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Fusion Engineering and Design 38 (1997) 159–188

ARIES-RS magnet systems

L. Bromberg a,*, P. Titus b, J.S. Schultz a, M. Sidorov a, S. Pourrahimi a,The ARIES Team

a Plasma Fusion Center, MIT, Cambridge, MA 02139, USAb Stone and Webster, Boston, MA 02215, USA

Abstract

A conceptual design of toroidal and poloidal field systems for the ARIES-RS reactor study is presented in thispaper. Means of designing the toroidal and poloidal field system for minimized size and cost, optimized structure andincreased access for maintenance are presented in the paper. Supports of the out-of-plane TF loads that do notinterfere with maintenance operation have been designed. Structural analyses of several cases that have the commonfeature of avoiding material in between the outer legs of the TF coil are presented in this paper. The implications onthe structural amount of material required are investigated. Methods of handling failure conditions in the toroidalfield coil due to unbalanced currents are studied. Optimization of the conductor in the poloidal and toroidal fieldsystems is carried out. Implications of the use of very fine superconducting strands for conductor stability andminimization of co-wound normal-conducting material are evaluated. Implications of novel schemes for magnetprotection such as internal dump, are described. © 1997 Elsevier Science S.A.

Keywords: ARIES-RS; Toroidal field; TF coil; Internal dump

1. Introduction

The ARIES-RS reactor is a conceptual com-mercial reactor based on modest extrapolationfrom the present tokamak physics and engineeringdata bases. In contrast with ARIES-I [1], with ahigh magnetic field and aggressive engineering,but with modest extrapolation of the physics,ARIES-RS operates with more conservative mag-net engineering. The engineering approach toevaluate the magnet issues for ARIES-RS is simi-lar to that of ARIES-II and ARIES-IV [2].

The magnets for ARIES-RS differ in severalways from the magnets design for ARIES-I [1]and from PULSAR [3]. On one hand, the designof the toroidal field magnet is less demandingthan that of ARIES-I due to the lower magneticfield. On the other hand, the problems in PUL-SAR with respect to relatively fast transients andthe fatigue problem with many pulses are notissues for ARIES-RS. The protection issues for allthese magnets, are however, comparable.

The objectives of the magnet work in ARIES-RS are to use conventional materials in the designof an unconventional magnet system that pro-vides large access for maintenance. However,within the present constraints, the design has uti-* Corresponding author.

0920-3796/97/$17.00 © 1997 Elsevier Science S.A. All rights reserved.

PII S0920 -3796 (97 )00115 -4

L. Bromberg et al. / Fusion Engineering and Design 38 (1997) 159–188160

lized credible but optimistic values. Superconduct-ing engineering margins that are required due topartial understanding of the superconducting be-havior have been minimized in order to explore,not a present-day conservative design, but anaggressive design.

In the ARIES-RS magnet study, four objectiveshave been pursued. First, the magnet cross sectionhas been optimized, using innovative solutions tominimize the cross section and the cost. The sec-ond objective in the ARIES-RS study has in-volved the investigation and support ofoff-normal magnet events. Due to lack of re-sources, these issues were summarily addressed inprevious studies. A failure analysis of the TF coilwas performed with structures as in the PULSARand in ARIES-II and ARIES-IV designs. Theresults of this investigation indicated that onemode of failure, a short across the leads to onecoil followed by dump of the rest of the system,results in large bending and deformation in theshorted coil as it tries to go into a circular shape.This failure mode is addressed in the presentstudy.

The third objective of the ARIES-RS study hasto do with access. In ARIES-RS the feasibility ofa maintenance approach that removes entire blan-ket/shield sectors was investigated. This approachrequires large access ports between the outer legsof the TF coil. The possibility of supporting thesesections of the coil without structure between legs(that would obstruct easy access, even if remov-able) was evaluated. Finally, the impact of theenlarged toroidal field coil on the poloidal fieldsystem has been investigated. This issue requiredthe modification of the TF coil.

The DEMO program was a dynamic processwhere self-consistent iterations were subsequentlyanalyzed. Due to the nature of the process, theanalysis referred to in this paper corresponds notto the latest case, but to the next-to-last straw-man. The parameters used in the analysis, how-ever, are self consistent.

The paper is divided as follows: in Section 2 theassumptions on the design of the TF coil, and itsoptimization, are described. In Section 3, the cal-culations that lead to the present design of the TFcoil superstructure are described. In Section 5, the

results of failure mode analysis are summarized,both for the toroidal and poloidal field systems.In Section 4, the PF coil system for the presentdesign is described. Conclusions are presented inSection 6.

2. Toroidal field system

The issues of the toroidal field coil are discussedin this section. First, the overall design of the coilis presented. Then the optimization of the crosssection, along the constraints in the optimization,are presented. Several analyses of the TF systemare then presented, and the bases for the changesrequired in the superstructure to satisfy themaintenance requirements, are indicated.

2.1. TF design

The reference ARIES-RS design uses multifila-mentary Nb3Sn superconductors with a peak fieldof 15.67 T at the coil. In this section,the TF coilconstraints and concepts are reviewed, and thecoil cross section is optimized. The ARIES-RSdesign optimizes the superconductor/copper, he-lium, insulation, and structural ratios of the coilin order to minimize the radial thickness of theTF magnet system subject to the following restric-tions: superconductor critical current, supercon-ductor stability, quench protection,superconductor strain, stress and strain in struc-tural materials, heat removal, pumping power,conductor fabrication, and magnet construction.

The conceptual magnet design has been basedupon isotropic structural materials. Fig. 1 shows a3-D representation of the TF magnets, the PFmagnets and the structure for ARIES-RS. Asshown in the figure, no intercoil structure isplaced between the outer legs of the TF coils.Also, the PF coils have been placed so that theydo not have to be removed during blanket/shieldand first wall maintenance operations, when entiresectors are removed between the legs of the TFcoil.

Advanced magnet design techniques have beenutilized that maximize the utilization of the struc-tural and the superconducting materials, minimiz-

L. Bromberg et al. / Fusion Engineering and Design 38 (1997) 159–188 161

Fig. 1. 3-D view of the toroidal and poloidal field magnets, and the super-structure.

ing the cross section of the coil. The ARIES-RSmagnet design concept uses structural forms withgrooves into which the conductor is wound, as inthe ARIES II–IV [2] and the ITER-Rebut [4]design. A thorough comparison between toroidal-shell and winding-in-case coils could not be car-ried out, since the elements that drive the designare buried in the details. However, due to im-proved capability to support the out of planeloads of the toroidal-shell magnets, it is expectedthat they are superior to magnets with a case, andto magnets with radially-directed plates. Thetoroidal shell magnet has been investigated inARIES-II and -IV [2], PULSAR [3], as well as inthe Rebut-ITER design [5]. A major differencebetween the proposed design and that of theRebut-ITER [6] is that the toroidal field leansagainst the bucking cylinder in ARIES-RS, asopposed to the Rebut-ITER, where it leansagainst the central solenoid. In the case of the

Rebut-ITER, the out-of-plane loads in the innerleg of the TF coil were reacted by the coilsthemselves [7], since the idea of transferring theout-of-plane loads through the ohmic solenoidwas rejected. In the case of ARIES-RS, the buck-ing cylinder is used to react the out-of-planeloads.

The ARIES-RS magnet design is similar to theARIES-II, -IV and PULSAR magnet designs,utilizing a winding method that eases the fabrica-tion of the magnet by removing the stiffest mate-rial from the winding process. Thus, rather thanwinding all of the materials in the magnet, onlythe conductor requires winding.

The use of layered-wound magnets in ARIESII, IV and PULSAR allows the use of multiplegrades to wind the toroidal field coil. This allowsus to optimize the conductor in each layer. Thiswould not be possible in designs with radial platesor with pancake winding, since then the number


of joints between grades would be prohibitive.The length of the conductor in each layered is onthe order of 1 km, matching the manufacturingability for Cable-in-Conduit Conductor (CICC).Even in the absence of grading, joints are requiredbetween adjacent layers. The depths of thegrooves in the shells are determined by the designof the conductor, and vary across grades. A dif-ferent depth, and therefore a different toroidalshell geometry, is required for each grade. Sincethe shells are prevented from sliding with respectto each other, large out-of-plane strength is ob-tained.

To prevent sliding of one shell with respect tothe other, which may damage the insulation or theconductor [8], it may be necessary to bolt theplates together or to e-beam weld the edges of theplates together. The disadvantage of the latterapproach is that the plates are at a commonvoltage, and therefore the insulation around eachconductor must be able to withstand the fullquench voltage. In the case of bolting the platestogether, it is not necessary to have the plates atthe same potential, and the potential of the platescan be adjusted in order to minimize the maxi-mum voltage across the insulation. The maximumvoltage is decreased to about the full quenchvoltage divided by the number of plates, decreas-ing the insulation thickness. The issues that arisedue to the use of bolts have to do with fatigue,but since the number of pulses in ARIES-RS islimited, fatigue doe not drive the design.

In ARIES-RS, an internal dump method isproposed, i.e. large sections of the coil are heatedabove superconducting state creating a large nor-mal zone in the magnet. Details of internalquenching are described in the Section 5.2. Thecharacteristics of the quench under these circum-stances, in particular the structural behavior dur-ing a quench, need to be investigated.

Layer winding a magnet not only simplifies theelectrical connections in the magnet, but also thecryogenic connections, since headers can be ap-plied at approximately 1 km intervals (betweenlayers). In addition, the winding process is sim-plified, since inserting the conductor inside theouter slot in the shell requires a simple one-di-mensional operation. For these reasons, the layer-

wound concept was found superior for toroidalfield magnets of this size. The manufacturing re-quires the handling of the full conductor/shellstructure.

The general procedure for winding a toroidalfield coil is described next. After winding the firstinnermost layer, the next shell is put in place. Toallow placement of the shell, the shells must havea break. After putting it in place, the structuralcontinuity of the shell may be achieved by care-fully welding the shell in place (with the use of alap-weld to prevent damage to the superconductorunder the shell weld). Alternatively, the largespace available in the outer regions of the magnetcould be utilized for mechanically attaching theends of the shells (however, this would affect theaccess for maintenance). The location of the breakin the shell can be changed from shell to shell inorder to minimize the stress concentration aroundthe region of the break. The process is repeateduntil all the layers are in place. This process issimilar to that proposed for the Rebut-ITER con-cept [9].

In spite of its superior mechanical propertiesand low coefficient of thermal expansion, Incoloy908 has not been selected as the base material forthe plates because of its increased cost [10]. Thesheath of the cable, however, is made out ofINCOLOY 908 [11], mainly because of the manu-facturing advantage of this material, which closelymatches the thermal cycle required for heat-treat-ment of the Nb3Sn superconductor [12]. For thestructure, an allowable membrane stress of 800MPa has been assumed and a Young’s modulusof 182 GPa. For Cu, the stabilizer material, anallowed stress of 200 MPa and a Young’s mod-ulus of 185 GPa are used.

In order to investigate the consequences of themost optimistic outcome, it is assumed that thesuperconductor shares the strain with the copper,the sheath and the structure. It is assumed thatthe superconductor and the copper start sharingthe strain during cooldown after heat treatment,at a temperature of 500°C. The thermal and me-chanical strains are included in the analysis. Asexpected, the mechanical load offsets the thermalstrain imposed on the superconductor to the pointthat it is almost strain free, thus increasing the


critical current in the superconductor. The as-sumed strain is discussed below.

The depths of the grooves in the plates aredetermined by the design of the conductor. Thisrequires a different radial shell thickness for eachgrade. For the ARIES-RS design, four gradeswere chosen.

A novel approach to the issues of the insulationis to use internal dump [13]. If internal dump isused in the case of magnet coil quench, the largedump voltages are avoided, and the insulation canbe made substantially thinner. This substantiallyincreases the allowed current in the copper duringquench. In addition, it simplifies the charging anddischarging of the coil, by minimizing the numberof current leads.

The superconductor material is chosen tomatch the maximum local field, by using severalgrades of superconductor. The Nb3Sn supercon-ducting material assumed in this study [14,15] issubstantially more aggressive that the materialpresently available for ITER [16,17].

2.2. Coil optimization

The reference ARIES-RS TF coil has a peakfield of 15.67 T. For this design, all of the super-conductor material in the TF coil is multifilamen-tary (NbTi and Nb3Sn) and the structuralmaterial is SS 316LN. Coil design parameters aregiven in Table 1. The thermal characteristics ofthe coils are given in Table 2, indicating themaximum value of the discharge voltage (20 kV),the value of the integral

Smax=& 150

4

cr

hdT

where c, r and h are the heat capacity, densityand resistivity of the conductor material, respec-tively and the integral is carried from liquid he-lium temperature to the allowed temperature aftera quench (150 K) [18]. It is assumed that theResistivity Resistance Ratio of the copper isRRR=200, which is possible due to the presenceof pure copper strands. This calculation assumesthat the copper is adiabatic, and that the copperstrand shares the current instantly.

Table 1TF coil design parameters for the ARIES-RS reactor

B case 15.67

n grades 4

Magnetic field, grade 1 (T) 3.5Magnetic field, grade 2 (T) 10.5Magnetic field, grade 3 (T) 13.5Magnetic field, grade 4 (T) 16

Pack current (kA) 51.2Total TF A-turns (A) 1.88×108

Number of turns/coil 226Stored energy/per ‘electric’ coil (GJ) 1.6

51Total stored energy (GJ)2.63×10−3Average current density (A m−2)

Effective moduli (GPa) 2002.81×10−3Overall strain562Mechanical stress (MPa)

Inner leg parametersOuter radii (m) 2.4Inner radii (m) 1.9

0.49Thickness (m)1.36Bucking cylinder inner radii (m)

Also shown are the contraction of the magnetmaterial between room temperature and 4 K, usedto calculate the mechanical state of the supercon-ductor at 4 K and no current. In this design, it isassumed that the mechanical strains are sharedwith the conductor strands. In this case, the me-chanical strain in the system is enough to compen-sate the thermally induced strain in thesuperconductor (due to the differences in thermalcontraction, as shown in Table 2).

Table 2Thermal contraction and quench assumptions for ARIES-RSTF coil design

Vmax (kV) 2011×1016Smax (A2 s m−4) (corresponding to DT�200 K)

Energy margin (J cm−3) 0.2Fraction of critical 0.8

Thermal contraction (293–4 K)2×10−3Superconductor3×10−3Stabilizer

Sheath 2×10−3

Plate 3×10−3

Insulation 2×10−3


Table 3Conductor-grading design parameters for ARIES-RS

43Grade 21

1613.5Field split 10.53.52.71×109 1.05×109 0.74×109Current density (A m−2) 4.7×109

7.779.41Critical temperature (K) 11.396.94Areas (10−4 m2)

0.489Superconductor 0.109 0.189 0.6911.96 1.96Copper 1.961.961.71Helium 1.69 1.70 1.71

4.364.16Total conductor 3.853.7611.36 9.98 9.19Structure 11.35

0.3450.360Insulation 0.3900.38814.5Total cross section 15.5 15.6 13.9

0.023Radial thickness (m) 0.027 0.026 0.024

The conductor grading information is shown inTable 3. The areas are shown for a single conduc-tor. The radial thickness refers to the thickness ofthe toroidal shells (the conductor is imbedded inthe shells). It is assumed that the current density inthe superconductor is 0.8 of critical. The supercon-ductor is ternary tin, and at the operating point iseffectively strain free (thermal and mechanicalstrains in the superconductor approximately bal-ance).

The description of the TF magnet using fourgrades of conductors with the characteristics inTable 3 are given in Table 4. The field and grade,the number of conducting elements per shell, andthe geometry of both the inner and the outer legsof a given row in the coils are shown.

In previous studies it was assumed that thesuperconductor cost decreased linearly with thesuperconducting to copper ratio [19]. However,recent experience with superconducting Nb3Sn in-dicates that the cost is not sensitive to the super-conductor/copper ratio. In this case, it isadvantageous to minimize the amount of copper inthe strand and add inexpensive normal conductingstrands adjacent to the superconducting strands.The copper is not as effective for conductor stabil-ity, and the energy margin of the superconductoris decreased. The stability of multiple strand super-conductor is not well understood. The energymargin (design value of 0.2 J cm3 for the toroidalfield coil) is about 20% of the design values inpresent day CICC designs, reflecting an improved

understanding of the present-day limitations thatmay result in improved future performance. Animportant difference between ARIES-RS andITER or TPX, is the long time allowed for tran-sients, since its operation is steady state and therest of the power plant (power cycle system, tur-bine, etc.), anyway, can not tolerate fast transients.

Although from an operation point of view itwould be useful for the toroidal field coil not toquench during disruptions, in practice a disruptionis a serious off-normal event that will requiremonitoring of the nuclear island to ensure that theevent did not harm the different systems in thepower plant. The transients in the TF magnetsystems due to a disruption have not been ana-lyzed, although due to the low energy margin it isexpected that it will result in quench. In the caseof a disruption, however, it may be necessary torecool the TF magnet.

Indeed, it may be possible to use the limitedenergy margin in the system to induce the internalquench by increasing the rate of transients, espe-cially during scram-type events.

The power-balance stability in the superconduc-tor in the ARIES-RS design is obtained by decreas-ing the size of the conductor, from about slightlyless than 1 mm for most designs, to about 0.4 mmin the ARIES-RS design. Thus the surface area perunit volume is substantially increased in this de-sign, allowing the use of individual strands thathave high superconductor to copper ratio, as de-scribed below.


Table 4Row by row design parameters for ARIES-RS TF coil

4 5Row number 61 2 33Grade 34 34 4

Inner leg11.2812.26Bmax (T) 15.74 13.2215.08 14.16

2.25Outmost radius (m) 2.35 2.33 2.30 2.28 2.23206Radial stress (MPa) 84 164 238 170 238

750682Combined stress (MPa) 596 717675 750

Outer leg9.259.22Innermost radius (m) 9.13 9.209.15 9.18

14 14Number of conductors 1414 14 14

1211Row number 107 8 92Grade 2 2 2 2 3

Inner leg6.57Bmax (T) 10.27 9.26 8.39 5.637.492.10Outmost radius (m) 2.21 2.18 2.15 2.13 2.08

216Radial stress (MPa) 225161 205177 192727Combined stress (MPa) 673 689 703 716 737

Outer leg9.35 9.38Innermost radius (m) 9.409.27 9.30 9.33

12Number of conductors 1214 1212 12

17Row number 13 14 15 18161Grade 12 12 1

Inner leg0.55Bmax (T) 4.67 3.68 2.66 1.62 0.001.94Outmost radius (m) 1.922.05 1.972.02 2.00

238Radial stress (MPa) 233 238 236 266238750Combined stress (MPa) 750744 750750 748

Outer leg9.54Innermost radius (m) 9.43 9.46 9.48 9.51 9.54

10Number of conductors 1012 1212 12

Quench protection is, to zeroth order, not af-fected by the copper segmentation [18].

Fig. 2 shows the TF conductor composition forthe several grades. For ARIES-RS, the conductorconsists of multifilamentary superconductingcomposite strands cowound with normal conduct-ing strands (pure copper). It is not necessary tosolder the conductors together since the contactunder pressure of the strands results in goodcurrent sharing amongst the superconductors andgood current transfer capabilities in the case ofoff-normal events. The structure and behavior ofthese heterogeneous conductors are being activelystudied at the present time. It has been found thatthe cost for the NB3Sn strands is not a strong

function of the superconductor to copper ratio.Therefore, substantial savings in the cost of thestrands can be achieved if the superconductor issegregated to strands with high SC/Cu ratio. ForARIES-RS, the ratio of the pure copper strand tothe total number of strands is 6:7 in the first gradeand decreases to 3:7 in the fourth (and highestfield) grade.

The winding rule for the TF conductors for thefirst and second grades (low field) is 7×3×3×3×3, while the third and fourth grades is 7×3×3×3×3×3, with have an additional windingstep to compensate the fact that the wire thicknessis about two-thirds that of the first and secondgrades. The first and second grades have, there-


Fig. 2. Different conductors for the TF coil.

heating of the coils. The effect of this heating hasnot been included in the analysis, and wouldresult in a small increase in the coil cross section.

Fig. 3 shows the cross section in: (a) inner; and(b) outer legs of the TF coil. The CIC conductoris circular, imbedded in toroidal shells. Thetoroidal width of the TF coil can therefore beeasily adjusted, as is done in the ARIES-II design,in order to minimize the toroidal shells. Thetoroidal width of the TF coil can therefore beeasily adjusted, as is done in the ARIES-II design,in order to minimize the toroidal extent of theouter legs of the TF coil (thus increasing access tothe blanket/shield). It is possible to carry thisfurther, by decreasing the toroidal distance be-tween the conductors, and therefore minimizingthe toroidal width of the coil. However, in ordernot to depart too much from conventional mag-net design and construction, this approach wasrejected.

The design is dominated by the structural mate-rial requirement. The overall thickness of the TFcoil (without protecting sheath, dewar and insula-tion) in the outboard leg is about 0.5 m.

The cooling of the toroidal field magnet pre-sents some interesting challenges. It is necessaryto keep the temperature as low as possible, espe-cially in the inner most conductor that receivesthe most radiation heating and is the region ofhighest field. It was assumed in the calculationsthat the maximum temperature in the conductorwould exceed 5 K.

The hydraulic path in the conductor, assumingthat each layer is independently cooled, is of theorder of 500 m. In order to minimize the heatingin the conductor, the structure will be cooleddirectly, to avoid the need to remove the heatingin the structure (due to eddy currents and toradiation) by the coolant in the conductor. Fi-nally, if necessary, the conductors will be cooleddirectly in the outboard section of the magnet, byflowing He through the plates, which are in closethermal contact with the conductors (due to thelarge surface area between them). This is an effec-tive manner of decreasing the thermal path, eventhough the hydraulic path is the same. This is anarea where additional work needs to be done.

fore, strands about 0.68 mm, while the third andfourth grades have a strand diameter of 0.4 mm.The reason for using a seven strand initial wind-ing stage is to increase the flexibility needed toprovide the appropriate superconductor to copperratio in each conductor grade.

The averaged fractions of different materialsfor the TF magnet are given in Table 5. Thisinformation has been used to determine the neu-tron and gamma ray fluxes in the design, and theresulting activation of the coils. The shield inARIES-RS is sufficiently thick to minimize the

Table 5Volume fraction in the TF coil, for the different grades

Grade 1 2 3 4

0.126Copper 0.125 0.136 0.142SC 0.0500.0330.007 0.012

0.108 0.108Helium 0.119 0.123Insulation 0.025 0.025 0.025 0.025Structure 0.735 0.728 0.685 0.659


Fig. 3. Cross section, in (a) the inner leg and (b) the outer leg of the TF coil.

3. TF magnet structural analysis

To improve scheduled and unscheduledmaintenance of the nuclear island, it would behighly desirable to find methods of supportingthe out-of-plane TF loads that do not interferewith maintenance operation. Even if thetoroidal field coil were extended outwardly toallow for maintenance of toroidally segmentedmodules, ideally there should not be structuresin between coils in the outboard legs of thetoroidal field coil. The number of these mod-ules is the same as the number of toroidal fieldcoils, and therefore one and only one module is

removed between a pair of adjacent TF coils.Structural analysis of several cases that havethe common feature of avoiding material in be-tween the outer legs of the TF coil are pre-sented in this section. The implications on theamount of structural material required are in-vestigated.

3.1. Introduction

The goal of this work is to investigate meth-ods that avoid the need for structural materialsfor coil support in the region between the outerlegs of the TF coil.


Since the ohmic transformer in designs for in-ductively-driven tokamaks (such as ITER andPULSAR) is between the TF coil and buckingcylinder, the TF coil in these designs can not usethe out-of-plane strength of the bucking cylinder.Furthermore, to minimize the pulse losses in thestructure it is necessary to electrically subdivideany out-of-plane structure (including the buckingcylinder). This substantially reduces the load-car-rying capacity of the overall structure.

For steady-state devices, with minimal plasmainductive requirements, it is possible to place thecentral solenoid stack inside the bucking cylinder.The TF coils can then lay directly on the buckingcylinder. The bucking cylinder can therefore beused to support the out-of-plane loads of the TFcoil. Similarly, in the top and bottom of thedevice a cap-like structure, similar to that of theprevious ARIES designs, can be used to minimizethe displacement of the out-of-plane loads.

This section discusses different methods of sup-porting the outer leg of the TF coil. Severalstructures have been analyzed. The first one,shown in Fig. 4, utilizes exclusively the outer legsof the TF coil and the bucking cylinder in theinboard side to support the out-of-plane loads.The system has a set of toroidally continuous capsthat are keyed to the bucking cylinder to supportthe top/bottom of the machine. The second casethat was analyzed, shown in Fig. 5, includedcontinuing the caps with a set of straps located inthe shadow of the outer legs of the TF coil. Thethickness of the straps in the shadow of the outerlegs of the TF coils was increased to about 0.60m, as shown in Fig. 6, at which point the stressesin the structural material became acceptable. Thefinal and fourth case, shown in Fig. 7, analyzedthe structural consequences of substantially in-creasing the height of the horizontal ports, inorder to accommodate the required plumbing thatthe blanket, shield and first wall require. Twoadditional changes were made between the thirdand the fourth models: the strap in the shadow ofthe TF coil was strengthened in order to minimizethe displacement and stress in the coil and super-structure, and the height of the coil was reduced,deviating the coil from a Princeton-D shape. Thelatter was done in order to reduce the size and

current of the PF coils, in particular the shapingcoils. These coils would carry an enormous cur-rent and produce high local fields if they wereplaced outside of a Princeton-D shape coil, asdescribed below in Section 4.

Results of modeling these four cases are pre-sented in this section. The sensitivity of thestresses to the height of the opening of the hori-zontal port is also reported.

3.2. Computational tools

The ARIES-RS structural behavior for out-of-plane loads has been analyzed using the ANSYScomputer code. The model is cyclically symmetric,consisting of a wedge which is 1/16 of the ma-chine. About the equatorial plane, the model issymmetric for in-plane loads and antisymmetricfor out-of-plane loads. The solution is non-linear.Gap elements are used at the coil/bucking cylin-der and the coil/shell interfaces. Crossed gaps atthe interface some model form of the out-of-planemechanical connection. However, as in ITER,friction is the major mechanism for transmissionof the out-of-plane forces [20].

The coordinate system uses local coordinates,with y as the coordinate along the TF coil (verti-cal in the throat), x as the coordinate across theTF coil (radial in the throat), and z as the toroidaldirection. Hoop stresses are calculated along ele-ment centroids in the direction along the TF coil.We have assumed that the current density in theTF magnet is uniform in each grade, i.e. thecurrent has been smeared over the conductor,structure, stabilizer, insulation, and coolant. Thetoroidal field is calculated using this current distri-bution.

The Lorentz forces are obtained from an analy-sis using grid points that are compatible acrossthe force analysis model and the ANSYS model.This compatibility of the nodal arrangement facil-itates communication of element centroid forcesbetween the force code and ANSYS.

For the FEM calculations, static equilibriumcalculations for different times in the scenario ofthe discharge were used. The out-of-plane loadsare calculated using the location and current ofthe coils indicated in Table 7, using finite cross-


Fig. 4. First model, with caps but no straps in the shadow of the outer leg of the TF coil.

section conductors. The entire scenario was inves-tigated for the third and fourth cases. The worstloading occurs at the End of Burn (EoB), and thepaper describes only this case.

Gaps between the TF coils ensure that the coilsdo not wedge against each other in the inner leg.Strong coupling exists between the coil and the

bucking cylinder, ensuring that the out-of-planeloads are transmitted from the TF coil to thebucking cylinder. This is made possible by thedirect contact between the TF coil and the buckingcylinder in the ARIES-RS design, unlike PULSARand ITER, where the ohmic transformer is betweenthe TF coil and the bucking cylinder.


Fig. 5. Second model which incorporates strap around the shadow of the outer leg of the TF coil.

Cyclic symmetry constraints are assumed. Theyare accomplished by coupling nodes from symme-try boundary to symmetry boundary with degreesof freedom specified in a cylindrical coordinatesystem. ANSYS STIFF-45 eight-node solid ele-ments are used throughout the model.

The calculated stresses are ‘smeared’ stresses.The local metal stress would require a more de-tailed modeling of cross-section. The finite ele-ment results confirm the simple treatments aboveand also show large bending stresses which resultfrom a deviation from the constant tension shape.


Fig. 6. Third model, with reinforced strap.

3.3. ARIES-RS parameters

The plasma major and minor radii are 5.24 and1.31 m, respectively. The plasma current is 10.4MA and the plasma elongation is 1.85 at theseparatrix. The number of TF coils is 16. Theseparameters are slightly different from those of the

final ARIES-RS design, since the numbers usedwere of the next to last iteration in the overalldesign.

The geometrical parameters of the PF systemsused in the calculations vary between models. Thecoil location and currents vary between models, asthey are optimized for each model separately. For


Fig. 7. Fourth model, with increased port height and further strap reinforcement.

models 1–3, the coil location and currents do notvary. However, for model 4 the coil location changessince the height of the TF coil has been decreased,allowing the placement of the divertor coils closerto the plasma. The geometry of the fourth case isdescribed in Section 4. In this paper, however, only

the currents are given for cases 1–3 in Table 6.The coil size is determined by the allowable

current density, which is in itself determined by themagnetic field and the loads. A detailed engineeringanalysis of the PF coil for the fourth model ispresented in Section 4.


In the ARIES-RS design, the outer leg of theTF coil has been moved sufficiently back so thatentire segments of the blanket/shield can be re-moved through the gap in-between the coils witha purely radial motion. The blanket and shieldhave been divided in as many segments as thereare TF coils. The width that determines the re-quired position of the outer leg of the TF coils isthe highest/lowest point in the blanket/shield.Both the opening between the TF coils and thetoroidal width of the blanket/shield decrease awayfrom the midplane, but the TF coil opening de-creases faster than the width of the blanket andshield.

The structural reaction to the out-of-planeloads requires special attention because the sec-tions that conventionally react the shears in thereactor were removed to simplify maintenance.Only normal conditions have been analyzed withthis model. Section 5 summarizes results for esti-mates of consequences of off-normal events.

3.4. Results

Four structural models have been investigated,with the hope of removing most material from theregion in between the outer legs of the TF coils.The four models are shown in Figs. 4–7. The firstmodel includes top and bottom caps (attached tothe bucking cylinder), but no additional structure(strap) in the outer legs of the TF coil. Thethickness of the caps is the same as the radial

thickness of the bucking cylinder. The caps areattached to the TF coil through a key that runsalong the TF coil midplane. In practice, the TFcoil would fit in a slot in the caps, in order totransfer the out-of-plane loads from the TF coilinto the caps. In this manner, the TF coil and thecaps are locked in the toroidal direction but thecoil is allowed to slide in the poloidal direction.

The peak stresses in the first model (assumingsmeared properties for the TF coils) are 1.4 GPain the area of the outer leg of the TF coil wherethe caps ceases to support the TF coil. The capsin this case are rotating toroidally, taking alongthe coil which has little capacity to prevent themotion. The outer leg of the TF coil thereforeexperiences a strain-controlled condition, and thecoil deforms into the typical S-shaped curve. Thecoil experiences very large stresses across the re-gion where the bending of the coil is maximum. Inaddition to the large loads in the region where thecaps end and the TF coil is self-supporting, thereare large inplane bending stresses at the innerchord of the coil near the top and bottom of themachine, as the coil and the structure react to thelarge displacement of the caps. These stresses are1.2 GPa. Both of these stresses are very large, andneed to be reduced.

The second model includes a reinforcement ofthe TF coil, by placing a strap structure in theshadow of the outer leg of the TF coil. Thepurpose of the strap is to increase the structuralmaterial in the outer leg of the TF coil in order toincrease the moment of inertia of the structure,and to decrease the rotation (toroidal displace-ment) of the caps. The thickness of the structurein the shadow of the TF coil is the same as thethickness of the caps (which is the same as thethickness of the bucking cylinder). The structurein the shadow of the TF coil is connected to boththe outer leg of the TF coil and to the upper andlower caps. Its main function is to prevent rota-tion of the caps with respect to each other bydeveloping shear stresses in the structure in theshadow of the TF coil.

In the second model it is assumed that theupper and lower caps are attached to the buckingcylinder for both toroidal rotation and tension.The caps/straps help to both support the inplane

Table 6Comparison between the optimized poloidal field systems forthe third model (bending-free TF coil) and the fourth model(non-bending-free coil); BoB

Model 4Model 3Coil

−3.456PF1 3.5868.7620.907PF22.528PF3 2.588

−9.460 −12.514PF4−11.768−42.447PF5

20.903PF6 −17.394PF7 −17.2611.556

18.292PF8PF9 1.580


tension of the coil and provide rigidity in case ofTF coil failure (short of a TF coil followed bydump of the rest of the TF coil, which results inlarge bending stresses in the fault-coil as it tries tobecome circular).

For the second model, the stresses in the coilnear the top and bottom of the machine are small,less than 60 MPa; the stresses in the caps/straps inthe top and bottom of the tokamak have practi-cally disappeared. The caps could be thereforethinned out in this region of low stresses. How-ever, the cap thickness has been kept constantbecause making it thinner would result in largerrotation of the outer legs of the TF coil, andtherefore larger stresses in that region. The capscould be made thinner, however, by making thecaps with truss elements, or by placing structurein between TF coils in the region underneath thecaps.

Similarly, the stresses in the region of the outerleg of the TF coil where the caps stop and thestraps begin (and where the stresses were 1.4 GPain the previous model), are reduced to about 600MPa. However, the peak stresses in the cap/strapstructure in the shadow of the TF coil are about2 GPa, in the region close to the caps stop.

In the second model, the region of large stressesis limited to the structure in the shadow of the TFcoil, adjacent to the region where the cap stopsand the strap begins. It is necessary therefore tofurther reinforce this section of the structure, as itis done in the third and fourth models. Thethickness of the structure in the shadow of theouter leg has been further increased, as shown inFigs. 6 and 7. For the model shown in Fig. 6, theheight of the caps is such that the inner compo-nents of the reactor can be removed by simpleradial motions, but does not leave space forplumbing for the blanket/shield and first wall. Afourth model, shown in Fig. 7 was developed witheven taller ports to accommodate the plumbing.

The stresses in the external coil structure for thethird model are acceptable with the exception of asmall region in the inboard section of the struc-ture in the outer leg, next to the place of disconti-nuity of the cap/strap structure. The stresses inthis region are in the order of 1.2 GPa (downfrom 2 GPa). This is still due to the large displace-

ment of the caps. If the port height were to beincreased, the radius of curvature needed to de-form the coil and the structure in the shadow ofthe coil would decrease, and so would the stresses.However, this would occur at the expense ofincreased displacement (rotation).

Fig. 8 shows the stresses in the TF coil for thefourth model. Again, the stresses peak in theregion above and below the coil, but the peakstresses are only 350 MPa. The stresses in theexternal coil structure (cap, bucking cylinder andcoil-shadow) are shown in Fig. 9. The stresseshave decreased to about 750 MPa in the structure.

The edge of the caps moves similarly to astrain-controlled deformation, and therefore in-creasing the height of the port would decrease thestresses, for constant thickness. The thickness ofthe strap and cap region near the strap wasreinforced in order to decrease the displacements.The corresponding displacements are shown inFigs. 10–12. These figures show the vertical, ra-dial and toroidal displacements for the case of thefourth model and EoB conditions. The displace-ments are significantly larger than for the thirdmodel, since increasing the height of the port(even though it decreases the stresses), increasesthe displacements of the cap and TF coil.

The device has very little radial displacement,less than 1 cm. The maximum vertical displace-ment is about 1 cm, occurring at the top andbottom of the machine. This is consistent with acoil that is shorter than a bending free coil. In thiscase, the coil wants to increase in height anddecrease its outer major radius (to become morelike a bending-free coil). The fact that the outerradius experiences a radial displacement of lessthan 1 mm (at the midplane) is due to the largestructure around the coil.

The toroidal displacements are substantiallylarger. For the case analyzed, the worst displace-ment is 0.05 m at the discontinuity region betweenthe caps and the strap, at the outer leg of themagnet. For comparison, for case (model 3) themaximum displacement was less than 0.04 m.There is no equipment in this region, since it hasbeen purposely left empty to allow for mainte-nance. Since maintenance will take place with themachine shut-down (and therefore with no elec-


Fig. 8. Stresses in the TF coil for the fourth model with 350 MPa maximum stress in the inboard side to the coil, top and bottom(EoB).

tromagnetic loads), it is not necessary to allow foradditional clearance for this operation.

3.5. Discussion

The support of the out-of-plane loads withoutstructure in the outer legs of the TF coils is madeespecially hard in the case analyzed because of thelarge distance between the poloidal field coils andthe plasma, which results in very large currents inthe divertor coils (PF5). The PF coil-to-plasmadistance is much larger than, for example, in

ITER, and as a consequence the currents in thecoils and the associated fields are large.

The aspect ratio of the plasma is high (A=4)compared with ITER, further increasing the rela-tive distance between the poloidal field coils andthe plasma.

In order to facilitate the design of a PF system,the toroidal field coil shape has been changedfrom a bending-free shape. The shape of the coilis compared in Fig. 13 with that of a bending freeshape, indicating that the coil half-height has beendecreased by about 1 m. Although the inplane


Fig. 9. Stresses in the external coil structure for the fourth model with 750 MPa maximum stress in the region where the cap/straphas discontinuity (EoB).

stresses are increased (due to bending), the out-of-plane shears have substantially decreased.

It has been determined that the decreasedheight increases the bending in the region wherethe vertical leg meets the rest of the coil, makingthe tension due to inplane loads across the regionof maximum binding (in the region between thevertical inner leg and the curved coil) vary from80–250 MPa. It is therefore possible to furtherdecrease the height of the toroidal field coil with-out compromising the TF coil or the superstruc-ture, and bring the poloidal field coils even closerto the plasma. This should be explored in the

future.The amount of material added to the structure,

even in the fourth model, is relatively small. Thethickness of the band in the outer leg of the TFcoil is comparable to that of the thickness of theTF coil. In addition, this structure can be used tosupport the outer PF coils.

It should be indicated that the calculationspresented below are smeared stresses. More workneeds to be done to determine stresses in individ-ual components. Of particular interest is thestresses in the insulation, which has not beendone. However, since the conductors are not sup-


Fig. 10. Vertical displacement for the fourth model with 0.01 m maximum displacement at the top and the bottom of the device.

posed to carry radial or out-of-plane loads, theinsulation stress should not be very large. How-ever, imperfections could result in larger stressesthat need to be addressed.

4. Poloidal field coils

In ARIES-RS all large PF coils are external tothe TF system and are superconducting usinginternally cooled, cable-in-conduit conductors ofternary Nb3Sn and NbTi, as in ITER [5]. Thereare small normal conducting internal coils for

providing plasma vertical stability. TernaryNb3Sn is favored over binary because it hashigher temperature margins even at the lowerfields [22]. NbTi is used in order to decrease thecost of the large ring coils which have relativelylow fields.

Several iterations were needed between the TFand PF designs in ARIES-RS, in order to come toan acceptable compromise. Increasing the outerleg of the TF coil increases, for a constant-tensioncoil, the height of the TF coil. This results in anincrease in the distance between the poloidal fieldcoils and the plasma, increasing the plasma shap-


Fig. 11. Radial displacement for the fourth model with less than 1 cm maximum displacement in the top/bottom of the coil.

ing difficulties. Indeed, it was found that for thesystem in Fig. 6, it was not possible to design arealistic poloidal field system [24].

Table 6 shows the coil currents for the optimalPF system for the third model (bending-free TFcoil), shown in Fig. 6, at Beginning of Burn(BoB). Very large currents are needed. The coilwith 40 MA, in particular, is very hard to design.This is due, of course, to the very large plasma-coil distance, making plasma shaping difficult.The largest current that could realistically bedriven in the coil system for the bending-free TFcoil was about 60% of the nominal currents.

Substantial reduction in coil currents wereneeded.

In order not to interfere with the maintenanceapproach, the solution of decreasing the coilheight, and increasing the number of coils, wasadopted. The TF coil is therefore no longer con-stant tension [23], but realistically (because of thefinite coil thickness and out-of-plane structure),the coils were never totally bending free [25].Several iterations were performed between the PFand TF designs, until a compromise between in-creased bending in the TF and increased PF cur-rents was reached. In this section, only the near


Fig. 12. Toroidal displacement for the fourth model with 0.05 m maximum displacement in the region where the cap stops beingtoroidally continuous.

final design is described. Table 6 shows the valueof the currents for the BoB of the optimized,non-bending-free TF coil.

It should be noted that the total current, mea-sured as MA-turns, is comparable for both cases.However, there is no need of a 40 MA current, asis the case of the third model (bending-free coil).

The PF-coil geometry is described in Table 7.The coil set is symmetric up-down, and only thevalues of the upper coils are given in the tables thatfollow. The center solenoid, however, has beensubdivided in such a manner that the first coil (PF1)

merges both the upper and lower coil, so insteadof having two PF1 coils, there is only one (PF1,UL).

Design allowables for the superconducting PFmagnets, similar to those assumed during theARIES study, are listed in Table 8 along withsuggested additional constraints on energy andpower-balance criteria for recovery from distur-bances. The ARIES-RS PF coils are more conser-vative than those of ITER as they are also designedfor energy margins \0.5 J cm3 and fractions ofcritical current in the well-cooled recovery regime.


Fig. 13. Comparison between the ARIES-RS TF coil and that of a bending-free coil.

The poloidal field system is designed usingmuch less aggressive assumptions than thetoroidal field system. The reason for this is thatthe rest of the nuclear island is less sensitive to theassumptions on the poloidal field coil than on thetoroidal field coil. The object is to demonstratefeasibility of this system, rather than perform acomplete optimization. The optimization of thePF system is limited to minimizing the supercon-dutor and copper, placing enough copper forquench protection in pure copper strands andoptimization of the coil location. The system wasself consistent, but resulted in coil that had a largefraction of He. Coils with large He fraction

should be decreased in size, resulting in a new setof currents and sizes. This final step was notperformed in this study, but it is expected that thedifferences between the case analyzed and thatwith lower helium fraction would not be verydifferent. This is due to the fact that the heliumfraction does not show up in the costing.

In Table 8, fc is the maximum allowable ratio ofconductor current to the critical current, ftr is themaximum ratio of conductor current to the cur-rent at the transition point between well-cooledand ill-cooled behavior during recovery from adisturbance, Tmax,dump is the maximum permissibletemperature in the winding pack following a coil


Table 7ARIES-RS PF-system winding-pack and plasma dimensions

NI (MAT)Coil R (m) Z (m) R1 (m) R2 (m) Z1 (m) Z2 (m)

0.91PF1, UL 1.60 0.00 1.36 1.84 14.910−0.9112.2141.82PF2, U 1.60 0.911.36 1.36 1.84

2.73PF3, U 1.60 2.27 1.36 1.84 1.82 5.0444.04PF4, U 1.60 3.39 1.36 1.84 2.73 14.5795.88PF5, U 12.5591.68 5.215.55 1.34 2.01

6.16 6.95PF6, U 17.6182.77 6.56 2.38 3.177.48PF7, U 4.50 7.09 4.11 4.89 6.70 17.2616.66PF8, U 24.4818.07 5.736.19 7.60 8.54

4.07 4.47PF9, U 10.27 4.27 10.07 4.52810.47

dump (K), Tmargin is the minimum permissibletemperature difference between the local conduc-tor temperature and the current-sharing tempera-ture during a scenario (K), and Emargin is theminimum permissible local energy margin, definedas the volumetric energy needed to be deposited inthe conductor metal that would heat the localhelium reservoir to the current-sharing tempera-ture (J cm3).

The fractions of critical current are calculatedat each point in a scenario including the effects ofstrain in the superconductor and heating of thesupercritical helium coolant. Critical fractions andmargins are also recalculated after disruption sim-ulations, because it is assumed that disruptionsare capable of occurring at any time during acycle.

Forced-flow, internally cooled cables have tworecovery regimes. In the well-cooled regime, al-most the entire local enthalpy of helium in theconduit at constant density, between the bathtemperature and the current-sharing temperature,

is available for conductor recovery. In theARIES-RS design it has been assumed that theconductor is in the well cooled regime.

4.1. Poloidal field scenarios

In Table 9, the parameters for the static equi-libria for a pseudo-scenario are presented. Thedifferent times of the scenario are: back bias (BB),end of ramp (EOR, low b, full current), beginningof burn (BoB), end of burn (EoB), and end ofpulse (EoP, which occurs at full current, low b atthe end of the cycle). The currents shown in Table9 were used to determine the stresses, to designthe poloidal field system, and to calculate the AClosses. The winding pack dimensions are given inTable 7. The winding characteristics of the PFsystem are described in Table 10.

The conductor type of the different coils hasbeen optimally determined for each coil. Thehigher field coils are made of ternary tin. Thelarge ring coils are made of NbTi. The number ofpancakes per coil, np, and the number of layers inthe coil, nl, are determined in order to limit thesize of the conductor current in the different coilsto about 70 kA. This choice decreases the amountof stabilizer required for quench protection andminimizes the cost of the winding process. Largercurrents would require expensive bus and powersupplies. The number of strands in the supercon-ductor, ns, has been arbitrarily set to 1125. Dstrand

is the corresponding strand diameter. Pw is thewetted perimeter of the strands. It has been as-sumed that the Resistivity Resistance Ratio

Table 8PF-magnet constraints used for ARIES-RS design

Vterminal (kV) 2066Icond (kA)

Bmax (T) 14sTresca membrane (MPa) 600fc 0.8Tmax,dump (K) 150

2.5Tmargin (K)1Emargin (J cm−3)

ftr 0.7


Table 9Currents in PF coils from ARIES-RS static equilibrium used for stress and scenario calculations (in MA)

EoPBB EoR BoB EoB

846.000Time (s) 0.000 266.000 446.000 1026.000

−3.35 8.11PF1, UL −3.45−14.91 7.8112.21PF2, U −6.49 11.92 8.76 8.86

2.65PF3, U −5.04 2.28 2.52 2.65−12.40 −14.26PF4, U −12.51−3.34 −14.57−11.72PF5, U −1.43 −12.55 −11.76 −12.43

−17.46−17.34PF6, U −17.39−2.20 −17.61−17.26 −17.20 −15.17PF7, U −3.24 −15.33

24.4818.29PF8, U 18.29−2.86 24.461.58PF9, U 0.94 −4.52 1.58 −4.52

−10.41Iplasma 0.00 −10.41 −10.41 −10.41

49.58Flux linkage 50.45−59.05 55.7453.19

(RRR) is 100 for the copper stabilizer for all thePF coils. The filament diameter, Df,eff, is 10 mm forthe ternary tin superconductor, and 5 mm for theNbTi. The effective cross-conductance length,Lp,eff, is 20 mm for all the conductors.

fCu,str represents the ratio of pure copper strandsin the conductor to the total number of strands inthe conductor. fCu,str is determined for each coilseparately. All the PF coils have a high percentageof strands that are made of pure copper, forquench protection. It is assumed that the com-posite strands have a Cu/non-Cu ratio of 1:1.

The corresponding conductor geometry isshown in Table 11. wconduit and hconduit are thewidth and height of the conduit for each of the PFcoils. ACu,cond, AnonCu, and Ass,cond correspond tothe copper, non-copper (i.e. superconductor), andsteel cross-sectional areas of the conductors of eachPF coil. AHe,cond is the corresponding He cross-sec-tional area, and Acondenv is the overall area of theconductor, including the insulation.

The current and voltage maxima for each of thePF coils is shown in Table 12. The size of eachconductor is adjusted in order to keep the maxi-mum current in each coil below 66 kA. Since themachine is steady state, the start-up and shut-downprocesses can be very slow, and the voltages can bemade arbitrarily low. It has been assumed that theramp-up and ramp-down are 4 min, and theheating/cooling phases are 3 min, unnecessarilyshort (but the resulting voltages and powers are

small even then).The peak values of the magnetic field in each coil

for the pseudo-scenario are shown in Table 13. Thepeak field on PF5 coils is about 14 T, and that ofPF6 is about 11 T. These coil sets are mainlyresponsible for shaping the plasma, and due to thelarge plasma-coil distance carries large currents.

Only PF9 coils have substantially lower fields,and use NbTI in order to minimize the cost of thesystem.

The performance of the PF system is shown inTable 14. fc varies throughout the coil and duringthe scenario. The value listed in Table 14 is themaximum value for a given coil. This ratio shouldbe less than about 0.7–0.8. The highest fraction ofcritical occurs for the divertor coils. Qmax¦ is themaximum critical heat flux from the conductor tothe helium bath. ftr is the ratio between the currentand the transition current beyond which the criticalenergy decreases dramatically. DThead is the maxi-mum temperature excursion between the operatingtemperature (which varies during the pulse) andthe current sharing temperature where the conduc-tor current exceeds the critical current. Similarly,Emar,min is the energy needed to raise the tempera-ture of the helium bath to the current sharingtemperature. fJprot is the ratio between the actualconductor current and that resulting in a finalconductor temperature of 150 K (after a 20 kVdump). Bmax is the maximum magnetic field in thecoils.


Table 10ARIES-RS PF conductor characteristics

fCu,strCoil SCtype np nl ns Dstrand (mm) Pw (mm) Df,eff (mm)

10PF1, Ul Ternary tin 30 8 1125 1.704 0.675.0160.7010PF2, U Ternary tin 19 3.89010 1125 1.321

3.908 10PF3, U Ternary tin 0.7510 10 1125 1.3270.7210PF4, U Ternary tin 27 3.90810 1125 1.327

10PF5, U Ternary tin 14 14 1125 1.276 3.758 0.5910PF6, U Ternary tin 16 16 1125 1.324 3.899 0.6710PF7, U Ternary tin 0.7516 3.89916 1125 1.324

3.899 10 0.70PF8, U Ternary tin 19 19 1125 1.3243.899 5PF9, U NbTi 8 8 1125 0.751.324

The average hoop, axial, and Tresca membranestresses in the conductor conduit were calculatedat each discrete point in the scenario. The totalload on each coil was also calculated followingeither current- or flux-conserving disruptions oc-curring at any time during the scenario. The PFmagnets are self-supporting against tensile loads.The amount of structure was added in order tokeep the stresses below permissable levels.

As a conclusion, the PF system is very robust,and the cryogenic losses from the normal opera-tion and the two model of disruption analyzed,are manageable. The coils can absorb the energyadiabatically.

5. Failure analysis

Faults in both the toroidal and poloidal fieldsystems were evaluated. The different fault scenar-ios involved [26,27]:� plasma disruption� a short across the leads of a toroidal field coil

followed by the dump of the rest of thetoroidal field coils

� fault of poloidal field coils (shorts during nor-mal and off-normal events).

5.1. Structural considerations

For the toroidal field coil analysis, it was deter-mined that failure of the poloidal field system ordisruption does not increase the out-of-planeloads substantially. It was found, however, that

the worst case scenario for the TF coil was duringa case with a coil shorted across its leads, fol-lowed with a dump of the other coils. In this case,the current in the shorted coil increases by mutualinductance, while the current in the other coilsdecrease. The system of coils therefore cease tobehave as a toroidal system (with D-shape coils asthe bending free structure). The over-driven coiltends to become circular. In the case of ARIES-II/-IV, ARIES-II and PULSAR, the coil hassmall capability of carrying bending (since it ismade of thin shells), and in the absence of a largesupport structure (present in ITER and inARIES-RS), the coil deformations will be verylarge. For the case of ARIES-RS, the large struc-ture to provide support for the out-of-plane loadsin the outer leg without the need of intercoilstructure restrains deformation. The calculationassumes a case with a failed TF with a current of120% nominal (shorted coil) coupled with the restof the coils operating at 80% nominal current(after a dump). These numbers are arbitrary, andhave been used only for providing guidance to theextent of the problem. It has been calculated (byscaling from ITER’s results) [28] that the defor-mation from this fault results in a radial deforma-tion of 2 cm for the shorted coil. This results in astrain of less than 0.2%.

For the case of the poloidal field faults, thefollowing cases have been analyzed: normal sce-nario, current-conserving disruption at any pointduring the scenario and short across the leads in acoil followed by disruption. The hoop stresses andthe vertical loads in the PF coils were monitored


Table 11ARIES-RS winding geometry

Acondenv (mm2)Coil wconduit (mm) hconduit (mm) ACu,cond (mm2) AnonCu (mm2) Ass,cond (mm2) AHe,cond (mm2)

2051PF1, U, I 59.3 59.3 428 84.6 3765120023611079PF2, U 46.6 81946.6 393 69.4

824 1323PF3, U 46.8 238146.8 204 29.123811089PF4, U 46.8 82446.8 401 65.3

575PF5, U 45.1 45.1 686 177.0 779 2218852PF6, U 46.7 46.7 581 116.2 822 2371929PF7, U 237146.7 82246.7 542 77.4

822 495PF8, U 46.7 46.7 922 2371131.71239PF9, U 46.7 46.7 263 46.4 2371822

to determine their largest value and to comparethem with the normal scenario. In addition, theout-of-plane loads on the toroidal field coil werealso monitored. It was determined that the faultconditions did not result in substantially largerloads (increased by less than 10–20%), as calcu-lated by scaling results for TPX [29,30].

5.2. Quench protection requirements

A critical component of protection of the mag-nets is that of quench monitoring. Prompt, error-free detection of the beginning of a quenchdecreases the requirement for subsequent quenchprotection. For ARIES-RS, advanced quenchmonitoring techniques will be needed [21]. Re-sponse times less than 1 s would be required.

Assuming that the copper strands share thecurrent with the superconducting ones instantly,and assuming adiabaticity, then with a value ofSmax=11×1016 A2 s m2, the maximum time forresponse is on the order of 1.5 s. Using 32 circuits(two per coil) with a maximum voltage acrossterminals of 20 kV and a pack current of 51 kA,gives a maximum dumping power of about 1 GWper circuit, or about 30 GW for all circuits. Witha total energy of about 50 GJ, the dump time ison the order of 2 s. Therefore, external dumpalone, is marginal for quench protection.

A novel approach to the issues of the insulationis to use internal dump. If internal dump is usedin the case of magnet coil quench, the large dumpvoltages are avoided, and the insulation can bemade substantially thinner. This substantially in-

creases the allowed current in the copper duringquench. In addition, it simplifies the charging anddischarging of the coil, by minimizing the numberof leads to the magnet.

For ARIES-RS, is was assumed that the samesensors used for quench detection could be usedto create a large normal zone that would absorbthe stored magnetic energy in the magnet, elimi-nating the need of a dump circuit. Enlarging thenormal zone in the magnet prevents the burn-outphenomena that would happen if the only normalzone is created by propagation of the initialquenched region. More energy is dissipated in themagnet (3–4 times more energy). This energyneeds to be removed with a refrigerator after thequench, before resumption of operation.

In the PULSAR and ARIES designs, the cur-rent in the copper was allowed to increase byincreasing the number of circuits in the TF mag-net, by electrically subdividing each TF coil.Thus, those TF magnet designs had twice as manyelectrical circuits and number of coils, decreasingthe stored energy per circuit that needed to bedischarged. Also, very high discharge voltageswere required, on the order of 15–20 kV. Thedesign of the coil, specifically the issues associatedwith the conductor insulation [8], and the com-plexity of the system, are substantially simplifiedby the use of internal coil dumps.

For high performance Nb3Sn, quench protec-tion dominates the conductor cross section. Sincethe cost of the conductor is a substantial fractionof the total cost of the coil, the cost of the quenchprotection could dominate the cost of the TF coil.


Table 12Peak currents and voltages on PF coils

Vterm,max (kA)Coil Icond,max (kA) Vterm,min (kA)Icond,min (kA)

0.08PF1, U, I 34.9 −0.09−64.3−0.070.06PF2, U −35.166.1−0.03PF3, U 14.5 −27.5 0.03

0.01PF4, U 0.0 −55.0 −0.010.05 −0.05PF5, U −62.00.00.13PF6, U 0.0 −66.4 −0.13

−0.120.10PF7, U −66.40.0−7.7 0.92 −0.96PF8, U 66.4

−66.4 −0.09PF9, U 0.0923.2

A method to obviate this is by physically separat-ing the superconductor from the quench protec-tion, as shown in Fig. 2. For ARIES-RS, as in theARIES and PULSAR designs, the material forquench protection is co-wound with the supercon-ductor. In this manner, inexpensive material (i.e.pure copper strands) could be used for this pur-pose. If the quench protection is incorporatedwith the superconductor in the manufacturingprocess of the superconductor strands, the cost ofthe conductor could increase substantially.

There is added advantage to separate thequench protection and the superconductor. It ispossible to obtain high RRR value strands if thestrands do not contain any superconducting mate-rial. It is still necessary to coat the strands with amaterial that will help control sintering of thestrands at the superconductor reaction tempera-ture, but the bulk of the conductor will remainrelatively impurity free (a thin chromium coatingis being considered for this, which could diffuseinto the bulk of the strands at the reaction tem-peratures of the superconductor).

For ARIES-RS the dump-time is determined tozeroth order by the maximum voltage duringdischarge Vdischarge (to prevent breakdown in theinsulation) and the pack current Ipack, since thereactive power during the dump phase of theenergy is given by Ipack×Vdischarge. Increasing ei-ther one will result in faster dump.

For ARIES-RS, the reference design assumedVdischarge=20 kV and Ipack=55 kA. The dischargevoltage is limited by breakdown of the insulation,while the current is the result of a compromise

between conductor thickness, power supply re-quirements, and coil size minimization.

For internal dump, the sensors imbedded in thecoil for quench monitoring are used for quenchingthe magnet, mainly by resistive heating. The en-ergy margin of the superconductor in the toroidalfield coil is in the order of 0.1–0.2 J cm3. This isthe energy that is required in order to turn theconductor normal. For a minimum initial propa-gation zone of 5 cm, the number required is in theorder 1–2 J site−1.

ARIES-RS, due to its high performance re-quirements and to the copper segregation, haslarge quench propagation speeds. The fast quenchpropagation decreases the peak temperature dur-ing a quench, either because it is easy to deter-mine the presence of a quench earlier, or becauseof internal dump occuring in a larger volume. Inorder to decrease the peak-to-average temperatureexcursion in the ARIES-RS, a large number ofnormal zones would be desirable.

To limit the temperature to about 150–200 K,the maximum energy deposited at the location ofthe hot spot is about 250–300 MJ m3 of metal[13]. If a substantial fraction of the TF coil energyis to remain in the coil, then the volume of theheated zone is in the order of 100 m3. This is asubstantial fraction of the coil volume, allowingfor small differences between the peak and aver-age temperatures. Rough estimates indicate thatthe large number of normal zones and the fastpropagation of them should keep the peak toaverage temperature to about 5 [31]. In order todo so, only about half of the energy can be


Table 13Peak field flux (T) on PF coils

EoBBB EoR BoB EoP

4.610PF1, U, I 9.196 6.963 7.2804.5788.5934.791PF2, U 4.7558.751 8.245

7.462PF3, U 7.546 8.548 8.5987.4508.407PF4, U 5.695 9.2217 8.498 8.948

13.460 13.637PF5, U 13.5202.664 13.81611.744 11.700 11.282PF6, U 2.275 11.415

8.703PF7, U 2.000 8.299 8.723 8.2448.762 10.907PF8, U 8.7611.123 10.904

1.730 1.732PF9, U 0.922 4.7114.716

50Flux linkage −59 53 5549

dumped in the magnet. The other half needs to beextracted, but with substantially lower voltagesthan indicated above. Ten kV or less should besufficient.

More detailed calculations of the temperaturedistribution during dump needs to be performed.

The amount of energy required to initiate thedump is small. If every turn is quenched, thenabout 3 kJ are required to initiate a quench inevery turn in the toroidal field coil.

6. Conclusions

In this paper, the design features of the magnetfor a DEMO-type tokamak reactor have beenreported. Design features that minimize the costof the reactor have been explored, while at thesame time the maintenance of the non-permanentsystems in the reactor has been improved. Alter-native methods for supporting the TF coils havebeen analyzed, and one design with ports largeenough to provide radial access to modularizedblanket/shield modules has been preliminarilyevaluated. The number of these modules is thesame as the number of TF coils.

The design is much more aggressive than thatfor ITER or TPX. The purpose of the study wasto determine the outcome of a credible, albeitaggressive design. The design has current closer tocritical, a superconductor with properties closer to

short sample, strain sharing between the super-conductor and the structure, advanced quenchprotection, higher stresses than allowed in ITER(due to steady state operation), advanced TFcooling techniques, among other things. Time lim-itations prevented the group to redesign ARIES-RS using ITER-type technology, but this will bedone in the near future.

A structure has been designed that includes acap, a strap outside the coil and the buckingcylinder. The structure is poloidally continuous,and should also be toroidally continuous. Thecaps will be keyed together, and they will beattached to the bucking cylinder with an arrange-ment similar to that in the ITER design (hole andpin).

Although an entire scenario was investigated,only normal load conditions were analyzed. Ab-normal conditions, representing single or multipleTF coil failures and disruptions, were scaled fromfailure studies in TPX and ITER. Additionalstudies need to be performed for off-normalevents, such as effect of disruptions on the TF coiland quench behavior.

Acknowledgements

This work was supported by US Department ofEnergy, Office of Fusion Energy Sciences.


Table 14ARIES-RS PF superconductor performance

Bmax (T)Coil fc Qmax¦ (W cm−2) ftr DThead (K) Emar,min (J cc−1) fJprot Ecoil (MJ)

0.302PF1, U, I 0.644 0.0142 0.341 6.547 8.047 9.80.4379.30.144PF2, U 0.709 0.0397 0.3510.405 6.745 4.237

0.310 0.100PF3, U 0.669 8.60.0045 0.212 7.023 11.8469.20.090PF4, U 0.736 0.0286 0.3670.351 6.832 16.546

0.084PF5, U 0.722 0.0624 0.455 4.296 4.298 0.253 13.80.181PF6, U 0.788 0.0536 0.427 5.456 10.459 0.558 11.70.211PF7, U 0.740 8.70.0419 0.7060.355 6.949 14.505

10.9PF8, U 0.811 0.0480 0.329 5.590 0.951 0.694 1.1310.176PF9, U 0.722 0.0249 0.693 2.497 10.883 0.292 4.7

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