Voltage Balancing on Three- Phase Low Voltage Feeder

VOLTAGE BALANCING ON THREE- PHASE LOW VOLTAGE FEEDER

A thesis submitted to The University of Manchester for the degree of Doctor of Philosophy in the Faculty of Engineering and Physical Sciences

2015

YUN LI

SCHOOL OF ELECTRICAL AND ELECTRONIC ENGINEERING Content

Content

Content ...... 2 List of Figures ...... 6 List of Tables ...... 10 Abbreviations ...... 11 Abstract ...... 13 Declaration ...... 14 Copyright Statement...... 15 Acknowledgement ...... 16 Publications ...... 17

Chapter 1 ...... 18 Introduction ...... 18 1.1 Background ...... 18 1.1.1 Electric vehicle ...... 19 1.1.2 Heat pump ...... 20 1.2 Motivation and objectives ...... 21 1.3 Contribution of the work ...... 22 1.4 Structure of the thesis ...... 25

Chapter 2 ...... 27 Voltage Imbalance in LV Networks ...... 27 2.1 Introduction ...... 27 2.2 Low voltage network ...... 28 2.2.1 LV network in the UK ...... 28 2.2.2 LV networks around the world ...... 35 2.3 Voltage imbalance ...... 39 2.3.1 Introduction ...... 39 2.3.2 Expressions ...... 40

Page | 2 Content

2.3.3 Associated standards ...... 44 2.3.4 Causes ...... 46 2.3.5 Consequences...... 48 2.4 Voltage imbalance mitigation...... 52 2.4.1 Conventional mitigating methods ...... 52 2.4.2 Modern mitigating methods ...... 55 2.5 Voltage limits ...... 59 2.6 Summary ...... 60

Chapter 3 ...... 62 Scott Transformer based Voltage Balancing Method ...... 62 3.1 Introduction ...... 62 3.2 Voltage regulation using transformers and tap changers ...... 63 3.2.1 Basic control ...... 63 3.2.2 Voltage magnitude regulation ...... 65 3.2.3 Voltage phase angle regulation ...... 70 3.2.4 Independent voltage magnitude and phase angle regulation ...... 73 3.3 Overview of the proposed method ...... 75 3.4 Scott transformer ...... 77 3.5 Phase regulating system ...... 86 3.6 Control algorithm ...... 88 3.7 Summary ...... 92

Chapter 4 ...... 94 Computer Simulation of ST based Voltage Balancing Method ...... 94 4.1 Introduction ...... 94 4.2 Modelling of ST based voltage balancing system ...... 95 4.3 Voltage balancing using tap changers ...... 98 4.3.1 Simulation methodology ...... 98 4.3.2 Three-phase balancing method ...... 99 4.3.3 Simulation results ...... 102 4.4 Application in a typical UK distribution network ...... 106 4.4.1 Simulation methodology ...... 106

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4.4.2 Distribution network model ...... 107 4.4.3 Simulation results ...... 109 4.5 Summary ...... 117

Chapter 5 ...... 118 Physical Test of ST based Voltage Balancing Method & Associated Simulation ...... 118 5.1 Introduction ...... 118 5.2 Experimental methodology ...... 119 5.3 Experimental plant ...... 119 5.3.1 Physical voltage balancing system ...... 120 5.3.2 Microprocessor based control system ...... 120 5.3.3 Low voltage feeder ...... 122 5.4 Experimental results ...... 122 5.5 Associated simulation study ...... 125 5.5.1 Simulation methodology ...... 125 5.5.2 Simulation results ...... 125 5.6 Summary ...... 128

Chapter 6 ...... 130 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages ...... 130 6.1 Introduction ...... 130 6.2 Impact of EVs and HPs on distribution networks ...... 131 6.3 Simulation methodology ...... 134 6.4 Statistical model of EV demand ...... 136 6.5 Monte Carlo simulation platform ...... 144 6.5.1 Adopted distribution network ...... 144 6.5.2 Residential demand profile generator ...... 145 6.5.3 HP electrical demand profile generator ...... 145 6.5.4 MC simulation algorithm ...... 147 6.6 Impact of EVs and HPs ...... 148 6.6.1 Deterministic study ...... 150 6.6.2 Monte Carlo study ...... 153

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6.7 Summary ...... 160

Chapter 7 ...... 162 Mitigation of Impact of EVs & HPs on LV Feeder Voltages ...... 162 7.1 Introduction ...... 162 7.2 DR based TOU tariff ...... 163 7.3 Simulation methodology ...... 166 7.3.1 Monte Carlo study ...... 166 7.3.2 Deterministic study ...... 166 7.3.3 Modelling of TOU tariff ...... 167 7.3.4 Modelling of ST based voltage balancing system ...... 167 7.4 Effectiveness of TOU tariff ...... 169 7.4.1 Evaluation of TOU tariff in deterministic study ...... 169 7.4.2 Evaluation of TOU tariff in MC study...... 171 7.5 Effectiveness of ST based voltage balancing method ...... 173 7.5.1 Evaluation of ST based voltage balancing method in deterministic study ..... 173 7.5.2 Evaluation of ST based voltage balancing method in MC study ...... 181 7.6 Cooperation of ST based balancing system and TOU tariff...... 184 7.7 Summary ...... 187

Chapter 8 ...... 189 Conclusion & Future Work ...... 189 8.1 Conclusion ...... 189 8.2 Future work ...... 193

References ...... 196

APPENDIX A ...... 206 APPENDIX B ...... 208 APPENDIX C ...... 210 APPENDIX D ...... 218 APPENDIX E ...... 219 APPENDIX F ...... 221

Word count: 52,256

Page | 5 List of Figures

List of Figures

Fig. 1-1 Evolution of the research work and the associated publications ...... 23 Fig. 2-1 Typical UK distribution system ...... 29 Fig. 2-2 Loop MV network arrangement ...... 30 Fig. 2-3 Ring main unit ...... 30 Fig. 2-4 Diagram of an urban single-transformer MV/LV substation ...... 31 Fig. 2-5 Typical LV network in urban areas ...... 32 Fig. 2-6 Typical LV network in rural areas ...... 33 Fig. 2-7 Service cable arrangements ...... 34 Fig. 2-8 American distribution system ...... 36 Fig. 2-9 Circuit diagrams of the LV supply around the world ...... 38 Fig. 2-10 Phase diagrams of balanced and non-balanced 3Φ voltages (a) balanced 3Φ voltage (b) non-balanced 3Φ voltage ...... 39 Fig. 2-11 Variations of U2, vuf1, vuf2 for a specified LVUR (a) LVUR=1% (b) LVUR=2% (c) LVUR=5% (d) LVUR=10% ...... 43 Fig. 2-12 Derating factor of induction machine with voltage imbalance ...... 45 Fig. 2-13 Typical ASD system ...... 51 Fig. 2-14 Effect of specially-connected transformers in voltage imbalance reduction ..... 54 Fig. 2-15 Typical inverter circuit used for voltage correction ...... 57 Fig. 2-16 Automatic phase selector ...... 58 Fig. 3-1 Essential of tap changing operation ...... 64 Fig. 3-2 Basic tap changer arrangement ...... 65 Fig. 3-3 Schematic diagram of line drop compensation ...... 66 Fig. 3-4 Single-phase type-A step voltage regulator ...... 67 Fig. 3-5 Single-phase type-B step voltage regulator ...... 67 Fig. 3-6 Magnitude regulating booster transformer (a) circuit diagram (b) phase diagram ...... 68 Fig. 3-7 Magnitude regulating booster transformer using autotransformer ...... 69 Fig. 3-8 Three-phase magnitude regulating booster transformer ...... 69 Fig. 3-9 Tertiary winding used voltage magnitude regulation ...... 70 Fig. 3-10 Voltage phase angle control (a) circuit diagram (b) phase diagram ...... 71 Fig. 3-11 Three-phase phase shifting transformer (a) circuit diagram (b) phase diagram ...... 72 Fig. 3-12 Three-phase quadrature booster ...... 73 Fig. 3-13 Independent voltage magnitude and phase control (a) circuit diagram (b) phase diagram ...... 74 Fig. 3-14 New introduced regulating transformer for voltage magnitude and phase angle control ...... 75 Fig. 3-15 Schematic diagram of the proposed voltage balancing system ...... 76 Fig. 3-16 Circuit diagram of a Scott transformer ...... 77

Page | 6 List of Figures

Fig. 3-17 Three-phase and two-phase voltage conversion ...... 79 Fig. 3-18 Three-phase and two-phase current conversion ...... 81 Fig. 3-19 Scott transformer II...... 83 Fig. 3-20 Phase regulating system ...... 87 Fig. 3-21 Operation of phase regulating system (a) injected voltage inversely

proportional to VM (b) injected voltage proportional to VM ...... 88 Fig. 3-22 Main control algorithm of proposed balancing method ...... 90 Fig. 3-23 Detailed control algorithm of action 1...... 91 Fig. 3-24 Detailed control algorithm of action 2...... 91 Fig. 4-1 Diagram of ST based voltage balancing system model in PSCAD/EMTDC ..... 96 Fig. 4-2 Simplified LV feeder used in the simulation ...... 98 Fig. 4-3 Three-phase balancing system (a) circuit diagram (b) simulation model in PSCAD/EMTDC ...... 100 Fig. 4-4 Control algorithm used in the 3Φ balancing system ...... 101 Fig. 4-5 The typical UK distribution network ...... 108 Fig. 4-6 Three phase voltages along the LV feeder with/without connection of EVs .... 110 Fig. 4-7 U2s when balancing system has different output magnitude ...... 114 Fig. 4-8 Phase-A voltage with balancing system installed at different locations ...... 115 Fig. 4-9 Phase-C voltage with balancing system installed at different locations...... 115 Fig. 4-10 U0 with balancing system installed at different locations ...... 116 Fig. 4-11 U2 with balancing system installed at different locations ...... 116 Fig. 5-1 Overview of the experimental plant ...... 121 Fig. 5-2 Three-phase voltage waveforms when balancing system is initiated ...... 123 Fig. 5-3 Three-phase voltage waveforms when load reduction occurs ...... 124 Fig. 5-4 Three-phase voltage waveforms when load rise occurs ...... 124 Fig. 5-5 Three-phase voltage profiles after the balancing system ...... 126 Fig. 5-6 U2 after the balancing system ...... 127 Fig. 6-1 MC simulation framework ...... 135 Fig. 6-2 Generic charging profile of lithium-ion EV battery ...... 136 Fig. 6-3 Framework for the determination of EV charging profiles ...... 137 Fig. 6-4 Traffic distribution in a typical workday ...... 139 Fig. 6-5 Probability distribution of the EV charging start time ...... 139 Fig. 6-6 Probability distribution of BEV battery capacity ...... 140 Fig. 6-7 Probability distribution of PHEV battery capacity ...... 140 Fig. 6-8 Probability distribution of daily travel distance ...... 141

Fig. 6-9 Charging profile when Qc is larger than ST ...... 143 Fig. 6-10 Charging profile when Qc is less than ST ...... 143 Fig. 6-11 Three EV charging profiles ...... 144 Fig. 6-12 Three random demand profiles ...... 145 Fig. 6-13 Three HP electric demand profiles ...... 146 Fig. 6-14 Detailed MC simulation algorithm ...... 148 Fig. 6-15 Voltage profiles at LV feeder end with no connection of EVs and HPs ...... 150 Fig. 6-16 Voltage profiles at LV feeder end in balanced scenario...... 151 Fig. 6-17 Voltage profiles at LV feeder end in non-balanced scenario ...... 152

Page | 7 List of Figures

Fig. 6-18 U2 profile at LV feeder end...... 152 Fig. 6-19 U0 profile at LV feeder end...... 153 Fig. 6-20 Mean profiles of phase-A voltage in balanced scenario ...... 154 Fig. 6-21 95th % profiles of phase-A voltage in balanced scenario ...... 154 Fig. 6-22 Valley values of phase-A voltage mean and 95th % profiles ...... 155 Fig. 6-23 Mean profiles of U2 in balanced scenario ...... 155 Fig. 6-24 95th % profiles of U2 in balanced scenario...... 156 Fig. 6-25 Peak values of U2 mean and 95th % profiles ...... 156 Fig. 6-26 Mean profiles of U0 in balanced scenario ...... 157 Fig. 6-27 95th % profiles of U0 in balanced scenario...... 157 Fig. 6-28 Peak values of U0 mean and 95th % profiles ...... 157 Fig. 6-29 Valley values of phase-A voltage in non-balanced scenario ...... 158 Fig. 6-30 Peak values of U2 in non-balanced scenario ...... 159 Fig. 6-31 Peak values of U0 in non-balanced scenario ...... 159 Fig. 7-1 Valley of Va 95th % profile and peak of U2 95th % profile at each LCP in case 5 ...... 168 Fig. 7-2 Investigated LV feeder F1 with ST based balancing system installed ...... 169 Fig. 7-3 Phase-A voltage profile with implementation of TOU tariff ...... 170 Fig. 7-4 U2 profiles with implementation of TOU tariff ...... 171 Fig. 7-5 U0 profiles with implementation of TOU tariff ...... 171 Fig. 7-6 Valley values of Va (a) mean and (b) 95th % profiles with implementation of TOU tariff ...... 172 Fig. 7-7 Peak values of U2 (a) mean and (b) 95th % profiles with implementation of TOU tariff ...... 173 Fig. 7-8 Voltage profiles at LCP14 with implementation of ST based balancing system ...... 175 Fig. 7-9 Voltage profiles at LCP23 with implementation of ST based balancing system ...... 176 Fig. 7-10 Voltage profiles at LCP21 with implementation of ST based balancing system ...... 176 Fig. 7-11 U2 profiles with implementation of ST based balancing system ...... 177 Fig. 7-12 U0 profiles with implementation of ST based balancing system ...... 177 Fig. 7-13 Tap position of the three tap changers (a) TAP1 (b) TAP2 and TAP3 ...... 178 Fig. 7-14 Total tap changing number and U2 at the feeder end ...... 180 Fig. 7-15 Valley of Va mean profiles at LCP21with/without implementation of ST based balancing system ...... 181 Fig. 7-16 Valley of Va 95th % profiles at LCP21 with/without implementation of ST based balancing system ...... 182 Fig. 7-17 Peak of U2 mean profiles at LCP21 with/without implementation of ST based balancing system ...... 182 Fig. 7-18 Peak of U2 95th % profiles at LCP21 with/without implementation of ST based balancing system ...... 183 Fig. 7-19 Peak of U0 mean profiles at LCP21 with/without implementation of ST based balancing system ...... 183

Page | 8 List of Figures

Fig. 7-20 Peak of U0 95th % profiles at LCP21 with/without implementation of ST based balancing system ...... 183 Fig. 7-21 Valley of Va mean profiles at LCP21 with cooperation of ST balancing system and TOU tariff ...... 184 Fig. 7-22 Valley of Va 95th % profiles at LCP21 with cooperation of ST balancing system and TOU tariff ...... 185 Fig. 7-23 Peak of U2 mean profiles at LCP21 with cooperation of ST balancing system and TOU tariff ...... 185 Fig. 7-24 Peak of U2 95th % profiles at LCP21 with cooperation of ST balancing system and TOU tariff ...... 186 Fig. 7-25 Peak of U0 mean profiles at LCP21 with cooperation of ST balancing system and TOU tariff ...... 186 Fig. 7-26 Peak of U0 95th % profiles at LCP21 with cooperation of ST balancing system and TOU tariff ...... 186

Page | 9 List of Tables

List of Tables

Table 2-1 Characteristics of the LV network in the UK ...... 31 Table 2-2 Values of investigated expressions when phase-C voltage is changed ...... 44

Table 4-1 Comparison between ST based balancing method and 3Φ balancing method ...... 102 Table 4-2 Three cases investigated in the simulation ...... 103 Table 4-3 Configuration for the phase regulating system ...... 103 Table 4-4 Configuration for the voltage magnitude regulating system ...... 103 Table 4-5 Voltage magnitude at two sides of ST based balancing system ...... 104 Table 4-6 Voltage magnitude at two sides of 3Φ balancing system ...... 104 Table 4-7 Values of U2s ...... 104 Table 4-8 Values of U0s ...... 105 Table 4-9 Tap changing number in case 1 ...... 105 Table 4-10 Tap changing number in case 2 ...... 105 Table 4-11 Tap changing number in case 3 ...... 105 Table 4-12 Three phase voltages in case I ...... 110 Table 4-13 The U0s and U2s in case I ...... 111 Table 4-14 Phase-A voltage in case II ...... 112 Table 4-15 Phase-C voltage in case II...... 112 Table 4-16 The U0 in case II...... 113 Table 4-17 The U2 in case II...... 113

Table 6-1 Details of four introduced EV charging modes ...... 138 Table 6-2 Number of EVs and HPs in 96-customer LV feeder ...... 149 Table 6-3 Periods of violations in non-balanced scenario ...... 160

Table 7-1 DR based technique types ...... 164 Table 7-2 Configuration of TAP1 (phase regulating system) ...... 174 Table 7-3 Configuration of TAP2 and TAP3 (magnitude regulating system) ...... 175 Table 7-4 Simulation results with TAP2 and TAP3 having 64 taps ...... 179 Table 7-5 Simulation results with TAP2 and TAP3 having 32 taps ...... 180

Page | 10 Abbreviations

Abbreviations

Abbreviations Definition ASD Adjustable speed drive AVC Automatic voltage control BEV Battery electric vehicle CHP Combined heat and power COP Coefficient of performance

CO2 Carbon dioxide CT Current transformer DNO Distribution network operator DR Demand response D-STATCOM Distribution static compensator Dyn delta-wye-ground e.m.f. Electromotive force ER Engineering Recommendation EU European Union EV Electric vehicle HV High voltage HP Heat pump IEEE The Institute of Electrical and Electronics Engineers IGBT Insulated-gate bipolar transistor LCP Load connection point LDC Line drop compensation LV Low voltage LVUR Line voltage unbalance rate MC Monte Carlo MV Medium voltage NEMA National Equipment Manufacturer’s Association

Page | 11 Abbreviations

OLTC On-load tap changer PCC Point of common coupling PHEV Plug-in hybrid electric vehicle pu Per unit PVUR Phase voltage unbalance rate PWM Pulse-width modulation RES Renewable energy source RMS Root mean square RMU Ring main unit ST Scott transformer TOU Time of use UKGDN United Kingdom generic distribution network VAr Voltage-ampere reactive VT Voltage transformer 1Φ Single-phase; 1Φ voltage refers to the phase-phase voltage in three-phase three-wire systems, but the phase-neutral voltage in three-phase four-wire systems

3Φ Three-phase; 3Φ voltage refers to the three phase- phase voltages in three-phase three-wire systems, but the three phase-neutral voltages in three-phase four- wire systems

Va Phase-A voltage, referring to the phase-A to neutral voltage in LV networks

Vb Phase-B voltage, referring to the phase-B to neutral voltage in LV networks

Vc Phase-C voltage, referring to the phase-C to neutral voltage in LV networks

Page | 12 Abstract

Abstract

Title: Voltage Balancing on Three-phase Low Voltage Feeder Candidate: Yun Li Institute: The University of Manchester Degree: Doctor of Philosophy (PhD) Date: February 2015

Voltage imbalance in low voltage (LV) networks is expected to deteriorate as low carbon technologies, e.g. electric vehicles (EVs) and heat pumps (HPs) are increasingly deployed. The new electrical demand attributable to EVs and HPs would increase the voltage magnitude variation, increasing the possibility of voltages moving outside the statutory LV magnitude limits. Moreover, the 1Φ nature of EVs and HPs, which will be connected via a 1Φ ‘line & neutral’ cable to a 3Φ four-wire LV mains cable buried beneath the street, further entangles this voltage management problem; the non-balanced voltage variations in the three phases boost the levels of voltage imbalance. Excessive voltage imbalance and magnitude variation need to be mitigated to limit their adverse effects on the electric network and connected plant. The voltage imbalance in LV networks is conventionally reduced by reinforcing the network, generally at a high cost. Some modern methods for voltage imbalance mitigation have been introduced in recent years. The power electronic converter based methods are inadequate due to the generation of harmonics, significant power losses and short lifetime. Besides, automatic supply phase selection and smart EV charging rely on an advanced smart communication system, which currently is not available. This project aims to develop alternative solutions that mitigate the voltage imbalance seen in LV networks. A voltage balancing method based on Scott transformer (ST) is proposed. This method does not generate harmonics and is independent of the smart communication system. Computer simulations demonstrated the proposed method is able to convert a non- balanced 3Φ voltage into a balanced 3Φ voltage at either a point on the LV feeder or a 3Φ load supply point with the predefined voltage magnitude. Besides, a physical voltage balancing system was created based on the proposed method and it was tested in an LV network in the laboratory. The test results show the balancing system is capable of maintaining a low level of voltage imbalance on the LV feeder by rapidly compensating for the voltage rises and sags caused by 1Φ load variations. This voltage balancing method is a potential solution for the network utilities to accommodate the significant penetration of low carbon technologies without breaching the network voltage limits. The impact of EVs and HPs on the LV network voltages is investigated based on a Monte Carlo (MC) simulation platform, which comprises a statistical model of EV charging demand, profiles generators of residential and HP electrical demand, and a distribution network model. The MC simulation indicates the impact of EVs and HPs is related to their distribution; when more than 21EVs and 13HPs are non-evenly distributed on a 96-customer LV feeder, the voltage limits are likely to be violated. Moreover, the effectiveness of the ST based voltage balancing method and the demand response based TOU tariff, implemented either alone or together, in mitigating the impact of EVs and HPs is investigated based on the established MC simulation platform. The results indicate the ST based balancing method alone is able to completely mitigate the voltage limit violations regardless of the penetration levels of EVs and HPs. Moreover, using both of the two investigated methods further enhances the balancing effectiveness of the ST based voltage balancing method.

Page | 13 Declaration

Declaration

No portion of the work referred to in the thesis has been submitted in support of an application for another degree or qualification of this or any other university or other institute if learning.

Page | 14 Copyright Statement

i. The author of this thesis (including any appendices and/or schedules to this thesis) owns certain copyright or related rights in it (the “Copyright”) and s/he has given The University of Manchester certain rights to use such Copyright, including for administrative purposes. ii. Copies of this thesis, either in full or in extracts and whether in hard or electronic copy, may be made only in accordance with the Copyright, Designs and Patents Act 1988 (as amended) and regulations issued under it or, where appropriate, in accordance with licensing agreements which the University has from time to time. This page must form part of any such copies made. iii. The ownership of certain Copyright, patents, designs, trade marks and other intellectual property (the “Intellectual Property”) and any reproductions of copyright works in the thesis, for example graphs and tables (“Reproductions”), which may be described in this thesis, may not be owned by the author and may be owned by third parties. Such Intellectual Property and Reproductions cannot and must not be made available for use without the prior written permission of the owner(s) of the relevant Intellectual Property and/or Reproductions. iv. Further information on the conditions under which disclosure, publication and commercialisation of this thesis, the Copyright and any Intellectual Property and/or Reproductions described in it may take place is available in the University IP Policy (see http://documents.manchester.ac.uk/DocuInfo.aspx?DocID=487), in any relevant Thesis restriction declarations deposited in the University Library, The University Library’s regulations (see http://www.manchester.ac.uk/library/aboutus/regulations) and in The University’s policy on Presentation of Theses.

Page | 15 Acknowledgement

Acknowledgement

First and foremost, I would express my gratitude to my PhD supervisor, Prof. Peter Crossley, for the invaluable guidance and constant encouragement he has given freely throughout the project. His exuberant enthusiasm and optimism for life and work deeply impress me, which will benefit my future career and the rest of my life. I also thank him for his relaxed and sagacious supervision.

Thanks to the staffs and colleagues working in the Ferranti Building at the University of Manchester, thanks for their help and the laughs.

I gratefully acknowledge The University of Manchester Division of Development and Alumni Relations and School of Electrical and Electronic Engineering for providing the financial support for my PhD.

Finally, I would like to express my deepest gratitude to my family, especially my parents for their encouragement and support.

Page | 16 Publications

Publications

1. Y. Li and P. A. Crossley, “Voltage control on unbalanced LV networks using tap changing transformers,” in Proc. 2012 IET 11th International Conference on Developments on Power System Protection (DPSP 2012), 23-26 April 2012, Birmingham, UK. (Conference paper and oral)

2. Y. Li and P. A. Crossley, “Voltage balancing in low voltage distribution networks using Scott transformers,” in Proc. 2013 22nd International conference on Electricity Distribution 2013 (CIRED 2013), 10-13 June 2013, Stockholm, Sweden. (Conference paper, oral and poster)

3. Y. Li and P. A. Crossley, “Scott transformer based voltage balancing in low voltage distribution networks,” in Proc. 2013 IEEE PowerTech Conference (POWERTECH 2013), 16-20 June 2013, Grenoble, France. (Conference paper and oral)

4. Y. Li and P. A. Crossley, “Investigation of voltage balancing in low voltage networks using Scott transformers,” in Proc. 2013 Cigre Study Committee B5 Colloquium, 25-31 August 2013, Belo Horizonte, Brazil. (Conference paper and poster)

5. Y. Li and P. A. Crossley, “Monte Carlo study on impact of electric vehicles and heat pumps on LV feeder voltages,” in Proc. 2014 IET 12th International Conference on Developments on Power System Protection 2014 (DPSP 2014), 31 March – 3 April 2014, Copenhagen, Denmark. (Conference paper and poster)

6. Y. Li and P. A. Crossley, “Impact of electric vehicles on LV feeder voltages,” in Proc. 2014 IEEE PES general meeting (PESGM 2014), 27-31 July 2014, National Harbor, USA. (Conference paper and poster)

7. Y. Li and P. A. Crossley, "Voltage balancing in low-voltage radial feeders using Scott transformers," IET, Generation, Transmission & Distribution, vol. 8, pp. 1489- 1498, 2014. (Journal paper)

8. Y. Li and P. A. Crossley, "Impact of electric vehicles and heat pumps on LV feeder voltages: analysis and mitigation," submitted to IEEE Transactions on Smart Grid, 2015. (Journal paper)

Page | 17 Chapter 1 Introduction

CHAPTER 1

INTRODUCTION

1.1 Background

Voltage is one of the most important elements in a power system. Its waveshape and magnitude directly influence the quality of the electricity received by customers and therefore affect the operation and safety of the equipment belonging to electric utilities and their customers. Since the power network operators only have the ability of controlling the voltage; they have no control over the currents that particular loads might draw, the expression “Power Quality = Voltage Quality” has been commonly accepted [1]. The proximity of low voltage (LV) distribution networks to end-users significantly highlights the importance of power quality in LV distribution networks. However, this is more vulnerable than the power quality in higher voltage networks due to uncontrolled human behaviour at the LV load connection points.

Each voltage level in an electricity network has three main characteristics – RMS voltage magnitude, frequency and the waveform [2]. In the UK, these characteristics in the LV network are set ideally at 240V RMS, 50 Hz frequency with a sinusoidal waveform [3, 4]. If these characteristics are maintained within a narrow range around the nominal values, the electrical appliances can achieve their best performance. A power quality problem may be any unintended variation from the normal voltage magnitude, its frequency or its

Page | 18 Chapter 1 Introduction waveform, which gives rise to undesirable performance of the equipment. The power quality issues include voltage magnitude variations, voltage imbalance, voltage fluctuations, waveform distortion, transients and frequency variations [1].

In order to accelerate the move towards a low carbon economy, the EU has set legally binding targets on carbon emission for its member states, including a commitment for a

20% reduction on the 1990 CO2 emission levels by 2020, a 40% reduction by 2030 and a 80% reduction by 2050 [5]. In order to achieve these environmental targets in time, the UK government has established various grants and incentives to encourage the deployment of low carbon technologies (LCTs), including electric vehicles (EVs) and heat pumps (HPs). Financial support is based on the consideration that the electrification of transport and heat would significantly reduce the carbon emission of a future decarbonized electricity supply. This will result in the increasing installations of LCTs in the LV networks.

1.1.1 Electric vehicle

EVs are the vehicles using electric motors for propulsion. The electricity powering the electric motor in EVs can be generated by an on-board electrical generator or absorbed from the power grid. Only the EVs drawing electricity from the power grid are considered in this thesis and these EVs can be divided into two types – one is the plug-in hybrid EV (PHEV) which operates on power from a combination of an on-board battery and a combustion engine, and the other is the battery EV (BEV) which operates purely on the battery power. Both BEVs and PHEVs receive the electricity by plugging into the power grid and save it in batteries. This is often referred to as Grid-to-Vehicle, where the EVs can be treated as active electrical loads. Besides, Vehicle-to-Grid is technologically possible, i.e. the EV batteries can inject the stored energy back into the grid. In this mode, the EVs are used as energy storage systems to provide ancillary services to the grid [6].

In 2013, transportation accounted for 25% of the UK’s CO2 emissions [7], and therefore, carbon reduction in the transportation sector plays an important role in the fulfilment of the environmental targets. When the EVs are driven using electric power, no petroleum based fuel is consumed and no CO2 is produced. For this reason, EVs have been treated

Page | 19 Chapter 1 Introduction

as one promising technology to reduce the CO2 emissions and help reduce the impact of increasing petroleum fuel prices on societies. Many automotive manufacturing companies have placed an emphasis on research and development of EVs, and a wide range of EVs are currently available as summarized in Appendix F. The number of EVs in the UK is estimated to rise to 12.8 million by 2030, equal to 41.3% of estimated UK 2030 household number [8].

1.1.2 Heat pump

Most of the energy used in our society is transferred to heat, half of which is consumed in residential space and water heating [9]. Currently, burning fossil fuels is the major way to obtain heat (around 80% from gas alone) [10]; as a result, space and water heating for buildings accounts for 19% of the UK’s CO2 emission[11]. This technological paradigm must change if deep CO2 emission cuts are to be achieved [9].

A heat pump (HP) is an appliance having the same principle of operation as a refrigerating machine. It takes low temperature heat energy and transfers it into higher temperature heat by consuming electrical energy. Based on the fact that liquids absorb a large amount of energy when evaporating and this energy is released when the vapour condenses back to liquid, HPs are able to absorb and release large amounts of energy [10]. The key parameter expressing the HP performance is the coefficient of performance (COP), which indicates how many units of heat are delivered by consuming one unit of electricity. Using HPs for domestic space and water heating is an effective way to reduce the CO2 emission [10].

According to the energy source, HPs are mainly divided into air source HPs, ground source HPs and water source HPs; these absorb energy from the air, the ground and the water respectively [10]. Both air source HPs and ground source HPs are commonly used, whilst a water source HP is only adopted when there is a river or lake around the dwelling. It is estimated that the uptake of HPs in the UK can reach up to 7.7 million in 2030, equal to 24.8% of the estimated UK 2030 household number [8].

Page | 20 Chapter 1 Introduction

1.2 Motivation and objectives

Both EVs and HPs will induce a substantial increase in the new electrical demand, imposing a threat on the thermal and voltage management of power systems, especially in local LV networks, where high concentrations of EVs and HPs could occur [12]. However, the network operators must connect these LCTs to the grid to facilitate the customers’ transition to a low carbon future, whilst keeping the system parameters within the statutory limits and maintaining an economic operation.

The additional electrical load, dominated by EVs and HPs, would significantly boost the voltage drop along the LV feeder, threatening to violate the statutory voltage magnitude limits. Since most appliances normally operate within a narrow band of voltage around the nominal value, the network operator must maintain the supplied voltage within the mandated limits. Maintaining the steady state voltage within acceptable limits is the most important issue when clustered EVs and HPs are connected to the LV network [2].

Moreover, the single-phase nature of these devices, which will be connected via a single- phase (1Φ) ‘line & neutral’ LV service cable (230V) to a three-phase (3Φ) four-wire LV mains cable (400V) buried beneath the street, will deteriorate the voltage imbalance, especially at downstream locations on radial LV feeders. Excessive voltage imbalance destroys the designed protective coordination and reduces the power delivery efficiency on LV feeders. Besides, 3Φ motors with a non-balanced 3Φ supply have a lower efficiency and a shorter lifetime.

Traditionally, the use of control, monitoring and communication equipment on distribution networks is limited and there is almost none in LV networks [13]. Therefore if no innovative solutions are found, network reinforcement would be necessary to ensure the LV feeder voltages remain within mandatory limits. This would result in significant network investment and delay the adoption of LCTs. The aim of the research described in this thesis was to design and develop alternative solutions, including new network technologies and flexible customer response, to mitigate the severe voltage imbalance and magnitude variation on LV feeders caused by a high penetration of EVs and HPs. To achieve this aim, the following objectives were defined:

Page | 21 Chapter 1 Introduction

 To propose a new voltage control technique – this technique should have the capability of voltage imbalance mitigation and voltage magnitude regulation.  To demonstrate the feasibility of the proposed voltage control technique – a demonstration should be carried out using both computer simulation and a laboratory based test facility.  To evaluate and quantify the impact of EVs and HPs on the LV feeder voltages – statistical models of EV and HP electrical demands should be established to evaluate their impact on LV feeder voltages from a probability perspective.  To evaluate and quantify the effectiveness of the proposed voltage control technique – this investigation should be based on a statistical study to understand the whole picture of the effectiveness of the proposed technique.

1.3 Contribution of the work

The research work presented in this thesis mainly evolved around a Scott transformer (ST) based voltage balancing method, and included its theory and operational design, demonstration using computer simulation, and demonstration using a physical test facility. In addition, a Monte Carlo (MC) simulation platform was established and used to investigate the impact of EVs and HPs on LV feeder voltages from the viewpoint of probability. The effectiveness of the ST based balancing method and a demand response (DR) based time-of-use (TOU) tariff, implemented either alone or together, in mitigating the adverse impact of EVs and HPs was analysed based on the established MC simulation platform. Fig. 1-1 illustrates the roles of the above subjects in the evolution of the research work as well as the associated publications.

Page | 22 Chapter 1 Introduction

ST based voltage balancing method

Simulation in a simplified LV network (Publication 1) MC simulation platform

Simulation in a typical UK distribution Impact of EVs on LV feeder network (Publication 7) voltages (Publication 5)

Simulation with regard to transient Impact of EVs and HPs on LV response (Publication 3, 4 ) feeder voltages (Publication 6)

Physical test (Publication 2, 7)

MC study on effectiveness of ST based voltage balancing method (Publication 8 )

Fig. 1-1 Evolution of the research work and the associated publications

The contributions of the thesis are summarized as below:

The proposal of the ST based voltage balancing method is one of the most important contributions of this thesis. This method is able to convert a non-balanced 3Φ voltage into a balanced 3Φ voltage with the predefined magnitude at either one point on the LV feeder or the supply point of a 3Φ load, helping to reduce the necessity for the network reinforcement for voltage imbalance mitigation. Moreover, compared with the recently proposed voltage imbalance mitigating methods using power electronic converters, the ST based balancing method, relying on transformers and taps changers, has many benefits, including no injection of harmonics, long lifetime, low power delivery losses and high reliability. For the DNOs, the ST based voltage balancing method is a potential solution to the excessive voltage imbalance and magnitude variation in future LV networks due to the significant penetration of LCTs.

The feasibility of the ST based voltage balancing method was firstly demonstrated in computer simulation. A model of the proposed balancing method was established in PSCAD/EMTDC. This balancing method was compared with another transformer and

Page | 23 Chapter 1 Introduction tap changer based balancing method using a simulation model of a simplified LV network, and the results illustrate its superiority. Besides, the proposed balancing system was implemented in a typical UK distribution network to investigate its performance with respect to the installation location and output voltage magnitude, and its transient response to the voltage disturbances caused by 1Φ load variations. Following the computer simulations above, a physical voltage balancing system was established based on the proposed balancing method. This balancing system is capable of automatically performing voltage balancing in LV networks, and was tested on an LV feeder in the laboratory under various different operating scenarios.

Both the computer simulation and physical test mentioned above demonstrate the feasibility of using the proposed balancing method on a real LV feeder. They provide the theory foundation for future site trials and ideas for commercial use of the ST based voltage balancing method.

The established MC simulation platform includes a statistical model of EV charging demand, profile generators of HP electrical demand and residential demand, and a distribution network model. This platform is flexible; by replacing the distribution network model and the residential demand with a specified network and its associated demand, the DNOs could evaluate the impact of EVs and HPs on the specified network and estimate the maximum acceptable penetration levels of EVs and HPs. Moreover, the MC simulation platform is extendible; the demand models of new loads or the generation models of DGs can be integrated into the platform so that the DNOs could evaluate the impact of the new loads and DGs. Moreover, the efficacy of potential network control technologies and possible network modifications can be evaluated using the MC simulation platform.

Although there have been numerous literature on the impact of EVs or HPs on power system, the studies investigating the impact of both EVs and HPs on LV networks are limited. This gap is filled in by the work presented in this thesis - evaluation of the impact of EVs and HPs on the LV feeder voltages - which was carried out based on the MC simulation platform. The results give the DNOs a generalized understanding of the impact of EVs and HPs on LV feeder voltage imbalance and magnitude variation.

Page | 24 Chapter 1 Introduction

Finally, the ST based voltage balancing system was installed on the investigated LV feeder in the MC simulation platform, and its effectiveness in mitigating the adverse impact of EVs and HPs was investigated from the viewpoint of probability. The statistical nature of the MC method is supposed to reflect the close-to-realistic scenario of a future LV network and deliver a global picture of the effectiveness of the proposed method. Therefore, the results of this study further enhance the theory foundation for the future site trial and commercial use of the ST based voltage balancing method.

1.4 Structure of the thesis

The thesis contains seven chapters after this one, and these were summarized as follows:

Chapter 2 describes the LV network, the voltage imbalance and the adopted voltage limits in detail to give a context for the research problem. Moreover, the literatures on both conventional and modern voltage imbalance mitigating methods are reviewed as a preface for the description of the ST based voltage balancing method in Chapter 3.

Chapter 3 introduces the ST based voltage balancing method. This method is the heart of this thesis, and this project revolves around it. In section 3.2, the transformer and tap changer based voltage regulation is summarized and presented in detail. Following this, the ST transformer based voltage balancing method is described in terms of its main components and the control algorithm.

Chapter 4 describes two computer simulation studies investigating the ST based voltage balancing method and both of them are carried out using PSCAD/EMTDC. The first study compares the proposed voltage balancing method with another transformer and tap changer based voltage balancing method. The second study investigates the performance of the ST based voltage balancing method on a typical UK distribution network.

Chapter 5 presents a physical test of the proposed voltage balancing method. A physical voltage balancing system is established based on the ST based voltage balancing method, and implemented on an LV feeder in the laboratory used for voltage regulation under

Page | 25 Chapter 1 Introduction different operating scenarios. In addition, a computer simulation study was used to model similar scenarios and verify the results of the physical test.

Chapter 6 describes an MC study investigating the impact of EVs and HPs on LV feeder voltages. The simulation methodology and the MC simulation platform are depicted; especially, the statistical model of the EV charging demand is presented in detail. Following this, the simulation results are analysed. This study is carried out in MATLAB and OpenDSS.

Chapter 7 investigates the effectiveness of the ST based voltage balancing method and the DR based TOU tariff in mitigating the adverse impact of EVs and HPs on the LV feeder voltages based on the established MC simulation platform. Both deterministic and probabilistic approaches are adopted in this study by using MATAB and OpenDSS. The effectiveness of the two investigated method, implemented either alone or together, is presented and discussed.

Chapter 8 concludes the thesis by summarizing the results of the work. Finally, possible future research ideas associated with this thesis are presented.

Page | 26 Chapter 2 Voltage Imbalance in LV Networks

CHAPTER 2

VOLTAGE IMBALANCE IN LV NETWORKS

2.1 Introduction

This chapter includes three parts: an introduction to the LV network, an introduction to voltage imbalance and a literature review on voltage imbalance mitigation. The main aim of the research project described in this thesis was to develop an alternative solution to excessive voltage imbalance and magnitude variation in LV networks caused by the penetration of EVs and HPs. The design and operation of the LV network and the resulting voltage imbalances are introduced in detail to give a context for the research problem. Moreover, the literature on both conventional and modern voltage imbalance mitigating methods are reviewed in section 2.4 as a preface for the description of the ST based voltage balancing method in Chapter 3. The limits with regard to LV network voltage imbalance and magnitude, as adopted in this project, are described in section 2.5. In addition, section 2.6 summarizes the content of this chapter and discusses the voltage imbalance mitigating methods reviewed in section 2.4.

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2.2 Low voltage network

LV network is defined as a network whose voltage upper limit is 1kV in the British Standard [3]. As the end delivery vehicle of an electrical power network, the LV network is directly extended towards the customers’ properties and supplies a massive number of dispersed small-scale loads; so it has the characteristics of small individual capacity but significant quantity. Because of the low voltage level, an individual LV feeder requires less finance in construction and improvement, than higher voltage feeders, but the large number of LV feeder, and consequently the huge amount of work could consume the majority of a utility’s capital [14]. Moreover, the service and appliances of LV networks are located close to human beings and animals, and therefore, strict government and utility regulations are necessary to cover their operational safety. The degree of the control imposed by such regulations, the specific situation, the load conditions and even the individual planner’s preference have a great influence on LV network design and cost [14].

2.2.1 LV network in the UK

The LV network in the UK is generally a 3Φ 4-wire system supplied from a single 3Φ Dyn MV/LV transformer and its nominal voltage RMS magnitude is 230V for phase- neutral voltage and 400V for phase-phase voltage [3]. The employment of 3Φ secondary feeder provides well-balanced loading on the service level, reducing losses and increasing utilization. Fig. 2-1 shows the circuit diagram of a UK distribution network, where the LV feeders are illustrated in detail [15, 16]. This layout is suitable for areas having high load densities.

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Bulk supply point 132/11kV substation Each LV customer represent many similar units, evenly distributed among the three phases 11/0.4 kV substation A B C 11kV feeder: 3Ф 3- Neutral wire system; it can Earthing Separate neutral and protective earth conductors also be supplied from a primary substation (33/11kV substation)

A B C Neutral

Common neutral and protective earth conductor A B C with multiple earths

LV customer

Fig. 2-1 Typical UK distribution system [15]

Secondary substation The 11kV/400V transformers in the secondary substations are generally Dyn connected and off-load tap changers are provided for voltage adjustment [17]; note: an off-load tap changer can only be changed when the transformer is de-energized. The optimum MV/LV transformer size varies between 0.5 to 2.0MVA in urban areas where there is a high load density [14].

In order to enhance the reliability of the power supply, a looped MV network, as illustrated diagrammatically in Fig. 2-2, is often used so that the secondary substation has two supplies [14]. Ring main units are usually used for the connection between secondary MV/LV transformers and the MV feeder [18]. It is a standard piece of switchgear in distribution systems, comprising switches for switching power cable rings and switches in series with fuses for the protection of distribution transformers, as portrayed in Fig. 2-3 [14]. The modern RMU is generally an enclosed unit, containing

Page | 29 Chapter 2 Voltage Imbalance in LV Networks the switching devices, and the SF4 gas is filled into the enclosure as the arc quenching medium [14, 19].

MV busbar

MV/LV transformer

Open point

Fig. 2-2 Loop MV network arrangement [14]

To MV To MV/LV To MV feeder transformer feeder Fig. 2-3 Ring main unit [14]

The detailed connection of a single-transformer secondary urban substation is portrayed in Fig. 2-4 [14]. The MV switches A are used for the protection of the distribution transformer and the MV cable by breaking a current higher than that determined by the thermal capability of the MV cable, or of the distribution transformer [14]. The switch fuse B is used to protect against the transformer overload [14]. In order to improve the service quality and investment optimisation, the utilities in UK tend to adopt circuit breaker with self-powered protection relay to replace the switch fuse [20]. The transformer LV switch C and the downstream LV fuses are used to protect against overloads and faults on the LV system. The LV unit substation, which consolidates the

Page | 30 Chapter 2 Voltage Imbalance in LV Networks incoming MV feeders, the MV/LV transformer and the LV distribution board as an integrated step-down substation, is the dominant form of a secondary substation [21, 22].

C A A A B

Fig. 2-4 Diagram of an urban single-transformer MV/LV substation [14]

LV network arrangement Most LV networks are arranged as multi-branched radial feeders, which comprise 4-wire (3Φ+Neutral) underground cables or overhead lines extending from the secondary substation [14]. The detailed network practice varies with the district concerned, and no ‘standardized’ values can be given. Table 2-1 summarizes the characteristics of the LV network in the UK [23].

Table 2-1 Characteristics of the LV network in the UK [23]

Area Urban Rural Min 2 1 Feeder number per Med. 6 3 transformer Max 16 6 Min 10 100 LV feeder length (m) Med. 100 250 Max 200 600 Min 150 500 Total line length per Med. 450 1600 transformer (m) Max 900 3200

In urban areas, it is usual to use underground cables for the LV system because of the high housing density. In this case, LV cables from neighbouring substations can

Page | 31 Chapter 2 Voltage Imbalance in LV Networks terminate close to each other, which permits a low-cost interconnection of the LV cables via an underground link box [17]. Since this arrangement provides a very flexible system; a complete substation can be taken out of service and the area, normally supplied from it, is fed by the surrounding substations via link boxes; i.e. looped networks as shown in Fig. 2-5 are usually adopted in urban areas [14]. Mains cables are normally laid along one side of the road and road crossings are employed to service properties on the opposite side. Double-sided mains can be used to accommodate large concentrations of load or as a means of reinforcing existing developments [17].

11kV

0.4kV

Fig. 2-5 Typical LV network in urban areas [14]

In rural areas, the housing density is low; the average distance between individual homes can be a few kilometres and occasionally only a single domestic or agricultural customer is supplied from a secondary substation. Thus it is not worth adopting looped LV systems or providing interconnection between substations, and a radial arrangement as shown in Fig. 2-6 is used. For customers, who require a higher level of reliability, two in-feeds, preferably from two independent MV sources or involving a diesel generator based standby supply can be adopted [17].

Page | 32 Chapter 2 Voltage Imbalance in LV Networks

kV 11

0.4kV

Fig. 2-6 Typical LV network in rural areas [14]

The majority of existing underground cables are paper insulated, lead covered, steel tape armoured construction. Most are 4-core (3 phase + neutral) but areas with 5-core (3 phase + street light + neutral), 2-core (1 phase + neutral) and 3-core (2 phase + neutral) also exist. A large proportion of existing overhead line is of bare wire construction [17]. In the UK, all new LV networks tend to adopt underground cables, except where the use of underground cables is not reasonably practical, which is often the case in rural areas.

Service connection The method use which customers are connected to the LV network varies as much as the LV network layout and depends mainly on the type of network (overhead or underground), the load conditions and the local regulations. A few methods of connecting detached or terraced houses to the underground cable network are schematically illustrated in Fig. 2-7 [14].

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(i)

(ii)

(iii)

(iv)

(v) (vi)

Disconnection box customer Underground fixed box

Fig. 2-7 Service cable arrangements [14]

The disconnection boxes are used in arrangement (i) and (ii) to provide facilities to link up the service cables. Since there is no linking connection between customers in arrangement (i), it is possible to provide fuses on each service cable supplying individual customers; whilst arrangement (ii) does not permit such individual good protection facilities as arrangement (i). Arrangement (iii) is suitable for individual large or remote customers. In arrangement (iv), fixed underground joints are used to connect up the service cables, which are cheaper than cable disconnection boxes or cabinets but selective protection mentioned earlier is not possible. Arrangements (v) and (vi) are variations of arrangement (iv) and they are selectively used according to the situation to achieve a lower cost if the local safety regulations are satisfied [14].

The number of customers connected to an LV network is limited by the voltage drop along the LV feeder and the power capacity of the network, for instance, in Central Network, UK, each LV feeder is allowed to supply at most 100 customers [17, 24]. Commercial and industrial properties normally have 3Φ services, whilst domestic properties have 1Φ service and these are distributed across the phases of the mains cable as evenly as possible to balance the load.

Page | 34 Chapter 2 Voltage Imbalance in LV Networks

In the past, service components and metering equipment were always located inside the consumer’s premise. However, the modern tendency is to install these components in a weatherproof housing outside the building unless it is prohibited by local planning regulators. The utility/consumer interface is often at the outgoing terminals of the meter, or sometimes, at the outgoing terminals of the installation main circuit breaker [15]. When the use of internal services is necessary, the LV circuits are terminated at a cut-out immediately inside a central service/metering position and lateral connection to the flats/rooms is provided and owned by the occupier [17].

In the UK, electricity mains was previously required to be delivered at 240V, within a tolerance of +6%/-6%, that is, within the range from 225.6V to 254.4V. However, in 1995, following the voltage harmonization across Europe, the law in the UK stated that the LV supply voltage must be within 230V +10%/-6%, which lays within 230V +/-10% stated by the European standard, and avoids the cost of replacing or adjusting all the electricity supply equipment in the UK [4, 25, 26]. Since there is almost no change in the allowable voltage range, DNOs made no change in network operation; they still supply the electricity at 240V in LV networks. In the practical application, voltages at the remote end of the LV feeder are generally kept above 230V, as the 6% below 230V is generally used for contingency, e.g. 11kV or LV back-feeding may use part of or the entire 6% voltage drop [17].

2.2.2 LV networks around the world

The LV network and its upper supplying system around the world have evolved into different forms, among which, the two most popular layouts: are referred to as ‘European’ and ‘American’ layouts [16, 27]. The UK LV network described above is actually the ‘European’ layout, adopted in almost all European countries. Fig. 2-8 shows the layout of the ‘American’ system.

Page | 35 Chapter 2 Voltage Imbalance in LV Networks

138/12.47kV substation

Each MV/LV transformer and its associated LV feeder represent many similar units

1Ф primary lateral A Neutral MV/LV transfromer: 1Ф 12.47kV feeder: 7.2/0.12-0.24kV 3Ф 4-wire system 3-wire LV feeder

B Neutral

A B C Neutral

LV customer

Fig. 2-8 American distribution system [15, 16]

The ‘American’ system is radially operated similar to the ‘European’ system. However, distribution in the ‘American’ system is effectively carried out at MV; the MV system is a 3Φ four-wire system from which many 1Φ primary laterals deliver power into various neighbourhoods, supplying numerous 1Φ MV/LV transformers. The result is that the MV/LV transformers are much smaller than those in ‘European’ systems – typically 25kVA or 50kVA 1Φ units – only supplying one or several premises, and the LV feeders are minimised. The LV feeder comprises of three wires; two ‘hot wires’ are extruded from the two ends of the 1Φ MV/LV transformer secondary at ±120V and a neutral wire from the earthed secondary midpoint, and therefore, the domestic appliances can be connected either at 120V from one hot wire to neutral, or at 240V between the two hot wires. The chief advantage of this design is the low capital cost when the load density is low. However, the low secondary voltage, about half of the European secondary voltage, seriously limits the extension of the LV network; at 120V, power can travel efficiently

Page | 36 Chapter 2 Voltage Imbalance in LV Networks only up to 200 feet (about 60m) [16]. A circuit with twice voltage can reach four times for any given load and voltage drop limitation. Moreover, since a 1Φ circuit has voltage drop in both phase and neutral conductors, whilst a balanced 3Φ circuit has voltage drop only in phase conductors, a balanced 3Φ circuit can reach twice as far as a 1Φ circuit. Thus an ‘European’ secondary can carry roughly eight times the burden of an ‘American’ system [16].

Please note the terms ‘European’ and ‘American’ systems are used to identify two distinctly different ways of laying out the lower section of the distribution system [16]. ‘European’ system means a 400V (phase to phase) 3Φ LV feeder from which a 1Φ 230V (phase to neutral) is provided to customers. ‘American’ system means 120/240V 1Φ service along with the conversion to 1Φ circuitry at the primary level, using 1Φ laterals in many areas [16]. However, it must be noted, 3Φ secondary services are widespread in many urban parts of America and there are areas of 1Φ distribution in rural parts of Europe [16].

Both ‘European’ and ‘American’ layouts are widespread beyond Europe and North America, and the usage typically follows colonial patterns with European practice being more widely adopted [27]. The distribution system in some regions of the world is a mixture, using bits of ‘American’ and bits of ‘European’ practices [27]. Fig. 2-9 shows a summary of the schematically circuit diagrams of the LV supply used around the world [15].

In Fig. 2-9, (a) refers to the 3Φ service and (b) refers to the 1Φ service of the ‘European’ system. They are commonly used all over the world to provide an electricity service of 220/380V, 230/400V, 240/415V or 250/440V. In some countries like the UK, generally only 1Φ service is available in domestic properties and the commercial and industrial properties have access to both 1Φ and 3Φ services; whilst in others, both 1Φ and 3Φ services are provided for domestic usage. Additionally, the ‘European’ system is mixed with a secondary voltage of 127/220V or 120/208V mainly in Middle Asia, Middle East and the USA [15]. This application can also be found in South Europe, but they are being or have been replaced by the 3Φ 4-wire 230/400V system for the voltage harmonization in Europe [15, 28]. Moreover, both circuit (c) and (d) are variations of circuit (a), and they are occasionally used in a few countries for domestic or commercial service.

Page | 37 Chapter 2 Voltage Imbalance in LV Networks

(a) 3Ф star; 4-wire: (b) 1Ф; 2-wire: (c) 3Ф; 3-wire (d) 2Ф star; 3-wire: Earthed neutral Earthed end of phase Earthed neutral

(e) 1Ф; 3-wire: (f) 3Ф star; 3-wire: (g) 3Ф delta; 3-wire (h) 3Ф delta; 4-wire: Earthed mid point Earthed neutral Earthed mid-point of one phase

(i) 3Ф open delta; 4- (j) 1Ф; 2-wire: (k) 1-wire: Earthed wire: Earthed mid Unearthed return point of one phase

Fig. 2-9 Circuit diagrams of the LV supply around the world [15]

The domestic LV service of ‘American’ system is illustrated by circuit (e) in Fig. 2-9, which is mainly employed in North America, Latin America and a few countries in Asia with a secondary voltage of 110/220V, 115/230V or 120/240V. The circuits (f), (g), (h) and (i) are mainly used in the commercial and industrial installations in North America [15, 29].

In Mexico, circuit (j) is adopted to provide a 120V supply in domestic premises [15]. Besides, circuit (k) – single-phase ‘earth return’ distribution – is suitable in very sparsely populated regions where the distance between customers is measured in kilometres and the requirement for only a single wire can significantly reduce the capital cost of the primary voltage system [16]. This type of supply system has been used in Australia, North Canada and parts of India and China [15, 16].

Page | 38 Chapter 2 Voltage Imbalance in LV Networks

2.3 Voltage imbalance

2.3.1 Introduction

A 3Φ power system is only called balanced or symmetrical when the 3Φ voltages and currents have the same magnitude and are phase shifted by 120° with respect to each other [1, 30]; If either or both of these requirements are not satisfied, the system is called non-balanced or asymmetrical. Voltage imbalance is a recognised power quality parameter, and in a 3Φ system it means the 3Φ voltages have unequal magnitudes or the phase differences between consecutive voltages are not 120°. Fig. 2-10 shows the phase diagrams of a balanced 3Φ voltage and an arbitrary non-balanced 3Φ voltage. The primary cause of voltage imbalance is the non-even distribution of 1Φ loads that can be continuously changing, so the perfect voltage balance can never be maintained. In the recent past, the voltage imbalance problem has triggered a growing concern for power quality issues in LV distribution networks due to the domestic adoption of LCTs and their non-uniform spacial and time distribution across an LV feeder.

Vb Va Va

Vc Vc (a) (b)

Fig. 2-10 Phase diagrams of balanced and non-balanced 3Φ voltages (a) balanced 3Φ voltage (b) non-balanced 3Φ voltage

Page | 39 Chapter 2 Voltage Imbalance in LV Networks

2.3.2 Expressions

Symmetrical components significantly simplify the analysis of non-balanced polyphase networks and balanced networks with non-balanced terminal conditions. Any non- balanced 3Φ voltages can be built up in an analytical manner from three fundamental sequence components: positive, negative and zero [31, 32]. Equation (2-1) and (2-2) shows the relationship between phase and sequence voltages in a 3Φ system, where Va,

Vb and Vc are the 3Φ voltages and V1, V2 and V0 are the positive (V1), negative (V2) and zero (V0) sequence voltage component. The rotation operator 푎 is given by (2-3) [31].

V0 1 1 1 Va 1 2 [V1] = ∙ [1 푎 푎 ] ∙ [Vb] (2-1) 3 2 V2 1 푎 푎 Vc

Va 1 1 1 V0 2 [Vb] = [1 푎 푎 ] ∙ [V1] (2-2) 2 Vc 1 푎 푎 V2

푎 = 푒푗∙120° (2-3)

The ratio of zero to positive sequence components (U0) and the ratio of negative to positive sequence components (U2) are employed in [33] to express the phase-to-neutral or phase-to-earth voltage imbalance in power system. U0 and U2, expressed as a percentage, are evaluated by (2-4) and (2-5) respectively. Both of them involve voltage magnitude as well as phase angle when calculating the voltage sequence components [34].

V U0 = 0 × 100% (2-4) V1

V U2 = 2 × 100% (2-5) V1

The U0 by definition is zero when the phase-to-phase voltages are measured which has no zero sequence component. However, the 3Φ four-wire system may have zero sequence component and both U2 and U0 can be used to express the voltage imbalance.

Page | 40 Chapter 2 Voltage Imbalance in LV Networks

For a three-phase-to-neutral voltage, its U2 is always equal to the U2 of the associated three-phase-to-phase voltage.

In order to avoid the use of complex algebra, (2-6) and (2-7) are introduced in [33] and

[35] separately to give a good approximation to U2; and Vij in (2-6) and (2-7) refers to the phase i to phase j voltages.

1− 3−6β V4 + V4 + V4 vuf1 = √ √ × 100% with 훽 = ab bc ca (2-6) 1+ √3−6β 2 2 2 2 (Vab+ Vbc+ Vca)

2 2 2 6×(Vab+ Vbc+ Vca) vuf2 = √ 2 − 2 (2-7) (Vab+ Vbc+ Vca)

The expression of voltage imbalance developed by National Equipment Manufacture’s Association (NEMA) is the line voltage unbalance rate (LVUR), given by (2-8) [36]. This definition assumes that the average is always equal to the rated value, and it works only with magnitudes, i.e. phase angles are not included.

max voltage deviation from the avg line voltage LVUR = ∙ 100% (2-8) avg line voltage

The Institute of Electrical and Electronics Engineers (IEEE) defines the voltage imbalance as the phase voltage unbalance rate (PVUR), given by (2-9) [37]. This expression is the same as LVUR; and again the phase angle information is lost since only magnitudes are considered. The only difference between the IEEE and NEMA definitions is that IEEE uses phase voltages rather than line-to-line voltages [34].

max voltage deviation from the avg phase voltage PVUR = ∙ 100% (2-9) avg phase voltage

Engineering Recommendation (ER) P29 ‘Planning limits for voltage unbalance in the United Kingdom’ states (2-10) can be used to estimate the value of U2 when the voltage imbalance is caused by a 1Φ load connected between phases. Moreover, in a power system supplying a high-speed electrified railway, (2-10) is usually used to predict the U2 due to 1Φ traction loads connected between two of the three phase lines [38-40]. The

Page | 41 Chapter 2 Voltage Imbalance in LV Networks

ratio only uses the apparent power of the load (SL) and the short-circuit power of the supply circuit (SSC).

S %voltage imbalance ≈ L ∙ 100% (2-10) Ssc

Among the above introduced expressions for voltage imbalance, U2, vuf1, vuf2 and LVUR involve the positive and negative sequence components of the 3Φ voltage, but ignore the zero sequence component. For a 3Φ voltage with a specified LVUR, assuming

Vab deviates the most from the average line-to-line voltage, the magnitude of Vab can be determined, and the magnitude of Vca can be obtained if the magnitude of Vbc is given

[34]. Moreover, assuming the phase angle of Vab is zero; the phase angles of Vbc and

Vcacan be obtained based on (2-11). In this way, the values of the U2, vuf1 and vuf2 corresponding to the specified LVUR can be obtained as shown in Fig. 2-11 [34].

푉푎푏∠0 + 푉푏푐∠휃푏푐 + 푉푐푎∠휃푐푎 = 0 (2-11)

Page | 42 Chapter 2 Voltage Imbalance in LV Networks

(a) LVUR = 1% (b) LVUR = 2% 1.15% 2.2%

2.1%

1.1% 2%

2.3%

1.05% 2.2%

2.1%

1% 2% 0.99 0.995 1 0.98 0.985 0.99 0.995 1 Vbc (p.u.) Vbc (p.u.) (c) LVUR = 5% (d) LVUR = 10% 5.8%

5.6% 11.5%

5.4% 11%

5.2% 10.5%

5% 10% 0.95 0.96 0.97 0.98 0.99 0.9 0.95 1 Vbc (p.u.) Vbc (p.u.)

U2 vuf1 vuf2

Fig. 2-11 Variations of U2, vuf1, vuf2 for a specified LVUR (a) LVUR=1% (b) LVUR=2% (c) LVUR=5% (d) LVUR=10% [34]

Fig. 2-11 shows the variations of U2, vuf1 and vuf2 when Vbc varies within the allowable range for 1%, 2%. 5% and 10% values of LVUR. It is clearly that for a given value of LVUR, there is a range of U2 and its approximations vuf1 and vuf2. Fig. 2-11(a) shows that the U2, vuf1 and vuf2 agree with each other when the LVUR is 1%. While for 2% LVUR shown in Fig. 2-11 (b), vuf2 starts to deviate slightly from U2. Fig. 2-11(c) and (d) show that as the LVUR increases, vuf2 deviates more from U2, and however, vuf1 is always the same as U2. Accordingly, vuf1 is a better approximation to U2 than vuf2, and vuf1 is able to successful quantify U2 even in the absence of angle measurement, irrespective of the severity of the voltage imbalance.

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Besides, the U2, vuf1, vuf2 , LVUR and PVUR were investigated in [41] by either changing the magnitude or the angle position of a phase-neutral voltage based on a balanced 3Φ voltage. Table 2-2 shows the values of these expressions when phase-C voltage is changed from 0.9 pu to 1.1 pu in ten steps. It’s clearly indicated that the U2, vuf1, vuf2, LVUR and PVUR are approximately in the ratio of 1:1:1:1:2 when the voltage imbalance is induced by varying only the voltage magnitude. Moreover, when the phase angle is changed to induce voltage imbalance, the PVUR is always zero, because the magnitudes of the phase voltages are not affected; whilst the U2, vuf1, vuf2 and LVUR almost have the same value.

Table 2-2 Values of investigated expressions when phase-C voltage is changed [41]

Phase A Phase B Phase C U2 vuf1 vuf2 LVUR PVUR Mag. Angle Mag. Angle Mag. Angle 1.0 0 1.0 -120 1.1 120 3.2258 3.2258 3.2258 3.2506 6.452 1.0 0 1.0 -120 1.08 120 2.5974 2.5974 2.5974 2.6136 5.195 1.0 0 1.0 -120 1.06 120 1.9608 1.9608 1.9608 1.9701 3.922 1.0 0 1.0 -120 1.04 120 1.3158 1.3158 1.3158 1.32 2.632 1.0 0 1.0 -120 1.02 120 0.6625 0.6622 0.6623 0.6633 1.325 1.0 0 1.0 -120 1.0 120 0 0 0 0 0 1.0 0 1.0 -120 0.98 120 0.6711 0.6711 0.6711 0.67 1.342 1.0 0 1.0 -120 0.96 120 1.3513 1.3513 1.3514 1.3467 2.703 1.0 0 1.0 -120 0.94 120 2.0408 2.0408 2.0408 2.03 4.082 1.0 0 1.0 -120 0.92 120 2.7397 2.7397 2.7397 2.7202 5.479 1.0 0 1.0 -120 0.9 120 3.4483 3.4483 3.4483 3.4170 6.897

2.3.3 Associated standards

The European Standard EN 50160 ‘Voltage characteristics of electricity supplied by public electricity networks’ [3] gives the limits of the voltage imbalance based on U2 as 2% for LV and MV systems and 1% for HV systems, and these values should be complied with for 95% of the 10-minute mean values during each period of one week under normal operating conditions. Additionally, it also states that in some areas with partial 1Φ or 2Φ connected users, imbalance up to 3% at 3Φ supply terminals could occur. Tighter limits are stated for HV systems because HV systems are designed to be used to their maximum capacity with a balanced 3Φ load; any imbalance could induce inefficient operation of a generally higher loaded transmission systems [30].

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Moreover, ER P29 states a U2 limit of 1.3% for systems with a nominal voltage below 33kV and 1% for other systems with a nominal voltage less than 132kV, which should not be exceeded for more than 5 minutes in every 30 minutes, and it also states the U2 should not exceed 2% when assessed over any one minute period. These limits are introduced specifically to consider the voltage imbalance attributable to proposed new loads and are not intended to be used as generalized network limits [42].

The American National Standards C84.1-2006 “Electric Power Systems and Equipment- Voltage Ratings (60 Hertz)”, developed by NEMA, recommends that electrical supply voltages should be designed and operated to limit the maximum voltage imbalance to 3%, and this value is measured at the electric-utility revenue meter when the supplied load is disconnected [43]. In this standard, the voltage imbalance is based on LVUR expressed by (2-8). The standard NEMA MG1-2009 “Motors and Generators” states that if the supply voltage imbalance defined by LVUR is over 1%, the rated horsepower of induction motors should be multiplied by the appropriate factor given in Fig. 2-12 [44, 45]. Besides, both the IEEE Standard 141-1993 “IEEE Recommended Practice for Electric Power Distribution for Industrial Plants” and the IEEE Standard 241-1990 “IEEE Recommended Practice for Electric Power Systems in Commercial Buildings” indicate that some electronic equipment may suffer problems if the voltage imbalance is over 2 or 2.5%. Moreover, they also advise not to connect 1Φ load to the 3Φ supply; when the supplied equipment are sensitive to phase-voltage imbalance, a separate circuit should be used instead [29, 46].

Fig. 2-12 Derating factor of induction machine with voltage imbalance [45]

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2.3.4 Causes

The DNOs always try to provide a balanced 3Φ voltage at the point of common coupling (PCC) between the distribution grid and a 3Φ customer’s internal network. Generally, the voltage at the PCC, which is the difference between the source voltage supplied from the generating plant and the voltage drop along the power network, is affected by three factors:- terminal voltages of the generators, impedance of power network and currents drawn by the loads connected to the power network [30]. Any dissimilarity of the above three factors would give rise to an imbalance of either the source 3Φ voltage or the voltage drops along the three phases of the power system, inducing the imbalance of the 3Φ voltage received by the customers.

Terminal voltages of generators The construction and operation of synchronous generators mean the output voltages of each generation site are highly symmetrical and the voltage from a large centralised power plant will be balanced. Although induction generators are often used in wind turbines, a balanced 3Φ set of voltages can be obtained at the output of the generation plant [30].

Nevertheless, the situation is different for small-scale DGs, and in the future they may be increasingly employed in LV networks and take up a significant share of electricity generation. In the UK they are expected to be used to satisfy the projected UK government targets [47]. In addition, and especially in the UK, most small-scale DGs are 1Φ connected in the LV network. The 1Φ nature of small-scale DGs, along with the fact that their employment is consumer-driven and not centrally planned, deteriorates the voltage imbalance in LV networks [48].

Impedance of power network For the 3Φ power network, the impedance of the electricity system components on each phase is rarely the same. The electric parameters of overhead lines always have different values for each phase due to the dependence of the phase impedances on the separation of the wires and the distance to the ground [49]. For 3Φ horizontal- or vertical-formation lines, the impedance of the centre phase is approximately 6-7% lower than the outer two phases [14]. Additional causes of power system impedance imbalance can be

Page | 46 Chapter 2 Voltage Imbalance in LV Networks asymmetrical transformer winding impedances (e.g., open wye and open delta transformer banks), blown fuses on 3Φ capacitor banks, 1Φ small defects (e.g., wire damages altering its section), degradation of isolators (e.g. humidity and dust deposit on insulator’s surface), etc [50, 51].

Non-balanced load current In HV and MV power systems, the loads are usually 3Φ connected and balanced. But non-balanced load conditions can arise in MV rural systems; 1Φ distribution transformers and spur-line supplies are tapped off the 3Φ network and the loads of these tappings are not balanced across the three phases [14]. Loads supplied in LV networks are mostly 1Φ connected, therefore load balance between phases is difficult to guarantee. On the planning stage of an LV network, much effort is paid in making the loading levels equal among the three phase; for instance, the use of one phase per floor of an apartment or office building or alternating connections for a row house. However, phase loads will never be exactly balanced due to human behaviour [52]. But with three or four LV distributors from each MV/LV substation, the overall effect will be less pronounced at the substation. On the MV feeder supplying a number of MV/LV substations, the overall imbalance, due to LV imbalance, will generally not be significant [14].

The continuous variation in load imbalance due to human behaviour, further deteriorates the voltage balancing problem in LV networks [53]. Moreover, energy saving schemes, such as adjustable speed drives (ASDs) can further increase the variations in customer loads. The non-linear property of ASDs, which generally contain a diode rectifier front- end draws varying non-sinusoidal currents, contributes to harmonic distortion. The combination of ASDs, with the proliferation of 1Φ nonlinear switch-mode power supply based loads, such as computers, can lead to non-balanced levels of distortion between phases which also makes the balancing process more challenging [54].

To achieve environmental targets, LCTs, including EVs, HPs and small-scale DGs, are expected to be increasingly used and connected to the LV networks. However, all of these low carbon technologies further deteriorate the voltage imbalance seen in LV networks, due to their single-phase nature.

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2.3.5 Consequences

A non-balanced 3Φ voltage has a hostile influence on the equipment both supplied by the network and used on the network itself. The imbalance could be further exacerbated since a small voltage imbalance can result in a disproportional larger imbalance in the phase currents [55], e.g. the currents drawn by a 3Φ induction motor at normal speed with non-balanced voltage supply will be greatly non-balanced in the order of 6 to 10 times the voltage imbalance [44]. Moreover, a balanced system is in a better position to respond to emergency load transfers, whilst non-balanced system would be less stable and can incur extra losses and the resulting heating effects [53]. However, the sensitivity of different types of electrical equipment to voltage imbalance differs significantly. An overview of the most common problems is presented in this subsection.

Induction machine Studies concerning the adverse effect of voltage imbalance on induction machines can be traced back to 1918, when Fortescue mathematically analysed the operation of induction machines with unsymmetrical supply using sequence components [31]. Based on sequence component analysis, Neumann pointed out that the negative sequence system of an induction machine in an unsymmetrical supply system tends to turn the motor in the opposite direction to that of the positive; the negative sequence system in fact introduces a braking torque and reduces the machine efficiency [56]. In 1954, Williams demonstrated that a reduction in efficiency of the induction motor could occur under non-balanced voltage conditions [57]. Then Gafford pointed out that the induction motor would suffer from extra temperature rise with non-balanced supply, which would reduce the machine’s life, after he carried out a research with regard to the temperature rise of induction motors under non-balanced conditions in 1959 [58]. In 1963, a method for the derating of induction motors was proposed by Berndt and Schmitz, where the rated stator current is taken to be maximum allowable current [59]. In the last two decades, research in this field has mostly focused on the control and protection strategy for induction machine.

Three-phase induction machines operate using an internally induced rotating magnetic field. With a balanced 3Φ supply, the supplied 3Φ voltage contains only positive sequence component, and the rotating magnetic field is circular. In case of a non-

Page | 48 Chapter 2 Voltage Imbalance in LV Networks balanced supply, negative sequence components induce a negatively rotating magnetic field, so the total rotating magnetic field becomes ‘elliptical’ [30]. Firstly, the machine is not able to reach its full torque as the inversely rotating magnetic field induces a negative braking torque which has to be subtracted from the base torque linked to the normal rotating magnetic field. At the same time, the full speed cannot be reached. Secondly, the full torque will be pulsating at double system frequency. The result is the production of speed pulsations, mechanical vibration, and consequently acoustic noise; and the bearing may suffer from mechanical damage [60]. Finally, the stator and, especially, the rotor would suffer from overheat, possibly leading to faster thermal aging; this heat is caused by the negative sequence current associated with the fast inverse rotating magnetic field as seen by the rotor. Moreover, the low negative sequence impedance further increases the negative sequence current induced by the negative sequence component of the non- balanced 3Φ voltage, increasing the machine losses [58]. At a normal operating condition, non-balanced voltages cause the line currents to be non-balanced in a relative order of 6 to 10 times the voltage imbalance [44]. Overall, the main effects of voltage imbalance on the operation of induction machine are reduced efficiency and decreased lifetime.

Synchronous generators Synchronous generators encounter similar problems as induction motors, when they are connected to a non-balanced system; but they mainly suffer from excessive heating. The negative sequence component of the non-balanced 3Φ stator current causes a double frequency current in the surface of the rotor. This current flows through the retaining rings, slot wedge, and field windings, causing high temperatures and possibly failures [61].

Capacity of power plants The negative sequence component causes the loading capacity of transformers, cables, and lines to be reduced. Generally, the operating limits of power system plants are determined by the RMS rating of the total current, which under non-balanced conditions includes the non-productive negative sequence components. This should be taken into account when setting the operating settings of protection devices, i.e. they should operate based on the total current. The maximum capacity of the power system plants, under non- balanced conditions, can be expressed by a derating factor, designed to ensure the plant is be capable of handling the load [30].

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The transformation of the negative sequence current through a transformer is identical to the positive sequence current. However the behaviour of the zero sequence voltage is related to the primary and secondary winding connections. This is especially true in the presence of a neutral conductor which affects the transformation of the voltage zero sequence component. For the Dyn MV/LV transformer, the zero sequence current flowing in the 3Φ four-wire LV feeder would be transformed into a circulating current in the delta-connected primary winding. The magnetic flux, associated with the circulating current, would flow though the construction parts of the transformer causing parasitic losses [30]. Moreover, since a small increase in the voltage of one phase, near the knee point of the H-B curve, significantly increases the width and hence the area of the hysteresis loop; consequently, the iron loss corresponding to this area rises [62]. Therefore, transformers exposed to a non-balanced 3Φ voltage have greater iron losses.

Power electronic converter Power electronic converters are widely used as a power supply interface for many devices such as ASDs, ACDC power supplies and other efficient devices. As shown in Fig. 2-13, the power electronic converter generally contains a diode rectifier front-end, a dc-link capacitor which is used to eliminate the ripples from the incoming current, and a pulse-width modulation (PMW) inverter to convert the dc voltage back to a variable 3Φ AC voltage [54]. With a balanced 3Φ voltage supply, the 3Φ converters with diode rectifier front-ends would draw non-sinusoidal currents rich in odd harmonics. The harmonic characteristic is determined by (2-12), where h is order of the harmonics, q is the number of pulses of the rectifier system.

h = kq ± 1 (k = 1, 2, 3, 4, … ) (2-12)

Under non-balanced conditions, non-characteristic harmonics would be drawn by the converter, i.e. the harmonics do not comply with (2-12) [63, 64]. The non-characteristic harmonics can lead to undesirable harmonic problems in the supply system and make the input currents significantly non-balanced; a test on a 5 horsepower (3.73kW) ASD shows that an increase in the supply voltage imbalance from 0.6% to 2.4% causes an increase in the input current imbalance from 13% to a maximum 52% [65].

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Diode DC link PWM inverter Rectifier capacitor (DC to AC)

Power Supply LOAD

Fig. 2-13 Typical ASD system

Moreover, the 3Φ diode bridge in ASD systems often operates as a 1Φ rectifier connected between the lines with the highest voltage difference when 1Φ voltage sag occurs in the 3Φ voltage supply. In this case, the RMS input current of the ASD increases because the output power has to be supplied via two phases; once the RMS input current exceeds the overcurrent protection threshold (usually, 1.2–1.5pu), the ASD system will trip [66]. Moreover, the increased current can also lead to excessive thermal degradation of the diodes and dc-link capacitor, decreasing their lifetime [64]. In addition, the variation of the dc-link voltage resulting from the non-balanced voltage supply further increases the occurrence of nuisance trips of the ASD system [66].

Protection and control equipment Voltage imbalance may incur nuisance trips of electric machines and ASD systems as discussed above. Moreover, the neutral connected relay for earth fault protection for MV/LV transformers could be tripped by a significant neutral current attributable to the voltage imbalance in the LV feeder, causing unnecessary interruption of supply. Most line drop compensation equipment operates from the load current in one phase and the phase voltage in the other two phases [14]. When the system is severely non-balanced, the equipment will incorrectly adjust the transformer tap position due to the assumption that the system is balanced.

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2.4 Voltage imbalance mitigation

Voltage regulation in distribution networks is mainly accomplished using OLTCs in HV/MV substations. In addition, voltage regulators and shunt capacitors are installed on MV feeders to reduce the voltage drop. Since these technologies have no capability of voltage balancing, especially in LV networks, they are not reviewed here. Conventional and modern voltage imbalance mitigating methods are reviewed and discussed in this section.

2.4.1 Conventional mitigating methods

Transmission line transposition Transposition of overhead line conductors at regular intervals used to be adopted in electrical transmission systems to keep the series inductive and the shunt capacitive reactance balanced, which reduces voltage imbalance. However, modern power lines are seldom transposed at regular interval to avoid the complexity of having a transposition tower, reduce the risk of lightning induced flashovers, and avoid the confusion of transposed lines when addressing emergency faults [49, 67, 68]. When considering a distribution network, phase loading levels are always changing, perfectly balanced load for each phase never exist due to many unpredictable factors e.g. customer behaviour, weather conditions and public events [69]. Therefore, line transposition rarely occurs in distribution networks, especially in the LV networks, for economic reasons and also because it maybe a physical impossibility [69].

Special transformers in railway supply system High-speed railway traction system draws significant 1Φ AC electrical power from the 3Φ power grid, and consequently is one of the main sources of voltage and current imbalance in power system. To reduce the imbalance, various specially-connected transformers have been employed in electric railway substations around the world; there are mainly three kinds of connection: single-phase connection (e.g. pure 1Φ, V-V connection), 3Φ connection (e.g. Wye-delta connection) and 3Φ/2Φ connection (e.g. Scott, modified-Woodbridge, Le Blanc connection) [70]. The 1Φ connection is employed in Italian railways [71, 72], French TGV [73], and New Zealand railways [74]. The V-V

Page | 52 Chapter 2 Voltage Imbalance in LV Networks connection is used in French TGV and British railways [39]. The Wye-delta connection is used in Chinese railways [70]. The Scott connection and the modified-Woodbridge connection are adopted in the Japanese Shinkansen [75, 76]. The Le Blanc connection is used in Taiwan [77]. All the above provide a 1Φ voltage supply on each railway track.

The effect of these special connections on the 3Φ supply voltage imbalance was examined and compared in [39, 40, 78]. Equation (2-13) and (2-14) are the derived imbalance-estimating formulas for different connections.

V-V and Wye-Delta connections: 2 1/2 VUF = (3k − 3k + 1) Sr (2-13)

Scott, Le Blanc and Modified-Woodbridge connections:

VUF = |1 − 2k|Sr (2-14)

Where, VUF refers to the voltage imbalance at the 3Φ supply side, Sr refers to the voltage imbalance calculated by using (2-10) for the 1Φ connection, 0 ≤ k ≤ 1 and k is relative to the train dispatch schedule.

The results are plotted in Fig. 2-14. It can be clearly seen that the V-V and Why-Delta connections help reduce the voltage imbalance in the 3Φ supply, compared with the 1Φ connection. The Scott, Le Blanc and Modified-Woodbridge connections are the most efficient in the reduction of voltage imbalance; when the tractions supplied in the two tracks by the same substation have equal demand, the 3Φ voltage at the primary side of the substation is then balanced.

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(i) Single-phase connection Sr (ii) V-V and Wye-Delta connections (i) (iii) Scott, Le Blanc and Modified- Woodbridge connections (ii) Sr 2 k is relative to the train dispatch schedule. Assuming traction load is (iii) St, the load on the two phases are k*St and (1-k)*St respectively. 0 0 0.5 1 k Fig. 2-14 Effect of specially-connected transformers in voltage imbalance reduction

Network reconfiguration Since a non-balanced load is the main cause of voltage imbalance in a distribution network, load redistribution in the network, according to the network situation, helps to minimize voltage imbalance. Distribution networks are generally radial feeders sectionalized by sectionalizing switches, and tie breakers are also installed for reconfiguration [79]. The configuration of distribution network can be varied with manual or automatic switching operations to transfer load between the circuits so that the system reaches a more balanced state. Afterwards, loads can be connected to a higher voltage level or to a relatively balanced point on the network. Several network reconfiguration algorithms using load estimation techniques have been proposed to give the optimal feeder switch positions [53].

The study in [69] shows that at the network planning or reconfiguration stage, assigning less load to the phase with a higher voltage drop and more load to the phase with a lower voltage drop reduces the voltage imbalance. In the ‘European’ LV networks, manual load transfer among the three supply phases is adopted after field measurements and software analysis designed to relieve the voltage imbalance. In the ‘American’ distribution systems, the connection phase of the distribution transformer to a primary feeder is conventionally changed to reduce the current imbalance in the primary feeder. The supply phase change of the LV load or the distribution transformer improves the phase voltage and current imbalance in some cases, but the trial and error approach used is time-consuming and the resultant balanced state may rapidly change because of the time and spacial variations in load [79]. A generic algorithm approach was developed in [79] to achieve a multi-objective optimization, including improving the primary feeder

Page | 54 Chapter 2 Voltage Imbalance in LV Networks voltage imbalance, by changing the phase arrangement between distribution transformers and a primary feeder. In addition, changing a 1Φ lateral from one phase to another is often carried out to reduce the imbalance in ‘American’ distribution systems [16].

LV network reinforcement Compliance with the voltage imbalance limit is one of the primary tasks in the LV network design, and therefore considerable efforts are made by electric utilities at the design stage to maintain a fair distribution of 1Φ loads among the three phases [80, 81]. However, perfect voltage balance never exits because of customers’ consumption habits and physical connection constraints. Additionally, the emerging connection of LCTs, e.g. 1Φ DGs, EVs and HPs, which were not considered at the design stage of LV networks, will deteriorate the voltage imbalance.

Conventionally, excessive voltage imbalance in LV networks is mitigated by reinforcing the networks, including increasing the feeder cross-section, paralleling feeders, installation of additional feeders and installation of capacitors [81, 82]. The work in [81, 82] shows that these traditional methods all help to reduce voltage imbalance in LV networks, but their disadvantages are expensive financial cost and the interruption of power supply to customers during their installation. These limitations severely limit their commercial use, especially in LV networks where the loads are becoming more volatile.

2.4.2 Modern mitigating methods

On-load tap changer (OLTC) Conventionally, the OLTC is only applied on HV/HV and HV/MV transformers, and due to its cost, perceived complexity and state of development, it is rarely incorporated in MV/LV transformers [83]. Hence, MV/LV transformers are generally equipped with off- load tap changers which are used to cope with long-term or seasonal electrical demand variations. In order to make the LV networks more flexible to accommodate the LCTs without violating the network constraints, several commercial MV/LV transformers with OLTC have recently become commercially available [84, 85]; this type of OLTC can only regulate the voltage magnitude and have no control of the voltage phase angle.

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The study presented in [81] investigated the impact of increasing PV and load penetration on LV network voltage and imbalance. The inclusion of an OLTC on the MV/LV transformer and network reinforcement including paralleling of the feeders and the installation of additional feeders were used to mitigate breaches of the network constraints. The simulation results indicate an OLTC MV/LV transformer is the most effective solution in improving the network’s voltage performance, especially when the 11kV supply voltage deviates significantly from 1.0 pu. The work in [86] addressed the impact of the OLTC MV/LV transformers on the LV network’s capability of accommodating the DGs. The simulations show that the investigated Dutch LV grid with the OLTC MV/LV transformer is able to accommodate the DGs without violating the voltage magnitude limits, and the OLTC MV/LV transformer helps to attenuate the voltage fluctuations of the MV grid and improve the power quality received by each customer. However, this study is based on the assumption that the network is balanced. The effectiveness of the OLTC MV/LV transformers in improving the LV network voltage magnitude was also demonstrated by the research reported in [83], but the result obtained from [83] also indicates the independent tap control of each phase can deteriorate the voltage imbalance because this type of control considers only the voltage magnitudes of the phase-neutral voltages. A detailed description of voltage magnitude and phase angle control using transformers and tap changers is presented in Chapter 3.

VAr compensation VAr compensation is one of the main voltage control techniques used in power system; it alters the circulation of the reactive power between the AC source and the load, and improves the voltage stability of the power system [87]. Methods based on independent VAr compensation in the three phases of a power network have been proposed in recent years to reduce system imbalance in either normal or faulted conditions [60, 82, 88, 89].

In [60], a series connected reactor in an existing motor controller, which in effect is a variable reactor, was controlled to compensate for the negative sequence component in the supply voltage; hence, reducing the supply voltage imbalance for an induction machine. This control technique was demonstrated to be able to reduce the negative sequence current by a factor of five or even more, but it will introduce harmonic losses due to the presence of solid-state switches [60]. A control algorithm for a static VAr compensator was proposed in [88] to inject the negative sequence current required by the

Page | 56 Chapter 2 Voltage Imbalance in LV Networks downstream network; this technique was installed in the upstream of a delta-wye MV/LV transformer, which already blocks the zero sequence current from the downstream network, so that a balanced system can be obtained as viewed from the upstream network. Moreover, a control method based on a distribution static compensator (D-STATCOM) is presented in [89] to mitigate the voltage sags under non-balanced fault conditions in industrial distribution networks.

In LV networks, the application of D-STATCOM in mitigating voltage imbalance is investigated in [82, 90], and the results indicate that the D-STATCOM can reduce the voltage imbalance considerably but its effectiveness depends on its installation location on the LV feeder. Moreover, an MC simulation in [82] shows that the mean of the U2 is reduced from 0.36% to 0.28% at the feeder beginning and from 1.84% to 1.41% at the feeder end when a 15kVA capacitor is installed at the 2/3 feeder length from the feeder beginning.

Voltage correction Voltage balancing can be achieved by injecting an additional voltage into one, two, or three phases of the non-balanced supply voltage to compensate for the voltage negative sequence component, and sometimes the voltage positive sequence component is altered for voltage regulation [91-93]. The injected voltage is generally generated by an inverter circuit as shown in Fig. 2-15, where the DC bus could be a DC generator [91], a battery source [93] or it could be obtained by rectifying the AC supply [92].

+ PWM Filter Vinj refers the V dc Inverter Circuit Vinj injected voltage -

Fig. 2-15 Typical inverter circuit used for voltage correction

The series voltage compensator for non-balanced 3Φ sources introduced in [92] was based on a 3Φ voltage source inverter with non-balanced gating patterns. This balancing system was connected in series with the supply via a 3Φ delta-wye transformer and proved to be feasible. It produced a balanced line-to-line load voltage, in both computer simulation studies and a laboratory test [92]. A dynamic voltage restorer based on a

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PWM inverter circuit and a storage unit was used in [93] to inject the necessary voltage to compensate for the voltage sage. The proposed control scheme was demonstrated to be superior to a conventional in-phase injection technique in terms of energy reduction and dynamic performance [93]. Moreover, an active line conditioner based on an IGBT inverter circuit was introduced in [91] to inject a voltage in one phase of a 3Φ system to provide a balanced 3Φ voltage for the downstream load and this control technique was shown to have a low volt-ampere demand and hence was cost-effective. Compared with VAr compensation based methods, these ‘voltage correction’ methods inject both real and reactive power into the distribution network.

Supply phase selection At present, DNOs occasionally reduce the voltage imbalance in LV feeders by manually changing the supply phases of some customers following field measurements and software analysis studies; this however, is time-consuming [94]. With the advent of artificial intelligence, telecommunication and power electronics equipment in distribution networks, automatic phase balancing can be introduced, which uses an automatic phase selector shown in Fig. 2-16 [94]. This selector consists of three switching devices, one for each phase, and their outputs are connected to the load. Only one switch is conductive at any one time, while the others are off; in this way, the load is always supplied by one phase of the LV feeder. A load transfer scheme based on the automatic phase selector was introduced in [95] for the voltage imbalance reduction in LV networks, and the automatic phase selector was used in [94] to minimize the LV power delivery losses.

A B C N Each switching device is composed Switch of anti-parallel Controller thyristors or a Triac

Residential load

Fig. 2-16 Automatic phase selector [94]

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Renewable energy source Recently, renewable energy sources (RESs) are increasingly being integrated into distribution networks in the form of 1Φ generators, e.g. photovoltaic and small wind turbines, or 3Φ generators, e.g. large wind farms. Because of their intermittent nature, RESs are expected to pose a threat to the network in terms of system stability and power quality. However, the advancement in power electronics and digital control technology makes it possible to actively control the RESs to enhance system operation with improved power quality [96]. The power electronic inverter interfacing the RES with the grid can generate the required reactive power into the grid, and at the same time control the real power transfer in each phase according to the non-balanced electrical demand. A control scheme for the grid-interfacing inverter of a RES was proposed in [96] to improve the power quality, including the current imbalance, at the PCC in a 3Φ four-wire system. Similarly, the energy storage system and its associated grid-interfacing inverter can be properly controlled to improve the power quality at the PCC; the energy storage system gives the necessary real power and the inverter generates the required reactive power [97].

2.5 Voltage limits

The limits of the LV network voltage magnitude and imbalance used throughout this project are described in this section. The voltage magnitude limit stated in ‘The Electricity Safety, Quality and Continuity Regulations 2002’ is adopted in this project, i.e. the LV supply should be within +10%/-6% around 230V [26]. Moreover, the U0 and U2, defined by (2-4) and (2-5), are used to express the voltage imbalance. The limit of U2 stated in [3], i.e. 95% of the 10-minute mean values of the U2 during each period of one week under normal operating conditions should be below 2%, is used to adjudge the voltage imbalance in this project.

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2.6 Summary

The LV network in the UK is generally a 3Φ four-wire radial feeder and it only provides 1Φ supply to domestic properties. Voltage imbalance in LV networks is mainly caused by the non-balanced distribution of the 1Φ domestic loads. In future, the increasing adoption of LCTs, e.g. EVs, HPs and small-scale DGs, will worsen the voltage imbalance and may cause unacceptable voltage magnitude variations.

The current voltage regulation technologies in distribution system, including the OLTC equipped on HV/MV transformers and the voltage regulators and shunt capacitors installed on MV feeders, have no capability of voltage balancing in LV feeders. The traditional methods to resolve voltage imbalance problems always involve network reinforcement, e.g. increasing the feeder cross-sections and fitting appropriate protective devices for the 3Φ machines, which are very expensive. These methods will become unacceptable when severe voltage imbalance occurs on a large scale and massive network reinforcements become necessary. Manually operated network reconfiguration and supply phase change need field measurements, software simulation and analysis; this trial and error approach is time-consuming and the satisfied balanced state may not be a long-term solution due to load variations.

These limitations in the capabilities of conventional voltage imbalance mitigating methods have encouraged researchers to investigate new mitigating methods. MV/LV transformers with OLTC can effectively keep the voltage profile along the LV feeder within the statutory limits, but the independent tap control of each phase can deteriorate voltage imbalance. Automatic network reconfiguration and supply phase selection rely on advanced communication and control system, which is currently not available. Moreover, installing an automatic phase selector at every customer’s PCC may be physically impossible, e.g. for the LV supplies as shown in Fig. 2-7(ii)(vi), it is impossible to install an automatic phase selector individually for each customer.

An alternative solution is to use a D-STATCOM to inject the necessary reactive power into the LV network to reduce the voltage imbalance. An inverter circuit in association with a DC generator, a battery source or a rectifier circuit, is able to mitigate the voltage

Page | 60 Chapter 2 Voltage Imbalance in LV Networks imbalance by injecting additional voltages into the network to counteract the voltage negative sequence component. With an appropriate control algorithm, the RES and its grid-interface inverter can inject the required active and reactive power into the network to improve the power quality, reducing the voltage imbalance. Moreover, a power electronics based AC-DC-AC solution is viable for voltage imbalance mitigation in LV networks, which effectively converts the non-balanced AC to DC and outputs a balanced 3Φ AC voltage. Power electronic converters have the benefits of flexibility and fast response time, but generation of harmonics and short converter lifetime seriously limit their commercial use. The non-linearity associated with the high-frequency converter switching would boost the current and voltage harmonics and reduce the efficiency of the power delivery process. Some of the components in a converter have a short lifetime; switches are mostly IGBT-transistors and on average operate for ten years, while energy storage capacitors begin to fail after five years [98]. This is significantly shorter than the life of traditional network components, e.g. transformers and cable, which are typically 50-70 years [98]. In addition, the performance of these components varies significantly with the operating conditions, which would weaken the reliability of the converters.

With the above consideration in mind, a new voltage balancing method based on transformers and tap changers is developed in this project. This method is able to convert a non-balanced 3Φ voltage into a balanced 3Φ voltage at either a downstream location on the LV feeder or at a 3Φ load supply point. Compared with the above mitigating methods, one using transformers and tap changers has a longer lifetime and doesn’t inject harmonics into the network. Moreover, this method doesn’t rely on advanced communication and control systems, external voltage supplies or energy storage. The details of the proposed voltage balancing method are presented in Chapter 3.

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CHAPTER 3

SCOTT TRANSFORMER BASED VOLTAGE

BALANCING METHOD

3.1 Introduction

This thesis focuses on the voltage imbalance in LV networks, which is predicted to be severely deteriorated due to the significant penetration of EVs and HPs. As discussed in Chapter 2, the conventional voltage imbalance mitigating methods are time-consuming and require huge financial cost. They are not able to tackle the voltage imbalance in an economic way because of the increasing volatility of the LV network loads. Accordingly, a number of studies concerning the voltage imbalance in LV networks were carried out recently and some modern mitigating methods have been proposed. However, some of these methods lean on an advanced communication and control system, which is not available presently; whilst the others resort to power electronic converters, which have the disadvantages of short lifetime and generation of harmonics.

The heart of the PhD project as well as this thesis is an LV radial feeder voltage balancing method using Scott transformers (STs); and the work accomplished in this project revolves around this balancing method. It converts a non-balanced 3Φ voltage into a balanced 3Φ voltage, with the specified magnitude, at either a downstream location on an LV feeder or a 3Φ load supply point. Compared with power electronic based

Page | 62 Chapter 3 Scott Transformer based Voltage Balancing Method voltage imbalance mitigating methods discussed in Chapter 2, this one based on transformers and tap changers provides a longer lifetime and would not inject harmonics into the networks. Furthermore, the operation of this method is independent of the communication system, and no energy storage or external supply is required.

In this chapter, the ST based voltage balancing method is proposed and its methodology is described. The tap changer based voltage regulation methods are firstly reviewed and they are categorized into magnitude regulation, phase angle regulation, and independent magnitude and phase regulation. Afterwards, the ST based voltage balancing method is introduced in the field of its main components, including the Scott transformers and the phase regulating system. Moreover, the control algorithm adopted in the proposed balancing method is presented. Finally, the characteristics of the proposed balancing method are discussed and summarized.

3.2 Voltage regulation using transformers and tap changers

As the most commonly used voltage regulation technology in power system, the transformer with tap-changing capability has a history of over 80 years [99]. Generally the tap changer has two types of applications: tap-changing transformers and voltage regulating transformers. Tap-changing transformers have two functions: voltage transformation and voltage control, whilst voltage regulating transformers only perform the latter, e.g. they buck or boost the voltage without changing the basic voltage level [100]. Both of these two kinds of transformers are able to control the voltage magnitude or the voltage phase angle, or both of them independently.

3.2.1 Basic control

Fig. 3-1 shows the essential feature of a tap-changing transformer; the voltage regulation is realised by changing the turn ratio of the transformers by tapping the windings to alter the number of turns [101]. Since the pu value of the voltage per turn in the high-voltage winding of a transformer is smaller than that in the low-voltage winding, variation of the turns in the high-voltage winding provides more choices of the voltage, and therefore tap

Page | 63 Chapter 3 Scott Transformer based Voltage Balancing Method changers are generally added in the high-voltage winding. One more advantage of this practice in a step-down transformer is that, at the time of light load when the secondary voltage has to be lowest, the maximum number of HV turns is included in the primary side, thus reducing the e.m.f. per turn, the flux-density and the core losses. Tap changers can be on-load or off-load. On-load tap changer operates when the transformer is carrying a load and it is able to take care of daily, hourly, and minute-by-minute variations of the power system conditions. However, off-load tap changer can only change the taps when the transformer is de-energised, and the settings of the off-load tap changing transformers are determined based on the estimation of the long-term load variation, or seasonal changes.

1 Neutral 2 3

Phase

Fig. 3-1 Essential of tap changing operation [101]

The tap-changing can be carried out manually or by an automatic voltage control (AVC) relay that senses the secondary busbar voltage and uses it to keep the secondary voltage level at a particular point within statutory limits by compensating for changes in the higher voltage level and for variations in the load. On-load tap changing transformers are always accompanied by an AVC relay, which in the simplest case as shown in Fig. 3-2 [102], raises or lowers the tap changer by comparing the voltage reference and the transformer secondary voltage so that the transformer secondary voltage equals the voltage reference, subject to the associated tolerance. Since the transformer tap changer alters the voltage in a discrete manner, a ‘deadband’ larger than one tap step size should be adopted in the tap changing operation to avoid the ‘hunting’ between the tap changer and the AVC relay. In common practice, the deadband of an AVC relay is just less than twice the transformer tap step size, e.g. the tap step size for HV/MV transformers in the UK is 1.67% and the corresponding deadband of the relay can be of the order of 2% [14].

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Moreover, a time delay is injected between the determination and the implementation of the tap operation as shown in Fig. 3-2 to prevent over-frequent tap changes, induced by short time voltage variations. This helps prevent excessive wear on the tap changer.

Supply Load

Tap changer

up down

AVC relay Delay times Voltage Deadband reference + _

Fig. 3-2 Basic tap changer arrangement [102]

3.2.2 Voltage magnitude regulation

Line drop compensation Line drop compensation (LDC), also known as voltage compounding, is an improved application of OLTCs; it controls the voltage not at the transformer secondary but at the remote load centre [14]. As shown in Fig. 3-3 [102], the LDC estimates the voltage drop along the distribution line by injecting a current proportional to the transformer secondary current (derived from a CT) through an impedance R’ and X’ adjusted to model the network impedance. The difference between the obtained voltage drop and the voltage proportional to transformer secondary voltage (derived from a VT) reflects the voltage at the remote load centre [14]. An AVC relay controls the transformer tap changer to keep the voltage at the load centre within the expected range. Despite the fact that a network with a single distribution feeder and only one load at its termination is rare, LDC is often used, even if it does not ideally suit the network [103]. Besides, the LDC can be applied to not only power transformers with tap changing capability as shown in Fig. 3-3, but also voltage regulating transformers.

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Distribution line CT impedance Supply Load VT Tap changer

up down

R’ X’ AVC relay - - +

Fig. 3-3 Schematic diagram of line drop compensation [102]

Step voltage regulator Step voltage regulator is defined as a regulating autotransformer which controls the voltage of the regulated circuit in steps by means of taps on the series winding without interrupting the load. The position of the tap in a step voltage regulator is generally controlled by an LDC. Based on whether the source supply or the regulated circuit is connected to the series winding via taps, there are type-A and type-B step voltage regulators.

In the type A step voltage regulator, the primary is directly connected to the shunt winding and the series winding is connected to the regulated circuit via tap as illustrated in Fig. 3-4 [104]. The regulated secondary voltage is essentially the addition of the primary voltage and the voltage across series winding. When the reversing switch selects point ‘R’, the secondary voltage would be higher than the primary voltage; when the reversing switch selects ‘L’, the secondary voltage would be lower than the primary voltage. A preventive autotransformer is usually used in the tap changing mechanism to limit the current during the transition period from one tap to the next [105].

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Reversing switch Primary L Preventive Circuit autotransformer R Secondary circuit Shunt winding Series winding

Fig. 3-4 Single-phase type-A step voltage regulator [104]

Fig. 3-5 shows a 1Φ type B step voltage regulator, in which the primary is connected to the series winding via tap and the regulated circuit is connected to the shunt transformer directly, contrary to the type A step voltage regulator. The regulated secondary voltage of the type B step voltage regulator is essentially obtained by subtracting the voltage across the series transformer from the primary voltage. When the reversing switch selects the point ‘R’, the secondary voltage would be higher than the primary voltage and be lower when the point ‘L’ is selected. Besides, since the shunt winding is across the unregulated primary circuit in the type A step voltage regulator, the transformer core excitation varies; whilst in the type B voltage regulator, since the shunt winding is across the regulated secondary circuit, the core excitation is constant. Hence, the type B step voltage regulator has less iron losses than the type-A step voltage regulator.

Reversing switch Secondary Preventive R circuit autotransformer Primary L Circuit

Series winding Shunt winding

Fig. 3-5 Single-phase type-B step voltage regulator [104]

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Three 1Φ step voltage regulators can be connected as a 3Φ voltage regulator in a variety of ways, including grounded-wye, open-wye, closed-delta and open-delta, to regulate the voltage in 3Φ systems. The capacity of the 3Φ voltage regulator varies significantly with the type of connection [106].

Voltage magnitude regulation using booster transformers The booster transformer, whose shunt transformer is connected between the phase and neutral, as shown in Fig. 3-6 (a) has the ability to control the voltage magnitude in power systems [101]. With a tap changer added to the shunt transformer, the series transformer is able to inject a variable voltage ΔVpn, either in phase or in anti-phase with Vpn, to Vpn as illustrated by the phase diagrams in Fig. 3-6 (b). The booster transformer introduced above is generally used when it is not desirable to equip the tap-changing gear to the main power transformer.

Series Phase transformer ∆Vpn ∆Vpn Shunt transformer ∆Vpn Vpn Vpn' Vpn Vpn' Vpn' reverser Vpn

Neutral (a) (b)

Fig. 3-6 Magnitude regulating booster transformer (a) circuit diagram (b) phase diagram

Generally, an autotransformer is used to replace the two-winding shunt transformer in Fig. 3-6 (a) and the obtained new voltage regulating transformer is shown in Fig. 3-7 [101]. Three 1Φ regulating transformers of this type are usually connected as shown in Fig. 3-8 to form a 3Φ regulating transformer. This is used to control the reactive power flow in transmission systems by changing the voltage magnitudes [107].

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∆Vpn

Vpn Vpn' reverser

Neutral

Fig. 3-7 Magnitude regulating booster transformer using autotransformer [101]

Fig. 3-8 Three-phase magnitude regulating booster transformer [107]

Fig. 3-9 shows another booster transformer based voltage magnitude regulating method. This method is analogous to the one shown in Fig. 3-6 (a); but a tertiary winding on the main transformer is used instead of a separate regulator unit [101].

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Series Phase transformer

Main transformer reverser

Tertiary

neutral

Fig. 3-9 Tertiary winding used voltage magnitude regulation [101]

3.2.3 Voltage phase angle regulation

Voltage phase angle control Phase shifting methods based on transformers and tap changers work by injecting a voltage component into the regulated phase; the injected voltage component and the regulated phase voltage differ in phase, generally having a phase shift of 90º [101]. Fig. 3-10 shows a simple method of obtaining a control of the phase angle between the voltages at the input and output sides by using a booster transformer [101]. As shown in ′ the corresponding phase diagram, the phase-A has injected it a voltage VBC, proportional to the phase-to-phase voltage VBC. The output phase-A voltage Va, differs from the input ′ phase-A voltage VA by the phase angle ϕ, which is determined by the magnitude of VBC. In the application of this phase shifting method, a tap changer is adopted and added on either the shunt regulating transformer or the series transformer to change the magnitude ′ of the injected voltage VBC, giving a corresponding control of the phase shift ϕ.

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V ' BC V ' A BC V A  Va B

VA VBC Va

VB VC C VBC N (a) (b)

Fig. 3-10 Voltage phase angle control (a) circuit diagram (b) phase diagram [101]

Phase shifting transformer Phase shifting transformers with the tap changing mechanism installed on the series transformer, can be obtained by special inter-connection of the windings of a 3Φ transformer. Fig. 3-11 (a) shows one possible arrangement; the three primary windings are delta-connected, the secondary windings are injected into the three phases and each of them is coupled with the primary winding connected between the other two phases [107]. In this way, a quadrature voltage is injected into each phase to change the voltage phase angle, e.g. a voltage ∆VA, which is a portion of the phase-to-phase voltage VBC, is injected into phase-A and the resultant phase-A voltage Va has a phase shift of ϕ with respect to the input phase-A voltage VA as shown in the phase diagram in Fig. 3-11 (b). The tap changer on the series winding is capable of varying the injected quadrature voltage to obtain a variable phase shift. This arrangement can also be viewed as a 3Φ step voltage regulator.

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VA Va VA Ta VA

VA 

Va V VC b Tb VC TC VB Vc VB (a) (b)

Fig. 3-11 Three-phase phase shifting transformer (a) circuit diagram (b) phase diagram [107]

Quadrature booster A quadrature booster is a phase shifting transformer with the tap changing mechanism installed on the shunt regulating transformers. It is constituted by combining two types of transformers: the shunt regulating transformer are used to generate quadrature voltages, and the series transformers are used to inject the quadrature voltages into the right phases, as illustrated in Fig. 3-12 [108]. Quadrature booster has the same operating principle as the phase shifting transformer shown in Fig. 3-11; the tap changers located on the shunt transformers can change the magnitudes of the injected quadrature voltages. Like other phase shifting transformers, quadrature boosters are mainly used in transmission systems for real power flow control. However, a quadrature booster was trialled in an UK distribution network in Ofgem low carbon fund project ‘Flexible Plug and Play’ and used to balance power flows in two parallel distribution lines supplying a large DG whose electricity export is constrained due to non-balanced load sharing on the two parallel lines [108, 109].

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Fig. 3-12 Three-phase quadrature booster [108]

3.2.4 Independent voltage magnitude and phase angle regulation

With proper design, a regulating transformer can independently control the voltage magnitude and phase angle. Fig. 3-13 illustrates one arrangement which is able to adjust these two variables separately [107]. The delta-connected exciting transformer has two sets of secondary windings; the voltage exciting the series transformer per phase is a combination of the portions of the three phase-to-phase voltages, e.g. the voltage exciting 1 the series transformer on phase-A consists of the voltage component ∆Va picked off from the left secondary in Fig. 3-13, in phase with the input phase-A voltage Va , and the 2 voltage component ∆Va obtained from the other secondary, which is 90º out of phase with Va. In this way, the voltage magnitude can be regulated by adjusting the magnitude 1 2 of ∆Va , whereas, the adjustment of ∆Va provides phase angle control.

Page | 73 Chapter 3 Scott Transformer based Voltage Balancing Method

V 1  V 2 VA a a Va A B C

This winding This winding adds the adds the voltage voltage 1 2 component Va component Va

(a)

1 2 Va Va

VA 1 2 Va  VA  Va  Va

(b)

Fig. 3-13 Independent voltage magnitude and phase control (a) circuit diagram (b) phase diagram [107]

Another arrangement using transformers and tap changers, which is capable of adjusting the voltage magnitude and phase angle independently, is introduced in [110] for power flow control in transmission systems. This regulating transformer as shown in Fig. 3-14 is a 3Φ transformer with three sets of secondary windings. The primary windings are wye-connected and the secondary windings of different phases are connected in series and injected into one of the three phase conductors respectively. By changing the taps on the secondary windings, a voltage with variable magnitude and phase angle can be

Page | 74 Chapter 3 Scott Transformer based Voltage Balancing Method injected into the regulated phase to realize the separate regulations of the voltage magnitude and phase angle.

A B C

Primary windings Secondary windings

Fig. 3-14 New introduced regulating transformer for voltage magnitude and phase angle control [110]

3.3 Overview of the proposed method

An LV radial feeder voltage balancing method using STs is proposed and the work carried out in this PhD project revolves around this method, which is also the heart of this thesis. This balancing method is able to convert a non-balanced 3Φ voltage into a balanced 3Φ voltage with the predefined magnitude at a downstream location on an LV feeder, where the downstream voltage imbalance is unacceptable, or at a 3Φ load supply point. Fig. 3-15 shows the schematic circuit diagram of the voltage balancing system using the proposed method. In this voltage balancing system, the non-balanced 3Φ voltage (V̅̅NcA̅̅̅̅, V̅̅NcB̅̅̅̅, V̅̅NcC̅̅̅̅) is firstly converted to a non-balanced 2Φ voltage (V̅̅T̅, V̅̅̅M̅) by use of Scott transformer I, and then the phase regulating system is implemented on the 2Φ system, consisting of phase-T and phase-M, to orthogonally balance the phase angles of the 2Φ voltage, i.e. ensure the voltage phase difference is 90º. Following this, Scott transformer II is used to convert the phase angle orthogonally balanced 2Φ voltage (V̅t, V̅̅m̅̅) into a balanced 3Φ voltage (V̅a, V̅̅b̅, V̅c ). As shown in Fig. 3-15, the proposed voltage balancing system mainly includes three parts: Scott transformer I, phase regulating system and Scott transformer II, and all of them are described in detail in the following.

Page | 75 Chapter 3 Scott Transformer based Voltage Balancing Method - e e c e g r V a h t b t l

o V d v e a

c e V s n a a l I h I a p

r B e m r o f s n 3 2 a P r P t A

e A t g t T T

a o c t i l c t o a S v r

m d e a s V t u a q h

V p e - l o g n m w

t a r

g e t r e d n s i s e e m t r y a c a s m l o s h r n f e

u i a o P s r f l g g n e s a e n a i n S b r t R t a a r l t u 1 g P e r A

T e s a h - P o T e w t g

V a M d T t M - - e l V e e c o s s n v

a a a l e I h h

a s r a b P P - e h n p m o r N o f s n a r t

S t N t - o e r r c e e e

r e r S

h m m g e t n r r i

s a o o t a d f f a l e s s e M o c n n T v r n a a

r r a o e t t t l s NcC c a a u V b h - d p NcB n n o o V NcA c

N l V a r t u e B C A N

Fig. 3-15 Schematic diagram of the proposed voltage balancing system

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3.4 Scott transformer

Scott transformer or Scott-connected transformer as shown in Fig. 3-16, is a type of transformers based circuit device used to convert a 3Φ supply to a 2Φ voltage, or vice- versa. Scott connection, i.e. the arrangement of the transformers in an ST, was firstly reported by Charles F. Scott in a paper presented at a meeting of the National Electric Light Association in March 1894, where he stated that: In considering the marked advantage of the two-phase system for distribution and of the three-phase system for transmission, it occurred to me that a combination of the two systems might secure the advantages of both, and I have worked out a simple and effective method of accomplishing this result [111]. Historically, STs were used to link 3Φ and 2Φ power systems with power flow in either direction. In more recent times, and mainly because most loads draw electrical energy from either 1Φ or 3Φ supply, STs become obsolete. However they are still used in railway power supply systems, to feed two 1Φ traction loads from a 3Φ supply. This reduces the voltage imbalance imposed by 1Φ traction loads on a 3Φ utility supply.

A I A IT

3 N V 1 NA Teaser 3 N and N refer transformer 1 2 VT the turns at the N N2 two sides of the 3 main transformer VNC N 6 1 IC C 0.5N S 0.5N B V 1 1 NB Main transformer I B I M N 2

Fig. 3-16 Circuit diagram of a Scott transformer

As shown in Fig. 3-16, a Scott transformer is made up of two 1Φ transformers: a teaser transformer and a main transformer. The primary (3Φ side) of the teaser transformer has

Page | 77 Chapter 3 Scott Transformer based Voltage Balancing Method

√3⁄2 (86.7%) times of the primary turns of the main transformer, and is connected between terminal A of the 3Φ supply and the primary mid-point S of the main transformer. The primary of the main transformer is connected between the terminals B and C of the 3Φ supply. In this way, the voltage across the teaser transformer primary becomes (√3⁄2)U assuming the line-to-line voltage of the 3Φ supply is U, and further across each turn of the teaser transformer, there will be a voltage exactly equal to that per turn of the main transformer. Thus if the turn numbers of the two secondaries are the same, the two secondary voltages would be equal in magnitude but in phase quadrature. Moreover, the point N on the primary of the teaser transformer in Fig. 3-16 refers to the neutral point of the 3Φ side. This point divides the teaser winding AS in a ratio of 2:1; the voltage V̅̅NA̅̅̅ across the winding NA and the voltage V̅̅NS̅̅̅ across the winding NS have the relationship shown in (3-1).

V̅̅NA̅̅̅ + 2V̅̅NS̅̅̅ = 0 (3-1)

The voltages V̅̅NB̅̅̅ and V̅̅NC̅̅̅, imposed by the terminal B and C of the 3Φ supply with respect to N have the following relationship:

V̅̅NB̅̅̅ = V̅̅NS̅̅̅ + 0.5 ∙ V̅̅CB̅̅̅ (3-2)

V̅̅NC̅̅̅ = V̅̅NS̅̅̅ − 0.5 ∙ V̅̅CB̅̅̅ (3-3)

where V̅̅CB̅̅̅ refers to the voltage across the winding CB. By adding the two sides of (3-1), (3-2) and (3-3) together, (3-4) can be obtained.

V̅̅NA̅̅̅ + V̅̅NB̅̅̅ + V̅̅NC̅̅̅ = 0 (3-4)

Equation (3-4) means that the three voltages, imposed by the three terminals of the 3Φ supply with respect to the neutral point N, have no zero sequence component, regardless of whether the 3Φ supply is balanced or not.

ST is able to convert a balanced 3Φ voltage to a balanced 2Φ voltage, and vice-verse, as shown in Fig. 3-17(a). However, if the supplied three- or two-phase voltage is non-

Page | 78 Chapter 3 Scott Transformer based Voltage Balancing Method balanced, the output voltage would be non-balanced as shown in Fig. 3-17(b), where N indicates the neutral point of the ST and Nc indicates the neutral conductor of the 3Φ & Neutral four-wire feeder. When a 3Φ voltage supply with zero sequence component supplies an ST, there would be a voltage difference between N and Nc. In Fig. 3-17(b), the three phase voltages NcA, NcB and NcC (solid lines), having zero sequence component, is converted to a non-balanced 2Φ voltage, whilst for the inverse conversion, this non-balanced 2Φ voltage is converted to the three phase voltages NA, NB and NC (dashed lines), having no zero sequence component.

A A

T T

N Nc N

M M C S B C S B

(a) (b)

Fig. 3-17 Three-phase and two-phase voltage conversion

Assuming the main transformer has N1 turns in its primary, the turn number of the teaser primary would be (√3⁄2)N1, and both these two transformers have N2 turns in their secondaries as shown in Fig. 3-16. The neutral point N divides the teaser transformer primary into the winding AN with (√3⁄3)N1 turns and the winding NS with (√3⁄6)N1 turns; moreover, the point S divides the main transformer primary into two windings with the equal turns of (1⁄2)N1 . With a 3Φ supply feeding the ST, the voltages on the primary and secondary sides have the following relationship.

V̅̅̅̅̅̅−V̅̅̅̅̅̅ ( 3⁄2)N NA NS = √ 1 (3-5) ̅V̅̅T̅ N2

V̅̅̅̅̅̅ N CB = 1 (3-6) ̅V̅̅M̅̅ N2

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Where, the parameters V̅̅T̅ and V̅̅M̅̅ refer to the voltages across the secondaries of the teaser and main transformers respectively.

From (3-5) and (3-6),

2N2 V̅̅T̅ = (V̅̅NA̅̅̅ − V̅̅NS̅̅̅) (3-7) √3N1

N2 V̅̅M̅̅ = V̅̅CB̅̅̅ (3-8) N1

From (3-1),

V̅̅NS̅̅̅ = −0.5 ∙ V̅̅NA̅̅̅ (3-9)

From (3-2) and (3-3),

V̅̅CB̅̅̅ = V̅̅NB̅̅̅ − V̅̅NC̅̅̅ (3-10)

Substitute (3-9) and (3-10) into (3-7) and (3-8) respectively, (3-11) can be obtained, from which the output 2Φ voltages from an ST can be obtained from the input 3Φ voltages.

Please note that V̅̅NA̅̅̅, V̅̅NB̅̅̅, V̅̅NC̅̅̅ in (3-7) refer to the three voltages imposed by the three terminals of the 3Φ supply with respect to the neutral point N instead of the neutral conductor.

V̅̅NA̅̅̅ ̅V̅̅ N [ T ] = 2 ∙ [2⁄√3 −(1⁄√3) −(1⁄√3)] ∙ [V̅̅̅̅̅] (3-11) N NB V̅̅M̅̅ 1 0 1 −1 V̅̅NC̅̅̅

In a transformer, the primary ampere-turns must balance the secondary ampere-turns when neglecting the magnetizing current. In Fig. 3-16, the parameters I̅A, I̅B, I̅C refer to the 3Φ currents in the supply, and I̅T, I̅̅M̅ refer to the currents in the secondaries of the teaser and main transformers, respectively. Since the neutral point N in Fig. 3-16 is not connected to the supply neutral, the teaser transformer primary current is I̅A , and its secondary current is I̅T. The relationship of the currents in the teaser transformer can be expressed by (3-12).

I̅A ∙ (√3⁄2)N1 + I̅T ∙ N2 = 0 (3-12)

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Since the teaser transformer is connected to the main transformer primary mid-point S, forming a return path, the teaser primary current would split into two parts equally at S, and the two halves flow in the opposite directions in the main primary winding, having no magnetic effect on the core and playing no part in balancing the main secondary ampere-turns. Therefore, the total main transformer primary current consists of two parts: the balancing current and the halves of the teaser primary current. In the main transformer, assuming the primary balancing current is I̅2, then the relationship between

I̅2 and I̅̅M̅ can be expressed by (3-13).

I̅2 ∙ N1 + I̅̅M̅ ∙ N2 = 0 (3-13)

The phase diagram in Fig. 3-18 illustrates the currents in the two sides of the ST. It can be clearly seen that the primary phase-B and phase-C currents, I̅B and I̅C , can be expressed by (3-14) and (3-15).

A T N  I2 I2 IT  0.5I  0.5I A A I M IC IB C B M

Fig. 3-18 Three-phase and two-phase current conversion

I̅B = −0.5 ∙ I̅A + I̅2 (3-14)

I̅C = −0.5 ∙ I̅A − I̅2 (3-15)

From (3-12),

N2 2 I̅A = − ∙ ∙ I̅T (3-16) N1 √3

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From (3-13),

N2 I̅2 = − ∙ I̅̅M̅ (3-17) N1

Substitute (3-16) and (3-17) into (3-14) and (3-15),

N2 1 N2 I̅B = ∙ ∙ I̅T − ∙ I̅̅M̅ (3-18) N1 √3 N1

N2 1 N2 I̅C = ∙ ∙ I̅T + ∙ I̅̅M̅ (3-19) N1 √3 N1

Therefore, (3-20) can be obtained, from which the supply 3Φ currents can be derived from the 2Φ load currents drawn from an ST.

I̅A −(2⁄√3) 0 N I̅ [I̅ ] = 2 ∙ [ ⁄ ] ∙ [ T ] (3-20) B N 1 √3 −1 ̅̅̅ ̅ 1 IM IC 1⁄√3 1

Scott transformer I The Scott transformer I (ST-I) in the proposed balancing system is used to convert the non-balanced 3Φ voltage (120º phase rotation) into a 2Φ voltage (90º phase rotation), and it’s an application of ST in the 3Φ four-wire LV feeder. For a non-balanced 3Φ supply with voltage zero sequence component, the neutral conductor Nc of the LV feeder and the neutral point N on the Scott transformer have different potential levels, and therefore, Nc and N are not connected as shown in Fig. 3-15. The relationship of the voltages and currents at the two sides of Scott transformer I can be expressed by (3-11) and (3-20).

Scott transformer II The Scott transformer II (ST-II) in the balancing system is an application of ST in the LV 3Φ 4-wire feeder for 2Φ-3Φ conversion. As shown in Fig. 3-19, two tap changers, TAP2 and TAP3 are added on the primaries of the teaser and main transformers respectively. Two parameters, β and γ, are used to reflect the effect of these two tap changers; there are (1 + β)N2 turns in the teaser transformer primary and (1 + γ)N2 turns in the main transformer primary connected into the circuit. In Fig. 3-19, V̅t, V̅̅m̅̅ refer the 2Φ supply

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voltages and V̅̅a̅, V̅̅b̅, V̅c refers the 3Φ output voltages of ST-II and their relationship can be expressed by the following equations.

푉̅ (1+β)푁 푡 = 2 (3-21) (3⁄2)̅푉̅푎̅ (√3⁄2)푁1

푉̅̅̅̅ (1+γ)푁 푚 = 2 (3-22) 푉̅̅̅cb̅̅ 푁1

V̅̅b̅ = −0.5 ∙ V̅̅a̅ + 0.5 ∙ V̅̅cb̅̅ (3-23)

V̅c = −0.5 ∙ V̅̅a̅ − 0.5 ∙ V̅̅cb̅̅ (3-24)

It TAP2  3 2N1 Ia

3 N 3 1 Teaser transfromer Vt 1 N 2 V 3 N a 6 1 Im Ib TAP3 0.5N1 Main transfromer V 1  N 2 b Vm 0.5N1 Ic

N1 Vc

Neutral conductor Fig. 3-19 Scott transformer II

From (3-21),

푁1 1 1 V̅̅a̅ = ∙ ∙ ∙ V̅t (3-25) 푁2 √3 1+훽

From (3-22) and (3-23),

1 1 푁1 1 V̅̅b̅ = − ⁄ ∙ V̅̅a̅ + ⁄ ∙ ∙ ∙ V̅̅m̅̅ (3-26) 2 2 푁2 1+훾

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From (3-22) and (3-24),

1 1 푁1 1 V̅c = − ⁄ ∙ V̅̅a̅ − ⁄ ∙ ∙ ∙ V̅̅m̅̅ (3-27) 2 2 푁2 1+훾

Substitute (3-25) into (3-26) and (3-27), (3-28) can be obtained, which expresses the relationship of the voltages at the two sides of ST-II.

1 0 √3(1+β) V̅̅a̅ N 1 1 V̅ [V̅̅̅] = 1 ∙ − ∙ [ t ] (3-28) b N 2√3(1+β) 2(1+γ) 2 V̅̅m̅̅ V̅ 1 1 c − − [ 2√3(1+β) 2(1+γ)]

In fig. 3-19, I̅t, I̅̅m̅ refer the 2Φ currents in the supply side and I̅a, I̅b, I̅c refer the 3Φ currents in the load side of ST-II and their relationship can be expressed by the following equations.

I̅t(1 + β)N2 + I̅a (√3⁄3)N1 − (I̅b + I̅c) ∙ (√3⁄6)N1 = 0 (3-29)

I̅̅m̅(1 + γ)N2 + I̅b(1⁄2)N1 − I̅c(1⁄2)N1 = 0 (3-30)

By combining (3-29) and (3-30) together, (3-31) can be obtained, which shows the relationship between the currents at the two sides of ST-II.

√3 √3 √3 I̅a I̅ N − [ t ] = 1 ∙ [ 3(1+β) 6(1+β) 6(1+β)] [I̅] (3-31) ̅̅̅ N 1 1 b Im 2 0 − 2(1+γ) 2(1+γ) I̅c

ST-II converts the phase angle orthogonally balanced 2Φ voltage into a balanced 3Φ voltage. The achievement of this goal is attributed to two properties of the ST.

Property One: A phase angle orthogonally balanced 2Φ supply for an ST means the output phase-B and phase-C voltages, V̅̅b̅ and V̅c, are equal in magnitude.

Based on (3-28), the magnitude of V̅̅b̅ and V̅c can be obtained as (3-32) and (3-33).

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̅̅̅ 푁1 푉̅푡∙ 푉̅푡 푉̅̅푚̅̅∙ 푉̅̅푚̅̅ 1 ̅ ̅̅̅̅ |Vb| = ∙ √ 2 + 2 − 푉푡 ∙ 푉푚 (3-32) 푁2 12(1+훽) 4(1+훾) 2√3(1+훽)(1+훾)

̅ 푁1 푉̅푡∙ 푉̅푡 푉̅̅푚̅̅∙ 푉̅̅푚̅̅ 1 ̅ ̅̅̅̅ |Vc| = ∙ √ 2 + 2 + 푉푡 ∙ 푉푚 (3-33) 푁2 12(1+훽) 4(1+훾) 2√3(1+훽)(1+훾)

When V̅t and V̅̅m̅̅ have a phase shift of 90º, the value of V̅t ∙ V̅̅m̅̅ becomes zero, and then

|V̅̅b̅| and |V̅c| are the same. Thus, if the 2Φ supply is orthogonally balanced in phase angle, the output voltages in phase-B and phase-C would have the same magnitude. Moreover,

|V̅̅b̅| and |V̅c| have the same value only when the value of V̅t ∙ V̅̅m̅̅ is zero, i.e. V̅t and V̅̅m̅̅ have a phase shift of 90º. Therefore the Property One is reversible, i.e. if the output phase-B and phase-C voltages, V̅̅b̅ and V̅c, from an ST are equal in magnitude, the 2Φ supply is orthogonally balanced in phase angle.

Property Two: The equality of the output 3Φ voltages in magnitude means that the output 3Φ voltage is balanced.

According to Property One, the equality of |V̅̅b̅| and |V̅c| means that the input phase-T and phase-M voltages have a phase shift of 90º, i.e. the value of V̅t ∙ V̅̅m̅̅ is zero. In this case, the magnitude of phase-B voltage can be expressed by (3-34).

̅̅̅ 푁1 푉̅푡∙ 푉̅푡 푉̅̅푚̅̅∙ 푉̅̅푚̅̅ |Vb| = ∙ √ 2 + 2 (3-34) 푁2 12(1+훽) 4(1+훾)

Based on (3-28), the magnitude of phase-A voltage can be obtained as expressed by (3- 35).

̅̅̅ 푁1 푉̅푡∙ 푉̅푡 |Va| = ∙ √ 2 (3-35) 푁2 3(1+훽)

Since |V̅̅a̅| and |V̅̅b̅| are equal, (3-36) can be obtained from (3-34) and (3-35).

1+ γ |V̅̅̅̅| = |V̅ | (3-36) m 1+ β t

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Moreover, assuming the phase angle of V̅t is θ; the phase angle of V̅̅m̅̅ would be (θ − 90°), substitute (3-36) into (3-28), (3-37) can be obtained, which shows the output 3Φ voltages in the case that they have equal magnitude.

ejθ √3 V̅̅a̅ jθ j(θ−90°) N1 e e |V̅̅̅t| [V̅̅b̅] = ∙ − + ∙ (3-37) N 2√3 2 1+β ̅ 2 Vc ejθ ej(θ−90°) − − [ 2√3 2 ]

Substitute (3-37) into (2-1), the sequence components of the output 3Φ voltages can be obtained as expressed by (3-38), which shows that the output 3Φ voltages are balanced.

V 0 0 ̅̅̅ N1 jθ √3 |Vt| [V+] = [e ] ∙ (3-38) N2 3 1+β V− 0

3.5 Phase regulating system

The phase regulating system in the proposed balancing system is used to orthogonally balance the 2Φ voltages in phase angle, i.e. ensure the 2Φ voltages have a phase shift of 90º. As shown in Fig. 3-20, the phase regulating system is essentially a booster transformer; consisting of a shunt regulating transformer and a series transformer. The primary of the shunt transformer is connected between phase-M conductor and the neutral conductor, and the tap changer TAP1 is installed on its secondary. One end of the series transformer primary is connected to the shunt transformer through TAP1 and the other through a single-pole three-throw switch, and its secondary is injected into the phase-T conductor. In Fig. 3-20, ̅V̅T̅, V̅̅M̅̅, I̅T, I̅̅M̅ indicate the voltages and currents at the input side of the phase regulating system, and V̅t, V̅̅m̅̅, I̅t, I̅̅m̅ indicate the voltages and currents at the output side.

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VM IT It

Series V VT transformer t IM Im

T(1) T(2) 1 Regulating 2 transformer T(n-1) V VM T(n) 3 m TAP1  IT Neutral conductor

Fig. 3-20 Phase regulating system

When the phase difference between V̅̅T̅ and V̅̅M̅̅ is exactly 90°, the switch would connect pin 2 in Fig. 3-20, so that no voltage is injected into phase-T. As shown in Fig. 3-21(a), when the phase difference between V̅̅T̅ and V̅̅M̅̅ is smaller than 90°, the switch would connect pin 3 to inject a voltage inversely proportional to V̅̅M̅̅ into V̅̅T̅, and when the phase difference between V̅̅T̅ and V̅̅M̅̅ is larger than 90°, the switch would connect pin 1 to inject a voltage proportional to V̅̅M̅̅ into V̅̅T̅, so that the resultant V̅t has a phase shift of 90° with respect to V̅̅M̅̅. The magnitude of the injected voltage is controlled by TAP1, and this effect is reflected by use of the parameter α; the phase regulating system injects a voltage

αV̅̅m̅̅ into phase-T conductor.

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Vt VT VT Vt

VM VM (a) (b)

Fig. 3-21 Operation of phase regulating system (a) injected voltage inversely proportional to V̅̅M̅̅ (b) injected voltage proportional to V̅̅M̅̅

The phase regulating system injects an appropriate voltage into phase-T and does not vary the phase-M voltage. The voltages at the two side of the phase regulating system can be expressed by (3-39).

V̅ 1 α V̅̅̅ [ t ] = [ ] ∙ [ T ] (3-39) V̅̅m̅̅ 0 1 V̅̅M̅̅

The shunt transformer in the phase regulating system draws current from the phase-M conductor; whilst the phase-T current remains constant. Based on the transformer current coupling theory, it is straightforward to show the current in the shunt transformer primary is αI̅t. The relationship of the currents at the two sides of the phase regulating system can be expressed by (3-40).

I̅ 1 0 I̅ [ T ] = [ ] ∙ [ t ] (3-40) I̅̅M̅ α 1 I̅̅m̅

3.6 Control algorithm

The operation of the proposed balancing system comprises voltage phase control and magnitude control. The phase regulating system changes the phase angle of the phase-T voltage to provide a phase angle orthogonally balanced 2Φ voltage for ST-II. The tap changers, TAP2 and TAP3, on ST-II change the output voltage magnitudes to obtain balanced 3Φ voltages.

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In the voltage phase control, the measurement of the phase angles of V̅t and V̅̅m̅̅ can be avoided. Based on the Property One of the ST, the phase angle difference between V̅t and

V̅̅m̅̅ can be obtained by comparing the magnitudes of V̅̅b̅ and V̅c: if V̅̅b̅ and V̅c have the same magnitude, the phase difference between V̅t and V̅̅m̅̅ is 90°; if V̅̅b̅ is larger than V̅c, the phase difference would be larger than 90°; if V̅̅b̅ is smaller than V̅c , the phase difference would be smaller than 90°.

The main control algorithm used in the proposed method can be divided into three operations, operations of TAP1, TAP2 and TAP3, as portrayed in Fig. 3-22, where, h1 and h2 are the acceptable tolerances and Vf is the desirable output voltage magnitude. Each of the three operations starts with the measurement of the magnitudes of the output

3Φ voltages, |V̅̅a̅|, |V̅̅b̅| and |V̅c|. The operation of TAP1, which aims at orthogonally balancing the input 2Φ voltages, works by comparing |V̅̅b̅| and |V̅c|. Action 1 refers to an increase in the phase difference between V̅t and V̅̅m̅̅ and action 2 a reduction in the phase difference. When the input 2Φ voltages have a phase shift of 90º (within an allowable tolerance), |V̅̅b̅| equals |V̅c|, the operation of TAP2 will be activated, this changes the position of TAP2 to make |V̅̅a̅| equal to Vf. Once the goal of TAP2 is achieved, the operation of TAP3 will be started to regulate both |V̅̅b̅| and |V̅c| to be Vf. When the goals of the above three operations are all achieved, i.e. |V̅̅a̅|, |V̅̅b̅| and |V̅c| are all equal to Vf within allowable tolerances, the output 3Φ voltages would be in a balanced state.

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Initial setting

Measure Va, Vb, Vc

Yes No Operation |Vb-Vc|>h1 Vb>Vc Action 1 of TAP1 Yes No Action 2

No Lower TAP2 by one tap size Yes Operation |Va-Vf|>h2 Va>Vn of TAP2 Lift TAP2 by one No Yes tap size

No Lower TAP3 by Yes one tap size Operation |Vb-Vf|>h2 Vb>Vn of TAP3 No Lift TAP3 by one Yes tap size

Fig. 3-22 Main control algorithm of proposed balancing method

The detailed control algorithms of action 1 and action 2 are illustrated in Fig. 3-23 and Fig. 3-24 respectively, where the switch pin 1, pin 2, pin 3 and the taps T(k) refer to the switch and the tap changer TAP1 in the phase regulation system shown in Fig. 3-20. Please note when TAP1 reaches its operating limit, i.e. a voltage larger than the maximum voltage provided by the phase regulating system needs to be injected into phase-T, the operation of TAP2 will be started directly. Likewise, when TAP2 reaches its limit, the operation of TAP3 would be started without equalling |V̅̅a̅| and Vf. Moreover, the operation would jump to measurement of the output 3Φ voltages once the TAP3 reaches its operating limit.

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Yes Switch to pin 2 Open T(1)

Pin 1, T(k) No Close T(k-1) k=1 Open T(k)

Tap and switch Pin 2 Close T(n) position Switch to pin 3

No Close T(k-1) k=1 Pin 3, T(k) Open T(k)

Yes Deliver an alarm: beyond regulation range; jump to the operation of TAP2

Fig. 3-23 Detailed control algorithm of action 1

Yes Deliver an alarm: beyond regulation range; jump to the operation of TAP2

Pin 1, T(k) No Close T(k+1) k=n Open T(k)

Tap and switch Pin 2 Close T(1) position Switch to pin 1

No Close T(k+1) k=n Pin 3, T(k) Open T(k) Yes Switch to pin 2 Open T(n)

Fig. 3-24 Detailed control algorithm of action 2

In this balancing method, all the operations are based on the output three phase voltage magnitudes, measured after each tap changer operation, and therefore the balancing system always tends to mitigate the latest voltage imbalance on the LV feeder. Once triggered by a voltage imbalance situation, the balancing system can capture and respond to the new voltage imbalance situation occurring during the ongoing balancing process.

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3.7 Summary

This chapter is centered on the ST based voltage balancing method, which is the heart of this project. Since this method employs transformers and tap changers, the voltage regulation methods based on tap changers and transformers are reviewed. Following this, the introduction of ST based voltage balancing method is split into four sections – overview, Scott transformer, phase regulating system and control algorithm - all of which are described at great length.

By combining (3-11), (3-28) and (3-39) together, the relationships between the voltages at the two sides of the balancing system can be expressed by (3-41).

2 √3α−1 √3α+1 − ̅̅̅ 3(1+β) 3(1+β) 3(1+β) ̅̅̅̅̅ Va VNA 1 1 1− 3α 1 1+ 3α [V̅̅̅] = − + √ − + √ ∙ [V̅̅̅̅̅] (3-41) b 3(1+β) 2(1+γ) 6(1+β) 2(1+γ) 6(1+β) NB ̅ ̅̅̅̅̅ Vc 1 1 1−√3α 1 1+√3α VNC − − + + [ 3(1+β) 2(1+γ) 6(1+β) 2(1+γ) 6(1+β) ]

By combining (3-20), (3-31) and (3-40) together, the relationships between the currents at the two sides of the balancing system can be expressed by (3-42).

2 1 1 − − 3(1+β) 3(1+β) 3(1+β) I̅A I̅a 3α−1 1 1− 3α 1 1− 3α ̅ √ + √ − + √ ̅ [IB] = 3(1+β) 2(1+γ) 6(1+β) 2(1+γ) 6(1+β) ∙ [Ib] (3-42) ̅ ̅ IC √3α+1 1 1+√3α 1 1+√3α Ic − − + + [ 3(1+β) 2(1+γ) 6(1+β) 2(1+γ) 6(1+β) ]

For the case that the input 3Φ voltage is balanced and all the tap changers are at the nominal positions, i.e. the values of α, β and γ equal zero, (3-41) and (3-42) can be simplified to (3-43) and (3-44) respectively.

2 1 1 − − V̅̅a̅ 3 3 3 V̅̅NA̅̅̅ 1 2 1 [V̅̅̅] = − − ∙ [V̅̅̅̅̅] (3-43) b 3 3 3 NB ̅ 1 1 2 ̅̅̅̅̅ Vc − − VNC [ 3 3 3]

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2 1 1 − − I̅A 3 3 3 I̅a 1 2 1 [I̅ ] = − − ∙ [I̅] (3-44) B 3 3 3 b ̅ 1 1 2 ̅ IC − − Ic [ 3 3 3]

The control methods introduced in section 3.2.4, are able to reduce the voltage imbalance in LV networks. However, compared with these, the ST based voltage balancing method has a simple physical construction and requires a basic control algorithm. A simulation study comparing the ST based balancing method with a conventional tap changer based balancing method has been carried out and is presented in Chapter 4.

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CHAPTER 4

COMPUTER SIMULATION OF ST BASED

VOLTAGE BALANCING METHOD

4.1 Introduction

This chapter presents two computer simulation studies related to the investigation of the ST based voltage balancing method introduced in Chapter 3. The proposed voltage balancing method is designed to convert a non-balanced 3Φ voltage into a balanced 3Φ voltage at a point on an LV feeder by using tap changers to regulate the voltages on a 2Φ system. The first simulation study aims to evaluate the feasibility of using this balancing method on an LV feeder. To confirm the advantages and disadvantages of this balancing method, it is compared to a different tap changer based voltage balancing method, which regulates the voltages on a 3Φ system. The methods are implemented on a simplified LV feeder with severe voltage imbalance. The second simulation study investigates the performance of the ST based voltage balancing method when used on a typical UK distribution network. The study investigates the effectiveness of this balancing method with variations in the installation location and the desirable output voltage magnitude. Both of the above simulation studies are carried out using the simulation software PSCAD/EMTDC [112]. In the following chapter, the test systems, simulation methodologies and the results of the two simulation studies are presented.

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4.2 Modelling of ST based voltage balancing system

PSCAD/EMTDC is a powerful simulation tool for studying the electric power system [113]. Its graphical user interface enables the users to schematically build a circuit, run the simulation and analyse the results in a single integrated environment. Moreover, it has a library of well-developed graphical models of various electric components, ranging from single passive elements to more complex models, e.g. electric machines [114]. This greatly helps the user to reduce his/her effort in the mathematical modelling, and concentrate on the analysis of the results. Simplicity of use is one of the most important features of PSCAD/EMTDC, achieved by the graphical nature of the employed modelling approach [113].

A model of the ST based voltage balancing system was established in PSCAD/EMTDC as shown in Fig. 4-1. This graphical model resembles the schematic circuit diagram in Fig. 3-15 and consists of Scott transformer I (ST-I), the phase regulating system and Scott transformer II (ST-II). Besides, the control algorithm described in section 3.5 is implemented by using the established digital and analogue control blocks in PSCAD/EMTDC. The three blocks at the bottom of Fig. 4-1 correspond to the three operations in the adopted control algorithm respectively.

As shown in Fig. 4-1, two 1Φ three-winding transformers are arranged in Scott connection to model an ST. At the 3Φ side of the ST, two of the three windings of the main transformer are connected in series to provide a terminal for the mid-point of the main transformer primary, and two of the three windings of the teaser transformer are connected in series to give a terminal for the neutral point of the ST; the turn number of each winding is determined according to the characteristics of the ST presented in section 3.4.

Page | 95 Chapter 4 Computer Simulation of ST based Voltage Balancing Method I I

r e m r o f s n a r t

t t o c S m e t s y s

g n i t a l u g e r

e s a h P I

r e m r o f s n a r t

t t o c S

Fig. 4-1 Diagram of ST based voltage balancing system model in PSCAD/EMTDC

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In the modelling of the phase regulating system, two series transformers are used; one is used to inject a voltage proportional to phase-M voltage into phase-T and the other is used to inject a voltage inversely proportional to phase-M voltage into phase-T. Moreover, the series transformer windings, injected into phase-T, are equipped with tap changers, which are controlled by the control signals ‘downtap’ and ‘uptap’ respectively. These tap changers correspond to the TAP1 in Fig. 3-15. During the operation of the balancing system, only one of the two series transformers is in operation, whilst the other is nullified by using the associated tap changer to reduce the turns injected into phase-T to zero. This arrangement helps to get rid of the complexity in modelling the single-pole three-throw switch used in the phase regulating system shown in Fig. 3-15.

On ST-II shown in Fig. 4-1, both the teaser and main transformers have a tap changer installed in their primaries, which are controlled by the control signals ‘tap2’ and ‘tap3’ respectively. These two tap changers correspond to the TAP2 and TAP3 in Fig. 3-15 respectively.

The tap changers in the model of the proposed balancing system are controlled by the implementation of the control algorithm illustrated in Fig. 3-22 using the existing control blocks in PSCAD/EMTDC. These generate the corresponding control signals, including ‘downtap’, ‘uptap’, ‘tap2’ and ‘tap3’. The block labelled ‘TAP1’ in Fig. 4-1 refers the operation of TAP1 in Fig. 3-22, which compares the phase-B and phase-C voltages and then generates the control signals, ‘downtap’ and ‘uptap’, to change the voltage phase angle on phase-T. The block labelled ‘TAP2’ refers to the operation of TAP2 in Fig. 3-22, which gives a control signal ‘tap2’ to control the magnitude of the output phase-A voltage. Its input signal ‘SOS’ represents the result of the operation of TAP1, based on which, the operation of TAP2 would be ignited once the operation of TAP1 has reached its goal or its operating limit. The block labelled ‘TAP3’ refers the operation of TAP3 in Fig. 3-22, which gives the control signal ‘tap3’ to control the magnitude of the output phase-B voltage. Its control signal ‘SOD’ shows the result of the operation of TAP2, which would activate the operation of TAP3 when the operation of TAP2 has reached its target or operating limit.

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4.3 Voltage balancing using tap changers

4.3.1 Simulation methodology

In order to demonstrate the feasibility of the ST based voltage balancing method, a simulation study is carried out by implementing this balancing method on a simplified LV feeder with severe voltage imbalance as shown in Fig. 4-2. The detailed network data of this LV feeder is presented in Appendix A.1. This LV feeder supplies a balanced 3Φ load from an MV/LV transformer via a 300 meter LV 3Φ four-wire cable tapped with non-balanced variable 1Φ loads (e.g. :- phase-a = 130kW, phase-b = 40kW and phase-c = 0kW). The imbalance of the loads on the three phase conductors of the LV feeder would induce non-balanced voltage drops, and then introduce a severe voltage imbalance at the end of the LV feeder. In this simulation study, the proposed balancing system is implemented at the end of the LV feeder to eliminate the imbalance caused by the non- balanced 1Φ loads; i.e. it converts the non-balanced 3Φ voltage to a balanced 3Φ voltage for the 3Φ load at the end of the LV feeder.

300m LV 3Ф cable Phase-A 130kW 5kW 11kV/433V Voltage Phase-B 40kW balancing 5kW system

Phase-C 0kW 5kW

Neutral conductor

Fig. 4-2 Simplified LV feeder used in the simulation

The ST based voltage balancing system works by using tap changers to regulate the voltages on a 2Φ system. However, the transformers and tap changers based control systems introduced in section 3.2.4 have the potential to complete the voltage balancing on LV feeders by regulating the voltages on a 3Φ system. A comparison between the ST based voltage balancing method and a tap changer used balancing method, which regulates the voltages on a 3Φ system, (called ‘3Φ balancing method’ in the following), is carried out in this study. The 3Φ balancing method, modelled in PSCAD/EMTDC and

Page | 98 Chapter 4 Computer Simulation of ST based Voltage Balancing Method implemented on the LV feeder shown in Fig. 4-2, is used to provide a balanced 3Φ voltage for the 3Φ load at the feeder end.

In this study, the simulations are carried out for three different non-balanced cases, obtained by transposing the three non-balanced 1Φ loads among the three phases of the LV feeder. The 1Φ loads are connected close to the end of the feeder to create the most severe voltage imbalance. The U2s, U0s, and magnitudes of the 3Φ voltages at the two sides of the balancing system, as well as the tap changing number required for each balancing method are recorded and analysed.

4.3.2 Three-phase balancing method

The transformer and tap changer based control systems introduced in section 3.2.4 are capable of independently regulating the voltage magnitude and phase angle, and potentially, could undertake voltage balancing on an LV feeder. However, since these control systems have a complex structure, including nine sets of tap changers, and require a complicated control algorithm, they are not investigated in this simulation study. The 3Φ balancing method is a simplified derivation of the control systems introduced in section 3.2.4 and this method is investigated and compared with the ST based balancing method.

Fig. 4-3(a) shows the circuit diagram of the 3Φ voltage balancing system. Similar to the ST based voltage balancing system, the 3Φ balancing system contains a phase regulating system and a magnitude regulating system. The phase regulating system consists of two booster transformers; the first one regulates the phase angle of the phase-B voltage so that the phase-B voltage lags the phase-A voltage by 120º and the second one is used to regulate the phase-C voltage phase angle so that the phase-C voltage leads phase-A by 120º. This phase regulating system ensures the phase angles of the 3Φ voltages are balanced, i.e. the sequence is A-B-C and the angular difference between each is 120º. In the voltage magnitude regulating system, three 1Φ tap changing transformers are connected in each phase of the LV feeder to control the corresponding voltage magnitude independently. Fig. 4-3(b) shows the graphical model of the 3Φ balancing system in PSCAD/EMTDC.

Page | 99 Chapter 4 Computer Simulation of ST based Voltage Balancing Method

C   B C  V  A B  V  A V m e t s y s

g n i t a l u g 3 5 4 e p p p r

a a a e T T T d u t i n g a M 2 p a T m e t s y s g n i t a l u g e r e s 1 a p h a P T

l A B C a r t u e (a) N (b)

Fig. 4-3 Three-phase balancing system (a) circuit diagram (b) simulation model in PSCAD/EMTDC

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The control algorithm adopted in the 3Φ balancing system is briefly illustrated by the flowchart in Fig. 4-4, where, Va, Vb, Vc, θa, θb, θc refer to the 3Φ voltage magnitudes and phase angles at the output of the balancing system, θs and h2 refer to the acceptable tolerances and Vf is the desirable output voltage magnitude. The operations of Tap1 and Tap2 change the phase angles of phase-B and phase-C voltages respectively to balance the phase angles of the 3Φ voltages. The operations of Tap3, Tap4 and Tap5 change the 3Φ voltage magnitudes respectively so they are all equal to the desirable value Vf. As shown in Fig. 4-4, the operation of Tap1 precedes the operation of Tap2, after which, the parallel operation of Tap3, Tap4, and Tap5 follows. Moreover, Va, Vb, Vc, θa, θb, θc are measured before every tap changing operation so that the balancing system always captures the latest voltage imbalance.

Initial setting

Measure Va, Vb, Vc, θa, θb, θc

Yes |θa-θb-120°|>θs Operation of Tap1 Operation of phase No regulating system Yes |θa-θc+120°|>θs Operation of Tap2

Fig. 4-4 Control algorithm used in the 3Φ balancing system

Table 4-1 shows a comparison between the ST based balancing method and the 3Φ balancing method in both its hardware requirement and complication of the control algorithm. The ST based balancing system has a simpler structure; it contains one booster transformer and three sets of tap changers, as compared to the two booster transformers and five sets of tap changers required by the 3Φ balancing system. Besides, the control

Page | 101 Chapter 4 Computer Simulation of ST based Voltage Balancing Method algorithm for the ST based balancing system only requires the measurement of the 3Φ voltage magnitudes; whilst the operation of the 3Φ balancing system relies on the measurement of the 3Φ voltage magnitudes and phase angles.

Table 4-1 Comparison between ST based balancing method and 3Φ balancing method Balancing system ST balancing method Three-phase balancing method Two Scott transformers, Three 1Φ tap changer transformers, Hardware One booster transformer, Two booster transformers, requirement Three sets of tap changers Five sets of tap changers

Control Measurement of the 3Φ Measurement of the 3Φ voltage algorithm voltage magnitudes magnitudes and phase angles

4.3.3 Simulation results

The ST based balancing system and the 3Φ balancing system are applied on the LV feeder shown in Fig. 4-2 in the three cases as indicated in Table 4-2. These cases are obtained by transposing the 1Φ loads amongst the three phases. Moreover, in order to compare the performance of the two investigated balancing systems, they are configured to have the same characteristics of tap changer control for the phase regulating system as shown in Table 4-3 (configurations of TAP1 in the ST based balancing system, and Tap1, Tap2 in the 3Φ balancing system), and for the voltage magnitude regulating system as shown in Table 4-4 (configurations of TAP2, TAP3 in the ST based balancing system, and Tap3, Tap4, Tap5 in the 3Φ balancing system). As shown in Table 4-3, the tap changers used for phase regulation in the 3Φ balancing system are specified with a tolerance of ±0.21º. However, since the phase regulating system in the ST based balancing system works by comparing the phase-B and phase-C voltage magnitudes, an equivalent tolerance of ±0.75V is used; with all the voltages at their nominal values in the balancing system, a phase shift of 90º±0.21º between phase-T and phase-M means a magnitude difference of 0.75V between the phase-B and phase-C voltages.

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Table 4-2 Three cases investigated in the simulation Case No. Description 1 Phase-A = 130kW, Phase-B = 40kW, Phase-C = 0kW 2 Phase-A = 0kW, Phase-B = 130kW, Phase-C = 40kW 3 Phase-A = 40kW, Phase-B = 0kW, Phase-C = 130kW

Table 4-3 Configuration for the phase regulating system Tolerances ±0.75V/±0.21º Max. tap changing No. 20 Range of injected voltage -23V to + 23V

Table 4-4 Configuration for the voltage magnitude regulating system Tolerances ±0.55V Max. tap changing No. 32 Tap changing range 0.8 p.u. to 1.2 p.u. Desirable output voltage 240V

Table 4-5 shows the magnitudes of the 3Φ voltages at the two sides of the ST based balancing system in the three investigated cases, and the equivalents of the 3Φ balancing system are shown in Table 4-6. In these tables, the magnitudes of the 3Φ voltages before the balancing system are outside the statutory limits; the voltage on the phase supplying the 130kW load drops below the lower voltage limit (217V) and the voltage on the phase supplying the 0kW load rises above the upper voltage limit (253V). This indicates the connection of non-balanced 1Φ loads along the LV feeder could cause violations of the voltage magnitude limits. Moreover, based on the fact that the secondary voltage of the MV/LV transformer is 250V (433⁄√3) phase-neutral, the disparity of the 3Φ voltages before the balancing system manifests that severely non-balanced voltage drops occur along the LV feeder. On the contrary, the voltages after the balancing systems are all within the voltage magnitude limits. Of special significance is the voltages after the ST based balancing system, shown in Table 4-5, which are all within 240V ± 1V. This indicates that the violations of the voltage magnitude limits are mitigated by both balancing methods.

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Table 4-5 Voltage magnitude at two sides of ST based balancing system Case Before After No. A B C A B C 1 210.4V 249.0V 265.9V 240.6V 240.2V 240.4V 2 265.9V 211.3V 247.9V 240.5V 240.1V 240.7V 3 247.6V 265.9V 211.8V 239.9V 240.2V 240.0V

Table 4-6 Voltage magnitude at two sides of 3Φ balancing system Case Before After No. A B C A B C 1 213.0V 248.6V 265.5V 234.3V 240.5V 239.9V 2 265.5V 213.1V 248.6V 239.8V 234.7V 239.9V 3 248.6V 265.5V 213.1V 240.1V 240.0V 234.7V

Table 4-7 and Table 4-8 show the U2s and U0s of the 3Φ voltages before and after the balancing systems in the three cases respectively. When non-balanced 1Φ loads are connected to the LV feeder, non-balanced voltage drops along the LV feeder induce severe voltage imbalance. The 3Φ voltages before the voltage balancing systems are non- balanced; the U2s are all above 2.4% and the U0s are all above 10% as shown in Table 4-7 and Table 4-8. For each investigated case, the U2 of the 3Φ voltage at the termination of the feeder is reduced to be around 1.2% by the 3Φ balancing system, or less than 0.2% by the ST based balancing system as shown in Table 4-7, and the U0 is reduced to around 2% by the 3Φ balancing system but zero by the ST based balancing system as shown in Table 4-8. These indicate that both balancing systems can reduce the voltage imbalance on the LV feeder, and moreover, the effectiveness of the ST based balancing system is superior to that of the 3Φ balancing system.

Table 4-7 Values of U2s ST based balancing system 3Φ balancing system Case No. before after before After 1 2.53% 0.11% 2.45% 0.93% 2 2.53% 0.16% 2.46% 1.26% 3 2.52% 0.06% 2.46% 1.59%

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Table 4-8 Values of U0s ST based balancing system 3Φ balancing system Case No. before after before After 1 11.5% 0 10.8% 0.73% 2 11.2% 0 10.7% 2.49% 3 11.0% 0 10.7% 2.83%

Table 4-9, Table 4-10 and Table 4-11show the tap changing numbers carried out by the investigated balancing systems for voltage balancing in the three investigated cases.

Table 4-9 Tap changing number in case 1 Subsystem ST based balancing system 3Φ balancing system Phase regulating Tap1 = 16 TAP1 = 8 system Tap2 = 18

Tap3 = 32 Voltage magnitude TAP2 = 4 Tap4 = 10 regulating system TAP3 = 6 Tap5 = 31

Table 4-10 Tap changing number in case 2 Subsystem ST based balancing system 3Φ balancing system Phase regulating Tap1 = 13 TAP1 = 0 system Tap2 = 20

Tap3 = 31 Voltage magnitude TAP2 = 9 Tap4 = 32 regulating system TAP3 = 7 Tap5 = 10

Table 4-11 Tap changing number in case 3 Subsystem ST based balancing system 3Φ balancing system Phase regulating Tap1 = 20 TAP1 = 18 system Tap2 = 14

Tap3 = 11 Voltage magnitude TAP2 = 2 Tap4 = 30 regulating system TAP3 = 5 Tap5 = 32

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Some tap changers in the 3Φ phase balancing system reach their limits without achieving the voltage balancing target. This is because the voltage imbalance severity is beyond its regulating range, as demonstrated by the tap changing number for Tap 3 in Table 4-9, Tap 2 and Tap 4 in Table 4-10, and Tap 1 and Tap 5 in Table 4-11, i.e. they are all at their full range values. However, the tap changers in the ST based balancing system do not reach their maximum values. Therefore, it can be concluded that with the same tap changing characteristics the ST based balancing system has a larger regulating range than the 3Φ balancing system. This also explains the superiority of the ST based balancing system over the 3Φ balancing system in the voltage balancing efficacy. Moreover, by comparing the total tap changing number of the two balancing systems in each investigated case, it can be found the ST based balancing system requires less tap change operations than the 3Φ balancing system for the same voltage imbalance, e.g. in Table 4- 9, the ST based balancing system performs 8 tap changes in phase angle regulation, less than 32 tap changes performed by the 3Φ balancing system; the ST based balancing system performs 10 tap changes in magnitude regulation, less than 73 tap changes performed by the 3Φ balancing system. Therefore the ST based balancing system would require less regulation time and introduce less transients into the network than the 3Φ balancing system.

4.4 Application in a typical UK distribution network

4.4.1 Simulation methodology

In the above simulation study the ST based balancing system was tested and verified on a simplified LV feeder, supplying several lumped loads. The results show the ST based balancing system can complete the voltage balancing on the LV feeder and it is superior to the 3Φ balancing system. However, the performance of the ST based balancing system is affected by both the balancing system itself, including its installation location on the LV feeder and the desirable output voltage magnitude, and the LV network situation. In this simulation study, a distribution network developed from the UK generic distribution network (UKGDN) [115] is used and the ST based balancing system is implemented on one of its LV feeders, which is modelled in detail and supplies 60 customers. To investigate how the performance of the ST based balancing system varies with its

Page | 106 Chapter 4 Computer Simulation of ST based Voltage Balancing Method installation location and output voltage magnitude; the following case studies are carried out: Case I: Impact of EV penetration Case II: Performance of balancing system with different output voltage magnitudes Case III: Performance of balancing system, when installed at different locations

4.4.2 Distribution network model

A distribution network developed from the UKGDN is used in this simulation study [115]. As shown in Fig. 4-5, the network includes a 33/11.5kV substation that supplies six 11kV feeders and each of the 11kV feeders supplies eight evenly distributed 11/0.433kV substations. Feeder F1 is used for the investigation in the simulation study; and it is supplied by the LV substation at the remote end of an 11kV feeder. For computational reasons, feeder F1 is modelled in detail, but the other feeders are simplified and modelled as lumped loads. The detailed network data of this network is presented in Appendix A.2.

Feeder F1 consists of a 150m 300mm² LV cable and a 150m 95mm² LV cable. It supplies via load connection points (LCPs), 60 customers, evenly distributed across the three phases and along the feeder. The distance between adjacent LCPs is 30m, and at each LCP, two customers are supplied per phase. Since this study aims to investigate the voltage imbalance mitigation, different types of resistive residential loads are connected on each phase of the feeder F1; every customer on phase-A has a load of 2kW; every customer on phase-B 1.3kW and every customer on phase-C 0.7kW. Moreover, the loads on the three phases of the feeder F2 are made unequal to model the background voltage imbalance for F1; all the other feeders have balanced 3Φ loads [116]. In addition, the cross-sections of the network feeders are designed suitably according to the network thermal limits and voltage drop limits. In the simulation study, the three phase-neutral voltages at each LCP along the LV feeder and their associated U0s and U2s were recorded and analysed.

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4.4.3 Simulation results

Case I: EVs are connected to the LV feeder F1 in a non-balanced manner designed to induce severe voltage imbalance. The results of this case are used as a reference for the following case studies, where the ST based balancing system is implemented on the feeder F1 to mitigate the voltage imbalance and magnitude variation caused by the EVs.

In case I, ten EVs are evenly connected to the phase-A conductor of Feeder F1, i.e. one EV is connected to phase-A at each LCP. All the EVs are in charging mode with a charging power level of 6kW, which is feasible using the existing UK domestic power supply infrastructure. With this penetration of EVs, the maximum current flowing in the LV cable is still below its continuous current limit.

The simulation is executed when the LV feeder F1 operates with and without the penetration of EVs. The per unit values (based on 230V) of the 3Φ voltages (Va, Vb, Vc) at the LCPs, obtained from the simulation, are shown in Table 4-12 and Fig. 4-6, in which, the LCPs are identified by their distance to the LV substation. Because of the load imbalance, Va, Vb and Vc diverge as the distance increases in the two investigated situations, and moreover, since the load on phase-B is in-between, Vb is always between Va and Vc. When the EVs are not connected, Va decreases by 0.034 pu, whilst Vc increases by 0.006 pu, along the length of the LV feeder. Nevertheless, Va, Vb and Vc are all within the statutory limits and above 1.0pu (230V). But when EVs are connected to the LV feeder, the divergence of Va, Vb and Vc is deteriorated; Va decreases by 0.099 pu and Vc increases by 0.023 pu. As a result, Va is lower than 1.0pu and Vc raises above the upper limit of 1.1pu (253V) at the end of the LV feeder. The voltage magnitude limits of the LV network are breached when the EVs are connected in a non-balanced manner.

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Table 4-12 Three phase voltages in case I Three phase voltages (pu) Distance Without EVs With EVs (meter) Va Vb Vc Va Vb Vc 0 1.070 1.078 1.080 1.062 1.078 1.081 30 1.067 1.076 1.081 1.048 1.078 1.084 60 1.063 1.074 1.082 1.036 1.079 1.087 90 1.059 1.072 1.083 1.026 1.079 1.090 120 1.056 1.070 1.084 1.017 1.079 1.093 150 1.053 1.069 1.085 1.009 1.080 1.095 180 1.047 1.066 1.085 0.994 1.079 1.098 210 1.043 1.063 1.086 0.981 1.079 1.100 240 1.039 1.061 1.086 0.972 1.079 1.102 270 1.037 1.060 1.086 0.966 1.078 1.103 300 1.036 1.059 1.086 0.963 1.078 1.104

3Φ voltages with/without connection of EVs

1.15 Va 1.1 Vb Vc 1.05 Va with EVs 1 Vb with EVs Vc with EVs

Voltage magnitude magnitude (pu) Voltage 0.95 0 30 60 90 120 150 180 210 240 270 300 Distance (m)

Fig. 4-6 Three phase voltages along the LV feeder with/without connection of EVs

The U0 and U2, at each LCP when the LV feeder operates with EVs, are always higher than those without EVs as shown in Table 4-13. When there is no EV connected, the U2s are less than 0.7% and the U0s are less than 2.2% along the entire LV feeder, whereas with EVs the U2 raises above 2% at 240m and the U0 reaches 6.42% at the end of the feeder. Accordingly this particular penetration of EVs worsens the voltage imbalance on the LV feeder. In the following case studies, the ST based balancing system is implemented on the feeder F1 with the EVs connected. Additionally, since the phase-B voltage magnitude is always medium across the 3Φ voltages, only the magnitudes of phase-A and phase-C voltages are analysed in the following case studies.

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Table 4-13 The U0s and U2s in case I The U0s and U2s Distance without EVs With EVs (meter) U0 U2 U0 U2 0 0.39% 0.23% 0.82% 0.77% 30 0.59% 0.28% 1.53% 0.95% 60 0.80% 0.33% 2.23% 1.11% 90 1.00% 0.37% 2.86% 1.25% 120 1.16% 0.41% 3.41% 1.38% 150 1.31% 0.44% 3.89% 1.49% 180 1.57% 0.51% 4.73% 1.70% 210 1.79% 0.56% 5.40% 1.88% 240 1.95% 0.60% 5.91% 2.01% 270 2.05% 0.63% 6.25% 2.10% 300 2.11% 0.64% 6.42% 2.15%

Case II: The balancing system is located just ahead of the LCP at 90m along the LV feeder. The performance of the balancing system is investigated with different output voltage magnitudes, including 1.043pu (240V), 1.052pu (242V), 1.061pu (244V), 1.07pu (246V), 1.078pu (248V) and 1.087pu (250V).

Table 4-14 and Table 4-15 show the phase-A voltages and phase-C voltages along the LV feeder respectively for different output voltage magnitudes. It is clearly that at 90m, both the phase-A voltage and the phase-C voltage are within tolerance (±0.56%) around the desirable output magnitude. However, when the desirable output voltage of the balancing system is 1.043pu, 1.052pu or 1.061pu, the phase-A voltage at the end of the feeder is below 1.0pu. In addition, when the desirable output voltage is 1.087pu, overvoltage occurs on phase-C; the phase-C voltage rises above 1.1pu at locations beyond 180m.

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Table 4-14 Phase-A voltage in case II Phase-A voltage (pu) Distance Output Output Output Output Output Output (m) 1.043 1.052 1.061 1.070 1.078 1.087 0 1.065 1.065 1.065 1.064 1.064 1.064 30 1.059 1.058 1.058 1.058 1.057 1.057 60 1.053 1.053 1.052 1.052 1.052 1.051 90 1.045 1.054 1.061 1.070 1.077 1.087 120 1.036 1.045 1.051 1.061 1.067 1.077 150 1.028 1.037 1.043 1.053 1.059 1.069 180 1.012 1.021 1.027 1.036 1.042 1.052 210 0.999 1.008 1.014 1.023 1.029 1.039 240 0.989 0.998 1.004 1.013 1.019 1.029 270 0.983 0.992 0.998 1.007 1.013 1.022 300 0.980 0.988 0.994 1.003 1.010 1.019

Table 4-15 Phase-C voltage in case II Phase-C voltage (pu) Distance Output Output Output Output Output Output (m) 1.043 1.052 1.061 1.070 1.078 1.087 0 1.081 1.081 1.081 1.081 1.081 1.081 30 1.081 1.081 1.081 1.081 1.081 1.081 60 1.080 1.080 1.080 1.080 1.080 1.080 90 1.046 1.055 1.064 1.072 1.081 1.091 120 1.048 1.058 1.067 1.074 1.084 1.094 150 1.050 1.060 1.069 1.076 1.086 1.096 180 1.053 1.063 1.072 1.080 1.089 1.099 210 1.056 1.066 1.074 1.082 1.091 1.102 240 1.058 1.067 1.076 1.084 1.093 1.104 270 1.059 1.069 1.078 1.085 1.094 1.105 300 1.059 1.069 1.078 1.086 1.095 1.105

Table 4-16 shows that the U0s at each LCP are almost the same for different output voltage magnitudes. The U0 is always 0.02% at 90m, which demonstrates that the output 3Φ voltages of the balancing system have negligible zero sequence components. Table 4- 17 and Fig. 4-7 show the U2s along the LV feeder. The U2, at each LCP located before the balancing point (90m), rises slightly as the output voltage magnitude increases. This is because the loads are all modelled as resistance (based on 230V) and higher output voltage would draw more non-balanced currents. The U2s at 90m are reduced to slightly different values due to the tolerances in the tap control as shown in Fig. 4-7. Moreover, the U2 is always below 1.3% along the whole LV network in all the scenarios. This case

Page | 112 Chapter 4 Computer Simulation of ST based Voltage Balancing Method study indicates that when the output voltage magnitude is 1.07pu or 1.078pu, both the voltage magnitude and imbalance along the LV feeder are within the statutory limits.

Table 4-16 The U0 in case II U0 Distance Output Output Output Output Output Output (meter) 1.043 1.052 1.061 1.070 1.078 1.087 0 0.31% 0.31% 0.31% 0.31% 0.31% 0.31% 30 0.47% 0.47% 0.48% 0.48% 0.48% 0.48% 60 0.54% 0.56% 0.58% 0.58% 0.58% 0.58% 90 0.02% 0.02% 0.02% 0.02% 0.02% 0.02% 120 0.59% 0.59% 0.59% 0.59% 0.59% 0.59% 150 1.10% 1.10% 1.10% 1.10% 1.10% 1.10% 180 1.98% 1.99% 1.99% 1.99% 1.99% 1.98% 210 2.71% 2.71% 2.70% 2.71% 2.70% 2.70% 240 3.24% 3.24% 3.24% 3.24% 3.24% 3.23% 270 3.60% 3.60% 3.60% 3.60% 3.60% 3.59% 300 3.78% 3.78% 3.77% 3.78% 3.77% 3.77%

Table 4-17 The U2 in case II U2 Distance Output Output Output Output Output Output (meter) 1.043 1.052 1.061 1.070 1.078 1.087 0 0.79% 0.80% 0.81% 0.82% 0.83% 0.84% 30 0.97% 0.99% 0.99% 1.01% 1.02% 1.03% 60 1.14% 1.15% 1.16% 1.18% 1.19% 1.21% 90 0.07% 0.08% 0.21% 0.11% 0.23% 0.26% 120 0.19% 0.21% 0.34% 0.23% 0.36% 0.38% 150 0.31% 0.32% 0.45% 0.35% 0.47% 0.49% 180 0.54% 0.55% 0.69% 0.57% 0.71% 0.73% 210 0.73% 0.74% 0.89% 0.76% 0.91% 0.92% 240 0.88% 0.89% 1.04% 0.91% 1.06% 1.07% 270 0.98% 0.99% 1.14% 1.01% 1.16% 1.17% 300 1.03% 1.03% 1.19% 1.06% 1.21% 1.22%

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U2s when balancing system has different output magnitudes 1.40% 1.20% 1.043 pu 1.00% 1.052 pu

0.80% 1.061 pu U2 0.60% 1.07 pu 0.40% 1.078 pu 0.20% 1.087 pu 0.00% 0 30 60 90 120 150 180 210 240 270 300 Distance (m)

Fig. 4-7 U2s when balancing system has different output magnitude

Case III: In this case, the performance of the ST based voltage balancing system is investigated when it is installed at different locations along the LV feeder. The output voltage magnitude of the balancing system is always set to be 1.07pu (246V). The investigated installation locations include 30m, 60m, 90m, 120m and 150m, and moreover, the balancing system precedes the LCP at each location.

The phase-A voltages and phase-C voltages at each LCP are illustrated in Fig. 4-8 and Fig. 4-9 respectively, which show that the voltage profiles vary significantly with the installation location of the balancing system. The phase-A voltage at the feeder end increases as the installation location moves away from the MV/LV transformer; in the contrary, the phase-C voltage at the feeder end decreases. Moreover, the phase-A voltage and phase-C voltage at the LCP following the balancing system are both regulated to be around 1.07pu, irrespective of the location of the balancing system. Nevertheless, the phase-A voltage at the end of the feeder drops below 1.0pu when the balancing system is installed at 30m or 60m on the LV feeder.

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Phase-A voltage with balancing system installed at different locations 1.08 1.07 1.06 at 30m 1.05 at 60m 1.04 1.03 at 90m 1.02 at 120m 1.01 Voltage (pu) Voltage 1.00 at 150m 0.99 0.98 0 30 60 90 120 150 180 210 240 270 300 Distance (m)

Fig. 4-8 Phase-A voltage with balancing system installed at different locations

Phase-C voltage with balancing system installed at different locations 1.095 1.090

at 30m

1.085 at 60m 1.080 at 90m 1.075 at 120m

Voltage (pu) Voltage at 150m 1.070 1.065 0 30 60 90 120 150 180 210 240 270 300 Distance (m)

Fig. 4-9 Phase-C voltage with balancing system installed at different locations

The U0s at different LCPs along the LV feeder are illustrated in Fig. 4-10. The U0 at the LCP immediately following the balancing system is always reduced to be nearly zero, regardless of the installation location, which again demonstrates that the balancing system can eliminate the voltage zero sequence component. Moreover, as the balancing system moves far away from the LV substation, the U0 at the feeder end declines.

Fig. 4-11 shows the U2 profile along the LV feeder is always within the statutory limits in all the investigated scenarios. The U2 at the LCP following the balancing system is reduced to a small value in difference scenarios; the inequality of these values is due to

Page | 115 Chapter 4 Computer Simulation of ST based Voltage Balancing Method the tolerances in the tap changer control. In addition, the U2 profile along the LV feeder varies significantly with the location of the balancing system. The U2, at the LCP preceding the balancing system, rises above 1.2% when the balancing system is installed at 120m or 150m on the LV feeder. Besides, when the balancing system is installed at 30m or 60m, the U2 at the end of the feeder exceeds 1.3%. Therefore, an appropriate installation location of the balancing system should be determined to ensure the U2s at the LCPs before and after the balancing system are at minimum.

U0 with balancing systen installed at different locations

6.00%

5.00% at 30m 4.00% at 60m 3.00% at 90m 2.00% at 120m 1.00% at 150m 0.00% 0 30 60 90 120 150 180 210 240 270 300 Distance (m)

Fig. 4-10 U0 with balancing system installed at different locations

U2 with balancing systen installed at different locations 1.80% 1.60% 1.40% at 30m 1.20% at 60m 1.00% at 90m 0.80% 0.60% at 120m 0.40% at 150m 0.20% 0.00% 0 30 60 90 120 150 180 210 240 270 300 Distance (m)

Fig. 4-11 U2 with balancing system installed at different locations

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4.5 Summary

This chapter presents two PSCAD/EMTDC simulation studies related to the ST based voltage balancing method.

In the first study, the ST based voltage balancing method is implemented on a simplified LV feeder with significant voltage imbalance, and its balancing performance is compared with results from a 3Φ balancing system. The simulation studies show both balancing systems can mitigate the excessive voltage imbalance and magnitude variation caused by non-balanced 1Φ loads. However, the ST based balancing system is superior to the 3Φ balancing system; with the same tap changing configuration, the ST based balancing system has a larger regulating range, and requires less tap changing operations. Besides, the ST based balancing system has a less complex physical structure and requires a simple control algorithm.

In the second study, the ST based balancing system is implemented in a typical UK distribution network and used to mitigate the excessive voltage imbalance and magnitude variation caused by EVs. It is shown that the connection of EVs in a non-balanced manner could induce unacceptable voltage imbalance and magnitude variation on the LV feeder. The performance of the ST based balancing system varies significantly with its output voltage magnitude; a high output voltage magnitude may make the voltage on the light-load phase rise above the upper voltage limit and a low output voltage magnitude may make the voltage on the heavy-load phase drop below the lower voltage limit. Accordingly, an appropriate voltage magnitude should be selected to deal with the divergence of the 3Φ voltages. Moreover, the performance of the balancing system also varies significantly with its installation location. The balancing system essentially divides the LV feeder into two sections; the balancing system can only control the voltages on the feeder following the balancing system. When the balancing system is located too far away from the LV substation, the voltages on the LCPs ahead of the balancing system may violate the voltage limits; whilst when the balancing system is too close to the LV substation, the voltages at the feeder end may breach the limits. Therefore, the location of the balancing system should be selected to ensure the voltages along the whole feeder comply with the limits.

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CHAPTER 5

PHYSICAL TEST OF ST BASED VOLTAGE

BALANCING METHOD &

ASSOCIATED SIMULATION

5.1 Introduction

In the previous chapter, the ST based voltage balancing method is modelled and tested in the computer simulation studies by use of PSCAD/EMTDC. The balancing performance of the ST based voltage balancing method on both a simplified LV feeder and a typical UK distribution network with the penetration of EVs is investigated, and the results show that this method is capable of completing the voltage balancing in the LV networks. However, the computer simulation models, by their nature, reflect the real-world power system in an artificial environment. In order to demonstrate the feasibility of the ST based voltage balancing method in a more convincing way, a laboratory based physical test is carried out in this chapter.

In this physical test, a small-scale physical voltage balancing system is built based on the ST based balancing method, and implemented on an LV feeder, with 1Φ loads tapped

Page | 118 Chapter 5Physical Test of ST based Voltage Balancing Method &Associated Simulation along the feeder. The transient response of the physical balancing system to the voltage variations on the LV feeder caused by the variable 1Φ loads is investigated. Moreover, a computer simulation study is carried out in accordance with the similar scenarios, investigated using the laboratory physical test system. In this chapter, the experimental methodology, the features of the physical balancing system and the experimental plant, a summary of the experimental results, and the corresponding computer simulation study are presented in detail.

5.2 Experimental methodology

A physical voltage balancing system was established in the laboratory based on the proposed ST based balancing method. All the tap changers in this balancing system are controlled by a microprocessor, programmed according to the control algorithm described in section 3.6. The system is capable of automatically changing the tap positions to balance the voltages on an LV feeder and to ensure they are within specified limits. The system is implemented on a laboratory based LV feeder, which supplies a balanced 3Φ load at its termination and non-balanced 1Φ loads tapped along the length. The physical test is carried out in three scenarios, created by three events - activating the balancing system, disconnecting a 1Φ load and reconnecting the load. During the test, the waveforms of the 3Φ voltages are recorded. The 3Φ voltages at the output of the balancing system have no zero sequence component, i.e. U0 is always zero, hence only the U2 is analysed, which reflects the balancing performance of the system.

5.3 Experimental plant

Fig. 5-1 shows an overview of the experimental plant, which comprises the physical voltage balancing system, the microprocessor based control system and the LV feeder. The details of the hardware used in this physical test are presented in Appendix C.

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5.3.1 Physical voltage balancing system The diagram of the established physical voltage balancing system shown in Fig. 5-1 resembles the circuit diagram shown in Fig. 3-15. It consists of six transformers TR1, TR2, TR3, TR4, TR5 and TR6, constructed by rewinding commercially available transformers according to the properties of the STs and the phase regulating system described in section 3.4 and 3.5. Transformers TR1 and TR2 constitute Scott transformer I; TR3 and TR4 constitute Scott transformer II; in addition, TR5 and TR6 constitute the phase regulating system.

Moreover, in this physical balancing system, miniature relays are used to vary the turn number of the transformers TR3, TR4 and TR5, realizing the function of the tap changers and the switch. There is a miniature relay connected in series in each tap position and the relays S1, S2 and S3 replace the single-pole three-throw switch. The tap changing is accomplished by opening the miniature relay in the previous tap position immediately after closing the miniature relay in the new tap position. Each miniature relay has an individual relay tripping circuit, and in this way, these relays are all controlled by the control signals from a microprocessor based control system.

5.3.2 Microprocessor based control system

The microprocessor based control system shown in Fig. 5-1 consists of a microprocessor, an analogue input circuit board, a miniature relay board and an interfacing board. The microprocessor is the heart of this control system. It is programmed according the control algorithm introduced in section 3.6 to measure the magnitudes of the three phase-to- neutral voltages exported from the balancing system and send control signals to either open or close the miniature relays. Since the microprocessor has specific input requirements, three analogue input circuits are established to convert the three phase- neutral voltages into acceptable signals respectively, and all these circuits are installed on the analogue input circuit board. Moreover, all the miniature relays used in the physical balancing system and their associated tripping circuits are installed on the miniature relay board. The interfacing board is used to interface the microprocessor with the analogue input circuits and the relay tripping circuits.

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5.3.3 Low voltage feeder

The LV feeder used in this physical test is supplied from a 400V 3Φ power source in the laboratory as shown in Fig. 5-1; a 150m 3Φ four-wire cable is used to draw power from the power source through isolation transformers and a 3Φ variac to supply three 40W bulbs (based on 230V), connected at the cable end. The isolation transformers are used to provide galvanic isolation between the LV mains supply and the experimental plant to avoid tripping other power supplies in the laboratory in case of faults occurring in the physical test. The 3Φ variac is used to change the magnitude of the 3Φ voltage supplied to the experimental plant, and in this physical test, the supplied voltage from the variac is 110V phase-to-neutral. Besides, a 200W (based on 230V) bulb is tapped on phase-A and a 60W bulb (based on 230V) is tapped on phase-C along the LV cable. These two bulbs are used to introduce non-balanced 3Φ voltages for the downstream bulbs. The physical balancing system, in the dashed rectangle in Fig. 5-1, is implemented on this LV feeder for voltage balancing and it is located before the three 40W bulbs to provide them with balanced 3Φ voltages.

5.4 Experimental results

Three cases are investigated in this physical test. A summary of the experimental results are presented in the following.

Case 1: This case investigates the balancing performance of the physical balancing system by comparing the 3Φ voltages at the output terminals of the balancing system before and after the system is started. As shown in Fig. 5-2, when the balancing system is switched off, the 3Φ voltages are severely non-balanced; the voltage on phase-A, supplying the 200W bulb, has the least magnitude, while the voltage on phase-B, with no load tapped along, has the largest magnitude, and the associated U2 is 9.52%. At 0.63s, the physical balancing system is activated, and in the following 200ms, the 3Φ voltages are gradually regulated to have the same magnitude. After the completion of the balancing process, the U2 becomes 1.42%.

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200 Voltage balancing system is started

150 200ms U2: 9.52% U2: 1.42%

100

0 Voltage (Volt) Voltage -50

Phase-A -100 Phase-B Phase-C -150 0.63

0.55 0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95 1

Time (Sec) Fig. 5-2 Three-phase voltage waveforms when balancing system is initiated

Case 2: The loads in the LV networks are always varying due to the uncontrolled human behaviour and occasionally the variations are significant and abrupt. This case aims at investigating the transient response of the balancing system to the voltage disturbance induced by the disconnection of the 200W bulb from phase-A. As shown in Fig. 5-3, U2 is initially 1.2% with the balancing system in operation. At 0.88s, the 200W bulb is disconnected, which induces an immediate 10% phase-A voltage rise and the 3Φ voltages become severely non-balanced. At the same time, the balancing system responds and after about 300ms, the 3Φ voltages are restored to a balanced state and the resultant U2 is 1.79%.

Case 3: This case investigates the transient response of the balancing system to a voltage distortion caused by the reconnection of the 200W bulb to phase-A. As shown in Fig. 5-4, U2 is initially 1.0% with the balancing system in operation. At 0.77s, the 200W bulb is reconnected to phase-A conductor, which immediately induces a 25% voltage drop on phase-A. At the same time, the balancing system is triggered, and after about 260ms, the balancing system restores the 3Φ voltages to a relatively balanced state and the resultant U2 is1.21%.

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200 Load reduction occurs on phase-A, 10% voltage rise happens to phase-A 300ms 150 U2: 1.2% U2: 1.79%

100

0 Voltage (Volt) Voltage -50

-100 Phase-A Phase-B Phase-C -150 0.88

0.7 0.8 0.9 1 1.1 1.2 1.3 Time (Sec)

Fig. 5-3 Three-phase voltage waveforms when load reduction occurs

200 Load rise occurs on phase-A, 25% voltagedrop happens to phase-A 150 260ms U2: 1% U2: 1.21%

100

0 Voltage (Volt) Voltage -50

Phase-A -100 Phase-B Phase-C -150 0.77

0.6 0.7 0.8 0.9 1 1.1 1.2 Time (Sec)

Fig. 5-4 Three-phase voltage waveforms when load rise occurs

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5.5 Associated simulation study

5.5.1 Simulation methodology

A computer simulation study is carried out in the similar scenario as the physical test. This simulation study is a continuance of the simulation study presented in section 4.4; the balancing system always outputs 240V (1.043pu) and is installed just ahead of the LCP at 300m, where the voltage imbalance is the most serious. The balancing system is initially at rest, i.e. all the taps are at their nominal positions; at 1s, the balancing system is started; at 2.5s, all the EVs in the LV network preceding balancing system are disconnected, and then they are reconnected to phase-A at 4s.

5.5.2 Simulation results

Fig. 5-5 shows the profiles of the 3Φ voltage RMS values during the simulation. The 3Φ voltages are originally severely non-balanced and remain unchanged until the start-up of the balancing system at 1s.

The balancing process, lasting from 1.0s to 1.6s, can be divided into three sections. The first section is from 1s to 1.3s; during this period, the phase-B and phase-C voltages are gradually regulated to have the same magnitude attributable to the operation of TAP1. The second section is from 1.3s to 1.46s; during this time, the phase-A voltage increases to 240V due to the operation of TAP2. Because the operation of TAP2 influences phase- B and phase-C, their voltage RMS values are also changed during this period as shown in Fig. 5-5. In the third section, from 1.46s to 1.6s, the operation of TAP3 draws phase-B and Phase-C voltages to 240V, after which the voltage balancing is completed.

At 2.5s, voltage distortion occurs to the 3Φ voltages as a result of the disconnection of the EVs. The balancing system responds immediately; 390ms later, the 3Φ voltages are restored to 240V. At 4s, due to the reconnection of the EVs, again the 3Φ voltages are distorted and the balancing system is triggered. After 340ms, the 3Φ voltages are all restored to 240V. These two balancing processes, occurring from 2.5s to 2.89s and from 4s to 4.34s respectively, also contain the three sections corresponding to the three operations in the control algorithm of the balancing system.

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The U2 profile during the simulation is shown in Fig. 5-6. The U2 of the 3Φ voltage is initially above 2%, and it is reduced to be close to zero at 1.6s attributable to the operation of the balancing system. At 2.5s and 4s, both the disconnection and the reconnection of EVs increase the U2 to be over 1%; whilst it is restored to be almost zero in a short time due to the balancing system. This indicates that the balancing system is able to maintain a low level of voltage imbalance by rapidly compensating for the voltage distortions induced by the 1Φ load variations.

248 Phase-A voltage Phase-B voltage 246 1.46s Phase-C voltage

244

242

240

238

Voltage RMS value (Volt) value RMS Voltage 236

234 1.3 1.6 2.89 4.34 232 1 1.5 2 2.5 3 3.5 4 4.5 Time (Sec)

Fig. 5-5 Three-phase voltage profiles after the balancing system

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1.3 Chapter 5Physical Test of ST based Voltage Balancing Method &Associated Simulation

0.025

0.02 2%

0.015 U2 0.01

0.005

0 1 1.5 2 2.5 3 3.5 4 4.5 Time (Sec) Fig. 5-6 U2 after the balancing system

The response time of the balancing system is influenced by the severity of the voltage imbalance and the sizes of the taps in each tap changer; both of which determine the required tap change number. A reduction in the size of each tap helps to reduce the U2 to a lower value, which improves the effectiveness of the balancing system but increases the response time. Because the tap sizes in the simulation study are much smaller than those used in the experimental facility, the U2 is reduced to a lower value in the simulation study.

Currently, most on-load tap changers are of a mechanical type and their operation relies on a motor drive unit [99]. However, mechanical tap changers have a low tap changing speed, requiring a long operating time; it takes about 100s for an on-load mechanical tap changer to move from tap 1 to tap 19 [117]. Voltage regulation of an LV feeder rich in flexible LCTs requires frequent tap changing operations. Mechanical tap changer would suffer significant wear to control the voltage fluctuations in LV feeders and requires repeated maintenance [118]. Accordingly, mechanical tap changers are not suitable to deal with the volatile voltage variations on the LV feeder. In contrary to mechanical tap changers, electronic tap changers have no moving parts and consequently have a higher performance, a lower maintenance cost and a longer lifetime [117]. In the LV application

Page | 127 Chapter 5Physical Test of ST based Voltage Balancing Method &Associated Simulation for voltage balancing and power quality improvement, a future “commercial” balancing system would have to be implemented with an electronic tap changer, probably based on IGBTs configured to switch the taps, provided the longevity and reliability of IGBT can be significantly improved and the problem of high cost of IGBT can be addressed.

5.6 Summary

This chapter presents the physical test of the ST based voltage balancing method as well as the associated simulation study. The transient response of the balancing system to the voltage disturbances caused by 1Φ load variations is investigated and the results demonstrate the feasibility of the proposed balancing method; they show that the ST based balancing system is capable of maintaining a balanced 3Φ voltage in the LV network by compensating for the voltage rises and sags rapidly. However, the balancing effectiveness and response time of the balancing system vary significantly with the tap size of the tap changers. An appropriate tap size should be determined for the balancing system by compromising the balancing efficacy and the response time.

The ST based voltage balancing system is supposed to have two applications in the future LV networks:- regulation of the steady-state voltage at a point along the LV feeder and improvement of the power quality at 3Φ load supply points. In the former application, the balancing system maintains a balanced 3Φ steady-state voltage with specified magnitude at the regulated point, ignoring the short-duration voltage sags and swells, which is similar to the conventional application of OLTC presented in section 2.1. However, since the LV network loads are becoming increasingly volatile, requiring more tap changing operations in a period, the conventional mechanical OLTC in this application would suffer from significant mechanical wear. In the latter application, the balancing system provides a balanced 3Φ supply by rapidly compensating for any voltage variations including voltage sags and swells. The OLTCs in this application are required to have a high operating speed, which is beyond the capability of conventional mechanical OLTCs. Therefore a future ‘commercial’ ST based balancing system could not employ mechanical OLTCs.

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The small-scale physical balancing system used in this physical test has a low power capacity. The tap changing operations are realized by using miniature relays, which have limited current capacity. To build a larger or ‘commercial’ balancing system based on the proposed method, an advanced electronic tap changing facility would be necessary. Therefore the site trial of the proposed balancing method in a real LV network is a part of the future work, and it can be carried out using electronic tap changers.

Page | 129 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages

CHAPTER 6

MONTE CARLO STUDY ON IMPACT OF EVS &

HPS ON LV FEEDER VOLTAGES

6.1 Introduction

In the UK’s decarbonisation journey, EVs and HPs are expected to be used to reduce carbon emissions in the transport and heat sectors. It is estimated that the number of EVs would increase to 12.8 million and the number of HPs would increase to 7.7 million in the UK by 2030 [8]. However, new electrical loads, resulting from the use of EVs and HPs, will challenge the voltage constraints of existing LV networks, which were designed according to the historical electricity demand data. The voltages received by the customers could drop below the lower statutory magnitude limit because of the increased voltage drop attributable to the rising electrical demand. Moreover, the volatility of LV network loads, further exacerbated by the penetration of EVs and HPs, as well as their 1Φ nature will worsen the problem of voltage imbalance on the LV feeder, especially in the downstream part of the feeder.

The impact of EVs and HPs on distribution networks, in terms of increased loading [119, 120], impact on efficiency [119, 121, 122] and deterioration of power quality [120, 121], have been separately investigated in studies using either deterministic or probabilistic

Page | 130 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages modelling approaches. In deterministic studies, only extreme scenarios are considered where the uncertainties are assigned with fixed values. However, probabilistic approaches create close-to-realistic scenarios and are capable of accurately capturing the diversity of electric loads in LV networks. Literatures on the impact of EVs and HPs on the electrical distribution system are reviewed in detail in section 6.2.

In this chapter, an MC simulation study was carried out to quantify and analyse the impact of both EVs and HPs on the LV feeder voltage magnitude and imbalance. A statistical model of EV charging demand is created in MATLAB based on the probability distributions of the uncertain parameters affecting EV charging. Two databases containing numerous residential demand and HP electric demand profiles respectively are imported into MATLAB to work as the profile generators. An MC simulation platform is established by combining the EV statistical model, the profiles generators of the residential demand and HP electrical demand, with a distribution network model established in OpenDSS. Based on this platform, the impacts of EVs and HPs are investigated in both balanced and non-balanced scenarios from the viewpoint of probability. Moreover, a deterministic study is performed based on the established MC simulation platform to straightforwardly elucidate the impact of EVs and HPs. In the following parts of this chapter, the literature on the impact of EVs and HPs, the MC simulation methodology, the details of the MC simulation platform and the simulation results are presented.

6.2 Impact of EVs and HPs on distribution networks

The study in evaluating the impact of EVs on power system can be traced back to 1978, when a pilot EV demonstration program was carried out in the USA [123]. Eight EVs of significantly different specifications were tested in a specified utility environment to evaluate the effects of different terrains (level, slight grade, and deep grade), traffic conditions (one, two, three, four stops in one mile and freeway) and payload. The benchmark data generated in this project provided a major input to establishing future industry roles in electric transportation research and development [123]. In 1993, the impact of EVs on the distribution network load profile was investigated in a simulation study presented in [124]. Based on the load data from the town of Blacksburg, USA, the

Page | 131 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages impact of different EV penetration levels and charging modes was studied. The results indicate an improvement of EV charging infrastructure is necessary to supply EVs if the penetration level is high.

More EV charging related characteristics, e.g. EV type, charging power level, charging start time and battery state of charge (SOC), were examined in recent studies to address the impact of EV charging in distribution networks in terms of increased load, impacts on efficiency, deterioration of power quality and loss of life of network assets. To deal with the uncertainty of the EV charging related characteristics, both deterministic and probabilistic approach were used in these studies.

In deterministic studies, the parameters of the power system are studied only in extreme scenarios where the uncertain parameters associated with EV charging are assigned with fixed values. In [6], a deterministic study was conducted to evaluate how the voltage profile on the UKGDN varies with EV penetration and aggregation levels. The result shows that the penetration of EVs at the LV segment level could cause violations of the voltage limits; whilst EVs connected at the micro-grid level do not breach the UK statutory voltage limits. A multi-period deterministic study based on the UKGDN was presented in [125], and it concludes that significant deployment of EVs could result in violation of the supply/demand matching capabilities and the statutory voltage limits, and power quality problems could be induced under certain conditions. The deterministic study presented in [119] based on a real data model shows that a high penetration level of EVs would result in loads exceeding the feeder current capacity. The study in [126] evaluated the impact of PHEVs on a typical distribution feeder in Blacksburg, USA, in which, a distribution transformer supplies five customers. The results show transformer overload only occurs when quick charging starts at 6:00 pm, coinciding with the evening peak load. The impact of EVs on an LV feeder in a suburban area in Dublin, Ireland was investigated in [127], and the result shows when 20-40% households own EVs, the test network reaches the limit of safe operation. The percentage depended on the connection points of the EVs along the LV feeder. Besides, the work in [122] evaluated the impact of different levels of PHEV penetration on distribution network investments and incremental energy losses based on two real larger-scale distribution networks. The obtained results show, depending on the charging strategies, investment costs can

Page | 132 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages increase up to 15% of the total actual distribution network investment costs, and energy losses can increase up to 40% in off-peak hours when 60% of total vehicles are PHEVs.

Compared with the deterministic approach, probabilistic approaches are capable of accurately capturing the diversity of EV charging related characteristics by using their probability distributions to create close-to-realistic situation. Numerous samples of the investigated power system parameters can be obtained in probabilistic studies, and show the impact of EVs from the viewpoint of probability. An MC simulation platform used for the investigation of the relationship between EV charging demand and the existing load peak was introduced in [128], and encountered the probability densities of the charging start time and the battery SOC. In [120], both deterministic and probabilistic approaches were used to evaluate the impact of EVs on British distribution networks in 2030. The probabilistic study considered the diversities of the residential load and the EV charging characteristics. The results indicate that network constraints, related to the LV network voltage and thermal loading, are likely to be violated when the penetration level is high. In addition, the LV cable losses increase to 10% for a high EV uptake level.

Both the deterministic and probabilistic studies show that the distribution networks could be significantly affected by high penetration levels of EVs and the associated uncoordinated EV charging. The impacts include the violation of network constraints, the decrease in power delivery efficiency and a reduction in the lifetime of the network assets.

The large-scale impact of solar heating options, including solar-assisted HPs, were studied in [129] and the results indicate that the deployment of solar assisted HPs can either increase or decrease the load factor depending on whether they displace electric- based or non-electric based heating. The electrical load characteristics of domestic HPs were investigated in [130], this indicated the daily demand profile of a HP comprises a base load proportional to the difference between room and external ambient temperature and a peaky pattern related to the need for domestic hot water. In [131], the impact of ground sourced HPs on annual, monthly and hour of day electricity consumption was investigated by analysing the results from a questionnaire survey. A simulation study, where an LV network with sequential increases in the penetration of HPs, was analysed in terms of transformer loading, steady-state voltage drop and transient voltages related

Page | 133 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages to the starting of HPs was presented in [132]. The results show if 20% customers install HPs, the network under investigation is not overloaded, but in many cases the voltage drops exceed the statutory limits. In [133], a series of simulation studies were carried out on typical UK urban and rural networks to assess the impacts of distributed combined heat and power (CHP) and HP on LV networks. The results indicated a mix of distributed CHP and HP could mitigate the negative impacts arising from the HP. In addition, a deterministic simulation study was carried out in [134] to investigate the impact of a large penetration of EVs and HPs on the remote LV feeder voltage magnitude. The simulation results show a 10% penetration of both EVs and HPs across the whole network causes the voltage at the remote end of the longest LV feeder to drop below the statutory lower limit.

According to the literature survey listed above, an exhaustive investigation into the impacts of both EVs and HPs on the LV feeder voltage magnitude and imbalance is still lacking. Consequently, in this chapter, an MC study was used to investigate the impact of both EVs and HPs on the LV feeder voltages.

6.3 Simulation methodology

The MC method is defined in the broad sense as a technique of model resolution using repeatedly random sampling to obtain numerical results; i.e. the simulation is executed many times to obtain efficient samples that reflect the probability distribution of the real results [135, 136]. This method is commonly used to deal with power system calculation problems involving uncertain parameters, for example, the reliability evaluation of a power network [136]. In the study presented in this chapter, the impact of EVs and HPs on the LV feeder voltages is affected by numerous uncertain factors, including the versatile HP electrical demand, the diversified residential demand and the EV charging related uncertainties, which include the EV type, battery capacity, charging start time and charging power level. Therefore, the MC method is used in this study.

When investigating the impact of EVs and HPs on the LV feeder voltages, it is crucial to determine the electrical demand of EVs and HPs, as well as the residential electrical demand. However, these are generally diversified and affected by uncertainties. In this

Page | 134 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages study, a statistical model of EV charging demand is established. In this model, all the uncertain parameters affecting EV charging are expressed in a probabilistic way. Consequently, this model is able to generate a random EV charging demand profile by randomly defining the uncertainties according to their presumed probability distributions. Moreover, a profile generator of the HP electrical demand is created based on a thermal demand database. Hence this generator is able to provide random HP electric demand profiles during the MC simulation. In addition, a residential demand profile generator is also established based on a database containing 10,000 residential demand profiles.

Fig. 6-1 shows the MC simulation framework employed in this study. The random electrical demand profiles, generated by the statistical model of EV demand and the profile generators of HP electrical demand and residential demand, are assembled in MATLAB, and the net demand profiles are imported into the distribution network model established in OpenDSS. The sequential power flow calculation is then executed in OpenDSS, and the results are returned to and analysed in MATLAB. In this MC study, the power flow calculation is accomplished in OpenDSS; whilst the generation of random demand profiles and the analysis of the simulation results are performed in MATLAB.

Power flow calculation Net demand profiles results

Statistical model of EV demand

Profile generator of HP electrical demand Distribution Analysis of

3 network simulation 2.5 2 MATLAB 1.5 model in results in 1 0.5 0 OpenDSS MATLAB

Profile generator of residential demand

3 2.5 2 1.5 1 0.5 0

Fig. 6-1 MC simulation framework

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The above procedure is repeatedly carried out and terminates when the convergence criteria is satisfied, i.e. the stopping rule of the MC method is an acceptable convergence for the sample means of the investigated parameters. In this study, the number of sequential power flow calculation is chosen as (N = 500); and consequently, the sample means of the investigated parameters all have a high degree of confidence in the specified confidence interval, as explained in Appendix E.

6.4 Statistical model of EV demand

EVs achieve zero carbon emission in transportation by consuming electricity, which can be generated by zero or low carbon energy source. However, the progressive replacement of conventional vehicles with EVs will introduce additional loads on existing distribution networks. The electrical energy drawn by an EV from a power outlet during the charging is initially stored as chemical energy in the EV battery, and then is converted back to electric energy and used to power the electric machines that operate the vehicle during the driving process. Hence, the electrical demand of an EV is essentially the electrical demand of its battery.

Nowadays, special attention has been given to the lithium-ion battery. It is reported by Frost & Sullivan that the use of lithium-ion battery for EVs is expected to experience significant growth from 2015 to 2020 [137]. Compared with other battery technologies, lithium-ion offers a combination of improved performance: greater energy-to-weight ratio, no memory effect and low self-discharge when not in use [138]. In this study, it is assumed that all the EVs employ lithium-ion batteries and the EV charging demand profiles are developed based on the generic charging profile of a lithium-ion battery as shown in Fig. 6-2 [139].

r e w

o P p

g n i g r a h C

T t1 t 2 Time

Fig. 6-2 Generic charging profile of lithium-ion EV battery

Page | 136 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages

As show in Fig. 6-2, the typical EV charging profile has a trapezoidal shape, consisting a rectangle and a triangle; the rectangular section refers to the normal charging section, where there is a constant charging power level; the triangular section refers to the slow charging section, where the charging power level decreases as the battery SOC increases. The charging profile in Fig. 6-2 is determined by the following factors:  Height of the trapezoid: charging power level P

 Lower base of the trapezoid: full charging time t2 determined by the energy required for EV charging Q

 Width of the rectangle: constant power charging time t1 determined by the energy required for EV charging Q  Position of the trapezoid on time axis: charging start time T

All the above parameters need to be identified to create the charging profile of an EV. Fig. 6-3 shows the framework for the creation of EV charging profiles employed in this study, where the uncertainties are defined by using the associated probability distributions, and they are introduced in detail below.

EV generic charging profile

80% Normal charging: 3KW Charging power level (P) 20% Fast charging: 7KW Start charging immediately Charging start time (T) EV after arriving home charging profiles EV battery capacity (Qc) Energy required for EV charging (Q) Energy consumed Daily travel distance during travelling (Qe)

Fig. 6-3 Framework for the determination of EV charging profiles

EV charging power 퐏 In the British Standard on EV charging system (BS EN 61851-1: 2011), four EV charging modes are introduced as shown in Table 6-1 [140]. All the modes, except mode 1, require communication between the vehicle and the charging supply [141]. Due to the lack of residual-current device protection, mode 1 is considered unsafe and will not be

Page | 137 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages allowed in the UK [142, 143]. In addition, since the UK residential properties generally have a 1Φ electric supply, allowing a peak load of around 12 kW (52A based on 230V) [17], the mode 2 in Table 6-1 is likely to be used for domestic EV charging in the UK. EVs are rarely charged at the maximum current at home because customers have adequate time throughout the night. Therefore, it is assumed that 80% of EVs are charged at 3kW (13A) and 20% at 7kW (30A).

Table 6-1 Details of four introduced EV charging modes [141] Charge Mode Charging accessories Max. Current/voltage Mode 1 Standardized socket-outlets 16A AC 250/480V

Mode 2 Standardized socket-outlets with 32A AC 250/480V in-cable control and protective

Mode 3 Dedicated EV supply equipment 63A AC 250/480V with in-cable control devices Mode 4 Off-board chargers with in-cable 400A DC 1000V control

Charging start time 퐓 The charging start time, i.e. the time when the EV is plugged for charging, is closely related to the electricity tariff utilised by the customer. Currently in the UK, the majority of the domestic customers are purchasing electricity under the unrestricted (flat-rate) tariff and a small number of customers use the Economy 7 tariff. For simplicity, only the unrestricted electricity tariff is considered in this study. Since the customers, using the unrestricted electricity tariff, have a single-rate electricity charge rate, it is assumed that they would charge their EVs immediately when they arrive home.

Fig. 6-4 shows the traffic distribution in a typical workday in the UK and each bar indicates the quantity of car driver trips in process in one hour [144]. There are two peaks in the traffic distribution profile: one occurs between 8:00 and 9:00, referring the peak of home to work journeys; the other occurs between 17:00 and 18:00, referring the peak of work to home journeys. Based on this data, the most likely time arriving home is assumed to be between 18:00 and 19:00 in this study. The distribution of the EV charging start time is modelled by a normal distribution, with the mean value of 18:30

Page | 138 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages and the standard deviation of one hour, i.e. T ~ N(18: 30, 1 hour), and the probability distribution curve is shown in Fig. 6-5.

Car driver trips per hour 300 250 200 150 100 50 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 The number of the hours in a day

Fig. 6-4 Traffic distribution in a typical workday

The probability distribution of the EV charging start time 0.5

0.4

0.3

0.2

0.1

0 15:00 16:00 17:00 18:00 19:00 20:00 21:00 22:00 time

Fig. 6-5 Probability distribution of the EV charging start time

EV battery capacity 퐐퐜 There are mainly two kinds of EV: PHEV and BEV. These two types of EVs have different characteristics in many aspects including EV battery capacity. Some currently available BEVs and PHEVs as well as their characteristics are summarized in Appendix F.

Based on the BEV battery capacities indicated in Appendix F.1, it is assumed that the battery capacity of BEVs in this study complies with the normal distribution N (22, 4),

Page | 139 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages where the mean value is 22kWh and the standard deviation is 2. Based on the PHEV battery capacities indicated in Appendix F.2, it is assumed that the battery capacity of PHEVs in this study complies with the normal distribution N (16, 4), where the mean value is 16kWh and the standard deviation is 2. Fig. 6-6 and Fig. 6-7 show the probability distribution curves for BEV and PHEV battery capacities, respectively.

Probability distribution of BEV battery capacity 0.25

0.2

0.15

0.1

0.05

0 16 18 20 22 24 26 28 battery capacity (kWh) Fig. 6-6 Probability distribution of BEV battery capacity

Probability distribution of PHEV battery capacity 0.25

0.2

0.15

0.1

0.05

0 10 12 14 16 18 20 22 battery capacity (kWh) Fig. 6-7 Probability distribution of PHEV battery capacity

Energy required for charging 퐐 The energy required for charging of an EV is determined by the energy consumed during travelling Qe, which can be obtained from the travel distance by using (6-1), where L refers to the travel distance and Ce refers to the energy coefficient, defined as the consumed electric energy of EV when travelling for one unit of distance.

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Qe = L ∙ Ce (6-1)

In this study, all the BEVs are assumed to have the Ce of 5.26 miles/kWh, computed by averaging the Ce s of all the lithium-ion battery based BEVs listed in Appendix F.1. Likewise, 3.52 miles/kWh, obtained based on the PHEVs in Appendix F.2, is used as the

Ce of all the PHEVs.

Moreover, the daily travel distance of an EV is modelled by using a beta distribution with the mean value of 18.7 miles and the maximum value of 100 miles, as shown in Fig. 6-8 [145].

The probability distribution of daily travel distance 0.04

0.03

0.02

0.01

0 0 20 40 60 80 100 Daily travel distance (mile) Fig. 6-8 Probability distribution of daily travel distance

In the MC procedure, a randomly generated L is assigned to each EV and the consumed energy Qe can be obtained by use of (6-1). Both PHEVs and BEVs have a minimum

SOC level of 20% to avoid the loss of life of batteries [138]. If the obtained Qe of an EV exceeds 80% of its battery capacity, this EV would require a full charging, i.e. the consumed energy of this EV is 80% of the battery capacity and the initial battery SOC is 20%. Therefore, the energy required for the charging Q can be calculated based on the

EV battery capacity Qc and the energy consumed Qe by using (6-2), where fe refers to the EV charger efficiency, and is assumed to be 90%.

푄푐 ∙ 80% ⁄ 푤ℎ푒푛 푄푒 > 푄푐 ∗ 80% 푓푒 푄 = (6-2) 푄푒 ⁄ 푤ℎ푒푛 푄푒 < 푄푐 ∗ 80% 푓푒 {

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Determination of 퐭ퟏ and 퐭ퟐ The area under the charging profile curve equals the energy required for charging. For the generic charging profile in Fig. 6-2, values of 푡1 and 푡2 vary with the type of the EV charging mode [139]. In this study, the ratio t1/t2 is assumed to equal 0.8 for a full charging. For the full charging shown in Fig. 6-2, the area of the trapezoid can be obtained by (6-3).

P ∙ t1 + 0.5 ∙ (t2 − t1) ∙ P = 0.8 ∙ Qc/fe (6-3)

By substituting (t1⁄t2 = 0.8) into (6-3), the values of t1 and t2 can be calculated as expressed by (6-4) and (6-5), from which, the base of the triangle △ t in Fig. 6-2 can be obtained as expressed by (6-6).

32푄푐 푡1 = (6-4) 45푃푓푒

8푄푐 푡2 = (6-5) 9푃푓푒

8푄 △ 푡 = 푐 (6-6) 45푃푓푒

The area of the triangle in Fig. 6-2, referring the energy required during the slow charging section, can be obtained as expressed by (6-7).

4 S = Q /f (6-7) T 45 c e

When the initial SOC of an EV battery is higher than 20%, the charging profile would be part of the full charging profile. If the energy required for charging is larger than ST, the EV charging profile would include both the normal charging section (the rectangular section in Fig. 6-2) and the slow charging section (the triangular section in Fig. 6-2) as shown in Fig. 6-9, where P is the charging power level, T is the charging start time and the red area refers to the section omitted from the full charge profile due to higher initial battery SOC than 20%.

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P t

Time axis T T  t1 T  t2

Fig. 6-9 Charging profile when Qc is larger than ST

The values of t1 and t2 in this case can be determined by (6-8) and (6-9) respectively.

푄−푆 푡 = 푇 (6-8) 1 푃

푡2 = 푡1 +△ 푡 (6-9)

Moreover, if the energy required for charging is smaller than ST, the EV charging profile would only include the slow charging section as shown in Fig. 6-10, where the red area refers to the section omitted from the full charge profile due to higher initial battery SOC.

The value of t2 in this case can be derived by use of (6-10).

2 푡 = √푄 ∙ 푆 (6-10) 2 푃 푇

P t

Time axis T T  t2

Fig. 6-10 Charging profile when Qc is less than ST

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Three EV charging profiles generated by this statistical model are shown in fig. 6-11.

8 6 4 2 0 15:00 16:00 17:00 18:00 19:00 20:00 21:00 22:00 23:00 Time Chargingpower (kW) Fig. 6-11 Three EV charging profiles

6.5 Monte Carlo simulation platform

The profile generators of residential demand and HP electrical demand are created in MATLAB [146], and together with the statistical EV demand model are combined with a distribution network model established in OpenDSS [147]. This is used to create an MC simulation platform.

6.5.1 Adopted distribution network

The distribution network used in Chapter 4, as illustrated in Fig. 4-5, is adopted in this study and is modelled in OpenDSS. The feeder F1 is used for the investigation of the impact of EVs and HPs on LV feeder voltages, so only the feeder F1 is modelled in detail and all the other feeders are simplified and modelled as lumped loads. In this study, 96 customers are supplied by the feeder F1, instead of the 60 customers used in Chapter 4. This is designed to reflect a realistic LV network as used in the UK. The two LCPs adjacent to the LV transformer each supply four customers per phase, while the other LCPs supply three customers per phase. During the MC simulation, the demand profiles, generated by the statistical model of EV demand and the profile generators of HP electric demand and residential demand, are assigned to the customers in Feeder F1.

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6.5.2 Residential demand profile generator

The power quality parameters shall be precisely measured as indicated in the British Standard [33]. For example, the RMS value of the voltage magnitude shall be measured over a 10-cycle time interval for a 50 Hz power system or a 12-cycle time interval for a 60 Hz power system. Therefore, it is preferable to use the demand profiles with high resolution. The currently accessible residential demand profile with the highest resolution is the one-minute resolution residential demand profile generated by the domestic electricity demand model in [148] and this is used in this study. The model in [148] is able to randomly generate a one-minute resolution demand profile according to the specified number of residents and the month of a year. A pool of 10,000 demand profiles is created by using this model for a February weekday with the number of residents per household defined according to [149]; i.e. households with one, two, three, four, and > five persons in the UK account for 30.58%, 34.1%, 15.57%, 12.88% and 6.86%, respectively [150]. In each MC simulation cycle, 96 residential demand profiles are randomly selected from the demand profile pool, and then imported to each customer in the investigated LV feeder. Fig. 6-12 shows three random profiles generated by this profile generator.

12 10 Profile 1 Profile 2 8 Profile 3 6

Power(kW) 4 2

0 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-12 Three random demand profiles

6.5.3 HP electrical demand profile generator

Similar to the residential demand profile generator, the HP electrical demand profile generator is a collection of numerous HP electrical demand profiles, which are converted

Page | 145 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages from the domestic thermal demand profiles provided in [151]. The database in [151] contains a large amount of one-hour resolution thermal demand profiles classified by occupancy types, seasons of the year, build periods and types of the houses. The thermal demand profiles categorized as winter period and the house types of detached, semi- detached and terraced are collected to create a pool of 840 thermal demand profiles. In this profile pool, the proportions of the above three house types are in the ratio of 1:2:2, complying with the current UK house stock distribution by type [152], while the houses specified by the other characteristics have equal proportions. Moreover, the COPs of all the HPs in this study are assumed to be 3 for simplicity. In order to sufficiently capture the characteristics of the residential demand and the EV demand, the granularity of the profiles is increased to one-minute by using the linear interpolation. In this way, a collection of 840 HP electrical demand profiles with the resolution of one-minute is obtained. In the MC simulation, a certain number of HP electrical demand profiles are randomly selected from the profile pool according to the HP penetration level and then imported to the customers in the investigated LV feeder. Fig. 6-13 illustrates three HP electrical demand profiles chosen from the profile pool; profile 1 indicates the occupancy type where the residents have part-time jobs in the morning; profile 2 refers to the occupancy type where the residents have full-time jobs and profile 3 refers to the occupancy type where there is one or more pensioners, disabled people or unemployed [151].

3 Profile 1 Profile 2 2 Profile 3

Power demand (kW) demand Power 0 0:00 3:00 6:00 9:00 12:00 15:00 18:00 21:00 Time

Fig. 6-13 Three HP electric demand profiles

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6.5.4 MC simulation algorithm

Fig. 6-14 shows the detailed MC simulation algorithm adopted in this study. A pool of 10,000 one-minute daily residential demand profiles and a pool of 840HP electrical demand profiles are firstly loaded, from which, each customer on the investigated LV feeder is assigned with a residential demand profile. According to the specified penetration levels of EVs and HPs and the investigated scenario, the owners of the EVs and HPs are selected; each owner of the HP is assigned with an HP demand profile, and moreover, the EV model generates a random EV charging profile for each owner of the EV. Following this, the final demand profiles are imported into the LV network model and the sequential power flow calculation for a one-day period is repeatedly executed until the convergence criteria is satisfied. The simulation results are numerous one- minute profiles of the voltages on the LV feeder as well as the associated U0 and U2. Because the limit of the voltage imbalance stated in [3] is based on 10-minute mean values, the granularity of the obtained profiles is reduced to 10-minute by calculating and collecting the 10-minute mean values.

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Load the residential demand profile pool (1440 one-minute time-step)

Load the HP demand profile pool (1440 one-minute time-step)

MC trial number N = 1

Randomly assign one residential profile to each customer

Allocate the EVs and HPs according to the EV and HP number and the studied scenario

Customers with HP NO YES Randomly assign one HP demand profile to each owner of an HP

Customers with EV NO YES Randomly define all the uncertainties

Generate the EV charging profile

Create the final profiles

Sequentially run power flow calculation for 1440 one-minute time step

Record results

N = 500 NO N = N + 1 YES End

Fig. 6-14 Detailed MC simulation algorithm

6.6 Impact of EVs and HPs

Based on the established MC simulation platform, a simulation study was carried out following the MC procedure shown in Fig. 6-14 and used to investigate the impact of EVs and HPs on the LV feeder voltages. The penetrations of EVs and HPs in the UK are estimated to reach 8.5% and 6.5% (based on the UK household number) in 2022, 21.9%

Page | 148 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages and 13.3% in 2026, and 41.3% and 24.8% in 2030 [8]. Since higher concentration of EVs and HPs could occur in some local networks, a few more presumptive penetration levels along with the estimated penetration levels as shown in Table 6-2 are investigated in this study. Moreover, both balanced and non-balanced scenarios are studied. In the balanced scenario, EVs and HPs are randomly located to all the customers following the uniform distribution, whilst in the non-balanced scenario, the EVs and HPs are sequentially located to customers in phase-A starting from the customers at the most remote end (assuming all the customers can own one EV and one HP at most) and then to customers in phase-B in the same manner as indicated in Table 6-2.

Table 6-2 Number of EVs and HPs in 96-customer LV feeder

Case EV HP Non-balanced scenario number1 8number (8.5%) 6 (6.5%) Phase-A: 8EVs, 6HPs 2 21 (21.9%) 13 (13.3%) Phase-A: 21EVs, 13HPs 3 40 (41.3%) 24 (24.8%) Phase-A: 32EVs, 24HPs; Phase-B: 8EVs 4 45 27 Phase-A: 32EVs, 27HPs; Phase-B: 13EVs 5 45 36 Phase-A: 32EVs, 32HPs; Phase-B: 13EVs, 4HPs 6 45 45 Phase-A: 32EVs, 32HPs; Phase-B: 13EVs, 13HPs 7 54 45 Phase-A: 32EVs, 32HPs; Phase-B: 22EVs, 13HPs 8 54 54 Phase-A: 32EVs, 32HPs; Phase-B: 22EVs, 22HPs

Only the voltage magnitude, U2 and U0 at the end of the LV feeder are analyzed because these are the extreme values on the whole LV feeder. This MC simulation study delivers numerous samples of the investigated parameters. To show the uncertainty of the results, both the mean and the corresponding 95th percentile profiles of the investigated parameters are analyzed.

Besides, to illustrate the impact of EVs and HPs in a straightforward way, a deterministic simulation study is carried out based on the MC simulation platform with predefined electric demand profiles. The case 5 in Table 6-2 is used as the background scenario; in the deterministic study, 96 residential demand profiles, 36 thermal demand profiles and 45 EV charging profiles are predetermined and imported into the investigated LV feeder. Following this, the one-day period sequential power flow is executed in both balanced

Page | 149 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages and non-balanced scenarios. Moreover, the case with no EVs and HPs connected to the LV feeder is also studied as a reference for comparison. The 3Φ voltages at the end of the LV feeder as well as the associated U2 and U0 are recorded during the simulation.

6.6.1 Deterministic study

Fig. 6-15 shows the RMS value profiles of the 3Φ voltages at the end of the LV feeder when no EVs or HPs are connected. The voltage magnitude varies within a small range around 245V – the minimum 239.1V is reached on phase-B at 17:20 and the maximum 247.7V is reached on phase-C at 8:10 – they are all within the statutory limits and above 230V. In addition, the 3Φ voltages remain almost constant from 24:00 to 6:00, indicating that the electrical load is seldom affected by human behaviour during the night.

260 253V 250 247.7V 8:10

240 239.1V 17:20 V 230 230V a Voltage(V) V b 220 V 217V c

210 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-15 Voltage profiles at LV feeder end with no connection of EVs and HPs

Fig. 6-16 and Fig. 6-17 shows the RMS value profiles of the 3Φ voltages at the end of the LV feeder in the balanced and non-balanced scenarios respectively. In the balanced scenario, the voltages are lowered around 19:00 when the EVs are most likely to be charged, compared to those in Fig. 6-15. The phase-C voltage reaches the minimum value 228V at 19:40; so no violation of the voltage magnitude limits occurs when the EVs and HPs are evenly distributed along the LV feeder. Moreover, the difference among the three voltages from 24:00 to 6:00 becomes relatively apparent, showing that

Page | 150 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages the adoption of EVs and HPs would increase the impact of human behaviour on the electrical demand during the night.

In the non-balanced scenario, as shown in Fig. 6-17, the phase-A voltage is lowered throughout the day, especially around 19:00, reaching the minimum 215V at 19:20; the phase-B voltage is lowered around 19:00, whilst the phase-C voltage is raised around 19:00, compared to the profiles in Fig. 6-15. The minimum occurs on phase-A because most EVs and HPs are connected to phase-A. Although the HPs would cause additional load throughout the day, only three HPs are connected to phase-B, having negligible effect on the phase-B voltage, and moreover, the 13 EVs would increase the voltage drop on phase-B when they are charged. Neither EVs nor HPs are connected on phase-C, but the phase-C voltage is raised around 17:00, reaching the peak 252.1 V at 17:00, because the EVs and HPs on the other two phases alter the neutral voltage.

260 253V 250

240

V 230 230V a Voltage(V) V 228V 19:40 b 220 V 217V c

210 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-16 Voltage profiles at LV feeder end in balanced scenario

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260 253V 252.1V 19:00 250

240

230 230V

Voltage(V) V a 220 V 217V b V 215V 19:20 c 210 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-17 Voltage profiles at LV feeder end in non-balanced scenario

The U2 of the 3Φ voltages at the end of the LV feeder is lifted by the connection of EVs and HPs as shown in Fig. 6-18, which illustrates the U2 profiles in the three investigated scenarios. The increase of U2 is relatively significant around 19:00, attributed to the clustered charging of EVs. In the balanced scenario, the U2 reaches the peak 0.74% at 21:30; whilst in the non-balanced scenario; the U2 reaches its peak 2.5% at 19:20, violating the voltage imbalance limit. Fig. 6-19 shows the U0 of the three voltages behaves in the same manner as the U2. The U0 rises to its peak 2.84% at 19:40 in the balanced scenario, and however, its peak in the non-balanced scenario is 6.9% occurring at 19:20.

2.5% 2.5% 19:20 without EVs&HPs with EVs&HPs evenly distributed 2% with EVs&HPs non-evenly distributed

1.5% U2 1% 0.74% 21:30

0.5%

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-18 U2 profile at LV feeder end

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8% 6.9% 19:20 without EVs&HPs with EVs&HPs evenly distributed 6% with EVs&HPs non-evenly distributed

4% U0 2.84% 19:40

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-19 U0 profile at LV feeder end

6.6.2 Monte Carlo study

The results of the MC simulation are vast samples of the investigated parameters, including the 3Φ voltages at the end of the LV feeder as well as their associated U0s and U2s. Their mean and 95th percentile profiles are analysed to illustrate the impact of EVs and HPs from a viewpoint of probability.

Balanced scenario The mean and 95th % profiles of the phase-A voltage at the end of the LV feeder in the balanced scenario are illustrated in Fig. 6-20 and Fig. 6-21 respectively, from which, the valley values of the profiles are obtained and shown in Fig. 6-22. In Fig. 6-20 and Fig. 6- 21, the blue-line profiles indicated as ‘reference’ refer to the case where no EVs and HPs are connected to the LV feeder. The phase-A voltage profiles in different cases are close to each other throughout the day, except for around 19:00, when they diverge from each other, caused by the clustered charging of the EVs.

The value of the phase-A voltage profile at the valley decreases as the penetration level of EVs and HPs increases from case 1 to case 8, because the increasing number of EVs and HPs would induce a larger voltage drop along the LV feeder. The valley of the mean profile is always above 230V for all the cases, but the valley of the 95th % profile drops below 230V in case 4 and reaches 227.4V in case 8, i.e. above (230V-6%) or 217V.

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Accordingly, there is no possibility of violating the voltage magnitude limits in the balanced scenario. Moreover, since the EVs and HPs are evenly distributed among the three phases, the phase-B and phase-C voltages are supposed to have the same characteristics as the phase-A voltage. Therefore, the penetration of EVs and HPs in the balanced scenario increases the voltage drop along the LV feeder, but doesn’t cause violation of the voltage magnitude limits.

250

245 reference 240 case 1 case 2 235 case 3

Voltage(V) case 4 230V case 5 230 case 6 case 7 225 12:00 15:00 18:00 21:00 24:00 3:00case 8 6:00 9:00 Time Fig.217V 6-20 Mean profiles of phase-A voltage in balanced scenario

250

245 reference 240 case 1 case 2 235 case 3 Voltage(V) case 4 230V 230 case 5 case 6 case 7 225 12:00 15:00 18:00 21:00 24:00 3:00case 8 6:00 9:00 Time Fig.217V 6-21 95th % profiles of phase-A voltage in balanced scenario

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mean valley 95th % valley

245

240 235 230 Voltage (Volt) Voltage 225 1 2 3 4 5 6 7 8 Case No.

Fig. 6-22 Valley values of phase-A voltage mean and 95th % profiles

The mean and 95th % profiles of the U2 at the end of the LV feeder in the balanced scenario are illustrated in Fig. 6-23 and Fig. 6-24 respectively, where the blue-line profiles indicated as ‘reference’ refer to the case where there is no EV and HP connected to the LV feeder. Fig. 6-25 shows the peak values of the U2 mean and 95th % profile for all the cases. The peak value of the U2 profile increases as the penetration level of EVs and HPs increases from case 1 to case 8, owning to the rising diversity of the electric demand imposed by EVs and HPs. However it is always below 2% without violating the voltage imbalance limit; the peak of the mean profile reaches 0.57% and that of the 95th % profile reaches 1.125% in case 8. Accordingly, although the U2 is increased by the penetration of EVs and HPs, there is no possibility of violating the voltage imbalance limit in the balanced scenario.

1.2% reference 1% case 1 case 2 0.8% case 3 case 4 U2 0.6% case 5 case 6 0.4% case 7 0.2% case 8

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-23 Mean profiles of U2 in balanced scenario

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1.2% reference case 1 1% case 2 case 3 0.8% case 4 case 5 U2 0.6% case 6 0.4% case 7 case 8 0.2%

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00

Fig. 6-24 95th % profiles of U2 in balanced scenario

mean peak 95th % peak 1.2% 1.0% 0.8% 0.6% U2 0.4% 0.2% 0.0% 1 2 3 4 5 6 7 8 Case No.

Fig. 6-25 Peak values of U2 mean and 95th % profiles

Fig. 6-26 and Fig. 6-27 illustrate the mean and 95th % profiles of the U0 at the end of the LV feeder respectively, and Fig. 6-28 shows the peak values of the U0 profiles. The U0 behaves in the same manner as the U2; the peak of the U0 mean profile increases to 1.8% in case 8; whilst the peak of the U0 95th % profile increases to 3.52% in case 8.

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4% reference case 1 3% case 2 case 3 case 4 2% U0 case 5 case 6 case 7 1% case 8

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-26 Mean profiles of U0 in balanced scenario

4% reference case 1 case 2 3% case 3 case 4 case 5 2% U0 case 6 case 7 1% case 8

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 6-27 95th % profiles of U0 in balanced scenario

mean peak 95th % peak 4%

2% U0 1%

0% 1 2 3 4 5 6 7 8 Case No. Fig. 6-28 Peak values of U0 mean and 95th % profiles

Page | 157 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages

Non-balanced scenario In the non-balanced scenario, the phase-A conductor supplies most of the EVs and HPs. Consequently, since this is the most vulnerable phase, only the phase-A voltage at the end of the LV feeder is analysed. Fig. 6-29 shows the values of the phase-A voltage mean and 95th % profiles at the valley in the non-balanced scenario. The results drop significantly from case 1 to case 3, and then change only slightly from case 3 to case 8. This is because in the non-balanced scenario, for case 4 onwards, the extra EVs and HPs are connected to the LCPs on the phase-A conductor close to the LV transformer or on the phase-B conductor, which has minimal effect on the phase-A voltage. The valley of the mean profile is above 230V only in case 1, but is above 217V in all the cases; whilst the valley of the 95th % profile is above 217V only in case 1. Therefore, the phase-A voltage is likely to drop below 230V and the voltage magnitude limits could be violated in the non-balanced scenario, especially when the LV feeder supplies more than 21EVs and 13HPs (estimated penetration level in 2026).

mean valley 95th % valley

235

225 215

205 Voltage (Volt) Voltage 195 1 2 3 4 5 6 7 8 Case No.

Fig. 6-29 Valley values of phase-A voltage in non-balanced scenario

The peak values of the U2 and U0 mean and 95th % profiles in the non-balanced scenario are shown in Fig. 6-30 and Fig. 6-31 respectively. The peak of the U2 profile increases apparently from case 1 to case 3, and changes little after case 3. The peak of the 95% profile is below 2% only in case 1 and reaches 3.01% in case 8, indicating there is a possibility of breaching the statutory 2% voltage imbalance limit in the non-balanced scenario when at least 21EVs and 13HPs are connected to the LV feeder. The peak of the mean profile rises above 2% in case 3 and reaches 2.42% in case 8, which manifests that the voltage imbalance limit is very likely to be violated when at least 40EVs and24HPs

Page | 158 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages

(estimated penetration level in 2030) are connected to the LV feeder in a non-balanced way.

mean peak 95th % peak 4% 3%

2% U2 1% 0% 1 2 3 4 5 6 7 8 Case No.

Fig. 6-30 Peak values of U2 in non-balanced scenario

Moreover, the peak of the U0 profiles, as shown in Fig. 6-31 increases significantly from case 1 to case 2, and then stays almost constant from case 2. The peak of the mean profile varies around 6.3% and the peak of the 95th % profile varies around 8.2% from case 2 to case 8, much higher than the equivalent in the balanced scenario in Fig. 6-28.

mean peak 95th % 10%

6% U0 4% 2% 0% 1 2 3 4 5 6 7 8

Case No. Fig. 6-31 Peak values of U0 in non-balanced scenario

Table 6-3 shows the time periods when the 95th % profiles of Va and U2 violate the associated limits in the non-balanced scenario. All the violations occur between 18:00 and 20:40, when the EVs are most likely to be charged.

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Table 6-3 Periods of violations in non-balanced scenario

Time periods Case No. Va below 217V U2 over 2% 1 None None 2 18:40-19:50 18:20-20:00 3 18:20-20:10 18:10-20:30 4 18:20-20:20 18:10-20:30 5 18:20-20:20 18:10-20:30 6 18:20-20:20 18:00-20:40 7 18:20-20:20 18:00-20:40 8 18:20-20:20 18:00-20:40

6.7 Summary

A literature study was undertaken to investigate the impacts of either EVs or HPs on distribution networks in terms of increased loading, impact on efficiency and deterioration of power quality. This chapter presents an MC study on the impacts of both EVs and HPs on the LV feeder voltage magnitude and imbalance. To complete this study, an MC simulation platform was established by combining a statistical model of the EV charging demand, the profile generators of the HP electrical demand and associated residential demand, and a distribution network model based on the UKGDN.

The MC study delivers numerous samples of the investigated parameters, which are analysed from the viewpoint of probability. The simulation results indicate that the impact of EVs and HPs on the LV feeder voltages is related to the distribution of EVs and HPs on the LV feeder. In the balanced scenario, the penetration of EVs and HPs increases the voltage drop and deteriorates the voltage imbalance along the LV feeder. However, the voltage imbalance limit is not breached and the voltage magnitude is always above 217V, regardless of the penetration level of EVs and HPs. In the non- balanced scenario, when 21EVs and 13HPs or more are connected the LV feeder, both the voltage magnitude and imbalance limits are possible to be violated, and moreover, the voltage imbalance is very likely to be violated when the penetration reaches 40EVs and 24HPs or more. In addition, the violations of the voltage limits all occur between 18:00 and 20:40, when the EVs are most likely to be charged, implying that the

Page | 160 Chapter 6 Monte Carlo Study on Impact of EVs & HPs on LV Feeder Voltages violations could be mitigated by spreading the EV charging throughout the night and avoiding the coincidence with the residential load peak.

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CHAPTER 7

MITIGATION OF IMPACT OF EVS & HPS

ON LV FEEDER VOLTAGES

7.1 Introduction

The MC study in Chapter 6 discussed the level of penetration of EVs and HPs that cause unacceptable voltage imbalance and magnitude variations on UK LV feeders. Statutory voltage limits are likely to be breached if the number of EVs connected to a 96-customer LV feeder is greater than 20 and the number of HPs is greater than 12, and assuming the connections are made in accordance with a non-balanced scenario, where the first EVs and first HPs are connected to the phase-A and the remainder are connected to the phase- B. In this chapter, the ST based voltage balancing method and the DR based TOU tariff are used, either individually or together, to mitigate the adverse impact of EVs and HPs on the LV feeder voltages.

Traditionally there have been limited control, monitoring and communication facilities on distribution networks with none applied on the LV networks; therefore, it is difficult or impossible to mitigate the severe voltage imbalance and magnitude variations without adopting the necessary control, mitigation and network reinforcement measures.

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Enhancement of the network infrastructure is conventionally used to deal with severe voltage imbalance on LV feeders, but this generally involves a high financial cost. According to the literature reviewed in Chapter 2, several new methods for mitigating the voltage imbalance on LV feeder have been proposed. However, the employment of power electronic converters would limit their commercial spread due to the increased harmonics and lifetime limitations. The ST based voltage balancing system employing transformers and tap changers does not inject harmonics into the network and its feasibility in balancing the LV network has been demonstrated in Chapter 4 and Chapter 5.

Apart from the network technologies above, a number of demand response (DR) based solutions have been proposed to minimize the adverse impact of EVs and HPs on LV distribution networks. These solutions help prevent transformers from overloading, provide demand-generation balancing, help the primary frequency response, and minimize the energy costs. The study in Chapter 6 show most violations of the voltage limits occur around 19:00, when the EVs are most likely to be charged. Therefore, the DR based TOU tariff is destined to be a solution, since it encourages customers to shift selected flexible loads, e.g. EV charging, to low-consumption periods.

With the above consideration in mind, this chapter evaluates the effectiveness of the DR based TOU tariff and the ST based voltage balancing method in mitigating unacceptable voltage imbalance and magnitude variations. The simulation models use both deterministic and MC methods applied to the MC simulation platform established in Chapter 6. An introduction to the DR based TOU tariff is presented in section 7.2. Following this, the simulation methodology and results are discussed in detail.

7.2 DR based TOU tariff

The term DR was originally used in economic theory to define the short-term relationship between price and quantity with the actions and interactions of substitutes and complements taken into consideration [153]. In [154], DR was mentioned in the context of finding an appropriate price policy that leads to the correct amount of power network capacity and its efficient utilization: It is common for writers in this field to view

Page | 163 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages the problem of demand response as a series of shifts in short-run demand curves in response to a change in prices. In [155], DR was defined as ‘changes in electric use by demand-side resources from their normal consumption patterns in response to changes in the price of electricity, or to incentive payments to induce lower electricity use at times of high whole sale market prices or when system reliability is jeopardized.’ Currently, the DR based techniques can take the form of customers shifting their electricity demand from periods of peak demand and typically high prices to periods of low demand and typically lower prices [156]. This is generally attributed to a wide range of control signals such as prices, resources availability and network security. The current available DR based techniques can be categorized into either incentive-based or time-based techniques as shown in Table 7-1 [155]. DR has been identified as a potentially cost- effective way to maintain secure and sustainable energy supplies and aid the transition to a low carbon economy by reducing demand and making better use of existing generation.

Table 7-1 DR based technique types [155]

Incentive-Based Techniques Time-Based Techniques  Demand Bidding and Buyback  Critical Peak Pricing with Control  Direct Load Control  Critical Peak Pricing  Emergency Demand Response  Peak Time Rebate  Interruptible Load  Real-time Pricing  Load as Capacity Reserves  Time-of-use Pricing  Non-Spinning Reserves  System Peak Response  Regulation Service Transmission Tariff  Spinning Reserves

The studies reviewed in section 6.2 show that the distribution networks could be significantly affected by the high penetration of uncoordinated EV charging; and the impacts include violations of network constraints, decrease in power delivery efficiency and a reduction in the lifetime of the network assets. Besides, the MC study presented in Chapter 6 indicates that the penetration of EVs and HPs could induce violations of the voltage limits, especially when the EVs are charging. However, all the above adverse impacts are supposed to be largely alleviated by smart coordinated charging [120, 126]. Various coordinated charging strategies using different optimizing methods have been proposed in recent studies to achieve different objectives [157]. Linear programming was

Page | 164 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages used in [158] to maximize the total amount energy that can be delivered to all EVs. A decentralized control strategy based on integer optimization was proposed in [159] to maximize the total number of EVs that can be charged under the given system capacity. Sequential quadratic optimization was employed in [139] to minimize the variations of national power demand over time. Both quadratic and dynamic optimizations were adopted in [121] to minimize the power losses and to maximize the main grid load factor. Albeit these optimal charging strategies have been demonstrated to be feasible, they all rely on an advanced two-way communication infrastructure, connecting the control centre, sensors and smart agents, which, however, is currently not available.

TOU tariff is one important DR based technique and its effectiveness is closely related to the response of customers, independent of the smart grid communication and control system. According to the variations in electricity generation, transmission and distribution costs throughout the day, the TOU tariff is created to provide incentives for consumers to shift some of their electricity consumption to times when the cost of generating electricity is lower. Since the violations of the voltage limits mainly occur when the EVs are most likely to be charged as shown in Chapter 6, the DR based TOU tariff is supposed to be a potential way to mitigate the adverse impact of EVs and HPs by encouraging the EV owners to disperse their EV charging across the night.

The results of a series of residential TOU tariff experiments conducted in the US in 1970s, aiming at evaluating the benefits of the TOU tariffs, were summarized in [160]. Although, many of these experiments had serious flaws limiting their usefulness for estimating the price elasticity of interest, the available data show that peak-period elasticity estimates have range of something like -0.2 to -0.8, and off-peak elasticity estimates range from about -0.1 to -0.8, but since different analytic approaches were implemented in different experiments, the estimated values of the elasticity were of limited theoretical interest [160]. Moreover, the experiments were either mandatory or voluntary; the final participants only accounted for 20% of the primary sample population in the voluntary experiment in Ohio, USA, indicating a low level of participation [160]. In recent years, numerous trials have been or are being carried out around the world to test TOU tariffs; these are generally predetermined and have two (peak and off-peak) or three (base, peak and night) different prices applied to fixed periods during the day [161]. The results of these trials summarized in [161] show that

Page | 165 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages customers do shift their electricity demand in response to the TOU tariff, and the size of demand shift varies between 0% and 22% depending on tariff types and trials. Increased deployment of EVs and HPs, both controllable loads, are expected to increase the effectiveness of DR techniques, and especially the TOU tariff.

7.3 Simulation methodology

7.3.1 Monte Carlo study

A series of MC studies were undertaken on the simulation platform established in Chapter 6 and used to investigate the effectiveness of the DR based TOU tariff and the ST based voltage balancing system. The MC simulation study was based on the MC algorithm described in Fig. 6-14 but the two mitigating methods were implemented either individually or together. Since the voltage limits are likely to be violated in the non- balanced scenario, only the non-balanced scenario is studied here, and moreover, all the cases listed in Table 6-2 are investigated. The MC simulation results are numerous samples of the investigated parameters, showing the effectiveness of the two investigated mitigating methods from a viewpoint of probability. The Va, U2 and U0 at the remote end of the LV feeder are recorded and analysed; note: the remote end is the location where the voltages are at a minimum. However, when the ST balancing system is applied, the parameters at the LCP just ahead of the balancing system are also recorded and analysed. These parameters reflect the minimum voltages in the non-controlled section.

7.3.2 Deterministic study

Apart from the MC study, the effectiveness of the two mitigation methods is investigated using a deterministic approach designed to elucidate the results straightforwardly. This deterministic study is similar to that presented in section 6.6.1; carried out based on the established MC simulation platform but with predefined demand profiles. Case 5 in Table 6-2 is used as the background scenario. This involves the 96 residential demand profiles, 36 thermal demand profiles and 45 EV charging profiles, as used in section 6.6.1. These profiles are imported into the investigated LV feeder, and then a one-day

Page | 166 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages period sequential power flow is executed. The simulation results are the one-day profiles of the investigated parameters.

7.3.3 Modelling of TOU tariff

To maintain the thermal comfort of a domestic customer, the time shift of the thermal demand requires either heat storage or building insulation improvements, but the financial cost normally makes DR techniques unattractive [130]. In this study, thermal consumption is assumed not to vary when the TOU tariff is implemented, i.e. the customers only respond to TOU tariff by deferring EV charging and there is no time shift in the HP electrical demand. The effectiveness of the TOU tariff is related to the response of the customers, including the proportion of the customers willing to respond and the actions performed by them. This is difficult to predict and in this study, three participation levels (the ratio of the EV owners who respond to the TOU tariff to the total EV owners) - 10%, 20% and 30%, obtained based on literature review in section 7.2 - are assumed and investigated. Moreover, the start time of the deferred EV charging is modelled by a normal distribution with a mean of 00:00 and a standard deviation of one hour.

7.3.4 Modelling of ST based voltage balancing system

The ST based voltage balancing system shown in Fig. 3-15 is modelled in OpenDSS by assembling the specially configured transformers. Whilst, the associated control algorithm shown in Fig. 3-22 is realized by the use of MATLAB. With an ST based balancing system installed, the LV feeder is divided into two sections; the section following the balancing system is the controlled section, where the voltages can be regulated by the balancing system; whilst the section ahead of balancing system cannot be controlled. To ensure the voltages along the entire length of the LV feeder are always within statutory limits, an appropriate installation location needs to be determined for the balancing system.

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The valley values of the Va 95th % profile and the peak values of the U2 95th % profile at each LCP along the LV feeder in the case-8 non-balanced scenario are obtained from the MC study in Chapter 6, as shown in Fig. 7-1, where LCP12 refers to the LCP closest to the LV transformer and LCP21 refers to the LCP at the remote end of the LV feeder. The valley of the Va 95th % profile drops below 230V at LCP15 and the corresponding peak of the U2 95th % profile is 1.79%. In consideration of the limited operating range of the balancing system and the uncontrollability of the voltages ahead of the balancing system, the ST based balancing system is installed at LCP15 in this study, as shown in Fig. 7-2. This is feeder F1 of the investigated distribution network shown in Fig. 4-5. The customers, originally connected at LCP15, are now supplied by the output voltage of the balancing system, i.e. referred to as LCP23.

240

235

230 3%

225 2.5%

220 2% Peak of U2 95th % profile % U295th of Peak Valley of Va 95th % profile % 95th VaValley of 215 1.5%

210 1% 12 13 14 15 16 17 18 19 20 21 LCP

Fig. 7-1 Valley of Va 95th % profile and peak of U2 95th % profile at each LCP in case 5

Page | 168 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages

ST based voltage balancing system

LCP12 LCP14 LCP16 LCP18 LCP20

F2 500kVA F3

F4 LCP13 LCP15 LCP23 LCP17 LCP19 LCP21

Fig. 7-2 Investigated LV feeder F1 with ST based balancing system installed

7.4 Effectiveness of TOU tariff

7.4.1 Evaluation of TOU tariff in deterministic study

The profiles of Va at the end of the LV feeder when the TOU tariff is implemented is shown in Fig. 7-3, where ‘No deferral’ refers to the reference case without the implementation of the TOU tariff, and the profiles indicated as ‘10% deferral’, ‘20% deferral’ and ‘30% deferral’ refer to the cases where 10%, 20%, or 30% of EV owners defer their EV charging respectively. With the implementation of the TOU tariff, the valley value of the profile increases with the participation level to the TOU tariff, and it is always above the valley of the ‘No deferral’ profile, indicating the implementation of the TOU tariff helps increase the valley value of Va profile and higher participation level would achieve better result. The valley of ‘No deferral’ profile is 215 V, violating the voltage magnitude limit (217V); whilst with a participation level of 10%, the valley becomes 218.1V, which is above the lower voltage magnitude limit, and it is further increased to 222.6V with a penetration level of 30%. Moreover, the deferred EV charging occurs around 24:00, resulting in a new drop of Va around 24:00, which nevertheless is acceptable as shown in Fig. 7-3.

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250

240

230 No deferral

Voltage(V) 222.6V 19:20 10% deferral 220 20% deferral 218.1V 19:20 30% deferral 215V 19:20 210 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-3 Phase-A voltage profile with implementation of TOU tariff

Fig. 7-4 shows the U2 profiles when the TOU tariff is implemented with different participation levels. It is clearly that the TOU tariff helps to decrease the peak of the U2 profile and higher participation level would result in smaller peak of the U2 profile. Without the implementation of the TOU tariff, the peak of U2 is 2.5%, i.e. higher than the statutory voltage imbalance limit (2%). Even if the participation level is 20%, the peak of U2 (2.1%) is still above the limit. It only becomes lower than the limit when the participation level is 30%. Therefore, the TOU tariff helps to relieve severe voltage imbalance, but a high participation level is required to prevent U2 exceeding the voltage imbalance limit. The U0 behaves in the similar manner as U2, as shown in Fig. 7-5. The peak of U0 is 6.9% without the TOU tariff, while it is lowered to be 5.4% when the TOU tariff is implemented with a participation level of 30%.

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3% no deferral 2.5% 19:20 2.5% 10% deferral 2.3% 19:00 20% deferral 2.1% 19:00 2% 30% deferral 1.96% 19:00

1.5% U2

0.5%

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-4 U2 profiles with implementation of TOU tariff

8% 6.9% 19:20 no deferral 10% deferral 6% 20% deferral 5.4% 19:00 30% deferral

4% U0

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-5 U0 profiles with implementation of TOU tariff

7.4.2 Evaluation of TOU tariff in MC study

Fig. 7-6(a) shows the valley values of the Va mean profiles with different participation levels (10%, 20%, 30%) in the TOU tariff. The values indicated by ‘No deferral’ are the reference cases, i.e. the TOU tariff is not implemented. Fig. 7-6(b) shows the valley values of the Va 95th % profiles with the implementation of the TOU tariff. In each case, the valley values of Va mean and 95th % profiles increase slightly as the participation level increases from 10% to 30%. With the implementation of the TOU tariff, the valley value of the Va mean profile is above 217V in all the cases. For the

Page | 171 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages participation level of 10%, the valley value from case 3 to case 8 varies around 220V. This increases to 221V for a participation level of 20% and 223V for a participation level of 30%. However, the valley value of the Va 95th % profile is always below 217V from case 3 to case 8, even with a participation level of 30%. Accordingly, the implementation of the TOU tariff helps to alleviate the impact of EVs and HPs on the voltage magnitude on the LV feeder. However, the voltage magnitude limits are likely to be violated if the number of EVs and HPs connected to the LV feeder is ≥40 and ≥24 respectively.

No deferral 10% deferral 20% deferral 30% deferral

240

230 220

210 Voltage (Volt) Voltage 200 1 2 3 4 5 6 7 8 Case No. (a)

No deferral 10% deferral 20% deferral 30% deferral

240 230 220 210

Voltage (Volt) Voltage 200 190 1 2 3 4 5 6 7 8 Case No. (b) Fig. 7-6 Valley values of Va (a) mean and (b) 95th % profiles with implementation of TOU tariff

Fig. 7-7 shows the peak values of the U2 mean and 95th % profiles with different participation levels (10%, 20%, 30%) in the TOU tariff. As shown in Fig. 7-7(a), the peak values of the U2 mean profiles are above 2% from case 5 to case 8 when only 10% EV owners defer their EV charging; whilst only the peak values in case 7 and case 8 are above 2% when 20% EV owners defer their EV charging, and in all cases the peak values

Page | 172 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages are below 2% with a participation level of 30%. However, as shown in Fig. 7-7(b), the peak values of the U2 95th % profiles are always above 2% from case 3 to case 8, showing that the voltage imbalance limit might be violated when more than 40EVs and 24HPs are connected to the LV feeder. In addition, the simulation results show the peak of the U0 profiles drop slightly when the TOU tariff is applied; note:U0 changes in the same way as U2.

No deferral 10% deferral 20% deferral 30% deferral 3.0% 2.5%

2.0%

1.5% U2 1.0% 0.5% 0.0% 1 2 3 4 5 6 7 8 Case No. (a)

No deferral 10% deferral 20% deferral 30% deferral 4%

2% U2 1% 0% 1 2 3 4 5 6 7 8 Case No.

(b) Fig. 7-7 Peak values of U2 (a) mean and (b) 95th % profiles with implementation of TOU tariff

7.5 Effectiveness of ST based voltage balancing method

7.5.1 Evaluation of ST based voltage balancing method in deterministic study

The balancing process of the ST based balancing method mainly includes the operation of TAP1 for phase regulation and the operation of TAP2 and TAP3 for magnitude

Page | 173 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages regulation, and its effectiveness is directly related to the configuration of the tap changers. In this study, the ST based balancing system, was configured to have different tap step sizes as shown in Table 7-2 and Table 7-3. The balancing system was implemented on the investigated LV feeder and its balancing effectiveness was investigated.

For the ST based balancing system used in this study, the maximum voltage injected into phase-T by the phase regulating system is designed to be 11.5 V based on the assumption that the nominal value of the phase-M voltage equals 230V. The total tap number associated to each tap step size of TAP1 can be obtained as shown in Table 7-2. Moreover, the regulating range of the tap changers TAP2 and TAP3 are designed to be from 0.9 p.u. to 1.1 p.u., and the total tap number associated with each tap step size can be obtained as indicated in Table 7-3. In both the phase regulating system and the magnitude regulating system, the deadbands of the tap changers are set to be 1.5 times the tap step sizes. The tap changer TAP1 operates to equal the output phase-B and phase-C voltage magnitudes, and the tolerances associated with each tap step size – allowable magnitude difference (h1) – are listed in Table 7-2. Besides, the tolerances

(h2), used in the magnitude regulating system aiming to make the output phase-A and phase-B voltages equal to the desirable voltage, are shown in Table 7-3.

Table 7-2 Configuration of TAP1 (phase regulating system)

Tap step size (V) 0.575 1.15 2.3 4.6 9.2

Total tap number (excluding 40 20 10 4 2 the nominal position) Allowable magnitude 0.374 0.747 1.5 2.99 5.97 difference h1 (V)

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Table 7-3 Configuration of TAP2 and TAP3 (magnitude regulating system)

Tap step size (V) 0.71875 1.4375

Total tap number (excluding 64 32 the nominal position)

Tolerance h2 (V) 0.539 1.078 Desirable voltage (V) 245

The deterministic simulation results, for the case that TAP1 has a tap step size of 1.15V and both TAP2 and TAP3 have 64 taps, are shown in Fig. 7-8 to Fig. 7-13. Fig. 7-8 shows the 3Φ voltage profiles at the LCP14, i.e. the LCP before the balancing system on the LV feeder, where the voltages are all within the statutory limits and above 230V; the minimum 237.5V occurs on phase-A at 19:30. The voltages at LCP23, which is immediately following the balancing system, vary slightly around 245V as shown in Fig. 7-9. Moreover, the voltages at the end of the feeder, LCP21, are all within the statutory limits; the minimum 231.4V occurs on phase-A at 19:20, as shown in Fig. 7-10. By comparing with the voltages without the implementation of the ST based balancing system shown in Fig. 6-17, it can be concluded that the balancing system helps to ensure the voltages along the whole LV feeder comply with the voltage magnitude limits.

250

245

V 240 a V Voltage(V) 237.5V 19:30 b 235 V c

230 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-8 Voltage profiles at LCP14 with implementation of ST based balancing system

217V

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250

245

V 240 a V

Voltage(V) b 235 V c

230 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-9 Voltage profiles at LCP23 with implementation of ST based balancing system

250 217V 249.8V 19:00

245

V 240 a V

b Voltage(V) 235 V c

231.4V 19:20 230 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-10 Voltage profiles at LCP21 with implementation of ST based balancing system

The U2 profiles of the217V 3Φ voltages at LCP14, LCP23 and LCP21 are illustrated in Fig. 7- 11. The U2 at the three LCPs are always below 2%, in contrary to the ‘With EVs&HPs non-evenly distributed’ profile shown in Fig. 6-18. The U2 at LCP14 reaches the peak 1.24% at 19:20 and the peak of U2 at LCP21 is 0.99% at 19:00; whilst the U2 at LCP23 is much lower and varies around 0.1% throughout the day. Therefore, no violation of voltage imbalance limit occurs on the LV feeder with the balancing system implemented. Moreover, Fig. 7-12 shows the U0 profiles of the 3Φ voltages at LCP14, LCP23 and LCP21. The peaks of the U0 at LCP14 and LCP21 are 1.1% and 3.56% respectively and both of them occur around 19:00. Whilst the U0 at LCP23 is almost zero throughout the day, reflecting the elimination effect of voltage zero sequence component attributed to the ST based balancing system.

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1.5% 1.24% 19:20 LCP14 LCP23

1% 0.99% 19:00 LCP21 U2 0.5%

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-11 U2 profiles with implementation of ST based balancing system

4% 3.56% 19:00 LCP14 LCP23 3% LCP21

2% U0 1.1% 18:20 1%

0% 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time Fig. 7-12 U0 profiles with implementation of ST based balancing system

Fig. 7-13 shows the variations of the tap changers TAP1, TAP2 and TAP3 throughout the simulated day. In total, TAP1 has 53 tap changing operations, and TAP2 and TAP3 have 79 and 65 tap changing operations respectively. As shown in Fig. 7-13, most tap changing operations occur around 19:00 and are designed to deal with the variable electrical demand caused by EV charging.

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0.8 TAP1

0.6

0.4

Tap positionTap 0.2

12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time (a)

1.01

0.99 TAP2 0.98

Tap positionTap TAP3 0.97

0.96 12:00 15:00 18:00 21:00 24:00 3:00 6:00 9:00 Time (b) Fig. 7-13 Tap position of the three tap changers (a) TAP1 (b) TAP2 and TAP3

More simulation studies are carried out with the configurations listed in Table 7-2 and Table 7-3, which are successively applied to the tap changers in the ST based voltage balancing system. Table 7-4 shows a summary of the results obtained when both TAP2 and TAP3 have 64 taps and different configurations are applied to TAP1; the results include the peaks of U2 and U0, the valley of Va and the peak of Vc at the end of the LV feeder, the tap changing numbers of TAP1, TAP2 and TAP3 during the simulated day as well as their summation. Moreover, the results for the case that both TAP2 and TAP3 have 32 taps and TAP1 is configured with different settings are summarized in Table 7-5.

As shown in Table 7-4, the peak of U0 at the end of the LV feeder changes little with the

Page | 178 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages increasing tap step size of TAP1, because the ST based balancing system always output a 3Φ voltage without zero sequence component regardless of the tap settings, and the U0 at the feeder end is only related to the non-balanced load in the downstream network of the balancing system. This is also demonstrated by the results in Table 7-5, showing that the peaks of U0 are almost equal to those in Table 7-4 and do not vary with the increasing tap number of TAP2 and TAP3.

The peak of Vc increases with the increasing tap step size of TAP1 in both Table 7-4 and

Table 7-5 due to the increased h1. However, since h1 would not directly affect Va, the peak of Va stays almost constant, despite the configuration variation of TAP1. Moreover, when the tap number of TAP2 and TAP3 decreases from 64 to 32, h2 is increased by 0.539V, which is so small that this tap configuration change has little effect on the output voltage. Therefore, the valley of Va and the peak of Vc in Table 7-4 are almost the same to those in Table 7-5 for each TAP1 configuration.

It is clear the tap changing number of TAP1 during the investigated day decreases with the increasing tap step size of TAP1 as shown in Table 7-4 and Table 7-5. Moreover, the tap changing numbers of TAP2 and TAP3 in Table 7-5 are smaller than those in Table 7- 4. This indicates the increase of the tap step size contributes to less tap changing operations.

Table 7-4 Simulation results with TAP2 and TAP3 having 64 taps

Tap step Peak Peak Valley of Peak of No. of No. of No. of Total size (V) of U2 of U0 Va (V) Vc (V) TAP1 TAP2 TAP3 tap No.

0.575 0.88% 3.56% 231.4 249.3 2018 79 73 2170 1.15 0.99% 3.56% 231.4 249.8 53 79 65 197 2.3 1.10% 3.57% 231.4 250.3 17 83 72 172 4.6 1.65% 3.59% 231.3 252.5 8 81 77 166 9.2 2.38% 3.61% 231.6 255.3 3 81 80 164

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Table 7-5 Simulation results with TAP2 and TAP3 having 32 taps

Tap step Peak Peak Valley of Peak of No. of No. of No. of Total size (V) of U2 of U0 Va (V) Vc (V) TAP1 TAP2 TAP3 tap No.

0.575 0.88% 3.57% 231.2 249.2 2020 28 23 2075 1.15 1.00% 3.57% 231.1 249.8 55 28 26 109 2.3 1.13% 3.58% 231.1 250.3 17 28 27 72 4.6 1.68% 3.59% 231.1 252.5 8 28 33 69 9.2 2.49% 3.63% 230.9 255.4 2 28 35 65

Fig. 7-14 shows the variations of the U2 at the feeder end and the total tap changing number when the tap size of TAP1 increases from 0.575V to 9.2V with both TAP2 and TAP3 having 64 taps; data obtained from Table 7-4. This figure shows the total tap changing number required during the simulated day decreases with the increasing tap step size of TAP1. However, larger tap step size of TAP1 induces higher peak values of U2, reducing the balancing effectiveness. This is also demonstrated by the results in Table 7-5, which shows the case with both TAP2 and TAP3 having 32 taps. Therefore, the determination of the tap configuration is a compromise between the required tap changing number and the balancing effectiveness, i.e. the tap changers should be appropriately configured to satisfy the conflicting goals of minimizing the tap changing number and maximizing the balancing effectiveness.

3% 3000

2% 2000 U2

1% 1000 Totaltapchanging No.

0 0 0.575 1.15 2.3 4.6 9.2 Tap step size of TAP1 (V) Fig. 7-14 Total tap changing number and U2 at the feeder end

Page | 180 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages

7.5.2 Evaluation of ST based voltage balancing method in MC study

The deterministic study in 7.5.1 indicates a tap step size of 1.15V for TAP1 and the use of 64 taps for both TAP2 and TAP3 deliver an appropriate total tap changing number and an acceptable balancing result. With this configuration, the ST based balancing system was implemented on the LV feeder in the MC simulation platform, and a simulation study was carried out to investigate its effectiveness. The 3Φ voltages, the associated U2s and U0s at LCP14 and LCP21 were analyzed below.

Fig. 7-15 illustrates the valley values of the Va mean profiles at LCP21 with/without the use of the ST based voltage balancing system. Without the balancing system, the valley value is above 230V only in case 1 and varies slightly around 218V from case 3 to case 8. When the balancing system is applied, the valley values are all above 230V. Fig. 7-16 indicates activation of the ST based balancing system significantly lifts the valley values of the Va 95th % profiles at LCP21; the valley values are now all above 225V. The simulation results show the valley of Va 95th % profile at LCP 14 is above 230V when the ST based balancing system is activated. Therefore, when the ST based balancing system is implemented on the LV feeder, the voltage magnitude limits are unlikely to be violated.

without ST balancing system with ST balancing system 240 235

230 225 220 215 210 Voltage (V) Voltage 205 1 2 3 4 5 6 7 8 Case No. Fig. 7-15 Valley of Va mean profiles at LCP21with/without implementation of ST based balancing system

Page | 181 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages

without ST balancing system with ST balancing system 240 230 220 210

Voltage (V) Voltage 200 190 1 2 3 4 5 6 7 8 Case No. Fig. 7-16 Valley of Va 95th % profiles at LCP21 with/without implementation of ST based balancing system

When the ST based balancing system is activated, the peak values of the U2 mean profiles at LCP21 are all lower than 1% as shown Fig. 7-17, and moreover, the peaks of the U2 95th % profile at LCP21 is below 1.2% in all the cases shown in Fig. 7-18. Additionally, the simulation results show peaks of the U2 95th % profiles at LCP14 are all lower than 2% with the balancing system activated. Accordingly, the ST based balancing system is capable of mitigating a severe voltage imbalance.

The peak of the U0 mean profile at LCP21 varies around 6% from case 2 to case 8, when the balancing system is not activated. However, the peak varies around 3% when activated as shown in Fig. 7-19. Likewise, the peak of the U0 95th % profile is reduced from around 8% to 4% in case 2 to case 8 when the balancing system is applied as shown in Fig. 7-20.

without ST balancing system with ST balancing system 3.0% 2.5%

2.0%

1.5% U2 1.0% 0.5% 0.0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-17 Peak of U2 mean profiles at LCP21 with/without implementation of ST based balancing system

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without ST balancing system with ST balancing system 4.0%

3.0%

2.0% U2 1.0% 0.0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-18 Peak of U2 95th % profiles at LCP21 with/without implementation of ST based balancing system

without ST balancing system with ST balancing system 8%

4% U0 2% 0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-19 Peak of U0 mean profiles at LCP21 with/without implementation of ST based balancing system

without ST balancing system with ST balancing system 10% 8%

U0 4% 2% 0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-20 Peak of U0 95th % profiles at LCP21 with/without implementation of ST based balancing system

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7.6 Cooperation of ST based balancing system and TOU tariff

The effectiveness of using both of the ST based balancing system and the TOU tariff in mitigating the adverse impacts of EVs and HPs was evaluated here. Based on the MC simulation platform established in Chapter 6, the ST based voltage balancing system was implemented on the investigated LV feeder together with the TOU tariff and three participation levels (10%, 20% 30%) to the TOU tariff are considered.

Fig. 7-21 shows the valley values of the Va mean profiles at LCP21. The values indicated by ‘No referral’ refer to the results obtained when the ST based balancing system is implemented alone and these are used for reference. Moreover, ‘10% deferral’, ‘20% deferral’ and ‘30% deferral’ refer to the cases where both investigated mitigating methods are implemented with the participation level to TOU tariff being 10%, 20% and 30% respectively. Fig. 7-21 shows, from case 5 to 8, the valley is around 233V with the ST based balancing system implemented alone, and it is around 234V with a participation level of 10%, around 235V with a participation level of 20%, and around 236V with a participation level of 30%, indicating co-operative use of the two mitigation methods raises the valley value, and a higher participation level in the TOU tariff induces a higher valley value. Fig. 7-22 shows the valley of the Va 95th % profile at LCP21, behaves in a similar manner as the Va mean profile. With a participation level of 30%, the valley of the Va 95th % profile is above 230V in all the cases studied.

No deferral 10% deferral 20% deferral 30% deferral 240 238 236 234 232

Voltage (V) Voltage 230 228 1 2 3 4 5 6 7 8 Case No. Fig. 7-21 Valley of Va mean profiles at LCP21 with cooperation of ST balancing system and TOU tariff

Page | 184 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages

No deferral 10% deferral 20% deferral 30% deferral 236

234 232 230 228

Voltage (V) Voltage 226 224 1 2 3 4 5 6 7 8 Case No.

Fig. 7-22 Valley of Va 95th % profiles at LCP21 with cooperation of ST balancing system and TOU tariff

In contrast to Va, U2 and U0 reduce when the TOU is implemented together with the ST based balancing system. The peaks of the U2 mean and 95th % profiles at LCP21 with co-operative use of the two mitigation methods are shown in Fig. 7-23 and Fig. 7-24 respectively. The peak decreases with an increasing participation level in the TOU tariff; for example, with the balancing system implemented alone, the peaks of the U2 mean and 95th % profiles vary around 0.8% and 1.1% respectively, from case 2 to case 8. However, when the participation level is 30%, they are reduced to below 0.7% and 1% respectively. Similarly, the peak of U0 mean and 95th % profiles at LCP21 declines with increasing TOU participation, as shown in Fig. 7-25 and Fig. 7-26.

No deferral 10% deferral 20% deferral 30% deferral 1.0% 0.8% 0.6% U2 0.4% 0.2% 0.0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-23 Peak of U2 mean profiles at LCP21 with cooperation of ST balancing system and TOU tariff

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No deferral 10% deferral 20% deferral 30% deferral 1.2% 1.0%

0.8%

U2 0.6% 0.4% 0.2% 0.0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-24 Peak of U2 95th % profiles at LCP21 with cooperation of ST balancing system and TOU tariff

No deferral 10% deferral 20% deferral 30% deferral 4.0%

3.0%

2.0% U0 1.0% 0.0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-25 Peak of U0 mean profiles at LCP21 with cooperation of ST balancing system and TOU tariff

No deferral 10% deferral 20% deferral 30% deferral 5.0% 4.0%

3.0% U0 2.0% 1.0% 0.0% 1 2 3 4 5 6 7 8 Case No. Fig. 7-26 Peak of U0 95th % profiles at LCP21 with cooperation of ST balancing system and TOU tariff

The simulation results show Va, U2 and U0 at LCP14 behave in the same manner as the equivalents at LCP21, and they are all within the statutory limits. Above all, the cooperation with the TOU tariff helps to enhance the efficacy of the ST based balancing system; further reducing the voltage magnitude variation and the imbalance along the LV

Page | 186 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages feeder.

7.7 Summary

In this chapter, the DR based TOU tariff and the ST based voltage balancing method were used to mitigate the adverse impact of EVs and HPs on the LV feeder voltages. Their effectiveness was investigated by using both deterministic and probabilistic approaches based on the MC simulation platform established in Chapter 6.

The results show the application of the TOU tariff contributes to relieve the adverse impacts of EVs and HPs. However, its effectiveness is limited and it relies on the response of the customers; higher participation levels would achieve greater effectiveness. The simulation results indicate that even if 30% EV owners defer their EV charging, the voltage limits still have a possibility to be violated when 40EVs and 24HPs or more are connected to the LV feeder.

The efficacy of the ST based voltage balancing system is related to the configurations of the tap changers; smaller tap step size helps to obtain lower level of voltage imbalance, but requires more tap changing operations, which should be reduced to increase the balancing speed and introduce less transients into the network. Therefore, the configuration of the tap changer should be determined to reach the contradictory goals of maximum balancing effectiveness and minimum tap changing number. Moreover, since the balancing system has a limited operating range and the voltages on the upstream feeder of the balancing system are uncontrollable, an appropriate installation location should be determined.

This chapter demonstrated how the ST based voltage balancing system can completely mitigate the violations of the voltage limits regardless of the penetration level of EVs and HPs. The valley value of the voltage magnitude is always over 225V and the peak value of the U2 is always below 1.2% at LCP21. When the TOU tariff and the balancing system are implemented together on the LV feeder, the voltage quality on the LV feeder is further improved. This is very important for the voltage regulation on a long LV feeder; because the ST based balancing system alone may not be able to solve the breach

Page | 187 Chapter 7 Mitigation of Impact of EVs & HPs on LV Feeder Voltages of the voltage limits occurring too far away. Besides, the TOU tariff could help to mitigate the violations occurring in the uncontrolled section, when the originally designed installation location is not appropriate for the new situation.

Page | 188 Chapter 8 Conclusion & Future Work

CHAPTER 8

CONCLUSION & FUTURE WORK

8.1 Conclusion

Increasing penetration of low carbon technologies, including EVs and HPs, are expected to deteriorate the voltage imbalance and magnitude variations seen on UK LV networks. However, existing voltage control in LV networks relies on on-load tap charging transformers, shunt capacitors and voltage regulators located in the MV and HV networks, but these devices are not capable of solving LV network voltage imbalance problems. The project of this thesis was motivated to investigate the solution of mitigating the adverse impact of EVs and HPs on LV feeder voltages. This aim has been successfully achieved through the work presented in this thesis which is summarized as follows.

The literature on mitigation of the voltage imbalance is thoroughly reviewed, including both conventional and modern voltage imbalance mitigating methods.  Conventional methods to deal with the voltage imbalance in LV networks include increasing the LV cable cross-section, installing a new LV cable and installing additional protective equipment, which are not economically beneficial. An alternative is manually changing the supply phase to deal with voltage imbalance. However, this method is time-consuming; and requires field measurements and/or

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software simulation and analysis, and generally does not provide an effective long term solution.  In recent years, alternative methods have been proposed to deal with voltage imbalance. The installation of on-load tap changers on MV/LV transformers was shown to be able to maintain the voltage profile along the LV feeder within the statutory limits, but the independent control of each phase may worsen the voltage imbalance. The automatic supply phase selection using static switches relies on an advanced communication and control system, which is currently not available. Moreover, it could be physically restricted in some networks. Several voltage imbalance mitigating methods, based on power electronic converters, have been proposed. The reactive power from a D-STATCOM can be controlled to reduce voltage imbalance. The inverter circuit in combination with either a DC generator, a battery source, an ACDC rectifier, or even the RES, can be controlled to inject the active and reactive power necessary to reduce the voltage imbalance. In addition, the power electronics based AC-DC-AC solution is a viable option to resolve voltage imbalance. However, the power electronic converters would inject harmonics into the network, reducing the power delivery efficiency. Moreover, the main components of the converters, e.g. IGBT-transistors and capacitors, generally have a short lifetime. Hence, due to these drawbacks, commercial use of power electronic converter based solutions is likely to be limited.

The Scott transformer (ST) based voltage balancing method was proposed in this project. The performance of this method was investigated using both computer simulation and a physical test system.  Compared with the modern voltage imbalance mitigating methods, the ST based voltage balancing method relying on transformers and tap changers does not introduce harmonics into the networks and has a long operating lifetime. Moreover, the operation of this method does not rely on an advanced communication and control system, external supplies or energy storage. In addition, some existing tap-changing transformers based arrangements, capable of independently controlling the voltage magnitude and phase angle, are supposed to be able to mitigate voltage imbalance, but the ST based voltage balancing method is simpler to construct and uses a less elaborate control algorithm.

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 The ST based voltage balancing method and the 3Φ voltage balancing method were compared in a simulation study, where both methods were implemented on a simplified LV feeder with severe voltage imbalance. The simulation results show that these two methods can reduce voltage magnitude variations and imbalance. However, the ST based balancing method appears to be superior to the 3Φ balancing method; since using the same tap changing configuration, the ST based balancing method has a larger regulating range and requires less tap changing operations for the same voltage imbalance.  The performance of the ST based voltage balancing method in a typical UK distribution network was investigated. The balancing system essentially divides the LV feeder into two sections; the balancing system controls the voltages on the feeder downstream of the balancing system, whilst the feeder upstream of the balancing system is beyond its control. The performance of the balancing system is closely related to its output voltage magnitude and the installation location. When the output voltage magnitude is too high, the voltage on the light-load phase is likely to rise above the upper limit; in the contrary, when the output voltage magnitude is too low, the voltage on the heavy-load phase may drop below the lower limit. Accordingly, the output voltage magnitude of the balancing system needs to deal with the divergence of the 3Φ voltages. Moreover, since the LV feeder ahead of the balancing system is non-controllable, the voltage in this section may violate the limit when the balancing system is located too far away from the MV/LV substation. Since the balancing system has a limited regulating range, the voltage at the feeder end may violate the limits when the feeder following the balancing system is too long. Therefore, the location of the balancing system should be appropriately determined to ensure the voltages along the whole LV feeder comply with the limits.  The feasibility of the ST based balancing method was further demonstrated using physical test facilities. Based on the ST based balancing method, a small-scale physical voltage balancing system was established. This includes a microprocessor based control system and is able to automatically complete the voltage balancing. This physical balancing system was implemented on an LV feeder in the laboratory and used for voltage balancing under different operating scenarios. The test results show that the balancing system is capable of rapidly

Page | 191 Chapter 8 Conclusion & Future Work

restoring the balance state on the LV feeder after the voltage disturbances caused by 1Φ load variations.

An MC simulation platform was established. Based on this platform, the impact of EVs and HPs on the LV feeder voltage imbalance and magnitude variation was analysed and quantified. Moreover, the effectiveness of the ST based balancing system and the DR based TOU tariff, implemented either alone or together, in mitigating the adverse impact of EVs and HPs was evaluated.  The MC simulation platform contains:- a statistical model of the EV charging demand, the profile generators of the HP electrical demand and the residential demand, and a distribution network model which was based on the UKGDN. The statistical EV demand model was created by modelling the uncertain parameters based on their associated probability distributions, and this model was able to generate a random EV charging profile during the MC simulation. Moreover, the profile generators of HP electrical demand and residential demand were created based on a database containing numerous domestic thermal demand profiles and a database containing numerous residential demand profiles, respectively.  The impact of EVs and HPs on the LV feeder voltages is closely related to the distribution of EVs and HPs on the LV feeder. When the EVs and HPs are evenly distributed between the three phases, the voltage drop and imbalance along the LV feeder are deteriorated, but still comply with the statutory voltage limits, despite the penetration level of EVs and HPs. However, when the number of EVs and the number of HPs connected to a 96-customer LV feeder in a non-evenly manner are ≥21 and ≥13 (estimated penetration level in 2026), both the voltage magnitude and imbalance limits might be breached. When the penetration level reaches 40 EVs and 24 EVs (estimated penetration level in 2030), the estimated mean of U2 violates the limit, indicating the voltage imbalance limit is likely to be violated.  The effectiveness of the ST based voltage balancing system in mitigating the impact of EVs and HPs on the LV feeder voltages was investigated based on the established MC simulation platform. The results show the effectiveness of the balancing system compared to the step size of the tap changers; i.e. a smaller tap step size obtains lower levels of voltage imbalance but requires more regulating

Page | 192 Chapter 8 Conclusion & Future Work

time. With a proper installation location and tap step size, the ST based voltage balancing system is able to completely mitigate the violations of the voltage limits caused by the penetration of EVs and HPs.  The DR based TOU tariff relieves the adverse impact of EVs and HPs on the LV feeder voltages. However, its effectiveness is related to the response of the customers in the TOU tariff; i.e. the effectiveness of the TOU tariff increases with the number of the customers who respond. The recent trials testing the TOU tariff show the size of the load shift attributable to the TOU tariff varies between 0% and 22% depending on tariff types and trials. The study in this project show that even with 30% EV owners deferring their EV charging, the voltage limits can still be violated when 40EVs and 24HPs or more are connected to the LV feeder.  The cooperation of the ST based voltage balancing system and the DR based TOU tariff enhances the effectiveness of the ST based balancing system; the valley of the voltage magnitude is further increased and the peak of the U2 is further decreased. Since the ST based balancing system has a limited regulating range, the enhancement of the balancing effect attributable to the TOU tariff is beneficial to the application of the ST based balancing system, especially on a long LV feeder. In addition, when the originally designed installation location of the balancing system is not suitable for the new situation, the application of the TOU tariff could help to avoid the cost of the relocation of the balancing system.

8.2 Future work

All the primary objectives have been successfully completed in this project whose aim has been to evaluate and mitigate the adverse impact of EVs and HPs on the LV feeder voltages. An ST based voltage balancing method was developed and tested in both computer simulation and physical experiment. An MC simulation platform was established; based on this platform, the impacts of EVs and HPs on the LV feeder voltages as well as the associated mitigation using the ST based balancing system and the DR based TOU tariff were quantified and analysed. With the above in mind, the future work of this project is presented in this section.

Page | 193 Chapter 8 Conclusion & Future Work

The future development of the research presented in this thesis should further investigate the solutions to the adverse impacts of LCTs on LV networks and promote the practical application. To realize this, the MC simulation platform needs to be improved to give a more comprehensive understanding of the impact of EVs and HPs on LV networks, and effectiveness of the mitigating methods. Following this, the site trial of the ST based voltage balancing method can be carried out.

Besides EVs and HPs, small-scale DGs, and especially photovoltaics and CHP plants, are expected to be increasingly connected to LV networks. Recent studies [115, 162] have shown that significant penetration of small-scale DGs will challenge voltage regulation in LV networks, but the literature on the impact of small-scale DGs together with EVs and HPs is limited. Therefore, the future work includes developing the models of the small-scale DGs and integrating these models into the established MC simulation platform to reflect the real electrical demand scenario in expected future LV networks. In this way, the impact of EVs, HPs and small-scale DGs can be evaluated, which will deliver the close-to-realistic challenges faced by future LV networks. Moreover, when small-scale DGs are included, the effectiveness of the ST based voltage balancing method can be more accurately evaluated.

Apart from small-scale DGs, the MC simulation platform can be further improved from two aspects: - the first is to use the real network data and the second is to improve the modelling of the HP electrical demand. The distribution network model in the MC simulation platform is developed from the UKGDN, which is a standard distribution network model created for research. However, the LV network is always diversified, adding to this the fact that the ST based voltage balancing method has a limited regulating range and its performance is influenced by its installation location, network models developed from the real LV networks should be adopted. The HP electrical demand profiles used in the MC simulation platform were directly converted from the one-hour domestic thermal demand profiles. The one-hour resolution is much lower than the one-minute resolution of the residential and EV demand profiles, and this reduces the accuracy of the simulation results. In addition, the intermittent operation pattern of HP is not considered; for example, with a high level of building insulation, an HP can be turned off for several hours whilst the room temperature is maintained within a specified range. Therefore, a detailed model of HP should be developed, and it can provide high-

Page | 194 Chapter 8 Conclusion & Future Work resolution HP electrical demand profiles. These two improvements of the MC simulation platform would help to deliver more accurate results.

The site trial of the ST based voltage balancing system is an important part of the future work, which is vital to the practical application and the commercial exploitation of the balancing system. To complete the site trial, a full-size prototype of the balancing system should be established. In this project, a small-scale prototype of the balancing system was established and used for the physical test. The tap changing operation in the small- scale balancing system was accomplished using miniature relays, with a current capacity that is inadequate for a commercial tap changing facilities. Moreover, since the conventional mechanical tap changer has a low tap changing speed and suffers severe mechanical wear, it is not suitable to solve the voltage disturbance problems caused by volatile LV network loads. Therefore, a full-size prototype of the balancing system and a future ‘commercial’ balancing system needs to adopt electronic tap changers, using semiconductor devices, IGBTs, to change the tap position.

In addition to the technical feasibility investigations presented above, the financial costs over the lifetime of the ST based balancing system, 3Φ balancing system, LV feeder reinforcement and other power electronics based balancing solutions should be investigated and compared in future work. A brief analysis of the investment costs of ST based balancing system, LV cable replacement and power electronics based AC-DC-AC solution is presented in Appendix D, and a detailed analysis should be carried out in future work. Power losses of power system equipment increase due to non-balanced 3Φ voltage supply, and the raising harmonics introduced by power electronic devices further increase the power losses of power system equipment and reduce their lifetimes. The no- load and load losses of ST based balancing system and other balancing solutions, as well as the associated costs, should be studied in detail in future work.

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References

1. R. C. Dugan, M. F. McGranahan, and H. W. Beaty, Electrical power systems quality. New York: McGraw-Hill, 1996, pp. 5, 12, 23. 2. V. Gosbell, S. Perera, and V. Smith, "Power quality in future low voltage electric y networks," University of Wollongong, Dec. 2012. [Online]. Available: http://w ww.copper.com.au/copper/wcms/en/home/projects/energy_efficiency/network-lo sses/power-quality-reliability/Power-Quality-in-Future-Low-Voltage-Electricity- Networks-Gosbell.pdf 3. Voltage characteristics of electricity supplied by public distribution networks, British Std. BS EN 50160:2011, 2010. 4. Nominal voltages for low votlage public electricty supply systems (Implementation of HD 472 S1), British Std. BS 7679:1993, 1993. 5. "A Roadmap for moving to a competitive low carbon economy in 2050," Eurppea n Union, 2011. [Online]. Avaliable: http://eur-lex.europa.eu/resource.html?uri=ce llar:5db26ecc-ba4e-4de2-ae08-dba649109d18.0002.03/DOC_1&format=PDF 6. P. Papadopoulos, L. M. Cipcigan, N. Jenkins, and I. Grau, "Distribution networks with Electric Vehicles," in Proc. 2009 44th Universities Power Engineering Int. Conf.. 7. "UK Greenhouse Gas Emission - 1st Quarter 2014 Provisional Figures," Departm ent of Energy & Climate Change, UK, 2014. [Online]. Available: https://www.go v.uk/government/uploads/system/uploads/attachment_data/file/328695/GHGQuar terly_statistical_release.pdf 8. "UK future energy scenario," National Grid, UK, 2011. [Online]. Available: http:/ /www.nationalgrid.com/NR/rdonlyres/86C815F5-0EAD-46B5-A580-A0A51656 2B3E/50819/10312_1_NG_Futureenergyscenarios_WEB1.pdf 9. A. Hawkes, L. Munuera, and G. Strbac, "Low carbon residential heating," 2011. [Online]. Available: https://workspace.imperial.ac.uk/climatechange/Public/pdfs/ Briefing%20Papers/Low%20carbon%20residential%20heating.pdf 10. T. Fawcett, "The future role of heat pumps in the domestic sector," 2011. [Online]. Available: http://www.eci.ox.ac.uk/publications/downloads/fawcett11b.pdf 11. "Delivering UK Energy Investment: Networks," Department of Energy & Climate Change, UK, 2015. 12. A. Harris, "Charge of the electric car - [power electric vehicles]," Engineering & Technology, vol. 4, pp. 52-53, 2009. 13. P. Kadurek, J. F. G. Cobben, and W. L. Kling, "Future LV distribution network design and current practices in the Netherlands," in Proc. 2011 IEEE Innovative Smart Grid Technologies Int. Conf.. 14. E. Lakervi, and E. J. Holmes, Electricity distribution network design. London: Peter Peregrinus Ltd, 1989, pp. 176-210. 15. "Connection to the LV utility distribution network, in Electrical installation guide, " Schneider Electric, 2009. [Online]. Available: http://www.schneider-electric.co m.au/documents/electrical-distribution/en/local/electrical-installation-guide/EIG- C-connection-LV-architecture-guide.pdf 16. H. L. Willis, Power distribution planning reference book. New York: Marcel Dekker, 1997, pp. 236-239.

Page | 196 References

17. T. Haggis, "Network Design Manual," E.On Central Network, 2006. [Online]. Available: http://www.eon-uk.com/downloads/network_design_manual.pdf 18. C. D. Poole, and T. E. Marks, Electrical distribution in buildings. Oxford: Blackwell Scientific, 1994, pp. 39-40. 19. "ABB AS, Power Products Division, SF6-insulated ring main unit, safering 36 an d SF6-insulated compact switchgear safeplus 36," ABB product catalogue. [Onlin e]. Available: http://www05.abb.com/global/scot/scot235.nsf/veritydisplay/6e82f 84352cc9b32c12577ed004ce454/$file/catalogue%201vdd006114%20gb%20janu ar%202014.pdf 20. J. M. Biasse, Y. Chollot, and J. Tobias, "Circuit-breaker RMU improves MV/LV transformer protection," in 20th international Conference and Exhibition on Electricity Distribution, 2009. 21. A. A. Regotti, and J. G. Trasky, "What to look for in a low-voltage unit substation," IEEE Trans. Industry and General Applications, vol. IGA-5, pp. 710- 719, 1969. 22. IEEE Standard Definitions for Power Switchgear, IEEE Std C37.100-1992, 1992. 23. S. Massucco, and F. Silvestro, "Dispower distributed generation with high penetration of renewable energy sources, ENK5-CT2001-00522 Appendix - Structure and data concerning electrical grids for Italy, Germany, Spain, UK, Poland," 2003. 24. "Low voltage network design," Ergon Energy, 2003. [Online]. Available: https:// www.ergon.com.au/__data/assets/pdf_file/0018/64233/Section-15-LV-Network- Design.pdf 25. CENELEC standard voltages, British Std. BS EN 60038, 2011. 26. The Electricity Safety, Quality and Continuity Regulations, UK Legislation, 2002. [Online]. Available: http://www.legislation.gov.uk/uksi/2002/2665/pdfs/uksi_20 022665_en.pdf 27. (Webpage) http://electrical-engineering-portal.com/north-american-versus-europe an-distribution-systems, 2011. 28. J. J. Bodson, "Electric current abroad," US Department of Commerce Internationa l Trade Administration, 2002. [Online]. Available: http://www.ita.doc.gov/media/ publications/pdf/current2002final.pdf 29. IEEE Recommended Practice for Electric Power Distribution for Industrial Plants, IEEE Std. 141-1993, 1994. 30. J. Driesen, and T. V. Craenenbroeck, "Voltage Disturbances, Introduction to Unb alance," Cooper Development Association, 2002. [Online]. Available: http://admi n.copperalliance.eu/docs/librariesprovider5/power-quality-and-utilisation-guide/5 13-introduction-to-unbalance.pdf?sfvrsn=4&sfvrsn=4 31. C. L. Fortescue, "Method of Symmetrical Co-Ordinates Applied to the Solution of Polyphase Networks," American Institute of Electrical Engineers, Trans. of the, vol. XXXVII, pp. 1027-1140, 1918. 32. D.A.K. Redman, "Phase-sequence component" Students' Quarterly Journal, vol. 3, pp. 170-173, 1993. 33. Electromagnetic compatability (EMC) Testing and measurement techniques, Power quality measurement methods, British Std. BS EN 61000-4-30, 2009. 34. P. Pillay and M. Manyage, "Definitions of voltage unbalance" IEEE Power Engineering Review, vol. 22, pp. 49-50, 2002. 35. Voltage characteristics of electricity supplied by public distribution networks, British Std. BS EN 50160, 2007. 36. Motors and Generators, ANSI/NEMA Std. MG1-1993, 1993.

Page | 197 References

37. IEEE Standard Test Procedure for Polyphase Induction Motors and Generators, IEEE Std. 112, 1991. 38. G. Burchi, C. Lazaroiu, N. Golovanov, and M. Roscia, "Estimation of voltage unbalance in power system supplying high speed railway," Electrical Power Quality and Utilisation, vol. XI, 2005. 39. C. Bin-Kwie, and G. Bing-Song, "Three phase models of specially connected transformers. Power Delivery," IEEE Trans. Power Delivery, vol. 11, pp. 323- 330, 1996. 40. T. H. Chen, and H. Y. Kuo, "Network modelling of traction substation transformers for studying unbalance effects," IET Generation, Transmission and Distribution, vol. 142, pp. 103-108, 1995. 41. A. K. Singh, G. K. Singh, and R. Mitra, "Some observations on definitions of voltage unbalance," in Proc. 39th North American Power Symposium, 2007. 42. Planning limits for voltage unbalnce in the United kingdom, Energy Networks Association, Engineering Recommendation P29, 1990. 43. Electric Power Systems and Equipment—Voltage Ratings (60 Hertz), ANSI Std. ANSI C84.1-2006, 2006. 44. Motors and Generators, ANSI/NEMA Std. MG1-2009, 2009. 45. "Matching utilization voltages to motor nameplates," Electric Power Research Ins titute, 2007. [Online]. Available: http://www.epri.com/abstracts/Pages/ProductAb stract.aspx?ProductId=000000000001017014 46. IEEE Recommended Practice for Electric Power Systems in Commercial Buildings, IEEE Std 241-1990, 1991. 47. "The Carbon Trust & DTI renewables network impacts study final report" Mott MacDonald Ltd, 2003. [Online]. Available: http://webarchive.nationalarchives.go v.uk/20100919181607/http:/www.ensg.gov.uk/assets/ct_impact_of_re_final_repo rt.pdf 48. P. Trichakis, P. C. Taylor, L. M. Cipcigan, and P. F. Lyons, "An investigation of voltage unbalance in low voltage distribution networks with high levels of SSEG," in Proc. 2006 Int. Universities Power Engineering Conf., pp. 182-186. 49. J. J. Grainger, and W. D. Stevenson, Power System Analysis, New York: McGraw-Hill, 1994, pp. 64-65. 50. S. D. Rubira and M. D. McCulloch, "Control method comparison of doubly fed wind generators connected to the grid by asymmetric transmission lines," IEEE Trans. Industry Applications, vol. 36, pp. 986-991, 2000. 51. D. R. Smith, H. R. Braunstein, and J. D. Borst, "Voltage unbalance in 3- and 4- wire delta secondary systems," IEEE Trans. Power Delivery, vol. 3, pp. 733-741, 1988. 52. Y. Ruifeng, and T. K. Saha, "Voltage variation sensitivity analysis for unbalanced distribution networks due to photovoltaic power fluctuations," IEEE Trans. Power Systems, vol. 27, pp. 1078-1089, 2012. 53. R. P. Broadwater, A. H. Khan, H. E. Shaalan, and R. E. Lee, "Time varying load analysis to reduce distribution losses through reconfiguration," IEEE Trans. Power Delivery, vol. 8, pp. 294-300, 1993. 54. A. V. Jouanne and B. Banerjee, "Assessment of voltage unbalance," IEEE Trans. Power Delivery, vol. 16, pp. 782-790, 2001. 55. L. Ching-Yin, "Effects of unbalanced voltage on the operation performance of a three-phase induction motor," IEEE Trans. Energy Conversion, vol. 14, pp. 202- 208, 1999.

Page | 198 References

56. R. Neumann, Symmetrical component analysis of unsymmetrical polyphase systems. London: SIR ISAAC PITMAN & SONS, LTD, 1996, pp. 124-133. 57. L. E. Williams, "Operation of 3-Phase Induction Motors on Unbalanced Voltages," Power Apparatus and Systems, Part III. Trans. of the American Institute of Electrical Engineers, vol. 73, pp. 125-133, 1954. 58. B. N. Gafford, W. C. Duesterhoeft, and C. C. Mosher, "Heating of induction motors on unbalanced voltages," Power Apparatus and Systems, Part III. Trans. of the American Institute of Electrical Engineers, vol. 78, pp. 282-286, 1959. 59. M. M. Berndt, and N. L. Schmitz, "Derating of polyphase induction motors operated with unbalanced line voltages," Power Apparatus and Systems, Part III. Transactions of the American Institute of Electrical Engineers, vol. 81, pp. 680- 683, 1962. 60. E. Muljadi, R. Schiferl, and T. A. Lipo, "Induction machine phase balancing by unsymmetrical thyristor voltage control," IEEE Trans. Industry Applications, vol. IA-21, pp. 669-678, 1985. 61. M. W. Davis, R. Broadwater, and J. Hambrick, "Modeling and testing of unbalanced loading and voltage regulation final report," NREL/SR 581-41805, National Renewable Energy Laboratory: Golden, Colo, 2007. 62. M. G. Davoudi, A. Bashian and J. Ebadi. "Effects of unsymmetrical power transmission system on the voltage balance and power flow capacity of the lines," in Proc. 2012 11th Int. Conf. Environment and Electrical Engineering. 63. K. Lee, G. Venkataramanan, and T.M. Jahns, "Source current harmonic analysis of adjustable speed drives under input voltage unbalance and sag conditions," IEEE Trans. Power Delivery, vol. 21, pp. 567-576, 2006. 64. S. S. Deswal, R. Dahiya and D. K. Jain, "Effect of votlage unbalance on adjustable speed drives and its mitigation using supercapacitor," Int. Journal of Electronics Communications and Electrical Engineering, vol. 3, 2013. 65. "Input performance of ASD's during supply voltage unbalance," EPRI Power Ele ctron, 1996. [Online]. Available: http://www.epri.com/search/Pages/results.aspx? k=Input%20Performance%20of%20ASDs%20During%20Supply%20Voltage%2 0Unbalance 66. J. L. Duran-Gomez, P. N. Enjeti, and W. Byeong Ok, "Effect of voltage sags on adjustable-speed drives: a critical evaluation and an approach to improve performance," IEEE Trans. Industry Applications, vol. 35, pp. 1440-1449, 1999. 67. S. Sivanagaraju, Electric power transmission and distribution. India: Dorling Kindersley, 2009, pp. 39-40. 68. E. T. B. Gross, and A. H. Weston, "Transpositions and unbalance of high-voltage lines," Electrical Engineering, vol. 71, pp. 606-606, 1952. 69. Y. Ruifeng, and T. K. Saha, "Investigation of voltage imbalance due to distribution network unbalanced line configurations and load levels," IEEE Trans. Power Systems, vol. 28, pp. 1829-1838, 2013. 70. Q. Z. Li, J. S. Zhang and Q. Q. Qian, "Optimization design on series tuning filtering and reactive compensation used in traction systems," in Proc. 1989 Main Line Railway Electrification Int. Conf. 71. G. Astengo, C. Rossi, and G. Sciutto, "Single-phase 25kV electrification of Sardinia railway: some particular problems related to an electrified railway located in an island," in Proc. 1989 Main Line Railway Electrification Int. Conf.. 72. R. Cella, G. Giangaspero, A. Montepagano, P. Pozzobon, M. Ruscelli, and M. Vanti, "Measurement of AT electric railway system currents at power-supply

Page | 199 References

frequency and validation of a multiconductor transmission-line model," IEEE Trans. Power Delivery, vol. 21, pp. 1721-1726, 2006. 73. "Power transformers and reactors," Alstom. [Online]. Available: http://www.alsto m.com/Global/Grid/Resources/Power%20Transformers/Documents/Power%20tra nsformers%20and%20reactors%20Brochure%20GB.pdf 74. A. G. Williams, "New Zealand Railways: an overview of the North Island Main Trunk electrification," in Proc. 1989 Main Line Railway Electrification Int. Conf.. 75. Y. Mochinaga, and T. Uzuka, "Development of single phase feeding power conditioner for Shinkansen Depots," Quarterly Report of RTRI, vol. 41, pp. 154- 158, 2000. 76. Y. Mochinaga, Y. koyama, and T. Takwuchi, "Insulation level reduction of three- winding Scott connected transfromer on Shinkansen at feeder circuit," IEE Japan Trans. Industry Applications, vol. 110, pp. 570-579, 1990. 77. S. H. Case, A. M. Hayes, J. H. McNeil, R. E. Morrison, and E. E. Riches, "Electrification of Taiwan main-line railway from Keelung to Kaohsiung," IEE Proc. Electric Power Applications, vol. 130, pp. 289-301, 1983. 78. T. H. Chen, "Comparison of Scott and Leblanc transformers for supplying unbalanced electric railway demands," Elsevier Electric Power Systems Research, vol. 28, 1994. 79. T. H. Chen and C. Jeng-Tyan. "Optimal phase arrangement of distribution transformers connected to a primary feeder for system unbalance improvement and loss reduction using a genetic algorithm," in Proc. 1999 IEEE Power Industry Computer Applications Int. Conf.. 80. V. Levi, M. Kay, M. Attree, and I. Povey, "Design of low voltage networks for premises with small scale embedded generators," in Proc. 2005 Electricity Distribution Int. Conf. and Exhibition. 81. S. C. Vegunta, P. Twomey, and D. Randles, "Impact of PV and load penetration on LV network voltages and unbalance and potential solutions," in Proc. 2013 Electricity Distribution Int. Conf. and Exhibition. 82. F. Shahnia, A. Ghosh, G. Ledwich, and F. Zare, "Voltage Unbalance reduction in low voltage distribution networks with rooftop PVs," in Proc. 2010 Universities Power Engineering Conf.. 83. N. Efkarpidis, C. G., T. Wijnhoven, D. V. Dommelen, T. D. Rybel and J. Driesen, "Technical assessment of on-load tap changers in Flemish LV distribution grids," in Proc. 2013 int. workshop in integration of Solar Power into Power Systems. 84. "FITformer REG The adaptable distribution transformer," Siemens. [Online]. Ava ilable: http://www.energy.siemens.com/mx/pool/hq/power-transmission/Transfor mers/Distribution%20Transformers/fitformer-reg/fit-former-reg-adaptable-distrib ution-transformer.pdf 85. "The system solution for votlage regulated distribution transformers," M.R., 2012. [Online]. Available: http://www.reinhausen.com/XparoDownload.ashx?raid=763 16 86. P. Kadurek, J. F. G. Cobben and W. L. Kling, "Smart MV/LV transformer for future grids," in Proc. 2010 Power Electronics Electrical Drives Automation and Motion Int. Symposium. 87. J. Dixon, L. Moran, J. Rodriguez, and R. Domke, "Reactive power compensation technologies: state-of-the-art review," in IEEE Proc., vol. 93, pp. 2144-2164, 2005. 88. C. Jen-Hung, L. Wei-Jen and C. Mo-Shing, "Using a static VAr compensator to balance a distribution system," in Proc. 1996 IEEE Industry Applications Conf..

Page | 200 References

89. S. H. Hosseini and A. Nazarloo, "Voltage unbalance reduction by distribution static compensator in a case study network," in Proc. 2011 Electrical and Electronics Engineering Int. Conf.. 90. N. Efkarpidis, T. Wijnhoven, C. González, and T. De Rybel and J. Driesen, "Coordinated voltage control scheme for Flemish LV distribution grids utilizing OLTC transformers and D-STATCOM’s," in Proc. 2014 IET 12th international Conference on development in power system protection. 91. V. B. Bhavaraju and P. N. Enjeti, "An active line conditioner to balance voltages in a three-phase system," IEEE Trans. Industry Applications, vol. 32, pp. 287-292, 1996. 92. A. Campos, G. Joos, P. D. Ziogas, and J. F. Lindsay, "Analysis and design of a series voltage unbalance compensator based on a three-phase VSI operating with unbalanced switching functions," IEEE Trans. Power Electronics, vol. 9, pp. 269- 274, 1994. 93. D. M. Vilathgamuwa, A. A. D. R. Perera, and S.S. Choi, "Voltage sag compensation with energy optimized dynamic voltage restorer," IEEE Trans. on Power Delivery, vol. 18, pp. 928-936, 2003. 94. D. V. Nicolae, M. W. Siti, and A. A. Jimoh, "LV self balancing distribution network reconfiguration for minimum losses," in Proc. 2009 IEEE PowerTech Conf.. 95. F. Shahnia, P. Wolfs and A. Ghosh, "Voltage unbalance reduction in low voltage feeders by dynamic switching of residential customers among three phases," in Proc. 2013 IEEE PES General Meeting. 96. M. Singh, V. Khadkikar, A. Chandra, and R. K. Varma, "Grid interconnection of renewable energy sources at the distribution level with power-quality improvement features," IEEE Trans. Power Delivery, vol. 26, pp. 307-315, 2011. 97. K. H. Chua, Y. S. Lim, J. Wong, P. Taylor, E. Morris, and S. Morris, "Voltage unbalance mitigaion in low voltage distribution networks with Photovoltaic systems," Journal of Electronic Sicence and Technoloy, vol. 10, 2012. 98. T. Kaipia, P. Peltoniemi, J. Lassila, P. Salonen, and J. Partanen, "Power electronics in smartgrids - impact on power system reliability," in Proc. 2008 IET-CIRED SmartGrids for Distribution. 99. D. Dohnal, "On-load tap-changers for power transformers a technical digest," M. R. publication. [Online]. Available: http://promad-srl.ro/index.php?option=com_k 2&view=item&task=download&id=69_cead4348ccbeeb756cf3bf9e6e014984&It emid=140 100. M. R. Salem, L.A. Talat and H.M. Soliman, "Voltage control by tap-changing transformers for a radial distribution network," IEE Generation, Transmission and Distribution, vol. 144, pp. 517-520, 1997. 101. M. G. Say, Performance and design of alternating current machines. London: Sir Isaac Pitman & Sons, Ltd, 1958, pp. 66-68, 72-77. 102. M. Thomson, "Automatic voltage control relays and embedded generation," Power Engineering Journal, vol. 14, pp. 71-76, 2000. 103. J. H. Harlow. “Load tap changing control,” 1996. [Onine]. Available: https://ww w.beckwithelectric.com/docs/tech-papers/loadcntl.pdf 104. Power Transformers Part 21: Standard Requirements, Terminology, and Test Code for Step-Voltage Regulators, IEC 60076-21:2011, IEEE Std C57.15, 2011. 105. L. H. Hill, "Transformer tap changing under load," Journal of the A.I.E.E., vol. 46, pp. 1214-1221, 1927.

Page | 201 References

106. M. T. Bishop, J.D. Foster and D.A. Down, "The application of single-phase voltage regulators on three-phase distribution systems," in Proc. 1994 Rural Electric Power Conf.. 107. O. I. Elgerd, Electric energy systems theory: an introduction. New Delhi: Tata McGraw-Hill, 1975, pp. 138-142. 108. A. Al-Riyami, K. Burt, P. Dyer, P. Pretlove, G. Manhangwe, and S. Georgiopoulos, "The protection of a quadrature-booster connected on the distribution network," in Proc. 2014 IET Developments in Power System Protection Int. Conf.. 109. "Flexible Plug and Play Quadrature-booster Report," UK Power Network, 2013. [Online]. Available：http://innovation.ukpowernetworks.co.uk/innovation/en/Pro jects/tier-2-projects/Flexible-Plug-and-Play-%28FPP%29/Project-Documents/Qu ad+Booster+Report+28+Oct_FINAL.pdf 110. M. O. Faruque and V. Dinavahi, "A tap-changing algorithm for the implementation of "Sen" Transformer," IEEE Trans. Power Delivery, vol. 22, pp. 1750-1757, 2007. 111. J. C. Brittain, "Charles F. Scott: a pioneer in electrical power engineering," IEEE Industry Applications Magazine, vol. 8, pp. 6-8, 2002. 112. (Webpage) https://hvdc.ca/pscad/, Manitoba HVDC Research Centre, 2014. 113. O. Anaya-Lara and E. Acha, "Modeling and analysis of custom power systems by PSCAD/EMTDC," IEEE Trans. Power Delivery, vol. 17, pp. 266-272, 2002. 114. "User's guide on the use of PSCAD," Manitoba HVDC Research Centre, 2005. 115. S. Ingram, S. Probert and K. Jackson, "The impact of small scale embedded gener ation on the operating parameters of distribution networks," Department of Trade and Industry (DTI), UK, 2003. [Online]. Available: http://webarchive.nationalarc hives.gov.uk/20100919181607/http://www.ensg.gov.uk/assets/22_01_2004_phas e1b_report_v10b_web_site_final.pdf 116. "Limits for voltage unbalance in the electricty supply system," Regulation and Supervision Bureau for the Water and Electricity Sector of the Emirate of Abu Dhabi, 2005. [Online]. Available: http://www.aadc.ae/img/bfc668e4-34b2-41e9- b417-111c607d94ac.pdf 117. J. Faiz and B. Siahkolah, "Implementation of a low-power electronic tap-changer in transformers," IET Electric Power Applications, vol. 2, pp. 362-373, 2008. 118. G. R. Chandra Mouli, P. Bauer, T. Wijekoon, A. Panosyan, and E. M. Baerthlein, "Design of a Power Electronic Assisted OLTC for Grid Voltage Regulation," IEEE Trans. Power Delivery, vol. PP, 2014. 119. J. Taylor, A. Maitra, M. Alexander, D. Brooks, and M. Duvall, "Evaluation of the impact of plug-in electric vehicle loading on distribution system operations," in Proc. 2009 IEEE PES General Meeting. 120. P. Papadopoulos, S. Skarvelis-Kazakos, I. Grau, L. M. Cipcigan, and N. Jenkins,"Electric vehicles' impact on british distribution networks," IET Electrical Systems in Transportation, vol. 2, pp. 91-102, 2012. 121. K. Clement-Nyns, E. Haesen and J. Driesen, "The impact of charging plug-in hybrid electric vehicles on a residential distribution grid," IEEE Trans. Power Systems, vol. 25, pp. 371-380, 2010. 122. F. Pieltain, T. G. S. Roman, R. Cossent, C. M. Domingo, and P. Frias, "Assessment of the impact of plug-in electric vehicles on distribution networks," IEEE Trans. Power Systems, vol. 26, pp. 206-213, 2011. 123. R. J. Ferraro, J. H. Barnett and C. R. Furman, "EPRI/TVA pilot electric vehicle demonstration: Update 1980," in Proc. 1980 IEEE Vehicular Technology Conf..

Page | 202 References

124. S. Rahman and G.B. Shrestha, "An investigation into the impact of electric vehicle load on the electric utility distribution system," IEEE Trans. Power Delivery, vol. 8, pp. 591-597, 1993. 125. G. A. Putrus, P. Suwanapingkarl, D. Johnston, E. C. Bentley, and M. Narayana, "Impact of electric vehicles on power distribution networks," in Proc. 2009 IEEE Vehicle Power and Propulsion Conf.. 126. S. Shengnan, M. Pipattanasomporn, and S. Rahman, "Challenges of PHEV penetration to the residential distribution network," in Proc. 2009 IEEE PES General Meeting. 127. P. Richardson, D. Flynn, and A. Keane, "Impact assessment of varying penetrations of electric vehicles on low voltage distribution systems," in Proc. 2010 IEEE PES General Meeting. 128. H. Sikai, and D. Infield, "The impact of domestic Plug-in Hybrid Electric Vehicles on power distribution system loads," in Proc. 2010 Power System Technology Int. Conf.. 129. D. H. Kinloch, J. M. Parmelee, F. E. Wicks, M. Becker, and S. Yerazunis, "Impacts of Solar Heating Options Upon Electric Power Systems," IEEE Trans. Power Apparatus and Systems, vol. PAS-101, pp. 1271-1280, 1982. 130. P. Boait, and A. Stafford, "Electrical load characteristics of domestic heat pumps and scope for demand side management," in Proc. 2011 21st Int. Conf. on Electricity Distribution. 131. J. Campillo, F. Wallin, I. Vassileva, and E. Dahlquist, "Electricity demand impact from increased use of ground sourced heat pumps," in Proc. 2012 IEEE PES Innovative Smart Grid Technologies Conf.. 132. M. Akmal, B. Fox, D. J. Morrow, and T. Littler, "Impact of high penetration of heat pumps on low voltage distribution networks," in Proc. 2011 IEEE PowerTech Conf.. 133. P. Mancarella, G. Chin Kim, and G. Strbac., "Evaluation of the impact of electric heat pumps and distributed CHP on LV networks," in Proc. 2011 IEEE PowerTech Conf.. 134. J. Yi, P. Wang, P. C. Taylor, P. J. Davison, P. F. Lyons, S. Brown, and D. Roberts, "Distribution network voltage control using energy storage and demand side response," in Proc. 2012 IEEE PES Innovative Smart Grid Technologies Int. Conf.. 135. N. Hadjsaid, and J. C. Sabonnadiere, Electrical Distribution Networks. London: ISTE Ltd, 2011, pp. 333-343. 136. M. Panteli, "Analysis and preventation of power system blackouts," PhD thesis, in School of Electrical and Electronic Engineering, University of Manchester, 2013. 137. "Global Analysis of the Electric Vehicle Lithium-ion Batteries Chemicals and Materials Market," Frost & Sullivan, 2014. 138. F. Marra, G. Y. Yang, C. Traholt, C. N. Rasmussen, and Y. Shi, "Demand profile study of battery electric vehicle under different charging options," in Proc. 2012 IEEE PES General Meeting. 139. P. Zhang, C. Zhou, B. G. Stewart, D. M. Hepburn, "A methodology for optimization of power systems demand due to electric vehicle charging load," IEEE Trans. Power Systems, vol. 27, pp. 1628-1636, 2012. 140. Electric vehicle conductive charging system, British Std. EN 61851-1, 2011. 141. "EV charging connection," Ratio Electric, 2012. [Online]. Available: http://www. ratio.nl/files/Ratio_EVcharging.pdf

Page | 203 References

142. Plugs, socket-outlets, vehicle connectors and vehicle inlets — Conductive charging of electric vehicles, British Std. EN 62196-1, 2012. 143. Guide to electric vehicle infrastructure. 2012, Association for the British Electrotechnical Industry （BEMA）. 144. "National Travel Survey statistics," Department for Transport, 2012. [Online]. Av ailable: https://www.gov.uk/government/uploads/system/uploads/attachment_data /file/243957/nts2012-01.pdf 145. R. Herman, and C.T. Gaunt, "A practical probabilistic design procedure for LV residential distribution systems," IEEE Trans. Power Delivery, vol. 23, pp. 2247- 2254, 2008. 146. MATLAB. Available from: http://www.mathworks.co.uk/products/matlab/. 147. (Webpage) http://smartgrid.epri.com/SimulationTool.aspx, Electric power Research Institute, 2011. 148. I. Richardson, M. Thomson, D. Infield, and C. Clifford, "Domestic electricity use: A high-resolution energy demand model," Energy and Buildings, vol. 42, pp. 1878-1887, 2010. 149. "Population and Household Estimates for the United Kingdom, March 2011," Office for National Statistics, 2013. 150. A. Navarro, L.F. Ochoa, and D. Randles, "Monte Carlo-based assessment of PV impacts on real UK low voltage networks," in Proc. 2013 IEEE PES General Meeting. 151. Demand profiles generators. [Online]. Available: http://www.esru.strath.ac.uk/Ea ndE/Web_sites/06-07/Carbon_neutral/tools%20folder/Demand%20Profile%20Ge nerators.htm. 152. "United Kingdom housing energy fact file," Department of Energy and Climate C hange, 2012. [Online]. Available: https://www.gov.uk/government/uploads/syste m/uploads/attachment_data/file/201167/uk_housing_fact_file_2012.pdf 153. I. Lampropoulos, W. L. Kling, P. F. Ribeiro and J. van den Berg, "History of demand side management and classification of demand response control schemes," in Proc. 2013 IEEE PES General Meeting. 154. P. O. Steiner, "Peak loads and efficient pricing," The Quarterly Journal of Economics, vol. 71, 1957. 155. "Assessment of demand side response and advanced metering," Federal Energy R egulatory Commiccion, 2012. [Online]. Available: http://www.ferc.gov/legal/staff -reports/12-20-12-demand-response.pdf 156. "Demand side response," Ofgem. [Online]. Available: https://www.ofgem.gov.uk /ofgem-publications/57026/dsr-150710.pdf 157. E. Sortomme, M. Hindi, S. MacPherson, and S. S. Venkata, "Coordinated Charging of Plug-In Hybrid Electric Vehicles to Minimize Distribution System Losses," IEEE Trans. Smart Grid, vol. 2, pp. 198-205, 2011. 158. P. Richardson, D. Flynn and A. Keane, "Optimal charging of electric vehicles in low-voltage distribution systems," IEEE Trans. Power Systems, vol. 27, pp. 268- 279, 2012. 159. Z. Kan and C. Lin, "Randomized PHEV charging under distribution grid constraints," IEEE Trans. Smart Grid, vol. 5, pp. 879-887, 2014. 160. D. J. Aigner, "The residential electricty time-of-use pricing experiments: what have we learned?" [Online] Available: http://www.nber.org/chapters/c8372.pdf 161. "Demand Side Response in the domestic sector- a literature review of major trials, " Department of Energy and Climate Change, UK, 2012. [Online]. Available: http

Page | 204 References

s://www.gov.uk/government/uploads/system/uploads/attachment_data/file/48552/ 5756-demand-side-response-in-the-domestic-sector-a-lit.pdf 162. F. Demailly, A. Even and N. Oliver, "Influence of embedded generation on voltage limits in LV systems - statistical approach," in Proc. 2003 17th Int. Conf. on Electricty Distribution.

Page | 205 Appendix A

APPENDIX A: Network model data

A.1 The simplified LV feeder

The information given here relates to the simplified LV feeder illustrated in Fig. 4-2, which is used in the simulation study presented in section 4.3.

Table A-1 Model data of the simplified LV feeder 11kV voltage source 3Φ ideal voltage source, f = 50Hz

11/0.433kV substation 500kVA, delta-wye-ground connected

0.4kV feeder 3Φ four-wire LV cable, 300m, (z = 0.1+0.073) Ω/km for phase wire; (z = 0.164+j0.011) Ω/km for neutral wire

A.2 The typical UK distribution network

The typical UK distribution network shown in Fig. 4-5 is adopted in the simulation studies presented in section 4.4 and Chapter 6. Table A-2 shows the network model data used in the simulation study presented in section 4.4.

Table A-2 Model data of the typical UK distribution network 33kV voltage source 3Φ ideal voltage source, f= 50Hz

33/11.5kV substation 15MVA, YY0, X = 10%, X/R = 20 Supplying six 11kV feeders

11kV feeder 3Φ three-wire radial feeder with the length of 4km Supplying eight evenly distributed 11/0.433kV transformers 300mm² XLPE cable (z = 0.08+j0.087) Ω/km for phase wire

11/0.433kV substation 500kVA, ΔYn, Z = 4%. X/R = 1.5 Supplying four 0.4kV feeders

0.4kV feeder 3Φ four-wire radial feeder with the length of 300m Supplying 20 evenly distributed customers per phase Consisting of a 150m 300mm² cable and a 150m 95mm² cable 300mm² LV cable (z = 0.1+0.073) Ω/km for phase wire; (z = 0.164+j0.011) Ω/km for neutral wire 95mm² LV cable (z = 0.32+0.075) Ω/km for phase wire; (z = 0.32+j0.016) Ω/km for neutral wire

Page | 206 Appendix A

The network model data shown in Table A-2 is also adopted in the simulation study presented in Chapter 6 except for the 11kV feeder and the 0.4kV feeder. The 11kV feeder used in Chapter 6 consists of a 2km 185mm2 cable and a 2km 95mm2 cable; the 0.4kV feeder used in Chapter 6 consists of a 150m 185mm2 cable and a 150m 95mm2 cable, and the data of these cables are listed in Table A-3.

Table A-3 Characteristics of the 11kV cables and the 0.4kV cables used in Chapter 6 11kV 185mm2 cable 3Φ, triplex cable, XLPE insulation Stranded aluminium conductor Positive sequence impedance: (Z1= 0.22+j0.102) Ω/km Zero sequence impedance: (Z0 = 0.53+j0.102) Ω/km

11kV 95mm2 cable 3Φ, triplex cable, XLPE insulation Stranded aluminium conductor Positive sequence impedance: (Z1= 0.415+j0.112) Ω/km Zero sequence impedance: (Z0 = 0.988+j0.236) Ω/km

0.4kV 185mm2 cable 3Φ, 3core sector shaped, extruded XLPE insulation Solid aluminium conductors, plain copper waveform cables Positive sequence impedance: (Z1= 0.164+j0.074) Ω/km Zero sequence impedance: (Z0 = 0.656+j0.05) Ω/km

0.4kV 95mm2 cable 3Φ, 3core sector shaped, extruded XLPE insulation Solid aluminium conductors, plain copper waveform cables Positive sequence impedance: (Z1= 0.32+j0.0735) Ω/km Zero sequence impedance: (Z0 = 1.28+j0.047) Ω/km

Page | 207 Appendix B

APPENDIX B: Model of ST based voltage balancing system

This section introduces the model of the ST based voltage balancing system. Please note in the following T1 and T2 refer to the teaser and main transformers of Scott transformer I respectively; T3 and T4 refer to the teaser and main transformers of Scott transformer II respectively, and T5 and T6 refer the shunt regulating transformer and the series transformer in the phase regulating system separately.

B.1 Model of ST based voltage balancing system in PSCAD/EMTDC Table B-1 shows the characteristics of each transformer in the PSCAD/EMTDC model of the ST based voltage balancing system.

Table B-1 Characteristics of PSCAD/EMTC model of ST based voltage balancing system T1 1Φ, 3-winding transformer (two windings in the primary), 0.254kV/0.127kV/0.23kV

T2 1Φ, 3-winding transformer (two windings in the primary), 0.22kV/0.22kV/0.23kV

T3 1Φ, 3-winding transformer (two windings in the secondary), 0.23kV/0.254kV/0.127kV

T4 1Φ, 3-winding transformer (two windings in the secondary), 0.23kV/0.22kV/0.22kV

T5 1Φ, 2-winding transformer, 0.23kV/0.23kV

T6 1Φ, 2-winding transformer, 0.023kV/0.23kV

B.2 Model of ST based voltage balancing system in OpenDSS The program code modelling the ST based voltage balancing system in OpenDSS is presented below.

! Configuration of T1 New transformer.T1 phases=1 windings=3 buses=[15.1.6 15.6.5 22.1.4] kvs=[0.23 0.115 0.23] kvas=[100 100 100] %Rs=0 XHL=1 XHT=1 XLT=1

Page | 208 Appendix B

! Configuration of T2 New transformer.T2 phases=1 windings=3 buses=[15.2.5 15.5.3 22.2.4] kvs=[0.1992 0.1992 0.23] kvas=[100 100 100] %Rs=0 XHL=1 XHT=1 XLT=1 ! Configuration of T6 New transformer.T6 phases=1 windings=2 buses=[22.1.6 22.3.5] kvs=[0.0115 0.23] kvas=[50 50] %Rs=0 XHL=1 XHT=1 XLT=1 ! Configuration of T5 New transformer.T5 phases=1 windings=2 ~ wdg=1 bus=22.2.4 kv=0.23 kva=50 ! this is for the case with tap size 1.15V ~ wdg=2 bus=22.3.5 kv=0.23 kva=50 NumTaps=20 Maxtap=1.0 Mintap=-1.0 Tap=0 ! this is for the case with tap size 4.6V !~ wdg=2 bus=22.3.5 kv=0.23 kva=50 NumTaps=4 Maxtap=0.8 Mintap=-0.8 Tap=0 ! this is for the case with tap size 9.2V !~ wdg=2 bus=22.3.5 kv=0.23 kva=50 NumTaps=2 Maxtap=0.8 Mintap=-0.8 Tap=0 ~ %Rs=0 XHL=1 XHT=1 XLT=1 ! Configuration of T3 New transformer.T3 phases=1 windings=3 ~ wdg=1 bus=22.6.4 kv=0.23 kva=100 NumTaps=64 Maxtap=1.1 Mintap=0.9 Tap=1 ~ wdg=2 bus=23.1.4 kv=0.23 kva=100 ~ wdg=3 bus=23.4.5 kv=0.115 kva=100 ~ %Rs=0 XHL=1 XHT=1 XLT=1 ! Configuration of T4 New transformer.T4 phases=1 windings=3 ~ wdg=1 bus=22.2.4 kv=0.23 kva=100 NumTaps=64 Maxtap=1.1 Mintap=0.9 Tap=1 ~ wdg=2 bus=23.2.5 kv=0.1992 kva=100 ~ wdg=3 bus=23.5.3 kv=0.1992 kva=100 ~ %Rs=0 XHL=1 XHT=1 XLT=1

Page | 209 Appendix C

APPENDIX C: Hardware used in the physical test

This section introduces the details of the hardware used in the physical test of ST based balancing system.

C.1 ST based voltage balancing system Please note all the transformers presented below were established by rewinding existing commercial toroidal transformers. The nominal value of turns per volt is 10.98. In this section, TR1 and TR2 refer to the teaser and main transformers of ST-I respectively; TR3 and TR4 refer to the teaser and main transformers in ST-II respectively; and TR5 and TR6 refer to the shunt and series transformers in phase regulating system respectively.

 TR 1 Description: 2188 turns/1044 turns, isolation transformer. A 1/3 point exists at its primary winding. Fig. C-1 shows the layout of TR1.

s n r u t V V

2 3 8 . 9 5 . 9 4 4 9 1 9 1

s s n n r r u u t t

4 8 4 8 s 0 1 n 1 2 r u t

0 3 7

Fig. C-1 Layout of TR1

 TR2 Description: 2526 turns/1044 turns, isolation transformer. A 1/2 point exists at its primary winding. Its layout is shown in Fig. C-2.

Page | 210 Appendix C s n r u t V

2 V 3 0 9 6 . 3 2 4 2 1

s s n n r r u u t

6 s 4 2 n 4 r 5 0 u 2 t 1

3 6 2 1

Fig. C-2 Layout of TR2

 TR 3 Description: 1044 turns/2188 turns, isolation transformer. A 1/3 point exists at its secondary side. In the primary side, there are 10 taps above the nominal point and 10 taps below the nominal point, and the tap size is 20 turns/tap. The layout of this transformer is shown in Fig. C-3.

s p a t

0 1 s V n s 3 r p . a u t 9

0 9 1 8 1 V

5 2 s 4 9 n 1 . r 4 u t 9

8 s 8 n r 1 s u 2 n t r

u 4 t

4 0 0 3 1 7

Fig. C-3 Layout of TR3

 TR 4 Description: 1044 turns/2526 turns, isolation transformer. A 1/2 point exists on its secondary side. In the primary side, there are 10 taps above the nominal point and 10 taps below the nominal point, and the tap size is 20 turns/tap. The layout of this transformer is shown in Fig. C-4.

Page | 211 Appendix C s p a t

0 s 1 n s r p u a t t

0 0 1 3 3 V 6 2 2 2

9 1 s . n 4 r 9 u

t s

s n 6 r n 2 r u 5 t u

t 2

4 3 4 6 0 2 1 1

Fig. C-4 Layout of TR4

 TR 5 Description: 2525 turns/1400 turns, isolation transformer. Four taps divides the secondary winding into five equal sub-windings. Its layout is shown in Fig. C-5.

5 ˣ 280 turns

T1 S2

s T2 n

r S1 u t T3 5 2 5 2 T4

Fig. C-5 Layout of TR5

 TR 6 Description: 2525 turns/596 turns, standard isolation transformer. The layout of this transformer is shown in Fig. C-6.

Page | 212 Appendix C

s s n n r r u u t t

5 6 2 9 5 5 2

Fig. C-6 Layout of TR6

Fig. C-7 shows a picture of the established balancing system.

TR 1 TR 6 TR 3

TR 2 TR 5

TR 4

Fig. C-7 Picture of the established balancing system

Page | 213 Appendix C

C.2 Microprocessor based control system

 Analogue input circuit The microprocessor based control system monitors the output 3Φ voltage magnitudes of the balancing system. In the physical test, the RMS values of the monitored voltages are around 90V. However the microprocessor can only accept signals between 0V and 5V. The analogue input circuits were established to interface the microprocessor and the monitored low voltage feeder. Fig. C-8 shows the layout of the designed analogue input circuit.

Operational amplifier: C1 LM741CN - IC INPUT R1 Isolation amplifier: ISO124 3.9MΩ R4 R3 C2 RMS-DC converter: 200KΩ R2 AD637 Buffer in Buffer out NC Vin OUTPUT Common NC Output offset +Vs CS -Vs DEN input RMS output C3 dB output Cav

Fig. C-8 Layout of the designed analogue input circuit

In Fig. C-8, the resistor R1 and R2 make up a voltage divider, which can provide an attenuating ratio of 40:1. The isolation amplifier is used to provide the electrical isolation between the monitored LV feeder and the microprocessor. Following the isolation transformer, R3, R4, C1, C2 and an operational amplifier are connected as a low pass filter to remove the 500kHz carrier ripples from the isolation transformer. At the end of the circuit, an RMS-DC converter is used to output a DC signal, which equals the RMS value of the attenuated AC signal.

Page | 214 Appendix C

Fig. C-9 below shows the connection of the DC power supply of the isolation amplifier, operational amplifier and RMS-DC converter. In this supply circuit, only one 15V DC supply is used. The +/-15V supplies in the right side of Fig. C-9 were used to supply the operational amplifier and the RMS-DC converter.

15V GND

1 2 5 6 7

DCP01515DB DCP01515DB

7 6 5 2 1

0.47uF

Vin

GND Vin V- V+ -15V

Input Output +15V section ISO124 section

V+ V- Vo GND

+15V

-15V Vout

Fig. C-9 DC power supply connection

The main components used in the analogue input circuit are listed in Table C-1.

Table C-1 Main components used in the analogue input circuit

Name Component Isolation amplifier TEXAS INSTRUMENTS - ISO124P - IC Operational amplifier NATIONAL SEMICONDUCTOR - LM741CN - IC DC-DC converter TEXAS INSTRUMENTS - DCP011515DBP-U - IC RMS-DC converter ANALOG DEVICES AD637

Page | 215 Appendix C

 Relay driving circuit Fig. C-10 shows the circuit diagram of the adopted relay driving circuit. A diode is connected in parallel with relay coil to protect the relay contacts. A transistor is used to adjust the current flowing through the relay coil. Moreover, an LED is used to indicate the switch position of the relay.

+12V A

Miniature Diode relay

a +5V Transistor LED Microprocessor R1

Fig. C-10 Layout of the relay circuit

The main components used in the relay driving circuit are listed in Table C-2.

Table C-2 Main components used in the relay driving circuit

Name Component Miniature relay TE CONNECTIVITY - IME06GR Transistor MULTICOMP - BC337 LED AGILENT – HLMP-1600

Moreover, the microprocessor used in this study is ‘MICROCHIP PIC18F4520’.

Fig. C-11 shows a picture of the established microprocessor based control system.

Page | 216 Appendix C

Analogue input circuits

Relay driving circuits Interfacing board Microprocessors

Fig. C-11 Picture of the established control system

Page | 217 Appendix D

APPENDIX D: Investment costs

A brief analysis of the investment costs of ST based balancing system, LV cable replacement and power electronics based AC-DC-AC solution is presented in this section. All these solutions are used to solve the problem of voltage imbalance in the LV feeder F1 introduced in section 4.4.2. The ST based balancing system and power electrics based AC-DC-AC solution are assumed to be installed at the LCP of 120m and they are designed to have a power delivery capacity of 100kVA.

 ST based voltage balancing system The cost of transformer is closely related to its capacity. The investment cost of a 11kV/0.4kV 500kVA substation in suburban areas, including the cost of transformer, RMU and LV switchgear and the installation cost, is around £20k. The total capacity of the transformers used in the 100kVA ST based balancing system is 220kVA, assuming the capacities of the shunt and series transformers are both 10kVA. According to the transformer capacity, the investment cost of a 100kVA ST based balancing system is estimated to be £8.8k.

 LV cable replacement The cost of 3Φ LV feeder in suburban areas, including the cable cost and the installation cost, is around £30k/km. The investment cost of replacing 300m LV cable in suburban areas is estimated to be £9k.

 Power electronics based AC-DC-AC solution The cost of 100kVA AC-DC rectifier is around £3k, and the cost of 100kVA DC-AC inverter is around £22k. Therefore, the cost of 100kVA AC-DC-AC system is £25k, which excludes the installation fee.

Based on the above analysis, the ST based balancing system has the lowest investment cost. However, a detailed financial cost analysis of the above solutions over the lifetime should be carried out in future work. Besides the investment cost, the cost related to the power losses and the maintenance cost should be considered.

Page | 218 Appendix E

APPENDIX E: MC stopping criteria

MC procedure produces a lot of samples and there is no guarantee that the deviation error of the sample mean value from the true mean value will decrease if a few more simulations are considered. However, the error bound does decrease as more simulations are carried out. Therefore, it is important to decide the necessary simulation times for the overall MC procedure. The convergence or stopping criteria is used to determine simulation times to obtain acceptable accurate results. The degree of confidence C is used as the convergence criteria in this study:

X̅+ L C = P(X̅ − L ≤ μ ≤ X̅ + L) = g(x)dx (E-1) ∫X̅−L Where,

σ L = Za (E-2) ⁄2 √n and g(x) is the Gaussian distribution, X̅ is the sample mean, σ is the sample standard deviation, L is the confidence interval, μ is the true population mean and n is the number of samples [136]. The value Za representing the point on the standard normal ⁄2 distribution curve such that the probability of observing a value greater than Za is equal ⁄2 to p is known as the upper critical value of the standard normal distribution. Za and p ⁄2 are closely related to the confidence level C: for a confidence interval with confidence level C, the value p is equal to (1-C)/2 as shown in Fig. E-1.

Area = C

Area = (1- C)/2 Area = (1- C)/2

 Za 2 Za 2

Fig. E-1 Relationship between confidential level C and Za ⁄2

Page | 219 Appendix E

The degree of confidence C expresses the probability of the sample mean lying within the interval μ ± L. For the given confidence interval and level that allows us to say that the results meet the specified accuracy, the required number of simulation can be obtained from:

2 2 σ Za n = ⁄2 (E-3) L2

Since the sample standard deviation is unknown, it is not possible to calculate n before the simulation starts. The above equation indicates that the sample size increases with the confidence interval, but is inversely proportional to the square of the confidence interval [136].

σ In the MC studies presented in this thesis, the values of all the investigated parameters √n are monitored; these values all have decreased to a low value when the MC trial is executed for 500 times, indicating that the sampled means of the investigated parameters all have a relatively high degree of confidence in the specified confidence interval. For example, the 3Φ voltages, U2 and U0 at the feeder end, obtained from the study investigating the impact of 54EVs and 54HPs on the LV feeder voltages in the non- balanced scenario, have high confidence levels in the specified confidence intervals as shown in Table E-1.

Table E-1 Confidence levels and intervals of voltages, U2 and U0

σ Investigated Confidence Confidence parameters √n interval L (V) level C Va 0.09 0.5 > 99.99% Vb 0.214 0.5 98.06% Vc 0.212 0.5 98.18% U2 0.016% 0.1% >99.99% U0 0.047% 0.1% 96.44%

Page | 220 Appendix F

APPENDIX F: Current available EVs

F.1 List of BEVs

EV name EV characteristics Battery Charging properties MyCar Top Speed: 40mph Lead Gel Deep Charge time between 6 and (NICE: 2008) Range: 40 miles Cycle: 8 hours 8.5kWh Ze-0 MPV Top speed: 50mph AGM sealed lead Charge time between 5 and (NICE: 2008) Range: 45 miles acid-Lithium-ion 10 hours option: 18kWh Quiet Car 2 Top speed: 80km/h Lithium-ion battery: 6 hours for a full charge (2008) (50mph) 15.5kWh from 13A socket with 3kW Range: 60 miles (4.8 miles/kWh) battery charger Think City Range: 99 miles Lithium-ion battery: 9.5-10 hours for a full (Think Global: Top speed: 68mph 23 kWh charge from a standard 2008 -2012) (5.4 miles/kWh) 230V, 10 or 16A supply Mega City Top speed: 40mph 10kWh 8 to 10 hours for a full (Aixam: 2009) Range: 60km charge from any 220V and 16A socket Nissan Leaf Top speed: 93mph Lithium ion battery: 8 hours for a full charge (2010) Range: 73 to 75 24kWh from 220/240V outlet miles (3.9 miles/kWh) The MIA (MIA Top speed: 66mph Lithium Ion 5 hours for a full charge electric: 2011) Range: 70 to 80 Phosphate from any 220V and 16A miles (LiFePO4): 12kWh socket Bolloré Bluecar Top speed: 81mph Lithium polymer (2011) Range: 160 miles (LMP): 30kWh Renault Fluence Top speed: 84 mph Lithium-ion battery: 10-12 hours for a full Z.E. Range: 115 miles 22kWh charge from a standard (2011) (6.5 miles/kWh) electric outlet; fast charge options are available. Tazzari EM1 Top speed: 62mph Latest generation 8 hours for a full charge (2012) Range: 87 miles lithium: 15kWh from an ordinary socket; (estimated) MULTIFAST options are available. BMW i3 Top speed: 93mph Lithium-ion battery About 4 hours for a full (expected in Range: 80 to 100 22kWh charge from 240V charging 2013) miles (5.7 miles/kWh) unit Lightning GT Top speed: 120mph Lithium-titanate 15 hours for a full charge (2013) Range: 150 miles battery: 22kWh from any standard 13A socket; fast charge options are available BYD e6 Top speed: 87mph Lithium battery: Variable charge options are (testing, china) Range: 186 mile 48kWh available

Page | 221 Appendix F

F.2 List of PHEVs

EV name EV characteristics Battery Charging properties BYD F3DM All-electric range: 40- Lithium ion 7 hours for a full charge (2008-2013) 60 miles phosphate battery: from the 220V socket in 16kWh the USA Chevrolet Volt Top electric only speed: Lithium-ion battery: 3 hours for a full charge (General 100mph 16kWh from a standard 230V Motors: 2011) All electric range: 25 to (3.9 miles/kWh) outlet 50 miles Fisker Karma All-electric range: 50 Lithium-ion battery: Charging time varies (2011) miles 20.1kWh between 6 and 14 hours (3.1 miles/kWh) Toyota Prius Top electric only speed: Lithium-ion battery: 1.5 hours for a full Plug-in Hybrid 62mph 4.4kWh charge from a standard (2012) All-electric range: 11 (3.1 miles/kWh) European 230V miles household outlet Mitsubishi Top electric speed: Lithium-ion battery: 4 hours for a full charge Outlander 75mph 12kWh from 250V/15A socket; PHEV All-electric range: 37 (3.9 miles/kWh) fast charge options are (2013) miles available. Ford Fusion Top electric speed: Lithium-ion battery: 2.5 hours for a full Energi (2013) 85mph 7.6kWh charge from 240V socket All-electric range: 21 (3.5 miles/kWh) miles V60 PHEV Top electric speed: Lithium-ion battery: Variable charge options (Volvo: 2013) 70mph 12 kWh are available All electric range: 31 (3.2 miles/kWh) miles

Page | 222