Conﬁguration and engineering design of the ARIES-RS tokamak power plantaries.ucsd.edu/LIB/REPORT/JOURNAL/FED/97-tillack.pdf · 2001-07-09 · M.S. Tillack et al.:Fusion Engineering

Fusion Engineering and Design 38 (1997) 87–113

Configuration and engineering design of the ARIES-RStokamak power plant

M.S. Tillack *, S. Malang, L. Waganer, X.R. Wang, D.K. Sze, L. El-Guebaly,C.P.C. Wong, J.A. Crowell, T.K. Mau, L. Bromberg, The ARIES Team

Uni6ersity of California, San Diego, Mail Code 0417, La Jolla, CA 92093-0417, USA

Abstract

ARIES-RS is a conceptual design study which has examined the potential of an advanced tokamak-based powerplant to compete with future energy sources and play a significant role in the future energy market. The design is a1000 MWe, DT-burning fusion power plant based on the reversed-shear tokamak mode of plasma operation, andusing moderately advanced engineering concepts such as lithium-cooled vanadium-alloy plasma-facing components.A steady-state reversed shear tokamak currently appears to offer the best combination of good economic performanceand physics credibility for a tokamak-based power plant. The ARIES-RS engineering design process emphasized theattainment of the top-level mission requirements developed in the early part of the study in a collaborative effortbetween the ARIES Team and representatives from U.S. electric utilities and industry. Major efforts were devoted todevelop a credible configuration that allows rapid removal of full sectors followed by disassembly in the hot cellsduring plant operation. This was adopted as the only practical means to meet availability goals. Use of an electricallyinsulating coating for the self-cooled blanket and divertor provides a wide design window and simplified design.Optimization of the shield, which is one of the larger cost items, significantly reduced the power core cost by usingferritic steel where the power density and radiation levels are low. An additional saving is made by radialsegmentation of the blanket, such that large segments can be reused. The overall tokamak configuration is describedhere, together with each of the major fusion power core components: the first-wall, blanket and shield; divertor;heating, current drive and fueling systems; and magnet systems. © 1997 Elsevier Science S.A.

Keywords: ARIES-RS; Reversed-shear tokamak; Plasma-facing components; Fusion power plant

1. Introduction

The goal of the ARIES-RS engineering effortwas to provide a self-consistent conceptual designwhich satisfies the mission requirements and pro-vides enough detail such that the major features

and R&D issues can be assessed. In the context ofa product development program, this is a prelimi-nary design concept which can be used to guideR&D programs to further improve the productand eventually lead to more detailed designs.

Prior to initiating the design phase of theARIES-RS study, various classes of engineeringdesign options were examined and their potentialto meet the power plant requirements assessed.

* Corresponding author. Tel.: +1 619 5347897, fax: +1619 5347716, e-mail: [email protected]

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

PII S0920 -3796 (97 )00113 -0

M.S. Tillack et al. / Fusion Engineering and Design 38 (1997) 87–11388

These options included material choices for thestructure, breeder and coolant. The design spacefor an attractive tokamak fusion power core isnot unlimited; previous studies have shown thatadvanced low-activation ferritic steel, vanadiumalloy, or SiC–SiC composites are the only viableestablished candidates for the primary in-vesselstructural material. In order to provide a frame-work for this assessment, these three materialclasses were used to distinguish engineering designchoices [1,2].

While there is no unique design concept guar-anteed to succeed, the ARIES Team chose toexamine a tokamak power plant operating with areversed-shear plasma and plasma-facing compo-nents using high-performance vanadium-alloystructures cooled by lithium. The major deviceparameters are summarized in Table 1. The blan-ket and shield designs were based on ARIES-II[3]. The choice to develop an existing design,rather than explore entirely new design concepts,was made partly in order to evolve a higher levelof detail and sophistication in the design, whichthen allows a much firmer basis for evaluating theattributes and a clearer understanding of the is-sues and R&D needs.

The engineering design process was driven bythe need to meet the top-level system require-ments. Meeting economic goals while maintainingthe safety and environmental attractiveness possi-ble with a fusion energy source was a majorthrust. For example, minimization of the use of

vanadium to only those areas where its use isnecessary for performance and safety reasonsmade a significant impact on the COE withoutcompromising the safety advantages. Another sig-nificant factor in the cost of electricity is the plantavailability, which is especially important withhigh capital-cost facilities such as a fusion plant.The layout of the power core, as well as thedetailed design of the internal components, wasstrongly influenced by the goal to reduce thedown-time for sector repairs and replacement toabout one month.

Attempts were made to integrate both physicsand engineering constraints in the development ofdesign solutions. For example, the requirementsfor both passive and active plasma stabilizationsystems were incorporated into the sectors fromthe initial strawman and revised as necessary dur-ing the design process. The divertor is anotherarea where efforts to develop physics solutionsconsistent with engineering limitations, and engi-neering solutions consistent with physics con-straints, were made throughout the design study.

In the following subsections, the overall designfeatures of the power core are summarized andthe principle features and conclusions from theindividual components described. These includethe first wall and blanket, radiation shielding,divertor, magnet systems, and current drive, heat-ing and fueling systems.

2. Fusion power core configuration

2.1. Power core layout

The configuration of the fusion power core isstrongly influenced by the availability goals. Theavailability of a power station has a large impactnot only on the cost of electricity but on theattractiveness of the entire plant, since utilitieswould not tolerate frequent shutdowns or ex-tended replacement periods. In any power plant,the availability is set by the reliability of thecomponents (mean time to failure, which sets thefrequency of unscheduled maintenance), their life-time (which sets the frequency of scheduledmaintenance), and the time required for replace-

Table 1Major parameters of the ARIES-RS tokamak power plant

Aspect ratio 4.005.52Major radius (m)1.38Minor plasma radius (m)

Plasma vertical elongation (x-point) 1.70Plasma current (MA) 11.32

7.98Toroidal field on axis (T)Toroidal beta 0.05Average neutron wall load (MW m−2) 3.96

2170Fusion power (MW)Total thermal power (MW) 2620Net electric power (MW) 1000

0.46Gross thermal conversion efficiency0.38Net plant efficiency

Mass power density (kWe tonne−1) 66.70

M.S. Tillack et al. / Fusion Engineering and Design 38 (1997) 87–113 89

Fig. 1. Cutaway of the ARIES-RS power core.

ment. In addition, it is prudent to ensure that themaintenance scheme for scheduled maintenancecan be used as much as possible for unscheduledmaintenance.

Any fusion power plant study tries to maximizecomponent reliability and lifetime by employingsuitable concepts and materials for the first wall,breeding blankets and divertor plates. Self-cooledlithium blankets with a vanadium alloy (V-4Cr-4Ti) as structural material were selected forARIES-RS, in part because they provide a goodbasis for high reliability. This blanket conceptenables a simple design with a small number ofcooling channels and low mechanical stresses inthe blanket structure. This structural materialshows promise for an exceptionally long life-timesince the allowable neutron fluence is probablylimited by irradiation creep only, and not byswelling or helium embrittlement [4].

In addition to these ingredients for high reli-ability and long lifetime, the ARIES-RS plant is

configured for quick replacement of the powercore components. The overall goal was a designallowing replacement of a sector in less than onemonth and parallel maintenance of all sectors tobe replaced during a shutdown. This scheme alsominimizes the need for in-situ repair of the com-ponents. The failed components are replaced andthe power plant is brought on line while the failedunit is removed to a hot cell for repairs. Suchquick replacement is not possible with a verticalreplacement scheme, as shown by previous studies[5–7]. Therefore, ARIES-RS is designed for hori-zontal insertion and withdrawal of entire sectorsalong rails.

The overall layout of the ARIES-RS fusionpower core is shown in Fig. 1. The figure showsthe in-vessel sectors together with the magnets,vacuum vessel and cryostat. The power core isdivided into 16 sectors corresponding to the 16TF coils. Each sector has its own horizontalmaintenance port, allowing replacement of the


Fig. 2. Cross section of the fusion power core.

entire sector without opening the cryostat or dis-assembling other components such as the coilsystem. Each maintenance port is sealed by twodoors that enclose a separate port vacuum. Theinner door is formed by the outboard low temper-ature shield, while the outer door is a separatecomponent located at the radius of the cryostat(see Fig. 2).

The entire power core is cooled by liquidlithium with an inlet temperature of 330°C and anexit temperature of 610°C. For the replacement ofa sector, six to eight coolant access tubes must bedisconnected and reconnected to the new sector.All coolant access tube connections are made inthe ante-chamber inside the port. A separate vac-uum is created inside this ante-chamber, such thatany coolant leak or dirt caused by handling oper-ations will not contaminate the plasma chamberor the building atmosphere. This feature providesthe flexibility to employ either welded or mechan-ical connectors for the coolant access tubes as well

as for the maintenance doors, depending on theshortest maintenance time. More detail on themaintenance procedure is found below and inRefs. [8,9].

The most severe penalty of single-piece sectorsis the increased size of both the TF and PF coilsystems, needed to allow adequate space for sec-tor removal. The relative increase in capital costas compared with a design having TF coils closelyfitting the outboard blankets is only about 5% ofthe power core cost. Since the cost of electricity isproportional to the availability, the net savingscan be large. A more quantitative comparison ofthis trade-off requires more detailed studies onreplacement time and availability.

A similar scheme also was proposed in theARIES-II–IV study [3,10]. ARIES-RS uses thesame general concept, but considerable technicalwork has been done to show the feasibility anddesirability of this scheme and to make substan-tial improvements.


Fig. 3. Sector geometry, including permanent shields and vacuum vessel.

2.2. Sector integration

Another distinctive feature of this design is theintegration of the sectors. The first wall, blanket,divertor, parts of the shield and stability shells forman integral unit within each sector.

Fig. 3 shows an individual sector and Fig. 4highlights the ‘replacement units’ and the mainte-nance separation line. The integrated sector con-struction eliminates time-consuming maintenanceoperations inside the plasma chamber and providesa very sturdy continuous structure able to withstandlarge loads. The unit is attached only to the bottomstructure of the vacuum vessel and can be connectedand disconnected by working from the port area.Sectors are disassembled and reuseable parts main-tained in hot cells after the plant returns to opera-tion. Furthermore, no rewelding is needed forelements located within the radiation environment.

Each sector contains the following elementswhich are integrated into one removable unit:

(a) Inboard region1. First wall–blanket (front zone of blanket)2. Reflector (rear zone of blanket)

(b) Divertor region3. Upper and lower divertor structure4. Divertor plates

(c) Outboard region5. First wall–blanket (front zone of blanket)6. Reflector–blanket (rear zone of blanket)7. High temperature shield8. Low temperature shield

In order to maximize the useful lifetime of allelements as well as to minimize the waste stream,all power core elements are subdivided into radialzones characterized by different lifetimes. No weldsare used between elements of different lifetimeclasses, thus allowing easy disassembly in a hot celland reuse of elements which are not at the end oftheir lifetime.

Three lifetime classes were selected: 2.5 FPY, 7.5FPY and 45 FPY. The first class with the shortestlifetime is composed of the plasma facing compo-nents—the first wall and divertor plates. Under-neath those components are the elements withintermediate lifetime. These elements include:� the inboard reflector (back zone of the inboard

blanket)� the upper and lower divertor structure


Fig. 4. Elevation view of the sector, showing the maintenance separation line. The divertor plates and inner FW–blanket elementshave a lifetime of 2.5 FPY, the elements comprising the structural ring have a lifetime of 7.5 FPY, and the shield elements arelife-of-plant.

� the outboard reflector–blanket (back zone ofthe outboard blanket)

They constitute an integral hardback supportstructure, or ‘skeleton ring’ capable of withstand-ing large loads caused by gravity and disruptionforces. Attached to this skeleton ring at the innerside are the plasma facing components, and at theouter side of the outboard region are the high andlow temperature shields.

The two shielding zones on the outboard sideare life-of-plant components, but are also part ofthe replacement unit in order to facilitate replace-ment of a power core sector. All zones which arenot at the end of their lifetime at scheduledmaintenance intervals will be separated in the hotcell and reused. High- and low-temperature

shields in the inboard region as well as in thedivertor region are life-of-plant components too,and can remain at their normal position duringthe removal of a replacement unit.

Another advantage of radial segmentation isthe thermal-hydraulic decoupling of the zones. Inboth the blanket and divertor, the rear zone (far-thest from the plasma) is used to maximize thecoolant bulk outlet temperature. The coolant al-ways enters in the regions of highest surface heatflux, where surface temperatures are difficult tomaintain within their design limits, and exits inregions heated exclusively by volumetric heating.This configuration works especially well with self-cooled blankets which absorb volumetric heatingdirectly in the coolant.


The primary disadvantage of radial segmenta-tion is the reduction in radial heat conduction,which is important in loss-of-coolant accidents.Conduction to the shield and vacuum vessel is animportant mechanism to limit the peak tempera-tures in the first wall. Detailed analyses haveshown that the safety requirements can be meteven in worst-case loss-of-coolant accidents [11].

3. Maintenance features for high availability

The power core configuration and integral sec-tor design have a major impact on the availabilityand maintainability, and have been discussedabove. In this section, additional design featuresadopted in order to achieve high availability arediscussed, including:1. Connection and assembly of the removable

power core sectors,2. Alignment and attachment of the sectors, and3. Transfer of power core sectors to the hot cells.

3.1. Connection and assembly of the remo6eablepower core sectors

The connections between the skeleton ring andthe different elements to be attached must bedesigned in such a way that a constant distance inthe radial direction is maintained, but differentialthermal expansion in the poloidal and toroidaldirection is possible without causing large thermalstresses. These connections must be designedstrongly enough to transfer disruption and gravityforces, but also must allow disconnection andreconnection in the hot cell without welding. As afirst estimate, the assumption has been made thatthe total force acting in the horizontal direction isof the order of 200 tons, which is 10 times largerthan the gravity force acting on the first zone (FWplus front part of blanket). This assumption leadsto a design of the links with a strength compara-ble to the stiffness of the outer walls of thesegments which are 10 mm thick. If future disrup-tion calculations determine larger forces, both theouter walls of the segments as well as the connect-ing links will have to be reinforced.

The divertor plates are made of a 5 cm thickvanadium plate with small cooling channels at theplasma facing side. The requirements for the at-tachment of these plates to the support structureare:� Precise alignment of the plates must be possi-

ble,� Differential expansion between the plates and

support structures must be allowed withoutlarge thermal stresses,

� Attachment should not be damaged by largedisruption forces, and

� Replacement of the plates must be possible inthe hot cell without cutting and rewelding ofhighly irradiated parts.

These requirements can be met with threadedhollow bolts which are flexible enough to allowfor differential thermal expansion between theplates and support structure without excessivelarge bending stresses. They can radiate the volu-metric heat to the surroundings at a temperaturebelow 700°C—the assumed temperature limit forthe vanadium alloy. The number and location ofthe bolts is determined by disruption forces. Scop-ing calculations showed that about 10 bolts arerequired for each plate.

A left-hand thread is located at one end of eachbolt and a right-hand thread at the other end. Thedistance between the plates and support structurecan be adjusted by turning the bolt. This can bedone either from the backside of the supportstructure, or, if necessary, from the plasma-facingside. The second approach would require a smallhole in the plasma-facing side for inserting aturning tool.

An additional requirement on the connectionsbetween the different elements of a power coresector is the possibility of separation in the hotcell without cutting and rewelding, allowing thereuse of some of the elements. This is possibleonly if there are no coolant connections inside theradiation environment which would require weld-ing. For this reason, each element has separatecoolant access tubes which are connected insidethe maintenance port. Lead-throughs are requiredfor those coolant tubes penetrating the vacuumvessel. The coolant tubes must have sufficientflexibility to allow for differential thermal expan-


Fig. 5. Coolant pipe passage through the vacuum vessel.

sion. The proposed design of these lead-throughsis shown in Fig. 5.

3.2. Alignment and attachment of the power coresectors

Each removable sector has a weight of approxi-mately 170 metric tonnes, and must be alignedinside the plasma chamber within a few millime-ters for proper function. Since the gravity forcescan be reacted directly at the base of the module,it has been decided to attach the removable sec-tors to the bottom structure of the vacuum vesselonly. Each module will expand equally in thevertical and radial direction if heated up to oper-ating conditions.

Since rolling or sliding support elements wouldnot survive the operating environment inside theplasma chamber, they will have to be removable.The proposed solution is to insert a hydraulicsystem for vertical and horizontal alignment ofthe sector during installation. Low friction sur-faces will be provided for ease of movement andto avoid seizing. A clearance gap in the railsallows for precise alignment. This gap will befilled with a suitable meltable alloy that can bemelted for alignment but which will be solidifiedwhen cooled down to the operating condition ofthe power core structure. After solidification, the

sector will be firmly supported in all directions.Fig. 6 illustrates a rail system for sector removaland support.

3.3. Transfer of power core sectors to hot cells

In case one of the in-vessel elements fails or isat the end of its lifetime, an entire power coresector is transferred to a hot cell, rather thanremoving a single internal element. This transferwill require protection from both the high levelgammas from the components and radioactivespecies such as tritium or activated dust in thechamber. Any connection between the plasmachamber or sectors to be removed and the build-ing atmosphere has to be avoided in order toprevent the spread of any radioactive materials.After a very short operating time, already hands-on maintenance is impossible and all handling ofthe power core must be performed with remote orrobotic methods.

To meet these requirements, the proposed solu-tion is to employ large transfer flasks for themovement of power core sectors to the hot cell.These flasks are attached to the outer flange of amaintenance port prior to opening the doors ofthe port. In this way, any spread of radioactivematerials from the plasma chamber (e.g. tritiumor dust) is avoided.


Fig. 6. Support and alignment rail concept.

After attachment, the outer door is opened andthe equipment including the rail system for theremoval of the sector is removed from the flaskinto the maintenance port. All coolant accesstubes can be disconnected and joined to a tempo-rary decay heat removal system prior to openingthe inner door of the port. The rail system isinstalled to support the power core sector, whichis then detached from the bottom structure bymelting a low-temperature metal alloy in the gapsof the rails. The used power core sector then canbe moved into the flask.

There is an important choice to be made inregard to the size of the transfer flask. Thesmallest size accommodates a single sector butrequires two docking operations to the port forthe removal of the used sector and the installationof a new one. One docking operation for theremoval of two neighboring sectors is sufficient ifthe flask can accommodate two new sectors. Thiscase is shown as an example in Fig. 7. Theoptimum choice between the two concepts de-pends on the time required for docking as well ason the number of sectors to be replaced during areactor shut-down.

The overall goal is to maximize the availabilityof the power station. This requires a suitablestrategy for the scheduled replacement of thepower core sectors. An attempt should be made tocombine sector replacement with other general

maintenance operations into one shut-down pe-riod. This consideration favours a yearly replace-ment of a part of the power core since a completereplacement of all sectors requires either too longa shut-down or parallel handling of too manysectors.

Unfortunately it is not possible at this stage toobtain well-grounded estimates for the requiredshut-down time to replace power core sectors.Estimates range from 4 to 16 days if parallelhandling of all sectors to be replaced is assumed.The time estimate for the repetitive tasks if sectorsare replaced in series ranges from 1.5 to 12 days.These estimates are not precise enough to selectthe transporter concept (single or multiple sectorflasks) or to decide on the number of sectors to bereplaced during a shut-down period. However,they already give some confidence that a reason-able availability can be achieved with the selectedmaintenance approach.

4. First wall and blanket

The ARIES-RS blanket uses a self-cooledlithium design with vanadium alloy as the struc-tural material. The V-alloy has low activation,low afterheat, high temperature capability andcan handle high heat flux. A self-cooled liquidlithium blanket is simple, and with the develop-


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Fig. 8. Outboard blanket and shield.

ment of an insulating coating, has low operatingpressure. Also, this blanket gives excellent neu-tronic performance and the potential for highreliability and long lifetime.

With the assumption of reliable insulating coat-ings, the MHD pressure drop is no longer a majorconcern. The design of the first wall, blanket andshield (FW/B/S) can be optimized to improve heattransfer and to simplify the configuration. Thefirst wall and breeding blanket use a simple box-like structure, with lithium coolant flowing insimple poloidal paths. The outboard cross sectionis shown in Fig. 8. At present, CaO is the refer-ence insulator material and is maintained on allcoolant channel interior surfaces by adding 0.5%Ca in the flowing Li. This material system isunder study currently in the U.S., although othersystems may also prove feasible. The developmentof insulating coatings is at a very early stage andmuch more R&D is required. The improvementsand design flexibility that coatings provide forself-cooled liquid metal blankets make this a veryhigh leverage item.

An important feature of the blanket and shieldis the radial segmentation into four zones: twoblanket zones and two shield zones. The blanketis divided into two regions to maximize the oper-ating life of the structures, reduce the replacementcost, and minimize the waste stream. Scheduledreplacement occurs in each zone when the V-alloyreaches 200 dpa. This occurs after 2.5 and 7.5FPY for the front and rear zones, respectively.The rear portion also serves as the structural ring,which provides poloidal continuity to the sectorsand attachment points for the inner blanket seg-ments. Location of the coolant connections out-side the vacuum vessel allows for easy disassemblyof the segments. Radial segments creates safetyconcerns, since radial heat transport pathways arecritical in loss-of-coolant scenarios. Design solu-tions have been proposed; however blanket re-sponse to coolant loss remains an importantconcern.

Table 2 summarizes the heat loads and peaktemperatures in the blanket and divertor, andFigs. 9 and 10 show the distribution of neutron


Table 2Power flows and peak temperatures

2092Multiplied neutron heating (MW)Total transport power (MW) 431

56Bremstrallung power (MW)25Core line radiation (MW)

341Power reradiated in divertor (MW)First wall surface heating (MW)1 165Divertor total surface heating (MW)2 348Divertor particle power (MW) 88

610Blanket bulk outlet temperature (°C)Divertor bulk outlet temperature (°C) 610Peak V temperature in first wall (°C) 700

681Peak V temperature in divertor (°C)

1 Radiation only.2 Radiation plus particle loads.

Fig. 10. Surface heat flux along the first wall from bremstral-lung, line radiation and enhanced impurity radiation in thedivertor slots.

wall loading and surface heat flux used in theanalyses. The neutron wall loading is determinedfrom Monte Carlo neutron transport analysis ofthe power core, whereas the surface heat flux isbased on a combination of the core plasma radia-tion distribution plus thermal radiation transportemanating from the radiative divertor slots.

Multiple flow passes in the blanket provide thecapability for removing a minimum of 0.5 MWm−2 of surface heat flux. The full coolant flow ispassed first through the front zone, where thesurface heat flux creates large temperature gradi-

ents, and then through the back zones where thebulk temperature can be raised by volumetricheating without exceeding any structure tempera-ture limits. Segmentation of the shield into a hotand cold zone allows partial utilization of the heatdeposited, and also provides further capability forsuperheating the coolant away from the high heatflux region.

ARIES-RS uses both active and passive stabi-lization systems for vertical displacement andkink-mode stabilization. These systems were inte-grated into the sectors (see Figs. 3 and 4). Passive,radiatively-cooled tungsten shells are located be-tween the blanket and shield on the inboard andoutboard sides and actively-cooled coils areplaced outboard between the low-temperatureshield and the vacuum vessel for vertical stability.A thickened vanadium ‘second wall’ behind thefirst wall cooling channel was shown to providesufficient conductivity to stabilize kink modes.

The design of the primary loop and powerconversion system are critical to the attractivenessof the power plant. This system is responsible forefficient power conversion, isolation of radioac-tive products within the nuclear island, and reli-able operation of the power plant. The choice ofan advanced Rankine cycle offers 46% gross ther-Fig. 9. Neutron wall load distribution along the first wall.


mal conversion efficiency. High thermal efficiencyis desirable to partially offset the high capital costof fusion. A double-walled IHX with a Na sec-ondary loop is used to isolate the activated Liprimary coolant from the steam side. The IHX isalso the location where the transition from V toSS is made. The piping which connects the blan-ket to the IHX uses a double-walled structurewith a thin V liner to minimize the added cost ofvanadium.

The tritium recovery process proposed forARIES-RS is based on a cold trap process. Forthe liquid lithium system, the hydrogen solubilityat the cold trap temperature of 200°C is 440appm, which is far above the design goal of 1appm. For this reason, cold traps have been re-jected as a candidate process for recovering tri-tium from lithium. However, the conceptdeveloped here is modified to add protium in thelithium so that the total hydrogen concentrationin the lithium is higher than the 440 appm satura-tion value. At 200°C, Li(H+T) will be supersatu-rated and precipitate out together. The Li(T+H)can be separated from lithium using a ‘meshlesscold trap’ process which was developed by thebreeder program to separate NaH from Na bygravitational force. The Li(T+H) can then beheated up to 600°C for decomposition. The hy-drogen stream will then be fed to the main Iso-tope Separation System (ISS) to separate tritiumfrom protium.

While many attractive features led to the adop-tion of this blanket concept, there remain a num-ber of engineering issues which must be resolved.Some of the more critical issues are:1. V-alloy material development2. V-alloy industry development3. Insulating coating development4. Power cycle development5. Tritium recovery6. Reliability and replacement assessment.

5. Radiation protection and shielding

Detailed shielding analysis has been performed,and identified no serious difficulties with theARIES-RS shields. A dedicated effort was

devoted to the bulk shield in particular, as itrepresents a major cost item for advanced toka-mak designs. Neutronics and economics consider-ations were used iteratively to guide the shieldtoward an optimal configuration while maintain-ing the attractive safety features of the design.Significant savings in shield cost were obtained byimplementing several cost-effective improvements.These are:1. The use of vanadium alloy is limited to those

regions where high temperature operation isabsolutely necessary: (a) the low-temperature(LT) shield, vacuum vessel, and other externalcomponents (having low levels of nuclear heat-ing that can be dumped without significantlyaffecting the power balance) can employ acheaper stainless steel as the main structuralmaterial, instead of vanadium; (b) the use ofsteel filler (rather than vanadium filler) in theshield reduces the cost tremendously. Fillershave no structural role and thus have lowerunit costs compared to structures.

2. Segmentation of the shield into high tempera-ture (HT) and low temperature (LT) zonesoffers a major reduction in the shield cost.

3. The nuclear heat deposited in the HT shield(�20% of the total heating) is recovered ashigh grade heat to enhance the overall powerbalance of the machine.

4. Careful attention is paid to the arrangement ofthe high performance materials (normally ex-pensive) within the shield. Neutronically effi-cient, but expensive materials, such as WC andB4C, are used only in the space-constrainedinboard side to reduce the overall size and costof the machine while less efficient, cheapermaterials are used in the divertor and out-board sides.

5. Designing the shield to last for the entire plantlife without replacement due to radiation dam-age considerations is essential to reduce theoverall cost and to minimize the radwastestream.

6. The shield is a moderate-complexity compo-nent and thus is not costed on the same basisas the blanket and plasma facing components.The shield has lower fabrication, installation,inspection, and quality assurance costs.


Fig. 11. Divertor geometry.

The high-performance cost-optimized shield ofARIES-RS emphasizes the attainment of the top-level requirements. For instance, the bulk shieldprovides lifetime protection for the vacuum vesseland magnet, costs B200 M$, produces 20% ofthe power, generates only Class C low-level waste,and suffers no damage in case of LOCA. More-over, the bulk shield has a simple configurationwith few attachments and plumbing. It is thusexpected to be reliable, and has the ability to bequickly maintained. Qualitative shielding assess-ments indicated no streaming problems for thedivertor and CD penetrations. All penetrationsinclude 1–2 bends and–or are not in direct line-of-sight with the plasma to minimize streamingeffects.

The torus and its ancilliary system within thebioshield are maintained with robotics and remotehandling system. Personnel access into thecryostat is not feasible at any time after shut-down. The development of radiation hardenedtools is a necessity for the maintenance activity ofARIES-RS. Many new materials are currentlycoming into the market with improved radiationresistance. While the development of rad-hardcomponents is ongoing for the experimentaldevice ITER, an aggressive R&D program is still

needed to develop more radiation resistant com-ponents for ARIES-like power plants.

6. Divertor systems

The divertor region of the sector, highlighted inFig. 11, consists of two principal parts: the targetplates and the structures. The structures fulfilseveral essential functions: (1) mechanical attach-ment of the plates through adjustable screw-boltswhich provide for module alignment and offerlateral flexibility for thermal expansion togetherwith strong support against EM events (e.g. dis-ruptions); (2) shielding for the magnets; (3)coolant routing paths for the plates as well as theinboard blanket and replaceable shield; (4) addi-tional heating of the coolant to optimize surfaceheat removal while maintaining high outlet tem-perature; and (5) a contribution to the breedingratio, since the coolant is Li.

The target plates include three pieces: inboard,outboard and ‘dome’ plates. The plasma flowsthrough the scrape-off layer and enters the diver-tor, where enhanced line radiation from injectedneon impurity allows much of the power to bedistributed along the plates and also partially


redirected out to the first wall. Most of the unra-diated particle energy strikes the outboard plates,but the peak surface heat flux has been main-tained below 6 MW m−2. The strike points arelocated close to the coolant inlet in order tomaintain the vanadium structures below 700°C.

The target plates are shown in Fig. 12. A 2-mmthick catellated W coating is applied to thecoolant channel front surface, which is only 1-mmthick V to satisfy temperature limits. Thermalstresses are reduced by using a relatively thicksolid back on the target plates. This thick solidback contributes to induced currents and forcesduring electromagnetic transients, but adds stiff-ness and strength as well. Therefore, stresses areexpected to decrease even under transient condi-tions.

The plates are connected to the rear zone viastrong adjustable screw-type attachments. These

attachments can be designed to react the full forceof disruptions and also accommodate thermal ex-pansion. They also permit precise alignment toadjacent surfaces and removal of individual platesin the hot cells.

Vacuum pumping ducts are placed behind thedome near the strike points for efficient exhaust.Radial channels then direct the gas to a single setof cryopumps at the bottom of the machine.Top-to-bottom conductance connecting both di-vertors is achieved by using the inter-sector vesselvolume underneath TF coils.

7. Electromagnetic disruption analysis

Scoping calculations were performed to esti-mate the electromagnetic forces and stresses in theARIES-RS device resulting from plasma disrup-tion events. These loads then were used to guidethe design of individual components and theirattachments. For simplicity, this scoping analysisconsidered only 2D axisymmetric cases withtoroidal plasma currents, although the evolutionof the current and motion of the plasma havebeen included. The influence of the induced cur-rents back on the plasma evolution also has beenneglected. This level of analysis does not guaran-tee that all design margins will be met, but ratherprovides order-of-magnitude estimates used in thedesign process.

Several components in the reactor, such as thefirst wall, blanket and shield modules are, inreality, segmented. In these components, currentmust pass through the support structure to flowfrom one module to the next. Current loops willform in the modules due to non-uniformity of theinduced electric fields. These current loops crossthe toroidal magnetic field in the sides of themodules and may exert significant torques on themodules. In this analysis, it was assumed that thepath resistance is large enough to neglect thiscurrent. A more sophisticated model would berequired to compute these currents. Only thetoroidally continuous components—the vacuumvessel, vertical stabilizing shells, and buckingcylinder—along with the divertor plates wereFig. 12. Divertor plate and attachment.


considered. All other structures were assumed tohave no electrical conductivity.

The poloidal field coils remain energized duringthe disruption, maintaining the currents set at thebeginning of the burn cycle when currents andfields are near their peak. It is also assumed thatthe equilibrium plasma current is distributed overa circular region. Within that circle, the currentdensity is assumed to be given by:

J=Jmax�

1−r2

r2minor

�aJ

For the analysis, rminor=1.31 m, Iplasma=10.43MA, and the shape factor aJ=0.374. During theplasma current decay phase of a disruption, thisequilibrium distribution is linearly ramped downto zero at all points. Vertical displacements of theplasma are considered by reducing the size of theplasma to prevent contact with the walls (see Fig.13).

In the analysis described here, the induced cur-rents and forces were computed using the com-mercial finite element code ANSYS©. Axisym-metric quadratic finite elements capable of model-ing the radial and axial magnetic field andtoroidal currents in the structure were used. Thefinite element model is shown in Fig. 14. Severalscenarios of fast plasma current quenches andvertical displacement events were considered.These were considered to be representive of themost severe disruption events seen in tokamakexperiments.

The divertors were modeled as three 4.4-cmthick, toroidally continuous plates. In fact, thedivertor plates are not toroidally continuous butinstead are segmented into 16 modules. Currentsinduced in these plates must flow through thesupport structure to complete a toroidal circuit.To consider the additional resistance of this path,the resistivity of the divertor plates in the modelwas selected so that the total resistance in themodel matched the total resistance through theplates and supports. Current loops within individ-ual divertor plates were not considered. The elec-trical resistivities of the divertor plates and allother components in the model are shown inTable 3.

Fig. 13. Vertical displacements of the plasma are treated byreducing the minor radius of the plasma.

The divertor supports were assumed to be1.5 m long with 14 cm2 cross-sectional areaon each side of each plate. Stresses in the plateswere computed by plate theory, assuming no dis-placement and free rotation of the plateboundary.

Three scenarios were chosen because they wererepresentative of the most severe seen experimen-tally [12]. The first is a fast plasma current quenchof 10 ms where the plasma remains stationaryduring the disruption. The second is a simulta-neous plasma current quench and vertical plasmadrift of 1.5 m occurring over 10 ms. The last is aslow vertical drift phase of 1.5 m lasting 2500 ms


Fig. 14. Finite element model used to compute induced currents in the structure.

followed by a rapid plasma current quench of 10ms. A current quench is modeled as a linear currentramp from 10.42 MA to zero.

The analysis results indicate that all componentsare able to withstand their electromagnetic loads ifadequately supported. The currents and forcesgenerally were greatest at the end of the disruptionevents, and were essentially independent of theduration of the disruption, except in the divertors.The quench times considered here are sufficientlyshort that all components, except the divertorplates, experience a roughly linearly rise in currentwith time. The higher resistance of the divertorplates causes their current to level off before the endof the disruption. For other components, the rateof the linear rise in current is inversely proportional

to the duration of the disruption. Thus, their totalcurrent and load is essentially independent of theduration of the disruption.

For the 10 ms stationary plasma disruption, theeddy currents peak at 10 ms with a total of 9.9 MA.The 100 ms disruption with plasma drift generatesthe same peak eddy currents at the end of thedisruption as well. The 2500 ms drift and fastquench scenario induces a maximum of 9.8 MA atthe end of the quench. The maximum magneticpressures and stresses for all three scenarios areshown in Tables 4 and 5.

The maximum current and loads in the vacuumvessel occur at the end of the disruption in allscenarios. The greatest current density appears onthe inboard side on the midplane. In all cases, the


Table 3Electrical resistivities of the materials in the model

Thickness (cm)Component Resistivity (mV–m)Material

4.3Divertor plates 32*100% vanadium0.242035% tenelon, 40% B4C, 20% WC, 5% voidInboard vacuum vessel0.63Divertor region vacuum vessel 57% borated tenelon, 35% tenelon, 3% WC, 5% void 20

0.66Outboard vacuum vessel 70% borated tenelon, 25% tenelon, 5% void5 0.18@500°C100% tungstenInboard stabilizing shell6Outboard stabilizing shell 100% tungsten 0.18@500°C

0.6927100% tenelonBucking cylinder

* Effective resistivity after considering current path through support structure.

greatest electromagnetic pressure appears on theinboard side near the midplane (the magnetic fieldis greater there than at the midplane). Since thevacuum vessel is cylindrical in this region, theresulting hoop stress is computed assuming thatthe vacuum vessel is a thin cylinder with uniformpressure. Hence, this stress is computed by multi-plying the electromagnetic pressure by the vac-uum vessel’s radius and dividing by its thickness.

The loads and stresses for the outboard diver-tor plate are summarized in Table 5. The loads onthe other two divertor plates were smaller thanthose shown. These tables list total radial andaxial forces on each plate over 360° (divide by 16to compute the force for a single module). Posi-tive radial force is away from the machine’s cen-terline; positive axial force is away from themidplane. The maxima are for either upper orlower divertor plates. The calculated stressesnever exceed 10 MPa, much less than the allow-able primary bending stress for vanadium up todivertor plate operating temperatures (165 MPa).

The loads on the tungsten vertical stabilizingshells are larger than the other components, dueto their low electrical resistance and proximity tothe plasma. In this idealized model, the loading istoroidally uniform, and if the shells were free toexpand, they would act as free-standing rings andwould not react radial loads into their supportstructure. However, non-uniform loads couldcause the shells to buckle, so that supports will beneeded. The maximum loads occur on the outershell and are shown in Table 5. These loads aresubstantial. Periodic spacing of supports would

cause unacceptable bending stresses in the shells.Therefore, continuous support structure along thenear-midplane edge is needed for these compo-nents.

8. Current drive, heating and fueling systems

With 88% bootstrap current, the current driverequirements for the reversed shear plasma aremodest. However, three RF current-drive systems,operating in different frequency regimes, are re-quired to drive and control the equilibrium cur-rent density profile needed to maintain MHDstability at the design point. The power from thesesystems also is expected to heat the plasma fromstartup to its final operating conditions. The threeRF systems make use of ICRF fast waves (98MHz), high-frequency fast waves (1.0 GHz), andlower hybrid waves (3.0–4.6 GHz) to drive seedcurrents in the on-axis, off-axis (inside the RSregion) and the edge regions, respectively. Thetotal power delivered to the plasma from thesesystems is 102 MW, of which 81 MW is used indriving useful currents. Table 6 summarizes thekey parameters of the RF heating and currentdrive systems.

The wave launcher and transmission systemsare designed and configured so as to minimizeintrusions into the periodic blanket and withinshield structures. All RF launchers can fit within asingle blanket sector occupying 0.58% of the firstwall area. As a result, the engineering impacts aremodest, and the effects on shielding and waste


Table 4Maximum magnetic pressures and stresses in the vacuum vessel

Maximum magnetic pressure (MPa)Case Maximum stress (MPa)Peak total current (MA)

0.66@10 ms10 ms quench 9.3@10 ms4.6@10 ms9.2@100 ms0.66@100 ms4.9@100 ms100 ms quench with drift

3.6@2510 ms 0.74@2510 ms 10@2510 ms2500 ms drift then 10 ms quench

disposal are manageable. Fig. 15 shows the lay-out of the launchers, which are located aroundthe outboard mid-plane. The transmission linesand waveguides are routed behind the shieldingstructures to minimize neutron streaming and ir-radiation of out-vessel components.

The ICRF launcher system consists of atoroidal array of six quarter-wavelength foldedwaveguides [13], having a width of 1.53 m and aheight of 0.51 m. It is located above the out-board equatorial plane with its aperture flushwith the first wall. As shown in Fig. 16, eachfolded waveguide is a resonant TE101 rectangu-lar cavity having 10 folds (9 vanes), and thephase difference between guides is 60° so as tolaunch the desired wavelength of 1.53 m. Theradial thickness of the module, including thecoupling region at the back end, is reduced to0.88 m by introducing a capacitive diaphragmnear the peak field region that acts like animpedance transformer. The design is accom-plished using a conservative assumption of oper-ation at 50% of the maximum power handlingcapability (40 MW m−2). The antenna directiv-ity is estimated to be 90% and the coupling effi-ciency is 95%.

The lower hybrid launcher system is com-posed of five separate modules, each radiating adifferent wavelength for current profile controlover a wide plasma edge region. These modulescan be located above and below, but not too farfrom the outboard equatorial plane. The baseunit of the system is a reflector waveguide–hy-perguide combination modelled after the ITER-EDA design [14] (see Fig. 17). Each reflectorwaveguide consists of a toroidal array of eightactive–passive overmoded (TE60) rectangularwaveguides that couple directly to the plasma.

The reflector waveguide is fed by a 3×2 arrayof TE30 waveguides, with a 180° phase differ-ence between rows, via the TE60 hyperguideunit. Each individual LH module is then madeup of a toroidal array of four reflector waveg-uides, with a width of 0.68–0.79 m and a heightof 0.24–0.27 m. The radial thickness is deter-mined to be 22.5 cm. The directivity of theseunits is projected to not exceed 70%, and apower handling capability of 60 MW m−2 isassumed.

There is considerable uncertainty in thelauncher configuration for the HFFW system. Itis envisaged that this might take the form of amatrix of combline structures [15] or an array ofso-called ‘leaky’ waveguides that radiate thebulk of their power along the toroidal length.Clearly, innovative launcher concepts need to bedeveloped in this frequency range. Using pru-dent assumptions, a port of 1.6 m width and 0.5m height, located below the outboard equatorialplane, has been set aside for this launcher sys-tem. The thickness of the combline structure isof the order of a few centimeters.

The structural material for the in-vessel RFsystem components is the same V-4Cr-4Ti alloyas that of the blanket–shield. To minimize theRF wall losses, the inside surfaces of the struc-tures are coated with dispersion strengthenedcopper, GlidCop AL-25, varying in thicknessfrom 8 mm in LH systems to 59 mm in ICRFsystems. Loss of copper conductivity over theRF component lifetime due to neutron-inducedtransmutations does not appear to significantlydegrade the system performance efficiency. Theability to coat vanadium with copper must beestablished.


Table 5Maximum stresses and total radial and axial forces

Maximum stressPeak total currentCase Peak axial forcePeak radial force(MN) (MPa)(MN)(MA)

1. Outboard divertor plate9.5@10 ms0.320@10 ms10 ms quench −0.57@10 ms0.038@10 ms

0.083@100 ms100 ms quench with drift 0.006@100 ms 1.8@100 ms−0.079@100 ms−0.98@2510 ms2500 ms drift then 10 ms −1.8@2502 ms 7.50.12@2510 ms

quench

2. Outboard stabilizing shell10 ms quench 1.7@10 ms −47@10 ms −13@7 ms

−12@50 ms100 ms quench with drift −46@100 ms1.6@100 ms3.2@2510 ms −89@2510 ms 33@2500 ms2500 ms drift then 10 ms

quench

The structures are cooled by liquid lithium at500°C at the first wall midplane. Detailed coolinganalysis has not been carried out for the RFstructures. However, a rough calculation in whichthe ICRF launcher is represented by a flat vana-dium plate yields an acceptable volume fraction ofthe coolant to be 40% at a coolant flow velocity of2 m s−1. Volume neutron heating was found to bethe dominant heat source.

The transmission lines (coax and waveguides)

are routed from the launchers and bent behind theshielding structures in order to protect the win-dows from direct neutron bombardment. Underthis shielded environment, the ceramic materialfor the windows is chosen to be BeO for its highspecific heat, high thermal conductivity and excel-

Fig. 15. RF plug-in unit that contains all RF launchers in oneblanket sector.

Table 6RF system summary

Total delivered power (MW) 102.1Current-drive efficiency (A/W) 0.017Normalized efficiency (A/m2 W) 1.63×1020

2.53Total first-wall penetration (m2)Driven current (MA) 1.37ICRF (Ion Cyclotron Resonant Frequency)

98Frequency (MHz)Wall-plug-to-plasma efficiency (%) 75Wall-plug power (MW) 23.2Power delivered to plasma (MW) 17.4

15.7Current-drive power (MW)HFFW (High-Frequency Fast Wave)

Frequency (GHz) 1.061Wall-plug-to-plasma efficiency (%)

Wall-plug power (MW) 61.637.6Power delivered to plasma (MW)32.1Current-drive power (MW)

LHW (Lower Hybrid Wave)4.6 and 3.5Frequencies (GHz)46Wall-plug-to-plasma efficiency (%)102.4Wall-plug power (MW)

Power delivered to plasma (MW) 47.1Current-drive power (MW) 33.0


Fig. 16. Folded waveguide cavity with capacitive diaphragm and coax feed of the ICRF launcher system.

lent thermal stress resistance. Beyond the win-dows, the lines are pressurized with an inert gassuch as nitrogen. If required, the center conductorof the coax near the ICRF launcher module

should be supported at intervals along its lengthby ceramic spacers made of SiC, in order towithstand the high neutron fluence in that region.

The amount of copper in the RF structures isan important concern from the perspective ofwaste disposal rating. The bulk of the copper isused as the structural material for passive wave-guides of the LHW launchers facing the plasma,and to provide sufficient conductive cooling fromthe surface heat flux. The total amount of copperused in the three RF system components (surfacecoating and passive waveguides) has been esti-mated to be about 0.2% of the total structurevolume, and therefore should not pose a wastedisposal issue if the RF systems are replacedtogether with the blanket sectors.

8.1. Fueling systems

For fueling the reversed-shear region ofARIES-RS, the use of pellet injection is mar-ginally adequate. A two-stage pneumatic pelletinjector is used, although further development willbe needed to achieve the required injection veloc-ity of 5 km s−1. The system is expected to createminimal interference with the plant layout and

Fig. 17. Toroidal array of two reflector waveguides of theLHW launcher system.


Fig. 18. TF and PF coil systems.

maintenance scheme. Transport tubes can be longenough to place the fueling equipment outside thecontainment building.

For the system design, a set of criteria havebeen used to ensure the plasma remains undis-turbed and stable to the pellet ablation. Using a50/50 DT pellet of 0.4 cm diameter, the requiredinjection velocity is 5 km s−1 at a repetition rateof 2.9 Hz. The injector is located at the outboardmidplane, pointing horizontally along the equato-rial plane. This requires 1.0 cm diameter tubesrunning with shallow bends through the mainte-nance ports.

For this set of parameters, the particle inven-tory fraction for each pellet is 4.3% and themaximum local density perturbation is calculatedto be 23%. The pellet penetrates as deep as 19 cmfrom the edge, with the peak deposition occurringat 2 cm inside the reversed shear region. Theimpact of off-axis fueling profile on plasma stabil-

ity should be modest with careful control. Exten-sion to an injection speed of 10 km s−1 willprovide greater flexibility in the fueling profile,and necessitates advanced technologies, such asrail guns and compact toroid injection.

9. Magnet systems

The ARIES-RS toroidal field coil set consists of16 coils using multi-filamentary Nb3Sn and NbTisuperconductors with a peak field of 15.8 T at thecoil. The PF coil set consists of 18 coils: eightform the center stack, and the remaining tenelongate the plasma, provide equilibrium andform the divertor magnetic configuration. Fig. 18shows the overall configuration of both the TFand PF coils and structures.

Four goals have been pursued in the design ofthe TF coil systems for ARIES-RS.


1. The magnet cross section has been optimized,using innovative solutions to minimize thecross sectional area and the cost.

2. The effect of off-normal events was evaluated.Failure modes of the TF coil were investigatedunder off-normal events, and it was realizedthat one mode of failure, namely that of ashort across the leads of one TF coil followedby a dump of the rest of the TF magnetsystem, results in large bending and deforma-tion in the shorted coil as it tries to go into acircular shape.

3. The consequences of enlarging the toroidalfield coils in order to allow for maintaining theblanket–shield by removing entire sectorswere investigated. This maintenance approachrequires large access ports between the outerlegs of the TF coil. The possibility of support-ing these sections of the coil without structurebetween legs (that would obstruct easy access,even if removable) was evaluated.

4. The impact of the large toroidal field coil onthe poloidal field system has been investigated,and modification to the TF coil have beenevaluated in order to minimize the impact onthe PF system.

9.1. TF coil internal design

Advanced magnet design techniques were usedto maximize the utilization of structural materials,as in ARIES-II and in an early ITER design[16,17]. The magnet design concept uses toroidal-shell structural forms with grooves into which theconductor is wound (see Fig. 19). Due to im-proved capability to support out of plane loads ofthe toroidal-shell magnets, it is expected that thisconfiguration will be superior to that of casemagnets, and to that of magnets with radially-ori-ented plates. A main difference between the pro-posed design and that of the early-ITER is that inthe ARIES-RS design the toroidal field coil leansagainst the bucking cylinder, as opposed to theearly-ITER, where it leans against the centralsolenoid. In the case of the early-ITER, the out-of-plane loads in the inner leg of the TF coil werereacted by the coils themselves, since the idea oftransferring the out-of-plane loads through the

ohmic solenoid was rejected. In the case ofARIES-RS, the bucking cylinder is used to reactthe out-of-plane loads.

With this toroidal shell approach, the toroidalfield coils are layer-wound, thus making possiblethe grading of the conductor. If the conductorwere pancake wound, a very large number ofjoints between the grades would be needed, mak-ing this approach impractical. Layer winding amagnet not only simplifies the electrical connec-tions in the magnet, but also the cryogenic con-nections. Furthermore, the winding process issimplified.

The coil design optimizes the superconductor–copper, helium, insulation, and structural ratiosof the coil in order to minimize the radial thick-ness of the TF magnet system subject to thefollowing restrictions: superconductor critical cur-rent, superconductor stability, quench protection,superconductor strain, stress and strain in struc-tural materials, heat removal, pumping power,conductor fabrication, and magnet construction.There are four grades of conductor, with differentcopper-superconductor-He-structure ratios. It isassumed that the superconductor shares the me-chanical strain with the structure. When com-bined with the thermal strains, the net strain inthe superconductors is very small, with resultingincreased critical current capabilities.

Two design choices are included in the designof the conductor. The first choice segregates thecopper required for quench protection into purecopper strands, with much reduced cost thanwhen the copper is included in the composite withthe superconductor. The copper is not as effectivefor conductor stability, and the energy margin ofthe superconductor is decreased, but sufficient.The second design choice uses internal, ratherthan external, dump of the magnet stored energyduring an off-normal event. If internal dump isused in the case of magnet coil quench, the largedump voltages are avoided, and the insulation canbe made substantially thinner. Internal dump sub-stantially increases the allowed current in the cop-per during quench. In addition, it simplifies thecharging and discharging of the coil, by minimiz-ing the number of current leads. It was assumedthat some of the sensors used for quench detec-


Fig. 19. TF coil internal geometry.

tion could be used also to create a large normalzone for the internal dump. More energy isdissipated in the magnet (3–4 times moreenergy), requiring longer recool times after amagnet dump.

9.2. TF coil structures and shape

The bucking cylinder therefore can be used tosupport the out-of-plane loads of the TF coil. Inaddition, in the top and bottom of the device acap-like structure, similar to those of the previousARIES designs [16], can be used to minimize thedisplacement due to the out-of-plane loads. Thecaps are reinforced with a set of straps located in

the shadow of the outer legs of the TF coil. Fig.20 shows the superstructure, consisting of buckingcylinder, the top–bottom caps and the strap inthe shadow of the toroidal field coil. The amountof material added to the structure, is relativelysmall. The thickness of the band in the outer legof the TF coil is comparable to that of the thick-ness of the TF coil. In addition, this structure canbe used to support the outer PF coils.

Because of the choice of maintenance approach,the outboard leg of the toroidal field coil is at alarge major radius. The height of the coil with aconstant tension shape [18], would push thepoloidal field coils far away from the plasma. Theshaping coils, mainly those near the top and


bottom of the TF coil, become very large and itwas difficult to design a self-consistent set of coils.The shape of the toroidal field coil was changed inthe reference ARIES-RS design, mainly by de-creasing the height of the coil. As there is plentyof structure surrounding the coil, the increasedlocal stresses are acceptable.

The decreased coil height increases bending inthe region where the vertical leg meets the rest ofthe coil, making the tension due to inplane loadsacross the region of maximum bending (in theregion between the vertical inner leg and thecurved coil) vary from 80 to 250 MPa. It istherefore possible to further decrease the height ofthe toroidal field coil without compromising theTF or the superstructure, and bring the poloidalfield coils even closer to the plasma. This shouldbe explored in the future.

9.3. Fault analysis

For the toroidal field coil analysis, it was deter-mined that failure of the poloidal field system ordisruption does not increase the out-of-plane

loads substantially. The worst case scenario con-sists of a coil shorted across its leads, followedwith a dump of the other coils. In this case, thecurrent in the shorted coil increases (by induc-tance coupling), while the current in the othercoils decrease. The system of coils therefore ceaseto behave as a toroidal system (with D-shape coilsas the bending free structure). The over-drivencoil tends to become circular. In the case ofARIES-II and PULSAR [15], the coil has littlecapability of carrying bending (since it is made ofthin shells), and in the absence of a large supportstructure (present in ITER and in ARIES-RS),the coil deformations would be very large. Thelarge cap–strap structure serve the dual purposeof minimizing the strains during these faults andto support the out-of-plane loads without theneed of intercoil structure in between the outerlegs of the TF coil.

9.4. PF coils

The poloidal field system, although large due tothe large toroidal field coils, does not presentunusual design issues. The system utilizes NbTiand Nb3Sn conductors. The peak field in the PFsystem is 13.5 T, occurring near the divertor coils.

10. Summary and conclusions

The final design concept for ARIES-RS offershope that a tokamak can achieve both the eco-nomic as well as safety and environmental charac-teristics required to provide a competitive productin the future energy marketplace. Economic at-tractiveness depends on both low capital cost andhigh availability. By optimization of the in-vesselpower core components, the high performanceaspect of Li coolant and V-alloy structure wereexploited, while use of ferritic steel and maximum‘burn-up’ of the structures reduces the impact ofthe high unit cost of vanadium. High availabilityrequires reliable operation as well as rapid re-placement of components which fail or reach theend of their useful life. The configuration choicesmade for ARIES-RS accepted small increases inthe plant capital cost in order to make a credibleFig. 20. TF coil detail.


case for replacement of sectors in under a month.One of the most important remaining uncertain-ties, and indeed, one of the most critical issues forfusion as an energy source is the ability of in-ves-sel components to operate reliably. However,confidence in reliability estimates requires testing.Until a serious program of testing and componentdevelopment is undertaken for these advancedtechnologies, it will be impossible to make a com-pelling case that these systems will meet theiravailability goals.

Refinement of the major fusion power corecomponents has helped to optimize their designsand to establish design limits for the key deviceparameters, such as neutron wall load and surfaceheat flux. The first wall and blanket were adaptedfrom the ARIES-II design study and furtherevolved. Using a very simple, and therefore hope-fully reliable geometry, the power density createdby this advanced, high-beta plasma appears to behandled in the blanket without violating any de-sign constraints. The entire feasibility of thepower core depends on successful development ofeffective electrically insulating coatings which canwithstand exposure to the operating environment,including intense neutron irradiation. Acceptanceof the risk of not finding such a solution wasmotivated by the tremendous advantage of liquidmetal coolants over helium-cooled ceramicbreeder blankets in their ability to handle veryhigh power densities. Demonstration of accept-able mechanical properties of the vanadium alloyunder irradiation is also a critical issue, althoughthe relatively low system pressure is hoped toallow the material to survive at least as long asthe scheduled maintenance interval of the powercore.

Engineering design of the divertor is moststrongly influenced by the power and particleloads on the target plates. Demonstrating a meansto remove the core transport power is one of themost critical issues for fusion power plants, be-cause current estimates predict localized heatingrates far exceeding the ability of solid structuresto remove the heat under the simultaneous con-straints imposed by efficient power conversion,safety and waste disposal. In ARIES-RS, the pos-sibility to radiate sufficient power from the core

or edge plasma to the first wall appeared to beimpossible, such that a means to distribute thepower within the divertor was sought. The finaldesign uses neon impurity injected in the throat ofthe relatively open divertor slots, such that signifi-cant power is radiated back into the vessel andthe remaining power is spread over an area largerthan the scrape-off layer (projected onto the an-gled target surfaces). The effect of the neon on thecore plasma is known to be critically important,but difficult to predict in the absence of adequateedge transport models.

Target plate erosion is also known to be acritical issue, as it may limit the operating life ofthe divertor to less than the scheduled replace-ment time for the in-vessel components. The re-sult would be a potentially large impact on theplant availability. ARIES-RS uses tungsten as theplasma-facing material because it has the possibil-ity of very low erosion rate if the plasma edgetemperature is sufficiently low. However, furthertesting under prototypical plasma edge conditionsand better modeling tools will be needed to obtainconfidence. In-vessel components are also stronglyinfluenced by disruptions and similar off-normalplasma events (such as VDE’s). 2D scoping calcu-lations suggest that the design can handle theelectromagnetic forces, although much more de-tailed 3D analysis will be required. The top-levelplant requirements stipulate that events whichterminate the plasma unexpectedly are simply nottolerable, such that plasma solutions absolutelymust be found to virtually eliminate such events.In that case, the primary engineering concern is toallow safe shutdown and component replacementin this unlikely event.

Plasma heating and current drive are essentialaspects of a fusion power plant. The importanceof these systems is sometimes overlooked; they areplasma-facing components exposed to the sameharsh environment as the first wall, subject to thesame requirements on safety and reliability, andalso providing a set of operating requirements oftheir own to satisfy their primary function. Forthese reasons, ARIES-RS devoted a significantamount of resources to exploring these systems.Three radio-frequency systems were included.Through design optimization it was possible to


incorporate all of the unique hardware into asingle special sector, with minimal impact on theplant performance parameters (like tritium breed-ing). Due to the high bootstrap fraction, the totalpower to run systems is modest.

Magnet systems remain one of the most devel-oped and well-understood systems in the powercore, but also the most expensive. The designfeatures required for full-sector maintenance wereshown to be feasible, but serve to exacerbate thisproblem. Support of out-of-plane loads is pro-vided without intercoil structure in the outer legsof the TF coils, using caps and outer straps. Thesecap and strap structures also have been shown toaccommodate off-normal events (e.g. a single coilshort during a dump) and the bending stressesdue to the shaping. Large, but inexpensive, super-structure is needed, consisting of a toroidally andpoloidally connected bucking cylinder, top andbottom caps, and straps in the shadows of thetoroidal field coil. The ARIES-RS design opti-mizes the conductor using four grades of conduc-tor, and utilizes internal dump and segregatedcopper strands for quench protection, thereforeminimizing the cost of the magnet system.

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Documents

Conﬁguration and engineering design of the ARIES-RS tokamak power plantaries.ucsd.edu/LIB/REPORT/JOURNAL/FED/97-tillack.pdf · 2001-07-09 · M.S. Tillack et al.:Fusion Engineering