FACULTEIT INGENIEURSWETENSCHAPPEN

________________

KATHOLIEKE

UNIVERSITEIT

LEUVEN

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Eindwerk voorgedragen tot het behalen van

het diploma van Burgerlijk werktuigkundig-

elektrotechnisch ingenieur (Master in de

ingenieurswetenschappen: energie)

Bram Van Eeckhout

Promotor:

Prof. Dr. Ir. Ronnie Belmans

Dagelijkse begeleiding:

Ir. Dirk Van Hertem

2007 – 2008

ii

© Copyright K.U.Leuven Zonder voorafgaande schriftelijke toestemming van zowel de promotor(en) als de auteur(s) is overnemen, kopiëren, gebruiken of realiseren van deze uitgave of gedeelten ervan verboden. Voor aanvragen tot of informatie i.v.m. het overnemen en/of gebruik en/of realisatie van gedeelten uit deze publicatie, wend U tot de K.U.Leuven, Departement Elektrotechniek – ESAT, Kasteelpark Arenberg 10, B-3001 Heverlee (België). Telefoon +32-16-32 11 30 & Fax. +32-16-32 19 86 of via email: info@esat.kuleuven.be. Voorafgaande schriftelijke toestemming van de promotor(en) is eveneens vereist voor het aanwenden van de in dit afstudeerwerk beschreven (originele) methoden, producten, schakelingen en programma’s voor industrieel of commercieel nut en voor de inzending van deze publicatie ter deelname aan wetenschappelijke prijzen of wedstrijden. © Copyright by K.U.Leuven Without written permission of the promotor(s) and the author(s) it is forbidden to reproduce or adapt in any form or by any means any part of this publication. Requests for obtaining the right to reproduce or utilize parts of this publication should be addressed to K.U.Leuven, Departement Elektrotechniek – ESAT, Kasteelpark Arenberg 10, B-3001 Heverlee (Belgium). Tel. +32-16-32 11 30 & Fax. +32-16-32 19 86 or by email: info@esat.kuleuven.be. A written permission of the promotor(s) is also required to use the methods, products, schematics and programs described in this work for industrial or commercial use, and for submitting this publication in scientific contests.

iii

PREFACE

Finishing a master thesis brings feelings of relief. The months of work to achieve the end result are

behind. It is also a milestone in a persons life and new horizons arise in the form of the start of a career as

a young engineer. It is therefore an ideal moment to thank the people who made all of this possible.

First of all, I would like to thank my thesis promotor Prof. Dr. Ir. R. Belmans. It is a great honor to realize

a master thesis together with a Professor with such a knowledge, experience and inspiring thoughts. I

need to thank him especially for giving me the opportunity to directly cooperate with ABB. This

cooperation added a connection between this thesis and the industrial reality, which I personally

experienced as extremely valuable.

Directly connected with this, I am very thankful to the people at ABB Corporate Research Center in

Västerås, Sweden, for the given opportunities. I especially want to thank my supervisors Dr. Muhamad

Reza and Dr. Kailash Srivastava at CRC. The three months I worked in Västerås helped me to better

understand the subject and added an extra dimension, the one of industrial experience, to this thesis work.

Many thanks go as well to Ir. Dirk Van Hertem, my daily supervisor at KU Leuven. The discussions that

we held on the subject, either early or late for one of us, or either face-to-face or online, brought me each

time new ideas and insights, and the will and courage to persist.

Last but not least, I want to express my gratitude to the people who always supported me from the

sideline, not only during the work for this master thesis, but during my whole study period at KU Leuven.

First of all my parents, thank you for letting me become the engineer that I always wanted to become.

Thanks as well to my sister Fien for showing me the way to Leuven. And then, there are many precious

friends to thank, especially Julie, for always being there for me. I really hope we all stay in touch,

whatever ways our future careers might bring us, and hence I hope the bonds we have now yield a

‘friendship for life’.

iv

ABSTRACT

This thesis investigates the use of Voltage Source Converter HVDC for the connection of offshore wind

farms with the onshore transmission grid and makes a comparison with HVAC on a techno-economic

basis. A typical future wind farm is proposed, rated at 300 MW and situated 50 km from the Point of

Common Coupling to compare both technologies. The main criteria for the technical comparison are the

losses, the realization of possible grid code requirements and the black start capability needed to start up

the offshore wind farm. The black start capability and variable frequency operation of VSC HVDC make

soft startup of the wind farm possible and give opportunities for wind farm topology optimization and

simplification. This allows for the use of directly connected (without converter) induction generators.

Variable speed operation is achieved by varying the frequency in the wind farm grid. The economic

comparison takes into account the investment costs, losses and annual maintenance. The main wind

turbine topologies are incorporated in the economic analysis. The sensitivity of the results obtained in the

economic comparison is investigated to variation of input parameters (cable length, converter loss,

offshore converter volume, cost of energy). This resulted in an estimation of cable break-even lengths for

the comparison of VSC HVDC with HVAC topologies.

v

SUMMARY

The planned offshore wind farms of today have reached power ratings of several hundred MW and are

situated tens of kilometers from shore. These power ratings and distances require a separate transmission

system at an elevated voltage compared to the wind farm collection grid voltage. The traditionally used

transmission technology between an offshore wind farm and the onshore grid is High Voltage Alternating

Current (HVAC). Voltage Source Converter (VSC) HVDC is proposed by two companies (ABB and

Siemens) as an alternative solution. This thesis compares the use of VSC HVDC and HVAC for the

connection of offshore wind farms with the onshore transmission grid.

VSC HVDC has several technical advantages over HVAC. The important features in this context are the

control of active and reactive power, the possibility to connect grids with asynchronous voltages and the

black start capability. Grids with differing frequencies can be connected using VSC HVDC. Several wind

turbine topologies are still under consideration for offshore use, others are reconsidered in combination

with VSC HVDC. The characteristics of VSC HVDC make simplification or economic optimization of

the offshore wind farm possible for each wind turbine topology.

A typical future wind farm is used in the comparison, rated at 300 MW and situated 50 km from the Point

of Common Coupling (PCC). VSC HVDC is first compared with HVAC on a technical basis. The main

criteria for the comparison are the losses of both systems, the realization of the grid code requirements

and the black start capability needed to start up the offshore wind farm. The losses are higher for VSC

HVDC (4,45% of Annual Produced Energy (APE)) than for HVAC (3,31% of APE) for this cable length.

The analysis for other cable lengths shows a break-even distance for the losses of 80 km. The probability

distribution of wind speeds is taken into account in the loss calculation.

Grid code requirements on power factor control, frequency response and voltage dip ride-through are

discussed in this thesis. No additional equipment is needed with VSC HVDC to achieve power factor

control. The HVAC option needs additional reactive compensators, which results in an elevated

transmission system cost. The use of VSC HVDC results in improved frequency response and voltage dip

ride-through capability for an offshore wind farm with any topology. The realization of these grid code

requirements for a wind farm connected to the shore with HVAC depends on the used wind turbine

topologies and is not enhanced by the transmission link.

The black start capability of VSC HVDC makes soft startup of the wind farm possible and gives

opportunities for energy output optimization. It is possible to operate the wind farm grid at variable

frequency during low wind speed periods, hereby implementing variable-speed operation of the wind

turbines in a cost-efficient way. The use of HVAC results in a fixed frequency operation of the wind

farm. Variable-speed operation is then only achieved after the installation of a power electronic converter

vi

in each wind turbine.

VSC HVDC is compared with HVAC on an economic basis as well. Only cost data from public domain

are used in the economic comparison. The economic comparison is based on a Discounted Cash Flow

(DCF) calculation taking into account the difference in investment costs and annual costs and benefits.

Two economic comparisons are performed.

In the first case, the investor of the transmission system is assumed to be another party than the investor

of the wind farm. This situation is for example applicable in Germany, where the transmission system

operator is responsible for the connection of offshore wind farms. The possible advantages of VSC

HVDC resulting in a more optimal wind farm are irrelevant for the transmission system investor. VSC

HVDC and HVAC are therefore evaluated only on their electricity transmission function. The investment

costs for both systems are compared, together with the annual monetary costs of the losses and the

maintenance of the transmission system. VSC HVDC is fount not economically feasible for a 300 MW

wind farm situated 50 km from the PCC.

In the second case, the wind farm investor and the transmission system investor are the same party and a

total system economic optimum is looked for. The wind farm topology is therefore optimized according

to the technical advantages of VSC HVDC. The use of directly connected induction generators in a

variable frequency wind farm, governed by the VSC HVDC link, is chosen in this thesis. This system is

compared to the main variable-speed topologies (DFIG, DDPMSG and GPMSG) in combination with

HVAC. The solution with VSC HVDC is found to be economically feasible compared to HVAC/DFIG

and HVAC/DDPMSG. HVAC/GPMSG is more cost-efficient than the VSC HVDC solution.

The sensitivity of the results obtained during the economic comparison is investigated for variation of

input parameters (cable length, cost of energy, converter loss, offshore converter volume). This resulted

in an estimation of break-even lengths between VSC HVDC and HVAC cables. In the first case, with

different investors, the break-even distance is 80 km. In the second case, the break-even distance depends

on the used topology with HVAC: 32 km for DFIG, <25 km for DDPMSG and 68 km for GPMSG. For

cable lengths longer than the break-even distance, the VSC HVDC solution becomes more economically

feasible, whereas HVAC is more feasible for shorter cable lengths.

The cost of energy is an uncertain parameter in the economic analysis. It is used to monetize the losses in

the transmission system and the differences in annual energy output for the different wind farm

topologies. The result is therefore recalculated for different values (40, 80, 120 and 160 €/MWh).

VSC HVDC is still a young technology compared to HVAC. Technological progress can be expected on

the fields of converter losses and offshore substation volume. This has a positive effect for VSC HVDC in

the economic comparison with HVAC.

vii

TABLE OF CONTENTS

PREFACE.................................................................................................................................................... I

ABSTRACT .............................................................................................................................................. IV

SUMMARY.................................................................................................................................................V

TABLE OF CONTENTS........................................................................................................................VII

LIST OF FIGURES ...................................................................................................................................X

LIST OF TABLES ..................................................................................................................................XII

LIST OF SYMBOLS AND ABBREVIATIONS.................................................................................XIII

1 INTRODUCTION ...............................................................................................................................1

1.1 PURPOSE & SCOPE...........................................................................................................................1 1.2 STRUCTURE .....................................................................................................................................2

2 TECHNICAL INTRODUCTION TO VSC HVDC..........................................................................3

2.1 INTRODUCTION ...............................................................................................................................3 2.2 COMPARISON OF VSC HVDC WITH HVAC AND LCC HVDC.......................................................3

2.2.1 Comparison between HVAC and HVDC cable systems .........................................................3 2.2.2 Comparison between LCC HVDC and VSC HVDC ..............................................................5

2.3 OPERATIONAL ASPECTS OF VSC HVDC.........................................................................................6 2.3.1 Control of AC voltage amplitude and frequency.....................................................................7 2.3.2 Control of active and reactive power.......................................................................................8

2.4 TECHNICAL ADVANTAGES ............................................................................................................10 2.4.1 Power flow control ................................................................................................................10 2.4.2 Grid voltage support and power system stability ..................................................................11 2.4.3 Black start capability .............................................................................................................11 2.4.4 Connection of asynchronous grids.........................................................................................13 2.4.5 Voltage dip ride-through and power quality..........................................................................14 2.4.6 Cable choice ..........................................................................................................................14

2.5 CONCLUSION .................................................................................................................................14

3 OFFSHORE WIND TURBINE TOPOLOGIES AND THEIR USE WITH VSC HVDC..........15

3.1 INTRODUCTION .............................................................................................................................15 3.2 OFFSHORE WIND TRENDS ..............................................................................................................15 3.3 WIND TURBINE TOPOLOGIES .........................................................................................................19

3.3.1 Fixed-speed wind turbines.....................................................................................................19 3.3.2 Variable-speed wind turbines ................................................................................................21

3.4 COMBINATION OF TRADITIONAL WIND TURBINE TOPOLOGIES WITH VSC HVDC .......................27 3.4.1 Squirrel Cage Induction Generators ......................................................................................27 3.4.2 Doubly-Fed Induction Generators .........................................................................................30 3.4.3 Direct-Drive Synchronous Generators ..................................................................................31

3.5 CONCLUSION .................................................................................................................................32

4 ENERGY OUTPUT OF OFFSHORE WIND TURBINES ...........................................................33

4.1 INTRODUCTION .............................................................................................................................33 4.2 POWER PROBABILITY DENSITY FUNCTION ....................................................................................33

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4.3 ENERGY OUTPUT OF WIND TURBINES............................................................................................35 4.4 MULTI TURBINE FREQUENCY APPROACH ......................................................................................36 4.5 DRIVE-TRAIN EFFICIENCY .............................................................................................................39 4.6 CONCLUSION .................................................................................................................................40

5 DESIGN AND OPERATIONAL ASPECTS OF VSC HVDC AND HVAC................................41

5.1 INTRODUCTION .............................................................................................................................41 5.2 DIMENSIONING OF THE TRANSMISSION SYSTEM FOR AN OFFSHORE WIND FARM .........................41

5.2.1 VSC HVDC...........................................................................................................................42 5.2.2 HVAC....................................................................................................................................43

5.3 BLACK START OF A WIND FARM ....................................................................................................46 5.3.1 VSC HVDC...........................................................................................................................46 5.3.2 HVAC....................................................................................................................................47

5.4 REACTIVE POWER REQUIREMENTS ON THE OFFSHORE NODE .......................................................48 5.4.1 Directly connected induction generators ...............................................................................48 5.4.2 Generators connected via a converter....................................................................................49

5.5 GRID CODE COMPLIANCE AT THE POINT OF COMMON COUPLING ...............................................50 5.5.1 Reactive power or power factor control ................................................................................50 5.5.2 Frequency response ...............................................................................................................54 5.5.3 Voltage-dip ride through capability.......................................................................................55

5.6 LOSSES ..........................................................................................................................................58 5.6.1 VSC HVDC...........................................................................................................................58 5.6.2 HVAC....................................................................................................................................60 5.6.3 Influence of length of transmission cable..............................................................................66

5.7 CONCLUSION .................................................................................................................................68

6 ECONOMIC COMPARISON OF VSC HVDC AND HVAC FOR THE CONNECTION OF

OFFSHORE WIND FARMS............................................................................................................69

6.1 INTRODUCTION .............................................................................................................................69 6.2 TRANSMISSION SYSTEM INVESTMENT COST .................................................................................70

6.2.1 VSC HVDC...........................................................................................................................70 6.2.2 HVAC....................................................................................................................................72 6.2.3 Comparison of investment cost .............................................................................................74

6.3 ANNUAL COSTS .............................................................................................................................75 6.3.1 Losses in the transmission system to shore ...........................................................................75 6.3.2 Maintenance costs of the transmission system ......................................................................75

6.4 DISCOUNTED CASH FLOW ANALYSIS – SCENARIO 1 ....................................................................76 6.5 WIND FARM INVESTMENT COSTS ..................................................................................................76

6.5.1 Generator cost........................................................................................................................77 6.5.2 Gearbox cost ..........................................................................................................................77 6.5.3 Converter cost........................................................................................................................77

6.6 ANNUAL COSTS AND REVENUES OF AN OFFSHORE WIND FARM....................................................78 6.6.1 Energy output of wind farm...................................................................................................78 6.6.2 Annual maintenance cost wind farm .....................................................................................78

6.7 DISCOUNTED CASH FLOW ANALYSIS – SCENARIO 2 ....................................................................79 6.7.1 VSC HVDC with SCIG versus HVAC with DFIG ...............................................................80 6.7.2 VSC HVDC with SCIG versus HVAC with DDPMSG........................................................80 6.7.3 VSC HVDC with SCIG versus HVAC with GPMSG...........................................................81 6.7.4 Discussion..............................................................................................................................81

6.8 RELEVANCE OF THE RESULTS .......................................................................................................82 6.9 CONCLUSION .................................................................................................................................82

7 SENSITIVITY ANALYSIS ..............................................................................................................83

7.1 INTRODUCTION .............................................................................................................................83

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7.2 DISTANCE BETWEEN WIND FARM AND PCC .................................................................................83 7.2.1 VSC HVDC compared to HVAC – Scenario 1 .....................................................................83 7.2.2 VSC HVDC with SCIG compared to HVAC with other topologies – Scenario 2 ................84

7.3 COST OF ENERGY...........................................................................................................................86 7.3.1 VSC HVDC compared to HVAC – Scenario 1 .....................................................................86 7.3.2 VSC HVDC with SCIG compared to HVAC with other topologies – Scenario 2 ................87

7.4 CONVERTER STATION LOSSES OF VSC HVDC .............................................................................89 7.5 CONVERTER VOLUME....................................................................................................................89 7.6 CONCLUSION .................................................................................................................................90

8 CONCLUSIONS................................................................................................................................91

9 REFERENCES ..................................................................................................................................93

10 APPENDICES..................................................................................................................................101

10.1 APPENDIX A – MATLAB MODEL ..............................................................................................101 10.2 APPENDIX B – ROTATIONAL SPEED RANGE.............................................................................102 10.3 APPENDIX C – CAPACITY FACTORS.........................................................................................103 10.4 APPENDIX D – HVAC OFFSHORE SUBSTATION .......................................................................104

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LIST OF FIGURES

FIGURE 2.1 COMPARISON OF COSTS FOR AC AND DC TRANSMISSION CABLES....................4 FIGURE 2.2 COMPONENTS OF VSC HVDC TERMINAL....................................................................6 FIGURE 2.3 SCHEME OF 2-LEVEL 3-PHASE VOLTAGE SOURCE CONVERTER BRIDGE ..........7 FIGURE 2.4 2-LEVEL CONVERTER SINGLE PHASE OUTPUT VOLTAGE......................................7 FIGURE 2.5 SCHEME OF 3-LEVEL 3-PHASE VOLTAGE SOURCE CONVERTER BRIDGE ..........8 FIGURE 2.6 3-LEVEL CONVERTER SINGLE PHASE OUTPUT VOLTAGE......................................8 FIGURE 2.7 POWER CONTROL OF VSC WITH PHASE REACTOR ..................................................9 FIGURE 2.8 P,Q-CAPABILITY CHART OF VSC HVDC LINK ..........................................................10 FIGURE 3.1 OFFSHORE WIND FARM DEVELOPMENT IN EUROPE .............................................16 FIGURE 3.2 FIXED-SPEED INDUCTION GENERATOR WIND TURBINE TOPOLOGY................19 FIGURE 3.3 COEFFICIENT OF PERFORMANCE AS FUNCTION OF � WITH � AS A

PARAMETER......................................................................................................................22 FIGURE 3.4 POWER OUTPUT OF FIXED-SPEED AND VARIABLE-SPEED WIND TURBINE AS

A FUNCTION OF WIND SPEED.......................................................................................23 FIGURE 3.5 DOUBLY-FED INDUCTION GENERATOR WIND TURBINE TOPOLOGY ...............24 FIGURE 3.6 DIRECT-DRIVE SYNCHRONOUS GENERATOR WIND TURBINE TOPOLOGY .....25 FIGURE 3.7 PROBABILITY DENSITY FUNCTION OF WIND SPEED OFFSET WITH DISTANCE

FROM MASTER TURBINE D AS A PARAMETER........................................................29 FIGURE 4.1 SITUATION OF EURO PLAT FORM MEASUREMENTS IN THE NORTH SEA.........34 FIGURE 4.2 PROBABILITY DENSITY FUNCTION AND CUMULATIVE DENSITY FUNCTION

FOR WIND SPEEDS IN OFFSHORE WIND FARM.......................................................34 FIGURE 4.3 POWER PROBABILITY DENSITY FUNCTION FOR OFFSHORE WIND FARM.......35 FIGURE 4.4 OPTIMAL ROTATIONAL SPEED OF 5MW WIND TURBINE AS FUNCTION OF

WIND SPEED......................................................................................................................37 FIGURE 4.5 POWER NOT TAKEN FROM WIND DUE TO MULTI TURBINE SPEED CONTROL

..............................................................................................................................................38 FIGURE 4.6 EXPECTED VALUE OF POWER NOT TAKEN FROM THE WIND AT EACH WIND

SPEED IN THE VARIABLE SPEED ZONE DUE TO MULTI TURBINE SPEED CONTROL...........................................................................................................................38

FIGURE 5.1 XLPE 3-CORE CABLE AND SINGLE PHASE CABLE AND REINFORCED SUBMARINE XLPE 3-CORE CABLE ..............................................................................43

FIGURE 5.2 MAXIMUM TRANSMISSION CAPACITY PER 3-CORE CABLE AT 150 KV ............45 FIGURE 5.3 EQUIVALENT SECTION OF HVAC CABLE..................................................................47 FIGURE 5.4 CAPABILITY CHART OF OFFSHORE VSC (RECTIFIER) ...........................................49 FIGURE 5.5 POWER FACTOR REQUIREMENT FOR WIND FARM IN UK.....................................51 FIGURE 5.6 CAPABILITY CHART OF ONSHORE VSC (INVERTER) .............................................52 FIGURE 5.7 HVAC SYSTEM OVERVIEW WITH COMPENSATION AT BOTH CABLE ENDS ....54 FIGURE 5.8 VOLTAGE DIP RIDE-THROUGH CHART FOR NGC, E.ON-NETZ AND SVENSKA

KRAFTNÄT ........................................................................................................................55 FIGURE 5.9 WIND FARM FREQUENCY AND ROTATIONAL SPEED RESPONSE ON FAULT...57 FIGURE 5.10 TOTAL TRANSMISSION SYSTEM LOSSES FOR CROSS SOUND CABLE PROJECT

..............................................................................................................................................59 FIGURE 5.11 VSC HVDC LOSSES FOR 300 MW OFFSHORE WIND FARM ....................................59 FIGURE 5.12 REACTIVE CURRENT DISTRIBUTION IN THE CHOSEN 3-CORE HVAC CABLE .61 FIGURE 5.13 ESTIMATED CURRENT DISTRIBUTION IN HVAC CABLE 1 AND 2 .......................61 FIGURE 5.14 CONDUCTOR TEMPERATURE AS FUNCTION OF CURRENT THROUGH CABLE

..............................................................................................................................................62 FIGURE 5.15 CABLE RESISTANCE AS A FUNCTION OF CURRENT THROUGH CABLE............63 FIGURE 5.16 HVAC TRANSMISSION SYSTEM LOSSES....................................................................65 FIGURE 5.17 LOSS PERCENTAGE FOR VSC HVDC AND HVAC AS FUNCTION OF CABLE

xi

LENGTH..............................................................................................................................66 FIGURE 6.1 VSC HVDC M5 SUBSTATION: ONSHORE AND OFFSHORE STATION ...................71 FIGURE 6.2 RELATIVE EVOLUTION OF COPPER PRICE 2002 – 2008...........................................72 FIGURE 6.3 INVESTMENT COST COMPARISON: VSC HVDC VERSUS HVAC ...........................75 FIGURE 6.4 COST BREAKDOWN FOR 5MW DFIG OFFSHORE WIND TURBINE........................77 FIGURE 6.5 ANNUAL MAINTENANCE COST FOR 5 MW DFIG WIND TURBINE .......................79 FIGURE 7.1 TOTAL INVESTMENT COSTS AND DCF: VSC HVDC VERSUS HVAC....................84 FIGURE 7.2 FINANCIAL COMPARISON: VSC HVDC WITH SCIG VERSUS HVAC WITH DFIG

..............................................................................................................................................84 FIGURE 7.3 FINANCIAL COMPARISON: VSC HVDC WITH SCIG VERSUS HVAC WITH

DDPMSG .............................................................................................................................85 FIGURE 7.4 FINANCIAL COMPARISON: VSC HVDC WITH SCIG VERSUS HVAC WITH

GPMSG................................................................................................................................85 FIGURE 7.5 SENSITIVITY OF DCF RESULT TO VARIATION OF COST OF ENERGY: VSC

HVDC VERSUS HVAC......................................................................................................86 FIGURE 7.6 SENSITIVITY OF DCF RESULT TO VARIATION OF COST OF ENERGY: VSC

HVDC WITH SCIG VERSUS HVAC WITH DFIG...........................................................87 FIGURE 7.7 SENSITIVITY OF DCF RESULT TO VARIATION OF COST OF ENERGY: VSC

HVDC WITH SCIG VERSUS HVAC WITH DDPMSG ...................................................87 FIGURE 7.8 SENSITIVITY OF DCF RESULT TO VARIATION OF COST OF ENERGY: VSC

HVDC WITH SCIG VERSUS HVAC WITH GPMSG ......................................................88 FIGURE 7.9 SUM OF DRIVE TRAIN AND TRANSMISSION SYSTEM LOSSES FOR COMPARED

TOPOLOGIES .....................................................................................................................88 FIGURE 7.10 SENSITIVITY OF DCF RESULT TO REDUCTION OF CONVERTER LOSSES..........89 FIGURE 7.11 SENSITIVITY OF DCF RESULT TO REDUCTION OF OFFSHORE CONVERTER

VOLUME.............................................................................................................................90 FIGURE 10.1 DETAIL OF POWER-SPEED CURVES FOR DIFFERENT WIND TURBINE

TOPOLOGIES ...................................................................................................................102 FIGURE 10.2 PROBABILITY DENSITY FUNCTION FOR DIFFERENT WIND FARM LOCATIONS

............................................................................................................................................103 FIGURE 10.3 HORNS REV TRANSFORMER STATION ....................................................................104 FIGURE 10.4 INSTALLATION OF OFFSHORE TRANSFORMER STATION...................................104

xii

LIST OF TABLES

TABLE 2-1 COMPARISON HVAC, LCC HVDC AND VSC HVDC FOR CABLE TRANSMISSION................................................................................................................................................6

TABLE 3-1 PLANNED AND CONSIDERED OFFSHORE WIND FARMS IN NORTHERN EUROPE ..............................................................................................................................18

TABLE 3-2 CHARACTERISTICS OF STUDIED WIND TURBINE TOPOLOGIES..........................32 TABLE 4-1 ENERGY OUTPUT COMPARISON BETWEEN FIXED- AND VARIABLE-SPEED

WIND TURBINES ..............................................................................................................36 TABLE 4-2 EFFICIENCIES OF DIFFERENT COMPONENTS OF DRIVE TRAIN...........................39 TABLE 4-3 ANNUAL ENERGY OUTPUT AND CAPACITY FACTOR FOR DIFFERENT

VARIABLE SPEED WIND TURBINE TOPOLOGIES.....................................................39 TABLE 5-1 HVDC LIGHT® MATRIX..................................................................................................42 TABLE 5-2 LOSS COMPARISON BETWEEN HVAC AND VSC HVDC AS TRANSMISSION

OPTION FOR A 300 MW OFFSHORE WIND FARM......................................................67 TABLE 6-1 DISCOUNTED CASH FLOW VSC HVDC VERSUS HVAC...........................................76 TABLE 6-2 INVESTMENT COSTS FOR DIFFERENT WIND FARM TOPOLOGIES......................78 TABLE 6-3 DISCOUNTED CASH FLOW VSC HVDC/SCIG VERSUS HVAC/DFIG ......................80 TABLE 6-4 DISCOUNTED CASH FLOW VSC HVDC/SCIG VERSUS HVAC/DDPMSG...............80 TABLE 6-5 DISCOUNTED CASH FLOW VSC HVDC/SCIG VERSUS HVAC/GPMSG..................81 TABLE 10-1 WIND SPEED DATA........................................................................................................103 TABLE 10-2 ENERGY OUTPUT FOR DIFFERENT LOCATIONS ....................................................103

xiii

LIST OF SYMBOLS AND ABBREVIATIONS

Symbol Explanation Standard unit

A Scale parameter Weibull distribution [m/s]

C Shape parameter Weibull distribution [-]

Ccable Cable capacitance [F/m]

Cp Coefficient of performance [-]

dc Conductor diameter [m]

D Wind farm dimensional parameter [m]

fgrid Grid frequency [Hz]

fswitch Converter switching frequency [Hz]

I Current [A]

Iactive Active cable current [A]

IC Charging current [Ar]

Icable Cable current [A]

Ireactive Reactive cable current [Ar]

Iturb Turbulence intensity [%]

J Turbine blades inertia [kg.m2]

l% Transmission system loss percentage [%]

L Cable length [m]

Lcable Cable inductance [H/m]

nsynchronous Synchronous rotational speed [rpm]

p Number of pole pairs [-]

P Active power [W]

Pconverter Active power converter [W]

Pel Electric active power [W]

Pgen Active power generator [W]

PJoule Joule losses [W]

Ploss Losses [W]

Pmech Mechanical power [W]

Pnom Nominal power [W]

Protor Active power rotor [W]

Pshield+armor Shield and armor losses [W]

Pstator Active power stator [W]

xiv

Pwf Active power wind farm [W]

pdf(vwind)A,C Probability density function of vwind [-]

PF = cos � Power factor [-]

Q Reactive power [Var]

Qcable Reactive power produced by cable [Var]

Qconverter Reactive power converter [Var]

Qdyn,STATCOM Dynamic reactive power range STATCOM [Var]

Qind Reactive power inductive compensator [Var]

Qrotor Reactive power rotor [Var]

Qstator Reactive power stator [Var]

Qwf Reactive power wind farm [Var]

Q+STATCOM STATCOM capacitive power rating [Var]

Q-STATCOM STATCOM inductive power rating [Var]

r Conductor radius [m]

R Resistance [Ohm]

R Radius of turbine blades [m]

RAC, �˚C AC resistance at temperature � [Ohm]

Rcable Cable resistance [Ohm/m]

Rcable,eff Effective cable resistance [Ohm/m]

RDC, �˚C DC resistance at temperature � [Ohm]

�Rshield+armor Resistance increase due to shield and armor [Ohm/m]

s Distance between conductors [m]

t Time [s]

Û Machine voltage [V]

UAC AC grid voltage [V]

Ucable Cable line voltage [V]

Uconv Converter output voltage [V]

voff Wind speed offset [m/s]

vwind Wind speed [m/s]

x Position on cable [m]

xs Resistance effect factor [-]

X Reactance [Ohm]

X Geometric parameter shield losses [-]

XC Capacitive reactance [Ohm]

xv

XL Inductive reactance [Ohm]

XPFC Power factor correcting reactance [Ohm]

yp Proximity effect factor [-]

ys Skin effect factor [-]

� Temperature coefficient [-]

� Pitch angle [deg]

� Voltage angle [deg]

� Conductor temperature [˚C]

� Tip-speed ratio [-]

�armor Armor loss increment [-]

�i �-corrected tip-speed ratio [-]

�shield Shield loss increment [-]

� 3,14159… [-]

air Density of air [kg/m3]

φ Machine magnetic flux [Wb]

Rotational speed wind turbine [rad/s]

off Rotational speed offset [rad/s]

opt Optimal rotational speed wind turbine [rad/s]

rotor Rotational speed rotor [rad/s]

RSC Rotational speed applied by RSC [rad/s]

stator Rotational speed stator [rad/s]

Abbreviation Explanation

ABB Asea Brown Boveri

AC Alternating Current

AFE Active Front End

CCC Capacitor Commutated Converter

cdf Cumulative density function

CSC Current Source Converter

CSCP Cross Sound Cable Project

DC Direct Current

DCF Discounted Cash Flow

DDPMSG Direct-Drive Permanent Magnet Synchronous Generator

DFIG Doubly-Fed Induction Generator

xvi

EBITDA Earnings Before Interests, Taxes, Depreciation and Amortization

EPF Euro Plat Form

FSIG Fixed-Speed Induction Generator

GB Gearbox

GenSC Generator Side Converter

GPMSG Geared Permanent Magnet Synchronous Generator

GridSC Grid Side Converter

HVAC High Voltage Alternating Current

HVDC High Voltage Direct Current

IGBT Insulated Gate Bipolar Transistor

InvC Investment Cost

LCA Life Cycle Analysis

LCC Line Commutated Converter

NGC National Grid Company

NPV Net Present Value

PCC Point of Common Coupling

pdf Probability density function

PF Power factor

PMSG Permanent Magnet Synchronous Generator

PR Phase Reactor

PWM Pulse Width Modulation

RSC Rotor Side Converter

SCIG Squirrel Cage Induction Generator

SG Synchronous Generator

SSC Stator Side Converter

STATCOM Static Synchronous Compensator

SVC Synchronous Var Compensator

TF Transformer

TSO Transmission System Operator

VSC Voltage Source Converter

XLPE Cross-Linked Poly Ethylene

1

1 INTRODUCTION

1.1 Purpose & scope

Voltage Source Converter (VSC) HVDC, also known as HVDC Light® (ABB) or HVDC Plus®

(Siemens), was introduced in 1997 with the commissioning of the 3 MW technology demonstrator at

Hellsjön, Sweden [1]. Until then, HVDC transmission systems were mainly installed as Line Commutated

Converter (LCC) HVDC, also known as HVDC Classic®. After ten years of operational experience and

development, it is clear that in a number of situations, VSC HVDC is a viable alternative for LCC HVDC

and even for HVAC connections. One of those contexts is the connection of an offshore wind farm with

the onshore transmission grid.

Wind energy has come to the stage of large-scale wind farms situated at sea several kilometers from

shore. The importance of the transmission system to shore increases with increasing cable length and

power rating of the wind farm. The transmission of the generated electric power to the onshore grid is

traditionally performed using HVAC technology. The main purpose of this thesis is to study the techno-

economic feasibility of VSC HVDC for the connection of an offshore wind farm with the main grid and

compare VSC HVDC with HVAC.

A first aim is to give an overview of the technical features of VSC HVDC. These features bring several

advantages when VSC HVDC is used as the transmission solution between an offshore wind farm and the

onshore grid. The economic value of these technical advantages is studied further in this thesis.

Several wind turbine topologies are still under consideration for offshore use. A second purpose is

therefore to give an overview of the different wind turbine topologies and highlight their pros and contras

for offshore use. Suggestions are made on how the topology of an offshore wind farm can be optimized

using VSC HVDC instead of HVAC. The main wind turbine topologies are therefore reviewed in

combination with VSC HVDC.

A 300 MW wind farm connected with a 50 km cable is used as a generic example to make the techno-

economic comparison between VSC HVDC and HVAC based on numerical calculations. This proposed

wind farm (rating and distance from shore) is highly relevant in the current offshore wind industry

climate. Both a VSC HVDC and HVAC transmission system are dimensioned to the needs of this wind

farm. The technical comparison comprises the following aspects. The startup of the wind farms is

discussed. Possible grid code requirements, stated by the transmission grid operator, are taken into

account. The losses of both transmission systems are calculated and compared. The necessary calculations

are done in Matlab/Simulink.

The decision for the one or the other technology is based on economic considerations. An economic

comparison is therefore performed as well. Various investor relations, which might have an impact on the

decision, are considered. The systems are economically compared with a discounted cash flow

2

calculation, taking into account differences in investment costs and annual costs and benefits. The annual

costs and benefits considered are maintenance costs, transmission system losses and wind farm energy

output. Reliability and possible differences in insurance costs are not taken into account.

A tool is developed in MS Excel, in order to achieve quick insight in the impact of input parameter

variation on the DCF result. Only cost data from public domain are used. Several assumptions are made

in the economic comparison. The sensitivity of the result of the economic analysis to variation of some

important parameters is therefore investigated.

1.2 Structure

Chapter 1 Introduction (this section) describes the purpose and scope of this thesis and gives an

overview of the structure.

Chapter 2 Technical introduction to VSC HVDC discusses the relevant features of VSC HVDC for the

connection of an offshore wind farm.

Chapter 3 Offshore wind turbine topologies and their use with VSC HVDC discusses first the trends

in the offshore wind industry. The important wind turbine topologies are described. Suggestions are made

for topology optimization and simplification in case VSC HVDC is used as transmission system between

the wind farm and the onshore grid.

Chapter 4 Energy output of offshore wind turbines discusses the difference in energy yield for the

various types of offshore wind farms, either connected with HVAC or VSC HVDC.

Chapter 5 Design and operational aspects of VSC HVDC and HVAC discusses the dimensioning of

VSC HVDC and HVAC for a wind farm rated at 300 MW and a cable length of 50 km. Black start

capability is needed to start up the wind farm. The startup procedure is therefore discussed for both

transmission options. The compliance to possible grid code requirements and the losses of both

technologies are studied as well.

Chapter 6 Economic comparison of VSC HVDC and HVAC for the connection of offshore wind

farms compares both systems for a 300 MW wind farm situated 50 km from the Point of Common

Coupling on an economic basis. Only data from public domain are used. Different wind turbine

topologies are compared in combination with both transmission systems as well.

Chapter 7 Sensitivity analysis investigates the sensitivity of the results of chapter 6 to variation of some

important input parameters in the financial analysis.

Chapter 8 Conclusions summarizes the main conclusions of this thesis.

Additional information can be found in the Appendices.

3

2 TECHNICAL INTRODUCTION TO VSC HVDC

2.1 Introduction

The technical aspects of VSC HVDC are discussed in this chapter. VSC HVDC is first generally

compared with HVAC and LCC HVDC in 2.2 based on literature survey. The main operational aspects of

VSC HVDC are highlighted in 2.3. It will be clear by then that VSC HVDC has several technical

advantages over both HVAC and LCC HVDC. Those technical advantages are treated in 2.4. For the

purpose of this thesis, special attention is given to the important advantages for the connection of an

offshore wind farm with the onshore transmission grid.

2.2 Comparison of VSC HVDC with HVAC and LCC HVDC

The main motivation for the development of DC technology for power transmission was transmission

efficiency, as the power losses of a DC line are lower than those of an AC line with corresponding power

rating. However, historically High Voltage Alternating Current (HVAC) was primarily chosen as the

appropriate technology for electricity transmission [2],[3],[4]. Its main advantage is the straightforward

transformation of voltages, by means of transformers. The invention of the high-voltage mercury arc

valve provided the development of High Voltage Direct Current (HVDC) systems, embedded in AC

grids. This made it possible to integrate HVDC links in AC networks for connections where HVDC

shows more favorable characteristics.

2.2.1 Comparison between HVAC and HVDC cable systems

For long distance underground or submarine transmission of power, HVAC is more difficult to

implement, leaving HVDC as an alternative. An AC cable can be modeled as a long cylindrical shunt

capacitor [5]. This capacitance gives rise to a reactive charging current which increases linearly with the

frequency, the length of the cable and the line voltage. At voltages used for long distance transmission of

electric power, the reactive power related to this charging current becomes considerable and reactive

compensation is necessary at one or both ends of the cable or at suitable intervals. The submarine

connection between an offshore wind farm and the onshore grid is in this case. The investment and

installation costs of compensation equipment add up to the costs of the cable system and make undersea

transmission with HVAC technology less feasible for longer transmission distances [6]. HVDC cables do

not exhibit a steady-state charging current as HVAC cables do.

Furthermore, HVAC cables are commonly installed in a three-phase configuration. The need for three

cables or one complex 3-core cable makes the investment costs per unit length higher for HVAC systems.

Transposition and cross bonding of the cables are necessary to keep the voltage system symmetrical and

to compensate the induced sheet voltages [7], which again adds up to the installation cost of submarine

HVAC systems. These extra costs are not necessary for HVDC links.

4

The substation or terminal costs are generally higher for HVDC because of the presence of one additional

AC/DC power electronic converter per substation. Furthermore, this converter causes additional losses

compared to HVAC systems. Pure cable losses are nevertheless higher for HVAC. The losses in HVAC

cables consist of 4 components [5]:

- RI2 losses in the conductor, increased by skin and proximity effect

- RI2 losses in the metallic shield (current is induced in the shield by the current in the conductors);

shield losses can be in the order of one-third of conductor losses

- RI2 losses in the steel wire armor (current is induced in the armor by the current in the conductors);

armor losses can be in the order of one-half of conductor losses

- Dielectric losses, which are relatively small

The losses in HVDC cables are lower for several reasons. Conductor losses are lower due to the absence

of skin and proximity effect. As there is no alternating current, no armor or shield currents are induced

and the related losses are thus absent as well. The current in HVDC cables is neither augmented due to a

charging current as it is the case for HVAC cables [8].

A general cost comparison between HVDC and HVAC cables is shown in Figure 2.1. There is a break-

even distance between HVAC and HVDC, generally considered between 40 and 80 km for cables [9].

Figure 2.1 Comparison of costs for AC and DC transmission cables [9],[10]

The reader has to be aware of the variations that are possible in Figure 2.1. The choice between VSC

HVDC and LCC HVDC for the DC option influences the result of the comparison significantly.

Furthermore, the power and voltage ratings of the compared transmission systems have an impact on the

outcome of Figure 2.1. The environmental conditions in the immediate surrounding of the cable

(underground or submarine) have an influence on the installation costs as well. For offshore wind farms,

one of the converters of the HVDC link will be placed on an offshore platform, which increases the costs

considerably. Figure 2.1 must be seen with these remarks in mind and specific economic analysis is

needed for each case.

5

2.2.2 Comparison between LCC HVDC and VSC HVDC

Most HVDC schemes in commercial operation today employ LCC HVDC, accumulating to an installed

capacity of more than 60 GW by the end of 2004 [3]. The first commercial LCC HVDC link was

commissioned in 1954 and purposed to connect the Swedish main grid with the island of Gotland [11].

Since its commercial introduction, LCC HVDC has known a great technological development,

particularly in its switching components and control systems. LCC HVDC uses thyristors in a Current

Source Converter (CSC) topology. Thyristors can only switch off when the current through them becomes

zero. The commutation process depends on the normal operation of the surrounding AC grid. The delayed

firing of the thyristors makes the current always lag the voltage. Reactive power is therefore absorbed by

a LCC HVDC link. A CSC is characterized by the unipolar direction of the current. The power direction

is changed by reversing the DC voltage, a time consuming operation. LCC HVDC is especially feasible

for long distance transmission of large amounts of electric power at very high voltages (e.g. +800 kV

claimed by ABB [12], test installation currently built at STRI, Ludvika, SE) or for long submarine

interconnections. For those systems, the economic benefits of low line losses outweigh the extra

investment costs of the AC/DC converter stations.

Voltage Source Converters are a known technology on an industrial basis for many years in a lower

voltage scale. In this context, it is known as Active Front End (AFE) for motor drives providing fast and

continuous control of the frequency and voltage magnitude. The use of the VSC topology for

transmission system purposes is however relatively new [1]. The VSC scheme uses Insulated Gate

Bipolar Transistors (IGBT) which can be switched on and off several times each power frequency cycle

by an external signal. This makes VSC advantageous over LCC because the VSC valves are independent

of the zero crossings of the current and the operation of the surrounding AC grid. Furthermore, the

reactive power, either capacitive or inductive, is controlled autonomously and reactive compensation is

not required. This gives the VSC HVDC an extra advantage over LCC HVDC in terms of power

controllability. Two other advantages of VSC HVDC are the absence of inverter commutation failures

and the limited injection of low-order harmonic currents [1],[13]. However, the numerous switching

operations with VSC HVDC lead to higher losses in the converter compared to LCC. This is a serious

drawback in bulk power transmission, since transmission losses represent high capital, making VSC

HVDC economically less interesting [13]. The losses in a converter station are approximately 1,6-1,8% of

the nominal power rating for each converter station for VSC HVDC and 0,8% for LCC HVDC [14].

Other drawbacks are the limited experience with this new technology, especially in higher power ratings

and the more expensive converter stations. A summarizing comparison between HVAC, LCC HVDC and

VSC HVDC for cable transmission is given in Table 2-1.

6

HVAC LCC HVDC VSC HVDC

Maximum voltage level 150 kV installed 245 kV claimed

e.g. NorNed: + 450 kV

+ 150 kV installed + 300 kV claimed

Substation volume Smallest size Biggest size Medium size Cable installation Complex

(transpositioning, cross-bonding, …)

Simple Simple

Installation cost

Substation Cables

Low High

High Low

Highest Low

Compensation needed Yes Yes No Losses

Substation [% of Pnom] Cables

0,3% High

0,8% Low

1,6-1,8% Low

Offshore experience Many small installations

No 1 project (oil platform)

Active power control No Yes Yes Reactive power control No No Yes Grid interconnections Synchronous Any Any Power flow reversal Fast Slow Fast Black start capability Yes No Yes, elaborate

Table 2-1 Comparison HVAC, LCC HVDC and VSC HVDC for cable transmission

2.3 Operational aspects of VSC HVDC

Figure 2.2 shows the general topology of a VSC substation. The main components are indicated. The

Converter Unit uses Pulse Width Modulation (PWM) to build the appropriate waveform out of the DC

voltage. The phase reactor (PR) is a voluminous component as it is an air coil. By controlling the voltage

angle over the PR, the desired active and reactive power exchange is achieved. The AC filter is used to

filter out the harmonics introduced by the switching of the converter. The interface transformer adapts the

voltage magnitude to the level of the Point of Common Coupling (PCC).

Figure 2.2 Components of VSC HVDC terminal [13]

7

2.3.1 Control of AC voltage amplitude and frequency

The fact that the power electronic switches in a VSC can be switched on as well as off, is of great

importance for the controllability of a VSC transmission link. By using PWM, the desired output voltage

amplitude is controlled by modulating the width of constant DC voltage pulses by setting the modulation

index [15],[16]. This makes the output voltage of the VSC HVDC link independent of the AC grid,

enabling unenergized network start and power delivery to weak AC grids. In Figure 2.3, a 2-level 3-phase

bridge is depicted [13]. Free-wheeling diodes are added in parallel with the switching devices to ensure

reverse current capability and to prevent the application of reverse voltages. The IGBTs are switched at a

fixed multiple (typically higher than 20) of the fundamental grid frequency. For a 2-level bridge, the AC

voltage waveform is built out of positive and negative DC voltages as shown in Figure 2.4 (red). The DC

capacitance is assumed to be infinite for the purpose of this figure (no DC voltage ripple). The voltage

measured behind the phase reactor and after filtering is depicted in blue.

Figure 2.3 Scheme of 2-level 3-phase Voltage Source Converter bridge

Figure 2.4 2-level converter single phase output voltage (fswitch = 21 times fundamental)

In order to lower the harmonic distortion of the output voltages, higher level converter bridges are used. A

3-level neutral point clamped bridge is commonly used for practical installations of VSC HVDC e.g.

Murray Link (Australia) [17] and Cross Sound Cable (USA) [18]. The scheme of a 3-level VSC bridge is

shown in Figure 2.5. A typical output voltage waveform of a 3-level Voltage Source Converter can be

8

seen in Figure 2.6 (red). The voltage waveform after the phase reactor and after filtering is shown in blue.

Figure 2.5 Scheme of 3-level 3-phase Voltage Source Converter bridge

Figure 2.6 3-level converter single phase output voltage (fswitch = 21 times fundamental)

Higher level converter topologies (e.g. 5-level, 7-level…) are able to lower the harmonic distortion

further and can lower the number of switching operations per IGBT. These higher level converter

topologies promise lower switching losses. However it becomes clear that the system complexity

increases rapidly with the number of voltage levels. This involves higher investment costs, which is why

practical schemes are limited to three levels upto now. The HVDC Plus® system by Siemens is based on

a multi-level approach, but no practical installations are operational yet [19]. Lower converter losses are

nevertheless expected.

The frequency of the AC voltage wave is controllable in two ways:

− The frequency of the oscillator that controls the valve firing can be controlled and varied.

− When the VSC feeds into an AC system with active generators, the converter station can participate in

the primary frequency control by regulating the active power it takes from or delivers to the AC

system.

2.3.2 Control of active and reactive power

Power flow control is independently possible at each converter [20]. In order for the DC link voltage to

9

remain constant, the following equation must hold

lossinvrect PPP += (2-1)

One of the converters is responsible for keeping the DC voltage on the link at its nominal level. This is

done by controlling the net power going into the DC capacitance of the link. The other converter sets the

amount of active power through the link.

The phase reactors play an important role in the control of the active and reactive power each converter

station supplies to the grid. All VSC transmission systems possess a series inductance separating the

power electronic converter from the AC grid. A simplified system is shown in Figure 2.7. The active

power and reactive power injected in the AC grid are given by

X

UUP convAC δsin

= (2-2)

( )

X

UUUQ ACconvconv δcos−

= (2-3)

X represents the series reactance of the phase reactor and the transformer in the converter station. �

represents the phase angle between the voltage waveform at the converter and the voltage in the AC grid.

A voltage measurement is performed in the AC grid. Furthermore, a phase locked loop measures the

frequency and phase angle in the grid in order to control �.

Active power control is performed by controlling the phase angle of the fundamental frequency

component of the converter voltage. The lagging or leading of the converter voltage compared to the AC

grid voltage sets the amount of active power transfer to or from the grid.

Figure 2.7 Power control of VSC with phase reactor

The reactive power transfer is controlled by the converter voltage amplitude. If the converter output

voltage is higher than the AC grid voltage, the converter acts as an overexcited synchronous generator

and pushes reactive power to the grid. When the converter output voltage is lower than the AC grid

voltage, the converter acts as an under-excited synchronous machine and consumes reactive power from

the grid. The possibility to consume as well as produce reactive power in a controlled manner is a unique

feature of VSC HVDC [22],[13]. The control of active and reactive power is furthermore almost

instantaneous.

The capability of VSC HVDC to absorb and inject active and reactive power is generally represented in a

P,Q-capability chart (Figure 2.8). The desired active power through the link is normally the first concern.

For example for an offshore wind farm, it is the produced amount offshore that needs to be evacuated.

10

Active power through a link can be specified by the electricity market or by loss minimization strategies

when incorporated in a meshed grid as well. The amount of available reactive power will depend on the

amount of active power being already transmitted, depending on the MVA rating of the converters.

Therefore, the P,Q-capability characteristic is typically a circle with a radius equal to the maximum MVA

rating of the converters. When the voltage in the AC grid varies, the converters are restricted by the

current rating of the power electronic switches and the capability circle scales accordingly. The reactive

power capability of the converter is more limited on the capacitive side of the P,Q-capability chart. The

voltage is raised above the AC grid voltage (2-2) to inject reactive power. The converter voltage is

however restricted to the maximum value of the power electronics and this makes the P,Q-capability chart

more limited on this side for higher AC voltages.

Figure 2.8 P,Q-capability chart of VSC HVDC link

A LCC HVDC link solely consumes reactive power. The uncompensated reactive power demand at the

terminals of a LCC HVDC converter is about 50-60% of the active power flow. Some LCC schemes use

another topology, e.g. Capacitor Commutated Converter (CCC) HVDC with a capacitor to provide

reactive power compensation. This produces a characteristic in the P-Q diagram that is more reactive

power neutral. The dynamic control of reactive power is not possible with LCC HVDC.

2.4 Technical Advantages

It is clear that VSC HVDC has, because of its high controllability, some specific technical advantages

over HVAC and LCC HVDC. These technical advantages of VSC transmission can offer, next to power

transfer, other system benefits. This can bring an economic value to VSC HVDC justifying the higher

investment costs and converter losses. Depending on their importance for the connection of an offshore

wind farm, they are discussed in more detail.

2.4.1 Power flow control

The operator can control the amount of active power through a VSC HVDC link. This makes it possible

11

to set the power flow as desired by the electricity market. Loop flows, because of different impedances on

different parallel connections can be opposed. Typical examples of loop flows in Central Europe are the

flows present on windy days in the north of Germany [23]. The major load centra in Germany are situated

in the south of the country, whereas the production is mainly in the north on windy days. Because of the

limited connections between the north and south on German territory, the flows search other, low

impedance routes through neighbouring countries. These currents tend to overload the connections in The

Netherlands, Belgium and France on the one side and in Poland and Czech Republic on the other side.

Another advantage of power flow control is the possibility to reduce overall network losses in a meshed

grid with VSC HVDC. The results of optimal power flow algorithms can be practically implemented due

to this feature.

This feature might seem less relevant for an offshore wind farm, as all the produced electricity needs to be

evacuated to the shore. The control of active power can nevertheless be used to improve the frequency

response of the wind farm as will be discussed later in Chapter 5.

2.4.2 Grid voltage support and power system stability

The voltage level in the grid is related to the reactive power. A VSC station can absorb or supply reactive

power almost independent of the active power. By choosing the appropriate working point in the P,Q-

capability chart, the transmission system operator can control the voltage level in the grid close to the

converter station. This improves the voltage stability in the grid [24]. Furthermore, the possibility to

operate the transmission system at a slightly higher voltage reduces the losses in the surrounding grid.

VSC HVDC can use this advantage especially when connected to a weak AC system [25]. This is often

the case for the connection of an offshore wind farm. Grids in the neighbourhood of coasts often need

reinforcements (e.g. extra transmission capacity, reactive power generators, …) when large generation

blocks such as offshore wind farms are connected (e.g. Germany, Belgium, …). The use of VSC HVDC

as a transmission solution between the wind farm and the onshore grid reduces the need for these

reinforcements.

2.4.3 Black start capability

The term black start capability originates from the operation of the electricity transmission grid. Under

normal operating condition, the transmission system is characterized by a sinusoidal voltage waveform

with a certain amplitude and frequency. It is fair to state that at any considered point in the grid, the

voltage amplitude is at a relatively fixed value, close to the rated voltage. The frequency is the same

throughout the whole grid and is controlled very strictly at a fixed value. When the transmission grid

experiences a blackout, the entire system de-energizes and all generation units come to a stop. Afterwards

it has to be re-energized or black started. In this way, a system exhibits black start capability when it can

bring a power system from shutdown condition to the initial stages of power system restoration, enabling

the return to normal condition [26]. For this purpose, the black start solution must be able to deliver

power to a load network with an independently created sinusoidal voltage waveform, with given voltage

12

amplitude and frequency. The term black start capability is nowadays used in two contexts.

In the traditional meaning, the term is used for generating units who can start a grid after blackout, based

on another power source than the electricity grid itself. Usually, black start generating units have backup

power (e.g. diesel-driven synchronous generators) or an independent energy supply (hydro or pumped

storage plants). They are used in case of total electricity grid blackout to create the voltage waveform. By

creating a rotating field, it becomes possible to start up or connect other generators [27]. The transmission

grid is then re-energized progressively, first of all supplying to smaller “power islands”, later on

connecting these islands to larger subsystems and finally accomplishing the restoration of the entire

system [28]. It is common to use synchronous generators as electrical machines in black start generating

units. Induction generators have reactive power requirements which make them unsuited for black start

purposes [28]. The black start generators must be able to cope with steep load changes, both active and

reactive. Active load blocks are added progressively to the generators in order to supply more and more

demand. Furthermore, during transmission system recovery, long lines between generating units are

energized. These lines represent important reactive power loads. The supply of this demanded active and

reactive power is possible with synchronous generators through the control of the applied torque by the

primary energy source and the excitation of the synchronous machine. The control of active and reactive

power output is also considered as part of the black start capability.

A new concept is the black start capability of transmission systems. Black start capability of transmission

links can be defined as the possibility to deliver the appropriate voltage waveform, independently of any

other device, at one end of the link, based on normal grid operation at the other end of the link. The power

source for the black start generation is thus the electricity grid itself and the transmission line can only use

its black start capability if the sending end does not experience blackout. Furthermore, the active and

reactive power through the link is controlled by the link operator.

Voltage Source Converter (VSC) HVDC is in this case. When the HVDC link is charged, the DC voltage

is thus present, the AC voltage waveform at the receiving end can be controlled by switching the IGBTs

in the converter station [15],[16]. The DC link disconnects the dead grid from the operating grid and can

start up the dead grid progressively after blackout [29]. By controlling the load angle, the active and

reactive power is set and the link is put in any point on the P,Q-chart, almost instantaneously. These

features are very useful in the power delivery to remote locations, the start-up of a network after blackout

or the startup of an offshore wind farm. An example of a VSC HVDC, with black start capability enabled,

is the Estonian Converter of the Estlink (Finland – Estonia). An auxiliary power winding is added to the

transformer for self-supply during “house-load” operation. In the unlikely event of a black-out in the

Estonian network, the Estlink transmits power from the asynchronous Finnish grid (part of Nordel) to the

Estonian grid and restores grid operation there [30],[31]. Likewise, the Cross Sound Cable Project

(CSCP) is a VSC HVDC link between New England and New York. It consists of a 40 km pair of cables

rated at 330 MW [18]. During the restoration following the August 14, 2003 blackout, this cable was

instrumental in restoring electric power to customers on Long Island. The CSC transported 330 MW to

13

Long Island, enough power to restore electric service to over 300,000 homes. The AC system voltage

stability in Connecticut and on Long Island was highly enhanced during and subsequent to the network

restoration by CSCPs AC voltage control feature. The fact that VSC HVDC does not need an active grid

in order to become connected allows it to be available almost instantly after blackout [24]. Since the

consequences for society differ significantly between a blackout of 15 min and a blackout of 6 hours, the

speed and robustness of VSC HVDC can be very valuable.

The definition of black start capability is different for transmission systems than for electricity generating

units. It makes sense to define black start capability only for transmission systems that form a connection

between two separated networks. Examples of separated networks are asynchronous grids or grids with

only one connection between them (e.g. wind farm grid – onshore grid, small island – onshore grid, …).

The transmission line possesses black start capability if it can feed from an operational network into a

‘load only’ network.

Line Commutated Converter (LCC) HVDC does not have black start capability because the switching of

the thyristors depends on the presence of the AC grid voltage. Normal AC grid operation on both sides of

the link is an absolute requirement for LCC HVDC to become operational. HVAC transmission cables

and lines can connect a ‘load only’ network to a generating network and make it part of the total network.

The control of active and reactive power is nevertheless not possible with HVAC, without the use of

auxiliary devices. The voltage amplitude can neither be controlled without additional equipment and the

frequency is fixed to the grid frequency. HVAC inherently has black start capability, but in a more limited

way than VSC HVDC as it is not controllable.

Black start capability is an absolute requirement for an offshore wind farm connection. A wind farm

behaves as a ‘load only’ network during startup. A voltage waveform with appropriate amplitude and

frequency has to be applied to bring the generators online. Depending on the chosen type of generators

and wind turbine topologies in the wind farm, the wind farm has typical active and reactive power

requirements. This will be discussed further in this thesis. The lack of black start capability is a serious

drawback for LCC HVDC and makes it unsuited for the link between an offshore wind farm and the

onshore grid. It will be shown that only VSC HVDC has full black start capability for a wind farm,

whereas HVAC always needs additional equipment.

2.4.4 Connection of asynchronous grids

Opposed to HVAC, HVDC can connect two asynchronous grids or even grids with different frequencies.

With VSC HVDC, it is even possible to independently control the frequency in the grid. This can be

useful in the case of an offshore wind farm, an isolated load or two independent systems where mutual

frequency support is possible. VSC HVDC does not restrict the frequency to the main grid frequency and

can even handle variable frequency operation [25]. By varying the frequency in a wind farm, it is possible

to control the rotational speed of the wind turbines. This approach is discussed further in Chapter 3. An

example of a VSC HVDC connection with variable frequency supply to a load is the connection to the

14

Norwegian rig Troll A [25],[32],[33]. The VSC HVDC system is used to transport power from the

Norwegian grid to the load side 70 km from the coastline. The load is an electric motor. The rotational

speed of this motor is controlled by the VSC HVDC link through variation of the load side frequency.

2.4.5 Voltage dip ride-through and power quality

A VSC HVDC link decouples two grids in such a way that faults in one grid do not propagate to the other

grid. VSC HVDC therefore delivers the required voltage dip ride-through capability for an offshore wind

farm. The rapid AC voltage control is also used to improve the power quality, including the reduction of

flicker and other transient disturbances [20]. Faults on the DC side of the link can not be handled by the

converter stations. AC grid breakers have to open and clear this kind of faults. This is one of the reasons

why VSC HVDC uses preferably cables instead of overhead lines, because of the reduced exposure to

lightning.

2.4.6 Cable choice

VSC HVDC uses simple solid extruded insulation cables. The change of power direction is achieved

through reversal of the current direction and not by polarity switching of the voltage as with LCC HVDC.

Solid extruded insulation cables cannot withstand a polarity reversal and fluid-filled or mass impregnated

cables must therefore be used with LCC HVDC. They are more complicated to handle, especially in a

submarine environment [20]. Solid extruded insulation cables avoid the leak of oil present for fluid-filled

cables [21].

Furthermore, HVDC cables are experienced to have a lower environmental impact than HVAC cables.

The need for three phases increases the number of cables or makes the cables more voluminous. More

trenches and more boat runs are needed to install HVAC leading to a greater seabed disturbance. The

isolation of DC cables can be thinner than for HVAC cables as well. With HVAC the peak voltage is a

factor 2 higher than cableU , whereas this factor is not present for HVDC links. More power per mm2

conductor is transferred with HVDC cables than with HVAC cables.

2.5 Conclusion

The technical aspects of VSC HVDC were discussed in this chapter. It became clear that VSC HVDC has

interesting qualities to form the connection between an offshore wind farm and the onshore transmission

grid. The most important features are the control of active and reactive power, the possibility to connect

two asynchronous grids with even differing and varying frequencies and the black start capability. The

lack of black start capability makes LCC HVDC unsuited for the connection of offshore wind farms.

HVAC has several disadvantages compared to VSC HVDC but is nevertheless often used in this context.

The higher investment costs for short cable lengths, and the losses in the converter stations are a

drawback for VSC HVDC compared to HVAC.

15

3 OFFSHORE WIND TURBINE TOPOLOGIES AND THEIR USE

WITH VSC HVDC

3.1 Introduction

More and more countries, especially in Europe, investigate the installation of offshore wind farms. The

major trends in the offshore wind industry are discussed in this chapter. It is shown that the wind farms

for the future are expected in a power range of several hundred MW and situated tens of kilometers from

the coast (3.2). A discussion on wind turbine topologies is given further in this chapter (3.3). The

investigated topologies are based on standard onshore topologies for wind turbines. The specific problems

and constraints for offshore installation are highlighted. The different wind turbine topologies are

reviewed in combination with a VSC HVDC link (3.4). It is shown that the use of VSC HVDC allows for

optimization or simplification of the offshore wind farm. The discussion held in this chapter is qualitative.

Interesting topologies for offshore use are taken further in this thesis and discussed in a later chapter in a

more quantitative way.

3.2 Offshore wind trends

Mankind faces an increasing electricity demand for the future. Fossil fuels are running out and their use

for electricity generation is harming the environment and affecting the climate [34]. This makes it

possible for alternative technologies to enter the electricity production market. Governments all over the

world stimulate electricity production out of renewable resources. A promising option, if not proven, in

this context is electricity out of wind energy. The installed wind capacity is growing at a very high annual

growth rate worldwide. According to the World Wind Energy Association, 93,8 GW of wind power was

installed at the end of 2007 [35] and more than 160 GW is expected at the end of 2010 [36]. As an

extension of the proven wind farm technology onshore, wind turbines are increasingly installed offshore.

Several reasons exist for moving wind turbines away from the land.

A first advantage at sea is that wind conditions are better than onshore. First of all, the average wind

speed is significantly higher. The power in the wind is proportional to the cube of the wind speed and the

increase in annual energy output by moving wind turbines offshore is considerable. The higher average

wind speed thus has a positive effect on the capacity factor of the wind farm. As an example, the Danish

Nysted offshore project showed a capacity factor of 47% in 2005 (an average wind-speed year for

Denmark) [37]. The capacity factor for offshore wind farms is generally expected to be around 40%,

whereas for coastal regions 32% and for inland regions 25% can be expected [38]. Secondly, the wind

speeds are less turbulent due to the smooth surface of the surrounding water. The wind speeds neither

increase as much with the height above sea level as they do above land. This makes it possible to use

lower towers (for the part above the sea level) for wind turbines located offshore [39].

Another reason for installing wind farms on the sea is that wind turbines need space. The number of

16

lasting suitable wind sites on land is decreasing, especially in densely populated areas such as Western

Europe [40]. Furthermore, the social acceptance for onshore wind energy reaches its limits in some

countries.

The possibility to build large scale wind farms offshore is important as well. Onshore wind projects are

often limited in power rating and more considered as distributed generation. Aggregation of wind turbines

in a large wind farm brings important scale advantages to the wind farm owner as well as to onshore grid

operators. The larger scale of the wind farm helps the owner to produce electricity at more competitive

cost compared to other generating units [41]. A possible advantage for grid operators is the connection of

larger wind farms to a higher voltage at the Point of Common Coupling (PCC). This relieves the

distribution grids on lower voltages, which are now sometimes highly loaded on windy days. The

centralized connection of wind energy makes it easier to account for it and to take grid reinforcing

measures that might be needed.

Another reason is that Europe strives to produce 20% of its electricity demand out of renewable resources

in 2020 and counts on offshore wind energy to achieve this goal [42]. Offshore wind farms are therefore

supported by the European governments through subsidies, grants and certificates [43]. The governmental

financial support is often higher for offshore wind projects than for onshore projects [44].

Historically, the first offshore wind farms are built near shore. The generated electricity is transmitted to

the shore on the voltage level of the offshore grid. This voltage level is typically not higher than 36 kV.

From a transmission grid side of view, there is no significant difference with other distributed generation

by wind turbines onshore. Those wind farms are mostly small scale (<50MW). Leading countries in the

development of small scale offshore wind production units are Denmark, UK (Round 1 Projects), The

Netherlands and Sweden [37]. The evolution of offshore wind farm development is shown in Figure 3.1.

At the end of 2007 a total installed capacity of 1,1 GW was realized offshore. The expected additional

offshore capacity that will be commissioned in 2008 is higher than any year before.

Figure 3.1 Offshore wind farm development in Europe [37]

17

Today, more and more wind farms are planned on a larger scale and further from shore. Power ratings of

several 100 MW and transmission distances to shore of tens of km are no longer exceptions. Several

reasons exist to push wind farms to higher power ratings and further from shore.

The further from shore, the better the wind conditions are. The wind speed is considered to be higher on

average and the wind is less turbulent. This leads to an increased energy output of the wind turbines

compared to nearshore wind farms.

Another reason is based on juridical considerations in the procedure to get the necessary building permits.

It is socially not always acceptable to put offshore wind farms too close to shore in some countries and

governmental or juridical decisions are taken to get them out of sight [45]. Furthermore, coastal waters

are often frequented by ships to access numerous ports on the coasts of Europe. Other waters are used for

fishing or military purposes [46]. Those reasons make the number of suitable near shore sites limited.

For wind farms situated at several tens of km from shore, it is no longer attractive to use the collection

grid voltage of the wind farm to transport the electricity to shore. The losses along the cable depend on

the current through the cable. It is advantageous to set up an explicit transmission system between the

wind farm and the onshore grid, rated at a higher voltage in order to reduce the current and the related

losses. A substation is installed on an offshore platform to collect the power from the individual turbines

and transform the voltage to higher, more suitable values.

The cost of the installed platform and substation offshore is divided over the different turbines. The more

turbines connected to the substation, the smaller the impact of this transmission system cost is in the total

cost of the electricity produced offshore. For this reason, wind farms built at larger distances from shore

contain more individual turbines and have thus a higher power rating.

It is nevertheless not feasible to put an offshore wind farm too far from shore, and the majority of the

planned and built wind farms are therefore situated within 60 km from shore. Examples are the planned

Belgian wind farms on the Thornton (30 km from shore, commissioning 2008) and Bligh Bank (45 km

from shore, commissioning 2010) and the Dutch Q7 project (23 km from shore, commissioning 2007).

Several reasons exist for the preference of not going too far from shore [37].

First of all the wind farm owner wants to reduce the transmission distance for the generated electricity.

Whatever cable technology is chosen, the cable investment and installation costs, as well as the losses,

will increase with the cable length and tend to nullify the possible gain in capacity factor. Furthermore,

constructional reasons such as larger depth of the sea and longer transportation distance of building

material increase the installation costs when moving the wind farms further from shore. The work

conditions are harsher during installation and maintenance of the turbines [47].

An exception to this trend is the German planned Borkum 2 wind farm. This wind farm will have a rated

capacity of 400 MW and will be situated over 200 km from the point of common coupling (PCC) in the

German grid (node at Diele) and 128 km from shore (commissioning 2009). It is part of the first phase of

18

a larger project, which is scheduled to have a capacity of 1500 MW when completed. The large wind

project fits in the German strategy to achieve the Kyoto targets [48],[49]. The total capacity will be built

on several sandbanks in the immediate surroundings of Borkum 2. The transmission system operator pays

for the connection between wind farms and the onshore grid in Germany. This makes the distance from

shore less important for private wind farm developers.

After this discussion, it is concluded that typical offshore wind farms are expected to be built between 30

and 60 km from shore. As mentioned in 2.2.1, this is about the expected break-even distance between

HVAC and HVDC for submarine cable connections. The power rating is expected in a 200-500 MW

range per wind farm. Table 3-1 shows some planned wind farms in Northern Europe with their respective

power ratings. Approximately 20-25 GW of offshore wind power is expected in the German North and

Baltic Sea by 2030 [47]. Some projects are planned farshore (>100 km). The projects in the UK are the

closest to shore (10-20 km) and therefore in a smaller power rating.

Country/Wind farm Rating [MW] Country/Wind farm Rating [MW]

Belgium

Thornton Bank zonder naam Bligh

300

180-250 330

Ireland

Arklow Kish Bank

520 250

Denmark

Horns Rev 2 Omø Stålgrund/Nysted 2

200 200

Netherlands

Egmont aan zee Ijmuiden Western Scheldt mouth

108 100 100

Sweden

Kriegers Flak Lillgrund Utgrunden 2

600

96-144 90

Germany

Borkum Riffgat Borkum Riffgrund Borkum 3 Borkum 4 Borkum Riffgrund West Butendiek Dan-Tysk Helgoland 1-3 Möwensteert Nordergründe Nordsee AWS Pommersche Bucht Sandbank 24 Schleswig-Holsteinische Nordsee

130

600-1000 60

400 up to 1800

240 1500

800-1000 210

150-600 500-1000

200 1000

400 500-1000

United Kingdom Barrow Burbo Cromer Gunfleet Kentish flats London Array Shell falt Solway Firth Southport Svarweather sands Tesside

90 90 90 90 90

4x275 270 180

90 90 90

Table 3-1 Planned and considered offshore wind farms in Northern Europe [50]

It has to be noted that the length of the transmission cable is the distance from the offshore substation to

the Point of Common Coupling (PCC) in the onshore grid and not just the distance to shore. A wind farm

of several 100 MW has to be connected to the grid on a suitable voltage level. A grid of this level is not

always present at the coast (e.g. Germany, Belgium…) and the additional onshore distance to a suitable

grid node adds up to the final length of the transmission cable. Once onshore, it might be possible to use

overhead lines which are much cheaper, but in general it is virtually impossible to construct these lines in

19

Europe because of public opposition.

3.3 Wind turbine topologies

Basically, three major wind turbine topologies are distinguished. The first generation wind turbines were

asynchronous machines directly connected to the grid. They are discussed in more detail in paragraph

3.3.1. In the search for variable-speed operation of wind turbines, two other topologies were introduced.

The first one, still based on an asynchronous generator, uses a converter to feed the rotor. The second one,

based on a synchronous generator, uses a full power converter. They are discussed in more detail in 3.3.2.

The discussed topologies originate from on land wind turbines. For offshore wind turbines some specific

constraints hold that are less important for land based wind turbines. First of all, the installation of wind

turbines offshore is more complicated than onshore. A reduction in weight and volume of the different

components or in the number of components per wind turbine has a higher impact on the total installation

cost of the wind farm offshore than onshore. Another important topic is the amount of maintenance

needed for the turbine components. Offshore maintenance is difficult and more costly than onshore.

Furthermore, the access to the wind farm is limited to some months of the year due to the weather

conditions. This diminishes the overall capacity factor of the offshore wind farm. As one of the reasons

for moving wind turbines offshore was the expected higher capacity factor, a minimum of maintenance

and a high reliability are certainly advantages for an offshore wind turbine topology.

3.3.1 Fixed-speed wind turbines

The first wind turbines were equipped with an asynchronous generator operated at fixed rotational speed.

The topology is known under the name Fixed-Speed Induction Generator (FSIG) topology and was very

popular in the early years of wind turbine development. It is still used by a high portion of the installed

wind capacity. As an example, the majority of existing land-based wind turbines still used this topology

in 2003 in the United Kingdom [51]. The topology is also called the ‘Danish model’ because of its high

penetration in the Danish grid. The outlook of the FSIG topology is shown in Figure 3.2.

Figure 3.2 Fixed-Speed Induction Generator wind turbine topology

The generator is a brushless squirrel-cage induction generator. A typical output voltage for this kind of

generators is 690V which explains the presence of a transformer for the connection to the grid [51].

20

Generators of this type are available up to 5 MW in power. The output voltages are higher for higher

power ratings (up to 11 kV) [52]. Some of the main advantages are the robustness, the mechanical

simplicity and the low cost of this configuration. This is mainly due to the limited amount of components

to form an operational wind turbine. In the following paragraphs the major disadvantages are discussed.

Asynchronous generators are normally produced with a low number of pole pairs (typically p = 2 or 3),

which involves a higher synchronous speed (1500 or 1000 rpm) than the typical optimal rotational speed

of the turbine blades (~10 rpm). A gearbox is therefore required in wind turbines with asynchronous

generators. Gearboxes are expensive and heavy components. Generally, gearboxes represent about 15-

20% of the total turbine cost for large scale wind turbines (� 2MW) [53],[54]. For offshore wind turbines,

this gearbox imposes an extra drawback because of the high maintenance costs (e.g. need for lubrication).

Generators with a higher number of pole pairs (p>10) are unacceptable because of their low power factor

(PF = 0,6 without compensation) [35].

Due to the single direct connection to the grid, the generator is obliged to operate at fixed speed. It is

advantageous for the energy output of a wind turbine to operate at variable rotational speed as will be

discussed in the following paragraph. The FSIG speed varies by only a few percent (variations in slip

speed) caused by sudden changes in wind speed. The steep torque-speed characteristic around the

synchronous speed of induction machines makes these variations minimal. Fluctuations of the mechanical

power due to wind gusts are therefore quickly transferred to the grid [55]. There is no possibility to use

the inertia of the turbine blades to absorb wind speed fluctuations. This puts higher mechanical stresses on

the blades than in a variable-speed configuration. The audible noise is higher than for variable-speed wind

turbines as well. This is nevertheless of minor concern for offshore wind farms.

Another drawback is the constant consumption of reactive power by induction machines. The generator

can only be operational when it is connected to a grid that delivers this reactive power demand. The

machine consists mainly of copper windings and massive iron. Therefore, seen from the grid, an

induction generator behaves as an inductance with a lagging power factor. A typical value for the power

factor is 0,89 during nominal operation [55]. The induced magnetic field rotates at synchronous speed,

determined by the number of poles and the frequency of the grid. To act as a generator the rotor has to be

driven at a speed higher than this synchronous speed. Because of the relative motion between stator and

rotor (slip), a current is induced in the rotor bars. The electromagnetic interaction of the rotor and stator

field results in a torque on the rotor.

The constant demand for reactive power can cause detrimental effects on the surrounding grid voltage

level, which would make the induction generators absorb even more reactive power [51]. This can lead to

voltage instability in grids with high wind penetration. Transmission grid operators released more

stringent grid codes for the connection of wind turbines because of this reason, e.g. [56],[57]. It is now

necessary to install additional power factor correcting capacitors at each wind turbine or a common

capacitor bank for larger wind farms. These capacitors are typically rated at 30% of the wind farm

21

capacity [51].

A very high inrush current is drawn from the grid when the generators are started. This transient current,

which can be up to 8 times the rated current, occurs because of the magnetization of the machine. The

effect is similar to the switching of an RL-circuit. The large inrush currents of a wind farm can produce

severe voltage sags in the grid. A soft-start device is often installed to reduce the starting current [35].

Because of the mentioned disadvantages, the FSIG topology is normally not installed in power ratings

higher than 2 MW [58] and it is more and more abandoned for onshore use. The main disadvantages hold

as well for a large scale offshore wind farm connected to the onshore grid with HVAC. The wind farm

will operate at a constant frequency of 50 Hz in Europe and the rotational speed of the turbines will

therefore not be variable. The variation in total reactive power demand of the wind farm as a function of

the wind speed makes the connection with HVAC complicated and infeasible as well. The voltage

stability of the total system is poor for large projects situated several kilometers from shore. The newest

grid code requirements on voltage dip ride-through capability and power factor control can not be

fulfilled without expensive additional equipment such as SVC or STATCOM. This topology is therefore

no longer under consideration for large offshore wind farms connected to the onshore grid with HVAC.

3.3.2 Variable-speed wind turbines

Because variable-speed wind turbines have the potential for increased energy capture, wind turbine

manufacturers have begun to explore this possibility in the nineties [59]. The performance of a wind

turbine is expressed by a coefficient of performance pC . The power captured by a wind turbine at a

certain wind speed is given by

( ) 32,21

windpairmech vRCP ⋅⋅⋅⋅⋅= πβλρ (3-1)

The coefficient of performance depends on the tip-speed ratio � and the pitch angle �. The tip-speed ratio

is defined by

windv

R ωλ

⋅= (3-2)

with ω the rotational speed of the blades and R the radius of the rotor.

The coefficient of performance is approximated for a three-blade turbine by [113]

( ) ieCi

pλβ

λβλ

5,12

54,0116

22,0,−

���

����

�−−= (3-3)

with

1

035,008,0

113 +

−+

=ββλλi

(3-4)

22

pC as a function of �, with � as a parameter is shown in Figure 3.3. The theoretical maximum for the

coefficient of performance is equal to the Betz limit (= 0,593). The maximum achievable coefficient of

performance for each � is shown in red. By varying the rotational speed (variable speed operation), � is

controlled in order to optimize pC and reach a higher output power at low wind speeds. The coefficient

of performance can as such be held longer at an optimum in Figure 3.3.

Figure 3.3 Coefficient of performance as function of � with � as a parameter

A Matlab/Simulink model of a 5 MW wind turbine is developed for this thesis. It is not the main purpose

of this thesis to develop a wind turbine model, but the results are used throughout this thesis. Information

on this model is given in Appendix A. The possible gain in energy capture can be seen in Figure 3.4. The

red line shows the power output of a fixed-speed wind turbine as a function of wind velocity. The blue

line depicts the same for a wind turbine operated at variable rotational speed. The nominal power of both

wind turbines was chosen at 5 MW. The wind turbines assumed in this thesis stop operation because of

safety reasons at 30 m/s (cut-out wind speed).

Variable-speed wind turbines always need a power electronic converter (Voltage Source Converter) to

facilitate the variable-speed operation of the generator. This converter poses an additional cost on the

wind turbine. The converter introduces extra losses in the system and the extra energy captured from the

wind is partially lost. The choice for a converter is made upon economical considerations.

A disadvantage of the presence of a converter is that, depending on the relative rating of the converter

compared to the total wind turbine rating, the contribution of the turbine to the short-circuit capacity of

the grid is reduced. Large disturbances cause large fault currents both at the stator and the rotor. Voltage

Source Converters are not very tolerant to high currents and they will block during severe disturbances

and not contribute to the short circuit current [55]. This drawback is only important in grids with a high

penetration of generation connected through a converter (e.g. wind turbines, photo-voltaic cells, …).

23

Another disadvantage is the harmonic injection into the power system, originating from the high

switching frequencies used in the converter. Harmonics cause extra losses and have a negative effect on

system components (transformers, compensators, …) in the surrounding grid.

Figure 3.4 Power output of fixed-speed and variable-speed wind turbine (5 MW)

as a function of wind speed (based on matlab wind turbine model)

There are several other reasons, next to the extra energy output, why more and more wind turbines are

equipped with a converter in the nacelle.

First of all, the converters can control the overall reactive power consumption of the wind turbine

electrical system. Within certain ranges, it is possible to control the absorption and production of reactive

power. Therefore, a power factor of 1 can be achieved without additional equipment. This is an important

quality in order to meet the grid code requirements of transmission system operators without using

additional power factor correcting equipment.

Another advantage of converter based topologies is that currents are tightly controlled. This means that

the torque is held more constant and rapid fluctuations in mechanical power can be temporarily stored as

kinetic energy in the inertia of the turbine (variation of rotational speed). This results in a less variable

output power, active as well as reactive, and enhances the power quality in the grid (e.g. reduction of

flicker) [55]. The mechanical stresses on the blades and other components are lower for this reason. As

said before, variable-speed wind turbines have a lower output of audible noise than fixed-speed wind

turbines [43].

Two important types of variable-speed wind turbines exist. The first uses a partial power electronic

converter rated at about one third of the wind turbine nominal power rating to feed the rotor of a wound

asynchronous generator. This topology is well known under the name Doubly-Fed Induction Generator

(DFIG) and will be discussed in section 3.3.2.1. The second option uses a full power converter and is

24

known as Direct-Drive Synchronous Generator (3.3.2.2).

3.3.2.1 Doubly-Fed Induction Generators

The Doubly-Fed Induction Generator (DFIG) topology is shown in Figure 3.5. Nowadays, this

configuration is the dominant solution for newly built wind turbines [61]. It is available in power ratings

up to 5MW and higher power ratings are under consideration. A transformer is needed for the connection

to the grid.

Figure 3.5 Doubly-Fed Induction Generator wind turbine topology

The generator is connected to the grid via two paths [51]. The first one is the direct connection of the

stator to the grid. The rotor is fed via a double power electronic converter. A variable frequency voltage

waveform is imposed on the rotor via the Rotor Side Converter (RSC). The difference between the

frequency applied to the rotor and the frequency on the stator allows the variable-speed operation of the

turbine [55].

statorRSCrotor ωωω =+ (3-5)

rotorω is the mechanical rotational speed of the rotor, statorω is the rotational speed of the stator field

defined by the grid frequency.

The rating of the converter is typically 20-30% of the nominal generator power and depends on the

desired speed range [61]. The converter only handles the rotor power rotorP , whereas the stator power is

directly passed to the grid. The following formulas hold

rotorstatorgen PPP += (3-6)

genrotor

statorstator PP

ω

ω= (3-7)

genrotor

statorrotorrotor PP

ω

ωω −= (3-8)

25

The sizing of the converter is based on economic considerations. A higher speed range (higher machine

slip) enables a higher energy output, but also needs a more expensive converter. A trade-off is normally

found for a converter able to vary the rotor speed about +30% of the synchronous speed [61].

The RSC controls the reactive power to and from the rotor. The Stator Side Converter (SSC) controls the

DC voltage level on the link between the RSC and SSC. Furthermore the SSC can deliver and control the

reactive power to the grid [62]. For unity power factor operation, the SSC supplies the same amount of

reactive power as the stator absorbs. Reactive compensation is therefore unnecessary in the turbine.

Compared to the FSIG topology with squirrel-cage induction generator, the rotor has an external feed

here. Slip rings are necessary to connect the RSC to the rotor. The necessary maintenance of those slip

rings poses a serious drawback on this topology especially for offshore wind turbines. As can be seen in

Figure 3.5, because of the direct connection between the stator and the grid, a gearbox is needed in this

topology for the same reasons as in the FSIG topology. This gearbox also needs important maintenance

during the lifetime of the wind turbine.

Induction machines with higher numbers of pole pairs have been investigated (e.g. p = 40) [64]. The

problem of poor power factor is partially solved by the reactive power control of the converter. The

gearbox is reduced from a three-stage type to a single-stage gearbox. The topology was compared with

others (DDSG, DFIG) in [64] and is considered as a promising option for the future.

3.3.2.2 Direct-Drive Synchronous Generators

Gearboxes are expensive components in wind turbines. For offshore wind turbines they account for 11%

of the turbine investment cost, whereas for onshore turbines their share can be up to 20% [46],[54].

Furthermore, their need for maintenance represents a high annual cost, especially for offshore wind farms

[64]. Next to that, their failure rate is not zero, and they bring down the reliability of the wind turbines

[65]. In order to avoid the presence of a gearbox in the nacelle, wind turbine manufacturers propose a

drive train based on a synchronous generator connected to the grid through a full power frequency

converter. Although DFIG based topologies are still very popular, the trends for new wind turbines more

and more concentrate on this configuration as power electronics are becoming cheaper. The topology is

shown in Figure 3.6.

Figure 3.6 Direct-Drive Synchronous Generator wind turbine topology

26

Synchronous generators rotate synchronously with the frequency applied to them. The relation between

the frequency applied to the generator and the synchronous rotational speed is given by

60

ssynchronougrid

npf

⋅= . (3-9)

p is the number of pole pairs of the generator, n represents the rotational speed of the machine in rpm.

The Voltage Source Converter decouples the synchronous generator from the grid. The frequency applied

to the generator can therefore differ from the grid frequency and be variable as well. The turbine blades

are directly coupled to the generator without gearbox; the generator rotates at the same rotational speed as

the blades.

The wind speed defines an optimal rotational speed for the turbine blades, based on the coefficient of

performance of the blades. Typical rotational speeds for 5 MW wind turbines are around 10 rpm

(6,9..12,1 rpm). Therefore a high number of pole pairs is needed to bring the frequency to workable

values without using a mechanical gearbox. A value of p = 60 is possible but already high. This means

that the applied frequency to the generator is in the order of 10 Hz. The Generator Side Converter

(GenSC) produces the appropriate voltage waveform to tune the generator to the optimal blade speed. The

Grid Side Converter (GridSC) controls the DC voltage on the link between the two converters and

controls the power factor of the total system (Reactive Power Control). The total active power is passed

through the converters from the generator to the grid.

A synchronous generator needs a DC excitation on the rotor in order to generate electricity. Two types of

excitation are possible. An external excitation system can be used, but this consumes power and decreases

the overall efficiency of the wind turbine by a few percent. By controlling the excitation system, the

reactive power output of the generator, and thus the output voltage, can be controlled. This is however a

small advantage because the reactive power output is controlled by the power electronic converter and

does not depend on the generator. The external excitation system is either fit in the rotor or connected via

slip rings. This type of generators is voluminous and has not been installed offshore so far.

Another option is to use permanent magnets as excitation system. Permanent Magnet Synchronous

Generators are known for their high efficiency and highly reliable power generation [66],[67]. In addition,

the high-power-density PMSGs are smaller in size, which reduces the voluminous generator at least

partially. A disadvantage of PMSGs is their high cost because of the large amounts of expensive

permanent magnet material. They are promising for the future as rare earth permanent magnet materials

are expected to drop in cost [68]. Compared to the asynchronous generators discussed before, PMSGs are

still large and therefore difficult to install offshore.

The major advantage of the direct-drive topology is the absence of a gearbox. No maintenance on the

gearbox has to be performed and an important cost factor in the wind turbine is omitted in this way.

Nevertheless, the necessary full power converter is an extra cost component and the high number of pole

27

pairs needed for the operation of this topology makes the generator expensive, voluminous and heavy

[64].

As a compromise, a synchronous generator with a smaller number of pole pairs has been proposed. A

single-stage gearbox is put between the turbine blades and the generator, instead of a 3-stage gearbox in

conventional induction generator configurations [69]. This system is known under the name Multibrid®

system (WinWind, Finland) [64].

3.4 Combination of traditional wind turbine topologies with VSC HVDC

The technical possibilities of VSC HVDC were discussed in Chapter 2. The most relevant characteristic

for the connection of offshore wind farms is the decoupling between the offshore grid and the onshore

transmission grid. This decoupling makes it possible to operate the wind farm at a different, even

variable, frequency and voltage amplitude. The black start capability needed to start up the wind farm is

delivered by the HVDC link itself. The reactive power is controlled by both converters at the ends of the

cable. Only active power is transported over the link.

In this section, the technical advantages this technology brings to an offshore wind farm are treated. The

topologies from section 3.3 are therefore reviewed in the following paragraphs in the context of an

offshore wind farm connected to the shore using VSC HVDC. Only a qualitative discussion is held in this

chapter.

3.4.1 Squirrel Cage Induction Generators

The topology of fixed-speed induction generators for wind turbines is often considered as first generation

technology. Manufacturers and wind farm developers increasingly leave this configuration because of the

disadvantages stated in paragraph 3.3.1 [70]. For offshore wind farms connected with VSC HVDC

however, some of these disadvantages do not hold anymore and some of the advantages have a higher

impact. It is therefore important to review this topology for offshore use with VSC HVDC, and not to

dismiss it right away.

First of all, the low cost of the topology is a major advantage over other topologies anytime. The limited

number of components needed for operation is an even more important advantage for offshore use. The

generator is limited in size and needs no maintenance compared to DFIG topologies. No converter is

needed compared to other variable speed topologies. This reduces the installation time and cost for

offshore wind turbines and keeps maintenance cost lower than for the DFIG topology during lifetime. The

problems mentioned in 3.3.1 for a wind farm with turbines of this topology are mainly solved when it is

connected to the onshore grid with VSC HVDC. The two most important drawbacks were the lack of

variable speed operation and the high reactive power requirements. The VSC HVDC link decouples the

offshore grid from the onshore grid and can, at least partially, take over the role of the individual

converters in variable-speed topologies and solve these problems. The absence of converters in the wind

turbines also reduces the losses in the individual turbines.

28

3.4.1.1 Variable speed operation with VSC HVDC

It is possible to control the AC voltage and frequency of the wind farm with the offshore Voltage Source

Converter. By varying the frequency of the wind farm according to the present wind speed conditions, the

synchronous rotational speed of all the turbines in the wind farm is adjusted. All turbines will have an

equal synchronous rotational speed but their individual rotational speed might differ slightly due to

variations in slip. The control of this speed can not be optimized from turbine to turbine as in classic

variable-speed topologies and a group optimum has to be found for the whole wind farm [61]. Neither can

the control be as fast as in variable-speed topologies because of the different conditions at each individual

turbine. Nevertheless, an important increase in energy output can be expected compared with HVAC

fixed-speed operation [70].

A strategy could be to use one of the ‘central’ turbines in the wind farm as a master turbine. The

frequency of the total wind farm is set optimal to the average wind speed conditions at this master turbine.

The other slave turbines accept the same synchronous speed. Wind speeds might differ throughout the

wind farm but they are nevertheless highly correlated and a spatial ‘memory effect’ exists due to the

propagation of wind gusts through the wind farm. The spatial variation of wind speeds was studied by

Nørgaard and Holttinen [71]. The various wind speeds at the individual turbine units will at any specific

time be distributed around the average wind speed of the wind farm (here assumed to be the wind speed at

the master turbine). As an approximation, the distribution is suggested to be a normal distribution. The

normalized standard deviation depends on two parameters: the spatial dimension D of the wind farm (the

distance from each turbine I to the master turbine) and the wind turbulence intensity Iturb.

The turbulence intensity can be empirically verified for a wind farm project. For offshore wind farms, the

turbulence intensity is small because of the extreme flatness of the surrounding waters. A value of Iturb =

10% is used to determine the normalized standard deviation in [71]. The distribution of wind speeds for a

turbine at 5 km, 15 km and 30 km from the master turbine is shown in Figure 3.7. A typical spacing

between individual turbines in a wind farm is 5-7 times the rotor diameter [74]. For a 5 MW wind turbine,

a typical rotor diameter is 126 m [72] leading to a turbine spacing of 630-882 m. According to the total

rating of the wind farm the dimensional parameter D will vary, but a large wind farm can already be

installed with a value equal to D = 5 km. The wind speed variation is then rather small.

The wind speed offset shown in Figure 3.7 is the offset of the average wind speed at turbine I compared

to the average wind speed at the master turbine. Short term wind speed variations, so called wind gusts,

are not taken into account. The overall wind farm frequency can not respond to these short term

variations. However, even with individual variable speed control per turbine, the response of the

rotational speed of the blades to wind gusts is limited. Due to the wind speed offset at the individual

turbines, the rotational speed is not optimal at each wind turbine. This will be discussed in more detail in

the next chapter, when the energy output of different wind farm topologies for a 300 MW wind farm is

compared.

29

Figure 3.7 Probability density function of wind speed offset with distance from master turbine D as

a parameter

3.4.1.2 Frequency range in the wind farm

From a VSC HVDC point of view, the minimum and maximum frequency of the wind farm can be

chosen in a relatively wide range. As such, the VSC HVDC link is not the limiting factor. Mainly, the

wind farm consists of iron components: transformers and generators. Those components are designed to

allow for a certain flux. When the flux in iron machines becomes too high, saturation will occur which

should be avoided by any means. The flux amplitude φ̂ in a transformer leg for example is determined by

the following formula

nf

U

grid ⋅⋅=

44,4φ̂ (3-10)

where U is the rms voltage, fgrid represents the frequency and n the number of turns on the transformer leg.

The flux in an induction machine is determined by the same f

U-ratio. In order to avoid saturation, this

ratio should be kept constant. Transformers are generally designed to operate at a frequency equal to 50

Hz in Europe. In a variable frequency wind farm they are operated at variable frequency. Saturation is

only avoided when the voltage is varied accordingly to the frequency.

When lowering the voltage with the frequency, the current through the transformers and generators will

increase proportionally. The low frequencies are used at low power output of the wind farm and the

currents remain within the ratings of the equipment. The maximum frequency can not be chosen too high

for two reasons. The gearboxes in the wind turbines would become very expensive because of the high

30

speed ratio between blades and generator synchronous speed. A second reason is the limited maximum

speed of the generators when increasing the frequency. To maintain the same f

U-ratio, the voltage is

raised as well. Nevertheless, the voltage ratings of the generator limit the possible increase in voltage

amplitude. Field weakening effects then limit the resulting speed range [61].

A demonstration project was installed at Tjaereborg in Denmark [73]. The installed VSC HVDC is able to

vary the frequency in a range between 30 and 65 Hz. The voltage has to vary accordingly. The frequency

in an 8 MVA wind farm was successfully varied between 30 and 50 Hz. This is enough to take the most

important energy output gain with variable speed operation.

3.4.1.3 Reactive power requirements

The specific reactive power requirements of the wind farm based on directly coupled induction generators

can be fulfilled by the offshore VSC HVDC converter. The fast control of reactive power makes it

possible for VSC HVDC to operate a wind farm without any additional expensive components. The

offshore converter station behaves as a strong grid node for the induction generators and guarantees the

voltage stability in the offshore grid. The sharp variations in reactive power demand lead to fast voltage

variations, also known as flicker, when using HVAC to connect to the onshore grid. No reactive power is

transported over a HVDC link and the reactive power variations are not present in the onshore grid and

the onshore grid is not subject to flicker.

3.4.1.4 Disadvantages of this topology

The rotational speed of the wind turbines is still controlled by the grid and not by the wind speed at each

turbine. This yields higher mechanical stresses on the drive train components [70]. The gearbox is still

present to convert the low speed of the blades to a higher speed for the generator. The wind farm voltage

is decreased during periods of lower wind speed. This will cause higher currents than in the case with

fixed voltage. This approach is therefore assumed to have a negative effect on the losses in the collection

grid of the offshore wind farm.

3.4.2 Doubly-Fed Induction Generators

Doubly-Fed Induction Generators were discussed in section 3.3.2.1. This is currently the most popular

topology to implement variable speed operation [61]. A 5 MW topology is planned to be installed on the

Belgian North Sea in a 300 MW wind farm on the Thornton Bank [74].

3.4.2.1 Reduction of rotor power or extension of speed range

It was made clear in section 3.3.2.1 that the rating of the converter to connect the rotor was defined by the

desired speed range of the turbine blades. It is advantageous to use the variable frequency as discussed for

squirrel cage induction generators in this topology as well. The offshore grid frequency is varied in the

same way to a group optimum for all the wind turbines. The individual converters deliver individual wind

31

turbine speed adjustment. The needed machine slip is limited in this way and converters of smaller ratings

can be used. Specific control strategies can be used to limit maximum or mean slip of the connected wind

turbine generators [61]. Another possibility is to enlarge the speed variation range with the same

converter ratings and capture more energy by optimizing the tracking characteristic. The reduction of the

converters decreases the investment costs of the wind farm. An economic optimum can be found

depending on the number of turbines in the wind farm, the power rating of the individual units and the

wind conditions [61].

3.4.2.2 Disadvantages of this topology

The direct connection of the stator to the grid makes the use of a gearbox unavoidable. The slip rings on

the generator increase the generator cost and need maintenance as well. The VSC HVDC link takes over

most of the variable speed control of the turbines and it should be investigated if the extra investment in a

more complicated generator, annual slip ring maintenance and a, albeit small, converter are worth the

limited advantages they bring compared to the topology with directly connected generators.

3.4.3 Direct-Drive Synchronous Generators

The generators in this topology are separated from the grid by a full frequency converter. The frequency

of the offshore grid does not have an influence on the rotational speed of the generators and the turbine

blades. The need for frequency variation according to the wind conditions is thus less clear. On the other

side, the frequency of the offshore grid is decoupled from the onshore grid by the VSC HVDC link. A

fixed frequency can be chosen for the offshore wind farm. This frequency can differ from the onshore 50

Hz in Europe. The choice of the optimal frequency is a complicated problem and it is not the purpose of

this thesis to go to deep into details. Only some suggestions and remarks for future investigation are

made.

The only equipment present in the AC offshore grid (bordered by the converter of the VSC HVDC and

the converters of the wind turbines) are transformers and cables. The frequency should be chosen in order

to optimize the total lifetime cost of these components. The following remarks can be made:

− The volume (amount of iron) and cost of transformers depends strongly on the frequency. A higher

frequency downsizes the transformers due to the lower flux. The choice for a higher frequency can

thus reduce the investment and installation costs of the wind farm.

− The losses of transformers increase with increasing frequency. Detailed loss calculations for variable

frequency transformers are treated in [75].

− The charging currents in the cables of the wind farm collection grid depend linearly on the frequency.

Higher charging currents involve higher losses.

As an example, the use of 60 Hz (standard in the US) is put upfront. The amount of iron in the

transformers is reduced by 17%. A high frequency wind farm was proposed by Meier, Norrga and Nee

32

[76]. The frequency in the wind farm of [76] is chosen to be 500 Hz.

3.5 Conclusion

The trends in the offshore wind industry are discussed in the first paragraphs of this chapter. Future wind

farms are found to have a power rating higher than 200 MW and to be situated several tens of kilometres

from shore. The use of a distinct transmission system is then unavoidable.

The different wind turbine topologies were discussed in this chapter as well. Their advantages and

disadvantages were explained especially for offshore use. The topology with directly connected squirrel

cage induction generators (SCIG) is still under consideration in combination with VSC HVDC, whereas it

is found unfeasible in combination with HVAC. The DFIG and direct-drive topologies are possible with

both VSC HVDC and HVAC. A synchronous generator with permanent magnets as excitation is

preferable for offshore use in direct-drive topologies. A hybrid topology which compromises for

generator and gearbox is still under consideration. The most important characteristics of the different

topologies are summarized in Table 3-2. The characteristics are shown for three main components in the

drive train of the topologies: generator, gearbox and converter. The other components of the turbine units

are considered to be the same in each topology.

SCIG DFIG DDPMSG GPMSG

Generator Complexity Weight/Cost

Simple Low

Brushed Medium

Very complex Very high

Complex High

Gearbox Complexity Weight/Cost

3-stage High

3-stage High

No gearbox None

Single-stage Medium

Converter Rating

None

~30% of Pnom

Pnom

Pnom

Table 3-2 Characteristics of studied wind turbine topologies

In order to come to a more quantitative analysis in this thesis a typical wind farm is put upfront to

investigate. The power rating of the wind farm is chosen to be 300 MW and the length of the transmission

link is chosen at 50 km. The individual turbines are rated at 5 MW each.

33

4 ENERGY OUTPUT OF OFFSHORE WIND TURBINES

4.1 Introduction

Opposed to conventional electricity generation units, the output of a wind farm depends on an

uncontrollable input: the wind. As seen in Figure 3.4, a wind turbine’s power output is wind speed

dependent. This means that the output power of an offshore wind farm will also be wind speed dependent.

The electric power transported over the transmission cable will thus be variable in time as well. It is

useful for this reason to have a look at the wind speed distribution over time for a typical offshore wind

farm. Keeping the power speed curve of a wind turbine in mind, it is possible to calculate a power

probability density function (4.2).

The energy output of a wind turbine depends on several parameters. The use of variable speed operation

results in an increased energy output (4.3). The wider the speed range, the more energy can be captured

from the wind, especially at low wind speeds. A multi-turbine speed control was proposed for the SCIG

topology in combination with a VSC HVDC link. The effect on the annual energy output of the multi-

turbine frequency approach is investigated in section 4.4. Furthermore, the studied topologies have

differences in the efficiencies of their drive train components (Table 3-2). The efficiency of each topology

is discussed in section 4.5. The energy yield of each topology is calculated.

4.2 Power probability density function

Wind speeds are distributed in time and commonly described by a Weibull distribution. A Weibull

probability density function is defined by two parameters, which define the shape I and scale (A) of the

distribution. The probability density function is given by

( )

Cwind

A

vC

windCAwind e

A

v

A

Cvpdf

��

���

�−

−

��

���

���

���

�=

1

, . (4-1)

For offshore wind farms, the two parameters are calculated out of a long time series of wind speed

measurements at sea. The values used in this thesis are A = 9,8 m/s and C = 2,1. These values are

calculated out of measurements at the Euro Plat Form (EPF) in the southern North Sea (see Figure 4.1)

[77] and are applicable values for an offshore wind farm about 50 km from shore. The resulting

probability and cumulative density functions are shown in Figure 4.2.

For each wind speed value exists an according wind turbine power output as shown in Figure 3.4.

Combining the power-speed curve with the wind speed probability function defines a probability function

for the power output. Several operational possibilities exist. When the wind speed is beneath the cut-in

wind speed of the wind turbine, the power output is 0 MW. Between the cut-in wind speed and the

nominal wind speed, the wind turbine operates below nominal output power. At the nominal wind speed,

the power output reaches the nominal output power and maintains this value due to pitch angle regulation

34

until the cut-off wind speed is reached. Above the cut-off wind speed, the wind turbine is stopped because

of safety reasons.

.

Figure 4.1 Situation of Euro Plat Form measurements in the North Sea [77]

Figure 4.2 Probability density function and cumulative density function for wind speeds

in offshore wind farm (Average wind speed = 8,68 m/s)

The speed range of the turbine blades influences the power-speed curve for each topology at low wind

speeds. The effect is discussed in Appendix B. The annual energy yield difference between the topologies

under study is small and the effect is not taken into account further. The power-speed curve of a DFIG

turbine is assumed to be valid for both the SCIG and direct-drive topologies as well. The minimum

rotational speed of the turbine blades is 6,9 rpm and the maximum is set at 12,1 rpm [72]. The resulting

power probability density function is shown in Figure 4.3 and compared with a fixed-speed turbine

topology. Fixed-speed turbines are assumed to operate only at nominal rotational speed of 12,1 rpm.

The fixed-speed wind farm is discussed first. For more than 18% of the time, the wind farm is not

35

operational because of lack of wind (cut-in wind speed = 4,33 m/s). For about 25% of the time, the wind

farm operates at nominal output power (nominal wind speed = 11,5 m/s to cut-out wind speed = 30 m/s).

The remaining 57% of the time is distributed over the power range with a higher probability in lower

power zones due to the cubic relation of power to wind speed.

For a variable-speed wind farm, a serious increase in operation time and energy output of the wind farm

can be seen. The turbines are connected to the grid when the blades reach a rotational speed of 6,9 rpm

and can operate in a continuous speed range up to 12,1 rpm [74]. The time the wind farm is not

operational due to lack of wind is reduced to 8%. This is because variable-speed operation allows wind

turbines to capture energy from the wind at lower wind speeds (cut-in wind speed = 3 m/s). The extra

time of operation is distributed over the different power zones. The nominal wind speed is lowered to

11,45 m/s. The cut-out wind speed is set at 30 m/s as well.

0%

5%

10%

15%

20%

25%

30%

0% 0-10% 10-20% 20-30% 30-40% 40-50% 50-60% 60-70% 70-80% 80-90% 90-100% 100%

Power output P_mech [% of P_nom]

Pro

bab

ilit

y [

%]

Fixed-speed

Variable-speed

Figure 4.3 Power probability density function for offshore wind farm

(based on matlab wind turbine model)

Figure 4.3 is important for the transmission cable between the wind farm and the onshore grid. The wind

farm power probability is the probability of the transported power through the cable. The active power is

proportional to the captured wind power. The reactive power depends on the power factor of the

generating units, the type of cable and the action of compensators at the endings of the cable. The

probability of the active power is used in the next chapter to calculate the total current and the related

losses in the transmission system.

4.3 Energy output of wind turbines

Given the wind conditions discussed in section 4.2 (Figure 4.2), it is possible to calculate the expected

annual energy yield of a wind turbine. The annual energy output can be found by multiplying the Weibull

probability distribution function of Figure 4.2 with the power-speed curve of Figure 3.4 and integrate this

function over a whole year. The results of this calculation are shown in Table 4-1. The numbers in Table

4-1 only show the energy taken at the blades of the turbines. The efficiency of the drive train and the

losses in the transmission system to shore are not taken into account so far because they depend on the

36

topology.

Turbine Energy

output (5MW)

Wind farm Energy

output (300 MW)

Capacity

Factor

Fixed-speed turbines 18 962 MWh 1 137 697 MWh 43,29%

Variable-speed turbines 20 516 MWh 1 231 020 MWh 46,84% Table 4-1 Energy output comparison between fixed- and variable-speed wind turbines

The energy output of a wind farm with variable speed operation is 8,20% higher than a wind farm with

fixed speed operation. Other authors state net energy gains of 9-18% with variable-speed wind turbines,

e.g. [59],[60]. The difference between the calculated gain with the model and the stated gains by other

authors is due to the use of offshore wind data, whereas [59]-[60] use onshore wind data. For offshore

wind turbines, the average wind speed is higher and the largest energy gains that are made at the lowest

wind speeds are less important in the total energy output. This is discussed in more detail in Appendix C.

4.4 Multi turbine frequency approach

The proposed topology with VSC HVDC and SCIGs uses a common rotational speed control for all the

turbines, whereas the other topologies have individual speed control (3.4.1.1). The effect of this approach

on the annual energy yield of the wind farm is investigated in this section. The approach with one master

turbine in the wind farm and the rest being slave turbines will be used. The rotational speed of the slave

turbines will not be optimal to their respective wind speeds. The optimal rotational speed as a function of

wind speed for a 5MW variable speed turbine is shown in Figure 4.4. The speed range is taken from the

Repower 5M topology (DFIG, 5MW, = 6,9..12,1 rad/s) [115].

The optimal rotational speed for each 5MW wind turbine, using this speed range, is shown in Figure 4.4.

With optimal rotational speed is meant, the speed which results in the highest coefficient of performance

and thus power output. The rotational speed of a slave turbine can only be not optimal (not equal to its

optimal value) in the wind speed range in which variable speed operation is used. It can be seen on Figure

4.4 that this range is situated between 7,24 m/s and 12,63 m/s. Although this speed range seems to be

small, the wind speeds in the proposed wind farm are within this range for 40% of the time (Figure 4.2).

The output power of each individual turbine is limited to the nominal output power by the pitch angle

regulation for wind speeds higher than the nominal wind speed of 11,45 m/s.

The probability of the wind speed offsets (difference between master and slave turbine speed) was shown

in Figure 3.7 and depends on the spatial distance between the slave turbines and the master turbine. The

optimal placement of the turbines to form a wind farm will not be discussed in this thesis but, given that

60 turbines of 5 MW are needed to form a 300 MW wind farm, it is fair to state that the maximum

distance between master and slave turbine will be around 5 km.

37

Figure 4.4 Optimal rotational speed of 5MW wind turbine as function of wind speed

(based on matlab wind turbine model)

As an example, a slave turbine at 5 km from the master turbine is now considered. The situation for a

master turbine wind speed of 9 m/s is shown in Figure 4.5. As an approximation, the wind speed offset

between slave and master is taken not to be higher than 0,5 m/s. As seen in Figure 3.7, this covers more

than 99% of the situations. The offset of the rotational speed of the wind turbine compared to its optimal

value is shown in blue in Figure 4.5 and is taken from Figure 4.4. The use of a non-optimal rotational

speed results in a difference between the actual power taken from the wind and the theoretically

maximum power that can be taken at this wind speed if the turbine would rotate at its optimal speed. The

power not taken from the wind due to this non-optimal rotational speed is calculated for a 5 MW wind

turbine model and depicted in red. Taking into account the wind speed offset distribution of Figure 3.7 the

expected value of the power not taken from the wind at the slave turbine is 0,484 kW for a master turbine

wind speed of 9 m/s. The same reasoning can be carried out for every wind speed in the variable speed

range of the master turbine. The result is shown in Figure 4.6. This curve is aggregated with the wind

speed probability function of Figure 4.2 to calculate the annual energy loss due to the multi turbine

frequency approach. This is found to be 1,578 MWh for a slave turbine situated 5 km from the master

turbine. Compared to the expected annual energy yield of a variable speed turbine of 20 GWh, this loss is

negligible.

Knowing that the majority of the turbines are situated closer to the master turbine (< 5km) than the

turbine studied above, the total energy loss due to the multi turbine frequency approach can be neglected

for the wind farm considered in this thesis. The energy taken at the turbine blades is approximately equal

for each variable-speed topology under consideration.

38

Figure 4.5 Power not taken from wind due to multi turbine speed control

(turbine at 5km from master turbine; wind speed at master turbine = 9 m/s)

Figure 4.6 Expected value of power not taken from the wind at each wind speed in the variable

speed zone due to multi turbine speed control (turbine at 5km from master turbine)

39

4.5 Drive-train efficiency

The different wind turbine topologies under consideration have each a specific drive train assembly. To

simplify the loss analysis, only the efficiencies of the three main components are considered: the gearbox,

the generator and the converter. The loss data for these components were found for 3 MW topologies in

[64]. All components have lower efficiencies when not loaded at their nominal ratings, which is often the

case for wind turbines. The data are therefore average values for wind turbines. The data from [64] are

extrapolated for the 5 MW wind turbines used in this thesis.

SCIG DFIG DDPMSG GPMSG

Generator 98,4% 98,0% 95,9% 97,9% Gearbox 93,8% 93,8% --- 96,9% Converter --- 98,33% 95,4% 95,4% TOTAL 92,3% 90,4% 91,4% 90,4%

Table 4-2 Efficiencies of different components of drive train [64]

The data given in Table 4-2 show that the SCIG topology has the most efficient drive train. This is mainly

because of the simple, efficient induction generator and the absence of a converter. The connection of the

rotor via slip rings makes the induction generator in the DFIG topology less efficient. The DDPMSG has

the lowest efficiency mainly due to its large size (high number of pole pairs) [35].

It is clear from Table 4-2 that the gearbox, if present, is an important loss component in the drive train.

The converter efficiency of the DFIG topology is given for the nominal power rating of the turbine. The

converter in a DFIG topology processes around 30% of the total power. This explains the differences with

the converters in the DDPMSG and GPMSG topologies.

The efficiencies are used in Chapter 6 to compare the different topologies on an economic basis. The

given drive-train efficiencies are combined with the numbers given in Table 4-1 for a variable speed

turbine. The resulting annual energy output and capacity factors are shown in Table 4-3.

Turbine Energy

output

Capacity

Factor

SCIG 18 935 MWh 43,23%

DFIG 18 541 MWh 42,33%

DDPMSG 18 761 MWh 42,84%

GPMSG 18 549 MWh 42,35% Table 4-3 Annual energy output and capacity factor

for different variable speed wind turbine topologies

40

4.6 Conclusion

The output of wind turbines depends on the wind speed. Wind speeds are Weibull distributed in time.

This results in a distributed output power for a wind farm. The distribution of this output power was

calculated in this chapter for an offshore wind farm situated 50 km from shore. This distribution will be

used in the next chapter to calculate the transmission cable losses for a 300 MW offshore wind farm.

Several wind farm topologies were proposed in Chapter 3. One of the topologies uses VSC HVDC in

combination with directly connected induction generators. Variable speed operation is achieved with

variable frequency operation of the offshore wind farm collection grid. This involves a multi turbine

frequency approach which results in one common rotational speed for all the turbines. The spatial

distribution of wind speeds throughout the wind farm will result in a non-optimal rotational speed. The

effect of this approach on the annual energy yield was investigated in this chapter and found to be

negligible.

The drive-train efficiencies of the different wind turbine topologies were given in this chapter. This

resulted in capacity factors for the different topologies between 42,33% and 43,23%.

41

5 DESIGN AND OPERATIONAL ASPECTS OF VSC HVDC AND HVAC

5.1 Introduction

A wind farm rated at 300 MW and situated 50 km from the Point of Common Coupling (PCC) is

investigated in this thesis. Two possible transmission technologies are compared for the connection of the

wind farm to the onshore grid: VSC HVDC and HVAC. Both transmission systems are dimensioned to

the needs of the proposed wind farm in this chapter. An absolute requirement for the transmission system

is black start capability to start the generators in the wind farm.

Wind farms built nowadays need to fulfill grid code requirements. Those grid codes incorporate power

factor control, frequency response and voltage dip ride-through capability. A part of this chapter treats the

compliance of the wind farm to the latest grid code requirements in countries with a high wind energy

penetration in the electricity portfolio (e.g. Germany and UK). For HVAC as transmission technology,

additional measures have to be taken to fulfill those grid code requirements.

The losses of both transmission systems are compared in the last part of this chapter. The power

probability function calculated in Chapter 4 is used to calculate the annual energy loss for a 300 MW

offshore wind farm for both transmission technologies.

5.2 Dimensioning of the transmission system for an offshore wind farm

A transmission system is designed according to the power it is expected to transport. The apparent power

output of a wind farm is given by

22wfwfwf QPS += . (5-1)

The active power Pwf of a wind farm depends on the wind speed and a nominal value of 300 MW is taken

in this thesis. The reactive power output of the wind farm depends on the power factor of the wind farm

and is calculated as

PF

PFPPPQ wfwfwfwf

21cossin

tan−

===φ

φφ . (5-2)

The wind farms considered with HVAC have power factor controlling converters in their topologies. The

power factor (PF) is kept close to unity. The wind farms considered with VSC HVDC can either consist

of directly connected generators or generators connected via a converter. For a wind farm based on

directly connected asynchronous machines, a typical PF in nominal operation is 0,89 [55]. This leads to a

reactive power requirement of 153,69 MVAr (lagging) during nominal operation. In order to keep the

voltage at its nominal level in the offshore grid, the reactive power will be delivered by the offshore

converter. This results in an apparent power rating of 337,08 MVA for the offshore converter. For the

converter-based wind turbine topologies, the power factor will be kept close to unity when connected to

42

the grid with VSC HVDC.

Grid code requirements on power factor control are stated in countries with a high portion of wind power.

A more thorough discussion on grid code requirements will be given later in this chapter but some

indicative figures are necessary here. The power factor onshore must for example be controllable down to

0,95 [61]. This results in an apparent power rating of 315,79 MVA for the onshore cable end bus.

The calculated power ratings will now be used to dimension the transmission cable for both options: VSC

HVDC and HVAC. Both systems are assumed to be coupled to an onshore grid node at 400 kV (Point of

Common Coupling) and an offshore wind farm node at 33 kV.

5.2.1 VSC HVDC

ABB offers several ratings of VSC HVDC links. In order to come to a more standardized approach, the

possibilities are ordered in a matrix. The matrix is shown in Table 5-1 [22].

Table 5-1 HVDC Light® matrix

Nine different modules are proposed, numbered M1 to M9. The possibility exists to choose between 3

levels of DC voltage and 3 current ratings. A choice for one of the topologies is decisive for a lot of

parameters because of the standardized production of these modules. A first constraint is the MVA rating

of the cable. Only active power is transmitted over the DC cable but the converter stations handle the

reactive power as well. Therefore the above proposed apparent power is used to choose a module. It is

best practice not to overrate the chosen module too much, because the operational features depreciate

significantly when a module is operated below nominal conditions.

A good choice for the proposed wind farm (300 MW) is the M5 module of Table 5-1 which has a base

power of 376 MVA. For a 50 km cable, the output power at the receiving end is expected to be 361 MVA

due to the losses in the converter stations and the cables. This option operates at a nominal current of

1140 A and uses 2 DC cables (+ 150 kV). The maximal DC current rating is 1233 A [22]. The cross

section of the copper conductor is 1200 mm2.

The voltage on the AC side of the converter is 195 kV and a transformer is needed to transform the

collection grid voltage of the wind farm to this value. A transformer is included in the converter station

onshore as well, to connect the cable system to the PCC at 400 kV.

Although the operational experience with VSC HVDC is limited (around 10 finished projects), M5 is one

of the more popular modules and experience was gained with this module in successful projects, e.g. in

USA (Cross Sound Cable [18]), Australia (Murray Link [17]) and Estonia (Estlink [30]). Nevertheless, a

converter of this rating has not been built on an offshore platform yet. The only experience with offshore

43

converters gained so far is in the Troll A project in Norway [32]. This project has a power rating of 2 x 40

MW and proved to be very successful.

5.2.2 HVAC

The design of a HVAC cable is less standardized and leaves more variables to decide on than a VSC

HVDC system. The design is performed based on choices made in existing or planned wind farms and on

logical reasoning. The data used, are provided by ABB in [78].

A first important choice to make is the choice between 3 single-core cables or a 3-core cable. The system

with 3 single-core cables has one cable per phase whereas a 3-core cable combines the 3 phases in one

cable. A picture of a 3-core cable and a single-core cable is shown in Figure 5.1 (left). The advantage of 3

single-core cables is that a higher current rating is achieved with the same amount of copper conductor.

The amount of copper is an important part of the price of the cable. The construction of the cables is also

simpler than for a 3-core cable. Single-core cables are thus less expensive to purchase than 3-core cables

with the same power rating. Nevertheless, more cables have to be laid and to keep the system

symmetrical, transposition of the cables is needed after certain intervals. With transposition is meant that

the single-phase cables have to change in relative position to keep the system balanced. The screens of the

cables have to be cross-bonded as well to eliminate sheath circulating currents [79]. The cost of

transposition is high, especially at sea where it is a real technical challenge and increases the installation

cost of the cables. For long submarine cables it is less expensive to put more 3-core cables than to bother

about transposition of single-phase cables. Therefore it is chosen to use the 3-core cable. This is a

common choice for submarine transmission systems and is used for example for the connection of the

Belgian wind farm on the Thornton Bank [74].

Figure 5.1 XLPE 3-core cable and single phase cable (left) and reinforced submarine XLPE 3-core

cable (right) (Courtesy of ABB)

Another choice to make is the insulation material of the cables. More and more manufacturers advise to

use Cross-Linked Poly Ethylene insulated cables (XLPE). They have a high level of performance and are

suitable for submarine application. The shunt capacitance is lower than for oil-impregnated cables used

before and they are considered to be more environmentally friendly [80]. The armor is made of steel,

44

whereas for land cables non-magnetic materials are used. This is done to give the cables extra strength in

the sea (e.g. impact of ship anchors). The reinforced submarine version of the 3-core HVAC cable is

shown in Figure 5.1 as well (right).

The conductor material is chosen to be copper for its high conductivity (lower losses) and to reduce the

number of cables needed to transport the electric energy. Alternatively, aluminum cables would be less

expensive to purchase but more difficult and costly to install. The trade-off between copper and

aluminum is strongly dependent on the market price of these materials.

The maximum power transported depends on the voltage rating and the current rating of the cables. The

voltage level can be chosen up to 245 kV for 3-core cables. No joints have been constructed for extruded

AC cables for higher voltages than 245 kV [50]. The choice of the transmission voltage is a trade-off

between losses and cable cost. The choice for a higher voltage leads to lower currents in the cable for the

same transported power and thus lower losses. The reactive charging current produced by the cable is

voltage dependent as well and increases with the voltage magnitude. A higher charging current increases

the losses of the cable system. The insulation requirements become higher for higher voltages and the

cable is more expensive. The transformer in the wind farm is also larger and more expensive if a higher

voltage step has to be performed. It has to be mentioned that if the PCC node onshore is at a reasonable

voltage (e.g. 150 kV) for the submarine transmission, this node voltage might be decisive for the cable

voltage and facilitate the voltage choice. The onshore transformer can be omitted in this way. The voltage

levels used in the onshore grid (and available in coastal regions) are country dependent and thus vary

from case to case. The PCC voltage is assumed to be 400 kV in this thesis and a transformer is necessary

onshore. The 245 kV cable is claimed by ABB but higher voltages than 150 kV have not been used so far

in practical submarine installations [81]. The chosen line voltage is therefore 150 kV. This is a standard

choice for submarine cables of this rating in Europe [62]. This voltage level is also used for the

connection of the wind farm on the Thornton Bank (300 MW) [74] and the Danish Horns Rev wind farm

(160 MW) [82].

After the voltage level is chosen, it is possible to decide on the conductor cross section according to the

current rating of the cable [78]. The nominal power of 300 MW gives rise to a maximum active current

through the cable system of

AU

PI

cable

wfactive 7,1154

3== . (5-3)

This can not be transported with one cable and at least two 3-core cables are needed in parallel. Two

cables with a conductor cross section of 500 mm2 per phase are chosen. The current rating is 655 A per

cable [78].

HVAC cables represent a shunt capacitance in the grid. This capacitance causes a reactive charging

current when the cable is operational. A capacitance of C = 0.17 F/km is reported in [78] for the chosen

cable. This capacitance exists between the phases and the ground. Therefore, the phase voltage is used

45

instead of the line voltage in the calculation of the charging current

kmArkmFkV

HzCU

fI cablecable

gridC /6,4/17,03

150502

32 =⋅⋅⋅⋅=⋅⋅⋅⋅= µππ (5-4)

A charging current of 4,5 Ar/km is reported after measurement in [78]. A shunt capacitor, and as such

also a HVAC cable, is a producer of reactive power (leading power factor). In order to keep the power

factor within a reasonable range, this reactive power should be compensated. The charging current

becomes considerable for longer cables and reduces the current rating left for transport of active power.

The situation is shown in Figure 5.2 for the chosen cables with only compensation onshore as well as with

compensation at both cable ends. For a 300 MW wind farm with 2 parallel cables, the critical lengths are

67 and 136 km respectively.

Figure 5.2 Maximum transmission capacity per 3-core cable at 150 kV (cross section = 500 mm²)

The reactive power produced by the cable can be used to contribute to the Var generation needed to

realize the grid code requirements. Nevertheless, the reactive power of HVAC cables is not controllable

and extra reactive compensation equipment will be needed to achieve a controllable power factor at the

cable ends.

Next to the capacitance, the resistance and inductance of the cable play an important role for cable

calculations. The resistance is reported in [78] to be 0,0366 �/km (DC resistance). The inductance can be

calculated as [79]

kmmHr

sLcable /ln2,005,0 ⋅+= (5-5)

with s the distance between the conductor axes and r the conductor radius. The distance between the

conductors s is assumed to be 4r. This results in an inductance of 0,33 mH/km.

46

5.3 Black start of a wind farm

During startup, a wind farm behaves as a ‘load only’ network. A voltage waveform with appropriate

amplitude and frequency has to be applied to the wind farm in order to bring the generators online.

Depending on the chosen type of generators and wind turbine topologies, the wind farm has typical active

and reactive power requirements. A transmission system that can deliver these requirements without the

help of additional equipment is said to have black start capability in this context.

5.3.1 VSC HVDC

HVDC connections did not play a major role in the connection of wind farms so far. One of the reasons is

the black start capability needed to start up a wind farm. The required sinusoidal voltage waveform to

connect the generators has to be created offshore out of the DC voltage on the link. The traditional LCC

HVDC does not have this black start capability. LCC HVDC needs a voltage waveform for the

commutation of the thyristors in its converter [83] and additional equipment is required to bring the wind

farm to operation. Possible solutions are the installation of a STATCOM with black start capability [25]

or a diesel-driven generator [8]. The investment and operation costs of these additional components made

LCC HVDC too costly to implement for wind farm applications. Another reason is the considerable size

of the converter station and its lack of reactive power control.

With the introduction of VSC HVDC, black start capability became possible for DC transmission systems

as was explained in 2.4.3. When starting the wind farm as a black net, the DC link is energized from the

onshore converter station. The offshore converter station is energized on the DC side. When the power

electronic switches in the offshore station are deblocked, it is possible to determine both voltage

amplitude and frequency in the wind farm. The AC voltage can be smoothly ramped up by the VSC. In

this way, transient over-voltages and inrush currents in the wind farm cable system can be prevented [73].

This operation mode is called ‘passive net operation’ [84]. Because both AC voltage amplitude and

frequency are controlled by the offshore converter, the onshore converter has to control the DC voltage.

The onshore converter can control the reactive power from and to the onshore grid as well. The passive

net operation is maintained for several minutes before the wind turbines are connected, to assure the

stable operation of the wind farm grid.

When the wind farm voltage is considered constant, the wind turbine blades are allowed to speed up

under the influence of the wind. When they reach the required speed of operation, they are connected to

the AC wind farm grid. At the point where induction generators start to deliver power, they draw a

reactive current from the grid to magnetize the machine [58]. This reactive current can be several times

the nominal current of the machine. The reactive power is delivered by the offshore converter by

controlling the voltage difference over the phase reactor of the station (AC voltage control) in case of

directly connected induction generators. When the generator is connected in a DFIG or full frequency

converter configuration, the converter takes over the reactive power control of the generator.

The active power produced by the wind turbines is absorbed by the offshore converter and transported to

47

the grid. There is no difference for a wind turbine being connected to a strong AC system or to a VSC

HVDC converter station. The VSC HVDC offshore converter behaves as a slack bus in the offshore grid,

absorbing the produced active power of the wind turbines, hereby controlling the offshore frequency and

delivering the required reactive power to control the AC voltage [84].

5.3.2 HVAC

The AC transmission cable between the onshore grid and the wind farm is used to transmit the voltage

waveform to startup the wind farm. The voltage amplitude might differ offshore from the amplitude

onshore and there might be a phase shift between both voltages as well, but more or less the voltage

waveform offshore will be the same as onshore. A long HVAC cable can be approximated by the

concatenation of several equivalent sections. A single-phase equivalent section is shown in Figure 5.3.

Figure 5.3 Equivalent section of HVAC cable

To calculate the voltage waveform in the wind farm during black start, a concatenation of equivalent

sections of 1 km each is used. The node at the PCC onshore is brought at 150 kV by the strong grid

connected to this node. The PCC node is taken as a slack bus for the wind farm and the voltage angle is

set to 0˚. At the offshore end of the cable the current is 0A (open circuit) because the generators are not

connected yet. The voltage along the cable is calculated as

( ) ( )xIZxU

xU W ⋅+⋅⋅= γγ sinhcosh3

)( 11 (5-6)

The current in each point of the cable is given by

48

( ) ( )xZ

UxIxI

W

⋅+⋅= γγ sinh3

cosh)( 11 (5-7)

The transmission coefficient γ and the surge impedance WZ are given by

( ) 34,244 −+−=+= ejejXjXR CLγ km-1 (5-8)

66,772,44 jjX

jXRZ

C

LW −=

+= � (5-9)

To calculate the voltage amplitude at the offshore cable end (x = 50 km in this case), the current at the

onshore node is needed. Although the wind farm does not produce any power during the black start

procedure, this current is not equal to 0 A due to the charging current in the cable. For a cable without

compensation equipment (total charging current provided by onshore grid), this results in an offshore

voltage wave with amplitude 151 kV shifted 0,14˚ in phase angle compared to the onshore voltage wave.

Because of the higher voltage offshore in this situation, the reactive power produced by the cable is

pushed to the onshore grid. The current at the PCC node is calculated at 232,3 A and purely reactive. This

was expected because the offshore node behaves as an open-circuit during start-up and does not take any

power. This number corresponds with the value of 4,6 A/km of the charging current mentioned before.

Having all the reactive power produced by the cable available at the onshore node, is an unfeasible

situation. To achieve a better balanced cable system, reactive compensation will be added at both ends of

the cable to redistribute the produced reactive power of the cable. At the moment the generators are

connected to the grid, active power will be produced and transported through the link. The reactive power

balance of the AC cables varies with the transferred power. The low inductance between the ends of the

cable makes the solution sensitive to these variations. This will have repercussions on the voltage

amplitudes and phase angles at the ends of the cable. These values should be kept within acceptable

ranges with the help of reactive compensation. A further discussion on reactive compensation at the ends

of the HVAC cable will follow after the grid code requirements are discussed. It will be possible then to

dimension the reactive compensation (5.5). The current distribution in the HVAC cables will be discussed

together with the cable losses in section 5.6.

5.4 Reactive power requirements on the offshore node

5.4.1 Directly connected induction generators

As explained in section 3.3.1, a wind farm based on directly connected induction generators has a reactive

power demand over the whole range of operation. As mentioned before, during nominal operation a

typical power factor of 0,89 can be expected [55]. This represents a total amount of 153,69 MVAr as was

calculated in section 5.2.

During startup of induction generators, a very high reactive ‘inrush’ current is drawn from the grid. This

inrush current yields a lower momentary power factor during startup (close to 0) [85] and is a transient

49

effect. Due to the fact that not all wind turbines start operation at exactly the same time and the transient

nature of these currents, the power factor varies more gently for a large wind farm. It is assumed that the

highest reactive power demand exists during nominal operation of the wind turbines.

The SCIG wind farm topology is only considered with VSC HVDC in this thesis. The reactive power

requirement of the asynchronous generators puts a serious burden on the offshore converter station. At

nominal power output, the reactive power mounts up to 0,4 pu of the converter rated power for the chosen

M5 module. The PQ-diagram of the rectifier (offshore converter) is shown in Figure 5.4. The grey zone

represents the possible reactive power demands of the wind farm and is within the ratings of the chosen

VSC HVDC link at nominal voltage in the wind farm.

The PQ-diagram in Figure 5.4 shows the limits in active and reactive power of the offshore rectifier. To

deliver Q to a grid, the output voltage of the converter is raised above the grid voltage as was discussed in

section 2.3.2. The limit on Q imposed on the capacitive side is determined by the voltage limits of the

power electronic switches. In case the reactive capability of the offshore converter would not be enough

to deliver the reactive power demand of the wind farm, extra switchable capacitors can be added in each

wind turbine. This would increase the cost of the wind farm but lower the losses in the offshore converter

as well.

Figure 5.4 Capability chart of offshore VSC (Rectifier)

(grey zone = possible wind farm P,Q-demand)

5.4.2 Generators connected via a converter

Wind farms based on generators connected to the grid via a converter do not have explicit reactive power

demands. Their power factors are kept close to unity by the converters.

The reactive power control of VSC HVDC is less exploited for these topologies. Only small variations in

reactive power are needed at the offshore station in order to control the AC voltage in the wind farm.

HVAC cables are a source of reactive power when a voltage is put at their ends. Both parallel cables will

deliver part of this reactive power to the wind farm. Compensation will be needed in the offshore wind

50

farm as the cables themselves are an uncontrollable source of reactive power. The reactive compensation

is discussed together with the grid code requirements in the next paragraph.

5.5 Grid Code compliance at the Point of Common Coupling

As the total installed capacity of wind energy in the European grid increases, the reliable and secure

operation of the transmission network needs further consideration. In the early stages of wind energy

development, wind turbines especially influenced the distribution grids. With the introduction of large

wind farms, the focus has switched from small scale distributed generation to large scale, centralized

connection of wind farms in the grid. The transmission grids are therefore increasingly influenced. Grid

operators tend to keep the power factor around unity everywhere in the grid. Induction generators are a

sink of reactive power and bring down the system voltage. Therefore, in countries with a high penetration

of wind energy (e.g. Germany, Denmark, UK,…), transmission system operators have released more

stringent ‘grid codes’ concerning the connection of wind farms. The main objective is to limit the effects

of large wind power parks on network quality and stability [86]. From large wind farms, also offshore, it

is expected that they provide reactive power control, frequency response and voltage dip ride-through

capability [87]. Not all European grid operators have these grid codes for wind farms, but it can be

expected that they will impose them in the upcoming years when more wind farms are connected. The

trend of the release of more stringent grid codes is followed in this thesis (UK and Germany).

5.5.1 Reactive power or power factor control

The reactive power control is described by transmission system operators in power factor charts.

According to the active power output of the wind farm, the power factor at the point of connection in the

onshore grid should be controllable. An example of the range in which the power factor should be

variable for the UK grid (National Grid Company) is shown in Figure 5.5 [87],[57]. Comparable power

factor requirements were stated by E.ON-netz in Germany [88],[56]. The grid operator will ask the wind

farm operator for a certain power factor according to the grid conditions (e.g. voltage level). A minimum

requirement in most countries is a power factor close to unity anytime. The power factor chart of Figure

5.5 is used in this thesis.

51

Figure 5.5 Power factor requirement for wind farm in UK (National Grid Company)

5.5.1.1 VSC HVDC

No reactive power is transported over a DC link. The onshore converter station of the VSC HVDC link

can control the reactive power injected in or absorbed from the grid. Therefore, limited by the MVA

ratings of the link, any amount of Q can be absorbed or delivered and the power factor can be controlled

accordingly. The M5 module has a power rating of 376 MVA. Therefore, with the grid voltage at 1 pu

and the wind farm operating at nominal power (300 MW), there are still 226 MVAr (PF = 0,7979) left for

reactive compensation. This amount of reactive power can not be reached on the leading side of the P,Q

diagram as could be seen in Figure 2.8. When the VSC wants to inject reactive power in the grid, the

converter output voltage, and thus the voltage over the IGBTs, is raised above the AC grid voltage. Due

to voltage constraints on the power electronic components the amount of deliverable reactive power is

constrained when the AC grid voltage is already high. Especially when the AC voltage in the grid is

higher than 1 pu, the capability of VSC HVDC to generate additional Q becomes limited. However, the

higher the voltage in the AC onshore grid, the less probable is the extra need for reactive power. When

the AC grid voltage is low, the VSC has full reactive power capability. On the lagging side, there is no

other specific constraint than the MVA rating of the converter. To achieve a power factor of 0,95 as

defined by the NGC grid code in Figure 5.5, the maximum amount of reactive power that must be

delivered by the converter is 98,61 MVAr. This is within the ratings of the onshore VSC station as can be

seen in Figure 5.6. The grey zone depicts the range in which the reactive power should be variable. It can

be concluded that VSC HVDC does not need any additional equipment to fulfill the most stringent grid

code requirements on power factor control at the Point of Common Coupling.

52

Figure 5.6 Capability chart of onshore VSC (inverter) (grey zone = grid code requirement)

5.5.1.2 HVAC

HVAC cables produce reactive power. This is caused by the dominant shunt capacitance of the cables. It

makes the interaction of the transmission cable with the surrounding power system more complicated

than for systems using AC overhead lines or HVDC. The reactive charging current redistributes over the

cable length depending on the voltage amplitudes and angles at both ends of the cable. The traditional

approximate formulas for P (2-1) and Q (2-2) based on the voltage amplitudes and phase angles on the

ends do not longer hold for HVAC cables due to this shunt reactance. The current in the cables and the

voltage along the cables are calculated with the formulas from section 5.3.2.

Reactive compensation is chosen to be installed at both ends of the cable to minimize the losses due to the

reactive current. The reactive power produced by the cable is furthermore assumed to be equally

distributed over both cable ends. Half of the total reactive power of the cable is thus present at the

onshore node during operation, whereas the other 50% is present at the offshore cable end. The total

reactive power produced by a cable is calculated as

MVArLCU

fUQ cablecable

gridcablecable 8,593

23 =⋅⋅= π (5-10)

Two parallel HVAC cables are needed for the connection of the 300 MW wind farm. A total reactive

power of 59,8 MVAr is thus present at both cable ends and should be compensated. Offshore this is done

by two on/off inductive reactors both rated at 29,9 MVAr. These reactors are switched together with their

respective cable. The same reactors (2 x 29,9 MVAr) are installed at the onshore node as well. Therefore,

a power factor unity is achieved for the transmission system with all the reactors switched on with their

respective cables. Nevertheless, the onshore node has to fulfill the power factor requirements of Figure

5.5 and the power factor should therefore be variable and controllable. There are several possibilities to

achieve variable reactive power compensation onshore.

53

The simplest solution is a combination of switched passive elements, i.e. switched capacitors and

inductors. Although simple and inexpensive, there are several disadvantages related to this technique

[86]. The available reactive power varies quadratic with the applied voltage. During voltage sags, these

reactors lose most of their capability. Depending on the amount of switched component branches, the

reactive power can only be altered in relatively large steps. Due to the limited switching speed of the

reactors, the dynamic performance of this solution is reduced. Furthermore, switching transients have to

be accepted and regular maintenance of the mechanical breakers is needed. This solution is found,

although inexpensive, not feasible for the purpose of dynamic power factor control.

A well known and widely used reactive power compensator with better dynamic performance than the

switched reactors described above is the Static Var Compensator (SVC). The SVC combines thyristor

controlled capacitors and inductances. With the use of a SVC, a smooth variation of reactive power over

the complete installed power range is possible. This dynamic performance is sufficient to meet the

reactive power requirements shown in Figure 5.5 [86]. The mechanical switches are replaced by

electronic switching components which reduces the need for maintenance. Disadvantages are the strongly

reduced capabilities during voltage dips. The SVC can neither start its operation when no grid voltage is

present due to the thyristors. Another disadvantage is the considerable size of the total SVC station, and

its related environmental impact.

A third option considered is a Static Synchronous Compensator (STATCOM). The dynamic performance

is similar to SVC but the full capability is maintained for lower grid voltages [86]. A STATCOM uses

IGBTs as power electronic switches and has theoretically black start capability when sufficient backup

power is available (e.g. battery). The control is faster than for SVC because the IGBTs are switched

several times per frequency cycle. The STATCOM station is also considerably smaller than the SVC

station. The major disadvantage is the high cost of this technology.

As a compromise between dynamic controllability and cost, a combination of switched passive elements

with a smaller STATCOM is chosen in this thesis. A maximum reactive power of 98,6 MVAr is needed

to fulfill the grid code requirement at nominal output power. The cables’ capacitive reactive power can be

used on the leading, capacitive side of Figure 5.5 by switching of the inductive compensators. The

STATCOM rating at the leading side is therefore

MVArQQ cableSTATCOM 8,388,596,9826,98 +=−=−=+ (5-11)

Half of the reactive power on the lagging side is decided to be provided by a switched inductor. This

inductor is rated at

MVArMVAr

Qind 3,492

6,98−=

−= (5-12)

The other 50% is provided by the STATCOM (Q-STATCOM = -49,3 MVAr). The power factor can now

safely be varied over the total range for each power output of the wind farm and the grid code

54

requirements are fulfilled. An overview of the complete HVAC system with compensation is shown in

Figure 5.7.

Figure 5.7 HVAC system overview with compensation at both cable ends

5.5.2 Frequency response

The frequency in the European grid is held as constant as possible at 50 Hz. The frequency stays at a

constant value if the electricity demand plus the losses in the grid equal the total production of electricity.

Due to the rapid variations in demand and the sometimes unpredictable output of production units,

continuous control of the frequency is needed. Wind farms must respond appropriately to frequency

changes in the grid. E.ON-netz states in its grid code that wind turbines must be able to operate in a grid

frequency band of 47,5-51,5 Hz [88]. Similar requirements were stated by Eltra in Denmark [89].

5.5.2.1 VSC HVDC

VSC HVDC decouples the onshore grid from the wind farm grid. If the frequency in the onshore grid

changes, the onshore VSC station continues to deliver active power to the grid at the new frequency. The

frequency of the wind farm grid is not influenced by the frequency change onshore. A wind farm

connected with VSC HVDC can thus continue operation within the frequency band stated in the grid

codes.

The VSC HVDC link can deliver primary frequency control as well [73]. When the onshore grid

frequency becomes higher than 50 Hz, the transmission grid operator can ask the wind farm to lower its

active power output. Tripping of turbines should be avoided, because of the time needed to restart and

reconnect them afterwards and the high inrush currents that occur during startup. An option, only possible

with VSC HVDC, is to change the power transported over the link for a limited amount of time. The

offshore converter changes its control mode from frequency control to active power control and delivers a

smaller amount of active power as asked by the TSO. A part of the generated active power will now no

longer be taken from the offshore grid and the connected wind generators. This will make the turbines

accelerate. As a consequence, the offshore grid frequency will increase as well. The energy is captured

and stored in the inertia of the turbine blades. This energy can be released to the grid at a later time. The

operation in this regime is limited in time due to the speed limits of the blades. If the blades reach their

55

maximum allowable speed, the wind turbines will start to trip due to the action of their individual speed

limiters. Nevertheless, the use of this regime allows the wind farm to participate in the frequency control

of the onshore grid in a limited way.

5.5.2.2 HVAC

HVAC does not decouple the offshore wind farm grid from the onshore grid. Frequency variations

onshore are passed directly to the offshore wind turbines. The response of the turbine units on frequency

deviations depends on the used topology (action of converter). The rotational speed of Doubly-Fed

Induction Generators is directly influenced by a frequency change in the grid. The converter can be used

to oppose this rotational speed change in a limited range. Direct-Drive Synchronous Generators are

decoupled from the grid with a full frequency converter. This converter continues operation in the

proposed frequency band. The generator is not directly influenced by a frequency change in the grid.

5.5.3 Voltage-dip ride through capability

Wind turbines are not allowed to trip during short voltage dips due to faults. This requirement became

essential due to the fact that thousands of MW wind power were at risk of being lost during faults. TSOs

therefore state voltage dip ride-through capability charts for wind turbines. Examples of E.ON-netz, NGC

and the Swedish grid operator are shown in Figure 5.8 [87], [88]. The voltage dip ride-through capability

is now investigated for the different topologies under discussion.

Figure 5.8 Voltage dip ride-through chart for NGC, E.ON-Netz and Svenska Kraftnät

5.5.3.1 Squirrel cage induction generators

This topology is only under consideration with VSC HVDC. It is impossible to fulfill the grid code

requirements with this topology and a HVAC connection. The wind turbines would immediately trip

during a voltage dip. If the onshore voltage is not reduced to zero due to the sag, the VSC HVDC station

can continue to deliver power at this lower voltage, only limited by the current rating of the power

56

electronics. If for example a voltage dip of 50% occurs during a period where the wind farm is operated at

50% of its nominal power, the VSC HVDC can double the current to 100% and continue operation.

Severe voltage dips often occur due to a fault in the grid. High short-circuit currents are related to faults

and these should be avoided in the converter valves, in order to protect the power electronic equipment.

When short-circuit currents are measured in the onshore converter station, the power electronic switches

will be blocked (no current is passed). The offshore converter continues to keep the AC voltage at

reference value. As the power transmission to the shore is stopped, the energy taken from the wind will be

stored in the rotational inertia of the turbine blades. The rotational speed increases and the wind mills trip

when their overspeed protection reaches its limits. A maximal overspeed of 15% of the nominal rotational

speed is allowed before the overspeed protection trips the wind turbines [74]. The power control of the

wind turbines is properly coordinated by the VSC HVDC station, acting as a strong voltage source

ramping up the frequency.

The rotational speed of the turbine blades is defined by the equation of motion

elmech PPdt

dJ −=

ωω (5-13)

During the fault, no electric power (Pel) is taken from the blades and the captured power (Pmech) goes into

the rotating mass of the blades, increasing the rotational speed (and the wind farm grid frequency). J in

the above formula represents the inertia of the turbine and the coupled drive train. This inertia is mainly

dominated by the inertia of the turbine blades and the contribution of the generator and the drive train

components (gearbox, joints, …) will be neglected. For a turbine with 3 blades, J can be approximated by

the following expression [90].

26,2)2(9486,0

RRJ ⋅⋅= (5-14)

For a 5 MW turbine diameter, 2R, of 126 m [74], the rotor inertia is 6,2.107 kg m2.

The acceleration of the blades influences the tip-speed ratio and thus the coefficient of performance as

was shown in Figure 3.3. The captured mechanical power will therefore vary as well. It is possible to

estimate the increase of the rotational speed and the wind farm frequency by taking into account this

effect. The worst case scenario is an input of a very high wind speed just below cut-off wind speed of the

turbines (~30 m/s). The turbine blades are rotating at nominal speed (12,1 rpm) before the fault occurs.

The rotational speed increases as shown in Figure 5.9. The increase of the frequency in the wind farm is

the same. The calculation of this example was performed at a wind speed of 29,9 m/s.

If the nominal frequency in the wind farm is chosen to be 50 Hz, an approximate frequency increase of 1

Hz/s is found from Figure 5.9. This is a small increment because of the large wind turbines (high J) used

in this simulation. If smaller turbines are used (e.g. 2 MW, 2R � 75 m), the frequency increment adds up

to more than 5 Hz/s. Most faults are cleared within 250 ms after the occurrence of the fault and the

57

frequency increase in the wind farm stays within acceptable ranges. The extra energy stored in the

additional rotational speed of the blades can be released after the fault.

An extra measure to limit the power input from the wind is the enlargement of the pitch angle during

faults. As could be seen in Figure 3.3, an increase in pitch angle lowers the coefficient of performance

and thus the mechanical power captured by the blades. The response of the pitch angle control is

nevertheless slow.

Figure 5.9 Wind farm frequency and rotational speed response on fault (event at t = 1s)

(based on simulink wind turbine model)

5.5.3.2 Doubly-Fed Induction Generators

VSC HVDC can be used in the same way as for Squirrel Cage Induction Generators. The offshore wind

farm grid is decoupled from the onshore grid and the voltage dip onshore is not present in the wind farm.

The transfer of power is possibly blocked by the VSC HVDC during faults. The produced energy is then

temporarily stored in the inertia of the turbine blades while the frequency in the wind farm ramps up.

The voltage dip is directly sensed by the generating units for Doubly-Fed Induction Generators connected

with HVAC. The capability of the wind farm to ride-through the voltage dip will depend on the individual

generating units. Most DFIG topologies use a crowbar protection against overcurrents. An example of a

Doubly-Fed Induction Generator wind turbine is the Vestas V80. It was found by Dahlgren et al [91] that

these units disconnect when the voltage dips below 85% of the nominal value. The units reconnect after 7

seconds if no further fault occurs.

The connection with VSC HVDC brings advantages shortly after the faults as well. When a decreased

voltage is noticed by an asynchronous generator, the machine will start to demagnetize. After the fault is

cleared, the voltage is brought back to its nominal value and the machine will draw reactive current to

remagnetize again. This can cause severe reactive currents (2 to 3 times the nominal value), extending the

58

duration of the voltage dip [92]. The voltage dip ride-through capability is clearly enhanced by the VSC

HVDC link.

5.5.3.3 Direct-drive topologies

The VSC HVDC decouples the offshore grid from the onshore grid and onshore voltage dips are not

sensed in the offshore wind farm. However, the generators are now individually decoupled from the

offshore grid as well. This is true for the connection with HVAC cables as well. Direct-drive topologies

provide good voltage dip ride-through on their own. An example is the Enercon E66/70 investigated by

Dahlgren et al [91]. The units are not affected in any particular way during voltage dips and they are

quickly recovered after the voltage dip.

5.6 Losses

An economic comparison between VSC HVDC and HVAC is made in the next chapter. The losses of

both cable systems are an important input in this economic comparison. They are discussed now for both

options. The power probability density function of Figure 4.3 will be used to accumulate for the variation

of wind speeds over time.

Losses in VSC HVDC transmission systems are generally considered higher than in HVAC systems. This

is mainly because of the losses in the converters. The pure line losses are nevertheless higher in HVAC

cables. For a certain cable length, the level of HVAC system losses will surpass the losses in the DC

system. The distance where the nominal HVAC losses surpass the losses in the VSC HVDC link for the

chosen 300 MW wind farm will be shown in this paragraph as well.

5.6.1 VSC HVDC

VSC HVDC losses are divided in two components: converter station losses and cable losses. The losses

for a VSC HVDC converter depend on many parameters (switching frequency, IGBT losses, snubber

circuits, etc.) and are complicated to calculate in a model. For the purpose of this thesis, it is chosen to use

the losses measured in a real project. Data are taken from the Cross Sound Cable Project (330 MW, 42

km) [18] which is similar to the cable used in this thesis (300 MW, 50 km). The data are shown in Figure

5.10. The red line shows the estimated losses before the cable was installed (ABB loss model), the blue

line depicts the actual measured losses after installation.

The converter station losses are given as a percentage of the nominal output power. Part of the converter

losses consists of ‘no-load’ losses which are independent of the power through the link. The other

converter losses depend on the power sent through the converters in a quadratic way. The power

independent ‘no-load’ losses are taken from Figure 5.10 at 0 MW inverter output power. During nominal

operation, the converter losses are 1,8% per converter for the CSC Project [18].

59

0

4

8

12

16

20

24

0 20 40 60 80 100 120 140 160 180 200 220 240 260 280 300 320 330

Inverter Output Power [MW]

Lo

ss

es

[M

W]

Actual Loss

Estimated Loss

Figure 5.10 Total transmission system losses (incl. station and cable losses)

for Cross Sound Cable Project (USA) [18]

The cable losses are Joule losses and depend on the power sent through the link. The Joule losses can be

calculated with the resistance per km value of the cables [22] and vary in a quadratic way with the power

output of the inverter. The data of Figure 5.10 were scaled to the cable length of 50 km. Only the cable

losses depend on the length of the cable.

Due to the evolution in converter technology for VSC HVDC, the converter losses were 1,6% for the

Estlink Project (300 MW, 2006) instead of the 1,8% of the Cross Sound Cable Project (2002) [30]. This

evolution was taken into account in the scaling of the VSC HVDC losses. The resulting losses for the 50

km wind farm cable with nominal wind farm power of 300 MW are shown in Figure 5.11. The red curve

of Figure 5.11 is used for the 300 MW wind farm under consideration in this thesis.

Figure 5.11 VSC HVDC losses for 300 MW offshore wind farm (cable length = 50 km)

60

5.6.2 HVAC

The line losses in HVAC cables consist of Joule losses in the copper conductors, shield losses, armor

losses and dielectric losses. The dielectric losses are neglected in the following discussion.

5.6.2.1 Conductor losses in AC 3-core cables

The most important component of the line losses are the Joule losses in the conductors. They are

calculated as

dxIRPL

cablecableJoule � ⋅=0

23 (5-15)

The current in a HVAC cable varies throughout the entire length of the cable. The total current consists of

an active and a reactive component.

22reactiveactivecable III += (5-16)

The active current per cable (2 in parallel) depends on the power output of the wind farm.

cable

wfactive

U

PI

32= (5-17)

The shunt capacitance of the cable causes reactive current. Reactive compensation is added according to

section 5.5.1.2. The current distribution in the cables is calculated with the formulas given in section

5.3.2. The reactive current (Ireac) and total current (Icable) distribution in the cables are shown in Figure

5.12 and Figure 5.13 respectively. The reactive current is dominant in the total current during periods of

low wind farm output power (low wind speed). It is therefore chosen to use only one cable during these

periods to minimize the losses in the total system. The second cable is optimally switched on when the

wind farm produces more than 9% of its nominal output power. A hysteresis switch sheme can be used to

avoid numerous switching operations. The second cable can be switched on at 11% of the nominal power

and switched off at 7% for example.

61

Figure 5.12 Reactive current distribution in the chosen 3-core HVAC cable (Pwf = 0 MW)

Figure 5.13 Estimated current distribution in HVAC cable 1 (left) and 2 (right)

Rcable is calculated taking into account temperature, skin and proximity effect.

5.6.2.1.1 Temperature effect

The resistance of copper cables varies with the temperature of the conductor following

( )( )20120, −+= ° θαCDCDC RR (5-18)

with the temperature coefficient of the resistivity of copper and � the operating temperature of the

conductor. A per phase DC resistance of 0,0366 �/km at 20°C is given in [79] for the chosen cables. The

temperature of the conductor depends on the steady-state current through the cable. It is assumed that the

surrounding temperature of the sea bed is 10°C. Data from [79] are used to calculate the temperature of

the conductor as a function of the current through the cable. The result is shown in Figure 5.14. When the

cable is fully loaded, the conductor temperature rises to around 80°C.

62

Figure 5.14 Conductor temperature as function of current through cable

To know the losses during AC operation, the AC resistance should be calculated. This is done by taking

the skin effect (ys) and proximity effect (yp) into account.

)1( psDCAC yyRR ++= (5-19)

5.6.2.1.2 Skin effect

If a cable is conducting high alternating currents, the distribution of current is not evenly dispersed

throughout the cross section. The center portion of the conductor experiences a higher magnetic flux due

to surrounding current than portions at the outside of the conductor. This forces the current to redistribute

to the outside portions of the conductor because of a greater back-emf in the conductor center. Due to this

redistribution, the cross section is not used evenly and the experienced resistance of the cable is higher

than for DC currents. The skin effect factor is calculated following IEC 60287 [93]. The values given are

valid for a conductor temperature of 20°C and vary slightly with temperature.

0585,01928,0 4

4

=+⋅

=s

ss

x

xy (5-20)

with

433,3108 7

2=

⋅⋅⋅=

−

DC

grids R

fx

π (5-21)

5.6.2.1.3 Proximity effect

The proximity effect is related to the interaction of conductors being close to each other. When the

current in two conductors flows in the same direction, the current is repelled to the outside. The currents

in two conductors with currents in opposite directions are attracted to each other. This effect gives rise to

a redistribution of the current and augments the experienced resistance of the cable. The proximity effect

63

factor is calculated following IEC 60287 [93]. The values given are valid for a conductor temperature of

20°C and vary slightly with temperature.

0537,0

27,01928,0

18,1312,0

1928,04

4

22

4

4

=

�����

�����

�

++

+��

���

�⋅⋅�

�

���

�⋅

+=

p

p

cc

p

p

p

x

xs

d

s

d

x

xy (5-22)

with

433,3108 72

=⋅⋅⋅=−

DCp R

fx π (5-23)

cd is the conductor diameter in mm = 25,23 mm

s is the spacing between conductor centres in mm = 2dc (estimated)

Because of the relation between the current and the conductor temperature, the resistance of the HVAC

cable is plotted as a function of current as well. The result is shown in Figure 5.15.

Figure 5.15 Cable resistance as a function of current through cable

An extra increase in resistance can be justified due to the stranded conductors. They are helicoidally

bonded together which increases the length of the conductor by +2%. This effect is neglected.

5.6.2.2 Shield and armor losses

The currents in the conductors induce currents in the metallic shields and the armor. These currents lead

to extra losses due to the resistance of the shields and the armor. They are commonly represented as an

increase in cable resistance

( )armorshieldcableeffcable RR λλ ++= 1, (5-24)

64

This increase in resistance depends on many factors such as distance between different phases,

temperature of the cable and the material properties of shields, armor and insulation. Shield and armor

losses are stated to be possibly up to 33% and 50% respectively of the conductor losses in [5] and the

high variations in shield losses (1,0-123,0% of PJoule) depending on the properties of the cables are

discussed in [94] as well. The calculation of this increment is based on IEC 60287 [93] in this thesis. For

the shield losses holds

2

1

5,1

��

���

�+

=

X

RR

R

Scable

Sshieldλ (5-25)

RS represents the resistance of the shield [�/m]. This value depends on the resistivity of the shield

material (copper in this case), the cross section and the temperature of the shield. The temperature of the

shields depends on the heat dissipated in the conductors and the thermal resistivity of the insulation

(XLPE). This temperature value is calculated in a thermal analysis of the cable, which would lead too far

for the purpose of this thesis. It is therefore assumed that the shield temperature is constant at 40°C. X

depends on the geometry of the cable. Depending on the current through the cable, shieldλ varies between

0,28 and 0,35.

The increment for the armor losses is given by

261077,2

1

123,1

���

����

�+

=

ω

λ

Acable

Aarmor

RR

R (5-26)

The armor material is steel, which has a negative effect on the armor losses due to its magnetic

characteristics and relatively high resistance. Once again RA depends on the armor temperature, which is

the result of a thermal analysis. It is assumed here that the armor temperature is equal to 20°C. The armor

consists of steel wires bonded around the three phases. The thickness of the wires is assumed to be 1 mm

and they are placed close against each other. armorλ varies between 0,56 and 0,7.

5.6.2.3 Losses in transformers and compensation coils

In order to find the total losses in the HVAC system, the losses in the transformers and the passive

compensators are incorporated. Large power transformers usually have efficiencies around 99,8% [95].

The accumulated losses for transformers and compensation coils are linearly approximated to 1,2% for

nominal load with a no-load value of 0,4% in [80]. The system under study in [80] is a 3-core cable

system connecting a 500 MW wind farm situated 100 km from shore and with a line voltage of 245 kV.

This system differs significantly from the system under study in this thesis. The reactive currents for a

100 km cable at 245 kV are higher and the compensators are of higher ratings. The losses for transformers

and compensators are assumed to be 0,8% of the nominal power of the transmission system with 1/3 of

65

this value being no-load losses.

5.6.2.4 Losses in STATCOM

The STATCOM consists of a reactive compensator connected to the grid via a Voltage Source Converter.

The losses in a STATCOM station are in the same range as those for a VSC HVDC substation (1,6%-

1,8% of Pnom during nominal operation). The compensation equipment is nevertheless dimensioned in

such way that the STATCOM is not needed (switched off) for a power factor equal to 1. Only for the

control of power factors differing from unity, the STATCOM is switched on. It depends on the grid

conditions (e.g. present voltage level) in the neighborhood of the Point of Common Coupling which

power factor is demanded by the TSO. It is assumed in this thesis that the dominantly demanded power

factor is equal to 1 and the losses of the STATCOM are therefore omitted in the economic analysis. They

can easely be incorporated if case-specific data on which power factor is demanded by the TSO is

inserted. The total losses of the HVAC transmission system (50 km) are shown in Figure 5.16 as a

function of the wind farm power output.

Figure 5.16 HVAC transmission system losses (300 MW, 50 km)

Figure 5.11 for the VSC HVDC connection is compared with Figure 5.16 for the HVAC option. The

losses in the VSC HVDC system are higher over the total power range. Taking into account the wind

speed probability distribution, and the related power probability distribution, the annual energy loss can

be calculated for both technologies.

The results of this calculation are shown for a variable speed wind farm in Table 5-2. The annual energy

losses for both technologies are compared with the annual energy output of a variable speed wind farm

(Table 4-1, drive train losses not included) to result in an annual transmission system loss percentage

66

( )( ) hpP

hpPl

N

i iigen

N

i iiloss

⋅⋅

⋅⋅=��

,

,

% (5-27)

ilossP , represents the transmission losses at wind speed I, igenP , is the power generated by the wind farm at

wind speed I, N is the number of wind speeds considered in the model, ip is the probability to have wind

speed I, h is the number of hours in a year. The availability of the wind farm is assumed to be 100% in the

made calculations.

For the VSC HVDC option 4,45% of the annual produced energy (APE) is lost in the transmission

system, whereas this is only 3,31% for the HVAC system.

5.6.3 Influence of length of transmission cable

The losses for the VSC HVDC and HVAC link were investigated so far for a fixed cable length of 50 km.

%l is shown in Figure 5.17 for cable lengths upto 130 km (� critical length for the HVAC option). The

break-even distance for the losses of VSC HVDC and HVAC is slightly below 80 km for the 300 MW

wind farm studied in this thesis.

Figure 5.17 Loss percentage for VSC HVDC and HVAC as function of cable length

(Pwf = 300 MW)

67

En

erg

y L

oss

HV

DC

[M

Wh

] 0

6 22

0

3 84

0

3 38

0

2 98

0

2 74

0

2 60

0

2 51

0

2 45

0

2 41

0

25 6

00

54

72

0

Av

era

ge

VS

C H

VD

C

Lo

sses

[M

W]

0

3,65

1

3,94

2

4,35

9

4,90

3

5,57

3

6,37

0

7,29

3

8,34

3

9,51

9

11,5

86

En

erg

y L

oss

HV

AC

[M

Wh

] 0

1 69

0

1 31

0

1 41

0

1 50

0

1 62

0

1 76

0

1 89

0

2 03

0

2 16

0

25 4

00

40

74

0

Av

era

ge

HV

AC

Lo

sses

[M

W] 0

0,99

2

1,34

3

1,82

6

2,47

9

3,30

6

4,31

5

5,51

4

6,91

4

8,52

8

11,4

73

Ho

urs

/yea

r

699,

5

1 70

4,0

975,

2

774,

8

607,

1

491,

4

407,

5

343,

7

293,

5

253,

1

2 21

0,1

8 7

60

,0

Pro

ba

bil

ity

[%] 7,

99%

19,4

5%

11,1

3%

8,85

%

6,93

%

5,61

%

4,56

%

3,92

%

3,35

%

2,89

%

25,2

3%

10

0,0

0%

Win

d F

arm

Po

wer

ra

nge

[%]

0%

0-1

0%

10

-20

%

20

-30

%

30

-40

%

40

-50

%

50

-60

%

60

-70

%

70

-80

%

80

-90

%

90

-10

0%

TO

TA

L

Table 5-2 Loss comparison between HVAC and VSC HVDC as transmission option for a 300 MW

offshore wind farm (cable length = 50 km)

68

5.7 Conclusion

The technical aspects of VSC HVDC and HVAC are compared in this chapter. The choice of a HVDC

Light® module for the connection of a proposed offshore wind farm is straight forward, based on the

available standard modules. More parameters are open in the dimensioning of a HVAC system. Each

wind farm project needs a proper dimensioning and the outcome depends on many case-specific

parameters. A 300 MW wind farm is proposed as a relevant example in this thesis. The cable length is

chosen at 50 km. A M5 HVDC Light® is chosen as an appropriate module for this wind farm. Two

parallel 3-core HVAC cables rated at 150 kV are chosen for the HVAC option. The unavoidable

compensation equipment for the AC cables is dimensioned based on grid code requirements in the UK

and calculations on the reactive power production of the cables.

The proposed VSC HVDC system is of great help in fulfilling the grid code requirements on power factor

control, frequency response and voltage dip ride-through. Additional equipment is needed for a HVAC

system to fulfill the grid code requirement on power factor control. This comes at additional cost as will

be discussed in the following chapter. The frequency response and voltage dip ride-through of the wind

farm depends on the used wind turbine topologies. VSC HVDC shows nevertheless important advantages

over HVAC in these fields.

The losses of the VSC HVDC link in this thesis are based on the losses in the Cross Sound Cable Project.

A loss calculation is performed to approximate the losses of the HVAC cables. For a wind farm situated

at 50 km from the Point of Common Coupling, the losses in the VSC HVDC link (4,45% of APE) are

found to be considerably higher than in the HVAC system (3,31% of APE) mainly due to the losses in the

converter stations. The influence of the cable length is investigated. The pure line losses per km are found

to be higher in the HVAC cables than in the HVDC cables. This results in an approximate break-even

distance for the losses of 80 km.

69

6 ECONOMIC COMPARISON OF VSC HVDC AND HVAC FOR THE

CONNECTION OF OFFSHORE WIND FARMS

6.1 Introduction

The technical and operational aspects of the chosen offshore wind farm are discussed in the previous

chapters. Two possibilities are considered for the transmission system between the offshore wind farm

and the onshore grid: VSC HVDC and HVAC. The two options are compared on a technical basis. VSC

HVDC shows technical supremacy over the HVAC option on many aspects important for the connection

of offshore wind farms but has higher losses as well. Nevertheless, the choice for a technology is

normally based on an economic comparison. The results of the economic comparison are discussed in this

chapters.

The methodology used for the economic comparison is a discounted cash flow (DCF) analysis. The result

of this calculation is the difference between the Net Present Values (NPV) of the compared technologies.

A DCF analysis incorporates the initial investment costs and the discounted annual costs and revenues for

both technologies. A tool is developed in MS Excel to perform the economic comparison and to allow for

quick variation of input parameters.

An important question in the economic comparison is: who is deciding on the investment and what are the

investors objectives? This thesis investigates two different scenarios. In a first comparison (Scenario 1)

both VSC HVDC and HVAC are only used for the purpose of transmitting electricity. The investor is

assumed to be another party than the wind farm investor and has no benefits of possible wind farm

optimization. The wind farm is seen as a black box. Scenario 2 describes the situation in which the wind

farm developer and the investor for the transmission cable are the same party. The investment for wind

farm and transmission system is now seen as one project and is economically optimized as such. Possible

advantages of using VSC HVDC for the wind farm are incorporated in the economic analysis.

The investment costs for both transmission systems are discussed in section 6.2. The annual costs used in

the first comparison are the losses and the maintenance costs of the transmission system (6.3). The results

and conclusions of the DCF for scenario 1 are discussed in section 6.4. The differences in investment

costs for the wind turbine topologies under study are described in section 6.5. The differences in energy

output and annual maintenance costs are taken into account (6.6). The discounted cash flow for scenario 2

is shown and discussed in section 6.7. A VSC HVDC link with a wind farm with directly connected

SCIGs is compared with a HVAC link with DFIGs, DDPMSGs or GPMSGs.

The cost data used must be put in the right perspective. Data from 2002 to 2008 have been found and

uncertainty exists on their correctness nowadays. The economic comparison is performed in Euro. Some

data were found in other currencies (American Dollar, Canadian Dollar, British Pound) and had to be

converted to Euro. There was a high variability on these currency rates during the last years. The copper

70

price varied significantly during the last years and cable prices were therefore variable as well. It is

proposed to the reader to take the data mentioned in this comparison as indicative.

6.2 Transmission system investment cost

6.2.1 VSC HVDC

The cost of a VSC HVDC connection to an offshore wind farm is broken down into the following

components: substation cost, onshore land use, DC cable cost and offshore rig. The costs attained in this

paragraph refer to the M5 module, chosen in the previous chapter for the 300 MW wind farm.

Several sources were used to attain accurate data. A study of VSC HVDC as an alternative for the

Vancouver Island Transmission Reinforcement (2005) is described in [98]. The VSC HVDC studied, is a

double M5 module for a connection of 2 x 300 MW. A study for the Cape Wind Project (2003) compares

a VSC HVDC option with a HVAC cable [99]. The VSC HVDC studied in [99] is rated at 300 MW. A

study held by Montana Alberta Tie Ltd. Is described in [100]. A study for the grid reinforcement of North

Auckland with a M5 HVDC Light® is described in [101]. A similar study was conducted for the

Middletown-Norwalk Transmission Project in USA [102]. A study on the feasibility of VSC HVDC in

Randstad is described in [103]. Furthermore, prices of comparable finished projects such as Murraylink

(220 MW for 176 km land cable, 2002), Cross Sound Cable (330 MW for 40 km submarine cable, 2002)

and Estlink (350 MW for 210 km submarine cable, 2006) and planned projects such as the Borkum 2 link

[104] (420 MW for 200 km land/submarine cable, 2009) are used as a guideline.

6.2.1.1 Substation cost

As discussed before, a substation for VSC HVDC is more complicated than a HVAC substation. The

most important components are the power electronic converter, the phase reactor, AC harmonic filters, a

transformer and switchgear whereas a HVAC substation is mainly a transformer with the necessary

switchgear. Few authors report explicitly on the cost of converter substations for VSC HVDC. The prices

are often incorporated in larger contracts, depend from project to project and are often confidential. It is

neither always clear if the prices refer to HVDC Light® (ABB) or HVDC Plus (Siemens). The substation

cost depends on the power rating of the converter and is commonly expressed in €/kW.

A first reference on substation cost is found in [98]. The author states a cost of 177,72 €/kW for the

substation (2005). The study for the Cape Wind Project [99] states a cost of 151,87 €/kW for the

substation (2003). A substation cost of 127,62 €/kW is used in [100] (2002). For the M5 module, rated at

376 MVA, this would lead to a substation cost of about 48,000,000 – 66,800,000 €/converter.

Three references are used in which ABB is mentioned as the source [101],[102],[103]. The converter

prices stated in these references vary between 43 and 50 M€ per converter station. The first 2 references

are given for a M5 module HVDC Light® whereas the highest price is found for an 1100 MW M9

module. These references are considered as more relevant and realistic.

71

For finished projects, only the total cost of the project (cable and installation included) is available on

public domain. The Estlink cost 110 M€ for a total length of 210 km DC cable [105]. The Murraylink cost

165 M$ [100] (165 M€ 2002) and the Cross Sound Cable cost 130 M$ (130 M€ 2002). The data of the

Estlink is more relevant because it is a more recent figure. Nevertheless, it has to be noticed that the

converter used in Estlink is a 2-level converter whereas more complicated (more IGBTs) 3-level

converters were used in Murraylink and Cross Sound Cable. An interesting project to compare with is the

German 420 MW wind farm near Borkum 2. The total project cost is estimated at 260 M€ of which half is

assumed to comprise converter cost and offshore installation [104].

As a compromise between the found data, a converter cost of 45 000 000 €/converter station is assumed

in this thesis. The two converters thus need a total investment of 90 000 000 €.

6.2.1.2 Onshore land use

The onshore converter station land use is standardized and can be found in [22]. For the M5 module, the

needed land is 80x25 m2. A drawing of both the onshore and offshore station is shown in Figure 6.1. The

three voluminous cylindrical phase reactors are visible in the center. Other voluminous components are

the converter stacks (white blocks), the switchgear and transformer. Cooling devices are placed outside.

The cost of land is country and region dependent. It is nevertheless a minor cost in the whole project and

is estimated at 125 000 € for this comparison.

Figure 6.1 VSC HVDC M5 substation: onshore (left) and offshore (right) station

6.2.1.3 DC cable cost

A cable cost of 439 600 €/km was stated in 2005 by [98] and 429 000 €/km was reported by [99] in 2003.

Both reported costs include the two cables needed for the bipolar DC cable pair. A cable cost for DC land

cables of 290 000 €/km is stated by [102] in 2004. For a 50 km cable, a purchase cost of around 22 000

000 € could be expected in 2005. The price of copper cables is nowadays higher than a few years ago due

to the increase in value of copper. The evolution of the copper price on international metal markets is

72

shown in Figure 6.2. The red boxes show the variations during various years used in the economic

analysis and an average value is deducted. Between 2005 and 2008, the copper price has more than

doubled (+110%). The 2008 DC cable price is estimated at 600 000 €/km adding up to a total cable cost

of 30 000 000 € for the 50 km HVDC cable pair.

Figure 6.2 Relative evolution of copper price 2002 – 2008

The installation cost of the DC cables has to be added to the investment cost of the transmission system.

Both cables are installed in one boat run and buried 1 m deep in the seabed to protect them from the

submarine environment. Only [99] reports on installation cost for DC cables. An installation cost of 215

000 €/km is used, leading to a total installation cost of 10 750 000 €.

6.2.1.4 Offshore rig

An offshore platform is necessary to install the M5 converter station offshore. The cost of a top site

structure offshore is unclear. The experience for offshore wind farms is limited and platform building

companies are reluctant to provide cost data because they vary strongly from situation to situation. The

cost of an offshore installation platform is dependent on the depth of the water and the weight and volume

of the installation. It will be expressed in €/m3 here. The size of the offshore substation is 30x40x20 m3 as

provided by ABB in [106]. It is reported in [99] that the additional cost for a VSC HVDC rig compared to

a HVAC substation rig is 8,600,000 €. Compared to a HVAC offshore substation a HVDC offshore

station is about 85% larger [82]. This gives a platform cost of 792,12 €/m3 or 19 000 000 € for the needed

total volume. This number is uncertain and after communication with ABB, it is decided to use 1 000

€/m3 or 24 000 000 €.

6.2.2 HVAC

The investment cost of HVAC can be calculated in the same way as for VSC HVDC. The cost is broken

down in substation cost, onshore land use, AC cable cost and offshore rig cost. For HVAC systems an

extra cost component has to be taken into account. Reactive compensation of the cable is taken as an

absolute requirement to bring the system online (grid code). The cost of this reactive compensation will

73

be included in the HVAC investment cost.

The study for the Cape Wind Project [99] reports on HVAC and these data are used. A comparison of

different HVAC cables is made in a study by the US National Renewable Energy Laboratory (2006) and

described in [107]. The Danish Altener Project by Seawind also reports on HVAC cost data [108].

6.2.2.1 Substation cost

The substation cost consists of electrical equipment needed to transform the voltage from the transmission

cable level (150 kV) to the wind farm grid level (33 kV) offshore and to the transmission grid level (400

kV) onshore. Transformer and switchgear are the main components. A busbar and additional control and

instrumentation equipment are assumed to be present as well.

The Cape Wind Study reports a cost of $12 000 000 (12 M€ 2002) for the electrical equipment in an AC

substation rated at 420 MW. The Danish Altener Study reports 15 000 000 € for an offshore AC

substation (rig included) rated at 240 MW. The cost of a transformer, which is an important cost in a

HVAC substation, is found to be approximately 12 000 €/MW in [107]. This leads to an offshore

transformer cost of almost 4 000 000 € for the proposed wind farm in this thesis. The total substation cost

is assumed at 10 000 000 €. Aggregated for both substations, this results in an investment cost of 20 000

000 €.

6.2.2.2 Onshore land use

The onshore land use is more limited for an AC substation than for a HVDC substation. The cost of

onshore land use is estimated at 50 000 € but is a minor cost in the total investment cost.

6.2.2.3 Cable cost

High voltage submarine AC cables are mainly provided by three companies in Europe: ABB, Nexans and

Prysmian. Prices often differ from company to company and from project to project.

The Cape Wind Study mentions an AC cable cost of 390 000 €/km in 2002 for a 3-core cable with 800

mm2 as conductor cross section (ABB and Prysmian (former Pirelli) proposal for Horshoe Shoal wind

farm). A cable cost of 550 000 €/km can be derived from Nexans’ proposal for the Sheringham Shoal

wind farm [117]. In 2006, a cable cost is reported between 620 000 €/km and 710 000 €/km depending on

the manufacturer [107]. The steep rise in cable cost is due to the high increase in copper price during the

last years. The prices of 2006 are more relevant than the prices of 2003. Since 2006, the copper price has

increased by 20% as can be seen in Figure 6.2. The impact of the copper price on the cable cost is higher

for AC cables than for DC cables because the relative amount of copper is higher. The 2008 HVAC cable

price is estimated at 750 000 €/km or 37 500 000 € per cable. For the two parallel cables this means

75 000 000 €.

The cable installation cost is higher for AC cables than for DC cables. For cables as long as 50 km, one

74

boat run per cable is necessary. The handling and unrolling of the cable is more complicated because of

the higher cross section compared to DC cables. Extra costs can be expected as well due to the complex

jointing of the three phases in a submarine environment. An installation cost of 170 000 €/km is reported

by [99] leading to a total installation cost for 2 AC cables of 17 000 000 €.

6.2.2.4 Offshore rig

A typical offshore wind farm connected to the grid by HVAC is the Horns Rev Danish wind farm [82].

The offshore substation used there has extra components such as a helicopter pad, control and

instrumentation equipment, man over board boat and an emergency diesel generator inclusive 2 x 50

tonnes of diesel. As an example the size of the offshore substation of the Horns Rev (160 MW, Appendix

D) wind farm is 20x28x15 m3. This is assumed to be somewhat higher for a 300 MW wind farm

(35x25x15 m3). Given the cost of 1000 €/m3 for an offshore platform used for the VSC HVDC platform,

the HVAC offshore rig cost is estimated at 13 125 000 €.

6.2.2.5 Reactive compensation

The need for reactive compensation of the HVAC cable system was discussed in 5.5.1.2. It was chosen to

use individual compensating shunt coils for the 4 cables. Furthermore, to fulfill the grid code

requirements, an extra inductive compensator is installed plus a STATCOM. The prices of compensation

equipment are not readily available. They are typically expressed in €/MVAr.

The cost of a STATCOM is based on the dynamic reactive power range

STATCOMSTATCOMSTATCOMSTATCOMdyn QQQQ ,,, −+ −=∆= (6-1)

For the STATCOM used in this thesis, a dynamic range of ~90 MVAr is found. The operational

experience with STATCOM is limited so far. An existing project at Austin, Texas (Holly STATCOM)

cost 11 726 000 € [109]. The dynamic range of this STATCOM is 190 MVAr (+95/-95 MVAr) leading to

a STATCOM cost of 62 000 €/MVAr. A STATCOM cost of 70 000 €/MVAr is stated by [110]. An

average value of 66 000 €/MVAr is used in this thesis. This results in a STATCOM cost of 5 940 000 €.

The cost of compensation coils is not found in public domain. Compensation coils are voluminous

apparatus and a circuit breaker is needed for their connection to the system. Their specific installation

offshore adds up to the cost. A cost of 10 000 €/MVAr is estimated in this thesis. A total of ~170 MVAr

of inductive compensators is needed for the wind farm under consideration. This leads to a compensation

cost of 1 730 000 €.

The total cost for compensation equipment is estimated at 7 870 000 €.

6.2.3 Comparison of investment cost

The used data are compared in Figure 6.3. The dominant cost component for VSC HVDC is the converter

stations whereas it is the cable cost for HVAC. VSC HVDC is more expensive to purchase compared to

75

HVAC with a total investment difference around 22 000 000 € based on the data used in this thesis. The

VSC HVDC costs 154 875 000 € compared to a total cost of 132 845 000 € for the HVAC connection.

0,0

20,0

40,0

60,0

80,0

100,0

120,0

140,0

160,0

180,0

Substation Cables Cable installation Reac Comp Rig & Land Total

Co

st

[M€]

HVDC Light

HVAC

Figure 6.3 Investment cost comparison: VSC HVDC versus HVAC (300 MW, 50 km)

6.3 Annual costs

6.3.1 Losses in the transmission system to shore

The system losses for both VSC HVDC and HVAC were discussed in 5.6. The values of Table 5-2 are

used in the economic comparison. The annual energy losses are given in MWh. To give a monetary value

to these losses in the discounted cash flow, a cost of energy in €/MWh is needed.

The following reasoning is used in this thesis. Losses need to be generated additionally to the market

demand. The cost to produce the energy of the losses is therefore considered as the production cost

(€/MWh) of the marginal energy unit in the power exchange market. This is simply the electricity price

on the power exchange. The electricity price on the power exchanges is however not a fixed value. It

varies during the day and during the year. It depends on decisions of market players on the exchange,

weather conditions, prices of primary energy resources, country, …. A value of 40 €/MWh is used in this

thesis as a base value. No account is given to possible extra subsidies for renewable energy sources that

can be in effect, and add considerably to the profitability of the wind farm. The sensitivity of the final

result upon variation of this parameter is therefore discussed in the next chapter.

6.3.2 Maintenance costs of the transmission system

The annual maintenance cost for VSC HVDC is estimated at 0,5% of the capital cost of the components

[98]. The onshore land use and converter platform are not included in this capital cost. This leads to an

annual maintenance cost of 650 000 €. The complexity of the converter stations gives rise to this high

cost. Enough spare components are required for the stations to keep their reliability high enough.

The lifetime maintenance cost for the HVAC equipment in the substations is estimated at 15% of the

investment cost [96],[97]. This results in an annual maintenance cost of 150 000 € for a wind farm with a

lifetime of 20 years. Extra maintenance is needed for the STATCOM onshore. The STATCOM is a

similar VSC installation as the ones used in the VSC HVDC stations. The rating is nevertheless smaller

76

and the STATCOM is situated onshore. The STATCOM maintenance is estimated at 50 000 € per year

[95].

6.4 Discounted Cash Flow analysis – Scenario 1

A Discounted Cash Flow analysis is shown in this paragraph for the first scenario. The wind farm is seen

as a black box, identical for both compared transmission options, and is not taken into account in the

DCF. The result is shown in Table 6-1.

(DCF analysis in €) Year 0 Year 1 Years 1-20

Losses 0,00 -559 276,75 -559 276,75

Maintenance cost 0,00 -450 000,00 -450 000,00

EBITDA 0,00 -1 009 276,75 -1 009 276,75

Depreciation 0,00 1 101 500,00 1 101 500,00

Benefit before taxation 0,00 -2 110 776,75 -2 110 776,75

Taxation (40%) 0,00 -844 310,70 -844 310,70

Benefit after taxation 0,00 -1 266 466,05 -1 266 466,05

Depreciation 0,00 1 101 500,00 1 101 500,00

Net investment -22 030 000,00 0,00 0,00

Net cash flow -22 030 000,00 -164 966,05 -164 966,05

Discounted cash flow -22 030 000,00 -157 110,53 -2 055 841,64

TOTAL -24 085 841,64

Table 6-1 Discounted Cash Flow VSC HVDC versus HVAC

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

The choice for HVAC is 24 000 000 € less expensive than the choice for VSC HVDC on a lifetime basis.

The difference is mainly due to the difference in investment costs. The losses and maintenance costs are

higher for the VSC HVDC system as well resulting in an annual loss.

6.5 Wind farm investment costs

Four different wind turbine topologies are chosen to compare in chapter 3. The use of VSC HVDC allows

to implement a simpler wind farm or optimize the wind farm topology. Suggestions for simplification or

optimization are given for eacht of the discussed topologies. It is chosen here to use the combination of

VSC HVDC with directly connected SCIGs. The operation of the wind farm grid at variable frequency

allows to run the wind turbines at variable speed. This topology is compared with HVAC and the standard

variable-speed topologies (DFIG, DDPMSG and GPMSG).

The chosen topologies are now incorporated in the financial analysis. The use of the discounted cash flow

method avoids the need for the total investment cost of the wind farm. Only the differences between the

topologies are needed in the calculation. The following assumptions are made. The offshore wind farm

collection grid is assumed to be the same for each topology. The towers, foundations, nacelles and rotor

blades are considered equal as well. The differences in investment costs, needed for the DCF, are

assumed to be present only in the drive train. Only three components of this drive train have been studied

further: the gearbox, the generator and the converter. The wind turbine units under consideration are rated

77

at 5 MW each. Accurate cost data on these three components are not readily available in public domain.

The most accurate data are found for DFIG wind turbines, because this is the only topology built offshore

in a large scale wind farm so far. The assumed investment cost values in the following paragraphs include

the costs for installation of the components in an offshore environment. Every extra unit of mass or

volume that needs to be installed offshore comes at a high cost. To give the reader an idea of the

proportions of the cost components for a DFIG offshore wind turbine, Figure 6.4 has been derived from

several references [54],[46].

Figure 6.4 Cost breakdown for 5MW DFIG offshore wind turbine

6.5.1 Generator cost

Information on cost of generators was found in [64] for the DFIG, DDPMSG and GPMSG topologies.

Nevertheless, only the material and construction cost is taken into account in [64]. The purchase cost and

installation cost were not reported. The ratios between the different generator costs found by [64] are as

follows. If the DFIG generator cost is equal to 1, then the cost for the GPMSG generator is 1,5 and the

cost for the DDPMSG generator is 5. The cost for a DFIG 5MW generator is assumed to be around

400 000 € (80 €/kW) [74],[111]. The generator costs for DDPMSG and GPMSG are assumed at 2 000

000 € (400 €/kW) and 600 000 € (120 €/kW) respectively, given the ratios of [64]. The cost of a squirrel-

cage induction generator is found to be 250 000 € (50 €/kW) in [111].

6.5.2 Gearbox cost

The cost of a 3MW three-stage and single-stage gearbox is given by [64]. For a three-stage gearbox a cost

of 367 000 € (73,33 €/kW) is assumed for the 5 MW turbines in this thesis, whereas 200 000 € (40 €/kW)

is used for the single stage gearbox in the GPMSG.

6.5.3 Converter cost

The cost of a power electronic converter is derived from [64] as well. 40 000 €/MW is assumed in this

thesis. A DFIG topology is assumed to use a converter rated at 30% of the total turbine power rating. This

gives a converter cost of 60 000 €. A converter cost of 200 000 € is used for the direct-drive topologies.

The different investment costs for a 5 MW turbine are summarized in Table 6-2. The wind farm

investment costs of DFIG, DDPMSG and GPMSG are used for the HVAC option. The investment cost

for the SCIG is used in combination with VSC HVDC as was discussed in chapter 5.

78

(cost in k€) SCIG DFIG DDPMSG GPMSG

Generator 250,0 400,0 2 000,0 600,0

Gearbox 366,7 366,7 0,0 200,0

Converter 0,0 60,0 200,0 200,0

TOTAL 616,7 826,7 2 200,0 1 000,0

Table 6-2 Investment costs for different wind farm topologies

6.6 Annual costs and revenues of an offshore wind farm

6.6.1 Energy output of wind farm

The energy output of the wind farms based on the different topologies depends on the drive train

efficiencies given in Table 4-2. The respective annual energy outputs were given in Table 4-3. The same

value for the cost of energy as was used for the losses in the transmission system (40 €/MWh) is used to

monetize the differences in energy output. No account is given to possible governmental support for

green energy.

6.6.2 Annual maintenance cost wind farm

The different topologies are discussed in chapter 3. An important difference between the topologies is the

need for maintenance. Maintenance is expensive for wind farms situated several kilometers from shore. A

boat is commonly used to access the wind farm. This access might be limited due to harsh weather

conditions during certain periods of the year. A helicopter can then be used to access the wind turbines

(e.g. Horns Rev). However this is again at elevated cost.

The experience for offshore wind farms is limited, with only a few relevant projects operational for

several years. Those relevant projects (e.g. Horns Rev) are based on a DFIG topology. The annual

maintenance demand of the present generation of DFIG offshore wind turbines is in the order of 40 to 80

hours [47]. The gearbox requires the major part of the maintenance. Large amounts of lubricant are

needed in the nacelle. The generator of a DFIG turbine requires maintenance too. The brushes of the rotor

connection for example need at least 2 refurbishments a year due to wear. It is assumed that the annual

maintenance cost for a wind turbine increases linear with distance from shore. Several reports and

publications give estimations for the annual maintenance cost of a DFIG wind turbine (e.g.

[46],[81],[118]-[121]). The used data are plotted in Figure 6.5 as black dots. A linear regression is made

through the data found in the references to find the maintenance cost as a function of distance from shore.

The result of this linear approximation is shown in blue. The costs for onshore maintenance are mainly

for people, time and components. The per km increase for offshore maintenance is due to the needed boat

or helicopter hours and the special training for the maintenance workers.

79

0

50

100

150

200

250

0 10 20 30 40 50 60 70 80 90 100

Distance from shore [km]

An

nu

al

Co

st

[k€

]

Figure 6.5 Annual maintenance cost for 5 MW DFIG wind turbine (black dots = data; blue line = linear regression)

There is no relevant experience for other topologies offshore. Some qualitative information was found for

onshore installations [52]. The generators of the other topologies are almost maintenance free compared

to the DFIG topology [122]. The gearbox is still present in the SCIG (three-stage) and GPMSG topology

(single-stage). An educated assumption is made on the ratio between the maintenance cost for the other

topologies and the maintenance cost for a DFIG wind turbine. The maintenance cost for a SCIG topology

is assumed to be 70% of the cost for DFIG, whereas 30% and 50% are used for DDPMSG and GPMSG

respectively. The wind farm is assumed at 50 km from shore and 60 turbines rated at 5 MW are installed.

6.7 Discounted Cash Flow analysis – Scenario 2

The discounted cash flow calculation of 6.4 showed the comparison between VSC HVDC and HVAC

with the wind farm considered as a black box. The choice for VSC HVDC is found to cost ~24 M€ more

than the HVAC option. This result is valid if the advantages VSC HVDC can bring to a wind farm

(chapter 3) are not relevant for the investor of the transmission link.

If the investor of the wind farm is the same party as the investor of the transmission system, an economic

optimum for the total solution is looked for. The advantages of VSC HVDC for the SCIG topology

(3.4.1) are monetized in this thesis. The incorporation of the wind farm topology improvements for DFIG

and direct-drive topologies is proposed as future work. The discounted cash flow calculations below will

therefore discuss the, for this thesis relevant, comparison between VSC HVDC with SCIG wind turbines

and HVAC with DFIG, DDPMSG and GPMSG. The results are shown in Table 6-3, 6-4 and 6-5.

80

6.7.1 VSC HVDC with SCIG versus HVAC with DFIG

(DCF analysis in €) Year 0 Year 1 Years 1-20

Energy output (wind farm) 0,00 945 756,45 945 756,45

Maintenance (wind farm) 0,00 2 996 482,36 2 996 482,36

Losses (cable) 0,00 -561 058,68 -561 058,68

Maintenance (cable) 0,00 -450 000,00 -450 000,00

EBITDA 0,00 2 931 180,13 2 931 180,13

Depreciation 0,00 471 500,00 471 500,00

Benefit before taxation 0,00 2 459 680,13 2 459 680,13

Taxation (40%) 0,00 983 872,05 983 872,05

Benefit after taxation 0,00 1 475 808,08 1 475 808,08

Depreciation 0,00 471 500,00 471 500,00

Net investment (wind farm) 12 600 000,00 0,00 0,00

Net investment (cable) -22 030 000,00 0,00 0,00

Net cash flow -9 430 000,00 1 947 308,08 1 947 308,08

Discounted cash flow -9 430 000,00 1 854 579,12 24 267 762,87

TOTAL 14 837 762,87

Table 6-3 Discounted Cash Flow VSC HVDC/SCIG versus HVAC/DFIG

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

6.7.2 VSC HVDC with SCIG versus HVAC with DDPMSG

(DCF analysis in €) Year 0 Year 1 Years 1-20

Energy output (wind farm) 0,00 416 699,28 416 699,28

Maintenance (wind farm) 0,00 -3 995 309,81 -3 995 309,81

Losses (cable) 0,00 -544 522,26 -544 522,26

Maintenance (cable) 0,00 -450 000,00 -450 000,00

EBITDA 0,00 -4 573 132,79 -4 573 132,79

Depreciation 0,00 -3 648 500,00 -3 648 500,00

Benefit before taxation 0,00 -924 632,79 -924 632,79

Taxation (40%) 0,00 -369 853,12 -369 853,12

Benefit after taxation 0,00 -554 779,68 -554 779,68

Depreciation 0,00 -3 648 500,00 -3 648 500,00

Net investment (wind farm) 95 000 000,00 0,00 0,00

Net investment (cable) -22 030 000,00 0,00 0,00

Net cash flow 72 970 000,00 -4 203 279,68 -4 203 279,68

Discounted cash flow 72 970 000,00 -4 003 123,50 -52 382 155,44

TOTAL 20 587 844,56

Table 6-4 Discounted Cash Flow VSC HVDC/SCIG versus HVAC/DDPMSG

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

81

6.7.3 VSC HVDC with SCIG versus HVAC with GPMSG

(DCF analysis in €) Year 0 Year 1 Years 1-20

Energy output (wind farm) 0,00 928 919,80 928 919,80

Maintenance (wind farm) 0,00 -1 997 654,91 -1 997 654,91

Losses (cable) 0,00 -561 058,68 -561 058,68

Maintenance (cable) 0,00 -450 000,00 -450 000,00

EBITDA 0,00 -2 079 793,78 -2 079 793,78

Depreciation 0,00 -48 500,00 -48 500,00

Benefit before taxation 0,00 -2 031 293,78 -2 031 293,78

Taxation (40%) 0,00 -812 517,51 -812 517,51

Benefit after taxation 0,00 -1 218 776,27 -1 218 776,27

Depreciation 0,00 -48 500,00 -48 500,00

Net investment (wind farm) 23 000 000,00 0,00 0,00

Net investment (cable) -22 030 000,00 0,00 0,00

Net cash flow 970 000,00 -1 267 276,27 -1 267 276,27

Discounted cash flow 970 000,00 -1 206 929,78 -15 793 063,44

TOTAL -14 823 063,44

Table 6-5 Discounted Cash Flow VSC HVDC/SCIG versus HVAC/GPMSG

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

6.7.4 Discussion

The energy output of a wind farm using SCIG is higher than any other topology. This is possible due to

the high efficiency of the drive train (Table 4-2). Compared to the topologies with DFIG and GPMSG, the

gain in energy output is enough to outweigh the monetary losses due to the use of VSC HVDC as

transmission system, whereas this is only partly the case in the comparison with DDPMSG.

The maintenance cost of the wind farm represents a high annual cost. The differences between the studied

topologies are significant and have a strong impact on the annual net cash flow. As the costs for

maintenance were based on limited experience and assumptions, further validation of these figures is

advised.

The combination of VSC HVDC with SCIG is more cost-efficient (~15 000 000 €) to install than the

combination of HVAC with DFIG for a 300 MW wind farm situated 50 km from the PCC. This is an

important result as most offshore wind farms, built or planned up to now, use HVAC with DFIG (e.g.

Horns Rev and Thornton). The lower investment costs of the wind farm together with the higher

efficiency of the SCIG topology and the lower need for maintenance are responsible for this result.

The combination of VSC HVDC with SCIG is more cost-efficient (~20 500 000 €) to install than the

combination of HVAC with DDPMSG as well. This result is due to the considerable higher investment

costs for DDPMSG wind turbines. The annual costs are higher for the SCIG topology.

GPMSGs are considered as promising options for the future. They find a compromise between an

expensive generator in the DDPMSG topology and a gearbox with related maintenance in the DFIG

82

topology. This can be seen in the DCF results as well. The combination of VSC HVDC with SCIG is less

cost-efficient (-15 000 000 €) than the combination of HVAC with GPMSG. The investment costs for

both options are almost break-even but the annual operational costs are higher for the option with VSC

HVDC and SCIGs

6.8 Relevance of the results

VSC HVDC is compared with HVAC in two ways. A first comparison considers the wind farm as a black

box. The benefits of VSC HVDC for the wind farm are not taken into account in this comparison. The

result of this comparison is relevant from a transmission system point of view. The transmission system

operator is responsible for the connection of the wind farm with the onshore grid in some countries (e.g.

Germany). The owner of the wind farm is another party than the owner of the transmission system. The

benefits achieved in the wind farm due to the use of VSC HVDC do not result in economic value for the

investor of the transmission link.

The second comparison has a look at the total system (wind farm + transmission link). This result is

relevant from a society point of view because the economic optimization is not influenced by several

different investor objectives. This case is valid in countries where the developer of an offshore wind farm

is responsible for the connection to the onshore grid as well.

6.9 Conclusion

A 300 MW wind farm situated at 50 km from the Point of Common Coupling as taken as a base example

to compare VSC HVDC with HVAC as connection between an offshore wind farm and the onshore grid.

Both transmission systems are compared with the wind farm considered as a black box. VSC HVDC is

found not to be a cost efficient option in this case, with a DCF result of -24 000 000 € compared to the

HVAC option.

A wind farm based on SCIG wind turbines connected with VSC HVDC was compared with the

conventional options for offshore wind farms (DFIG, DDPMSG and GPMSG) connected with HVAC. It

is concluded that the use of GPMSG with HVAC is more cost efficient than SCIG with VSC HVDC.

SCIG with VSC HVDC is nevertheless a less expensive option than DFIG with HVAC or DDPMSG with

HVAC. The topology with DFIG and HVAC is the current industry standard for large scale offshore wind

farms.

83

7 SENSITIVITY ANALYSIS

7.1 Introduction

An economic comparison between VSC HVDC and HVAC is made in the previous chapter for different

wind farm topologies. The results are based on a 300 MW wind farm situated 50 km from the PCC. The

outcome varies when a parameter is changed in the model. The sensitivity of the result to the variation of

the following parameters is investigated (the values between brackets are the base scenario values):

− Wind farm distance from PCC or cable length (50 km)

− Cost of energy (40 €/MWh)

VSC HVDC is a relatively new transmission technology compared to HVAC. The technological

evolution is still going on as we speak. The influence of further progress in two characteristics of the VSC

substation is investigated:

− Losses in converter station (1,6%)

− Volume of offshore converter station (100%)

7.2 Distance between wind farm and PCC

The comparison of two transmission systems often results in a break-even distance between the

investigated solutions. It is generally accepted that HVAC is more feasible for short cable lengths,

whereas HVDC becomes more viable for longer transmission distances. The statement of a break-even

distance between two technologies is nevertheless a delicate matter and each case will have a different

outcome. VSC HVDC was compared with HVAC in this thesis in a submarine environment. The results

of the analysis depend on a high number of variable parameters and presumptuous conclusions on the one

or the other technology should be avoided. Every wind farm project requires its own technical and

financial optimization. The results nevertheless give an indication of the proportions between HVAC and

VSC HVDC. The length of the cable is varied from 50 km to 25 km and 75 km to get an idea of the

sensitivity of the result to variation of the distance between the offshore wind farm and the PCC. The

financial data are adapted to these cable lengths and the losses and compensation equipment are

recalculated for the new cable length.

7.2.1 VSC HVDC compared to HVAC – Scenario 1

VSC HVDC and HVAC are first compared without looking at the wind farm itself. The results are shown

in Figure 7.1. The investment costs are plotted as a function of cable length for both technologies. The

difference between the investment costs of both technologies is shown as a yellow column for the three

investigated cable lengths. The per unit length investment costs are higher for HVAC than for VSC

HVDC. The total investment costs for HVAC become higher for a cable length of 75 km. The economic

value (DCF result) of VSC HVDC compared to HVAC increases with cable length. Nevertheless, the

84

DCF result (depicted in red columns) is still negative due to the higher losses and maintenance costs for a

cable of 75 km. VSC HVDC is therefore not a feasible solution for connection of an offshore wind farm

closer than 75 km from the PCC from this perspective (break-even distance � 80 km).

-50,0

0,0

50,0

100,0

150,0

200,0

25 50 75

Cable length L [km]

M€

� InvC DCF HVAC InvC HVDC Light® InvC

Figure 7.1 Total investment costs and DCF: VSC HVDC versus HVAC

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

7.2.2 VSC HVDC with SCIG compared to HVAC with other topologies – Scenario 2

The solution with VSC HVDC and SCIG wind turbines is now compared to HVAC with DFIG,

DDPMSG and GPMSG. The results are shown in Figure 7.2-4. The difference in total project investment

costs (wind farm + transmission cable) between the option with VSC HVDC and the option with HVAC

is shown in blue. The results of the discounted cash flow calculation are shown in green.

-50,0

-25,0

0,0

25,0

50,0

25 50 75

Cable length L [km]

M€

� InvC Project DCF

Figure 7.2 Financial comparison: VSC HVDC with SCIG versus HVAC with DFIG

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

85

-25,0

0,0

25,0

50,0

75,0

100,0

25 50 75

Cable length [km]

M€

� InvC Project DCF

Figure 7.3 Financial comparison: VSC HVDC with SCIG versus HVAC with DDPMSG

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

-50,0

-25,0

0,0

25,0

50,0

25 50 75

Cable length [km]

M€

� InvC Project DCF

Figure 7.4 Financial comparison: VSC HVDC with SCIG versus HVAC with GPMSG

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

A longer transmission cable has a positive effect on the economic value of the VSC HVDC option with

SCIG compared to the HVAC option with other topologies. The results for 25 and 75 km make it possible

to estimate the break-even distance between the different topologies.

A topology based on VSC HVDC in combination with SCIG results in a net annual benefit compared to

the option with HVAC and DFIG. This makes the DCF result higher than the difference in investment

costs. The break-even distance is estimated around 32 km.

A wind farm based on VSC HVDC with SCIG requires a considerably lower investment than a wind farm

based on HVAC and DDPMSG. The total system efficiency is lower for the wind farm with VSC HVDC

for shorter cable lengths and the maintenance requirement is higher. This results in a net annual monetary

loss for the wind farm operator with the VSC HVDC option. The discounted value of these annual losses

is nevertheless smaller than the difference in investment cost for cable lengths above 25 km and a wind

86

farm lifetime of 20 years. The break-even cable length is situated below 25 km.

A wind farm based on GPMSG promises an optimum between efficiency and maintenance and this for a

more limited investment cost than the DDPMSG. The economic comparison between a wind farm based

on VSC HVDC with SCIG and a wind farm based on HVAC with GPMSG show an annual loss for the

VSC HVDC option. The break-even cable length is situated around 68 km.

7.3 Cost of energy

The cost of energy is an arguable parameter. It depends on too many factors (e.g. country, weather,

competition in the market,…) to be represented by a fixed value for. The sensitivity of the DCF result to

variation of this parameter needs to be investigated. The base scenario value was set at 40 €/MWh. The

parameter is varied up to 80, 120 and 160 €/MWh. There are two reasons for assuming a higher value.

First of all the demand for energy is expected to grow in the coming years [37]. As the reserves of the

primary energy sources (especially oil) tend to decrease, an increase in energy cost can be expected.

Another reason is the possible governmental support for green energy. The Belgian government supports

the first 216 MW offshore wind farm with 107 €/MWh and 90 €/MWh for capacity above 216 MW [44].

These prices are given in the form of green certificates. Although the Belgian values are rather extreme, a

certain amount can be expected and should be incorporated in the cost of energy a wind farm developer

experiences.

The cost of energy influences the final result via two ways: the energy output of the wind farm and the

energy lost in the transmission cable. The energy output of the wind farm depends on the chosen topology

and its efficiency. The losses in the transmission cable depend on the length of the cable. The sensitivity

will therefore first be discussed for the comparison between VSC HVDC and HVAC with the wind farm

as a black box. It will be discussed for a wind farm based on VSC HVDC with SCIG wind turbines

compared to a wind farm based on HVAC with other wind turbine topologies in section 7.3.2.

7.3.1 VSC HVDC compared to HVAC – Scenario 1

-80,0

-70,0

-60,0

-50,0

-40,0

-30,0

-20,0

-10,0

0,0

40 80 120 160

Cost of energy [€/MWh]

DC

F r

esu

lt [

M€]

25 km

50 km

75 km

Figure 7.5 Sensitivity of DCF result to variation of cost of energy: VSC HVDC versus HVAC

87

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

An increase in cost of energy has a negative effect on the DCF result for VSC HVDC compared to

HVAC. The explanation is the higher losses of VSC HVDC, which become more important when the cost

of energy is higher. The effect is less negative for longer transmission distances. The losses in a HVAC

cable system increase more with increasing cable length than in a VSC HVDC cable system (Figure

5.17). A smaller difference in losses between both technologies (higher cable length) results in a less

sensitive result. The sensitivity becomes zero when the break-even distance for the losses is reached. This

is calculated in chapter 5 to be at 80 km.

7.3.2 VSC HVDC with SCIG compared to HVAC with other topologies – Scenario 2

-20,0

-10,0

0,0

10,0

20,0

30,0

40,0

50,0

60,0

70,0

40 80 120 160

Cost of energy [€/MWh]

DC

F r

esu

lt [

M€]

25 km

50 km

75 km

Figure 7.6 Sensitivity of DCF result to variation of cost of energy:

VSC HVDC with SCIG versus HVAC with DFIG (wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

-20,0

-10,0

0,0

10,0

20,0

30,0

40,0

50,0

60,0

40 80 120 160

Cost of energy [€/MWh]

DC

F r

esu

lt [

M€]

25 km

50 km

75 km

Figure 7.7 Sensitivity of DCF result to variation of cost of energy:

VSC HVDC with SCIG versus HVAC with DDPMSG (wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

88

-40,0

-30,0

-20,0

-10,0

0,0

10,0

20,0

30,0

40 80 120 160

Cost of energy [€/MWh]

DC

F r

esu

lt [

M€]

25 km

50 km

75 km

Figure 7.8 Sensitivity of DCF result to variation of cost of energy:

VSC HVDC with SCIG versus HVAC with GPMSG (wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

The net energy output difference at the onshore PCC defines the sign of the sensitivity in the comparisons

for the wind farm and transmission system. The connection with VSC HVDC compared to the use of

HVAC as transmission cable has a negative effect due to the higher losses. The difference in losses

between both transmission technologies becomes again smaller for a higher cable length. The wind farm

based on SCIG is more efficient than any of the other topologies under study (Table 4-2). The sum of the

drive train losses and the transmission losses are shown in Figure 7.9 for the studied topologies. The

difference of this sum for the compared topologies defines the sensitivity to the variation of the cost of

energy. The result is insensitive to the cost of energy when the difference is zero. The distances where

this happens can be read from Figure 7.9. These values correspond with Figure 7.6 to 7.8.

Figure 7.9 Sum of drive train and transmission system losses for compared topologies

89

7.4 Converter station losses of VSC HVDC

One of the main disadvantages of VSC HVDC is the high loss percentage of the converters. Since its

commercial introduction in 1997, loss percentages have nevertheless followed a decreasing trend due to

technological evolution in the converter. The converter losses are nowadays assumed at 1,6% per

converter [30] but some progress can be expected in this domain for the upcoming years. The DCF result

is therefore recalculated for converter losses down to 1% in steps of 0,2 %-points. The DCF values are

shown in Figure 7.10. As could be expected, the reduction in converter losses has a positive effect on the

DCF result in the comparison between VSC HVDC and HVAC. The discounted cash flow increases with

1,8 M€ per 0,2%-points reduction in converter loss. This result is valid for any chosen wind farm

topology because the converter losses do not have an influence on the wind farm topology. The same

conclusion can be drawn for any cable length as shown in Figure 7.10.

-50,0

-40,0

-30,0

-20,0

-10,0

0,0

10,0

1,6% 1,4% 1,2% 1,0%

Loss per converter [%]

DC

F r

esu

lt [

M€]

25 km

50 km

75 km

Figure 7.10 Sensitivity of DCF result to reduction of converter losses

(wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

7.5 Converter volume

An important cost factor in the investment cost for VSC HVDC is the cost for the offshore platform. A

reduction in converter volume offshore reduces this cost component and is expected to have a favorable

effect for VSC HVDC on the DCF result. The offshore converter volume is therefore reduced in steps of

10% of the original volume to a minimum of 70%. The DCF results are shown in Figure 7.11. The

discounted cash flow result increases with 1,8 M€ per 10% reduction in offshore converter volume. This

result is valid for any chosen wind farm topology because the offshore converter volume does not have an

influence on the wind farm topology. The same conclusion can be drawn for any cable length as shown in

Figure 7.11.

The results for reduction in losses and converter volume are important for VSC HVDC manufacturers. It

is seen that a 10% reduction of converter volume brings an equal economic benefit as a converter loss

90

reduction of 0,2 %-points. This figure can be used to decide in which field the efforts for technological

development are best put.

-50,0

-40,0

-30,0

-20,0

-10,0

0,0

10,0

100% 90% 80% 70%

Offshore converter volume [%]

DC

F r

esu

lt [

M€]

25 km

50 km

75 km

Figure 7.11 Sensitivity of DCF result to reduction of offshore converter volume (wind farm lifetime = 20 years; taxation rate = 40%; discount factor = 5%)

7.6 Conclusion

The sensitivity of the results derived in Chapter 6 to variation of some important input parameters was

investigated in this chapter. A wind farm connected with VSC HVDC was always compared with a

system based on HVAC technology. The effect of the length of the transmission cable was investigated

first. This made it possible to derive break-even distances between two options. The results depend on the

topologies used in the comparison. A break-even distance of 80 km is found if only the transmission

function of VSC HVDC is compared with HVAC. VSC HVDC can influence the topology and operation

of a wind farm, resulting in wind farm simplification. The comparison between VSC HVDC with SCIGs

and HVAC with variable speed topologies resulted in break-even distances between 20 and 68 km

depending on the topology used with HVAC.

The cost of energy is an uncertain parameter. Its influence on the financial result depends on the chosen

wind turbine technology and the transmission distance. The result is more sensitive if the difference in

energy output at the onshore PCC is higher between the compared topologies.

Technological progress is expected for VSC HVDC in the form of a reduction of the losses in the

converter stations and a reduction of the volume of the offshore converter station. This progress will have

a positive effect on the feasibility of VSC HVDC.

91

8 CONCLUSIONS

Future wind farms are found to have a power rating higher than 200 MW and to be situated several tens

of kilometres from shore. The use of a separate transmission system is then unavoidable. This thesis

makes a comparison between VSC HVDC and HVAC for the connection of an offshore wind farm to the

onshore grid. HVAC is traditionally used, whereas VSC HVDC is a rather new technology with

promising characteristics in this context. The most important features of VSC HVDC are the control of

active and reactive power, the possibility to connect two asynchronous grids with even differing and

varying frequencies and the black start capability.

The different wind turbine topologies are discussed. Their advantages and disadvantages are explained,

especially for offshore use. The topology with directly connected squirrel cage induction generators

(SCIG) is still under consideration in combination with VSC HVDC, whereas it is found unfeasible in

combination with HVAC. The DFIG and (geared) direct-drive topology are possible with both VSC

HVDC and HVAC. The annual energy output of each wind turbine topology is investigated. An important

increase in energy output is achieved by using variable-speed operation of the wind turbines. The

operation of the offshore wind farm grid at variable frequency makes it possible to run directly connected

induction generators at variable speed. A multi-turbine frequency approach is used in this context. It was

shown that the effect on the annual energy output of this approach is negligible compared to individual

speed control per turbine.

In order to come to a more calculation-based analysis in this thesis, a typical wind farm is put upfront to

investigate. The power rating of the wind farm is chosen at 300 MW and the length of the transmission

cable at 50 km. The individual turbines are rated at 5 MW each. A M5 HVDC Light® is chosen as an

appropriate VSC HVDC transmission system for this wind farm. Two parallel 3-core HVAC cables rated

at 150 kV are chosen for the HVAC option. The compensation for the AC cables is approximately

dimensioned based on grid code requirements in the UK and calculations on the charging current in the

cables. The proposed VSC HVDC system succeeds in fulfilling the grid code requirements on power

factor control, frequency response and voltage dip ride-through without additional equipment. The

frequency response and voltage dip ride-through of the wind farm depends on the used wind turbine

topologies (converter) in the HVAC case. For a 300 MW wind farm situated at 50 km from the Point of

Common Coupling, the losses of the VSC HVDC link are found to be 4,45% of the Annual Produced

Energy whereas only 3,31% for HVAC cables. The influence of the cable length on the losses is

investigated. The pure line losses per km are found to be higher for HVAC cables than for HVDC cables.

This resulted in an approximate break-even distance for the losses of 80 km.

Both systems are compared on an economic basis. A first comparison is held with the wind farm

considered as a black box. VSC HVDC is found not to be a cost-efficient option in the case of a 300 MW

wind farm with cable length 50 km, with a DCF result of -24 000 000 € compared to the HVAC option.

92

VSC HVDC can possibly influence the topology and operation of a wind farm. The different wind turbine

topologies are therefore incorporated in the economic analysis. A wind farm based on SCIG wind

turbines connected with VSC HVDC was compared with the conventional options for offshore wind

farms (DFIG, DDPMSG and GPMSG). It is concluded that the use of GPMSG with HVAC is more cost

efficient than SCIG with VSC HVDC. SCIG with VSC HVDC is nevertheless a less expensive option

than DFIG with HVAC or DDPMSG with HVAC.

The sensitivity of the DCF result to variation of the cable length is investigated. A break-even distance of

80 km is found when the wind farm is taken as a black box. The comparison between VSC HVDC with

SCIGs and HVAC with variable speed topologies resulted in break-even distances between 20 and 68 km

depending on the topology used with HVAC. Progress is expected in the field of VSC HVDC in the form

of a reduction of the losses per converter station and a reduction of the volume of the offshore converter.

A reduction of the converter station losses with 0,2 %-points results in an increase of the DCF result with

1,8 M€. A reduction of the offshore converter volume with 10% results in an increase of the DCF result

with 1,8 M€. The influence of variations in cost of energy is investigated. The outcome depends on the

chosen wind turbine topology.

93

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10 APPENDICES

10.1 Appendix A – Matlab model

The calculations in this thesis require a wind turbine model. A model is therefore developed in

Matlab/Simulink. The model is based on [126] and [127].

The main assumptions of the model are:

Nominal power: Pnom = 5 MW;

Weibull parameters: A = 9,8 m/s; C = 2,1;

Air density: �air = 1,225 kg/m³;

Blade radius: R = 63 m;

Pitch angle: � = 0..40°;

Wind speed range: vwind = 3..30 m/s; (cut-in to cut-out)

Variable speed operation is used at low wind speeds. The rotational speed range of the blades (min..max)

is an input parameter of the matlab model. The Matlab model calculates the optimal settings for an

instantaneously responding wind turbine. It returns the optimal rotational speed and pitch angle, given a

certain input wind speed.

The Simulink model allows for time-domain analysis. The inertia of the turbine blades prohibits the

immediate response of the wind turbine. This inertial response is implemented via the equation of motion.

A control loop for the pitch angle is implemented to track the optimal pitch angle setting. The simulink

model is for example used for the time-domain analysis in section 5.5.3.1.

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10.2 Appendix B – Rotational speed range

The range in which the rotational speed of the blades is varied in variable-speed topologies, differs from

topology to topology. A typical nominal rotational speed for a 5MW wind turbine is 12,1 rpm [74]. For

the option based on SCIG controlled by the VSC HVDC link, the data from the Tjaerborg project are

taken [73] to define the rotational speed range of the turbines. The frequency was variable between 30

and 50 Hz, which leads to a minimum rotational speed of 7,26 rpm. It is not certain if 30 Hz is the

absolute minimum frequency (and thus rotational speed) achievable with this configuration. The

connection to the Troll A platform is used as a motor drive and can lower the frequency offshore to

almost 0 Hz. For a DFIG turbine, data are taken from the REpower 5M turbine [115]. The rotational

speed is varied between 6,9 and 12,1 rpm. The speed range of a direct-drive topology depends on the

capability of the converter. E.g. the Enercon E-112 rated 4,5 MW is designed to lower its speed to 8 rpm

[123]. Rotational speeds down to 6,2 rpm were nevertheless achieved during operation with this turbine.

A higher speed range allows the turbine to better optimize the coefficient of performance of the blades.

Therefore a higher power output is achieved in the low speed ranges. A detail of the power-speed curves

is shown in Figure 10.1. It was calculated that the annual produced energy of a direct-drive topology will

be 0,47% higher than for a SCIG topology and 0,28% higher than for a DFIG topology. The overall effect

on the annual energy yield is not taken into account further. The differences are small and the data on the

speed ranges are uncertain.

Figure 10.1 Detail of power-speed curves for different wind turbine topologies

(based on matlab wind turbine model)

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10.3 Appendix C – Capacity factors

Weibull parameters for wind farm on different locations:

A (m/s) C (-) Average Wind Speed (m/s)

Offshore Wind Farm 9,8 2,1 8,68

Coastal Wind Farm 8 1,9 7,10

Onshore Wind Farm (non coastal)

7,2 1,8 6,40

Table 10-1 Wind speed data

Figure 10.2 Probability density function for different wind farm locations

Annual Energy output (MWh)

Capacity Factor

(%)

Fixed speed Variable speed

Increment due to variable speed (%)

Fix Var

Offshore Wind Farm 1 055 000 1 142 000 8,25 40,14 43,46

Coastal Wind Farm 725 280 816 650 12,60 27,60 31,07

Onshore Wind Farm (non coastal)

584 890 674 190 15,27 22,26 25,65

Table 10-2 Energy output for different locations

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10.4 Appendix D – HVAC offshore substation

Figure 10.3 Horns Rev transformer station (160 MW)

Figure 10.4 Installation of offshore transformer station (UK)