ANL/FPP/TM-128
THE IMPACT OF ALTERNATE FUSION FUELS OK FUSION REACTOR TECHNOLOGY
An Initial Assessment Study
FUSION POWER PROGRAM
Aiftnit MctioMl Latentary 97M Suva Uss AVMW U. S. Bflpirtwun if Ewf|y ArfMM, !tf TW8 SOCgf^lT IS TABLE OF CONTENTS Pa ABSTRACT vi 1.0 Introduction ...... 1
1.1 Study Objectives 2 1.2 Study Methodology > 2 1.3 Brief Review of Advanced Fuel Cycles 3
2.0 Plasma Engineering 9
2.1 Introduction. 9 2.2 Tokamak Concepts Studies...... 9
2.2.1 Background and Motivation. 9 2.2.2 The Tokamak Global Code 10 2.2.3 Field and Current Limitations. 11
2.2.4 Cat-D Tokamaks 13
2.3 FRM Concept Studies 15
2.4 Cyclotron Radiation Considerations 21
2.4r.l Introduction 21 2.4.2 Effects of Nonunifortn Plasma Fro files 22
2.4.3 Effects of Holes in the Wall 32
2.5 Review of the p-6Li Fusion Chain Reaction 34
2.5.1 Introduction 34 2.5.2 The p-5Li Fusion Chain Reaction 35 2.5.3 Energy Balance 36 2.5.4 Neutron and Radioactive Ash Yields 39 2.6 Summary of Typical Power Splits and Sensitivity
Studies 42
2.6.1 Typical Power Splits 42
3.0 Engineering/Technology Considerations...... 49
3,1 Materials, First-Wall and Blanket Considerations. .... 49 3.1.1 First-Wall Surface Effects 50 3.1.2 Structural Materials 52 3.1.3 Candidate Coolants » . . . 52 3.1.4 General First-Wall/Blanket Design Considerations . 53 Table of Contents (cont'd.)
Page
3.2 Nuclear Analysis 54
3.2.1 Scope of Analysis 56 3.2.2 Radiation Damage to the First Wall 61 3.2.3 Blanket and Shield Performance 64 3.2.4 Radiation Damage to Superconducting Magnets .... 71 3.2.5 Neutron Energy Spectra in Alternate Fuel Systems. . 74 3.2.6 Reactor Activation and Environmental Impact .... 74 3.2.7 Biological Shielding 84 3.2.8 Major Penetration Shielding . 87 3.3 Tritium and Fuel Processing Considerations 92
3.3.1 Fuel Cycle Considerations 92 3.3.2 Vacuum Pumping 96 3.3.3 Fuel Processing and Tritium Safety. . 96 3.3.4 Fuel Supply 98 3.3.5 Costs . 99 3.3.6 Conclusions 100
3.4 Magnet Design Considerations for Alternate Fuel Tokamak Reactor 100
3.4.1 Geometrical Limits on Plasma Current and Toroidal Field 100
3.4.2 Irradiation Effects on Nb3Sn Superconductors. . . . 102
3.5 Safety 102
4.0 Summary and Conclusions 110 List of Figures
Number Page
1-1 Reaction parameters and cross sections for various fusion reactions. The reaction parameter is average over a Maxwellian ion distribution. The curves shown for p-6Li, p-9Be, and p-1:lB contain large uncertainties. Five strong D-6Li reactions occur with different
Number Page
3-7 A comparison of alternate fuel systems on neutron spectrum in blanket structural material: 316 stainless steel ...... 75 3-8 A comparison of alternate fuel systems on fraction spectrum in blanket structural material: 316 stainless steel 76 3-9 A comparison of alternate fuel systems on biological hazard potential structural material: 316 stain- less steel 80 3-10 A comparison of advanced fuel systems on biological hazard potential structural material: V-15Cr-5Ti 81 3-11 A comparison of alternate fuel systems on biological hazard potential structural material: T14381 82 3-12 Isotopic contribution to BHP-alr in Cat-D system structural material: 316 stainless steel 83 3-13 A comparison of alternate fuel systems on 63Ni isotope contribution to BHP-air, structural material: 316 stainless steel ..... 85 3-14 Effect of alternate fuel systems on biological shield requirement on normal concrete, structural material: 316 stainless steel 86 3-15 A comparison of alternate fuel system on biological dose at reactor shutdown, structural material: 316 stainless steel 88 3-16 A comparison of alternate fuel systems on penetration shield requirement, structural material: 316 stain- less steel 90 3-17 Effect of alternate fuel systems on biological shielding during reactor operation 91 3-18 Fuel cycle scenario for Cat-D 93 3-19 Tritium facility scenario for Cat-D .... 94 3-20 Fuel processing cycle for D-3He 95 3-21 Maximum plasma current and toroidal fields for alternate fuel tokamak reactors 101 3-22 Relative effect of Ti4381 structural material on different alternate fuel systems 104 3-23 Relative effect of V-15Cr-%ti structural material on different alternate fuel systems. 3-24 Relative effect of 316 stainless steel sti -tural material on different alternate fuel systems ..... 105 List of Tables
Number Page
1-1 Fuel Characteristics 5 1-2 Fuel/Confinement Alternatives 7 2-1 Reference Tokamak Reactor Designs 12 2-2 Reference FRM Reactor Designs 20 2-3 Neutron and Radioactive Ash Yields as a Function of Temperature for p-6Li 41 2-4 Power Splits for Tokamaks 44 2-5 Power Splits for FRM 44 2-6 Power Splits for Cat-D Tokamak 46 3-1 Comparison of Power Splits and Allowable Wall Loadings for Alternate Fuel Cycles „ 49 3-2 Thermal Stress Factors for Candidate Structural Alloys . . 53 3-3 Characteristics of the Alternate Fuel Systems Studied. . . 58 3-4 Structural Material Composition ... 59 3-5 System Dimensions and Material Compositions ...... 60 3-6 Nuclear Radiation Response Rates at the First Wall .... 62 3-7 Shielding Performance of Alternate Fuel Systems for a Total Integral Wall Load of 30 MW-yr/m2 69 3-8 System Energy Multiplication of Advanced Fuel Systems per Source Neutron ...... 70 3-9 Maximum Response Rates in Superconducting Magnets .... 72 3-10 A Comparison of Short-Term Radiological Impacts of Alternate Fuel Systems 78 3-11 Fuel Processing Requirements for Alternate Fusion Fuel Cycles 97 3-12 Tritium Inventories (g) 98 3-13 Estimated Tritium and Vacuum Costs ($M) for Various 2500 MWth Reactors 99 ABSTRACT
The initial results of a study carried out to assess some of the technology implications of non-D-T fusion fuel cycles are presented. The primary emphasis in this report is on D-D, catalyzed-D and D-3He fuel cycles. Tokamaks and field-reversed mirrors have been selected as sample confinement concepts. A new technique of employing neutronic computer codes to study the transport of cyclotron radiation for cases of non-uniform density and temperature profiles is described. The technology areas con- sidered include first wall design considerations, shielding require- ments, fuel cycle requirements and some safety and environmental considerations. Conclusions resulting from the study are also presented. 1.0 Introduction
This report describes the initial results of a program to examine the fusion reactor technology impacts of non-deuterium/tritium fusion fuel cycles. A variety of elements other than deuterium and tritium can undergo fusion; examples include D-D, D-3He, 3He-3He, p-6Li, p-7Li, D-6Li, and p-11B. These fuels have, in varying degree, the generic features of reduced neutron pro- duction, increased fusion energy carried by charged particles, and the elimi- nation of a need for tritium breeding. On the other hand, the combination of lower cross sections, higher plasma temperature, lack of availability of some fuels (3He), and increased radiation losses make efficient confinement (i.e., high energy multiplication) more difficult. A key consideration then is whether or not the advantages are indeed sufficient to justify the develop- ment of alternate fuel power plants. A quantitative evaluation of the tech- nology involved must be undertaken.
Deuterium-based fuels have the advantage of operating at relatively low temperatures but involve more neutron and tritium production via "side" D-D reactions. Lower temperature deuterium based fuels, being easier to burn, are compatible with a wide range of confinement concepts; for example, tokamaks could burn catalyzed deuterium and D-3He. (Catalyzed deuterium refers to burning deuterium such that the reaction products of tritium and 3He are also burned at a rate equal to its birth rate. The resulting energy release allows operation at lower temperatures.) But, unfortunately, the attribute of "cleanliness" is not as fully achieved in the deuterium-based fuels as the more ideal proton-based fuels. The proton fuels, on the other hand, due to a combination of the high temperature required and modest energy released per reaction, appear much harder to successfully burn, forcing the use of more advanced confinement approaches. A more extensive review of the different fuel cycles is present in Sec. 1.3 below.
The next two subsections briefly describe the objectives and general methodology used in this study. Section 2.0 describes various plasma engi- neering considerations regarding alternate fusion fuels while Sec. 3.0 reports on the reactor technology considerations of this study which include first wall materials, neutronics and shielding requirements, fuel cycle considera- tions, magnet considerations and some safety issues. Section 4.0 summarizes the main conclusions from the study. 1.1 Study Objectives
The basic objective of the study reported herein was to carry out a preliminary assessment of the impact on various fusion reactor technologies of alternate fusion fuels. Here, "alternate" or "advanced fuels" refer to any fusion fuel cycle other than the deuterium-tritium-lithium fuel cycle. It is not the objective of this study to compare various fusion confinement concepts in terms of their application to different fusion fuel cycles. In general, the emphasis in this study is to examine generic reactor technology issues common to most magnetic confinement concepts. It is recognized that there are limitations to this approach and that more detailed reactor studies of particular concepts are required to assess the usefulness of each concept for the burning of alternate fusion fuels.
1.2 Study Methodology
The basic approach of the study was to develop representative parameters for the distribution of reactor power into neutrons, charged particles and electromagnetic radiation for two sample confinement concepts. These power split distributions were then used to examine some key reactor technology areas such as first wall design considerations, shielding requirements, environmental impact of activated materials, and fuel cycle design consider- ations.
In order to study "representative" combinations of confinement concepts and fuel cycles that are illustrative of a wide variety of possible combina- tions, we have selected for study a matrix of alternate fuels along with a modest and high 3 confinement systems. Since catalyzed-D is the only "self- sufficient" (only requires naturally available deuterium) low-temperature fuel, it is selected as the base fuel. However, due to its potential attractiveness, D-3He is also considered despite the complication of 3He breeding. Since tokamaks are presently the most highly developed confinement system, and since they (or at least modest 3 versions) can burn these low temperature fuels, they are selected as a representative modest 6 (10-20%) confinement system. (These 6 values are about a factor of two larger than predicted by most current theories, but are considered to be reasonable extrapolations for purpose of the study.) Concurrently, the Field-Reversed Mirror (FRM) is se- lected as a representative high-3 (> 70%) device. This is acceptable for low-temperature fuels, but may not be adequate for the higher temperature fuels where even more innovative confinement approaches such as Surmac may be necessary.
For purposes of technology considerations discussed in Sec. 3.0, the most important parameter is the fraction of total power carried by neutrons. This is relatively constant for the various cases presented in Sec. 2.0, thus it is possible to examine certain first wall/blanket/shield problems without specifying a particular type of device. This is the general approach taken in this study when the fuel cycles are characterized in terms of neutron power fractions by the following representative values:
D-T - D-D -x< 40% cat-D-D -v. 40% D-3He -v 1% .
1.3 Brief Review of Advanced Fuel Cycles
A variety of elements other than deuterium and tritium can undergo fusion with reasonable cross sections at energies less than 2 MeV. As illu- strated by the cross-section plot of Fig. 1-1, prominent examples include D-D, D-3He, 3He-3He, p-6Li, p-?Li, D-6Li, and p-llB. These "advanced fuels" have, in varying degree, the generic advantages of reduced neutron produc- tion, increased fusion energy carried by charged particles, and the elimina- tion of a need for tritium breeding. These features suggest practical advan- tages in a fusion power plant including: reduced neutron activation, reduced tritium handling/storage and reduced thermal heat rejection, improved safety by elimination of hazards associated with large tritium and liquid metal re- servoirs, simpler blanket design and maintenance, increased first wall and blanket lifetime as a result of the reduced neutron flux, and a larger number of options for efficient energy conversion. These benefits should result in an improved environmental compatibility, increased siting for flexibility and more ready public acceptance. On the other hand, the combination of lower cross sections, higher plasma temperature, lack of availability of some fuels (3He), and increased radiation losses make efficient confinement (i.e., high energy multiplication) more difficult. Consequently, the advantages noted above may be partially offset by the special features associated with i |
Iff16
tr LU h- I0"17 L±J
Q_
10',-IB
10"1-19 10 I02 PLASMA TEMPERATURE, keV
Figure 1-1. Reaction parameters and cross sections for various fusion reac- tions. The reaction parameter is averaged over a Maxwellian ion distribution. The curves shown for p-6Li, p-9Be, and p-uB contain large uncertainties. Five strong D-6Li reactions occur with different
To better understand the choices that must be considered in this pro- cess, we must review some additional aspects of alternate fuels and possible confinement approaches for them. A possible division of the fuels is into "deuterium-based" and "proton-based" mixtures. D-D, D-3He, and D-6Li into the former class, while p-6Li, p-?Li and p-HB are proton based. As outlined in Table 1-1, deuterium-based fuels have the advantage of operating at relatively low temperatures but involve more neutron and tritium production via "side" D-D reactions. These observations introduce a dilemma. The lower temperature deuterium-based fuels, being easier to burn, are compatible with a wide range of confinement concepts; for example, tokamaks with g - 12% could burn catalyzed deuterium and D-3He. (Catalyzed deuterium refers to burning deuterium such that the reaction products of tritium and He are also burned at a rate equal to its birth rate. The resulting energy release allows operation at lower temperatures.) But, unfortunately, the attribute of "cleanliness" is not as fully achieved in the deuterium-based fuels as the more ideal proton-based fuels. The proton fuels, on the other hand, due to a combination of the high temperature required and modest energy released per reaction, appear much harder to successfully burn, forcing the use of more advanced confinement approaches. Superimposed on these problems is the quandry of D-3He. It offers an almost ideal combination of minimal neutron production and relative ease of
Table 1-1. Fuel Characteristics
Proton Based Deuterium Based
Plasma Temperature High (MOO keV) Relatively Low (< 100 keV)
Neutron and Tritium Production minimal some burning. However, 3He is not naturally abundant, forcing the use of another reactor, for example a semi-catalyzed deuterium reactor, to breed 3He. Consequently, the concept of small satellite D-3He plants that receive 3He from larger, less flexible "nuclear park" type, semi-catalyzed deuterium plants has been proposed.
The concept of small satellites brings out an important feature of the confinement approach. The higher the temperature required, the more radia- tion will be generated by the plasma. Consequently, a relatively large plasma is required to minimize losses by reducing the surface to volume ratio. For this reason, the proton-based fuels tend to be naturally matched with larger devices whereas D-3He, and to a lesser extent catalyzed deuterium, are candidate fuels for smaller devices. This introduces yet another quandry. The advantages of relatively clean alternate fuels are best featured by making it possible to locate a plant near a user. This in turn suggests the use of smaller plants rather than large central power stations, or perhaps a combination of the two. However, based on present concepts, it is not clear how to use the high temperature proton fuels in a desirable small plant. At any rate an evaluation must consider a matrix of fuel types and various alternative confinement approaches.
As indicated in Table 1-2, in order to study "representative" combina- tions that illustrate extremes, we have selected for study a matrix of high- and low-temperature fuels along with a modest and high 3 confinement systems. Since catalyzed-D is the only "self-sufficient" (only requires naturally available deuterium) low-temperature fuel, it is selected as the base fuel. However, due to its unique attractiveness, D-3He is also considered despite the complication of 3He breeding. Since tokamaks are presently the most highly developed confinement system, and since they (or at least higher B versions) can burn these low tenperature fuels, they are selected as a representative modest 8 (10-20%)* confinement system. Concurrently, the FRM is selected as a representative idgh-3 (>70%) device. This is acceptable for the low-temperature fuels, but may not be adequate for the higher temperature fuel where even more drastic confinement approaches such as Surmac may be necessary. During the present year it was only possible to begin studies of p-6Li which was selected as the "representative" high
* These 3 values are about a fraction of two larger than predicted by most current theories, but are not considered to be reasonable extrapolations for purpose of the study. 6 Table 1-2. Fuel/Confinement Alternatives
Low Temperature Catalyzed-D •D-3lie
High Temperature p-6Li
Modest 6 Tokamak
High 3 Field Reverse Mirror
temperature fuel. This choice was made on the basis that, of the various high-temperature, proton-based fuels, p-eLi appears to come closest to being ignitable in magnetically confined devices.
The tokamak sets a reference standard and provides a good basis for
comparison of D-T and low-temperatures alternate fuel technology. Indeed it is a widely accepted view that the D-T tokamak will emerge as a first gener- ation fusion device. A natural question then is: assuming it is feasible, how desirable is it to develop an alternree fuel version of the D-T tokamak, A previous University of Illinois study demonstrated that catalyzed deuterium and D-3He tokamaks should be achievable and hence represent an obvious and important route. This study assumed a reasonable scaling of beta (8 - 12%) and what was then thought to be a reasonable confinement scaling, a trapped ion scaling. With these assumptions, catalyzed deuterium and D-3He tokamaks could operate in an ignited mode if they were large. As a part of the present study we have attempted to reevaluate this conclusion in light of revised confinement scaling laws based on more recent experiments, e.g. empirical energy scaling based on PLT experiments and on Alcator data. Both scalings predict improved confinement. The resulting impact on the alternate fuel systems is dramatic. For example, with B = 12%, ignited catalyzed deuterium and D-3He tokamaks are approximately one- fourth the volume for Alcator scaling as opposed to trapped ion scaling. While these new results are still uncertain due to uncertainties about the accuracy of nT scaling at the high temperatures involved (30-40 keV), they demonstrate a plausible and potentially important alternate-fuel tokamak approach. Consequently, this approach will be considered in the present technology assessment. Even if the confinement system is similar to the D-T approach, e.g., a catalyzed deuterium tokamak, important differences occur. For example, startup, first-wall design and energy conversion are unique and not a simple extrapolation or substitution of D-T technology. Additional differences include: • Blanket technology differs drastically due to the elimination of lithium or lithium compounds in the blanket. • The first wall criteria are considerably different since wall loading limits are set by surface heating from radiation. Also the wall life time may be largely determined by high-energy fusion product bombardment. • The large fraction of energy in charged particles highlight the need for direct convertor technology effort. However, direct conversion technology has not been considered in this effort. • Startup heating is a critical aspect of alternate fuels which requires new approaches such as burn propagation starting from a D-T core. Additional work on this subject is anticipated in the future. 2.0 Plasma Engineering
2.1 Introduction
This section presents some results on a number of plasma engineering issues regarding alternate fusion fuels. Characteristics of tokatnaks and field-reversed mirrors are developed as representative confinement concepts for some alternate fuels. The resulting power splits between charged particles, neutrons and radiation are reviewed. Special consideration is given to cyclotron radiation models because of their importance in the power balance of high temperature plasmas. A brief survey of the p-6Li reaction is also presented.
2. 2 Tokamak Concepts Studies
2.2.1 Background and Motivation
A previous University of Illinois study has identified reactor parameters for alternate-fuel tokamaks. Tokatnaks are relatively low-6 de- vices and are therefore best suited to burning the low temperature deuterium based alternate fuels. Consequently, the study examined D-D and D-3He toka- maks. A number of variations of D-D tokamaks were examined. The distinc- tion between the D-D tokamaks involve the fusion products which are burned. In the semi-catalyzed D tokamak, the fusion product tritium is assumed to be burned in-situ, however, the fusion product 3He is assumed to not be burned. In the fully-catalyzed D tokamak, both the fusion product tritium and 3He are assumed to be burned in-situ. In the present study, we have re-examined deuterium-based alternate fuel tokamaks. Fully catalyzed deuterium tokamaks have been examined in the most detail, because, of the D-D tokamaks, the fully catalyzed deuterium tokamak is most easily ignited. The D-3He tokamak has received less attention despite its relatively low ignition temperature, because of the difficulties involved in obtaining 3He as fuel.
We have been motivated to re-examine deuterium based alternate fuel tokamaks for two reasons. First of all, the previous University of Illinois study accepted as reasonable those reactors for which the confinement time necessary to achieve ignition was less than the confinement time for the reactor as calculated from trapped ion scaling. In light of revised scaling laws based on more recent experiments, e.g., empirical energy scaling based on PLT experiments and Alcator data, we have reexarained the possible oper- ating parameters of alternate-fueled tokamak reactors. For example, the re- actor would have been considered unreasonable if the necessary confinement time for ignited operation greatly exceeded the confinement time calculated for trapped ion scaling. In the present study, the same reactor would be considered reasonable if the necessary confinement time for ignited opera- tion was less than, or approximately equal to, the confinement time calculated for Alcator scaling. Although it is certainly not clear that Alcator scaling can be extrapolated from present experiments to high-temperature alternate fuel tokatnaks, it is reasonable to examine the tokamaks which are feasible with improved energy confinement scaling.
A second motivation for re-examining deuterium-based alternate fuel tok- amaks has arisen as a consequence of the present technological assessment. The present technological assessment has provided curves of the maximum tor- oidal magnetic field (B^.) and plasma current (I) versus major radius (R)
for alternate fuel tokamaks based on design stress limits in ma'y.iets (see Sec. 3.4 for further details). Re-examination of alternate fuel tokamaks from the previous University of Illinois study indicates that stress limits were in some cases exceeded. In the present study, we have assumed that B.j and I are at their maximum values for a given fuel cycle and major radius.
Whereas relaxing the restriction on the plasma confinement time has allowed us to consider smaller ignited alternate fuel tokamaks, choosing a combination of B^., I, and R from the field and plasma current limitation curves has in general reduced reactor power densities.
2.2.2 The Tokamak Global Code
The tokamak global code solves continuity and energy balance equations to determine the size of the reactor and the particle confinement time which are necessary to produce a specified thermal power with a specified plasma multiplication Q. Bremsstrahlung is treated as a power loss per unit volume in typical fashion, while the effects of reabsorption and reflection of cyclotron radiation are treated with the Krajcik model.^ ' The temp- eratures of the electrons and ions are specified. Consequently, neither auxiliary plasma heating nor fusion product heating are treated in detail. Credit is not taken for fusions induced by superthermal fusion products.
10 The restrictions on B-p, I, and R affect the reactor design through consideration of MHD stability. The plasma major radius which is necessary to obtain a specified total reactor power is calculated from the fusion power density. However, the fusion power density is proportional to the densities of the fusing species. The densities of the fusing species are in turn proportional to the plasma pressure, which is calculated from the
definition of the poloidal beta (3p). However, the poloidal beta and the
poloidal magnetic field (Bp) are related to BT, I, and R through equa- tions (1-3). Considerations of MHD stability, impose two commonly cited requirements as follows: the safety factor at the edge of the plasma, q(a), is greater than approximately 2.5 and that the poloidal beta is less than the aspect ratio, A, of the tokamak. Specifically, we have used equations 1 and 2 which are essentially definitions of the poloidal beta and the safety factor q(a) in conjunction with equation (3) which simply expresses a relationship between the plasma current, I, and the poloidal magnetic field as derived from Ampere's Law:
« = p (2(q A)2+ 1 + K2) (1) P (1 + K2)
B /2 B - T C 1 + K¥ (2) B P" ^s ~n
/2~ where K is the ratio of the major to minor axis of the elliptical plasma cross-section, the plasma current is in mega-amperes, the toroidal and poloidal magnetic fields are in Tesia, and the plasma minor radius (a) is in meters.
2.2.3 Field and Current Limitations
In the present study we have designed a Cat-D tokamak reactor so that the fields and currents do not exceed the limits specified in Fig. 3-21 of Sec. 3.4. For example, Table 2-1 presents parameters for a Cat-D tokamak with R - 12 m, B - = 8.7 T, and 1-51 MA. For comparison, from Fig. 3-21 of Sec. 3.4, the field and current limits are R =• 12 m, B . = 8.7 T (Bmav = 14.5 T), and
11 Table 2-1. Reference Tokamak Reactor Designs
Parameter Cat-D
R(m) 12
a(m) 4
b(m) 8
Plasma-Wall Sep (m) 0.2
6(%) 10
Bpl(T) 8.6 r(sec) - required 15 x(sec) — Alcator scaling 52 nD (#/cm3) 1.8 x 1011* ne (#/cm3) 2.4 x 1011*
Pth (GW) 7.9
Power Splits (GW)
Leaking Particles 2.3
Brems s trahlung 1.8
Cyclotron radiation to walls 0.3
Cyclotron radiation to holes 0.4
Neutrons 3
Electric Power 4.2
12 I = 50 MA. The above example illustrates a second aspect of our study; i.e., less stringent criteria for the necessary confinement time. For example, for the R = 12 in Cat-D tokamak reactor the necessary confinement time to achieve ignition is approximately 1/3 of the confinement time calculated from Alcator scaling. Thus, with Alcator scaling as a basis of comparison, this represents a reasonable reactor. However, had we compared with trapped ion scaling the reactor may have been rejected as unreasonable, since the confinement time to achieve ignition exceeds the trapped ion confinement time by a factor of 1.5.
2.2.4 Cat-D Tokamaks
Thus far, a relatively large Cat-D tokamak which is consistent with field and current limitations and Alcator scaling has been described. In this section, we describe smaller Cat-D tokamaks and the assumptions that have been made in the study. Using equation (1) and specifying 3 and K , we have calculated a maximum value for q, with the constraint that &„ = 3. Next, we have chosen a value for q less than the previously calculated maximum value, and with equations (2) and (3) have determined, for a specified major radius, a point on the field and current limitation curves
which satisfy these equations. We have maintained 3p ^ 3 at the expense of allowing q to vary. However, we have not allowed q to drop much below ^2.5. Figure 2-1 illustrates points on the field and current limitation
curves which satisfy equations (1-3) with 8p ^ 3 and q^'2.5 for 3 = 10%. We have calculated parameters characterizing tokamaks at these points. Beside the points on Fig. 2-1 we have listed four key items which describe the corresponding tokamaks:
(1) the reactor thermal power (Pth) (2) the ratio of the necessary confinement time to achieve ignition to the confinement time calculated from Alcator scaling,(x /T . ). req ale (3) the plasma safety factor q, and (4) the plasma ellipticity K.
Note that — < 1 for all but the R = 8 m tokamak and that even for this Talc case,Trec. exceeds Ta^c by only a factor of 2. Thus, all the reactors displayed on Fig, 2-1 may be considered reasonable on the basis of Alcator scaling.
13 100 FSL-79- 220
10
Bp, (T)
Figure 2-1. Ignited fully-catalyzed deuterium tokamaks which are consistent with field and current limitations and Alcator scaling for e = o.io.
We have also applied the MHD equations (1), (2) and (3) in a less opti- mistic fashion. Using equation (1) and specifying K = 1.6, q = 3.0, and
3D = 3.0 we have calculated $ = .06. Then, we have used equations (2) and (3) to determine points on the field and current limitation curves which are consistent with the specified values. Figure 2-2 illustrates these points.
We have listed Pth, Treq/Talc, and K beside the points. Note that we have not denoted reactors for R £ 10m. This is because the reactors with R < 10m were not ignited, due to the low 3. However, note that both the R =
12 and R * 14m tokamaks have req 14 100 FSL-79-219 10 P = 6.9 C.U, -iSl = .6, q - 3.0, K = 1.6, I: = .06 ale V = 3.5 GW, —-^ = 1.3, = 3.0, >; = 1.6, !'. » .06 10 12 (T) Figure 2-2. Ignited fully-catalyzed deuterium tokamaks which are consistent with field and current limitations and Alcator scaling for e = 0.06. 2.3 FRM Concept Studies The J[ield-jteversedffl.rror (FRM) has been chosen as representative of high-beta devices for alternate fuels. High-beta devices are especially appropriate for alternate fuels because the thermonuclear power per unit volume scales a 3 -p- for a given magnetic field, so high 3 operation can compensate for the smaller reactivities of alternate fuels. Also the power per unit volume emitted in cyclotron radiation scales as T2 (-^i^). Alternate e p fuel reactors must operate at high ion kinetic temperatures. For most fusion plasmas, Te — 7^ due to rapid energy exchange between ions and electrons. Thus, if the magnetic field is not excluded from the plasma, the cyclotron power emitted per unit volume is large. High-3 operation not only decreases the cyclotron power emitted per unit volume, but it also leads to increased-reabsorption of cyclotron radiation. Consequently, alternate fuel FRM's have for their fuel and temperature of operation relatively small power loss due to cyclotron radiation. Therefore, a larger fraction of the 15 energy leaving the plasma is in the form of charged particles. The FRM concept can take advartage of a high plasma leakage rate because it is well suited to efficient direct conversion of the kinetic energy of the leaking plasma and charged particles. The closed field region is surrounded by an open field region that allows energetic particles to be effectively diverted out the ends of the mirror. A majority of their kinetic energy can thus be directly converted to electrical energy. For most energy loss mechanisms, for example classical diffusion, the energy confinement time generally increases as the size of a plasma increases. Thus, the plasma multiplication can be increased simply by increasing the size of the system. Unfortunately, stability considerations limit the radius of the FRM to a few ion gyroradii. Therefore, the alternate fuel FRMs are not ignited reactors. However, despite their small size, the Q values of the alternate fuel FRMs are substantial due to the assumed near-classical diffusion in the closed field region. Thus, to summarize, alternate fuel FRMs can be categorized as small, efficient reactors with relatively large power densities and moderated Q. The FRM's basic concept is similar to Astron; however, the FRM uses in- ternal plasma currents, rather than single gyroradius high-energy electron or ion beams, to maintain the magnetic field reversal (Fig. 2-3). This makes initiation of reversal by neutral-beam injection at sub-MeV energies feasi- ble. We have termed our FRM concept SAFFIRE (Self-Sustained Advanced-Fuel- I[ield- Reversed Mirror) to stress the fueling and heating techniques that have been incorporated in an attempt to achieve steady-state operation without high-energy injection used in other FRM design. To facilitate our study, a plasma model based on the Hill's vortex ' field configuration (Fig. 2-4) has been developed. This model is similar to to that of Condit, et al. ^ \ however, it provides a measure of self- consistency through use of the vortex field-equilibrium to represent the reversed-field configuration. This analogy allows an explicit check on the field-reversal requirement and, in addition, provides an equilibrium plasma density-profile which is then used to reduce the steady-state particle and energy balance equations to average O-dimensional form. 16 PLASMA AND AZIMUTHAL CURRENT REVERSED MAGNETIC FIELD CONFIGURATION Figure 2-3. The FRM confinement scheme displaying both closed and open field line regions. FSL-78-52 EXTERNAL MIRROR CLOSED FIELDS COLD PLASMA SHIELD REGION Figure 2-4. Magnetic field model for a field-reversed mirror. The closed field region essentially has a toroidal shape and is embedded in an external open-type mirror field. 17 The SAFFIRE concept rests on the intimate relation between three aspects of the FRM; namely, fueling, fusion-product heating, and stability. To obtain an attractive energy multiplication, neutral-beam injection must be held to a minimum and steady-state or long-pulse operation must be achieved. Indeed, calculations show that once reversal is attained, dlmagnetic currents can supply most of the current required to maintain reversal provided that cross-field diffusion, i.e., the radial density profile, can be maintained. To do this, cold fueling during operation is necessary. Fueling is a difficult problem for toroidal devices, but due to the small size and outward "pushing" of field lines in the FRM, fueling is greatly simplified. Since ions trapped at any point will rapidly spread along magnetic field surfaces, it is only necessary to deposit fuel between the outboard side and the null point, a mere 6-cm distance in present design. This can be done using "state-of-the-art" neutral-beam injectors, operating at < 1 amp and < 10 keV, aimed at an angle to the field surfaces to provide good matching with the desired profile. In addition to fueling, the plasma must be continuously heated to make up for energy losses and to bring the cold fuel up to temperature. Hopefully, a bulk of the heating can be supplied by fusion products, supplemented by some auxiliary ion-cyclotron type heating. Fortunately, as shown in Fig. 2-5, although many of the fp's escape the closed-field volume, they are confined such that their orbits pass through both opened an closed regions. Thus their energy can still be deposited in the plasma, i.e., a 20-cm plasma could still retain as much as 70% of the a energy and 20% of the proton energy. The remaining crucial aspect is stability. Beyond the basic question of whether a stable region exists at all, there is the question of whether sufficiently large sizes are permitted to be attractive, i.e., to allow good fusion product heating and reasonable total power. The basic configuration is MHD unstable, but stability is thought to arise from finite ion gyro- (7 8} (9} '•adius effects. Experiments ' ' and theoretical studies suggest that a stable region occurs for a S = L/p. (plasma radius/ion gyro- radius) of about 5 or less. In practice, values of "v 10 are desirable for good performance with Cat-D and D-3He, so precise limits are a crucial objective for future theory and experiment. 18 100 FSL-78-54R — 3.67 MeV ALPHA •-I4.I MeV PROTON CONFINED WITHIN - THE CLOSED FIELD REGION —- - 600 840 1080 1320 1560 1800 aBo (m-Tesla)xlO3 Figure 2-5. Containment of fusion products in the FRM as a function of plasma radius (a) and vacuum magnetic field (B ). A final distinctive feature of SAFFIRE is the use of a cold plasma that is introduced at one end so as to flow along the open-field lines. ' This plasma serves to shield the reversed-field region from wall impurities, limit fusion-product ash buildup, reduce charge exchange with background neutral gas, and couple the diverted plasma to a dump or direct collector. Indeed, based on a fairly detailed study, we have concluded that the cold plasma region can be sustained with fusion product energy, that this region is quite effective for the purposes already indicated, and that it, there- fore, enables the FRM to operate In a truly steady-state mode- Having described our FRM design SAFFIRE in some detail, we now present parameters characteristic of SAFFIRE and the SAFFIRE plasma with Cat-D and D-3He as fuel. We defer a discussion of p-6Li and other high temperature fuels to Section 2.5, because the FRM may not be the most suitable device for high temperature fuels due to its inherently small size. Table 2-2 compares D-3He and Cat-D SAFFIRE parameters for two similarly sized systems (^5 MW gross single cell thermal output). It shows 19 Table 2-2. Reference FRM Reactor Designs Parameter D-He3 Cat-D Radius (cm) 26.7 28.9 Volume (liters) 239 303 e 1.06 1.06 Elongation 3.0 3.0 Stability Factor 15.0 15.0 T (sec) 8.6 particle 16.3 17.5 Tclas.(sec) 34.5 T. (keV) ions ' 68 80 T .. (keV) 66 71.2 electrons Ave. Electron Density (#/cm ) 7.64 x 1014 6.29 x 1014 Ave. Ion Density (#/cm ) 5.08 x 1014 5.51 x 1014 Fractions of Ion Density: electron 1.51 1.14 deuterium .43 .76 tritium 1.5 x 10 ~3 7.2 x 10 ~3 helium-3 .43 0.05 alpha .08 0.09 proton .07 0.08 Gross Power (MW) 4.8 4.7 Gross Power Fraction: bremsstrahlung .30 .20 cyclotron .03 .04 leaking particles .24 .20 fusion products charged particles .40 .20 neutron .03 .37 Net Elec. Power (MW) 2.5 1.8 20 that the Cat-D plasma is slightly larger (300 vs 240 £); however, this ne- cessitates little change in the size of the vacuum chamber and magnets. The most significant difference is in the distribution of the output power. Due to the lower Z, the Cat-D system has reduced radiation losses, but these are canceled by an increase in fusion energy carried by neutrons. This latter effect leads to a decrease in both self-heating and net power for the Cat-D system, which thus has an overall efficiency of 38% as compared to 52% for the D-3He case. In summary then, Cat-D SAFFIRE reactors appear feasible, although less clean and efficient than their D-3He counterparts. These characteristics are, however, offset by their more attractive fuel cycle. A D-3He fuel cycle depends upon obtaining 3He, from either a D-T system having excess tritium breeding (subsequent decay to 3He), or a larger Semi Cat-D system that generates He directly; while the Cat-D cycle is fuel self- sufficient. 2.4 Cyclotron Radiation Considerations 2.4.1 Introduction Electron cyclotron radiation poses a serious power loss problem for fusion reactors employing the D-D reaction and other alternate fuel reac- tions. ' ' It is also important from the standpoint of diagnostics, since the power loss spectrum depends only on the electron density profile, the electron temperature profile, the magnetic field, and the reflectivity properties of the wall. According to an approximate calculation by Trubnikov, the cyclotron power loss per unit area from a cylindrical plasma of radius &_ is given by Pcyc = (2 x 10-") B? Tj njr a? (1-R)T , (4) where T is the electron temperature, n the electron density, R the re- flectivity of the wall, and the constant is chosen so that all quantities are in inks units except for T , which is in keV. Similar results hold for a slab and a torus. The Trubnikov formula, however, is limited to uniform plasma profiles, and it obviously is based on the assumption of homogeneous wall properties. 21 In this section we investigate how the power loss is affected, first, by nonuniform plasma profiles and, second, by the presence of holes in the wall. The first issue is studied by means of two transport codes, ANISN and a ray-tracing code; the second issue is studied with the ray-tracing code. 2.4.2 Effects of Nonuniform Plasma Profiles If one works with sufficiently simple geometry and assumes uniform plasma profiles, the numerical calculation of the cyclotron power loss is not a difficult problem. Significant complications enter, however, when the profiles can no longer be considered uniform. In this section we describe two computational approaches to this problem, built around deter- ministic transport codes. The physical model is a plasma contained in a long circular cylinder, with the electron temperature and electron density allowed to be arbitrary functions of the distance from the axis. The magnetic field is parallel to the axis and, for simplicity, uniform (it could be allowed to depend on r with no conceptual changes). The walls of the cylinder are allowed to absorb an arbitrary fraction of the incident radiation and to mix the polarization states in an arbitrary fashion; for definiteness, the ingoing and outgoing angles are taken to be related by the condition of specular reflection. In order to establish the concepts and notation, it will be useful to begin with a few words about the transport equation for cyclotron radiation. In general geometry it has the form [Q.V + aa(r,n,u>)] Ia(r,Q,a>) = na(r,fi,ai), (5) where I (r,fi,w) denotes the energy flux of cyclotron radiation having frequency u, polarization state a, direction of propagation Q, and spatial location r. The quantities m and n are the coefficients of absorption and emission, respectively. Some years ago Trubnikov presented a theory which determines these coefficients on the basis of the linearized Vlasov equation. At frequencies large compared to the electron plasma frequency, which is a sufficient range for our purposes, the absorption coefficient given by his theory has the functional dependence aa = dif aa(Te/mec2' sin 8' 5 22 where u is the electron plasma frequency, JJ_ = eB/m is the electron gyro- frequency, and 9 is the angle between the direction of propagation and the magnetic field. At relatively low temperatures (T £10 keV) , the absorption coefficient peaks in the first several cyclotron harmonics, and as the temperature is raised it spreads over an increasingly large spectral range. The emission is related to the absorption by Kirchhoff's law, ou2T (7) the coefficient of proportionality being the Rayleigh-Jeans intensity. The quantity a , an extremely complicated function of the indicated ° (15) arguments, has been studied numerically by Tamor and is available in a code written by him. This code is used to set up the transport equation for the ANISN computations. The ray-tracing computations make use of an approxi- , . * ~ (16) mate analytic expression for a . We turn now to the solution of Eq. (5), and hence the computation of the power loss, via the two transport codes. Computations with ANISN The cyclotron radiation transport equation has the form of the neutron transport equation without scattering and without fission. This raises the possibility of adapting existing neutronics codes to the cyclotron transport problem. One such code is ANISN, in which the neutron flux is allowed to depend on only a single spatial variable but in which the propagation vector lies in three dimensions. The directions of propagation are discrete but arbitrarily numerous; in the angular selection Sg, which we use here, there are 96 angles on the unit sphere. Since in this problem there is no scattering, the frequency dependence can be handled by a single energy group. The adapatation of ANISN to cyclotron transport gives rise to two general problems, one minor and the other more serious. First, there is the circumstance that the absorption coefficient depends on the angle between the direction of propagation and the magnetic field. ANISN, however, does not allow for an angle-dependent absorption, as that dependence is unlikely to occur in neutron transport problems. We handle this problem by noting 23 that, since there is no scattering, each ray remains within a given cone during its transit of the cylinder, and therefore sees an absorption which varies only with r. If different groups are associated with the various cones, therefore, the problem is circumvented. A related difficulty is the presence of the polarization index. It can be handled in similar fashion if we recall that a ray maintains its polarization during a single transit. (This property would fail if the absorption coefficient were a matrix in the polarization indices.) Thus it suffices to extend the group labeling to include polarizations as well as cones. The second, more serious, problem is associated with the boundary condi- tion. We shall allow a ray to suffer an arbitrary loss in intensity and an arbitrary degree of polarization mixing at the boundary, with the ingoing and outgoing angles related by the condition of specular reflection. The possi- bility of reflection with polarization mixing is essential, but the specular relation between the angles breaks down if the wall has blisters comparable in size to the wavelength of the radiation. The specular condition, however, is a reasonable first approximation in general circumstances. Needless to say, ANISN does not directly permit such a general boundary condition. We therefore adopt the stratagem of a response-function calculation. The outgoing and ingoing fluxes at the boundary are always related by an equation of the form ut n i° (e,*) = Ta(e,+) i* (e,) + sa(e,«, (8) where Figure 2-6. Labeling of first few a2imuthal angles for polar angle an incoming flux in one particular (9) The aim of the response-function calculation is to compute T..(6 ) and S?(9 ) via ANISN. This proceeds in two steps: 1. Consider a transport problem with no emission term and with the incoming flux at the boundary restricted to a single azimuthal angle 2. Restore the emission term, remove the shell source, and maintain vacuum boundary conditions. Now the computed outgoing flux deter- mines the source function: I (St'^ti ^ = ^"^k^* 25 Having done these computations, we have one equation (9) which relates the outgoing and incoming fluxes for any choice of boundary conditions. A second equation is provided by the boundary condition itself, which in the present problem can be written: W^ " I Roa- where R -(6k) is the reflectivity matrix, assumed to be given. Again the frequency dependence has been suppressed. Note that the aximuthal angles have been chosen so that Combining Eqs. (9) and (10), we obtain the outgoing and incoming fluxes: ja-o" where we have introduced the matrix and its inverse U~^ defined by The power spectrum P(OJ) deposited on the wall, per unit area, follows from Eq. (8). We have a4) PC.) - 4 Iwki sinek cos*k. I tia(ek,C) - iff(V*£>l. where w. ^ is the angular quadrature weight and the factor of 4 arises from the cylindrical symmetry of the system. To obtain the total power loss P eye 26 per unit area, we carry through this procedure for a number of frequencies and form the integral Pcyc / Above 15 keV or so, the integrand is a smooth function where it has its largest values, and there is no difficulty in estimating the integral. As an example, consider a plasma contained in a cylinder of radius a = 2 m, with a uniform axial field B = 10 T. The electron density profile and temperature profile are taken as 2 2 n 2 e (r) = (5.5 x 10 ° m-3) [1 - Aa ] , r 2 2 (16) T (r) = (45 keV) [1 - (f) 1 , c a. giving a central beta of 10%. We consider two model boundary conditions. The first is characterized by the reflectivity matrix /0.90 0.05\ R \0.05 0.90/ for each cone and frequency; the second corresponds to a perfectly absorbing boundary, i.e., each entry of R vanishes (naturally this case could be handled without the response-function formalism). We choose a uniform spatial mesh of 0.01 m and work with the angular package S8, which has four cones in each half space. The first step is to prepare a table of absorption coefficients using the code OPAKE. For each frequency, the number of entries in the table is 200 x 2 x 4, corresponding to 200 spatial points (electron temperatures), 2 polarizations, and 4 cones. The computation time increases in proportion to the number of spatial points and cones. It increases in a complicated way with the number of frequencies, since the code employs an algorithm which handles many frequencies in the same loop. For a package of 30 frequencies ranging from 2 to 12 cyclotron harmonics, the running time on the IBM 370/195 was in the neighborhood of 1-4 s per temperature and per cone. Having assembled the absorption coefficients, we make two ANISN runs, corresponding to steps (1) and (2) above, at each frequency. These two runs, which together consume about a minute of computational time, are of course 27 independent of the boundary condition to be imposed on the flux and need be made only once. Finally we combine this output with the chosen boundary conditions and solve for P(w) as given in Eq. (14). This last step requires only a few seconds. The results for P(to) are shown in Fig. 2-7, along with some ray-tracing results to be discussed in the next section. The integrals for the total cyclotron power loss are: '0.90 0.05\ Pcyc = 6'53 x 10 2 • R = 0.90/ (17) P i = 1.89 x 105 WattS , R = 0 . eye 2 Because of difficulties in obtaining the absorption coefficients, the curves have not been extended below the second cyclotron harmonic. This region, however, is not expected to contribute more than 1 or 2 percent to the power loss. A good way of checking the ANISN computations is to apply them to the case of a uniform plasma profile, since in this case the power loss can be expressed very simply in terms of angular integrals over the absorption coefficients. Specifically, one can show that the transmission and source functions of Eq. (8) become 6 To(9,« = e-V '^' (18) s (e,<|>) •• — [l - T (e, where 2aaa(6) cos* V6'*>=—iini (20) and as usual the frequency dependence is suppressed. Taking for simplicity a reflection mtmatrix of the form R <(0) = 6 ..R, we obtain for the power loss spectrum: 2 uT T -%ir ,JJJTT P(u) = 2- (1-R) If d9 sin26 / d<(. cos* 2ir3c2 a o o 28 1 1 I 1 NONUNIFORM PROFILE 2.4 2O 3 2 ne=(5.5xl0 rn" )[l- (r/o) ] 3 = 10 T 2 Z = 2m 2.0 _ Te=(45 keV) [l-(r/o) /V. • ANISN, R=0 e 1.6 — * / \ > O / \ ~ 12 _ / \ ^RAY-TRACING, R = 0 3 Q_ / ^\\ / \ 0.8 - / \ /090 0.05 \ ~ \i J\rANIS Nl R= lo.O5 0.90/ 0.4 - / /'%RAY-TRACING, R = 0.95 7 / ^ 1 10 15 20 25 w/ftD Figure 2-7. Power loss spectra for plasma with nonuniform profiles. We approximate the integral as a sum over the S8 angles with appropriate weights, and call this the semianalytic spectrum. Figure 2-8 shows this result for a system with n and T given by their peak values in the previous (nonuniform) case and with reflection coefficient R = 0.95. This is compared with the spectrum obtained by running the ANISN program for this system, with the same mesh and angular selection as before. (Also shown is a ray- tracing spectrum discussed in the next section.) As can be seen, the agree- ment is quite good over the entire frequency range. The maximum discrepancy of 15% occurs in the region of peak power loss and probably can be attributed to coarse-mesh effects. The total power losses, 1.74 x 105 watts/m2 for the semianalytic result and 1.93 x 105 watts/m2 for the ANISN result are similarly close to one another. The Trubnikov power loss formula Eq. (4), follows from Eq. (21) upon insertion of an analytic approximation for the absorption coefficient followed by an approximate evaluation of the angular and frequency integrals. Applied to the case just considered (n = 5.5 x 1020 m~3, T = 45 keV, B = 10 T, a - 2 m, R » 0.95), this formula gives a power loss of 6.37 x 10s watts/m2, which exceeds the semianalytic figure by a factor of 3.7 and the ANISN figure 29 1 1 1 1 1.4 ~~ UNIFORM PROFILE ~ 20 3 ne = 5.5 x I0 m" B = IO T 1.2 ~~ Te = 45 keV a = 2 tn R = 0.95 — ANISN _ T l0 SEMIANALYTIC^^^ //~" \ f* 0.8 3 0.6 " / ^ a. 0.4 ~ / V 0.2 — / / ^ RAY-TRACING 1 20 25 Figure 2-8. Power loss spectra for plasma with uniform profiles. by a factor of 3.4. In this case, therefore, it can be concluded that the Trubnikov formula is high by a factor of 3 or 4. It is also instructive to compare the Trubnikov formula with the power loss obtained earlier in the system with a nonuniform profile. Naturally an unambiguous comparison is not possible in this case, since the formula pertains only to uniform profiles. Suppose, though, that we define an average electron density n and an average electron temperature T as a certain fractioneof their peak values, (22) T = eT (0) , 6 £ and determine e by requiring the Trubnikov power loss, calculated with n __ e and T , to agree with the ANISN figure. We find e = 0.45 in the case with a reflecting boundary and e = 0.40 in the case with an absorbing boundary. Thus the cyclotron losses agree with the Trubnikov estimate if the electron density and temperature are assigned uniform values equal to somewhat less than half of their peak values. This fraction seems low, confirming the suspicion that the Trubnikov formula overestimates the power loss. 30 Computations with a Ray-Tracing Code In this approach the transport equation is integrated along a ray across the cylinder, and the power loss is obtained by an angular integral of the result. Neglecting polarization mixing we can express the power loss as * * „ , I (JJ)(n«R) P(«) = I (1-R ) I — s— dU , (23) 1-R e~Vfi> GO ft *\ y\ where n is the outward normal to the wall and the quantities T (£2) and I (G) are Riven by the following ray-tracing integrals across the cylinder in direction ft: Ia(«) = j n0 exp(-/ cta ds') ds, (25) o s with S, the total chord length in the given direction. An important simplifying feature of these last two integrals is the fact that, along the paths of inte- gration, the integrands depend only on the distance from the center of the cylinder. The ray-tracing integrals were evaluated by Runge-Kutta techniques for a first order differential equation. The integration over angles was accomplished by dividing the unit sphere into 400 angular zones and tracing the ray which was located at the center of each zone. The result was then multiplied by the appropriate geometrical factor and the integral was ob- tained by summing the individual contributions. Several approximations, in addition to the neglect of polarization mixing, were made in order to speed up the computations. The angular dependence of the absorption coefficient was ignored, and the absorption coefficient was approximated by its value at 90", which, for the extra- ordinary mode, is given approximately by to2 , A i I i ^ 31 exp [-p(A3-l +|-X"3)], (26) 31 where u = m c2/T and X = 9 W/2US3,,. This expression is within 2% of the e e r> more precise result in the temperature range of interest. Only the extra- ordinary mode was considered, and the resulting power loss was doubled. As a first test of this code, the uniform-profile plasma of the previous section was considered, with the diagonal reflectivity taken as R = 0.95. The resulting power loss spectrum, shown in Fig. 2-8, is lower than the ANISN and semianalytic results and is also shifted somewhat to higher frequencies, but generally seems comparable with these spectra. The total power loss is 75% of the ANISN figure. The code was then applied to the more realistic problem with profiles given by Eq. (16) of the previous section. Diagonal reflectivities of R = 0 and R = 0.95 were considered. The results (Fig. 2-7) bear about the same relation to the ANISN results as in the uniform-profile case: the spectra are lower and shifted to the right, but some similarities with the ANISN spectra can be seen. The total cyclotron power losses in watts/m2 are 1.43 x 105 (for R = 0) and 2.28 x 101* (for R = 0.95), which are to be compared with the ANISN results of 1.89 x 105 and 6.53 x 101*, respectively. 2.4.3 Effects of Holes in the Wall The wall surrounding a fusion plasma will have a variety of holes for devices such as neutral beam injectors and vacuum pumps. Since these holes do not reflect the outgoing radiation back into the plasma, it is necessary to take them into account in evaluating cyclotron power losses. The simplest approximation is to use a linear interpolation between the power loss from a plasma surrounded by a reflecting wall and the power loss from a plasma surrounded by a vacuum (or perfectly absorbing wall). This is given by P = (l-f)P „ + fP , (27) eye wall vac ' where f is the fraction of the wall taken up by holes, P is the power which would be lost if there were no holes, and P is the power which would vac be lost if there were no wall. 32 In order to test the validity of this approximation, a version of the ray-tracing code was applied to the uniform-profile plasma considered earlier. Each ray was traced from the detector point through a number of reflections, until it intersected a hole, where it was terminated. The intensity at the detector point is then 1(5) =IRJ[l-e-^>] I |^»V (28) RJ i=0 where 1 is the Rayleigh-Jeans intensity, R is the reflectivity, n is the number of reflections, and i(fl) = a£(fi), with £(n) the chord length. Detector points were selected to include both hole areas and wall areas, and care was taken to assure that sufficient detector points were chosen in each region. The power losses to the wall and the holes are A A A / I(fi) n-fi dQ , (29) Jhole and the total cyclotron power loss is the integral of J ,, + J. , over all wall hole frequencies. Figure 2-9 shows the calculated power loss as a function of the hole fraction. The straight line gives the linear interpolation from Eq. (27), which is seen to be accurate to within 25%. This is sufficient for many purposes. We conclude, therefore, that the hole fraction should be taken into account in calculations of the cyclotron power loss and that the linear interpolation formula seems to give a good first approximation of the effect. In other words, one can use the methods described in the previous section, to obtain P ... and P, .. , and then combine them according to Eq. (27) . 33 LINEAR INTERPOLATION 0.5 1.0 HOLE FRACTION f Figure 2-9. Power loss (arbitary units) as a function of hole fraction. Same plasma as in Fig. 3, with R = .95. 2. 5 Review of -the p-6Li Fusion Chain Reaction 2.5.1 Introduction The p-^Li fuel cycle has recently received considerable interest in the fusion community because of the relative abundance of Li and the greatly reduced tritium handling requirements. With almost all of its fusion product energy residing in charged particles and significantly reduced neutron yields, the p-^Li cycle may prove to be the ultimate goal of alternate-fuel fusion. However, severe radiation losses may result in extremely high temperature (^ 500 k.eV) operation. Hopefully, the chain reactive features of the p-6Li fuel cycle will balance out the enhanced radiation losses and allow for ignited operation in a suitable confinement system. Current papers^ »19^ by McNally and Conn are valuable overviews of this problem. 2.5.2 The p-6Li Fusion Chain Reaction The reaction of primary interest in our studies is the catalyzed and fully propagating p-6Li cycle (see Fig. 2-10) which involves the reactions^ * p + 6Li -»• 3He(2.3MeV) + ^HeCl.? MeV) p(11.3 MeV) + 2 "*He(1.43 MeV) p(6.4 MeV) + ^(7.5 MeV) + ^(1.4 MeV) 3He + Li D(0.4 MeV) + 7Be(0.1 MeV) p + n + 7Be Q = -2.1 MeV. The completely catalyzed p-6Li reaction promotes superthermal 3He and p ions which react in chain propagation fashion with background 6Li ions £ + 6Li - « + 3^E* + 4.02 MEV I i i i 1 6Li + * - p* + a + a* +16,9 MEV I 5|_i + • - P* + 9BE -2.1 MEV ; 6Ll + t * DD* BE 7 6LI +T - P* + Li •«••- INDICATES A HIGH-ENERGY FUSION PRODUCT CAPABLE OF UNDERGOING FURTHER REACTION. Figure 2- 10. Chain reaction for p-6Li. 35 supplemented by chain branching reactions involving p + 6Li, D + 6Li, 6Li + 6Li, 3He + 6Li, 3He + 3He, etc. Indeed, there may be up to 80 side reactions which involve p, D, T, 3He, '•He, 6Li, 7Li, 7Be, and thirteen exothermic 6Li + 6Li reactions which produce a significant amount of neutrons as well as the radioactive nuclei 7Be and 12C. The fast p and 3He produced in the p-6U cycle have substantial probability of reacting 6 with Li prior to thermalization if the electron temperature, Te, exceeds 100 keV. It is desirable to increase the reaction probability by increasing Te; however, since bremsstrahlung and cyclotron radiation losses also increase PF with higher Te, the ratio of fusion power to radiation power (M = ) must be optimized with respect to Te. With the high electron temperatures required for the p-6Li cycle, radiation losses become quite severe. Since cyclotron radiation is basically system dependent it will not be included in power balance considerations. The bremsstrahlung radiation power, on the other hand, can be minimized with respect to the fusion power by careful optimization of the background fuel composition. Since sLi has such a large nuclear charge compared to the single electronic charge of a proton, it is advantageous from a radiation point of view to have a fuel mixture lean in 6Li and rich in p in order to minimize bremsstrahlung losses. The optimization of the fuel composition also affects the slowing-down reaction probabilities, since practically all such reactions are propor- tional to the background density of 6Li. As one can see, ignition of a p-6Li system depends heavily upon a carefully optimized electron temperature and fuel composition. 2.5.3 Energy Balance The two primary mechanisms for the chain reaction release of fusion power are 16 1 16 and v •36 '""r" '36 n (l)p w. (2) U F16 F36; U 46*16 * 36 3 6 where 1, 3, 4, 6 subscripts denote p, He, ^He, and Li; Q16 = 4.02 MeV, (1) Q36 = 16.9 MeV, and Q46 = -2.1 MeV; and P36, P46, and P16(Pi6 and P^g^' refer to different superthermal protons) are the reaction probabil- ities of the superthermal ions with background 6Li ions as they thermalize. The propagating chain reaction features of the p-6Li cycle are included in the power production terms by making use of the following expansion -=— = 1 + x + x2 + x3 1-x where x = P16 P36 or x " P46 P36 P16" The above power production terms are written in terms of the background proton density, n^. The Y and a factors are background density ratios of the multi-charged fuel species with the proton density and can be defined as n3 n6 v = — and a = — . nl nl The factor A(% 4), when multiplied by the The superthermal reaction probabilities p^g can be expressed as P.6 = 1 - e where f3/ 2 T.1 3/i'2 Ti. P = n (ffv) dt i6 J 6 i6 " Jn6(av).6dE/(dE/dt). E E The reaction probabilities can be approximated over certain regions of energy away from resonances in cross sections by the following expression 37 where Tf /T is the ratio of the fusion time of superthermal ions to their slowing-down time. These reaction probabilities neglect nuclear elastic scattering reactions which may increase the fusion reactivity thru the promotion of additional supertherraal ions from knock-on collisions. If we include the fusion power term from the six exothermic channels of the 6Li + Li side reactions and include a bremsstrahlung radiation term which is corrected for relativistic effects and electron-electron brems- 2 strahlung, we have the following power balance which is normalized to n, : X36 36 46^46 P16 Q46 a 1/2 = 4.705x10 i=l J6 J 2T 1 + 1 + 1 - M c z2 + YZ3 + az2 + T /tn c2)2 (1 e e A plot of the normalized fusion power and bremsstrahlung versus temperature is given in Fig. 2-11. As one can see for Te = T-j_/3 and a = y = .8, ignition against bremsstrahlung losses cannot be achieved. However, as f 18} pointed out by R. Conn^ 't who included little known cross section data concerning the 3He + 6Li •*7 D + 7Be reaction into his calculations, there may be a sizable equilibrium density of D in the background plasma which may significantly increase the fusion power from D-3He side reactions. Accordingly, Conn reports that marginal ignition against bremsstrahlung losses (Ti ^ 500 keV, Te ^ 300 keV) is possible for optimized values of Y and a. It appears reasonable to expect that the added D in the thermalized background could significantly increase the fusion power since the D-3He reaction does have a large reactivity at the temperatures of interest. 38 FSL-79-157 © Fusion Power (p-6Li, 3H^-6Li, 6Li-6Li) © Classical Bremsstrahlung (D Total Relativistic Bremsstrahlung I 10" w 28 c°- I0" 10" 200 400 600 800 1000 Tj (keV) Figure 2-11. Normalized fusion and Bremsstrahlung power vs. temperature. However, even including side reactions such as D-3He, Conn could only achieve the power ratio 2.4 'Br Thus, the ignition of the p-6Li fusion chain reaction is still marginal and will require a much more rigorous analysis including heretofore neglected effects such as nuclear elastic scattering to fully evaluate its potential as an alternate fuel. 2.5.4 Neutron and Radioactive Ash Yields Some of the current claims made in favor of the p-6Li fuel cycle as compared to the D-T-Li cycle are grossly reduced neutron production and 39 practically nonexistent radioactive inventory. In light of the recently published reactivities ' of various 6Li + 6Li reaction channels, these claims appear to be overly optimistic. At T^ %, 500 keV, there appears to be significant production of low energy neutrons (see Table 2-3). The majority of these low energy neutrons come from the following reactions: 6Li + 6Li -*a(.342 MeV) + n(1.370 MeV) + 7Be(.196 MeV) 6Li + 6Li + n(8.665 MeV) + nC(.788 MeV). The 6Li(eLi, na) 7Be channel comprises about 40% of the total 6Li-6Li reactivity while the 6Li(6Li,n) X1C channel comprises about 1.5% of the total reactivity. At T± ^500 keV, the neutron production rate is 7.3 x 1011 neutrons/cm -sec, and the corresponding radioactive nuclei production rate is 1.0mCi/cm3-sec. The neutron production rate and wall loading (^1.9 watts/cm^ for p-^Li and 1.8 watts/cm^ for D-Tle) are comparable to those of a D-3He system operating at T^ % 100 keV and may cause significant problems for tnultipole reactors which have been proposed to burn p-6Li. Therefore, neutronically speaking, a p-6Li system may be roughly equivalent to a D-3He system and would hence require a modest shield. The neutron and radioactive nuclei production can be minimized by vary- ing the background fuel composition i.e., by varying y and a. However, this would have to be done at the expense of lowering the fusion power. With a radioactive nuclei production of l.OmCi/cm -sec, the radioactive inventory, though modest compared to D-T systems, would still be significant. This calculation does not include the T and 10Be produced in other side reactions. Although certain amounts of 7Be could be burned up in side reactions with 6Li, there would appear to be significant amounts of residual radioactive nuclei in the plasma exhaust that would have to be handled by appropriate methods. In conclusion, it appears that the evaluation of p-6Li as a viable alternate fuel requires a more rigorous analytical treatment,, A multi-group Fokker-Planck treatment should be employed to incorporate the effects of nuclear elastic scattering on the fusion reactivity. More in depth calcula- tions should be done to determine neutron fields and radioactive ash produc- tion. Better and more complete cross-section data is needed to determine 40 Table 2-3. Neutron and Radioactive Ash Yields as a Function of Temperature for p~ Li. n"~ = 3x1014 cm"3 P 14 -3 ri3He = n6Li = 2.4x10 cm P Activity * Activity ** kw FP n's Produced Produced T(keV) ( 1 FPP-+L± I. 3) 3 6 6 6 cm He- Li ( Li- Li) cm -sec mCi/cm -sec nCi/cm -sec 100 .002 .99 .001 1.2xl09 2.6xlO~3 3.4xlO"4 300 .016 .97 .031 1.7xlOn .33 .10 500 .035 .93 .069 7.3xlOn 1.0 .47 700 .053 .91 .090 1.5xlO12 1.5 .96 900 .069 .88 .12 2.2xlO12 1.6 1.4 100 .076 .87 .13 2.6xl012 1.7 1.7 FP = fusion products *Denotes fusion products ^C, 7Be **Denotes fusion product T the contribution of numerous branch reactions to the overall fusion power. From this preliminary work, however, it appears that the p-6Li cycle has no significant advantages over the D-3He cycle in terms of neutron and tritium production levels. 41 2.6 Summary of Typical Power Splits and Sensitivity Studies 2.6.1 Typical Power Splits It is useful to briefly review the factors which influence the power splits. First of all, the power splits are strongly influenced by the choice of fuel cycle. For example, one would expect to have more power in 14 MeV neutrons for a D-T reactor than for a D-D reactor. However, there would be some 14 MeV neutrons from a D-D reactor as well, because D-D reactions pro- duce tritium ions which consequently fuse. Thus D-T, D-D, and D-3He reactors all contain D, T, and 3He ions in different concentrations. The different fuel cycles are classified by the source of fuel to the system, rather than the concentrations of different species. Thus, a fully catalyzed D-D reactor is classified as a reactor with a source of D, T, and 3He. The D is fed into the reactor to exactly makeup the D which is burned, while the T and He are fed into the reactor at the net rate at which they are produced. The fuel cycles are not classified by the concentration of species in the system because the concentration of species in the system is a function of the reactor operating temperature. Thus, we would expect the power splits to be a strong function of the fuel cycle and the temperature at which the reactor is operated. The plasma temperature also affects the amount of radiation produced per unit volume. Although almost all bremsstrahlung escapes the plasma, the cyclotron radiation is reabsorbed in the plasma. The extent to which the cyclotron radiation is reabsorbed depends upon the density of the plasma, which in turn depends upon 3, and the size of the plasma. The finite size of the plasma affects the power splits in conjunction with the confinement scal- ing. The most obvious effect is that as the confinement is improved, either due to increased reactor size or improved confinement scaling, the power lost in leaking plasma decreases. A secondary effect on the power splits is due to the reaction of superthermal fusion products, which is in turn related to the finite geometry of the system. The FRM code accounts for the reactions of superthermal fusion products; however, the present tokamak global code does not. The tokamak code is being modified to include these effects. 42 In summary, the power splits are affected by fuel cycle, plasma tempera- ture, plasma density (g and magnetic field strength), plasma size and plasma confinement scaling. It is, in general, difficult to isolate the effects upon the power splits of variations in the items mentioned above for large variations. This is because other considerations couple changes in one parameter to changes in other parameters for reasonable reactor designs. For example, in this study we have considered tokamaks and FRMs. The tokamaks are large reactors with relatively low 3. On the other hand, FRMs are small reactors with relatively high 6. As a second example, consider the relation between fuel temperature and fuel cycle. In general, one attempts to operate a fusion reactor at the temperature which maximizes the fusion power density, which is proportional to jZ . For the D-T fuel cycle, this optimum temper- ature Is around 13 keV while for the Cat-D fuel cycle the optimum temperature is around 45 keV. Finally, if one is considering a high temperature reactor, then one should also consider a high $ device in order to minimize cyclotron radiation power losses. Thus, fuel-cycle, fuel-temperature, 3, and device size are intimately related. Recognizing the interrelations among the above factors, we will present power splits for realistic combinations of these factors, i.e., the alternate fuel tokamaks and FRMs discussed in Sections 2.2 and 2.3. Then, we will examine the sensitivity of the power splits to small variations in parameters about the reactor designs. Table 2-4 presents the power splits for the reference semi-catalyzed D, fully catalyzed D, and D-3He tokamaks from the previous University of Illinois study.^ ' Note that the fraction of power in neutrons is less by almost an order of magnitude for D-3He than for semi-catalyzed D. The fraction of power in neutrons is less by about 10% for fully-catalyzed D than for semi-catalyzed D. This is, of course, because the fusion product 3He is burned in the fully-catalyzed D reactor. Although the fractional power leaving the plasma in neutrons is least for D-3He, the fractional power leaving the plasma in the form of bremsstrahlung and cyclotron radiation is greater for D-3He than for the other deuterium based fuels. The relationships among the relative power fractions for the fuel cycles are not much different for FRMs than for the tokamaks, as is illustrated 43 Table 2-4. Power Splits for Tokamaks Semi-Cat-D Cat D D-3He Power Splits Leaking Particles .20 .26 .31 Brerasstrahlung .26 .25 .46 Cyclotron Radiation .11 .11 .16 Neutrons .43 .38 .05 in Table 2-5 . Table 2-5 presents the relative power fractions for fully catalyzed D and D-^He FRMs. Again, the fractional power in neutrons is less by approximately an order of magnitude for D-3He relative to Cat-D, while the fractional power in bremsstrahlung and cyclotron radiation is greater for D- He than for Cat-D. The fractional power loss in cyclotron radiation is smaller for the FRM than for the tokamak, despite the FRM's smaller size, because of the FRM's high g. Table 2-5. Power Splits for FRM Power Splits Cat-D D-3He Leaking Particles .23 .23 Charged Fusion Products .19 .39 Bremsstrahlung .18 .26 Cyclotron .03 .08 Neutrons .35 .03 44 Having compared the power splits for two different devices, FRMs and tokamaks, let us now examine the sensitivity of a single device, e.g., a tokamak, to changes in system parameters. Table 2-6 lists the relative power fractions for some of the tokamaks denoted in Figures 2-1 and 2-2 of Section 2.2. The tokamaks with major radius at R = 8m and R = 10m have identical 3. Note that for the larger tokamak, the relative power in cyclotron radiation is reduced as the increased reabsorption in the plasma outweights the in- creased magnetic field. The relative power in leaking particles is greater for the larger tokamak, because, although the necessary confinement time de- creases, the increase in the volume of the system leads to a larger relative power fraction in leaking particles. The R = 8m and 10m tokamaks in Table 2-6 are for 6 = 0.1. The power frac- tions for the 6 = 0.1, R = 12m tokamak exhibit the same trends discussed above when compared with the R = 8 and 10m devices. However, the 3 = »06, R = 12m power fractions are different from the (3 = .1, R = 12m power fractions. For the 3 = .06 tokamak, the relative power lost in cyclotron radiation is in- creased because of less reabsorption and more magnetic field in the plasma. Also, the power fraction in leaking particles is smaller because the neces- sary confinement time is longer. In fact, the power in leaking particles will be determined by the actual confinement time. Thus, the sensitivity of the power fractions to R and 3 should be viewed with some caution. In summary, the power splits are most sensitive to fuel cycle, less sen- sitive to reactor type (i.e., FRM or tokamak), and least sensitive to para- meter variations for a device. Regarding parameter variations for a tokamak, the power fractions are more sensitive to changes in 3 than changes in R. For purposes of technology considerations discussed in Section 3, the most important parameter is the fraction of total power carried by neutrons. This is relatively constant for the various cases discussed above, thus it is possible to examine certain first wall/blanket/shield problems without specifying a particular type nf device. This is the general approach taken in Section 3 where the fuel cycles are characterized in terms of neutron power fractions by the following representative values: DT 80% DD 40% Cat DD 40% D-3He 1% 45 Table 2-6. Power Splits for Cat-D Tokamak 3 = 10* R = 8 R = 10m Power Splits Leaking Particles .25 .29 Bremsstrahlung .24 .23 Cyclotron .13 .11 Neutrons .38 .37 R = 12m 3 = .10** 3 = .06*** Leaking Particles .29 .16 Bremsstrahlung .23 .26 Cyclotron .10 .20 Neutrons .38 .38 * q = 2.4, K = 1.6, A = 3 ** q = 2.8, K = 2.0, A = 3 *** q = 3.0, K = 1.6, A = 3 References 1. C. K. Choi, et al,, "Exploratory Studies of High-Efficiency Advanced- Fuel Fusion Reactors," Electric Power Research Institute, EPRI-581 (November, 1977). 2. R. A. Krajcik, Nucl. Fusion, 13_ (1973) 2. 3. E. Morse, "High Beta, Low Aspect Ratio Plasmas," University of Illinois, COO-2218-41 (1976). 4. W. C. Condit, G. A. Carlson, R. S. Devoto, J. M. Doggett, W. S. Neef and J. D. Hanson, "Preliminary Design Calculations for a Field-Reversed Mirror Reactor," Lawrence Livermore Laboratory, UCRL-42170 (1976). 5. G. H. Miley, et al., "Fueling Requirements for a Self-Sustained Field- Reversed Mirror," Trans. Am. Nuol. Soa.3 28, 42 (1978). 6. M. Y. Wang, G. H. Miley and L. S. Wang. "Alpha Particle Effects on the Reversed-Field Mirror," Trans. Am. Nucl. Soa., 27, 93 (1977). 7. W. C. Condit, T. K. Fowler and R. F. Post. "Status Report on Mirror Alternatives," Lawrence Livermore Laboratory, UCRL-52008 (1976). 8. R. Linford, LASL, private communication (March, 1978). 9. R. N. Sudan and M. N. Rosenbluth, "Stability of Field-Reversed Ion Rings in a Background Plasma," Phys. Rev. Lett., 26, 972 (1976). 10. J. Gilligan, et al., "Divertor Design for the D-3He Field-Reversed Mirror (FRM),' Third ANS Topical Meeting on Fusion, Santa Fe, New Mexico (May, 1978). 11. B. A. Trubnikov and A. E. Bazhanova, in Plasma Physics and the Problems of Controlled Thermonuclear Reactions (Pergamon Press, London, 1959), Vol. Ill, Pg. 141. 12. W. E. Drummond and M. N. Rosenbluth, Phys. Fluids, §_, 276 (1963). 13. B. A. Trubnikov, in Bevievrs of Plasma Physics (Consultants Bureau, N. Y., 1979), Vol. VII, pg. 345. 14. B. A. Trubnikov, in Plasma Physics and the Problems of Controlled Thermonuclear Reactions (Pergamon Press, London, 1959), Vol. Ill, Pg. 122. 15. S. Tamor, Nucl. Fusion, 18, 229 (1979). 16. B. A. Trubnikov and V. S. Kudryantsev, in Proc. U. N. Intern. Conf. Peaceful Uses of Atomic Energy (Geneva, 1958), Vol. 31, Pg. 93. 17. W. W. Engle, Jr., "A Users Manual For ANISN," Oak Ridge National Laboratory, K-1693 (1967). 47 References (Continued) 18. J. Rand McNally, Jr., "Some Thoughts on the P(6Li)3He(6Li,p)8Be Chain Reaction," Proc. IEEE International Conference on Plasma Science, Montreal, Canada, June 4-6, 1979. 19. R. Conn and G. Shuy, "Advanced Fuel Cycles and the Propagating p-6Li Cycle," University of Wisconsin, UWFDM-262 (1978). 20. J. M. Dawson, "CTR Using the p-11B Reaction," University of California, Los Angeles, PPG-273 (August, 1976). 21. L. Ruby and T. P. Lung, "Effect of 6Li + 6Li Reactions in a Fusion Reactor Operating on bLi(p,3He)"»He," Nuel. Set. S Eng. _66, 107 (1978). 3.0 Enginee ring/Technolo gy ConsIderat ions 3.1 Materials, First-Wall and Blanket Considerations Major differences between the alternate fusion systems (D-D and D-3He) and the D-T system that affect materials selection and first-wall and blanket design considerations are: (1) no tritium breeding requirement (2) differ- ences in neutron energy and flux, (3) higher particle transport and radiation heat loads on the first wall, and (4) higher magnetic fields. Elimination of the tritium-breeding requirement reduces the number of materials compati- bility considerations in the blanket, and eliminates the whole tritium-breeding system including the need for lithium in the blanket and the related tritium transport problems. The greatest impact is probably on the coolant selection since coolant/breeder compatibility is no longer important. Variations in the power splits or ratios of neutron, radiation and particle transport energy losses from the plasma have an important impact on materials selection and first-wall/blanket design. Table 3-1 gives the estimated power splits used for assessment of the impact on materials considera- tions. In general, the radiation losses are assumed to be deposited on the Table 3-1. Comparison of Power Splits and Allowable Wall Loadings for Alternate Fuel Cycles Power Split (%) D-T D-D D-3He Neutrons 80 38 5 Radiation 8 35 63 Transport 12 26 32 Total Wall Loading (P (max) = 0.6 MW/m2) 3.0 0.97 0.63 49 surface of the first wall and the neutron energy is deposited within the first wall and blanket. Unless the particle flux is diverted, the particle transport energy losses from the plasma will also be deposited on the surface of the first wall. From Table 3-1 it is apparent that the combined radiation and transport losses for the D-D and D- He are considerably higher than for the D-T fuel cycle. As will be discussed in more detail in subsequent para- graphs, the maximum allowable surface heat flux on the first wall is an important design limitation, particularly if the plasma burn is cyclic in nature. An example of the total allowable wall loading for the three cycles is given in Table 3-1 for a constant surface heat flux limit of 0.6 MW/m2. Since the plasma particles are similar in the D-T and two alternate-fuel sys- tems, plasma-wall interactions such as sputtering and blistering will be similar. As indicated in Table 3-1, the fraction of fusion power released in the neutrons is considerably smaller in the alternate fuel cycles. Also, the helium and hydrogen production rates from the 2.45 MeV neutrons from the D-3He reaction are much lower in the candidate structural materials than for the 14.1 MeV neutrons produced in the D-T reaction. Resultant effects on mechanical properties of the structures should be less severe for the lower energy neutrons. The higher magnetic fields generally considered for the alternate fuels concepts will enhance electromagnetic and magneto-hydrodynamic effects. 3.1.1 First-Wall Surface Effects In general, the types of plasma-wall interactions in the D-D and D-3He fuel cycles will be similar to those in the D-T fuel. Physical sputtering of the first-wall materials has been identified as a major impurity source for a D-T plasma. The sputtering source terms for D-T reactor analyses have typically been based on sputtering yields for characteristic edge tempera- tures of 50-1000 eV. If the characteristic edge temperature for an alternate fuel cycle is higher than that of a D-T cycle because of the higher plasma central temperature, the sputtering yields for most wall materials will in- crease. The importance of any increase is dependent on the effect of the impurity on plasma performance. Since the sputtering yields are produced primarily by the fuel ions, the masses of the fuel ions are important. 50 For the same ion flux to the wall, the D-D fuel should produce a slightly lower yield than D-T, whereas the D-3He should produce similar yields if the energies are the same. Expressions for energy dependent sputtering yields have been developed in the U.S. and in Germany. An expressior that agrees fairly well with most experimental data is given by 50 Where (M. +M,)2 Eth E , = — 2 U for M. > M, , th Mj o 1 2 S in atoms per incident ion, E is the incident particle energy in eV, U is the surface binding energy in eV, and Z and M are the atomic and mass numbers, respectively, with the subscripts 1 and 2 referring to the incident particle and target atom, respectively. Since low-Z sputtered wall materials are expected to have less effect on plasma performance than higher-Z impurities, the low-Z coating concept proposed for the ANL Experimental Power Reactor may also prove useful for the alternate fuel concepts. The helium-blistering mechanism may be more critical for the D-3He fuel cycle because of the higher helium flux to the wall. This may be an important consideration in the selection of first wall materials. The surface morphology of the first wall is more critical for the alter- nate fuel concepts because of the enhanced cyclotron radiation produced in the higher temperature plasmas. High reflectivity of the cyclotron radiation from the wall is necessary to maintain an acceptable plasma energy balance. Although data for specific conditions anticipated are limited and further work is warranted, some general observations can be made. Surface rough- ness produced by blistering should not significantly affect the reflectivity of cyclotron radiation with wavelengths of the order of 1 mm. However, the electronic structure of the wall is important for good reflectivity. An D, L. Smith, J. Nucl. Tech. 75, 20 (1978). 51 estimate of the reflectivity can be made from the electronic conductivity of metals. In general, high conductivity metallic surfaces should provide superior reflectivity and ceramic insulators may be unacceptable. Impurity interactions, such as oxidation of the first wall, may be particularly important in the alternate concepts since the electronic structure of the wall could be signifi- cantly affected. 3.1.2 Structural Materials Probably the most important difference related to the structural material for the alternate fuel concepts relates to reduced neutron effects. Radiation damage produced by the 2.45 MeV neutrons from the D-3He reaction should be similar to the radiation damage that occurs in fission reactors. These low energy neutrons produce nearly as much displacement damage, however, helium and hydrogen generation rates are much lower. Therefore, swelling rates and mechanical property effects predicted from fission reactor data should be more reliable for the lower energy neutrons. As indicated earlier, larger fractions of the fusion energy is deposited on the first wall in the alternate fuel cycles than in the D-T system. As a result, thermal stresses induced in the first wall will likely be more important in the alternate fuel cycles. This will be particularly critical for operation with cyclic plasma burns. Therefore, the thermal stress factor may be a critical parameter in the selection of first-wall structural materials. High thermal conductivities and low thermal expansion coefficients, which lead to low thermal stress factors, are beneficial. Table 3-2 lists the thermal stress factors for the six alloy classes most often considered for D-T fusion reactor structural materials. These results indicate that the refractory alloys of niobium and vanadium should withstand higher surface heat loads, and therefore, may be the pre- ferred structural materials. 3.1.3 Candidate Coolants Since most of the coolant requirements for the alternate fuel concepts are similar to those for the D-T reactor, the proposed coolants for D-T reactors (helium, water, liquid metal, and molten salt) are the primary candidates. The major considerations that differ for the alternate fuel 52 Table 3-2. Thermal Stress Factors for Candidate Structural Alloys «n m or. Thermal Stress* Alloy Temperature, C „ ._ ,„ J v ' Factor, MPa*m/W 316 stainless steel 400 0.220 Inconel 625 500 0.218 Ti-6242 400 0.128 Fe-9Cr-lMo 400 0.105 V-15Cr-5Ti 500 0.055 Nb (FS-85) 500 0.028 * aE •yin A f\ t? ^ *~ k(l-u) concepts relate to the high surface heat load, the absence of tritium breeding, and the higher magnetic fields. It is generally concluded that the allowable wall loading in a D-T reactor with helium coolant is limited. Therefore, helium will have a significant disadvantage for the alternate concepts be- cause of high pumping power requirements. The liquid metals, e.g., sodium, and lithium, are good low-pressure coolants. However, the MHD effects are a major concern in D-T reactors and they may be prohibitive in the higher magnetic fields proposed for the alternate concepts. Electromagnetic effects in molten salts may also be more important in the higher magnetic fields. Water, although it requires high pressure and is limited to relatively low temperatures, may be the best choice. A major problem relates to its chemical compatibility with the candidate structural materials that have i.he lowest thermal stress factors. Concerns regarding reactivity with lithium are not important in the alternate fuel cycles since tritium breeding is not required. However, the expense of tritium processing from water coolant may still be necessary. 3.1.4 General First-Wall/Blanket Design Considerations Because of the nature of the plasma energy deposition, the blanket region or volume of sensible heat generation may be considerably less for the alter- nate fuel concepts than for the D-T system. The blanket region will probably consist primarily of structure and coolant, although other moderator and absorber materials may be used to minimize the blanket thickness. This 53 advantage again emphasizes the need for an efficient coolant. It may also be beneficial to divide the blanket into relatively thin layers to obtain acceptable thermal stresses. Massive solid structures will also be more susceptible to differential swelling effects. Although the liquid lithium blanket possesses some disadvantages, it is generally considered to be resistant to radiation damage. Absorbing neutrons (and energy) directly into a liquid has some inherent advantages. Because of the reduced neutron radiation damage characteristic of the alternate fuel cycles, design of a blanket for full reactor life should be a primary goal. From a radiation damage response, this appears to be more probable for the D-3He fuel cycle than for the D-D fuel. Since fatigue problems associated with a cyclic plasma burn and the high first-wall heat flux may be the life-limiting criteria, steady-state operation may be more critical to the development of a viable alternate fuel reactor than for a D-T fueled reactor. 3.2 Nuclear Analysis (1-3) Potential advantages of alternate fuel fusion systems include elimination of the need for a tritium breeding blanket, reduction of radia- tion damage to reactor components, and reduction in the radioactive material inventories. These potential merits associated with the use of alternate fuel cycles are based on two characteristics of those cycles; (1) low neutron power fraction or low neutron wall load, and (2) softer neutron energy spectrum. The second characteristic reflects the fact that in alternate fuel systems such as D-D or D-3He, the 2.45 MeV neutrons induced by the D(D-3He)n reaction comprises a substantial portion of the neutron source, as contrasted to all of the source neutrons being of 14 MeV in D-T systems. Figure 3-1 compares the intrinsic attenuation characteristics of the 2.45 MeV and 14 MeV source neutrons in a typical fusion blanket system. Shown is a nuclear radiation dose in an epoxy-base superconducting magnet insulator for a given integral neutron wall loading. It is seen that the difference in the source energies (or in the degree of spectrum softening) has a very slight design impact on the radiation shielding. Due to the fact that there are more 54 10 0.6 07 O.B 0.9 1 11 12 FIRST WALL/BLANKET/SHIELD THICKNESS (m) First Wall: 10 mm SS (Density Factor = 1.0) Blanket: 0.3 m SS (Density Factor = 0.85) Shield: 50% SS + 50% B4C (Density Factor = 0.85) Figure 3-1. Importance of the source neutron energies on dose in epoxy - insulator. 55 2.45 MeV neutrons than 14 MeV neutrons per unit energy (or per unit wall load), alternate fuel systems whose neutron power is made up in a large portion by the 2.65 MeV source neutrons may have a higher neutron wall flux than a D-T system with the same total wall loading. Figure 3-2 represents another aspect of the design impact of the difference in the source neutron energies. Shown in Fig. 3-2 is the bio- logical hazard upon inhalation (BHP . ) associated with stainless steel air blanket/shield systems for the two neutron energies. One finds that beyond 30 yr, in particular, times larger than 50 yr after shutdown, the BHP . in the soft neutron spectrum system exceeds that in the D-T source system. The results of Fig. 3-2 stem from the fact that many of the long-term radioactive isotopes, particularly 63Ni in this case, are induced to an appreciable degree through (n,y) reactions which are favored by low-energy neutrons. Most of the analysis in this report is done for the same total wall loading. The reason for this is that one is then comparing systems with the same total energy (or power) production. It is believed that this is the most valid basis of comparison. 3.2.1 Scope of Analysis The analysis is oriented toward quantifying advantages as well as disadvantages of alternate fuel reactor designs in order to make more pre- cise comparisons with mainline D-T fuel reactor designs. The alternate fuel systems that are studies in this section include catalyzed-D-D (Cat-D) and lean-D-3He (D-3He). Also included in the study are two reference D-T fuel systems with and without a tritium breeding blanket. Characteristics of these systems are given in Table 3-3. The selection of materials for the primary structure, coolant (and breeding material in the case of D-T cycle) is one of the central design issues that must be developed from comprehensive design interactions and trade-offs. The trade-offs will involve careful considerations of the material compatibility, thermal hydraulic performance, material radiation 56 "ca I m 0.1 1 10 100 1000 TIME AFTER REACTOR SHUTDOWN (YR) Neutron Wall Load: 1 MW/nT Reactor Power/Unit Length: 1 MW 2 Yr Reactor Operation Structural Material: 316 SS First Wall: 10 mm SS (Density Factor = 1.0) Blanket: 0.3 m SS (Density Factor = 0.85) Shield: 1.2 m, 50% SS + 50% B4C (Density Factor = 0.85) Figure 3-2. Importance of the source neutron energies on biological hazard potential. 57 Table 3-3. Characteristics of the Alternate Fuel Systems Studied Wall Loading (MW/m2) Source Neutron Split (%) In Number Total Neutron In Energy 14 MeV 2.45 MeV 14 MeV 2.45 MeV DT-Reference 1.0 0.80 100 100 Cat-DD 1.0 0.40 50 50 85 15 D-3He 1.0 0.01 25 75 66 34 damage, neutronics performance, etc., as well as the economical impact. In this section several candidate materials are tentatively chosen for the purpose of neutronics comparisons. These are: (1) Structural Material • Type 316 stainless steel (SS) • V-15Cr-5Ti alloy (V15Cr5Ti) • Ti-4Al-2.5V-8Sn-0.5Si alloy (T14381) ** (2) Coolant Material • Helium gas (for nonbreeding D-T, Cat-D and D-3He) • Liquid lithium (for breeding D-T) (3) Tritium Breeding Material • Liquid lithium (for breeding D-T blanket only) (4) Shielding Material • 50% structural material and 50% boron carbide (B^C) (5) Biological Shielding Material • 5% reinforced (Fe) normal concrete. ' See Table 3-4 for compositions. **The shield is assumed to be cooled by a helium coolant for all the cases. 58 Table 3-4. Structural Material Composition Type 316 Ti-4 Al-2.5 V-8 Sn-0.5 Si<6) Stainless Steel^ ' V-15 Cr-5 TiVJ;1 (Ti 4381) Element wt-% atom/b-cm wt-% atom/b-cm ut% atom/b-cn C 0.058 2.286(-4) 0.02 6.118(-5) 0.01 2.267(-5) N 0.007 2.366C-5) 0.05 1.311(-4) 0.008 1.555(-5) 0 0.05 1.148(-4) 0.065 1.106 (-4) Al 0.004 5.447(-6) 4.0 4.038C-3) Si 0.460 7.752C-4) 0.03 3.924(-5) 0.5 4.848(-4) P 0.026 3.974(-5) 0.01 1.186(-5) S 0.011 1.624(-5) Ti 0.040 3.953C-5) 5.00 2.835(-3) 84.9 4.825(-2) V 79.794 5.754(-2) 2.5 1.336(-3) Cr 16.70 1.520C-2) 15.00 1.060(-2) 0.01 5.236(-6) lln 1.430 1.232C-3) 0.0025 1.239C-6) Fe 64.44 5.462C-2) 0.01 6.578(-6) 0.02 9.752(-6) Co 0.030 2.410(-5) Ni 13.90 1.12K-2) 0.001 6.258C-7) 0.005 2.319(-6) Cu 0.060 4.470(-5) 0.01 4.286(-6) Ga 0.01 5.270(-6) Zr Nb 0.0025 4.996(-7) Mo 2.840 1.40K-3) 0.008 3.164(-6) 0.005 1.419C-6) Sn 8.0 1.835C-3) Ta 0.003 1.149(-6) W 0.0075 1.523(-7) Density 7.86 (g/cc) 6.10 4 52 The analysis focuses on the following: (1) radiation damage to the first wall; (2) blanket and shield performance; (3) radiation damage to the super- conducting magnet; (4) reactor activation and environmental impact; (5) bio- logical shielding; and (6) penetration shielding. The particle transport calculations were carried out using the one- V dimensional ANISJT ' and the two-dimensional DOT ' codes with the S8-P3 approximation. The system dimensions used in the present analysis are given in Table 3-5. The 46 neutron-group/21 gamma-group cross section library for the particle transport calculations was generated from the DLC37 library which is based on the ENDF/B-IV version. The nuclear response functions (12) were processed from MA.CKLIB-IV.v ' The neutron induced activation calcula- tions were performed by the general radioactivity calculation code, RACC with the associated 46 neutron group cross section and decay data libraries, RACCXLIB and RACCDLIB/14) Table 3-5. System Dimensions and Material Compositions Breeding DT Non-Breeding DT, Cat-DD, D~3He Region Radii (m) Material Composition Radii (m) Material Composition Plasma 0 -3.50 0 -3.50 Scrape-Off 3.50-4.00 3.50-4 .00 First Wall 4.00-4.01 100% Structure 4.00-4.01 100% Structure /90%Li |85%,Structure Blanket 4.01-a 4.01-4.31 + 110% Structure Il5% He [42.5% Structure (42.5% Structure Shield a-b •{42.5?(XC 4.31-b -442.5%TiBitC [15% He (15% He a. 4.71 316 SS 4.61 V-15 Cr-5 Ti/Ti 4381 b. Variable 60 3.2.2 Radiation Damage to the First Wall Prior to the radiation damage analysis, computations were made to assess the tritium breeding blanket thicknesses required to attain a breeding ratio (BR) of * 1.3. The following results were obtained. Lithium Blanket Thickness Structural Material (m) BR 316 SS (DT70) 0.70 1.28 V15Cr5Ti (DT60) 0.60 1.33 T14381 (DT60) 0.60 1.32 These thicknesses will be used throughout this section for the D-T systems. Table 3-6 compares the four fuel systems of breeding D-T (DT60 or DT70), nonbreeding D-T (DTOO), Cat-D and D-3He in terms of nuclear response rates in the first wall. The response rates are shown for the three candidate struc- tural materials. On comparing the DT6O/DT7O and DTOO systems, one finds that the differ- ence in the blanket material compositions in these systems has a slight influ- ence on the first wall damage caused by the fusion neutrons. It is observed that the Cat-D systems yield relatively high nuclear heating and atomic dis- placement rates, when compared to the D-T systems that have a neutron wall load as much as twice that of Cat-D's. There are two reasons for this phenomenon, i.e., (a) the relatively smooth variation of the associated response functions over the source neutron energies (14 to 2.45 MeV); and (b) the fact that there are more source neutrons in the Cat-D systems than in the D-T case. Such a high heat generation rate in the Cat-D first wall may pose a difficult heat removal problem since, in addition to the neutron/photon heat, a great amount of heat deposited by charged particles and X-ray transport must be retrieved. Figure 3-3 illustrates the effect of the bremsstrahlung radiation on the first wall/blanket heating in the stainless steel structure of the Cat-D and D-3He systems. It is assumed that the ratios of bremsstrahlung to neutron powers in Cat-D and D-3He are 0.65 and 10, respectively, and a monochromatic X-ray of 45 keV is emitted by the radiation. The high heating rate as well as the appreciably deep heat penetration into the first wall/blanket region must be accounted for in the design. 61 Table 3-6. Nuclear Radiation Response Rates At The First Wall FUEL CYCLE DT7O/6O DT00 Ca t-DD D-3He Total Wall Load (MW/m2) 1.0 1.0 1.0 1.0 Neutron Wall Load (MW/m2) 0.8 0.8 0.4 0.01 Nuclear Heating (MW/m3) Neutron 3.66 3.62 1.73 0.041 316 Stainless Steel Gamma 5.08 7.28 4.25 0.126 Total 8.74 10.90 5.98 0.166 Neutron 1.58 1.80 1.00 0.028 V-15 Cr-5 Ti Alloy Gamma 2.68 4.30 2.76 0.088 Total 4.26 6.10 3.76 0.117 Neutron 1.57 1.66 0.87 0.023 Ti 438r ' Alloy Gamma 2.34 3.49 1.89 0.052 Total 3.91 5.15 2.76 0.075 Gas Production (appm/vr) Hydrogen 451 427 190 4.0 316 Stainless Steel Helium 122 117 50 1.0 Hydrogen 214 205 87 1.7 V-15 Cr-5 Ti Alloy Helium 50 49 21 0.4 Hydrogen 174 169 72 1.4 Ti 4381W Alloy Helium 107 105 45 0.9 Atomic Displacement (dpa/yr) 316 Stainless Steel 9.10 10.3 7.08 0.24 V-15 Cr-"5 Ti Allov 8.76 10.5 7.43 0.26 Ti 4381(*^ Alloy (**) W Ti-4 Al-2.5 V-8 Sn-0.5 Si (**) Currently Not Available 62 100 10 H -D with Bremsstrahlunq o Cat-D without Bremsstrahlunq t—1 \ \ % \ \ H o \ D-3He with Brensstrahlunq \ \ D-' He without Brentnstrahiung 01 10 20 30 40 50 DISTANCE FROM FIRST WALL (mm) First Wall/Blanket = SS Bremsstrahlunq Wall Load Cat-D: 0.26 MW/m2 D-3He: 0.10 MW/m2 Figure 3-3. Effect of Bremsstrahlung radiation on nuclear heating structural material: 316 stainless steel. 63 On the other hand, the radiation damage due to gas production is more severe in the D-T systems because the associated damage cross sections have high threshold energies, so that the damage processes are very sensitive to the source neutron energies. There is a substantially smaller radiation damage burden on the first wall in the lean D-3He systems. In particular, the gas production rates in the D-3He systems are reduced by more than 100 times relative to the D-T system for all the structural systems investigated. Comparing the three structural materials, it is found that the nuclear heating rates in the vanadium alloy systems are less than those in the stain- less steel systems by a factor of 1.5 to 2, being reduced by another 10 to 20% in the titanium alloy systems. These results are due primarily to the large differences in both the neutron and gamma kerma factors between those structural materials rather than the difference in the respective particle energy spectra in the first wall. The variation of gas production rates in the three candidate, first wall materials exhibit a trend more or less similar to those for the nuclear heating. Due to the lack of dpa data for titanium in the current version of MACKLIB-IV, the complete comparison is not possible. Although the difference in dpa between stainless steel and V15Cr5Ti is not appreciable, it is observed that the V15Cr5Ti dpa's tend to exceed the stainless steel dpa's in the alternate fuel systems because of the larger V15Cr5Ti dpa cross section in the intermediate energy range of ^ 4 MeV to ^ 2 keV in which the effect of the 2.45 MeV neutron source is most dominant. The atomic displace- ment energies of V, Cr, Fe and Ni are all assumed to be 40 eV. 3.2.3 Blanket and Shield Performance In order to determine the required blanket/shield thickness for each fuel cycle, an analysis was made on the radiation shielding performance necessary for protection of the superconducting magnets which enclose the bulk shield. The magnet protection criterion used here is a nuclear radiation dose limit >x in an epoxy-base insulator. The radiation-induced insulation deterioration is considered to be the only irreversible damage process in the magnet. Therefore, usually the magnet lifetime, and hence, the reactor lifetime itself largely de- pends upon the degree of this deterioration. Nuclear radiation dose rates of 64 107 Gy (109 rad) and 108 Gy (1O10 rad) in the insulator can be considered (although the radiation damage data on the organic insulators at liquid helium temperatures are lacking) to be the upper and lower bounds of the protection criterion. In this section, a dose limit of 5 x 107 Gy (5 x 109 rad) is assumed. Figures 3-4 through 3-6 show the variation of the nuclear doses as a function of the bulk shield thickness for the four fuel-cycle systems and for the three structural materials. The dose variation is represented as a plant lifetime in terms of MW-yr/m2 of total integral wall loading. It is seen in the figures that the Cat-D and D-3He systems offer an improved nuclear performance, compared to the DT systems, from the magnet shielding point of view. Obviously the best performance among those systems investigated in the present analysis is offered by the D-3He cycles in which the neutron power fraction is only 1%. The most significant difference in the shielding requirement is brought about when the tritium fuel multiplication is enforced in a D-T blanket, particularly in the inboard section of tokamak D-T reactors. If the breeding blanket is designed to attain a BR of ^1.3 as the case in this study, the increase in the DT blanket/shiald thickness must be "\< 0.4 m and ^ 0.7 m relative to the Cat-D and D-3He systems, respectively. These substantial reductions in the required blanket/shifild thickness will have important impacts on the overall reactor performance, magnet design, cost of the reactor, etc. For example, under the condition of a plant lifetime of 30 MW-yr/m2 total integral wall load, the required blanket/shield thickness for each system is calculated and shown in Table 3-7, Table 3-8 summarizes the results of system energy deposition and multiplication in the four fusion fuel systems. The energy deposition is defined as the energy absorbed in the entire system (including the first wall, blanket and shield) per source neutron of 14.1 MeV, 8.3 MeV and 5.4 MeV for D-T, Cat-D and D-3He, respectively. The energy multiplication represents the ratio of the energy deposited to the average source neutron energy. For all the fuel cycles studied, the heat depositions in the 10 mm thick 316 stain- less steel first wall result in *> 10% of the total, compared to "» 5% in both 65 Q < 100 O -I 0 o W E-< I*, h-1 0.4 0.6 0.8 1 12 1.4 16 REQUIRED BLANKET/SHIELD THICKNESS (m) Dose Limit in Epoxy-Insulation: 5 X TO7 Gy Figure 3-4. Effect of alternate fuel systems on shielding require- ment structural material: 316 stainless steel. 66 100 H-l IS o y / He Breedinqing) ) / IDT60 10 cat-n f. /moo ui J.M I / \ 0.4 0.6 II0.8 1 \Z 1.4 1.6 REQUIRED BLANKET/SHIELD THICKNESS (m) Oose Limit in Epoxy-Insulation: 5 x 10 Gy Figure 3-5. Kffect of alternate fuel systems on shielding require- ment structural material: V-15Cr-5Ti. 67 2 100 t•^—B- -B- •/• 3 o 22 I / DTOO . D-3He i Cat-D Breediinq) 10 •:/:: X.i I ...l. / _5 0.4 0.6 OB 1 \Z 1.4 1.6 REQUIRED BLANKET/SHIELD THICKNESS (m) Dose Limit in Epoxy-Insuiation: 5 x ID Gy Figure 3-6. Effect of alternate fuel systems on shielding require- ment structural material: T14381. 68 Table 3-7. Shielding Performance of Alternate Fuel Systems for a Total Integral Wall Load of 30 MW-yr/m^ FUEL CYCLE DT70/60 DT00 Cat-DD D-3He Total Integral Wall Load (MW-yr/m2) 30 30 30 30 Integral Neutron Wall Load (MW-yr/m2) 24 24 12 0.3 Required First Wall/Blanket/Shield Thickness (m) for Epoxy Insulator Dose Limit of 5 x 109 rad Structural Material 316 Stainless Steel 1.45 1.04 0.99 0.73 V-15 Cr-5 Ti Alloy 1.48 1.16 1.10 0.77 Ti-4 Al-2.5 V-8 Sn-0.5 Si Alloy 1.64 1.38 1.29 0.92 of V15Cr5Ti and Ti4381 first walls. The nuclear heating reactions taking place in the blanket regions yield 80-85%, 75-85% and 60-85% of the system heating rates in 316 stainless steel, V15Cr5Ti and T14381, respectively. Due to the exothermic reaction or 6j'i(n,a)t (the reaction Q value of 4.8 MeV) in liquid lithium, the nuclear heating in the breeding blanket systems is dominated by the energy release associated with neutron reactions, which contrast to the nuclear heating due primarily to gamma interactions in the nonbreeding D-T, Cat-D, and D-3He systems. The relatively high nuclear heating rates in the V15Cr5Ti and Ti4381 shield regions are related to the less effective radiation attenuation characteristics of these alloy materials as already shown in Table 3-7. It is worthwhile to note that the system energy multiplication per source neutron in the alternate fuel systems is remarkably large. Although the neutron power fractions in alternate fuel systems are small when compared to D-T systems, the results shown in Table 3-8 indicate that the overall 69 Table 3-8. System Energy Multiplication of Advanced Fuel Systems Per Source Neutron FUEL CYCLE DT70/60 DT00 Cat-DD D-3He Average Source Neutron Energy (MeV) 14.06 14.06 8.26 5.43 Energy Deposition (MeV) First Wall 1.54 1.93 1.24 0.90 Blanket 15.4* 15.42 11.20 9.15 316 Stainless Steel Shield 1.00 1.77 1.60 1.51 Total 17.96 19.12 14.04 11.56 First Wall 0.75 1.08 0.78 0.63 Blanket 14.16 12.06 9.30 7.95 V-15 Cr-5 Ti Shield 1.47 2.57 2.13 1.92 Total 16.38 15.70 12.21 10.51 First Wall 0.69 0.91 0.57 0.41 Blanket 14.90 10.23 6.80 5.13 Ti 4381 Shield 1.86 4.38 3.65 3.29 Total 17.45 15.53 11.02 8.82 System Neutron Energy Multiplication 316 Stainless Steel 1.28 1.36 1.70 2.13 V-15 Cr-5 Ti 1.17 1.12 1.48 1.94 Ti 4381 1.24 1.10 1.33 1.62 L 70 system energy multiplication in alternate systems, particularly in Cat-D systems, are significant. For example, the Cat-D system using the 316 SS structural material yields a system energy multiplication (based on the source energy of neutron plus radiation transport) of 1.70 which is compared to 1.28 for the corresponding 316 SS D-T breeding systems. 3.2.4 Radiation Damage to Superconducting Magnets Based on the blanket/shield thicknesses given in Table 3-7, the maximum nuclear response rates in superconducting magnets are calculated as shown in Table 3-9. Each group of the radiation response rates among the four fuel cycles and the three candidate structural materials shows an agreement within a factor of ^ 3. The largest difference can be found in the gas production rates (induced by the threshold reactions) which are sensitive to the high energy portion of neutron flux, and hence to the bulk shield thickness. The response rates other than the gas production do not differ more than 'v 50% from one system to another. Note that these response rates rapidly decrease inside the magnets. Assuming a total magnet volume to be of the order of 1000 m^, the anticipated heat load in the magnets becomes much less than 100 kWth. Based on a power requirement for the heat removal of about 300 W /W h» the refrigeration power consumption is only a few percent of the reactor power output. The resistivity increase induced by the nuclear irradiation over a 30 yr reactor operation is estimated to be about 3% of the saturation values for either copper or aluminum stabilizer. A resistivity increase of 71 Table 3-9. Maximum Response Rates in Superconducting Magnets FUEL CYCLE DT70/60 DTOO Cat-DD D-3He Total Wall Load (MW/m2) 1.0 1.0 1.0 1.0 Neutron Wall Load (MW/m2) 0.8 0.8 0.4 0.01 Heating (MW/m3) 316 Stainless Steel 7.38(-5) 7.24(-5) 7.71 (-5) 6.97(-5) V-15 Cr-5 Ti 9.08(-5) 9.68(-5) 1.00(-4) 1.02(-4) Ti 4381 8.79(-5) 8.46(-5) 1.00(-4) 8.07(-5) Hydrogen Production (appm/yr) 316 Stainless Steel 7.38(-4) 5.92(-4) 4.68(-4) 2.27(-4) V-15 Cr-5 Ti 5.67(-4) 5.59(-4) 4.63(-4) 3.49(-4) Ti 4381 5.33(-4) 4.73(-4) 4.88(-4) 3.50(-4) Helium Production (appm/yr) 316 Stainless Steel 1.25(-4) 1.02(-4) 8.08(-5) 3.97(-5) V-15 Cr-5 Ti 9.36(-5) 9.01(-5) 7.49(-5) 5.76(-5) Ti 4381 8.79(-5) 1.02(-4) 1.06(-4) 7.9K-5) dpa in Structure (dpa/yr) 316 Stainless Steel 4.54(-5) 3.73(-5) 3.39(-5) 3.02(-5) V-15 Cr-5 Ti 4.77(-5) 4.13(-5) 3.55(-5) 3.90(-5) Ti 4381(a) dpa in Cu - Stabilizer (dpa/yr) 316 Stainless Steel 6.52(-5) 5.38(-5) 4.95(-5) 4.60(-5) V-15 Cr-5 Ti 5.96(-5) 5.24(-5) 4.49(-5) 4.90(-5) Ti 4381 6.240-5) 5.06(-5) 5.33(-5) 4.76(-5) 72 Table 3-9. Maximum Response Rates in Superconducting Magnets (cont'd.) FUEL CYCLE DT70/60 DTOO Cat-DD D-3He dpa in Al - Stabilizer (dpa/yr) 316 Stainless Steel 9.93(-5) 8.58(-5) 8.11(-5) 8.09(-5) V-15 Cr-5 Ti 9.08(-5) 7.75(-5) 6.72(-5) 7.98(-5) Ti 4381 9.93(-5) 7.52(-5) 7.98(-5) 7.58(-5) Resistivity Increase ' in Cu - Stabilizer (fi-cm x 10"8) 316 Stainless Steel 1.08(0) 8.95(-l) 8.25(-l) 7.67(-l) V-15 Cr-5 Ti 9.90(-l) 8.72(-l) 7.49(-l) 8.16(-1) Ti 4381 1.04(0) 8.43(-1) 8.87(-l) 7.93(-l) Resistivity Increase in Al - Stabilizer (fi-cm x 10~8) 316 Stainless Steel 2.86(0) 2.47(0) 2.34(0) 2.33(0) V-.15 Cr-5 Ti 2.61(0) 2.24(0) 1.94(0) 2.30(0) Ti 4381 2.86(0) 2.17(0) 2.30(0) 2.19(0) (a) dpa data not available based on 3 x 10~7 [l-exp(-563 dpa)] £2-cm (20) based on 8 x 10"7 [l-exp(-366 dpa)] fl-cm 20) 73 3.2.5 Neutron Energy Spectra in Alternate Fuel Systems The blanket/shield performance of a fusion reactor strongly depends on the extent to which the neutron energy moderation takes place in the system. The essential difference in the nuclear performance between different fusion fuel systems must be understood based on a st-idy of the characteristic neu- tron energy spectra. In Fig. 3-7, the blanket neutron spectra at 0.15 m from the first wall are plotted for DT70, DTOO, Cat-D and D-3He systems with 316 SS. Figure 3-8 shows the corresponding fractional spectrum of neutrons with energies > E_ as a function of a given energy, EQ. From these figures, it is found that the spectrum profiles of nonbreeding DT(DTOO), Cat-D and D-3He are almost identical at energies below 2.45 MeV where the source neutrons from the D(D-3He) reaction are generated. The difference in their absolute fluxes corresponds approximately to the difference in the number of source neutrons. The implication is that, except for threshold-type reactions in which the flux of neutrons with energies > 2.45 MeV is important, the system nuclear characteristic is largely determined by the total number of source neutrons and is less dependent upon the source neutron composition. The most remarkable difference can be seen in the breeding blanket spectrum which is substantially harder than the other cases. In fact, at ^ 1 keV the fractional spectrum of DT70 reaches i> 100%, compared to "v 80% in the other fuel systems. The flux dip around ^ 250 keV in the DT70 blanket is due to the resonance reactions of lithium. 3.2.6 Reactor Activation and Environmental Impact In order to achieve an environmentally acceptable reactor concept, the safety and radiological issues must be taken into account early in the design effort. One of the safety and environmental impacts of fusion power reactors, as envisioned today, are related to the inventories of neutron-induced acti- vation of reactor components and radioactive tritium (particularly in D-T systems). This makes a vivid contrast to the hazard concerns of fission reactors in which fuel actinides and fission products represent more than 99.9% of the substantially high level of radioactivity inventories.^23^ 74 > n E- 10 10u 101 102 103 104 105 10' NEUTRON ENERGY (eV) Neutron Wall Load: MW/nT DT7O/DTnO: 0.8 Cat-D: 0.4 D-3He: 0.01 Figure 3-7. A comparison of alternate fuel systems on neutron spectrum in blanket structural material: 316 stainless steel. 75 o o t—I u 0J3- 10 10 10' 10" 10' 10^ 10' 10' 10' NEUTRON ENERGY EQ (eV) Neutron Wall Load: MW/m2 DT70/DT00: 0.8 Cat-D: 0.4 D-3He: 0.01 Figure 3-8. A comparison of alternate fuel systems on fraction spectrum in blanket structural material: 316 stainless steel. 76 The reactor activation problem is investigated in two aspects, i.e., (1) short-term radioactivity and (2) long-term radioactivity. Primary con- cerns regarding the former are (a) decay afterheat problems and (b) potential of accidental release of activated materials into the environment. The latter problem of reactor activation is concerned with long-term radwaste storage requirements and with material recycling considerations. 3.2.6.1 Short-Term Activation and Afterheat in First Wall The impact of accidental release of radioactive materials is studied elsewhere in this report. The analysis here is focused on other short-term radiological impacts, in particular, on the effect of radioactive decay on afterheat. Table 3-10 compares several fuel systems in terms of radioactivity, biological hazard due to inhalation (BHP-air) and decay afterheat at two time points of zero time and one day after shutdown following a two-year reactor operation. It has turned out that the Cat-D systems, particularly the vanadium alloy structured Cat-D system, yield higher radioactivity and BHP-air levels than the D-T systems. Furthermore, the difference in these radioactivity parameters between D-T and D-^He is not proportional to the difference in their neutron wall loads. Due to the relatively large number of soft neutrons in the alternate fusion systems, the radiological importance per unit source energy in these systems becomes more significant than in D-T systems. The decay heat load in the first wall immediately after scheduled or unscheduled reactor shutdown appears to be sizable. Although this level of heat load should not pose any difficult design constraints, it remains to be further studied to draw a definite conclusion. This problem is particularly important as to whether an emergency or auxiliary cooling system should be provided for the first wall in alternate fuel reactors. For all of the short-term radioactivity measures studied here, the effect of structural material choice is not so appreciable. It is worthwhile to point out that when compared to typical thermal or fast breeder fission reactors, all of the candidate fusion systems seem to be substantially safer in terms of the potential hazard in case of accidents. In particular, D-3He systems enhance such an attractive feature of the alternate fusion fuel cycles. 77 Table 3-10. A Comparison of Short-Term Radiological Impacts of Alternate Fuel Systems Heating Rate at Afterheat at First Wall Structural Fuel Radioactivity BHP-AIR First Wall (% of Operating Nuclear Material Cycle (Ci/MWth) (km3/MWth) (% of Reactor Power) Heating Rate) Time After Reactor Shutdown 0 1 Day 0 1 Day 0 1 Day 0 1 Day DT70 3.6(5) 2.2(5) 9.4(4) 8.0(4) 8.0 (-2) 2.K-2) 1.8(0) 4.8(-l) 316 Stainless Cat-DD 6.1(5) 2.9(5) 1.0(5) 8.0(4) 5.0(-2) 9.2(-3) 1.7(0) 3.0<-l) Steel D-3He 1.3(4) 7.1(3) 2.3(3) 1.7(3) 1.3(-3) 1.9(-4) 1.6(0) 2.3(-l) DT60 2.0(5) 3.9(4) 6.2(3) 4.2(3) 4.8 (-2) 8.6(-3) 2.2(0) 4.0(-l) 00 V-15 Cr-5 Ti Cat-DD 1.8(6) 6.9(4) 1.9(4) 1.3(4) 1.4(-1) 3.8(-3) 7.4(0) 2.0(-l) D-3He 6.1(4) 1.8(3) 6.4(2) 4.5(2) 4.9(-3) 8.K-5) 8.3(0) 1.4(-1) DT60 1.5(5) 8.1(4) 4.7(4) 4.2(4) — — Ti 4381 Cat-DD 2.0(5) 1.2(5) 5.6(4) 5.1(4) — — D-3He 4.8(3) 2.6(3) 1.1(3) 9.9(2) Fission Reactor 5.6(6)a 1.4(6)a 5.7(6)b — 7.3(0)a 5.7(-l)a — ~ • Based on a specific power generation rate of 1 MW/cm in the axial direction of infinite cylindrical model. • Reactor shutdown after 2 yr reactor operation. • Total wall load is 1 MW/m2. • Based on 1100 MWe PWR Core (37.5 MWth/metric ton of 3.3% enriched U; 82 metric tons of U) after 293 day irradiation. •The afterheat data is for the system total. •Based on an LMFBR 3.6.2.2 Long-Term Activation Problem Considering the current concerns regarding the siting problems of storage of radioactive waste materials discharged from nuclear power plants, the long-term radwaste storage requirement for fusion reactors is one of the important issues to be studied in the early stage of design effort. At the same time, a deliberate assessment of the reactor materials recyclability must be made, taking into account the current attention to the availability of limited reactor materials resources. Both of these two problems, i.e., radwaste storage and material recycling fall into a common consideration of how reactor components will be handled or processed after the reactor has been decommissioned. The objective of this section is to provide the basic information on long-term radioactive material inventories in the candidate fusion fuel sys- tems for the future analysis of reactor decommissioning. Figures 3-9 through 3-11 show comparisons of the fusion systems in terms of BHP-air as a function of time after reactor shutdown following two-year reactor operation. Over the time period from five years to 50 years, the titanium and vanadium systems exhibit three to four orders of magnitude lower BHP's than the corresponding 316 stainless steel systems. As shown in Fig. 3-12, this is due largely to radioactive isotopes ^^Co (half-life of 5.3 yr, decay — (?fi^ mode of B ), and 63Ni (100 yr, g~) induced directly or indirectly from a series of nickel isotopes contained in 316 stainless steel as the primary constituents. These nickel isotopes appear only as impurities in both V15Cr5Ti and T14381. At times larger than 50 yr after shutdown, the BHP's are determined mostly by the impurity levels, in particular, by Nb and Mo that lead to the formation of 93mNb (16.6 yr, IT), 9l*Nb (20,000 yr, 0~) and 93Mo (3,000 yr, EC). Note that 93mNb contributes to the long-term activation through a two-step decay process of 93Mo -+ 93mNb -> 93Nb where the 93Mo decay half-life is much longer than the 93raNb. The only exceptions in which the primary structural material elements represent the major sources for long- lived radioactive isotopes are: (1) 26A1 (7.2 x 105 yr, 6 /EC) induced by the 27Al(n,2n) reaction in TiA381 and (2) 63Ni (100 yr, g~) induced by the 62Ni (n,y) and 61 79 10-1 10 100 1000 TIME AFTER REACTOR SHUTDOWN (YR) Specific: Power/IJnii: Length: 1 MW Neutron Wall Load: MW/n2 DT70: 0.80 Cat-D: 0.4 D-3He: 0.01 2 Yr Reactor Operation Figure 3-9. A comparison of alternate fuel systems on biological hazard potential structural material: 316 stainless steel. 80 6 Si a a. CQ 1 10 100 1000 TIME AFTER REACTOR SHUTDOWN (YR) Specific Power/Unit Length: 1MW Neutron Wall Load: MW/m2 DT60: O.flO Cat-D: 0.40 D-3He: 0.01 2 Yr Reactor Operation Figiire 3-10. A comparison of advanced fuel systems on biological hazard potential structural material: V-15Cr-5Ti. 81 10-4 1 10 100 1000 TIME AFTER REACTOR SHUTDOWN (YR) Specific Power/Unit Lenqth: 1 MW Neutron Wall Load: MW/m2 DT60: 0.80 Cat-D: 0.40 D-3He: 0.01 2 Yr Reactor Operation Figure 3-11. A comparison of alternate fuel systems on biological hazard potential structural material: T14381. 82 7? V o K-l O < 01 10 100 1000 TIME AFTER REACTOR SHUTDOWN (YR) Reactor Shutdown following 2 Yr Reactor Operation Figure 3-12. Isotopic contribution to BHP-air in Cat-D system structural material: 316 stainless steel. 83 Comparing the different fusion systems for a given structural material, one finds that the Cat-D systems always have a higher hazard potential than the D-T systems. The difference in BHP-air even between D-3He and D-T reduces to only a factor of 10 or so beyond 50 yr after shutdown. Particularly in the case of the 316 stainless steel systems, the D-T BHP decreases steeply after 50 yr reflecting the rapid decreases of 60Co and 55Fe (2.7 yr, EC). On the other hand, both of the alternate fuel systems show a relatively slow decrease of BHP beyond the same time period. As shown in Fig. 3-13, the importance of the 63Ni isotope on the BHP in the alternate systems is much larger than that in the D-T system over the entire time span considered here. Obviously the higher degree of importance of the 63Ni isotope in the alternate systems stems from the more dominant low energy neutron fluxes in these sys- tems, which tend to induce more 63Ni through the 62Ni(n,y) reaction. 3.2.7 Biological Shielding In this section the analysis is oriented toward the neutron-induced activation of the reactor containment building. The building is assumed to consist of normal concrete which is reinforced by iron by 5% in volume. A (27) biological dose of 2.5 mrem/hr (following the NRC guideline) is tentatively chosen as the limit criterion assuming an occupational radiation exposure of 5 rem/yr in a 40 hr/week - 50 week/yr work in a restricted area. Using the bulk shield thicknesses determined in Section 3.2.3 (see Table 3-7) based on a plant lifetime of 30 MW/yr/m2 total wall load, the biological doses at the outermost point of the reactor building during reactor operation and at shutdown are calculated as shown in Fig. 3-14. It is seen that the dose during operation (contributed mostly by neutrons) is about three-fold high relative to the dose at shutdown. In addition, once the bulk shield thickness has been determined based on the magnet insulation damage, the impact of different fusion cycles on the biological shielding becomes less significant. In the present case, the required reactor building thick- ness is slightly less than 2 m for all the fuel systems examined. However, the actual biological shield design should evolve from a detailed trade-off study involving the bulk shield design. Increase of the bulk shield thickness has the following obvious impacts: (1) substantial reduction in the reactor building volume and the associated cost; (2) decrease of biological dose inside 84 II E- U < Pi 0.1 10 100 1000 TIME AFTER REACTOR SHUTDOWN (YR) Reactor Shutdown following 2 Yr Operation Figure 3-13. A comparison of alternate fuel systems on 63Ni isotope contribution to BHP-air, structural material: 316 stainless steel. 85 Cat-D • >i DurHnq: Operation 1 S l g 102 8 ill i I Mil S 10 o o 4 UiUM-OMt PQ 10° D -1 10 4 -H i i i W * • I • • 0.5 1 15 2 2.5 BIOLOGICAL SHIELD THICKNESS (m) Specific Power/Unit Length: 1 MW Neutron Wall Load: MW/m2 DT70: 0.80 Cat-D: 0.40 D-3He: 0.01 Structural Material: 316 SS Figure 3-14. Effect of alternate fuel systems on biological shield requirement of normal concrete, structural material: 316 stainless steel. 86 the reactor building; and (3) cost increase of the bulk shield. Figure 3-15 shows the biological dose at the end of bulk shield as a function of the shield thickness. When the dose limit of 2.5 mrem/hr is again adopted, the bulk shield thicknesses given in Table 3-7 are not thick enough to allow a personal access to the reactor building without an adequate remote handling equipment. Even if the dose decays for one or two days, the shield thickness required for "hands-on" maintenance, for example, must be ^ 0.6 m larger than what Table 3-7 indicates. Further study is needed to determine whether such a sizable increase in the bulk shield thickness is feasible or at least attractive from the economical and engineering standpoint. Of course, the "hands-on" or "contact" maintenance is not the only conceivable maintenance mode. Perhaps there will be a strong economic incentive for a maintenance scenario that combines ( oft) contact, semi-remote and remote operations. Depending upon the degree of the remote handling system incorporated, the bulk shield thickness must be varied. At any event, it appears that the biological dose criterion for the reactor maintenance is a more restrictive condition to meet than the magnet insulation damage criterion in the bulk shielding design. 3.2.8 Major Penetration Shielding An important aspect common to fusion blanket/shield designs is that (29) they must accommodate a variety of major penetrations. ' The major penetrations are required for: (a) plasma chamber evacuation (vacuum pumping); (b) supplementary plasma heating (neutral beam or radio-frequency heating); (c) impurity control (e.g., divertors); (d) access for diagnostics and mainte- nance; and (e) cooling. These various types of penetrations not only bring about design difficulties from the magnet shielding standpoint, but become a potential major source for activation of components in the reactor buildings. The analysis in this section is focused on the impact of the candidate fuel systems on the penetration shielding requirement. The shielding require- ment is studied in two aspects: (1) insulation damage to the superconducting magnet; and (2) duct wall activation associated with neutron streaming. A comprehensive parametric study of the penetration shielding for D-T reactors 87 10 10 12 1.4 . 1.6 18 Z ZZ 2.4 2.6 BLANKET/SHIELD THICKNESS (m) Specific Power/Unit Length: 1 MW 2 Neutron Hall Load: MW/m DT70: 0.80 Cat-D: 0.40 D-3He: 0.01 Reactor Shutdown following 2 Yr Reactor Operation Figure 3-15. A comparison of alternate fuel system on biological dose at reactor shutdown, structural material: 316 stainless steel. 88 has been done elsewhere. Making use of the same two-dimensional model as in Ref. (29), the computations here were performed by the DOT-III code. In the present analysis, the penetration duct dimension is fixed at: (1) duct length: 4 m; (2) duct opening: circular cross section of 0.8 m in diameter; and (3) duct connection: perpendicular to the bulk blanket/shield. The duct is surrounded by an 0.3 m thick 316 stainless steel followed by a shield of 50% B4C + 50% SS. Figure 3-16 shows the maximum absorbed dose in the epoxy-insulator of superconducting magnets vs. penetration shield thickness. A dose limit of 5 x 107 Gy (5 x 109 rad) is for a plant lifetime of 30 MW-yr/m2 total wall load. The dose variation is almost identical in the DT70 and Cat-D systems, and both systems require a total penetration shield thickness of 0.5 m (0.3 m 316 SS plus 0.2 m 50% B4C + 50% SS) to meet the insulator protection criterion. The D-3He system requires a less shield thickness by 0.15 m for the same dose criterion. The major penetrations seem to cause a difficult design constraint to solve in terms of the reactor activation. As shown in Fig. 3-17, the biological dose decreases only by a factor of 100 along the 4 m duct wall. This slow decrease contrasts to a rapid gamma-ray attenuation in the bulk blanket/shield region without penetration which is also shown in Fig. 3-17 for the sake of comparisons. Even if one accounts for a dose reduction due to reactor shut- down by a few orders of magnitude (relative to the dose during reactor opera- tion) , the anticipated biological dose along the duct wall seems to be no less than 106 mrem/hr for all the systems investigated. This implies that shortly after shutdown, the reactor can be accessed only by a remote system or by personnel only if a sufficient shield is afforded around the penetration duct. It appears that the shielding requirement for the major penetration (as for the bulk shielding) is a more binding condition to meet from the reactor activation standpoint than from the magnet protection standpoint for a broad range of biological doses of interest. In particular, the high level of dose near the end of duct where a neutral beam injector, for instance, is located, may post a difficult shielding design problem. This observation is particularly important when one considers a large volume of the injector housing to be shielded. A further detailed investigation is needed to deter- mine what kind of maintenance mode (as to whether partially remote or fully 89 10 0.30 036 040 045 050 056 oao PENETRATION SHIELD THICKNESS (m) Integral Neutron Wall Load: MW-yr/m DT70: 24 Cat-D: 12 D-2He: 0.3 Penetration Shield Material: 0.3 m SS followed by (50% B4C + 50% SS) Duct Length: 4 ra Duct Diameter: 0.8 m Duct Opening: Circular Figure 3-16. A comparison of alternate fut1. systems on penetration shield requirement, structural material: 316 stainless steel. 90 10' IN BLflNKET/SHIELD WITHOUT PENETRfiTION 1.5 2.5 3.5 DISTANCE FROM FIRST WALL (m) Neutron Wall Load: MW/nT DT70: 0.80 Cat-D: 0.40 D-3He: 0.01 Structural Material: 316 SS Duct Diameter: 0.80 m Duct Length: 4 m Duct Opening: Circular Figure 3-17. Effect of alternate fuel systems on biological shielding during reactor operation. 91 remote) is economically attractive and is feasible from the engineering point of view in conjunction with the major penetration shield design. 3.3 Tritium and Fuel Processing Considerations Tritium, vacuum, and fuel processing considerations were evaluated for deuterium-based alternate fuel fusion reactors. Two types of alternate fuel reactors, Cat-D and D-3He, were evaluated and compared to a reference case commercial D-T fusion reactor. Fuel cycles, vacuum pumping and fuel handling issues will be discussed. 3.3.1 Fuel Cycle Considerations Fuel cycle scenarios were developed for deuterium-based alternate fuel fusion reactors. A schematic of the fuel processing cycle for catalyzed- D-D is illustrated in Fig. 3-18. Whether the reactor has a divertor or not, the exhaust from the plasma will eventually be pumped by the vacuum system, most likely by compound cryopumps. Upon regeneration of the vacuum pumps, the fuel is processed for chemical purification, isotopic enrichment, storage, and refueling. The chemical purification subsystem is designed to remove all condensible impurities (CD , D2O, N2, etc.). The helium is then separated from X the D2 fuel by a falling film condenser. The possibility exists that the helium, which is pumped separately, may be recovered from the vacuum system such that it is sufficiently free of tritium that it can be processed directly. The helium is then isotopically enriched and 3He is a product, which is a valuable fuel for a D- He reactor. Isotopic separation of helium isotopes (30 31) appears to be straightforward, using cryogenic distillation. ' Separa- tion factors are reportedly higher than for cryogenic distillation of hydrogen (32) isotopes. The hydrogen isotopes are then isotopically distilled at "v 20 K. It is necessary only to remove a small amount of the protium as HD from the top of the column, as the tritium is recycled as fuel. Qualitatively, the components of the fuel processing cycle are very similar to a D-T fuel pro- (33) cessing cycle (Fig. 3-19) with the additional feature of isotopic enrich- ment of helium. The D-3He fuel processing cycle (Fig. 3-20) is quite similar to that of the D-D cycle, except that 3He is used as a fuel and tritium is not recycled. 92 D-D(CAT) D I D2 + DT + He He VAC sep 0. N, C, etc. 1 rl»He BRIS SEP\ HD Pr) D (yVaste Pr) I S T Figure 3-18. Fuel cycle scenario for Cat-D. 93 REACTOR BUILDING | TRITIUM FACILITY BUILDING 1. Plasma Chamber 15. Tritiated Waste Treatment 2. Limiter Plates 16. Water/Tritium Recovery Unit 3. Debris Separator 17. Helium (tritium-free) 4. D-T Cryocondensation Pump 18. Tritiated Waste — Liquids 5. Helium Pump and Solids 6,7. Regeneration Pumps 19. Detritiated Gasses: N2, O2, 8. Metal Bellows Pumps C02> Ar 9. Breeder Blanket 20. Isotopic Separation Unit 10. Electrolysis Unit 21. D2 Supply 11. Fueling Cleanup Unit 22. D2 Storage 12. Tertiary Enclosures 23. DT and T2 Storage 13. Emergency Air Detritiation 24. T2 Shipment/Receiving System 25. Fuel Blender 14. Secondary Enclosures, Purge 26. Gas Fueling Streams 27. Pellet Fueler Figure 3-19. Tritium facility scenario for Cat-D. 94 D-3He Figure 3-20. Fuel processing cycle for D-3He. 95 The hydrogen isotope separation unit then must separate D-T as well as HD. Although deuterium is plentiful, 3He is not. Potential sources of 3He will be discussed in Sec. 3.3.4. Mass flow rates in the fuel processing cycle have been determined for a number of D-D and D-3He fusion reactors. The results (Table 3-11) include both tokamak reactors and small field-reversed mirrors. Comparison of Cat-D parameters with comparable cases show that vacuum pumping requirements (a major cost item) are comparable.- The D-3He cycles require significantly higher helium pumping speeds. As shown in (Table 3-12), Cat-D and particularly D-3He achieve substantial reductions in tritium inventories. The alternate fuel reactors do not need to breed tritium and thus tritium control in heat transfer fluids is also simplified. Specific vacuum pumping and fuel handling issues are discussed below. 3.3.2 Vacuum Pumping The vacuum pumping requirements for the cases considered are shown in Table 3-11. The required pumping speeds for hydrogen isotopes and helium for Cat-D cases are quite similar to the reference D-T, while the D-3He requires significantly higher helium pumping speeds. The ability to pump helium in a mixture of hydrogen isotopes presents a significant technological challenge. It has been shown that hydrogen isotopes can prevent cryosorptlon pumping of helium and it is, therefore, necessary to utilize a compound pump, in which the hydrogen isotopes are removed by cryocondensation and the helium is then pumped separately. Recently, compound cryopumps of various designs are being developed for the TSTA. The candidate designs all utilize cryo- sorption pumping for hydrogen isotopes with various concepts for He pumping. Recent results show that with a compound cryosorption pump, the He can be pumped and regenerated such that it will be essentially free of tritium and hyd §e; isotopes. 3.3.3 Fuel Processing and Tritium Safety The required fuel processing rates are shown in Table 3-11. From these rates, tritium inventories (Table 3-12) were calculated. The "vulnerable" inventory is that portion of the tritium which is located in the reactor building. Tritium physically located in the separate fuel processing facility 96 Table 3-11. Fuel Processing Requirements for Alternate Fusion Fuel Cycles Fusion Fractional Fuel Burned Plasma Exhaust Required Pumping 3 Power Burnup per dav (B) (g/day) Speed (m /s) (MW) D He D T ^He H D T He \ ••He 0 HDT He Tokamaks 1 Reference D-T 2500 .05 230 345 - 0.46 4,370. 6,560. - 460. 2 .3 1.8 x 10 * 9 5 x 102 Cat-DD-Case 1 (Burn h of 3He 2500 .05 50 654 178 89 90.5 12,400. * 89.2 357. 6 .5 2.6 x 10* 9 8 x 103 Cat-DD-Case 2 (Burn none of 3He) 2500 .05 00 814 244 - 83. 15,500. * 244. 325. 8 .1 3.2 x 101* 1 3 x 103 Cat-DD-Case 3 (Burn all of 3He) 2500 .05 1 .00 481 120 ; 120 41. 9,140. * - 321. 4 .8 1.9 x lO* 6 6 x 102 10 D- 3He 3 1 (lean D) 2500 .10 .05 221 - 331 111. 1,990 .5. 6,290. 441. 2 .2 4.6 x 10 1 .8 x 10 * i ! Field Reversed i Mirrors Cat-DD Case 1 1.3 .28 .48 0.42:0.12; 0.057 .06 1.07 .024 .062 .20 t 004 2.24 0 53 Case 2 3.9 .29 .45 1.22 0.29 0.16 .11 3.0 .052 .19 .60 •012 6.32 1 59 D- 3He Case 1 .,1.0 .24 .20 0.11 - 0.15 .052 .33 .0008 .61 .20 001 0.81 1 89 Case 2 2.5 .33 .28 0.26 1 " 0.36 .125 .53 .0016 .95 .50 003 1.45 3 .28 Case 3 4.6 • 36 .30 0.48 - 0.66 .23 .84 .0028 1.51 .90 005 2.42 5 .43 Reference D-T 5.0 .40 - 0.51 0.77 .001 .77 1.15 - 1.02 005 2.86 1 .90 *Recycled Table 3-12. Tritium Inventories (g) MWth Total "Vulnerable" Tokaraaks D-T 2500 8000. 2000. Cat-DD 2500 100. a- 20. D-3He 2500 10. ~ 2. FRM's D-T -•). 2. 0.50 Cat-DD 2. 0.10 0.02 D-3He 3. * 0.005 0.001 building or the storage vault has a much lower potential for release to the environment. The alternate fuel fusion reactors have inventories estimated to be two to three orders of magnitude lower than comparable D-T fusion reactors. It is noted that one gram of tritium is 9600 Curies, which, represents a significant safety concern. Therefore, at least for tokamaks, many similar safety systems such as the emergency air detritiation system will be required. 3.3.A Fuel Supply The reference D-T reactor requires a source of tritium, i.e., the breeding blanket. Alternate fuel cycles do not require tritium breeding. In addition, deuterium is abundant and no resource problems are foreseen. Therefore, the Cat-D reactor does not present fuel availability concerns. However, the relatively "clean" D-3He reactor requires a source of 3He, and as noted previously, deuterium is abundant while 3He is not. Potentially available U.S. resources of 3He are estimated to be 200-500 kg/36-1 A 2500 MWth D-3He reactor would burn 330 g/day or 120 kg/yr. Since natural resources of 3He are not sufficient, the fuel must be supplied either from a D-D reactor or from decay of tritium. 98 As shown in Table 3-11, a 2500 MWth D-D reactor burning no 3He will produce 244 g 3He per day. This is not sufficient to supply one comparable D-3He reactor. Further, since some 3He will be burned in the Cat-D reactor, the support ratio will be less than ^0.5. Another potential source is tritium decay. To produce 120 kg/yr of 3He, a tritium inventory of 2150 kg is required. A storage facility for this amount of tritium is estimated to cost $200 million. To produce this amount of tritium from fission reactors, it is estimated to cost $2 x 1010. However, since 5.6% decays per year, the annual fuel cost if all the helium were supplied from tritium is $1.2 x 109. Another option may be excess breeding from D-T reactors. Assuming the excess tritium breeding ratio is 0.1, it would require about 170 D-T reactor-years to produce 2150 kg of tritium. It would require about 10 such D-T reactors to supply tritium to make up for the decay. It appears that utilization of "clean" D-3He requires the use of D-D reactors or possibly D-T or some combination of both. 3.3.5 Costs Capital costs of tritium and vacuum systems for the three fuel systems, assuming a 2500 MWth fusion reactor are shown in Table 3-13. The reduced tritium inventories for the alternate fuel cases result in significantly simplified fuel processing and, therefore, lower fuel processing costs. However, the total costs are significantly affected by vacuum pumping costs and as a result, total costs are not significantly different. Table 3-13. Estimated Tritium and Vacuum Costs ($M) for Various 2500 MWth Reactors Emergency Tritium Vacuum Fuel Air Recovery System Processing Detritiation Total Reference D-T 2.0 13.3 14.8 7.1 37.2 Cat-D (burn 1/2 of 3He 0.0 31.4 4.8 4.9 40.9 D-3He (lean D) 0.0 27.9 4.0 3.9 35.8 99 3.3.6 Conclusions Preliminary scoping studies of tritium handling, fuel processing, and vacuum pumping requirements for the alternate fuel systems for Cat-D and D-3He were completed and comparisons made to corresponding D-T reactors. Significant conclusions to date are: 1. Fuel processing for Cat-D and D-3He is generally similar to that for D-T and presents no major technological difficulties. 2. Tritium inventories are reduced by about two to three orders of magnitude. 3. 3He availability is severely limited, and it does not appear possible to have a fusion economy based on D-3He. The best source of 3He identified was D-D reactors and the support ratio only 0.5, or less. 4. Total tritium and vacuum costs for deuterium based alternate fuel fusion reactors do not appear to be significantly different from those for D-T reactors. 3.4 Magnet Design Considerations for Alternate Fuel Tokamak Reactor 3.4.1 Geometrical Limits on Plasma Current and Toroidal Field An alternate fuel tokamak reactor, optimized for some plasma or economic property, may require toroidal magnetic field (TF) and plasma currents far higher than those contemplated in design studies for D-T tokamak reactors. These larger fields and high currents may require that the reactor be very large. The high plasma current requires a large flux swing in the primary ohmic- heating (OH) coil system to initiate it. With superconducting OH coils, which could be limited to a field swing of + 8 I, the required area within the OH solenoid is roughly proportional to the desired plasma current. Similarly, a high toroidal field requires thick superconducting TF coils, with significant amounts of copper or aluminum for thermal stability and stainless steel or aluminum for support against the magnetic forces. As the central support sylinder and coil thickness increase, the radius of • curvature of the coil increases, which demands even more support and an even larger coil. 100 A quantitative estimate of these considerations has been undertaken, (37) using the magnet model developed at ANL. Figure 3-21 shows the combi- nations of plasma current and maximum toroidal field that can be obtained for different major radii of the tokamak. The reactor was assumed to have an aspect ratio of 3, a burn time of 30 minutes, and inboard and.outboard blanket and shield thicknesses of 0.7 m. The TF coil was assumed to operate at 4.2 K and to have copper stabilizer and stainless steel support material. The OH solenoid was assumed to have a field level of + 8 T and a current density of 1000 A/cm2. No particular superconductor was assumed. The range of fields shown in Fig. 3-21 indicate peak fields of the conductor of 12 to 20 T. This entire range is above the fields producible with NbTi and the upper half of it is above the field producible with Nb3Sn. The figure is intended to show che constraints on plasma currents and toroidal field, independent of the choice of superconductor. i i i i i i i i i i r i i i PEAK FIELD AND PLASMA CURRENT LIMITATIONS FOR VARIOUS MAJOR RADII I i i I i I i I 14 15 16 17 18 19 20 MAXIMUM TUROIDAL FIELD (T) Figure 3-21. Maximum plasma current and toroidal fields for alternate fuel tokamak reactors. 101 3.4.2 Irradiation Effects on NbgSn Superconductors A survey was made of the available measurements which have been performed to delineate the effects of neutron irradiation on the performance of Nb3Sn multifilament superconductor. The published data are obtained from experi- ments done at fission reactors and, although total fluences are in the range of 1019-1022 neutrons/m2 no significant flux with neutron energies above 5 MeV is obtainable from such reactors. Measurements of the critical current I exist up to 5 T, and only recent data include irradiations made at the ambient temperature of an actual operating TF magnet. If the low field data is extrapolated to the 10-12 T region required for TF coil design, then the measured relative degradation is: AI 22 D = —£• = (-4.4 + 1.2 x 10 *) B + 42.0 + 3.5 x 10~ 1022 < This degradation expression shows that at 10 T, influences to 1022 neutrons/m2 are unimportant, but at 1 x 1023 neutrons/m2 the degradation is 45%. In this study those conductors which showed enhanced performance from neutron damage were not considered as they generally showed poorly optimized critical current performance for high field application. Also, since annealing to 450°C for up to several hundred hours fails to reverse the neutron induced degradation of the critical properties of Nb3Sn, it was concluded that radia- tion damage must be considered permanent and cumulative. In addition to making preliminary conceptual design analyses for alternate fuel tokamak concepts, first-order safety assessments are being performed. This is being done so that potential safety hazards and environmental impacts may be identified as early as possible and either be avoided or minimized. Hazards to those persons who will be working in the fusion power plant, as well as to the general public, are being considered. 102 The presence of radioactive tritium, which has a half live of 12.33 years and emits a weak beta particle (E =18.6 keV), is one of the safety hazards under consideration. The design of the tritium recovery and pro- cessing systems for the plasma will be developed in order to minimize the total amount of tritium involved. The amount of tritium to be kept in storage will also be minimized. The total tritium inventory varies depending on the fuel combination involved. The reference D-T system would require 8000 g of tritium inventory, Cat-D, 100 g, and D-3He only 10 g. Of those totals the amount of tritium considered to be vulnerable is 2000 g for D-T, about 20 g for Cat-D and about 2 g for D-3He. Each gram of tritium correponds to about 9600 Ci. The vulnerable portions are that part of the tritium which could conceivably be released in the case of an accident (see Sec. 3.3). The induced radioactivities in the first wall and the structural materials in the blanket and shield are the other major sources of radioactivity presently being considered. Several candidate materials have been considered for all three fuel systems. The primary candidates are 316 stainless steel, Ti4381 and V-15Cr-5Ti alloys. The normalized activities (Ci/MWth) of these alloys have been calculated for preliminary alternate fuels design conditions both during reactor operation (from 1 second to 2 years) and following reactor shutdown (from time zero to 1000 years). See Sec. 3.2 for the details of those calculations. The total decay heat, the beta decay heat, and total biological hazard potential in air have also been calculated for a number of time steps over the same range of values as for the normalized activities. The atmospheric (x/Q) dispersion of tritium and of structural material was examined for a hypothetical accident based on a short-term uniform release from a point source. A simple model is utilized in order to compare the relative external dose from the candidate alloys dispersed in a Gaussian plume. The impact of tritium leaking from the reactor building following a hypothetical accident was also calculated using the same methodology. The relative external doses due to a given structural material in different alternate fuel systems being volatilized by a hypothetical reactor accident and its leaking out of containment are shown in Figs. 3-22 through 3-24. 103 APPROXIMATE LEVEL TOR ESTABLISHING EXCLUSION RADIUS (CORRESPONDS TO 20 rem/2 hr) I I I I I 1 I I 1 I I If I If 10 10° 10 DISTANCE FROM REACTOR, m Figure 3-22. Relative effect of Ti4381 structural material on different alternate fuel systems. 10 1 = ' 'II' II I 1 1 1 11 1' 1 APPROXIMATE LEVEL FOR I ESTABLISHING EXCLUSION f .6* =—^s: RADIUS (CORRESPONDS TO = SO rem/2 hr) - u- io7 '•HI M 1 i mil l ^SijsjCAT D id° - DT SO*5*^ -| id9 DHe3\^ I id10 x i iitiii IIIIII I \^ I IIIIII il i i i i | i il i i IIIIII 10 10 DISTANCE FROM REACTOR, m Figure 3-23. Relative effect of V-15Cr-5Ti structural material on different alternate fuel systems. 104 APPROXIMATE LEVEL TOR ESTABLISHING EXCLUSION RADIUS (CORRESPONDS TO -=j 20 rem/2 hr) i - J_L±JJ_LlJ I I_1_1_UJ1.! .l._.J...J ! 1-LU '0 10 10 DIST< NCE FROM REACTOR, m Figure 3-24. Relative effect of 316 stainless steel structural material on different alternate fuel systems. The same information could have been plotted for a given alternate fuel system to show the relative effect of different structural materials. However, be- cause of the nature of assumptions made regarding the design and power outputs, the method chosen to display the results is believed to be more appropriate. Based on the results, the Cat-D system produces somewhat higher total external doses following a hypothetical accident than does a D-T system; while the D-3He system produces much lower doses. The relative doses for D-T compared to D-3He are about 10 times as great for 316 stainless steel, 35 times for the vanadium alloy and about 60 times for the titanium alloy. The impact of the absolute value of these doses relates to the size of exclusion boundary that would be required around a fusion power plant. Based on current NRC guidelines of 20 rem whole body dose for a two-hour exposure following an accident to establish the exclusion radius, a corresponding value of approxi- mately 10~6 rem/s/MWth has been shown on the plots. The intersect of that line and each curve would represent the approximate exclusion radius. This value is used at the Construction Permit (CP) state, 25 rem is used later. 105 The relative effect of the vulnerable tritium inventory in the reactor building was also calculated for each of the alternate fuels systems. The normalized activities of tritium would be: 7.72 x 103 Ci/MWth for D-T, 77.2 Ci/MWth for Cat-D, and 7.72 Ci/MWth for D-3He. The activity of tritium when compared to the activity of structural material is generally quite small (less than one percent), except for the case of a D-T fueled system where the relative tritium-to-structural material activities are 2.7% for 316 stainless steel, 7.4% for Ti4381 and 17.5% for the vanadium alloy. The external whole body tritium doses from the plume at 200 meters from the reactor and at one hour following an accident would be 1.9 x 10~9 rem/s-MWth for D-T, 1.9 x 10"11 for Cat-D, and 1.9 x 10~12 for D-3He. The whole body doses due to tritium compared to the doses due to struc- tural materials are very small (on the order of 10~3 percent), except for D-T fueled systems where the ratio for tritium dose to vanadium dose is 0.135%. The biological hazard potential (BHP) of the accidently released tritium in air would be 39 km3 air/MWth and in water would be 2.6 x 10"3 km3 water/MWth for the D-T fueled system. The respective values for Cat-D would be less by a factor of 100 and the values for D-3Ke would be less than for the D-T sys- tem by a factor of 1000. The BHP of tritium in air is much less than that of the associated structural material. No comparison of the BHP of tritium in water to that of structural ma- terial was made because there is insufficient information to calculate those values for the structural materials. 106 References 1. "Proceedings of the Review Meeting on Advanced-Fuel Fusion," EPRI-ER-536-SR (September, 1977). 2. "Explanatory Studies of High-Efficiency Advanced-Fuel Fusion Reactors," EPRI-ER-581 (November, 1977). 3. "Explanatory Study of High-Efficiency Advanced-Fuel Fusion Reactors," EPRI-ER-919 (December, 1978). 4. G. J. Zeman and D. L. Smith, "Low Cycle Fatigue Behavior of Types 304 and 316 Stainless Steel Tested in Sodium At 550°C," Nuel. Technot., ^2, 82 (1979). 5. D. L. Smith, Argonne National Laboratory, Personal Communication (October, 1979). 6. J. W. Davis, McDonnell Douglas Astronautics Company, Personal Communica- tion (October, 1979). 7. A. E. Profio, Radiation Shielding and Dosimetry, A. Wiley - Interscience Publication, John Wiley and Sons, N. Y., pp. 414-418 (1979). 8. "ANISN-ORNL: Multigroup One-Dimensional Discrete Ordinates Transport Code With Anisotropic Scattering," RS1C/CCC-254, Oak Ridge National Laboratory (1973). 9. W. A. Roades and F. R. Mynatt, "The DOT III Two-Ditnensional Discrete Ordinates Transport Code," ORNL-TM-4280, Oak Ridge National Laboratory (1973). 10. D. M. Plaster, R. T. Santoro and W. D. Ford, III, "Coupled 100-Group Neutron and 21-Group Gamma-Ray Cross Sections for EPR Calculations," ORNL-TM-4872, Oak Ridge National Laboratory (1975). 11. D. Garken (compiler), "ENDF/B Summary Documentation," BNL-17541, Brookhaven National Laboratory (1975). 12. Y. Gohar and M. A. Abdou, "MACKLIB-IV: A Library of Nuclear Response Functions Generated with the MACK-IV Computer Program from ENDF/B-IV," ANL/FPP/TM-106, Argonne National Laboratory (1978). 13. J. Jung, "Theory and Use of the Radioactivity Code RACC," ANL/FPP/TM-122, Argonne National Laboratory (1979). 14. J. Jung, "Multigroup Neutron Activation Cross Section Library RACCXLIB and Decay Chain Data Library RACCDLIB for the Radioactivity Calculation Code RACC," Argonne National Laboratory (to be published). 15. See for example, D. J. Rose and M. Clark, Jr., Plasmas and Controlled Fusion, Chap. 11, The M.I.T. Press, Massachusettes Institute of Technology and John Wiley and Sons, Inc., N. Y. (1961). 16. D. G. Doran and N. J. Graves, "Displacement Cross Sections and PNA Spectra: Tables and Applications," HEDL-TMG-78-70, Hanford Engineering Development Laboratory (1976). 17. See for example, C. A. M. van der Klein, "The Organic Insulation in Fusion Reactor Magnet Systems," RCN-240, Reactor Centrum Nederland (1975). 107 References (cont'd.) 18. G. H. Miley, H. Towner and N. Ivich, "Fusion Cross Sections and Reactiv- ities," COO-2218-17, University of Illinois, C-U Campus (1974). 19. M. A. Abdou, "Radiation Considerations for Superconducting Fusion Magnets," ANL/FPP/TM-92, Argonne National Laboratory (1977). 20. M. A. Abdou, "Nuclear Design of the Blanket/Shield System for a Tokamak Experimental Power Reactor," Journal of Nual. Technol., 29_, 7 (1976). 21. T. Okada et al,, "Neutron Irradiation Effects on Superconducting Nb-Ti Alloys in the Magnets for Fusion Reactors," USERDA-CONF-750989, Vol. II, p. 436 (1976). 22. F. S. L. Hsn and J. E. Kunzler, "Magnetoresistance Probe for Measuring Magnet Field Intensity in a Small Space," The Review of Scientific Instruments, J34_, 297 (1963). 23. G. Kessler and G. L. Kulcinski, "Radioactive Inventories of Reactor Economics," Chapter 5 of Fusion and Fast Breeder Reactors by W. Hafele et. al., RR-77-8, International Institute for Applied Systems Analysis, Luxenburg, Austria (1976). 24. G. G. Eichholz, Environmental Aspects of Nuclear Power, Ann Arbor Science, Michigan, p. 159 (1977). 25. M. L. Tobias, "Inventory and Relative Biological Hazard of LMFBR Core Materials," ORNL/NUREG/TM-165, Oak Ridge National Laboratory (1978). 26. C. M. Lederer and V. S. Shirley (Editors), Table of Isotopes, 7th edition, John Wiley and Sons, Inc., N. Y. (1978). 27. USNRC, "Standards for Protection Against Radiation," USNRC Rules and Regulations, Title 10, Chapter I, Part 20 (1975). 28. M. A. Abdou, G. Fuller, E. R. Hager and W. F. Vogelsang, "Shielding and Maintainability in an Experimental Tokamak," Section 11-03 of Abstracts of 8th Symposium on Engineering Problems of Fusion Research. San Francisco, California (November, 1979). 29. See for example, J. Jung and M. A. Abdou, "Radiation Shielding of Major Penetrations in Tokamak Reactors," J. of Nucl, Teahnol., Q, 71 (1978). 30. W. R. Wilkes, "Distillation of He Isotopes," Proc. of ICEC-4, Eindhoven, 1972, IPC Science and Technology Press, p.p. 119-121 (1972). 31. W. R. Wilkes, 3He - **He Distillation Apparatus, Mound Laboratory, Report MLM-2005, p. 15,(March 21, 1973). 32. W. R. Wilkes, Mound Laboratory, Private Communication. 33. J. L. Anderson and R. H. Sherman, "Tritium Systems Test Assembly, Design for Major Device Fabrication Review," Los Alasmos Scientific Laboratory, Report LA-5855-P (June, 1977). 34. J. S. Watson and P. W. Fischer, "Cryosorption Vacuum Pumping Under Fusion Reactor Conditions," Proc. Seventh Int. Vacuum Congreas and Third Int. Conf. on Solid Surfacing. Vienna, Austria (September, 1977). 108 References (cont'd.) 35. D. 0. Coffin and C. R. Waltheus, "Vacuum Pumping of Tritium in Fusion Power Reactors," 8th IEEE Svmp. on Engineering Problems of Fusion Research, San Francisco, California and references therein (November 13-16, 1979). 36. T. Ben Rhinehammer and Layton J. Wittenberg, "An Evaluation of the Fuel Resources and Requirements for the Magnetic Fusion Energy Program," Mound Laboratory, Report MLM-2419, p. 55 (October 31, 1978). 37. L. R. Turner and M. A. Abdou, "Computational Model for Superconducting Toroidal Field Magnets for a Tokatnak Reactor," Seventh Svmp. of Engineer- ing Problems of Fusion Research, pp. 762-766, Knoxville (1977). 109 4.0 Summary and Conclusions This report presented the results of an initial study to assess some of the fusion reactor technology requirements of alternate or non-D-T fusion fuels. The basic approach of the study was to develop representative parameters for the distribution of reactor power into neutrons, charged particles and electro- magnetic radiation for two sample confinement concepts. These power split distributions were then used to examine some key reactor technology areas such as first wall design considerations, shielding requirements, environmental impact of activated materials, and fuel cycle design considerations. A variety of elements other than deuterium and tritium can undergo fusion; prominent examples include D-D, D-3He, 3He-3He, p-6Li, p-7Li, D-6Li, and p-nB. These fuels have, in varying degree, the generic features of reduced neutron production, increased fusion energy carried by charged parti- cles, and the elimination of a need for tritium breeding. On the other hand, the combination of lower cross sections, higher plasma temperature, and increased radiation losses make efficient confinement (i.e., high energy multiplication) more difficult. A key consideration then is whether or not the advantages are indeed sufficient to justify the development of alternate fuel power plants. A quantitative evaluation of the technology involved must be undertaken. Deuterium-based fuels have the advantage of operating at relatively low temperatures but involve more neutron and tritium production via "side" D-D reactions. Lower temperature deuterium based fuels, being easier to burn, are compatible with a wide range of confinement concepts; for example, tokamaks could burn catalyzed deuterium and D-3He. But, unfortunately, the attribute of "cleanliness" is not as fully achieved in the deuterium-based fuels as the more ideal proton-based fuels. The proton fuels, on the other hand, due to a combination of the high temperature required and modest energy released per reaction, appear much harder to successfully burn, forcing the use of more advanced confinement approaches. In order to study "representative" combinations that are illustrative of a wide variety of possible combinations, we selected for study a matrix of alternate fuels along with a modest and high (3 confinement system. Since 110 catalyzed-D is the only "self-sufficient" (only requires naturally available deuterium) low-temperature fuel, it was selected as the base fuel. However, due to its potential attractiveness, D-3He was also considered despite the complication of 3He breeding. Since tokamaks are presently the most highly developed confinement system, and since they (or at least modest 8 versions) can burn these low temperature fuels, they were selected as a representative modest 3 (10-20%) confinement system. (These 3 values are about a factor of two larger than predicted by most current theories, but are considered to be reasonable extrapolations for purpose of this study.) Concurrently, the Field-Reversed Mirror (FRM) was selected as a representative high-3 (> 70%) device. This is acceptable for the low-temperature fuels, but may not be adequate for the higher temperature fuels where even more innovative confine- ment approaches may be necessary. The tokamak sets a reference standard and provides a good basis for comparison of D-T and low-temperatures alternate fuel technology. As a part of this study, we evaluated the characteristics of alternate fuel tokamaks in light of confinement scali...^ laws based on recent experiments, e.g., empirical energy scaling based on PLT experiments and on Alcator data. For example, with 3 = 10%, ignited catalyzed deuterium and D-3He tokamaks are approximately one-fourth the volume for Alcator scaling as compared to trapped ion scaling. While these results are still uncertain due to uncertainties about the accuracy of nT scaling at the high temperatures involved (30-40 keV), they demonstrate a plausible alternate fuel tokamak approach. As an example, an ignited cat-D tokamak using Alcator scaling with 6 = 10% and B = 8.7 T would have a major radius of 12 m and a minor radius of 4 m; the thermal power level is approximately 7900 MW. With a modest improvement in confine- ment, ignited cat-D tokamaks as small as 1000 MW may be possible. The FRM has been chosen as representative of high-beta devices for alternate fuels. High-beta devices are especially appropriate for alternate fuels because high 3 operation can partially compensate for the smaller reactivities of alternate fuels. High-3 operation also decreases the cyclo- tron power emitted per unit volume, and it also leads to increased reabsorption of cyclotron radiation. Consequently, alternate fuel FRM's have for their fuel and temperature of operation relatively small power loss due to cyclo- tron radiation. For most energy loss mechanisms, the energy confinement time 111 generally increases as size of plasma increases. Thus, the plasma multipli- cation can be increased simply by increasing the size of the system. Un- fortunately, stability considerations limit the radius of the FEM to a few ion gyroradii. Therefore, the alternate fuel FRM's are not ignited reactors. However, despite their small size, the Q values of the alternate fuel FRM's are substantial due to the assumed near-classical diffusion in the closed field region. Thus, alternate fuel FRM's can be categorized as small (5 MW), efficient reactors with relatively large power densities and moderate Q. The power splits are affected by fuel cycle, plasma temperature, plasma density (3 and magnetic field strength), plasma size and plasma confinement scaling. Table 4-1 presents the power splits for the reference semi-catalyzed D, fully catalyzed D, and D-3He tokamaks. Note that the fraction of power in neutrons is less by almost an order of magnitude for D-3He than for semi- catalyzed D. The fraction of power in neutrons is less by about 10% for fully-catalyzed D than for semi-catalyzed D. This is, of course, because the fusion product 3He is burned in the fully-catalyzed D reactor. The relationships among the relative power fractions for the fuel cycles are not much different for the FRM than for the tokamak, as is also illus- trated in Table 4-1. Table 4-1. Power Splits for Tokamaks and FRMs Power Splits for Tokamaks Power Splits (%) Semi-cat-D Cat-D D-3He Leaking Particles .20 .26 .33 Bremsstrahlung .26 .25 .46 Cyclotron Radiation .11 .11 .16 Neutrons .43 .38 .05 Power Splits for FRMs Power Splits (%) Cat-D D-3He Leaking Particles .25 .24 Charged Fusion Products .19 .39 Bremsstrahlung .18 .26 Cyclotron Radiation .03 .08 Neutrons .35 .03 112 For purposes of technology considerations discussed in this report, the most important parameter is the fraction of total power carried by neutrons. This is relatively constant for the various cases, thus it is possible to examine certain first wall/blanket/shield problems without specifying a particular type of device. This is the general approach taken in this study when the fuel cycles are characterized in terms of neutron power fractions by the following representative values: D-T *> 80% D-D •>, 40% cat D-D -\. 40% D-3He % 1% (lean D mixture). Since electron cyclotron radiation poses a potentially serious loss mechanism for alternate fuel reactor concepts, this study examined the problem by employing two techniques: adopting neutronic computer codes (ANISN) to the problem of cyclotron radiation transport, and the use of ray-tracing techniques. It was found that previously used expressions by Trubnikov (ref. 13, Sec. 2) overestimates the power lost by factors of 3 -> 4. Ray- tracing techniques indicated total power losses somewhat lower (^ 75%) com- pared to ANISN results. While this study concentrated mainly on deuterium-based fuels, some attention was given to the p-6Li cycle. The evaluation of p-6Li as a viable alternate fuel requires a more rigorous analytical treatment. A multi—group Fokker-Planck treatment should be employed to incorporate the effects of nuclear elastic scattering on the fusion reactivity. More in- depth calculations should be done to determine neutron yields and radioactive ash production. Better and more complete cross-secticn data is needed to determine the contribution of numerous branch reactions to the overall fusion power. From this preliminary work, however, it appears that the p-6Li cycle has no significant advantages over the D-3He cycle in terms of neutron and tritium production levels. One of the major technological considerations of fusion reactors is the first wall facing the plasma and the blanket region. Major differences between the alternate fusion systems and the D-T system that affect materials selection and first-wall and blanket design considerations are, no tritium breeding requirement, differences in neutron energy and flux, higher particle 113 transport and radiation heat loads on the first wall, and possible higher magnetic fields. Elimination of the tritium-breeding requirement reduces the number of materials compatibility considerations in the blanket, and elimi- nates the whole tritium-breeding system including the need for lithium in the blanket and the related tritium transport problems. The greatest impact is probably on the coolant selection since coolant/breeder compatibility is no longer important. Important considerations for the first wall include allowable surface heat fluxes, plasma/wall interactions, reflection of cyclotron radiation and neutron damage. For a typical case of a stainless steel wall operating under cyclic conditions, surface heat loading limitations reduce the total wall loading by a factor of three for D-D and a factor of five for D-3He compared to D-T. Surface roughing will probably not significantly reduce reflection of cyclotron radiation; however, there is very little experimental data. On the other hand, the use of low-Z, non-conducting surfaces (e.g., oxide coatings) may not be acceptable to insure proper cyclotron radiation re- flection, thus, perhaps, making impurity control more difficult for alternate fuels. It is important to note that radiation damage parameters (atomic displacements and gas production rates) are reduced by about two orders of magnitude for D-3He compared to D-T for the same total wall loading. D-D offers about a factor of two reduction compared to D-T. The diffeiance in neutron energy spectrum has a minor impact compared to the vastly different neutron fluxes for these fuel cycles. The nuclear analysis of shielding requirements examined balk, penetration and building shielding requirements, as well as some environmental consider- ations regarding neutron activation of reactor components. It is clear that maintenance of reactor internal components (e.g., first wall, blankets, etc.) for any deuterium-based fuel cycle will have to be by remote methods. Bulk shielding requirements (e.g., thicknesses) to protect superconducting magnets are reduced by about a factor of two for D-3He compared to D-T. For D-D compared to D-T for the same total wall loading, 114 the shield thickness is reduced by about 30%. Typical thicknesses are 70 cm for D-3He, one meter for D-D, and 1.4 tn for D-T with a tritium breeding blanket. One finds similar differences between the various fuel cycles for penetration shielding requirements. With these bulk reactor shielding thicknesses, there is little difference between the various fuel cycles in the required biological shielding in the reactor building. Typical thicknesses are about 2.meters of concrete material. It is likely, then, that for the same size building, one can expect little difference in the cost of the building for the various deuterium-based fuel cycles. For a case where 316 stainless steel is used as the blanket structural material, the long-term BHP for activated material for D-D is similar to D-T. After about 100 years, the difference between D-3He and D-T is small, i.e., less than a factor of ten. For other materials, such as vanadium and titanium, BHP is significantly lower for all fuels compared to stainless steel. Again, D-D is similar to D-T for the same total wall loadings. The major conclusions resulting from an analysis of fuel cycle considera- tions are that the tritium throughput for D-D is about a factor of 15 less than for D-T for the same total power. For cat-D, the reduction is 100 over D-T. For D-3He, the reduction is about 1000. However, there is little difference in the total cost of the fuel handling/vacuum system between D-T, D-D and D-3He. The only feasible source of 3He seems to be a D-D reactor; support ratios of D-3He to D-D will be at least 1:2. Finally, it is noted that fuel processing for cat-D and D-3He is generally similar to that for D-T and presents no major technological difficulties. With inspect to general safety considerations, it was found that D-3He reactors may result in lower radiation doses (10-100X lower) under possible reactor accident conditions compared to D-T, but D-D reactors may produce higher total external doses compared to D-T. The whole body doses due to tritium compared to the doses due to structural materials are small. 115 Distribution for ANL/FPP/TM-128 Internal: M. Abdou P. Finn J. Roberts R. Arnold Y. Gohar H. Schreyer C. Baker J. Jung D. L. Smith C. Boley W. Kann R. P. Smith A. Bolon S-H. Kim H. Stevens J. Brooks R. Kustom L. Turner R. Burke V. Maroni S-T. Wang F. Cafasso R. Martin R. Weeks R. Clemmer R. Mattas FP Program (20) J. de Paz B. Misra ANL Contract File D. Ehst J. Norem ANL Libraries (5) K. Evans E. G. Pewitt TIS Files (6) External: DOE-TIC, for distribution per UC-20 (117) Manager, Chicago Operations and Regional Office, DOE Chief, Office of Patent Counsel, D0E-C0R0 President, Argonne Universities Association Applied Physics Division Review Committee: P. W. Dickson, Jr., Westinghouse Electric Corp. R. L. Hellens, Combustion Engineering, Inc. K. D. Lathrop, Los Alamos Scientific Lab. ' B. Loewenstein, Electric Power Research Inst. R. ". Redmond, Ohio State U. R. Btier, Stanford University D. B. Wehmeyer, Detroit Edison Co. N. Amherd, Electric Power Research Institute, Palo Alto, California J. Anthony, General Electric Company, Schnectady, New York R. E. Aronstein, Bechtel National, Inc., San Francisco, California S. Baron, Burns & Roe, Oradell, New Jersey J. Baublitz, OFE/DOE, Washington, D.C. L. A. Berry, Oak Ridge National Laboratory, Oak Ridge, Tennessee Bibliotheque, Seine, France Bibliothek, Garching, West Germany R. Blanken, OFE/DOE, Washington, D.C. R. Botwin, Grumman Aerospace Corporation, Bethpage, New York S. Burnett, General Atomic Company, San Diego, California G. Carlson, Lawrence Livermore Laboratory, Livermore, California G. Casini, Commission of the European Community, Ispra, Italy C.E.A. Fusion, Cedex, France F. F. Chen, University of California, Los Angeles, California J. F. Clarke, OFE/DOE, Washington, D.C. D. Cohn, Massachusetts Institute of Technology, Cambridge, Massachusetts R. Conn, University of California, Los Angeles, California J. Crocker, EG&G Idaho, Inc., Idaho Falls, Idaho E. Dalder, OFE/DOE, Washington, D.C. R. Davidson, Massachusetts Institute of Technology, Cambridge, Massachusetts J. Davis, McDonnell Douglas Astronautics Company, St. Louis, Missouri S. Dean, Science Applications, Inc., McLean, Virginia J. Decker, OFE/DOE, Washington, D.C. 116 D. DeFreece, McDonnell Douglas Astronautics Company, St. Louis, Missouri D. Dingee, Battelle Pacific Northwest Laboratory, Richland, Washington H. Dreicer, Los Alamos Scientific Laboratory, Los Alamos, New Mexico H. Forsen, Exxon Nuclear Company, Bellevue, Washington T. K. Fowler, Lawrence Livermore Laboratory, Livermore, California H. Furth, Plasma Physics Laboratory, Princeton, New Jersey J. Hancox, Culham Laboratory, England C. Head, OFE/DOE, Washington, D.C. R. L. Hirsch, Exxon Research & Engineering Company, Florham Park, New Jersey J. Holmes, Hanford Engineering Development Laboratory, Richland, Washington D. Jassby, Plasma Physics Laboratory, Princeton, New Jersey T. Kammash, University of Michigan, Ann Arbor, Michigan A. Knobloch, Max-Planck Institut fur Plasmaphysik, West Germany J. Kokoszenski, The Ralph M. Parsons Company, Pasadena, California R. A. Krakowski, Los Alamos Scientific Laboratory, Los Alamos, New Mexico D. Kummer, McDonnell Douglas Astronautics Company, St. Louis, Missouri T. Latham, United Technologies Research Center, East Hartford, Connecticut C.-S. Li, Chinese Academy of Sciences, Peking, China Librarian, Culham Laboratory, England Library, Centre de Recherches en Physique des Plasma, Switzerland Library, FOM - Inst. voor Plasma-Fysica, The Netherlands L. M. Lidsky, Massachusetts Institute of Technology, Cambridge, Massachusetts F. J. Loeffler, Purdue University, Lafayette, Indiana D. G. McAlees, Exxon Nuclear Company, Bellevue, Washington G. Miley, University of Illinois, Urbana, Illinois R. Moir, Lawrence Livermore Laboratory, Livermore, California D. B. Montgomery, Massachusetts Institute of Technology, Cambridge, Massachusetts 0. B. Morgan, Oak Ridge National Laboratory, Oak Ridge, Tennessee K. Moses, TRW, Redondo Beach, California M. Murphy, OFE/DOE, Washington, D.C. T. Ohkawa, General Atomic Company, San Diego, California D. J. Paul, Electric Power Research Institute, Palo Alto, California F. Powell, Brookhaven National Laboratory, Upton, Long Island, New York P. Reardon, Plasma Physics Laboratory, Princeton, New Jersey T. Reuther, OFE/DOE, Washington, D.C. F. Ribe, University of Washington, Seattle, Washington M. Roberts, OFE/DOE, Washington, D.C. A. Robson, Naval Research Laboratory, Washington, D.C. P. Rose, Mathematical Sciences, North West, Bellevue, Washington M. N. Rosenbluth, Institute for Advanced Study, Princeton, New Jersey R. Rutherford, Plasma Physics Laboratory, Princeton, New Jersey J. L. Scott, Oak Ridge National Laboratory, Oak Ridge, Tennessee R. Scott, Electric Power Research Institute, Palo Alto, California Z. Shapiro, Westinghouse Electric Corporation, Pittsburgh, Pennsylvania W. M. Stacey, Jr., Georgia Institute of Technology, Atlanta, Georgia R. Staten, OFE/DOE, Washington, D.C. D. Steiner, Oak Ridge National Laboratory, Oak Ridge, Tennessee F. H. Tenney, Plasma Physics Laboratory, Princeton, New Jersey Thermonuclear Library, Japan Atomic Energy Research Institute, Japan C. Trachsel, McDonnell Douglas Astronautics Company, St. Louis, Missouri A. Trievelpiece, Science Applications, Inc., La Jolla, California K. Zwilsky, OFE/DOE, Washington, D.C. 117