RESOLUTION OF FISSION AND FUSION TECHNOLOGY …€¦ · be based on plasma stability, confinement,...

RESOLUTION OF FISSION ANDFUSION TECHNOLOGY INTEGRATIONISSUES: AN UPGRADED DESIGNCONCEPT FOR THE SUBCRITICALADVANCED BURNER REACTOR

W. M. STACEY,* C. L. STEWART, J.-P. FLOYD, T. M. WILKS, A. P. MOORE,A. T. BOPP, M. D. HILL, S. TANDON, and A. S. ERICKSON

Georgia Institute of Technology, Nuclear & Radiological Engineering ProgramAtlanta, Georgia 30332-0745

Received June 19, 2013

Accepted for Publication September 10, 2013

http://dx.doi.org/10.13182/NT13-96

The conceptual design of the subcritical advancedburner reactor (SABR), a 3000-MW(thermal) annular,modular sodium pool–type fast reactor, fueled by metallictransuranic (TRU) fuel processed from discharged lightwater reactor fuel and driven by a tokamak D-T fusionneutron source based on ITER physics and technology,has been substantially upgraded. Several issues related tothe integration of fission and fusion technologies havebeen addressed, e.g., refueling a modular sodium poolreactor located within the magnetic coil configuration of atokamak, achieving long-burn quasi-steady-state plasmaoperation, access for heating and current drive powertransmission to a toroidal plasma surrounded by a

sodium pool fast reactor, suppression of magnetohydro-dynamic effects in a liquid metal coolant flowing in amagnetic field, tritium self-sufficiency in a TRU transmuta-tion reactor, shielding the superconducting magnets fromfusion and fission neutrons, etc. A design concept for aSABR that could be deployed within 25 years, based on theIFR/PRISM metal-fuel, sodium pool fast reactor techno-logy and on the ITER fusion physics and technology, ispresented. This design concept can be used for realistic fuelcycle, dynamic safety, and other performance analyses of aSABR.

Note: Some figures in this paper may be in color only in the electronicversion.

I. INTRODUCTION

The concept of combining fusion, which can producecopious neutrons, with assemblies of fissionable or fertilematerial, which can use those neutrons to produce copiousamounts of energy via fission or to produce newfissionable material via neutron transmutation, has beenaround since the early days of nuclear power.1–6 Morerecently, in the United States, the concept of usingfusion neutrons to fission the actinides remaining inused nuclear fuel7–10 has attracted some attention, but

the neutron transmutation of fertile material intofissionable material also remains of interest11,12 for thelonger term.

At Georgia Institute of Technology (Georgia Tech),we are pursuing the concept of combining the leadingmetal-fuel, sodium-cooled IFR/PRISM fast reactor tech-nology13–16 with the leading tokamak fusion technologybeing incorporated in the ITER fusion experimentalpower reactor17,18 in the subcritical advanced burnerreactor7,8 (SABR), which would utilize as fuel theprocessed transuranics (TRUs) remaining in used fuelremoved from conventional nuclear power reactors. Webelieve that subcritical operation of a fast burner reactorwould have certain advantages relative to critical

FISSION REACTORS

KEYWORDS: subcriticalreactor, fast burner reactor,transmutation reactor

*E-mail: [email protected]

NUCLEAR TECHNOLOGY VOL. 187 JULY 2014 15


mailto:[email protected]

operation, namely, that in a subcritical reactor (a) themuch larger reactivity margin to prompt critical wouldallow the use of 100% TRU fuel (thus requiring fewertransmutation reactors) and (b) since the burnup reactivitydecrement would be offset by increasing the fusionneutron source strength, the number of reprocessing stepscould be minimized by leaving the fuel in the reactor untilit reached the radiation damage limit (thus requiring fewerbatches of fuel, less refueling downtime, and fewer high-level-radioactive-waste repositories).

An initial design concept for SABR was developed afew years ago,7 and initial dynamic safety,19 fuel cycle,20,21

and materials scenario22 studies of the SABR concept havebeen performed to examine (a) the peaking of powerdistributions, fuel composition evolution, and neutrondamage with irradiation; (b) the dynamic response of thesystem to various reactivity insertions and coolant systemmalfunctions; and (c) the reduction achievable in decay heatand required high-level-waste-repository capacity for fueldischarged from conventional reactors by using it as fuel forsubcritical fast burner reactors rather than directly burying it.

However, these studies have also identified a numberof fission-fusion technology integration issues—refuelingthe fast reactor located within the tokamak magnet system,access for the plasma heating and current drive (H&CD)systems with an annular fission core outboard of theplasma, the magnetohydrodynamic (MHD) effects on theliquid metal sodium coolant flow in the tokamak magneticfield, the effect of the structure required within the tokamakmagnet system on the neutron multiplication of the fastreactor and on the tritium self-sufficiency, etc.—that needto be defined in somewhat greater detail. The purpose ofthis paper is to describe the revised and more detailedSABR concept that has been developed to address theseand other fission-fusion technology integration issues.

Follow-up research has been initiated to (a) repeatand extend the fuel cycle, dynamic safety, and decay heat/repository capacity studies for the revised SABR conceptand (b) compare the overall transmutation systems thatwould be needed with critical and subcritical (SABR) fastburner reactors for closing the back end of the nuclear fuelcycle. This research will be published in separate papersover the next few years.

II. DESIGN OVERVIEW

II.A. Design Objectives

The purpose of SABRs is to fission the TRUs in boththe legacy used fuel and in the used fuel that will bedischarged from the existing and near-future fleet ofconventional light water reactor (LWR) nuclear powerreactors, thereby reducing by at least an order of magnitudethe storage capacity requirements for long-term high-level-waste repositories, while also extracting the energy contentin that used fuel. The dominant guiding principle of the

previous SABR design7,8 and of this design upgrade is theutilization of state-of-the-art fast reactor and tokamakfusion physics and technology that will [by ITER (Refs. 17and 18) in the case of fusion] or could (by a demonstrationfast reactor and fuel reprocessing/refabrication system inthe case of fast reactors) be demonstrated on a reactor scalein the near term to make deployment of the first SABRfeasible within 25 years. Retention of the inherent safetycharacteristics that have been demonstrated for sodiumpool–type fast reactors with metal fuel and achievement oftritium self-sufficiency are two other primary designobjectives. It is our intent that this SABR design upgradebe based on plasma stability, confinement, noninductivecurrent drive, and other plasma parameters in the upperrange of the established database and confront the mainfission-fusion technology integration and radiation shield-ing issues to provide a realistic basis for future dynamicsafety and fuel cycle performance analyses of the SABRconcept. Based on (a) the EBR-II operating experience, thedesign of the IFR and PRISM reactor concepts, and theassumption that a prototype fast reactor and associated fuelreprocessing and refabrication facilities will be operatedwithin the next 25 years and (b) the operation of ITER as aprototype fusion neutron source in the 2020s, we adopt75% availability for a 40-year lifetime as a design objectivefor SABR for the purpose determining shielding require-ments and materials radiation damage.

II.B. Configuration

The SABR is a metal-fuel, modular sodium pool–type, annular fast reactor located just outboard of theplasma and within the toroidal field (TF) magnet coils of asodium-cooled tokamak neutron source, as depicted inFig. 1. The total power output is 3000 MW(thermal),mostly from fission but including several hundredmegawatts thermal from fusion and exoergic neutronreactions. Ten 300-MW(thermal) modular sodium pools,each containing 80 fuel assemblies and an intermediateheat exchanger (IHX), comprise the annular fast reactorcore. The toroidal and poloidal ring coil magnet systemsare identical to those in ITER (Refs. 17 and 18), and otherfusion technology is adapted from ITER.

II.C. Major Parameters and Materials

Table I describes the major parameters and materialsof the upgraded SABR design. A detailed discussion ofthe choice of these parameters and materials is provided inthe following sections.

III. FAST BURNER REACTOR

III.A. Fuel Assembly

The SABR fuel is based on the 40Zr-10Am-10Np-40Pu TRU fuel being developed at Argonne National

Stacey et al. RESOLUTION OF FISSION AND FUSION TECHNOLOGY INTEGRATION ISSUES

16 NUCLEAR TECHNOLOGY VOL. 187 JULY 2014

Fig. 1. (a) Perspective view of SABR configuration; (b) radial build of SABR configuration.



TABLE I

Major Parameters and Materials of the SABR Design

Fast reactor coreTRU fuel composition 40Zr-10Am-10Np-40PuBOL TRU mass 15 104 kgBOL keff 0.973Specific power 198.6 W/g HMFuel assembly 800Fuel pin 469 per assembly, 375 200 totalPower density 256 kW/,Linear fuel pin power 12.3 kW/mSodium coolant mass flow rate 16 690 kg/sCoolant temperature (Tincool; Toutcool) 628 K; 769 KFuel and clad temperature (Tmaxfuel; Tmaxclad) 1014 K; 814 KClad and structure ODS MA957Electric insulator SiCFuel/clad/bond/insulator/duct/coolant/wire (vol %) 22.3/17.6/7.4/6.5/9.3/35.3/1.5%

Tritium blanketTritium breeder Li4SiO4

Sodium coolant mass flow rate 1 kg/s (3.8 m/s)Minimum and maximum blanket temperatures 450uC, 574uCBOL TBR 1.12Startup T required 0.7 kg

ReflectorMaterials—reflector assembly in-core (vol %) ODS steel (58.1%), SiC (6.6%), Na (35.3%)Materials—graphite reflectors (vol %) Graphite (90%), Na (10%)

ShieldMaterials Graphite, tungsten carbide, boron carbide, Na

PlasmaMajor radius 4.0 mPlasma radius 1.2 mElongation 1.5Toroidal magnetic field (on-axis) 5.6 TPlasma current 10 MAInductive current startup 6.0 MANoninductive current drive 4.5 MABootstrap current fraction 0.55H&CD power 110 MW (70 EC, 40 LH)Confinement factor, H98 1.2Normalized bN 3.2%Safety factor at 95% flux surface 3.0Maximum and BOL fusion power 500 MW and 233 MWMaximum fusion neutron source strength 1.8|1020 n/sFusion gain (Qp 5 Pfusion/Pextheat) 4.6

Superconducting magnetsCentral solenoid Adapted from ITER CS system

Maximum field 13.5 TFlux core radius 0.69 mCS coil thickness 0.3 mInductive flux 59 V?s

TF coils ITER TF systemNumber 10 (18 on ITER)Maximum field 11.8 TTF bore dimensions 16.5 m (height)|9 m (width)

PF ring coils ITER PF system



Laboratory (ANL) and is discussed in detail in Ref. 7. TheSABR fuel assemblies are adapted from the S-PRISMassemblies.16 There is one significant modification madeto the assembly to reduce the MHD pressure drop:a 3-mm-thick flow channel insert (FCI) made of SiC(Ref. 23) is placed within the duct to prevent current loopsfrom completing through the duct wall. There is a small gapbetween the FCI and the duct to prevent stresses betweenthe two different materials and which is filled with thesodium coolant; however, this coolant does not need toflow to remove heat from the core, as there is no fuel inthis region. The FCI design is derived from those proposedfor the ITER liquid metal coolant tritium breedingmodules.23 The dimensions of the assembly cross sectionare given in Table II. With the exception of the axialshielding in each pin, which has been reduced to 15 cm,and the fission gas plenum, which retains the plenum-to-fuel volume ratio of 2.5 from S-PRISM, the nonfuelaxial partitions of the assemblies are retained from theS-PRISM designs. The location of control rods will bedetermined later when the fuel cycle and dynamic

safety analyses are carried out. Table III shows theaxial partition data for the fuel assemblies using theS-PRISM notation.16

Plasma H&CD systemEC current drive efficiency 0.025 A/WEC power 70 MW

EC upper quadrant modules (7.25 MW each) 4 (30 MW total)EC equatorial modules (20 MW each) 2 (40 MW total)

LH current drive efficiency 0.035 A/WLH power 40 MWLH number of modules (20 MW each) 2

DivertorMaterials Tungsten, CuCrZr, Na cooledPower to divertora 105 MW; 31.5 MWHeat flux 1 to 8 MW/m2

Sodium coolant mass flow rate 0.09 kg/s

First wallMaterials Be, CuCrZr, ODS steelThickness 8.1 cm (1 cm Be, 2.2 cm CuCrZr, 4.9 cm ODS steel)14-MeV fusion neutron power flux (average) 1.7 MW/m2

Fission z scattered fusion neutron power flux 0.6 MW/m2

Surface heat fluxa 0.44 MW/m2; 0.74 MW/m2

Sodium coolant mass flow rate 0.057 kg/s

Sodium poolNumber of modular pools 10Mass of fuel per pool 1510.4 kgMass of Na per pool 22 067 kgPower per pool 300 MWMass flow rate per pool 1669 kg/sNumber of pumps per pool 2Pumping power per pool 20 MW

aThe first number is for a 50/50 split between charged particles and radiation across the separatrix, and the second number is for a15/85 split.

TABLE II

Assembly Cross-Section Description

Pin diameter 0.539 cmPin pitch 0.6346 cmPin pitch/diameter 1.177Fuel diameter 0.370 cmCladding thickness 0.0559 cmCladding material ODS MA957Bond material NaFuel smear density 75%Assembly pitch 16.142 cmDuct outer across flats 15.710 cmDuct thickness 0.394 cmGap between duct and FCI 0.200 cmFCI outer across flats 14.522 cmFCI thickness 0.300 cm

TABLE I(Continued)



Each assembly houses a hexagonal lattice of 469 fuelpins. These pins are taken from the ANL zero conversionratio (CR 5 0) core design carried out for the AdvancedFuel Cycle Initiative.15 The CR 5 0 core had three pindesigns with varying assembly fuel volume fractions to

flatten the radial power profile; SABR uses only the largestof these, because the external neutron source and highleakage factor reduce power peaking and therefore the needfor power shaping. The outside of the cladding of each pin iscoated in a thin layer of SiC for electrical insulation overwhich is a layer of oxide dispersion strengthened (ODS)steel to prevent insulation breakdown in the event of crackformation in the SiC (Ref. 24). These layers, which preventcurrent loops from completing through the fuel pins, are ofthe order of a few microns thick, and the outer metal layer isthin enough that it has sufficient resistance. Figure 2 showsthe fuel assembly cross section, and Fig. 3 shows thedimensions of the fuel pins.

III.B. Sodium Pools

The fissionable material is contained in ten modularsodium pools, which form a fission annulus within the TFcoils, as depicted looking down from above in Fig. 4. Thepools are sized to be removable radially in and outbetween the TF coils in two locations (at pools 1 and 6) tofacilitate refueling. The vertical separation of 4.4 mbetween the two most equatorial poloidal field (PF) coilssets the maximum height of each modular sodium pool.The minimum separation between TF coils is 4.4 m at thepool location.

Fig. 2. SABR fuel assembly configuration.

TABLE III

Fuel Assembly Axial Parameters

Assembly Axial Partition Height (cm)

Upper handling socket 30.480Handling socket – duct overlap 0.686Duct standoff 4.445Pin upper end plug 2.540Fission gas plenum 121.875Core 65.0Lower shielding and end plug 15.0Pin-grid overlap 0.914Pin support grid 5.182Grid–nose piece overlap 1.727Duct–nose piece overlap 3.048Nose piece 33.020Pin (total) 204.415Duct (total) 215.135Assembly (total) 274.901



The ten pools occupy *88% of the annulus, as shownin Fig. 4. The remainder of the annulus consists of twoH&CD ports at locations 180 deg apart to bring power intothe plasma. On either side of each of these H&CD ports arefour small wedge-shaped tritium blankets (TB5 throughTB8). The locations of the TF coils are indicated by thespokes extending radially outward in Fig. 4.

The sodium pool vessels, like the cladding, are madeof ODS MA957 (Ref. 25) the inner surface of each vesselis coated with electrical insulator. A stress analysis wasperformed in Autodesk Inventor Professional 2013 toestimate the necessary thickness of the vessel walls. Thepressure from the sodium and the weight of the fuel,reflector, and core grid plate provided the primaryloading. Because the outside of the vessel is not exposedto air, the thermal gradient across the wall is very small,and thermal stresses are negligible.

According to the American Society of MechanicalEngineers (ASME) code, the allowable limits for nuclearpressure vessels can be expressed as

spƒSmt tð Þ

and

spzsthƒ1:5Sm tð ÞKtSt tð Þ

�,

where

Smt(t) 5 lower of Sm(t) and St(t)

Sm 5 lowest of one-third of the yield or ultimatetensile strength at room or operatingtemperature

St(t) 5 lowest of the minimum stress necessary tocause 1% total strain in time t, two-thirds ofthe minimum stress to cause creep rupture intime t, or 80% of the minimum stress tocause the onset of tertiary creep in time t

sp 5 hoop stress in the vessel wall

sth 5 thermal stress in the vessel wall.

The quantity Kt is

Kt~1z1

21{

sp

St

� �:

The yield and ultimate tensile strengths of ODS MA957 at700uC are both 400 MPa (Ref. 26); at the operatingtemperature of 650uC, these strengths are higher. Creeprupture and strain data25,27 were estimated at a temper-ature of 650uC for a time of t 5 10 years. From these,values of Sm 5 133 MPa, St 5 150 MPa were calculated.

For a vessel wall thickness of 1 cm, the maximumcalculated stress was 100.5 MPa, which was sufficientlylow to meet the ASME criteria. The pool’s upper end cap,which supports the structure, is 25 cm thick. The vesselis shaped such that it follows the curvature around theplasma first wall as closely as possible to maximize thesolid angle of the fission core with respect to the fusionneutron source. The vessel on which the stress analysiswas performed has rounded corners and an elliptically

Fig. 3. SABR fuel pin configuration.



Fig. 4. Layout of the sodium pools, H&CD ports, and wedge-shaped tritium blankets (TB5 through TB8).

Fig. 5. SABR modular sodium pool configuration.



tapered bottom; these curves are omitted from Figs. 4and 5 for simplicity.

Each modular sodium pool contains 80 fuel assem-blies in four radial layers of (from in to out radially) 19,19, 21, and 21 assemblies; there is one layer of 21reflector assemblies outside the fourth radial layer of fuelassemblies. The layout of the fuel assemblies and the IHXwithin each modular pool is shown in Fig. 5. Eachsodium pool also contains two electromagnetic pumps.

We plan in future dynamic safety studies to examinethe neutronic coupling of the modular sodium pools forthe possibility of power oscillations. We anticipate thatthe modules will be tightly coupled and any incipientpower oscillations would be strongly damped, but we planto examine this matter.

III.C. SABR Refueling

The sodium pool removal ports are located at poolsnumbered 1 and 6 in Fig. 4. After removing pools 1 and6, the remaining pools in each bank can be rotatedthrough the annulus to their respective removal positionand extracted. Pools 1 through 5 are removed throughrefueling port 1, and pools 6 through 10 are removedthrough refueling port 6. Reloading the modular pools isthe reverse of the removal process.

The removal motion can be broken into two distinctphases: the rotation around the annulus and the extractionfrom within the coils. Because the entire reactor resideswithin the cryostat under vacuum, there is no access to asecondary heat sink while a pool is in motion. The sodiumpool must be disconnected from the secondary coolant andany piped reactor vessel auxiliary cooling system (RVACS).Thus, the only heat removal is by radiative cooling, which isinsignificant compared to initial decay heat levels.Therefore, to prevent cladding temperatures from becomingtoo high while the pool is being moved, the secondarysodium is used to reduce the pool temperature to slightlyabove the sodium freezing point (378 K) before disconnect-ing the secondary feed to the IHX. The thermal capacity ofthe pool is then used to absorb the decay heat while the poolis being moved.

The maximum duration of the rotation phase occurswhen pools 5 and 10 are being moved. It is estimated thatthese rotations will take *1 h. The maximum decay heatrate that could be tolerated without any heat sink for 1 hwithout clad damage would be 3.18 MW (*1% of fullpower). We estimate that reduction to 1% full poweroccurs in *104 s, so it should be safe to disconnect thesecondary system by 3 h after shutdown. Thus, reductionof decay heat to acceptable levels should not cause anydelay in refueling.

Once each pool is rotated around to its extractionpoint, the secondary coolant feed previously connected topool 1 (or pool 6) is attached to reduce the temperaturewithin the pool. When a sufficient thermal buffer is re-established, the feed is disconnected and the extraction is

performed. Once a pool is removed from within the coils,it is reattached to a peripheral secondary coolant feed(away from the reactor) to help remove decay heat. Thepool is then relocated to a hot cell where the assembliesare removed and the fuel can be sent for reprocessing.Circulation of an inert gas against the sodium pool vesselalso becomes available to cool the modular sodium poolsonce they are outside the cryostat.

It should be noted that neither the heating/currentdrive apparatus nor the wedge-shaped tritium blanketsneed to be able to move to facilitate refueling.

Because of the high fast neutron fluence to the plasma-facing pool wall, once every 10 to 15 years, the fresh fuelwill be loaded into new pools, which are inserted into thereactor. The exact lifetime of the sodium pool will dependon the irradiation time of each core batch and fuel shuffling,which determine the fusion power. The irradiation timewill also influence the amount of cooling time necessary toreduce decay heat levels to 1% of full power.

III.D. Core Neutronics Performance

For the neutronics design analysis, the reactor wasrepresented with a multigroup, finite difference, discreteordinates approximation in R-Z geometry since all of thestructures except for those in the fission annulus have360-deg symmetry about the toroidal axis. All of thestructures within the modular sodium pools were homo-genized and given smeared volume fractions that includedthe core (or reflector) assembly, sodium, vessel wall, andthe gaps between the banks of pools. The volumefractions for each radial pool zone are given in Table IVfor the original region of the ring, showing how much steel,sodium, and void were smeared into those regions.Because smearing around the annulus effectively dilutesthe fuel with small amounts of structure and coolantand spreads it over a larger volume than it truly occupies, itis expected that the spectrum will slightly harden andthe reactivity of the fuel slightly increase when modeledin a more precise manner, but this effect is expected tobe small.

The R-Z model is shown in Fig. 6, with dimensionsgiven in Fig. 1b. The neutronics calculations were carriedout using ERANOS (European Reactor ANalysisOptimized calculation System), a fast reactor code28

originally developed to model the Phenix andSuperPhenix reactors. ERANOS employs the EuropeanCell COde (ECCO) to collapse 1968-group JEFF2.0 crosssections within each reactor lattice cell to the 33 groupsused in core calculations, ranging from 20 MeV down to0.1 eV. The core calculations were carried out using an S8

quadrature with 170 radial and 292 axial meshes. Thefusion neutron source was distributed among 24 plasmaregions and grouped into 5 source strengths to approx-imate the fusion density distribution of a tokamak plasma.

The beginning-of-life (BOL) reactor core character-istics are shown in Table V. The parameter ksource is the



number of neutrons produced on average from eachfusion source neutron. Thus, the fusion power requiredto drive the subcritical fission reaction at a given powerlevel is

Pfus~Pfis1{ksourceksource

� �n

Efus

Efis

� �, ð1Þ

where n is the number of neutrons per fission and E is theamount of energy released per fission or fusion. Table Valso shows the more familiar parameter keff, the number

TABLE V

BOL Core Performance Parameters

Parameter Value

keff 0.973ksource 0.756Fission power [MW(thermal)] 3000Fusion power [MW(thermal)] 233Average linear power (kW/m) 12.30Power peaking (volumetric) 1.27

TABLE IV

Sodium Pool Radial Ring Volume Fractions for R-Z Model

Radial Zone Percent Assembly Percent ODS Steel Percent Sodium Percent Void

Inner/outer pool wall 0 87.944 0 12.056Radially inner sodium 0 2.947 84.997 12.056Core ring 1 76.167 2.868 8.909 12.056Core ring 2 74.015 2.787 11.143 12.056Core ring 3 79.557 2.710 5.677 12.056Core ring 4 77.429 2.638 7.878 12.056Reflector ring 0 58.829 24.176 12.056Radially outer sodium 0 2.261 85.683 12.056

Fig. 6. SABR R-Z neutronics model.



of neutrons that would be produced on average from eachfission neutron if the neutron fission distribution was thesame as would exist in a critical reactor, which isdifferent from the neutron fission distribution in asubcritical reactor with a neutron source.

Results from the previous SABR fuel cycle ana-lysis20,21 indicate that the limiting factor on fuel residencetime with a 500-MW(thermal) neutron source is radiationdamage to the cladding. Because the present SABRdesign includes both a thicker first wall and the sodiumvessel wall between the neutron source and the fissioncore, it may be necessary to extend the fusion neutronsource somewhat beyond 500 MW(thermal) to provide3000 MW(thermal) of fission power to an end of cycledetermined by cladding radiation damage limits. (Asdiscussed in Sec. IV, the fusion power level could beincreased without changing the design.) The goal of thefuel cycle analysis is to optimize fuel burnup subject toradiation damage to the clad, tritium production, andfusion power constraints. The challenge is in finding abalance among these constraints while matching therefueling window with both the first-wall replacement andthe sodium pool replacement intervals to maximize thecapacity factor.

III.E. Fast Reactor Thermal Performance

In each of the ten modular sodium pools,300 MW(thermal) of power is generated. This poweris removed by ten shell-and-tube IHXs (one per sodiumpool) located beside the fuel assemblies in the sodiumpool. The IHX is based on the S-PRISM (Refs. 16and 29) design and has been shortened to fit into thepool, while preserving the heat transfer rate. Eachpool has its own heat exchanger with sodium as thesecondary coolant. There are two independent second-ary loops, each containing steam generators, turbines,and heat rejection systems that service five of theten modular sodium pools. The intermediate, steamgenerator, and heat rejection loops have not beenexamined in any detail at this stage, because it isassumed that these systems will exist far outside thetokamak and thus will not deviate from existing fastreactor plant designs.

Steady-state thermal-hydraulic calculations for thecore and IHX were made using RELAP5-3D (Ref. 30).RELAP solves the multidimensional fluid energy andmomentum equations as well as the heat conductionequation to generate the temperature distributions withinthe core and heat exchanger. A summary of core thermalparameters is given in Table VI, and a summary of heatexchanger parameters is given in Table VII.

One of the purposes of this SABR design activity isto provide a basis for a comprehensive dynamic safetyanalysis, which will be performed in the near future.However, based on the core design, some preliminary

natural circulation calculations were made using themethodology outlined by Todreas and Kazimi.31 Thisrequires solving Eq. (2) for the maximum heat removalrate Q from each pool:

Q~cp2bDLr2

R

� � 12{nð Þ

DTð Þ3{n2{nð Þ , ð2Þ

where

cp 5 average specific heat capacity of Na

b 5 thermal expansion coefficient of Na

g 5 acceleration due to gravity

DL 5 height difference between core center and heatexchanger center

r 5 cold-leg Na density

R 5 hydraulic resistance coefficient

TABLE VI

Modular Fast Reactor Thermal Performance Parameters

Thermal power per pool 300 MWMass flow rate per pool 1669 kg/sCoolant NaMass of sodium per pool 22 067 kgPumping power per pool 20 MWNumber of pumps per pool 2Pump type ElectromagneticPower density 256.7 kW/,Average linear power 12.3 kW/mPower peaking 1.27Core inlet/outlet temperature 628/769 KFuel maximum/allowable

temperature1014/1200 K

Cladding maximum/allowabletemperature

814/973 K

Coolant maximum/allowabletemperature

787/1156 K

TABLE VII

Intermediate Heat Exchanger Parameters

Secondary mass flow rate 1573 kg/sSecondary fluid NaSecondary inlet/outlet

temperature (IHX)590/739 K

Length 3.5 mNumber of tubes 5700Inner/outer radii of tubes 8.7/9.5 mm



n 5 flow regime (n 5 1 for laminar, n 5 0.2 forturbulent)

DT 5 temperature difference across the core.

Natural circulation removes a maximum of 79.4 MWper pool when two conditions are met: (a) the primarysodium in the hot leg is allowed to reach near-boilingtemperature of 1150 K and (b) the mass flow rate on thesecondary side is maintained at its normal steady-statevalue. This ensures SABR is passively safe for onlycertain types of transients. To maintain passive safety inall conditions, a RVACS needs to be designed.

IV. TOKAMAK NEUTRON SOURCE

IV.A. Plasma Physics Parameters

A D-T tokamak based on the physics and technologyof the ITER design17,18 is used as the neutron source forSABR. A systems code for fusion parameter trade-offstudies, based on engineering and physics constraints32

(and used in previous design iterations of the SABRneutron source33) was used to explore the SABR plasmaphysics design parameter space. The target maximumfusion power for the tokamak neutron source waschosen to be similar to the ITER design value of500 MW(thermal). With a target plasma current of10 MA (chosen based on the previous SABR design), amajor radius of 4.0 m was found to satisfy the variousengineering and physics constraints, as well as provide aflexible operating physics parameter range.

Systems studies were performed to determine thevalues of normalized beta (plasma to magnetic pressure

ratio), auxiliary H&CD power, and confinement H-factor[multiplier on the empirical ITER IPB98(y,2) confinementscaling] required to achieve various fusion power outputsat R 5 4.0 m, I 5 10 MA, and BQ 5 5.6 T. Thenormalized ratio of plasma pressure to the magneticpressure bN was varied between a conservative value of3% (routinely exceeded in present experiments) and amoderately improved value of 4% (often exceeded inpresent experiments). The upper limit of achievable bN isdetermined by plasma MHD stability limits with respectto ballooning modes. The confinement scaling parameterH-factor is a measure of the enhancement of energyconfinement time compared to the ITER IPB98(y,2)empirical scaling law of present tokamak confinementvalues as of 1998 and was varied in the conservativerange of 1.0 to 1.4 that has been achieved since that time.As shown in Fig. 7, there is a rather wide range of plasmaphysics operating space within which the target designparameters can be achieved. We select as reference designparameters [500 MW(thermal) fusion power, H-factor 51.2, bN 5 3.2%], for which the reference auxiliary heatingpower is 110 MW(thermal), with a corresponding Qp 5Pfus/Paux 5 4.6.

Looking at the operational space in a different way,Fig. 8 displays the major radius, plasma current, andfusion power phase-space that can be accessed for aplasma with minor radius a 5 1.2 m, H-factor 5 1.2, andbN 5 3.2%.

The design parameters and constraints are summar-ized in Table VIII. The values in parentheses indicate therange of these parameters for which a neutron sourcedesign meeting the SABR requirements could beachieved, based on Figs. 7 and 8.

Fig. 8. Major radius and plasma current phase-space for SABR.(The contours of constant fusion power are shown bythe lines going from upper left to lower right, while theH&CD power contours go from lower left to upperright. BQ 5 5.6 T.)

Fig. 7. H-factor, bN, auxiliary power, and fusion poweroperational space for SABR. (The horizontal linesindicate the fusion power, and the slanting lines indicatethe H&CD power in megawatts. BQ 5 5.6 T.)



These parameters are for the most part conservative,in that they have already been achieved and will bedemonstrated in ITER operation. The bootstrap currentfraction was taken to be constrained at 55%, which is areasonable value according to previous analysis for ITER(Ref. 34), ARIES (Ref. 35), or other advanced plasmascenarios.36 Perhaps the most aggressive design assump-tion in this analysis is that 40% of the noninductivecurrent drive will be obtained through electron cyclotron(EC) and radio-frequency (RF) current drive technology,which is discussed in Sec. IV.B.

IV.B. Plasma H&CD Systems

IV.B.1. H&CD System Design Considerations

The objective of the H&CD systems design is to usethe same H&CD systems that will be demonstrated onITER, insofar as possible. The unique geometric restric-tions imposed on SABR by the presence of the annularfission core within the TF coil system, along with economicconsiderations, discouraged the use of the large neutral beamand ion cyclotron (IC) H&CD systems, thus guiding thedesign toward the utilization of the smaller EC and lowerhybrid (LH) H&CD systems. LH and EC systems areknown to drive current efficiently,36–38 providing thenecessary noninductive current drive for steady-stateoperation and startup assistance.39

The design analysis of the H&CD system began withan examination of the constraints and limits that would befaced in its implementation. The main challenge for aH&CD system on SABR is the limited midplane plasmaaccess on the outboard side of the tokamak because of thepresence of the surrounding annular reactor core withinthe TF coils. The ITER H&CD system design specifiesmodular, horizontally inserted port plugs interfacing withthe plasma at the outboard midplane that can carry EC,LH, and IC systems, plus additional EC systems in theupper outboard quadrant,40 a location much more easilyaccommodated in the SABR design than midplane access.LH systems, on the other hand, have not been shown tohave launch angle flexibility and conservatively haveouter midplane plasma access.

In SABR, two H&CD access ports (*1-m toroidalarc length and full plasma height at the first wall) willbring the H&CD power into the plasma. These ports willbe 180 deg apart toroidally and will pass betweenadjacent Na pools (see Fig. 4). The radial extent of theNa pool and access requirements for the associatedrefueling equipment also limit the azimuthal launch anglesof the outboard midplane H&CD systems.

There are two requirements on the H&CD system.The first is to start up the 10 MA of plasma current, andthe second is to maintain that plasma current indefinitelyin steady state. The current can be started up by acombination of (a) inductive magnetic flux swingsprovided by the central solenoid (CS) and the PF coilsand (b) by noninductive startup assist provided by theH&CD system.

An important requirement of the H&CD system forthe SABR design is to noninductively drive a plasmacurrent of 4.5 MA, which supplements the 5.5 MA ofbootstrap current to provide 10 MA of plasma current.Analyses like those leading to Figs. 7 and 8 show that theauxiliary power input levels necessary to operate at asufficient fusion power are within the capabilities of anITER-like H&CD system,40 allowing a focus here onintegration of current drive capabilities into the SABRdesign. An ideal current drive system would also drivecontrollable current profiles and be capable of stabilizingneoclassical tearing modes41 (NTMs), which requiressteerable EC launchers.

IV.B.2. Current Drive Efficiency and Profile Control

An expression for the current drive efficiency of a RFsystem commonly used in the literature is

cCD~n20RmIA

Pw, ð3Þ

wheren20 5 electron density of the plasma (1020/m3)

Rm 5 major radius (m)

TABLE VIII

Tokamak Neutron Source Plasma Physics DesignParameters and Constraints

Variable Value

TF in plasma 5.6 TFlux swing in CS 26.3 V?sCS magnetic field 13.5 TElongation, k 1.5Triangularity, d 0.4Safety factor at 95% flux surface 3.0Safety factor at centerline 1.0Fraction of Greenwald density 0.75Power output 500 MWAuxiliary heating power 110 MWBootstrap current fraction 0.55Minor radius 1.2 mFusion power gain, Q 4.6Major radius 4.0 (3.5 to 4.5) mPlasma current 10 (8.0 to 12.0) MAConfinement H-factor 1.2 (1.0 to 1.4)Normalized beta bN 3.2% (3.0% to 4.0%)Average density 1.65|1020 m{3

Average temperature 12 keVNeutron source rate ƒ1.8|1020 n/sInductive startup flux fraction 0.60



IA 5 driven plasma current (amps)

Pw 5 RF power (W).

The SABR neutron source plasma requires 110 MW ofauxiliary power to drive 4.7 MA of current in the plasma.In analyses of the DEMO (Ref. 38) and ITER (Ref. 37)facilities, it has been shown that current drive efficiencyvalues of *0.035 and 0.025 A/W can conservatively beexpected from LH and EC systems, respectively, duringoperation in advanced plasma scenarios that will beinvestigated on ITER. This research suggests that asystem of 40 MW of LH power plus 70 MW of ECpower could meet the noninductive H&CD requirementsof SABR, based on the design value of average plasmaelectron density of *1.65|1020 m{3.

One common goal of current drive systems is toenable a controllable current drive profile, and model-ing36,40 suggests that this may be realistically achievableto some degree with a combination of EC and LH H&CD.Lower hybrid current drive and bootstrap current tend toconcentrate toward the plasma boundary,36,37 peaking thecurrent drive profile. Analysis of ITER scenarios40 andscenarios of proposed research facilities36 show that theradial deposition of current drive from EC systems can bemodified by changing the azimuthal launch angle of theEC wave, at the cost of efficiency. At peak efficiency andwith a bootstrap current fraction of 55%, the current driveof the system proposed in the previous paragraph woulddrive an additional 200 kA of current above the nominal4.5 MA required driven plasma current of the design.Although this excess current may have some benefits, atrade-off could be made between overcapacity andconfiguration of the current drive profile by altering theazimuth of the incoming EC waves, and driving currentless efficiently at more favorable radial locations, asillustrated by Fig. 9 in Ref. 36. This typeof design consideration could mitigate the peakingassociated with the current drive of LH systems and thebootstrap phenomena.

IV.B.3. H&CD System Plasma Access andGeometric Analysis

The EC and LH H&CD systems use the modulesdeveloped for use in ITER (Ref. 40), with a minimal levelof modification required. The 40 MW of LH H&CD willcomprise two 20-MW LH units with launchers centeredabout or slightly below the outer midplane of the twoaccess ports 180 deg apart. This system will comprise thesame power supply, transmission line, and launcher unitsas those of the modular LH H&CD port plug designed foruse in ITER. Also, 40 MW of the EC H&CD power willbe provided by two 20-MW systems using the ITERequatorial port plug components, launching through thetwo main H&CD access ports. The EC units will beinstalled above the LH ports, and the launchers will

vertically follow the curvature of the plasma and first wall,with a poloidal position ranging from above the near-midplane LH H&CD units through the upper quadrantlocation. This location is beneficial for geometricconsiderations and provides some ancillary benefit tothe EC current drive efficiency, as analysis38 suggests thatefficiency is higher when the waves are launched from theupper quadrant.

A redesign of the ITER EC H&CD systems toconform to the geometry of the main SABR H&CDaccess ports instead of those for ITER could allow the fullamount of EC H&CD power envisioned for SABR to belaunched from the main access ports, if this shouldprove to be beneficial. Although the minimization of thenumber of upper quadrant ports is desirable in theSABR design, it may be important to include at leasttwo steerable upper quadrant launchers (located toroid-ally between the main access ports) for use in stabilizingNTMs. For the LH and EC H&CD launchers located inthe main access ports, we anticipate some reduction inthe necessary geometric size of the concatenation of thetwo launchers due to a consolidation and reduction inthe amount of shielding.

Since the SABR design has a smaller plasma and theplasma access is limited by the surrounding annularfission core, a reexamination of the design of the ITERport plugs and upper quadrant plugs should be undertakenfor the purpose of minimizing their size. The ITERequatorial port plugs are designed to accommodate largeICRF launchers, which are not planned for SABR.Moreover, with the comparatively higher launcher powerdensity capabilities of the EC and LH H&CD technolo-gies, it is reasonable to believe that the size of the modularequatorial port plugs could be reduced to fit the mainSABR port geometries and still support a RF systemconsisting solely of LH and EC H&CD. The ITER upperquadrant EC H&CD port width tapers significantlytoward the plasma, and it would be favorable to theSABR design, but not mandatory, if this port design wasrefined so that the maximum cross section was moreclosely approximated by the section nearest to the plasma,rather than the large support plate in the ITER design. Thegeometric implementation of the RF H&CD systems forSABR will have to be engineered when the overall designis developed in greater detail, but the specificationsof the SABR system seem to be well within realisticparameters achievable with ITER LH and EC H&CDsystem components.

The upper quadrant port designs for ITER incorporatesteerable 7-MW EC H&CD systems,40 and the remainingEC H&CD power not launched from the midplane accessports will be evenly distributed toroidally in the upperoutboard quadrant positions between the main accessports to provide steerable H&CD and NTM stabilizationcapability. Using the ITER components for all launchers,this would suggest that four 7-MW upper quadrant



launchers would be included, with two in each areabetween access ports. To accommodate the upper quadrantplasma access restriction provided by the modular sodiumpools in SABR, some reconfiguration of the geometry ofcomponents in the 7-MW EC launchers may be needed,e.g., reducing the angle between the axes of the CS and thelauncher, while maintaining the same poloidal location ofthe launching window as in the ITER design. One possibleadaptation of the ITER H&CD system is shown in Fig. 9.

IV.B.4. H&CD System Noninductive Startup Assistance

Noninductive startup of tokamak plasmas has longbeen a goal of fusion researchers, due to the geometric,economic, and performance benefits that would resultfrom the reduction or even removal of the CS magnetsystem. In present tokamaks, the PF coils and CS providethe magnetic flux required to start up and drive thetoroidal current (which heats the plasma ohmically). TheCS of the SABR tokamak neutron source provides *60%of the total magnetic flux necessary to initiate full plasmacurrent, requiring the assistance of the PF coils and thenoninductive H&CD system to bring the plasma currentto its full value of 10 MA.

Thus, another function of noninductive H&CD is toassist with current startup by reducing the amount ofmagnetic flux that would otherwise be required from themagnet systems to achieve 100% inductive startup. Thisuse of the H&CD system also improves confinementduring low-confinement mode and facilitates the low tohigh confinement mode transition by shaping the safetyfactor profile.42 The LH H&CD system plays a crucial

role in startup ‘‘flux saving’’ and also has other, morespecific, benefits, such as enhancing plasma controlduring startup.43 It is especially efficient at current driveearly in the discharge, after dissociation and before thedensity increases to operational values, driving currentnear the plasma center and favorably impacting stability.44

An investigation by Kim et al.43 found that in ITER-like plasmas, a large majority of plasma current will bedriven by the LH system until almost half of the 15-MAITER plasma current is achieved at 38 s, where the LH-driven plasma current reaches its maximum duringstartup. This analysis was performed with 20 MW ofLH power. Hogweij et al.44 estimate that for fluxconsumption ‘‘a reduction of *15% can be reached,’’while in the analysis by Kim et al.,43 early application of LHpower resulted in a flux savings of 43 Wb out of 124 Wbfor the control discharge without LH startup assist. The fluxthat is expected to be provided from the ITER CS is*118 Wb (not including the contribution from the PFcoils40,45).

Results from the examination of using EC H&CD instartup control are also promising,39 and it has been usedsuccessfully for various purposes on low-aspect-ratiodevices, e.g., in a simulation of the planned NSTX-Uduring startup.46,47 It also seems to be effective inassisting startup by minimizing flux consumption,increasing plasma control, and enabling better hybridscenarios on larger machines like DIII-D (Ref. 39), Tore-Supra (Ref. 48), and in ITER scenarios.42

The application of similar analyses to the plasmaregimes of SABR remains to be done, but given the aboveresults, the ability of the LH and EC H&CD systems of

Fig. 9. Possible H&CD configuration. [On the left are one of two 20-MW LH equatorial (bottom) and 20-MW EC ‘‘high-equatorial’’ launchers in the H&CD port. In the center are two of the four 7-MW EC upper quadrant launchers shownaccessing the plasma vertically from above. For clarity, only half of the plasma chamber, sodium pool annulus, and TF coilsare shown. The shielding, structure, and tritium blankets are omitted.]



SABR to complement the flux generation of the PF andCS and enable full plasma current startup seems feasible.

For the SABR tokamak neutron source, the CS isdesigned to induce 6 MA of current during startup, andthe noninductive current drive will be expected tocontribute *4 MA. This inductive and noninductiveplasma current driven during startup totals the 10 MA ofplasma current specified in the design. As fusion heatingoccurs and the bootstrap current engages, the noninduc-tive current drive system will ramp to its full capacity. Theflux contributed by the current ramp in the PF coils willalso provide a significant inductive contribution to currentstartup, which is not taken into account in the aboveconsiderations. This excess current startup capability willallow the noninductive current drive system to furthertailor plasma safety factor profiles, as is done in the hybridstartup scenarios,42 or will allow the noninductive systemto stand by to provide additional current as needed.Several important parameters of the H&CD system arelisted in Table IX.

IV.C. Magnet Systems

IV.C.1. Toroidal Field Coils

Tokamak reactors confine plasma using superimposedtoroidal and poloidal magnetic fields. Large TFs arenecessary to stabilize the plasma against MHD ‘‘kink’’instabilities and other instabilities that would be deleteriousto plasma performance. These large fields are generated bya series of D-shaped TF coils outside of the plasma.

The ITER superconducting (SC) TF coil system49

shown in Fig. 10 has been used, with very fewmodifications, for the upgraded SABR design. The samecoil dimensions, ‘‘cable-in-conduit’’ conductor design,coil structural design, etc., are used. Each ITER TF coil iscapable of carrying 9.13 MA of current and generating amaximum field of 11.8 T at the conductor. The SABR TFcoils are the same size as the ITER coils but are reduced innumber from 18 to 10 (to provide access for refueling)and are located at a smaller major radius (because of thereduction in size of the CS for SABR relative to ITER).The SABR TF coils produce the same 11.8 T maximum

field at the conductor as in ITER with a slightly largercurrent of 11.2 MA per coil. The SABR coils, like theITER coils, are helium-cooled through channels insidethe conductor windings and are thermally insulated fromthe reactor by actively cooled thermal shields.

Since SABR incorporates only 10 TF coils, compared to18 TF coils in ITER, it might be expected that the field rippleon the outboard of the plasma would be larger than in ITER.However, since the SABR plasma diameter of 2.4 m is muchless than the ITER plasma diameter of 4 m, the outboardsurface of the plasma is 1.6 m farther from the outboard legof the TF coil in SABR than in ITER. This probably morethan compensates for the increase in field ripple at the outerplasma surface that would be produced by reducing thenumber of TF coils from 18 to 10. If not, ferromagneticinserts can be added on the inboard side of the outer leg of theTF coil50 to correct any significant field ripple.

IV.C.2. Central Solenoid and PF Coils

The poloidal field that confines the plasma isgenerated by a current in the plasma itself. In ITER andmost modern tokamak designs, this current is initiatedand, to some extent, controlled using a CS and a set of PF

TABLE IX

SABR H&CD System Parameters

EC power 70 MW, two 20-MW unitsand four 7.25-MW units

LH power 40 MW, two 20-MWunits

EC current drive efficiency 0.025 A/W LH current drive efficiency 0.035 A/WEC frequency 170 GHz LH frequency 5 GHzAverage plasma, ne 1.65|1020 m{3 Required steady-state current drive 4.5 MAMaximum required noninductive

current drive for startup4 MA Excess steady-state current drive

for profile shaping200 kA

Maximum available inductivecurrent drive for startup

6 MA (CS only) Total steady-state current drive 4.7 MA

Fig. 10. The ITER SC magnet system.18



coils. These coils, in addition to a series of smallercorrectional coils around the plasma, are also instrumentalin controlling the shape and vertical position of theplasma.

The CS design of SABR is adapted from the ITERdesign.49 In SABR, the CS is designed to provide 60% ofthe inductive flux swing necessary for startup. Theremaining 40% will be provided by a combination ofthe inductive flux swing from the PF coils andnoninductive current drive methods. The SABR CS isdesigned to produce 59 Wb to induce 6.0 MA ofinductive plasma current for startup. With a maximumfield of 13.5 T in the CS magnet, this flux swing requiresa flux core radius of 0.69 m. To satisfy a stress limit of200 MPa, the CS coil thickness must be 0.3 m thick.

The PF coil system is anticipated to be essentiallyidentical to that of ITER (Ref. 49) shown in Fig. 10,although the currents and perhaps the locations may bealtered to accommodate the smaller SABR plasma andplasma current. The vertical separation of the ITER PFcoils nearest the outboard midplane allows a verticaldistance of 4.4 m for removal of the fission core Na poolmodules. The pool modules in SABR were designed to beno more than 4.4 m tall, which is the approximate distancebetween PF3 and PF4 on ITER. It should be possible toadjust the ring coil positions to achieve a somewhat largervertical opening for Na pool removal, if necessary.

IV.D. Divertor and First Wall

Energetic particles leaving the last closed flux surfaceof the plasma are swept downward toward either theinboard or outboard divertor targets and deposited over a

relatively small area. Energetic photons from bremsstrahl-ung, line and recombination radiation, and atomic physicsprocesses in the plasma are distributed uniformly over theplasma chamber first wall as a surface heat flux. Inaddition, there is a power flux of 14-MeV neutrons andlower-energy fission neutrons through the first wall, someof which are deposited as a volumetric heat source. At fullfusion power of 500 MW and full H&CD power of110 MW, there will be 210 MW of non-neutron poweremerging from the plasma. Assuming that this emergingpower is equally distributed between photons (radiation)and charged particles, this will constitute a power flow tothe divertor of 105 MW and an average surface heat fluxto the first wall of 0.44 MW/m2. It is possible to increasethe radiation fraction of the power leaving the plasma byinjection of impurity ions in the edge, and ITER uses thisscheme to achieve a 15/85 split between charged particleand radiation. This same split would result in a powerflow to the divertor of 31.5 MW and an average surfaceheat flux on the first wall of 0.74 MW/m2. The averagepower flux of 14.1-MeV fusion neutrons through the firstwall is 1.7 MW/m2, and the scattered fusion neutron plusfission neutron power flux through the first wall is0.6 MW/m2.

IV.D.1. Divertor

The toroidal divertor for SABR was adapted7 fromthe ITER divertor (Fig. 11) to sodium coolant and asomewhat smaller plasma major radius. It comprises aseries of divertor cassettes that allow for the radialremoval of individual divertor modules through horizontaldivertor ports below the sodium pools for maintenance

Fig. 11. ITER divertor plasma heat sink.18



and replacement. The SABR divertor and first wallare cooled by the same 10-mm coolant channel structureas in the ITER design, but with sodium coolant. Thesechannels are lined with thin layers of SiC to mitigate MHDpressure drops.

Detailed heat removal calculations previously carriedout for the ITER divertor cooled with sodium7 indicatedthat the heat removal capability was more than adequate.The heat removal system was modeled using Fluent,51

a three-dimensional fluid program that solves the energyequation coupled with the Navier-Stokes equation. TheFluent model and mesh were created for a single coolingchannel. For an inlet coolant temperature of 293uC,the maximum CuCrZr temperature is 756uC, well belowthe 1500uC melting point of CuCrZr.

The smaller total surface area of the divertor in theSABR design than in the ITER design, together with similaranticipated incident total energy loads, will increase the heatload per unit area of the divertor (and on the first wall) forSABR relative to ITER. Based on the previous analysis,7

we believe that heat removal from the SABR divertor willbe adequate, but this will be checked again in the future.

IV.D.2. First Wall

The first wall consists of the plasma-facing compo-nents of the containment vessel with their associatedcoolant channels. The materials used in the first wall mustbe resilient against extreme heat and particle fluxes.Although most current tokamaks incorporate carbon tilesin the first wall, ITER will use 8.1-cm-thick beryllium-clad ODS steel. While reductions in first-wall thick-ness are anticipated in post-ITER tokamaks, SABRincorporates the ITER first-wall design but with sodiumas the coolant.

The SABR first-wall design anticipates average heatloads of *0.44 MW/m2, again, based on a roughly50/50 split between radiated power and charged particlesleaving the plasma and assuming 500-MW fusion power.As in ITER, certain panels used as limiters duringstartup will be exposed to significantly higher loads andwill make use of enhanced panels. Although higher thanaverage ITER heat loads of *0.25 MW/m2, this averageheat load is well below the design tolerances of theapproved ITER first-wall panel designs (2 MW/m2 fornormal panels and 4.7 MW/m2 for enhanced panels). Themelting point of beryllium imposes an upper limit on theallowable temperature under accident conditions of1200uC. As with the divertor, the increased heat andparticle flux per unit area over the first wall relative toITER will be mitigated by the fact that the tiles will becooled with sodium rather than water. Detailed heatremoval calculations for the previous SABR design7

indicate more than adequate thermal performance withsodium coolant, with surface temperatures well below the1200uC design limit necessary to avoid Be melting.

V. TRITIUM BREEDING BLANKET

SABR must at least be tritium self-sufficient (andideally a tritium producer for the startup of futureSABRs), which requires lithium-containing tritium breed-ing blankets near the fusion source and the neutronmultiplying subcritical fast reactor modules. The eightblanket modules used to generate tritium for SABR are(a) the ‘‘inboard’’ blanket (TB1) located just outside thefirst wall of the plasma chamber on the inboard side;(b) the ‘‘plasma’’ blanket (TB2) located above the plasmacore; (c) the ‘‘core’’ blanket (TB3) located below thefission core in regions not occupied by the divertormodule replacement ports; (d) the ‘‘outboard’’ blanket(TB4) located radially outboard from the modular sodiumpool but inside of the shielding; and (e) four wedge-shaped blankets (TB5 through TB8) that occupy spacebetween the modular sodium pools in the fissionannulus, as shown in Fig. 4.

V.A. Configuration and Materials

The blanket is composed of solid lithium orthosilicate(Li4SiO4), chosen because of its high lithium atomicdensity, low chemical reactivity, and relatively wideoperating temperature range for tritium recovery. A highlithium atomic density is desirable because of the6Li(n,a)T reactions that are utilized to breed tritium forthe plasma neutron source. The cross section for the 6Lireaction is *650 b for thermal neutrons52 and muchsmaller for 7Li; therefore, the analysis performed fortritium production was with a blanket design enriched in6Li to 90%. The tritium produced from these (n,a)reactions migrates through the grains of the solid breedingblanket into helium purge channels that carry the tritiumaway to purification and processing facilities.53–56 Thepurge and sodium coolant channels are arranged in anarray format within the tritium breeding blanket, with apitch-to-diameter ratio of 10 for both types of channels.The operating temperature window for tritium recovery52

of Li4SiO4 is Tmin 5 325uC v T v Tmax 5 925uC.

V.B. Tritium Breeding Ratio

Calculations for lithium capture and correspondingtritium production in the blanket were performed for theBOL conditions using the ERANOS fast reactor code28

in conjunction with the neutronics model discussion inSec. III. The coolant and purge channels were neglectedin these calculations due to the small volume fractioncompared to the blanket material (v1%), and thecalculation was based on one tritium atom produced forevery 6Li or 7Li capture. A useful parameter tocharacterize tritium self-sufficiency is the tritium breedingratio (TBR), which is defined as



TBR~_NNb

_NNp

, ð4Þ

where _NNb is the rate of tritium production in the blanketand _NNp is the rate of tritium burning in the plasmachamber. The TBR must be w1 by some margin to allowfor decay, losses in the processing systems, and restartafter maintenance and refueling shutdowns. A typicalvalue used to satisfy this criterion in previous fusionreactor analysis56 is 1.15, and a TBR value of 1.12 wascalculated for the SABR blanket configuration shown inFigs. 1 and 4. These blankets will be shown subsequentlyto fulfill the tritium inventory requirements for SABR.

The ERANOS S8 discrete ordinates calculation of theTBR was checked by comparison with the Monte Carloneutron transport code57 (MCNP) calculation described inSec. VI. The MCNP calculation of the TBR was v5%larger than the ERANOS calculation, with the differencesbeing attributable to slight geometrical differences in themodels and to slight differences in the definition of thefusion neutron source.

V.C. Thermal Performance

Tritium migration rates through the solid lithiumorthosilicate blanket to the purge channels are dependentupon the operating temperature of the blanket. Mostmigration occurs for blanket temperatures between 400uCand 900uC (Ref. 52), which is significantly under themelting temperature of 1250uC. A single cell of thesodium coolant tube matrix for the plasma blanket TB2

(Fig. 12), was analyzed using RELAP5 (Ref. 30).Table X summarizes the parameters for the RELAP5model. Similar calculations were carried out for thesimilar plasma blanket TB2 and the taller inboard (TB1)and outboard (TB4) blankets.

The coolant channel cell model, with appropriatedimensions, and the thermal-hydraulic parameters inTable X were used to calculate heat removal andtemperatures for the four blanket modules TB1 throughTB4. The total power generated in each blanket wasdetermined from the ERANOS neutron plus gammaheating calculation, which was assumed to be uniformlydeposited in a smaller fraction of the respective blanketclose to the neutron source (either plasma or fission) toperform a conservative heat removal analysis. The ‘‘heatdeposition region’’ is located at the bottom for the‘‘plasma’’ module adjacent to the plasma neutrons, the topfor the ‘‘core’’ module close to the fission neutron source,and assumed to be uniform over the entire blanket for the‘‘inboard’’ and ‘‘outboard’’ modules due to their smallthicknesses. Assuming that all the power was deposited inthis ‘‘heat deposition region’’ as shown for the plasmablanket as an example in Fig. 12, all four blanket moduleswere determined to be well within the tritium migrationwindow constraint of Tmin 5 325uC v T v Tmax 5925uC, as shown in Table XI.

Because of the inappropriateness of the current two-dimensional model representation of the wedge-shapedblanket zones TB5 through TB8, heat removal analysiswas not performed for these blankets. Although theseblankets contribute just under 6% of the tritiumproduction, they are relatively large and therefore imposeless stringent heat removal constraints. Since the smallerblankets with large power densities, such as the inboardtritium blanket (TB1), are shown to be properly cooledwith sodium, it is assumed that the wedge-shaped blankets(TB5 through TB8) can also be sufficiently cooled withthe present blanket cooling configuration.

V.D. Tritium Recovery and Self-Sufficiency

The value of the TBR is a preliminary check of theSABR capability to produce more tritium than it

TABLE X

Thermal-Hydraulic Parameters for RELAP5 Model

Parameter Value

Na inlet temperature 450uCHydraulic diameter 0.02 mDiameter of blanket cell 0.2 mNa mass flow rate 1 kg/sNa velocity 3.815 m/s

Fig. 12. Schematic of blanket and coolant channel cell for theplasma blanket.



consumes. However, the time dependency of the tritiuminventory must be examined to ensure that smallunplanned outages of processing systems and largeplanned shutdowns can be tolerated. Previous analysis56

has defined general analytical models to determine thetritium inventory in plasma fusion reactor systems. TheSABR system model has been simplified to that shown inFig. 13 with model parameters defined in Table XII.

Once the fuel cycle is determined, the time-dependentfusion neutron source requirement will be known and canbe used in the calculation. For now, we estimate based onpast experience that the average fusion power level over the

first year would be *260 MW and carry out the inventorycalculation at this fixed fusion power level over the year.

Equations (5) describe the tritium inventory for theblanket I1 the processing system I2, and the storagecontainment I3. The requirement for the initial inventoryin storage to start the fusion reactor was determined by theconstraint that I3 w 0:

_II1~L _NN{I1T1

{lI1 ,

I1(0)~0 ,

TABLE XI

Heat Removal Parameters for the Tritium Blanket Modules

Parameter Plasma Blanket Core Blanket Outboard Blanket Inboard Blanket

Maximum temperature 574uC 462uC 451uC 491uCMinimum temperature 450uC 450uC 450uC 450uCTotal power removed 26.1 MW 25.5 MW 9.5 MW 24.1 MWPower per tube 12 757 W 7550 W 21 316 W 670 378 WNumber of tubes 2050 3373 447 36Length of blanket 0.319 m 1.9 m 6.3 m 6.3 mLength of heat deposition 3.98 cm 23.75 cm Uniformly UniformlyNa outlet temperature 451uC 451uC 451uC 479uC

Fig. 13. Schematic of time-dependent tritium inventory model with parameters described in Table XII.



_II2~1{b

b_NNz

1{f

T2

I2{lI2 ,

I2(0)~0 ,

_II3~1{e

T2

I2{_NN

b{lI3 ,

and

I3(0)~I0 : ð5Þ

Solutions to Eqs. (5) with zero initial tritium concentra-tion for the blanket and the processing system are shownin Fig. 14. With these parameters, the tritium doublingtime for the system is *470 days, and the reactor regainsthe initial inventory amount I0 again in the storagecontainment in 250 days. If the reactor runs for 1 yearand shuts down for refueling for 90 days, there will be*2.3 kg left, which is enough to restart the reactor withthe requirement of 0.7 kg.

The above modeling is more simplistic than inRef. 53 because plasma exhaust, tritium breeding blanket,tritium breeding blanket coolant, first-wall coolant, and

cleanup and separation processing units are lumped into

one processing compartment with a composite residencetime. During operation, any one of these compartments

may be shut down for maintenance, requiring a minimum

tritium inventory Imin in the storage vessel such that thereactor can continue operation. This minimum inventory

depends most on the plasma fueling rate and fractionaltritium burnup in the plasma. Since SABR burns *15 kg

T/year, the minimum inventory was calculated to be

*1.6 kg of tritium for a 2-day processing systemshutdown. The minimum inventory is a reserve for the

processing systems and should be taken into account with

the initial condition for startup of the reactor, but ex postfacto inclusion of this minimum amount to the initial

storage containment condition did not have significant

effect on analysis, except for the doubling time because ofthe larger total inventory. Therefore, the minimum

inventory was superimposed with the initial inventory

required for startup determined by solving Eq. (5) andcalculated to be 2.3 kg.

Fig. 14. Tritium inventories in the blanket, processing systems,and storage containment for 1 year after startup.

TABLE XII

Tritium Inventory Model Parameters

Parameter Description Value

_NN T burn rate in plasmaa 9.38|1019 s{1

L TBRa 1.12T1 Residence time in blanketb 10 daysT2 Residence time in T processingb 1 daye Nonradioactive loss in processingb 0.1%l T decay constant 1.7841|10{9 s{1

f T leakage from blanket and other nonradioactive lossb 1.1%b T burn fraction in plasmab 5%I1 Steady-state T inventory in blanketc 0.451 kgI2 Steady-state T inventory in processingc 0.811 kgI3 (1 year) T inventory in storagec after 1 year 1.071 kgI0 Required startup inventoryc 0.7 kg (excluding minimum inventory)Imin Minimum inventory 2.3 kg

aERANOS (Ref. 26).bPrevious analysis.53

cEquation (5) and Fig. 12.



VI. SHIELDING

Shielding around the plasma and the modular sodiumpools, but inside the TF coils, is required to protectsensitive equipment, ensure the SABR design lifetime,and maintain safety barriers. In particular, the SC magnetsneed to be extensively shielded to limit radiation damageto the SC material and to the electrical insulating material.Because of the geometric complexity and consequentdifficulty of replacement, these magnets are designed tolast the lifetime of the plant. Thus, the radiation shieldingmust allow for 40 years of operation at 75% availability,or 30 effective full-power years (FPY) (full power means500 MW of fusion power plus *3000 MW of fissionpower), without reaching radiation damage limits on thesuperconductor or the insulator.

VI.A. Radiation Damage Limits

Radiation damage to the organic insulator in theSC winding pack can cause mechanical failure and istherefore a major concern for limiting the TF magnets’lifetime. Radiation damage to the organic insulatorscauses elongation, resulting in stresses that will eventuallylead to failure of the TF magnets. There is disagreementconcerning what values of dose organic insulatorscan withstand, ranging from 5|108 to 1010 rads.Originally, glass epoxy was the standard organicinsulator; however, as more research comes out, manydesigns are moving to Kapton and glass-filled poly-imides. The SABR design uses the glass-filled poly-imide with a generally accepted radiation damagedose limit of *1|109 rads (Refs. 58 and 59), which isthought to be the lower dose limit of glass-filled polyimideorganic insulators.

Radiation damage to the SC material is anotherconcern. The SC magnets must be kept below a criticaltemperature for superconductivity, and irradiation cancause the critical temperature for superconductivity todecrease in the SC material. The radiation design limit canbe specified in terms of neutron fluence, since photondose seems to have little effect by comparison on thesuperconductive material. The fast neutron fluence(w0.1 MeV) limit from extensive past studies is set to1.0|1019 n/cm2 (Ref. 58).

Another concern is the radiation effects on thecopper stabilizer. The copper stabilizer when irradiatedundergoes swelling and an increase in electricalresistivity. Calculations on the forced-cooled, cable-in-conduit conductors have shown that stability of the coilscan be maintained if the copper stabilizer resistivity(accounting for radiation damage and magnetoresistiv-ity) is *1.2 nV?m or lower.58 The initial resistivity forthe SABR design is an engineering conservative valueof 0.55 nV?m (similar to the initial resistivity value tothe conductor used in the MARS design60). In addition,

a design limit of 80% of the stabilizer resistivity limitwas used to determine when annealing was necessary.When the resistivity reaches the design limit, annealingat room temperature, which has been found to remove*80% of the radiation induced resistivity,58,60 wouldbe performed.

Although not a radiation damage issue, nuclearheating must be kept within a limit to ensure the cryostatsystem can keep the SC magnets below their criticaltemperature. The limit was determined by analyzing thecooling of the field coils under normal and plasma swingconditions with a cryogenic system maximum refrigera-tion capacity limit of 100 kW at 4 K (Ref. 58). Thecryogenic system was divided into four units, each havinga capacity of 25 to 30 kW. The design of the SABRcryogenic system for the SC magnets was kept to ITERspecifications (Table XIII).

VI.B. Design Process and Modeling

The MCNP5 code57 was used to determine theshielding required to meet the design specifications. TheMCNP model shown in Fig. 15 and described inTable XIV is a toroidally symmetric (R-Z) approximationof the actual geometry. Therefore, gaps due to discon-tinuities between the modular sodium pools constitutingthe annular fission core and the separations between theouter legs of the ten separate TF magnets were notmodeled. In addition, to simplify some of the modeling,material homogenization of the fission core, of the coolantthrough the shielding, divertor case, and of the SCwinding pack was performed. For simplicity and aconservative lower shielding thickness, only the divertorstructure was considered in homogenizing the divertorregion. All calculations were performed with initial BOLmaterial compositions.

The TF magnets are closest to the plasma and fissioncore and thus likely to receive the highest neutron andphoton flux of any of the magnets. Therefore, theshielding is modeled to ensure the lifetime againstradiation damage of the TF magnets. If the TF magnets

TABLE XIII

Radiation Design Limits*

Radiation Limit Value

Insulation dosage 1|109 radsNuclear heating rate 1.0 to 5.0 mW/cm3

Neutron fluence(w0.1 MeV)

1.0|1019 n/cm2

Copper stabilizerresistivity

1.2 nV?m

*From Refs. 58 and 63.



are sufficiently shielded, it is safe to assume that both theCS and the PF magnets are as well.

Neutrons cause most of the irradiation damage to thecopper and SC material; however, the photon dose,especially to the organic insulator, plays a large role aswell. Many photons are emitted from the plasma core dueto bremsstrahlung, line radiation, and recombination;however, due to their emitted energies and the total shieldthickness, these can be considered negligible compared tothe photon dose to the SC magnets that accumulate fromphotons produced from neutron inelastic scatteringcollisions and radiative capture. There are*1.8|1020 D-T reactions per second, each emittinga neutron at 14.1 MeV, when the fusion coreis operating at full power (500 MW). The fusionneutrons cause additional neutron sources from fissionand (n,2n) reactions that are also accounted for inthe calculation.

VI.C. Toroidal Field Coil Representation

Because of the complex geometry of the SC magnets,some assumptions were made to form a homogenousmedium to be used for calculations. The TF magnet wassplit into three regions: a steel case, an effective organicinsulator layer, and the toroidal winding pack (housing

the SC and the copper stabilizer). Table XV shows thevolumetric percentage of each material that was used forhomogenization to make a single TF winding packregion. The maximum dose to the organic insulator isobtained with an effective insulator layer on the insideof the TF magnet winding pack (closest to theplasma). Similar methodology for determining the doseto the organic insulator model is discussed in Refs. 64through 68.

The TF magnets will use the glass-filled polyimidefor the organic insulator to ensure that the TF coils will beable to meet the radiation dose limit. The glass-filledpolyimide design has *5% to 10% more radiationresistant properties than the glass-filled epoxy.61,65–68

VI.D. Shielding Results

The radiation design limits that determined the overallshield thickness were the organic insulator dose, thecopper resistivity limit, and the fast neutron fluence limitto the superconductor. The nuclear heating rate conditiondiscussed above appears to be satisfied for all configura-tions that satisfied the organic insulator dose, the copperresistivity, and the fast neutron fluence limits over the 30-FPY design objective for SABR. As shown in Table XVI,the present shield design protects the TF coils against

Fig. 15. MCNP model for shielding calculation of SABR.



radiation damage to the superconductor and to theinsulator for times well beyond the 30-FPY designobjective, but the copper stabilizer would exceed theradiation design limit before 30 FPY. However, bywarming the TF coils up to room temperature, many ofthe displacements can be removed (annealed). Figure 16illustrates that by annealing once, the 30-FPY objectivecan be reached without the copper resistivity exceeding80% of the limiting value and by annealing twice, the

copper stabilizer is able to achieve 40 FPY without thecopper resistivity exceeding 80% of the radiation designlimit. By annealing twice, the copper stabilizer canachieve 40.2 FPY without the copper resistivity exceed-ing 80% of the radiation damage limit.

The 120 nV?m stabilizer resistivity design limit usedin the above calculations was initially derived for theMARS conductor design operating at 1.8 K and 10 T. Anassumption that was made during the MARS calculation

TABLE XIV

Shielding Material Descriptions

Name (See Fig. 15) Materials Thickness (cm)

TF (toroidal field magnet) See Table XV 58.12Ins. (organic insulator) Effective layer of glass-filled polyimide 4.42TF case SS316LN-IG (stainless steel)

Outer side 7.08Inner side (next to plasma) 20.48

VV 50 vol % ODS steel, 50 vol % He 14.35Graphite Graphite with 10 vol % Na 7FW (first wall) part 1 Beryllium 1FW part 2 Mix of ODS steel, Na, and CuCrZr 2.2FW part 3 80 vol % ODS steel, 20 vol % Na 4.9OB4C B4C with 5 vol % Na 6.35OShield-1 WC with 5 vol % Na 36OShield-2 WC with 5 vol % Na 32.4OShield-3 WC with 5 vol % Na 18OShield-4 WC with 5 vol % Na 33IB4C-1 B4C with 10 vol % Na 6.5IB4C-2 B4C with 10 vol % Na 7IB4C-3 B4C with 10 vol % Na 6IB4C-4 B4C with 10 vol % Na 10IShield-1 WC with 10 vol % Na 12IShield-2 WC with 10 vol % Na n/aIShield-3 WC with 10 vol % Na 10IShield-4 WC with 10 vol % Na 10Trit-1 Li4SiO4 6.7Trit-2 Li4SiO4 31.9Trit-3 Li4SiO4 Occupies the space under

the Na pool not used for thedivertor replacement ports

Trit-4 Li4SiO4 28

TABLE XV

Approximate TF Winding Pack Material Representation

MaterialApproximate Volume

Percent Density (g/cm3)

Nb3Sn (SC, non-Cu) 6.4 3.6 (Ref. 72)Copper 13.91 8.96Helium (coolant) 13.1 0.000179SS316LN-IG (stainless steel) 48.32 7.99Glass-filled polyimide (insulation) 18.27 1.42



was a heat removal transfer limit of 1 W/cm2 of areawetted by liquid helium.60 This limit could varysomewhat when changing the winding pack design orthe cryogenic system. This 120 nV?m stabilizer resistivitydesign limit is considered to be the standard for cable-in-conduit SC coils in general. Additional information onhow radiation affects the copper resistivity and thepreradiation resistivity values of copper processed bydifferent methods can be found in Ref. 60.

To calculate the copper stabilizer radiation damage,the incoming neutron flux through the TF winding packwas distributed into energy bins. These fluxes were thenmultiplied by copper resistivity damage rates for eachneutron group, to give the average additional copperresistivity per year.60 The minimum FPY time until thefirst annealing cycle is required was calculated using 80%of the resistivity design limit and an initial resistivity of0.55 nV?m to receive *16.5 years (Ref. 60). The timesafter the subsequent anneal that a second and then thirdanneal would be required are *13.2 and 10.5 years,respectively, as indicated in Fig. 16.

It should be noted that these calculations were madefor the maximum 500 MW of fusion power and thatSABR would be expected to operate at fusion powers of

200 to 500 MW, so that the actual FPY before reachingthe design limit on fast neutron fluence and gamma dosescould be 50% to 100% greater than those shown inTable XVI. In any case, we have confidence that the SCmagnets would survive against radiation damage for the30-FPY design lifetime.

The nuclear heating rate is found assuming that allenergy deposition from neutrons and photons in the TFmagnet is converted to heat. The maximum rate of nuclearheating was determined to be *0.064 mW/cm3, which iswell below the design limit. Additional heat from neutronactivated isotopes will result in a slightly larger thanpredicted nuclear heating rate. Some of the isotopes thatcould result in additional heating are 16N, 62Cu, 64Cu,56Mn, and more.61,62 Activation will not greatly affectthe nuclear heating because the neutron flux in the TFmagnet is very small, and the magnet materials werechosen to minimize the activation.62,69 Additionalinformation supporting the activation data and calcula-tions can be found in Refs. 70 and 71; Ref. 72 providesfurther information on flux jump limitations to radiationin magnets.

VII. SUMMARY

The SABR is a 3000-MW(thermal) modular sodiumpool fast reactor of the IFR/PRISM type fueled bymetallic TRU fuel processed from discharged LWR fuel.The neutron source is a D-T fusion tokamak based onITER physics and technology. Several issues related tothe integration of fission and fusion technologies havebeen addressed in an upgraded conceptual design ofSABR. The status of the physics and technology base forSABR is such that it would be technologically feasible tocarry out the additional supporting research and devel-opment and deploy a first SABR within 25 years.

Refueling a sodium pool fast reactor located within themagnetic coil configuration of a tokamak is a majorengineering challenge. We propose a design concept inwhich the annular fast reactor consists of ten 300-MW(thermal) modular Na pools, each containing 80 fuelassemblies configured in four layers (or rows) radially to

TABLE XVI

Minimum FPY Until Radiation Design Limit Is Met for Any Location on the TF Coil Section

TF SectionFPY Limited by Fast

Neutron FluenceFPY Limited byInsulator Dose

FPY Limited by Stability Resistance forCryostability (wt% Anneal)

Inside 107.8 61.6 26.1Top 97.5 39.1 26.3Outside 1|106 1|106 3|104

Bottom 143.7 56.8 35.6

Fig. 16. Copper stabilizer resistivity at 10 T in the TF magnetcable-in-conduit coils of the SABR design as afunction of FPY.



accommodate up to a four-batch fuel cycle, and an IHX.An entire Na pool module would be disconnected from thesecondary heat removal system and removed radiallythrough a refueling port centered at the plasma midplanewith dimensions 4.4 m in the vertical direction betweenadjacent poloidal ring coils and w4.4 m in the toroidaldirection between adjacent TF coils. The removed modulewould be loaded into a transport cask and relocated to a hotcell for refueling and subsequent reloading into SABR.Other Na pool modules would be disconnected fromthe secondary heat removal system, rotated toroidallyto the position of the refueling port, and similarlyremoved to the hot cell for refueling and subsequentreloading. During operation, the refueling port wouldbe filled with a shielding plug, which would be removedfor refueling.

Achieving the reliable long-burn (quasi-steady-state)plasma operation necessary to meet the 75% availabilitydesign objective is a major physics challenge that is beingvigorously addressed in the worldwide tokamak researchprogram. The SABR design is based on plasma stability,confinement, and other performance parameters in theupper range of those that are routinely achieved today andthat will be further demonstrated in the operation of ITER inthe 2020s. The long-burn plasma operation will beaccomplished with noninductive drive of 4.5 MA of plasmacurrent plus ‘‘bootstrap’’ self-drive of 5.5 MA of plasmacurrent by pressure gradients within the plasma.Electromagnetic wave H&CD power will be used for thispurpose (70-MW EC frequency power at 170 GHz and40-MW LH frequency power at 5 GHz). Access for thiselectromagnetic H&CD power to the toroidal plasmasurrounded by a modular sodium pool fast reactor annulusis a substantial engineering challenge that was addressed byinserting small (*1 m toroidally) H&CD access portsbetween the modular Na pools at two locations separated by180 deg toroidally. The location of these H&CD accessports adjacent to opposite TF coils allows the rotation of Napool modules for refueling discussed above.

The ITER SC magnet system was used for the SABRdesign with only minor modification. The solenoidalmagnet core was reduced because the inductive fluxrequirement for SABR is less than for ITER, and thenumber of TF coils was reduced from 18 to 10 (whichrequired &10% increase in coil current to obtain the same11.8 T field at the conductor as in ITER).

The ITER first-wall and divertor systems were scaleddown for the smaller SABR plasma and converted to Nacoolant, using the same cooling channels. Suppression ofMHD effects on the flow of a liquid metal coolant in amagnetic field was accomplished by lining coolant tubeswith SiC to eliminate any current paths into metalstructures. A similar technique was employed to preventcurrent flows in fuel pins, fuel assembly walls, etc.

Tritium breeding Li4SiO4 blankets are locatedinboard of and above the plasma, below the fast reactor,

outboard of the Na pool, and in wedge-shaped modulesbetween the modular sodium pools in the fission annulus.This configuration resulted in a BOL TBR 5 1.12.Dynamic tritium inventory calculations based on thisBOL configuration suggest that SABR would be tritiumself-sufficient, even allowing for downtime for refuelingand estimated tritium losses, and could even produceadditional tritium to start up subsequent SABRs.

Radiation shielding calculations at a constant 500-MW fusion power indicate that the SC magnets couldachieve the lifetime design objective of 40 FPY at 75%availability, with one annealing of the copper resistivityincrease. Since the actual fusion power will vary from*200 to 500 MW (and the high-energy fusion neutronsare primarily responsible for the damage), these resultsindicate that the SC magnets probably would achieve thedesign lifetime of SABR without annealing.

The purpose of this upgraded conceptual design ofSABR is to provide a realistic basis for fuel cycleoptimization analyses, dynamic safety analyses to assurethe passive safety of both the fission core and the fusionneutron source, and other performance analyses that willcontribute to assessing the deployment of a SABR to helpclose the back end of the nuclear fuel cycle. We plan tocarry out such analyses in the near future and hope thatothers will do likewise.

ACKNOWLEDGMENTS

The work reported in this paper was performed by studentsin a graduate design course in the Nuclear & RadiologicalEngineering Program at Georgia Tech under the guidance ofprofessors W.M.S. and A.S.E. The input and advice of manyprofessionals in the field is gratefully acknowledged, inparticular, E. A. Hoffman of ANL, whose comments guidedthe fuel assembly design process; M. A. Abdou of University ofCalifornia, Los Angeles, whose comments guided theselection of SiC for electrical insulation of metal structures;and B. E. Nelson of Oak Ridge National Laboratory, whoguided us to current information on the ITER design. One ofthe authors (C.L.S.) gratefully acknowledges the support of theU.S. Department of Energy Office of Nuclear Energy’s NuclearEnergy University Programs.

REFERENCES

1. B. R. LEONARD, JR., ‘‘A Review of Fusion-Fission(Hybrid) Concepts,’’ Nucl. Technol., 20, 161 (1973); http://dx.doi.org/10.13182/NT73-1.

2. L. M. LIDSKY, ‘‘Fusion-Fission Systems: Hybrid,Symbiotic and Augean,’’ Nucl. Fusion, 15, 151 (1975); http://dx.doi.org/10.1088/0029-5515/15/1/016.

3. A. A. HARMS, ‘‘Hierarchical Systematic of Fusion-FissionEnergy Systems,’’ Nucl. Fusion, 15, 939 (1975); http://dx.doi.org/10.1088/0029-5515/15/5/023.





http://dx.doi.org/10.1088%2F0029-5515%2F15%2F1%2F016




4. H. A. BETHE, ‘‘The Fusion Hybrid,’’ Nucl. News, 21, 7, 41(May 1978).

5. R. W. MOIR, ‘‘The Fusion Breeder,’’ J. Fusion Energy, 2,351 (1982); http://dx.doi.org/10.1007/BF01063686.

6. R. W. MOIR et al., ‘‘Fusion Breeder Reactor DesignStudies,’’ Nucl. Technol./Fusion, 4, 4, part 2, p. 589 (1983).

7. W. M. STACEY et al., ‘‘A TRU-Zr Metal Fuel SodiumCooled Fast Subcritical Advanced Burner Reactor,’’ Nucl.Technol., 162, 53 (2008); http://dx.doi.org/10.13182/NT08-1.

8. W. M. STACEY, ‘‘Georgia Tech Studies of Sub-CriticalAdvanced Burner Reactors with a D-T Fusion Tokamak NeutronSource for the Transmutation of Spent Nuclear Fuel,’’ J. FusionEnergy, 38, 328 (2009); http://dx.doi.org/10.1007/s10894-009-9195-0.

9. M. KOTSCHENREUTHER et al., ‘‘Fusion-FissionTransmutation Schemes—Efficient Destruction of NuclearWaste,’’ Fusion Eng. Des., 84, 83 (2009); http://dx.doi.org/10.1016/j.fusengdes.2008.11.019.

10. W. M. STACEY, ‘‘Principles and Rationale of the Fusion-Fission Hybrid Burner Reactor,’’ AIP Conf. Proc., 1442, 31(2012).

11. R. W. MOIR et al., ‘‘Axisymmetric Magnetic MirrorFusion-Fission Hybrid,’’ LLNL-TR-484071, LawrenceLivermore National Laboratory (2011).

12. C. L. STEWART and W. M. STACEY, ‘‘The SABrRConcept for a Fission-Fusion Hybrid 238U-to-239Pu FissileProduction Reactor,’’ Nucl. Technol., 187, 1 (2014); http://dx.doi.org/10.13182/NT13-102.

13. C. E. TILL, Y. I. CHANG, and W. H. HANNUM, ‘‘TheIntegral Fast Reactor—An Overview,’’ Prog. Nucl. Energy, 31,3 (1997); http://dx.doi.org/10.1016/0149-1970(96)00001-7.

14. Y. I. CHANG et al., ‘‘Advanced Burner Test ReactorPreconceptual Design Report,’’ ANL-ABR-1, Argonne NationalLaboratory (2008).

15. E. A. HOFFMAN, W. S. YANG, and R. N. HILL,‘‘Preliminary Core Design Studies for the Advanced BurnerReactor over a Wide Range of Conversion Ratios,’’ ANL-AFCI-177, Argonne National Laboratory (2008).

16. A. E. DUBBERLEY et al., ‘‘Super-PRISM Oxide andMetal Fuel Core Designs,’’ Proc. 8th Int. Conf. NuclearEngineering (ICONE 8), Baltimore, Maryland, April 2–6,2000.

17. M. SHIMADA et al., ‘‘Chapter 1: Overview andSummary,’’ Nucl. Fusion, 47, S1 (2007); http://dx.doi.org/10.1088/0029-5515/47/6/S01.

18. ITER Web Site: www.iter.org.

19. T. S. SUMNER, W. M. STACEY, and S. M.GHIAASIAAN, ‘‘Dynamic Safety Analysis of the SABRSubcritical Transmutation Reactor Concept,’’ Nucl. Technol.,171, 123 (2010); http://dx.doi.org/10.13182/NT10-4.

20. C. M. SOMMER, W. M. STACEY, and B. PETROVIC,‘‘Fuel Cycle Analysis of the SABR Subcritical Transmutation

Reactor Concept,’’ Nucl. Technol., 172, 48 (2010); http://dx.doi.org/10.13182/NT10-3.

21. C. M. SOMMER et al., ‘‘Transmutation Fuel CycleAnalyses of the SABR Fission-Fusion Hybrid Burner Reactorfor Transuranic and Minor Actinide Fuels,’’ Nucl. Technol.,182, 274 (2013); http://dx.doi.org/10.13182/NT11-55.

22. V. ROMANELLI et al., ‘‘Comparison of the WasteTransmutation Potential of Different Innovative DedicatedSystems and Impact on the Fuel Cycle,’’ Proc. Int. Conf.Engineering Nuclear Energy Systems (ICENES), San Francisco,California, May 15–19, 2011.

23. M. ABDOU, ‘‘Introduction to MHD and Applications toThermofluids of Fusion Blankets,’’ Lecture, Institute for PlasmaResearch, Gandhinagar, India (2007); http://www.fusion.ucla.edu/abdou/Presentation%20Web%20Page.html#2007.

24. M. ABDOU, University of California, Los Angeles,Personal Communication (2013).

25. M. L. HAMILTON, et al., ‘‘Fabrication TechnologicalDevelopment of the Oxide Dispersion Strengthened AlloyMA957 for Fast Reactor Applications,’’ Pacific NorthwestNational Laboratory (Feb. 2000).

26. D. T. HOELZER et al., ‘‘Creep Behavior of MA957 and14YWT (SM10 heat),’’ Oak Ridge National Laboratory;http://web.ornl.gov/sci /physical_sciences_directorate/mst/fusionreactor/pdf/Vol.48/3.2_p50-59_Hoelzer.pdf.

27. B. WILSHIRE and T. D. LIEU, ‘‘Deformation and DamageProcesses During Creep of Incoloy MA957,’’ Mater. Sci. Eng. A,386, 81 (2004); http://dx.doi.org/10.1016/j.msea.2004.07.047.

28. G. RIMPAULT et al., ‘‘The ERANOS Code and Data-System for Fast Reactor Analyses,’’ Proc. Int. Conf. NewFrontiers of Nuclear Technology: Reactor Physics, Safety andHigh-Performance Computing (PHYSOR), Seoul, Korea,October 7–10, 2002, American Nuclear Society (2002).

29. C. BOARDMAN, ‘‘GE Advanced Liquid Metal ReactorS-PRISM,’’ Proc. Advisory Committee on NuclearSafeguards Workshop, Rockville, Maryland, June4–5, 2001, U.S. Nuclear Regulatory Commission (2001).

30. ‘‘RELAP5-3D Home Page’’: http://www.inl.gov/relap5/,Idaho National Laboratory (Feb. 2013).

31. N. TODREAS and M. KAZIMI, Nuclear Systems,Volume 2, p. 77, Hemisphere Publishing Corporation (1990).

32. W. M. STACEY, Fusion Plasma Physics, Wiley-VCHVerlag, Berlin, Germany (2005).

33. J.-P. FLOYD et al., ‘‘Tokamak Fusion Neutron Sourcefor a Fast Transmutation Reactor,’’ Fusion Sci. Technol., 52,727 (2007); http://dx.doi.org/10.13182/FST07-1.

34. C. GORMEZANO et al., ‘‘Chapter 6: Steady StateOperation,’’ Nucl. Fusion, 47, S285 (2007); http://dx.doi.org/10.1088/0029-5515/47/6/S06.

35. F. NAJMABADI et al., ‘‘The ARIES-AT AdvancedTokamak, Advanced Technology Fusion Power Plant,’’Fusion Eng. Des., 80, 3 (2006); http://dx.doi.org/10.1016/j.fusengdes.2005.11.003.



http://dx.doi.org/10.1007%2FBF01063686


http://dx.doi.org/10.1007%2Fs10894-009-9195-0

http://dx.doi.org/10.1007%2Fs10894-009-9195-0

http://dx.doi.org/10.1016%2Fj.fusengdes.2008.11.019




http://dx.doi.org/10.1016%2F0149-1970%2896%2900001-7

http://dx.doi.org/10.1088%2F0029-5515%2F47%2F6%2FS01


www.iter.org





http://www.fusion.ucla.edu/abdou/Presentation%20Web%20Page.html#2007



http://web.ornl.gov/sci/physical_sciences_directorate/mst/fusionreactor/pdf/Vol.48/3.2_p50-59_Hoelzer.pdf

http://web.ornl.gov/sci/physical_sciences_directorate/mst/fusionreactor/pdf/Vol.48/3.2_p50-59_Hoelzer.pdf

http://dx.doi.org/10.1016%2Fj.msea.2004.07.047

http://www.inl.gov/relap5/

http://dx.doi.org/10.13182/FST07-1





36. V. S. CHAN et al., ‘‘Physics Basis of Fusion DevelopmentFacility Utilizing the Tokamak Approach,’’ Fusion Sci.Technol., 57, 66 (2010); http://dx.doi.org/10.13182/FST10-1.

37. J. DECKER et al., ‘‘Calculations of Lower Hybrid CurrentDrive in ITER,’’ Nucl. Fusion, 51, 073025 (2011); http://dx.doi.org/10.1088/0029-5515/51/7/073025.

38. E. POLI et al., ‘‘Electron-Cyclotron-Current-DriveEfficiency in DEMO Plasmas,’’ Nucl. Fusion, 53, 013011(2013); http://dx.doi.org/10.1088/0029-5515/53/1/013011.

39. G. L. JACKSON et al., ‘‘Noninductive Plasma Initiationand Startup in the DIII-D Tokamak,’’ Nucl. Fusion, 51, 083015(2011); http://dx.doi.org/10.1088/0029-5515/51/8/083015.

40. ITER Technical Basis, Sec. 2.5, ‘‘Additional Heating andCurrent Drive,’’ International Atomic Energy Agency (2001).

41. R. J. LA HAYE et al., ‘‘Cross-Machine Benchmarking forITER of Neoclassical Tearing Mode Stabilization by ElectronCyclotron Current Drive,’’ Nucl. Fusion, 46, 451 (2006); http://dx.doi.org/10.1088/0029-5515/46/4/006.

42. F. IMBEAUX et al., ‘‘Current Ramps in Tokamaks: FromPresent Experiments to ITER Scenarios,’’ Nucl. Fusion, 51,083026 (2011); http://dx.doi.org/10.1088/0029-5515/51/8/083026.

43. S. H. KIM et al., ‘‘Lower Hybrid Assisted Plasma CurrentRamp-Up in ITER,’’ Plasma Phys. Controlled Fusion, 51,065020 (2009); http://dx.doi.org/10.1088/0741-3335/51/6/065020.

44. G. M. D. HOGEWEIJ et al., ‘‘Optimizing the CurrentRamp-Up Phase for the Hybrid ITER Scenario,’’ Nucl. Fusion,53, 013008 (2013); http://dx.doi.org/10.1088/0029-5515/53/1/013008.

45. V. E. LUKASH et al., ‘‘Simulation of ITER PlasmaScenarios Starting from Initial Discharge of Central Solenoid,’’Proc. 38th Conf. Plasma Physics, Strasbourg, France, June 27–July 1, 2011, European Physical Society.

46. A. EJIRI et al., ‘‘Non-Inductive Plasma Current Start-Up byEC and RF Power in the TST-2 Spherical Tokamak,’’ Nucl.Fusion, 49, 065010 (2009); http://dx.doi.org/10.1088/0029-5515/49/6/065010.

47. R. RAMAN et al., ‘‘Non-Inductive Plasma Start-Up andCurrent Ramp-Up in NSTX-U,’’ Proc. 54th Annual Mtg. APSDivision of Plasma Physics, Providence, Rhode Island, October29–November 2, 2012, American Physical Society.

48. F. G. RIMINI et al., ‘‘Electron Cyclotron Heating andCurrent Drive Studies During Current Ramp-Up in Tore-Supra,’’ Plasma Phys. Controlled Fusion, 47, 869 (2005); http://dx.doi.org/10.1088/0741-3335/47/6/009.

49. ITER Technical Basis, Sec. 2.1, ‘‘Magnets,’’ InternationalAtomic Energy Agency (2001).

50. K. TOBITA et al., ‘‘Reduction of Energetic Particle Loss byFerritic Steel Inserts in ITER,’’ Plasma Phys. Controlled Fusion45, 133 (2003); http://dx.doi.org/10.1088/0741-3335/45/2/305.

51. ‘‘FLUENT/GAMBIT Version 6.1.22/s.12,’’ ComputationalFluid Dynamics, Fluent, Inc. (2004).

52. W. M. STACEY, Fusion, 2nd ed., Chap. 10, Wiley-VCH,Weinheim (2010).

53. P. ANDREW et al., ‘‘Tritium Retention and Clean Up inJET,’’ Fusion Eng. Des., 47, 233 (1999); http://dx.doi.org/10.1016/S0920-3796(99)00084-8.

54. C. H. SKINNER et al., ‘‘Studies of Tritiated Co-DepositedLayers in TFTR,’’ J. Nucl. Mater., 290–293, 486 (2001); http://dx.doi.org/10.1016/S0022-3115(00)00643-7.

55. A. A. HAASZ and J. W. DAVIS, ‘‘Hydrogen Retention inand Removal from Carbon Materials,’’ Springer Series inChemical Physics, Vol. 78, Springer Publishing Company,Berlin, Germany (2005).

56. M. A. ABDOU et al., ‘‘Deuterium-Tritium Fuel Self-Sufficiency in Fusion Reactors,’’ Fusion Technol., 9, 250 (1986).

57. X-5 MONTE CARLO TEAM, ‘‘A General Monte Carlo N-Particle Transport Code, Version 5, Volume I, MCNP Overviewand Theory,’’ LA-UR-03-1987, Rev. October 3, 2005, LosAlamos National Laboratory (2003).

58. C. HENNING, ‘‘Magnet Design Technical Report—ITERDefinition Phase,’’ UCID-21681, Lawrence Livermore NationalLaboratory (1989).

59. ‘‘Annex to Design Requirements and Guidelines Level 1(DRG1): ITER Structural Design Criteria for MagnetComponents (SDC-MC),’’ N 11 FDR 50 01-07-05 R 0.1,ITER.

60. M. W. GUINAN, ‘‘Radiation-Effects Limits on Copper inSuperconducting Magnets,’’ UCID-19800; CONF-830466-4,Lawrence Livermore National Laboratory (1983).

61. H. Y. ATTAYA, Y. GOHAR, and D. SMITH, ‘‘US–ITERActivation Analysis,’’ CONF-901007-49, Argonne NationalLaboratory (1990).

62. S. J. PIET, ‘‘Exploratory Study of Burn Time, Duty Factor,and Fluence on ITER Activation Hazards,’’ EGG-FSP-10396,EG&G Idaho (1992).

63. Proc. Mtg. Electrical Insulators for Fusion Magnets,Germantown, Maryland, 1981, U.S. Department of Energy (1981).

64. M. A. SAWAN, ‘‘Application of FENDL Nuclear Data toITER Neutronics and Shielding Analysis,’’ UWFDM-1006,University of Wisconsin (1996).

65. R. R. COLTMAN, JR., C. E. KLABUNDE, and C. J.LONG, ‘‘The Effect of a 100 MGy (1010 rads) Gamma-RayDose at 5 K on the Strength of Polyimide Insulators,’’ SSD 81-10, Oak Ridge National Laboratory (1981).

66. M. IWATA et al., ‘‘Properties of Newly DevelopedThermoplastic Polyimide and Its Durability Under Proton,Electron, and UV Irradiations,’’ Proc. 11th Int. Symp. Materialsin a Space Environment, Aix-en-Provence, France, September15–18, 2009.

67. N. J. SIMON, ‘‘A Review of Irradiation Effects on Organic-Matrix Insulation,’’ NISTIR-3999, National Institute ofStandards and Technology (1993).

68. W. MAURER, ‘‘Neutron and Gamma Irradiation Effects onOrganic Insulating Materials for Fusion Magnets,’’Kernforschungszentrum Karlsruhe (1985).



http://dx.doi.org/10.13182/FST10-1
















http://dx.doi.org/10.1016%2FS0920-3796%2899%2900084-8

http://dx.doi.org/10.1016%2FS0920-3796%2899%2900084-8

http://dx.doi.org/10.1016%2FS0022-3115%2800%2900643-7

http://dx.doi.org/10.1016%2FS0022-3115%2800%2900643-7

69. D. G. CEPRAGA et al., ‘‘Neutronics and ActivationCalculation for ITER Generic Site Safety Report,’’ FusionEng. Des., 63, 193 (2002); http://dx.doi.org/10.1016/S0920-3796(02)00130-8.

70. M. E. SAWAN and H. Y. KHATER, ‘‘Initial NuclearPerformance Evaluation of the FIRE Ignition Device,’’ Proc.

18th Symp. Fusion Engineering, Albuquerque, New Mexico,October 25–29, 1999, IEEE (1999).

71. M. E. SAWAN and H. Y. KHATER, ‘‘Nuclear Analysis ofthe FIRE Ignition Device,’’ Fusion Technol., 39, 2, 393 (2001).

72. V. V. KASHIKHIN, ‘‘Flux Jumps in Nb3Sn Magnets,’’TD-03-005, Fermilab (2003).



http://dx.doi.org/10.1016%2FS0920-3796%2802%2900130-8

http://dx.doi.org/10.1016%2FS0920-3796%2802%2900130-8

RESOLUTION OF FISSION AND FUSION TECHNOLOGY …€¦ · be based on plasma stability, confinement,...

Documents

Transcript of RESOLUTION OF FISSION AND FUSION TECHNOLOGY …€¦ · be based on plasma stability, confinement,...