IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 28, NO. 1, JANUARY 2013 253

A 15-kVA High-Temperature SuperconductingPartial-Core Transformer—Part II: Construction

Details and Experimental TestingAndrew Lapthorn, Member, IEEE, Pat Bodger, and Wade Enright

Abstract—A new 15-kVA, 230-230 V, high-temperature super-conducting, partial-core transformer has been designed, built, andtested. The transformer utilizes a unique core design called par-tial core, consisting of a central laminated slug of core steel only.The windings are layer wound with first-generation Bi2223 HTS.In part 1 of this paper, amodel is used to predict the performance ofthe transformer as well as the ac losses of theHTS. In this part, a se-ries of electrical tests was performed on the transformer, includingopen circuit, short circuit, resistive load, overload, ac withstandvoltage, and fault ridethrough tests. The test results are comparedwith the model. The transformer was found to be 98.2% efficientat rated power with 2.86% voltage regulation.

Index Terms—High-temperature superconductors, partial core,transformers.

I. INTRODUCTION

A S STATED in part 1 of this paper, the past decade has seenthe development of a number of service-ready prototype

high-temperature superconducting (HTS) transformers [1], ofwhich the majority have been traditional full-core transformerswith first-generation (1G) HTS wire [2]–[4]. An alternative ap-proach was taken at the University of Canterbury (UoC) in theform of a high-temperature superconducting partial-core trans-former (HTSPCT).A prototype HTSPCT was developed at UoC [5], [6]. The

prototype failed under full-load conditions [7]. A new windingdesign has since been developed incorporating lessons learnedfrom the first failure and was successfully tested. Part 1 of thispaper presented the modelling used in the design of the trans-former. This, the second part of this paper, presents the construc-tion details of the new prototype HTSPCT. Experimental resultsfrom a series of tests are also presented with comparisons madeto the modelling data from the Part 1 paper.

II. TRANSFORMER DESIGN

The design of the HTSPCT involved computer modeling andempirical experimentation using a program developed from tra-ditional transformer design theory and a reverse-as-built design

Manuscript received January 25, 2012; revised May 17, 2012; accepted Oc-tober 18, 2012. Date of publication December 11, 2012; date of current versionDecember 19, 2012. Paper no. TPWRD-00097-2012.The authors are with the Electrical and Computer Engineering Department,

University of Canterbury, Christchurch 8140, New Zealand.Color versions of one or more of the figures in this paper are available online

at http://ieeexplore.ieee.org.Digital Object Identifier 10.1109/TPWRD.2012.2226479

TABLE IDESIGN SPECIFICATIONS OF THE HTSPCT

approach [8], [9]. The computer program uses a combination ofdesign theory and finite-element analysis to determine the com-ponents of the Steinmetz ‘exact’ transformer equivalent circuit.The design specifications of the transformer are given in Table I.The design and construction of the new HTSPCT was based

upon the original HTSPCT which failed under high load [7].Despite the transformer failure, there were a number of aspectsof the original design that were successful, for example, thevacuum Dewar, the main assembly, the copper lead-out design,and the electrical performance prior to failure. As such, a lot ofthese aspects of the design can be found in [7].The core was designed as a parallel-stacked circular core, 347

laminations of 0.23 mm, high-permeability, grain-orientated sil-icon steel (ABB product code 23JGSD085). The core was boundwith Vidatape S, a woven high-shrink polyester tape, and hotdipped twice in an electrical baking varnish. A hole in the centerof the core enables a 1240-mm-long G10 fibreglass 5/8 UNCthreaded rod to enable correct positioning of the core inside thewarm bore tank wall relative to the HTS windings. The core has

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254 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 28, NO. 1, JANUARY 2013

Fig. 1. Photograph of the new interlayer insulation design.

been cut into eight sections for experiments into the heat distri-bution of the core during operation.The windings were wound with 1G wire from Trithor GmbH

(Now Zenergy Power GmbH). This tape is a Bismuth-based su-perconductor (Bi2223), where many su-perconducting filaments are encased in a silver alloy matrixusing a powder-in-tube process. The manufacturer’s wire spec-ification of the critical current at 77 K is 136.7 A in a self field.However, the alternating magnetic fields present in the trans-former result in a reduction of critical current [10], [11]. Forthis reason, the “rated” current for the windings was set to 65Arms. The HTS wire was insulated from turn-to-turn short cir-cuits by the HTS manufacturer with a spiral wrap of an elec-trical-grade-stretched polymer foil [12].

A. Insulation Design

One of the most significant contributing factors of the orig-inal HTSPCT’s failure was the interlayer insulation design. Thewindings were bound in several layers of NOMEX insulationpaper. This prevented heat generated from the HTS conductorfrom being cooled sufficiently by the .The new HTSPCT presented in this paper uses an alternative

winding insulation design consisting of sheets of G10 fibreglass,2 mm thick. The sheets had 3-mm-wide vertical channels ma-chined into the material every 3 mm, as shown in Fig. 1. Thedepth of the cuts was 1.5 mm, leaving 0.5 mm of the fiber sheet.In addition to this, 85-mm-long slots were cut into the verticalchannels with slots on adjacent channels offset. The result ofthis machining left the fiberglass very flexible in the horizontaldirection while still retaining strength, freely allowing the flowof around the HTS wire, and any nitrogen gas bubbles thatform to escape.

B. Transformer Construction

The windings were layer wound onto a fiberglass former thatwas 115 mm in diameter and 402 mm long. The former wasmade to slide over the warm bore tank of the transformer mainassembly and was held in place with a clamping ring also madeof fiberglass. The purpose of this arrangement was to allow forthe removal of the windings so that other winding arrangementscould be tested in the future, including a proposed fault currentlimiter (FCL), and a 2G YBCO winding. The first layer was

Fig. 2. Photographs of the winding process for the HTSPCT. (a) The firstlayer being wound directly onto the former. (b) The second layer being wound,showing the interlayer insulation. (c) The force gauge used to measure the wiretension. (d) Photograph of the finished winding.

wound directly onto the fiber former as shown in Fig. 2(a), fol-lowed by a layer of the interlayer insulation material which isheld in place by the next layer of HTSwire as shown in Fig. 2(b).The wire tension was controlled using a spring system and mea-sured to ensure the wire tension remained below 15 N as shownin Fig. 2(c), as too much strain can cause damage to the super-conductor. The final layer was wrapped in fiber-reinforced tapeto hold the windings in place as shown in Fig. 2(d).Copper lead-outs were used to connect the HTS windings

to the transformer bushings as shown in Fig. 3. The lead-outsconsisted of two 1.5 5-mm copper conductors and were po-sitioned radially around the transformer. The lead-outs wereconnected to the bushing busbars using a compression joint asshown in Fig. 3(a), which was located above the level.The bends in the lower section of the lead outs photographed inFig. 3(b) were present to allow for thermal expansion and con-traction of the copper without pulling or pushing on the HTSwindings. The copper lead outs were soldered to the HTS wireusing Indium-based solder, (97% 3% ), and Ersin RedJelly flux paste. The joint was made by sandwiching approxi-mately 10 cm of HTSwire between two specially formed copperlead-outs soldered with a temperature-limited soldering iron asshown in Fig. 3(c). Too high of a temperature during this processcan result in the HTS wire delaminating; the soldering tempera-ture was limited to 160 C using a variac to control the solderingiron voltage.

LAPTHORN et al.: A 15-kVA HIGH-TEMPERATURE SUPERCONDUCTING PARTIAL-CORE TRANSFORMER—PART II 255

Fig. 3. Details of the copper lead outs connecting the transformer bushings tothe HTS windings. (a) The compression connection used to connect the copperlead outs to the bushing busbars. (b) Bends in the copper lead outs to allow forthermal expansion and contraction. (c) One of the copper lead outs about to besoldered to the HTS wire.

C. Cooling System

In order for the HTS wire to become superconducting, itneeded to be cooled below the critical temperature. This wasachieved by using , which has a boiling point of 77 K.The use of the fiberglass vacuum Dewar for containing the

restricted the outside temperature from evaporating the. However, boil off of the still occurred even when

the transformer was not operating and increased with increasedtransformer load. To prevent the transformer running dry, asystem was developed to monitor the level inside theDewar and to top it up from a reservoir when necessary.The system consists of a float mechanism inside the Dewar to

gauge the level. The float mechanism extends through thetop of the transformer main assembly inside a sight glass. Twooptical sensors attached to the sight glass are used to measurethe level of the float. The optical sensors connect to a cryogenicrated solenoid valve via a controller circuit. The controller opensthe solenoid valve, filling the chamber until the uppermostoptical sensor is triggered by the float, the controller then closesthe solenoid valve. When the boil off of lowers the floatbelow the lower optical sensor, the controller opens the solenoidvalve again to resume filling. The failure mode of the solenoidvalve is the open state so in the case of a controller failure, theHTS windings will not run dry.This system was used during the testing of the transformer

and worked well.

D. Quench Detection System

A quench detection system was used to protect the trans-former from being damaged if the HTS conductor was to quenchduring operation. Quenching can occur in the HTS wire oper-ating at high current density. Localized heating can increasethe temperature in a section of the wire by a few degrees andmay result in that section of wire changing to the normal resis-tive state. In the normal state, ohmic losses cause more local-ized heating and result in an unstable escalating process. Thisquenching process can be seen as an increase in resistance ofthe HTS wire [13]–[16].The increase in the winding resistance during quenching will

result in an increase in the series impedance of the transformer.As noted in [7], the quenching process results in a rapid col-lapse of the secondary voltage of the transformer. This can bemeasured as an increase in the percentage voltage regulation.The quench detection system monitors the primary and sec-ondary voltages and calculates the voltage regulation. When thevoltage regulation reaches a threshold level, the quench detec-tion system opens a circuit breaker (CB) and de-energizes thetransformer. The typical response time from quench detectionto breaker operation was measured at 200 ms.This system was implemented during the testing of the trans-

former and worked as expected.

III. EXPERIMENTAL

A series of electrical tests was performed on the transformerto ascertain its operating characteristics. These tests includedinsulation resistance, dc resistance of the windings, open circuit,short circuit, and resistive load testing. During this testing, thewindings were submerged in .

A. Filling Procedure

Before the commencement of the electrical testing, thevacuum Dewar was filled with . Because of the combus-tion noted in [7], it was suspected that there could be oxygenpresent in the . To prevent the ingress of oxygen into the

chamber, the filling procedure was performed as follows.• Close the two nitrogen-venting valves.• Apply a vacuum to the chamber through the purgingline.

• Hold the vacuum for 24 h.• Fill the chamber through the dry nitrogen gas line.• When positive pressure is obtained, open the normal ni-trogen-venting valve.

• Fill with through the filling tube.

B. Test Results

The first test performed was an insulation-resistance (IR)test. This was performed with a S1-5005 insulation tester fromMegger. The insulation was tested between the inside andoutside windings from the bushings both prior to and afterfilling the chamber. Since the main transformer assemblywas made from G10 fiber and is an insulator, IR tests were notperformed between the windings and the tank. The test voltagewas 500 V for 60 s. The results from the IR testing were 128G for the dry test and 9.1 G for the wet test. This was a pass

256 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 28, NO. 1, JANUARY 2013

TABLE IIELECTRICAL TEST MEASUREMENTS

in both tests although the significant reduction in resistance inthe wet test was noted. It is not understood why this occurred.The second test performed was a dc resistance test for each

winding. This test was performed both prior to and after fillingthe chamber with a MPK254 digital micro-ohmmeterfrom Megabras. Measurements were taken from between thewinding’s bushings. The results for the dry test were 4.133and 4.741 for the inside and outside windings, respectively.The dc resistance under was measured as 2.3 m for theinside winding and 2.5 m for the outside winding. This resultindicates that the HTS wire was superconducting, and thatthe combined resistance of the copper leadouts, busbars, andbushings was small.Open circuit, short circuit, and resistive load tests were the

next tests to be performed. Measurements of the inside and out-side windings’ voltage and current were taken using calibratedFluke 41B Power Meters. The measurements were taken every10 s throughout the open, short, and resistive load tests. A sum-mary of the test results is given in Table II. The model showsgood agreement with the measured results, with the largest dif-ference being in the calculation of the excitation current.1) Open-Circuit Test Results: For the open circuit test, the in-

coming supply was connected to the inside winding (primary),with the outside winding (secondary), left in open circuit. Thevoltage was slowly increased from 0 V to just over the ratedvoltage while measurements were taken. Fig. 4 is a plot of the

Fig. 4. Open-circuit voltage against excitation current.

Fig. 5. Comparison between the measured and modeled open-circuit powerloss.

open-circuit voltage against the excitation current. This is a typ-ical result for a PCT with a linear increase in excitation currentwith applied voltage. No onset of core saturation is evident inthe results with the% total harmonic distortion (THD)measured2.5%. This was expected as the designed peak flux density was

0.7 T, well below the saturation level of the core steel.Fig. 5 compares the calculated losses with those measured for

the open-circuit test. The model shows good agreement with themeasured data.2) Short-Circuit Test Results: For the short-circuit test, the

outside winding was shorted using 95 mm cable to ensure alow-impedance connection. The supply was connected to theinside winding and the voltage increased while monitoring thesupply current. Fig. 6 is a plot of the measured losses of theshort-circuit test compared to those calculated with the model.The model shows good agreement with the measured data.3) Resistive Load Test Results: For resistive load testing, the

outside winding was connected to an adjustable resistance andthe inside winding to the supply. The test was performed overseveral hours with the load varied from a light load up to the fulltransformer rating. Fig. 7 compares the measured losses from

LAPTHORN et al.: A 15-kVA HIGH-TEMPERATURE SUPERCONDUCTING PARTIAL-CORE TRANSFORMER—PART II 257

Fig. 6. Comparison between the measured and modeled short-circuit powerloss.

Fig. 7. Comparison between the measured and modeled resistive load-testpower loss.

the resistive load testing with those calculated with the model.The model shows good agreement with the measured data.The plot in Fig. 8 shows the relationship between the supply

and the load currents for the HTSPCT under load, and comparesit to the measured power factor at the supply end. At low loads,the magnetizing current dominates. This current is in quadratureto the load current and, as a result, the power factor is poor. Asthe load is increased, the difference between the supply and theload current decreases and the power factor improves.4) Overload Test Results: The transformer was tested for an

extended period of time with a 1.1-p.u. resistive load. The orig-inal purpose of the test was to perform a high-load endurancetest of the transformer. However, the losses at this operatingpoint were such that the test could only be run for 9 h. Coolingof the windings during the test was from a continuous-flow-typecryostat system, where the level was maintained using thecooling system of Section II-C fed from a 160 l Dewar.Approximately half of the was used to fill the HTSPCTDewar for the test; the rest was consumed during the testing ofthe transformer through mechanical and electrical losses. Thetransformer performance was stable throughout the testing and

Fig. 8. Supply current (crosses) and the measured power factor (circles) againstthe load current.

TABLE IIIOVERLOAD TEST MEASUREMENTS

it is believed that with an adequate cooling system, the trans-former could run continuously at this load level. A summary ofthe test results for the overload test is given in Table III.5) Separate Source ACWithstand Voltage Test: This test was

performed in accordance with IEC 60076-3. The secondary sideof the transformer was tied to earth and the primary windingconnected to the test voltage. The test voltage was 3029 V, andthe test duration was 60 s. The test voltage was supplied from a230-V to 11-kV distribution transformer fed from a variable acsource. The test frequency was 50 Hz. Although not necessaryfor this test, the high-voltage (HV) current was not able to bemeasured due to the very low capacitance of the HTSPCT insu-lation. The low-voltage (LV) side current was measured insteadand converted to the HV current via the distribution transformerturns ratio. The LV current was measured at 109 mAwhich con-verted to 2.27 mA on the HV side. This gives an estimated in-sulation capacitance of 2.4 nF.6) Fault Ridethrough: The final test performed on the HT-

SPCT was a fault ridethrough test. In this test, the HTSPCTwas connected to a 1-p.u. resistive load and rated voltage ap-plied. The secondary was then short circuited via a manuallyoperated switch. Fault currents were detected via a relay and acontactor was opened, cutting power to the circuit. Oscilloscopetraces of the primary voltage and current are shown in Fig. 9,and

258 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 28, NO. 1, JANUARY 2013

Fig. 9. Oscilloscope traces of the primary voltage, in volts, current, and in am-peres during the fault ridethrough test.

Fig. 10. Oscilloscope traces of the secondary voltage, in volts, current, and inamperes during the fault ridethrough test.

the secondary voltage and current in Fig. 10. Fault levels wereapproximately 350 A rms for about six cycles. The transformersuffered no noticeable damage as a result of this test.The short-circuit reactance of the transformer was measured

at 0.315 . At 230 V, the current through this impedance wouldbe more than 700 A rms. This suggests that fault current hasbeen limited during the fault ridethrough tests. The source ofthis limitation is probably due to the feeding source impedancealthough some may be due to quenching of the superconductor.It would be expected that if the transformer was connected toa stronger network, the fault currents would be significantlyhigher. However, this was not able to be proven with the testingfacilities available.

IV. DISCUSSION

A. Test Results

The testing of the new HTSPCT provided some useful re-sults. The change in insulation resistance between the dry andwet tests was not insignificant and needs further research to de-termine the cause. The testing over a wide load range for an

TABLE IVCALCULATED LOSSES (IN WATTS)

extended period of time demonstrates the validity of the newinsulation design. This design provides greater contact of theHTS wire with the coolant which was believed to be oneof the contributing factors to the original HTSPCT failure.

B. HTS Losses

Table IV lists the losses calculated from the model for theopen circuit, short circuit, and resistive load test. The calculatedlosses for the core and each winding are given as well as thelosses calculated for each part of the HTS ac loss model acrossthe entire winding.The data in Table IV show that the open-circuit results have

a strong radial component of the magnetic field, as seen in thecalculation, while the current-carrying component of the

losses is small, reflected in the and calculations. Alter-natively, the short-circuit calculations show a strong axial com-ponent as well as larger current-carrying components, seenin the increases of , , and compared to the open-cir-cuit calculations. These results are not unexpected. The resistiveload calculations are a combination of the open- and short-cir-cuit cases with the individual loss components almost a sum ofthe open- and short-circuit calculations. The one notable excep-tion to this is the , which makes up nearly half of the totalcalculated losses. This is due to the combination of both highcurrents and high radial magnetic field.Fig. 11 illustrates the total calculated HTS ac losses for each

turn in the HTSPCT where turn 1 to 80 is top to bottom, andlayers 1 to 8 are from closest to the core outwards. The lossesare greatest at the ends of the windings closest to the core. Thisis where the strongest radial flux is found.It is apparent from this analysis that the radial component of

the magnetic field is a significant contributing factor to the aclosses of the HTSPCT. Also, the quantity of losses developedat temperatures is too high to make this design viable in acommercial environment when the “cooling penalty” (approx-imately 20 times) is taken into account. For these reasons, itwould be advantageous to research ways to reduce the radialcomponent of the field, for example, incorporating the use offlux diverters into the design [17].

C. Model Limitations

There are some limitations to the modeling used in this paper.While aspects of the model were based on measured data, suchas the magnetizing losses and the resistive losses , otherparts of the model were unable to be measured. These includethe parameters and in the dynamic resistance model and

LAPTHORN et al.: A 15-kVA HIGH-TEMPERATURE SUPERCONDUCTING PARTIAL-CORE TRANSFORMER—PART II 259

Fig. 11. Total HTS ac losses calculated for each turn in the HTSPCT.

and in the self field losses model. and were estimatedto be 0.95 and 0.94, respectively, and and were adjustedto give a reasonable fit with the measured data with 1.5and 45 600. Since the dynamic resistance and self-fieldlosses made up a significant portion of the calculated losses, thisis not the preferred method.Another limitation with the model lies with the finite-element

analysis of the magnetic field. The nonlinear effects of the coreand of the screening currents in the HTS itself were not mod-elled. The effect of these aspects on the global flux distributionwas assumed to be negligible but this may not actually be thecase. This could be a reason for the difference in calculated andmeasured excitation current since is the main contributingfactor to this current magnitude.Despite these limitations, the model has been able to provide

a reasonable approximation to the measured data across a widerange of operating conditions. This suggests that although thecalculated losses may not be exactly right for their individualcomponents, the overall sum of the losses is an acceptable ap-proximation.

V. CONCLUSION

This paper has presented a new design for a 15-kVA 230-VHTSPCT. The new design offers improvement over the orig-inal HTSPCT with an updated interlayer insulation design thatenables greater contact between the coolant and the HTSconductor. The transformer was tested over this range of oper-ating conditions and compared with the model. There was goodagreement. A discussion on the source of ac losses in the HTSwindings found that the radial flux was likely to be the maincontributor to the measured losses.Due to the design having a partial core, the excitation cur-

rent is much larger than an equivalent rated full-core design,approximately 30% of the rated load current. Despite this, theelectrical performance of the HTSPCT is comparable to a simi-larly rated full-core equivalent. This is shown by the differencebetween the supply current and the load current and the highpower factor that is shown in Fig. 8 at high loading. This is dueto the excitation current being in quadrature to the load current.

However, at light loading, the high magnetizing current dom-inates, and the performance is poorer. For optimal results, theHTSPCT should be used in applications where the loading isexpected to be consistently high.

ACKNOWLEDGMENT

The authors would like to acknowledge the support of Indus-trial Research Ltd. for the duration of this research. Their con-tributions of materials, scholarships, measurements on the HTSsamples, and knowledge were invaluable to the completion ofthis project.

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[17] M. Al-Mosawi, C. Beduz, Y. Yang, M. Webb, and A. Power, “The ef-fect of flux diverters on AC losses of a 10 kVA high temperature super-conducting demonstrator transformer,” IEEE Trans. Appl. Supercon-duct., vol. 11, no. 1, pp. 2800–2803, Mar. 2001.

Andrew Lapthorn (M’11) received the B.E. (Hons.)degree in electrical engineering from the Universityof Canterbury, Christchurch, New Zealand, in 2007,where he is currently pursuing the Ph.D. degree.His research interests are applications of high-tem-

perature superconductors in partial-core trans-formers.

Pat Bodger received the B.E. (Hons.) and Ph.D. de-grees in electrical engineering from the University ofCanterbury, Christchurch, New Zealand, in 1972 and1977, respectively.Currently, he is the Chair of Electric Power En-

gineering, Department of Electrical and ComputerEngineering, University of Canterbury. He is also aDirector of the Electric Power Engineering Centre.From 1977 to 1981, he worked for the ElectricityDivision, Ministry of Energy, New Zealand.His research interests are electroheater and trans-

former design, high voltage, energy system modeling, and renewable energysystem design.

Wade Enright received the B.E. (Hons.) and Ph.D.degrees in electrical and electronic engineering fromthe University of Canterbury, Christchurch, NewZealand, in 1992 and 1995, respectively.Currently, he is a Senior Lecturer at the University

of Canterbury. He also offers electrical engineeringservices to the industry via his own company “Viva.”He specializes in power transformers and highvoltage. During 1996, he worked for the ManitobaHVdc Research Centre, Winnipeg, MB, Canada.