Nuclear Engineering and Design Volume 39 issue 1 1976.pdf

Nuclear Engineering and Design 39 (1976) 215-238 © North-ttolland Publishing Company

ON THE SAFETY OF TOKAMAK-TYPE, CENTRAL STATION FUSION POWER REACTORS

D. OKRENT, W.E. KASTENBERG, T.E. BOTTS, C.K. CHAN, W.L. FERRELL, T.H.K. FREDERKING, M.J. SEHNERT and A.Z. ULLMAN. Department of Energy and Kinetics, School of Engineering and Applied Science, University of Cali]ornia, Los Angeles, California 90024, USA

Received 20 April 1976

A preliminary examination is made of several potential safety questions for Tokomak-type, central station fusion power reactors, using UWMAK-I as a reference concept. Larger tritium inventories than previously reported are estimated. The divertor loss-of-flow accident appears to require fast plasma shutdown. The large helium inventory in the cryogenic magnets affords a mechanism for generating large pressures within the containment. Lithium-air and lithium-concrete reactions have the potential for generating large amounts of heat, significant pressures, and active chemical species capable of volatilizing structural material such as a radioactive first wall.

1. Introduction

A considerable body of experience and expertise now exists in the design, construction, operation and safety analysis of fission power reactors. On the other hand, only in the last few years has fusion reactor design gone from preliminary concepts [1-3] and preliminary designs [4 -6 ] to conceptual designs o f central station fusion power reactors [7 -10] . In addition, only elementary safety analysis and limited environmental studies [11-17] have been carried out on fusion reactors. Comparative examination of fusion and fission reactor safety have thus far usually dealt with gross considerations such as total radioactive inventory, long- lived activity present in waste products, and hypothet- ical whole core disruptive accidents [18-22] .

Postulation of specific accident chains has generally not been done for fusion reactors, nor have the possible accident sequences been examined using failure modes and effects analysis or event trees.

This paper contains a summary of preliminary safety studies focusing on Tokamak-type fusion power reactors. Although an attempt was made to keep the study as generic as possible, the UWMAK-I design [8] was used as a basis because it was reasonably well docu- mented at the time of initiation of the study.

A basic objective of these studies was to identify potential safety questions at an early stage o f fusion reactor development so that the appropriate steps can be taken in research or design. No estimates have been made of the probabilities of the postulated events studied herein. In the next section the scope of the study as well as a summary of findings is presented. This is followed in turn by a brief description o f portions of UWMAK-I, a delineation of event trees, a treat. ment of tritium inventory questions, an examination of loss-of-flow accidents, a discussion of some magnet- related safety considerations, a brief study of the loss- of-helium accident, and an examination of l i thium-air and l i thium-concrete reactions.

2. Scope of study and summary

The scope of this study can be embodied in the following questions: (1) Are there materials present in a fusion reactor

which are potentially hazardous to the health and safety of the public?

(2) Are there mechanisms for making these materials mobile (or volatilizing them)?

(3) Are there mechanisms for breaching containment and releasing these materials?

215

216 D. Okrent et al. / Safety o f Tokamak fusion power reactors

Table 1 ttazardous material

Radiological Non-radiological

Radioactive structure beryllium material boron

chromium Radioactive copper corrosion products lead

lithium mercury

Tritium nickel

Prelir dnary results of an at tempt to identify potential sore ~'es of hazard to the public health and safety resultc , first in the list of materials given in table 1.

The non-radiological hazards appear to require at least preliminary consideration because of the very large quantities in use or in inventory at the fusion power reactor site.

In an at tempt to answer the second and third questions, three approaches were considered. First, although the designs of CTRs are still in a state of flux, an attempt was made to identify various accidents of potential generic concern to Tokamaks. This a t tempt pro- duced the list shown in table 2.

Secondly, it is possible to classify accidents in various categories, such as the following (some of which

Table 2 Some potential accidents of probable generic concern to Tokamaks

A. Plasma related 1. overpower transients 2. confinement instabilities 3. runaway electrons 4. power oscillations 5. localized reactions

B. Blanket/coolant/heat-removal-related 1. loss-of-flow 2. afterheat removal 3. loss-of-coolant 4. lithium air fires and reaction with concrete 5. steam generator or secondary system failure

C. Magnet-related 1. loss of superconductivity 2. magnetic forces 3. arcing 4. rupture of helium system

D. Other - tritium fire or escape

are overlapping): (1,) site-related (earthquake, flood, tornado); (2) missile generation (external and internal); (3) component failure; (4) chemical reaction; (5) over-pressurization; (6) over-heating; and (7) loss of power.

Thirdly (and the approach taken here), it is possible to postulate various initiating events, construct event trees which describe the sequence following the initiating event, and then examine those branches which may significantly alter a tree and/or lead to a breach of containment or release of hazardous material (or both).

Event trees have been constructed for the following postulated accident sequences: (1) loss-of-flow of blanket lithium; (2) lithium spill within containment; (3) plasma density - temperature transient; (4) superconducting magnet quench; (5) loss-of-flow to one or more divertor pumps; and (6) loss of blanket coolant.

As a result of the approach outlined above, the major findings can be summarized as follows. The potential hazard of tr i t ium is related, in part, to the total inventory of tr i t ium which resides in the blanket, in the fuel system, and in the heat transfer and energy con- version systems. In addition, an amount of tritium must be maintained in storage for continuous operation of the plant. As part of this study, tritium inventory considerations were examined. It appears that previous estimates of plant inventories may be too small because hold-up in fuel reprocessing systems has been ignored and tri t ium reserves (in the event of one or more of the plant 's reprocessing systems becoming inactive) have been neglected. Depending on the reprocessing system used and the amount required in storage, it appears that approximately 5 0 - 1 0 0 kg of t r i t ium may be in inventory at the site. This corresponds to about 109 Ci.

The radioactivity of the first wall material of a large scale, magnetic confinemeDt fusion reactor may be quite large; it is of the order of 109 Ci for the UWMAK-

I design using a stainless steel first wall. The possibility

of volatilizing a significant port ion of the wall, and

hence of the radioactive materials in that port ion of the wall, is therefore of considerable interest in the as-

D. Okrent et al. / Safety o f Tokamak fusion power reactors 217

surance of public health and safety. Lithium-air or lithium-concrete reactions seem to

represent possible means of large scale volatilization of the first wall materials in a major accident, should such reactions occur on a large scale despite the liners on the concrete and the evacuated containment. In both cases, flame temperatures may be reached which can melt some potential first wall materials, such as stainless steel or vanadium, but not materials such as niobium, molybdenum, and TZM. Furthermore, significant pressures of highly reactive chemical species, such as O, H and N, are formed. The air fire can give rise to oxygen atom pressures of about 100 Pa (1 Pa -~ 10 .5 atm) and the concrete reaction about 10 Pa. The reaction probability of atomic oxygen with many of the refractory metals at high temperatures and low pressures is of the order of 0 .1-1 , so the ablation rate of first wall materials could be equivalent to a vapor pressure of 10-100 Pa for the air reaction, and about 1 -10 Pa for the concrete reaction. By comparison, the vapor pressures of molybdenum and niobium reach 10 Pa at 3000 and 3200 K, respectively, considerably above their respective melting points. A first wall vaporization hazard, then, for a lithium-air or lithium-concrete fire may arise both from the high temperatures reached and the high chemical reactivity of the resulting chemical species. An effective vaporization flux of 10 Pa for molybdenum or niobium at 2400 K could produce a recession rate of about 0.3 cm/sec of the metal wall material. Since the wall material is typically less than 1 cm thick, vaporization times for the reaction with flame materials would be only a few seconds under the most reactive conditions postulated. In such an event, essentially all of the radioactive materials in the first wall could be volatilized due to the reactivity and high temperature of the flame. Lithium reactions were therefore examined during these studies.

The radioactivity present in corrosion products is expected to be highly system dependent; it was not examined in this study.

Principle threats to containment integrity include: (1) gross failure of sections of the magnet, under stan-

dard operating conditions, or more particularly under conditions of rapidly changing field;

(2) gross rupture of the helium cryogenic system and a release of helium into the primary containment;

(3) lithium-concrete reactions; (4) lithium-air fires; and

(5) missile generation due to mechanical failure of vacuum system.

3. A brief description of UWMAK-I

UWMAK-1 is the first of a series of fusion power reactor conceptual designs based on the Tokamak confinement concept, studied by the University of Wisconsin. Thermal power from this reactor is 5000 MW and electrical power output is 1500 MW(e). An effort was made to use existing technologies and industries whenever possible.

The D - T plasma radius is 5 m and the torus radius is 13 m, giving characteristic reactor dimensions of 25 m radius and 40 m height (fig. 1).

The reactor duty cycle consists of a steady-state burn for 5400 sec followed by 600 sec for recycling and start-up. Neutral beam injection is used to maintain the plasma during the burn.

Surrounding the plasma is a 316 stainless steel blanket through which lithium flows as the heat removal and tritium breeding material. Corrosion of the stainless steel limits the peak lithium temperature to 500°C. Behind the lithium is a stainless steel neutron moderator and reflector, followed by a helium cooled magnet shield and the magnets.

Superconducting solenoids are used to maintain the magnetic fields necessarily for plasma confinement. There are 12 D-shaped toroidal field coils, 12 transformer coils, and 8 vertical field coils. These coils are kept below their transition temperature with 250 000 1 of liquid helium. All of the solenoids are designed to be totally cryostatically stable.

After the lithium loop, heat extraction is supplied by a secondary liquid sodium loop leading to a conven- tional steam cycle. The path through the coolant loops and heat exchangers is responsible for most of the steady state tritium release, 10 Ci/day.

The plasma operates with a divertor utilizing a double null point. Particle extraction occurs both above and below the plasma where liquid lithium, flowing down steel plates, traps and removes charged particles.

Calculations to determine inventories of radioactive corrosion products, radioactive structural material, and tritium were carried out. Tritium and corrosion product extraction systems are included in both the lithium and sodium loops.

218 D. Okrent et aL / Safety o f Tokamak fusion power reactors

TRANSFORMER COILS

>|VE RTOR COILS

SHIELD

BLANKET

TOROIOAL COILS

i i ii i iili iii!iii!i!~!i

Fig. 1. UMAK-I, a conceptual fusion power reactor design from the University of Wisconsin.

4. Event trees

The initiating events listed below were selected because the potential release of biologically hazardous material or the ability to tap large sinks of energy which might violate containment integrity were involved. Only the reactor system was considered herein as a source of initiating events. Other parts of the plant, as well as natural occurrences such as earthquakes, remain to be pursued in later studies.

Large amounts of energy can be related to the superconducting solenoids, molten lithium, and the plasma. Events with the potential for tapping these sources seem to include magnet and plasma quenches, thermal excursions (including the liquid helium system), lithium spills, and a divertor pump loss-of-flow accident (LOFA).

Event trees have been constructed for the following postulated accident sequences: (1) loss of flow of blanket lithium; (2) lithium spill within the primary containment; (3) plasma density - temperature instability transient; (4) superconducting magnet quench; (5) loss of flow to one or more divertor pumps; and (6) loss of blanket coolant.

BLA' IKET I LOF ~ IN 1 SE ;TION

I

WALL REMAINS

AFTEI ~IEAT FAILS

Fig. 2. Event tree for a blanket loss of flow.


I BLANKET LOFA I & BREACH OF I PLASMA CHAMBER I

I CONTAINMENT HOLDS THE REACTANTS

I BLANKET LOCA I & RELEASE OF Li IN P.C.

I ? I

MOLTENL~ I S ,LLEO,N I (UNPROTECTED) J

I DIVERTOR LOFA I CAUSING Li RELEASE

J I Li-CONCRETE NO REACTION | REACTION DUE O 8 BREACHING I TO BREACHED OF CLADDING I CLADDING

TOO SLOW CONTAINMENT THROUGH TO OVER- OVERPRESSURE CONCRETE COME HEAT FAILURE CONTAINMENT LOSSES

~ . ~ ~ M I S S I L E S TRITIUM ~

Fig. 3. Event tree for a lithium spill within the primary containment.

Event trees characterizing these events are given in figs. 2 - 7 . The trees are to be read time sequentially from top to bo t tom and are presented here in outline rather than detailed form, except for that dealing with magnet quench. A detailed description of these event trees can be found in ref. [34].

5. Tritium inventory

5.1. Mathematical model

The determination of the tr i t ium inventory is necessary for analysis of the consequences of a plant ac-

220 D. Okrent et al. / Safety of Tokamak fusion power reactors

FUEL FEED I RATE FLUCTU- ATiO N

J FUEL MIXTURE FLUCTU- ,ATION

I

d/ex Zex

I Io,- i . . . . N (CONTAIN- d t X ~- ' MENT TIME)

where

(2)

Iex = I 0 , a t t = 0 , (3)

and N - is the burn-up rate of tritium. In the above equation we can ignore r if the condi-

tion exists that

TEMPER- ATURE

T > Tmi n = (I/7 N-) + 1 , (4)

where T = N+/N - is the breeding ratio and

I=tb +rex. (5)

The solutions to these equations are

I b = N+X(1 - exp(- t /X)] , (4)

and

Fig. 4. Event tree for a plasma density-temperature instability transient.

Iex = N - [ ( T - 1)t - TX(1 - exp(-t/X))] + I 0 . (5)

The minimum external operating inventory for this model can be shown to be given by

1+,0 cident and the potential release during steady-state operation. This inventory is highly system dependent; however, models have been developed which attempt to quantify the total tritium inventory based on production rate, burn-up rate, and the mean residence times of the tritium in various systems.

One such model [24] divides the tritium supply into two distinct inventories, the blanket inventory I b and the external inventory lex consisting o f the tritium removal system, fuel injection system, fuel in the plasma, and operating reserve and the fuel accumulated for use in other reactors.

For the blanket, the rate equation is

dlb/dt = N + - (I/X + l / r ) I b , (1)

where A~ is the rate of tritium production, X is the mean residence time of tritium in the blanket, and r is the mean-life for radioactive decay of tritium.

For the external inventory,

Then we may define

A l = I o - I e x m i n = N - X [ 1 - ( T - 1 ) l n ( T T - ~ l ) l , (7)

where A1 is a start-up reserve needed to maintain operation of the plant until the bred fuel becomes available.

As an example, i f I b sat--- 12.5 kg, N - = 0.7 kg/day, and T--- 1.3, then X ~ 14 days and AI = 5.5 kg. As- suming Iex min = 10 kg, the start-up inventory will be

I 0 = AI + Iex min = 15.5 kg .

This model is lacking, however, in its determination of the total tritium inventory in that it does not include a fuel processing system. This sytem can affect the total plant inventory in that a substantial amount of fuel must flow through this system, and the reten- tion time may be significant.

The previous model can be modified to include a processing system as follows [23].


I A SUPER- ] CONDUCTING MAGNET QUENCHES

iNoo is0L

EDDY [ CURRENTS FIELD ENERGY HEAT LHa TRANSFERRED TRANSFER TO U AND LINES STRUCTURE

L~ ] L~ FLOW LO~S-OF- MAINTAINED FLOW

NO EFFECT

DEPOSITED FIELD ENERGY OR NEW FIELD CONFIGUR- ATION DISRUPTS PLASMA

I FIELD ENERGY TRANSFERRED TO OTHER MAGNETS

[ I-- OTHER NEARBY SOLENOIDS SOLENOIDS ARE CONTINUE DRIVEN OPERATION NORMAL

~FE RUPTURE

He OVERFRESSURE IN DEWAR

J DISRUPTION DAMAGE TO OF PLASMA BLANKET, VACUUM SHIELD, OR DEWARS

I BOIL-OFF SYSTEM CONTAINS Ho VAPOR

io- j RUPTURES OR EXPLODES

Fig. 5. Event tree for a superconducting magnet quench.


I SYSTEM CON- [ TINUES SAFE OPERATION

LOSS OF FLOW OF Li TO ONE OR MORE DIVERTOF PANELS +

I CHARGED PARTICLES I HEAT Li STILL ON PANEL I SYSTEM SAFELY

SHUTS DOWN

I ALL Li RUNS I OFF THE PANEL BEFORE ABLATING (RAPIDLY)

I SOME Li IS [ VAPORIZED BY CONTINUING HEAT LOAD

IS Li ABLATION RAPID AND SELF SPUTTERING CONSIDERABLE?

FAILED PANEL I AFFORDS PATH TO PRIMARY CONTAINMENT

POSSIBLE PATH J TO Li-CONCRETE REACTION

PLASMA O N , I PANEL MELTS

I NO LOSS OF I VACUUM

I

LOW VACUUM ~'~ CHAMBER PRES- / SURE MAINTAINED I

l VACUUM CHAMBER I PRESSURE RISES [

LiPRESSURE ] IN PLASMA CHAMBER GREATER THAN BACKGROUND

-PATH IS DESIGN DEPENDENT

ICOO LANT I BEH,ND PANEL t ENTERS PL*SMA I CHAMBER ]

L ' 'ON"Es ' t lF ,RSTWALLI [ L''NTERACTS , ESNTCROSS, IMELTSLO ALLYI W,THFL MA I sEPERATR'x I ,N,T,AT,NG INSTABILITIES

Li IN PRIM. ] CONTAINMENT IF VACUUM PUMP FAILS

Fig. 6. Event tree for a divertor loss of flow.


Li El ITERS PLA~

THE PLASMA PRIM. CONT. CHAMBER THROUGH VAC,

PUMP OR PIPE

I

() I

PRIMARY CONTAINMENT

I i I WITH CONCRETE REACTION: TAKES PLACE P. T, % PRODS.

CONTAINED CONTAINMENT LOCALLY FAILS

I

I I L'ENTEBS I SECONDARY

CONTAINMENT

i.o=.,o. I I WITH CONCRETE REACTION: TAKES PLACE P, T, & PRODS.

IS CONTAINMENT CONTAINED FALLS

Fig. 7. Event tree for a blanket loss of coolant.

For the blanket, as before

dIb /d t = N + - (/b/X) • (8)

For processing system:

d/p Ib Ip = - - - + a N - - - - (9)

dt X tp '

where Ip is the processing system tritium inventory, tp is the mean residence time of tritium in the processing system, and a N - is the amount of unburned fuel from the burn-cycle.

Finally, for the fueling and storage system

d/fs _ lp ( a N - + N - ) (10) dt tp

where Ifs is the fuel injection and storage system in-

ventory. The quantity ( N - + a N - ) is the amount of fuel injected into the burn-cycle.

Solving these three coupled equations yields

I b = N+X[ 1 - exp ( - t /X) ] , (11)

--N+tPX~ [exp(X - t / tp ) - exp(- t /X)] Ip (X -- t p )

+ (N + + a N - ) [1 - e x p ( - t / t p ) ] t p , (12)

N+X I f s - (X - tp)[X exp(--t/X) - tp e x p ( - t / t p ) ]

+ ( N + + a N - ) tp e x p ( - t / t p ) + I 0

+ ( N + - N - ) t - N+(X + tp) - a N - t p . (13)

224 D. Okrent et al. / Safeo' o f Tokamak fusion power reactors

w z

4O

w Z z_o

O m I-- L Z O

6O

N" = .69kg /DAY N + = 1.031e/DAY %OF FUEL BURNED = 5 X IS IN DAYS rio IS IN DAYS X = 1 4 . 4 ~ ~l IS m kg x = m . 0 ~ ' ~ , ~

50 x = 6.4 ~ 1 = 7 ~ x = ! 4 5

10

0 . . . . . . , , I J , , ~ . . . . . . I , , L , L L , , , I . . . . . . . . . I

0 1 2 3 4 MEAN PROCESSING TIME l ip DAYS)

Fig. 8. Start-up inventory versus mean processing residence time for various blanket mean residence times.

120

110

IO0

8 90

. o

1 I-- 70

;r- , :~ so O w

m O > :E 40 = ~ z'3 .-,i z

20

10

0

MAXIMUM BLANKET INVENTORY = 12.5 k9 T -- BURN-UP RATE = .7 k g / D A Y

BREEDING RATIO = 1.3

Y//J

j l l l l , , , I . . . . . . . . . I , , , j . . . . . I . . . . . . . . . I 1 2 3 4 MEAN PROCESSING TIME (tp DAYS)

Fig. 9. Start-up inventory versus mean processing time for various plant breeding ratios.

The time to minimum total operating inventory can be found by minimizing Its, since all tr i t ium ulti- mately enters or leaves here and the other inventories build to a saturation level:

dt min 0 = [exp(- tmm/tp) - exp(-tmin/X)]

- ( N + + a N - ) e x p ( - t m i n / t p ) + (N* - N ) . (14)

Solving this will define tmin, i.e. I 0 is set such that at train,/fs is at some minimum operating level:

N+X I0 = Its min . - - - [ X exp(-tmin/X ) A - tp

- tp e x p ( - t m i n / t p ) ] - ( N + - a N - ) t p e x p ( - t m i n / t p )

- (N + - N - ) t m i n + N+()t + tp) + cffV--tp. (15)

If we take the previous example, assuming 5% burn-

up and tp = 1 day we find AI = 19.2 kg instead of only 5.5 kg.

The results of sensitivity studies for AI are given in figs. 8--10 in which AI is given as a function of processing time tp for various blanket residence times, breeding ratios and percent burn-up, respectively.

5. 2. B l a n k e t m e a n residence t imes f o r U W M A K - I

In the UWMAK-I reference reactor [8], trit ium is burned at a rate of 0.689 kg/day and the breeding ratio is 1.5 with a tri t ium concentration build-up of 1.2 wppm/day. Tritium is removed from the lithium blanket via an yt t r ium extraction bed which is located at the intermediate heat exchanger exit. The tri t ium partial pressure corresponding to the exit temperature of 283°C is estimated to be 1.48 × 10 -11 torr by ex- trapolation from an equilibrium Y - Y H 2 relation. Using the Seiverts relation with K s = 0.346 torrl/2/a.f. the concentration of LiH in lithium is 1.1 × 10 -5 mole

D. Okrent et al. / Safety of Tokamak fusion power reactors 225

60

50

MAXIMUM BLANKET INVENTORY " 12.5 kg T 5% FUEL BURN-UP BURN.UP RATE = .7 kg/DAY

,9,

0 40 :E <

I--

, o

w ~

f- l ,u

=>_= 20 Q.:E

<[=c

10

BR = 1.1 BR = 1.3--~,

BR " 1.5 BR - 1.7

0 ~ 0 1 2 3 4

MEAN PROCESSING TIME (tp DAYS)

Fig. 10. Start-up inventory versus mean processing residence time for various fuel burn-up percentages.

fraction. This represents a total tritium blanket inventory I b = 8.7 kg.

If there are another 2 kg of tritium in the beds, the mean residence time of tritium in the blanket is about

/total X- - 10.3 days.

T N -

5.3. Fuel e f f luent reproeessing mean residence t ime fo r UWMAK-I

No purification and isotope separation system is presented in the UWMAK-I design, therefore a preliminary conceptual system was formulated using a cryogenic fractional distillation process.

A reasonable mean residence time of tritium in the distillation system is estimated to be 1.5-2.5 h.

In the UWMAK-I design an additional time must be included to allow for the lithium divertor system, since

the majority of the reactor effluent is removed via the divertor lithium circuits using yttrium beds at 200°C. The amount of tritium contained in the beds is estimated to be 3.5-4.0 kg.

Assuming a burn-up of 10% and a trapping efficiency of tritium in the divertor lithium of 96%, one can obtain a mean residence time in the divertor lithium of about 0.5 day and a total processing residence time of 0.6 day. This leads to a AI of about 12 kg for UWMAK-I, where 3.2 was calculated by the previous model.

5. 4. Other inventory consMerations

In the operation of any commercial power reactor, it is desirable to maintain continuous plant operation if at all possible, even though certain non-essential systems are down. In evaluating reserve fuel needs, one should determine what component within the tritium fuel system can be non-functional while still maintain- ing plant operation.

Of primary importance is the fuel injection system. If this system or a back-up system fails, then continued plant operation would be impossible since fueling of the plasma would cease.

In the blanket, a failure of the tritium removal system would not necessitate shutdown of the reactor. The primary restriction may be the allowable release of tritium to the environment.

In UWMAK-I the blanket system is composed of 12 modules, each of which has its own lithium circuit and tritium removal system. To examine this aspect we might make the following assumptions about the behavior of the tritium extraction system:

(a) the normal plant release is 1/2 of the allowed yearly release;

(b) tritium release occurs primarily through the steam cycle such that at steady state the release is linear in concentration of tritium in the blanket;

(c) the removal time of tritium from the blanket is one day;

(d) once the reprocessing and extraction system re- commences operation, the tritium inventory can be restored rapidly to its normal level.

Then the maximum number of consecutive days N O of operation with one removal system down is


given by

N o

[~x + ~-(nx/12)] + (365 - N o ) ( ~ x ) = 365x, (16) n=0

where x is the average allowable release rate, giving N O = 93.5 days.

The daily release rate at the end of 93.5 days for x = 10 Ci/day is ~ 44 Ci/day or 4.4 times the average daily steady state release rate.

If all removal systems are failed, then N O = 27 days. At the end of this period the release rate is 16.3 Ci/day (forx = 10 Ci/day).

The above calculations assume that there is no limit on the one day allowable release to the environment as long as the maximum yearly release is not exceeded. This may or may not be the case. A one day maximum release limit might be imposed.

The excess fuel needed for continued operation will depend on allowable releases and the number of removal systems failed. For a burn-up rate of 0.7 kg and a breeding ratio of 1.5 in UWMAK-I, the tritium production rate is 1.05 kg/day and consequently four removal systems would have to be failed before excess tritium would be needed. If all 12 units were down, then the amount of excess fuel needed would be between 3.2 kg and 18 kg depending on the maximum allowable releases. It appears, then, that the chosen minimum inventory of 10 kg may be satisfactory in this regard.

One reprocessing system exists for the entire UWMAK-I divertor, rather than modules as is the case in the blanket lithium circuit. A failure in the fuel effluent removal system would result in the saturation of the divertor lithium with Li(D + T) after 60 cycles. This would result in precipitation of the hydride in the cold end of the heat exchanger. Based on an operating inventory of 3.4 X 104 kg and a mass flow rate of 480 kg/sec, saturation would occur after approximately 1 hr, leaving little time to attempt repairs during operation. This would indicate that a reserve fuel inventory is of little consequence in a failure of the divertor lithium circuit fuel effluent removal system, since shutdown of the reactor will probably be necessary.

In the remainder of the effluent processing system, however, no component is essential to the continued operation of the plant. A failure in the purification

and isotope separation system would not necessitate shutdown of the plant. However, a rather large flow rate of tritium exists here, and to maintain continued operation of the plant an equivalent amount of fresh fuel would have to be made available to equal the 7 kg/day throughput of tritium in the processing system. The above value of the minimum operating inventory would be inadequate in this circumstance.

Based in part on the above, and if one considers the possibility of low burn efficiencies and other factors [4], it appears that the maximum total inventory at the plant site could be a few times greater than the stated inventory for UWMAK-I.

6. Loss-of-flow accidents in lithium-cooled CTR devices

Loss-of-flow accidents (LOFAs) for the divertor and blanket regions of lithium-cooled CTR devices have been considered to elucidate the potential safety hazards of such accidents and to delineate any differences in operating behavior that might be expected in com- paring these two cooling systems.

SHIELD

BLANKET

TOROIDAL COILS

Fig. 11. Coolant loops for UWMAK-I.


LITHIUM IN

1 Tx'C

GUTTER REGION

(a-a CROSS-SECTION)

Fig. 12. Schematic of one divertor liquid lithium collector plate.

Fig. 11 gives a schematic concept of UWMAK-I which was used as the model for the analysis. The blanket region completely surrounds the plasma. The li thium enters and leaves the blanket through pipes.

In the double-nulled divertor concept that is shown in fig. 11, there are divertors on the top and bo t tom of the chamber. These are stainless steel plates with a film of flowing li thium on each side of the divertors onto which charged particles, including bo th unburnt fuel and impurities, impinge. The particles are guided by the magnetic field lines along the separatrix. The divertors provide the major part of the vacuum pumping in the plasma chamber (95%) and remove the heat carried by the impinging particles.

Fig. 12 gives a schematic picture of one o f the divertor plates. The li thium enters the divertor from the top and flows down the stainless steel plate into a gutter from which it is pumped out. In a LOFA, it is assumed that the source of l i thium at the top of the divertor is terminated, while the plasma continues to burn uninterrupted. The lithium film that is left on the divertor plate will continue to drain off. In the absence of prompt plasma shutdown, ions will continue to impinge the divertor plate.

To model the divertor LOFA, the material, energy, and pressure balance equations shown in table 3 were used. In the first equation the rate of change o f the lithium film thickness is determined by the rate of lithium vaporization and the rate of l i thium outflow. In

Table 3 Material, energy and pressure balance equations

d6 Li/dt = -h - Q/L W,

Ein U(8 Li) PLi/~AHvap =

= [PwSwCp, w + 6LiCp, LiPLiU(6 Li)] (dT/dt) ,

V(dP/dt)

5Li h

0

L W

U(SLi)

OLi AHvap Pw 6 w Cp, w Cp, Li T

V P

+ SP = rU(8 Li) •

lithium film thickness recession rate of the lithium surface rate of discharge of lithium due to flow off of the wall length of the divertor wall width of the divertor wall rate of energy deposition on the divertor wall step function, I while lithium remains and 0 when the lithium is totally consumed. lithium density latent heat of vaporization of lithium density of wall material thickness of wall material heat capacity of wall material heat capacity temperature of the divertor wall and lithium film volume of the plasma chamber pressure of vaporizedlithium in the plasma chamber pumping speed of the remaining vacuum pumps vaporization rate for lithium

the second equation the rate of energy storage in the lithium film and stainless steel plate is determined by the energy deposited and the energy loss by vaporization. In the third equation the rate of change of pressure in the plasma chamber is determined by the rate

at which the pumps decrease the pressure plus the rate at which li thium vaporization increases the pressure. These equations were solved numerically.

The sequence of events occurring at the divertor after the loss of flow are shown in fig. 13 where the temperature, pressure and film thickness are given as a function of time. More than half of the l i thium drains off of the divertor wall in the first half second. Approx- imately 4 sec after LOFA initiation, l i thium vaporization becomes the major gas load in the chamber (>2600 tort 1/see). The temperature continues to rise

228 D. Okrent et aLI Safety o f Tokamak fusion power reactors

"°1,"! "1- !! / I ,o Jl i i / / - i°. '11 i 'D / 4"- :!, 2 I

i / / i i -1 [ ~ ~ " l I - - , - - FILM THICKNESS| l'~ ~-"

"; X DIVERTER LOFA: 1 o.~ F " ~ i ! INITIAL FILM TEMPERATURE: ~o~I - I .oo

l / " - , . J " I INITIAL FILM THICKNESS: 1 MM / l / / 7 ~ I DIVERTER HEIGHT: 100 CM /

I / i -'\_ I i 0 • ~ I 0 600

0 4 8 12 16 20 TIME (SECONDS)

Fig. 13. Lithium film thickness, pressure, and divertor temperature for a divertor LOFA.

causing more rapid evaporation of the lithium film. As- suming the lithium vapor can fill the entire chamber and that the plasma maintains steady operation, a peak pressure of roughly 4 × 10 -3 torr occurs at around 10.5 sec, at the time of dry-out of the divertor. If the plasma is still running, some fraction of the lithium entering the chamber will be ionized and then may be contained in only a part of the chamber volume, because of the magnetic fields present. For such a case, greater local pressures may be expected. After divertor dry-out, there will be no lithium source, and the pressure will closely approach its operating value (-~10 -5 torr) within 1 sec. Continued plasma operation will melt the divertor wail within 30 sec if adiabatic conditions are assumed.

In the region near the failed divertor it is of importance that a pressure pulse is expected between 4 and 10.5 sec after the initial loss of flow. This time range may then be taken as an upper limit on the time available to respond to the loss of flow, by either restoring lithium flow or shutting off the influx of energetic particles onto the divertor, assuming that the influx of lith-' ium does not quench the plasma. If no response occurs, the pressure pulse may interact with the plasma causing its energy to be deposited onto the first wall. This could cause portions of the radioactive first wall to be ablated and transported to other parts of the reactor.

The other accident which was analyzed is the LOFA for the blanket region. An event tree for the blanket LOFA is shown in the first event tree (fig. 2). The case where plasma reaction shutdown occurs (the more be- nign case) will not be discussed. Where shutddwn does not occur, the major cooling mechanism to prevent melting of the first wall is natural convection; however, it does not appear that natural convection will occur throughout the entire blanket because of gravitational and magnetic forces which inhibit lithium motion. Melting or failure of the first wall will lead to lithium entering the vacuum chamber. This will cause the plasma to be quenched, if this has not already occur- red. The collapsing magnetic field leads to the possibility of large structural forces. Also, since lithium enters the vacuum system, it might cause mechanical failure of the Roots-blowers of UWMAK-I, resulting in missiles.

7. Some safety considerations o f superconducting magnet systems

Large scale superconducting magnet systems are used in all present conceptual designs of Tokamak- type central station power reactors. In considering safety related aspects of such systems, a premise is


made that the related hazards are associated with body forces, stored energy, and magnet mass, but not with the potential release of magnet materials to the environment. Energy is stored both as magnetic field energy and enthalpy of the liquid helium or supercriti- cal helium used to cool the magnets. Body forces related to magnets are due to time varying magnetic fields, magnetic field gradients, and magnet weight. Superconducting-normal transitions leading to magnetic field collapse (magnet quenches) will result in induced forces and currents in magnets and conductors present, as well as an increased rate of helium boil-off in the liquid-cooled case. To estimate the rate of field collapse, two scenarios have been investigated. First, a magnet loss-of-coolant-accident (LOCA) is considered. A fraction of the magnet mass then is allowed to adiabatically thermalize with a fraction of the energy stored in the magnetic field of the failed solenoid. In the second case the thermal history of a solenoid is investigated for the case of a local hot spot being formed; several physical properties of the stabilizer material are considered to be degraded.

Stored energy in magnet systems gives a measure of the amount of work available from a magnet failure. The magnetic field energy associated with the entire UWMAK-I magnet system is 350 GJ (1GJ = 1 giga joule) and with the entire PPPL (Princeton Plasma Physics Laboratory) design magnet system [9] is 250 GJ. For UWMAK-I the difference inenthalpy between 4.2 and 300 K for the entire helium inventory within the primary containment (250 000 1) is 84 GJ. These values are not directly comparable as very different physical mechanisms are involved for the two different cases. One of the main reasons for more detailed investigations of accident paths is to estimate the efficiency and ability to localize stored energy as it is transferred in an accident sequence. Superconducting solenoids will experience considerable body forces throughout the burn and refueling. cycles. Toroidal field (TF) coils in present conceptual designs weigh on the order of hundreds of metric tons. For UWMAK-I [8] ~ach of the 12 TF coils weighs 900 MT and for the PPPL [4] design each of the 48TF coils weighs 130 MT. Vertical field and ohmic heating coils are generally shaaller in these designs, but of comparable size.

Several simplifying assumptions have been made for the LOCA analysis. It is assumed that there is a com-

plete loss of cryocoolant from the magnet dewar. A partial loss of helium could represent a more severe initial condition, as helium heated isochorically from a 4.2 K liquid under its saturated vapor to 300 K has a final pressure of 700 atm. As such, in the event of a partial loss of venting, there is a possibility of coupling of a partial LOCA to a dewar overpressure and rupture. A further assumption is that all of the magnetic field energy to be transferred to the magnet structure does so adiabatically. Such an assumption neglects conduc- tion to supports and any helium remaining in the dewar. The supports should be thermally isolated to minimize refrigeration costs. Conditions imposed upon the magnet are that the windings are in series and that the magnet operates in a persistent mode. This re- quirement implies that current leads to the power supply remain intact, the power supply is of low impe- dance and is not being driven, and any safety system has failed. Further simplifications include the use of cylindrical rather than toroidal g~eometry, and neglect- ing coupling to any nearby conductors.

Prior to the loss of coolant an event or series of events must take place to initiate the process. One such

event could be a loss or degradation of dewar vacuum or partial loss of superinsulating material leading to an increased boil-of rate for liquid cooled systems. A missile generated by some other accident sequence could impede the flow of coolant or rupture the dewar. Earthquake motion may also be considered as a possible initiating event for a LOCA as it could degrade a dewar vacuum or induce currents, along with the already present magnetic fields, in magnet and dewar structural material, increasing the rate of boil-off and possibly failing transfer lines.

Effects are also considered which would tend to localize the deposition of magnetic field energy in the event of a magnet LOCA. An increase in stabilizer re- sidual resistivity due to neutron induced displacement damage would tend to deposit more field energy at point where the quench begins. An increase in the electrical and thermal resistivity of the superconductor- stabilizer interface or locally degraded superconductor pinning properties would also tend to localize and en- hance the effects of a LOCA.

Magnet current and temperature as a function of time and final temperature as a function of magnet mass being heated are calculated. A simultaneous solu- tion of energy balance and circuit equations provides


the time dependence of the current and temperature. Conservation of energy can be stated in the form

12R(T) = m C p ( T ) ( d T / d t ) , (17)

where I is the magnet current, R(T) is the stabilizer re- sistance, m is the mass of the magnet being heated, and Cp(T) is the total magnet heat capacity. The circuit equation is given by

L(dl/dt) + I R ( T ) = O, (18)

where L is the self-inductance of the magnet. Persistent mode is assumed, necessitating the homogeneous equa-

tion. Cp is taken to be proportional to T 3 below an average debye temperature and a constant at higher temperatures. Electrical resistivity is taken to be constant below 20 K and to vary linearly with temperature above 20 K.

To obtain a value for the final temperature of the heated fraction of the magnet mass a further assumption was used, namely that the solenoid is an infinitely long right circular cylinder. The stored energy is then

field energy/unit length = (rrD2]4)(B2/l~O), (19)

where D m is the effective solenoid diameter and B is the magnetic induction. Then, setting D m equal to D, the mean solenoid diameter, the total temperature change of this ideal model is given by

d~ T~ = ~ (B2 D/I a o~P Cpfid), (20)

where ~b is the solenoid thickness, pis the mean solenoid destiny, Cp is the mean specific heat, and lid is the fraction of the solenoid mass adiabatically thermal- ized by the field energy. At= is thus proportional to B 2 and DIG B 2 is less than 100 T 2 for NbTi, D/4) is greater than 10 due to physical constraints on the solenoid structure and is taken to be 20 for the current analysis.

Final temperatures T,~ have been determined as a function of the mass fraction lid [26]. Allowing for the possibility of phase transmons in extreme situations [27] T= may be evaluated at T= = T~ (eid), where eid is the specific energy related to the solid mass of the ideal model, i.e. the quantity of field energy in kJ deposited in a kilogram of magnet mass. Then the mass

fraction fi d may be expressed as

lid = ~ (B2D/IJoOPeid) • (21)

Numerical values of the related final temperatures have been determined on the basis of Cu (fig. 14) as a function of specific energy expressed as l/lid.

For the calculations of thermal histories a modifica- tion of the ideal model has been introduced to relate more closely to design parameters of the UWMAK toroid system. Thermal response functions have beei. evaluated assuming that insulation between windings and structural support prevents efficient heat transfer for the duration of the thermalization process (subscript id replaced by subscript Cu). The related specific energy parameter ecu is the energy (LIg/2) per windings mass affected (the mass has been calculated on the basis of an average turn). Fig. 15 shows the in- stantaneous dissipation rate versus time for various specific energies ecu. Peaks occur as stabilizer resistivity increases with time as the temperature rises, while de- creasing current finally dominates. Characteristic times

T~ K

1000

500,

300

100

50

(MP) cu / , - - /

l l / i t / I B,T / / I / I / /

I / / / , / / , ' , % " ~ - ,

i 5 I / I

, , , ; / ,"-... I I ,,//

I / / " I ~ 7 I / 1 ' I i

I I i I i , / I i / / "p/ / i / / / I

~#SSS

i i , i i i I 10

I | i a I ! I I I I I I | i

fl

Fig. 14. Temperature excursion of adiabatic thermalization based on the simplified energy estinaate eq. (20); (thermodyn- amic properties used are those of copper; both structural supports and windings are heated); (MPcu is the melting point of copper).

D. Okrent et al. I Safety o f Tokamak fusion power reactors 231

12R W

109

108

234

I I I

l !

!

#

I I I I-' i: ,!

~cu KJ/Kg

, ' - 117

59

/ • 29 / j~%'% "- .%

/ l # . / • ,,, %% % % % i I i • % % •

I / ', % l l % % ,¢ % %

I% %% I / i I

50 100 t,m¢

•%

1SO

Fig. 15. Ins tan taneous dissipation rate versus t ime for various specific energies eCu.

required to reach the final temperature are of the order of tens of seconds. AT= ~ T= due to the small initial temperature.

Final temperature as a function of the reciprocal magnet mass faction f~ 1 is presented in fig. 14. For the case of a large fraction of the magnet absorbing the field energy, final temperatures below room temperature are predicted. When very large values of induction or values of lid of the order of a few per cent are taken, final temperatures will be reached which might be expected to destroy superconducting filaments or insu- lating material. Accordingly, the energetics of magnet scenario 1 yield high temperatures To~(eid ) for lid < < 1. The related non-uniformity of the spatial temperature distribution is considered in scenario 2, which is discussed below.

In addition, it is noted that for fcu = 1 the functions allow an assessment of the Cu-stabilizer size required in order to prevent intolerable T= values during nearly uniform thermalization (proposed maximum adiabatic temperature excursion criterion [26]). This point is il- lustrated in fig. 16 which allows a comparison of function T=(eid) of the ideal model with function T=(efu ) of the UWMAK system calculation. An example of the design criterion ('MATEC') is included which presumes that organic materials of the magnet system tolerate only a maximum temperature 7 ' , = 350 K. Then the maximum ecu value for the sizing of the copper stabilizer is about 23 kJ/kg.

The second scenario for a magnet failure involves the

/

K / " T®(%U) MATEC

/ _ / _ _ / / / / -/.. .,:o' //d 300 /"

s s

I I , I I I I I 10 30 50 70

KJIKI

Fig. 16. Final temperature T~ as a funct ion of the specific energy eid of the ideal model and Too as a func t ion of ecu of the windings; (MATEC designates an example of the adiabatic temperature excursion criterion with To. = 350 K).

description of a very localized event. At some point along the selenoid windings either a manufacturing flaw or greater than anticipated neutron damage is postulated to degrade the conductor. Degradation involves a decrease in the superconducting-stabilizer interface thermal and electrical conductivities. Any normal metal of low electrical conductivity placed around filaments to decrease coupling losses is neglected.

The initiating event is a catastrophic flux jump in a single filament heating all of the filaments above their transition temperature at one point along the conductor. A time of 10/asec is taken for the time to transfer electrical current to the stabilizer.

A temperature profile of the form A exp( -bx 2) is imposed on the conductor at t = 0. For a half width of. 10 cm the peak temperature is 200 K.

Total cryostatic stability is designed into the present generation of magnet systems [4,8,28]. Such a stability criterion requires that a normal region in the conductor does not propagate. However, if the stabilizer electrical conductivity, thermal conductivity, or helium bath cooling power are degraded, such may not be the case. Therefore, as these properties may be expected to degrade over the life of the reactor, the stabilizer electrical and thermal conductivities and helium bath cooling power alone and in combination are assumed to be degraded by a factor of ten.

To calculate the time dependence of the peak temperature and solenoid current, the diffusion equation is solved in two dimensions for heat flow in the sole-


noid. Dimensions for the solenoid are those of magnet D-I of UWMAK-1 [1]. The diffusion equation

1 0 T _ - A (22) V2T D at k '

is solved numerically, where D is the thermal diffu- sivity, k is the thermal conductivity, and A is the source term. D, k and A are all temperature dependent. The source term is the joule heating less the helium cooling. Boiling heat transfer to the helium bath is taken to be that presented by Brentari et al. [29] and accounts for film boiling. Current decay is calculated by performing an energy balance at the end of each time step.

Fig. 17 presents the peak conductor temperature as a function of time. In all cases the current decreases less than 10 A from an initial current of nearly 10 kA during the times considered. As an absolute upper limit on the temperature range over which the analysis may be applicable, calculations were continued until the peak temperature rose above the melting point of the stabilizer material. Clearly, insulator breakdown, thermal strain, or some other mechanism will come into play earlier. However, as the peak temperature generally rises quickly at higher temperatures, the

1400

1 2 ~

w

u m

; 8~

• -0.1x HELIUM COOLING POWER AND COPPER ELECTRICAL AND THERMAL CONDUC.

~ ) • TIVITIEE [ ] O -0,1x COPPER ELECTRICAL

AND THERMAL CON- • DUCTIVITIES

O [ ] • -0.1x HELIUM COOLING POWER AND COPPER

• • THERMAL CONDUC. [ ] TIVITY

[ ] -0.1x COPPER THERMAL O • CONDUCTIVITY

[ ] • ~ , l x HELIUM COOLING POWER AND COPPER

• • ELECTRICAL CON- O [ ] DUCTIVITY

• [ ] /k -0.1x HELIUM COOLING • POWER

• [] • ~ ) . l x COPPER ELECTRICAL O CONDUCTIVITY

• [ ] ~7 - AS PER UWMAK-1 OEEIGN 6110 [ • • O PARAMETERS)

4 0 0 ~ g , ' O Orl . . . . ' • "

18 I ~ V I O C ] - - - ' ' • A • A A A AAa m

o l i J i ' ~ J ~. 0 20 40 60 80 100 120

TIME (SEC)

Fig. 17. Peak conductor temperature as a function of time following a local thermal excursion in a solenoid suffering from varying degrees of magnet property degradation.

time to gross failure and field collapse will be of the order of the time taken to melt the stabilizer. It will be noted that times to catastrophic failure and resul- tant more rapid current decay cover roughly the same range of times as those calculated for the adiabatic thermalization analysis.

Each of the three stabilizer properties which are assumed to degrade causes a very different system response. A factor of ten decrease in stabilizer electrical conductivity merely increases by roughly 30 sec the time needed for recuperation. A factor of ten decrease in helium bath cooling power leads to a very slowly rising peak temperature coupled with a growing normal zone. A factor of ten decrease in stabilizer thermal conductivity leads to rapid local heating and a more slowly growing normal zone. In combination the time needed to obtain very high local temperatures may be even less.

A safety analysis based on a slightly different dif- ferential equation has been presented by Yeh [30]. His major results agree with the present work. An ex- tended variation in the transport properties of copper (possibly brought about by neutron induced defects) permits a finite, non-vanishing probability for the occurrence of anomalies.

In summary, some comments may be made based on the two previous accident analyses. Given present plans to apply total cryostatic stabilization criteria, adiabatic thermalization of itself will lead to severe magnet damage in the event that the field energy is localized to a few percent of the total magnet mass. In the event of a LOCA, current decay will be no more rapid than that proposed for UWMAK-II [28] (10 sec) except for very small values off . In the event of a local thermal excursion coupled with degraded physical properties, a magnet may suffer local damage with re- sultant thermal strains leading to gross magnet damage.

8 Containment pressurization from helium pipe rupture

As pointed out in section 2, there is a potential for overpressurization of a Tokamak reactor containment due to the release of the helium in the superconducting magnet. The abbreviated term for this loss-of-helium accident is LOHA.

Current fission reactors have some kind of confine-


ment to minimize the consequence of the release of radioactive materials to the environs [31 ]. A similar design concept is utilized for a fusion reactor. In the preliminary UWMAK-I design [8] the reactor is completely enclosed by the primary containment. The containment consists of a reinforced concrete toroidal roof which is supported by one center column and 12 peripheral columns. The wall of the containment is a free-standing circular steel cylinder which is supported on a track to allow the containment wall to revolve. There is a removable hatch in the cylinder to permit removal of the torus magnet module. To minimize the diffusion of the tritium, the primary containment is to be run at a vacuum. The design pressure load of the containment is given as 15 psig in the inward direction and 10 psig in the outward direction.

Fig. 18 shows the schematic arrangement of the helium system. The toroidal, divertor and transformer magnets are housed inside double walled, stainless steel dewars. The inner vessel is filled with helium which completely submerges the magnet. The 20 cm vacuum gap between the inner and outer walls is f'flled with super-insulation material. The heat generated by the magnet during the normal power cycle will boil off part of the liquid helium. The helium vapor is led away from the dewar through the vent pipe. The vent pipe

DIVERTER TRANSFORMER TOROIDAL MAGNET DEWARS

I

l ' I i I i I PRIMARY I

CONTAINMENT I " - 1

REFRIGERATOR I LIQUEFIER

I ° - - ] STORAGE DEWAR

Fig. 18. Schematic diagram for the liquid helium cycle of a Tokamak-based power reactor.

runs through the primary containment walls and feeds the helium vapor into the liquefying system. The liquid helium is then fed into a 200 000 1 storage tank. The cycle is completed when the liquid helium is dis- tributed back to the magnet dewars. The total helium inventory in the proposed 5000 MW(th) fusion power plant is 450 000 1 with 250 0001 within the primary containment boundary.

There are various possible ways that helium can be released to the primary containment. Abnormal heat generation in the magnet or heat leakage in the insulation system reading to rapid boiling of helium is one possibility. If the vapor relief system is unable to cope with the pressure rise, then a pipe may break or even a dewar rupture can occur. The dewar can also be ruptured by external forces such as those arising from an earthquake. If the dewar ruptures, the helium at first flashes to a two-phase mixture. A small amount of heat transferred either from the magnet or from the containment structure would superheat the helium.

In calculating the pressure and temperature inside the containment following a hypothesized LOHA, 250 000 1 of helium, the total volume within containment, was assumed to be released. In accordance with fig. 2 (Vol. 1, of ref. [8]), the free volume of the containment is 1.6 × 105 m 3. The basic equation for the analysis is the energy equation [32] which states that the internal energy U of the helium in the containment is equal to its enthalpy H plus the heat transferred to the helium Q, i.e.

U = H + Q / M , (23)

where M is the mass of helium. For constant volume V and for superheated helium

vapor, the equation of state relating pressure P and temperature T is

P V = M R T , (24)

where R is the gas constant per unit mass. Three cases were considered. In the first case, one-

third of the energy stored in the magnet is assumed to be converted into the thermal energy due to joule heating, which in turn heats the helium. For this case, the peak pressure inside the containment is found to be 4.1 atm, and the peak temperature is 1027 K. This case should represent a very conservative estimate for the parameters and conditions postulated. For the second and third case, instead of assuming the release


Table 4 Pressure and temperature inside a Tokamak containment following a LOHA

Case Volume = Volume = Volume = 1.6X 10 sm 3 1.8X 104m 3

P(atm) T(K) P(atm) T(K)

1 Magnet energy release 4.1 1027 36 1016 2 Final temp. of 300 K 1.2 300 10.6 300 3 Final temp. of 500 K 2.0 500 17.8 500

of th . magnet energy, a maximum final temperature, is assumed. The final temperature then represents an assumed thermal equilibrium temperature inside the containment. For final temperatures of 300 and 500 K, the pressures are 1.2 and 2.0 atm, respectively. By comparison, the design pressure of the containment is 1.68 atm. It is obvious from these calculations that the heat transfer mode could play an important role in the containment pressurization.

A comparison of Fig. 2 (Vol. 1 of ref. [8]) with data in Vol. 2 of the same report reveals a discrepancy of a factor of 2 in the linear dimension of the primary containment. The free volume of the containment in accordance with Vol. 1 is 1.6 × 105 m3,whi leVol . 2 suggests the volume is 1.8 × 104 m 3. If the smaller volume is used, the maximum temperature and pressure for the magnet energy release cases are 1016 K

and 36 atm. The pressure for the cases of 300 K and 500 K final temperature are 10.6 and 17.8 atm, respectively. The results of these different six cases are summarized in table 4. The free volume of the containment obviously plays an important role in establishing the de- gree of overpressurization.

The results should not be used as a direct assessment of the safety of the UWMAK-I containment design because the design is still in a very preliminary stage and various conservative assumptions were in-

volved in the analysis.

9. Some potential safety hazards of liquid lithium

9.1. In troduct ion

Fusion power reactors operating on a d e u t e r i u m - tri t ium ( D - T ) cycle may contain unprecedented quan-

titles of l i thium in one of several forms. Lithium will be used as the tritium-breeding material and, depending on the form in which lithium is present, may also be used as the reactor coolant. Several current reference designs, such as UWMAK-I [8] have adopted liquid lithium metal as the breeding material and the coolant. The high chemical reactivity of liquid l i thium poses a potential safety hazard in the CTR environment that has been assessed under two conditions of interest.

Lithium is known to burn readily in air with an in- tense flame. The primary containment building in UWMAK-[ is maintained in an inert condit ion under normal operating conditions, but events can be postulated in which this containment is breached. Under such conditions if there is a lithium spill within the containment a lithium air fire might result. A model has been developed to estimate the results of such a lithium-air fire in terms of (a) peak flame temperature; (b) chemical composition of the flame; and (c) heat release. In essence, each of these factors represents a mode by which radioactive or toxic materials might be released from the primary containment region of the CTR.

The primary containment of a CTR such as UWMAK-I is constructed of reinforced concrete clad with steel. It is known that li thium and other alkali metals, such as sodium, can attack concrete, leading to the release of water from the concrete and its sub- sequent reaction to form hydrogen and other products. While the steel liner on the concrete serves to reduce

h ~ 2 0 m / ~ . 4 x l O 4 m 3

i~ - A R E A A ~

Fig. 19. Model for a reactor containment vessel used in model- ling an accidental spill of lithium.

D. Okrent et al. / Safety of Tokamak fusion power reactors 235

the opportunity for a lithium spill to contact the concrete containment, the possibility that there will be a partial breach of this cladding material still exists. For example, thermal stresses induced by lithium at 800 K, the coolant temperature in UWMAK-I, on a 300 K steel liner of the dimensions in UWMAK-I, could lead to buckling and the possibility of local failure of the linear. Similarly, seismic events might lead to both a lithium spill and failure of the steel liner. Unlike the case of a lithium-air fire, it is not necessary to posit the prior breach of containment for the occurrence of a lithium-concrete interaction. Hence the possibility of overpressurization of containment must be assessed for the lithium-concrete interaction. The lithium is presumed to form a pool above the breached section of the steel liner, as shown in fig. 19.

9.2. Quasi-equilibrium models for the l i thium-air and

l i thium-concrete interactions

Relatively little is known about the kinetics of product formation in the lithium-air and lithinm-concrete interactions, which precludes doing a fully dyna- mic calculation of the course of such an interaction. The assumption will be made, therefore, that the con- sumption of lithium is a rate process, while the products formed by the reaction represent an equilibrium distribution. The model provides an estimate of the temporal history of the reaction.

A series of constraints must be placed on the course of the reaction, namely, (a) reactant stoichiometry; (b) chemical equilibrium between the product species;

and (c) overall adiabatic behavior. These three constraints arise from the conservation of mass and energy, and from the virtually adiabatic nature of the containment over the time scale of the interactions. In addition, a pressure constraint is required. Because of the physical models this constraint is taken as a fixed total pressure for the lithium-air fire, since the containment is breached, and as the lithium pressure for the lithium-concrete reaction, since this will be maintained in equilibrium with the liquid lithium pool.

The temperature of the lithium pool is not predic- table simply from the equilibrium state of the system, but rather is determined by various initial conditions

and thermal transport rates. The initial temperature will be the lithium temperature from the location at which the spill occurs. The temperature of the pool may decrease, due to effects such as thermal transport to the concrete or the endothermicity of the dehydra- tion reaction of concrete, or it may increase due to exothermicity of the reaction with water and subse- quent heat transfer to the lithium pool. In fact, during the course of the interaction it is expected that the temperature of the pool would first decrease, and later increase.

9. 3. Lithium - air fire

The calculation of the course of a lithium-air fire was performed as described in the previous section. The results of the analysis show that the fire may produce (a) a peak flame temperature of 2400 K; (b) a heat release of 0.3-0.7 TJ/% UWMAK-I lithium

inventory; and (c) a partial pressure of highly reactive species greater

than 10 -3 atm. The peak flame temperature is significant in that it is sufficiently high to melt and mobilize all proposed first wall materials other than niobium and TZM. The heat release during the course of the fire coresponds to the thermal output of UWMAK-I for 100 sec, for each 1% of the UWMAK-I inventory of lithium consumed. Lastly, the production of such large amounts of highly reactive materials, primarily oxygen stores, can chemically mobilize approximately 0.1 kg of wall materials such as stainless steel, niobium, or TZM, for each 1 kg of lithium consumed.

9.4. Li thium-concrete interaction

The analysis of the lithium-concrete interaction allows one to compute the course of interaction as a function of the amount of concrete reacted. To develop some temporal sense from such a calculation requires a model for th depth of penetration of lithium into concrete. Efforts to locate this information in the literature have been unsuccessful. Consequently, re- course has been taken to data for the sodium-concrete reaction [33], which is likely to be quite similar to the lithium-concrete reaction due to the extensive chemical similarities between the two metals. On this


11111

1500OK 0o

i 0.1 ~

0.01

"T

x 0.001

! ~ O.OO01

0.00(101 PER CENT REACTIVE FLOOR AREA

~4 ' I ' I ' I ' I 100%

1 10 .2 10 0 102 10 4 10% I ' I i I i I i I

10"4 10 .2 10 0 10 2 104 1%

lOO ,k ,k 0.1% I J I i i

10, 1oO TIME (HOURS)

Fig. 20. Percent UWMAK-I lithium consumed as a function of time following a lithium spill.

basis, we may estimate the depth of penetration of lithium into concrete d as d = dot1~4 which is not expected to be particularly sensitive to the temperature of the lithium pool. Comparison of figs. 20 and 22 will show that, depending on the lithium pool temperature, in excess of 1 TJ of work may be extracted for each 1% of the UWMAK-I lithium inventory consumed.

The total pressure within the containment building may be evaluated either at the reaction temperature or after cooling on the various heatsinks within the containment building. These results are shown in fig. 21 taking a minimum heat sink temperature of 400 K. In this case, the results are not as sensitive to the lithium pool temperature, so no effort has been made to estimate its value for a realistic model. Of critical importance is that the pressure may reach the design level of the UWMAK-I primary containment, 1.7 atm, in about 1 h even if only 1% of the concrete floor area is active towards lithium. In fact, the potential pressures are sufficiently high that overpressurization of the primary containment after reaction product cooling on

heat sinks to 400 K is a realistic possibility, as is suc-

10,111111

O 10

0.1

0,01

/22" . R E = U R E A T =cr.,o.,/ , ,

y 4~OK

PER CENT REACTIVE FLOOR AREA

I ' i , i , i , i 100% 1 ~ lO -z lO 0 102 104

10% I ' 1 J I ' I ' !

10 4 10 -2 10 0 102 104 I ' I w I ' 1 I I 1 %

10 4 10 .2 10 0 10 2 104 0,1% T J I i f , I

10-2 10 0 10 2 10 4

TIME (HOURS)

Fig. 21. Containment pressure as a function of time following a lithium spill.

11111

P_

¢~ 10

ID ~ 0.1 < > <

0.01

j 1

-- ~ / TEMPERATUR

PER CENT REACTIVE FLOOR AREA

1 ~ lO 2 lO o 102 104 I ' I ' I ' I ' I 10%

10"4 10-2 100 102 104 I ~ I ' I ' I ' I 1 %

lo'4 1~ z lO 0 102 104 I ' I ' I ' I 0 . 1 %

10"2 10 0 10 2 10 4

TIME (HOURS)

Fig. 22. Available Canot work as a function at time foUowing a lithium spill.


cessive overpressurization of both primary and secondary containments. However, if 77 can be kept small,

for example less than 0.1%, then the times required to reach these conditions become extremely long, amounting to weeks or months.

During the course of the interaction hot reaction products can deposit their energy in the colder parts

of the CTR, such as the containment structure. In doing so work may be extracted, for example to damage and create missiles from the concrete structure. The amount of such work available is shown in fig. 22. Again, the result corresponds to the thermal output of UWMAK-I for 100 sec for each 1% of the UWMAK-I inventory of lithium consumed. Lastly, the production of such large amounts of highly reactive materials, primarily oxygen atoms, can chemically mobilize approximately 0.1 kg of wall materials such as stainless steel, niobium or TZM, for each 1 kg of lithium consumed.

Finally, it is noted that a considerable amount of LiH may be formed, and the toxicity of this material may require consideration in accident evaluation.

9.5. Conclusions

The effects of l i th ium-ai r and l i th ium-concre te interactions have been considered within the confines of a fusion power reactor such as UWMAK-I. The consequences of these interactions appear to present circumstances of a potentially serious nature, with (a) large heat and potential work release; (b) the formation of toxic species; (c) the potential for initiating secondary events with

related results. These circumstances may therefore bear more care-

ful study or consideration in future reference designs

for lithium-cooled fusion power reactors.

Acknowledgement

This reasearch was supported in part by the Electric

Power Research Institute.

R e f e r e n c e s

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[15] R.V. Anderson, AI-AEC-13136, Sept. (1974). [16] D. Steiner, ORNL-TM-3094 (1970). [17] F.N. Flakus, At. Energy Rev. 13 (3) (1975) 587-614. [18] W. Hafele and C. Starr, A perspective of the fusion and

fission breeders, Unpublished Report, Kernforschungs- zentrum, Karlsruhe, W. Germany, Feb. 1973.

[19] J.E. Draley and S. Greenberg, Some features of the impact of a fusion reactor power plant on the environment, Texas Symp. on the Technol. of Controlled Thermo- nuclear Fusion Exp. and the Eng. Aspects of Fusion Reactors, Austin, Texas, Nov. 1972.

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[21] J.E. Draley, V.A' Maroni, T.A. Coultas and R.A. Krakowski, An environmental impact study of a reference theta-pinch reactor (RTPK), Proc. 1st Topical Meeting on the Technol. of Controlled Nucl. Fusion, San Diego, California, Apr. (1974).

[22] J.B. Watson, Tritium containment and recovery in fusion reactor systems, Topical Meeting on the Technol. of Controlled Nucl. Fusion, San Diego, California, Apr. (1974).

[23] M.J. Sehnert, Tritium inventory considerations in con-


trolled thermonuclear reactors, M.S. Thesis, School of Eng. and Appl. Sci. UCLA (Sept. 1975) Contract ULCA- 7589.

[24] W.F. Vogelsang, Nucl. Technol. 15 (1970) 470. [25] R.W. Conn, C.L. Kulcinski and C.W. Maynard, Prelimi-

nary design aspects of the UWMAK-II conceptual Tokamak power reactor, Proc. Amer. Nucl. Soc., Winter Meeting, 16-21 Nov. 1975, San Francisco, p. 59.

[26] R.C. Amar, T.H.K. Frederking and W.E. Kastenberg, Cryogenic Engr. Conf., Kingston, Ontario, July 1975, paper B-7.

[27] A. Bejan, T.A. Keim, J.L. Kirtly, Jr., J.L. Smith, Jr., P. Thullen and G.L. Wilson, Advan. Cryos. Eng. 19 (1974)

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Physics and Controlled Nuclear Fusion Research, 5th IAE IAEA Conf., Nov. (1974).

[29] E.G. Brentari, P.J. Giarratano and R.V. Smith, NBS Technical Note No. 317, U.S. Department of Commerce, National Bureau of Standards, Washington, D.C. (1965).

[30] H.T. Yeh, Thermal aspects of a superconducting coil for fusion reactor, 6th Fusion Technol. Conf. (1975).

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