An Integrated Transformer-Shunt Reactor Operation Device for The Connection of Off-Shore Wind-Farms

Abstract—An innovative design of a device embedding the functions of both power transformer and shunt reactor is presented. The device is thought for installation onboard substation platforms of long HV/EHV AC cable connected off-shore wind plants. The compact design saves much space on the hosting offshore platform, which can be sized smaller and cheaper. Machine sizing as well as cost saving considerations are exposed in this paper.

Keywords—Active and reactive power, electrical machine design, finite elements analysis, high voltage cable, offshore wind plant, shunt reactor, substation platform, transformer.

I. INTRODUCTION

V OR EHV AC submarine cable lines (SCLs) are an efficient and cost-effective solution for shore connection of offshore wind-farms (OWFs), in a broad

range of both transmission distances and generation capacities. Long HV/EHV AC SCLs generally require a significant amount of inductive shunt compensation, usually provided by line-connected shunt reactors (SRs). As a rule of thumb, for an OWF-to-shore connection usually a 90% of shunt compensation degree can be assumed, equally shared between the line terminals. This means that at least one three-phase shunt reactor has to be installed in the offshore substation site.

H

Present-day OWFs use a MV (usually 36kV) AC interarray cable network to collect power from individual turbines. The collected power is delivered to the offshore substation. The substation platform also hosts a MV/HV (or MV/EHV) step-up transformer (SUT) (analogous to a generator step-up unit), which adds to the SR unit. Since footprint and weight of typical HV SRs are comparable to those of transformers with the same MVA rating, the aggregate weight and footprint of offshore-located SR and SUT significantly boost up size and cost of the

substation platform, which in turn accounts for a non-negligible share of the overall capital cost of the project. With these premises, the integration of the two machines (SUT and SR units) in a single device looks promising in terms of weight and footprint savings, and finally of total cost savings, since smaller and cheaper substation platforms could be used. Moreover, SRs have to remain connected at all times, to prevent dangerous overvoltages in case of load rejection and to ensure correct operation of HV circuit breakers when switching off the unloaded cable line. So, embedding the SR's function in the substation transformer also yields a safer operation.

In the present paper, the authors explore the feasibility of a machine performing the twofold task of transformer operation (step-up voltage raising and active power delivery to the SCL) and shunt-reactor operation (reactive power absorption from the SCL). Section II of this paper describes the SR and SUT and their ratings. Section III introduces and discusses a possible combined transformer-shunt reactor (TSR) structure and its operating principle. Section IV shows the results of FEM calculations for an applicative example of TSR to a typical OWF connection scheme. Finally, Section V shows a preliminary cost analysis and a comparison between the proposed solution and the conventional approach using two distinct machines.

II. TSR CONCEPT AND WINDING POWER RATINGS

The single-line diagram of Fig. 1 shows the locations of the SUT and SR units in a typical wind-farm connection scheme. The SUT and SR obviously fulfill different tasks and have different ratings: the SUT should be able to match the aggregate OWF MV-side capability (also taking into account the interarray cable network and the MV shunt compensation, if any), whereas the SR is rated to compensate the SCL reactive surplus, at full active loading and in optimal operating conditions [1].

Wind turbines

100 MVA

33 kV 155 kV

Export cable

35 MVAR

Receivingsubstation

HV network

Fig. 1: Single-line diagram of an offshore windfarm with conventional step-up transformer and shunt reactor.

F. Palone, M. ReboliniEngineering Dept.

TERNA Rete Italia s.p.a.Rome, Italy

(email: francesco.palone@terna.it )

C. Bruzzese, F. Ielpo, S. LauriaDept. of Astronautical, Electrical, and Energy Eng.

University of Rome – La SapienzaRome, Italy

(email: claudio.bruzzese@uniroma1.it)

A. VitielloEngineering Dept.

Getra PowerMarcianise(CE), Italy

(email: angelo.vitiello@getra.it)

SUT

SR

SCL

Wind turbines

33 kV 155 kV

Export cable

35 MVARReceiving substation

Network

Fig.2 : Single-line diagram of an OWF using a TSR combined unit.

An alternative single-line diagram including the proposed TSR is shown in Fig. 2. In this case the reactive power surplus of the HV cable, connected to the primary winding of the TSR, is compensated by the TSR itself. The MV collection network is connected on the primary side of the TSR, which also acts as a step-up transformer.

It follows that the power ratings of primary and secondary windings are different: in fact the secondary (MV) winding of the TSR in Fig. 2 acts as the SUT's correspondent winding in Fig. 1, and must have the same rating, to accommodate the full OWF capability.

Given an OWF active power rating Pn,w and maximum reactive capability Qn,w=–0.3Pn,w, the secondary winding rating STSR,MV should be at least 1.05Pn,w.

S TSR ,MV=Pn , w2

Q n ,w2

≈1.05 Pn ,w (1)

On the other hand, in order to deal with the SCL reactive power surplus, the SR is located on the primary (HV) side of the SUT. It follows that TSR's primary winding, which also fulfills the role of the SR with reactive rating Qshunt, needs a higher rating than the secondary. The rating of the HV winding of the TSR, STSR,HV can be evaluated as follows:

S TSR , HV =Pn , w2

Q shunt2

(2)

As a first approximation, Qshunt can be assumed as the SCL reactive surplus Qc. The latter depends on the electrical characteristics of the SCL, as well as on its loading [1]. For an homogeneous SCL, assuming a maximum transmission power Pmax (max active load in optimal operating conditions) matching the OWF capacity Pn,w, Qc will be:

2,

2wnzc PSQ −= (3)

at each SCL terminal. Neglecting losses, the latter value can be estimated by [2]:

( )LKP

SPQ

c

zcPc ⋅′′−= tan

2

22

max, (4)

where Pc and Sz (MVA) are respectively the surge impedance loading (SIL) and the apparent power at ampacity of the SCL, both calculated for a given voltage U, Kʺ (km-1) is the cable propagation constant and L (km) is the cable length. More conservatively, one can consider the no-load reactive power

QcNL, again assumed equally shared between the line terminals and neglecting losses:

⋅′′=

2tan

LKPQ ccNL

(5)

The two quantities differ little for lossy homogeneous SCLs. Note that an asymmetrical shunt compensation could improve the maximum active power transmission for inhomogeneous lines such as those composed of submarine plus land cable stretches [3].

Assuming that the cable has the minimum required size, i.e. Pmax≅Pn,w, and also considering Pmax≅0.9Sz, it is Qc(Pmax)≅0.435Sz≅0.5Pn,w. Since the OWF is operated, at steady state, near the unity power factor condition at the MV substation busbars [4], it follows that the primary winding has only a slightly higher rating than the secondary, say, about 115%.

III. TSR OPERATING PRINCIPLE

A core-type construction is commonly used for both HV transformers and shunt reactors. The latter use gapped phase legs in order to increase the reluctance and the reactive power rating.

Fig. 3-a schematically shows the reluctive paths for the fluxes in a single-phase two-winding symmetrically-gapped core. Fig. 3-b shows the correspondent magnetic network.

Fig. 4 shows the equivalent electric circuit derived from Fig. 3-b by using the duality principle of electric and magnetic networks.

N2 N1 N1 N2R’0R’’0

R’/2

R’’

R’/2

A B

R’0

R’

R’’0 R’’

F1 F2

f1 f2

a) b)

Fig. 3. Simplified representation (a) and magnetic circuit (b) of a single-phase, two-winding transformer. R’0 is the air gap reluctance, R’/2 and R” are core reluctances. R”0 is the leakage reluctance of the magnetic path between the two concentric winding (insulation layer).

TSRintegrated unit

1/(R’+R’0) 1/R’’

1/R’’0 I2

K1=N1K2=1/N2

V1 V2

I1 R1R2

Fig. 4. Equivalent circuit derived from the magnetic network of Fig. 3-b, by applying the principle of duality of electric and magnetic networks.

The circuit of Fig. 4 shows that the air gap in the phase leg of the transformer allows increasing the no load current from typical values of modern large power transformers (0.1% of rated current using step-lap core construction) up to values typical for shunt reactors.

Multiple air gaps can be designed to reduce the fringing effects of the magnetic flux in the core. As a consequence, the structure of the TSR's core is more similar to that of a variable shunt reactor (VSR) rather than to a normal power transformer's core. VSRs (Fig. 5) are common components in HV and EHV networks and are generally equipped with two (main and regulating) windings. Moreover, VSRs have windings with different ratings, which is also true for the proposed TSR device, as shown in Section II. However, the ratio of rated powers of main and regulating windings of a typical VSR is generally lower than 1:2, whereas the ratio of power ratings of primary and secondary windings of the TSR considered in Section II is only about 115 %.

In the following a typical windfarm application is considered for TSR application; the aggregated rated power of the windfarm generators is 100MVA, the export cable voltage is 150kV and the cable length is 50km. The reactive power surplus of the export cable can be estimated as follows:

Qc=U r22 π f c L=74MVAR (6)

where Ur is the nominal voltage of the export cable (150kV), f is the network frequency (50Hz), c is the cable capacitance per unit of length and L is the cable line length. Considering a 90% of shunt compensation degree and a symmetrical compensation scheme, the rated no-load current of the TSR can be chosen as 35% on a 100MVA base. The other nameplate data for the example application of the TSR are reported in table I.

TABLE I. NAMEPLATE DATA OF THE TSR FOR THE EXAMPLE APPLICATION

Rated Voltages Ur1/Ur2 155/33kVRated power of

primary winding Sr1106MVA

No-load current I0 (primary side, 100MVA base)

35%Rated power of

secondary winding Sr2100MVA

IV. 3-D FEM TIME-STEPPING SIMULATIONS

In order to assess the proposed solution, extensive 3D FEM simulations have been carried out, both at steady state and transient regime. The nameplate data of Table I have been used for a preliminary sizing of a single phase unit, in order to simplify the FEM analyses. The final three-phase plant is obviously made up of three single-phase units.

Considering the axial and planar symmetries of the single-phase construction, only 1/8 of the TSR unit has been represented using Maxwell V14 (Fig. 6). The following material constraints have been considered for sizing the TSR:

• Maximum flux density Bm = 1,457T, in order to have a linear V-I characteristic up to 120% of rated voltage as prescribed by IEC 60076-6 [3].

• Windings current density Jw=3A/mm2, in order to ensure short circuit withstand capability.

• No-load losses P0=25kW.• I2R losses Pcc=84kW.

(a) (b)Fig 5. 95MVAR, 70%÷100% regulation range, 420/√3kV, single phase variable shunt reactor manufactured by Getra for TERNA [5]. (a): Reactor core before winding installation. (b): Completed active part.

Fig. 6 : TSR representation in ANSOFT

At first, steady state no-load and short-circuit simulations have been carried out. The no-load flux density map is reported in Fig. 7; the maximum value of Bm=1,457T is respected. Fig. 7 also shows that the stray flux in the windings during the no-load operation (i.e. shunt reactor operation) is almost nil, notwithstanding the presence of air gaps in the phase leg.

The short-circuit condition is simulated in Fig. 8. There is a significant flux also in the return leg (outermost leg). In fact, the TSR has been fed by the HV (external) winding, so that the return leg acts like a flux shunt for the leakage flux due to its very low reluctance. The central leg, instead, is shielded by the innermost winding (short circuited).

The behavior of the TSR has been also verified by simulating normal (rated) operating conditions, i.e. with nominal voltage; the results are shown in Fig. 9. Again, there is no significant stray flux on inner winding coming from the gapped core. The operation of the TSR's windings under normal operating conditions is similar to that of a normal transformers.

Transient analyses have also been carried out to verify the behavior of the TSR during switching maneuvers. In fact the energization of long, shunt compensated cable lines can lead to troublesome harmonic resonances and delayed zero crossing [6]. A time-stepping FEM technique, using Maxwell V14, has been used. The windfarm export cable has been represented with its exact-pi equivalent; the main data of the submarine cable line are reported in Table II.

Fig. 7. No load simulation. Magnetic flux density magnitude map.

Fig. 8. Short circuit simulation, TSR fed by HV (outermost) winding. Magnetic flux density magnitude map.

Fig. 9. Rated voltage and current operation simulation. Magnetic flux density magnitude map.

TABLE II. EXPORT CABLE LINE PARAMETERS

Resistance 61.25mΩ/km Reactance @ 50Hz 120mΩ/kmCapacitance 210nF/km Length 50km

The HV network has been represented by a lumped short circuit impedance, yielding a very low short circuit power of 250MVA, in order to represent a critical network condition. The most critical circuit breaker closing time (corresponding to zero voltage) has been at first studied. Results are shown in Figs. 10 and 11.

Fig. 10 shows that due to the unfavorable closing time the flux density reaches full saturation values (B>2T). The resulting primary current, shown in Fig. 11, exhibits the typical 2nd harmonic distortion due to the current inrush; the peak value however is less than 1kA, i.e. less than 2.7 p.u.. The expected electromagnetic forces on the primary winding do not pose mechanical problems.

In order to obtain the desired values of no-load current and short circuit impedance, a parametric study has been adopted, varying the phase column gaps length and the radial insulating distance between windings. Results are shown in Figs. 12 and 13. As shown, by slightly varying the physical size, the target values of I0 and zcc are easily obtained. The linear regression coefficient is close to 100%; small deviations around the regression lines are due to precision limits related to the used finite element mesh. It shall be noted that the mean square error is below 0,5%, one order of magnitude lower than the standard tolerances of 5% [7].

Fig. 10. TSR energization at zero voltage. Magnetic flux density magnitude map. The peak flux values are reported.

Fig. 11. Inrush current in the TSR, energized at a zero-voltage instant.

V. COST COMPARISON

Using the TSR allows for significant economic advantages as follows:

• Cost reduction for the machinery;• Cost reduction for the off-shore platform.

The first is due to the evident advantage of using a single machine instead of two, as also substantiated in the following. The second cost saving is also more important than the first, since the cost of the platform and marine foundations (planning, construction, materials, etc.) is much higher than that of the machines placed on them.

A first comparison can be performed between the classical solution (using both SR and SUT) and the proposed one (only one machine: the TSR) which has been considered in Section IV. For a given rated power SF (VA or VAR), the cost of step-up transformer and shunt reactor can be estimated as [8]:

Cost LF € =ALFBLF⋅S FC LF (7)

with ALF=-0.136*106, BLF=241.5, and CLF=0.447.The weight of the machines can be estimated as [8]:

Fig. 12. No load current plotted on air gap length (parametric analysis).

Fig. 13. Parametric analysis of short circuit impedance zcc, as function of the inner diameter of the HV winding.

W k g = 5.95 S F /1000 0.795

⋅B I L[ k V ]0.244 (8)

The offshore platform incremental cost, for increasing weight of the payload (machines), can be estimated as 20€/kg [8].

The total costs and weights for the classical (Fig. 1) and proposed (Fig. 2) solutions are reported in Table III. For evaluating weight and costs of the TSR, the HV winding rating, i.e. 106 MVA, (higher than the 100MVA LV one) has been chosen as SF.

The weight of the TSR has also been verified by considering the proposed design studied in section IV (Fig. 6), yielding about 67tons for each phase, thus confirming the preliminary estimation.

TABLE III. COST COMPARISON BETWEEN CLASSICAL AND PROPOSED SOLUTION

Classical solution (SR+SUT)

Proposed solution (TSR)

Rated power (MVAR) 35+100=135 106Weight (t) 86.5+193=279.5 202Cost (M€) 0.44+0.78=1.22 0.8

Platform cost (M€) 5.6 4.0Total cost (M€) 6.8 4.8

TIME(s)

CU

RR

EN

T(s

)

Table III shows that the total cost (electro-magnetic machines and platform) is 2M€ lower using TSR instead of the classical solution. The overall cost saving with respect to the classical solution using SR and SUT is around 30%. Only a small part (0.4M€) of this economic advantage is directly related to the cost saving due to the use of an integrated machines: the most significant benefit is the huge reduction of platform weight (78t) and of incremental cost (1.6M€).

CONCLUSION

The integrated transformer-shunt reactor (TSR) device is conceived as an effective solution for shore connection of offshore windfarms, from both the functional and economical point of view. The TSR combines a step-up transformer and a shunt reactor in the same unit, using the same, well established, technology available for variable shunt reactors.

Accurate 3D steady-state and time-stepping FEM simulations confirmed the feasibility of the proposed machine, showing that its performance is analogous to the overall performance of a normal transformer paired with a shunt reactor.

Cost estimations, based on current prices for offshore wind farm equipments, show significant economical advantages with respect to the classical solution, especially due to the large weight reduction on the offshore platform. The TSR can also be used for island connection with submarine HVAC cables, especially with small room or constrained installation spaces.

For these reasons TERNA s.p.a. is currently evaluating the possible use of TSR units in future evolutions of the Italian National Power Transmission Grid.

REFERENCES

[1] F. M. Gatta, A. Geri, S. Lauria, and M. Maccioni, "Steady-state operating conditions of very long EHVAC cable lines," Electric Power Systems Research, vol. 81 n. 7, pp. 1525-1533, July 2011.

[2] F. M. Gatta, A. Geri, S. Lauria, and M. Maccioni, "Steady-state operating conditions of very long EHVAC cable lines: Two case studies," Electric Power Systems Research, vol. 83 n. 1, pp. 160-169, Feb. 2012.

[3] S. Lauria, F. Palone, "Optimal operation of long inhomogeneous AC cable lines: The Malta–Sicily interconnector," accepted for publication on IEEE Trans. on Power Delivery, DOI: 10.1109/TPWRD.2013.2293054.

[4] H. Guo, K. Rudion, C. Heyde, and Z. A. Styczynski, “Stability studies of offshore wind farms” Fifth International CRIS conference on Critical Infrastructures, Bejing 2010.

[5] GETRA POWER, “GETRA nella rete elettrica di trasmissione e la sfida delle energie rinnovabili”, presented at Naples University ”Federico II”, Naples 29 nov. 2012, available online on www.getra.it/getdoc.php?ID=326.

[6] L. Colla, M. Rebolini, and F. Iliceto “400 kV AC new submarine cable links between Sicily and the Italian mainland - Outline of project and special electrical studies” CIGRE 2012 Session, paper C4-116, Paris (France) 24-29 Aug. 2008.

[7] International standard IEC 60076-6 “Power Transformers – Part 6: Reactors” edition 1.0, Dec. 2007

[8] G. Guidi and O. B. Fosso, “Investment cost of HVAC cable reactive power compensation offshore”, IEEE International Energy Conference and Exhibition (ENERGYCON), Florence (Italy), 9-12 Sep., 2012.