac grid current, the dc-bus current, and the circulating current. In the normal operation, the dc-bus voltages of both positive and negative poles are synthesized as one-half of the rated dc-bus voltage as ∗ vdc

p

∗ vdc

n

Vdc,rated . 2 Vdc,rated . = 2 =

Fig. 7. Three equivalent circuits for the universal MMC model under generalized nature of dc bus. (a) Equivalent circuit for the ac grid current. (b) Equivalent circuit for the dc-bus current. (c) Equivalent circuit for the circulating current.

(4) Fig. 8. Block diagram of the converter total capacitor energy controller of the MMC that operates in the VR mode.

(5) d (Exu − Exl ) dExΔ = dt dt ∗ ∗ + vdc p − vdc n ixo .

PxΔ = Pxu − Pxl = B. Arm Capacitor Energy Control Arm capacitor energy control is one of main concerns of the MMC control. The arm capacitor energy control includes the converter total capacitor energy control and the arm capacitor energy balancing among six arms. To control the arm capacitor energy, the power flowing into each arm in the same leg should be considered. The sum and difference of the power flowing into upper and lower arms in the same leg of “x” phase are calculated as (6) and (7), respectively, neglecting the fluctuation terms PxΣ = Pxu + Pxl = ≈

∗ vdc

p

∗ + vdc

d (Exu + Exl ) dExΣ = dt dt idc ∗ + ixs . i − vxs xo,cir n 3

(6)

∗ ixo,cir ≈ −2vxs

(7)

In VSC-HVDC application, one MMC operates either as a voltage regulator (VR) to regulate the dc transmission line voltage or as a power dispatcher (PD) to control the transmission line power flow [30]. The control of the MMC that operates as the VR is only taken into consideration in this paper due to the limit of the paper size. The converter total capacitor energy control strategy of the VR converter is presented briefly as an example, and the further detailed discussion on the PD converter can be referred to [23] and [31]. For the VR, the total capacitor energy is regulated by controlling the coupled ac grid-side active power as shown in Fig. 8. The power drawn by the HVDC transmission line is employed as a feed-forwarding term to improve dynamic performance. In this paper, it is assumed that the grid voltage

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Fig. 10. Upper and lower arm capacitor energy balancing—common difference management. (a) Principle. (b) Block diagram.

Fig. 9. Leg capacitor energy balancing. (a) Principle of the leg capacitor energy balancing. (b) Block diagram of the leg capacitor energy balancing controller.

vector is oriented to q-axis of the synchronously rotating d–q reference frame. The balancing control of arm capacitor energies among six arms of the MMC includes the balancing of leg capacitor energies among three phases and the balancing of upper and lower arm capacitor energies at each phase. From (6), a dc component of the circulating current can be induced instantaneously to balance three-phase leg capacitor energies as shown in Fig. 9(a). The block diagram of the leg capacitor energy-balancing controller is shown in Fig. 9(b). In this method, the average value of three-phase leg capacitor energies is calculated online, and a dc component circulating current is induced inside the converter to drive three-phase leg capacitor energies to their average value. It should be noted that since a considerable twice line frequency component exists in leg capacitor energy ExΣ , notch filters are employed to cut off the alternating power components. Upper and lower arm capacitor energies should also be balanced at each phase. From (7), a line frequency ac circulating current can be temporarily injected to eliminate the difference between upper and lower arm capacitor energies, namely ExΔ . A positive sequence ac circulating current can independently drive the common component of the energy differences, it means that the positive-sequence ac circulating current enables internal energy transfer of the MMC among the upper three arms and the lower three arms as shown in Fig. 10(a) [23]. In addition, a negative-sequence ac circulating current can independently drive the differential component of the energy differences as shown in Fig. 11(a) [23]. In the control method of upper and lower arm capacitor energy balancing, both the positive- and negative-sequence line frequency ac circulating currents are injected temporarily inside the MMC to eliminate arm capacitor energy differences, namely EuΔ , EvΔ , EwΔ . Block diagrams of

Fig. 11. Upper and lower arm capacitor energy balancing—differential difference management. (a) Principle. (b) Block diagram.

the upper and lower arm capacitor energy-balancing controllers are presented in Figs. 10(b) and 11(b). In summary, in the normal operation mode, the hybrid MMC converter is topologically identical to the HBSM-based MMC. A comprehensive arm capacitor energy-balancing strategy is possible entirely by circulating current injection. The circulating current would only be injected temporarily in case of arm capacitor energy unbalance and would vanish automatically in the steady state. III. POLE-TO-POLE DC SHORT-CIRCUIT FAULT A. Operation Principles After a pole-to-pole dc short-circuit fault occurs, the dc-bus voltage of the MMC should be synthesized to null as soon as the


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Fig. 12. DC pole-to-pole short-circuit FRT of the hybrid MMC. (a) Operation scheme of the upper arm. (b) Operation scheme of the lower arm. (c) Principle of the operation.

fault is detected to extinguish the fault current. In the pole-topole FRT mode, a bipolar output voltage of the FBSM is required for both the upper arm and the lower arm. Then, the switches Su 2 and Sl2 of the HBSMs of both upper and lower arms are normally turned ON, while the switches Su 1 and Sl1 are normally turned OFF as shown in Fig. 12(a) and (b). In such case, the HBSMs are bypassed and the circuit of the hybrid MMC is topologically equivalent to the double star-connected cascade H-bridge (CHB) converter. The principle of pole-to-pole shortcircuit FRT operation is abstractly shown in Fig. 12(c). Even though the number of the submodules in the active operation is decreased to one-half of that in the normal operation mode, the converter is still able to generate the voltage to control the ac grid-side current due to bipolar output voltage capability of FBSMs. In the pole-to-pole FRT mode, the dc-bus voltages of both positive and negative poles of the converter are synthesized as null as (8) and (9). For the bypassed HBSMs, the conducting switches and the gate drivers would continuously consume power and discharge the submodule capacitors. One may argue that it would considerably decrease HBSM capacitor voltages and lead to difficulty of the recovery process. However, since the nonpermanent fault recovery time including fault current extinguishing and deionization is only hundreds of milliseconds [10], the HBSM capacitor voltage variation during the FRT process is negligible ∗ vdc

p

∗ vdc n

= 0

(8)

= 0.

(9)

Since the dc transmission line is a low-damping system, it would take a long time to extinguish the fault current naturally especially in the overhead line application where a considerable inductance exists. An active damping can be employed in pole-to-pole dc short-circuit FRT to absorb the inductive energy actively to accelerate the fault current extinguishing process, and the absorbed inductive energy would be transferred to the ac grid in the meantime actively without leading to increase

Fig. 13. Leg capacitor energy balancing during pole-to-pole short circuit FRT—voltage injection. (a) Principle. (b) Block diagram.

of the submodule capacitor voltage since the total capacitor energy of the hybrid MMC is simultaneously regulated by controlling the ac grid-side active power. Then, the dc-bus voltages of both poles of the MMC should be synthesized as (10) and (11) to implement an active-damping scheme with equivalent active resistance Ractive ∗ vdc

p

∗ vdc

n

Ractive idc 2 Ractive idc . = 2 =

(10) (11)

B. Arm Capacitor Energy Balancing In case of a pole-to-pole short-circuit fault, the dc transmission line voltage is collapsed to null and the system is unable to transmit active power. Nevertheless, the hybrid MMC converters are still able to provide reactive power to the coupled ac grids, which would conspicuously improve the transient stability during fault recovery [10]. During the FRT period, the arm capacitor energies of six arms are still required to be balanced. The arm capacitor energy-balancing strategy in the normal operation mode stated in Section II is valid for the pole-to-pole short-circuit FRT except the leg capacitor energy balancing. During the FRT period, only the capacitor energy of the FBSMs is required to be balanced since all the HBSMs are bypassed. It is obvious that the arm capacitor energy of each arm should be reduced to one-half of that in the normal operation since only one-half of the submodules are needed to be considered. To implement a seamless mode transition between the normal

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operation mode and the FRT mode, the arm capacitor energy of each arm is calculated as twice of the capacitor energies of FBSMs. As shown by (6) and Fig. 9(a), the leg capacitor energy balancing by injection of the dc circulating current is only valid when the converter synthesizes the dc-bus voltage. During the pole-to-pole short-circuit FRT, a dc component of the circulating current is not able to transfer energy among three legs due to dc-bus voltage collapse. To balance leg capacitor energies of three phases, a differential component of active powers of three phases should be generated intentionally. Two solutions are proposed in this paper, one is based on the voltage injection while the other one is based on the current injection. At first, the voltage injection-based leg capacitor energy-balancing method is introduced. Without loss of generality, the MMC ac current (only includes the reactive current) can be represented as (12). From (6), a differential component of active powers of three phases can be introduced, while a common mode voltage ∗ ∗ as (13) is injected in the output ac voltage vxs . The vxs,com generated differential power terms are calculated as (14). By (14), a common mode output ac voltage can be injected in accordance with the MMC ac current to exchange energy among three legs as shown in Fig. 13(a). The block diagram of the voltage injection method is presented in Fig. 13(b) ⎧ ius = Is sin (ωt + ϕ) ⎪ ⎪ ⎪ ⎪ ⎪ ⎪ ⎨ 2π ivs = Is sin ωt + ϕ − 3 ⎪ ⎪ ⎪ ⎪ 2π ⎪ ⎪ ⎩ iws = Is sin ωt + ϕ + 3 ∗ vxs,com = Vcom sin (ωt + α) ⎧ 1 ⎪ PuΣ,v inj = Vcom Is cos (ϕ − α) ⎪ ⎪ 2 ⎪ ⎪ ⎪ ⎨ 1 2π Σ Pv ,v inj = Vcom Is cos ϕ − α − 2 3 ⎪ ⎪ ⎪ ⎪ ⎪ ⎪ ⎩ PwΣ,v inj = 1 Vcom Is cos ϕ − α + 2π 2 3

(12)

(13)

.

(14)

Besides the voltage injection method, a leg capacitor energybalancing method based on negative-sequence grid current injection method can be employed. Without loss of generality, the ac grid voltage can be represented as (15). From (6), a differential component of active powers of three phases can be introduced, while a negative-sequence current is injected into the ac grid. The generated differential power terms that are contributed by the injected negative-sequence grid current represented as (16) are calculated as (17) ⎧ vug = Vg sin (ωt) ⎪ ⎪ ⎪ ⎪ ⎪ ⎪ ⎨ 2π vvg = Vg sin ωt − 3 ⎪ ⎪ ⎪ ⎪ 2π ⎪ ⎪ ⎩ vwg = Vg sin ωt + 3

(15)

Fig. 14. Leg capacitor energy balancing during pole-to-pole short circuit FRT—current injection. (a) Principle. (b) Block diagram.

⎧ ius,neg = Is,neg sin (ωt + β) ⎪ ⎪ ⎪ ⎪ ⎪ ⎪ ⎨ 2π ivs,neg = Is,neg sin ωt + β + 3 ⎪ ⎪ ⎪ ⎪ 2π ⎪ ⎪ ⎩ iws,neg = Is,neg sin ωt + β − 3 ⎧ 1 ⎪ ⎪ PΣ = Vg Is,neg cos (β) ⎪ ⎪ u ,c inj 2 ⎪ ⎪ ⎪ ⎨ 1 2π PvΣ,c inj = Vg Is,neg cos β − 2 3 ⎪ ⎪ ⎪ ⎪ ⎪ 1 2π ⎪ ⎪ ⎩ PwΣ,c inj = Vg Is,neg cos β + 2 3

(16)

.

(17)

By (17), a negative-sequence current can be temporarily injected into the coupled ac grid in accordance with the grid voltage to exchange energy among three phases as shown in Fig. 14(a). The block diagram of the current injection method is presented in Fig. 14(b). The advantage of the voltage injection method is that it would not affect the coupled ac grid, while the drawback comes that it is only valid when the converter imports the reactive current into the ac grid to provide the reactive power. On the other side, the current injection method is generally valid as long as the coupled ac grid does not collapse. The drawback of the current injection method is that it instantaneously injects the negative-sequence current into the ac grid. Considering that the magnitude of the injected negative-sequence current is very small and vanishes as long as the three-leg capacitor energies are balanced, it can still be regarded as a promising solution for the dc pole-to-pole short-circuit FRT.


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Fig. 15. DC positive pole-to-ground short-circuit FRT of the hybrid MMC. (a) Operation scheme of the upper arm. (b) Operation scheme of the lower arm. (c) Principle of the operation.

IV. POLE-TO-GROUND DC SHORT-CIRCUIT FAULT A. Operation Principles While a pole-to-ground dc short-circuit fault occurs, the pole-to-ground voltage of the healthy pole would be increased to as high as twice the rated value, and the transformer secondary-side phase-to-ground voltages would get offset by half of the rated dc-bus voltage leading to a serious offset voltage between the transformer neutral point and the ground. Moreover, a significant common mode current would be induced in the dc transmission line, if the dc transmission system is grounded in the ac side and the arc would not be naturally extinguished [32]. Such serious problem would trip the ac-side circuit breaker and lead to total failure of power transmission. In this section, a novel FRT strategy is proposed to enable the converter to ride through the fault transmitting at most one-half of the rated active power. A positive-pole-to-ground short-circuit fault case is explained for instance. In case of a positive-pole-to-ground dc short-circuit fault, the positive pole dc-bus voltage of the MMC should be synthesized to null as soon as the fault is detected to extinguish the fault current, while the negative pole dc-bus voltage can be still maintained as its rated value, namely one-half of the dc transmission line rated voltage to keep transmitting the active power. Then, in the positive-pole-to-ground FRT mode, a bipolar output voltage of the FBSM is required for the upper arm, while it is not required for the lower arm. To operate in such case, the switch Su 2 of the HBSMs should be normally turned ON and the complementary switch Su 1 should be normally turned OFF in the upper arm as shown in Fig. 15(a), and the switch Sl6 of the FBSMs should be normally turned ON and the complementary switch Sl5 should be normally turned OFF in the lower arm as shown in Fig. 15(b). In such case, the HBSMs are bypassed in the upper arm and the circuit of the upper arm is topologically equivalent to that of a single-star CHB converter, and the lower arm is topologically kept identical to that of the hybrid MMC in the normal operation mode. The principle of the positive-

Fig. 16. Dynamics of the fault current in case of a positive-pole-to-ground dc short-circuit fault. (a) Emerging of the fault current. (b) Extinguishing of the fault current.

pole-to-ground short-circuit FRT operation is abstractly shown Σ Σ and vc,xl represent sums of submodin Fig. 15(c), where vc,xu ule capacitor voltages of upper and lower arms of “x” phase, respectively. In the positive-pole-to-ground FRT mode, the dcbus voltages of the positive and negative poles of the converter are synthesized as (18) and (19), respectively. By using the proposed FRT strategy, the dc negative pole-toground voltage is clamped to its rated voltage as it, before the fault occurs, and both the transformer neutral point-to-ground voltage and the dc positive pole-to-ground voltage are actively clamped to null. So, the arc in the positive pole can be extinguished naturally and the negative pole is still able to transmit the active power without leading to any overvoltage issue ∗ vdc

p

= 0

∗ vdc

n

=

Vdc,rated . 2

(18) (19)

The transient process of the fault current extinguishing is one of the interests in FRT. Fig. 16 illustrates the emerging process and the extinguishing process of the fault current which flows through the start point reactor grounding device. While a positive-pole-to-ground dc short-circuit fault occurs, a fault current is induced through the path shown in Fig. 16(a) and it is driven by the positive pole dc-bus voltage of the MMC, namely ∗ vdc p . After the fault is detected and the MMC transits into the

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TABLE I PARAMETERS OF THE SIMULATED SYSTEM MMC converter

Fig. 17. Common difference management part of the upper and lower arm capacitor energy balancing controller with the feedforwarding term of dc pole output power difference.

Number of HBSMs per arm Number of FBSMs per arm Rated submodule capacitor voltage Submodule capacitor Inductance of arm inductor Controller sampling frequency

108 108 2.2 kV 4.5 mF 15 mH 10 kHz

Smoothing reactor Inductance

10 mH

Star point reactor grounding device Inductance of three phase inductor Resistance of neutral point resistor

10 H 200 Ω

HVDC transmission line Rated voltage Rated current

Fig. 18.

400 kV 1 kA

Schematic diagram of the simulated hybrid HVDC system.

FRT mode, then, the fault current is extinguished by the resistor of the grounding device as shown in Fig. 16(b). Dynamics of emerging and extinguishing processes of the fault current are calculated as (20) and (21), respectively, which indicate that the time constant is mainly determined by parameters of the start point reactor grounding device

Vdc,rated RG (t − tfault ) 1 − exp − L G iG (t) = − Lo 2RG 3 + 3 + LSR

RG (t − tFRT ) iG (t) = iG (tFRT ) exp − L G Lo 3 + 3 + LSR

(20)

.

(21)

B. Arm Capacitor Energy Balancing During the FRT period, the arm capacitor energies of six arms are still required to be balanced. The arm capacitor energybalancing strategy in the normal operation mode stated in Section II is also valid for the pole-to-ground short-circuit fault FRT. For upper arms, only capacitor energies of the FBSMs are required to be regulated since all the HBSMs are bypassed. Arm capacitor energies of upper arms should be reduced to half of those in the normal operation since only one-half of the submodules (the FBSMs) are needed to be considered. To implement a seamless mode transition between the normal operation mode and the FRT mode, the arm capacitor energy of each upper arm is calculated as twice of the capacitor energies of FBSMs inside it. Since the positive pole dc-bus voltage is null, only the negative pole transmits the active power during the FRT period and it leads to unbalance of dc pole output powers of the positive and negative poles. Since the active power that is drawn from the ac grid flows equally into upper three arms and lower three arms, the active power exchange between upper three arms and lower three arms is required throughout the FRT period to

Fig. 19. Simulation results of the pole-to-pole short-circuit FRT without employing the proposed voltage injection- or current injection-based leg capacitor energy-balancing method.

maintain upper and lower arm capacitor energy balance. The common difference management part of the upper and lower arm capacitor energy-balancing controller illustrated in Fig. 10 would automatically induce a positive-sequence circulating current to compensate the dc pole output power difference between the positive pole and the negative pole. However, it would lead to poor dynamics since the integral regulator of the controller shown in Fig. 10(b) takes time to fully compensate the dc pole output power difference. To deal with this issue, the dc pole output power difference is employed as a feedforwarding term as shown in Fig. 17, by which dynamic performance can be improved conspicuously. During the FRT period, the transmission system is able to transmit at most one-half of the rated active power since the pole-to-pole voltage is reduced to one-half of that in the normal operation. While the arm capacitor energy of each arm has been already regulated well, the relationship between the ac grid current and the positive-sequence circulating current can be deduced as (22) and (23) from (6) and (7). Assuming that the converter only draws the active power during the FRT period and the voltage drop across the arm inductor is negligible


Fig. 20. Simulation results of the pole-to-pole short-circuit FRT by employing the voltage injection-based leg capacitor energy-balancing method.

comparing to the ac grid voltage (which is reasonable in the HVDC application), (24) can be deduced Vdc,rated idc ∗ i P¯xΣ ≈ − vxs xs = 0 2 3 Vdc,rated idc ∗ P¯xΔ ≈ −2vxs =0 ixo,cirAC,p os − 2 3 ⎧ ixs ⎪ ⎪ i =− ⎪ ⎪ xo,cirAC,p os 2 ⎪ ⎪ ⎨ idc ixs idc + + ixo,cirAC,p os = ixu = ⎪ 2 3 3 ⎪ ⎪ ⎪ ⎪ i i idc ⎪ xs dc ⎩ ixl = − + + ixo,cirAC,p os = − ixs 2 3 3

(22) (23)

.

(24)

Several interesting properties can be observed from (24). In the upper arm current, only the term contributed by the dc-bus current is included and the ac component vanishes. In the lower arm current, the ac component is as large as the MMC ac-side current, while it is only one-half of the MMC ac-side current in the normal operation. However, it would not lead to additional current stresses of the power semiconductors since the MMC ac-side current does not exceed one-half of its rated value due to active power reduction during the FRT period. The proposed pole-to-ground short-circuit fault FRT strategy enables the transmission system to extinguish the fault current

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Fig. 21. Simulation results of the pole-to-pole short circuit FRT by employing the current injection-based leg capacitor energy balancing method.

and to transmit at most one-half of the rated active power without additional power semiconductor stress. Moreover, this strategy can be applied with ground return in case of a permanent failure of one pole of the transmission line, and the HBSMs corresponding to the faulted pole can be bypassed by mechanical switches to reduce losses. Recently, a discussion on the MMC based on various submodule circuits was reported [9]. The dc short-circuit FRT method proposed in this paper (including the pole-to-pole short-circuit FRT and the pole-to-ground short-circuit FRT) is valid not only in the HBSM/FBSM hybrid MMC but also in other type of MMC topologies. And, its application can be extended to other MMC topologies which can generate the bipolar arm output voltage in sufficient range, e.g., the FBSM-based MMC. For the HBSM/FBSM hybrid MMC, theoretically, mean capacitor voltages, and numbers of the HBSM and FBSM in each arm should coincide (25) and (26) neglecting the capacitor voltage ripple and the voltage drop across the arm inductor NFBSM Vcap,FBSM > Vg NFBSM Vcap,FBSM + NHBSM Vcap,HBSM > Vg +

(25) Vdc,rated . 2 (26)

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Fig. 24. Photograph and schematic diagram of the experimental setup. (a) Photograph of the experimental setup. (b) Schematic diagram of the experimental setup. TABLE II PARAMETERS OF THE PROTOTYPE MMC SYSTEM UNDER TEST MMC converter

Fig. 22.

Simulation results of the pole-to-ground short-circuit FRT.

Number of HBSMs per arm Number of FBSMs per arm Rated submodule capacitor voltage Submodule capacitor Inductance of arm inductor Controller sampling frequency

3 3 50 V 5.4 mF 4 mH 10 kHz

Smoothing reactor Inductance

4 mH HVDC transmission line

Rated voltage Rated current

Fig. 23. Transient of the fault current that flows into the ground in the case of positive pole-to-ground dc short-circuit fault.

V. SIMULATION RESULTS To verify the validity of the proposed FRT strategy, fullscale simulation has been performed. Schematic of the simulated system is illustrated in Fig. 18. The hybrid MMC has 216 submodules per arm including 108 HBSMs and 108 FBSMs. A Thevenin equivalent circuit method is employed to reduce simulation time without sacrificing accuracy [33]. Two load resistors are employed to emulate the positive and negative poles of the dc-bus of the other hybrid MMC which operates in the PD mode to control the transmission line power flow. Parameters of the simulated system are shown in Table I. Fault detection technique is a crucial issue in fault management, however, it is

300 V 10 A

out of scope of this paper. For simplicity of the simulation, the fault type is preassumed in the MMC controller in each scenario and the fault detection is done only by monitoring the dc-bus current. The controller judges it as a dc-side short-circuit fault while the dc-bus current exceeds 1.5 kA. The fault is cleared at t = 2.15 s in each scenario and the MMC controller restarts the normal operation mode automatically at t = 2.2 s. A. Pole-to-Pole DC Short-Circuit Fault Fig. 19 shows the simulation results illustrating dynamic performance while neither the voltage injection- nor the current injection-based leg capacitor energy-balancing strategy proposed in this paper was employed. Before the fault occurred at t = 2.0 s, the transmission system delivered 400-MW active power and the hybrid MMC provided 200-MVAR reactive power to the ac grid. It is clearly shown that with injection of the line frequency ac circulating current, upper and lower


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arm capacitor energy difference was still eliminated during the pole-to-pole dc short-circuit FRT. However, even though the leg capacitor energy-balancing controller employed in the normal operation (in Fig. 9) kept injecting the dc circulating current, it was not able to balance leg capacitor energy because the dc-bus voltage was collapsed to null during FRT and the submodule capacitor voltage could not be regulated to the rated voltage. Fig. 20 shows the simulation results illustrating dynamic performance in case of a pole-to-pole dc short-circuit fault while the voltage injection-based leg capacitor energy-balancing strategy was employed. Before the fault occurred at t = 2.0 s, the transmission system delivered 400-MW active power and the hybrid MMC provided 200-MVAR reactive power to the ac grid at the same time. To realize fast fault current extinguishing, the active damping technique was employed in the FRT period and the active damping resistance was set as 4 Ω. It is shown that the hybrid MMC transited to the FRT mode as soon as the fault was detected and the dc line fault current was extinguished in 40 ms. The submodule capacitor voltage was regulated at its rated value during the FRT period successfully by employing the proposed FRT method. The circulating current vanished as soon as the submodule capacitor voltages were balanced and the hybrid MMC was able to continuously provide 200-MVAR reactive power during the FRT period. At t = 2.2 s, the fault was cleared and the hybrid MMC transited into the normal operation mode. Fig. 21 shows the simulation results illustrating dynamic performance while the current injection-based leg capacitor energybalancing strategy was employed. Before the fault occurred at t = 2.0 s, the transmission system delivered 400-MW active power and the hybrid MMC provided no reactive power. The submodule capacitor voltage was regulated at its rated value during the FRT period successfully by employing the proposed current injection method even at no apparent power condition. At t = 2.2 s, the fault was cleared and the hybrid MMC transited into the normal operation mode. B. Pole-to-Ground DC Short-Circuit Fault Fig. 22 shows the simulation results illustrating dynamic performance in case of a positive-pole-to-ground dc short-circuit fault. Before the fault occurred at t = 2.0 s, the transmission system delivered 400-MW active power and the hybrid MMC provided no reactive power. After the fault occurred, the healthy negative pole-to-ground voltage was clamped to double of its rated value and a common mode current was induced in the HVDC transmission line. As soon as the fault was detected, the hybrid MMC transited into the FRT mode and SFRT was closed to emulate the FRT mode of the other hybrid MMC which operates as the PD. It is shown that the fault current was extinguished successfully by the proposed strategy and the transmission system was able to continuously transmit 200-MW active power during the FRT period without leading to additional voltage stresses on arresters or insulation of the transformer and the healthy pole. Due to the temporary reduction of the active power, the MMC ac grid-side current decreased to one-half of that before the fault during the FRT period. As discussed in

Fig. 25. Experiment results of the pole-to-pole short-circuit FRT without employing the proposed voltage injection- or current injection-based leg capacitor energy-balancing method. (a) Leg capacitor energy. (b) Difference of upper and lower arm capacitor energy. (c) Upper arm FBSM capacitor average voltage. (d) Lower arm FBSM capacitor average voltage. (e) Circulating current.

Section IV, the ac component in the upper arm current almost decreased to zero and the ac component in the lower arm current almost kept the same as that before the fault. By the proposed FRT strategy, the submodule capacitor voltage was regulated at its rated value successfully. At t = 2.2 s, the fault was cleared and the hybrid MMC transited into the normal operation mode, and the SFRT was opened to emulate recovery of the PD.

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Fig. 23 shows the transient process of emerging and distinguishing of the fault current flowing through the grounding device into the ground. The theoretical result calculated by (20) and (21) is illustrated by dashed lines. It coincides with the simulation result with reasonable accuracy and the validity of fault current transient analysis is verified. VI. EXPERIMENT RESULTS To verify the validity of the proposed FRT strategy, the downscale experiment has been performed. The photograph and schematic of the experimental setup are illustrated in Fig. 24. The hybrid MMC has six submodules per arm including three HBSMs and three FBSMs. Two load resistors are employed to emulate the positive and negative poles of the dc bus of the other hybrid MMC which operates in the PD mode to control the transmission line power flow. The ac side is grounded by the neutral point of the three-phase ac voltage instead of a start point reactor grounding device since it is difficult to manufacture a grounding device with such high-inductance reactor in the laboratory. Parameters of the simulated system are shown in Table II. For simplicity of the experiment, as done in the simulation, the fault type is preassumed in the MMC controller in each scenario and the fault detection is by only monitoring the dc-bus current. The controller judges it as a dc-side short-circuit fault while the dc-bus current exceeds 13 A. In the experiment investigating performance of pole-to-pole dc short-circuit FRT, the controller restarts the normal operation mode automatically 1.0 s later after the fault is detected. In the experiment investigating performance of pole-to-ground dc short-circuit FRT, the controller restarts the normal operation while the dc-bus current decreases lower than 7 A. A. Pole-to-Pole DC Short-Circuit Fault Fig. 25 shows the experiment results illustrating dynamic performance while neither the voltage injection- nor the current injection-based leg capacitor energy-balancing strategy proposed in this paper was employed. Before the fault occurred, the transmission system delivered 3-kW active power and the hybrid MMC provided 3-kVAR reactive power to the ac grid. It is clearly shown that with injection of the line frequency ac circulating current, upper and lower arm capacitor energy difference was eliminated during the pole-to-pole dc short-circuit FRT. However, even though the leg capacitor energy-balancing controller employed in the normal operation (in Fig. 9) kept injecting the dc circulating current, it was not able to balance leg capacitor energy because the dc-bus voltage collapsed to null during FRT and the submodule capacitor voltage could not be regulated to the rated voltage. Fig. 26 shows the experiment results illustrating dynamic performance in case of a pole-to-pole dc short-circuit fault while the voltage injection-based leg capacitor energy-balancing strategy was employed. Before the fault occurred, the transmission system delivered 3-kW active power and the hybrid MMC provided 3-kVAR reactive power to the ac grid at the same time. It is shown that the hybrid MMC transited to the FRT mode as

Fig. 26. Experiment results of the pole-to-pole short-circuit FRT by employing the voltage injection-based leg capacitor energy-balancing method. (a) Hybrid MMC ac grid-side current. (b) Circulating current. (c) Upper arm current. (d) Lower arm current. (e) Leg capacitor energy. (f) Difference of upper and lower arm capacitor energy. (g) Upper arm FBSM capacitor average voltage. (h) Lower arm FBSM capacitor average voltage.


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Fig. 27. Experiment results of the pole-to-pole short circuit FRT by employing the current injection-based leg capacitor energy-balancing method. (a) Hybrid MMC ac grid-side current. (b) Circulating current. (c) Upper arm current. (d) Lower arm current. (e) Leg capacitor energy. (f) Difference of upper and lower arm capacitor energy. (g) Upper arm FBSM capacitor average voltage. (h) Lower arm FBSM capacitor average voltage. (i) DC-bus voltage and current.

soon as the fault was detected and the dc line fault current was extinguished. The submodule capacitor voltage was regulated at its rated value during the FRT period successfully by employing the proposed method. The circulating current vanished as soon as the submodule capacitor voltages were balanced and the hybrid MMC was able to continuously provide 3-kVAR reactive power during the FRT period. Fig. 27 shows the experiment results illustrating dynamic performance while the current injection-based leg capacitor energybalancing strategy was employed. Before the fault occurred, the transmission system delivered 3-kW active power and the hybrid MMC provided no reactive power. The submodule capacitor voltage was regulated at its rated value during the FRT

period successfully by employing the proposed current injection method even at no apparent power condition. B. Pole-to-Ground DC Short-Circuit Fault Fig. 28 shows the experiment results illustrating dynamic performance in case of a positive-pole-to-ground dc short-circuit fault. Before the fault occurred, the transmission system delivered 3-kW active power and the hybrid MMC provided no reactive power. As soon as the fault was detected, the hybrid MMC transited into the FRT mode. It is shown that the fault current was extinguished successfully by the proposed strategy and the transmission system was able to continuously transmit

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Fig. 28. Experiment results of the positive pole-to-ground short-circuit FRT. (a) Hybrid MMC ac grid-side current. (b) Circulating current. (c) Upper arm current. (d) Lower arm current. (e) Leg capacitor energy. (f) Difference of upper and lower arm capacitor energy. (g) Upper arm FBSM capacitor average voltage. (h) Lower arm FBSM capacitor average voltage. (i) Difference of upper and lower arm capacitor energy without feedforward of dc pole output power difference. (j) DC-bus voltage and current.

1.5-kW power during the FRT period. Due to the temporary reduction of the active power, the MMC ac grid-side current decreased to one-half of that before the fault during the FRT period. The ac component in the upper arm current decreased to a very small value just to compensate the reactive power consumed by the arm inductor, and the ac component in the lower arm current almost kept the same as that before the fault. By the proposed FRT strategy, the submodule capacitor voltage was regulated at its rated value successfully. Comparing Fig. 28(f) and (i), it is shown that performance of the upper and lower arm capacitor energy balancing can be conspicuously improved

by employing the feedforwarding term of dc pole output power difference shown in Fig. 17. VII. CONCLUSION In this paper, a comprehensive FRT strategy was proposed for the HBSM/FBSM hybrid MMC to deal with both pole-to-pole and pole-to-ground dc short-circuit faults. In case of a pole-topole short-circuit fault, the hybrid MMC is able to continuously operate as a STATCOM to support the ac grid during the FRT; while in case of a pole-to-ground short-circuit fault, the hybrid


MMC is able to extinguish the fault current and continuously operate to transmit at most one-half of the rated active power. By employing the proposed strategy, the power transmission gap in case of the fault can be minimized and the transmission system transient stability can be highly enhanced. In the meantime, the submodule capacitor voltage can be regulated at its rated value during the FRT period. The validity of the proposed method is verified by both the full-scale simulations and the down-scale experiments extensively.

[19] [20] [21]

[22] [23]

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tems” IEEE Trans. Power Del., vol. 30, no. 6, pp. 2468–2476, Feb. 2015. S. Inoue and S. Katoh, “Modular multilevel converter with DC fault protection,” U.S. Patent 0 148 392, Jun. 12, 2013. S. Norrga, “Modified voltage source converter structure,” U.S. Patent 0 195 084, Aug. 15, 2012. M. Davis and H. Gambach, “Method for eliminating a fault on a high-voltage dc line, system for transmitting an electric current via a high-voltage dc line, and converter,” U.S. Patent 0 308 235, Nov. 21, 2013. R. Zeng, L. Xu, L. Yao, and B. W. Williams, “Design and operation of a hybrid modular multilevel converter,” IEEE Trans. Power Electron., vol. 30, no. 3, pp. 1137–1146, Apr. 2014. S. Cui, S. Kim, J. Jung, and S. K. Sul, “A comprehensive cell capacitor energy control strategy of a modular multilevel converter (MMC) without a stiff DC bus voltage source,” in Proc. IEEE Appl. Power Electron. Expo., 2014, pp. 602–609. A. Antonopoulos, L. Anguist, and H. P. Nee, “On dynamics and voltage control of the modular multilevel converter,” in Proc. Eur. Power Electron. Appl., 2009, pp. 1–10. S. Debnath, J. Qin, B. Bahrani, M. Saeedifard, and P. Barosa, “Operation, control, and applications of the modular multilevel converter: A review,” IEEE Trans Power Electron., vol. 30, no. 1, pp. 37–53, Mar. 2014. M. Hagiwara and H. Akagi, “Control and analysis of the modular multilevel cascade converter based on double-star chopper-cells (MMCCDSCC),” IEEE Trans. Power Electron., vol. 26, no. 6, pp. 1649–1658, Oct. 2010. K. Sekiguchi, P. Khamphakdi, M. Hagiwara, and H. Akagi, “A grid-level high-power BTB (back-to-back) system using modular multilevel cascade converters without common DC-link capacitor,” IEEE Trans. Ind. Appl., vol. 50, no. 4, pp. 2648–2659, Nov. 2013. J. Kolb, F. Kammerer, and M. Braun, “Straight forward vector control of the modular multilevel converter for feeding three-phase machines over their complete frequency range,” in Proc. IEEE Annu. Ind. Electron. Soc., 2011, pp. 1596–1601. M. Perez, R. Lizana, S. Bernet, J. Espinoza, and J. Rodriguez, “Control of arm capacitor voltages in modular multilevel converters,” IEEE Trans. Power Electron., vol. 31, no. 2, pp. 1774–1784, Apr. 2015. L. Tang, “Control and protection of multi-terminal DC transmission systems based on voltage-source converters,” Ph.D. dissertation, Dept. Elect. Comput. Eng., McGill University, Montreal, QC, Canada, 2003. S. Cui, J. Jung, Y. Lee, and S. K. Sul, “A novel control strategy of a modular multilevel converter (MMC) based VSC-HVDC transmission system,” in Proc. IEEE Appl. Power Electron. Expo., 2015, pp. 972–979. C. Zhao, T. Li, L. Yu, Y. Huang, L. Li, and X. Li, “DC pole-to-ground fault characteristic analysis and converter fault recovery strategy of MMCHVDC,” Proc. Chin. Soc. Elect. Eng., vol. 34, pp. 3518–3526, 2014. (in Chinese) U. Gnanarathna, A. Gole, and R. Jayasinghe, “Efficient modeling of modular multilevel HVDC converters (MMC) on electromagnetic transient simulation programs,” IEEE Trans. Power Delivery, vol. 26, no. 1, pp. 316–324, Oct. 2010.

Shenghui Cui (S’13) received the B.S. degree from Tsinghua University, Beijing, China, and the M.S. degree from Seoul National University, Seoul, Korea, in 2012 and 2014, respectively, both in electrical engineering. Since March 2015, he has been with the Institute for Power Generation and Storage System, E.ON Energy Research Center, RWTH Aachen University, Aachen, Germany, as a Research Associate. His Master’s thesis deals with the modeling and control of modular multilevel voltage source converters for high-voltage dc applications. His research interests include high power converters for medium and high voltage applications.

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Seung-Ki Sul (S’78–M’87–SM’98–F’00) was born in Korea, in 1958. He received the B.S., M.S., and Ph.D. degrees in electrical engineering from Seoul National University, Seoul, Korea, in 1980, 1983, and 1986, respectively. From 1986 to 1988, he was an Associate Researcher with the Department of Electrical and Computer Engineering, University of Wisconsin, Madison, WI, USA. From 1988 to 1990, he was a Principal Research Engineer with LG Industrial Systems Company, Korea. Since 1991, he has been a member of faculty of the School of the Electrical and Computer Engineering, Seoul National University, where he is currently a Professor. He contributed to PWM technology. He is one of pioneers in the area of carrier-based PWM technology applied to the control of the power converters. From 2005 to 2007, he was the Vice Dean of the Engineering College of Seoul National University. In addition, from 2008 to 2011, he was the President of the Electrical Engineering Science Research Institute funded by the Korean Government. He has more than 130 IEEE journal papers and a total of more than 300 international conference papers in the area of power electronics. He was the program chair of IEEE Power Electronics Specialists Conference 2006 and General Chair of IEEE ECCE-Asia, International Conference on Power Electronics, 2011. He has been actively involved in various industry projects sponsored by many Korean, Japanese, and American companies. For his sabbatical year from 2003 to 2004, he was an acting Director of the research center of Yaskawa Electric Company, Japan. His current research interests include power electronic control of electrical machines, electric/hybrid vehicles and ship drives, high-voltage dc transmission based on the modular multilevel converter, and power-converter circuits for renewal energy sources. From 1998 to 2003, he was a board member of Hyundai Elevator Co., which is the largest elevator maker in Korea. From 2011 to 2014, he was a Board Member of LS Industrial System Co., which is the largest power electronics-related product maker in Korea. He holds 14 USA patents, 7 Japanese patents, 11 Korean patents, and granted 41 Ph.D. under his supervision. For last three years, Dr. Sul has served as Editor-in-Chief of the Journal of Power Electronics, which is a SCIE registered journal, published by the Korean Institute of Power Electronics (KIPE), Seoul, Korea. He has been the President of the KIPE since 2015.