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l

4:. f NEACRP-A-975

Fuel Management Effects on the Inherent Safety of

Modular High Temperature Reactor .

Yukinori HIROSE +, Peng Hong LIDI, Eiichi SUETO>lI,

Tohru OBARA and Hiroshi SEKIMOTO

Research Laboratory for Nuclear Reactors,

Tokyo Institute of Technology l

+ Present address : Nippon $izomic Indcstry Gro>~.!r Co.,'Ltd., 1.

to be published in J. of~Nuc1. Sci. Technol.

ABSTRACT

Analysis was performed on the effects of fuel loading

schemes and fuel materials on the inherent (passive) safety

characteristic of the modular pebble-bed type high-temperature

reactor against depressurization accident involving loss of

helium forced circulation. Two extreme fuel loading schemes, the

infinite-velocity multipass and Once-Through-Then-Out (OTTO),

were evaluated for both uranium and thorium fuel cycles. The

results of th2 analysis show that the maximum core temperatures

attained following the accident xere much lover for the infinite-

velocity multipass scheme than for the OTTO scheme. For both

schemes, the thorium cycle showed slightly higher maximum peak

temperatures, compared to uranium cyc1.e.

KEY \YORDS : fuel management. modular high-tenserature reactor,

inherent safety, depressurization accident, loss of helium forced

circuletion, OTTO, infinite-velocity mUltipaSS. fUel:CyCle,

equilibrium condition, maximum core temperature

15200002 i

I. INTRODUCTION

As an inherently safe reactor, the modular high temperature

reactor was proposed through development of the pebble-bed

reactor. The inherent or passive safety for the high-temperature

reactor can be stated as the capability of the reactor to limit

the maximum fuel temperature to the values for which the fission

.product release will increase the fission product inventory in

the primary circuit by factors, but not by order of magnitude (1)

The well known TRISO particles with silicon-carbide (Sic) layer

comfily \i-ith this requirement for temperatures up to 15OO'C. in

other tvords, providing the modular HTR design with passive safety

property means limiting the maximum fuel temperature below 16OO'C

during the accidents.

One essential determinant of maximum fuel temperatures for a

pebble-bed reactor is the in-core fuel management. This involves

numerous parameters such as choice of fuel loading scheme, fuel

material, etc.

In the present paper, the above two parameters were

considered. Concerning the first, two extreme fuel leading

schemes, infinite-velocity multipass and Once-Through-Then-Out

(OTTO), !vere considered. For the fuel material parameter, lolc

enriched uranium fuel and thorium with 233. L' fuel, were compared.

II. FUEL LOADIXG SCHEYES AND ACCIDEXT COSDITIONS ;

In pebble-bed reactors, fuel balls are loaded from the top

2

of the reactor, flow through the core and are discharged from the

bottom. The fuel ball loading schemes commonly used for the

pebble-bed reactor may be classified into multipass and OTTO

schemes. For the multipass scheme, the fuel balls flow through

the reactor numerous times before being finally discharged. But

for the OTTO scheme, the fuel balls transit the core only once.

The OTTO scheme may attain a nearly exponential axial power

profile with the aim of fuel temperature flattening. For both

0 multipass and OTTO schemes, the ball flow velocities affect the

nuclide density distributions, neutron flux, burn-up, .and in turn

the power and temperature profiles. With slower ball flow

velocities, the power peak shifts to the upper part of the core.

This effect is greater for the OTTO scheme. If the velocities

become infinite for the multipass scheme, the power peak becomes

broader and occurs at the center. For the usual multipass

scheme, the ball velocity is slow and the power peak appears in

the upper part of the core showing the characteristics between

the OTTO and infinite-velocity multipass schemes.

0 From the above, investigations were carried out on the OTTO

'.. L and infinite-velocity multipass fuel loading schemes. The

inherent safety characteristics mere evaluated foi each scheme

during the depressurization accident involving 10~'s of helium

forced circulation after the equilibrium condition. YFurthermore,

two kinds of fuel material - low enriched uranium dioxide and

thorium with 233U dioxide - were also investigated'for each

fuel loading scheme.

15200004

It should be noted that for the infinite-velocity multipass

loading scheme, the ball flow velocities are large enough that

they can be assumed to be infinite. For this loading scheme the

nuclide densities can be treated to be spatially uniform. The

accuracy of this assumption was investigated at the start of the

present study. For the OTTO loading scheme, all nuclides are

treated space-dependently, since the ball flow velocities are

relatively slow. The details of the OTTO cycle burn-up

simulation are presented in reference (2) .

After the accident, the reactor becomes subcritical by its

negative temperature coefficient reactivity feedback, but the

cooling power of the helium goes down through the

depressurization of the helium coolant. Decay heat is removed

only by conduction inside and betii.een fue~l balls, radiative

cooling, natural convection of helium, and conduction through the

graphite reflector into surrounding sections.

0

III. CALCULATIONAL PROCEDURE

Based on the conditions stated in the preceding section, the 0

following calculations were performed for both OTTO and infinite-

velocity multipass schemes.

1. The Equilibrium Condition of the Core

Search of the equiiibrium condition of the core’(density

distribution for each nuclide, neutron flux distribution for each

4

15200005

energy group, and power distribution) is necessary to obtain

information about the initial conditions prior to the accident.

Considering the fuel ball motion through the core, the

simulated burn-up equation for the OTTO cycle can be written as

fallows(2) :

dN. --v--l +- c A. ai'AiNi'+ c C,ui',a,gPi'-ti,gNi'6g + ds it I' g 1

C 2 z I

G~,,~,~~~,+N~,$~ - hiNi - C ~~ a ,NiGjo= 0 0 g I90 0 (1)

where

Xi(s) : Atomic density of isotope i at position s

S : Distance measured along fuel ball stream line

v : Ball speed

‘i G.

1 ,a,g a. l'+i

T., I -fl,g

@g(s)

Decay constant of isotope i

Absorption cross-section of isotope i for energy group g

Probability that decay of isotope i"produces isotope i

Probability that neutron absorption in isotope i'

produces isotope i

Yield of isotope i due to fission in isotope i' 1.

Neutron flux in energy group g at position s'

Fo: the infinite-velocity multipass scheme, the first term of the

Eq. (1) is removed, and Xi(s) becomes independent of, s.

The neutron transport problem ~2s treated 2s err r-z two-

dimens,ional four group diffusion problem. The group' constants

and their self-shielding'factors 2s a function of temperature

5

15200006

and atomic density were prepared using a part of VSOP code

system(3' (ZUT-DGL. THERMOS, and GM).

2. Steady State Temperature Distribution

After the nuclide densities and neutron flux distributions

for the equilibrium cycle were calculated, thermal-hydraulic

calculation was performed to obtain the temperature distribution

in the core for this equilibrium condition, which-was used as an

initial condition in the accident analysis.

In the normal condition, helium coolant enters the core from

: the upper part of the core, flows throug:l the voids among fuel

bails, and exits from the bottom part of the core.

Considering the spherical shape of the fuel and the void

fraction of the core, an effective coolant flow was modelled. In

this model, within the core the coolant flow through the virtual

flow channels without cross-flow, and in a particular virtual

flow channel the mass flow is assumed to be constant.

For the momentum equation, Ergun equation (4) , which is based

on Slake-Kozeny eqiiation for laminar flow and Burke-Plummer :' i

equation for turbulent flow, was used : r'

where

P': Coolant pressure

G : Mass flow rate for virtual flow channel

6

0, : Coolant mass density

n : Coolant viscosity

g : Gravitational acceleration

c : Void fraction of the core

D P

: Fuel ball diameter

In the core the heat transfer processes involved are (1)

forced convective heat transfer between fuel ba-11s and heliun

0 coolant:. (2) radiative heat transfer between the adjacent fuei

balls, (3) heat conductions inside the fuel balls and ($1 heat

convections in x:he helium coolant. Considering these heat

transfer’processos the oyerall effective heat transfer

coefficient between helium coolant and fuel balls was obtained

and used for the energy balance equation. The energy balance

equation in the fuel can be expressed as :

I a L-k

F cl-&) --- r ai DiJ

h (Ts-T,,) - Q = 0

(3)

where

0 kp : Effective conductivity inside the fuel bell

h : Forced convective heat transfer coefficient ,between .

fuel b2lls and helium coolent

4 : ?Owei density

T. 0 : Eelium bulk temperature

TS : Fuel ball surface temperature

For the coolant, the energy balance equation can be

5

expressed similarly as :

a -i-G T- dz Tb- ’ (lD-‘) h (Ts-Tb) = 0 (4)

where P

k b,r : Coolant effective conductivity in the radial

directicn

k b,z : C0012;;t effec*iv= * - conductivity in the a;iel

direction

Using the equilibrium condition as the initial condition,

the fission products and othe: radiative nuclides distributions,

and the corresponding decay heat’distribution e~fter a

depressurization accident involving loss of helium forced

circulation were calculated. For the fission-products yield

calculation, JSDC FP Decay and Yield Data xere used (3)

After the depressurization accident, the helium

flow velocity was assumed to be practically zero and helium

pressure was also assumed to be equal to the atmospheric

siressure.

The temperature for the mixture of helium coolant and fuel

balls was calculated from the helium temperature and the surface

temperature of the balls.

P&T~~ (1 -&) + pbCpbTpb& Tmx =

PpCpp (1 -E) +PbCpb&

(5) vii th

where

T ms

*P

Pb C

FP C pb T PC T

Pb k

q T =T+- PZ s GOk .DP

: Nisture temperature

: >!ass density of the ball

: Ness density of helium

: Specific heat for cosstanz pressure of the bali

: Specific heat for co:,stac: pressure of helium

: Average temperature cf th? ball

: Bulk temperature of heiium

: Heat conductivity of xhe ball

Finally, xhe heat transfer equation to be solved was

expressed as the follorving :

0 where

k’e : Effective conductivity of the core during the

accident

9 : Decas- heat production during the accident

IV. C.ALCtiLATED RESULTS ASD DISCUSSION

(6)

The reactor configuration used in the calculation is shown

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15200070

in Fig. 1. The main design parameters, as well as the respective

calculation results for the equilibrium condition for each fuel

loading scheme and fuel material, are shown in Table 1. The core

diameter is much smaller than the usual HTR designs in order to

enable both reactor-shutdown by only reflector rods and decay-

heat removal without any active cooling system. For both fuel

loading schemes, the saze fissile enrichments 2nd moderation

-. re~los were used. Tine burnup and conversion r2tios for the

infinite-velocity ~ultipass scheme are relatively higher then for

t;ire OTTO scheme for both uranium 2nd thorium cycles. T'ne Thorium

fuel cycle ge:-e sligtly higher values of bUrnUp 2nd conrersion

r2tio.

Before coGparIng the two extreme fuel loading schemes, we

discuss the effects of the spetial distribution of the nuclide

densities resulted from the finiteness of fuel ball velocity for

the multipass scheme. Even for finite-velocity multipass scheme,

the densities of large half-life nuclides do not depend strong11

on the sp2tial distribution of neutron flux, when the bell

velocir? is high enough. For short half-life nuclides, however,

the dependency on the speti21 distribution of the neutron flux

increeses 2nd these nuclide densities 2ctually exhibit sp2tial

distributions. Since these sp2ti2l distributions 2re essential

in the present enalysis, they were investig2ted for sore

importenr shorthalf-life nuclides, 13'x;e, 14'sm1, 23gsp 2nd

233Pa. and compared with the results for the infinite-velocity

multipass scheme. Discrepancies of not more than 1%. 3% and 1%

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15200011

respectively were detected in group neutron fluxes. densities of r

short half-life nuclides. and burn-up values. For the accident

analysis concerned, these will not contribute significantly to

the final results.

The poser distributions for each fuel loading scheme and

fuel cycle 2re sho1r-n in Figs. 2 through 5. The poxer

distiibuticns of the OTTO scheme shoed peaks on the upper pert

of the core. In the OTTO schtne, since the fuel balls tr2nsit

the core only once, the upper part of the core contains zort

iresh fuel 2nd the lo.<er part cons2ins highly 'burned-up fuel.

For this fuel distribution, the,fission re2ctions mostly occur in

the upper pert of the core, 2nd determine the power density

profiles with 2 sharp peak in this part. These effects appear to

an extreme degree in the present design, since the core radius

is small 2nd axial coupling .of the neutron flux distribution is

weak. The power distribution of the-infinite-velocity nultipass

scheme becomes bro2der and ins peak locates close to the center.

Using the neutronic calculation results for the ecuilibrium

condition, the thertaal hydrzulic calculation rvere done. These

results ere shown in Table 2. hll of the nexinun temperatures of

heliuz and fuel bell occur 2~ the center botton of the core.

Under these masinun temgeretures, the reactor ten be operzted

szfely concerning the fission product release. The masimum

helium temperatures are higher for the thorium cycle than for the

uranium cycle. It is attributed to the larger outlet temperature

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15200012

mismatch among coolant channels for the thorium cycle caused by

the steeper radial power profile.

The accident analysis was begun with the calculation of

decay heat production of fission products, actinides and other

radiative nuclides ,subsequent to the accident. The results for

the OTTO scheme, Figs. 6 and 7, gave distributions of the decas

heax production in similar profiles, ivith the poxer distributions

in the esuilibrium condition. For the infinite-velocity mulxigass

scheme, since the nuclide densities xere uniform, the calculated

results of tine-dependent deca, hear production shoic-n in Fig. 8

are considered valid through a11 core regions.

Esing the tergerature distributions in the equilibrium

condition as initial conditions, and the time dependent decay

heat production of fission product after the accident, the axial

temperature profiles were calculated as shown in Figs. 9 through

12. In these figures, the axial temperature profiles in the

equilibrium condition, lrhich served as initial conditions, are

also show in dashed lines. These initial temperature profiles

are more flat for the OTTO scheme than the infinite-velocity

multipass scheme as rve expected. The maximum tezperetures

attained for each loading scheme and fuel cycle are summarized in

Table 3.

At the beginning of the accident, the temperature peaks for

the OTTO scheme occurred in the upper part of the core. On the

other hand, the peaks for the infinite-velocity multipass scheme

occurred at the lower part of the core. Eoth peaks moved toward

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15200013

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the center of the core with the passage of time.

The masimum temperature of the OTTO scheme esceeded 1800°C

in 23.3 hours for uranium, and 22 hours for thorium cycles

respectively. The maximum temperatures of the infinite-velocity

multipass scheme were much lower than the 1600°C, further, their

temperature transients Kere n?lch slower. The differences in the

mesin:um peak temperatures bei~~=,~ ---1~ the t',i'o schemes exceeded 3OO'C;

considerably large from the reactor design point of view.

Clearly', the distributions of the fission product as well as the

decay heat production for the txo fuel loading schemes, can

explain the above results. The distributions of the fission

product and actinide for the infinite-velocitymultipass scheme

do not produce any peak, so the heat production density is

distributed more uniformly throughout the core.

However, in the usual multipass fuel loading scheme, the

fuel ball velocity is not infinite, and if velocity is very slow

the fission products' distribution becomes similar to the

distribution for'the OTTO fuel loading scheme: the peaks of the

distribution will then shift to the upper part of the core.

Slight differences in the maximum temperatures xere found

betxeen uranium and tho,l -lum fuel cycles. They were attributed to

the different radial power profile in the equilibrium condition

rvhich determined the decay heat productions.

In conclusion, the multipass fuel loading scheme is

relatively safer than the OTTO scheme for the depressurization

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1?200014

accident, thouph the OTTO scheme aims to obtain the higher gas

outlet temperature with lower maximum fuel temperature at

operation condition. By use of the infinite-velocity multipass

loading scheme, a reactor can avoid the 1600°C by a safety

margin of about 300°C.

ACKXO\CLEDG?IEXT

The authors are very grateful to Dr. >I. Aritomi of Tokyo

Institute of Technology and Dr. T. Watanabe of Kawasaki Heavy

Industries, Ltd., for their edvice in performing the thermal

analyses.

REFEREXCES

(1) REL‘TLER, H.. LO:S\'E,PT, G. H.: Sucl. TeChnOl., 52, 2?(1933).

(2) SExI>iOTO, H., ei al.: J. of .\‘ucl. Sci. Technol., 24,

0 755(1S87).

(3) TEUCSERT, E., es al.: Jul-1649, (lOSO)

(4) BIRD, R.

Sons, 196

(3) IH.1R.4, IJ,.

(

B., et al.: "Ti-znspori Phenomena", John Kilev &

1960).

et al.: JAERI-?! 9715(1981)

15200016

.-.

Thermal ~owc,' (M\VI;II)

Core di.amcter (ITI)

Core height (III)

Mean power dcns,lLl~y (W/cc)

Multiplica1:j.o~~ ,Y:~cI:or

Fuel. lcmdlng SCII(!III(!

i;‘u c I. mate IT I ill

Enrichment (X)

Moderation R~~I;.i~o

Burnup (M\Vl3/tolr IIM)

Conversion ralA.0

concl~i.I::i,o~~

Ifelium tcmpcr;~I:urc: ( "C)

System pressure (atm)

Fuel load:l,nr: ~sc:l~cmc?

Fuel materid

Helium mass %‘:l.ow (kg/s)

Pressure ~11~01~ (i,tl.m)

Max. helj.um 1:omp. ( "C)

Max. Pucl surl'. I:c!mp. ("C)

. .

. . . .

1 r

O-

2 113 -

207 -

iO58-

1126-

-II69 -

i263-

(cm)

25 150 260 5 1 I

1 Reflector

Fig. 1 Reactor Configuration

(cm)

* . .

15

10

5.0

0.0

15

10

5.0

0.0

l

‘1 .‘Z.’

0

15200021

: .: .. : .‘.’ :Y:.:.=::s..

-. ..~.

0.30

T 0.25

3 0.20

z 5 0.15

2 0 g 0.10

0.30

z 0.25

g 0.20

z E 0.15 3 E E 0.10

2 2! 0.05

0

r-r ,

I I I I

T =l.h,

T =1Oh

. a,. ,

.

0 200 400 600 800 1000

DISTANCE FROM TOP OF THE CORE (cm)

iiz, 7 :-:d&i produciion aicer reactor s’hurdosa for OTTO ihoricm fuel c?cle 1520002~

IO O F-- ~~-t~~~-rrm~n-rrrrl~~~-nn”n’t-~~-r-r-l-i-l-r~~~-r I I I I I:

-

0' 10 2 IO3 IO 4. IO5 2

m (0 TfM,E AI?TEIR Sf-IUTDOWN (s)

IO 6 lOi

. - ,

.

0 c C’ w u

Fig. 13 lIeat production after reactor shutdoxn for inflnite-

velocity multipass uranium and thorium fuel cycles

2000

800

so0

400

1200

1000

800

600

400

r

I ! / 1 7’

I 4

DISTANCE FROM TOP OF THE CORE (cm)

1400

1200

t-

I I I I

T=63h (max)

i

DISTANCE FROM TOP OF TI-IE CORE (cm)

i 000

800

600

400

DISTANCE FROM TOP OF THE CORE (cm)

max

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