Superconducting AC Homopolar Machines for High-Speed Applications

energies

Concept Paper Superconducting AC Homopolar Machines for High-Speed Applications

Swarn Kalsi 1, Kent Hamilton 2 , Robert George Buckley 2 and Rodney Alan Badcock 2,*

1 Kalsi Green Power Systems; Princeton, NJ 08540, USA; [email protected] 2 Robinson Research Institute, Victoria University of Wellington, PO Box 33436, Lower Hutt 5046, New Zealand; [email protected] (K.H.); [email protected] (R.G.B.) * Correspondence: [email protected]; Tel.: +64-4463-3714

Received: 5 October 2018; Accepted: 25 December 2018; Published: 28 December 2018

Abstract: This paper presents a novel high-speed alternating current (AC) homopolar motor/generator design using stationary ReBCO excitation windings. Compact, lightweight, high-efficiency motors and generators are sought for a multitude of applications. AC homopolar synchronous machines are an ideal choice for such applications as these machines enable very high rotational frequencies. These machines include both AC armature winding and direct current (DC) excitation winding within the stationary part of the machine. The stationary excitation winding magnetizes a solid steel rotor, enabling operating speeds limited only by the mechanical stress limit of the rotor steel. The operating speeds are many multiples of conventional power 50/60 Hz machines. Significant cooling requirements limit machines of this type utilizing copper excitation windings to only a few kilowatts. However, megawatt ratings become possible when superconductor coils are used. This paper describes the design and analysis of an AC homopolar machine in the context of developing a 500 kW flywheel system to be used for energy recovery and storage in commuter rail subway systems. Different approaches are discussed for an AC armature employing conventional copper coils. Challenges of building and cooling both armature and field coils are discussed and preferred approaches are suggested. Calculations of the machine performance are then made.

Keywords: superconducting machines; superconductors; high-speed machines; AC homopolar machines; energy storage; ﬂywheel energy storage

1. Introduction The power of an electrical machine is proportional to both its rotor excitation field and rotational speed. Thus in order to achieve high-power density machines, both magnetic field and rotational speeds must be as large as possible. Speeds above 15,000 rpm are not possible with active coils attached to the rotor due to excessive centrifugal stresses, challenges of handling power, and cooling system interfaces. Additionally, copper windings cannot feasibly produce magnetic fields above ~1.5 T, so higher fields require superconducting coils [1]. A recent review [2] shows demonstrations of superconducting machines up to 10,000 RPM, with high-temperature superconductor (HTS) rotors demonstrated to 15,000 RPM. Machines employing permanent magnets (PM) on the rotor have been built for high speed applications. However, in such machines, PM are held on the rotor with a structural support. The permissible rotor speed is limited by the constraints of the structural support. For example, [3] describes a PM machine rated 45 kW at 32,000 RPM but its diameter is limited to 92 mm. However, the AC homopolar machine described in this paper is rated 500 kW at 25,000 RPM and a rotor diameter of 200 mm. Our rotor made of a single material—the rotor speed is only limited

Energies 2019, 12, 86; doi:10.3390/en12010086 www.mdpi.com/journal/energies Energies 2019, 12, 86 2 of 14 by centrifugal stresses in the rotor material. It is possible to achieve much larger machine with a rotor made of a single solid material. The recent emergence of commercially produced HTS wire has radically changed both the cost-metric and technical viability of superconducting machines [4]. HTS wires are superconducting at temperatures up to 93 K, enabling mechanical refrigerators to be used in conjunction with gas-exchange cooling. This eliminates many of the problems associated with rotating liquid cryogens, whilst the elevated operating temperature provides ‘thermal head-room’, further enhancing system stability. As a result, novel HTS generators and motors are currently a subject of hot global interest [5,6]. The key objectives are achieving compact and lightweight devices of the highest possible torque/weight and electrical efficiencies [7]. One typical application powers electric aircraft propulsion fans with energy generated from on-board generators directly coupled to high-speed gas turbines [8]. A second application is machines which act as both motor and generator integrated with high-speed flywheels for energy storage [9]. This system could be employed in transit systems as a method of fast energy recovery and delivery. Either on moving assets such as buses, or at stationary transport nodes such as subway stations [10,11]. Alternately, the flywheel energy storage device could act as a battery for storing surplus solar and wind energy for later use during lean periods of energy generation [12,13]. HTS exciter coils enable realizing compact and light-weight machines. An alternating current (AC) homopolar synchronous machine topology [14,15] enables high rotational speeds achieved by removing the requirement for active rotating coils, making it ideal for high power applications. Building such a machine using conventional copper excitation coils presents significant cooling challenges and limits the size of such a machine to a power rating of only a few kilowatts. However, megawatt rated machines become possible by replacing the direct current (DC) field excitation coil with a suitable superconducting coil. This paper describes design and analysis for machines employing a HTS DC excitation coil. Different approaches are discussed for building AC armature coils using conventional copper coil technologies. Challenges of building and cooling both armature and field coils are discussed, and possible approaches are suggested for handing them. The application example used is a stationary fast energy storage system for use in subway systems, however, much of what is discussed here is applicable for airborne applications such as those studied by Sivasubramaniam [16]. The machine described in this reference was designed for aerospace application with intention of keeping size and mass as small as possible, whereas the current machine is being designed for on ground application to mate with a flywheel for energy storage. An electromagnetic shield (Figure 3) has been incorporated for reducing AC losses in the superconducting coil and minimizing refrigerator cooling power.

2. Application Requirements In order to describe the design process, a ﬂywheel energy storage application is selected for storing 9 MJ of kinetic energy. This electric motor/generator machine coupled to a ﬂywheel is rated at 500 kW. Its requirements are summarized in Table1.

Table 1. Requirements for a flywheel energy storage device employing AC homopolar motor/generator.

Parameter Value Energy storage capability, MJ 9 Rotating machine power rating, kVA 500 Rotary speed, RPM 25,000 Line voltage, V 866 Rotor diameter, mm <200 Axial length of stator, mm <500 Axial length of rotor, mm <420 Overall stator coil current density, A/mm2 <3.0 Maximum ﬂux density in stator laminations, T <1.5 Energies 2018, 11, x FOR PEER REVIEW 3 of 15 Energies 2019, 12, 86 3 of 14 • The stator laminations are 0.1 mm thick JFE Steel Corporation, Chiba, Japan JNEX-Core (model 10JNEX900)This machine intended is designed to reduce using iron the following losses. assumptions:

• TableThe rotating 1. Requirements machine is synchronousfor a flywheel AC energy homopolar storage type device with fouremploying poles; AC homopolar • motor/generator.The field winding is located on the stator and is made from HTS; either rare-earth-barium-copper-oxide (ReBCO), or di-bismuth-strontium-calcium-copper-oxide (DI-BSCCO); Parameter Value • The operating temperature for HTS windings in 50 K for ReBCO and 30 K for DI-BSCCO; Energy storage capability, MJ 9 • The ﬁeld winding is cooledRotating using machine a suitable power cryocooler rating, kVA available off-the-shelf; 500 • The armature on the stator is a three-phaseRotary speed, winding RPM employing a suitable 25,000 Litz copper wire cable; 2 • The armature coils have current densities;Line voltage, 3 A/mm V (overall cross-section 866 including cooling tube) 2 and 6 A/mm (in copper); Rotor diameter, mm <200 • The armature winding is liquidAxial cooled; length of stator, mm <500 • The rotor is made of high-permeabilityAxial length Aermet of rotor, 310 mm [17 ] magnetic steel; <420 and • The stator laminationsOverall are stator 0.1 mm coil current thick JFE density, Steel A/mm Corporation,2 <3.0 Chiba, Japan JNEX-Core (model 10JNEX900)Maximum intended flux to reduce density iron in stator losses. laminations, T <1.5

3. Machine Machine Configuration Conﬁguration A cutaway diagram of the AC homopolar motor/generatormotor/generator is is shown shown in in Figure Figure 11.. TheThe shaftshaft andand ◦ rotor have not been sectioned so so as to show the fo four-poleur-pole layout, with 90° 90 rotationallyrotationally offset offset poles. poles. The three armature coil colors illustrate the three phases winding scheme. The HTS coil cryostat and its surroundingsurrounding insulationinsulation have have been been omitted omitted for for clarity, clarity, as has as an has electromagnetic an electromagnetic (EM) shield (EM) locatedshield locatedbetween between the armatures the armatures and the and superconducting the superconductin coils.g The coils. ferromagnetic The ferromagnetic poles are poles shaped are shaped as in any as insalient any polesalient machine pole machine so as to reduceso as to harmonic reduce harmonic and torque and ripple torque to acceptable ripple to levels,acceptable as described levels, as in describedthe next section. in the next section.

Figure 1. Sectioned view of the AC homopolar motor/generator. Figure 1. Sectioned view of the AC homopolar motor/generator. In the flywheel energy storage system, the homopolar motor/generator is located between two glassIn fiber the reinforcedflywheel energy polymer storage flywheels system, as shown the ho inmopolar Figure2 motor/generator. This arrangement is located reduces between the required two glassshaft fiber size by reinforced halving thepolymer power flywheels transferred as throughshown in any Figure one 2. part This of arrangement the shaft. This reduces assembly theis required housed shaftin a vacuum size by halving chamber, the with power the transferred rotating components through an supportedy one part onlyof the by shaft. magnetic This assembly levitating is bearings, housed inenabling a vacuum operation chamber, up with to 25,000 the rotating RPM. At components the high rotational supported speed only desired,by magnet housingic levitating the system bearings, in a enablingvacuum isoperation required up to reduceto 25,000 windage RPM. At friction the high that rotational would otherwise speed desired, result in housing high drag. the system Frictionless in a vacuumnon-contact is required magnetic to bearingsreduce windage are also friction required that as mechanicalwould otherwise bearings result would in high not bedrag. able Frictionless to operate non-contactcontinuously magnetic due to frictional bearings losses are also and required wear [18 as]. mechanical bearings would not be able to operate continuously due to frictional losses and wear [18].

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. . FigureFigure 2. 2.Flywheel Flywheel energy energy storage storage system system incorporating incorporating a a superconducting superconducting motor/generator. motor/generator. Figure 2. Flywheel energy storage system incorporating a superconducting motor/generator. TheThe designdesign ofof thethe flywheels,flywheels, vacuumvacuum chamber,chamber, andand bearingsbearings areare outsideoutside thethe scopescope ofof thisthis paper.paper. Discussion of the armature, rotor, and stator follows below. DiscussionThe design of the of armature, the flywheels, rotor, vacuum and stator chamber, follows and below. bearings are outside the scope of this paper. Figure3 shows the cross-section of the machine with main components. The stator core is DiscussionFigure of 3 showsthe armature, the cross-section rotor, and of stator the machine follows wibelow.th main components. The stator core is axially axially split into two equal parts and the superconducting excitation coil is sandwiched between splitFigure into two 3 shows equal theparts cross-section and the superconducting of the machine excitationwith main coil components. is sandwiched The stator between core them. is axially The them. The stator core is constructed using laminated magnetic steel. Sections3 and8 detail the splitstator into core two is equalconstructed parts and using the superconductinglaminated magnetic excitation steel. Sections coil is sandwiched 3 and 8 detail between the laminationthem. The lamination directions and the unconventional magnetic flux path these must support. Armature coil statordirections core andis constructed the unconventional using laminated magnetic magnetic flux path steel. these Sections must support. 3 and 8Armature detail the coil lamination sides are sides are shown laid axially on the stator. High rotational frequency combined necessitates the rotor be directionsshown laid and axially the unconventional on the stator. magneticHigh rotational flux path frequency these must combined support. necessitates Armature coilthe sidesrotor are be machined from a single piece, and the rotor’s role in augmenting the magnetic field supplied by the shownmachined laid from axially a single on the piece, stator. and theHigh rotor’s rotational role in frequency augmenting combined the magnetic necessitates field supplied the rotor by thebe exciter coil requires this piece to be magnetic steel. In Figure3 the rotor is shown with one solid-pole machinedexciter coil from requires a single this piece,piece toand be the magnetic rotor’s steelrole .in In augmenting Figure 3 the the rotor magnetic is shown field with supplied one solid-pole by the (on the right) and one void-pole (on the left) aligned with stator core—the void pole is located at the exciter(on the coil right) requires and one this void-pole piece to be (on magnetic the left) steelaligned. In Figurewith stator 3 the core—the rotor is shown void pole with is one located solid-pole at the mid-point between two adjacent solid-poles on the rotor and represents the part of the rotor volume (onmid-point the right) between and one two void-pole adjacent (on solid-poles the left) alignedon the rotor with and stator represents core—the the void part pole of the is located rotor volume at the where minimum magnetic flux is supported, i.e., where there is no magnetic steel. mid-pointwhere minimum between magnetic two adjacent flux is solid-poles supported, on i. e.,the where rotor thereand represents is no magnetic the part steel. of the rotor volume where minimum magnetic flux is supported, i.e., where there is no magnetic steel.

. Figure 3. Schematic cross-section of stator and rotor layout. A detailed image of the superconducting. coilFigure is shown 3. Schematic in Figure cross-section 10. of stator and rotor layout. A detailed image of the superconducting Figurecoil is shown3. Schematic in Figure cross-section 10. of stator and rotor layout. A detailed image of the superconducting 4. Sizing Analysis coil is shown in Figure 10. 4. Sizing Analysis A 2D approximation of the machine as shown in Figure4, is utilized for sizing analysis with 2D FEA.4. Sizing InA 2D this Analysis approximation approximation of the the rotor machine poles areas shown not offset in Figure by 90 degrees4, is utilized as the for behavior sizing analysis of interest with is the 2D capabilityFEA.A In 2D this ofapproximation theapproximation machine toof effectivelythe machinerotor poles route as shownare ﬂux not within inoffs Figureet the by rotor4,90 is degrees utilized and stator. as for the sizing Two behavior kinds analysis of armatureinterest with 2D is windingsFEA.the capability In this were approximation considered;of the machine (a) the coils torotor effectively in poles air-gap, are route and not (b)offs flux coilset withinby in 90 slots degrees the in rotor the as iron theand core.behavior stator. Two of interest kinds isof thearmature capability windings of the were machine considered; to effectively a) coils routein air-gap, flux andwithin b) coilsthe rotorin slots and in thestator. iron Two core. kinds of armature windings were considered; a) coils in air-gap, and b) coils in slots in the iron core.

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. . Figure 4. Representative cross-sections of the air-gap (a) and iron-slot (b) armature coils showing flux Figure 4. Representative cross-sections of the air-gap (a) and iron-slot (b) armature coils showing flux distribution as calculated using 2D FEA software FEMM [19]. Only (a) shows the armature windings distributionFigure 4. Representative as calculatedusing cross-sections 2D FEA softwareof the air-gap FEMM (a)[ and19]. iron-slot Only (a) shows(b) armature the armature coils showing windings flux as the section taken in (b) passes through the stator tooth (all dimensions in mm). asdistribution the section takenas calculated in (b) passes using through2D FEA thesoftware stator FEMM tooth (all [19]. dimensions Only (a) shows in mm). the armature windings as the section taken in (b) passes through the stator tooth (all dimensions in mm). Figure 44aa showsshows aa cross-sectioncross-section withwith armaturearmature coilscoils residingresiding inin thethe gapgap betweenbetween thethe statorstator andand rotor, i.e.,i.e.Figure there there 4a are areshows no stator a cross-section iron teeth. with Although armature a variety coils residing of armature in the winding gap between configurationsconfigurations the stator are and possible,rotor, i.e. aa uniqueuniquethere are single-layersingle-layer no stator (SL)iron (SL) windingteeth. winding Although design design is ais considered variety considered of forarmature for this this study. windingstudy. The The dimensions configurations dimensions of this ofare thiscross-sectionpossible, cross-section a unique have have been single-layer been selected selected to(SL) suit towinding thesuitspecifications the design specifications is considered of Table of Table1. for 1. this study. The dimensions of thisLikewise, cross-section FigureFigure have4 b4b showsbeen shows selected the the cross-section cross-section to suit the with specifications wi armatureth armaturecoils of Table coils housed 1.housed between between iron statoriron stator teeth. Ateeth. double-layer ALikewise, double-layer (DL) Figure winding (DL) 4b windingshows configuration the configuration cross-section is selected is wiselected here.th armature The here. radial The coils depthradial housed ofdepth the betweenarmature of the armatureiron coils stator is coilssameteeth. is as sameA for double-layer the as air-gapfor the air-gap winding(DL) winding winding coils (Figureconfiguration coils 4(Figurea). is4a). selected here. The radial depth of the armature coilsIn is each each same of of asthese these for simulationsthe simulations air-gap winding the the stator stator coils and and (Figure rotor rotor are 4a). are taken taken to be to unlaminated be unlaminated 1020 1020steel steelwith withnon- linearnon-linear InBH each curve BH of curve thesesaturating saturatingsimulations at approximately at the approximately stator and 2.8 rotor T. 2.8 Each are T. Each takenrepresentative representative to be unlaminated cross-section cross-section 1020 considerssteel considers with non-the shaftthelinear shaft material BH material curve to be tosaturating air be airand and the at the approximatelyrotor rotor has has been been 2.8sized sized T. soEach so as as notrepresentative not to to require require thecross-section the shaft shaft to to be be considers part part of of the the magneticshaft material circuit. to Although be air and the the design rotor of has a function functional been sizedal generator so as not which to require supplies the alternating shaft to be current part of at the themagnetic stator windings circuit. Although requires a the non-symmetric design of a function lobed al rotor, generator the 2D which approximations supplies alternating used for sizing current are at axisymmetricthe stator windings as the total requires magnetic a non-symmetric fluxflux is not affected lobed rotor, byby thethe the asymmetry.asymmetry. 2D approximations used for sizing are axisymmetricThe location as of ofthe thethe total fieldfield magnetic excitation excitation flux coil coil is is notis identical identical affected in inby both both the cases. asymmetry.cases. However, However, 1/2 1/2 the the active active length length of ofiron iron coreThe core is location longeris longer forof for the the the field DL DL design excitation design (85 (85 mmcoil mm is for for identical SL SL versus versus in 130both 130 mm mmcases. forfor However, DL).DL). ArmatureArmature 1/2 the coilscoils active for length both SLof andiron DLcore designs is longer employ for the liquid DL design cooling. (85 mmThe fortwo SL possible versus 130winding mm for configurations configurations DL). Armature are coils shown for both in FigureSL and 55.. TheDLThe designs lightweightlightweight employ SLSL layoutlayoutliquid oncooling.on thethe left,left, The andan twod thethe possible moremore conventionalconventional winding configurations DLDL layoutlayout on onare thethe shown right.right. in TheFigure coil colors5. The signifylightweight the three SL layout phases. on the left, and the more conventional DL layout on the right. The coil colors signify the three phases.

(a) (b)

Figure 5. Armature(a) windings, ( a) single layer (SL) and (b) double(b) layer (DL). Figure 5. Armature windings, (a) single layer (SL) and (b) double layer (DL). Harmonic content in the terminal voltage voltage for for the the two two types types of of windings windings is is shown shown in in Figure Figure 66.. Overall, both SL and DL designs have similar harmonic contents. The 3rd,rd 9th,th and 15thth harmonics Overall,Harmonic both SL contentand DL in designs the terminal have similarvoltage harmonic for the two contents. types of The windings 3 , 9 , isand shown 15 harmonicsin Figure 6. Overall, both SL and DL designs have similar harmonic contents. The 3rd, 9th, and 15th harmonics

Energies 2019, 12, 86 6 of 14 Energies 2018, 11, x FOR PEER REVIEW 6 of 15 cancelcancel out out in in a a three-phase three-phase system. system. The The 5th 5th and and 7th 7thharmonics harmonics are larger are larger for SL for than SL thanDL design, DL design, but thebut 11 theth and 11th 13 andth harmonics 13th harmonics are smaller aresmaller for the SL for than the SLthethan DL design. the DL Nevertheless, design. Nevertheless, harmonic harmonic contents incontents both designs in both appears designs to appears be within to be an within acceptable an acceptable range. However, range. However, during detailed during detailed design phase, design polesphase, will poles be will shaped be shaped further further to reduce to reduce harmonic harmonic contents, contents, just just as asis iscustomary customary for for salient salient pole pole synchronoussynchronous machines. machines.

Figure 6. Harmonic content in the terminal voltage of SL and DL designs.

The 2D representationsFigure 6. Harmonic presented content in Figure in the 4terminal each display voltage a of single SL and magnetic DL designs. circuit enclosing the cross-sectioned superconducting exciter coil. As current in the coil travels into the page the magnetic flux travelsThe 2D around representations the coil in presented a clockwise in direction, Figure 4 i.e.,each vertically display upwarda single throughmagnetic the circuit rotor, enclosing to the left theat the cross-sectioned upper rotor lobe superconducting and into the stator, exciter vertically coil. As downward current in through the coil the travels stator, theninto tothe the page right the to magneticre-enter the flux rotor travels at the around lower rotorthe coil lobe. in Thisa clockwise 2D approximation direction, i.e., was vertically used in sizing upward the requiredthrough ironthe rotor,components, to the left however at the ifupper the rotor rotor lobes lobe wereand into aligned the instator, this wayvertically the stator downward windings through would experiencethe stator, thenzero netto the change rightin to magnetic re-enter fieldthe rotor along at their the lengthlower rotor as the lobe. lobes This passed. 2D approximation was used in sizingTo the ensure required the iron stator components, windings experiencehowever if the an alternatingrotor lobes were magnetic aligned field in thethis rotorway the lobes stator are windingsrotationally would offset experience by 90 degrees, zero net as change illustrated in magn in Figureetic field1 and along below their in length Figure as7. the This lobes twists passed. the magneticTo ensure circuit the so each stator stator windings winding experience experiences an a changingalternating net magnetic fieldfield as the the rotor rotor lobeslobes pass.are rotationallyThe idealized offset magnetic by 90 circuit degrees, flux as path illustrated is illustrated in Figure in red 1 in and Figure below7. This in Figure path travels 7. This azimuthally twists the magneticthrough thecircuit central so each “bridge” stator section winding of theexperiences stator as a supported changing bynet the magnetic concentric field laminations as the rotor oflobes this Energies 2018, 11, x FOR PEER REVIEW 7 of 15 pass.part ofThe the idealized stator (Figure magnetic3). circuit flux path is illustrated in red in Figure 7. This path travels azimuthally through the central “bridge” section of the stator as supported by the concentric laminations of this part of the stator (Figure 3). Omitted from Figure 7 for clarity, a small portion of flux lines also cross through the void-pole to the stator core. Each half of the stator core experiences maximum magnetic field under a solid-pole and minimum under a void-pole. Nevertheless, field under all pole, facing each ½ stator core is unidirectional.

Figure 7. Illustration of the idealized flux pattern through the stator (in red), relative to. the rotor. The flux path is closed through the rotor (light grey) and encircles the field coils (dark grey). Figure 7. Illustration of the idealized flux pattern through the stator (in red), relative to the rotor. The flux path is closed through the rotor (light grey) and encircles the field coils (dark grey).

Coils with straight sides (spanning the whole stator core) are employed for armature winding located in the stator. Figure 4 shows the flux contours when each of the two configurations is salient. In this approximation the rotor poles are aligned with the stator core. This approximation is used for calculating the flux leakage through the interpole regions and stator. At any instant, an armature coil experiences maximum flux linkage (Φmax) in the part of the stator aligned with the rotor pole, and minimum flux linkage (Φmin) in the part of the stator aligned with the void pole. The net flux linkage (Φo) is the difference between these two values. The FEA code is utilized for calculating flux linkages in each half of the core. The net flux linkages are utilized for calculating induced voltage in a coil. Furthermore, location of Φmax is identified as d-axis (solid-pole) and location of Φmin is identified as q-axis (void-pole). The design analysis then assumes that the whole machine is represented by a half- core being acted upon by complete set of d- and q- poles, with Φo being experienced under each pole. On this basis, the induced voltage in an armature coil is as in Equation (1):

Eo = c·Φo·ωo (1) where: Eo = induced voltage/coil (V) c = proportionality constant Φo = net flux linkage with a coil (Wb) ωo = rotational frequency (rad/s) The excitation field coil is circular in shape which makes it simple for construction. It is sized to provide necessary ampere-turns for excitation of the armature coils. The SL design has a larger effective air gap between the rotor and stator as it does not include iron teeth, this necessitates more ampere-turns than what is required in the DL design. Since this coil is sandwiched between two halves of stator iron core, it does not experience magnetic forces due to interaction with the armature winding current. Figure 3 shows cross-section of the machine with locations of armature and field excitation coils. However, excitation coil side facing the armature winding experiences AC field due to currents in the armature coils. It is essential that this AC field be sufficiently attenuated for minimizing AC losses in the excitation coil. This is achieved by creating a separation between armature coils and excitation coils and installing an EM shield (copper or aluminum) to attenuate AC field experienced by the field coil. The location of the EM shield is identified in the Figure 3. In this design analysis, the excitation coil employs ReBCO superconductor. Preliminary design of the machines employing SL and DL armature windings are summarized in Table 2. Both machines are sized to generate 500 kW at a three-phase line voltage of 830–900 V. Total ampere-turns needed for SL design are much larger than DL design due to significantly larger air gap in the SL design—this results in more turns in the field coil for SL design than the DL design.

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Omitted from Figure7 for clarity, a small portion of flux lines also cross through the void-pole to the stator core. Each half of the stator core experiences maximum magnetic field under a solid-pole 1 and minimum under a void-pole. Nevertheless, field under all pole, facing each 2 stator core is unidirectional. Coils with straight sides (spanning the whole stator core) are employed for armature winding located in the stator. Figure4 shows the flux contours when each of the two configurations is salient. In this approximation the rotor poles are aligned with the stator core. This approximation is used for calculating the flux leakage through the interpole regions and stator. At any instant, an armature coil experiences maximum flux linkage (Φmax) in the part of the stator aligned with the rotor pole, and minimum flux linkage (Φmin) in the part of the stator aligned with the void pole. The net flux linkage (Φo) is the difference between these two values. The FEA code is utilized for calculating flux linkages in each half of the core. The net flux linkages are utilized for calculating induced voltage in a coil. Furthermore, location of Φmax is identified as d-axis (solid-pole) and location of Φmin is identified as q-axis (void-pole). The design analysis then assumes that the whole machine is represented by a half-core being acted upon by complete set of d- and q- poles, with Φo being experienced under each pole. On this basis, the induced voltage in an armature coil is as in Equation (1):

Eo = c·Φo·ωo (1) where:

Eo = induced voltage/coil (V) c = proportionality constant

Φo = net flux linkage with a coil (Wb) ωo = rotational frequency (rad/s) The excitation field coil is circular in shape which makes it simple for construction. It is sized to provide necessary ampere-turns for excitation of the armature coils. The SL design has a larger effective air gap between the rotor and stator as it does not include iron teeth, this necessitates more ampere-turns than what is required in the DL design. Since this coil is sandwiched between two halves of stator iron core, it does not experience magnetic forces due to interaction with the armature winding current. Figure3 shows cross-section of the machine with locations of armature and field excitation coils. However, excitation coil side facing the armature winding experiences AC field due to currents in the armature coils. It is essential that this AC field be sufficiently attenuated for minimizing AC losses in the excitation coil. This is achieved by creating a separation between armature coils and excitation coils and installing an EM shield (copper or aluminum) to attenuate AC field experienced by the field coil. The location of the EM shield is identified in the Figure3. In this design analysis, the excitation coil employs ReBCO superconductor. Preliminary design of the machines employing SL and DL armature windings are summarized in Table2. Both machines are sized to generate 500 kW at a three-phase line voltage of 830–900 V. Total ampere-turns needed for SL design are much larger than DL design due to significantly larger air gap in the SL design—this results in more turns in the field coil for SL design than the DL design. The SL machine is shorter in overall length and lighter than the DL machine. However, the DL machine is more efficient than the SL machine and uses much shorter length of ReBCO wire (185 m for SL versus 54 m for DL). This makes DL machine less costly to build than the SL machine. The table also lists preliminary component weights for the two machines; The DL machine weighs 27% more than the SL machine. The DL design is preferable over the SL design if size and weight are not of concern. In the subway energy storage applications, the heavier, more efficient DL design is desirable as the motor is ground based at the subway station, sufficient space is available for the larger machine size, and high efficiency is desirable. For this machine design the superconducting material is the largest single component cost. In the SL design the expected superconductor cost is $39,000, while in the DL Energies 2019, 12, 86 8 of 14 design much less superconductor is required, bringing the cost of the superconductor down to $4400. The superconductor cost for the SL design constitutes a large fraction of total expected cost of the machine, $100,000. The DL design is significantly more cost effective. Only the machine design employing the DL stator with ReBCO field coil is discussed further.

Table 2. Preliminary design summary for SL and DL machines.

Parameter Single Layer Double Layer Power Rating, MVA 524 540 Output power at full-load, MW 519 535 Line voltage, V-rms 900 866 Phase current, A-rms 336 360 Overall axial length, m 0.32 0.46 Overall diameter, m 0.44 0.44 Mass of the machine alone, kg 345 467 Mass of cryo-cooling system, kg 30 49 Total mass, kg 375 516 Efﬁciency at full-load, % 98.9 99.2 Cryocooler load, kW 0.80 1.36 Other Parameters of Interest Rated speed, RPM 25,000 25,000 Number of pole 4 4 Frequency, Hz 833 833 FIELD WINDING DETAILS Number of turns 204 24 Field winding current, A 252 317 Field winding Io/Ic ratio 0.54 0.49 HTS wire width, mm 3 3 HTS wire length, m 185 22 STATOR WINDING DETAILS Active length under each pole, mm 85 130 Number of armature turns/ph 20 8 Number of coils in armature 6 24 Number of turns/race coil 1 1 Machine component weight summary Shaft, kg 4 4 Rotor yoke, kg 42 59 Poles, kg 7 12 Stator case, kg 144 178 Cooling system, kg 30 49 Total machine mass, kg 345 467 Total system mass, kg 375 516 Torque density, N/kg 533 388

5. Armature Design Armature design employs single turn coils with embedded cooling tube. A Litz wire turn with an embedded cooling tube is shown in Figure8. The overall cross-section of each turn is 10 mm × 12 mm and the stainless-steel tube at the center has a diameter of 5 mm. The total cross-section of the Litz cable is equivalent to that of a single AWG 1/0 conductor. Table3 provides more details, including data from the New England Wire catalogue. Energies 2018, 11, x FOR PEER REVIEW 9 of 15

$100,000. The DL design is significantly more cost effective. Only the machine design employing the DL stator with ReBCO field coil is discussed further.

5. Armature Design Armature design employs single turn coils with embedded cooling tube. A Litz wire turn with an embedded cooling tube is shown in Figure 8. The overall cross-section of each turn is 10 mm × 12 mm and the stainless-steel tube at the center has a diameter of 5 mm. The total cross-section of the Litz cable is equivalent to that of a single AWG 1/0 conductor. Table 3 provides more details, Energies 2019, 12, 86 9 of 14 including data from the New England Wire catalogue.

. Figure 8. Coil turn cross-section with an embedded cooling tube (dimensions in mm). Figure 8. Coil turn cross-section with an embedded cooling tube (dimensions in mm). Table 3. Litz wire parameters.

ParameterTable 3. Litz wire parameters. Value Space available for Litz,Parameter mm2 Value100 Equivalent AWG size 1/0 Space available for Litz, mm2 100 Strand diameter, mm 0.32 Number ofEquivalent strands AWG size 665 1/0 Copper cross-section,Strand diameter, mm2 mm 54 0.32 Resistance of cable,Number Ohm/m of strands 0.00035 665 ResistiveCopper loss per coil,cross-section, W mm2 5154 Cooling tube diameter, mm 5 Water ﬂowResistance velocity, m/sof cable, Ohm/m 0.00035 1 Pressure drop overResistive two coils inloss series, per kPacoil, W 6.7 51 Difference in cooling tubeCooling wall and tube water diameter, temperature, mm K 0.34 5 Temperature rise of waterWater through flow 2 coils velocity, in series, m/s K 1.24 1 Pressure drop over two coils in series, kPa 6.7 Only 6.7 kPaDifference pressure drop in cooling is experienced tube wall for and a sustained water temperature, water ﬂow rateK of 0.34 1 m/s in a tube length (2.3 m) equal to twoTemperature coils connected rise of in water series. through The water 2 coils temperature in series, K rise is about1.24 1.2 K. Based on this data, it looks feasible that all coils of a phase could be connected in series to form a single cooling circuitOnly for each6.7 kPa phase. pressure A wave drop shape is windingexperienced arrangement, for a sustained shown water in Figure flow9, israte selected of 1 m/s as itin enables a tube constructionlength (2.3 m) of equal all coils to two of a phasecoils connected using a single in seri lengthes. The of water cable. temperature The total resistive rise is loss about in all1.2 armatureK. Based coilsonEnergies this is 1.22018data, kW, 11 it, whilex looksFOR PEER carrying feasible REVIEW thethat rated all coils load. of a phase could be connected in series to form a 10single of 15 cooling circuit for each phase. A wave shape winding arrangement, shown in Figure 9, is selected as it enables construction of all coils of a phase using a single length of cable. The total resistive loss in all armature coils is 1.2 kW while carrying the rated load.

. Figure 9. Phase-A coils connected in a wave fashion for four pole-pitches. Figure 9. Phase-A coils connected in a wave fashion for four pole-pitches. 6. Superconducting Excitation Coil Design 6. SuperconductingThe superconducting Excitation excitation Coil Design coil employs ReBCO wire cooled to 50 K. The coil is housed in a stainless-steelThe superconducting cryostat and excitation surrounded coil byemploys multi-layer-insulation ReBCO wire cooled (MLI) to foil50 K. as The shown coil inis Figurehoused 10 in. Toa stainless-steel reduce thermal cryostat conduction and surrounded the cryostat by multi-layer-insulation is not mounted to the (MLI) stator foil but as rathershown mountedin Figure 10. to theTo vacuumreduce thermal chamber conduction housing via the insulated cryostat is support not mounted rods that to protrudethe stator through but rather ports mounted in the stator.to the Betweenvacuum thechamber cryostat housing wall and via the insula statorted body support is a copper rods EMthat shieldprotrude [20]. through It is designed ports toin protectthe stator. the superconductorBetween the cryostat from thewall alternating and the stator ﬁeld body due to is current a copper in EM the armatureshield [20]. windings. It is designed It is attached to protect on the to superconductor from the alternating field due to current in the armature windings. It is attached on to the room-temperature bore side of cylindrical cryostat box. Any losses in it are removed using the means employed for cooling the room-temperature surface of the cryostat.

Figure 10. Schematic cross-section of HTS excitation coil cryostat. The HTS coil consists of two 12 turn ReBCO coils which remain static while magnetizing the rotating solid steel rotor.

The field winding, made of ReBCO, has 24 turns. It carries 317 A when the machine is operating at full-load. A different superconductor could be used for generating the total amp-turns equal or greater than 7600 (~24 turns * 317 A). Furthermore, the ratio of operating-current/critical-current (Io/Ic) for the superconductor is preferred to be less than or equal to ~50%. Current leads for the superconducting coil provide a path for thermal conduction from room-temperature to the cold environment of the superconducting coil, and require the exciter circuit to include non- superconducting elements which generate heating through ohmic losses. This thermal load is estimated to be ~25 W. A contactless exciter concept is available [21] that could charge the excitation coil without physical current leads spanning room-temperature and cryogenic environments, while enabling the exciter circuit to be almost entirely superconducting. With this exciter implemented, it would be possible to operate the flywheel continuously with a smaller cooling system as the conduction and ohmic heating loads are greatly reduced.

7. Rotor Design Figure 11 shows a three-dimensional view of the rotor. This configuration creates a four-pole field pattern as viewed from each half of the stator core. The rotor is manufactured from one solid

Energies 2018, 11, x FOR PEER REVIEW 10 of 15

Figure 9. Phase-A coils connected in a wave fashion for four pole-pitches.

6. Superconducting Excitation Coil Design The superconducting excitation coil employs ReBCO wire cooled to 50 K. The coil is housed in a stainless-steel cryostat and surrounded by multi-layer-insulation (MLI) foil as shown in Figure 10. To reduce thermal conduction the cryostat is not mounted to the stator but rather mounted to the

Energiesvacuum2019 chamber, 12, 86 housing via insulated support rods that protrude through ports in the stator.10 of 14 Between the cryostat wall and the stator body is a copper EM shield [20]. It is designed to protect the superconductor from the alternating field due to current in the armature windings. It is attached on theto the room-temperature room-temperature bore bore side side of of cylindrical cylindrical cryostat cryostat box. box. Any Any losses losses inin itit areare removedremoved usingusing the means employed for cooling the room-temperatureroom-temperature surfacesurface ofof thethe cryostat.cryostat.

. Figure 10. Schematic cross-section of HTS excitation coil cryostat. The HTS coil consists of two 12 turn ReBCO coils which remain static while magnetizing the rotating solid steel rotor. Figure 10. Schematic cross-section of HTS excitation coil cryostat. The HTS coil consists of two 12 turn TheReBCO field coils winding, which remain made ofstatic ReBCO, while hasmagnetizing 24 turns. the It carriesrotating 317 solid A whensteel rotor. the machine is operating at full-load. A different superconductor could be used for generating the total amp-turns equal or greater The field winding, made of ReBCO, has 24 turns. It carries 317 A when the machine is operating than 7600 (~24 turns * 317 A). Furthermore, the ratio of operating-current/critical-current (I /I ) for the at full-load. A different superconductor could be used for generating the total amp-turnso cequal or superconductor is preferred to be less than or equal to ~50%. Current leads for the superconducting greater than 7600 (~24 turns * 317 A). Furthermore, the ratio of operating-current/critical-current (Io/Ic) coil provide a path for thermal conduction from room-temperature to the cold environment of the for the superconductor is preferred to be less than or equal to ~50%. Current leads for the superconducting coil, and require the exciter circuit to include non-superconducting elements which superconducting coil provide a path for thermal conduction from room-temperature to the cold generate heating through ohmic losses. This thermal load is estimated to be ~25 W. A contactless environment of the superconducting coil, and require the exciter circuit to include non- exciter concept is available [21] that could charge the excitation coil without physical current leads superconducting elements which generate heating through ohmic losses. This thermal load is spanning room-temperature and cryogenic environments, while enabling the exciter circuit to be estimated to be ~25 W. A contactless exciter concept is available [21] that could charge the excitation almost entirely superconducting. With this exciter implemented, it would be possible to operate the coil without physical current leads spanning room-temperature and cryogenic environments, while flywheel continuously with a smaller cooling system as the conduction and ohmic heating loads are enabling the exciter circuit to be almost entirely superconducting. With this exciter implemented, it greatly reduced. would be possible to operate the flywheel continuously with a smaller cooling system as the 7.conduction Rotor Design and ohmic heating loads are greatly reduced. Figure 11 shows a three-dimensional view of the rotor. This configuration creates a four-pole field 7.Energies Rotor 2018 Design, 11, x FOR PEER REVIEW 11 of 15 pattern as viewed from each half of the stator core. The rotor is manufactured from one solid piece of CarpenterFigure 11 Technology shows a three-dimensional Corporation, Philadelphia, view of th USA,e rotor. Aermet This configuration 310 magnetic steelcreates or similara four-pole and fieldpiece pattern of Carpenter as viewed Technology from each Corporation, half of the Philadel stator core.phia, The USA, rotor Aermet is manufactured 310 magnetic from steel orone similar solid shrink-fitand shrink-fit onto onto the steelthe steel shaft. shaft. The The diameter diameter is selectedis selected based based on on the the structural structural strength strength of of selectedselected rotorrotor material.material. TheThe rotorrotor bodybody betweenbetween thethe twotwo setssets ofof solid-polessolid-poles carriescarries thethe fluxflux fromfrom solid-polessolid-poles onon thethe leftleft toto thethe solid-polessolid-poles onon right,right, thisthis magneticmagnetic circuitcircuit isis completedcompleted right to left through the stator, surroundingsurrounding the the exciter exciter coils coils (Figure (Figure7). The 7). fluxThe within flux within the rotor the is essentiallyrotor is essentially DC therefore DC saturation therefore ofsaturation the rotor of body the isrotor permissible. body is permissible.

. Figure 11. An isometric view of the solid rotor body on a shaft (all dimensions in mm). Figure 11. An isometric view of the solid rotor body on a shaft (all dimensions in mm). 8. Stator Design 8. StatorFigure Design 12 shows the arrangement of stator laminations. The end sections are radially laminated, while the bridge section is concentrically laminated, e.g., like paper wound onto a scroll. Armature Figure 12 shows the arrangement of stator laminations. The end sections are radially laminated, while the bridge section is concentrically laminated, e.g., like paper wound onto a scroll. Armature coils span the stator laminations parallel to the rotational axis. In the DL machine the armature windings are housed in the stator lamination teeth. The exciter coil is located between end sections and inside the bridge section as shown in Figure 1. The selected core material is the JFE Steel 10JNEX900. This material is available in 0.1 mm thickness. The core loss is 10 W/kg in 1 T field and 1 kHz frequency. Its saturation magnetic field is 1.8 T.

Figure 12. Arrangement of stator end section laminations.

The end sections are constructed of an axially laminated core block surrounded by radially laminated ring segment blocks. Slots and teeth are cut into an axially laminated core, as commonly found in conventional machine cores. On the outside perimeter of this core, circular segments make up balance.

9. Performance Parameters

Table 4. Machine parameters at rated frequency (833 Hz).

Parameters Value Rated line voltage, V 866 Rated current, A 360 Power rating, kVA 540

Energies 2018, 11, x FOR PEER REVIEW 11 of 15 piece of Carpenter Technology Corporation, Philadelphia, USA, Aermet 310 magnetic steel or similar and shrink-fit onto the steel shaft. The diameter is selected based on the structural strength of selected rotor material. The rotor body between the two sets of solid-poles carries the flux from solid-poles on the left to the solid-poles on right, this magnetic circuit is completed right to left through the stator, surrounding the exciter coils (Figure 7). The flux within the rotor is essentially DC therefore saturation of the rotor body is permissible.

Figure 11. An isometric view of the solid rotor body on a shaft (all dimensions in mm).

8. Stator Design Energies 2019, 12, 86 11 of 14 Figure 12 shows the arrangement of stator laminations. The end sections are radially laminated, while the bridge section is concentrically laminated, e.g., like paper wound onto a scroll. Armature coilscoils span span the the stator stator laminations laminations parallel parallel to to the the rotational rotational axis. axis. In In the the DL DL machine machine the the armature armature windingswindings are are housed housed in in the the stator lamination teeth.teeth. TheThe exciter exciter coil coil is is located located between between end end sections sections and andinside inside the bridgethe bridge section section as shown as shown in Figure in Figure1. The selected1. The selected core material core ismaterial the JFE is Steel the 10JNEX900. JFE Steel 10JNEX900.This material This is availablematerial is in available 0.1 mm thickness. in 0.1 mm The thickness. core loss The is 10 core W/kg loss in is 1 10 T ﬁeldW/kg and in 1 kHzT field frequency. and 1 kHzIts saturationfrequency. magneticIts saturation ﬁeld magnetic is 1.8 T. field is 1.8 T.

. Figure 12. Arrangement of stator end section laminations. Figure 12. Arrangement of stator end section laminations. The end sections are constructed of an axially laminated core block surrounded by radially laminatedThe end ring sections segment are blocks.constructed Slots andof an teeth axiall arey cutlaminated into an axiallycore block laminated surrounded core, asby commonly radially laminatedfound in conventionalring segment machineblocks. Slots cores. and On teeth the outsideare cut perimeterinto an axially of this laminated core, circular core, segmentsas commonly make foundup balance. in conventional machine cores. On the outside perimeter of this core, circular segments make up balance. 9. Performance Parameters 9. Performance Parameters Table4 includes parameters necessary for calculating machine performance under different loads and operating conditions as well as for interfacing it with inverters. Parameters are expressed in Table 4. Machine parameters at rated frequency (833 Hz). per-unit values; referenced to a base impedance of 1.39 ohm. Synchronous reactance of this machine is low due to the larger air-gap betweenParameters rotor and stator. Small synchronousValue reactance is usually beneficial for stable operation ofRated the machine line voltage, on an electricV grid. However, the 866 small value also creates very large fault current during suddenRated current, short-circuit A of the stator winding. 360 Attention needs to be paid to this aspect during inverterPower design. rating, No-load kVA and full-load field currents 540 are 278 A and 317 A, respectively. The field winding should be designed to carry 317 A with an Ic/Io of about 2. However,

it must be noted that if the machine operating at full-load suddenly loses load, the armature voltage will rise to 1.14 pu (= 866 × 1.14 = 987 V-rms). Since it is difficult to change HTS field winding current rapidly, it would be necessary to modulate the voltage electronically in the inverter. In the event of a quench of the superconducting field winding, it would be necessary to extract its magnetic stored energy as rapidly as possible. The field winding will be shorted at its terminal through a 0.32 Ω discharge resistor—the maximum voltage imposed (during the discharging operation) on the field winding is limited to 100 V. The time constant for the discharge current is 3.0 ms. Table5 summarizes performance data for the machine calculated using 2D finite element analysis. The performance has been calculated for a full-load current of 360 A at a power-factor of 0.99 lagging. A more accurate value for power-factor would be based on the inverter design but the value selected here is conservative. The load angle with respect to the terminal voltage is 1.34◦, which is quite small. However, in absence of a fixed terminal voltage, the value of load angle is of little interest. The nominal shaft torque for the machine is 208 Nm. But during short-circuit at the machine terminals, the torque will increase to 1487 Nm. Thus, it would be necessary to design the rotor shaft and the armature winding mechanical support to bear this torque with an acceptable safety margin, e.g., a safety factor of at least 2. Attractive forces between rotor and stator are 9282 N under a solid-pole and 246 N under the void-pole. Energies 2019, 12, 86 12 of 14

Table 4. Machine parameters at rated frequency (833 Hz).

Parameters Value Rated line voltage, V 866 Rated current, A 360 Power rating, kVA 540 Base impedance, Ω 1.39 D-axis synchronous reactance, pu 0.14 Q-axis synchronous reactance, pu 0.02 Armature resistance, pu 0.0023 Induced voltage with pf = 1 load, pu 1.14 Field current-no-load, A 278 Field current-full-load, A 317 Field winding inductance, mH 0.96 Maximum ﬁeld winding discharge voltage, V 100 Resistance of ﬁeld winding discharge resistor, Ω 0.32 Field winding discharge time constant, ms 3.0 Mutual inductance between Field and Armature, H 6.29 × 10−5

Table 5. Calculated performance.

Performance Value Rated line voltage, V 866 Rated current, A 360 Power rating, kVA 540 Field current-no-load, A 278 Power factor at full-load-lagging 0.99 Field current-full-load, A 317 Load angle at full-load, deg 1.3 Induced voltage at full-load, pu 1.14 Power generated at full-load, pu 1.01 Torque on armature at full-load, Nm 208 Torque on armature during a short-circuit, Nm 1487 Attractive force between rotor and stator Under a salient pole, N 9282 Under non-salient pole, N 246 Losses at Full-load Value Stator iron core loss, kW 1.7 Armature resistive loss, kW 1.2 Armature cooling system power, kW 0.13 Field winding cooler power, kW 1.37 Refrigerator Coefﬁcient of Performance (COP) 21 Total losses at full-load, kW 4.45 Efﬁciency at full-load, % 99.2 Total losses at no-load, kW 3.23

The Table5 also lists different loss components when the machine is carrying full-load. Stator core loss and armature winding resistive loss are two significant components. Cryogenic cooling system power is estimated to be 1.37 kW based on a refrigerator coefficient of performance (COP) of 21 at 50 K (COP is the ratio of power input to refrigerator to the heat power removed by the refrigerator). This COP value is obtained from Yuki Iwasa (MIT) book, 2nd Edition [22]. A cooler system selected by a manufacturer may have different COP. Furthermore, thermal load calculations for the HTS coil might be different as they are a function of HTS coil type, its construction and interface with the room-temperature systems. Based on the assumptions made here, total loss at full-load is 4.4 kW, which yields an efficiency of 99.2%. This looks attractive at this conceptual design process and could be used as a goal to strive for. Energies 2019, 12, 86 13 of 14

It must be noted that this machine experiences significant losses under no-load operation. For example, if the field current is kept at its full-load operating value, the machine thermal load (cryo-cooler power) and core loss will remain unchanged and total loss would be 3.2 kW. However, if the field current is turned off during a prolonged no-load operation then the only loss component (cryo-cooler power loss) would be less than 1.36 kW. Thus, it would be necessary to turn off the field current when the flywheel is turning and storing energy for a long period of time. With contactless current transfer to the HTS coil, the above-mentioned 1.36 kW loss would be reduced to 30% (~0.4 kW).

10. Summary For application of a ground-based fast energy storage system the DL concept looks attractive as compared with the SL design, which is more expensive due to the longer length of superconductor needed for the field winding. The DL solution is manufacturable using existing methods and technology, which will result in reduced construction time and cost, as well as shorter development time. The analysis presented here is based on employing ReBCO field coils, however, the comparison is equally valid for lower cost DI-BSCCO coils. DI-BSCCO coils are a commercially available technology with potential for further reducing development and construction time and cost. The concept design presented here is similar to a superconducting machine built and tested by General Electric [16] and, therefore, the authors are sure about the feasibility of the presented machine. One of the authors was also part of a team that built and tested such a machine in about 1978 using copper field windings. The design of a complete flywheel system suitable for fast energy recovery and delivery presents several challenges not within the scope of this paper. Magnetic bearings will be required to support the high rotational frequencies, as will a machine enclosure capable of holding partial vacuum to prevent significant windage losses. Design and development of flywheels and the machine shaft also present several challenges relating to the high rotational frequencies. We are currently working towards manufacturing a demonstrator DL homopolar motor. As part of this process we will address flywheel construction and magnetic bearing design including the potential to use passive superconducting bearings.

11. Patents BADCOCK, Rodney Alan; HAMILTON, Kent Anthony; KALSI, Swarn Singh. ‘Flywheel Energy Storage System.’ CHN.Patent ZL 2017210288283, June 1, 2018. BADCOCK, Rodney Alan; HAMILTON, Kent Anthony; KALSI, Swarn Singh. ‘Flywheel Energy Storage system.’ PCT Application PCT/NZ2018/050093.

Author Contributions: Conceptualization of the electrical machine: S.K.; conceptualization of the flywheel integration: R.A.B.; conceptualization of the integrated vacuum-contained flywheel system: K.H.; validation: S.K., R.A.B., and K.H.; formal analysis: S.K.; resources: Robinson Research Institute; writing—original draft preparation: S.K., R.A.B., and K.H.; writing—review and editing: S.K., R.A.B., K.H., and R.G.B.; visualization: K.H.; supervision: R.A.B.; project administration: R.A.B.; funding acquisition: R.A.B. Funding: This research was funded by the New Zealand Ministry of Business Innovation and Employment grant number RTVU1707. Acknowledgments: The authors wish to acknowledge the invaluable advice and assistance provided by James Storey and Chris W Bumby. Conflicts of Interest: The authors declare no conflict of interest.

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