LI et al.: CONTROL OF HVDC LIGHT SYSTEM USING CONVENTIONAL AND DIRECT CURRENT VECTOR CONTROL APPROACHES 3107Fig. 1. Conﬁguration of a HVdc light transmission system .under different conditions. Finally, the paper concludes with thesummary of the main points.II. HVDC LIGHT SYSTEM AND CONTROLAn HVdc light system primarily consists of three parts: arectiﬁer station; an inverter station; and high-voltage dc trans-mission cables or lines , , . Fig. 1 shows an equivalentmodel of an HVdc light system, in which there are two VSC sta-tions. Each station is connected to an ac system via an equivalentimpedance representing the converter transformer and reactorbetween the VSC and the ac system. On the dc side of eachstation, a capacitor bank is connected. Depending on the direc-tion of active-power ﬂow, one station works as a rectiﬁer whilethe other operates as an inverter. Each VSC station has two de-grees of control freedom. One degree is used for reactive-powercontrol, while the other degree is used for active power or dcvoltage control.The HVdc light system, as shown in Fig. 1, allows fullyindependent control of active and reactive powers within theoperating range of the design. Normally, each station controlsits reactive power independent of the other station. However,the ﬂow of active power in the dc transmission system mustbe balanced, which means that the active power entering theHVdc system must be equal to the active power leaving it, ifneglecting the losses in the dc transmission system . Toachieve this power balance, one of the VSC stations has tocontrol the dc voltage, while the other VSC station should bedesigned for active-power control. Both stations contain eitherreactive power or ac voltage support control functions , ,as shown in Fig. 1. In Fig. 1, Vdc1, V ∗dc1, Vdc2, and V ∗dc2 representmeasured and reference dc voltages of the VSCs on the left andright sides of the HVdc light system, Vac1, V ∗ac1, Vac2, and V ∗ac2represent measured and reference voltages of the ac systems onthe left and right sides, and P∗ac1, Q∗ac1, P∗ac2, and Q∗ac2 signifythe active and reactive power references of the two ac systemswith power ﬂowing into each VSC deﬁned as positive.III. HVDC LIGHT TRANSIENT AND STEADY-STATE MODELS INd–q REFERENCE FRAMEFig. 2 depicts the equivalent system model of a VSC stationconnected to an ac system. A capacitor is shunt connected acrossFig. 2. Equivalent system model of a VSC station.the dc side of the voltage source PWM converter, and the shuntresistor across the dc bus models the power loss in the VSCstation. The voltages va 1, vb1, and vc1 represent the three-phaseline-to-neutral voltages injected by the PWM converter ontothe ac system, and the voltages va , vb, and vc signify the three-phase ac-system line-to-neutral voltages at the point of commoncoupling (PCC). The converter transformer and reactor betweenthe VSC and the PCC are represented as a series combinationof a resistor R and an inductor L.In the d–q reference frame, the voltage balance equation atthe interconnection of converter and ac system isvdvq= Ridiq+ Lddtidiq+ ωsL−iqid+vd1vq1(1)where ωs is the angular frequency of the ac system voltages.Also, vd, vq , vd1, and vq 1 represent the d and q componentsof the PCC voltages and the VSC output voltages, respectively.The currents id and iq represent the d and q components of thecurrent ﬂowing between the ac system and the VSC.Equation (1) can be expressed by a complex equation (2) us-ing space vectors, in which vdq , idq , and vdq 1 are instantaneousspace vectors of PCC voltage, line current, and VSC output volt-age. Under the steady-state condition, (2) becomes (3), whereVdq , Vdq 1, and Idq stand for the steady-state space vectors of thePCC and VSC output voltages and line current.vdq = R · idq + Lddtidq + jωsL · idq + vdq1 (2)Vdq = R · Idq + jωsL · Idq + Vdq1. (3)In the PCC voltage orientation frame –, , the PCCd-axis voltage is constant and the q-axis voltage is zero. Thus,the instantaneous active and reactive powers transferred fromthe ac system to the VSC are proportional to the d- and q-axis
3108 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 25, NO. 12, DECEMBER 2010Fig. 3. Conventional d–q vector control structure (The VSC in this ﬁgure is used for dc-voltage control.)currents, respectively, as shown by (4) and (5).pac(t) = vdid + vq iq = vdid (4)qac(t) = vq id − vdiq = −vdiq . (5)In terms of steady-state conditions, Vdq = Vd + j0 if the d-axis of the reference frame is aligned with the PCC voltageposition. Assuming Vdq1 = Vd1 + jVq1 and neglecting the re-sistance, the current ﬂowing from the ac system to the VSCisIdq =Vdq1 − VdqjXL=Vd1 − VdjXL+Vq1XL(6)where XL = jωsL is the reactance of the converter transformerand reactor between the VSC and the PCC.Using passive sign convention for the converter, power ﬂow-ing from the ac system to the VSC is positive. Thus, the powertransferred to the VSC from the ac system can be computedas Pac + jQac = Vdq I∗dq = VdI∗dq . By solving this power equa-tion together with (6), (7) is obtained. According to (7), the acsystem active and reactive powers Pac and Qac are controlledthrough q and d components Vq 1 and Vd1 of the VSC voltageinjected onto the ac system. If the resistor R is considered, sim-ilar power control characteristics of the VSC still exist underboth steady-state and transient open-loop control conditions, asin Pac = −VdVq1XLQac =VdXL(Vd − Vd1) . (7)IV. CONVENTIONAL CONTROL DESIGNOF HVDC LIGHT SYSTEMConventional VSC control of an HVdc light system has anested-loop structure consisting of a faster inner current loop anda slower outer control loop that generates d- and q-axis currentreferences i∗d and i∗q to the current loop controller –,, . If the VSC is used for active-power control, thed-axis current reference, according to (4), isi∗d =P∗acVd(8)where P∗ac is a desired active power transferred from the ac tothe dc system. If the VSC is used for dc voltage control, how-ever, the d-axis current reference is generated by a dc voltage-loop controller , , . The q-axis current referencefor both VSC stations of the HVdc light system is generatedeither according to reactive power or PCC voltage support con-trol requirement. For reactive-power control, the q-axis currentreference, according to (5), isi∗q =−Q∗acVd(9)where Q∗ac is a desired reactive power of the ac system. For PCCvoltage control, the q-axis current reference i∗q is obtained basedon the error signal between the PCC voltage setpoint and theactual PCC voltage to be controlled , , , .Fig. 3 shows the overall traditional d–q vector control struc-ture , , , in which the VSC is used for dc voltagecontrol. The d and q reference voltages, v∗d1 and v∗q1, include thed and q voltages, vd and vq , from the current-loop controllersplus the compensation terms shown in (10) and (11). The tworeference voltages are then used to generate a set of three-phasesinusoidal reference voltages, v∗a1, v∗b1, and v∗c1, to control thePWM converter. Thus, this control conﬁguration actually regu-lates id and iq (i.e., the ac system active and reactive powers)using vd and vq , respectively . But, according to Section III,d-axis voltage is only effective for reactive power or iq control;q-axis voltage is only effective for active power or id control. Al-though, the ﬁnal control voltage applied to the PWM convertercontains the compensation terms, those compensation terms arenot generated by the PI controllersv∗d1 = −vd + ωsLiq+ vd (10)v∗q1 = −vq − ωsLid. (11)It is possible that d- and q-axis current and voltage refer-ences generated by the nested-loop controller may exceed thephysical constraints of the VSC. Although rate limiters couldbe applied to the reference inputs to limit the changing rates ofthe controller output signals, those rate limiters cannot preventthe current and/or voltage signals generated by the nested-loopcontroller from exceeding the physical constraints of the VSC.
LI et al.: CONTROL OF HVDC LIGHT SYSTEM USING CONVENTIONAL AND DIRECT CURRENT VECTOR CONTROL APPROACHES 3109Therefore, the following strategies are used in the design of thenested-loop control system.1) To prevent the converter from entering the nonlinear mod-ulation mode, a saturation mechanism is applied to theoutput voltage of the controller if the amplitude of thereference voltage generated by the controller exceeds thePWM saturation limit. The general strategy is to set a lim-itation on |vdq1|∗but keeps v∗dq1 unchanged, as shown by(12) , , where v∗d1n e wand v∗q1 new are the d and qcomponents of the modiﬁed controller output voltage, andVmax is the maximum allowable d–q voltage. It is foundthat any other saturation mechanisms could cause moresystem oscillation and unbalancev∗d1 new = Vmax · cos( v∗dq1)v∗q1 new = Vmax · sin( v∗dq1). (12)2) To prevent the converter from exceeding the rated cur-rent, the q-axis current reference is adjusted if the ampli-tude of the reference current generated by the dc voltageand reactive-power control loops exceeds the rated currentlimit. The general strategy is keeping the d-axis currentreference unchanged to maintain dc voltage control ef-fectiveness while modifying the q-axis current referenceto satisfy the reactive-power control demand as much aspossible, as shown by (13) , i∗d new = i∗d i∗q new = sign (i∗q ) · (i∗dq max)2 − (i∗d)2.(13)V. DIRECT CURRENT VECTOR CONTROL DESIGN OF HVDCLIGHT SYSTEMThe proposed direct current vector control strategy includesthree parts: 1) transformation from power control to currentcontrol; 2) development of a current-loop control scheme; and3) conversion from current control signals to voltage controlsignals.A. Transformation From Power Control to Current ControlThe proposed method ﬁrst transforms power control to currentcontrol based on (4) and (5). For example, according to (4),active power (Pac) is proportional to d-axis current (id) so thatPac can be controlled through the regulation of id. Therefore,for a reference and actual measured active powers P∗ac and Pac,a control strategy for id is developed in the following way. IfP∗ac > Pac, id should be increased. If P∗ac < Pac, id shouldbe decreased. Hence, id can be adjusted through a PI controlmechanism until Pac = P∗ac.B. Development of a Current-Loop Control SchemeIn the control of power converters for electric power applica-tions, a nested-loop control structure is normally used, in whicha fast current control loop is critical to assure a high power qual-ity in terms of harmonics and unbalance. Although direct powercontrol (DPC) strategies have been proposed recently, many dis-Fig. 4. Actual, reference, and tuning currents in a transient feedback controlcase study.advantages, such as poor power quality, have been found due tothe elimination of the current control loop –.Unlike DPCs, the proposed control structure employs a di-rect current control (DCC) mechanism after the power-loopcontroller, in which the output from the power-loop controller isused as the reference for the current-loop controller. However,instead of generating a d- or q-axis voltage based on a d- or q-axiscurrent error signal, as shown by the standard control structure(see Fig. 3), the proposed scheme outputs a current signal fromthe d or q current-loop controller. The output of the proposedcurrent-loop controller is a d or q tuning current, while the inputerror signal tells the controller how much the tuning currentshould be adjusted during the dynamic control process. There-fore, for a reference and actual measured d-axis current i∗d and id,for example, a control strategy is developed in the following way.If i∗d > id, the d-axis tuning current should be increased. If i∗d <id, the d-axis tuning current should be decreased. Thus, the tun-ing current can be adjusted through a PI control mechanism untili∗d = id. It is necessary to point out that this tuning current couldbe different from the actual measured current. Fig. 4 demon-strates a relationship among the actual, reference, and tuningd-axis currents for a case study conducted in Section VI.For the development of the proposed DCC method, intelli-gent control concepts have been adopted. The key goal is tominimize absolute or rms error between the desired and actualvalues of a controlled variable through an adaptive tuning strat-egy , . However, major challenge for an intelligent con-troller is that the learning is normally required for an unknownsystem, and incorrect learning, particularly in the presence ofsignal noise, could result in unstable system performance. Thedevelopment of the proposed control mechanism has also beenbased on predictive current control (PCC) notions –.However, instead of using a current-loop controller, the PCCsuse predictive current and voltage equations or approaches togenerate control voltages for the VSC so that the feedback con-trol advantages associated with the current-loop controller arelost in the PCCs –. To overcome the challenges, thepower control to current control relationship is implementeddirectly through the proposed DCC design with feedback.C. Conversion From Current Control Signalsto Voltage Control SignalsDue to the nature of the voltage source converter, the d and qtuning current signals generated by the current-loop controllersmust be transformed to the d and q voltage signals for the VSCcontrol. There are many different ways to accomplish the signal
3110 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 25, NO. 12, DECEMBER 2010Fig. 5. Direct current vector control structure (The VSC in this ﬁgure is used for dc-voltage control.)Fig. 6. Basic concept of fuzzy-PI based control mechanism.transformation, including artiﬁcial neural networks. However,a convenient way for a stable and reliable current-to-voltagetransformation can be achieved from (1), as shown by (14)and (15), which is equivalent to the transient d–q (1) after beingprocessed by a low-pass ﬁlter for the purpose to reduce the largeoscillations in the d and q reference voltages applied directly tothe converterv∗d1 = −Rid + ωsLiq + vd (14)v∗q1 = −Riq − ωsLid. (15)Fig. 5 shows the overall DCC structure, which consists of ad-axis current loop for dc voltage control and a q-axis currentloop for reactive power or PCC voltage support control. Signal-processing technology is applied to the measured voltages andcurrents to prevent the high-order harmonics from entering thecontrollers. The current-loop controller could also operate ona mechanism that combines PID, fuzzy, and adaptive controltechnologies , . Details about the fuzzy-tuned PI controlscheme are available in , in which the PI part operates ona direct target control principle, while the fuzzy and adaptiveparts adjust the PI parameters based on the error between thecontrolled variable and its target value, and the change in error,as illustrated by Fig. 6. However, similar to the results presentedin , it is found that the improvement of the fuzzy-tuned PIcontrol is limited. The initially tuned PI gains are still the mostimportant factors affecting the overall system performance.In addition, a nonlinear programming formulation, as shownin the following, is developed to prevent the resultant d–q cur-rent from exceeding the VSC rated current and to prevent theconverter from entering the nonlinear modulation mode, whereIrated is the rated phase rms current, Vconv is the phase rmsvoltage of the VSC output voltage, and Q∗ac is the referencereactive power absorbed by the VSC from the ac system. If theVSC is used for dc voltage control, the basic principle of thenonlinear programming formulation is that under the converterrated power and PWM saturation constraints, the system shouldbe operated to achieve the dc voltage control goal while mini-mizing the difference between the reference and actual reactivepowers. If the VSC is used for active-power control, the basicprinciple is that under the VSC rated power and PWM saturationconstraints, the system should be operated to achieve the active-power control goal while minimizing the difference between thereference and actual reactive powers.Minimize|Qac − Q∗ac| .Subject toVdc = V ∗dc or Pac = P∗ac,I2d + I2q3≤ Irated,Vconv =V 2d1 + V 2q13≤Vdc2√2.The nonlinear programming strategy is implemented in thefollowing way. If |i∗dq | generated by the dc-voltage/active-powerand reactive-power control loops exceeds the rated current limit,i∗d and i∗q are modiﬁed by (13). If |v∗dq1| generated by the currentcontrol loops exceeds the PWM saturation limit, v∗d1 and v∗q1are modiﬁed by (16). In fact, according to Section III, (16)represents an optimal control strategy keeping the q-axis voltagereference v∗q1 unchanged so as to maintain the dc-voltage oractive-power control effectiveness while modifying the d-axisvoltage reference v∗d1 to meet the reactive-power control demandas much as possiblev∗d1 new = sign(v∗d1) · (v∗dq1 max)2 − (v∗q1)2v∗q1 new = v∗q1. (16)
LI et al.: CONTROL OF HVDC LIGHT SYSTEM USING CONVENTIONAL AND DIRECT CURRENT VECTOR CONTROL APPROACHES 3111Fig. 7. HVdc light system with feedback control in SimPowerSystems.Fig. 8. Detailed components of a VSC station.VI. PERFORMANCE EVALUATION AND COMPARISONFor performance evaluation of the conventional and proposedcontrol mechanisms, an HVdc light system is simulated inMATLAB using SimPowerSystems (see Fig. 7). Detailed pa-rameters of the HVdc system are provided in Tables I and II ofthe Appendix. The PI gains of current and voltage controllers aregiven in Table III. A three-level neutral-point-clamped VSC (seeFig. 8) is used at both stations for improved power quality .Major measurements include PCC three-phase voltages, cur-rents, dc capacitor voltages at both VSC stations, and active andreactive powers of the two ac systems.In the ﬁrst case, HVdc control is evaluated under the condi-tion that the controller output voltage is within or around theconverter PWM saturation limit (see Figs. 9 and 10). The activepower ﬂows from ac system 1 to ac system 2. Therefore, theleft VSC operates as a rectiﬁer and the right VSC operates as aninverter. The initial active-power reference of the rectiﬁer VSCis 30 MW. The reactive-power reference supplied to ac systemsby both VSCs is set at 0 MVar. The PCC voltages are not con-trolled. At t = 4 s, there is a ramp change of the active-powerreference at the rectiﬁer station from 30 to 80 MW. At t = 8 s,the active-power reference changes to 50 MW. It can be seenfrom Figs. 9 and 10 that when the converter operates in its linearmodulation mode, both the conventional and the proposed con-trol mechanisms have similar performance, but the dc systemvoltage is more stable using the proposed control approach es-pecially when the converter operates around its PWM saturationlimit [see Figs. 9(c) and 10(c)]. The active power in Figs. 9(a)and 10(a) is positive, but negative in Figs. 9(b) and 10(b), im-plying that the power transfers from ac system 1 to the rectiﬁerFig. 9. Performance evaluation using traditional control approach (withinPWM saturation limit). (a) Active and reactive powers at PCC1 . (b) Activeand reactive powers at PCC2 . (c) Capacitor voltages at inverter and rectiﬁerstations. (d) Three-phase currents in ac system 2.station and from the inverter station to ac system 2. Due to thelosses in the HVdc system, there is a small difference betweenthe active powers at the two stations. The reactive power in bothac systems is around −40 MVar, demonstrating that the reactivepower is generated by the ac ﬁlter and the contribution by theVSCs is almost 0 MVar. Both approaches have similar powerquality [see (Figs. 9(d) and 10(d)]. It is necessary to point outthat the results, as shown in Figs. 9 and 10, are slightly differentbecause the PCC voltages are not controlled, and the converteroperates near PWM saturation limit between 4 and 8 s.In the second case, the HVdc system is evaluated for the con-dition that the controller output voltage may exceed the PWM
3112 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 25, NO. 12, DECEMBER 2010Fig. 10. Performance evaluation using proposed control approach (withinPWM saturation limit). (a) Active and reactive powers at PCC1 . (b) Active andreactive powers at PCC2 . (c) Capacitor voltages at inverter and rectiﬁer stations.(d) Three-phase currents in ac system 2.saturation limit evidently (see Figs. 11 and 12). This may hap-pen for a large reactive-power requirement. The active-powerreferences from ac system 1 to ac system 2 are 30 MW at t= 0 s, 80 MW at t = 5 s, 50 MW at t = 9 s, and 30 MWat t = 11 s. Again, the PCC voltages are not controlled. Thereactive-power reference of the rectiﬁer station is 0 kVar, whilethe reactive-power references of the inverter station are 0 MVarat t = 0, −30 MVar at t = 3 s (supplying reactive power to acsystem 2), and 0 MVar at t = 7 s.For the conventional control strategy, the HVdc system worksproperly for active and reactive power controls at the rectiﬁerstation, and for dc voltage and reactive power control at theinverter station if the controller output voltage does not exceedthe PWM saturation limit (before t = 3 s in Fig. 11). However,if the controller output voltage at either station is over the PWMsaturation limit (after t = 3 s in Fig. 11), the dc voltage of theHVdc system becomes uncontrollable. The more the controlleroutput voltage exceeds the limit, the more the dc voltage deviatesfrom the reference dc voltage. An interesting phenomenon is thatafter the controller output voltage exceeds the PWM saturationlimit, the dc system voltage ﬂoats with the active and reactivepower demands [see Fig. 11(c)].For the same condition of Fig. 11, the proposed DCC strat-egy demonstrates a superior performance (see Fig. 12). Shortlyafter the start of the system, the controllers at both the rectiﬁerFig. 11. Performance evaluation using traditional control approach (beyondPWM saturation limit). (a) Active and reactive powers at PCC1 . (b) Active andreactive powers at PCC2 . (c) DC voltages at inverter and rectiﬁer stations. (d)PCC voltages in ac systems 1 and 2.and inverter stations regulate: 1) the active power transferredfrom the rectiﬁer to the inverter stations at the desired value;2) the dc voltage at the target dc voltage level; and 3) reactivepower absorbed from the ac systems by both stations at thereactive-power references. As the reactive-power reference ofthe inverter station change to −30 MVar at t = 3 s, a conditionunder which the output voltage from the controller at the inverterstation exceeds the PWM saturation limit, the proposed controltechnique regulates the ac system reactive power according tothe optimal control rule, i.e., maintaining dc-link voltage as theﬁrst priority while compensating reactive power as much as pos-sible. At t = 5 s, as the active power transferred from ac system1 to ac system 2 increases gradually from 30 to 80 MW, thereactive power generated by the inverter station drops steadilyaccording to the optimal control rule [see Fig. 12(b)], while thevoltage dynamics in both ac systems 1 and 2 are not manifested[see Fig. 12(d)]. As the converter at the inverter station operateswithin the PWM saturation limit at t = 9 s, the proposed controlmechanism quickly returns to its normal operation. For eachactive or reactive power demands changing from one conditionto another, the proposed control technique can adjust the actualdc system voltage to the reference value quickly with muchreduced oscillation [see Figs. 12(c) and 11(c)].
LI et al.: CONTROL OF HVDC LIGHT SYSTEM USING CONVENTIONAL AND DIRECT CURRENT VECTOR CONTROL APPROACHES 3113Fig. 12. Performance evaluation using proposed control approach (beyondPWM saturation limit). (a) Active and reactive powers at PCC1 . (b) Active andreactive powers at PCC2 . (c) DC voltages at inverter and rectiﬁer stations. (d)PCC voltages in ac systems 1 and 2.Fig. 13. Illustration of d–q control voltage adjustment under an over modula-tion conditionThe difference between Figs. 11 and 12 can be explainedmore clearly by Fig. 13. Assume that the output voltage ofthe current-loop controller at the inverter station is at point Afor an over modulation condition (i.e., outside the circle of 1p.u. radius). Then, according to (12), the conventional approachadjusts the control voltage applied to the power converter topoint B, which reduces both d and q components of the controlvoltage. The decline of the q-axis voltage shrinks the powertransferred from dc to ac systems according to (7). Thus, thepower from ac system 1 cannot be effectively transferred to acsystem 2 so that the dc voltage increases. However, the proposedapproach, based on (16), changes the control voltage applied tothe power converter to point C, which reduces d-axis voltagebut keeps the q-axis voltage unchanged. Hence, active powerfrom ac system 1 is effectively transferred to ac system 2 sothat the dc system voltage is not affected by the adjustmentof the control voltage, while the reactive power is regulatedaccording to the optimal control rule. Meeting the optimizationneed requires that the q-axis loop for reactive-power controlstops changing (see Fig. 5), while the d-axis loop for activepower or dc voltage control operates when an overmodulationcondition appears. However, due to the competing d–q controlnature (see Fig. 3), the conventional method conﬂicts with thisoptimal control requirement. It is found that when this optimalcontrol strategy is applied to the conventional control structure(see Fig. 3), large oscillation and unbalance of the HVdc systemresult.VII. GRID VOLTAGE SUPPORT CONTROL EVALUATIONThe conventional and proposed control mechanisms are alsoevaluated for ac system voltage support control applications un-der balanced and unbalanced faults. Both VSC stations of theHVdc system can be used for voltage support control of the acsystem, to which a VSC is connected. The control objective isto maintain the PCC voltage at a desired value for any voltageﬂuctuation due to load changes or for any voltage sag due toa fault. Assume that there is a voltage sag in the ac system 2at a certain time, which causes a voltage drop at the PCC towhich the inverter VSC is connected. The extent of the voltagedrop depends on the location and type of the load, or fault in theac transmission system. For the voltage control application, theq-axis current reference, according to Section IV, is determinedfrom the error signal between the desired and actual PCC volt-ages to be controlled. Therefore, the alteration of the currentreference i∗q may cause more stability problems to the HVdcsystem.Figs. 14 and 15 compare the performance of the traditionaland proposed control approaches for PCC voltage support con-trol when there is a small voltage drop on the PCC bus of theac system 2, while the active power transferred from ac system1 to ac system 2 is 30 MW and the reactive-power referencesof the rectiﬁer VSC is zero. The voltage drop starts at t =2 s and ends at t = 4 s. Under the low-voltage drop condi-tion, the reactive power needed for voltage support is small sothat the converter operates within the PWM saturation limit dur-ing the voltage support control period. As shown by the ﬁgures,both the traditional and proposed control approaches, havingsimilar performance, are able to achieve the dc voltage and gridvoltage support control goals.However, if a voltage drop causes the output voltage from thecontroller exceeding the PWM saturation limit, the conventionalcontrol approach could result in critical problems in both the dcand the ac systems (see Fig. 16). During the voltage drop period,the conventional control technique is unable to maintain the dccapacitor voltage around the reference value depending on theextent of the voltage drop. After the voltage drop is cleared,large oscillations appear on the dc capacitor voltage of the HVdc
3114 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 25, NO. 12, DECEMBER 2010Fig. 14. Conventional control mechanism in grid voltage support control undera low-voltage drop condition. (a) DC voltages at inverter and rectiﬁer stations.(b) PCC voltage with voltage support control. (c) Active and reactive powers atPCC2 .Fig. 15. Proposed control mechanism in grid voltage support control undera low-voltage drop condition. (a) DC voltages at inverter and rectiﬁer stations.(b) PCC voltage with voltage support control. (c) Active and reactive powers atPCC2 .system, and the high oscillation and unbalance are found in theac system currents. These dynamics are manifested in the acsystem 1 to which the rectiﬁer station is connected (see Fig. 16).The greater the voltage drop is, the greater and longer lasting theoscillations found in the dc and ac systems after the voltage dropis cleared become, which would trigger the protection systemsto disable the HVdc system.Fig. 16. Conventional control mechanism in grid voltage support control fora high-voltage drop condition. (a) DC voltages at inverter and rectiﬁer stations.(b) AC system bus voltages at PCC1 and PCC2 . (c) Active and reactive powersat PCC1 . (d) Active and reactive powers at PCC2 .Fig. 17. Proposed control mechanism in grid voltage support control for ahigh-voltage drop condition. (a) DC voltages at inverter and rectiﬁer stations.(b) AC system bus voltages at PCC1 and PCC2 . (c) Active and reactive powersat PCC1 . (d) Active and reactive powers at PCC2 .
LI et al.: CONTROL OF HVDC LIGHT SYSTEM USING CONVENTIONAL AND DIRECT CURRENT VECTOR CONTROL APPROACHES 3115Fig. 18. Conventional method in grid voltage support control for a single-line-to-ground fault. (a) DC voltage at inverter and rectiﬁer stations. (b) Ac-tive and reactive powers at PCC1 . (c) Active and reactive powers at PCC2 .(d) Three-phase current at PCC2 .Compared to the traditional control approach, the proposedcontrol mechanism has much improved performance. Fig. 17shows the behavior of the proposed DCC approach for the samesystem condition used in Fig. 16. Unlike the conventional con-trol strategy, the proposed control method effectively regulatesthe HVdc system under the optimal control rule during the volt-age drop period by retaining the dc capacitor voltage constantwhile supporting the grid voltage as much as possible. After thevoltage drop is cleared, the proposed control strategy returns tonormal system operation effectively and stably, and the systemalways responses to a control condition transition smoothly withreduced oscillation (see Fig. 17).Figures 18 and 19 compare the performance of the traditionaland proposed control approaches for PCC voltage support con-trol when a single-line-to-ground fault causes an unbalancedvoltage drop on PCC2 bus. If the voltage drop does not causethe controller output voltage exceeding the PWM saturation,both the conventional and proposed control methods can stabi-lize the dc system voltage and effectively maintain the active andreactive power demands at both stations. Unlike Figs. 14 and 15,the three-phase currents in ac system are unbalanced because ofthe unbalanced voltage drop on PCC2 bus [see Figs. 18(d) and19(d)]. However, this unbalance phenomenon is not manifestedin ac system 1 due to the HVdc link that separates ac system 1from ac system 2 [see Figs. 18(b) and 19(b)].Fig. 19. Proposed method in grid voltage support control for a single-line-to-ground fault. (a) DC voltage at inverter and rectiﬁer stations. (b) Active andreactive powers at PCC1 . (c) Active and reactive powers at PCC2 . (d) Three-phase current at PCC2 .Fig. 20. Conventional method in grid voltage support control for a line-to-linefault. (a) DC voltage at inverter and rectiﬁer stations. (b) Active and reactivepowers at PCC1 . (c) Active and reactive powers at PCC2 . (d) Three-phasecurrents at PCC2 .
3116 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 25, NO. 12, DECEMBER 2010Fig. 21. Proposed method in grid voltage support control for a line-to-linefault. (a) DC voltage at inverter and rectiﬁer stations. (b) Active and reactivepowers at PCC1 . (c) Active and reactive powers at PCC2 . (d) Three-phasecurrents at PCC2 .Figures 20 and 21 compare the performance of the traditionaland proposed control approaches for PCC voltage support con-trol for an unbalanced voltage drop on PCC2 bus caused by aline-to-line fault. For the same short-circuit power per phase,a line-to-line fault has a higher possibility of causing the con-troller output voltage to exceed the PWM saturation limit thana line-to-ground fault. Beyond the PWM saturation limit, theVSC using the conventional control method would cause mal-function of the HVdc system (see Fig. 20), which is similar tothe VSC behavior under balanced fault condition (see Fig. 16).However, the oscillation of active and reactive powers as well asthe current waveform, as shown in Figs. 20(c) and (d), indicatesa current unbalance. But, under the same unbalanced fault, theproposed control method has superior performance, includingmore stable dc-link voltage, better isolation of the two ac sys-tems through the HVdc link, and improved system stability andreliability (see Fig. 21).VIII. CONCLUSIONThis paper evaluates and compares the conventional and anew direct current vector control design for control of an HVdclight system. The proposed control mechanism employs an op-timal control strategy and operates by either using PI controlmechanism or integrating PI, fuzzy, and adaptive control tech-niques together.For the rectiﬁer and inverter stations of an HVdc light sys-tem, if the converters operate within the PWM saturation limit,both the conventional and the proposed direct current vectorcontrol mechanisms can: 1) meet the reactive- and active-powercontrol demands; and 2) maintain the dc system voltage arounda reference value. But, the proposed control method has bet-ter dynamic performance, especially for variable active powertransferred from one ac system to the other. If the output voltageof the current-loop controller at either station exceeds the PWMsaturation limit, the conventional control mechanism could drivethe HVdc system into a state, in which the dc system voltage be-comes uncontrollable. However, the proposed control approachmanages the HVdc system properly in an optimal control modeif an over modulation condition appears.In the grid voltage support control, if a voltage drop does notcause converter operation beyond the PWM saturation limit,the HVdc system operates appropriately using both the con-ventional and the proposed control techniques. However, if avoltage drop causes the output voltage from the controller to ex-ceed the PWM saturation limit, it could force the conventionalcontrol approach into a state that results in critical stability prob-lems to the HVdc system, while the proposed control schemeoperates in an optimal control strategy and has much more su-perior performance than the conventional control approach forany voltage drop caused by either a balanced or an unbalancedfault.APPENDIXTABLE INETWORK DATA (SEE FIG. 7)
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3118 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 25, NO. 12, DECEMBER 2010 J. M. Zurada, Introduction to Artiﬁcial Neural Systems. Boston, MA:PWS Publishing Company, 1992. Y. A.-R. I. Mohamed and E. F. El-Saadany, “Robust high bandwidthdiscrete-time predictive current control with predictive internal model:A uniﬁed approach for voltage-source PWM converters,” IEEE Trans.Power Electron., vol. 23, no. 1, pp. 126–136, Jan. 2008. J. C. Moreno, J. M. Esp´ı Huerta, R. G. Gil, and S. A. Gonz´alez, “A robustpredictive current control for three-phase grid-connected inverters,” IEEETrans. Ind. Electron., vol. 56, no. 6, pp. 1993–2004, Oct. 2009. J. Rodriguez, J. Pontt, C. A. Silva, P. Correa, P. Lezana, P. Cortes, andU. Ammann, “Predictive current control of a voltage source inverter,”IEEE Trans. Ind. Electron., vol. 54, no. 1, pp. 495–503, Feb. 2007. J. Yen and R. Langari, Fuzzy Logic: Intelligence, Control, and Information.Upper Saddle River, NJ: Prentice-Hall, 1999. A. Luo, C. Tang, Z. Shuai, J. Tang, X. Xu, and D. Chen, “Fuzzy-PI-based direct-output-voltage control strategy for the STATCOM used inutility distribution systems,” IEEE Trans. Ind. Electron., vol. 56, no. 7,pp. 2401–2411, Jul. 2009. A. Bendre, S. Krstic, J. V. Meer, and G. Venkataramanan, “Comparativeevaluation of modulation algorithms for neutral-point-clamped convert-ers,” IEEE Trans. Ind. Appl., vol. 41, no. 2, pp. 634–643, Mar./Apr. 2005.Shuhui Li received the B.S. and M.S. degrees in elec-trical engineering from Southwest Jiaotong Univer-sity, Chengdu, China, in 1983 and 1988, respectively,and Ph.D. degree in electrical engineering from TexasTech University, Lubbock, in 1999.From 1988 to 1995, he was with the Schoolof Electrical Engineering, Southwest Jiaotong Uni-versity, where his research interest focused on themodeling and simulation of large dynamic systems,dynamic process simulation of electriﬁed railways,power electronics, power systems, and power systemharmonics. From 1995 to 1999, he was engaged in research on wind power,artiﬁcial neural networks, and applications of massive parallel processing. In1999, he joined Texas A&M University, Kingsville, as an Assistant Professor,where became an Associate Professor in 2003. In 2004 and 2006, he was withOak Ridge National Laboratory for simulation system development on super-computers. Since 2006, he has been an Associate Professor at the Universityof Alabama, Tuscaloosa. His current ﬁelds of interests include renewable en-ergy systems, power electronics, power systems, electric machines and drives,ﬂexible AC transmission systems (FACTS), intelligent control, microgrids, anddistributed generation.Timothy A. Haskew (M’86–SM’02) received theB.E.E., M.S., and Ph.D. degrees in electrical engi-neering from Auburn University, Auburn, Alabama,in 1987, 1988, and 1991, respectively.Since 1991, he has been with the University ofAlabama, Tuscaloosa, where he is currently a Profes-sor of Electrical and Computer Engineering, and alsothe Director of the Electromechanical Systems Lab-oratory and the Electrical and Computer EngineeringGraduate Program. He is the author or coauthor ofover 50 refereed publications and 2 book chapters.His research interests include electromechanical systems, electric machinery,power electronics, and control systems.Dr. Haskew was the Seminar Chair in the program committee for the IEEEApplied Power Electronics Conference.Ling Xu received the B.Eng. degree from HuazhongUniversity of Science and Technology, Wuhan,China, in 2007, and the M.Sc. degree from the Uni-versity of Alabama, Tuscaloosa, in 2009, both in elec-trical engineering. He is currently working toward thePh.D. degree at the University of Alabama.His research interest includes renewable energyconversion, power converter control, and VSC-HVdctransmission