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Control and Experiment of Pulse-Width-Modulated Modular Multilevel Converters Makoto Hagiwara, Member, IEEE, and Hirofumi Akagi, Fellow, IEEE Department of Electrical and Electronic Engineering Tokyo Institute of Technology, Tokyo, Japan E-mail: mhagi@akg.ee.titech.ac.jp, and akagi@ee.titech.ac.jp Abstract— A modular multilevel converter (MMC) is one of the next-generation multilevel converters intended for high- or medium-voltage power conversion without transformers. The MMC is based on cascade connection of multiple bidirectional chopper-cells per leg, thus requiring voltage-balancing control of the multiple floating dc capacitors. However, no paper has made an explicit discussion on voltage-balancing control with theoretical and experimental verifications. This paper deals with two types of pulse-width-modulated modular multilevel convert- ers (PWM-MMCs) with focus on their circuit configurations and voltage-balancing control. Combination of averaging and balancing controls enables the PWM-MMCs to achieve voltage balancing without any external circuit. The viability of the PWM- MMCs, as well as the effectiveness of the voltage-balancing control, is confirmed by simulation and experiment. Keywords— Medium-voltage power conversion, multilevel con- verters, voltage-balancing control. I. I High-power converters for utility applications require line- frequency transformers for the purpose of enhancing their volt- age or current rating [1]-[4]. The 80-MVA STATCOM (STATic synchronous COMpensator) commissioned in 2004 consists of 18 NPC (Neutral-Point-Clamped) converter-legs [4], where each of the ac sides is connected in series by the corresponding transformer. The use of line-frequency transformers, however, not only makes the converter heavy and bulky, but also induces the so-called dc magnetic-flux deviation when a single-line-to- ground fault occurs [5]. Recently, many scientists and engineers of power systems and power electronics have been involved in multilevel con- verters intended for achieving medium-voltage power conver- sion without transformers [6]-[8]. Two of the representatives are: 1) the diode-clamped multilevel converter (DCMC) [6][7], 2) the flying-capacitor multilevel converter (FCMC) [8]. The three-level DCMC, or a neutral-point-clamped (NPC) converter [9] has been put into practical use [10]. If a voltage- level number is more than three in the DCMC, inherent voltage imbalance occurs in the series-connected dc capacitors, thus resulting in requiring an external balancing circuit (such as a buck-boost chopper) for a pair of dc capacitors [11]. Furthermore, a significant increase in the clamping diodes required renders assembling and building of each leg more complex and difficult. Thus, a reasonable voltage-level number would be up to five from a practical point of view. As for the FCMC, the four-level PWM inverter is currently produced by one manufacture of industrial medium-voltage drives [12]. However, the high expense of flying capacitors at low carrier frequencies (say, lower than 1 kHz) is the major disadvantage of the FCMC [13]. A modular multilevel converter (MMC) has been proposed in [14]-[20], intended for high-power applications. Fig. 1 shows a basic circuit configuration of a three-phase modular multilevel inverter. Each leg consists of two stacks of multi- ple bidirectional cascaded chopper-cells and two non-coupled buffer inductors. The MMC is suitable for high- or medium- voltage power conversion due to easy construction/assembling and flexibility in converter design. Siemens has a plan of putting it into practical use with the trade name of “HVDC- plus.” It is reported in [19] that a system configuration of the HVDC-plus has a power rating of 400 MVA, a dc-link voltage of ±200 kV, and 200 cascaded chopper-cells per leg. The authors of [14]-[20], however, have made no detailed description of staircase modulation, especially about a crucial issue of how to achieve voltage balancing of 200 floating dc capacitors per leg. Moreover, no experimental result has been reported yet. This paper focuses on voltage-balancing control and oper- ating performance of a pulse-width-modulated modular multi- level converter (PWM-MMC) equipped with either two non- coupled buffer inductors or a single coupled buffer inductor per leg. The final aim of this paper is to apply the PWM-MMC to medium-voltage power converters in a power rating of 1 to 10 MVA, a dc-link voltage of 10 to 30 kV, and a switching frequency of 200 to 2000 Hz. Combining averaging control with balancing control enables to achieve voltage balancing of multiple floating dc capacitors without any external circuit. In addition, this paper proposes the dual MMC for low-voltage large-current power conversion. Each dc side of positive and negative chopper-cells possesses a common dc capacitor, whereas its ac side is connected in parallel via multiple buffer inductors. The similarity between the two MMCs exists in terms of circuit configuration and control method. The validity of the two MMCs is confirmed not only by simulated results, but also by experimental results. II. T M M C A. Classification from the Topologies Fig. 2 shows a classification of modular multilevel convert- ers based on single-phase half-bridge or full-bridge converter- i d i Pu i Nu i Zu i u i v i w v 1u v 8u v C1u v C4u v C5u v C8u v ju R L l l E v C ju C (j : 1 ∼ 8) u-phase v-phase w-phase cell 1 cell 4 cell 5 cell 8 u v w (a) (b) Fig. 1. Circuit configuration of a chopper-cell-type modular multilevel inverter: (a) The power circuit, and (b) The bidirectional PWM chopper-cell with a floating dc capacitor. cells. From their topologies, modular multilevel converters can be classified into 1) double-star-configured MMCs, 2) a star-configured MMC (Fig. 3(a)), 3) a delta-configured MMC (Fig. 3(b)), and 4) the dual MMC (Fig. 14). Moreover, the double-star-configured MMCs can be classified into 1) a chopper-cell-type MMC (Fig. 1), and 2) a bridge-cell-type MMC. B. Comparisons in Function and Application The double-star-configured MMC topology possesses the common dc-link terminals as shown in Fig. 1(a), which enable dc-to-ac and ac-to-dc power conversion. However, the star/delta-configured MMC topology has no common dc-link terminals as shown in Fig. 3. As a result, it has no capability of achieving dc-to-ac and ac-to-dc power conversion although it can control active power back and force between the three- phase ac terminals and the floating dc capacitors. This means that the star/delta-configured MMC topology is not applicable to industrial motor drives, but it is suitable for STATCOMs and energy storage systems [21]-[23]. This consideration is one of the most significant differences in function and application between the double-star-configured MMC topology in Fig. 1 and the star/delta-configured MMC topology in Fig. 3. The bridge-cell-type MMC replaces the chopper-cell in Fig. 1(b) with single-phase full-bridge converter-cells. Hence, Modular Multilevel Converters (MMCs) Double-star-configured MMCs Star-configured MMC (Fig. 3(a)) Delta-configured MMC (Fig. 3(b)) The dual MMC (Fig. 14) using half-bridge converter-cells (chopper-cell-type MMC (Fig. 1)) using single-phase full-bridge converter-cells (bridge-cell-type MMC) Fig. 2. A classification of modular multilevel converters. cell 2 cell n cell 2 cell n (a) (b) Fig. 3. Circuit configuration of MMCs: (a) Star-configured MMC, and (b) Delta-configured MMC. the dc-voltage source E can be replaced with a single-phase ac- voltage source [16]. The detail of the dual MMC is discussed in section V. In this paper, the chopper-cell-type MMC is referred to simply as “the MMC” because attention is paid to it exclusively. C. Definition of DC Loop Currents Fig. 1 shows a three-phase inverter based on the MMC. Each leg of the circuit consists of two stacks of four bidirectional chopper-cells and two non-coupled buffer inductors. Each chopper-cell consists of a floating dc capacitor and two IGBTs that form a bidirectional chopper. Attention is paid to the u- phase chopper-cells because the operating principle is identical among the three legs. The following circuit equation exists in Fig. 1(a) 1 : E = 8 j=1 v ju + l d dt (i Pu + i Nu ). (1) Here, E is a supply dc voltage, v ju is an output voltage of the u- phase chopper-cell numbered j, l is a buffer inductance, and i Pu and i Nu are positive- and negative-arm currents, respectively. The KVL (Kirchhoff’s voltage law) loop given by (1) is referred to as the “dc loop,” which is independent of the load. 1 The subscript symbol j means numbering of each chopper-cell. 1 — 2 v ∗ C ¯v Cu i ∗ Zu i Pu i Nu i Zu v C ju v ∗ C v ∗ Bju v ∗ Au K 5 ±1 current minor loop (j : 1 ∼ 8) +1 : i Pu , i Nu > 0 −1 : i Pu , i Nu < 0 K 1 + — s K 2 K 3 + — s K 4 (a) (b) Fig. 4. Block diagram of dc-capacitor voltage control: (a) Averaging control, and (b) Balancing control. The circulating current along the u-phase dc loop, i Zu can be defined as i Zu = i Pu − i u 2 = i Nu + i u 2 = 1 2 (i Pu + i Nu ). (2) Note that i Pu , i Nu , i u and i d are branch currents whereas i Zu is a loop current that is impossible to measure directly. III. C M MMC The voltage-balancing control of eight floating dc capacitors per leg in Fig. 1 can be divided into 1) averaging control, and 2) balancing control. A. Averaging Control Fig. 4(a) shows a block diagram of the averaging control. It forces the u-phase average voltage ¯v Cu to follow its command v ∗ C , where ¯v Cu is given by ¯v Cu = 1 8 8 j=1 v C ju . (3) Let a dc-loop current command of i Zu be i ∗ Zu , as shown in Fig. 4(a). It is given by i ∗ Zu = K 1 (v ∗ C − ¯v Cu ) + K 2 (v ∗ C − ¯v Cu )dt. (4) The voltage command obtained from the averaging control, v ∗ Au is given by v ∗ Au = K 3 (i Zu − i ∗ Zu ) + K 4 (i Zu − i ∗ Zu )dt. (5) When v ∗ C ≥ ¯v Cu , i ∗ Zu increases. The function of the current minor loop in Fig. 4(a) forces the actual dc-loop current i Zu to follow its command i ∗ Zu . As a result, this feedback control of i Zu enables ¯v Cu to follow its command v ∗ C without being affected by the load current i u . v ∗ Au v ∗ Au v ∗ Bju v ∗ Bju v ∗ u /4 v ∗ u /4 E/8 E/8 v ∗ ju v ∗ ju (j : 1 ∼ 4) (j : 5 ∼ 8) (a) (b) Fig. 5. Voltage command of each arm: (a) Positive arm, and (b) Negative arm. B. Balancing Control The use of the balancing control described in [21] forces the individual dc voltage to follow its command v ∗ C . Fig. 4(b) shows a block diagram of the u-phase balancing control, where v ∗ Bju is the voltage command obtained from the balancing control. Since the balancing control is based on either i Pu or i Nu , the polarity of v ∗ Bju should be changed according to that of i Pu or i Nu . When v ∗ C ≥ v C ju (j : 1 − 4) in the positive arm of Fig. 1(a), a positive active power should be taken from the dc power supply into the four chopper-cells. When i Pu is positive, the product of v Bju (= v ∗ Bju ) and i Pu forms the positive active power. When i Pu is negative, the polarity of v Bju should get inverse to take the positive active power. Finally, v ∗ Bju for j = 1 − 4 is represented as: v ∗ Bju = K 5 (v ∗ C − v C ju ) (i Pu > 0) −K 5 (v ∗ C − v C ju ) (i Pu < 0), (6) while v ∗ Bju for j = 5 − 8 is represented as: v ∗ Bju = K 5 (v ∗ C − v C ju ) (i Nu > 0) −K 5 (v ∗ C − v C ju ) (i Nu < 0). (7) Fig. 5 shows a voltage command of each chopper-cell v ∗ ju . The positive-arm and negative-arm commands are obtained as: v ∗ ju = v ∗ Au + v ∗ Bju − v ∗ u 4 + E 8 (j : 1 − 4) (8) v ∗ ju = v ∗ Au + v ∗ Bju + v ∗ u 4 + E 8 (j : 5 − 8), (9) where v ∗ u is an ac-voltage command for the u-phase load. Note that Fig. 5 includes the feedforward control of the dc supply voltage E. The voltage command v ∗ ju is normalized by each dc-capacitor voltage v C ju , followed by comparison with a triangular waveform having a maximal value of unity and a minimal value of zero with a carrier frequency of f C . The actual switching frequency of each chopper-cell, f S is equal to f C . The eight chopper-cells have the eight triangular waveforms with the same frequency but a phase difference of 45 ◦ (= 360 ◦ /8) each other for achieving harmonic cancellation and enhancing current controllability. As a result, the line-to- neutral voltage is a nine-level voltage waveform, and a line-to- line voltage is a 17-level voltage waveform with an equivalent switching frequency of 8f C . 10 0 -10 [kV] 200 0 -200 [A] [A] 200 0 -200 [kV] 2.5 2.0 [MW] 1.5 1.0 0.5 [A] 80 40 0 [A] 80 40 0 [V] 40 0 -40 [V] 40 0 -40 [kV] 2 1 0 40 ms v uv v vw v wu v uv i u i v i w i u i u i Pu i Nu i Nu i Pu i u v C1u v C5u v C1u v C5u p d i ∗ Zu i Zu v ∗ B1u v ∗ Au v ∗ 1u v ∗ 5u v ∗ 1u v ∗ 5u Fig. 6. Simulated waveforms obtained from Fig. 1 where f = 50 Hz, v ∗ C = 2.25 kV, and E = 9 kV. C. Simulated Results Fig. 6 shows simulated waveforms from Fig. 1. Tables I and II summarize circuit parameters and control gains used for simulation using a software package of the “PSCAD/EMTDC” [24]. The dc supply voltage E and the rated active power P are set as E = 9 kV and P = 1 MW, and therefore the nominal rated line-to-line rms voltage of the MMC is 5.5 kV. An intrinsic one-sampling delay occurs due to digital control. The dc-capacitor voltage command of each chopper-cell is v ∗ C = 2.25 kV (= 9 kV/4), while three-phase load voltage TABLE I C P U S. Rated active power P 1 MW Rated line-to-line rms voltage V 5.5 kV Rated rms current I 105 A Rated frequency f 50 Hz DC supply voltage E 9 kV Buffer inductance l 3 mH (3.1%) DC capacitance C 1.9 mF DC capacitor voltage V C 2.25 kV (= 9 kV/4) Unit capacitance constant [25] H 115 ms Carrier frequency f C 2 kHz Equivalent switching frequency 8 f C 16 kHz Load power factor cos φ 0.9 at 50 Hz Load inductance L 6 mH (6.2%) Values in () are on a three-phase 5.5-kV, 1-MW, and 50-Hz base. TABLE II C G U S. Proportional gain of averaging control K 1 0.5 A/V Integral gain of averaging control K 2 150 A/(V·s) Proportional gain of current control K 3 1.5 V/A Integral gain of current control K 4 150 V/(A·s) Proportional gain of balancing control K 5 0.35 commands of each phase are given as: v ∗ u = 0.5E sin 2π ft v ∗ v = 0.5E sin 2πft − 2 3 π v ∗ w = 0.5E sin 2πft − 4 3 π E = 9 kV f = 50 Hz. (10) Note that v ∗ u , v ∗ v , and v ∗ w are the three-phase line-to-neutral voltages. In Fig. 6, each chopper-cell is operated at unity modulation index. In Table I, a unit capacitance constant H is defined as: H = 3 × 8 × 1 2 CV 2 C P = 12 × 1.9 × 10 −3 × 2.25 2 × 10 6 10 6 = 0.115, (11) where V C is the rated dc voltage of each chopper-cell. Note that H is defined as a ratio of all electrostatic energy stored in dc capacitors with respect to rated active power [25]. Therefore, H has a unit of second 2 . Fig. 6 indicates that v uv is a 17-level line-to-line voltage, achieving voltage balancing of all the dc capacitors. The dc input power, p d is represented as: p d = E × i d , (12) where i d is a dc input current. The waveform of p d includes the following two frequency components: One is a 6th-frequency (300 Hz) component stemming from a three-phase full-bridge 2 H can be defined in the traditional 2-level converters as well. An optimization of H is left for the future work. i P i N i 0 v 0 L R v 1 v 2 v 3 v 4 v C1 v C2 v C3 v C4 l l E/2 E/2 C d C d (70 V) (70 V) v S C d = 20 mF L ac = 1.1 mH E = 140 V L ac E C d C d 250 VA MMC cell 1 cell 2 cell 3 cell 4 (a) (b) Fig. 7. Experimental circuit: (a) Half-bridge circuit, and (b) System configuration. v C1 v C4 clock(10 MHz) sampling signal 8 bit gate Controller DSP FPGA AD AD AD AD AD i P i N E cell 1 cell 4 v ∗ 1 v ∗ 4 Fig. 8. Digital control system used for experiment. converter, and the other is a fundamental-frequency (50 Hz) component stemming from an output frequency of 50 Hz. The detailed analysis of each waveform will be carried out in the next section. IV. E R A. System Configuration Used for Experiment Fig. 7(a) shows a half-bridge circuit based on the MMC, where the stack number of chopper-cells was selected as four per leg to confirm the basic operating principle, al- though the stack number of chopper-cells was eight in Fig. 1. Fig. 7(b) shows a system configuration used for experiment. The midpoint of the two dc input series-connected capacitors is connected back to the neutral point of the star-winding of the transformer, thus making the midpoint voltage stable. The dc supply voltage E is regulated at 140 V, and the capacitance value of C d is chosen as 20 mF. TABLE III C P U E. Rated active power P 250 W Rated rms voltage V 0 50 V Rated rms current I 0 5 A Rated frequency f 50 Hz DC supply voltage E 140 V Buffer inductance l 1 mH (3.2%) DC capacitance C 3 mF ±10% DC capacitor voltage V C 70 V Unit capacitance constant [25] H 118 ms Carrier frequency f C 8 kHz Equivalent switching frequency 4 f C 32 kHz Power factor cos φ 0.9 at 50 Hz Load inductance L 2 mH (6.4%) Values in () are on a single-phase 50-V, 5-A, and 50-Hz base. TABLE IV C G U E. Proportional gain of averaging control K 1 0.5 A/V Integral gain of averaging control K 2 80 A/(V·s) Proportional gain of current control K 3 1 V/A Integral gain of current control K 4 640 V/(A·s) Proportional gain of balancing control K 5 0.5 Fig. 8 shows the digital control system used for experi- ment. The system detects each dc-capacitor voltage v C j , both positive- and negative-arm currents i P and i N , and a dc supply voltage E as input signals to the A/D unit. The A/D unit consisting of seven A/D converters takes in the analogue signals, and then it converts them into seven 12-bits digital signals. A DSP unit using a 16-bit DSP (ADSP-2105) takes in the digital signals, and produces the voltage commands v ∗ j after completing the digital processing shown in Figs. 4 and 5. An FPGA unit has the following multi-functions: 1) generating carrier signals with appropriate phase differ- ences, 2) comparing v ∗ j with the corresponding triangular carrier signal, and 3) producing gate signals with a dead time of 2 µs. The FPGA unit produces 8-bit (= 2 × 4) gate signals in total, because each chopper-cell possesses two semiconductor switches. Furthermore, the FPGA unit sends back sampling signals to the DSP unit for realizing the so-called “syn- chronous sampling.” Attention is paid to a one-sampling delay occurring in the DSP unit. The experiment selected the carrier frequency as f C = 8 kHz, while each triangular carrier had a phase difference of 90 ◦ (= 360 ◦ /4). If a conventional synchronous sampling was adopted, the DSP unit would yield a time delay of 63 µs (= 1/(2f C )) because its sampling and com- mand renewal would be conducted at the peak value of each carrier waveform. The sampling delay can be reduced with the following sampling method for improving controllability of output voltages: As for chopper-cell 1, for instance, the sampling is conducted at the peak value of carrier signal 4, and then v ∗ j is changed at the peak value of carrier signal 1. As a consequence, the time delay is minimized to be 31 µs (= 1/(4f C )). A three-phase MMC can reduce the time delay further by phase-shifting of each leg by 120 ◦ . B. Operating Performance under a Steady-State Condition Tables III and IV summarize the circuit parameters and control gains used for experiment 3 . An R-L load with a power factor of 0.9 at 50 Hz was utilized. The experiment was carried out under a condition of P = 250 W. Fig. 9 shows the experimental waveforms when the dc- voltage command of each chopper-cell was set as v ∗ C = 70 V. The ac voltage command for the load was given by: v ∗ 0 = √ 2V 0 sin 2π ft V 0 = 50 V f = 50 Hz. (13) Fig. 7(a) yields the following equation: v 0 = 1 2 4 j=3 v j − 2 j=1 v j − l 2 d dt i 0 (14) v 0 = R + L d dt i 0 , (15) The waveform of v 0 is slightly different from that in a DCMC and an FCMC, due to the existence of the second term of the right hand in (14). Note that both DCMC and FCMC have a complete staircase waveform with a constant (not curved) voltage level. Substituting (14) into (15) yields: 1 2 4 j=3 v j − 2 j=1 v j = Ri 0 + L + l 2 d dt i 0 . (16) Hence, harmonic voltages included in the left hand in (16) appear across both L and l/2. If l/2 is dominant over L (L l/2), most of the harmonic voltages appear across l/2, bringing less harmonic voltages to v 0 . If L is dominant over l/2 (L l/2), most of the harmonic voltages appear across L to the contrary. Although the load current i 0 is sinusoidal with a fundamen- tal component of 50 Hz, the arm currents i P and i N contain non-negligible low-order harmonic currents. This interesting phenomenon occurs due to the effect of the balancing control. Fig. 9 indicates that the voltage command from the balancing control, v ∗ B1 is a discontinuous sawtooth waveform as expected from (6) and (7). As a result, v ∗ B1 brings a discontinuous cir- culating current to the dc loop, producing low-order harmonic currents in i P and i N . It is obvious that the fluctuations in i P and i N are accompanied by those in v ∗ B1 . The switching ripple currents contained in i P and i N are determined by the inductance value of the buffer inductors and the harmonic voltages resulting from the four chopper-cells. Note that switching-frequency components of 16 kHz are dominant in i P and i N . In other words, the ripple currents can be reduced by increasing the carrier frequency and the buffer inductance value. Carefully looking into v ∗ B1 and i P in Fig. 9 reveals that a subtle difference exists at the times of polarity change between 3 The experimental conditions are the same in H and l(per unit) as those in Fig. 6. 80 0 -80 [V] [A] 10 0 -10 [V] 100 70 40 [V]100 70 40 [A] 5 0 [V] 4 0 -4 [V] 4 0 -4 [V] 70 35 0 40 ms v 0 i 0 i P i N i N i P i 0 ¯v C v C1 v C3 v C1 v C3 i Z i ∗ Z i ∗ Z i Z v ∗ B1 v ∗ A v ∗ 1 Fig. 9. Experimental waveforms obtained from Fig. 7, where f = 50 Hz, v ∗ C = 70 V, and E = 140 V. the waveforms of v ∗ B1 and i P . The reason is that the 16-kHz switching ripple component contained in i P is not taken into the control circuit precisely, because the sampling frequency of DSP is set as 16 kHz. This phenomenon produces no effect on the balancing control because the harmonic current makes no contribution to forming any active power. Applying the averaging control forces the average voltage ¯v C to follow its command v ∗ C (= 70 V). The calculation of (3) has a function of reducing ac components in ¯v C . From Fig. 9, v C1 and v C3 contain 50-Hz components caused by i 0 . This ac component is inverse proportional to the fundamental frequency of i 0 . This is similar to flying capacitors in the FCMC. The dc voltage of each chopper-cell is kept balanced by the balancing control. Making reference to (8) simplifies the voltage commands of chopper-cells 1 and 2, v ∗ 1 and v ∗ 2 as follows: v ∗ 1 = v ∗ 2 − v ∗ 0 2 + E 4 , (17) where a reasonable approximation of v ∗ B1 = v ∗ B2 = v ∗ A 0 was made. Equation (17) implies that chopper-cells 1 and 2 were 80 0 -80 [V] [A] 10 0 -10 [V] 100 70 40 Voltage command was changed v 0 i 0 v C1 v C3 Fig. 10. Experimental waveforms when the voltage command was changed from 50 V to 25 V. 100 70 40 [V] [V] 80 0 -80 [A] 10 0 -10 Only the balancing control was disabled v 0 i 0 v C1 v C2 v C3 v C4 Fig. 11. Experimental waveforms when only the balancing control was disabled. operated under the same modulation index. As a result, v C1 was equal to v C2 in Fig. 9 because the chopper-cells 1 and 2 utilize a common arm current i P , and (17) leads to a relation of v ∗ 1 = v ∗ 2 . In a similar way, v C3 was equal to v C4 in Fig. 9. Experimental waveforms of i Z and i ∗ Z show that no steady-state error, even in a small control gain, existed between i Z and i ∗ Z in terms of their dc components because of an extremely low resistance along the dc loop. C. Operating Performance under a Transient-State Condition Fig. 10 shows experimental waveforms of the MMC when the voltage command was reduced to half, but the circuit parameters and the control gains were not changed. The transient voltage fluctuations in v C1 and v C3 were suppressed to less than 5% with respect to its rated voltage of 70 V. Fig. 11 shows experimental waveforms before and after the balancing control was disabled intentionally but the averaging control was enabled. The voltage imbalance was gradually expanding with the passage of time. Hence, the balancing control is indispensable for stable operation. i P i N i 0 a b c l ac = 1.0 mH (3.2%) l bc = 1.0 mH (3.2%) l ab = 4.0 mH (12.8%) Fig. 12. Coupled inductor used for experiment. 80 0 -80 [V] [A] 10 0 -10 [V] 100 70 40 [V] 100 70 40 [A] 5 0 [V] 6 0 [V] 4 0 -4 [V] 70 35 0 40 ms v 0 i 0 i P i N i N i P i 0 ¯v C v C1 v C3 v C1 v C3 i Z i ∗ Z i ∗ Z i Z v ∗ B1 v ∗ A v ∗ 1 Fig. 13. Experimental waveforms when a coupled buffer inductor was used. D. Operating Performance Using a Coupled Buffer Inductor The two non-coupled buffer inductors in Fig. 7 can be replaced with a single coupled buffer inductor intended for a size reduction in the magnetic components. Fig. 12 shows specifications of the inductor used for experiment. The termi- nals “a” and “b” are connected to the positive- and negative- arms respectively, while the terminal “c,” or the midpoint is directly connected to the load. The relation of l ab = 4l ac = 4l bc exists in Fig. 12. It should be noted that the inductor presents no inductance to i 0 because the magnetic fluxes produced by the fundamental frequency components in i P and i N are can- celed out each other. As a consequence, the inductor presents the inductance of l ab only to the circulating current i Z . The use of the coupled inductor results in bringing considerable i P i 0 v 0 i N v 1 v 2 v 3 v 4 v C1 v C2 RL i P1 i P2 i N1 i N2 l l l l E/2 E/2 C d C d (35 V) (35 V) cell 1 cell 2 cell 3 cell 4 Fig. 14. Half-bridge circuit using the dual modular multilevel converter. reductions in size, weight, and cost to the magnetic core 4 . Fig. 13 shows experimental waveforms when the coupled inductor was utilized. In Fig. 13, each chopper-cell was oper- ated at a modulation index of 0.83, while the balancing control using the load current was utilized (see the appendix.). Since the inductor produces no effect on i 0 , (14) can be rewritten as follows: v 0 = 1 2 4 j=3 v j − 2 j=1 v j . (18) Hence, v 0 is a staircase waveform with a constant voltage level as shown in Fig. 13. In other words, the MMC operates as a multilevel voltage source that is independent of i 0 . Comparison between Figs. 9 and 13 reveals that both have similar waveforms except for v 0 . V. T D MMC Fig. 14 shows a half-bridge circuit of the other MMC with a dual relation to Fig. 7. Each dc side of positive and negative chopper-cells possesses a common dc capacitor, whereas its ac side is connected in parallel via two buffer inductors 5 . A. Control Method The control method of the dual MMC is basically the same as that of Fig. 1 except for the following: Although Fig. 1 has a common dc loop to cascaded chopper-cells per leg, the dual MMC has multiple dc loops because the multiple chopper- cells forming an arm are connected in parallel. This means that the multiple current minor loops should be provided for the averaging control. The chopper-cell 1 in Fig. 14, for instance, should control a circulating current i Z1 included in i P1 . Here, i Z1 is obtained as i Z1 = i P1 − i 0 4 . (19) 4 An optimization of the inductor is left for the future work. 5 The dc capacitor of each chopper-cell can be floated like Fig. 7. 40 0 -40 [V] [A] 20 0 -20 [A] 10 0 -10 [V]100 70 40 [A] 5 0 [V] 4 0 -4 [V] 4 0 -4 [V] 70 35 0 40 ms v 0 i 0 i P i N i N i P i 0 i P1 i N1 i P1 i N1 v C2 v C1 v C1 v C2 i Z1 i ∗ Z i ∗ Z i Z1 v ∗ B1 v ∗ A1 v ∗ 1 Fig. 15. Experimental waveforms obtained from Fig. 14, where f = 50 Hz, v ∗ C = 70 V, and E = 70 V. B. Operating Performance under a Steady-State Condition Fig. 15 shows experimental waveforms obtained from Fig. 14. Here, the dc supply voltage was E = 70 V, and the dc voltage command of each chopper-cell was v ∗ C = 70 V. The ac voltage command of the load was given by: v ∗ 0 = √ 2V 0 sin 2π ft V 0 = 25 V f = 50 Hz. (20) The circuit parameters and control gains were the same as those in Fig. 9. It should be noted that the experiment was carried out using four non-coupled buffer inductors. Figs. 9 and 15 show that both MMCs have similar wave- forms. However, comparison reveals that the two waveforms of v 0 are different in harmonic voltage. The reason is that the buffer inductance in the dual MMC is smaller than that in Fig. 7 due to parallel connection of the two buffer inductors. Fig. 15 shows that i P1 and i N1 are halves of i P and i N , respectively. Therefore, the dual MMC is suitable for low- voltage large-current power conversion. VI. C This paper has dealt with two types of pulse-width- modulated modular multilevel converters (PWM-MMCs), proposing their control method and verifying their operating principle. Computer simulation using the “PSCAD/EMTDC” software package has confirmed the proper operation of the three-phase PWM-MMC. Experiments using a laboratory system has verified the viability and effectiveness of the single-phase PWM-MMC. The MMC is showing considerable promise as a power converter for medium-voltage motor- drives, high-voltage direct current (HVDC) systems, STAT- COMs, and back-to-back (BTB) systems. R [1] S. Mori, K. Matsuno, T. Hasegawa, S. Ohnishi, M. Takeda, M. Seto, S. Murakami, and F. Ishiguro, “Development of a large static var generator using self-commutated inverters for improving power system stability,” IEEE Trans. Power Sys., vol. 8, no. 1, pp. 371-377, Feb. 1993. [2] K. Kunomura, K. Yoshida, K. Ito, N. Nagayama, M. Otsuki, T. Ishizuka, F. Aoyama, and T. Yoshino, “Electronic frequency converter,” in Conf. Rec. IEEJ-IPEC 2005, pp. 2187-2191. [3] T. Uzuka, S. Ikedo, and K. Ueda, “A static voltage fluctuation compen- sator for ac electric railway,” in Conf. Rec. IEEE-PESC 2005, pp. 1869- 1873. [4] T. Fujii, S. Funahashi, N. Morishima, M. Azuma, H. Teramoto, N. Iio, H. Yonezawa, D. Takayama, and Y. Shinki, “A ± 80MVA GCT STATCOM for the Kanzaki substation,” in Conf. Rec. IEEJ-IPEC 2005, pp. 1299- 1306. [5] M. Hagiwara, P. V. Pham, and H. Akagi, “Calculation of dc magnetic flux deviation in the converter-transformer of a self-commutated BTB system during single-line-to-ground faults,” IEEE Trans. Power Elec- tron., vol. 23, no. 2, pp. 698-706, March 2008. [6] F. Z. Peng, “A generalized multilevel inverter topology with self volt- age balancing,” IEEE Trans. Ind. Appl., vol. 37, no. 2, pp. 611-618, March/April 2001. [7] J. Rodriguez, J. S. Lai, and F. Z. Peng, “Multilevel inverters: a survey of topologies, controls, and applications,” IEEE Trans. Ind. Electron., vol. 49, no. 4, pp. 724-738, Aug. 2002. [8] T. A. Meynard, and H. Foch, “Multi-level choppers for high voltage applications,” in Conf. Rec. EPE, vol. 2, no. 1, p. 41, March 1992. [9] A. Nabae, I. Takahashi, and H. Akagi, “A new neutral-point-clamped PWM inverter,” IEEE Trans. Ind. Appl., vol. IA-17, no.5, pp. 518-523, Sep./Oct. 1981. [10] H. Akagi, “Large static converters for industry and utility applications,” Proc. IEEE, vol. 89, no. 6, pp. 976-983, Jun. 2001. [11] H. Akagi, H. Fujita, S. Yonetani, and Y. Kondo, “A 6.6-kV transformer- less STATCOM based on a five-level diode-clamped PWM converter: system design and experimentation of a 200-V 10-kVA laboratory model,” IEEE Trans. Ind. Appl., vol. 44, no. 2, pp. 672-680, March/April 2008. [12] T. A. Meynard, H. Foch, P. Thomas, J. Courault, R. Jakob, and M. Nahrstaedt, “Multicell converters: basic concepts and industry applica- tions,” IEEE Trans. Ind. Electron., vol. 49, no. 5, pp. 955-964, Oct. 2002. [13] J. Rodriguez, S. Bernet, B. Wu, J. O. Pontt, and S. Kouro, “Multi- level voltage-source-converter topologies for industrial medium-voltage drives,” IEEE Trans. Ind. Electron., vol. 54, no. 6, pp. 2930-2945, Dec. 2007. [14] R. Marquardt, and A. Lesnicar, “A new modular voltage source inverter topology,” in Conf. Rec. EPE 2003, CD-ROM. [15] M. Glinka, and R. Marquardt, “A new single-phase ac/ac-multilevel converter for traction vehicles operating on ac line voltage,” in Conf. Rec. EPE 2003, CD-ROM. [16] M. Glinka, “Prototype of multiphase modular-multilevel-converter with 2 MW power rating and 17-level-output-voltage,” in Conf. Rec. IEEE- PESC 2004, pp. 2572-2576. [17] M. Glinka, and R. Marquardt, “A new ac/ac multilevel converter family,” IEEE Trans. Ind. Electron., vol. 52, no. 3, pp. 662-669, June 2005. [18] http://www.modernpowersystems.com [19] J. Dorn, H. Huang, and D. Retzmann, “Novel voltage-sourced converters for HVDC and FACTS applications,” in Conf. Rec. CIGRE, 314, OSAKA 2007. [20] S. Allebrod, R. Hamerski, and R. Marquardt, “New transformerless, scalable modular multilevel converters for HVDC-transmission,” in Conf. Rec. IEEE-PESC 2008, pp. 174-179. [21] H. Akagi, S. Inoue, and T. Yoshii, “Control and performance of a transformerless cascade PWM STATCOM with star configuration,” IEEE Trans. Ind. Appl., vol. 43, no.4, pp. 1041-1049, July/Aug. 2007. [22] L. Maharjan, S. Inoue, and H. Akagi, “A transformerless energy storage system based on a cascade multilevel PWM converter with star configu- ration,” IEEE Trans. Ind. Appl., vol. 44, no.5, pp. 1621-1630, Sept./Oct. 2008. [23] L. Maharjan, S. Inoue, H. Akagi, and J. Asakura, “A transformerless battery energy storage system based on a multilevel cascade PWM converter,” in Conf. Rec. IEEE-PESC 2008, pp. 4798-4804. [24] http://www.hvdc.ca/ [25] H. Fujita, S. Tominaga, and H. Akagi, “Analysis and design of a dc voltage-controlled static var compensator using quad-series voltage- source inverters,” IEEE Trans. Ind. Appl., vol. 32, no. 4, pp. 970–977, Jul./Aug. 1996. A A. Another Averaging Control Method This subsection describes another averaging control method different from Fig. 4(a), which is characterized by controlling independent average voltages of positive and negative chopper- cells, ¯v Cup and ¯v Cun , which are given as follows: ¯v Cup = 1 4 4 j=1 v C ju , (21) ¯v Cun = 1 4 8 j=5 v C ju . (22) The current command of the positive chopper-cells, i ∗ Zup is represented as: i ∗ Zup = K 1 (v ∗ C − ¯v Cup ) + K 2 (v ∗ C − ¯v Cup )dt, (23) while that of the negative chopper-cells, i ∗ Zun is represented as: i ∗ Zun = K 1 (v ∗ C − ¯v Cun ) + K 2 (v ∗ C − ¯v Cun )dt. (24) Here, the voltage command of the positive chopper-cells, v ∗ Aup is given by: v ∗ Aup = K 3 (i Zu − i ∗ Zup ) + K 4 (i Zu − i ∗ Zup )dt, (25) while that of the negative chopper-cells, v ∗ Aun is given by: v ∗ Aun = K 3 (i Zu − i ∗ Zun ) + K 4 (i Zu − i ∗ Zun )dt. (26) B. Other Balancing Control Methods The balancing control forms an active power between the output voltage of each chopper-cell, v Bju and the corresponding arm current. The following describes two other balancing control methods different from Fig. 4(b). 1) The method using the arm currents i Pu and i Nu : For the positive chopper-cells numbered 1 to 4, v ∗ Bju is given by: v ∗ Bju = K 5 (v ∗ C − v C ju )i Pu . (27) For the negative chopper-cells numbered 5 to 8, v ∗ Bju is given by: v ∗ Bju = K 5 (v ∗ C − v C ju )i Nu . (28) Equations (27) and (28) indicate that v Bju (= v ∗ Bju ) contains the same frequency components as i Pu or i Nu , if the ac components included in v C ju can be eliminated. Hence, v Bju contains dc and 50-Hz components. The dc component of v Bju forms an active power with that of i Pu or i Nu , and the 50-Hz component of v Bju with that of i Pu or i Nu . To avoid an undesirable effect on the control system, ac components included in v C ju should be eliminated fully by a low-pass filter. 2) The method using the load current i u : For the positive chopper-cells numbered 1 to 4, v ∗ Bju is given by: v ∗ Bju = K 5 (v ∗ C − v C ju )i u . (29) For the negative chopper-cells numbered 5 to 8, v ∗ Bju is given by: v ∗ Bju = −K 5 (v ∗ C − v C ju )i u . (30) Equations (29) and (30) indicate that v Bju (= v ∗ Bju ) contains the same frequency components as i u or −i u . Note that i u is in phase with i Pu , whereas −i u is in phase with i Nu . The 50-Hz component included in v Bju forms an active power with that of i Pu or i Nu . Makoto Hagiwara (M’06) was born in Tokyo, Japan, in 1979. He received his B.S. and M.S. and Ph.D. degrees in electrical engineering from the Tokyo Institute of Technology in 2001, 2003, and 2006 respectively. Since April 2006, he has been an Assistant Professor in the department of electrical and electronic engineering at the Tokyo Institute of Technology. His research interests are self-commutated converters for utility applications. Hirofumi Akagi (M’87-SM’94-F’96) was born in Okayama, Japan, in 1951. He received the B.S. degree from the Nagoya Institute of Technology, Nagoya, Japan, in 1974, and the M. S. and Ph. D. degrees from the Tokyo Institute of Technology, Tokyo, Japan, in 1976 and 1979, respectively, all in electrical engineering. In 1979, he was with the Nagaoka University of Technology, Nagaoka, Japan, as an Assistant and then Associate Professor in the department of electrical engineering. In 1987, he was a Visiting Scientist at the Massachusetts Institute of Technology for ten months. From 1991 to 1999, he was a Professor in the department of electrical engineering at Okayama University, Okayama, Japan. From March to August of 1996, he was a Visiting Professor at the University of Wisconsin, Madison and then the Massachusetts Institute of Technology. Since January 2000, he has been a Professor in the department of electrical and electronic engineering at the Tokyo Institute of Technology, Tokyo, Japan. His research interests include power conversion systems, ac motor drives, active and passive EMI filters, high-frequency resonant-inverters for induction heating and corona discharge treatment processes, and utility applications of power electronics such as active filters for power conditioning, self- commutated BTB systems, and FACTS devices. He has authored or coau- thored more than 80 IEEE Transactions papers, and two invited papers in Proceedings of the IEEE. According to Google Scholar, the total citation index for all his papers is more than 7,000. He has made presentations many times as a keynote or invited speaker internationally. He was elected as a Distinguished Lecturer of the IEEE Industry Ap- plications Society (IAS) and PELS for 1998-1999. He received two IEEE IAS Transactions Prize Paper Awards in 1991 and 2004, and two IEEE PELS Transactions Prize Paper Awards in 1999 and in 2003, nine IEEE IAS Committee Prize Paper Awards, the 2001 IEEE William E. Newell Power Electronics Award, the 2004 IEEE IAS Outstanding Achievement Award, and the 2008 IEEE Richard H. Kaufmann Technical Field Award. He served as the President of the IEEE PELS for 2007-2008. . Control and Experiment of Pulse- Width- Modulated Modular Multilevel Converters Makoto Hagiwara, Member, IEEE, and Hirofumi Akagi, Fellow,. control with theoretical and experimental verifications. This paper deals with two types of pulse- width- modulated modular multilevel convert- ers (PWM-MMCs)