superconducting fault current limiter with integrat …

SUPERCONDUCTING FAULT CURRENT LIMITER WITH INTEGRAT ED

VACUUM INTERRUPTER

A thesis submitted to The University of Manchester for the degree of

Doctor of Philosophy

in the Faculty of Engineering and Physical Sciences

2012

Xiaoze Pei

SCHOOL OF ELECTRICAL AND ELECTRONIC ENGINEERING

Content

2

Content

List of Figures..................................................................................................................7

List of Tables .................................................................................................................14

Nomenclature.................................................................................................................15

List of Abbreviations ....................................................................................................19

Abstract..........................................................................................................................21

Declaration.....................................................................................................................23

Copyright Statement.....................................................................................................24

Acknowledgments .........................................................................................................25

The Author.....................................................................................................................26

1 Introduction.......................................................................................................27

1.1 Background .....................................................................................................27

1.2 Project motivation and context........................................................................29

1.3 Organisation of thesis......................................................................................32

2 Literature Review .............................................................................................34

2.1 Superconductivity ...........................................................................................34

2.1.1 Elementary properties of superconductors..........................................34

2.1.2 Materials..............................................................................................39

2.1.3 Summary .............................................................................................42

2.2 Fault current limitation....................................................................................42

2.2.1 Introduction.........................................................................................42

2.2.2 Operating principles of fault current limitation ..................................43

2.2.3 Application of fault current limiters....................................................44

2.2.4 Conventional methods.........................................................................45

2.2.5 Non-superconducting fault current limiters ........................................46

2.2.6 Superconducting fault current limiters................................................50

2.2.7 Summary .............................................................................................62

Content

3

2.3 Vacuum interrupter .........................................................................................64

2.3.1 Introduction.........................................................................................64

2.3.2 Vacuum interrupter structure ..............................................................65

2.3.3 Vacuum arc [94]..................................................................................66

2.3.4 Vacuum breakdown ............................................................................67

2.3.5 Current breaking in a vacuum.............................................................69

2.3.6 Practical design for high current levels...............................................70

2.3.7 Main application fields........................................................................71

2.3.8 Summary .............................................................................................72

2.4 Vacuum interrupter actuator ...........................................................................73

2.4.1 Spring actuator ....................................................................................73

2.4.2 Solenoid actuator.................................................................................73

2.4.3 Permanent magnetic actuator ..............................................................74

2.4.4 Voice-coil type actuator ......................................................................75

2.4.5 Summary .............................................................................................76

2.5 Conclusions.....................................................................................................77

3 SFCL Coil and Experimental Test Rig...........................................................78

3.1 Introduction.....................................................................................................78

3.2 Coil former......................................................................................................78

3.3 Coil manufacturing process.............................................................................80

3.3.1 MgB2 wire current connections and winding process.........................80

3.3.2 MgB2 wire heat treatment ...................................................................83

3.4 Instrumentation ...............................................................................................84

3.4.1 Voltage signals ....................................................................................84

3.4.2 Current signals ....................................................................................85

3.4.3 Temperature signals ............................................................................85

3.5 Control ............................................................................................................87

3.5.1 High-current test circuit ......................................................................87

3.5.2 Low-current test circuit .......................................................................88

3.5.3 LabVIEW control programme ............................................................89

3.5.4 Cryostat ...............................................................................................90

3.6 Conclusions.....................................................................................................93

4 Experimental Investigation of SFCL Coils.....................................................94

Content

4

4.1 Introduction.....................................................................................................94

4.2 Single-strand SFCL Coil .................................................................................95

4.2.1 Instrumentation ...................................................................................95

4.2.2 Calibration of BAS16 diode temperature sensors...............................97

4.2.3 Temperature profile.............................................................................98

4.2.4 Frequency sweep.................................................................................99

4.2.5 Quench tests ......................................................................................100

4.2.6 Long duration quench tests ...............................................................109

4.2.7 Simulated fault test............................................................................113

4.2.8 Temperature rise test .........................................................................113

4.2.9 Summary ...........................................................................................114

4.3 Three-strand SFCL Coil................................................................................115

4.3.1 Instrumentation .................................................................................115

4.3.2 Temperature profile...........................................................................116

4.3.3 Frequency sweep...............................................................................117

4.3.4 Current sharing test ...........................................................................119

4.3.5 Quench tests ......................................................................................120

4.3.6 Long duration quench test.................................................................125

4.3.7 Summary ...........................................................................................127

4.4 Conclusions...................................................................................................128

5 Modelling of SFCL Coil .................................................................................129

5.1 Introduction...................................................................................................129

5.2 MATLAB model...........................................................................................130

5.2.1 Model description..............................................................................130

5.2.2 Comparison with short cycle quench tests ........................................134

5.2.3 Comparison with long duration quench tests ....................................136

5.2.4 Summary ...........................................................................................139

5.3 Finite element thermal model........................................................................140


5.3.2 Results and discussion ......................................................................144

5.3.3 Summary ...........................................................................................148

5.4 Prediction for three-second fault test ............................................................148

5.5 Conclusions...................................................................................................150

Content

5

6 Operating Actuator for Vacuum Interrupter .......... ....................................151

6.1 Introduction...................................................................................................151

6.2 Analytical model ...........................................................................................153


6.2.2 Magnetic field in the airgap ..............................................................155

6.2.3 Effect of the actuator coil current on the electromagnetic force.......158

6.3 Finite element model.....................................................................................160


6.3.2 Magnetic field distribution in the airgap...........................................162

6.3.3 Effect of the actuator coil current on the electromagnetic force.......167

6.3.4 Magnetic latch...................................................................................171

6.4 Design of the full operating actuator.............................................................172

6.4.1 Contact popping ................................................................................172

6.4.2 Contact bounce and rebound.............................................................173

6.4.3 Contact welding ................................................................................174

6.5 Design of prototype actuator and interrupter ................................................175

6.6 Construction of prototype operating actuator ...............................................178

6.6.1 Actuator stationary part.....................................................................178

6.6.2 Actuator moving part ........................................................................179

6.6.3 Complete prototype...........................................................................180

6.7 Conclusions...................................................................................................181

7 Design of the Actuator Control Circuit ........................................................182

7.1 Introduction...................................................................................................182

7.2 Control circuit ...............................................................................................182

7.2.1 Topology selection............................................................................182

7.2.2 Components selection .......................................................................183

7.3 Trigger signal and MOSFET drive circuit ....................................................185

7.3.1 Trigger signal circuit .........................................................................185

7.3.2 MOSFET drive circuit.......................................................................188

7.4 Conclusions...................................................................................................191

8 Experimental Investigation of an SFCL Coil with Integrated Vacuum

Interrupter ...................................................................................................................192

8.1 Introduction...................................................................................................192

Content

6

8.2 Testing of the vacuum interrupter operating actuator ...................................192

8.2.1 Actuator magnetic field and static force ...........................................192

8.2.2 Opening the vacuum interrupter at atmospheric pressure.................194

8.2.3 Opening the vacuum interrupter in a vacuum...................................198

8.2.4 Comparison of the tests at atmospheric pressure and in the vacuum202

8.3 Single-strand SFCL coil with and without the vacuum interrupter ..............203

8.3.1 Test rig diagram ................................................................................204

8.3.2 Quench test........................................................................................204

8.3.3 Simulated fault test............................................................................207

8.4 Conclusions...................................................................................................210

9 Conclusions and Further Research...............................................................211

9.1 Conclusions...................................................................................................211

9.2 Further research.............................................................................................214

References ....................................................................................................................215

Appendix A – Materials Data.....................................................................................228

Appendix B – Components for Operating Actuator................................................242

Total word count of thesis: 53732

List of Figures

7

List of Figures

Figure 1.1 Methods of fault current limitation [3].........................................................28

Figure 1.2 Potential applications of FCL devices [3] ....................................................28

Figure 1.3 Proposed resistive SFCL with vacuum interrupter.......................................31

Figure 2.1 Comparison of resistivity of normal conductors and a typical superconductor

[23] ..................................................................................................................................35

Figure 2.2 Exclusion of the magnetic field in a superconductor [25]............................36

Figure 2.3 Applied magnetic field: Type I (left) and Type II (right) superconductor [25]

.........................................................................................................................................37

Figure 2.4 Typical temperature profile of a Type II superconductor [25] .....................37

Figure 2.5 Configuration of SuperPower 2G HTS wire (wire type SCS4050) [30] ......40

Figure 2.6 Typical fault current waveforms with and without fault current limiting ....43

Figure 2.7 Solid-state FCL with a turn-off IGCT ..........................................................47

Figure 2.8 Solid-state FCL with thyristors.....................................................................47

Figure 2.9 Hybrid switching FCL ..................................................................................48

Figure 2.10 Thyristor-controlled resonant FCL.............................................................49

Figure 2.11 Thyristor-controlled series-parallel resonant FCL .....................................49

Figure 2.12 Thyristor-controlled series resonant FCL...................................................50

Figure 2.13 Resistive SFCL [2] .....................................................................................51

Figure 2.14 Bridge type SFCL.......................................................................................55

Figure 2.15 DC biased iron core SFCL..........................................................................57

Figure 2.16 Shielded iron core SFCL ............................................................................59

Figure 2.17 Fault current controller SFCL.....................................................................60

Figure 2.18 Flux-lock type SFCL ..................................................................................61

Figure 2.19 Typical structure of a vacuum interrupter [101].........................................65

Figure 2.20 Typical structure of a permanent magnetic actuator...................................74

Figure 3.1 Former specification (side view) [17]...........................................................79

Figure 3.2 Prototype former...........................................................................................80

Figure 3.3 Cryostat interior ............................................................................................81

Figure 3.4 Clamp used to connect wire to copper braid ................................................82

List of Figures

8

Figure 3.5 Top copper braid current connection............................................................82

Figure 3.6 Coil after heat treatment (left) and soldered joints (right) ............................83

Figure 3.7 Coil ready to be installed in the cryostat (left) and installed in the cryostat

(right)...............................................................................................................................84

Figure 3.8 Temperature monitor (top) and temperature controller (bottom) .................86

Figure 3.9 Cryocon S700 silicon diode temperature probes ..........................................86

Figure 3.10 BAS16 diode temperature sensor soldered onto the MgB2 wire ................87

Figure 3.11 High-current test circuit schematic [15, 17] ...............................................88

Figure 3.12 Low-current test circuit schematic [15, 17]................................................89

Figure 3.13 Screenshot of LabVIEW control programme.............................................90

Figure 3.14 Cryostat system: vacuum cryostat vessel with vacuum pump set (left),

vacuum pump (top middle), gauge (bottom middle) and cryocooler compressor (right)

.........................................................................................................................................91

Figure 3.15 Cryostat top plate........................................................................................92

Figure 4.1 Schematic showing locations of normal voltage taps (left), detailed voltage

taps and BAS16 diode temperature sensors (right).........................................................96

Figure 4.2 Temperature diode calibration curves when cooling down (1st) and warming

up (2nd) ...........................................................................................................................98

Figure 4.3 Temperature profile of the coil .....................................................................99

Figure 4.4 Impedance of the coil with varying frequency at 25K ...............................100

Figure 4.5 Total coil impedance with varying frequency at 25K ................................100

Figure 4.6 Coil response at 34K with a potential peak current of 311A......................102

Figure 4.7 Coil response at 34K with a potential peak current of 311A, highlighting the

point of quench..............................................................................................................102

Figure 4.8 Coil temperature response at 34K with a potential peak current of 311A..103

Figure 4.9 Coil response at 34K with a potential peak current of 372A......................104

Figure 4.10 Coil response at 34K with a potential peak current of 372A, highlighting

the point of quench........................................................................................................104

Figure 4.11 Coil temperature response at 34K with a potential peak current of 372A105

Figure 4.12 Coil response at 32K with a potential peak current of 622A....................105





List of Figures

9




Figure 4.18 Estimated quench currents versus temperature.........................................108

Figure 4.19 Coil current response during a ten-cycle quench test with a potential peak

current of 372A .............................................................................................................110

Figure 4.20 Coil voltage response during a ten-cycle quench test with a potential peak

current of 372A .............................................................................................................110

Figure 4.21 Coil first turn detailed voltage response during a ten-cycle quench test with

a potential peak current of 372A...................................................................................111

Figure 4.22 Coil temperature response during a ten-cycle quench test with a potential

peak current of 372A.....................................................................................................111

Figure 4.23 Coil current response during a fifty-cycle quench test with a potential peak

current of 372A .............................................................................................................112

Figure 4.24 Coil temperature response during a fifty-cycle quench test with a potential

peak current of 372A.....................................................................................................112

Figure 4.25 Coil current response to simulated fault at 34K with a potential peak

current of 316A .............................................................................................................113

Figure 4.26 Coil temperature response during continuous 200Apeak current test at 30K

for one hour ...................................................................................................................114

Figure 4.27 Picture of the three-strand coil (Courtesy of Hyper Tech) .......................115

Figure 4.28 Schematic showing locations of the voltage taps .....................................116

Figure 4.29 Temperature profile of the coil .................................................................117

Figure 4.30 Impedance of the coil with varying frequency at 25K .............................117

Figure 4.31 Total coil impedance with varying frequency at 25K ..............................118

Figure 4.32 Coil response with current flow at 30K....................................................119


Figure 4.34 Each coil strand voltage response of the first and fifth turns at 34K with a

potential peak current of 249A......................................................................................121





List of Figures

10



Figure 4.39 Each coil strand voltage response of the fifth turn at 30K with a potential

peak current of 529A.....................................................................................................124

Figure 4.40 Estimated quench currents versus temperature.........................................125


current of 249A .............................................................................................................126


current of 249A .............................................................................................................126

Figure 4.43 Each coil strand voltage response of the first and fifth turns during a ten-

cycle quench test with a potential peak current of 249A ..............................................127

Figure 5.1 Results comparison for a fault at 34K with a potential peak current of 372A

.......................................................................................................................................135


.......................................................................................................................................135


.......................................................................................................................................136

Figure 5.4 Results comparison for a ten-cycle fault at 34K with a potential peak current

of 372A .........................................................................................................................137

Figure 5.5 Results comparison for a fifty-cycle fault at 34K with a potential peak

current of 372A .............................................................................................................137


of 372A (considering the heat dissipated into the nitrogen) .........................................138


current of 372A (considering the heat dissipated into the nitrogen).............................138

Figure 5.8 Temperature comparison for a fifty-cycle fault at 34K with a potential peak

current of 372A with adiabatic and nitrogen boundary conditions...............................139

Figure 5.9 Flux2D FE model of the geometry .............................................................141

Figure 5.10 Mesh of the Flux2D FE model, showing detail of the coil in the slot......141

Figure 5.11 Instantaneous power loss density in the coil during a one-second fault at

34K with a potential peak current of 372A...................................................................143

Figure 5.12 Time variation of temperature in the centre of the coil for varying distances

from the former .............................................................................................................145

List of Figures

11

Figure 5.13 Temperature profile vertically down the former through the centre of the

coil after a one-second fault, with a gap of 1.5mm from the former ............................145

Figure 5.14 Temperature profile after a one-second fault, with a gap of 1.5mm from the

former............................................................................................................................146

Figure 5.15 Temperature response comparison for a one-second fault at 34K with a

potential peak current of 372A from the MATLAB model and the FE model.............147

Figure 5.16 Temperature profile through the centre of the coil for a one-second fault at

34K with a potential peak current of 372A from the MATLAB model and the FE model

.......................................................................................................................................148

Figure 5.17 Modelled current response for a three-second fault at 34K with a potential

peak current of 372A.....................................................................................................149

Figure 5.18 Temperature response comparison for a three-second fault at 34 K with a

potential peak current of 327A from the MATLAB model and the FE model.............149

Figure 6.1 DVS10CB vacuum interrupter ...................................................................151

Figure 6.2 Geometry of the actuator (the two steel walls in the front are not shown).154

Figure 6.3 Simplified model of the magnetic circuit ...................................................154

Figure 6.4 Actuator permanent magnet operating points on the 2nd quadrant B-H

characteristic (demagnetisation curve) at 20oC.............................................................155

Figure 6.5 Geometry of the actuator displayed in Vector Fields Opera: full view

without steel walls at the front (left) and plan view (right)...........................................160

Figure 6.6 Geometry of the actuator displayed in Vector Fields Opera (only a quarter is

shown using model symmetry) .....................................................................................161

Figure 6.7 N48 permanent magnet normal demagnetisation curve at 20ºC [143].......161

Figure 6.8 EN1A mild steel B-H curve [147] ..............................................................162

Figure 6.9 3-D plot of the flux density distribution produced by the magnets ............163

Figure 6.10 Flux density distribution in the airgap produced by the magnets.............163

Figure 6.11 Cross-sectional view of the flux density distribution with vectors produced

by the magnets...............................................................................................................164

Figure 6.12 Flux density distribution in the airgap produced by the magnets (along the

path noted in Figure 6.11) .............................................................................................165

Figure 6.13 Plan view of the flux density distribution with vectors produced by the

magnets .........................................................................................................................166


path noted in Figure 6.13) .............................................................................................166

List of Figures

12

Figure 6.15 3-D plot of the flux density distribution produced by the magnets and the

actuator coil carrying 50A in the anti-clockwise direction ...........................................168

Figure 6.16 Flux density distribution in the airgap produced by the magnets and the

actuator coil carrying 50A in the anti-clockwise direction (along the vertical path)....168


actuator coil carrying 50A in the clockwise direction ..................................................169


actuator coil carrying 50A in the clockwise direction ..................................................169

Figure 6.19 Actuator force versus current characteristics............................................170

Figure 6.20 Geometry of the actuator with magnetic latches (the two steel walls and

latches at the front are not shown) (left) and cross-sectional view (right)....................171

Figure 6.21 3-D structure of the operating actuator.....................................................176

Figure 6.22 Cross-sectional view of the vacuum interrupter actuator (all dimensions in

mm) ...............................................................................................................................176

Figure 6.23 Plan view of the vacuum interrupter actuator (all dimensions in mm).....177

Figure 6.24 Plan view of the stopper and carbon fibre plate with latch steel (all

dimensions in mm)........................................................................................................177

Figure 6.25 N48 Nd-Fe-B permanent magnet..............................................................178

Figure 6.26 Steel frame with permanent magnets........................................................179

Figure 6.27 Actuator coil on the fibreglass tube..........................................................179

Figure 6.28 Operating actuator with vacuum interrupter.............................................180

Figure 7.1 Schematic diagram of the actuator control circuit ......................................183

Figure 7.2 Schematic diagram of the precision full-wave rectifier [154] ....................186

Figure 7.3 Schematic diagram of the ‘open’ signal circuit for the vacuum interrupter187

Figure 7.4 Schematic diagram of the ‘close’ signal circuit for the vacuum interrupter

.......................................................................................................................................188

Figure 7.5 Schematic diagram of the MOSFET drive circuit ......................................189

Figure 7.6 Trigger signal and MOSFET drive circuit..................................................190

Figure 7.7 Vacuum interrupter with its actuator and control circuit............................190

Figure 8.1 Actuator opening and closing force versus current ....................................194

Figure 8.2 Opening operation of the vacuum interrupter with a capacitor voltage of

100V at atmospheric pressure .......................................................................................196

Figure 8.3 Opening operation of the vacuum interrupter with different capacitor

voltages at atmospheric pressure...................................................................................197

List of Figures

13

Figure 8.4 Operating actuator in the vacuum chamber (left) and external view of the

vacuum chamber (right) ................................................................................................199

Figure 8.5 External connectors for the vacuum interrupter and its actuator................199

Figure 8.6 Opening of the vacuum interrupter with different capacitor voltages in the

vacuum chamber ...........................................................................................................200


vacuum chamber with the current duration increased to 15ms.....................................201

Figure 8.8 Comparison of the opening of the vacuum interrupter at atmospheric

pressure and in the vacuum...........................................................................................202

Figure 8.9 Schematic of the high-current test circuit with the vacuum interrupter .....204

Figure 8.10 Quench test with a potential peak current of 324A with and without the

vacuum interrupter ........................................................................................................206

Figure 8.11 Temperature rise of the coil during a quench test with a potential peak

current of 324A with and without the vacuum interrupter............................................207

Figure 8.12 Simulated fault test without the vacuum interrupter.................................208

Figure 8.13 Simulated fault test with the vacuum interrupter......................................209

List of Tables

14

List of Tables

Table 2.1 Summary of active SFCL projects .................................................................63

Table 4.1 MgB2 wire specification.................................................................................95

Table 5.1 Temperature response comparison with fault time variation.......................147

Table 6.1 DVS10CB vacuum interrupter specification...............................................151

Table A.1 Thermal conductivity of alumina................................................................228

Table A.2 Specific heat capacity of alumina ...............................................................229

Table A.3 Density of alumina ......................................................................................231

Table A.4 Thermal conductivity of copper ..................................................................231

Table A.5 Specific heat capacity of copper .................................................................232

Table A.6 Density of copper ........................................................................................233

Table A.7 Resistivity of copper ...................................................................................233

Table A.8 Thermal conductivity of magnesium diboride ............................................234

Table A.9 Specific heat capacity of magnesium diboride............................................234

Table A.10 Density of magnesium diboride ................................................................235

Table A.11 Thermal conductivity of monel.................................................................236

Table A.12 Specific heat capacity of monel ................................................................237

Table A.13 Density of monel .......................................................................................237

Table A.14 Thermal conductivity of niobium..............................................................237

Table A.15 Specific heat capacity of niobium.............................................................238

Table A.16 Density of niobium....................................................................................239

Table A.17 Thermal conductivity of nitrogen..............................................................240

Table A.18 Specific heat capacity of nitrogen.............................................................240

Table A.19 Density of nitrogen....................................................................................241

Table A.20 Thermal conductivity of polystyrene ........................................................241

Table A.21 Specific heat capacity of polystyrene........................................................241

Table A.22 Density of polystyrene ..............................................................................241

Table B.1 Summary of components for prototype operating actuator.........................242

Nomenclature

15

Nomenclature

Symbol Meaning SI Units

Aac Cross-sectional area of the actuator coil m2

Acb Cross-sectional area of copper in the copper braid m2

Ag Area of the airgap m2

Am Permanent magnet arc surface area m2

Apeak Peak current A

Arms Root mean square current A

B Magnetic flux density T

Bc Critical magnetic flux density T

Bg Magnetic flux density in the airgap T

Bm Permanent magnet magnetic flux density T

Br Remanence in the permanent magnet T

Cp Specific heat capacity J/kg·K

d Distance between two electrodes m

di

dt Rate of change of current A/s

Adi

dt

Rate of change of current in strand A of the three-strand SFCL

coil A/s

Bdi

dt

Rate of change of current in strand B of the three-strand SFCL

coil A/s

Cdi

dt

Rate of change of current in strand C of the three-strand SFCL

coil A/s

sdi

dt Rate of change of current in the high-current test circuit A/s

dv

dt Rate of change of voltage V/s

E Back-emf V

F Force on the actuator coil N

Fc Actuator closing force N

Nomenclature

16

Fo Actuator opening force N

Fr Actuator reluctance force N

g Power loss density W/m3

gc Power generated in the SFCL coil per unit volume W/m3

gcb Power loss density in the copper braid W/m3

gcm Power loss density in the SFCL coil (calculated from

measurement) W/m3

Hc Critical magnetic field strength A/m

Hc1 Lower critical magnetic field strength A/m

Hc2 Upper critical magnetic field strength A/m

Hg Airgap magnetic field strength A/m

Hm Permanent magnet magnetic field strength A/m

I Current in the actuator coil A

IA Current in strand A of the three-strand SFCL coil A

IB Current in strand B of the three-strand SFCL coil A

Icm Measured current passing through the SFCL coil A

IC Current in strand C of the three-strand SFCL coil A

Is High-current test circuit current A

Iq Quench current A

Jc Critical current density A/mm2

k Thermal conductivity W/m·K

K Spring stiffness N/mm

Kp Permeance coefficient H/m

l Total length of the actuator coil m

lmf Length of the actuator coil in the magnetic field m

L Inductance of the actuator coil H

LA Self-inductance of strand A of the three-strand SFCL coil H

LB Self-inductance of strand B of the three-strand SFCL coil H

Lc High-current test circuit inductance H

Lcb Length of the copper braid m

LC Self-inductance of strand C of the three-strand SFCL coil H

lg Length of the airgap m

lm Length of the permanent magnet m

Nomenclature

17

M Mutual inductance H

MAB Mutual inductance between strand A and strand B H

MAC Mutual inductance between strand A and strand C H

MBA Mutual inductance between strand B and strand A H

MBC Mutual inductance between strand B and strand C H

MCA Mutual inductance between strand C and strand A H

MCB Mutual inductance between strand C and strand B H

M1 Mass of the moving part kg

M2 Mass of the stationary part kg

N Turns of the actuator coil -

P Pressure Pa

r Distance from the centre along the radial direction m

R Resistance of the actuator coil Ω

RA Resistance of strand A of the three-strand SFCL coil Ω

RB Resistance of strand B of the three-strand SFCL coil Ω

Rc High-current test circuit resistance Ω

RC Resistance of strand C of the three-strand SFCL coil Ω

Rcb Resistance of the copper braid Ω

Rfcl SFCL coil resistance Ω

t Time s

T Temperature K

Tc Critical temperature K

v SFCL coil volume m3

vc Velocity m/s

vcb Volume of the copper braid m3

vcm Measured voltage across the SFCL coil V

v10 Initial velocity of the moving contact before collision m/s

V Voltage of supply for the actuator coil V

VA Voltage across strand A of the three-strand SFCL coil V

VB Voltage across strand B of the three-strand SFCL coil V

Vc Voltage across the capacitor V

VC Voltage across strand C of the three-strand SFCL coil V

Vs Voltage of the high-current test circuit V

Nomenclature

18

V0 Voltage across the capacitor at time t = 0 V

x Distance along the SFCL coil length m

z Distance from the centre along Z-axis direction m

ρ Resistivity of copper Ω·m

ρd Density kg/m3

ρ0 Normal state resistivity of a superconductor Ω·m

µ0 Permeability of free space (4π×10-7) H/m

µm Relative permeability of the permanent magnet -

List of Abbreviations

19


1G First generation

2G Second generation

1-D One-dimensional

2-D Two-dimensional

3-D Three-dimensional

AC Alternating current

Al 2O3 Alumina

AMF Axial magnetic field

AMSC American Superconductor

ASL Applied Superconducting Limited

BIL Basic impulse level

BSCCO Bismuth strontium calcium copper oxide

CAS Chinese Academy of Sciences

CIEE Chinese Institute of Electrical Engineering

CTFF Continuous tube forming and filling

CuCr Copper chrome

DC Direct current

DoE Department of Energy

emf Electromotive force

EPR Electromagnetic repulsion plate

EU European Union

FE Finite element

FCL Fault current limiter

HTS High temperature superconductor

IBAD Ion beam assisted deposition

IGCT Integrated gate commutated thyristor

KEPCO Korea Electric Power Corporation

KEPRI Korea Electric Power Research Institute


20

LBCO Lanthanum barium copper oxide

LSIS LS Industrial Systems

LTS Low temperature superconductor

LVDT Linear variable differential transformer

MCP Melt, cast and processed

MFCL Matrix fault current limiter

MFP Mean free path

MgB2 Magnesium diboride

MOCVD Metal organic chemical vapour deposition

MOD Metal-organic deposition

MOSFET Metal-oxide-semiconductor field-effect transistor

MRI Magnetic resonance imaging

NbTi Niobium titanium

Nd-Fe-B Neodymium-iron-boron

OD Outer diameter

PIT Powder-in-tube

PLD Pulsed laser deposition

RABiTS Rolling assisted biaxial textured substrates

RMF Radial magnetic field

RMS Root mean square

SCE Southern California Edison

SFCL Superconducting fault current limiter

SF6 Sulfur hexafluoride

SRBL Synthetic resin bonded lamination

TRV Transient recovery voltage

U.K. United Kingdom

U.S. United States

U.S.A. United States of America

VI Vacuum Interrupter

YBCO Yttrium barium copper oxide

Abstract

21

Abstract

Fault current levels in land-based power systems are generally rising because of the

increase in renewable generation capacity. Once the fault current level exceeds the

capacity of the existing protection equipment, expensive upgrades become necessary. In

order to avoid excessively expensive equipment upgrades, many fault current limitation

techniques have been investigated.

This thesis presents the work conducted on the design, manufacture and testing of a

resistive superconducting fault current limiter (SFCL) with an integrated fast-acting

vacuum interrupter. The practical application of magnesium diboride (MgB2) in round

wire form was also investigated.

A single-strand MgB2 SFCL coil was investigated and demonstrated repeatable and

reliable current-limiting action. In practical power system applications, the development

of SFCLs needs a considerable scale-up of the current-carrying capability of the MgB2

wire samples. One option is to use parallel wires in order to carry current levels in the

kA range. The behaviour of a prototype three-strand MgB2 SFCL coil was assessed,

which showed that each of the three wire strands shared the current approximately

equally and demonstrated reliable and repeatable behaviour during testing. The MgB2

SFCL coil with multiple wire strands in parallel shows considerable potential as a

practical method for scaling-up the current levels required for power system

applications.

One of the significant operational issues for resistive SFCLs is the temperature recovery

time after a fault is cleared. A vacuum interrupter was integrated therefore into the

SFCL system to quickly remove the superconducting coil from the circuit during a fault

condition and allow the superconducting coil to recover whilst a bypass resistor acted as

a current limiting resistor. A fast-acting actuator and its control circuit were designed

and manufactured to control the operation of the vacuum interrupter. The SFCL with a

prototype vacuum interrupter was successfully tested to validate the design process. The

Abstract

22

energy dissipated in the superconducting coil was significantly reduced by the fast

operation of the vacuum interrupter and the recovery time significantly reduced. This

research demonstrates the potential of a cost-effective and compact SFCL for the power

system applications.

Declaration

23

Declaration

No portion of the work referred to in this thesis has been submitted in support of an

application for another degree or qualification of this or any other university or institute

of learning.

Copyright Statement

24

Copyright Statement

i. The author of this thesis (including any appendices and/or schedules to this

thesis) owns any copyright in it (the “Copyright”) and s/he has given The

University of Manchester the right to use such Copyright for any administrative,

promotional, educational and/or teaching purposes.

ii. Copies of this thesis, either in full or in extracts, may be made only in

accordance with the regulations of the John Rylands University Library of

Manchester. Details of these regulations may be obtained from the Librarian.

The page must form part of any such copies made.

iii. The ownership of any patents, designs, trade marks and any and all other

intellectual property rights except for the Copyright (the “Intellectual Property

Rights”) and any reproductions of copyright works, 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 Rights and Reproductions cannot and must not be made available for

use without prior written permission of the owner(s) of the relevant Intellectual

Property Rights 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://www.campus.manchester.ac.uk/medialibrary/policies/intellectual-

property.pdf), 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.

Acknowledgments

25

Acknowledgments

I am sincerely grateful to my academic supervisor, Professor Sandy Smith, for his

invaluable guidance and support during my Ph.D. research. I would therefore like to

take this opportunity to express my gratefulness and appreciation for his encouragement

and great patience.

I would like to thank the Dorothy Hodgkin Postgraduate Award, EPSRC and Rolls-

Royce Plc. for the financial funding of my Ph.D. research. Thanks are also due to Mark

Husband from Rolls-Royce for his guidance during the work.

I would like to thank all the academic staff, research associates and Ph.D. candidates of

the Power Conversion Group. A special thanks to Dr Paul Tuohy for his help

proofreading this thesis.

Thanks are also due to Malcolm Bailey and Paul Shaw from the School of Electrical

and Electronic Engineering’s electronics and mechanical workshops for manufacturing

some of the equipment used in the experimental test.

Finally, I would like to thank my family for their blessings and support, especially to

my husband Xianwu Zeng for his encouragement and accompanying me in the U.K.

The Author

26

The Author

Xiaoze Pei obtained her BEng (Hons) in Electrical Engineering from Beijing Jiaotong

University in 2006. Thereafter, she received her MEng (Hons) in Power Electronics and

Power Drive from Beijing Jiaotong University in 2008. The work described in this

thesis, which started in 2008, constitutes the author’s first major research project. The

following paper has been published:

Xiaoze Pei, Alexander C. Smith, Mark Husband, and Matthew Rindfleisch,

“Experimental Tests on a Superconducting Fault Current Limiter using Three-

strand MgB2 Wire”, in 22nd International Conference on Magnet Technology,

Marseille, France, 2011, pp 1-5, [Accepted by IEEE Transactions on Applied

Superconductivity].

Chapter 1: Introduction

27

1 Introduction

1.1 Background

The generation and transmission capacity of electric power systems is increasing to

meet growing demands [1, 2]. In addition, the interconnections within the power grids

increase to improve power quality and reliability. Distributed renewable energy

generation such as solar energy and wind energy, are also being connected into power

systems. All these connected systems cause the capacity of the power systems to

increase, which leads to the potential of increasing fault current levels. Increasing fault

current levels in transmission and distribution systems have become a serious problem

for land-based power systems.

The rising fault current levels will ultimately require a large amount of equipment in the

existing power systems such as circuit breakers and transformers, to be replaced. In

order to avoid large scale and expensive equipment upgrades, cost-effective fault

current limitation techniques are being examined [3].

In the past several decades, a variety of fault current limitation techniques have been

investigated, summarised in Figure 1.1 [3]. Generally, they fall into two broad

categories: ‘permanent impedance increase’ and ‘condition based impedance increase’

[3]. The first category includes techniques such as introducing high impedance

transformers and current limiting reactors, which add extra impedance during both

normal and fault conditions. Amongst the latter category, novel fault current limiters

(FCLs) using superconductors and/or semiconductors have been studied [2]. FCLs

introduce negligible impedance during normal operation and are almost invisible to the

power system. Once a fault occurs, however, the impedance of the FCLs increases and

reduces the fault current to a level that is acceptable to the existing protection equipment

in the system. Potential applications for FCLs in power networks are shown in Figure

1.2.


28

Figure 1.1 Methods of fault current limitation [3]

Figure 1.2 Potential applications of FCL devices [3]

The discovery of high temperature superconductors (HTS) in the 1980’s significantly

improved the potential for the practical application of superconducting fault current


29

limiters (SFCLs). SFCLs utilise the characteristics of the superconducting materials to

limit the fault current. Many SFCL design concepts have been evaluated [2, 4-7];

resistive SFCLs have the simplest and most compact structure [4, 8-11]. Bismuth

strontium calcium copper oxide (BSCCO) and yttrium barium copper oxide (YBCO)

have been widely researched for SFCLs applications. More recently, magnesium

diboride (MgB2) in simple round wire form has been tested and shown to have

significant potential as a low-cost resistive SFCL [12-15]. The critical temperature of

MgB2 is 39K and, for this thesis study, the MgB2 is operating in solid nitrogen (<63K).

The most serious problem with resistive SFCLs is overheating during a fault and the

long recovery time of the superconducting material. This research proposes therefore to

integrate a vacuum interrupter with a resistive SFCL to overcome this problem.

1.2 Project motivation and context

Resistive SFCLs have been previously investigated because of their simple structure

and making direct use of the non-linear characteristics of the superconducting materials.

MgB2 has been manufactured in round wire form which makes coil manufacture easier.

MgB2 also provides the potential to be a cost-effective and economic material for SFCL

applications because the raw materials are cheaper than BSCCO and the manufacturing

process is simpler than YBCO coated conductors [16]. A prototype single-strand MgB2

SFCL coil was manufactured and tested by Oliver [17] and the coil demonstrated

successful operation via an experimental test rig was built for the test of the prototype

coil. In addition, Oliver [17] implemented a finite element (FE) model for the thermal

design of the alumina former and a MATLAB model to simulate the behaviour of the

SFCL coil [15, 17]. This thesis builds on the research work reported by Oliver [17]

expanding the experimental studies, thermal modelling and simulation. Additionally,

this research also integrates vacuum interrupter mechanism.

As with the work of Oliver [17] a resistive SFCL using round MgB2 wire as the

superconducting material was initially investigated in this thesis. The wire had

improved manufacturing and mechanical properties to that used by Oliver.


30

For typical distribution system voltages of 11kV or 33kV, for example, the development

of an MgB2 SFCL needs considerable scale-up in term of the current-carrying capability

of the MgB2 wire. The two obvious options to achieve this are to increase the wire size

and use a multifilament core and/or use multiple wire strands. Large wire sizes in

monofilament superconductors may result in a condition referred to as ‘flux jumping’, a

phenomena that can lead to instability such that the material moves from a

superconducting state to a normal conducting state. Thus the wire diameter is limited

[18]. Typical distribution level currents of 1250Arms, for example, mean it is unlikely

that a single wire SFCL will be feasible [19]. In practical power system applications,

parallel wires will likely be required in order to carry current levels in the kA range. The

full quench current level will be reduced; however, if the individual wire strands (and

connections) are not closely identical in terms of their quench current level and

impedance. Ensuring each strand of wire carries the same current is therefore an

important issue for multi-strand wire solutions. The behaviour of a three-strand MgB2

SFCL coil is assessed following the investigation of a single-strand SFCL coil.

The major disadvantage of a resistive SFCL is that the superconducting material heats

up rapidly after quenching during a fault condition and may take several minutes to cool

down and recover after the fault is cleared. The purpose of this research project is

therefore to explore the potential of integrating a vacuum interrupter into the resistive

SFCL directly. The cryostat used to cool the superconducting material typically uses a

vacuum around the cold chamber as super thermal insulation. The objective is to mount

a vacuum interrupter into the vacuum chamber of the cryostat in series with the

superconducting coil. A bypass resistor is then connected in parallel with both of them

but external to the cryostat. The schematic circuit of the proposed resistive SFCL with

an integrated vacuum interrupter is shown in Figure 1.3. During normal operation, the

impedance of the vacuum interrupter and superconducting coil remain essentially zero

such that the bypass resistor conducts virtually no current. When a fault occurs, the

superconducting coil begins to quench and starts to develop a resistance. The voltage

that appears across the superconducting coil then increases and can be used as the

trigger signal for the vacuum interrupter. The vacuum interrupter needs to be designed

to open quickly and isolate the superconducting coil from the system. The bypass

resistor acts as the fault current limiting resistance after the vacuum interrupter opens.


31

The temperature rise of the superconducting coil is reduced as a result and additionally

the coil starts to recover whilst the bypass resistor continues to limit the fault current.

The bypass resistor would be in the circuit during the fault for typically a few hundred

milliseconds before the back-up circuit breaker isolates the fault. The bypass resistor is

therefore designed and rated for short-term operation which makes it compact and

relatively low-cost. The fast-acting vacuum interrupter will reduce the energy dissipated

in the superconducting coil and allows for a faster recovery time compared to a basic

resistive SFCL. It is demonstrated in this thesis that it is possible to operate the vacuum

interrupter within a half-cycle of the onset of the fault current and also possible to

reconnect the SFCL coil within seconds rather than minutes.

Figure 1.3 Proposed resistive SFCL with vacuum interrupter

A fast operating actuator for the vacuum interrupter is a key specification requirement

for this system and hence a main feature of this research project. The characteristics of

existing vacuum interrupters and actuator technologies are reviewed. A ‘voice-coil’ type

actuator with magnetic latches is proposed for this application. The actuator is expected

to provide fast opening and closing of the vacuum interrupter whilst magnetic latches

can hold the actuator in the open or closed position.

It is worth noting the differences between the proposed research and a hybrid SFCL. In

a hybrid SFCL, an interrupter is driven by the fault current through a parallel coil [20,

21]. The operating of the interrupter is therefore strongly affected by the fault current

level. When the fault current is low, the opening stroke of the interrupter is small and

the current in the superconducting coil may fail to interrupt because of insufficient


32

clearance distance between the contacts [22]. In addition, there is a risk of the

interrupter reclosing before the fault current is cleared. In this research project, the

vacuum interrupter is operated by a separate fast-acting actuator whose operation is

independent from the fault current level. If successful, this concept would represent a

technical step forward in realising a practical low-cost SFCL solution.

1.3 Organisation of thesis

Chapter 2 explains the characteristics of the superconducting materials related to this

research project. Three superconducting materials utilised in current SFCL research

projects are also covered. A review of fault current limitation theory, conventional

practical methods of fault current limitation, novel FCLs including non-superconducting

and superconducting FCLs, is reported. A summary of the characteristics of the vacuum

interrupter and its applications are also presented. In addition, four different structures

of vacuum interrupter actuators are compared.

Chapter 3 describes the manufacturing process of the MgB2 SFCL coil undertaken

during the study, which includes the current connections, winding the coil and the heat

treatment process. The instrumentation of voltage, current and temperature signals on

the prototype coil is introduced. High-current and low-current experimental test circuits

and a LabVIEW control system are also presented.

Chapter 4 presents the experimental investigation of a single-strand and three-strand

MgB2 SFCL coil. The three-strand MgB2 SFCL coil is evaluated to explore the potential

of increasing the current carrying capacity using multi-strand wires. The quench

behaviour of both SFCL coils is also evaluated.

Chapter 5 presents the work carried out to expand and improve the MATLAB and finite

element thermal model of the single-strand MgB2 SFCL coil presented by Oliver [17].

The development of the MATLAB model is to help understand the SFCL current-

limiting process and inform future design work. The FE thermal model is built to

calculate the temperature distribution around the coil and surrounding cryogen. The

results of both the MATLAB and FE models are compared with the experimental test

results to verify the accuracy of the models. The two models are then used to predict a


33

three-second quench test to check whether the SFCL coil can withstand a fault of this

duration.

Chapter 6 describes the design and manufacture of an operating actuator for the vacuum

interrupter. A ‘voice-coil’ type actuator is used due to its simple structure and quick

response time. An analytical model and FE model of the actuator are built to estimate

the magnetic field in the airgap and resultant force produced on the coil. The magnetic

latches in both the open and closed positions are included in the FE model. The design

of the high performance operating actuator for the vacuum interrupter investigates many

problems such as contact popping, welding, bounce and rebound.

Chapter 7 describes the design and manufacture of an actuator control circuit. A full-

bridge DC-DC converter is utilised because it allows bidirectional current flow. The

‘open’ and ‘close’ signals are produced by a trigger signal circuit which is used to

control the operation of MOSFETs using a drive circuit.

Chapter 8 presents the experimental test results of an MgB2 SFCL coil with an

integrated vacuum interrupter. The vacuum interrupter operating actuator is initially

tested at atmospheric pressure before placing it into a separate vacuum chamber.

Thereafter, the vacuum interrupter is connected into the SFCL system with a bypass

resistor in parallel. The experimental test results with and without the vacuum

interrupter are compared.

Chapter 9 summarises the conclusions drawn from the work reported in this thesis and

outlines additional work that could be conducted in the future.

Chapter 2: Literature Review

34

2 Literature Review

2.1 Superconductivity

Superconductivity is a phenomenon that occurs in certain materials below a critical

temperature, magnetic field strength and current density that exhibit zero DC resistivity.

In 1911, Heike Kamerlingh Onnes found that pure mercury exhibited zero resistance

when the temperature was reduced below 4K, which was the first discovery of

superconductivity [23]. A superconducting alloy composed of niobium and titanium

(NbTi) was discovered in the 1960’s and is still one of the popular superconducting

materials due to its high critical magnetic field strength and current density levels [18].

Before 1986, all the superconducting materials discovered exhibited superconductivity

below 25K and were termed low temperature superconductors (LTS). In 1986, Karl

Müller and Johannes Bednorz achieved superconductivity at around 30K in lanthanum

barium copper oxide (LBCO), which was the first high temperature superconductor

(HTS) [24]. Today, in high temperature superconductors the critical temperature of

YBCO is 90K whilst the critical temperature of BSCCO is 105K. This is important

because they exhibit superconductivity in liquid nitrogen which is cheaper and much

easier to produce than liquid helium.

2.1.1 Elementary properties of superconductors

2.1.1.1 Zero DC resistivity

The resistivity of a normal conductor and a typical superconductor is shown in Figure

2.1. In a normal conductor the current is carried by the electrons which are constantly

colliding with the atoms. During each collision some of the energy is dissipated and

converted into heat, which causes electrical resistivity in the conductor. When the

temperature decreases, the thermal vibrations of the atoms reduces and the electrons are

less frequently scattered by atoms, reducing the resistivity of the conductor. The

resistivity of a pure metal should approach zero as the temperature reduces towards 0K.

However, in practice a metal cannot be perfectly pure and will contain some impurities.


35

The electrons are scattered by these impurities which is independent of temperature. As

a result, there is a residual resistivity at 0K for impure metals [23].

The temperature at which a superconductor exhibits zero resistivity is called its critical

temperature. Above the critical temperature, the resistivity of a superconductor is

similar to a normal metallic conductor. Zero resistivity below the critical temperature

can be explained by ‘cooper pairs’, which formed part of BCS theory developed by

John Bardeen, Leon Cooper and John Schrieffer [24]. A ‘cooper pair’ is two electrons

that are bound together at low temperature. The velocity of the ‘cooper pair’ is low at

low temperature and therefore there is insufficient energy to break the pairing. The DC

resistivity is zero.

Res

isti

vit

y

Impu

re m

etal

Pure m

etal

Res

isti

vit

y

Temperature0

Super

cond

ucto

r

Tc

Figure 2.1 Comparison of resistivity of normal conductors and a typical superconductor

[23]

2.1.1.2 Meissner effect

A magnetic field would be fully excluded from the inside of a superconductor if it is

placed into a magnetic field of relatively low strength below the critical magnetic field

strength, Hc. This perfect diamagnetism in a superconductor, as shown in Figure 2.2

(left), is known as the ‘Meissner effect’. As the magnetic field strength increases above

Hc, the superconducting state will break down, which means that the magnetic field can

fully penetrate the superconductor.

Generally speaking, superconductors can be divided into Type I and Type II

superconductors. Type I superconductors are usually made from a single metallic

element and exhibit perfect diamagnetism below their critical magnetic field strength.


36

Type II superconductors are always alloy or compounds and they are able to remain in

the superconducting state at a higher applied magnetic field strength and current density

levels, which makes commercial applications possible.

Figure 2.2 Exclusion of the magnetic field in a superconductor [25]

Type I superconductors must expel all the applied magnetic field to stay in the

superconducting state, which means they can only be placed in a magnetic field of

relatively low strength. Figure 2.3 (left) shows that a Type I superconductor has only

two states: the superconducting state and normal state.

A Type I superconductor exhibits zero resistivity and zero internal magnetic flux

density when the applied external magnetic field strength, H, is below the critical

magnetic field strength, Hc. Once the applied magnetic field strength is higher than Hc, a

Type I superconductor would completely break down; the magnetic field can then

penetrate the superconductor and it would become a normal conductor. A Type I

superconductor has relatively low current carrying capacity because a transport current

will generate its own internal magnetic field and this seriously limits its application.

In Figure 2.3 (right), it can be seen that a Type II superconductor has three states: the

superconducting state, mixed state (vortex state) and normal state. When the magnetic

field strength is below the lower critical field strength, Hc1, it is in the superconducting

state with zero resistivity and internal magnetic field. Hc1 for a Type II superconductor

is equivalent to Hc for a Type I superconductor. When the applied magnetic field

strength increases between Hc1 and the upper magnetic field strength, Hc2, the


37

superconductor is in the mixed state, which allows some penetration of magnetic field.

In the mixed state, magnetic field penetrates the superconductor via a series of magnetic

flux filaments that pass through the material. The filaments are surrounded by

circulating super currents which screen the magnetic field from the rest of the

superconductor. These filaments together with the super currents are termed vortices.

Once the magnetic field strength increases above Hc2, the superconductor becomes a

normal conductor.

Ap

pli

ed m

agn

etic

fie

ld

Figure 2.3 Applied magnetic field: Type I (left) and Type II (right) superconductor [25]

2.1.1.3 Critical temperature

The temperature profile of a Type II superconductor can be divided into three regions:

the normal region, transition region and superconducting region, as shown in Figure 2.4.

20ρ

Res

isti

vit

y

Figure 2.4 Typical temperature profile of a Type II superconductor [25]


38

In the normal region, the resistivity of a superconductor is similar to a normal metallic

conductor. In the transition region, a superconductor transforms from the normal

conducting state to the superconducting state. Finally, in the superconducting state, a

superconductor exhibits zero DC resistivity. The normal state resistivity, ρo, is defined

as the resistivity at a temperature where the onset of superconductivity begins [25]. The

critical temperature is usually defined where the resistivity is half its normal state

resistivity (ρo/2).

2.1.1.4 Critical current density

The critical current density, Jc, is the third fundamental property of a superconductor

together with the critical magnetic field strength and critical temperature. The critical

current density is defined as the current density level which if exceeded will cause the

superconductor to lose its superconductivity. For a Type I superconductor, the critical

current density commonly can be taken as the level of current density which develops

the critical magnetic field strength on its surface.

A Type II superconductor behaves like a Type I superconductor for the magnetic field

strength produced by the transport current below the lower critical field strength, Hc1.

When the magnetic field strength produced by the transport current increases above Hc1,

a Type II superconductor allows magnetic field penetration via magnetic flux filaments.

As mentioned in section 2.1.1.2, these filaments are surrounded by circulating super

currents. Since the energy of the superconductor is not utilised in expelling magnetic

field, the critical current is generally much higher in a Type II superconductor than a

Type I. The increase of magnetic field penetration continues until the magnetic field

strength increases above Hc2 and the superconductor completely transitions to the

normal conducting state.

2.1.1.5 AC losses

Superconducting materials exhibit zero resistivity only when a DC current exists in it

because all the current is carried by super-electrons. However, it does not exhibit zero

resistivity to AC current because of alternating magnetic field and inertial mass of the

super-electrons. These AC losses translate into a resistivity. The AC resistivity may be

several orders of magnitude smaller than the normal state resistivity, ρo, but it is finite

and increases with frequency. For land-based power system frequency, AC losses are


39

largely due to the hysteretic motion of the vortices, i.e. flux creep, hysteresis and related

magnetic phenomena in a superconductor [23, 25].

2.1.2 Materials

Amongst all the superconducting materials, BSCCO, YBCO and MgB2 are investigated

and utilised in SFCL applications. They are discussed in the following sections.

2.1.2.1 BSCCO

BSCCO with a critical temperature of 105K was discovered in 1988 by Maeda and his

colleagues at the National Research Institute for Metals in Japan [26]. BSCCO can be

manufactured into both bulk and tape form. BSCCO tape, which has been commercially

available since the late 1990’s, is also known as first generation (1G) wire.

BSCCO tapes are manufactured using a powder-in-tube (PIT) process. Precursor

BSCCO powders are filled into a silver or silver alloy tube which is drawn down to a

small diameter monocore wire. In order to manufacture multifilament BSCCO tape, a

specific number of monocore wires are stacked closely into a second silver or silver

alloy tube, which is again extruded down in diameter and rolled into a flat tape. The

tape is then reacted at high temperature to form a dense and aligned multifilamentary

tape and to increase the current carrying capacity. This design is relatively expensive

because a large amount of silver or silver alloy is used [27].

2.1.2.2 YBCO

YBCO was discovered to have a critical temperature of 90K in 1987. YBCO can be

manufactured in either thin film form or as a coated conductor. The thin film of YBCO

is created on a substrate using a pulsed laser deposition (PLD) technique [28]. However,

YBCO thin film has two problems which limit its commercial application. Firstly,

YBCO is very sensitive to grain boundaries. If the grain boundary angle is greater than

5°, the performance of YBCO thin film will be seriously affected. Secondly, YBCO thin

film is very brittle, which is difficult to wind into coils, for example.

An YBCO coated conductor, which comprises multiple coatings on a base substrate, is

also known as second generation (2G) wire. Companies that currently produce 1G HTS

wire are migrating to 2G HTS wire. SuperPower Inc. manufactures 2G wire using an


40

ion beam assisted deposition (IBAD)/metal organic chemical vapour deposition

(MOCVD) method [29-31]. SuperPower can manufacture YBCO tape in excess of

500m with a critical current between 200 and 275A/cm-width [30]. The configuration of

the SuperPower 2G HTS wire type (SCS4050) is shown in Figure 2.5. The architecture

of 2G wire is totally different from 1G wire.

Figure 2.5 Configuration of SuperPower 2G HTS wire (wire type SCS4050) [30]

American Superconductor (AMSC) 2G wire is manufactured using a metal-organic

deposition (MOD)/rolling assisted biaxial textured substrates (RABiTS) method [9, 32].

This process produces a 4cm wide tape that is subsequently cut to a width of typically

4.4mm or 4.8mm. Currently AMSC manufactures YBCO tape in nominal lengths of 80

to 100m, with a critical current of 250A/cm-width in self-field at 77K [33]. AMSC

states that the 2G HTS wire they manufacture can carry over 100 times the electrical

current of copper wire of the same size, which significantly increases the current

capacity of power cables, for example [34].

Companies currently manufacturing HTS wire include Superconductor Technologies

Inc., Bruker Energy & Supercon Technologies, Inc., Zenergy Power plc., Nexans

Superconductors, Fujikura, Sumitomo, Furukawa and Showa.


41

YBCO tape provides improved performance in a magnetic field and improved

mechanical properties with potentially lower cost through using cheaper raw materials.

Additionally, the large surface area of the tapes may enable them to cool down faster

after a quench (transition to the normal state). The manufacturing process however is

extremely complicated and this may make it difficult to be commercially competitive.

The tapes also have a low resistivity when quenched due to the presence of a copper

stabiliser, which means that long lengths are required to develop the desired resistance

in FCL applications. Furthermore, a tape form superconductor makes them less

attractive for winding into coils and may lead to larger AC losses than round wires.

2.1.2.3 Magnesium diboride

The discovery of MgB2 which exhibits superconductivity below 39K has attracted great

research interest when it was discovered in 2001 [35]. MgB2 can be made into round or

square wire form and also flat tape. The wire form can be chosen to match the

requirement of a specific application, making coil manufacture easier.

Companies manufacturing MgB2 wires include Columbus Superconductor, Hyper Tech

Research, Inc. and Hitachi Research Laboratory. Columbus Superconductor produces

MgB2 wires using an ex-situ PIT process, which is a similar manufacturing process to

1G wire [36]. The mixture of the magnesium and boron powder is heat treated at 900°C

for 1 hour to form MgB2 powder which is then filled into a metal tube. The tube is

reduced in diameter and sintered between 800 and 1000°C. Columbus Superconductor

is manufacturing MgB2 wires for tests in magnetic resonance imaging (MRI) and FCL

applications.

Hyper Tech developed and patented the continuous tube forming and filling (CTFF)

process [37]. A metal strip is formed into a U-shape and filled continuously with a

mixture of magnesium and boron powder. The U-shape tube is then closed and reduced

in diameter by wire drawing. The wire is then heated to form MgB2 superconductor:

called an in-situ process [16, 36-38]. The advantage of the CTFF process is that it can

make continuous long lengths of wire compared to the PIT process, which requires

larger and larger tubes to obtain longer lengths of wire. Hyper Tech produces a


42

commercial multifilament wire consisting of 18 MgB2 filaments with a niobium barrier,

an inner copper monofilament sheath and a monel outer sheath [16].

The advantages of MgB2 wire are the use of cheaper raw materials compared with

BSCCO, a simpler manufacturing process compared with YBCO and reasonable

cooling costs compared with low temperature superconductors. In addition, MgB2 wires

can be manufactured with a variety of resistive sheaths. For a fault current limiter

application, for example, a short length of MgB2 wire can be used to insert a desired

resistance into the circuit by using a high resistivity sheath material such as stainless

steel. A large amount of work has been carried out on the fault current-limiting

properties of various MgB2 wires. All of these wire samples have demonstrated good

current-limiting properties with a fast transition from the superconducting to the normal

state. MgB2 wire therefore has become one of the strong candidates for the development

of SFCLs [12, 13, 15, 39].

2.1.3 Summary

The basic characteristics of superconductors including zero DC resistivity, critical

magnetic field strength, critical temperature, critical current density and AC losses have

been explained. Three superconducting materials namely BSCCO, YBCO and MgB2 are

commonly used for SFCL applications. Their raw materials, manufacturing processes

and properties have been summarised.

2.2 Fault current limitation

2.2.1 Introduction

Many factors can cause a fault in a power system. The fault current level can be

relatively large, which may damage equipment in the power system and even cause

permanent failure. Power systems have to be designed to withstand mechanical and

thermal stresses during a fault. Power system protection devices detect fault conditions

and operate circuit breakers and other devices to limit the damage. Today, fault current

levels in land-based distribution systems are of increasing concern because they are

generally rising due to the increasing capacity of connected distributed generation.

Increasing fault current levels will require expensive network investment in upgrading


43

equipment such as circuit breakers and transformers. There is a growing need therefore

for fault current limiting devices embedded into electrical networks to avoid a large

scale and expensive upgrade of existing switchgear. FCLs are expected to reduce fault

current levels without adding additional impedance during normal operation. The

capital cost of purchasing and installing FCLs must be less than the cost of upgrading

the existing equipment before they can be attractive for commercial applications.

2.2.2 Operating principles of fault current limitation

After a short circuit occurs in a power network, the fault current will increase rapidly.

The rate of current rise depends on the source voltage, source impedance and fault

phase angle. A typical prospective short circuit current is shown in Figure 2.6.

Prospective fault current

Limited fault current

Nominal current

Time

Current

Figure 2.6 Typical fault current waveforms with and without fault current limiting

The fault current would eventually be interrupted by a conventional circuit breaker. If

the first peak of the fault current is higher than the rating of the device in the power

network, it is possible that damage may occur. The simplest way to reduce the fault

current is to increase the source impedance; however, this ultimately results in

additional voltage drop, reactive power and potentially a high transient recovery voltage

(TRV).

A fault current limiter must be inserted directly into the system in order to reduce the

first several peaks of fault current level under this situation. An ideal limited current


44

waveform, as shown in Figure 2.6, is initially non-sinusoidal because the network

parameter is changed by inserting non-linear impedance.

The requirements for FCLs are listed as follows [2, 3]:

• Minimise impedance during the normal state. The resistance of the FCL during

normal state would produce extra heating and losses whilst the inductance would

induce voltage drop.

• Fast response. The FCL is able to detect the fault and reduce the fault current

below the required level before the first fault current peak.

• Quick and automatic recovery. It is desirable to reclose onto the power system as

soon as possible after a fault is cleared.

• Fail safe. The FCL would still limit the fault current even if it fails.

• Compact structure, light weight and low cost.

2.2.3 Application of fault current limiters

Possible applications for FCLs in power transmission and power distribution networks

are discussed in the following sections.

2.2.3.1 Transmission level voltage

Power transmission is a high voltage application which operates above 72kV. At the

transmission level voltage, FCLs are desirable for:

• Sub grid coupling. The distributed grids are often divided into sub grids to

manage the short circuit current. They are supplied by higher voltage levels via

separate transformers. These sub grids can be coupled together with FCLs,

providing higher power quality, fewer losses and less voltage drop.

• Busbar coupling. It is common to reduce the fault current by splitting of the

busbar. However, busbars coupled with FCLs are an excellent solution to reduce

fault currents to a reasonable level whilst inserting negligible impedance under

normal conditions.


45

2.2.3.2 Distribution level voltage

Power distribution is a medium voltage level application with voltage ratings from

approximately 6kV to 72kV. FCLs can be used in the following locations:

• Busbar coupling and transformer feeder. Busbars coupled with FCLs would

increase the short circuit capacity. An FCL also can be placed in each

transformer feeder to increase the short circuit capacity so that the transformer

can be designed with lower impedance.

• Coupling of dispersed generation. With the development of wind and solar

power, there are increased numbers of connections of distributed generation.

They can be connected to a higher voltage grid via separate transformers if there

is no extra short circuit capacity. However, an FCL can be used between a

distribution generation system and a distribution grid to avoid the need for an

expensive transformer.

• Generator feeder. Increased generation capacity of generators would increase the

short circuit current; an FCL in a generator feeder location would be more

practical instead of upgrading an old substation.

• Power plant auxiliaries. These auxiliaries are close to power plants; therefore, it

is possible to have a high short circuit current. Installing FCLs will reduce the

rating of switchgear and the cost in these locations.

2.2.4 Conventional methods

There are two general categories of techniques to reduce short circuit currents:

permanent impedance increase and condition based impedance increase [2, 3].

The permanent impedance increase technique increases the impedance for both normal

and fault conditions, which includes splitting into sub grids, splitting of busbar,

increasing the voltage level, introducing high impedance transformers and current

limiting reactors. The power grid is split into small sub grids to reduce the power

capacity of each sub grid. The potential fault current level therefore would be reduced.

Busbars can be split into substations to reduce the fault current by reducing the

interconnections between power networks. However, these two methods may reduce the

stability and reliability of power networks. Increasing the voltage level will reduce the


46

current that flows in the network for a given power capacity. However, this would lead

to increased system costs due to high voltage operation. High impedance transformers

and current limiting reactors are introduced to limit the fault currents by adding extra

impedance into the system. However, the inserted impedance will increase the voltage

drop across these devices, which will reduce the power quality of the system. All these

methods are not ideal since they reduce the performance of the system.

The condition based impedance increase technique only inserts impedance into the

network when a fault occurs. There are three traditional ways to achieve this: FCL

circuit breakers, fuses and Is-limiters. FCL circuit breakers made of baffle plates in

series are hard to design above 1kV because the arc voltage drop between the two plates

is only tens of volts [40]. Fuses and Is-limiters are widely used up to medium voltage

levels of 36kV. However, they have to be replaced manually after each fault operation.

Compared with these traditional methods, the new concepts of using superconductors

and/or semiconductors have the advantages of fast response time and automatic

recovery. They can also be designed for higher voltage levels.

2.2.5 Non-superconducting fault current limiters

2.2.5.1 Solid-state FCL

Solid-state FCLs perform fault current limiting by controlling semiconductors. The

schematic circuit of a solid-state FCL using a diode bridge circuit and an integrated gate

commutated thyristor (IGCT) is shown in Figure 2.7 [2, 41]. The operating principle is

as follows: during normal operation, AC current is commutated by the diode rectifier

and flows through the IGCT in one direction. Once a fault occurs, the IGCT usually

turns off in less than 1ms and the current is automatically diverted into a limiting

resistor. This topology only uses one turn-off IGCT. However, the current flows

through three semiconductors (two diodes and one IGCT) in series during normal

operation which causes on-state losses in the semiconductors.

A solid-state FCL without a turn-off device is shown in Figure 2.8 [2, 42]. During

normal operation, current flows through the main thyristors T1 and T2 in turn. When a

fault occurs, auxiliary thyristors T1a and T2a are usually triggered within 1ms, which


47

results in the pre-charged capacitors C1 and C2 starting to discharge. This discharge

current reduces the current through the main thyristors T1 and T2, and when the current

reduces to zero they are turned off. The fault current is then commutated into C1 and C2

and eventually into the limiting resistors R1 and R2 after C1 and C2 are fully

discharged. Thyristors T1a and T2a are turned off at the next natural current zero-

crossing. This design uses cheap thyristors instead of an IGCT and has low on-state

losses compared with the previous topology because current only flows through one

thyristor during normal operation. However, capacitors C1 and C2 must be sufficiently

large to give the main thyristors enough time to recover and regain their forward voltage

blocking ability.

Figure 2.7 Solid-state FCL with a turn-off IGCT

Vac

Load

Circuit

breaker

+

T2

T2a

C2 L2

D2 R2

+

C1

D1R1

L1 T1a

T1

Figure 2.8 Solid-state FCL with thyristors


48

2.2.5.2 Hybrid switching FCL

A hybrid switching FCL combines fast-acting mechanical switches with semiconductors

to reduce the number of semiconductor devices [2, 43], and is shown in Figure 2.9.

Figure 2.9 Hybrid switching FCL

During normal operation, all switches are closed and the current flows through the main

contact S1 because it has the lowest impedance with negligible losses. In the event of a

fault, S1 is opened (within 0.15ms) forming an arc across the contacts and the current is

diverted into IGCTs G1 and G2. G1 and G2 are turned off thereafter forcing the current

into a current limiting resistor R. S2 then opens without arcing to isolate G1 and G2

from the transient recovery voltage. This topology eliminates the on-state losses of the

semiconductors by placing a mechanical switch S1 in parallel. However, this system is

very complex and it is difficult to scale up to transmission level voltages.

2.2.5.3 Thyristor-controlled resonant FCL

A thyristor-controlled resonant FCL inserts a resonant inductor and capacitor or an

inductor to limit the fault current. Figure 2.10 shows the schematic circuit of a thyristor-

controlled resonant FCL [44, 45]. During normal operation, thyristors T1 and T2 are not

conducting and capacitor C provides series compensation in the power network. When a

fault occurs, thyristors T1 and T2 are triggered and inductor L is inserted into the

network in parallel with the capacitor. The parallel resonant LC circuit reduces the fault

current. This circuit may cause transient overvoltage, which is one of the concerns

relating to power quality.


49

Figure 2.10 Thyristor-controlled resonant FCL

A series-parallel resonant FCL is one of the modifications of the resonant FCL, and is

shown in Figure 2.11 [44]. Inductor L1 and capacitor C are tuned to resonate at the

supply frequency. During normal operation, thyristors T1 and T2 are in the off state and

the voltage drop across the FCL is negligible. When a fault occurs, thyristors T1 and T2

are fired and the parallel resonant circuit (inductor L2 and capacitor C) in series with

inductor L1 will reduce the fault current. This circuit performs differently from the

original resonant FCL because it is initially a serial circuit.

Figure 2.11 Thyristor-controlled series-parallel resonant FCL

A series resonant FCL shown in Figure 2.12 is another modification of the resonant

FCL [44, 46]. Inductor L and capacitor C are tuned to resonate at the supply frequency.

During normal operation, the thyristors are not conducting and the voltage drop across

the FCL is negligible. In the event of a fault, the thyristors are turned on, which short-

circuits the capacitor C, leaving the series inductor L to limit the fault current. These

two modifications of the resonant FCL demonstrate improved power quality and

removal of high frequency oscillations from the bus voltage [44].


50

Vac

C

T2

T1Circuit

breakerL

Load

Figure 2.12 Thyristor-controlled series resonant FCL

There are other non-superconducting FCLs such as polymer PTC resistor FCLs [47, 48]

and magnetic FCLs [49], but they will not be discussed here.

2.2.6 Superconducting fault current limiters

An SFCL is able to operate closer to an ideal FCL due to its non-linear characteristics.

Many different structures of SFCLs have been proposed and investigated after the

discovery of HTS. The various types are discussed in the following sections.

2.2.6.1 Resistive SFCL [2]

A resistive SFCL is the simplest and most compact SFCL design, which directly utilises

the characteristics of superconductors. The schematic circuit of a resistive SFCL is

shown in Figure 2.13. A resistor or inductor is normally placed in parallel with the

superconducting element to avoid overvoltage if the resistance of the superconductor

increases too rapidly after a fault occurs.

During normal operation, the superconducting element carries all the current and has

negligible resistance. Once a fault occurs, the current through the superconductor

increases quickly. When the current density in the superconductor exceeds the critical

current density level, the superconductor quenches and develops resistance, which limits

the current. The resistance of the superconductor increases quickly and therefore a high

percentage of the fault current is diverted into the parallel resistor or inductor, which

helps to limit the fault current. The fault will be completely cleared by the circuit

breaker usually within 200ms.


51

Advantages:

• Compact structure, simple design and light weight.

• The superconductor automatically quenches after a fault occurs and a trigger

mechanism is therefore not essential.

• It is intrinsically fail safe. The superconductor will burn out if it fails.

Disadvantages:

• AC losses occur in the superconductor during normal operation and Joule losses

are dissipated in the superconductor after it quenches.

• Hot spot problems. It is impossible to guarantee all parts of the superconductor

quench at the same time. The weakest points will start to quench first and may

develop into hot spots.

• Long recovery time. Due to the heat dissipated and temperature rise of the

superconductor, it may take several seconds to several minutes to recover.

• Current leads are needed between external ambient conditions and the cryogenic

interface, which produce extra losses and may introduce a thermal insulation

problem.

Figure 2.13 Resistive SFCL [2]

ACCEL/Nexans lead a project called CURL 10/110 to develop resistive SFCLs using

BSCCO 2212 bulk [4, 50, 51]. A BSCCO 2212 bifilar coil with a bypass metallic

conductor was manufactured using a melt, cast and processed (MCP) technique [50]. In


52

2004, Nexans installed a CURL 10 resistive SFCL in the grid for RWE in Germany for

testing and evaluation [51]. This system demonstrated effective current limitation

during the period of field testing [4]. From 2005, Nexans started to design a resistive

SFCL for high voltage power system applications. A concept design of a single phase

110kV/1850A resistive SFCL using BSCCO 2212 tubes with magnetic field assisted

quench was carried out [52].

Thereafter, Nexans designed and manufactured three resistive SFCLs for commercial

applications. Nexans manufactured a 12kV/2MVA resistive SFCL and Applied

Superconductivity Limited (ASL) designed the closed-loop cryogenic system. This

device was installed at Bamber Bridge, U.K. in 2009, which was the first commercial

system worldwide [53, 54]. In the same year, a 12kV/16MVA resistive SFCL was

installed in the auxiliary power supply at Boxberg, Germany, which was the first SFCL

installed in a power station [53, 54]. Another 12kV/9MVA resistive SFCL is presently

undergoing installation in Ainsworth Lane, U.K. Nexans also started to develop the

SFCL with YBCO coated conductors under the financial support of a German

government project called ENSYSTROB. A resistive 12kV/800A SFCL with YBCO

tape was successfully tested at the high power test lab in Berlin in September 2011 and

has been subsequently installed at the Boxberg power plant for field tests [8, 53, 55]. As

part of an European Union (EU) program called ECCOFLOW, a resistive 24kV/1005A

SFCL, which is designed for two applications, is presently under development [53, 55].

SuperPower developed a matrix fault current limiter (MFCL) using MCP BSCCO 2212

tube supplied by Nexans for transmission level voltages of 138kV and above [11, 56]. A

trigger matrix is connected in series with a current-limiting matrix. The MFCL uses the

trigger matrix to produce a magnetic field to quench the current-limiting matrix when a

fault occurs, which prevents hot spot problems. After the prototype was successfully

tested, SuperPower started to investigate a pure resistive SFCL with a YBCO coated

conductor [57]. SuperPower demonstrated that it is possible to develop this for both

distribution and transmission level voltages using a modular design. For example, a

distribution level SFCL of 11-15kV/800-2000A needs three SFCL modules, whilst a

transmission level SFCL of 138kV/1700A needs 14 SFCL modules [58].


53

CESI Ricerca in Italy has manufactured and tested a prototype 500kVA resistive SFCL

using multifilamentary Ni-sheathed MgB2 tapes supplied by Columbus Superconductor.

The prototype showed good current-limiting characteristics with no degradation after

repeated fault tests [59]. In 2010, a 9kV/4MVA SFCL using BSCCO 2223 was

developed and installed at the outgoing feeders of the San Dionigi substation in north

Italy for field testing. The three-phase SFCL has demonstrated a successful short circuit

test [10]. A 9kV/15MVA SFCL for the incoming feeders using a YBCO coated

conductor is currently under development. The design and simulation results suggest

that it is feasible [60].

Siemens initially investigated a resistive SFCL using YBCO thin films [61]. In 2005,

Siemens and AMSC formed a partnership to investigate YBCO coated conductor for

resistive SFCLs [62]. A single phase 7.5kV/300A SFCL using AMSC 344S 2G HTS

wire was developed and successfully tested. This SFCL utilised a bifilar coils design to

cancel the magnetic field, which then reduced the inductance [9]. A project to develop a

138kV SFCL ‘SuperLimterTM’ for a demonstration project at Southern California

Edison (SCE) power network has been sponsored by the U.S. Department of Energy

(DoE) [1]. A hybrid resistive SFCL using modular superconducting elements has been

proposed. The superconducting modules are connected in series with a fast operating

switch and a reactor is connected in parallel with the modules and the switch [1, 20].

The superconducting modules reduce the fault current for the first three cycles. After

three cycles of fault current, the fast operating switch removes the superconducting

elements from the circuit and the fault current is limited by the reactor. This method

reduces the amount of heating in the superconducting modules [54]. A single-phase

prototype has been tested successfully; however, the three-phase system has been

cancelled due to lack of funding [55].

Tokyo Denki University and the Japanese National Institute of Advanced Industrial

Science and Technology have been investigating an SFCL using a vacuum interrupter

driven by an electromagnetic repulsion switch [63, 64]. A vacuum interrupter is

connected in series with the superconducting element and a coil is connected in parallel

with both of them. When a fault occurs, the superconducting element starts to quench

and the fault current transfers into the coil. The coil has two functions: firstly, it acts as


54

a limiting coil; and secondly, it drives the electromagnetic repulsion plate (EPR) which

is fixed onto a moveable contact to open the vacuum interrupter quickly. This design

can reduce Joule losses dissipated in the superconducting element and hence reduce its

recovery time. A prototype SFCL using YBCO thin film was constructed and examined,

interrupting the fault current of the superconductor within a half-cycle after the

modification of the parallel coil [21]. The EPR would return to the closed position when

the fault current was completely interrupted by a back-up circuit breaker. An

electromagnetic repulsion switch for operating with the reclosing scheme, which

consists of an electromagnetic repulsion plate, cylinder, check-valve and intake vent,

was investigated. The reclosing time of the vacuum interrupter can be adjusted by the

intake vent. The recovery characteristics of AMSC 344S was examined and showed that

the recovery time increases as the voltage of the superconducting element increases.

The prototype with an electromagnetic repulsion switch demonstrated expected

adjustable reclosing behaviour [65, 66].

Hyundai and Yonsei University in Korea developed a 13.2kV/630A resistive SFCL

using AMSC 344S YBCO coated conductor. The superconducting coil effectively

limited the short circuit fault current without damage [67]. The Korea Electric Power

Research Institute (KEPRI) and LS Industrial Systems (LSIS) are leading a project to

develop a 22.9kV/630A hybrid resistive SFCL [68]. This hybrid resistive SFCL consists

of a superconducting coil, a fast operating switch, a driving coil and a current limiting

reactor [69]. In the event of a fault, the HTS becomes resistive and the current flows

through the driving coil, which produces a force to open the fast switch. The current is

subsequently diverted into the reactor to reduce the fault current. The hybrid resistive

SFCL reduces the amount of superconductors and hence the cost and volume of the

SFCL [70]. The SFCL was constructed and preliminary tested before placing it in the

Gochang Power Testing Centre of Korea Electric Power Corporation (KEPCO) grid in

November 2009 [70]. The SFCL is still under long-term operation and testing. The fault

and reclosing tests have been successfully conducted [71]. The protection parameters of

the network have been recalculated taking the SFCL into consideration. Short circuit

tests prove that an SFCL can be used in addition to the power system protection by the

adjustment of the parameters of the protection devices [72]. The final target is to

develop a 22.9kV/3kA hybrid resistive SFCL for general applications [73].


55

2.2.6.2 Bridge type SFCL [2, 5, 74]

A bridge type SFCL is made up of a bridge rectifier, a DC voltage source and a limiting

coil, as shown in Figure 2.14. The level of DC current through the limiting coil, which

is supplied by the DC voltage source, is designed to be higher than the peak nominal

current of the power system. During normal operation, the DC voltage source keeps all

four diodes forward biased. The inductor therefore is bypassed from the power system.

When a fault occurs, if the AC current increases higher than the DC current, diodes D3

and D4 or D1 and D2 would be reverse biased in the positive or negative half-cycle.

The limiting inductor will therefore be connected into the system to limit the fault

current.

The inductor does not have to be made of superconducting material, but

superconducting material can be used to minimise the losses. In addition, during normal

conditions, the inductor only carries DC current, which makes a superconductor an ideal

choice. Thyristors can be used to replace the diodes, so it is possible for them to turn off

the current at the next current zero-crossing after a fault occurs.

Figure 2.14 Bridge type SFCL

Advantages:

• No AC losses in the superconducting coil because it is operating with DC

current.

• Fast recovery after the fault clears because the coil remains in the

superconducting state during the fault.

• The trigger current level can be adjusted by the DC current source.


56

• Does not require a room temperature/cryogenic interface in the power line.

Disadvantages:

• AC losses in the semiconductors are relatively high.

• No fail safe mechanism. If one of the semiconductors fails and creates a short

circuit, the SFCL cannot limit the fault current.

The bridge type SFCL shown in Figure 2.14 was proposed by Boenig [5, 74]. The first

prototype 2.4kV/150A SFCL used BSSCO 2223 tape and was developed in 1993 [75].

After the success of the first prototype, another prototype 15kV/1.2kA SFCL was built

using thyristors instead of diodes. The thyristors can control the magnitude of the fault

current and interrupt the current completely. During high voltage testing in Southern

California Edison Centre Substation, a voltage insulation failure occurred in an auxiliary

piece of equipment and a three-phase test was therefore no longer possible [76].

The Chinese Institute of Electrical Engineering (CIEE) and the Chinese Academy of

Sciences (CAS) designed a 10.5kV/1.5kA bridge type SFCL using AMSC BSSCO 2223

tape [77, 78]. Four resistor/IGCT pairs were connected in series with a superconducting

coil instead of a traditional DC power supply for the bridge type SFCL. During normal

operation, the current flows through the bridge and the IGCT switches are turned on. In

the event of a fault, the first peak current is limited by the HTS coil. When the current in

the HTS coil reaches a pre-set level, the IGCTs are switched off, and the resistors are

inserted into the circuit. The fault current is then limited by the resistors and the HTS

coil [78]. After successful tests in the lab, the SFCL was installed in the Gaoxi

substation of Hunan in 2005 and subsequently demonstrated long-term reliable

operation in the real grid for one year. The SFCL reduced the fault current below 635A

successfully in a three-phase-to-ground short circuit test [78, 79].

2.2.6.3 DC biased iron core SFCL [2]

A DC biased iron core SFCL is shown in Figure 2.15. It consists of two iron coils,

which are driven into saturation by a DC supply. During normal operation, the iron

cores are fully saturated because the AC current is much lower than the DC current. The


57

inductances of L1 and L2 are small because they are similar to air core inductors. Once

a fault occurs, the increased AC current will drive coil L1 or L2 out of saturation in a

positive or negative half-cycle and the operating region of the core returns back to the

high permeability region. This causes the inductances of the coils to increase, which

reduces the fault current level.

Advantages:

• Inherently fail safe.



• No AC losses in the superconducting coil because it is operating with DC

current.

• Does not require a room temperature/cryogenic interface in the power line.

Disadvantages:

• The device is very bulky and heavy due to the iron core.

• AC losses in the primary windings and iron cores.

Figure 2.15 DC biased iron core SFCL

Zenergy Power has been investigating a DC biased iron core SFCL utilising BSCCO.

Under the support of the U.S. DoE, the first 15kV/1.2kA SFCL was installed and

successfully tested on the grid controlled by Southern California Edison since March

2009 [80]. The configuration of the iron core was assessed and several improvements


58

were undertaken, which led to a lighter and more compact design [54]. After successful

testing of the demonstrator SFCL, Zenergy Power built a 12kV/1250A SFCL for ASL,

which is presently installed in a CE Electric U.K. substation near Scunthorpe [19, 55]. A

138kV/1.3kA SFCL using the same concept was under development and was expected

to be installed in AEP Tidd substation belonging to American Electric Power in early

2012 [54]. However, the project was terminated in September 2011 due to several

administrative crises [55].

Innopower designed and tested a single-phase and a three-phase 380V prototype DC

saturated iron core SFCL in 2004 and 2005 respectively [6]. Innopower then developed

a 35kV/90MVA SFCL using the same structure [81, 82]. The SFCL consists of six

rectangular cores and a central core, which are all placed in a single cryostat. During

normal operation, a DC magnetisation coil drives the iron core into saturation. When a

fault occurs, the current in the DC magnetisation coil is removed by a switch. The

energy in the iron core is dissipated in an energy release circuit and the iron core returns

to its high permeability operating region. The DC magnetisation coil may be subjected

to a high back electromotive force (emf) due to the opening of the switch; a protective

circuit is introduced therefore to suppress the induced voltage [83, 84]. The SFCL was

installed in the Puji substation operated by the Southern China power grid in 2007 for

live grid operation [82, 85, 86]. The SFCL demonstrated encouraging current-limiting

performance. The artificial imposed fault test in the live grid was in good agreement

with the design expectations [87, 88].

2.2.6.4 Shielded iron core SFCL [2]

Figure 2.16 shows the scheme of a shielded iron core SFCL, which is made up of a

primary winding around an iron core with a superconducting cylinder in between. This

SFCL is also called an inductive SFCL because its structure is similar to a transformer

with a short circuit secondary winding. During normal operation, the current in the

superconducting cylinder is lower than its critical current and it screens all the flux from

the iron core. The impedance of the device, which consists of the resistance of the

primary winding and the stray inductance, is very low. In the event of a fault, the

current in the superconducting cylinder exceeds the critical current and the cylinder

starts to develop a resistance. The magnetic flux penetrates into the iron core, so the


59

inductance of the primary winding increases. The equivalent impedance of the device

becomes the inductance of the primary winding and the referred cylinder resistance to

the primary in parallel.

Superconducting cylinder

Primary winding Cryostat

Iron

core

Figure 2.16 Shielded iron core SFCL

Advantages:

• Fail safe. If the superconducting cylinder burns out, the magnetic flux will

penetrate into the iron core, which increases the inductance of the primary

winding to limit the fault current level.

• Does not require current leads for the room temperature/cryogenic interface.

Disadvantages:


superconducting cylinder, it may take several seconds to several minutes to

recover.

• Large size and heavy weight, which limits its scalability.

• AC losses in the primary winding.


60

2.2.6.5 Fault current controller SFCL [2]

A fault current controller SFCL consists of two anti-series connected thyristors, with

each thyristor connected with a superconducting coil inductor in parallel, as shown in

Figure 2.17.

Figure 2.17 Fault current controller SFCL

The thyristors are triggered at their respective peak load current and then constant DC

currents circulate through the parallel thyristors and inductors. The triggering signals are

removed after the thyristors are turned on. During normal operation, the AC current is

lower than the constant DC current, so that the thyristors are conducting and short-

circuit the inductors. When a fault occurs, the AC current rises above the constant DC

current. When the current is in the positive half-cycle, the current through thyristor T2

will go across zero and then T2 will turn off. Inductor L2 is inserted into the circuit to

reduce the fault current. Thyristor T1 will subsequently turn off at the negative half-

cycle and insert inductor L1 into the circuit.

Advantage:



Disadvantages:

• AC losses in the semiconductors and superconducting coil.

• No fail safe mechanism. If one of the semiconductors fails and creates a short

circuit, the SFCL cannot limit the fault current.


61

2.2.6.6 Flux-lock type SFCL[7, 89]

A flux-lock type SFCL consists of a flux-lock reactor with a superconducting coil L3

and a magnetic field coil. A schematic circuit of a flux-lock type SFCL is shown in

Figure 2.18. During normal operation, the superconducting element is in the

superconducting state, so the voltage across it is zero. The magnetic flux linkage

through the iron core is constant in a DC mode and therefore the voltages across these

three coils are zero, and the impedance of the SFCL is negligible. Once a fault occurs,

the superconducting element loses its superconductivity and develops a resistance,

which reduces the fault current. The magnetic flux then varies in the iron core and the

induced voltage across the coils changes. The current flows in the magnetic field coil

and the external magnetic field is applied to the superconducting element, which causes

the resistance of the superconducting element to increase faster and relatively evenly

along its length.

Figure 2.18 Flux-lock type SFCL

Advantages:

• During a fault condition, the magnetic field coil applies the magnetic field to the

superconducting element to make it quench equally, which prevents hot spot

problems in the superconducting element.

• The superconducting element is isolated from the power line.

Disadvantages:

• The device is very bulky and heavy due to the iron core.


62

• AC losses in the three windings.


superconductor, it may take several seconds to several minutes to recover.

Some modified versions of the flux-lock type SFCL have been investigated such as

using two co-axial air core coils as an alternate [90]. A flux-lock type SFCL with two

triggering current levels produces more effective current-limiting using a second

superconducting element in case the initial transient component of the fault current is

large [91].

A summary of the main SFCL projects which are believed to be currently active is

provided in Table 2.1.

2.2.7 Summary

As the capacity of power networks increases, fault current limitation devices are of great

interest. The operating principle and applications of fault current limitation devices were

briefly introduced. The conventional methods of fault current limitation were then

summarised. The topologies, operating principles and projects of non-superconducting

FCLs and SFCLs were discussed. The advantages and disadvantages of different

topologies were also summarised.


63

Table 2.1 Summary of active SFCL projects

Country Lead Company Type Material Phase Test Rating Test Year Ref

Germany ACCEL/Nexans Resistive BSCCO 2212 bulk 3-phase 12kV, 600A 2004 [4]

Germany/U.K. Nexans/ASL Resistive BSCCO 2212 bulk 3-phase 12kV, 100A 2009 [53]

Germany Nexans Resistive BSCCO 2212 bulk 3-phase 12kV, 800A 2009 [53]

Germany/U.K. Nexans/ASL Resistive BSCCO 2212 bulk 3-phase 12kV, 400A 2011 [53]

Germany Nexans Resistive YBCO tape 3-phase 12kV, 800A 2011 [53]

EU Nexans Resistive YBCO tape 3-phase 24kV, 1005A 2012 [53]

Germany/U.S.A. Siemens/AMSC Resistive YBCO tape 1-phase 7.5kV, 300A 2007 [9]

Germany/U.S.A. Siemens/AMSC Hybrid Resistive YBCO tape 1-phase 138kV, 1.2kA 2011 [55]

Italy CESI Ricerca Resistive BSCCO 2223 tape 3-phase 3.2kV, 220A 2005 [60]

Italy CESI Ricerca Resistive MgB2 tape 1-phase 397V, 96A 2006 [59]

Italy ERSE Resistive BSCCO 2223 tape 3-phase 9kV, 250A 2010 [10]

Italy ERSE Resistive YBCO tape 3-phase 9kV, 1kA 2012 [10]

U.S.A. SuperPower Resistive BSCCO 2212 bulk 1-phase 8.6kV, 800A 2004 [11]

U.S.A. Zenergy DC biased iron core BSCCO 2223 tape 3-phase 12kV, 1.2kA 2009 [80]

U.S.A. Zenergy DC biased iron core BSCCO 2223 tape 3-phase 12kV, 1.2kA 2011 [19]

China CAS Diode bridge BSCCO 2223 tape 3-phase 10.5kV, 1.5kA 2005 [78]

China Innopower DC biased iron core BSCCO 2223 tape 3-phase 35kV, 90MVA 2007 [92]

China Innopower DC biased iron core BSCCO 2223 tape 3-phase 220kV, 300MVA 2010 [10]

Korea Hyundai Resistive YBCO tape 1-phase 13.2kV, 630A 2007 [67]

Korea KEPRI/LSIS Hybrid resistive YBCO thin film 3-phase 22.9kV, 630A 2007 [69]

Japan Toshiba Resistive YBCO tape 3-phase 6.6kV, 72A 2008 [93]


64

2.3 Vacuum interrupter

2.3.1 Introduction

Air, oil, SF6 and vacuum are commonly used in circuit breakers. For medium voltage

applications, SF6 and vacuum have already replaced the previous air and oil solutions

and have become the most widely used types of circuit breaker. Vacuum as an

insulation medium for circuit breakers was first reported in 1926 by Royal Sorensen and

Mendenhall [94]. In 1962, General Electric developed the first power vacuum

interrupter capable of interrupting 12.5kA at 15.5kV. In the U.K., a vacuum interrupter

capable of interrupting 15.3kA at 13.2kV was tested in 1967. In the following year,

33kV vacuum interrupters were manufactured in both the U.S.A. and the U.K. Currently,

ABB commercially supplies a vacuum circuit breaker, the VD4, rated up to 12kV/4kA.

The rated breaking current is up to 63kA [95]. Schneider produces the Powersub™

vacuum substation circuit breaker which is rated at 15-38kV/600-4000A. This vacuum

interrupter can break a current up to 40kA and withstand a basic impulse level (BIL) of

110-150kV [96]. A 126kV single break vacuum interrupter with an axial magnetic field

contact is presently under research in China [97-99].

Vacuum interrupters are capable of carrying and breaking currents under both normal

and fault conditions. They are widely used in medium voltage applications because of

the following advantages [100]:

• Maintenance free design and high reliability.

• Self-contained and environment compatible. There is no risk of explosion or fire

and external effects during breaking.

• Compact and small volume due to high dielectric strength of vacuum.

• Comparatively low erosion of contacts leads to long electrical life span.

• Short contact stroke, which allows simple operating actuator and requires

relatively low mechanical energy to operate.

• Is capable of breaking a fault current with a high rate of change of current with

time (di/dt) and withstand transient recovery voltage because of the fast

dielectric recovery capability.


65

2.3.2 Vacuum interrupter structure

The structure of a vacuum interrupter is shown in Figure 2.19. It is usually made up of a

pair of contacts, bellows, shields and a vacuum enclosure [94, 101]. When the current

rating of the vacuum interrupter is lower than a few kA, cylindrical butt contacts are

commonly used [94, 101]. For higher current ratings, spiral contacts or contrate contacts,

which can force the arc to rotate, are usually preferred. Most vacuum interrupter

contacts are made of copper chrome (CuCr).

Metal bellows are usually used in order to allow the movement of one of the contacts

inside the vacuum. One end of the bellows is connected to one end plate of the vacuum

interrupter. The other end is attached to the shank of the moving contact.

Shields are placed around the contacts and over the bellows. They are used to prevent

the metal vapour of the arc from condensing onto the vacuum enclosure and bellows.

Various materials such as stainless steel, nickel, nickel/cobalt iron alloys and copper

have been used for the shields.

Figure 2.19 Typical structure of a vacuum interrupter [101]


66

The vacuum enclosure includes two end plates and insulation between them. The end

plates are usually made of stainless steel. The insulation between the two end plates in

early manufactured vacuum interrupters was made of glass. Nowadays, ceramic

materials such as alumina are generally the most commonly used material for insulation.

2.3.3 Vacuum arc [94]

When the current carrying contacts separate, a metal vapour arc is developed in the gap

between the two contacts. The diffuse mode and the constricted mode are the two main

forms of the vacuum arc [94].

When the current is relatively low, from a few amps to a few kA, the vacuum arc is in

the diffuse mode. A cathode spot is a small area that emits vaporised cathode contact

material. One or several spots, separated from each other, are spread on the cathode

contact. The number of cathode spots is determined by the magnitude of the current and

the contact materials. In a half-cycle of alternating current, the number of cathode spots

increases when the current rises and then decreases when the current reduces. When the

current approaches zero, only one cathode spot is left. The cathode spots are a conical

shape with the peak at the cathode and in continuous motion across the contact surface.

The cathode spots carry high current in small surface areas, which produce extremely

high electric fields and local temperatures and hence lead to field effect emission. The

gap between the contacts is filled with a neutral plasma column, which is made up of

metal vapour electrons and ions emitted from the cathode contact. Most of the metal

vapour is condensed on the anode, except some which escapes from the contact and

freezes on the shields [102].

The vacuum arc voltage is low in the diffuse mode, which is concentrated across the

cathode region. The arc voltage increases slowly as the current increases. The voltage

drop of the plasma column increases with the length of the gap and the current level.

The voltage drop of the plasma column is however much lower than the cathode voltage

drop, and therefore the length of the gap has little influence on the arc voltage. The

vacuum arc has a positive resistance characteristic, which allows several cathode spots

exist at the same time. The erosion of contacts is low in the diffuse mode because most


67

of the emission electrons and ions are collected and recombined on the anode. In

addition, the polarity of the cathode and anode contacts is alternating due to AC current.

As the current increases, the energy dissipated in the vacuum interrupter increases

because both the current and arc voltage are increasing. This increases the heat flux to

the contacts and the shields. Once an anode spot establishes due to the focusing of the

current, the vacuum arc is transitioned into the constricted mode. The current level at

which the transition occurs depends on the material and size of the contacts. During the

transition from the diffuse mode to the constricted mode, the arc voltage increases with

a significant noise component. After the anode spot develops, the arc voltage drops and

becomes stable. In the constricted mode, the current is concentrated in a limited surface

area on the contacts, which causes localised overheating and considerable vaporisation.

The plasma in the gap then contracts into one column. The anode spot is playing an

active role in the discharge. The high metal vapour density causes the pressure of the

gap to be close to atmospheric pressure and therefore the vacuum arc in the constricted

mode shows similar characteristics to a high pressure arc. The constricted mode of the

vacuum arc produces considerable erosion of both contacts; one or two orders greater

than the diffuse arc.

2.3.4 Vacuum breakdown

Breakdown in a vacuum interrupter often occurs following arc interruption. It is hoped

that the dielectric strength of the contact gap is recovering, so that it can hold off the

transient recovery voltage, which the circuit is trying to apply across the contact gap. If

there is no electric field in the gap, the atoms or molecules of a gas are in random

motion and are frequently colliding with each other. These collisions are elastic because

the energy is preserved during the collisions. A small number of charged particles are in

random motion and they may recombine when they collide. When an electric field is

applied, the charged particles receive a force from the electric field and they accelerate,

so that their kinetic energy increases. The average distance between collisions is defined

as the mean free path (MFP). As the strength of the electric field increases, the charged

particles obtain more and more energy. It is possible that with the long mean free path

the charged particles accumulate enough energy and will ionise atoms when they

collide. An electron will be removed from the atom during a collision, leaving a positive


68

ion behind. The electrons produced from the collisions will accelerate and ionise more

atoms, which can produce a large amount of electrons. This process is referred to as

Townsend avalanche, which may eventually lead to breakdown in gases. There are two

essential conditions for breakdown in gases: firstly, there needs to be enough atoms in

the gap to be ionised; and secondly, the mean free path needs to be long enough for the

charged particles to accumulate enough energy for ionisation.

In practical applications, circuit breakers are designed to circumvent one of the

conditions necessary to increase the breakdown voltage. For example, an SF6 interrupter

operates on the right side (high pd) of Paschen’s curve [103], where p is the pressure in

the interrupter and d is the distance between the two electrodes. The pressure of the SF6

gas is high and hence a large number of particles are in constant collision with each

other. They cannot gain sufficient energy to ionise atoms during collision however

because of the short mean free path. Therefore, the higher the pressure in the SF6

interrupter, the higher the breakdown voltage.

The vacuum interrupter is another example which operates on the left side (low pd) of

Paschen’s curve. The population of charged particles and atoms in the gap are very

limited. The charged particles have the long mean free path and can acquire enough

energy, but they may not collide with other atoms before they condense on the contacts.

Paschen’s curve shows that the dielectric withstand increases rapidly when the pd is

lower than 10-2Pa·m. The pressure of the vacuum interrupter is usually lower than

10-3mbar and the distance between the contacts can be designed from 1cm to 10cm.

Long path breakdown is possible for the vacuum interrupter because of the increase in

pd. It is necessary therefore to monitor the vacuum level in the vacuum interrupter in

order to ensure its reliability. Permanent monitoring of the vacuum level is not possible

because it has to be removed to do so.

The contact surface conditions of a vacuum interrupter can seriously affect the dielectric

withstand of the vacuum interrupter. The contacts of the vacuum interrupter experience

welding when closing and suffer from arcing when pulled apart whilst conducting

current. Most of the metal vapour and particles condense on the contacts when operating.


69

The surface conditions of the contacts therefore change after operation, which may

significantly affect the ability to support a high dielectric strength between the contacts.

2.3.5 Current breaking in a vacuum

A vacuum interrupter is normally carrying current in the closed position during most of

its life time. It is designed to interrupt the current on command. The current interruption

is in fact rapid removal of the plasma which is conducting the current in the arc. The

plasma consists of metal vapour, metal ions, electrons, gas molecules, metal droplets

and particles. Most of the metal vapour are ionised and when they arrive at the anode,

they recombine with electrons and become metal atoms. The cooling of cathode spots,

which depends on the thermal conductivity of the contact material and atom evaporation,

is very rapid. Un-ionised metal vapour is not directly influenced by the electric field and

will eventually condense onto a cool surface. The production of the metal vapour relies

on the rate of erosion from the cathode. It has been proved that the metal vapour exists

for about 1µs in a 1cm contact gap [94]. The rate of current decrease at supply

frequency is much slower than the extinction of the metal vapour. A vacuum interrupter

therefore can easily break the arc in the diffuse mode and rapidly recover its high

voltage blocking capacity.

In the diffuse mode, when a vacuum interrupter breaks current at supply frequency, the

number of cathode spots decreases as the current decreases. When the current

approaches zero, the energy is not enough to maintain a high temperature at the root of

the arc and the arc becomes unstable. The last spot extinguishes and the current is

interrupted abruptly, which leads to a current chopping phenomena. Current chopping

may cause overvoltage because the current in the inductive load cannot stop suddenly

and there is no free-wheel path through a vacuum interrupter for the current after

chopping. The current is then commutated into the capacitance of the network, which

causes overvoltage.

When the arc is in the constricted mode, the current interruption is changed by the

anode spot, which continues to emit metal vapour after a current zero. In addition, the

temperature of the anode spot is relatively high and the positive ions attack it. It is

possible to build the cathode spot on the previous anode when the current alternates to


70

the next half-cycle. The removal of the constricted arc is much slower than the diffuse

arc; the rate of change of voltage with time (dv/dt) capability of a vacuum interrupter

therefore is much lower in the constricted mode.

In conclusion, the diffuse arc is easier to interrupt than the constricted arc, which

suggests that it is better to maintain the arc in the diffuse mode whenever possible. The

practical methods to keep the arc in the diffuse mode are described in the next section.

2.3.6 Practical design for high current levels

The vacuum arc can be affected by the applied magnetic field. In practical applications,

both radial magnetic field (RMF) and axial magnetic field (AMF) technologies are in

use to prevent the arc from passing into the constricted mode and improve the current

breaking capacity of a vacuum interrupter [101, 104].

A radial magnetic field can cause a rapid rotational movement of the arc, so that the

energy is distributed onto a large surface area of the contact. The radial magnetic field is

produced by the path imposed on the current in the contacts. Spiral contacts or contrate

contacts are widely used to rotate the arc by the self magnetic field of the current. The

root of the arc does not stay in the same location for a long period, which eventually

postpones the appearance of an anode spot. In addition, the vapour is equally spread

throughout the volume of the vacuum interrupter when it starts to recover. Furthermore,

the heat and vapour are uniformly distributed on the shields due to rotation which

ensures the shields do not have hot spots and the vapour can continue to condense on

the shields.

An axial magnetic field delays the formation of an anode spot because of its confining

effect, which helps in maintaining the arc in the diffuse mode. An AMF significantly

reduces the arc voltage and hence the power dissipated in the vacuum interrupter. The

application of an AMF is very important from a practical point of view. Firstly, it can

change the appearance of the arc; the plasma from the cathode spots becomes more

columnar and distinct when an AMF is applied. Secondly, the application of an AMF

reduces the arc voltage sharply, which would delay the formation of the anode spot.

There are several methods used to produce the axial magnetic field using the current


71

being interrupted such as integrating the coils behind the contacts or using an external

coil to surround the inter contact zone [101].

Both RMF and AMF technologies can offer a high current breaking capacity. These two

techniques have their own advantages and disadvantages. RMF can carry high

continuous current because of its low contact resistance, whilst AMF offers high voltage

rating and contact endurance [104].

2.3.7 Main application fields

Medium voltage levels are the primary application field for vacuum interrupters. The

vacuum interrupter and SF6 interrupter dominates the medium voltage level market. The

advantages of a vacuum interrupter over an SF6 interrupter are summarised in [101]:

• A vacuum interrupter is an enclosed circuit breaker, which adopts maintenance

free design with high electrical endurance.

• The dielectric recovery of a vacuum interrupter is faster than an SF6 interrupter,

which is especially suitable for a severe initial transient recovery voltage.

• A vacuum interrupter requires less operating energy than an SF6 interrupter.

A vacuum interrupter also has several disadvantages compared to an SF6 interrupter:

• A vacuum interrupter uses butt contacts, which need high contact pressure to

minimise contact resistance and avoid separation of the contacts.

• The butt contacts of a vacuum interrupter are usually made of CuCr which has a

higher resistance than the silver-plated multiple contacts of an SF6 interrupter.

• A vacuum interrupter is restricted by overheating.

In low voltage level applications, a vacuum interrupter cannot compete with an air

circuit breaker. The disadvantages which limit the application of a vacuum interrupter in

low voltage systems include:


72

• A vacuum interrupter has a low arc voltage, which cannot limit the fault current.

An air circuit breaker however has a high arc voltage, which can significantly

reduce the fault current level.

• The continuous current and fault current level in low voltage systems is higher

than medium voltage level applications. The butt contacts with a high contact

pressure and relatively high contact resistance are not suitable.

In conclusion, an air circuit breaker is a simple and economical technology widely used

in low voltage level applications. A vacuum interrupter is employed only in special

situations where enclosed breaking can offer significant advantages such as polluted or

explosive environments [101].

In high voltage level applications, a vacuum interrupter cannot compete with an SF6

interrupter. There are two main limitations of a vacuum interrupter in high voltage level

applications [101]:

• A lighting impulse voltage level in a vacuum interrupter is limited to 123-145kV.

• A vacuum interrupter has the possibility of emitting X-rays when the voltage is

higher than a hundred kV. Therefore, shielded enclosures are necessary to

protect operators. A vacuum interrupter is not a good choice in high field

applications.

For extreme conditions of high di/dt and initial TRV, a hybrid circuit breaker which

consists of a vacuum interrupter and an SF6 interrupter connected in series has been

investigated and tested [105, 106]. The vacuum interrupter withstands the initial TRV

and the SF6 interrupter endures the peak of the TRV. A prototype hybrid circuit breaker

has successfully demonstrated interruption of a short circuit current of 63kA at 145kV

[105].

2.3.8 Summary

Vacuum interrupters are widely used in medium voltage level applications and extreme

situations in low voltage and high voltage level applications. The development and

basic operating principles of the vacuum interrupter have been summarised. The two


73

main forms of the vacuum arc, the diffuse mode and constricted mode, were described.

Thereafter, the principle of vacuum breakdown and current breaking were outlined.

Finally, the practical design of vacuum interrupters for high current applications was

summarised.

2.4 Vacuum interrupter actuator

There are three main types of vacuum interrupter actuator in use today: the spring

actuator, solenoid actuator and permanent magnetic actuator. The voice-coil type

actuator, which has the advantages of quick response and low moving mass also has the

potential to be used as a vacuum interrupter actuator.

2.4.1 Spring actuator

A conventional spring actuator was commonly used when the vacuum interrupter was

first invented. It is rarely manufactured now because it has several key disadvantages.

Firstly, the spring dissipates its energy and provides lower force during the travel of the

moving contact whereas a vacuum interrupter requires a higher force when approaching

the open or closed position to stop bouncing. Secondly, a spring actuator typically has

approximately 150 mechanical components, many of which are not standardised parts

[107, 108]. During one operation, up to 30 mechanical components are moving. The

spring actuator therefore is extremely complicated with numerous moving parts, which

can cause reliability problems and it also needs regular maintenance.

2.4.2 Solenoid actuator

A solenoid actuator consists of three parts: a coil which is energised by a DC supply; a

magnetic core to provide the path for the magnetic flux and a moving plunger. Both

slow and fast driving speeds are possible through control of the DC supply to control

the current through the coil. As the plunger of the solenoid approaches the open or

closed position, the force on the plunger increases because the plunger decreases the

reluctance of the magnetic flux path as it moves. This would help prevent the vacuum

interrupter from bouncing [108].

The solenoid actuator however suffers from problems of reliability and regular

maintenance requirements. In addition, in order to hold the movable contact in the open


74

or closed position, the solenoid coil is required to be energised continuously or have

additional mechanical latches installed. The former method dissipates electric energy

continuously. The latter method requires extra energy to release the latches and

therefore the structure is complex, which would further reduce the reliability of the

mechanism. Moreover, the dynamic characteristics of a solenoid actuator are very

difficult to predict due to magnetic flux leakage, eddy currents and saturation, which

makes the system complex and highly non-linear [109].

2.4.3 Permanent magnetic actuator

In 1987, Manchester University Energy Systems Group developed a new type of

permanent magnetic actuator for operating a medium voltage level vacuum interrupter

[110-112]. Thereafter, the permanent magnetic actuator has drawn great interest due to

its simple structure compared with the conventional type of spring actuator [107, 108,

112-115].

The structure of a permanent magnetic actuator is shown in Figure 2.20. It consists of

seven parts: a stationary iron core, a movable armature, two pieces of permanent magnet,

a shaft, a closing and an opening coil [107, 111]. Compared with a spring actuator, the

total number of parts in the permanent magnetic actuator is dramatically reduced. The

reliability therefore is significantly improved and the permanent magnetic actuator only

needs minor maintenance.

Figure 2.20 Typical structure of a permanent magnetic actuator


75

A vacuum interrupter with a permanent magnetic actuator is commercially available

today. For example, the VM1 vacuum interrupter with a permanent magnetic actuator

designed by ABB offers in excess of 100,000 operations [107, 116].

When the current is supplied to one of the two coils, the resultant magnetic field

produced by the current through the coil and the permanent magnets begins to change

and the force on the armature changes. The permanent magnetic actuator therefore will

open or close the vacuum interrupter. Energy is only required during the opening or

closing operation. After the actuator achieves its open or closed position, the power

supply can be removed. Without any current through the coil and any mechanical latch,

the armature can be held in the open or closed position by the magnetic force produced

by the permanent magnets.

Eddy currents however are induced in the armature of the permanent magnetic actuator

in response to the changes in the current flowing in the coil. The eddy currents generate

the magnetic flux that opposes the magnetic flux produced by the coil current. As a

result, the resultant magnetic flux in the actuator is reduced, which may significantly

degrade the overall performance of the system. Thin laminated sheets of electric steel

are widely used to minimise eddy currents in motor applications [117]. However, the

speed of the permanent magnetic actuator could be very high, so it is not possible to

laminate the armature because it may reduce its structural rigidity.

2.4.4 Voice-coil type actuator

The voice-coil type actuator is an electromagnetic actuator that generally consists of one

or more coils placed in a magnetic field. The Lorentz force is produced on the coil when

current flows through the coil. This force is proportional to the magnetic flux density,

current through the coil and the length of the coil in the magnetic field [118]. This

actuator type has been known for decades and has been used as the source of force in

loudspeakers and the drive mechanism for disk drive read heads, for example [119, 120].

This actuator type can be a fast and powerful means for operating a vacuum interrupter.

It has the following advantages:


76

• Compact structure with a small number of parts and maintenance free.

• Light weight of moving part compared with the permanent magnetic actuator.

• High acceleration rate and fast response time.

The major disadvantage of this actuator is the same as a solenoid actuator, i.e. the coil

does not have an inherent stable position when the power supply is removed. There is a

need therefore for a latch mechanism. Magnetic latches, which include latch magnets to

provide the latching magnetic field and latch steel pieces fixed on the moving part can

be installed in the actuator directly [118, 121]. The actuator can therefore be latched in

the open or closed position without energy from an external supply. When the vacuum

interrupter is opening or closing, the force on the coil is higher than the latching force,

so that the actuator releases automatically.

2.4.5 Summary

Four different types of vacuum interrupter actuator, i.e. the spring actuator, solenoid

actuator, permanent magnetic actuator and voice-coil type actuator were introduced and

compared. The structure of the spring actuator and solenoid actuator is too complex,

which causes problems in terms of reliability and maintenance. Permanent magnetic

actuators provide reliable operation with a compact structure. However, the speed is

limited by the heavy armature and eddy currents. The voice-coil type actuator is widely

used in loudspeakers and disk drive read heads and shows great potential for use in a

vacuum interrupter. The benefits of the voice-coil type actuator include its simple

structure and maintenance free design.


77

2.5 Conclusions

This chapter began with a description of the properties of superconductors, which

exhibit zero DC resistivity below a critical magnetic field strength, critical temperature

and critical current density. Superconducting materials such as BSCCO, YBCO and

MgB2 used for FCLs were introduced including their raw materials, manufacturing

processes and properties.

The chapter then presented the necessity for fault current limitation devices because of

the increase in the fault current level in modern power networks. Conventional methods

of fault current limitation were summarised and the topologies and operating principles

of non-superconducting FCLs and SFCLs were highlighted. The advantages and

disadvantages of different SFCL topologies were summarised and compared. In

addition, the related SFCL projects using the above techniques were detailed. A vacuum

interrupter integrated into the resistive SFCL was proposed to solve its overheating and

recovery problem.

The general principles of the vacuum interrupter operation were summarised including

the development and basic structures of the vacuum interrupter; two main forms of the

vacuum arc - the diffuse mode and constricted mode; and the principles of vacuum

breakdown and current breaking.

Actuation mechanism including the spring actuator, solenoid actuator, permanent

magnetic actuator and voice-coil type actuator were also compared. The voice-coil type

actuator shows great potential for a vacuum interrupter and will therefore be

investigated in this thesis.

Chapter 3: SFCL Coil and Experimental Test Rig

78

3 SFCL Coil and Experimental Test Rig

3.1 Introduction

It is necessary to build the MgB2 SFCL coil before any experimental investigation can

be conducted. The design of an early prototype former will be introduced. The

manufacturing process of the SFCL coil and heat treatment process will then be

described. The instrumentation of the SFCL coil will be presented. The experimental

test rig including the test circuits, control system and cryostat system will also be

discussed in this chapter. The prototype former, instrumentation connection circuit and

experimental test rig were designed and built previously by Oliver et al. [15, 17].

3.2 Coil former

A prototype alumina former was designed to support the MgB2 wire. It was

manufactured by Dynamic Ceramic Ltd. using Dynallox 96 (96% Al2O3). The

specification and drawing of the prototype former is shown in Figure 3.1. The prototype

former shown in Figure 3.2 has a blackened appearance due to the earlier heat

treatment.

The selection of the former material was an important aspect of the design. Alumina

was chosen because it has the following advantages:

• Wide temperature range. The heat treatment of the alumina former with the

MgB2 coil is conducted at 700ºC, whilst the operating temperature is around

-243ºC. The temperature band is almost 1000ºC. The maximum operating

temperature of alumina is 1700ºC [122], which is higher than the temperature of

the heat treatment. The minimum operating temperature of alumina is not

available, but previous tests show that the former had no problem at the

cryogenic temperature [17].


79

• High thermal conductivity. The thermal performance of the former is a critical

factor. The thermal conductivity of alumina is 25W/m·K [122]. Alumina

therefore is able to conduct heat quickly during a fault test due to its high

thermal conductivity. This can help to keep the temperature of the MgB2 wire

below its melting point in a fault test and to recover quickly after the fault is

cleared.

• High dielectric strength and volume resistivity. The dielectric strength of

alumina is 14.6kV/mm [123] and its volume resistivity is higher than 1014Ω·cm

at 25ºC [122]. Alumina therefore can provide excellent electrical insulation for

the coil, which is important for high voltage applications.

Furthermore, the thermal contraction/expansion of alumina over the operating

temperature range needs to match that of the MgB2 wire, to prevent strain being placed

on the wire. The thermal expansion coefficient of alumina is 7.8×10-6K-1 [122]. Hyper

Tech Research who supplied the MgB2 wire indicated that the thermal expansion

coefficient of the MgB2 wire is 10.38×10-6K-1. The difference between the two thermal

expansion coefficients will cause interference between the MgB2 wire and the former of

about 7.224×10-4, which corresponds to about 0.48mm across the circumference when

cooled down to 20K from room temperature. Alumina therefore is a good material

choice for the MgB2 wire.

14.00mm

5.00mm

134.00mm

5.00mm

14.00mm

3.00mm

Figure 3.1 Former specification (side view) [17]


80

Figure 3.2 Prototype former

3.3 Coil manufacturing process

There are two methods for fabricating the MgB2 SFCL coil: wind-and-react and react-

and-wind [16]. In the wind-and-react method, the MgB2 wire is wound onto the former

and then heat treated whilst in the react-and-wind method, the MgB2 wire is first reacted

and then wound onto the former. The first approach was used because any cracks

resulting from the winding process may be healed when heat treated. A spool of wire

containing 15 metres of 1.28mm MgB2 wire in un-reacted form was delivered by Hyper

Tech.

3.3.1 MgB2 wire current connections and winding process

The SFCL coil was made up of two interleaved solenoid coils. Both coils were wound

from top to bottom but in opposite current directions. This design effectively reduced

the inductance of the coil by cancelling the solenoidal field created by the two coils.

There were four current connections available on the cryostat bottom plate. These

connections were connected to the copper bar connections at the top of the internal

copper bucket in the cryostat through thermally-anchored leads. The copper bar


81

connections (numbered 1 to 4) and the annular working space within the copper bucket

are shown in Figure 3.3.

Figure 3.3 Cryostat interior

Copper braid was used to connect the MgB2 wire to the copper bars inside the cryostat.

The benefit of copper braid is that it can cope with any expansion/contraction of the

wire during heat treatment and test in nitrogen. It was necessary to hold the MgB2 wire

in place after winding; copper clamps were machined therefore to attach the MgB2 wire

onto the copper braid. The copper clamp would not only hold the wire in place but also

provide the electrical connection. The copper clamp was loosely placed on the copper

braid, as shown in Figure 3.4, and then fitted into the channel at the top of the former

and folded back, to pass out vertically through the grooves in the top lip, as shown in

Figure 3.5.

The two ends of the copper braid were bolted to the copper connections inside the

cryostat. This arrangement was used at the top of the former for each of the two MgB2

wires forming the interleaved coils. At the bottom of the former, the current connections

for the coils were attached to another copper braid by copper clamps. The copper braid

was then joined into a single loop using a clamp.


82

Figure 3.4 Clamp used to connect wire to copper braid

Figure 3.5 Top copper braid current connection

The MgB2 wire was placed through the copper clamp leaving several centimetres

protruding past the clamp, to allow a soldered joint to be made after heat treatment. The

connection was made by pressing the MgB2 wire and copper braid between the two

clamp plates and screwing the plates together tightly. The former with the copper braid

was then placed on a lathe. The MgB2 wire was wound onto the former by turning the

lathe slowly. The other end of the MgB2 wire was then passed through the copper clamp

at the bottom of the former, again leaving some of the MgB2 wire protruding for a

soldered connection. This process was repeated for the MgB2 wire forming the second

coil. After winding, the copper clamps were fastened against the MgB2 wire and the

protruding portions of the stainless steel screws used in the clamps were removed and

filed flush with the outside of the clamp plate.

A silica rope was wrapped around the copper braid and the MgB2 wire at the top of the

former, and finally a stainless steel jubilee clip was placed over the silica rope and

clamped tightly. This helped to maintain the tension on the MgB2 wire during heat

treatment so that the complete former could be heat treated with the MgB2 wire in the

coil slots. The silica rope was pre-heated to burn out the cellulose.


83

3.3.2 MgB2 wire heat treatment

Hyper Tech specified that the MgB2 wire should be held at 700ºC for 20 to 40 minutes

to react the elemental magnesium and boron to form MgB2 [13, 16]. This heat treatment

process would also help to repair any cracks that may have occurred during winding.

Dynamic Ceramic advised that a maximum temperature gradient of 150ºC/hour should

be used for the ceramic former. A low temperature gradient is normally specified to

avoid an uneven heating that might cause the alumina to expand more rapidly on one

side than the other, which may create cracks. From previous experience, there was no

problem heating the former with a temperature gradient of 200ºC/hour. The former was

heated therefore in a vacuum oven with a gradient of 200ºC/hour until it reached 700ºC,

and then held constant in the oven at 700ºC for 30 minutes. After 30 minutes, the oven

was cooled down to 600ºC, and then flowing argon gas was used to cool the oven down

to room temperature with a gradient of 200ºC/hour.

Hydrocarbon impurities in the alumina reduced to carbon during the heat treatment

process, resulting in the blackened appearance. The former and coil after heat treatment

is shown in Figure 3.6 (left).

Figure 3.6 Coil after heat treatment (left) and soldered joints (right)

The protruding MgB2 wire ends either side of the copper clamps were then soldered to

the copper braid to provide extra connection length and strength. The soldered joints are

shown in Figure 3.6 (right).


84

Instrumentation connections were also mounted onto the coil. Voltage taps and BAS16

diode temperature sensors were soldered onto the MgB2 wire in several places. Four

Cryocon S700 commercial temperature probes were mounted on a piece of fibreglass

board before being attached to the former. The coil instrumentation is described in more

detail in section 3.4.

Kapton tape was wrapped round the entire former to provide voltage insulation between

the coil and the cryostat copper bucket. Heat shrink sleeving was also placed over the

exposed copper braid, again to provide voltage insulation between the braid and the

cryostat copper bucket. The manufactured prototype coil and former, ready to be

installed in the cryostat, is shown in Figure 3.7 (left). The former was finally installed

into the cryostat by lowering it into the annular working space and bolting the copper

braid connections onto the cryostat copper bar connections. A picture of the prototype

coil installed in the cryostat, along with the instrumentation wiring, which was brought

out of the top of the cryostat, is shown in Figure 3.7 (right).

Figure 3.7 Coil ready to be installed in the cryostat (left) and installed in the cryostat

(right)

3.4 Instrumentation

3.4.1 Voltage signals

Each of the two interleaved coils has three full turns and one part turn. The voltage of

each turn was measured by soldering voltage taps directly onto the MgB2 wire. The

exact locations of the voltage taps are described in the next chapter.


85

The connections for the voltage taps were brought out from the cryostat through a

socket on the top plate and connected to a multi-channel high precision linear

differential amplifier built by the electronics workshop. Each voltage channel was

provided with a variable gain setting which could be adjusted by changing a resistor.

High gain was used for the coil when it was superconducting and low gain was used

when the coil was expected to quench. The output from the voltage amplifier was

connected to a LabVIEW interface card. The instantaneous voltage signals were

recorded by LabVIEW during testing.

3.4.2 Current signals

Two Hall Effect current sensors were used to measure the current passing through the

coil. One was rated up to 100A, to provide an accurate current reading for low current;

the other one was rated up to 1000A, to measure high current. A power supply was

manufactured by the electronics workshop at the university to provide ±12V to the

sensors. Appropriate measurement resistors were connected to the outputs to convert the

current signals to voltage signals and then connected to the LabVIEW interface card.

3.4.3 Temperature signals

Twelve Cryocon S700 temperature probes were placed in the cryostat. These probes use

an industrial small bobbin package to provide good thermal contact to the diode

junction. The diameter of each probe is around 8mm. Four of these probes were located

on the four internal current leads connecting the external supply to the current

connections. Instrumentation wires for these four sensors were brought out of the

cryostat through the socket on the bottom plate and connected to a Cryocon 14

temperature monitor. Another four sensors were located on the outside surface of the

copper bucket at the top, middle, bottom and on the cryocooler cold head. These four

probes were connected to a Cryocon 34 temperature controller. The controller and

monitor front panels are shown in Figure 3.8.

The last four probes, which were mounted onto a fibreglass board, are shown in Figure

3.9. The board was then mounted onto the side wall of the former. The probes on the

left and right of the board fitted onto the top and bottom copper braids and MgB2 wire

joints, whilst the other middle two measured the temperature of the former.


86

The connections to these four probes were brought out of the cryostat through the socket

on the top plate and connected directly to the Cryocon 14 temperature monitor. The

monitor displays the four temperature readings on the front screen and provides one

temperature as an output. This output was connected directly to the LabVIEW interface

card.

Figure 3.8 Temperature monitor (top) and temperature controller (bottom)

Figure 3.9 Cryocon S700 silicon diode temperature probes

Accurate temperature measurement of the MgB2 wire is important in preventing the

wire from overheating. The Cryocon temperature probes were too large to fit into the

slot housing the wire. It was found however that a common diode could be used to

measure temperatures from 20K to 300K [124-126]. The forward voltage drop in the

p-n junction of a diode depends on the temperature at constant current. A constant

current power supply of 10µA was designed for the diodes and built by the electronics

workshop. The BAS16 diode temperature sensor with a SOT23 package was used to

measure the temperature of the MgB2 wire. The BAS16 diode was chosen due to its


87

low-cost and small size. One spare pin on the diode package was soldered onto the

MgB2 wire to provide good thermal contact. One of the diodes soldered onto the MgB2

wire in the slot is shown in Figure 3.10. BAS16 diodes from the same manufacturer had

similar characteristics with temperature. However, the accuracy can be significantly

improved by individual calibration [124, 125]. All the diodes therefore were calibrated

before carrying out any tests.

The connections to these sensors were brought out of the cryostat again through the

socket on the top plate and connected to an isolated operational amplifier circuit. The

output from the amplifier was connected to the LabVIEW interface card. The

temperature was recorded by LabVIEW during and after testing.

Figure 3.10 BAS16 diode temperature sensor soldered onto the MgB2 wire

3.5 Control

3.5.1 High-current test circuit

A controllable high current supply was built to test the quench behaviour of the SFCL

coil. A schematic of the high-current supply is shown in Figure 3.11. Switch control and

data acquisition was provided by the LabVIEW system. The switch was controlled

directly from LabVIEW to turn on and off. Voltage, current and temperature signals

were also monitored and recorded using the LabVIEW data acquisition system.

A variac was directly connected to the laboratory supply. The output of the variac was

connected to a current step-up transformer with a turns ratio of 4:1 through the point-on-

wave switch. This switch was manufactured by the electronics workshop. The switching

point was at current zero during each test. The switching operation was performed using

a back-to-back thyristor arrangement. A gate signal was sent by LabVIEW to close the

switch. The switch would be turned off at the next current zero-crossing after removal


88

of the gate signal. Furthermore, the switch outputted a single pulse for each cycle,

which was read into LabVIEW to control the number of current cycles passed through

the coil during testing.

Figure 3.11 High-current test circuit schematic [15, 17]

A variable load resistor was placed in series with the SFCL coil in the secondary side of

the transformer to represent the system impedance during a fault. This resistor could be

varied from 0 to 0.45Ω. The maximum current supplied to the coil, with the variac set at

100% and the load resistance at 0Ω, was estimated to be 1053Apeak. It should be noted

however that this level of current would not be reached in practice because the coil

would quench before this current level was reached.

The test supply was constant voltage rather than constant current. The coil would act

therefore as an SFCL when transitioning from the superconducting state to the normal

resistive state. The impedance of the coil was assumed to be negligible in the

superconducting state. The ratio of the load resistor to the quenched coil resistance

would directly determine the current limitation.

3.5.2 Low-current test circuit

A low-current supply was made by the electronics workshop to provide a constant low

current (up to 3Apeak) signal. A schematic of the low-current supply is shown in Figure


89

3.12. The low-current supply produced a variable frequency and waveform signal. This

could be used for a frequency sweep to determine the MgB2 coil inductance.

Figure 3.12 Low-current test circuit schematic [15, 17]

The low-current supply was controlled by a gate signal from LabVIEW, which turned

the supply on and off at the next current zero-crossing. Again the unit outputted a single

pulse for each cycle of supply that was read into LabVIEW to control the number of

current cycles passed to the coil.

3.5.3 LabVIEW control programme

All the data acquisition and control was performed using a LabVIEW embedded

system. The graphical user interface is shown in Figure 3.13. The number of current

cycles to pass through the coil could be set by ‘Current cycle to test’ on the left top

corner of the screen. When the ‘Run’ button was pressed, LabVIEW sent an ‘On’ signal

to the switch and the current supply was connected to the coil at the next voltage zero-

crossing.

When the programme was running, the ‘cycle counter’ output from the switch was read

in by LabVIEW and incremented. All the voltage channels, current sensors and diode

temperature sensor signals were also read in at a rate of 2000 samples per second during

a test. These signals were displayed in charts on the screen and saved to a text file on

the PC for post signal processing.


90

When the number of current cycles was equal to the set value, LabVIEW sent an ‘Off’

signal to the switch. Current flow in the coil would then be turned off at the next current

zero-crossing. All the signals were still read in and recorded at a rate of 10 samples per

second until the ‘Stop post test recording’ button was pressed. The diode temperature

sensor signals from ‘Temperature post test’ would be used as an indication of the

thermal recovery of the MgB2 wire.

Figure 3.13 Screenshot of LabVIEW control programme

3.5.4 Cryostat

The cryostat system was designed and built by Scientific Magnetics [127]. This system

included five major parts: a vacuum insulated cryostat, a cryostat top lid, a cryocooler

compressor, a vacuum pump set, temperature monitors and a controller.

The cryostat consisted of an aluminium vacuum vessel surrounding a cylindrical copper

bucket. As shown in Figure 3.3, a central copper cylinder was mounted inside the

copper bucket at the bottom. The annular space between them was the working space

for the former. A polystyrene cylinder was placed inside the central copper cylinder to

reduce the amount of liquid nitrogen required. A fibreglass neck tube was connected


91

between the copper bucket and the cryostat top flange. The fibreglass tube is not

designed to be placed in compression and could be damaged therefore if the vacuum is

not established before filling with liquid nitrogen. The vacuum level of the vacuum

vessel should be better than 1×10-3mbar before the liquid nitrogen could be poured into

the bucket [127].

Figure 3.14 Cryostat system: vacuum cryostat vessel with vacuum pump set (left),

vacuum pump (top middle), gauge (bottom middle) and cryocooler compressor (right)

The copper current connections were located around the top lip of the copper bucket, as

shown in Figure 3.3 (numbered 1 to 4), to which the coil current connections were

made. The internal current connections were connected to the external current

connections on the bottom plate by internal leads through the vacuum vessel. These

leads were thermally anchored to the cryocooler cold head to prevent heat transfer from

ambient temperature. The bottom plate of the cryostat contained connections for

instrumentation such as temperature probe connections, heater control, and the four

external current connections. A 50W heater was located at the top of the cryostat copper

bucket. This heater could be controlled to maintain a given temperature at any of the

four temperature sensors connected to the temperature controller. The heater power was

controlled by a PI control loop.


92

Figure 3.15 Cryostat top plate

The top lid was made up of the top plate and insulation baffles. The top plate, shown in

Figure 3.15, contained four connectors for test instrumentation wiring, pressure relief

valve, nitrogen vapour vent check valve, lifting points and a port through which liquid

nitrogen was poured when the cryostat was closed. These four connectors were used for

the diode temperature sensors, normal voltage taps, detailed voltage taps, and

commercial temperature probe wire connections.

Cryogenic temperature was reached using a Gifford–McMahon AL230 single-stage

cryocooler, as shown in Figure 3.14 (right). Two pipes from the front panel of the

cryocooler carried compressed helium to the cold head and low pressure helium back to

the cryocooler. The cryostat was manufactured so that the copper bucket sat on top of

the cold head, cooling the copper bucket rapidly due to its high thermal conductivity.

The performance of the cryocooler is strongly influenced by the quality of the vacuum

in the vessel. The vacuum pump and gauge are shown in Figure 3.14 (middle). The

vacuum was created using a Leybold PT70 B compact turbo molecular pump with

forepump. The pump line was connected to a valve at the bottom of the cryostat, along

with a pressure gauge.


93

3.6 Conclusions

Material selection and geometry of the prototype former to support the MgB2 wire was

first introduced. The wind-and-react method was used to manufacture the MgB2 SFCL

coil. Copper clamps were manufactured to attach the MgB2 wire onto the copper braids

for the current lead connections. The MgB2 wire was then wound onto the alumina

former and heat treated in a vacuum oven. The instrumentation consisting of voltage,

current and temperature signals was also described. The high-current and low-current

experimental test circuits and the LabVIEW control system were also presented. Details

of the cryostat system including a vacuum insulated cryostat, a cryostat top lid, a

cryocooler compressor, a vacuum pump set, temperature monitors and a controller were

included.

Chapter 4: Experimental Investigation of SFCL Coils

94

4 Experimental Investigation of SFCL Coils

4.1 Introduction

A single-strand MgB2 SFCL coil with a wire diameter of 1.28mm was initially

investigated. The MgB2 wire sample was delivered on a spool from Hyper Tech. Two

sections of MgB2 wire were used, each over three metres long, for winding the

interleaved design coil on the ceramic former. To react the elemental magnesium and

boron to form MgB2 and anneal possible cracks in the MgB2 wire during the winding

process, the heat treatment process described in the previous chapter was carried out at

Bodycote Heat Treatment Ltd. in Stockport. The instrumentation probes were mounted

onto the coil afterwards. The main purpose for the experimental tests on the single-

strand coil was to determine whether MgB2 in simple round wire form would be suitable

as a resistive SFCL.

A three-strand MgB2 SFCL coil was then tested to determine the current capacity for

multi-strand wire. The full quench current level will be reduced if the quench current

level and impedance of each wire strand is not identical; therefore, ensuring each strand

carries the same current is one of the most important aspects. The diameter of each

individual wire was 0.63mm, and the three wires were hand braided with a pitch of

18.4mm. The smaller diameter wire was used because of the limitation in the slot size of

the former. Two braids were fabricated, hand wound onto the former, heat treated by

Hyper Tech and delivered to Manchester afterwards. The instrumentation probes again

were mounted onto the coil.

The specification of both wires is shown in Table 4.1. The chemical composition of the

two wires was the same; only the percentage of each composition is slightly different.


95

Table 4.1 MgB2 wire specification

MgB2 wire Strand B source Powder Barrier Sheath

1 1617 SMI-97/98 MgB2 Niobium Monel

2 1617-3-B SMI-97/98 MgB2 Niobium Monel

MgB2 wire OD (mm) Mg:B % Powder % Barrier % Sheath

1 1.28 1:2 30.7 17.1 52.2

2 0.63 1:2 30.6 14.8 54.4

4.2 Single-strand SFCL Coil

4.2.1 Instrumentation

Two series of voltage taps were mounted onto the single-strand coil: the normal voltage

taps were connected to each turn; and the detailed voltage taps were equally distributed

on the first turn. There were eight voltage channels on the whole coil, which is shown

schematically in Figure 4.1 (left). Each interleaved coil had three full turns and one part

turn. Voltage channels 1 to 3 and 6 to 8 were full turns, whilst voltage channels 4 and 5

were part turns. In order to monitor the voltage distribution starting from one end of the

coil, eight detailed voltage channels 11 to 18 were connected to the first turn, as shown

in Figure 4.1 (right). The notation A, B and C corresponds to the first two turns in

Figure 4.1 (left).

These two series of voltage channels shared the same voltage amplifier external to the

cryostat; they were used in separate tests for different purposes. The normal voltage

channels were first used to get an overview of the voltage distribution across the coil

and then the detailed voltage channels were used to get the voltage of each section in the

first turn.


96

Figure 4.1 Schematic showing locations of normal voltage taps (left), detailed voltage

taps and BAS16 diode temperature sensors (right)

The distribution of temperature in the cryostat was measured using four Cryocon S700

commercial temperature probes mounted around the former. Six BAS16 diode

temperature sensors were directly mounted onto the wire to measure the temperature of

the wire. The locations of the six temperature diodes, which are denoted as D1 to D6,

are shown in Figure 4.1 (right). Diodes 1 to 4 were mounted on the coil at the middle of

the detailed voltage channels 11, 13, 15 and 17, whilst diodes 5 and 6 were mounted on

the coil in the first and second half of the second turn. The output voltage of each diode

was connected to the LabVIEW data acquisition board.

All the diodes were calibrated before carrying out any tests to improve their accuracy;

the diode voltage could then be easily converted to the corresponding temperature by

interpolation of the calibration data. This recorded diode voltage would indicate the

temperature rise after quench and the recovery process of the coil. It is important to

ensure that the temperature of the coil returns to the normal operating temperature

before the next quench test can be conducted to prevent any possible damage or

degrading of the coil.


97

4.2.2 Calibration of BAS16 diode temperature sensors

Temperature is one of three important parameters for the MgB2 wire to remain in the

superconducting state. As mentioned in section 3.4.3, BAS16 silicon diode temperature

sensors were used to measure the coil temperature. The characteristics of each diode

however might be a little different. Individual calibration of each diode therefore would

significantly improve their accuracy.

After the cryostat was filled with liquid nitrogen, the temperature inside the cryostat was

measured at approximately 77K after several minutes. The cryocooler compressor was

then turned on to bring the temperature inside the cryostat down to 20K. One of the

Cryocon temperature sensors mounted on the fibreglass board was used to provide a

calibrated temperature reading to which the diodes were calibrated. The diode voltages

were recorded using the LabVIEW calibration programme per each change in

temperature, i.e. one Kelvin.

After the entire test was completed, the cryostat was then warmed up to room

temperature by turning off the cryocooler and using the heater inside the cryostat to

warm it up more quickly. The temperature controller was set higher than the measured

temperature in the cryostat, step by step, to ensure the full transition of nitrogen across

the two phase boundaries during warming up. The LabVIEW programme again

recorded the diode voltages per each change in temperature.

The voltage of each diode with respect to temperature when cooling down (1st) and

warming up (2nd) is shown in Figure 4.2. The results show that the voltages of the six

diodes have similar characteristics. It is clear that each diode voltage when warming up

and cooling down is similar at the same temperature. The measured diode voltages in

any subsequent tests therefore could be converted to the equivalent temperatures easily

using simple interpolation. The BAS16 diodes have a closely constant sensitivity of

2.7mV/K between 30K and 300K, but below 30K the sensitivity increases to 3.9mV/K.


98

0.4

0.6

0.8

1

1.2

1.4

1.6

20 90 160 230 300

Temperature (K)

Dio

de

Vo

ltag

e (V

)

Ch.1 1st

Ch.2 1st

Ch.3 1st

Ch.4 1st

Ch.5 1st

Ch.6 1st

Ch.1 2nd

Ch.2 2nd

Ch.3 2nd

Ch.4 2nd

Ch.5 2nd

Ch.6 2nd

Figure 4.2 Temperature diode calibration curves when cooling down (1st) and warming

up (2nd)

4.2.3 Temperature profile

The objective of the temperature profile test was to obtain the impedance of the coil at

different temperatures and to determine if the coil was superconducting below the

critical temperature.

The temperature of the cryostat was slowly controlled from 20K to 40K. The low-

current supply was used to pass two cycles of low level current through the coil at

different temperatures. A gain of 981 was used for the voltage amplifier when the coil

was superconducting, whilst a gain of 51 was used when the coil transitioned to the

normal resistive state. The impedance of each turn was determined using the root mean

square (RMS) voltage divided by the RMS current.

Figure 4.3 shows the impedance per metre of the MgB2 wire based on the average

impedance of each turn. It is obvious that the impedance can be divided into three

sections: below 36.2K, the wire is superconducting and the impedance (inductance) is

constant; between 36.2K and 37.3K, the coil is in the transition zone in which the

impedance increases quickly with temperature; and above 37.3K, the impedance

increases steadily, which is similar to a normal conductor with increasing temperature.

The resistance of the coil was 0.27Ω/m at room temperature (293K), reducing to


99

0.0327Ω/m at 37.3K. The critical temperature is defined as the temperature at which the

resistivity is half of the normal state resistivity. The critical temperature for the coil is

approximately 36.7K.

0

0.01

0.02

0.03

0.04

20 24 28 32 36 40

Temperature (K)

Imp

edan

ce p

er m

etre

(Ω

/m)

Figure 4.3 Temperature profile of the coil

4.2.4 Frequency sweep

A frequency sweep test was carried out to determine the impedance of the coil with

varying frequency in the superconducting state.

The low-current supply was used to pass two cycles of low level current through the

coil with the frequency varying from 10Hz to 100Hz at 25K. The current and voltage

across each turn was recorded by LabVIEW. Again the impedance of each turn was

determined using the RMS voltage divided by the RMS current. Figures 4.4 and 4.5

show the individual impedance of each turn and total impedance of the coil with varying

frequency. It can be seen that the impedance of each turn is similar and they have a

linear relationship to frequency. This indicates that the coil is dominated by its

inductance in the superconducting state.

The interleaved coil design was employed to minimise the inductance; however, there is

still a relatively small and finite coil inductance, which will exceed the internal


100

inductance of the MgB2 wire. A first order equation for the trend line of impedance

gives an estimated coil inductance of 6.26µH.

0

0.0002

0.0004

0.0006

0.0008

0.001

0 20 40 60 80 100Frequency (Hz)

Imp

edan

ce p

er m

etre

(Ω

/m) Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Figure 4.4 Impedance of the coil with varying frequency at 25K

y = 0.0000088x + 0.0000186

R2 = 0.9996766

0

0.0002

0.0004

0.0006

0.0008

0.001

0 20 40 60 80 100

Frequency (Hz)

Imp

edan

ce p

er m

etre

(Ω

/m)

Average

Linear(Average)

Figure 4.5 Total coil impedance with varying frequency at 25K

4.2.5 Quench tests

Quench tests were carried out for a number of reasons: firstly, to determine whether the

coil would limit a fault current; secondly, if it did, whether it would recover to the

superconducting state afterwards; thirdly, what effect would the operating temperature


101

have on the quench current level; and fourthly, whether the coil would have any

noticeable deterioration after successive quench tests.

The high-current test circuit was used to pass two cycles of high current and the

LabVIEW programme recorded all the instrumentation channels. Quench tests were

first carried out at a temperature of 34K, to reduce the quench current level and hence

reduce the heat dissipated in the coil during quench. The potential peak current level

was gradually increased by adjusting the set point of the variac and the load resistor.

The potential peak current was calculated based on the coil in the superconducting state,

which had negligible impedance. In practice, when the potential peak current was higher

than the quench current level, the calculated peak current could not be achieved because

the coil already quenched before reaching the peak current level.

The potential peak current was increased gradually until the coil started to quench.

Figures 4.6 to 4.8 show the coil response during a short cycle quench test at 34K with a

potential peak current of 311A. It is clear that the coil is quenching (but not

completely). The peak current is limited to 270A in a quarter-cycle. It is interesting that

the whole coil does not quench at the same current level. Figure 4.7 clearly shows that

channels 1, 4, 5, 7 and 8 quench earlier than the other channels.

Figure 4.8 shows the temperature response of the six sections of the coil using the

BAS16 diode temperature sensors. The temperature rise of the first section is about 5K,

whilst the other sections do not show any observable temperature rise. There will be

heat generated if any coil section is quenching, leading to a temperature rise in that coil

section. This demonstrates that the wire is partially quenching. The quench test only

lasted for 50 milliseconds; however, the coil took 50 seconds to recover to the operating

temperature of 34K. The temperature of the coil was checked to make sure the coil was

fully recovered before the next test was carried out, to prevent possible damage to the

coil.


102

-6

-4

-2

0

2

4

6

0 0.01 0.02 0.03 0.04 0.05

Time (s)

Vo

ltag

e (V

)

-300

-200

-100

0

100

200

300

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current

Figure 4.6 Coil response at 34K with a potential peak current of 311A

0

1

2

3

0.005 0.007 0.009 0.011 0.013 0.015

Time (s)

Vo

ltag

e (V

)

0

100

200

300

Cu

rren

t (A

)Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current

Figure 4.7 Coil response at 34K with a potential peak current of 311A, highlighting the

point of quench


103

32

34

36

38

40

0 10 20 30 40 50 60

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6

Figure 4.8 Coil temperature response at 34K with a potential peak current of 311A

The potential peak current was further increased to 372A. In Figure 4.10, channels 1, 4,

5, 7 and 8 quench earlier than the other channels, which is the same as in the previous

quench test. However, channel 3 also quenches in the second test. This indicates that as

the potential peak current increases, more sections of the coil are quenching.

It is clear that the voltage across the first turn rises as the potential peak current

increases. The Joule losses in the first turn have increased because of the high current

and voltage; this explains why the temperature rise of the first section is about 6.5K in

Figure 4.11, which is higher than the previous test. As a result, the recovery time for the

coil is also longer. The BAS16 diode temperature sensors are clearly very important in

indicating the temperatures of the wire sections.

From the previous tests, it can be seen that channels 1, 4, 5 and 8 quench earlier than the

other channels. The particular feature of these channels is that they are the end turns of

the coil. There are three reasons which may explain why the end turns quench first.

Firstly, the end of the coil was connected to the copper braid in the cryostat with the

copper braid connected to the copper bar and then to the external connections. The

temperature of the copper bar would therefore be slightly higher than the MgB2 coil.

Secondly, the copper braid incurs Joule losses when current is flowing through it, which

causes the temperature to rise. This would mean that the coil ends were slightly warmer


104

than the coil centre leading to a reduction in the critical current. Thirdly, the magnetic

field around the end of the coil may be different than at the middle. The magnetic field

can affect the quench of superconductors as discussed in section 2.1.1.2. These explain

why the coil begins to quench initially at the coil ends.

-9

-6

-3

0

3

6

9

0.01 0.02 0.03 0.04 0.05 0.06

Time (s)

Vo

ltag

e (V

)

-300

-200

-100

0

100

200

300

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current


0

1

2

3

4

5

0.02 0.022 0.024 0.026 0.028 0.03

Time (s)

Vo

ltag

e (V

)

0

100

200

300

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current


the point of quench


105

32

34

36

38

40

42

0 10 20 30 40 50 60

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6


Short cycle quench tests were repeated at 32K and 30K, to evaluate the variation in the

quench current with temperature. A sample of 32K and 30K test results are shown in

Figures 4.12 to 4.17. The voltage amplifier had a maximum output voltage of 10.5V and

‘clipped’ the signals that went above this level. As shown in Figure 4.12, the voltage

channel 5 is clipped by the amplifier. It is worth mentioning that in Figure 4.14 the

temperature rises in both the first and third sections of the first turn. This indicates more

sections of the first turn are quenching due to the higher current.

-12

-8

-4

0

4

8

12

0 0.01 0.02 0.03 0.04 0.05

Time (s)

Vo

ltag

e (V

)

-600

-400

-200

0

200

400

600

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current



106

0

2

4

6

8

10

0 0.002 0.004 0.006 0.008 0.01

Time (s)

Vo

ltag

e (V

)

0

100

200

300

400

500

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current


the point of quench

30

35

40

45

50

55

0 30 60 90 120

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6



107

-12

-8

-4

0

4

8

12

0 0.01 0.02 0.03 0.04 0.05

Time (s)

Vo

ltag

e (V

)

-700

-500

-300

-100

100

300

500

700

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current


0

2

4

6

8

10

12

0.005 0.007 0.009 0.011 0.013 0.015

Time (s)

Vo

ltag

e (V

)

0

100

200

300

400

500

600

700

Cu

rren

t (A

)Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Current


the point of quench

It is clear from all the tests conducted that the coil successfully recovered to the

superconducting state after quench. The repeated quench tests show that the quench

behaviour was consistent.


108

28

30

32

34

36

38

40

0 10 20 30 40 50 60

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6


From the tests, the quench currents at 34K, 32K and 30K are 245A, 433A and 639A

respectively. As shown in Figure 4.18, the quench current level increases linearly as the

temperature reduces. This demonstrates that a higher quench current level can be

achieved by reducing the operating temperature of the coil. However, the quench

current level along the wire tends to be less uniform when operating at a lower

temperature. Furthermore, reducing the operating temperature would increase the

equipment costs, so there is a compromise between the quench current level and the

normal operating temperature.

200

300

400

500

600

700

30 31 32 33 34

Temperature (K)

Qu

ench

Cu

rren

t (A

)

Figure 4.18 Estimated quench currents versus temperature


109

4.2.6 Long duration quench tests

After the short cycle quench tests were complete, long duration quench tests were

conducted to determine if the coil was capable of limiting the current over a longer time

period because the protection equipment in a power network may take up to several

hundred milliseconds to clear the fault.

The operation of the coil during a ten-cycle quench test with a potential peak current of

372A is illustrated in Figures 4.19 to 4.22. The coil begins to quench at a current of

approximately 276A, and then as the quench progresses rapidly throughout the coil, the

current is limited to 110Apeak after ten cycles. Figure 4.20 shows the voltage of each

turn during a quench test for ten cycles. During this test, all the channels quench except

channel 6.

It was observed that the coil started to quench in the end turns. It was assumed that the

coil started to quench from the end, propagating along the coil. As detailed in section

4.2.1, eight voltage taps were mounted equidistance along the first turn of the coil, but

they shared the same voltage amplifier channels with the normal voltage taps. After it

was confirmed that the coil could survive in a ten-cycle quench test, this quench test

was repeated and the detailed voltages were recorded. Figure 4.21 shows that detailed

voltage channels 11 and 12 start to quench at the first peak current and channels 13 and

14 start to quench in sequence afterwards. This confirms the hypothesis that the coil

starts to quench from the end and propagates along the coil. This also demonstrates that

the coil is partially quenched, and more and more sections are quenching as the

temperature rises. The coil is however still not fully quenched during this test.

The temperature rise of the coil is shown in Figure 4.22. Again, as detailed in section

4.2.1, the first diode temperature sensor was mounted onto the coil between detailed

voltage channel 11 and the second sensor was mounted onto the coil between detailed

voltage channel 13. Figure 4.22 shows that both channels are quenched; the temperature

in these sections therefore increases. It is clear from Figure 4.22 that the temperature

rise of the first and the third sections is 30K and 8K respectively. The voltage on

detailed voltage channel 11 is higher than channel 13; the power dissipated is therefore

higher in the first section. This explains why the temperature rise of the first diode


110

temperature sensor is much higher than the second. Figure 4.22 also shows that the coil

takes more than two minutes to recover.

-300

-200

-100

0

100

200

300

0 0.05 0.1 0.15 0.2 0.25

Time (s)

Cu

rren

t (A

)


current of 372A

-9

-6

-3

0

3

6

9

0 0.05 0.1 0.15 0.2 0.25

Time (s)

Vo

ltag

e (V

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8


current of 372A


111

-2

-1

0

1

2

0 0.05 0.1 0.15 0.2 0.25

Time (s)

Vo

ltag

e (V

)

Voltage Ch.11

Voltage Ch.12

Voltage Ch.13

Voltage Ch.14

Voltage Ch.15

Voltage Ch.16

Voltage Ch.17

Voltage Ch.18

Figure 4.21 Coil first turn detailed voltage response during a ten-cycle quench test with

a potential peak current of 372A

30

40

50

60

70

0 30 60 90 120

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6

Figure 4.22 Coil temperature response during a ten-cycle quench test with a potential

peak current of 372A

A one-second quench test with a potential peak current of 372A was conducted to

determine whether the coil can survive for a one-second fault. The operation of the coil

is shown in Figures 4.23 and 4.24. The current is limited to 300Apeak in the first quarter-

cycle then reduces to 63Apeak after one second. This test demonstrates the coil is able to


112

carry the fault current for one second without any damage or degradation. The highest

temperature recorded on the coil is 107.2K during the test, as shown in Figure 4.24.

-300

-200

-100

0

100

200

300

0 0.2 0.4 0.6 0.8 1

Time(s)

Cu

rren

t (A

)

Figure 4.23 Coil current response during a fifty-cycle quench test with a potential peak

current of 372A

30

50

70

90

110

0 30 60 90 120

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6

Figure 4.24 Coil temperature response during a fifty-cycle quench test with a potential



113

4.2.7 Simulated fault test

A simulated fault test was conducted to determine the coil behaviour when a real fault

occurs. The coil initially carried rated current in the superconducting state when the

current was suddenly increased to simulate a fault. The objective of this test was to

access the current-limiting performance of the coil in a more realistic condition.

Initially, the coil operated with a peak current of 40A. The fault was made by manually

closing a switch which was in parallel with the load resistor, to produce a short circuit

fault. The potential current was 316Apeak without any coil resistance. As shown in

Figure 4.25, in practice the current was limited to 252Apeak in a quarter-cycle, reducing

to 115Apeak over eight cycles. This test demonstrates that the coil can effectively limit a

fault current in a relatively short period of time.

-300

-200

-100

0

100

200

300

0 0.1 0.2 0.3 0.4 0.5

Time (s)

Cu

rren

t (A

)

Figure 4.25 Coil current response to simulated fault at 34K with a potential peak

current of 316A

4.2.8 Temperature rise test

An SFCL would have to operate at or below its rated current most of its life in a

practical application. The purpose of a temperature rise test was to determine whether

the AC losses in the superconducting coil would cause a slow increase in temperature,

leading to an unexpected quench. The normal steady state temperature of the coil was


114

recorded with the coil operating at 30K with a current of 200Apeak for more than one

hour.

Figure 4.26 illustrates the recorded temperature during the test. It is clear that at this

current level there is no obvious temperature rise after one hour of steady operation. A

steady state temperature rise of less than 0.1K was recorded. This demonstrates that an

SFCL manufactured with a similar design to the coil in this test would be able to operate

at rated conditions without experiencing an unwanted quench.

29

29.5

30

30.5

0 1000 2000 3000 4000

Time (s)

Tem

per

atu

re (

K) Temp Ch.1

Temp Ch.2

Temp Ch.3

Temp Ch.4

Temp Ch.5

Temp Ch.6

Figure 4.26 Coil temperature response during continuous 200Apeak current test at 30K

for one hour

4.2.9 Summary

The experimental tests showed that the coil successfully operated as an SFCL with

consistent characteristics. The tests proved that the quench current increased nearly

linearly as the temperature was reduced. Furthermore, the tests confirmed that the coil

was dominated by its inductance in the superconducting state and the resistance

increased quickly and became dominant after the coil started to quench.

The quench process started from the ends of the coil and extended throughout the coil as

the fault current increased. A more uniform quench would be produced when the coil

was operated with a fault current which significantly exceeded the quench current of the


115

coil. The long duration quench test showed that the coil was able to limit the fault

current for one second without any damage.

The coil heated up quickly after quench and took several minutes to recover to the

superconducting state. Some measures therefore need to be taken to protect the coil

from heating up, to reduce the recovery time. One option included placing a fast-acting

vacuum interrupter in series with the SFCL coil.

A steady state temperature rise of less than 0.1K was observed during normal

superconducting operation after the coil was tested for one hour, which demonstrated its

thermal stability in continuous operation.

4.3 Three-strand SFCL Coil

In a practical power system application, parallel MgB2 wires will likely be required in

order to carry current levels in the kA range. Ensuring each strand of wire carries the

same current is a very important issue for multi-strand wires because the critical current

level will decrease if the critical current and impedance of each wire strand is not

closely identical. This was achieved by braiding the strands to equalise the impedance

of each parallel wire path. The three strand wires were braided together, as shown in

Figure 4.27.

Figure 4.27 Picture of the three-strand coil (Courtesy of Hyper Tech)

4.3.1 Instrumentation

It was impossible to measure the current in each strand directly using conventional

current transducers because of the braiding. The voltage taps mounted onto a section of

each strand however would be a good indication of the current distribution. The

locations of the voltage taps are shown in Figure 4.28. Each coil on the former was

made up of three full turns and one part turn. The solid dots in Figure 4.28 indicate the


116

locations of the voltage taps. The wire strands themselves were un-insulated and the

sheaths were in random electrical contact along the length of the coil. Insulation paper

was used to separate and insulate each strand where individual voltage taps were

required on the first turn of the two interleaved coils. The voltage taps on each strand

therefore could be measured and used to monitor each strand current in the braid.

Figure 4.28 Schematic showing locations of the voltage taps

4.3.2 Temperature profile

The impedance of the coil from 20K to 40K using the same method as described in

section 4.2.3 is shown in Figure 4.29. The resistance at 37.3K and at room temperature

are 0.0408Ω/m and 0.38Ω/m for the three-strand coil, whilst they are 0.0327Ω/m and

0.27Ω/m for the single-strand coil.

It is clear that the resistance of the three-strand coil is higher than the single-strand coil.

The chemical composition of the two coils is similar; the resistance therefore is mainly

determined by the cross-sectional area. The cross-sectional area of the single-strand coil

is about 1.37 times the three-strand coil. From the resistance and known cross-section

area, the electrical resistivity of the single-strand coil is 3.47×10-7Ω·m and the three-

strand coil is 3.55×10-7Ω·m at room temperature.


117

0

0.01

0.02

0.03

0.04

0.05

20 24 28 32 36 40

Temperature (K)

Imp

edan

ce p

er m

etre

(Ω

/m)

Figure 4.29 Temperature profile of the coil

4.3.3 Frequency sweep

The impedance of the coil was obtained using the same method as described in section

4.2.4. The low-current supply was used to pass two cycles of low level current through

the coil with the frequency varying from 10Hz to 100Hz at 25K. The voltage channels 1

to 3 and 6 to 8 show similar voltages. The voltages on channels 1 and 6 were used

therefore to calculate the impedance for the first and fifth turns. Figures 4.30 and 4.31

show the individual impedance of each turn and the total impedance of the coil with

varying frequency. Again the results show that the coil is inductance dominated in the

superconducting state. The inductance of the coil is approximately 3.65µH from the

trend line in Figure 4.31. This reduction (when compared to the single-strand coil)

requires further investigation.

0

0.0001

0.0002

0.0003

0.0004

0.0005

0.0006

0 20 40 60 80 100Frequency (Hz)

Imp

edan

ce p

er m

etre

(Ω

/m)

Voltage Ch.1

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.9

Voltage Ch.10

Figure 4.30 Impedance of the coil with varying frequency at 25K


118

y = 0.0000052x + 0.0000110

R2 = 0.9990691

0

0.0001

0.0002

0.0003

0.0004

0.0005

0.0006

0 20 40 60 80 100

Frequency (Hz)

Imp

edan

ce p

er m

etre

(Ω

/m)

Average

Linear(Average)

Figure 4.31 Total coil impedance with varying frequency at 25K

The total voltage drop across each strand is related to its resistance and inductance,

expressed as Equation 4.1. VA, VB and VC are the total voltages across each strand. These

voltages are identical because the coil ends are connected to the copper braid. The

resistances RA, RB and RC consist of the strand resistance and the contact resistance

between each strand and the copper braid. IA, IB and IC are the currents flowing through

each strand. The self-inductance of each strand is denoted as LA, LB and LC. MAB, for

example, represents the mutual inductance between strand A and strand B.

/

/

/

A A A AA AB AC

B B BA B BC B B

CA CB CC C C C

R I di dt VL M M

R I M L M di dt V

M M LR I di dt V

+ =

(4.1)

In the superconducting state, the resistance of each strand can be influenced strongly by

the contact resistance and it is difficult to ensure that the contact resistance of each

strand is identical. It is important in these applications therefore to make sure that the

inductance of each coil strand is sufficient to dominate the individual strand impedance.

The coil inductance is clearly the major factor in helping to ensure each strand carries

the same current. The self-inductance of each strand is determined by the physical size

of the former and the position of each strand on the former. Each individual strand is

closely linked magnetically to the other strands because they have been braided


119

together. Their coefficients of coupling are close to unity therefore and the mutual

inductances will be closely related to the self-inductances. The three strands are braided

therefore to ensure the impedances of each individual strand are closely matched: this is

a common technique used in winding coils in large generators [128], for example.

4.3.4 Current sharing test

In practice, there have been quite a lot of problems in ensuring the uniform distribution

of the current in each strand [129, 130]. The objective of this experimental test was to

determine the current distribution between each of the three strands.

The high-current test circuit was used to pass five cycles of current through the coil at

an initial temperature of 30K. The voltage and current signals were recorded and are

shown in Figure 4.32. This figure show a 90º phase shift between the coil voltage and

current, which again demonstrates that the coil is inductance dominated. It also clearly

shows that the voltages in the first and fifth turns of each strand show no observable

difference, which indicates that the current is the same in each wire strand. In the

superconducting state, the coil inductance clearly plays an important role in current

sharing in a multi-strand wire. This result confirms that with careful coil design, it is

possible to scale-up the current levels using multi-strand wire.

-3

-2

-1

0

1

2

3

0.1 0.12 0.14 0.16 0.18 0.2

Time (s)

Vo

ltag

e (V

)

-90

-60

-30

0

30

60

90

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Voltage Ch.9

Voltage Ch.10

Current

Figure 4.32 Coil response with current flow at 30K


120

4.3.5 Quench tests

Two cycles of high current were passed through the coil to test the coil quench

behaviour. The voltage and current were recorded by LabVIEW.

Figure 4.33 shows the current and voltage waveforms of a quench test at 34K with a

potential peak current of 249A. It is clear from Figure 4.33 that all the turns have

quenched but only partially quenched. The current was limited to 178Apeak within a

quarter-cycle.

Figure 4.34 highlights the voltage of each strand on the first and fifth turns. It can be

seen that the voltages of channels 1 to 3 are higher than channels 6 to 8. This may be

explained by the temperature variation within the cryostat test chamber. Channels 6 to 8

were near the bottom of the copper bucket in the cryostat and close to the cold head.

The temperature therefore would be slightly lower at the bottom of the coil compared to

the connections at the top of the coils. As a result, the quench current would be lower on

the first turn compared to the fifth turn and this is where the quench would be expected

to commence. Figure 4.34 also shows that the voltage of each strand is still exactly the

same after quench, which again demonstrates that the quench current of each strand is

identical. This further confirms that it is possible to use multi-strand MgB2 wire to

scale-up the current levels for power system applications.

-6

-3

0

3

6

0 0.01 0.02 0.03 0.04

Time (s)

Vo

ltag

e (V

)

-200

-100

0

100

200

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Voltage Ch.9

Voltage Ch.10

Current



121

-6

-3

0

3

6

0 0.01 0.02 0.03 0.04

Time (s)

Vo

ltag

e (V

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8


potential peak current of 249A

Figure 4.35 shows the current and voltage waveforms of a quench test at 32K. Again it

is clear that all the turns are partially quenched. The voltages of channels 1 to 3 still

show no noticeable difference, as are channels 6 to 8. The voltages of channels 1 to 3

are still higher than channels 6 to 8, which is the same as the previous test.

-12

-6

0

6

12

0.01 0.02 0.03 0.04 0.05 0.06

Time (s)

Vo

ltag

e (V

)

-400

-200

0

200

400

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Voltage Ch.9

Voltage Ch.10

Current



122

-12

-6

0

6

12

0.01 0.02 0.03 0.04 0.05 0.06

Time (s)

Vo

ltag

e (V

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8



Figures 4.37 to 4.39 present the results of the quench test carried out at 30K with a

potential peak current of 529A. It can be seen in Figure 4.37 that channels 5 and 10 do

not quench whilst the other channels are partially quenched. The voltages of channels 1

to 3 are still the same, but the voltages of channels 6 to 8 show differences now.

Channels 6 to 8 still quench at the same time but the voltage of channel 6 is lower than

channels 7 and 8 in the first cycle and later the voltage of channel 8 is lower than

channel 7 as well. This is possibly an indication of some minor material variations in the

MgB2 wire. It is also worth pointing out that the voltage difference between channel 6

and channel 7 is smaller in the second cycle compared to the first. This shows that the

parallel strands will try to equalise the currents naturally.


123

-10

-5

0

5

10

0.01 0.02 0.03 0.04 0.05 0.06

Time (s)

Vo

ltag

e (V

)

-500

-250

0

250

500

Cu

rren

t (A

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Voltage Ch.9

Voltage Ch.10

Current


-10

-5

0

5

10

0.01 0.02 0.03 0.04 0.05 0.06

Time (s)

Vo

ltag

e (V

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8




124

-1.5

-1

-0.5

0

0.5

1

1.5

0.01 0.02 0.03 0.04 0.05 0.06

Time (s)

Vo

ltag

e (V

)Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Figure 4.39 Each coil strand voltage response of the fifth turn at 30K with a potential


The quench test was repeated and the coil demonstrated repeatable and reliable current-

limiting properties with no detectable degradation of the MgB2 wire performance during

the quench process.

The quench currents at different temperatures, summarised from repeated tests, are

shown in Figure 4.40. The quench current level has a nearly linear relationship to

temperature, which is the same as the single-strand coil. At the same temperature, the

quench current level for the three-strand coil is lower than the single-strand coil. This is

because the equivalent cross-sectional area of MgB2 is smaller than the single-strand

coil. The diameter of the MgB2 wire used in the single-strand coil and the three-strand

coil are 1.28mm and 0.63mm respectively. The portion of MgB2 is 30.7% in the single-

strand coil and 30.6% in the three-strand coil. The equivalent cross-sectional area of

MgB2 is therefore 0.395mm2 and 0.286mm2. It is easy to determine that the quench

current density for the single-strand coil and the three-strand coil are about 620A/mm2

and 531A/mm2 at 34K with a self-field of 50Hz. Hyper Tech suggested that the MgB2

wire, which they manufactured for MRI scanners, had a critical current density variation

between 15% and 20%. Therefore, the quench current density difference of the two

coils was acceptable.


125

0

100

200

300

400

30 31 32 33 34

Temperature (K)

Qu

ench

Cu

rren

t (A

)

Figure 4.40 Estimated quench currents versus temperature

4.3.6 Long duration quench test

Following the short cycle tests, the three-strand coil was subjected to a ten-cycle quench

test to check whether the MgB2 wire could survive a long duration fault typical of a

power system application.

The behaviour of the coil during a ten-cycle quench test at 34K with a potential peak

current of 249A is illustrated in Figures 4.41 to 4.43. The peak current is limited to

177A in a quarter-cycle and further reduces to 69A over ten cycles, as the coil resistance

increases with temperature. It can be seen from Figure 4.42 that all the turns have

quenched but not fully quenched. From Figure 4.43 it is also clear that the voltages

across the parallel strands are closely identical during the ten-cycle test. This test

demonstrates again that the three-strand SFCL coil shares the current equally and can

limit the fault current for 0.2s without any degradation.


126

-200

-100

0

100

200

0 0.05 0.1 0.15 0.2 0.25

Time (s)

Cu

rren

t (A

)


current of 249A

-8

-4

0

4

8

0 0.05 0.1 0.15 0.2 0.25

Time (s)

Vo

ltag

e (V

)

Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.4

Voltage Ch.5

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Voltage Ch.9

Voltage Ch.10


current of 249A


127

-8

-4

0

4

8

0 0.05 0.1 0.15 0.2

Time (s)

Vo

ltag

e (V

)Voltage Ch.1

Voltage Ch.2

Voltage Ch.3

Voltage Ch.6

Voltage Ch.7

Voltage Ch.8

Figure 4.43 Each coil strand voltage response of the first and fifth turns during a ten-

cycle quench test with a potential peak current of 249A

4.3.7 Summary

The initial investigation of a three-strand SFCL coil has shown that the coil successfully

performed repeatable and reliable operations as an SFCL. Each of the three wire strands

shared the current equally during the current sharing test and demonstrated closely

identical responses during quench current tests at 34K and 32K. However, the voltage

across each strand of the fifth turn was different during quench current test at 30K,

which suggested that the coil tend to be less uniform at lower operating temperature.

It is believed that an SFCL using MgB2 in wire form with multiple strands in parallel

shows considerable potential as a practical method for scaling-up the current levels

required for power system applications.


128

4.4 Conclusions

The experimental investigations conducted on the two coils showed reliable and

repeatable current-limiting properties in short cycle and long time duration quench tests.

Tests on both coils demonstrated that the quench current increased nearly linearly as the

temperature reduced. The quench current density of the single-strand coil was about

16.7% higher than the three-strand coil at 34K with a self-field of 50Hz but in general

this was considered acceptable.

The inductance of the two coils was relatively small but finite, determined by the

physical structure of the former. Tests proved that the coil was dominated by inductance

in the superconducting state; the resistance became dominant after the coil started to

quench.

A significant problem for the resistive SFCL coil is that there are high Joule losses in

the coil after it quenches, so the temperature of the coil rises quickly and takes a long

time to recover to the superconducting state. Solutions are needed therefore to protect

the coil from heating up, to reduce the recovery time. One option is placing a fast-acting

vacuum interrupter in series with the SFCL coil, which will be investigates in the

chapters 6 to 8.

Chapter 5: Modelling of SFCL Coil

129

5 Modelling of SFCL Coil

5.1 Introduction

There are two interdependent aspects of the MgB2 SFCL coil that need to be considered

carefully: the electrical characteristics and the thermal properties. Resistivity of a

superconductor is strongly dependent on temperature, whilst the temperature rise of the

coil is caused by resistance Joule losses. A MATLAB model will be developed to

simulate the behaviour of the Hyper Tech 1.28mm MgB2 wire. There are two main

purposes for this MATLAB model: firstly, the simulation results will be compared with

the experimental results to validate the theory of the SFCL operation; secondly, the

validated MATLAB model will be used to predict and provide guidance for future

SFCL design work.

To acquire more accurate thermal information of the SFCL coil, a finite element (FE)

thermal model will be built using Flux2D [131]. The thermal analysis will be performed

using a transient thermal two-dimensional (2-D) FE solution. Power dissipated in the

coil will be calculated from the experimental results obtained. The FE thermal model

will therefore be used to simulate the thermal response of the coil using the measured

power dissipation. The temperature from the FE thermal model will be first compared

with the experimental results. After validation, the FE thermal model will then be used

to predict the thermal response of the coil for a three-second fault.

The results from the MATLAB model using adiabatic boundary conditions will estimate

the highest temperature; whilst the FE model, assuming perfect thermal contact between

the coil and nitrogen, will estimate the lowest temperature. The practical temperature

rise of the coil will fall between these two boundaries if there are no hot spots in the

coil. This would provide useful design guidance for further development.


130

5.2 MATLAB model

The SFCL coil cannot be modelled as a stand-alone device. If an SFCL coil is

connected to a network where the quenched resistance of the coil is dominant, the

current level is limited effectively and then the resistance develops slowly. On the other

hand, if the quenched coil resistance is not dominant, the current-limiting effect is not as

significant and the resistance develops quickly in the coil. The behaviour of the coil

therefore is highly dependent on the network or system impedance and has to be

modelled as part of the network. It was decided that the high-current test circuit should

be used as the network for the MATLAB model of this SFCL coil.

There are three operating regions for a superconductor: the superconducting region, the

transition region and the normal resistive region. In the superconducting region, the

resistance is negligible and in the MATLAB model, the resistance therefore is taken as

zero. In the transition region, the current density rises above the lower critical current

density and the coil starts to develop resistance. The resistance increases due to the

increasing current density. The coil temperature increases as a result due to the Joule

losses. The current density then starts to reduce because of the increasing resistance.

Due to the dissipated losses however, the temperature of the coil either increases above

the critical temperature or the current increases above the reduced upper critical current

density. The coil then transitions to the normal resistive region. In the normal resistive

region, the resistance is mainly dominated by the temperature. The resistance of the coil

is similar to a normal conductor with increasing temperature.

5.2.1 Model description

As described in the previous chapter, the high-current test circuit consisted of a constant

voltage source and a variable load resistor. The load resistance, connecting lead

resistance, coil inductance and transformer impedance were taken into consideration in

the MATLAB model. The fault current was reduced after the coil quenched, which

reduced the Joule losses and temperature rise in the coil. The limited current level

depended on the resistance of the coil following quench compared with the load

resistance.


131

In this MATLAB model, the SFCL coil was divided into arbitrary equal sections along

one of the interleaved coils. The MATLAB model time solution used a small time step,

for example, 200µs for a short cycle simulation and 500µs for a long duration

simulation. The calculation was repeated at each time step until the simulation time was

complete. At each time step, the MATLAB model undertook the following procedures:

• Set up the initial conditions such as temperature and resistance of each coil

section.

• Calculate the voltage and current variation in the circuit.

• Calculate the heat generated and temperate rise of each coil section using the

heat equation.

• Calculate the resistance of each coil section with updated current and

temperature.

The details of the calculations at each step are explained in the following section. The

first step is setting up the initial conditions. At the beginning of the simulation, an

operating temperature of 34K, for example, was specified as the initial temperature of

the central coil sections with a linear increase in temperature specified, so that the coil

ends were 0.1K hotter. The initial resistance of all the sections were set to be zero. From

the second time step, the initial conditions were obtained from the simulation result of

the previous time step.

The second step is calculation of the voltage and current variation. The high-current test

circuit consisted of a constant voltage source connected across a load resistor in series

with the SFCL coil through a transformer. The expression for the current variation is

given by:

( )s s c fcls

c

V I R Rdi

dt L

− += (5.1)

where: sdi

dt = Rate of change of current in the circuit (A/s)

Vs = Voltage of supply (V)


132

Is = Circuit current (A)

Rc = Circuit resistance, including load, transformer and lead resistance (Ω)

Rfcl = SFCL coil resistance (Ω)

Lc = Circuit inductance, including transformer and SFCL inductance (H)

The third step is calculation of the heat generated and temperature rise of each coil

section. The coil was taken as a solid conductor with adiabatic boundary conditions,

with heat only conducted along the conductor to simulate heat conduction along the coil

[118]. The heat was generated from the resistance Joule losses. The temperature rise

was calculated from the standard one-dimensional (1-D) heat conduction equation

[132]:

2

2

( , ) 1 ( , )d pc

CT x t T x tg

x k k t

ρ∂ ∂+ =∂ ∂

(5.2)

where: T = Temperature (K)

x = Distance along the coil length (m)

t = Time (s)

k = Thermal conductivity (W/m·K)

gc = Power generated in the coil per unit volume (W/m3)

ρd = Density (kg/m3)

Cp = Specific heat capacity (J/kg·K)

where:

2fcl s

c

R Ig

v= (5.3)

where: v = Volume of the SFCL Coil (m3)

It should be noted that the values of the specific heat capacity and thermal conductivity

vary with temperature. It was difficult however to obtain reliable material data at

cryogenic temperatures. The specific heat capacity and thermal conductivity of the

materials used from 20K to 500K therefore were sourced from [133-136], detailed in


133

Appendix A. The data was superimposed into single average curves and saved into

tables with respect to temperatures. In the MATLAB model, they were determined

using a lookup table with interpolation between the two nearest temperature points

[118].

The heat conduction equation shows that the rate of thermal energy generation equals

the rate of increase of internal energy, minus the net rate of heat gain by conduction

[132]. A partial differential equation solver was used to solve the above equation. The

MATLAB model however took about ten minutes to simulate one cycle (20ms), which

was too slow to integrate into a power system analysis. A performance test on the

MATLAB model showed that the majority of the processing time was taken by solving

the partial differential equation. Some modifications were therefore necessary to

optimise the MATLAB model. If the heat transfer along the coil was neglected, the heat

equation could be rearranged as:

c

d P

gdT

dt Cρ= (5.4)

If c

d P

g

Cρ is taken as constant during each time step, the equation therefore can be

expressed as:

c

d P

g tT

Cρ∆∆ = (5.5)

The MATLAB model simulation time was significantly reduced to several seconds for

one cycle using this approximation. The results from both MATLAB models will be

compared in the next section.

The fourth step is calculation of resistance. The resistance of the SFCL coil is highly

non-linear. In the MATLAB model it is divided into the following four regions: (1) If

the temperature is greater than 37.3K, the coil is in the normal resistive region and the

resistance of the coil is similar to a normal conductor with increasing temperature. The


134

R-T relationship determined from the test results is used to calculate the resistance. (2)

If the temperature is between 36.2K and 37.3K, the coil is developing resistance as a

function of temperature. The resistance was determined therefore from the temperature

profile. (3) If the temperature is lower than 36.2K and the current density exceeds the

critical current density, the coil is developing resistance due to current. An S-curve

approximation was used for the current-resistance profile, as detailed by several papers

[137-139]. (4) If it is none of the above, this means that the coil is in the

superconducting region and the resistance is therefore zero.

5.2.2 Comparison with short cycle quench tests

The MATLAB models including and neglecting heat conduction along the coil were

used to simulate quench behaviour at 34K, 32K and 30K with the same potential peak

current as the experimental tests. Current waveforms from the MATLAB models

compared to the experimental tests are shown in Figures 5.1 to 5.3.

In these graphs, the trace in blue is the experimental test result on the coil; the trace in

red is the MATLAB modelled result including heat conduction along the coil, and the

trace in pink is the MATLAB modelled result without considering heat conduction. It

should be noted that the red and pink traces overlap, so the red trace cannot be seen. The

results demonstrate that the MATLAB model with heat conduction along the coil gives

practically the same results as the one without heat conduction. Heat conduction along

the wire therefore does not affect the SFCL coil behaviour for a short cycle simulation.

The simulation time is significantly reduced by neglecting heat conduction along the

coil.

These results show a good correlation between the MATLAB model and the

experimental results for the first cycle. This confirms that the equations used to model

the quench process of the SFCL coil are reasonable.

In the MATLAB model, the resistance was not allowed to reduce in value during this

simulation, to avoid large oscillations. It is therefore suitable for fault current levels that

are much higher than the quench current level.


135

-200

-100

0

100

200

300

0 0.005 0.01 0.015 0.02

Time (s)

Cu

rren

t (A

)

Experimental

Modelled withheat transferalong the wire

Modelledwithout heattransfer


-400

-200

0

200

400

600

0 0.005 0.01 0.015 0.02

Time (s)

Cu

rren

t (A

)

Experimental





136

-400

-200

0

200

400

600

800

0 0.005 0.01 0.015 0.02

Time (s)

Cu

rren

t (A

)

Experimental




5.2.3 Comparison with long duration quench tests

Figures 5.4 and 5.5 show the comparison between the experimental and modelled

results for ten-cycle and fifty-cycle quench tests. It can be seen that the MATLAB

model shows a good correlation with the experimental results for the first cycle.

However, for subsequent cycles the MATLAB model shows a slightly lower current

than the experimental test. This would indicate that the modelled results give a higher

resistance. A possible reason for this is that the MATLAB model assumes adiabatic

boundary conditions. In practice the coil was wound on the alumina former and

surrounded by the solid nitrogen. There would be heat transfer therefore to the alumina

former and/or the solid nitrogen.

The MATLAB model was then revised to include the heat conduction into the nitrogen.

The coil therefore was modelled as a long straight conductor surrounded by the

nitrogen. The temperature of the coil was used as the boundary temperature between the

coil and the nitrogen. The temperature of the solid nitrogen at a certain distance, i.e. two

and half times the radius from the centre of the coil, for example, was taken as the

ambient operating temperature. This approximation will be checked using an FE

thermal model described in section 5.3.2. The solid nitrogen from the coil boundary to

the distance specified for ambient was divided into twenty annular segments. The heat

conduction equation was used to calculate the temperature variation of each segment by


137

setting ‘gc’ to zero, i.e. there is no heat generated in the nitrogen. The temperature rise

of the coil for each step was then calculated as the energy generated by the resistance

minus the heat dissipated into the nitrogen. The heat dissipated into the nitrogen was

calculated as the sum of the increase in the nitrogen internal energy.

-300

-200

-100

0

100

200

300

0 0.05 0.1 0.15 0.2

Time (s)

Cu

rren

t (A

)

Experimental

Modelled


of 372A

-300

-200

-100

0

100

200

300

0 0.2 0.4 0.6 0.8 1

Time (s)

Cu

rren

t (A

)

Experimental

Modelled


current of 372A


138

Figures 5.6 and 5.7 show comparisons of the experimental and modelled results

considering the heat dissipated into the nitrogen for ten-cycle and fifty-cycle quench

tests. It is clear that the model now shows an excellent correlation with the experimental

results during the period measured. However, the simulation time using this MATLAB

model was significant: 24 hours for a fifty-cycle quench.

-300

-200

-100

0

100

200

300

0 0.05 0.1 0.15 0.2

Time (s)

Cu

rren

t (A

) Experimental

Modelled withheat transferto nitrogen


of 372A (considering the heat dissipated into the nitrogen)

-300

-200

-100

0

100

200

300

0 0.2 0.4 0.6 0.8 1

Time (s)

Cu

rren

t (A

) Experimental

Modelled withheat transferto nitrogen


current of 372A (considering the heat dissipated into the nitrogen)


139

The limitation in the MATLAB model is that it can only calculate heat transfer in one

direction with a constant thermal conductivity in each section at each time step. The

MATLAB model therefore only took the nitrogen into consideration. However, the

alumina former has a higher thermal conductivity; it can conduct the heat quicker than

the nitrogen but this MATLAB model did not take it into consideration.

The temperature response from both MATLAB models is shown in Figure 5.8. The

temperature after a one-second quench test from the adiabatic model is 170K, whilst for

the model taking the nitrogen into consideration, it is 135.9K. The measured

temperature of the coil after the quench test is 107.2K. The temperature from the

MATLAB model is clearly higher than the test result. This is caused possibly because

of the thermal resistance between the surface of the coil and the temperature sensor.

0

40

80

120

160

200

0 0.2 0.4 0.6 0.8 1

Time (s)

Tem

per

atu

re (

K)

Modelledadiabatic

Modelledwith heattransfer tonitrogen

Figure 5.8 Temperature comparison for a fifty-cycle fault at 34K with a potential peak

current of 372A with adiabatic and nitrogen boundary conditions

5.2.4 Summary

A MATLAB model of the SFCL coil temperature response has been developed. This

MATLAB model successfully predicted the SFCL response over a range of

temperatures. The MATLAB model can be used to predict the SFCL behaviour in a

power system analysis and also to aid future designs. It is useful to indicate the


140

maximum temperature of the SFCL coil with adiabatic boundary conditions because

this is the worst case situation.

The results from the MATLAB model considering heat conduction along the coil were

almost identical to the model ignoring heat conduction for a short cycle faults. Heat

conduction along the coil therefore was neglected to reduce the simulation time.

5.3 Finite element thermal model

After the MATLAB model was validated, a transient thermal FE model was used to

perform a thermal analysis of the coil. Flux2D provides a 2-D numerical approximation

to the following equation [140]:

1d P

T T Trk k g C

r r r z z tρ∂ ∂ ∂ ∂ ∂ + + = ∂ ∂ ∂ ∂ ∂

(5.6)

where: g = Power loss density (W/m3)

r = Distance from the centre along the radial direction (m)

z = Distance from the centre along axis direction (m)


To build the FE thermal model, it is necessary to define the geometry and the materials

used. A copper spacer was placed in the centre of the cryostat working space. The

middle of the copper spacer was filled with a polystyrene space-filler to save volume

and reduce the required amount of liquid nitrogen. The alumina former with the MgB2

coil wrapped around it was placed in the cryostat annular working space and the coil

was connected to the copper connections using copper braid. Nitrogen took up all the

remaining working space. The geometry was entered as an axisymmetric model into

Flux2D. The FE model geometry and mesh are shown in Figures 5.8 and 5.9. The outer

three surfaces were defined as imposed temperature boundaries, to keep the external

surface of the cylinder at a constant operating temperature of 34K.


141

Copper

braid

MgB2

coil

Alumina

former

Polystyrene

space-filler

Nitrogen

Copper

spacer

Figure 5.9 Flux2D FE model of the geometry

Figure 5.10 Mesh of the Flux2D FE model, showing detail of the coil in the slot

The specific heat capacity and thermal conductivity of the materials detailed in

Appendix A also were used in the FE model. The thermal contact resistances were not

taken into consideration in the FE model because of the lack of reliable data. During the


142

winding process, gaps between the coil and the former were common. However, the size

of these gaps was difficult to estimate precisely. The gap therefore would be taken as a

variable as part of a sensitivity analysis. The distance of the coil from the former varied

from 0mm to 1.5mm.

The amount of thermal energy dissipated in the coil and copper braid connections were

calculated from the experimental results obtained from the SFCL coil. The FE thermal

model therefore was built to simulate the thermal response of the coil using the

measured power loss. Heat generated in the coil and copper braid was calculated as

follows. The coil was dominated by resistance after quenching. The instantaneous

power loss density in the coil was then determined using the expression:

cm cmcm

V Ig

v=

(5.7)

where: gcm = Power loss density in the coil (W/m3)

Vcm = Measured voltage across the coil (V)

Icm = Measured current passing through the coil (A)

The instantaneous power loss density in the coil during a one-second fault is shown in

Figure 5.11. It is clear that the power loss density in the coil is effectively reduced due

to the limited fault current. This is one of the reasons why the coil was able to survive

during a long duration fault.


143

0

1

2

3

4

5

6

7

0 0.2 0.4 0.6 0.8 1

Time (s)

Po

wer

den

sity

(G

W/m

3 )

Figure 5.11 Instantaneous power loss density in the coil during a one-second fault at

34K with a potential peak current of 372A

The cross-sectional area of copper in the copper braid is 16.08mm2, so the resistance of

the copper braid was calculated from:

cbcb

cb

lR

A

ρ= (5.8)

where: Rcb = Resistance of the copper braid (Ω)

ρ = Resistivity of copper (Ω·m)

lcb = Length of the copper braid (m)

Acb = Cross-sectional area of copper in the copper braid (m2)

The circuit was equivalent to two parallel copper braids; each one carried half the

current of the coil. The instantaneous power loss density in the copper braid connections

was determined using the following expression:

2

2cm cb

cbcb

I Rg

v =

(5.9)

where: gcb = Power loss density in the copper braid (W/m3)


144

vcb = Volume of the copper braid (m3)

The power loss density in the coil and copper braid was specified using the time varying

values by writing the data into two separate files [141]. In the FE model, the power

density for the coil and copper braid automatically linked to the two files to calculate

the temperature response. For any simulation time between two time points, the power

loss density was calculated using linear interpolation between the two nearest points.

5.3.2 Results and discussion

The temperature response with different gaps using the measured power loss density is

shown in Figure 5.12. The temperature drops slightly near one second because the test

only lasted for 49 cycles (0.98s) and there is no heat generated after this. The FE model

used the space average power dissipated in the coil. It is therefore difficult to say what

the peak temperature in the coil was during testing. However, it is predicted that the

average temperature during a one-second quench test would be in the range from 69.2K

to 122.7K. The temperature rise of the coil wound on the former without a gap is

dramatically lower than the other sizes of gaps. This confirms that the coil would

benefit by minimising the thermal contact resistance between the coil and the former.

The variation in temperature rise in other cases, except perfect contact, is small. The

distance of the coil from the former was therefore set at 1.5mm for the remaining

simulations.

The temperature profile vertically down the former through the centre of the coil with a

gap of 1.5mm from the former is shown in Figure 5.13. The temperature rise focuses in

the coil and drops dramatically around the coil. The temperature profile showing the

detail of the coil in the slot is shown in Figure 5.14. The temperature at the outer

boundary, which is two and half times the radius from the coil centre, is about the same

as the operating temperature of 34K. This confirms that in section 5.2.3, the selection of

this distance as the ambient operating temperature in the MATLAB model is reasonable.

It can be seen from Figure 5.14 that the temperature of the surrounding nitrogen drops

dramatically the further from the coil. In the MATLAB model, the nitrogen was divided

into 20 annular sections between the coil boundary and the nitrogen at 34K. This

sections number was chosen as a compromise between computational time and accuracy.


145

0

20

40

60

80

100

120

140

0 0.2 0.4 0.6 0.8 1

Time (s)

Tem

per

atu

re (

K)

0.0mm

0.1mm

0.2mm

0.3mm

0.5mm

0.8mm

1.0mm

1.5mm

Figure 5.12 Time variation of temperature in the centre of the coil for varying distances

from the former

0

20

40

60

80

100

120

140

0 20 40 60 80 100 120

Path length (mm)

Tem

per

atu

re (

K)

Figure 5.13 Temperature profile vertically down the former through the centre of the

coil after a one-second fault, with a gap of 1.5mm from the former


146

Figure 5.14 Temperature profile after a one-second fault, with a gap of 1.5mm from the

former

The temperature response from both numerical models with adiabatic boundary

conditions and including the nitrogen is shown in Figure 5.15. The FE model shows a

good correlation with the temperature response predicted by the MATLAB model. The

temperature rise of the numerical models including the nitrogen is about 40K lower than

the simple adiabatic boundary conditions. The temperature rise of the MATLAB model

is slightly higher than the FE model. Under adiabatic boundary conditions in the FE

model, it did not solve correctly if the thermal conductivity of the nitrogen was set to

zero; a low value of 10-3W/m·K therefore was specified. This explains why the FE

model had a slightly lower temperature rise under adiabatic boundary conditions.

The temperature variation following a quench for 0.04, 0.2 and 1 seconds from the

experimental test results, and the two numerical models are summarised in Table 5.1.

The temperature from the test results is consistently lower than those predicted by the

numerical models. The thermal resistance between the surface of the coil and the

temperature sensor may cause lower temperature measurements., Additionally, it is


147

possible that the coil might be close to the former, which would produce a lower

temperature rise. Furthermore, both numerical models do not consider phase change of

nitrogen, which absorbs energy without temperature rise.

0

40

80

120

160

200

0 0.2 0.4 0.6 0.8 1

Time (s)

Tem

per

atu

re (

K)

MATLABadiabatic

FE adiabatic

MATLAB withheat transferto nitrogenFE with heattransfer tonitrogenExperimental

Figure 5.15 Temperature response comparison for a one-second fault at 34K with a

potential peak current of 372A from the MATLAB model and the FE model

Table 5.1 Temperature response comparison with fault time variation

Quench

time (s)

Experimental

temperature

(K)

MATLAB

temperature

(adiabatic)

(K)

MATLAB

temperature

(nitrogen)

(K)

FE

temperature

(adiabatic)

(K)

FE

temperature

(nitrogen)

(K)

0.04 41.2 51 48.2 50.4 47

0.2 56.3 89.3 75.6 86.7 71.7

1 107.2 178 135.9 174.6 132

Figure 5.16 shows the temperature profile through the coil centre after a one-second

fault from both numerical models. The temperature in the coil and surrounding nitrogen

is very similar in both models. It shows that the temperature rise is centralised in the

coil, probably due to the relatively poor thermal conductivity of the surrounding

nitrogen. This again indicates the importance of measuring the coil temperature as close

to the coil as possible in order to obtain accurate measurements.


148

0

30

60

90

120

150

0 0.8 1.6 2.4 3.2

Length (mm)

Tem

per

atu

re (

K)

MATLAB

FE

Figure 5.16 Temperature profile through the centre of the coil for a one-second fault at

34K with a potential peak current of 372A from the MATLAB model and the FE model

5.3.3 Summary

The FE model showed a reasonable correlation with the experimental test results,

demonstrating that the data for the thermal properties of the materials specified in the

FE model were reasonable. The FE model predicted an average temperature in the range

from 69.2K to 122.7K in the SFCL for a fifty-cycle fault with a gap between the coil

and the former between 0mm to 1.5mm.

Comparing the results from the FE model and the MATLAB model with the

experimental test results, the temperature rise for different time variation demonstrated a

reasonable correlation. The two numerical models are useful therefore for predicting the

quench behaviour and the temperature rise of the coils with a similar design.

5.4 Prediction for three-second fault test

After the MATLAB and the FE models were validated using the experimental test

results, they were used to estimate the quench behaviour for three seconds. Figure 5.17

shows the predicted current response from the MATLAB model. The current level is

limited from 300Apeak to 40Apeak in a three-second fault condition. Figure 5.18 shows

the predict temperature response from both numerical models. The predicted


149

temperature from the MATLAB model with adiabatic boundary conditions is 290.3K,

which would be the highest temperature if the coil quenches uniformly. Using the power

loss density from the MATLAB model, the predict temperature from the FE model with

the coil surrounded by nitrogen is 212K. The predicted highest temperature is much

lower than the melting point of the SFCL coil; this coil therefore can withstand a fault

for three seconds.

-300

-200

-100

0

100

200

300

0 0.5 1 1.5 2 2.5 3

Time (s)

Cu

rren

t (A

)

Figure 5.17 Modelled current response for a three-second fault at 34K with a potential


0

50

100

150

200

250

300

0 0.5 1 1.5 2 2.5 3

Time (s)

Tem

per

atu

re (

K)

Matlabadiabatic

FE with heattransfer tonitrogen

Figure 5.18 Temperature response comparison for a three-second fault at 34 K with a

potential peak current of 327A from the MATLAB model and the FE model


150

5.5 Conclusions

The MATLAB model successfully predicted the SFCL coil response over a range of

temperatures and fault current levels. This validates the theoretical model developed for

the quenching process of the superconducting coil.

However, there are several limitations in the MATLAB model: firstly, it is not as

accurate for faults in which the superconducting material remains in or near the

transition region; secondly, the MATLAB model including heat transfer to the

surrounding nitrogen requires a long simulation time; and thirdly, the MATLAB model

can only consider heat transfer in one direction, for example, along the coil or into the

nitrogen.

The FE thermal model also showed a reasonable correlation with the experimental test

results. The FE thermal model considered the physical structure and materials

surrounding the SFCL coil during test. It provided a detailed temperature profile around

the coil. It was concluded that the coil winding should be manufactured carefully to

minimise the gap between the coil and former, in order to minimise the coil temperature

rise.

The FE thermal model was 2-D axisymmetric using space average power loss density

values in the coil. It was not possible therefore to investigate other thermal issues such

as localised hot spots.

After the two numerical models were validated with experimental test results, the

MATLAB model was initially used to predict the SFCL behaviour for a three-second

fault and the FE thermal model was then used to predict the temperature profile using

the power loss density from the MATLAB model. These two numerical models would

provide reliable predictions for future designs.

Chapter 6: Operating Actuator for Vacuum Interrupter

151

6 Operating Actuator for Vacuum Interrupter

6.1 Introduction

The advantages of integrating a vacuum interrupter into the resistive SFCL system were

discussed in the previous chapters. It is important to design and build a fast-acting

actuation mechanism to operate the DVS10CB vacuum interrupter, as shown in Figure

6.1: the specification is outlined in Table 6.1.

Figure 6.1 DVS10CB vacuum interrupter

Table 6.1 DVS10CB vacuum interrupter specification

Item Rated value

Maximum operating voltage 1.5kV

Rated AC current 320A

Maximum interrupting current 3.2kA

Contact holding force for 3.2kA 2.0kgf

Length 64mm

Diameter 50mm

Contact stroke 2.0mm

Allowed contact wear 1.0mm


152

The operating actuator for the vacuum interrupter has features that are clearly different

from that required for other types of circuit breakers such as air, oil and SF6. A vacuum

interrupter requires a comparatively short contact stroke and is required to compress the

butt contacts together with a considerable force, to prevent the contacts from separating

[94]. The requirements of the operating actuator for this application are summarised

below:

• The opening time should be less than 10ms. Once a fault occurs, the contacts

should be opened before the next current zero-crossing to prevent the SFCL coil

from heating up.

• The operating actuator should provide adequate holding force in the closed

position to reduce the contact resistance.

• The operating actuator should provide latching in the open and closed positions.

• The total movement of the actuator should be 5mm, which is 3mm further than

the stroke of the vacuum interrupter. This could provide extra holding force and

eliminate contact wear loss.

• It should be strong enough to withstand the mechanical stress during opening

and closing.

• The overall design should be of reasonable size and weight. A simple structure

and low cost will be a preference.

After comparing the advantages and disadvantages of different actuator types in section

2.4, the voice-coil type actuator was chosen because it has a simple structure and fast

response. An analytical model was developed first to calculate the approximate

magnetic field levels in the airgap. The current through the actuator coil and the force

produced on the coil were then taken into consideration. After the initial calculation

from the analytical model, a 3-D FE model was built using Vector Fields Opera

software [142]. This model was used to simulate the magnetic field distribution in the

airgap from the permanent magnets. Current flowing through the actuator coil in both

directions was then used to produce a vertical electromagnetic force on the actuator coil.

Magnetic latches were employed to hold the actuator in the open or closed position. The

latching force was also calculated using the FE model.


153

The complete operating actuator was designed to avoid contact popping, bounce,

rebound and welding. After the size of all the components was finalised, 2-D and 3-D

designs of the operating actuator were drawn using VariCAD. A prototype operating

actuator was then built by the mechanical workshop.

6.2 Analytical model

A preliminary analytical model was built to determine the magnetic field in the airgap

and the electromagnetic force using a classical magnetic circuit model, which was

simplified by ignoring steel reluctance, leakage flux and fringing effects.


The voice-coil type actuator is based on the simple principle of the loudspeaker. A light,

hollow coil is suspended in a strong magnetic field to allow free movement along the

axial direction [119, 120]. As soon as current flows in the actuator coil, an

electromagnetic force is produced on the coil.

The actuator was made up of steel blocks, permanent magnets and a copper-wound coil.

It was impossible to obtain an annular permanent magnet of suitable size and with a

radially magnetised field. Four pieces of permanent magnet with an arc surface facing

the centre therefore were used to keep the airgap small and equal. Four steel walls, a

bottom steel block and a steel cylinder inside the actuator coil provided a low reluctance

return path for the magnetic flux produced by the magnets. Four magnets were fixed to

the outer steel walls and the actuator coil placed in the airgap between the magnets and

inner steel cylinder. The geometry of the actuator is shown in Figure 6.2. It should be

pointed out that the two steel walls in the front of the figure are hidden to show the

detail at the middle of the actuator. The blocks in grey and black represent the steel and

permanent magnets respectively, whilst the ring in magenta represents the actuator coil.

N48 neodymium-iron-boron (Nd-Fe-B) rare-earth permanent magnets were sourced

from Arnold Magnetic Technologies Corporation. The remanence and coercivity are

1.4T and 975kA/m respectively [143]. The height of the actuator coil was designed to

be 5mm shorter than the permanent magnets, so that full utilisation of the magnetic field

could be achieved. The height of the permanent magnets was 80mm and the height of


154

the actuator coil was 75mm. The diameter of the copper wire including polyurethane

insulation was 0.95mm; the actuator coil was made up of 79 turns.

Figure 6.2 Geometry of the actuator (the two steel walls in the front are not shown)

Figure 6.3 Simplified model of the magnetic circuit


155

Figure 6.3 shows a simplified actuator model in the cross section through the centre of

one of the magnets. The approximate magnetic flux density in the airgap could be

calculated using Ampere’s Law along the red line, which denotes the mean flux path of

the magnetic circuit.

6.2.2 Magnetic field in the airgap

The magnetic field in the airgap is produced by the magnets when there is no current

flowing through the actuator coil. The permanent magnets operate in the second

quadrant of the hysteresis loop, which is called the demagnetisation curve.

The magnetic characteristic of one of the permanent magnets can be evaluated by

adding the actuator open-circuit load line to the magnet 2nd quadrant B-H characteristics

(drawn from the origin) with a gradient equal to the negative permeance coefficient

[144-146], as shown in Figure 6.4. The intersection of these two curves, point P0, is

called the operating point. Note, for this design, the permanent magnet 2nd quadrant B-H

characteristic is non-linear at 20oC. This material was chosen because of the high

remanence at this temperature.

Figure 6.4 Actuator permanent magnet operating points on the 2nd quadrant B-H

characteristic (demagnetisation curve) at 20oC


156

The operating point of the permanent magnet in the model shown in Figure 6.3 can be

further simplified using the following assumptions: the permeability of steel is assumed

to be infinite and hence the reluctance of the steel can be neglected; and there are no

flux leakage and fringing effects in the magnetic circuit.

Because there is no current flowing in the actuator coil, according to Ampere’s Law:

0m m g gH l H l+ = (6.1)

where: Hm = Permanent magnet magnetic field strength (A/m)

lm = Length of the permanent magnet (m)

Hg = Airgap magnetic field strength (A/m)

lg = Length of the airgap (m)

Based on the assumptions that there is no flux leakage and fringing effects, all the flux

from the permanent magnets goes through the airgap, therefore:

g g m mB A B Aφ = = (6.2)

where: Bg = Magnetic flux density in the airgap (T)

Ag = Area of the airgap (m2)

Bm = Permanent magnet magnetic flux density (T)

Am = Permanent magnet arc surface area (m2)

The permeability of free space is µ0, so that:

0g gB Hµ= (6.3)

Substituting Equations 6.2 and 6.3 into Equation 6.1 yields:

0

0g mm m m

g

l AH l B

Aµ

+ =

(6.4)


157

Equation 6.4 can be rewritten as:

0gm

m m p mg m

AlB H K H

l Aµ

= − = −

(6.5)

where: Kp = 0gm

g m

Al

l Aµ

, the permeance coefficient

The expression for the second quadrant linear section is given by [144, 145]:

0m m m rB H Bµ µ= + (6.6)

where: µm = Relative permeability of the permanent magnet

Br = Remanence in the permanent magnet (T)

Combining Equations 6.5 and 6.6 yields:

0m

m m rp

BB B

Kµ µ

= − +

(6.7)

Equation 6.7 can be rearranged as:

01 1

r rm

m g mm

pm g

B BB

l AK l A

µ µµ

= = + +

(6.8)

The arc surface area of the magnet is assumed to be the same as the airgap. Equation 6.8

can be simplified therefore as:

1

rm

gm

m

BB

l

lµ

=

+

(6.9)


158

From Equation 6.9, it can be seen that it is necessary to keep /g ml l as low as possible to

achieve a large magnetic flux density in the airgap. The actuator coil was placed in the

airgap requiring a tube to support it. The following factors were considered for the

length of the airgap: the distance from the steel cylinder to the tube, the thickness of the

tube, the diameter of the copper wire for the actuator coil, the thickness of the epoxy to

hold the actuator coil onto the tube and the distance from the epoxy to the permanent

magnets. The length of the airgap was initially set to be 5.8mm. The thickness of the

permanent magnets at the centre was 15.2mm. It was therefore determined using

Equation 6.9 that the magnetic flux density between the magnets and the steel cylinder

in the airgap was approximately 1T.

6.2.3 Effect of the actuator coil current on the electromagnetic force

When current flows through the actuator coil, the magnetic field produced by the

magnets may be increased or reduced, depending on the direction of the magnetic field

produced by the actuator coil. The operating point of the permanent magnets in the

second quadrant will move accordingly. For example, in the position where the

magnetic field produced by the actuator coil is in the same direction as the permanent

magnets, the operating point will move up the curve, as shown by P1 in Figure 6.4, and

if the field is in opposite direction, the operating point will move down the curve, as

shown by P2.

The actuator coil of N turns carrying a current I is then considered using the same

approximation. Ampere’s Law can be expressed as:

g gH l NI= (6.10)

where: N = Turns of the actuator coil

I = Current in the actuator coil (A)

Thus the magnetic flux density produced by the actuator coil in the airgap is given by:

0gg

NIB

lµ= (6.11)


159

The actuator coil was made up of 79 turns with an estimated current of 50A. The

maximum flux density produced by the actuator coil in the airgap is therefore 0.85T.

The actuator can be regarded as a linear motor which produces a linear force along its

axial direction. When the actuator coil is stationary, it can be simplified to a resistor and

inductor in series. It is also assumed here that the inductance L remains constant. When

the actuator coil starts to move, a back-emf is produced across the actuator coil, whose

polarity is the reverse of the input voltage. Therefore, when the actuator coil is supplied

from a power supply, the instantaneous current flowing through the actuator coil is

given by [118]:

diV RI L E

dt= + + (6.12)

where: V = Power supply voltage (V)

R = Resistance of the actuator coil (Ω)

L = Inductance of the actuator coil (H)

di

dt = Rate of change of current in the actuator coil (A/s)

E = Back-emf (V)

where:

g mf cE B l v= (6.13)

where: lmf = Length of the actuator coil in the magnetic field (m)

vc = Velocity of the copper coil (m/s)

The electromagnetic force is produced on the actuator when it carries a current in the

magnetic field. The Lorentz force acts along the axial direction to open and close the

vacuum interrupter. This force is linearly proportional to the current and the magnetic

flux density in the airgap:

g mfF B Il= (6.14)


160

The length of the actuator coil in the magnetic field is 15.8m and the total length of the

actuator coil is 35m. Therefore, if the actuator coil carries a current of 50A, the total

force on the actuator coil would theoretically be 790N. However, due to flux leakage

and fringing effects, the average magnetic flux density would be lower than 1T, and

therefore the force would be lower than 790N.

6.3 Finite element model

The magnetic field distribution in the actuator is non-linear and three-dimensional,

which makes it difficult to evaluate accurately using analytical methods. The actuator

therefore was analysed using a 3-D FE model. The FE model was used to determine the

steady state magnetic field distribution and the static electromagnetic force.


The 3-D FE model of the actuator was built and the geometry of the model is shown in

Figure 6.5: on the left is the actuator with the two steel walls at the front removed to

show the actuator coil and the permanent magnets in the middle; on the right is a plan

view of the actuator. Different colours are used to represent different materials. Green

represents steel and purple represents the permanent magnets. The actuator coil was

modelled is a red annular conductor.

Figure 6.5 Geometry of the actuator displayed in Vector Fields Opera: full view

without steel walls at the front (left) and plan view (right)


161

Due to the symmetric structure with respect to the YZ and ZX planes, the model was

reduced to a quarter, as shown in Figure 6.6. The mesh and computational time

therefore significantly reduced. It should be noted that symmetry cannot be used on the

actuator coil in the FE model.

Figure 6.6 Geometry of the actuator displayed in Vector Fields Opera (only a quarter is

shown using model symmetry)

The characteristics of the magnetic materials were specified by user-defined B-H

curves. The demagnetisation curve of N48 permanent magnet material was specified, as

shown in Figure 6.7 [143]. The magnetisation direction of the magnet was defined as

pointing perpendicular to the steel wall. A non-linear B-H curve, as shown in Figure

6.8, was used for modelling the EN1A steel blocks [147]. EN1A low carbon mild steel

was chosen to manufacture the steel blocks because it is easy to machine and has a

reasonable permeability.

0

0.5

1

1.5

-1000 -800 -600 -400 -200 0

Magnetic field strength (kA/m)

Mag

net

ic f

lux

den

sity

(T

)

Figure 6.7 N48 permanent magnet normal demagnetisation curve at 20ºC [143]


162

0

0.5

1

1.5

2

0 5 10 15 20

Magnetic field strength (kA/m)

Mag

net

ic f

lux

den

sity

(T

)

Figure 6.8 EN1A mild steel B-H curve [147]

6.3.2 Magnetic field distribution in the airgap

The current density for the actuator coil was initially set to be zero to simulate the

magnetic field produced by the permanent magnets.

Figure 6.9 shows a 3-D plot of the flux density distribution produced by the permanent

magnets. The flux density at the corner of the steel blocks where the steel walls and the

cylinder were attached to the bottom steel block is large.

Figure 6.10 shows the flux density at the midway point of the airgap between the steel

cylinder and the magnets. It is clear that the flux density is focused at the four sections

where the permanent magnets are located. The maximum flux density in the airgap is

0.965T, which confirms that the maximum flux density calculated from the analytical

model of 1T is reasonable. The average flux density across the whole airgap surface is

0.45T.


163

Figure 6.9 3-D plot of the flux density distribution produced by the magnets

Figure 6.10 Flux density distribution in the airgap produced by the magnets


164

Figure 6.11 shows the detailed cross-sectional view of the flux density in the actuator.

The flux vectors (cones) represent the direction of the magnetic field, whilst the size

represents the strength of the magnetic flux density. It is clear that most of the magnetic

flux goes through the steel path and airgap; only a small portion near the ends of the

permanent magnet does not. Figure 6.12 shows the radial flux density in the airgap

along the path noted in Figure 6.11. The path is located at the centre of the arc of one of

the magnets from top to bottom and at the midway point in the airgap between the

magnet and the steel cylinder. The flux density, as expected, is relatively constant at the

centre of the magnet but reduces slowly at both ends due to flux leakage and fringing

effects. The flux density at the centre is 0.965T and the average flux density along the

path is approximately 0.9T.

Figure 6.11 Cross-sectional view of the flux density distribution with vectors produced

by the magnets


165

0

0.2

0.4

0.6

0.8

1

0 10 20 30 40 50 60 70 80

Length (mm)

Flu

x d

ensi

ty (

T)


path noted in Figure 6.11)

Figure 6.13 shows the plan view of the flux density distribution. It can be seen that the

airgap flux density around the magnet circumference is concentrated at the magnets.

Figure 6.14 illustrates the radial magnitude of the flux density along the magnet

circumference at the middle of the airgap.

The path is shown in Figure 6.13, and starts from the middle of one of the magnets and

continues to the middle of the adjacent magnet. The maximum flux density occurs at the

mid-point of the magnets, as expect, whilst the minimum flux density occurs at the mid-

point between the magnets again, as expect. At the edge of the magnets, the flux density

is approximately 0.5T, which is about half of the maximum flux density. The magnetic

field at the edge is reduced because of the flux leakage and fringing effects. The average

flux density along this path is 0.48T.


166

Figure 6.13 Plan view of the flux density distribution with vectors produced by the

magnets

0

0.2

0.4

0.6

0.8

1

0 10 20 30 40 50 60 70 80 90 100 110

Length (mm)

Flu

x d

ensi

ty (

T)


path noted in Figure 6.13)


167

6.3.3 Effect of the actuator coil current on the electromagnetic force

The actuator coil was modelled as a solid annular conductor with a cross-sectional area

of 67.5mm2 (height of 75mm and thickness of 0.9mm). It was assumed that the

maximum current flowing through the actuator coil was 50A. The current density of the

conductor was therefore specified as 58.5A/mm2 in the FE model.

The direction of current through the actuator coil was initially set to be anti-clockwise.

The magnetic field in the inner steel cylinder is increased because the field from the

actuator coil is in the same direction as that from the magnets. The actuator coil will be

subjected to a downward force along the negative Z-axis and this provides the force to

open the vacuum interrupter. Figure 6.15 shows the resultant radial flux density

distribution produced by the magnets and the actuator coil. It is clear that the magnetic

field in the steel is significantly increased. The corner between the steel cylinder and the

bottom steel block clearly becomes saturated.

Figure 6.16 shows the flux density in the airgap along the vertical path denoted in

Figure 6.11. At the top of the path the magnetic field is considerably increased because

of the field produced by the actuator coil; however, at the bottom of the path the

magnetic field is slightly reduced because part of the steel blocks are saturated by the

increased magnetic field produced by the actuator coil and the magnetic circuit

reluctance increases. The overall average flux density along the vertical direction is

increased. The actuator coil initially receives a downward force of 882N at the top

position. The force produced for a 5mm displacement varies from 882N to 875N. This

confirms that it is an excellent linear actuator providing nearly constant force along its

whole displacement.

The FE model was resimulated with the current direction changed to clockwise. The

flux density distribution with the actuator coil carrying 50A in the clockwise direction is

shown in Figure 6.17. The radial flux density along the vertical path in the airgap from

top to bottom is shown in Figure 6.18. It is clear that the magnetic field in the airgap at

the top of the path is now significantly reduced by the current in the actuator coil and is

slightly increased at the bottom. There is a vertical force of 547N, which tends to move


168

the actuator coil upwards. The contact gap in the vacuum interrupter would be closed by

this force.


actuator coil carrying 50A in the anti-clockwise direction

0

0.2

0.4

0.6

0.8

1

1.2

0 10 20 30 40 50 60 70 80

Length (mm)

Flu

x d

ensi

ty (

T)


actuator coil carrying 50A in the anti-clockwise direction (along the vertical path)


169


actuator coil carrying 50A in the clockwise direction

0

0.2

0.4

0.6

0.8

1

0 10 20 30 40 50 60 70 80

Length (mm)

Flu

x d

ensi

ty (

T)


actuator coil carrying 50A in the clockwise direction


170

-1000

-800

-600

-400

-200

0

200

400

600

800

1000

0 10 20 30 40 50

Current (A)

Fo

rce

(N)

Fc - Fr

Closing force, Fc

Reluctance force, Fr

Fo - Fr

Opening force, Fo

Figure 6.19 Actuator force versus current characteristics

The FE model was simulated with the current in the actuator coil varying from 10A to

50A, in steps of 10A, for both the clockwise and anti-clockwise directions. The force on

the actuator coil is illustrated in Figure 6.19, showing that for the same current, the

magnitude of force to open the actuator, Fo, is higher than that to close the actuator, Fc.

The difference between the opening force and closing force is due to asymmetry in the

structure of the actuator that results in a reluctance force component, Fr, which is also

shown in Figure 6.19. The reluctance force shown in Figure 6.19 is calculated by the FE

model by setting the permanent magnet material property to that of air. For the chosen

coordinate definition, the reluctance force is always along the negative Z-axis, is

independent of the current direction and increases closely to the square of the coil

current, as would be expected.


171

6.3.4 Magnetic latch

The actuator does not have an inherent stable position when the power supply to the

actuator coil is removed. It is important that if the contacts are in the open or closed

position, the actuator is safely secured in that position. Magnetic latches at both ends

were employed therefore to hold the actuator in position [118, 148], as shown in Figure

6.20. The top magnetic latch was made from a steel plate, a magnet and a latch steel.

The steel plate was fixed at the top of the outer steel wall, forming part of the low

reluctance path for the top latch. The top latch magnet was placed beneath the steel

plate. The latch steel was then placed between the top latch magnet and the main

magnet, as shown in Figure 6.20. The latching force at the top (closed position) was

provided by the top latch magnet, whilst the latching force at the bottom (open position)

was provided by the main magnet.

Figure 6.20 Geometry of the actuator with magnetic latches (the two steel walls and

latches at the front are not shown) (left) and cross-sectional view (right)

The latch magnetic circuit is complex and non-linear; it was almost impractical

therefore to calculate the flux density distribution and force accurately using an

analytical model. The magnetic latches were added into the FE model to simulate the

magnetic flux density distribution in the airgap between the magnet and latch steel and

to estimate the latching force. It was found that the magnetic field in the airgap between


172

the magnet and the latch steel is higher when the latch steel is at the top position. The

latch steel therefore was designed to have a bigger area at the bottom than the top, as

shown in Figure 6.20 (right), to increase the latching force at the bottom position.

The force on the latch steel was calculated using Maxwell stress tensor over the surface

surrounding it [118, 148]. The latching force from one latch at the top was 52N. Two

top latches were employed to keep the actuator mechanically balanced. They would

provide a force of 104N in total, which was more than enough to hold the interrupter in

the closed position. Four sets of latches at the bottom position were used: two of these

contained the top latch magnets. The latching force from those with the top latch

magnets was 14N, whilst for the latches without the top latch magnets it was 24N;

therefore, the total latching force in the open position would be 38N.

Coil current is necessary to release the magnetic latches when operating the vacuum

interrupter. This current creates a force on the actuator coil in a direction opposing to

the latching force. The resultant force on the moving part is therefore the force from the

actuator coil minus the magnetic latching force. The force from the actuator coil has to

be larger than the magnetic latching force to operate the vacuum interrupter.

6.4 Design of the full operating actuator

To design a high performance operating actuator for the vacuum interrupter, many

factors need to be considered such as contact popping, bounce, rebound and welding

[94]. The reasons that cause these problems and how to design the operating actuator to

satisfy the requirements will be discussed in the following sections.

6.4.1 Contact popping

Butt contacts are widely used in vacuum interrupters for low interrupting current levels

(<7kA) [94, 149]; the actual contact area is considerably smaller than the apparent area

of the contacts due to macro-roughened surfaces. The magnetic field of the current

passing through the contacts produces an electromagnetic repulsive force tending to

force the contacts apart, called contact popping. This repulsive force is more significant

when high current flows through the closed contacts. When the contacts separate,

popping would lead to arcing in the limited space inside the vacuum interrupter, and this


173

could cause severe erosion of the contact surface. In order to prevent contact separation,

a minimum static contact holding force is usually required. The common way to provide

this force is to place a spring along the driving shaft near the vacuum interrupter. The

spring is pre-compressed (commonly called the wipe spring) in order to provide a

considerable force immediately after the contacts are closed. During normal operation,

the contacts would gradually wear and erode. The compression of the spring and

holding force are then reduced, so the pre-compression of the spring helps minimise this

reduction [94].

The required contact holding force for a DVS10CB vacuum interrupter is 2kgf when the

contacts carry 3.2kA. The wipe spring with a stiffness of 2N/mm therefore is placed

next to the movable contact. The spring is pre-compressed by 10mm, so a force of 20N

is placed on the closed contacts. To provide the holding force reliably, the stroke of the

actuator is designed to be 5mm, which is 3mm further than the stroke of the vacuum

interrupter. An extra force of 6N therefore exists on the contacts when the actuator is in

the closed position.

6.4.2 Contact bounce and rebound

Contact bounce on closing is undesirable because, like popping, it can cause arcing

when the contacts separate. It is necessary to analyse the contact dynamics to avoid

contact bouncing. When closing the vacuum interrupter, the movable contact is

travelling towards the stationary contact at a high velocity. The ideal condition would be

that the two contacts move together at the same velocity after collision and slow down

gradually by transmitting the kinetic energy. To control contact bounce, one should

minimise the kinetic energy that must be dissipated and maximise the rate of

dissipation. This kinetic energy could be transmitted into a supporting structure and

eventually dissipated or stored in other components. A flexible supporting spring can be

placed next to the stationary contact of the vacuum interrupter to minimise bounce [94,

150]. The stiffness of the spring has to comply with the following equation [94]:

2 22 1 2 1

21 10

[(1 / ) 1](1 / )M M M M FK

M v

+ − +< (6.15)


174

where: K = Spring stiffness (N/mm)

M1 = Mass of the moving part (kg)

M2 = Mass of the stationary part (kg)

F = Resultant force on the vacuum interrupter in the closed position (N)

v10 = Initial velocity of the moving contact before collision (m/s)

A compression spring with a stiffness of 5N/mm was placed therefore between the

interrupter stationary contact and the supporting plate, forming a flexible support

structure to absorb the kinetic energy and avoid contact bouncing.

Contact rebound on opening is a similar problem to contact bounce. If the rebound

brings the contacts close to each other again, the dielectric recovery of the vacuum

interrupter could be compromised. The movable contact therefore must be stopped and

held at the end of the opening stroke. To ensure the movable contact does not rebound,

the kinetic energy has to be absorbed at impact. Building some dissipative element into

the mechanical stopper is feasible. In this prototype, contact opening rebound is

overcome using the magnetic latches, to provide an adequate holding force and a rubber

pad as a damper.

6.4.3 Contact welding

Arcing caused by contact popping and bouncing may melt the surfaces of the contacts,

and this may form a weld when the contacts come together. Although the expected

interrupting current level is much lower than its maximum value, contact welding is still

unavoidable. It is necessary therefore to build a mechanism to break any contact weld

when opening the vacuum interrupter. A snatch bracket is placed between the actuator

and the wipe spring to fracture contact welding. During the process of closing the

vacuum interrupter, the actuator travels further after the contacts are closed, to compress

the wipe spring and provide the holding force to avoid popping. When the vacuum

interrupter opens, the actuator has to travel back the same distance before the contacts

separate. When the contacts start to separate, a significant amount of kinetic energy

therefore has already accumulated in the moving part of the mechanism. This energy is

used to fracture the welding by snatching the moving contact and pulling it away from

the stationary contact.


175

6.5 Design of prototype actuator and interrupter

The component properties for the prototype operating actuator such as material and size

are summarised and listed in Table B.1 in Appendix B. After the size of each

component was finalised, the structure of the prototype unit was designed in VariCAD.

The geometry of the 3-D structure of the vacuum interrupter and operating actuator is

shown in Figure 6.21. The complete prototype unit consisted of an actuator, a snatch

bracket, a driving shaft, a wipe spring and a supporting structure.

The actuator includes the steel blocks and cylinder, permanent magnets, an actuator coil

and magnetic latches. The design of the vacuum interrupter is shown in Figures 6.21 to

6.24. A thin fibreglass tube was used to support the actuator coil and prevent it from

deforming. The actuator coil was wound tightly on the fibreglass tube. The steel

cylinder in the middle acts as a guide for the fibreglass tube to move up and down along

its axial direction. A carbon fibre plate was attached on the fibreglass tube to deliver the

force from the actuator coil. Four latch steel blocks were fixed on the carbon fibre plate

to hold the moving part in the open or closed position. Two steel plates were mounted at

the top edge of the steel walls and the latch magnets were placed beneath them. They

provided the magnetic latch in the closed position. Mechanical stoppers are also

designed at the end of both strokes.

The snatch bracket was made from a hollow aluminium cylinder and was fixed onto the

carbon fibre plate by a plastic screw. A brass bolt was used as a shaft to deliver the

force. A wipe spring was placed surrounding the shaft between the snatch bracket and

the movable contact of the interrupter. A supporting spring was placed between the

stationary contact of the interrupter and the supporting plate. The supporting plate was

made of paxolin, a type of synthetic resin bonded lamination (SRBL). The supporting

plate was fixed to the steel walls by four aluminium studs. The supporting plate has two

functions: firstly, it can hold the supporting spring; and secondly, it can provide

electrical insulation terminals for the actuator coil power supply and the vacuum

interrupter.


176

Figure 6.21 3-D structure of the operating actuator

Figure 6.22 Cross-sectional view of the vacuum interrupter actuator

(all dimensions in mm)


177

Figure 6.23 Plan view of the vacuum interrupter actuator


Figure 6.24 Plan view of the stopper and carbon fibre plate with latch steel



178

6.6 Construction of prototype operating actuator

6.6.1 Actuator stationary part

N48 permanent magnets of the size listed in Table B.1 were manufactured by Arnold

Magnetic Technologies Corporation. Nd-Fe-B is susceptible to corrosion, so nickel

plated magnets were used to protect them from corrosion. A picture of one of the

permanent magnets is shown in Figure 6.25. It can be seen that it has a surface arc to

keep the airgap radial length constant.

Figure 6.25 N48 Nd-Fe-B permanent magnet

The steel blocks and the cylinder were manufactured from EN1A mild steel because it is

easy to machine. The steel cylinder was bolted at the middle of the bottom steel block.

Each magnet was placed on the steel wall with great care and held in place by a tufnol

square frame. The four steel walls with magnets mounted were then bolted onto the

bottom steel block. The steel frame with the permanent magnets is shown in Figure

6.26.


179

Figure 6.26 Steel frame with permanent magnets

6.6.2 Actuator moving part

A carbon fibre plate and fibreglass tube was supplied by Tri-cast Composite Tubes Ltd.

The plate was machined round with a diameter of 156mm and then cut to produce four

flat edges, as shown in Figure 6.24. An annular slot with a depth of 2mm was cut into

the plate to place the fibreglass tube. This would provide more contact surface area

between the plate and tube. AWG19 copper wire with a diameter of 0.912mm was used

for the actuator coil. The temperature rise of the actuator coil was not a problem because

the actuator current was only a pulse of a few tens of milliseconds duration with a

maximum current of 50A. The carbon fibre plate and fibreglass tube were placed on the

lathe and 79 turns of copper wire were wound onto the tube. Two small holes were

drilled in the plate to allow both connection ends of the copper wire to pass through.

Four blocks of latch steel were screwed onto the flat edges on the carbon fibre plate.

The top plate and actuator coil on the fibreglass tube is shown in Figure 6.27.

Figure 6.27 Actuator coil on the fibreglass tube


180

6.6.3 Complete prototype

The magnetic latches and stoppers in the closed position were assembled at the top edge

of the steel walls. The snatch bracket, shaft, wipe spring, vacuum interrupter, supporting

spring and plate were assembled in turn afterwards. The ends of the actuator coil were

then connected to the terminals on the supporting plate. The assembled prototype

operating actuator with the vacuum interrupter is shown in Figure 6.28.

Figure 6.28 Operating actuator with vacuum interrupter


181

6.7 Conclusions

The structure of the vacuum interrupter actuator was analysed and calculated using the

analytical and the FE models. The analytical model was used to calculate the maximum

magnetic field in the airgap. The actuator was taken as an idealised linear model without

considering the reluctance of the steel, flux leakage and fringing effects.

The FE model was built to analyse the flux density distribution and the static force on

the actuator coil with current flow through it. The flux density distribution from the FE

model demonstrated that the analytical model was reasonable. The simulation shows

that the force to open the vacuum interrupter was higher than to close it. This was

because the reluctance force between the actuator coil and the steel frame has the same

direction as the force to open the interrupter. This satisfied the objective that it is more

important to have a higher force to open the actuator than to close it. The force along the

displacement was found to be nearly constant. Magnetic latches in both the open and

closed positions were also added into the FE model. However, the FE model did not

consider actuator coil movement, back-emf, material hysteresis and eddy currents, but it

still provided useful guidance for the design.

A high performance operating actuator were designed and built. The wipe spring,

supporting spring, stopper with rubber pad, magnetic latches and snatch bracket were

included to overcome contact popping, bounce, rebound and welding.

Chapter 7: Design of the Actuator Control Circuit

182

7 Design of the Actuator Control Circuit

7.1 Introduction

The actuator control circuit was designed for two purposes. Firstly, after the SFCL coil

quenches, the voltage across it will increase rapidly due to the increasing coil resistance.

When the voltage is higher than a pre-set threshold voltage level, an ‘open’ signal must

be automatically generated to trigger the vacuum interrupter to open. Secondly, after the

fault is cleared, a ‘close’ signal needs to be generated to close the vacuum interrupter

once the ‘close’ button is pressed. The force required to open and close the vacuum

interrupter is generated from the electromagnetic force on the actuator coil. The

direction of the magnetic field is unchanged; the force direction therefore is changed by

reversing the polarity of the voltage and current.

7.2 Control circuit

7.2.1 Topology selection

The polarity of output voltage and current can be controlled using a full-bridge DC-DC

converter [151]. A single phase full-bridge DC-DC converter, consisting of two legs,

was chosen. Each leg was made up of two switches and their anti-parallel diodes. Figure

7.1 shows a schematic diagram of the actuator control circuit. MOSFETs (G1 to G4)

were selected because they are suitable for low voltage and low current applications.

Their integral body diodes are labelled D1 to D4. The DC bus capacitor C1 was pre-

charged to a voltage between 50V and 100V by a DC power supply, to provide the

energy required to open and close the vacuum interrupter. Resistor R1 was used to limit

the surge current and protect the capacitor [151].

The operating principle for a full-bridge DC-DC converter is explained below. It should

be noted that the positive direction for the coil voltage and current is defined as shown

in Figure 7.1. When MOSFETs G1 and G4 are turned on, G2 and G3 are turned off, a


183

positive voltage therefore is imposed on the actuator coil, which causes the current in

the actuator coil to flow in the positive direction. The coil current interacts with the

magnetic field produced by the magnets to produce a downward force and open the

vacuum interrupter. When MOSFETs G1 and G4 are turned off, the current cannot

immediately decay to zero due to the coil inductance. The current continues to flow

through diodes D2 and D3 until it is fully discharged. During this period, a negative

voltage is imposed on the actuator coil through the diodes, which forces the current to

decay to zero. When MOSFETs G2 and G3 are turned on, G1 and G4 are turned off, the

direction of the voltage and current reverses. The coil therefore generates an upward

force to close the vacuum interrupter. When MOSFETs G2 and G3 are turned off, the

current flows through diodes D1 and D4 until it decreases to zero. Controlling the

polarity of the output voltage and current of the full-bridge DC-DC converter therefore

controls the opening and closing of the interrupter.

+C1

G1-G G3-G

R1

Actuator coil

G2-G G4-G

G1-S

G2-S

G3-S

G4-S

G4G2

G3G1

+ -i v

D1

D2

D3

D4

GND

50V

-100

V

Figure 7.1 Schematic diagram of the actuator control circuit

7.2.2 Components selection

The components required for the control circuit include MOSFETs, a capacitor and a

current limiting resistor.

The maximum voltage on the capacitor was 100V. It was necessary to work out the

actuator coil resistance to estimate the maximum current level. The resistance of the

actuator coil is given by:


184

ac

lR

A

ρ= (7.1)

where: l = Total length of the actuator coil (m)

Aac = Cross-sectional area of the actuator coil (m2)

The total length of the actuator coil was 35m, the cross-sectional area was 0.64mm2, and

the resistivity of copper was 1.68×10-8Ω·m at 20ºC. The resistance of the actuator coil

therefore was 0.92Ω. The total resistance of the actuator coil circuit including the

connection wires was about 1Ω. The maximum current through the actuator coil would

be 100A when it stops moving. MOSFETs were chosen therefore for a rating greater

than 120V and 120A; about 1.2 times the maximum operating levels to guarantee a

voltage and current safety margin [152]. IXFH120N20P N-channel enhancement mode

power MOSFETs with a rating of 200V/120A were chosen for the control circuit [153].

The criteria used to select the capacitor were the voltage level and capacitance. The

maximum voltage on the capacitor was 100V, which required the capacitor to be rated

greater than 120V (using a 20% safety margin). When the capacitor was driving the

actuator coil, the capacitor and actuator coil can be regarded as a simple RC circuit. The

voltage across the capacitor, which is time dependent, can be expressed by using

Kirchhoff's current law:

0c c

ac

dV VC

dt R+ = (7.2)

where: C = Capacitance (F)

Vc = Voltage across the capacitor (V)

Solving Equation 7.2 yields:

0( ) ac

t

R CcV t V e

−= (7.3)

where: V0 = Voltage across the capacitor at time t = 0 (V)


185

The voltage across the capacitor has an exponential decay, so the worst case situation

was considered to guarantee the capacitor can provide enough current to drive the

actuator coil for a certain period of time, i.e. 10ms. The capacitor was pre-charged with

50V; the requirement was that the voltage across the capacitor was no less than 30V

after discharging through the actuator coil for 10ms. From Equation 7.3 the minimum

capacitance was therefore 23mF. Three 8.25mF/450V aluminium electrolytic capacitors

from SIC-SAFCO were connected together in parallel to produce a total capacitance of

24.75mF.

When the capacitors were fully discharged, the maximum voltage difference between

the power supply and capacitor was 100V. A resistor of 20Ω was chosen to limit the

surge current to 5A.

7.3 Trigger signal and MOSFET drive circuit

7.3.1 Trigger signal circuit

The method to determine whether the SFCL coil has quenched or not, was to compare

the coil voltage with the pre-set threshold voltage level. When the coil is in the

superconducting state, the amplitude of the coil voltage is lower than the pre-set

threshold voltage level. Once a fault occurs, the coil starts to quench and the amplitude

of the voltage across the coil increases quickly and will exceed the pre-set threshold

voltage level, which then generates a signal to open the vacuum interrupter. However,

the voltage across the coil is sinusoidal and therefore an absolute value circuit was

necessary to convert it into a positive output signal.

A precision full-wave rectifier circuit was used, as shown in Figure 7.2 [154]. When the

coil voltage is positive, the output of operational amplifier (op-amp) (U3) is negative, so

diode D5 is reverse biased and diode D6 is forward biased. This closes the feedback

loop around op-amp U3 through resistor R21 and forms an inverting amplifier. Op-amp

U4 sums the negative of twice the output of op-amp U3 and the negative of the input

coil voltage, leaving the output of op-amp U4 the same as the coil voltage. When the

coil voltage is negative, diode D5 is forward biased and diode D6 is reverse biased,

which closes the feedback loop around op-amp U3. Op-amp U4 inverts the coil voltage


186

resulting in a positive output. Thus, the output of op-amp U4 is a positive voltage that

represents the absolute value of the coil voltage, no matter whether it is positive or

negative.

R13

10K C31220pF

Vcoil

C7

0.1uF+15V

GND

GND-15V

R1820K

D5

1N4148

D61N4148GND R19

10K

R2110K

R17

10K

+15VC8

0.1uF-15V

GND

R20

20K

3

26

1 5

74

U3741

3

26

1 5

74

U4

741

-15V

+15V

Abs_Vcoil

Figure 7.2 Schematic diagram of the precision full-wave rectifier [154]

The circuit to generate the ‘open’ signal for the vacuum interrupter is shown in Figure

7.3. The absolute value of the coil voltage is compared with the pre-set threshold

voltage level by the voltage comparator (U5). The pre-set threshold voltage level, which

is connected to the positive input of voltage comparator U5, could be adjusted from 0V

to 7.2V by a potentiometer. When the coil is superconducting, the absolute coil voltage

is lower than the pre-set threshold voltage level and the output of voltage comparator

U5 is held at a high voltage level by a 10kΩ pull-up resistor. After the coil starts to

quench, the output of voltage comparator U5 transitions to a low voltage level until the

coil voltage exceeds the pre-set threshold voltage level. The output of voltage

comparator U5 returns to the high voltage level when the coil voltage drops below the

pre-set threshold voltage level. The output of voltage comparator U5 is connected to the

trigger input of LM555 timer (U2) through a capacitor. Timer U2 is used to build a

monostable circuit [154, 155]. Normally the trigger input of timer U2 is held at the high

voltage level by a 2.4kΩ pull-up resistor. The output of timer U2 stayed at the low

voltage level. The trigger pulse must be of shorter duration than the intended output

pulse to avoid it being triggered again. When the output of voltage comparator U5

transitions to the low voltage level, the trigger pulse changes to a short pulse through

capacitor (C9). When the trigger input falls below a third of the supply voltage, a pulse

is generated for a certain period of time. The duration of the output pulse is determined

by the time constant of an RC network, which consists of a capacitor (C17) and a


187

resistor (R9). The output pulse terminates when the charge on capacitor C17 equals two

thirds of the supply voltage. The period of output pulse t is given by:

1.1 9 17t R C= × (7.4)

The initial period for the ‘open’ signal is set at 10ms. The resistor R9 and the capacitor

C17 were chosen to be 91kΩ and 0.1µF. The output pulse width could be lengthened or

shortened during testing by adjusting the values of resistor R9 and capacitor C17.

3

21

84

U5A

LM293R1110K

R10

5.1K

C22

0.1uF R1210K

+15V

C230.1uF

GND

GND

TRIG2

Q3

R4

CVolt5 THR 6

DIS 7

VC

C8

GN

D1

U2

555

OPEN

C150.1uF

GND

R9

91K

+15V+15V

C14

0.1uF+15V GND

+15V

GND

R144.7K

C12

0.1uF+15V GND

C9

0.1uF

R162.4K

C24220pF

R15

1M

Abs_Vcoil

Figure 7.3 Schematic diagram of the ‘open’ signal circuit for the vacuum interrupter

The signal to close the vacuum interrupter is produced by another LM555 timer (U1)

monostable circuit, as shown in Figure 7.4. Normally the trigger input of timer U1 is

held at the high voltage level by a 2.4kΩ pull-up resistor and the output of timer U1

stays at the low voltage level. When the ‘close’ button is pressed, the voltage across the

button becomes low. The low pulse changes to a short low pulse by passing through

capacitor (C6), which triggers timer U1 to produce the output pulse. The duration of the

‘close’ signal was initially set at 10ms, which was the same as the ‘open’ signal. A

ceramic capacitor was placed in parallel with the button to eliminate contact bouncing

in the button.


188

TRIG2

Q 3R4

CVolt5 THR 6

DIS7

VC

C8

GN

D1

U1

555

CLOSE

C100.1uF

GND

C11

0.1uF

R8

91K

+15V

R72.4K

S1

CLOSE

C4

0.1uF

+15V

GND C6

0.1uF

R210K

C50.1uF

Figure 7.4 Schematic diagram of the ‘close’ signal circuit for the vacuum interrupter

7.3.2 MOSFET drive circuit

The gate drive circuit is an important interface between the control signals and the

power MOSFETs. The gate signal must be positive with respect to the source terminal

in order to switch the MOSFET. It is easy to implement a driver for a low-side

MOSFET, i.e. MOSFET G2 or G4, for example, because its source terminal is

connected to ground. However, the source terminal of a high-side MOSFET, i.e.

MOSFET G1 or G3, for example, can be floating between ground and the DC bus

voltage level depending on the states of the MOSFETs in the same leg.

In order to switch the high-side MOSFET, it is necessary to employ a level shifting

circuit and a floating power supply. Level shifting is used to convert a control signal to

a floating signal. There are four potential options to achieve this [151, 156, 157]:

transformer level shifting, opto-couples level shifting, fibre optic link level shifting and

electronic level shifting. The last approach does not provide isolation, but it is a low

cost solution and can be integrated onto a simple chip. International Rectifier

Corporation has developed a series of gate drive ICs utilising this technique [158]. A

floating supply can be achieved using one of the following methods. The simplest way

is to adopt a transformer isolated supply. The second method is ‘charge-pump supply’,

which overlays the voltage of one supply onto another. It is not commonly used for the

driver supply due to its complexity [159]. The last method is a very common and

popular technique called ‘bootstrap supply’, which is made up of a diode and a storage

capacitor. The diode anode is connected to the same power supply as the low-side driver.


189

The capacitor is charged through the diode when the low-side MOSFET is turned on.

When the low-side MOSFET is turned off, the high-side driver circuit can be powered

from the storage capacitor.

The MOSFETs in the full-bridge DC-DC converter were driven by two IR2113 IC

drivers [160]. These are high voltage and high speed drivers with independent high-side

and low-side referenced output channels. The bootstrap technique is widely used

together with IR2113 IC drivers for high frequency power converters, where switching

devices in the same leg are turned on in sequence. However, it is not suitable for this

application because the capacitor in the bootstrap technique can only be charged when

the low-side MOSFET is turned on. For the actuator control circuit, the MOSFET pairs

G1 and G4 or G2 and G3 only operate occasionally for 10ms, to open or close the

vacuum interrupter. An isolated high frequency DC converter TMA1515S [161] from

Traco Power therefore was used to provide the floating power supply for the high-side

driver. A 20kΩ dummy load resistor was connected to the output of the TMA1515S to

limit any surge voltages. A MOSFET driver for one leg is shown in Figure 7.5. The

structure of the driver for the other leg was the same except the connections for the

‘open’ and ‘close’ signals were swapped. Pull-down resistors were connected to the

high-side and low-side inputs to keep the inputs at the low voltage level when there was

no signal. Four 20Ω resistors, which were recommended in the IXFH120N20P

datasheet, were used as the driving resistors.

LO 1

COM 2

VCC 3

NC 4

VS 5

VB 6

HO 7

VDD9

HIN10

SD11

LIN12

NC8

NC14

VSS13

U6

IR2113

R265.1K

C18

0.1uF

C13 0.1uF

GND

+15V

OPEN

CLOSE

R22 20

R24 20

GND

G2-G

+15V

+C26 10uF

G1-S

G1-G

C19 0.1uF

+C29 10uF

+C25

10uF

R5 10K

R3 10K

GND

Vcc1

GND2 +Vout 4

-Vout 3U8

TMA1515SGND

+15VG1-S

VB1

VB1

R25 20K

Figure 7.5 Schematic diagram of the MOSFET drive circuit


190

The trigger signal and MOSFET prototype drive circuit were built on a veroboard, as

shown in Figure 7.6. To assemble the MOSFETs onto a heat sink, a thin layer of

silicone paste was placed at the back of the MOSFETs to provide good thermal contact.

The MOSFETs were then screwed onto the heat sink with insulation cloth placed

between the two surfaces. A picture of the vacuum interrupter actuator and its control

circuit is shown in Figure 7.7.

Figure 7.6 Trigger signal and MOSFET drive circuit

Figure 7.7 Vacuum interrupter with its actuator and control circuit


191

7.4 Conclusions

A full-bridge DC-DC converter was designed to open the vacuum interrupter

automatically once the SFCL coil quenched and to close the vacuum interrupter after a

fault was cleared.

An absolute value circuit was designed to ensure the coil voltage was positive. This was

then compared with a pre-set threshold voltage level using a simple comparator. Once

the coil quenched, the absolute coil voltage rose above the pre-set threshold voltage

level. This caused the comparator to transition to a low logic level, triggering a LM555

timer monostable circuit to generate a pulse for 10ms. After a fault had cleared, a close

button was pressed to trigger another LM555 timer to produce a ‘close’ signal. These

pulses were sent to two IR2113 IC drivers to drive the corresponding MOSFETs. The

control circuit, trigger signal and drive circuit were designed, built and successfully

tested.

Chapter 8: Experimental Investigation of an SFCL Coil with Integrated Vacuum Interrupter

192

8 Experimental Investigation of an SFCL Coil with Integrated

Vacuum Interrupter

8.1 Introduction

The vacuum interrupter operating actuator and its control circuit was designed and

constructed as described in the previous chapters. It is necessary to test the operating

actuator separately and validate the results with the design models before mounting it

into the SFCL system. The actuator was initially tested at atmospheric pressure. The

ultimate aim was to integrate the whole unit into the vacuum chamber in the cryostat.

The actuator therefore was assembled into a separate vacuum chamber and tested. After

confirming that the actuator and its control circuit worked correctly in the vacuum

chamber, the vacuum interrupter was then placed into the SFCL system for testing.

Quench tests with and without the vacuum interrupter were compared. Simulated fault

tests were carried out to test the SFCL behaviour during a fault with and without the

vacuum interrupter. A fault was simulated experimentally by manually closing a switch

to short-circuit a load resistor.

8.2 Testing of the vacuum interrupter operating actuator

8.2.1 Actuator magnetic field and static force

The magnetic field of the actuator plays an important role in the behaviour of its

operation. The magnetic flux density in the airgap between the steel cylinder and the

permanent magnets was measured using a Gauss meter. The magnetic flux density at the

middle of the arc surface covering the four permanent magnets varied from 0.95T to

1.04T. The maximum magnetic flux density at each side was slightly different because

the airgap in each direction is not absolutely constant. The magnetic flux density

obtained from the analytical model and the FE model was 1T and 0.965T respectively.

The results from both models are within ±5% of the experimental measurements, which

is a good correlation. The experimental measurements also revealed that the magnetic


193

flux density at the middle of the permanent magnets was higher than at both edges,

which was the same as obtained from the FE model results.

The static opening force and closing force were measured using an S-type load cell,

Model 615, from Tedea Huntleigh [162]. The load cell was assembled between the top

plate and the carbon fibre plate, which was attached to the actuator coil. The force from

the actuator coil was transmitted to one end of the load cell through the carbon fibre

plate. The other end of the load cell was fixed to the top plate. The reading of the load

cell was obtained using a Maywood 2000 Digital Indicator, which was calibrated by an

Agilent 34420A NanoVolt/MicroOhm meter.

A series of tests were carried out on the actuator, which was supplied by a DC power

supply for ten seconds. The measured force was recorded as the applied current level

was increased from 10A to 50A. A second series of tests were conducted with the same

current level but in the opposite direction. The measured opening and closing forces

together with the predicted forces from the FE model are shown in Figure 8.1 from

which it is clear that the measured opening and closing forces exhibit a good correlation

with the FE model predicted forces. The measured closing force is approximately 10%

lower than the predicted force, which is within the measurement error band and

considerations associated with the mechanical implementation of the force sensing

transducer (i.e. transducer mass and stiffness).

The force measurement demonstrates that the opening force is higher than the closing

force as a consequence of the actuator reluctance. This additional reluctance force

designed for the opening direction of actuation is a desirable feature in terms of the

project objective where opening the vacuum interrupter in as shorter time period as

possible is a primary performance target.

An indirect method was used to measure the latching force in both the open and closed

positions. The current through the actuator coil was increased from 1A, in steps of 1A,

until the actuator coil started to move. The current level can be converted to force, from

the measured force versus current profile. The critical current that caused the actuator

coil to move was 3A in the open position and 8A in the closed position. The latching


194

force in the open and closed positions therefore were about 36N and 96N, respectively,

approximately 8% lower than predicted by the FE model. However, it was enough to

hold the vacuum interrupter in both positions.

0

200

400

600

800

1000

0 10 20 30 40 50

Current (A)

|Fo

rce|

(N

)

Opening force

Measured opening force

Closing force

Measured closing force

Figure 8.1 Actuator opening and closing force versus current

8.2.2 Opening the vacuum interrupter at atmospheric pressure

Opening the vacuum interrupter is more critical than closing. If the vacuum interrupter

could open faster when a fault occurs, less energy would be dissipated in the

superconducting coil and the coil could recover in a shorter period of time after the fault

clears. The opening operation of the vacuum interrupter therefore was carefully

evaluated in this section.

During the opening operation, the actuator was activated using its control circuit. The

voltage across the actuator coil was measured using a Lecroy AP300 differential voltage

probe. The current through the actuator coil was measured using a CP150 current probe

and the displacement of the actuator coil was measured by a DFg5 unguided miniature

linear variable differential transformer (LVDT) from Solartron Metrology. The LVDT

is a type of electric transformer, which is made up of three solenoidal coils around a

tube and a ferromagnetic core [163]. The stroke of the DFg5 LVDT is ±5mm. All the

signals were recorded using an oscilloscope and the data were plotted out in figures


195

using MATLAB. The response time constant of the DFg5 LVDT was about 3ms [164],

so the displacement is time-shifted by 3ms in the figures.

The opening operation was tested with the voltage of the storage capacitor charged to

50V. The test was repeated by increasing the capacitor voltage every 10V up to 100V.

The actuator was initially tested at atmospheric pressure. The current duration for

opening the vacuum interrupter was initially set at 10ms. Figure 8.2 shows the actuator

coil voltage, current and linear displacement when the vacuum interrupter is opening

with a capacitor voltage of 100V. The opening operation can be divided into the

following phases:

• 0-t1: The current increases quickly but limited by the inductance of the actuator

coil. When the current is less than 8A, the actuator coil is stationary because the

force on the actuator coil is less than the top latching force.

• t1-t2: Once the current exceeds 8A, the actuator coil starts to move. A back-emf,

which is proportional to velocity, appears across the actuator coil. The current

continues to increase but the rate of change of current reduces due to the back-

emf. The velocity of the actuator coil and back-emf are both increasing at the

same time. After 0.0025s, the current starts to drop slightly as the back-emf

increases.

• t2-t3: When the actuator coil achieves a 3mm displacement, the contacts of the

vacuum interrupter start to separate. They may start to open earlier but it is not

easy to determine the exact point they start to open. The current continues to

drop as the velocity increases. When the required force to open the vacuum

interrupter exceeds the force that the actuator coil can provide, the actuator coil

starts to decelerate and the back-emf starts to decrease. The current in the

actuator coil starts to increase steadily as the back-emf drops. When the current

is at its maximum level, the contacts are fully open and the back-emf across the

coil becomes negligibly small.

• t3-t4: The current drops slightly and becomes constant. The moving part of the

actuator moves slightly further compressing the rubber pad on the stopper. A

nearly constant current flows through the actuator coil after it is stopped and

holds the moving part in the open position.


196

• t4-t5: The MOSFETs are turned off after 10ms but the actuator coil current

cannot instantly drop to zero due to the coil inductance. The current therefore

freewheels through the MOSFET diodes. A negative voltage is then applied to

the actuator coil, forcing the current to decrease nearly linearly. After the current

decreases to zero, i.e. the energy stored in the inductance is dissipated, the

voltage across the actuator coil returns to zero.

Figure 8.2 Opening operation of the vacuum interrupter with a capacitor voltage of

100V at atmospheric pressure


197

During the 10ms operation, the voltage across the actuator coil is reducing because

energy is dissipated in the actuator coil. It should be pointed out that although the stroke

of the actuator was designed to be 5mm, the real stroke increased to 5.5mm because the

rubber pad on the stopper was compressed.

Figure 8.3 shows a comparison of the opening operation with different capacitor charge

voltages. It can be seen that it takes around 9.7ms, 8.5ms, 7.9ms, 7.3ms, 6.5ms and 6ms

respectively, for capacitor voltages from 50V to 100V to fully open the vacuum

interrupter. It is clear that a higher voltage on the capacitor produces a higher current in

the actuator coil and opens the vacuum interrupter in a shorter period of time.

Figure 8.3 Opening operation of the vacuum interrupter with different capacitor

voltages at atmospheric pressure


198

8.2.3 Opening the vacuum interrupter in a vacuum

The ultimate objective of this research project was to integrate the vacuum interrupter

and its actuator into the vacuum vessel within the cryostat. It was impossible to place

the unit directly into the existing cryostat because it was not initially designed for this.

To emulate this, the vacuum interrupter and its actuator were mounted into an external

vacuum chamber.

The actuator was held in place using a steel bracket inside the chamber, as shown in

Figure 8.4 (left). The connections to the vacuum interrupter and its actuator coil were

made of solid copper rods passed through a socket on the wall of the vacuum chamber.

Two copper rods with a diameter of 3mm were used for the actuator coil connections,

whilst another two copper rods with a diameter of 6mm were used for the vacuum

interrupter connections. A wooden block with four holes in the correct position for the

copper rods was manufactured. These four solid copper rods were pulled through the

wooden block and then a copper tube which had the same diameter as the socket on the

vacuum chamber, as shown in Figure 8.5. The wooden block and copper tube were held

together using two clamps. Sealing paste was pushed into the gap between the copper

rods and the copper tube to hold the copper rods in position and left until it cured. The

wooden block was then removed; 3M adhesive epoxy was filled into the rest of the

copper tube and left overnight until it cured to provide a good vacuum seal. Four copper

rods were then held in the correct position to ensured good electrical insulation between

each rod and socket. The bare copper rods were covered by heat shrink sleeving to

provide electrical insulation. The external and internal connectors were connected to

these four copper rods. The external connectors are shown in Figure 8.5.

The LVDT power supply and feedback signal were connected through a signal socket.

The vacuum chamber is shown in Figure 8.4 (right). A vacuum pump unit, which was

the same as the one used for the cryostat, was used to pump the air out from the

chamber. The pressure in the chamber was reduced to 0.014mbar during testing.


199

Figure 8.4 Operating actuator in the vacuum chamber (left) and external view of the

vacuum chamber (right)

Figure 8.5 External connectors for the vacuum interrupter and its actuator

The opening test in the vacuum chamber with a capacitor voltage from 50V to 100V, in

steps of 10V, is shown in Figure 8.6. There was a serious problem however when

analysing the displacement of the coil. When the initial voltage on the capacitor was

50V, the actuator coil rebounded by approximately 1mm. When the initial voltage on

the capacitor was 60V and 70V, the actuator rebounded and closed again after it opened.

When the initial voltage on the capacitor was 80V to 100V, the actuator operated

correctly.

The reason for this problem was the lack of air resistance in the vacuum chamber and

the instantaneous velocity of the actuator coil was too high to stop when it collided with


200

the stopper. That is why the rebound was more serious with a capacitor voltage of 60V

and 70V than 50V. When the capacitor voltage was increased, the opening velocity was

higher and the actuator opened faster. For example, the opening time is 5ms when the

capacitor voltage is 100V, then the actuator is held in the open position for 5ms before

the current is removed and this help to reduce any rebound. This explains why the

actuator works correctly with a capacitor voltage of 80V to 100V.


vacuum chamber


201

The duration of the current through the actuator coil was increased to 15ms to solve the

rebound problem. Figure 8.7 shows the opening of the vacuum interrupter with a current

duration of 15ms. It is clear that the actuator works correctly as expected with a

capacitor voltage between 50V to 100V. The contact rebound problem has been

overcome by increasing the current duration. The duration of the current pulse for

opening the vacuum interrupter therefore was set at 15ms.


vacuum chamber with the current duration increased to 15ms


202

8.2.4 Comparison of the tests at atmospheric pressure and in the vacuum

The opening of the vacuum interrupter at atmospheric pressure and in the vacuum with

a capacitor voltage of 100V is compared in Figure 8.8. It presents the actuator coil

voltage, current and linear displacement when the vacuum interrupter is opening.

Figure 8.8 Comparison of the opening of the vacuum interrupter at atmospheric

pressure and in the vacuum


203

The opening operation can be separated into the following periods:

• 0-t1: The voltage and current are the same in both conditions because the

actuator coil has not moved.

• t1-t2: The actuator coil starts to move and accelerates faster in the vacuum than

at atmospheric pressure because of the lack of air resistance. In the vacuum, a

higher velocity induces a higher back-emf and hence a lower current in the

actuator coil. The displacement of the actuator coil in the vacuum is therefore

higher.

• t2-t3: In the vacuum, the contacts start to separate when the displacement is

3mm. The coil current in the vacuum reduces quicker compared to atmospheric

pressure due to the higher velocity and hence higher back-emf.

• t3-t4: At atmospheric pressure, the contacts start to separate. In order to provide

the force to open the vacuum interrupter, the velocity of the actuator coil

decreases and the coil current increases steadily. The coil current in the vacuum

increases faster than at atmospheric pressure because the back-emf is decreasing

faster. The contacts are fully opened in the vacuum when a displacement of

5mm has been achieved and the current increases to the maximum level.

• t4-t5: At atmospheric pressure, the contacts are fully open when the current

increases to its maximum level.

• t5-: The current drops to a steady state level before the MOSFETs are turned off.

At atmospheric pressure, the current in the actuator coil lasts for 10ms. In the

vacuum, the current duration increases to 15ms to keep it in the open position

for a longer period of time and reduce any rebound.

It is clear from the displacement that the opening in the vacuum is faster than at

atmospheric pressure. The opening time in the vacuum is approximately 5ms, compared

to 6ms at atmospheric pressure.

8.3 Single-strand SFCL coil with and without the vacuum interrupter

After the operating actuator demonstrated successful opening operation, the vacuum

interrupter was placed into the SFCL system. A resistor was connected in parallel but


204

external to the cryostat, to act as the fault limiting resistance once the vacuum

interrupter opened. The purpose of the test was to prove whether the operation of the

vacuum interrupter would prevent the SFCL coil heating up and improving recovery

times.

8.3.1 Test rig diagram

The high-current test circuit was modified to carry out the test with the integrated

vacuum interrupter (Figure 8.9). The single-strand SFCL coil was connected in series

with the vacuum interrupter still mounted in the separate vacuum chamber. The bypass

resistor was placed in parallel with them. The load resistor was 0.45Ω and the bypass

resistor was 0.3Ω. A switch was placed in parallel with the load resistor to simulate a

fault by manually closing the switch and short-circuiting the load resistor.

Figure 8.9 Schematic of the high-current test circuit with the vacuum interrupter

8.3.2 Quench test

A quench test was carried out to compare the differences between the SFCL system

with and without the vacuum interrupter. A quench test without the vacuum interrupter

was carried out initially. The actuator control circuit was disabled by discharging the

storage capacitor and turning off the power supply. The amplitude of voltage across the

coil was approximately 0.2V with normal current passing through it. The pre-set


205

threshold voltage level therefore was set at 2V during the test. The quench test was

repeated with the vacuum interrupter and the capacitor charged to 100V.

The fault current, coil current, coil voltage and vacuum interrupter voltage during both

quench tests are shown in Figure 8.10. When the vacuum interrupter does not operate, it

is clear that the SFCL coil quenches and limits the fault current to 240Apeak in the first

quarter-cycle and then reduces the current to 160Apeak over the next two cycles. The coil

current is lower than the fault current because after the coil quenches, the bypass

resistor shares some of the fault current.

The voltage across the coil increases after the coil quenches. When the coil voltage is

greater than the pre-set threshold voltage level of 2V, the actuator control circuit

triggers the actuator and the actuator coil starts to move. The vacuum interrupter

contacts start to separate 2.5ms later. After the contacts separate, an arc is drawn

between the contacts inside the vacuum interrupter. The coil voltage suddenly reduces

due to this arc voltage. The coil current and voltage continue to reduce and the arc

extinguishes naturally at the next coil current zero-crossing. The current through the coil

then becomes practically insignificant and the fault current is fully diverted into the

bypass resistor.

The voltage across the vacuum interrupter is then determined by the voltage across the

resistor. The measured arc voltage of the vacuum interrupter was about 12V. This

appeared reasonable because the arc voltage of the vacuum interrupter is made up of

metal vapour and electrons coming from the electrodes. The arc voltage depends on the

contact material and is usually below 80V [94].

Figure 8.11 presents the temperature rise of the coil in a quench test with and without

the vacuum interrupter. When the vacuum interrupter is disabled, 32.3J energy is

dissipated in the coil for two cycles. The temperature rise of the SFCL coil is 13K and

takes approximately 55s to recover to the operating temperature of 34K. When the

vacuum interrupter operates, the energy dissipated in the coil is reduced to 2.7J, which

is approximately 8% of the value without the vacuum interrupter. There is no

discernible temperature rise in the coil. This demonstrates that the energy dissipated in


206

the coil is significantly reduced by opening the vacuum interrupter as soon as possible

and as a result, the temperature rise of the coil is practically eliminated.

The SFCL coil has demonstrated successful quench behaviour, and shown that MgB2

wire is a potential candidate material for future SFCL development. With the integrated

vacuum interrupter, the fault current was fully diverted into the bypass resistor at the

next current zero-crossing after the fault was imposed. This significantly reduced the

energy dissipated, temperature rise and recovery time of the coil.

Figure 8.10 Quench test with a potential peak current of 324A with and without the

vacuum interrupter


207

30

35

40

45

50

0 20 40 60 80

Time (s)

Tem

per

atu

re (

K)

SFCLwithout VI

SFCL withVI

Figure 8.11 Temperature rise of the coil during a quench test with a potential peak

current of 324A with and without the vacuum interrupter

8.3.3 Simulated fault test

After operation of the SFCL with the vacuum interrupter was demonstrated successfully,

a simulated fault test was carried out. The number of AC cycles supplied was set to be

20 cycles in the LabVIEW control program. Before the test, the switch in parallel with

the load resistor was opened so that normal current flowed in the circuit and the SFCL

coil showed no measurable resistance. The switch was then manually closed to short-

circuit the load resistor and simulate a fault.

The operation of the SFCL coil without the vacuum interrupter was tested initially.

Figure 8.12 shows the fault current, coil current, coil voltage and vacuum interrupter

voltage from the simulated fault test. Normal operating current was 50Apeak and the

switch was manually turned on at approximately 0.25s. The current increases rapidly

and makes the coil quench. In this test it is clear that the SFCL coil quenches and limits

the current to 225Apeak in the first quarter-cycle, then further reduces the current to

75Apeak over the next seven cycles. The voltage across the coil increases after quenching

and 81.7J energy was dissipated in the coil.


208

Figure 8.12 Simulated fault test without the vacuum interrupter

Figure 8.13 shows the result from the simulated fault test with the vacuum interrupter

connected. Normal operating current was 50Apeak and the switch was manually closed at

approximately 0.18s to simulate the fault. The coil quenches due to the rising fault

current and the coil voltage increases over the pre-set threshold voltage level of 2V,

which triggers the actuator to open the vacuum interrupter. The vacuum interrupter

starts to open approximately 2.5ms later and an arc is drawn between the contacts. The

arc extinguishes naturally at the next coil current zero-crossing. The bypass resistor then

works as a limiting resistor because the fault current is fully diverted. The energy

dissipated in the coil is reduced to 4.5J, which is approximately 5.5% of the value

without the vacuum interrupter.


209

Figure 8.13 Simulated fault test with the vacuum interrupter


210

8.4 Conclusions

The actuator static force profile was measured using a load cell and was close to that

predicted by the FE model. This demonstrated that the FE model could provide accurate

prediction for the actuator force.

The vacuum interrupter actuator and its control circuit demonstrated successful

operation at atmospheric pressure. The ultimate aim was to integrate the actuator with

the interrupter into the vacuum vessel inside the cryostat. For practical reasons they

were assembled into a separate vacuum chamber for testing. The opening velocity of the

actuator in the vacuum was higher than at atmospheric pressure due to the lack of air

resistance. One operational issue was found during testing in the vacuum: the actuator

rebounded into the closed position due to the high instantaneous velocity. This problem

was solved by increasing the current duration to 15ms, which therefore held the actuator

in the open position longer.

The quench test and simulated fault test were carried out to compare operation with and

without the vacuum interrupter. The SFCL coil limited the fault current to a safe level as

shown in previous tests. The fast-acting actuator opened the vacuum interrupter before

the next current zero-crossing. The fault current was then diverted into the bypass

resistor and the current through the SFCL coil became practically zero. The energy

dissipated in the coil was significantly reduced with the vacuum interrupter. The SFCL

coil with the integrated vacuum interrupter demonstrated successful operation. This

system represents one method of reducing the temperature rise in the resistive SFCL

and hence improves the recovery times.

Chapter 9: Conclusions and Further Work

211

9 Conclusions and Further Research

9.1 Conclusions

This chapter details the contributions from this research project and summarises

conclusions drawn from the work described in the preceding chapters. Further areas of

research are then discussed. All of the aims and objectives outlined in chapter 1 have

been achieved:

• To investigate the behaviour of single-strand and three-strand MgB2 SFCL coils.

• To develop MATLAB and FE models of the single-strand SFCL coil and

compare the simulation results with the experimental test results.

• To design and manufacture a fast-acting operating actuator for a vacuum

interrupter.

• To design and manufacture a vacuum interrupter actuator control circuit.

• To test the vacuum interrupter operating actuator and validate the design process.

• To investigate the behaviour of the SFCL coil with and without an integrated

vacuum interrupter and compare the results.

The concept of the SFCL with an integrated vacuum interrupter is potentially a solution

for a low-cost and reliable SFCL operating in power network applications.

This research project has demonstrated that a resistive SFCL with an integrated vacuum

interrupter using round MgB2 wire shows significant promise for further research and

product development.

The inductances of the single-strand and three-strand SFCL coils were relatively small

but finite due to the interleaved coil design. Additionally, it was confirmed that both

coils were inductance dominated in the superconducting state but the resistance

increased quickly and became dominant after the coil started to quench.


212

Both SFCL coils showed reliable and repeatable current-limiting properties in short

cycle and long duration quench tests. In addition, both SFCL coils showed that the

quench current increased nearly linearly as the temperature was reduced from 34K to

30K. The quench current density of the single-strand coil was about 16.7% higher than

the three-strand coil at 34K with a self-field of 50Hz but in general this was considered

acceptable.

The single-strand coil operated with 200Apeak current for one hour with an observed

temperature rise of less than 0.1K, which indicated that the thermal stability of the

SFCL coil was good and that AC losses could be removed by the cryocooler.

Tests on the three-strand SFCL coil showed that each of the three wire strands shared

the current equally and demonstrated closely identical responses during normal

operation and fault conditions. Multiple strands of MgB2 wire showed potential to be a

practical method to scale-up the current level for SFCL applications.

During a quench test with a potential peak current of 372A for one second, a

temperature of 107.2K was measured on the single-strand SFCL coil, which resulted in

the coil taking more than two minutes to recover to the superconducting state.

A MATLAB model of the single-strand SFCL coil was developed to understand the

MgB2 wire quench process, which could be used to predict the SFCL behaviour in

power system analysis and inform future designs. The SFCL electrical and thermal

response from 34K to 30K was simulated and a good correlation was demonstrated with

the experimental test results.

An FE thermal model was developed to estimate more accurate temperature

distributions for the SFCL coil. The FE model predicted an average temperature of

69.2K to 122.7K in the SFCL coil for a one-second fault and with a gap between the

coil and the former varying between 0mm to 1.5mm. This suggested that the coil

winding should be manufactured carefully to minimise the gap between the coil and

former, in order to minimise the coil temperature rise.


213

The MATLAB and FE thermal models were used to predict the temperature profile of

the single-stand SFCL coil during a three-second fault. A temperature of 212K to 290K

in the coil was predicted, which demonstrated that the SFCL coil could withstand a

three-second fault test.

An analytical model and an FE model were developed to predict the magnetic field in

the airgap and the force on the coil of the voice-coil type actuator. The FE model

suggested that the force along the whole displacement was nearly constant. The actuator

static force measured by a load cell was identical to the force predicted by the FE model.

Both the FE model and measurement showed that the opening force was higher than the

closing force due to the reluctance force (between the actuator coil and the steel frame),

which caused by asymmetry in the structure of the actuator.

The voice-coil type actuator and its control circuit demonstrated successful operation at

atmospheric pressure. They were assembled into a separate vacuum chamber for testing.

The opening velocity in the vacuum was higher than at atmospheric pressure due to the

lack of air resistance.

The SFCL with an integrated vacuum interrupter demonstrated successful operation.

The fast-acting actuator opened the vacuum interrupter before the first half-cycle of the

fault current during the quench and simulated fault tests. The arc in the vacuum

interrupter extinguished at the current zero-crossing and interrupted the current. The

fault current was then diverted into a bypass resistor and the current in the SFCL coil

became negligible. The energy dissipated in the coil and recovery times were

significantly reduced. The SFCL coil with an integrated vacuum interrupter shows good

potential for further research and product development.


214

9.2 Further research

The three-strand SFCL coil demonstrated successful operation with the current shared

equally by each strand. More strands of parallel wires are worthy of being investigated

to determine whether there is a limitation to the number to multiple wire strands that can

be used.

In this research project, the fault detection strategy of an SFCL with an integrated

vacuum interrupter was achieved by comparing the SFCL coil voltage with a pre-set

threshold voltage level. Another method to detect a fault is to monitor the phase shift

between the current through the SFCL coil and the voltage across it. The combination of

these two detection methods would further improve the accuracy and reliability of

detecting a fault condition.

The SFCL with an integrated vacuum interrupter operated automatically when a fault

condition occurred. After the fault was cleared, the vacuum interrupter was closed

manually. A reclosing control scheme integrated into the grid protection systems is a

potential area of further work.

The SFCL with an integrated vacuum interrupter concept has been successfully tested at

low voltage and current levels. A prototype for an 11kV/33kV distribution network

could be designed and manufactured to further validate the concept.

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215

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Appendix A – Materials Data

228


A1 Alumina

Table A.1 Thermal conductivity of alumina

Temperature (K) Thermal conductivity (W/m ·K) Reference 20 23 77 150 194 48 273 35 373 26

[134]

20 24 30 47 35 63 60 150 85 160 95 140 200 50

[133]

84 125.9 126 48.1 249 35.1 100 133 150 77 200 55 250 43.4 273.2 39.7 300 36 350 30.7 400 26.4 500 20.2

[133]


229

Table A.2 Specific heat capacity of alumina

Temperature (K) Specific heat capacity (J/kg·K) Reference 30 2.58 35 4.31 40 6.78 45 10.20 50 14.63 55 20.30 60 27.25 65 35.51 70 44.94 75 55.56 80 67.61 85 80.88 90 95.07 95 110.04 100 125.94 105 142.59 110 160.08 115 178.11 120 196.73 125 215.69 130 235.02 135 254.51 140 274.22 145 293.93 150 313.63 155 333.38 160 353.05 165 372.41 170 391.70 175 410.74 180 429.28 185 448.10 190 466.09 195 483.67 200 501.66 205 518.82 210 535.55 215 551.87 220 568.19 225 583.67

[133]


230

Temperature (K) Specific heat capacity (J/kg·K) Reference 230 599.15 235 614.21 240 628.86 245 643.49 250 657.31 255 670.69 260 684.08 265 697.05 270 709.60 275 722.16 280 733.87 285 745.59 290 756.89 295 768.18 298.16 774.88 300 778.64 305 789.10 310 799.56 315 809.19 320 818.81 325 828.43 330 837.64 335 846.00 340 855.21 345 863.58 350 871.53 360 887.84 370 902.90 380 916.71 390 930.52 400 943.07 410 955.20 420 966.50 430 977.38 440 987.42 450 997.47 460 1006.67 470 1015.46 480 1023.82 490 1032.19 500 1040.14

[133]


231

Table A.3 Density of alumina

Density (kg/m3) Reference 3670 [122]

A2 Copper

Table A.4 Thermal conductivity of copper

Temperature (K) Thermal conductivity (W/m ·K) Reference 30 4300 35 2900 40 2050 45 1530 50 1220 60 850 70 670 80 570 90 514 100 483 150 428 200 413 250 404 273.2 401 300 398 350 394 400 392 500 388

[133]


232

Table A.5 Specific heat capacity of copper

Temperature (K) Specific heat capacity (J/kg·K) Reference 30 26.65 35 41.84 30 26.65 35 41.84 40 58.99 45 77.69 50 96.94 55 116.57 60 135.73 65 153.80 70 171.00 75 187.61 80 202.88 85 216.69 90 230.33 95 241.54 100 251.79 110 270.58 120 286.98 130 301.04 140 312.38 150 322.38 160 331.29 170 338.78 180 345.18 190 351.04 200 356.18 210 360.54 220 364.38 230 367.89 240 371.12 250 374.34 260 377.36 270 380.16 273.15 380.87 280 382.38 300 386.06 373 393.29 473 414.22 573 422.58

[133]


233

Table A.6 Density of copper

Density (kg/m3) Reference 8933 8950

[134]

Table A.7 Resistivity of copper

Temperature (K) Resistivity (Ω·m×10-8) Reference 30 0.00828 35 0.0147 40 0.0239 45 0.0358 50 0.0518 55 0.0727 60 0.0971 70 0.154 80 0.215 90 0.281 100 0.348 125 0.522 150 0.699 175 0.874 200 1.046 225 1.217 250 1.387 273.15 1.543 293 1.678 300 1.725 350 20.63 400 2.402 500 3.09

600 3.79

[135, 165]


234

A3 Magnesium diboride

Table A.8 Thermal conductivity of magnesium diboride

Temperature (K) Thermal conductivity (W/m ·K) Reference 25 9.04 39.8 15 50 17.3 65.7 20 100 22.5 150 22.88 200 23.27 250 24.33 300 26.06

[166, 167]

Table A.9 Specific heat capacity of magnesium diboride

Temperature (K) Specific heat capacity (J/kg·K) Reference 21.12 0.911

23.05 2.00

25.06 3.097

27.19 4.19

29.66 4.83

32.41 6.74

34.92 9.38

37.75 12.20

41.37 14.85

45.43 22.59

54.12 40.99

57.42 50.46

60.99 60.21

65.42 77.32

70.26 100.92

75.6 123.43

81.29 153.39

85.52 170.41

91.29 199.28

97.04 226.35

102.83 253.76

108.58 287.39

114.13 311.64

119.2 343.13

[133]


235

Temperature (K) Specific heat capacity (J/kg·K) Reference 124.4 368.28

129.79 393.42

136.05 439.32

140.35 461.91

154.92 538.48

173.53 625.09

183.37 675.29

194.73 726.34

208.21 782.83

219.27 828.43

236.77 897.89

238.69 899.98

246.14 920.89

248.39 928.85

254.87 948.09

259.61 974.04

266.88 990.35

274.04 1014.20

279.19 1024.66

286.69 1039.72

287.69 1035.54

298.81 1046.84

[133]

26.44 5.28

128.41 400.62

145.03 488.6

160.29 570.69

183.37 675.29

208.21 782.83

248.39 928.85

279.19 1024.66

300.14 1039.31

[133]

273.71 913.79 [133]

238.69 899.98

254.87 948.09

287.69 1035.54

304.22 1041.39

[133]

Table A.10 Density of magnesium diboride

Density (kg/m3) Reference

2600-2630 [168]


236

A4 Monel (Nickel alloy)

Table A.11 Thermal conductivity of monel

Temperature (K) Thermal conductivity (W/m ·K) Reference 25.43 9.08 40.44 12.4 55.54 14.3 70.52 15.2 84.36 16 98.90 16.5 113.26 17.1 127.12 17.5 142.56 18.1 156.84 18.6 171.86 19 198.90 19.8 213.90 20.2 229.03 20.6 245.02 21.1 263.14 21.7 282.81 22.3 295.2 22.6

[133]

25.98 5.65 40.92 9.5 55.32 11.8 70.02 13.3 86.51 14.4 101.15 15.4 115.14 16.1 130.22 17 143.22 17.7 160.39 18.3 175.46 18.8 204.02 19.7 220.29 20.1 239.98 20.8 258.13 21.3 288.14 21.8 361.2 28 392.2 30.5 416.2 30.1 432.7 31.4 465.2 32.2 467.7 31 484.7 33.1 500.2 33.5

[133]


237

Table A.12 Specific heat capacity of monel

Temperature (K) Specific heat capacity (J/kg·K) Reference 73 259.41

123 305.43

173 347.27

223 380.74

273 405.85

373 447.69

473 259.41

[133]

116 297.06 144 322.17 200 364.01 293 418.4 366 447.69 478 476.98 589 489.53

[133]

Table A.13 Density of monel


8840 [134]

A5 Niobium

Table A.14 Thermal conductivity of niobium

Temperature (K) Thermal conductivity (W/m ·K) Reference 20 229

25 187

30 145

35 116

40 97

45 84

50 76

60 66

70 61 80 58

90 56.3

100 55.2

150 53

200 52.6

[133]


238

Temperature (K) Thermal conductivity (W/m ·K) Reference 250 53

273.2 53.3

300 53.7

350 54.4

400 55.2

500 56.7

[133]

Table A.15 Specific heat capacity of niobium

Temperature (K) Specific heat capacity (J/kg·K) Reference 21.6 12.59 24.78 18.07 28.89 26.57 33.44 37.40 38.21 50.46 42.98 66.65 47.77 78.49 52.52 92.13 57.46 106.39 59.01 110.08 62.38 118.53 65.37 125.85 67.55 130.67 69.73 135.65 72.6 140.46 74.2 145.31 79.23 155.23 84.62 164.68 89.99 173.64 95.58 181.84 101.43 189.66 107.34 196.23 112.93 202.67 118 206.89 123.15 210.92 128.57 215.09 133.91 220.08 139.54 223.47 145.2 227.19 150.61 230.12 156.29 232.59 161.87 235.86 167.6 238.28 173.2 239.79 178.87 241.29

[133]


239

Temperature (K) Specific heat capacity (J/kg·K) Reference 184.5 243.43 189.92 246.06 195.2 247.11 198.91 247.73 200.81 249.03 201.16 248.74 204.13 249.74 206.51 250.04 209.84 250.87 211.68 251.83 212.25 251.58 215.09 252.76 216.5 253.09 217.49 253.47 220.8 254.81 221.21 254.14 222.91 254.76 225.98 255.64 228.03 255.98 231.91 256.69 233.76 257.48 237.37 257.90 238.98 258.15 243.73 258.32 244.35 259.32 248.53 259.74 249.2 259.99 253.6 260.87 254.94 260.79 259.03 261.22 260.72 261.21 265.3 262.09 265.8 262.21 270.51 261.5 270.64 263.01 273.15 269.03 373 272.38 473 275.89 573 279.62

[133]

Table A.16 Density of niobium

Density (kg/m3) Reference 8570 [169]


240

A6 Nitrogen

Table A.17 Thermal conductivity of nitrogen

Temperature (K) Thermal conductivity (W/m ·K) Reference 20 0.4 25 0.32 30 0.27 35 0.23 40 0.2 45 0.18 50 0.16 60 0.169 65 0.16 70 0.151 75 0.1413 80 0.1322 85 0.1231 90 0.1142 95 0.1053 100 0.0966 125 0.052 150 0.01385 160 0.01474 170 0.01562 180 0.01651 190 0.01739 200 0.01826 220 0.01989 240 0.02145 260 0.02298 280 0.02449 300 0.02598 400 0.03252 500 0.03864

[133]

Table A.18 Specific heat capacity of nitrogen

Temperature (K) Specific heat capacity (J/kg·K) Reference 25 970.73 30 1233.57 50 1481.48

[134]


241

Temperature (K) Specific heat capacity (J/kg·K) Reference 70 2001.20 100 14874.58 150 1039.42 200 1039.42 250 1039.42 298.15 1039.42 400 1043.91 500 1055.85

[134]

Table A.19 Density of nitrogen

Temperature (K) Density (kg/m3) Reference

0.15 1140 [134]

77.4 808.5 294.23 1.16

[170]

A7 Polystyrene

Table A.20 Thermal conductivity of polystyrene

Thermal conductivity (W/m ·K) Reference 0.057 0.045 0.037 0.034 0.033

[171]

Table A.21 Specific heat capacity of polystyrene

Temperature (K) Specific heat capacity (J/kg·K) Reference 77 380 173 710 273 1110 473 2150

[136]

Table A.22 Density of polystyrene


10-33 [171]

Appendix B – Components for Operating Actuator

242


Table B.1 Summary of components for prototype operating actuator

Item Material Number Size (mm)

Length 198

Width 30 Steel block for wall EN1A steel 4

Height 120

Length 228

Width 228 Steel block for bottom EN1A steel 1

Height 30

Radius 63 Steel cylinder EN1A steel 1

Height 88

Length 50

Width 14 Steel block for latch EN1A steel 4

Height 14

Length 50

Width 20 Steel block for latch EN1A steel 2

Height 6

Length 80

Width 80 Steel block for stopper EN1A steel 4

Height 6

Length 20

Width 20 Steel block for stopper EN1A steel 4

Height 6

Length 50

Thickness 15.3-20 Permanent magnet N48 Nd-Fe-B

magnet 4

Height 80

Length 50 Permanent magnet for

latch

N48 Nd-Fe-B

magnet

2

Width 10


243

Height 6

Length 200

Width 200 Moving plate Carbon fibre 1

Height 6

Inner diameter 63.5

Outer diameter 65.5 Thin tube for actuator

coil Fibreglass 1

Height 98

Turn 79

Radius 66 Actuator coil AWG19 copper

wire 1

Height 75

Natural length 35 Wipe spring Stainless steel 1

Diameter 12

Natural length 30 Supporting spring Stainless steel 1

Diameter 12

Radius 63 Non-magnetic cylinder Aluminium 1

Height 11

Inner diameter 18

Outer diameter 25 Snatch bracket Aluminium 1

Height 30

Diameter 8 Fixing pole Aluminium 4

Length 110

Diameter 8 Shaft Brass 2

Length 30

Length 180

Width 180 Supporting plate Paxolin 1

Height 10

Length 20

Width 20 Stopper damper Rubber 4

Height 2

Diameter 63 Stopper damper Rubber 1

Height 2

superconducting fault current limiter with integrat …

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