Assessment of communication-independent grid code compatibility solutions for VSC–HVDC connected offshore wind farms

Electric Power Systems Research 121 (2015) 38–51

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Assessment of communication-independent grid code compatibility solutions for VSC–HVDC connected offshore wind farms Sotirios I. Nanou ∗ , Georgios N. Patsakis, Stavros A. Papathanassiou School of Electrical and Computer Engineering, National Technical University of Athens (NTUA), 15780 Athens, Greece

a r t i c l e

i n f o

Article history: Received 18 July 2014 Received in revised form 29 October 2014 Accepted 2 December 2014 Keywords: Grid code compatibility Fault ride-through Frequency response VSC–HVDC Wind turbines Dynamic model

a b s t r a c t Offshore wind farms connected to the mainland through High Voltage DC links based on Voltage Source Converters (VSC–HVDC) are subject to Grid Code (GC) requirements, such as frequency response, Fault Ride Through (FRT) capability, and active and reactive power control. Due to the presence of HVDC links, GC compatibility becomes a challenging task from a control and regulation perspective, since the offshore network AC voltage and frequency are decoupled from the onshore grid voltage and frequency and hence grid-side disturbances are not directly experienced by the Wind Turbines (WTs). To address this challenge of modulating the response of offshore WTs according to the onshore grid conditions, without resorting to dedicated communication schemes, which may introduce delays and reliability concerns, three alternative control strategies are presented and evaluated for the offshore HVDC VSC, which are based only on local measurements. Each strategy can cope with onshore frequency response, FRT and power control requirements at the same time. A coordinated control approach between the WT and HVDC converters is also proposed by integrating a voltage dependent current modulation strategy into the default WT current controllers. Time domain simulations are carried out in DIgSILENT PowerFactory in order to verify the effectiveness of the proposed control schemes. © 2014 Elsevier B.V. All rights reserved.

1. Introduction As wind power penetration in electrical systems steadily rises and potential onshore wind sites become scarcer, offshore Wind Farms (WFs) constitute a steadily increasing portion of new wind power capacity. For large distances from shore, the state-of-theart High Voltage DC (HVDC) transmission technology [1,2] utilizing Voltage Source Converters (VSCs), is employed. Transmission System Operators (TSOs) impose technical requirements for the connection of offshore WFs, in order to ensure the secure operation of the mainland grid [3–5]. Among these requirements, particularly important are primary frequency support and Fault Ride-Through (FRT) capability, both requiring fast control of the WF active output power in response to grid frequency and voltage variations. In the case of WFs directly connected to the onshore grid, such requirements are now satisfied, using suitable control schemes [6–13]. In the presence of an HVDC link, however, the offshore WF network is decoupled from the onshore grid and therefore the WTs are unable to detect and respond to onshore grid disturbances, based only on local measurements.

∗ Corresponding author. Tel.: +30 6973931447. E-mail address: [email protected] (S.I. Nanou). http://dx.doi.org/10.1016/j.epsr.2014.12.002 0378-7796/© 2014 Elsevier B.V. All rights reserved.

An obvious solution to address this problem is the use of a dedicated communication link between the onshore grid end and the individual WTs in the offshore side [14–17]. However, remote communication links may present reliability and performance issues (complete loss of communication, latency or reduced data rate) [14,17,18], which may severely compromise the effectiveness of such a solution. Further, the response of WTs to commands received by their controller (e.g. power set-point commands) is much slower than what can be achieved when the WT responds to actual electrical disturbances. For these reasons, attention is focused on developing communication-independent control schemes, which rely only on local measurements at the offshore network side, as discussed in the following. As far as primary frequency support is concerned, a communication-independent solution widely proposed in the literature consists in varying the DC voltage reference of the HVDC link, regulated by the onshore VSC, as an indication of onshore frequency deviations. The offshore VSC detects the changes of DC voltage and effects an offshore frequency excursion that emulates the onshore conditions, triggering thus the offshore WT frequency response [18–24]. Concerning FRT capability, besides the fast voltage support that the onshore VSC should provide when a voltage dip occurs, the main challenge is to avoid DC over-voltages that might trip the

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Nomenclature 

difference operator (deviation from the predisturbance value) * reference phase angle provided to the VSC model  PLL,REC ,  PLL,WT PLL phase angle of the voltage at the REC and WT GRSC terminals ωC cut-off frequency of the washout filter of the SEC controller (Strategy II) WT rotor speed ωr Eth , Zth Thevenin equivalent for onshore grid (voltage and impedance) fgrid , fn onshore and nominal grid frequency ∗ ,f ,ω foff off off reference and measured offshore grid frequency

∗ ∗ , id,REC , iq,REC , iq,REC reference and measured d- and q-axis id,REC components of REC current id0,REC , iq0,REC unconstrained reference d- and q-axis components of REC current ∗ ∗ , id,WT , iq,WT , iq,WT reference and measured d- and q-axis id,WT components of WT GRSC current id,WT0 pre-disturbance value of d-axis component of WT GRSC current id0,WT , iq0,WT unconstrained reference d- and q-axis components of WT GRSC current REC rated current Imax iREC,abc , ioff,abc REC and SEC three-phase output currents iWT WT GRSC output current magnitude iWTg,abc , iWT,abc WT generator and GRSC three-phase output currents derivative gain of the WT current magnitude conKdi troller frequency droop gain of the WT controller Kdr Kf offshore frequency reference droop gain of the SEC controller (Strategies I, II) Kgc frequency droop gain imposed by the GC virtual inertia control gain of the WT controller Kin Kr reactive current regulation gain during onshore voltage dips, imposed by the GC offshore voltage reference droop gain of the SEC Kv controller (Strategy III) Kvdc DC voltage reference droop gain of the REC controller (mode 1) Kvi proportional gain of the WT current magnitude controller offshore voltage reference droop gain of the SEC Kw controller (Strategy II) L WT GRSC filter inductance m∗dq modulation indices (in the SRF) provided to the VSC model mi scaling factor for the WT output current magnitude PMPPT WT active power reference provided by the MPPT controller REC active power set-point imposed by the TSO p∗REC pREC , pSEC REC and SEC output active power WF output active power pWF p∗WT , q∗WT active and reactive power reference provided by the WT GRSC controller q∗REC , qSEC reference and measured REC output reactive power rdc HVDC line resistance ROCOFmax df/dt withstand capability of the offshore WTs d-axis voltage component provided by the PLL of the vd,WT WT GRSC controller

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Vdcmax , Vdcmin maximum and minimum DC voltage limits of the REC controller (mode 2) v∗dc,REC , vˆ dc,REC reference and estimate of the DC voltage at the REC terminals vdc,REC , vdc,SEC DC voltage at the REC and SEC terminals vdc,WT WT DC link voltage vgrid onshore grid voltage magnitude vgrid,abc , voff,abc onshore and offshore three-phase grid voltages v∗off , voff , voff 0 reference, measured and pre-disturbance offshore grid voltage magnitude Vthres voltage activation threshold for fast voltage support by the REC controller vWTg,abc , vWT,abc WT generator and GRSC three-phase grid voltages p∗WF,r power reserve command imposed by the TSO v∗dc,r DC voltage deviation reference signal for power reserve provision

entire HVDC link, induced by the fact that active power export to the onshore gird is drastically reduced. To avoid this risk, without resorting to the installation of costly high-power HVDC choppers, the active power generated by the WF needs to be immediately reduced, mainly by employing either of the following two concepts proposed in the literature: • “Frequency modulation” [24–27]: When the DC voltage exceeds a predefined threshold, the offshore VSC increases its frequency, leading thus the WTs to curtail their active output power. • “Voltage dip” [15,26–29]: DC voltage overshoots trigger a controlled voltage dip at the offshore grid, which will inevitably reduce the output power of the WF. So far, FRT and primary frequency response have been mostly studied in the literature independently from each other, when a communication-less scheme is assumed. However, this is not effective, because controllers designed and optimized for one type of disturbance may not be effective in another. For instance, the offshore frequency modulation approach, which can be implemented to provide frequency response, would not be effective in case of onshore voltage dips, where a rapid active power reduction is necessary, because the standard WT frequency controllers are designed to respond in the slow time scales relevant to large power systems. Further, rapid DC voltage changes may induce unrealistically high Rates Of Change Of Frequency (ROCOF) in the offshore grid, which might exceed the df/dt withstand capabilities of the WTs [25,30,31]. The voltage dip approach, on the other hand, enables fast WF active power reduction during onshore grid faults, however its applicability becomes questionable in case of onshore over-frequency events of a long duration, which would lead to sustained voltage dips in the offshore grid, possibly beyond the WT FRT capability limits. Employing independent FRT and frequency response controllers is not practical, when reliance on a communication link is not assumed, because it is difficult to distinguish the type of gridside disturbance that has initiated a DC voltage variation that is experienced by the offshore VSC, in order to activate the appropriate controller or use the correct set of control parameters. When relying on local measurements to achieve GC compatibility, a unified controller needs to be designed for the offshore VSC, which will successfully handle different types of onshore grid disturbances. Hence, the main objective of this paper is to propose alternative control schemes for VSC–HVDC connected WFs, which

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Fig. 1. Frequency dependent active power reduction [5].

achieve satisfactory response both for grid voltage and frequency disturbances, using the same controller structure and parameters, and are suitable for implementation in standard transient-stability type simulations. Three control strategies are identified and comparatively assessed for the offshore VSC, none of which relies on a dedicated communication link between the two HVDC terminals. The first strategy, used as the benchmark, follows directly from literature and employs offshore frequency modulation, in case of both voltage dips and frequency deviations at the onshore side. The second strategy involves a novel, unified voltage and frequency controller for the offshore VSC, via continuous voltage and frequency regulation of the offshore network. The third strategy is a new approach, employing offshore voltage modulation to achieve both FRT and frequency response. For this purpose, a voltage-dependent current modulation strategy is integrated in the standard WT current controllers, leading to offshore voltage variations of limited magnitude, which is important especially in case of prolonged over-frequency events. The paper is organized as follows: The main technical requirements imposed by present day GCs to offshore WFs are briefly discussed in Section 2. The study-case system is presented in Section 3. The modeling approach and the proposed control schemes for the offshore WF and the VSCs of the HVDC link are described in Sections 4 and 5 respectively. Time domain simulations are presented in Section 6 and the main conclusions are summarized in Section 7. Parameter values are provided in the Appendix. 2. Grid code requirements for offshore wind farms Technical requirements imposed to WFs, as well as other power stations, are common today for most GCs issued by TSOs [3–5]. Such requirements are now extending to offshore installations, e.g. as in the German GC, which includes a separate set of requirements pertaining to the grid connection of offshore WFs in seas [5]. These requirements stipulate the response of a WF, among other things, during abnormal onshore grid conditions, i.e. frequency or voltage deviations. In cases of frequency disturbances in the onshore grid, the interconnected WF should provide primary frequency response, which is typically perceived as a reduction in output power in case of overfrequency events (Fig. 1). According to [5], for instance, the WF is expected to remain connected in the frequency range 47.5–50.2 Hz and unrestrictedly feed the grid with its maximum available power for specified time intervals. In the range 50.2–52.7 Hz, the WF should reduce its active power output, as shown in Fig. 1, according to the characteristic: pWF = −Kgc fgrid

Fig. 2. Voltage-against-time profile for FRT capability [5].

of ±5% around the nominal value. Reactive current support up to 100% of the current rating is expected. Following a voltage dip, the WF must resume normal generation immediately after fault clearance, its active power increasing to the pre-disturbance value with a gradient between 10% and 20% of the rated power per second [5]. The aforementioned requirements are used as a reference in the rest of this paper, in order to assess the performance of the evaluated control strategies. 3. Study-case system The single-line diagram of the system under study is depicted in Fig. 4. The WF consists of Permanent Magnet Synchronous Generator (PMSG) based WTs, with a total installed capacity of 400 MW. The HVDC system comprises the offshore Sending-End Converter (SEC), the onshore Receiving-End Converter (REC) and a 100 km long submarine ±150 kV HVDC transmission line. To cater for the reactive power requirements imposed by the TSO, a power rating of 430 MVA has been assumed for the converters of the HVDC link. The onshore grid is represented by its Thevenin equivalent, with a short circuit capacity of 20,000 MVA. As an alternative, in order to test the frequency response of the offshore WF, the onshore grid is assumed to consist of 8 × 200 MW Synchronous Generators (SGs) and a 1500 MW load. The entire system is simulated in DIgSILENT PowerFactory. Specific parameter values are given in Table 1 of the Appendix [32,33]. For the SG case, the IEEE Type I automatic voltage regulator and excitation system model has been assumed and the IEEEG1 speed governor model. 4. Wind farm modeling and control Fig. 5 shows the typical configuration of a PMSG WT, with a full-scale frequency converter. Standard models from the literature are employed for the WT [10,34–37]. The aggregation method [11] is used to represent the entire WF, based on a 2 MW WT model [34,35]. Average-value converter models [38–40], combined with additional controller simplifications [41,42], are often employed in the relevant literature to represent converter dynamics, even when dealing with fast AC voltage transients, as they provide a good

(1)

where fgrid is the frequency deviation from a threshold value (e.g. 50.2 Hz). According to Fig. 2, for voltage dips within regions I and II of the diagram, the WF is expected to exhibit FRT capability and provide the necessary reactive power support to the system, as shown in Fig. 3, where a reactive current change relative to the pre-fault operation is expected when the voltage deviation exceeds a deadband

Fig. 3. Required reactive current for voltage support [5].

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Fig. 4. Study-case system. Onshore grid represented by (a) its Thevenin equivalent or (b) a synchronous generator.

compromise between accuracy, computational speed and ease of integration in transient stability analysis packages. In this paper, all high-frequency components related to the switching of power converters are neglected and the DC/AC converters are described by the fundamental frequency model of [38–40], while no important simplifications are made in the representation of the converter controllers. The Grid Side Converter (GRSC) of the WT controls the active and reactive power injected to the grid, as shown in Fig. 5. A threelayer control approach based on vector control [38,43] is adopted. The three phase grid voltages and output currents (vWT,abc , iWT,abc ) are transformed to the Synchronous Reference Frame (SRF) and vice versa using the Park transformation, where the grid voltage phase angle is measured by means of a SRF Phase Locked-Loop (PLL) [38]. The active power reference p∗WT is defined by the Maximum Power Point Tracking (MPPT) and frequency controllers (Fig. 5), whereas the reactive power reference q∗WT is set to zero. The MPPT and frequency response block, shown in detail in Fig. 6, first determines the maximum available power PMPPT , as a function of the measured rotor speed ωr and the MPPT curve [34,35], while the frequency controller superposes the necessary frequency-dependent term: p∗WT

= PMPPT − Kdr (foff − fn ) − Kin

dfoff dt

(2)

The next control level determines the current references id0,WT , iq0,WT of the GRSC as shown in Fig. 7. In this paper, an additional scaling is implemented for the reference currents: ∗ id,WT = mi id0,WT

(3a)

∗ iq,WT = mi iq0,WT

(3b)

where the scaling factor mi reduces the WT output current and therefore its output power whenever the offshore voltage drops below a predefined threshold, via a Proportional-Derivative (PD) controller: mi = 1 + Kvi voff + Kdi

dvoff dt

∗ ∗ Based on the current references id,WT , iq,WT , the modulation ∗ ∗ indices md , mq are then generated in the SRF, using the conventional current control scheme depicted in Fig. 7 [38,43]. The Generator Side Converter (GESC) in Fig. 5 regulates the DC link voltage and the AC voltage at the PMSG terminal to their nominal values, as in [34,36]. The same current control strategy as for the GRSC is adopted.

5. VSC–HVDC modeling and control 5.1. Receiving end converter As mentioned in Section 4, average-value models are employed for the representation of both HVDC converters (REC and SEC). The overall REC controller is depicted in Fig. 8. In normal operating conditions, the REC exports the incoming DC power to the onshore grid by regulating the HVDC voltage to its nominal value. During onshore voltage dips, the REC controller regulates the reactive current iq0,REC to the reference value of Fig. 3, implemented in the LVRT block of Fig. 8 by the deadzone-linear function: iq0,REC = Kr sign(vgrid ) max{|vgrid | − Vthres , 0}

Fig. 5. PMSG WT power converters and controller structure.

(4)

(5)

Fig. 6. Frequency response and MPPT controller of the WT GRSC (Fig. 5) [11].

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Fig. 7. Active and reactive power controller of the WT GRSC (Fig. 5).

The current magnitude limiter shown in Fig. 8 is activated during severe voltage dips, in which case the reactive current takes precedence over the active component. The active and reactive ∗ ∗ current references id,REC , iq,REC are given by: ∗ id,REC = sign(id0,REC ) min{|id0,REC |,



2

2 ∗ Imax − (iq,REC ) }

∗ iq,REC = sign(iq0,REC ) min{|iq0,REC |, Imax }

v∗dc,r =

Kvdc p∗WF,r Kgc

(8)

(6a) (6b)

The rate limiter shown in Fig. 8 ensures that, after fault clearance, the active power injected to the grid returns to the predisturbance level observing the gradient restrictions imposed by the TSO. The DC voltage controller regulates the HVDC voltage to its reference value v∗dc,REC . Following the approach of [19–24], onshore frequency deviations result in increased DC link voltages, by superimposing a signal v∗dc,REC on the default voltage reference (1 pu): v∗dc,REC = Kvdc fgrid

v∗dc,r , as in the case of onshore over-frequency events (see Fig. 8). Based on (1) and (7), v∗dc,r can be determined by:

(7)

where the gain Kvdc is chosen so as to maintain the DC voltage within acceptable limits (e.g. ±10%) even for severe onshore frequency deviations (see Table 2 of the Appendix). In normal onshore grid conditions (fgrid = 0), the DC voltage reference is set to 1 pu and therefore the REC active output power is regulated to the maximum available WF output power. Taking into account the ability of the REC controller to readjust the WF output power by modulating the DC voltage, a specified power reserve level p∗WF,r that might be imposed by the TSO, can also be maintained, by superimposing a DC voltage reference signal

The REC control scheme, denoted as Mode 1 in Fig. 8, does not treat regulation of the REC output power to a given set-point, that might be imposed by the TSO, which is a typical requirement stipulated by current GCs [4,5]. For this purpose, an alternative control concept, denoted as Mode 2, is also included in the REC controller of Fig. 8, which provides a closed-loop approach to directly control the REC active output power to a given reference p∗REC . In this case, a PI controller is utilized to determine the signal v∗dc,REC , while the Vdc,min and Vdc,max limits maintain the DC voltage variations within acceptable limits. In both cases, the resulting change in the DC voltage level establishes a virtual communication link to the SEC and then to the offshore WF, in order for the latter to regulate accordingly its output power.

5.2. Sending end converter In normal grid conditions, the SEC operates as a grid-forming power converter [15,38,43], regulating the voltage and frequency of the offshore network to their nominal values. During onshore grid events, the SEC detects the DC voltage deviations induced by the REC controller and modifies the offshore network voltage or frequency, to evoke the necessary response from the WF active power.

Fig. 8. REC controller (same for all evaluated control schemes).

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excessive ROCOF, possibly exceeding the df/dt withstand capabilities of state-of-the-art WTs [25,30,31]. These factors compromise the effectiveness of Strategy I when dealing with FRT incidents.

Fig. 9. SEC controller for Strategy I (offshore frequency regulation).

For this purpose, three alternative control strategies are described and implemented in the following. 5.2.1. Strategy I Control Strategy I, shown in Fig. 9, is largely based on concepts already introduced in the literature, either for frequency response [19–23] or for FRT [24–27]. In case of onshore over-frequency events or voltage dips, the REC will increase the DC voltage and the SEC controller will respond increasing the offshore network fre∗ , activating thus the frequency response capabilities quency by foff of the WTs. For frequency response purposes, an accurate estimate of the onshore frequency deviation is needed, realized via estimation of the onshore-side DC voltage vˆ dc,REC , using the following relation:

vˆ dc,REC = vdc,SEC −

pSEC rdc

vdc,SEC

(9)

where vdc,SEC and pSEC are local measurement of the offshore-side DC voltage and DC power. Then, the required offshore frequency ∗ is: deviation foff ∗ foff = Kf ˆvdc,REC

(10)

The active power reduction of each WT is determined by (2). Combining (2), (7) and (10), the eventual WF active power reduction, for a given onshore frequency deviation fgrid , will be: pWF = −Kdr Kf Kvdc fgrid

(11)

The controller gains Kdr and Kf must be chosen so that the power reduction meets the droop requirements of the GC. Based on (1) and (11), it must hold that Kgc = Kdr Kf Kvdc . As will be shown in Section 6, the controller could theoretically achieve FRT response using the same gain values, since a rapid DC voltage increase would trigger a rapid offshore frequency increase, to reduce the WF power flowing into the DC link. However, in real world applications, frequency measurement inevitably presents a lag [27], whereas the frequency response of WTs is rather slow, as it is normally tailored to the characteristics of large power systems. Further, a rapid frequency deviation could induce an

5.2.2. Strategy II A more reliable and fast manner to trigger an active power reduction of the offshore WF, in order to achieve FRT, is by inducing an offshore voltage dip [15,24,26–29]. Since all modern inverterinterfaced WTs have FRT capabilities, this course of action will effectively reduce their output power without leading to their disconnection. Hence, Strategy II, shown in Fig. 10, attempts to combine the principles of Strategy I with the application of controlled voltage dips. The deviation of the estimated onshore DC voltage is again used to activate the control of the SEC. In order to differentiate between frequency response and FRT incidents, a washout filter is added before the AC voltage controller and a ROCOF limiter is placed along the offshore frequency regulation path. When a fast power reduction is necessary to achieve FRT, a rapid offshore AC voltage reduction v∗off will be enforced, along with an offshore frequency deviation, which however will not exceed the specified ROCOF limits. On the other hand, during onshore over-frequency events, the offshore frequency modulation is primarily utilized to reduce the WF output power according to the GC requirements, whereas the voltage controller will be suspended by the washout filter, since the DC voltage variations will follow the relatively slow onshore frequency changes. Parameter values for this control scheme are provided in Table 3 of the Appendix. Conventional FRT controllers in WTs are equipped with a reactive current regulation block, whose purpose is to support the voltage of the grid during fault conditions (similar to the one shown in Fig. 3). Since Strategy II deliberately induces voltage dips in the offshore grid, the offshore voltage support operation of the WT should be preferably deactivated. Furthermore, in order to effect a large reduction of WT output power, when dealing with onshore faults, a severe offshore voltage dip will be needed, since the WT output power capability roughly tracks the terminal voltage. In order to avoid such severe dips, a coordinated control approach between the WTs and the HVDC converters is introduced, by integrating a voltage dependent current reduction strategy into the default WT current controllers. As already mentioned in Section 4 and illustrated in Fig. 7, the parameter mi in (4) implements an amplified reduction in the magnitude of the WT output current upon detecting a reduced terminal voltage, thereby leading to the fast decrease of the WT output power. Thus, effective curtailment of the WF active power becomes possible, without resorting to severe voltage dips, which increase stresses on the WTs and could potentially impact the operation of auxiliary equipment in the offshore network. The aforementioned control amendment is best suited to full-power converter WTs, as assumed in this work, which are

Fig. 10. SEC controller for Strategy II (combined voltage and frequency regulation).

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5.2.3. Strategy III The modification in the WT current controller implemented in the context of Strategy II, leading to effective active power curtailment with moderate voltage dips, could be utilized not only for FRT purposes, but also to achieve frequency response. In this way, the simple control scheme presented in Fig. 11 would suffice in both cases. The necessary offshore AC voltage reduction, v∗off , is given by:

Fig. 11. SEC controller for Strategy III (offshore voltage regulation).

inherently capable of effectively controlling their output current during transients of their terminal voltage. When dealing with Doubly-Fed Induction Generator (DFIG) WTs, the offshore voltage dip method could be augmented by control solutions proposed in the literature [15,28], in order to alleviate the transients induced by the voltage dips and the resulting stresses on the WT and HVDC converters.

v∗off = −Kv ˆvdc,REC

(12)

where the controller gain Kv is chosen so as to contain offshore AC voltage variations within acceptable limits (e.g. up to −15%) (see Table 3 of the Appendix). The offshore network frequency is maintained at the nominal value. If Mode 1 is selected for the REC controller (Fig. 8), where the REC output power is indirectly controlled, the WT controller gain Kvi in (4) must be chosen so as to fulfill the GC frequency response

Fig. 12. Response following a 30% step decrease of the load at t = 5 s, for the three alternative SEC control strategies. (a) Onshore grid frequency, (b) REC DC voltage, (c) SEC active output power, (d) offshore grid frequency, (e) offshore grid voltage, (f) WT output current magnitude.

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Fig. 13. Response of the three alternative SEC control strategies, following an onshore under-frequency event, while the WF operates with a power reserve level of 10%. (a) Onshore grid frequency, (b) REC DC voltage, (c) SEC active output power, (d) offshore grid frequency, (e) offshore grid voltage, (f) WT output current magnitude.

requirements. More specifically, for a given onshore frequency deviation fgrid , the WT active power deviation is expressed by: pWT = (voff 0 + voff )(mi id,WT 0 ) − voff 0 id,WT 0

(13)

Substituting (4), (7) and (12) into (13), the following expression is obtained (see Appendix II): pWT = −Kv Kvdc (PMPPT /voff 0 )(Kvi voff + 1)fgrid

(14)

Combining (1) and (14), the parameter Kvi should vary in operation according to: Kvi =

voff 0 Kgc − PMPPT Kv Kvdc voff PMPPT Kv Kvdc

(15)

According to (15), a coordinated control approach between the individual WTs and the HVDC converter controllers appears to be necessary, as the WT controller gain Kvi depends upon the SEC and REC controller gains Kv and Kvdc . This dependence can be effectively circumvented if Mode 2 is selected for the REC controller (Fig. 8), as the REC controller can then directly impose the active power

reference p∗REC necessary to achieve frequency response based on (1). At the offshore side, the SEC will regulate the WF output power accordingly and thus the WT controller gain Kvi setting need not be constrained. 6. Simulation results The system depicted in Fig. 4 is used as a study case. Specific parameter values adopted for the REC, SEC and WT controllers are presented in Tables 2–4 of the Appendix respectively, along with generic guidelines that can be followed to tune the parameters of the controllers. All time domain simulations are carried out in DIgSILENT Powerfactory. In all scenarios, the WF initially generates its rated power. 6.1. Frequency response capability In this section, the frequency response of the offshore WF is evaluated, for both onshore over- and under-frequency events.

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Fig. 14. Response following a 3-phase voltage dip at the 400 kV side of the REC transformer, for the three alternative SEC control strategies. (a) Onshore grid voltage, (b) REC reactive output current, (c) REC active output current, (d) SEC active output power, (e) REC DC voltage, (f) offshore grid frequency, (g) offshore grid voltage, (h) WT output current magnitude.

The onshore over-frequency transient is shown in Fig. 12(a). The REC controller, which is set to Mode 1, readjusts the DC voltage vdc,REC (Fig. 12(b)) at slightly increased levels (approximately 1.05 pu), corresponding to the deviation of the onshore frequency. The response of the SEC output power pSEC , shown in Fig. 12(c),

is identical, regardless of the control strategy employed for the SEC. In Strategies I and II the SEC imposes an offshore frequency fluctuation (Fig. 12(d)) that emulates the onshore grid conditions, triggering thus the frequency response of each WT. With

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Fig. 15. Effect of WT virtual inertia Kin when operating under SEC Strategy I. (a) REC DC voltage, (b) offshore grid frequency and (c) SEC active output power, following the 3-phase voltage dip of Fig. 14(a).

Strategy II, a voltage dip takes place, as observed in Fig. 12(e), which is soon eliminated due to the washout filter. The offshore frequency remains constant in Strategy III, whereas a voltage dip occurs at the offshore WF network to achieve the necessary active power reduction. The necessary depth of the voltage dip is only 10%, thanks to the WT control modification introduced in (4), which readjusts the WT output current accordingly (Fig. 12(f)). The system response in case of an onshore under-frequency event is shown in Fig. 13(a), while the WF is operated with a 10% power reserve, by increasing slightly the pre-disturbance DC link voltage (Fig. 13(b)). The onshore frequency drop leads the WF to release the active power reserve maintained under normal operation, as demonstrated in Fig. 13(c). As can be seen in Fig. 13(d) and (e), it is evident that operation at a specified power reserve requires only slight perturbations of the offshore frequency or voltage, depending on the SEC strategy employed. The pre-disturbance WT current (Fig. 13(f)) is higher in Strategy III, due to the reduced operating offshore voltage (0.96 pu), imposed by the SEC controller shown in Fig. 11. It

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Fig. 16. Effect of the derivative gain Kdi in the PD controller of each WT (Fig. 6) when operating under SEC Strategy III. (a) REC DC voltage, (b) offshore grid voltage and (c) SEC active output power, following the 3-phase voltage dip of Fig. 14(a).

is noteworthy that, in SEC Strategy III, the offshore voltage is regulated to acceptable levels throughout the under-frequency event, thus rendering this strategy suitable also for the provision of under-frequency response. 6.2. Fault ride-through capability In this section the FRT capability is assessed of the entire system comprising the HVDC link and the offshore WF. The network of Fig. 4(a) is used as a study case and the three-phase voltage dip of Fig. 14(a) at the 400 kV side of the REC transformer is simulated (voltage drop down to 15% for 200 ms, and gradual recovery within 500 ms). During the onshore voltage dip (0.5–0.7 s), the voltage support mode of the REC controller is activated, controlling its reactive output current iq,REC (Fig. 14(b)) to the reference imposed by the GC (−1.0 pu in this case), while the current magnitude limiter suppresses the REC active output current id,REC down to zero (Fig. 14(c)). After recovery of the voltage, the reactive output current returns to zero, whereas the active current is restored to its

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S.I. Nanou et al. / Electric Power Systems Research 121 (2015) 38–51

Fig. 17. Mode 2 operation of the REC controller. (a) REC active output power, (b) REC DC voltage, (c) offshore grid frequency, (d) offshore grid voltage, for step variations of the REC active power reference.

pre-disturbance value with a gradient of 0.2 pu/s, as imposed by the GC. In all three strategies, the SEC controller reduces the WF output power (Fig. 14(d)) while maintaining the DC voltage below 1.2 pu (Fig. 14(e)), achieving successful FRT performance. However, in Strategy I, the offshore frequency foff (Fig. 14(f)) exhibits ROCOF values of 20 Hz/s or more for approximately 100 ms, that exceed the typical df/dt withstand capability of state-of-the-art WTs (around 2–4 Hz/s [30,31]), a fact that could severely compromise the FRT capability of the offshore WF in real operating conditions. This essential drawback is effectively circumvented in Strategy II, where the ROCOF at the offshore network is constrained to 2 Hz/s by the rate limiter shown in Fig. 10. On the other hand, an offshore voltage dip of approximately 50% is enforced right after the occurrence of the onshore grid fault (Fig. 14(g)), to ensure a fast WF active power reduction. Due to the reduced DC voltage variations, the offshore voltage controller is soon suspended and the AC voltage dip in the offshore network lasts for a shorter time after the clearance of the fault. The frequency controller ensures then that the WF power tracks the REC active power, which is restored to its pre-disturbance value with a gradient of 0.2 pu/s, as imposed by the GC. Using Strategy III, the offshore frequency remains constant, while a more prolonged offshore voltage reduction is now required to achieve a successful FRT response, compared to Strategy II. It is observed that, even though the depth of the onshore voltage dip is 85%, the necessary depth at the offshore side is less than 40%, illustrating the positive effect of the proposed WT current reduction strategy. The response of the WT output current iWT in Strategy III, shown in Fig. 14(h), is fast enough to achieve the necessary active power reduction without a severe voltage dip. It is noteworthy that the DC voltage increase with Strategy II (Fig. 14(e)) is relatively lower, due to the derivative term in the washout filter in Fig. 10, which effects a steeper voltage dip. A similar effect can be achieved if the inertial response of the offshore WTs (control scheme in Fig. 6) is utilized in Strategy I, as well as in Strategy III if the derivative term of the PD controller in Fig. 7 is exploited. As it can be seen in Fig. 15(a) and (b), increased virtual inertia values in the WT frequency controller (gain Kin in

(2)) lead to substantially lower DC voltage and offshore frequency overshoots, as a steeper WT active power reduction can now be attained (illustrated in Fig. 15(c)). The duration of high ROCOF values in the offshore network also becomes shorter. Similarly, the effectiveness of Strategy III is further enhanced when higher Kdi gains are used in the PD controller of each WT (Eq. (4)), because the reduction of the WT output current is also determined by the rate of change of the offshore voltage. As can be seen in Fig. 16(a)–(c), it is evident that the required offshore voltage dip can be reduced even below 20% in order to achieve successful FRT response. 6.3. Active power control capability All previous simulations were performed using Mode 1 for the REC controller. In this section the response of the offshore WF is tested when the REC controller operates in Mode 2. As shown in Fig. 8, in this case the REC directly controls its output power to a given reference p∗REC , while the WF output power is regulated to the desired level, without any intervention to the SEC controller. In Fig. 17, the response of the WF is evaluated to consecutive step changes of the REC active power reference p∗REC . The REC output power successfully tracks its reference (Fig. 17(a)), whereas the DC voltage vdc,REC (Fig. 17(b)), the offshore frequency foff (Fig. 17(c)) and the offshore voltage voff (Fig. 17(d)) are properly adjusted in each strategy in order to control the WF active power. In all strategies, the offshore WF network frequency and voltage remain within reasonable operating limits, indicating that Mode 2 is a viable control alternative for the REC. 7. Conclusions In this paper, communication-less control schemes were developed for a VSC–HVDC connected WF with GC compatibility. Particular emphasis is placed on the SEC controller design, where three alternative control schemes were identified and proposed, in order for the WF to comply with typical FRT and frequency response requirements imposed by GCs. While all three alternative strategies

S.I. Nanou et al. / Electric Power Systems Research 121 (2015) 38–51

employ DC voltage variations to effectively transmit information on the onshore grid conditions to the offshore side, different SEC controller implementations rely on offshore frequency changes, voltage dips or both to effect the necessary WF power response. Fundamental operational constraints, such as the ROCOF and FRT withstand capability of the offshore WTs, as well as the maximum acceptable HVDC voltage limits, are properly accounted for. Time domain simulations were performed to assess GC compliance of the simulated WF, for all control schemes assessed. In Strategy I, which relies solely on offshore frequency modulation, the WF successfully provides frequency response, however its effectiveness in FRT incidents is questionable, because it requires high ROCOF values for non-negligible durations, which may trigger the WT protection, while it also presumes a very rapid response of the WT frequency control loop, which is also uncertain. Strategy II, on the other hand, relying on a combined application of frequency and voltage excursions, provides acceptable performance in terms of both FRT and frequency response, avoiding at the same time extreme offshore voltage and frequency excursions. With Strategy III, a new approach is introduced for the SEC controller, relying solely on offshore voltage regulation to achieve GC compatibility, along with a simple modification to the WT current controller, enabling prolonged active power reduction of the WF during onshore over-frequency events, while maintaining offshore voltage levels within acceptable limits. Comparing Strategies I to III, it can be concluded that Strategy II can meet GC requirements without a need to amend the controllers of commercially available WTs. However, the SEC controller structure becomes more complex in this strategy, while particular attention should be placed on the selection of the acceptable ROCOF limits, as well as of the cut-off frequency of the washout filter used to prevent prolonged offshore voltage dips. Furthermore, although the washout filter provides satisfactory performance for all examined disturbances, its optimized design warrants further attention to ensure that operating conditions and disturbances (e.g. unbalanced faults, possible electromechanical oscillations of higher frequency) that may excite higher frequency DC voltage oscillations, do not impair the effectiveness of SEC Strategy II. If a coordinated control approach can be established between the HVDC and WT controllers, Strategy III becomes also an attractive solution, since it can achieve GC compatibility with mild offshore voltage dips, while tuning of the controllers appears to be less challenging, as well. Acknowledgement The work of Sotirios Nanou is supported by IKY Fellowships of Excellence for Postgraduate Studies in Greece – Siemens Program. Appendix I. Study case system and control parameter values

49

Table A1 Study case system parameters. Onshore grid

Value

Nominal voltage Z , ∠Z ( * ) th th SG rating SG inertia constant SG droop constant Load demand

400 kV 8 , 80◦ 8 × 200 MW 6s 4% 1500 MW

HVDC cable

Value

Length Inductance Resistance

100 km 0.223 mH/km 0.027 /km

HVDC VSC

Value

Nominal power Nominal AC voltage Pole-to-pole DC voltage DC capacitance Phase reactor

430 MVA 150 kV 300 kV 400 ␮F 0.1 pu

*

For the grid topology shown in Fig. 4(a).

Table A2 REC controller parameters. REC parameter

Symbol

Value

Guidelines

Reactive current regulation gain during onshore voltage dips (Eq. (5)) DC voltage reference droop constant (Eq. (7), Fig. 8)

Kr

2

Explicitly imposed by the GC considered.

Kvdc

3.75

Vdcmax ,Vdcmin

0.1, −0.1

Calculated based on (7) in order to regulate the DC voltage levels below the maximum acceptable limit for normal operation (i.e. 1.1 pu), even for large onshore frequency deviations (up to 1.5 Hz). Selected to ensure that the DC voltage reference remains within acceptable limits for normal operation.

Maximum DC voltage reference deviations (Fig. 8)

Table A3 SEC controller parameters. SEC parameter

Symbol

Value

Guidelines

Offshore frequency reference droop constant (Eq. (10), Fig. 9) Offshore voltage reference droop constant (Strategy III, Eq. (12), Fig. 11)

Kf

0.28

Kv

2

Kw

50

Calculated based on (1) and (11) to meet the GC frequency response requirements. Calculated based on (12) in order to regulate the offshore voltage above the minimum acceptable limit in normal operation (0.85 pu), even for maximum DC voltage reference deviations (up to 1.1 pu). Selected to induce effective offshore voltage dips during fast DC voltage deviations (FRT events). Values in the range 40–60 provide acceptable performance. Higher values increase the under-shoot in the induced offshore voltage dip.

Tables A1–A4. Appendix II. Derivation of Eq. (15) (applicable to SEC Strategy III) For a given onshore frequency deviation fgrid , substituting (7) to (12), the induced offshore voltage dip voff is expressed by: voff = −Kv Kvdc fgrid

(16)

Combining (16) and (4), the steady-state value of the scaling factor mi becomes: mi = 1 − Kvi Kv Kvdc fgrid

(17)

Offshore voltage reference droop constant (Strategy II, Fig. 10)

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S.I. Nanou et al. / Electric Power Systems Research 121 (2015) 38–51

Table A3 (Continued)

Substituting (16) and (17) to (13), the final WT active power reduction is expressed by:

SEC parameter

Symbol

Value

Guidelines

Washout filter cutoff frequency (Strategy II, Fig. 10)

ωC

21.5 r/s

Selected to filter-out slow DC voltage variations during onshore frequency changes, thus imposing offshore voltage dips only during onshore grid faults. Values in the range 2␲0.5–2␲2.5 r/s provide satisfactory performance. Higher values increase the under-shoot in the induced offshore voltage dip. Selected to comply with df/dt withstand capabilities and protection settings of the WTs.

Maximum ROCOF (Strategy II, Fig. 10)

ROCOFmax

2 Hz/s

Table A4 WT GRSC controller parameters. WT parameter

Symbol

Value

Guidelines

Frequency droop control gain (Eq. (2), Fig. 6)

Kdr

20

Virtual inertia control gain (Eq. (2), Fig. 6)

Kin

Adjustable

Proportional gain of WT current magnitude controller (Strategy II, Eq. (4), Fig. 7)

Kvi

2

Proportional gain of WT current magnitude controller (Strategy III, Eq. (4), Fig. 7)

Kvi

Eq. (15)

Derivative gain of WT current magnitude controller (Eq. (4), Fig. 7)

Kdi

Adjustable

Reduces WT output power under SEC Strategy I. It is set equal to Kgc (explicitly imposed by the GC). Reduces the required offshore frequency excursion for FRT response under SEC Strategy I. Values in the range 0–20 provide satisfactory performance. Immunity to noisy measurements is achieved via proper selection of the low pass filter time constant (here T = 0.5 s). Higher values can lead to oscillatory behavior. Reduces the required offshore voltage dip for FRT response under SEC Strategy II. Calculated based on (4) in order to reduce the WT output current to zero for a 50% offshore voltage dip. A coordinated control approach is required between the WT and SEC controllers in SEC Strategy III, in order to meet GC frequency response requirements. Reduces the required offshore voltage dip for FRT response under SEC Strategies II and III. Values in the range 0–6 provide satisfactory performance. Immunity to harmonics or noisy measurements is achieved via proper selection of the low pass filter time constant (here T = 0.5 s).

pWT = −voff 0 id,WT 0 Kvi Kv Kvdc fgrid − id,WT 0 Kv Kvdc fgrid



+ Kvi id,WT 0 Kv Kvdc fgrid

2

(18)

Resubstituting (16) (Kv Kvdc fgrid = voff 0 − voff ) in (18) yields: pWT = −Kv Kvdc id,WT 0 (1 + Kvi voff )fgrid

(19)

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