design optimisation of the iter tf coil case and structures

Fusion Engineering and Design 75–79 (2005) 207–213

Design optimisation of the ITER TF coil case and structures�

Marco Ferraria,∗, Pietro Barabaschia, Cornelis T.J. Jonga, Yuri Krivchenkovb,Reinhard K. Maixc, Neil Mitchelld

a ITER International Team, Boltzmannstr. 2, D-85748 Garching, Germanyb Efremov Institute, Metallostroy, St. Petersburg 196641, Russia

c ITER VHTP, ATI Atominstitut Wien, Stadionallee 2, A-1020 Vienna, Austriad ITER International Team, 801-1 Mukouyama, Naka-machi, Naka-gun, Ibaraki-ken 311-0193, Japan

Abstract

The design status of the main ITER magnet structures and the toroidal field (TF) coil cases, is described. In the last 3 years,the major geometry of the TF coil case has remained unchanged, but significant improvements have been made to the structuralsupport, to the material options and main fabrication steps. The lack of access for inspection and repair means a high reliability isrequired, and appropriate codes and standards to control the manufacturing have been defined. The structural assessment showsthat the design and operational criteria can be satisfied with reasonable inspection criteria.© 2005 Elsevier B.V. All rights reserved.

Keywords: ITER; Toroidal field coil case; Intercoil structures

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. Introduction

The magnet system of ITER consists of 18 toroidaleld (TF) coils, a six-module central solenoid (CS),

� Assignment of copyright is subject to the provision of the ITERDA Agreement concluded among the European Atomic Energyommunity, Japan, the Russian Federation and the United States on1 July 1992, which continue to apply during the ITER Transitionalrrangements.∗ Corresponding author. Present address: Superconducting CoilsStructures Division, ITER Transitional Arrangements, Garching

oint Work Site, c/o Max-Planck-Institut fur Plasmaphysik, Boltz-annstr. 2, D-85748 Garching, Germany. Tel.: +49 89 3299 4147;

ax: +49 89 3299 4313.E-mail address: [email protected] (M. Ferrari).

RL: www.iter.org.

six poloidal field (PF) coils and three sets of correccoils. A TF coil is composed of a winding pack encloin a case. Wedged together in a torus and linked bouter intercoil structure (OIS), the TF cases aremain structural components of the magnet system[1].

Among the main optimisation items, the selecof a radial (instead of toroidal) insertion procedurewinding pack into the case allows a better matchof the winding pack at the critical front nose intface and reduces weld distortion. The OIS has bgreatly reduced in mass, allowing the use of a simwelded/forged structure, by the adoption of a “fricttype” shear joint which links the coils and providesnecessary electrical insulation barrier. Adjustmenthe layout of the inner poloidal keys have reduced p

920-3796/$ – see front matter © 2005 Elsevier B.V. All rights reserved.doi:10.1016/j.fusengdes.2005.06.266

208 M. Ferrari et al. / Fusion Engineering and Design 75–79 (2005) 207–213

stress levels by almost a factor of 2. The reference sys-tem of rings for pre-compressing the inner legs of theTF coils has been confirmed as uniaxial glass fibre andthe design has been modified to allow re-tighteningof the rings without the need to remove the centralsolenoid. For the fabrication, the main components (theTF coil case sub-assemblies before winding pack inser-tion and subsequent closure welding) have been zonedinto three material grades of austenitic stainless steel,limiting the volume of expensive highest quality steel.

Extensive stress analysis of the TF structures hasbeen performed[2]. The ITER magnets operate inunusual structural conditions, at cryogenic tempera-tures and with high cyclic loading. A dedicated designcode has been derived, based on ASME Section III[3] and API579[4]. This code allows advantage to betaken of the high yield stresses at low temperature, butapplies more conservative limits than ASME SectionIII to secondary stresses. The lack of in-service inspec-tion means that a defect-based fracture/fatigue assess-ment is required, derived from API579, with acceptabledetection levels an integral part of the manufacturingNDT specifications.

Structural assessments based on the code confirmthe acceptability of the predicted stress levels and theircompatibility with the proposed manufacturing and QAprocedures.

2. TF coil case main sub-assemblies

twot aveb howni eree com-p eter-m ion,t loca-t ingt thep eldsa ions,w ort).

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a very close matching as both surfaces are flat, (ii)closure welding, (iii) filling and impregnation of theassembly gap between the winding pack and case and(iv) final machining of external surfaces. In this way,the risk of both direct damage and stress concentrationinduced by closure welding on winding pack is virtu-ally eliminated and the presence of a weak layer of fillerat the load-bearing nose interface is avoided.

The use of castings as basic elements of TF coil casesub-assembly (originally envisaged as an option on thelow stress outboard leg) has been excluded due to thelack of compatibility with the TF coil case fabricationscheme (e.g., NDT of closure welds), uncertainties inthe stress allowable values and, in particular, difficul-ties in achieving an adequate level of defect detection(using ultrasonic inspection, X-rays being inadequatefor crack detection).

The reference design is, therefore, based only onthe use of plates and forgings as basic elements. Thematerial requirements vary with location depending onthe local stresses and to allow more cost-effective sup-ply, three material classes are defined inTable 1andallocated to the sub-units as shown inFig. 1.

.

The main TF coil case sub-assemblies andoroidal C-shaped symmetrical halves in 2001, heen replaced by two pairs of sub-assemblies as s

n Fig. 1, for a total weight of about 195 tonnes, whach pair consists of a U-shaped section and thelementing plate plasma side. This change has dined a new way to insert, along the radial direct

he winding pack into the case and the subsequention of long poloidal closure welds that are not joinhick sections (this benefit is partially off-set byresence of an additional two transverse closure wt the border between inboard and outboard leg reghere the thickness is high but the weld length is shThe new assembly procedure for the TF coil fo

ees the following main steps: (i) vertical insertionhe winding pack in the TF coil case positioned horizally onto the inboard leg (sub-assembly AU) achiev
Fig. 1. TF coil case sub-assemblies and material classes

M. Ferrari et al. / Fusion Engineering and Design 75–79 (2005) 207–213 209

Table 1Static requirements (guaranteed properties at 4 K) for TF coil casematerial classes (as forgings and plates)

Yield strength (MPa) Fracture toughness(MPa

√m)

Class-1 ≥1000 ≥200Class-2 ≥900a ≥200Class-3 ≥750 ≥150

a Yield strength≥290 MPa at 20◦C.

Typical candidate strengthened austenitic steels arebased on modified 316LN stainless steel with a high Ncontent (0.18–0.27 wt%) and in some cases a high Mncontent (5.5–11.0 wt%). The class-1 material requires,in particular, electro-slag refinement to achieve thespecified performance, and its use is kept to a mini-mum.

Fabrication of each single main sub-assembly (seeFig. 1) is obtained through butt-welding of basic seg-ments, where the first pass may be made by electronbeam welding (EBW) due to the low distortion leveland the high deposition rate. The EBW pass has to belimited to a thickness of about 50 mm due to instabilityproblems of the beam and gas bubbles caused by thehigh N content of the base metal. The remaining passesmay be completed either by submerged arc welding(SAW), which has a high deposition rate, but must becarried out only in one position (so that the weld is hor-izontal, requiring heavy handling equipment to movethe case sections), or by narrow gap gas tungsten arcwelding (NG-GTAW).

3. Inner intercoil structures (IIS) andpre-compression rings

The TF coil inboard legs are wedged all along theirside walls in operation and are structurally linked atthe top and the bottom by means of two upper andtwo lower pre-compression rings formed from unidi-r fourl coils cen-t esek omel cen-t andr .

Fig. 2. Inner intercoil structures and pre-compression rings.

In the earlier 2001 design, four full-length key slotswere present, each one hosting four M150 cylindri-cal keys. A stress concentration of over 1000 MPa wasoccurring in the first innermost key slot with the laterkeys not working effectively. The length of the keyshas been optimised to distribute the shear loads moreuniformly, so that the innermost key slots are shorter(50% and 75% of the coil thickness, respectively) anddo not penetrate to the inner bore. Besides reduction ofthe peak stress value, this arrangement also keeps thepeak stress region away from the final closure welds ofthe TF coil case.

The key slots are machined oversize and lined withtwo half-shells made of steel. The shells are machinedbased on a precise survey of the key slots after assemblyso as to absorb any mismatch between the coils andprovide an inner perfectly circular slot for the key. Theyalso provide the low voltage insulation between thecoils. The keys are segmented into four units alongtheir length (so that the first keyway contains 2 unitsand the second contains 3 units).

ectional bonded glass fibre and by four upper andower sets of poloidal shear keys, i.e., the inner intertructures, arranged in keyways normal to the coilreline (Fig. 2). When the TF coils are energized, theys tend to de-wedge radially outward and becoose. This effect increases the local stress conrations on the shear key insulation and key sloteduces substantially the life of these components


The use of pre-compression rings is foreseen, asshown inFig. 2, to increase the initial pre-load on theshear keys to maintain zero-gap during operation onthe shear keys, close to the ends of the key slot and toguarantee a gap <0.5 mm on central shear keys.

Two rings are installed, both at the top and at thebottom curved regions of the TF coil during initialassembly of the machine and each pair of rings is pre-loaded with 6 Inconel-718 bolts per TF coil, to providea radially inward pre-load of 35 MN at 4 K (per ringpair) per TF coil. The radial pre-load is applied to thickpre-compression flanges integral to the TF coil case.

The required pre-compression force has to beachieved within a limited space, particularly, alongthe radial direction. This space constraint limits themaximum local area available for the rings to about0.22 mm2 total. This, in turn, means that the hoop stressin the rings must be typically 500 MPa, applied at roomtemperature by a radial extension of the rings of about20 mm. The manufacturing procedure for the rings isbased on a glass fibre unidirectional epoxy compos-ite wound in the hoop (toroidal) direction with eachring made of a number of layers in the radial direc-tion. This is expected to offer a high strength but lowstiffness, making the pre-tensioning less sensitive tosettling effects in operation. The thermal expansionis lower than steel (resulting in a loss of pre-tension)but can be minimised by careful selection of the glasstype.

The fibre is coated with resin, wound and then cured.T inp sionl firstf spe-c ofj nut( tene g as

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rings and act as shear panels in combination with theTF coil cases.

The upper and lower OIS are located, respectively,above the upper ports and below the divertor ports ofthe vacuum vessel. The upper and lower intermedi-ate OIS are located, respectively, above and below theequatorial ports.

There are two types of OIS. The first one, used in theupper and lower OIS, is based on box structures consist-ing of two main shear panels linked midway betweenthe coils by insulated expansion bolts, which are ori-ented along the toroidal direction. These expansionbolts (the largest one is M120) withstand simultane-ously toroidal tension and shear loads.

The second type, used in the upper and lower inter-mediate OIS, is a “friction-joint” based on a singleshear panel, with a thickness of 100 mm, connectedto the side wall of the case through a Y-shaped flange(Fig. 3). In the earlier design[1], a massive box typestructure linked by bolts and keys was also an option.For several reasons (opening of the shear key slotsdue to lack of space for bolting, the cost of fabricating

he assembly of the rings occurs without the CSlace, but it will be necessary to check the ten

evel of the bolts, and retighten them, after theew thousand plasma pulses. This is achieved withial nut-tensioners[5] each equipped with a numberack-bolts, instead of a “classical” single large studFig. 2). There is sufficient access and room to retighach single jack-bolt, without CS removal, by usintandard torque wrench.

. Outer intercoil structures

In the outboard region, the out-of-plane supporovided by four sets of outer intercoil structures inrated with the TF coil cases and positioned arounerimeter within the constraints provided by the acucts to the vacuum vessel. The OIS form four toro
Fig. 3. Friction-joint intermediate OIS.


a massive structure without unacceptable defects andhigher stress concentrations), this option has beendropped.

In the “friction-joint”, at the top and bottom re-entrant corners where the Y-shaped flange connects tothe case sidewall, a large fillet radius is provided tomitigate stress concentration effects. The shear loadtransmission across the insulated joint in the panel isaccomplished by multiple-finger friction-joints, whichare welded to the two adjacent shear panels. Thefriction-joints are pre-loaded by bolts acting on six fin-gers separated by G10 insulated layers. This multiplefinger arrangement allows to multiply by a factor of5 the number of friction interfaces and therefore theshear load capability of the joint.

The friction-joint panel is assembled on site, oncethe TF coils are in position, by a weld on each side tothe shear panel extension attached to the TF coil cases.A potential issue is shrinkage when this in situ weldis performed. This shrinkage is expected to reduce thetoroidal dimension of the OIS. This toroidal shrink-age is accommodated by temporarily releasing the boltpressure to allow some sliding between the fingers ofthe joint. Shrinkage can also induce bending deforma-tions, which can be controlled only if the weld is donefrom both sides. Access to the inside of the shear panelis, therefore, essential and this requirement has beensatisfied by a new assembly procedure and sizes forthe vacuum vessel ports.

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less extensive than at room temperature and above.Secondary stresses are more significant and have alower allowable than at higher temperatures and (iii)in-service inspection of the magnets is not possible, andoperational monitoring of structural behaviour (dis-placements, for example) is insensitive. Leak-before-break is not generally a design option and in somecases, fast fracture may occur. Defect assessmentbased on component inspection during manufactur-ing is, therefore, required. For quality assurance, thispractically means combining ASME Section III[3]for design with Section XI (in-service inspection) orAPI 579 (fitness-for-service)[4] at the manufacturingstage.

The maximum allowable primary Tresca membranestress in the material (base metal), is defined as 2/3 ofthe 0.2% yield stress. In normal operation, the allow-able value is increased by a factor of 1.3 for combinedmembrane and bending stresses and 1.5 (i.e. up to yield)for combined primary and secondary. For welds, in sec-tions characteristic of the coil cases (i.e. a thicknessover 150 mm), these two values are decreased by a fac-tor of 0.8.

A fast fracture assessment has been performed. Thedesign stress intensity factor,Km should be comparedwith the fracture toughness,KIC at the design temper-ature.Km has been calculated using the expression (innormal operation):

K = Yσ(π a)1/2 <KIC

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. Design criteria, codes and standards

The structural design criteria[6] have been deriveo cover some special features of the ITER magnd structures that are not adequately coveredingle existing design code (the ASME pressureel code being a well-known example). Theseures are briefly summarised: (i) the magnets open the range 77–4 K, with most operational life cloo 4 K. Compared to high temperature componeield and ultimate strength are increased but fracoughness remains approximately similar. Fractuelatively more important as a failure mode than pic yielding (and safety factors are adjusted accngly), (ii) yield stress is increased well aboveypical level of residual fabrication stresses. Limng of stress peaks through local plasticity is m

m1.5

hereσ is the maximum principal tensile stress,Y thetress intensity factor anda is the crack size, calculatncluding growth due to fatigue effects. The facto.5 is a safety factor. Safety factors are also applie

he initial crack size and the rate of crack growth.There are two methods of fatigue assessment

rst is material-based and uses SN fatigue life cuo establish the characteristics of a componenthen applies a safety factor to define the allowife. The second is defect-based and uses linearic fracture mechanics to predict crack growth fromre-existing defect. Both procedures must generalpplied where there is the possibility of manufactuefects. The fatigue life (SN) procedure is applican its own where component testing can ensure ccteristic defects are included in the specimens usstablish the SN curve, or to highly localised st


concentrations where the absence of defects can beassured by detailed inspection.

Inspection of components, especially welds, bymethods that can reliably detect defects is, therefore,an essential part of the manufacturing process for mostcomponents. The main method is ultrasonic inspec-tion, to ASME inspection standards. The defect sizeto be used for structural assessment is determined bythe agreed accept/repair level of the ultrasonic signal,interpreted to a defect size through the use of calibra-tion blocks according to ASME XI.

6. Structural assessment

The success of the design optimisation and the appli-cation of the different fatigue criteria can be illustratedby a structural assessment of three regions, the poloidalbutt welds at the top and bottom of the inner straight leg,the outer friction-joint panel weld to the TF coil case,and the inner poloidal key slots. The TF coil case lifeis specified as 30,000 full power (15 MA DT plasma)pulses. Main results of the finite element stress analysesare shown inFigs. 4 and 5. Fig. 5 shows stress distri-bution in lower key slots at end-of-burn with 433 MPamaximum stress in the second slot.

6.1. LEFM assessment, TF back plate lowercurved region

riald ssi fetyfc t inbp

6a

upt ERm als berop DTs

Fig. 4. Lower intermediate OIS maximum cyclic stress.

6.3. SN assessment, third inner poloidal key slot

Stress range 358 MPa with maximum princi-pal stress 402 MPa (end-of-burn), triaxial stressesneglected, class-2 austenitic forged steel, SN data

Fig. 5. Maximum principal stress in IIS key slots.

Class-2 austenitic forged steel, ITER mateatabase for da/dN at 4 K, estimated residual stre

n base metal 50 MPa and in weld 200 MPa, saactors of 2 on number of cycles, 1.5 onKIC, 2 onrack area≥ maximum possible subsurface defecackplate to define NDT sensitivity = 190 mm2 and inoloidal closure weld 240 mm2.

.2. LEFM assessment, lower OIS friction panelttachment at location A

Cyclic stress = 230 MPa with maximum stresso 505 MPa, class-1 austenitic forged steel, ITaterial database for da/dN at 4 K, estimated residu

tress 50 MPa, safety factors of 2 on the numf cycles, 1.5 onKIC, 2 on crack area≥ maximumossible subsurface defect in weld to define Nensitivity = 120 mm2.


ASME Section III, Division 1, Appendix 1 at roomtemperature, SN curve contains mean stress correc-tion with Goodman and safety factors (2 on stresslimiting) ≥ maximum allowable stress range to sustain30,000 cycles according to SN curve is 690 MPa. Inall three cases, the static stresses are acceptable (theallowable direct stress is 2/3× 900 = 600 MPa). In thefirst two cases, the required defect detection level iscomfortably above the minimum that can be expectedin austenitic steels (about 25 mm2 according to ITERR&D) and in the third case the cyclic stresses are underhalf the allowable limit.

7. Conclusions

Design optimisation of the ITER TF coil cases andassociated structures has been successful to reducemanufacturing requirements on the coils (in termsof component thicknesses and NDT inspection crite-ria) while satisfying a rigorous design code suitablefor the operation of large metallic structures underintense fatigue conditions. Proposed manufacturingroutes (segmentation into basic forged units and weld-ing methods) have been defined although, freedom isleft to individual manufacturers to further optimisethese as long as overall standards and tolerances aremaintained.

Acknowledgements

This paper has been prepared as part of workundertaken within the framework of ITER transitionalarrangements (ITA). These are conducted by the par-ticipants: the European Atomic Energy Community,Japan, the People’s Republic of China, the Repub-lic of Korea, the Russian Federation and the UnitedStates of America, under the auspices of the Interna-tional Atomic Energy Agency. The views and opinionsexpressed, herein, do not necessarily reflect those ofthe participants to the ITA, the IAEA or any agencythereof. Dissemination of the information in this paperis governed by the applicable terms of the former ITEREDA Agreement.

References

[1] ITER DDD, N 11 DDD 178 04-06-04 R 0.1, EngineeringDescription, June 2004.

[2] ITER DDD, N 11 DDD 115 01-06-27 R 0.2, Structural Analysis,August 2004.

[3] ASME Boiler And Pressure Vessel Code, Section III, 2001 edi-tion.

[4] API Recommended Practice 579, Fitness-for-Service, first ed.,January 2000.

[5] Superbolt® Catalog by P&S Vorspannsysteme AG (St. Gal-lenkappel, Switzerland), 2001.

[6] ITER Magnet Structural Design Criteria, DRG1 Annex, August2004.

design optimisation of the iter tf coil case and structures

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