NAT_graphiteXenon.txt

GRAPHITE AND XENON BEHAVIOR
AND THEIR INFLUENCE ON
MOLTEN-SALT REACTOR DESIGN

DUNLAP SCOTT and W. P. EATHERLY Oak Ridge
National Labovatory, Nucleayr Division, Oak Ridge, Tennessee 37830

Received August 4, 1969
Revised October 2, 1969

 

Existing data on dimensional changes in gra-
phite have been fitled to parabolic lemperature-

sensitive curves. From these, the graphite life,
radiation-induced stresses, and permissible
geometries have been calculated. Il is concluded
existing materials can be utilized in a molten-salt
reactor which has a cove graphite life of about
four years, without serious cost penally.

Fission product xenon can be vemoved bY
sparging the fuel salt with helium bubbles and
removing them after envichment. With reasonable
values of salt-to-bubble transfer coefficient and
graphite permeability, the penalty lo breeding
ratio can be reduced to <0.5%. |

 

INTRODUCTION

One of the attractive aspects of the molten-salt
reactor concept is that even the most stringent of
the present materials or process limitations per-
mit reactor designs having acceptable economic
performance. In this paper we consider, as
examples, the effects of finite graphite lifetime
and xenon poisoning on MSBR design.
graphite lifetime implies periodic replacement of
the graphite; neutron economy requires the re-
moval of the bulk of '*°Xe from the core to keep
the xenon poison fraction below the target value
of 3%.

The major economic penalty associated with
graphite replacement would be the load factor
penalty associated with taking the reactor off-
stream. This cost can be circumvented by assur-
ing that the graphite will maintain its integrity for
at least the time interval between normal turbine

NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8

Finite

FEBRUARY 1970

 

KEYWORDS: molten-salt re-
actors, breeder reactors, ra-
diation effects, temperature,
graphite moderator, stresses,
porosity, xenon-135, poison-
ing, helium, bubbles, design,
economics, reactor core, MSBR

maintenance requirements; i.e., the downtime re-
quired for scheduled maintenance should coincide
with that for graphite replacement. Hence, to
avoid load-factor penalties associated with graph-
ite replacement, we have set as a minimum re-
quirement that the graphite have a life of two
years; at the same time, a longer life would be
desirable and consistent with the power industry’s
objective of increasing the time interval between
turbine maintenance operations. Thus, in the
reference design the reactor performance is con-
strained to yield a graphite life of about four
years, and this paper points out the basis for that
value.

The removal of *°Xe from the core also puts a
constraint on the graphite. The fuel salt must be
excluded from the graphite to prevent local over-
heating and also to decrease fission-product poi-
soning, and this, in turn, requires that the
graphite pore diameters not exceed one micron.
This, however, is not a limitation, for the xenon
removal will be shown below to require a gas
permeability of the order of 10”° cm?/sec, and
such a value requires pore diameters of ~0.1 p.

Even though xenon is excluded from the graph-
ite, it needs to be removed from the salt stream
if the desired neutron economy is to be attained.
This removal is accomplished by injecting helium
bubbles into the flowing salt, which transfers the
xenon from the salt to the bubbles and effectively
removes xenon from the core region.

In the following sections we shall discuss in
some detail: the method of analysis of existing
data on radiation damage to permit prediction of
the graphite lifetime in MSBR cores; the applica-
tion of these damage rates and the radiation-
induced creep to calculate induced stresses in
the graphite; the considerations involved in the
distribution and removal of !*°Xe and other noble
gases; and last, the method proposed for safely

179
Scott and Eatherly

collecting and disposing of these gases. We con-
clude that graphite and the removal of xenon
present no questions of feasibility, but require
only minor extensions of existing technology.

GRAPHITE LIFETIME

It has been recognized for several years that
under prolonged radiation exposure, graphite be-
gins to swell extensively, even to the point of
cracking and breaking into fragments. However,
data at high fluences and within the operating
temperature ranges anticipated for molten-salt
breeder reactors were largely nonexistent.
Nevertheless, by using existing data, it was possi-

ble to estimate graphite behavior over the range
of MSBR conditions. For a large group of

commercial graphites—including British Gilso-
graphite, Pile Grade A, and the American grades
AGOT and CSF-it was found that the volume
distortion, v, could be related to the fluence, &,
by a parabolic curve,

v =Ad + Bd® (1)

where A and B are functions of temperature only.
The fit of this equation to the experimental data is
excellent for the isotropic Gilso-graphite, but the
relation is approached only asymptotically for the
anisotropic graphites. We may write the fluence
as the product of flux and time, i.e., & = ¢f. The
behavior of v is to decrease to a minimal value,
v,, and then increase, crossing the v = 0 axis at
a defined time, 7, given by

0= Agr + Bl¢r)? . 2)

Clearly, in terms of v, and 7, Eq. (1) can be

rewritten as
v = 4y, t (1 - —t—)
T T

For isotropic graphite, the linear dimensional
changes will be given approximately by one-third
the volume change. If the graphite is anisotropic,
the preferred c-axis direction will expand more
quickly, the other directions more slowly. This
will induce a more rapid deterioration of the
material in the preferred c-direction. As a
consequence, we require that the graphite be iso-
tropic, and can rewrite Eq. (2a) in terms of the
linear distortion, G,

G=_1.V=4_Vflt(1__t.>
-

(2a)

~ 2
3 3 7T (2b)
and this relation will be used hereafter.

The best values for the parameters ¢7 and v,
were found to be

o7 = (9.36 - 8.93 X 10™° T) x 10°* nvt (E > to keV)

180

GRAPHITE AND XENON BEHAVIOR

NUCLEAR APPLICATIONS & TECHNOLOGY

and
v, =-12.0+8.92x107° T %

where T is the temperature in °C, valid over the
range 400 to 800°. At 700°C, these yield ¢r = 3.1 X
10*# with a 90% double-sided confidence limit of
+0.2 X 10*2, and v, = -5.8%, with the limit +0.2.
Figure 1 shows the behavior of the linear distor-
tion as a function of fluence with temperature as
a parameter.

As pointed out in the introduction, it is neces-
sary that the graphite exclude both salt and xenon.
The first requires the graphite to have no pores
larger than ~1pu in diameter; the second, as will
be seen below, requires the graphite to have a
permeability to xenon of the order of 10~® cm?2/
sec. Clearly, as the graphite expands and visible
cracking occurs (v > +3%), these requirements
will have been lost. Lacking definitive data, we
have made the ad hoc assumption that the pore
size and permeability requirements will be main-
tained during irradiation until the time when the
original graphite volume is reattained, namely,
when ¢ = 7 as defined above.

To estimate the lifetime of an MSBR core, we
must take into account the strong dependence of T

 

 

 

 

 

 

 

 

 

 

 

+2 [
850°C /800° /750°
/ 700°/
+1
|
S
o
Z o
o 0 / / / 650
-
@
o
-
: / /
a O
© 1 600/
< / / /
w
Z
- 700
- \;/// -
-3. }
0 1 2 3
FLUENCE ® x 10722 (E >50 ke V)
Fig. 1. Graphite linear distortion as a function of

fluence at various temperatures.

VOL. 8 FEBRUARY 1970
on temperature, and the temperature of the core
graphite will depend on the fuel salt temperatures,
the heat transfer coefficient between salt and
graphite, and the gamma, beta, and neutron heat
generation in the graphite. In all core designs
which have been analyzed, the power generation
(i.e., fission rate) has been found to vary closely
as sin(rz/1) along the axial centerline, where z2/1
is the fractional height, and I is the effective total
core height. Thus, the heat generation rate in the
graphite will vary as

q= 4o + 41 sin(rz/1) , 3)

where ¢, approximates the rate due to delayed
gamma, beta, and neutron heating, and ¢, approxi-
mates the maximum prompt heating. Represent-
ing the actual graphite core prisms as cylinders
with internal radius @ and external radius b, we
can readily calculate the internal temperature
distribution assuming g is not a function of radius.
The result is’

q 2 2 log(7 /a)
T = Ta - IR' (’V -a ) + —__mg(b/a)
X [Tb - T, + 74%{. (b* - az)] , (4)

where T, and T, are the surface temperatures at
a and b, respectively, and K the graphite thermal
conductivity. Since the axial heat flow is negli-
gible, the heat, @, that must cross the graphite
surfaces per unit area becomes

__9a  _ K _ d (p? _ g2
Q“—-2+alogb/a|:Tb T“+4K(b a)]
qb K

_ oo n _ 4 2 _ 2
Qb_-_2_+blogb/a [Tb Ta+4K(b a)](5)

and if % is the heat transfer coefficient between
the salt and graphite interface, then also

Q. =h(Ta - TO)
Qb =h(Tb = TO) ’ (6)

where T o is the bulk salt temperature. The coef-
ficient % is calculated from the turbulent-flow
Dittus-Boelter equation®

04 Ko

h = 0.023 Re®® Pr 57 (7)

where
Re is the Reynolds number
Pr is the Prandtl number
K, is the salt thermal conductivity.

We are assuming the coolant channels at the outer
surfaces of the cylinder will have the same effec-
tive hydraulic radius as the interior channel of

NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8

Scott and Eatherly

GRAPHITE AND XENON BEHAVIOR

radius ¢. Finally, since the heat capacity of the
salt is a weak function of temperature, in keeping
with the sinusoidal variation of power density
along the axis, the salt temperature, T, becomes

1 T2
T0=§[Tf+T,~-(T/-T,-)cos—l—:| (8)
in which T, and 7, are the entering and exiting
temperatures, respectively, of the salt.

Equations (3) through (8) completely determine
the temperatures in the graphite. As we shall see
below, the radiation-induced strain in the graphite
along the z-axis, i.e., A2/z, is given by

Az 2 b

24 _ T

~ (—z——z—b_a)faG( Yv dr
where G(T) is the damage function of Eq. (2b). A
similar expression applies to the radial strain. A
negligible error is introduced by a_pproximating
the right-hand side by G(T), where T is the aver-
age temperature over the cross section; there-
fore,

Az

— = G(T) . 9)

Since T varies with z/l, each point along the
cylinder will have a different life, 7(T), defined
by Az/z = 0. The minimum value of these 7(T)
thus defines the time at which the cylindrical
prism should be replaced based on our criterion.

The properties of the fuel salt and graphite
which are required in the above equations are
given in Table I. The graphite is assumed to be
similar to the British Gilso-graphite, although
there are other coke sources than Gilsonite which
lead to isotropic materials.

Calculations have been made on two core con-
figurations. Case No. CC-58 is the reference
design discussed elsewhere’ in this series of

TABLE 1

Materials Properties Used in Graphite Lifetime and
Stress Calculations

 

Fuel Salt

 

Thermal conductivity, W/(cm °C)| 1.29 X 10~2
Specific heat, Wsec/(g °C) 1.36

Viscosity, cP (7.99 x 107%) exp(4342/T °K)

 

 

Graphite

 

37.63 exp(-0.7T °K)
5.52 % 10~% + 1.0 X 10°(T °C)
1.9 x 10°
0.27
(5.3 -1.45x 107> T + 1.4
x 107°T2%) x 107"

k in psi~* n/(cm® sec)
T in °C.

Thermal conductivity, W/(cm °C)
Thermal expansion, (°C)™"
Young’s modulus, psi

Poisson’s ratio

Creep constant, k

 

 

 

 

FEBRUARY 1970 181
Scott and Eatherly

papers. This design optimizes the fuel conser-
vation coefficient® with the peak power density
constrained to a value of 63 W/cm?® to prolong
graphite life. Case No. CC-24 is the identical
core scaled to a smaller geometry (half the
volume of No. CC-58) and with the constraint on
graphite lifetime raised. In both cases the fuel
conservation coefficient is 15.1 [MW (th)/kg]?. The
damaging flux in case No. CC-24 is also more
typical of the case when the core is optimized
with no constraints. The pertinent core param-
eters® are given in Table II.

TABLE II

Relevant Core Characteristics for Calculation of
Graphite Lifetime and Stresses

 

 

 

 

 

Case No. Case No.
CC-24 CC-58

Peak flux (E > 50 keV), n/(cm® sec)| 5.15 x 10** 3.2 x 10"
Salt flow per unit area, g/(cm® sec)| 1.39 x 10° 8.21 x 10°
Heat generation, delayed, W/cm® 1.16 0.71
Heat generation, prompt, W/cm® 7.17 4.39
Salt inlet temperature, °C 550 950
Salt outlet temperature, °C 700 700

 

 

\Q= 3'2 X ]Ol“
4 L\"

GRAPHITE AND XENON BEHAVIOR

 

 

CORE LIFE AT 80% PLANT FACTOR (year)

 

 

 

 

 

 

3
2
0.60 0.75 0.90
INTERNAL RADIUS, a (in cm)
Fig. 2. Core lifetime as a function of graphite prism

dimensions for cores with peak damage fluxes
of 3.2 and 5.15 x 10™ nv (E > 50 keV). In all
cases the ratio of the radii, b/a, is 6.67.

182 NUCLEAR APPLICATIONS & TECHNOLOGY

Since the salt-to-graphite ratios in the core
are determined by nuclear requirements, and the
salt flow by cooling requirements, the only vari-
able left at this point is the absolute value of the
internal radius g, which scales the size of the
graphite prisms. Figure 2 shows the lifetime of
the central graphite prism as it is affected by the
radius a; there is an obvious decrease in core life
as the radius ¢ increases and the internal graphite
temperatures climb. We note that the ratio of
fluxes in the two cases is 1.61; ata= 0.9 cm, the
corresponding reciprocal ratio of lifetimes is
1.74, the additional gain in lifetime at the lower
flux being due to decreasing graphite tempera-
tures.

The latter case has been studied in more detail

since it corresponds to the current reference
design concept. For this design the equivalent
radii are

a = 0.762 cm

b = 5.39 cm .

The associated temperature distributions for the
central core prisms are given in Fig. 3, and the
local lifetimes 7(T) as a function of z// in Fig. 4.
The life of the prism, i.e., the minimal 7(7),
occurs at 2/l = 0.55 and has a value of 4.1 years
at 80% plant factor. The entire prism will change
length as given by the integral of the right-hand
side of Eq. (9) over the length of the prism; this
is shown as a function of time in Fig. 5. The
radial distortion for various times is shown in
Fig. 6, and gives the prism a double hourglass
shape toward the end of life.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

800
—
= 700 MAXIMUM INTERNAL P
s TEMPERATURE=_ |~ =
=
P INNER AND OUTER
S / 7z é SURFACE
g / 7 TEMPERATURE
5 00 A BULK SALT
=4 TEMPERATURE
=
500
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
FRACTIONAL HEIGHT (z/1)
Fig, 3. Temperatures associated with the central gra-

phite prism as a function of vertical position
for the reference core design.

VOL. 8 FEBRUARY 1970
7(T) AT 80% PLANT FACTOR (year)

Scott and Eatherly

 

 

 

N
™™

 

 

 

 

 

 

 

 

 

 

| \ /
N —
4
0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9
FRACTIONAL HEIGHT (z/1)
Fig. 4. Local lifetime 7(T) as a function of vertical

position for the central graphite prism of the
reference core design.

The cost of replacing graphite will include both
material procurement and labor. We have esti-
mated these costs based on an industry supplying
the order of ten or more reactors. The total
operating cost associated with graphite replace-
ment would amount to ~0.2 mill/kWh for a two-
year life, or 0.10 mill/kWh for a four-year life.
(Capital investment would also be required for the
replacement equipment and is included in the
capital cost estimates of Bettis and Robertson.”)
Thus, the cost penalty associated with graphite
replacement is not a crippling one, although it is
large enough to merit considerable effort on
graphite improvement.

Assuming pyrolytic impregnation, as discussed
by McCoy etal.,’® successfully excludes xenon
from the graphite, three significant material
requirements need to be met: the graphite must
be isotropic, have entrance pore diameters
<1 u, and have a radiation stability at least as
good as the Gilso-graphite. All of these require-
ments can be met with existing graphite tech-
nology, although all have not been met in a single
graphite of the dimensions required. Unquestion-
ably, there will be production difficulties in
initially producing such a graphite, but the prob-
lems will be in process control rather than in
basic technology.

NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8

FEBRUARY 1970

GRAPHITE AND XENON BEHAVIOR

 

0

 

/

0 1 2 3 4
TIME AT 80% PLANT FACTOR (year)

AXIAL DISTORTION (%)
'

N—_— |

 

 

 

 

 

 

Fig. 5. Axial distortion of the central graphite prism

as a function of time for the reference design.

INTERNAL STRESSES IN THE GRAPHITE

In the preceding section we have tacitly as-
sumed that the thermal and radiation-induced
stresses developed during the lifetime of the
graphite are not limiting. We turn now to validate
this assumption. We again consider the central
prism as in the preceding section, and have for
the constitutive equations,

€ 2% [6; - H(ff,' + 0,)] + k¢[0i '%(01' + Ok)] + &,

where (10)
e; = strain in the ¢’th direction
o, = stress
E = Young’s modulus
u = Poisson’s ratio
B = secondary creep constant
¢ = damaging flux
dots = time derivatives
g = dG/dt = damage rate function defined from

Eq. (2b).

In cylindrical coordinates, ¢ (r, 6,2). In the
absence of externally applied stresses and be-
cause of the vanishing of ¢ at the ends of the
core prism, the z component of strain will not be
9 function of #» and # (plane strain). We shall
assume E is a function of T, but will not let 2 or
¢ vary with » and 6. Then Eq. (10) can be inte-
grated in closed form. The resulting dimensional
changes in the prism are

Az _ Aa

Z a b (11)

183
Scott and Eatherly GRAPHITE AND XENON BEHAVIOR

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0
. ; l
\ /\48 mos. O, _)kl(b (Ago-BAg1)+ %Aglt . (13)

6 mos.

\‘\ // It can also be shown that g = g(7) with an error
= \ - not exceeding 2% for the parameter values of
= " \ N \ interest.

;C:’ \ \\ 42 mos. » . The thermal stresses have the same interre-
S \ lationships as the radiation stresses, and are
= \\ \( 36 mos. . given by
g‘ -2 \Q\ L/ — 0-00) = 0, 0gd) = 00)
& ;; 18 mos. and
24 mos. £
30 mos. 0z0 = T S (T -T()] et . (14)
-3 We may substitute this result into Eq. (12) to give

0 0102 03 04 05 0.6 0.7 0.8 0.9 1.0 the maximum Stress o, (b) throughout the entire
FRACTIONAL HEIGHT (z/1) life of the central prism.

For the reference design concept, the maxi-

mum stresses also occur at the point z/I = 0.55

Fig. 6. Radial distortion of the central graphite prism and are shown in Fig. 7. The stresses reach a

as a function of vertical height for the refer- maximum at the end of life and amount to only
ence design. Times are calculated at 80% plant 490 psi. Since the isotropic graphite which is
factor, presumably to be used in the reference design

would have a tensile strength in the range 4000 to

where G is the average value of G over the cylin-
drical cross section. Hence the prism behaves
locally as though it were at the average radiation
distortion G. 500 L

The tensile stresses are maximum at the out-
side surface of the cylinder, and are given by

0,(0) =0, 0y(d) = 0,(b)

 

 

 

 

 

400
and /
0. = 7o e [Jlg gt et + 00e™ (12) /
300
with
g = Ek ¢
2(1 - p) -

 

200 , ,

We may consider the initial stress o,, to be the
thermal stresses introduced as the reactor is
brought to power, and these anneal out exponen-
tially with time because of the radiation-induced 100
creep. We are interested only in large times ¢, /

and if we set
Ag =g -g0) . /

and remember g is a linear function of time

g=27=57<1‘7>=g°+g1t’ -60
0 1 2 3 4

then TIME AT 80% PLANT FACTOR (year)

SURFACE STRESSo OR o, (psi)

 

 

 

 

 

 

 

 

 

 

Ag = A + Ag,t
& &o &1 Fig. 7. Axial or tangential surface stress at z/I = 0.55

and Eq. (12) thus takes the asymptotic form as a function of time,

184 NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8 FEBRUARY 1970
5000 psi, there appears to be no reasonable proba-
bility that these stresses can cause graphite
failure.

XENON-135 BEHAVIOR

Because of its high-neutron-absorption cross
section, it is important to keep xenon out of the
eraphite and also out of the fuel salt. The **Xe
comes from two sources, directly from the fission
of uranium and indirectly from the g decay of the

fission products tellurium and iodine. Tellurium

probably exists as a free metal which is believed
to be insoluble in molten salt and tends to migrate
to the available surfaces such as the core vessel
walls, the heat exchanger, the graphite moderator,
and even any circulating gas voids which may be
present. However, “°Te has a relatively short
half-life, <3 min, and it is conservative for our
purposes to assume that most of it decays into
iodine before it can leave the fuel system. The
iodine forms stable iodides in the molten-salt fuel
and will remain with the fuel unless steps are
taken to remove it.

Since the half-life of **° is 6.7 h, it appears
possible to process the salt stream at a relatively
slow rate to remove a large portion of the iodine
before it decays into xenon. The equivalent yield
of **Xe in the processed salt can then be repre-
sented by

U _ ,U U _ 1
YXe yxe+ V1 <1 Tp >\I+ 1) ’

where
y)l(Je= direct yield of xenon from uranium of
mass U

yU = chain yield of T from uranium of mass U

decay constant of '*°1

Al

T, = time required to process the entire salt

inventory for the removal of iodine.

It is apparent that the effective yield cannot be
reduced by iodine processing alone to below ~1.1%
for 233U fission.! As shown in Fig. 8, a relatively
high processing rate of 62 gal/min for 1500 £t * of
fuel only reduces the effective yield of 133%e to
~0.0225. One scheme for processing a side
stream uses HF for converting I” to I, and then
removing the HF and I, by purging with helium or
H,. However, it would still be necessary to re-
move additional xenon, and so stripping the fuel
salt with helium is the preferred process for re-
moving “°Xe.

Xenon is very insoluble’ in molten lithium-
beryllium fluoride and obeys Henry’s law for
gases; at 650°C the pressure coefficient of solu-
pility is 3.3 X 107° moles(Xe)/[ecm® (salt) atm].

NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8

Scott and Eatherly

GRAPHITE AND XENON BEHAVIOR

 

 

 

 

6
235y yye = 0.0024 y1 = 0.0617
233U yye = 0.0111 yy = 0.0505
5
4 —
//
233 ’,/:::::::
3

 

L/

/

0 200 400 600 800
TIME TO PROCESS COMPLETE FUEL INVENTORY (min)

 

EQUIVALENT YIELD OF '33Xe (%)

 

 

 

 

62 gal/min FOR 1500 ft® FUEL INVENTORY

 

 

 

 

 

 

1000

Fig. 8. Effect of iodine removal on 135%e yield.

Because of the relatively high xenon partial pres-
sure at very low concentration, it will tend to
leave the salt through any free surface available
before the concentration becomes high enough to
make a significant contribution to the ***Xe poi-
soning. In the MSBR, the free surfaces of impor-
tance are those associated with the voids in the
graphite and the entrained helium bubbles. The
graphite planned for MSBR use, but without sur-
face impregnation, has a bulk void volume avail-
able to xenon of ~10% and could contain a large
inventory of xenon if there is an umimpeded flow
from the salt to these voids. Almost all the xenon
poisoning in the MSBR will result from the neu-
tron absorptions in the '*°Xe within the graphite,
and so the graphite should have those properties
which keep xenon concentration in graphite low.
The concentration of xenon in the graphite is
controlled by the concentration in the salt, the
mass transfer coefficient from the salt to the
graphite, the total surface of graphite exposed to
the salt, the diffusion coefficient of xenon in the
graphite, and the void fraction available to xenon.
The mass transfer coefficient from the salt to the
graphite is very strongly influenced by the char-
acteristics of the salt flow boundary film which,
together with the area of graphite exposed to the

FEBRUARY 1970 185

 
Scott and Eatherly

salt, are determined by the heat transfer condi-
tions required to cool the graphite. However, the
effective diffusivity and available volume of the
graphite can be controlled during manufacture by
impregnating the surface with a thin layer of
material having low permeability. The base
stock graphite which would be used in the core
will probably have properties which are dictated
by radiation damage and lifetime considerations
and not those which affect xenon concentration in
the voids. However, as will be illustrated below,
a sealed surface will act as an effective barrier
to the transfer of xenon into the graphite interior
if the value for the diffusivity of the sealed sur-
face can be controlled to <~0.25 x 10~ ¢m®/sec,
and the associated void volume is 1%. As dis-
cussed by McCoy et al,” experiments have been
performed which indicate that the application of
such a sealant is feasible. With the above condi-
tions of permeability and available void volume
met, the xenon poisoning in the core becomes
controlled by its concentration in the salt and
therefore by our ability to process the salt stream
for xenon removal.

The salt will be processed for xenon removal
by injecting small bubbles of helium into the fuel
salt stream and then removing them after they
have taken up a portion of the xenon present. The
rate of transfer of xenon to the bubbles is affected
by the concentration of xenon in the salt, the total
surface area of bubbles, the mass transfer coef-
ficient of the dissolved xenon from the salt to the
entrained bubbles, and the residence time of the
bubble in contact with the salt. We can control all
of these except the mass transfer coefficient to
the bubbles, and a good value for the coefficient
is difficult to determine. A review® made of
existing data on rates of mass transfer to gas
bubbles points to reliable equations for predicting
mass transfer under certain simple conditions
involving stationary liquids and rising bubbles. It
was also found by analysis that the mass transfer
to bubbles associated with turbulent liquid in a
pipe could vary by as much as a factor of 6 de-
pending upon the characteristics assigned to the
bubble-salt interface. An experiment is under way
to evaluate this effect; at present a conservative
value of 17 X 10~ cm/sec at the lower end of the
range for the mass transfer coefficient under
turbulent conditions is being used in the calcula-
tions to estimate the xenon poisoning in the MSBR.

The surface area of a given volume of en-
trained bubbles varies as the inverse of the
spherical diameter and therefore the void volume
needed in the circulating fuel can be kept small if
the bubbles are small. Methods of injecting the
bubbles into the salt are under study and one in
particular appears promising.’ In this latter

186 NUCLEAR APPLICATIONS & TECHNOLOGY

GRAPHITE AND XENON BEHAVIOR

method, the gas is injected into the throat of an
inverted venturi which is formed in the annulus
between a teardrop shape (installed on the pipe
axis) and the pipe wall. When tested with water,
bubbles of ~0.5 mm were produced, and we ex-
pect similar results when used with molten salt.
The bubbles in the salt are allowed to recirculate
through the system several times before removal.
The solubility of xenon in molten salt is such that
with helium-to-salt volume fractions as small as
7%, the bubbles can be recirculated ~10 times be-
fore the xenon concentration in the bubbles gets
high enough to significantly affect xenon poisoning.

The final step in the xenon processing of the
salt is the removal of the bubbles, and this can be
done with a centrifugal gas separator of a type
which was developed for another circulating fuel
reactor.”® Kedl has proposed'’’*? a model for de-
scribing the migration of the noble gases in a
molten-salt reactor and for estimating the poison-
ing resulting from this migration. This model
was evaluated in a summary experiment using a
gas containing a “’Kr tracer on the MSRE during
prenuclear operation; calculated results were in
good agreement with experiment. After the MSRE
had been operating at power, the calculated re-
sults for **Xe poison were not in complete agree-
ment with experiment results; in general the
poisoning was less than expected for the case
where the circulating bubble fraction was very
small and about as expected for the larger bubble
fractions. Studies are under way to determine the
source of the difference.

Figure 9 summarizes the predicted effect on
the poison fraction of varying the volume flow of
helium into the salt stream of the MSBR for two
different graphites in the core. Poison fraction
is the neutron absorptions in Xe relative to
fissile fuel absorptions. The values of the physi-
cal constants used are considered the most likely

135

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

> [MOLTEN-SALT BREEDER REACTOR-
., \SINGLE FLUID 1000 Mw(e)
x4 MASS TRANSFER COEFFICIENT T0
= BUBBLES, 17 X 10~® cm/sec
= 3 Xe DIFFUSION COEFFICIENT IN
S \ GRAPHITE, AS SHOWN
e N FRACTION OF BUBBLES REMOVED
- 2 S PER CYCLE, 10%
3 \\ BUBBLE SIZE, 0.5 mm AVERAGE
S 1 = D >0.25 x 10~% cm?/sec

P — —
0 D = 0.25 x 10"® cm?/sec
o 1 2 3 4 5 6 7 8 9 10
He INJECTION RATE (scfm)

Fig. 9. Effect of helium injection rate on 13Xe poison

fraction,

VOL. 8 FEBRUARY 1970
for the MSBR. The upper curve is for an uncoated
base graphite where the main resistance to xenon
transfer is in the fluid flow boundary and not in
the graphite. The lower curve represents the ad-
vantage gained by surface impregnation of the
base graphite to a xenon permeability of ~0.25 X
10~ cm’/sec. One important difference between
the two curves is in the volume of helium which
must be handled to get the desired results. The
conditions of the lower curve are such that only
low helium gas flows are needed to reach the
target of 0.5% Xe poison fraction. However, it
should be noted that some helium processing is
required even if a good coating is obtained.

OFF-GAS SYSTEM

The removal of **°Xe from the circulating fuel
by sparging with helium will also remove many
other fission products including the other noble
gases together with some of the relatively noble
metal fission products. Altogether these repre-
sent a significant radiation and heat source. It is
the purpose of the off-gas removal and disposal
system to safely collect and store these materials
to permit recycling the helium. The present ver-
sion of the off-gas system removes the helium
and certain fission products from the circulating
salt stream, and de-entrains any salt which
may be carried along; separates the noble gas
from the helium, and holds the '**Xe outside the
reactor for a total of 48 h, and then reinjects a

Scott and Eatherly

GRAPHITE AND XENON BEHAVIOR

major portion of the gas stream into the fuel salt
for recycle. The remainder of the stream is fur-
ther processed for the removal of the longer-
lived fission products so that the clean helium can
be used in the helium purge stream (i.e., along the
fuel pump shaft). This system is described below.

Figure 10 is a flow diagram which illustrates
the considerations involved in the off-gas system
design. Ten percent of the total salt flow from
the primary salt pump is bypassed through the
133Xe removal system, which includes the gas
separator and the bubble generator. About 9 scfm
of helium with fission products is removed from
the salt as it passes through the bubble separator.
This gas, together with some salt, goes to the
entrainment separator where the salt is removed
and returned to the primary circulating system
through the pump expansion tank. The remaining
gas, together with ~2 scim of purge gas from the
pump expansion tank, then goes to the particle
trap and 1-h decay tank. The pipe lines from the
bubble separator to and including the entrainment
separator are cooled by entrained primary salt;
in addition a secondary coolant salt flowing in an
annulus around the section of gas line from the
entrainment separator to the particle trap will
remove the heat released to the pipe wall. At the
particle trap and 1-h decay tank the entrained
fission-product particles are removed and allow
the short-lived xenon and Kkrypton isotopes to
decay for 1h before entering the absorption
charcoal beds where the xenon is held up for 47 h.

 

 

   
    
  
  

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2 scfm_ CLEAN PURGE He
PRIMARY
SALT PUMP 11 scfm_ |
ENTRAINMENT
SEPARATOR He SUPPLY
18 MW ACCUMULATOR
3.2 MW,
-
1
| PRIMARY 'finguée |
- SALT
N SEPARATOR CHARCOAL
| REACTOR BUBBLE BED
l, | VESSEL { GENERATOR
| COMPRESSOR l 9 scfm
0.27
| 85 vy
90 day (Y1 TRI'%UQND
HOLDUP ' REMOVAL
CHARCOAL
SED SYSTEM
Fig. 10. Molten-Salt Breeder Reactor Off-Gas System.
NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8 FEBRUARY 1970 | 187
Scott and Eatherly

The decay of the noble gases in the 1-h volume
delay tank reduces the heat load on the head end of
the charcoal bed by a factor of 6, making it easier
to keep the temperature of the charcoal low
enough to maintain an acceptable adsorption ef-
ficiency. In the decay tank the fission-product
particles will be removed from the gas stream
by washing with a liquid coolant which will also
cool the gas and all surfaces of the tank walls.
This coolant stream will be pumped through a
heat exchanger where the heat will be transferred
to a closed steam system which in turn dumps its
heat to a natural-draft air-cooled condenser.
About 18 MW of energy is released in the decay
tank. One of the primary objectives in the design
of this system is to make the heat-rejection sys-
tem independent of external power needs such that
a power failure will not jeopardize the system.
The coolant stream will be processed to keep the
concentration of fission products below the satu-
ration level to prevent deposition on the walls.

The gas goes from the decay tank to the char-
coal adsorption beds where the xenon is held up
for 47 h and the krypton is held for 4 h. This
holdup is by adsorption of these heavy gases on
the charcoal which does not delay the helium
carrier gas. The total delay of 48 h in the two
holdup sections is sufficient to reduce the “*°Xe
concentration returning to the circulating salt
system to 23% of its value when it left the salt.
About 3.2 MW are released in this adsorption
holdup system.

After leaving the charcoal bed the flow splits
into two streams. One system of ~2 scfm goes to
the low-flow charcoal beds where the xenon is
held for 90 days, after Wthh most of the activity
has decayed. However, ®*Kr (10. 4-year half-life)
and tritium (12-year half life) remain in the gas
stream and are removed by appropriate collector
beds using hot titanium beds for tritium removal
and molecular sieves for °°Kr removal. The
resulting clean helium is then compressed into
the accumulator for reuse in the cover-gas sys-
tem and the pump shaft purge.

The other stream of ~9 scfm is simply com-
pressed and returned to the bubble generator for
recycle to the "°’Xe removal system.

SUMMARY AND CONCLUSIONS

The previous sections indicate that although
radiation damage to the graphite and “**Xe poison-
ing of the core introduce design complexities into
the reactor plant, accommodation of these factors
requires neither significant new technology nor
excessive cost penalties. We specifically con-
clude the following.

188 NUCLEAR APPLICATIONS & TECHNOLOGY

GRAPHITE AND XENON BEHAVIOR

1. Radiation damage in graphite will limit
core life, but existent materials are adequate to
provide a core life of the order of four to five
years without excessively penalizing reactor per-
formance. Although a specific grade of graphite
that fulfills all our requirements is not currently
available commercially, only minor extensions
of existing technology, including the experience
factor, are required to produce such a material.

2. Under the conditions existing in the central
region of the core, stresses induced internally in
the graphite are small and do not limit the per-
missible graphite exposure based on the present
design.

3. Although not specifically discussed here,
we have had suff1c1ent experience with pyrolytlc
carbon impregnation® of graphite to feel confident
the process can be developed for use with MSBR
graphite. The times and process conditions used
to impregnate small graphite samples are practi-
cal for scale-up, and in at least one case the per-
meability of the graphite was not significantly
affected by neutron irradiation.

4. Xenon-135 can be removed successfully
from the circulating salt by sparging with helium
bubbles. However, several parameters such as
salt-to-gas mass transfer coefficients are only
approximately known. Nevertheless, it appears
that the parameter values are in a range where
the sparging will be effective.

5. Helium bubble insertion and extraction have
been demonstrated successfully in an air/water
system. Although previous experience with mol-
ten-salt systems has confirmed the ability to
transform fluid flow results on water to a molten-
salt system, it remains to be demonstrated that
the helium-bubble/salt system performance can
be predicted using results from the air/water
system.

6. An off-gas system has been conceptually
designed which permits removal of the fission
products from the helium, recirculation of the
cleaned helium, and control of the fission-product
heat. Such a system appears practical, based on
the capabilities of present technology.

ACKNOWLEDGMENTS

This research was sponsored by the U.S. Atomic
Energy Commission under contract with the Union
Carbide Corporation. Oak Ridge National Laboratory is
operated by the Union Carbide Corporation Nuclear Di-
vision for the U.S. Atomic Energy Commission.

VOL. 8 FEBRUARY 1970
REFERENCES

1. H. S. CARSLAW and J. C. JAEGER, Conduction of
Heat in Solids, 2nd ed., Oxford, Clarendon Press (1959).

2. W. H. McADAMS, Heat Tvransmission, 3rd ed.,
McGraw-Hill, New York (1954).

3. E. S. BETTIS and R. C. ROBERTSON, ‘‘The Design
and Performance Features of a Single-Fluid Molten-Salt
Breeder Reactor,”’ Nucl. Appl. Tech., 8, 190 (1970),

4, A, M. PERRY, ‘‘Reactor Physics and Fuel Cycle
Analyses,’’ Nucl. Appl. Tech., 8, 208 (1970).

5. H. C. McCOY, W. H. COOCK, R. E. GEHLBACK,
J. R. WEIR, C. R. KENNEDY, C. E. SESSIONS, R. L.
BEATTY, A. P. LITMAN, and J. W, KOGER, ‘*Materials
for Molten Salt Reactors,”’ Nucl. Appl. Tech., 8, 156
(1970).

6. C. B. BIGHAM, A. OKAZAKI, and W. H. WALKER,
““The Direct Yield of Xe-135 in the Fission of U-233,
U-235, Pu-239, and Pu-241, Tvans. Am. Nucl. Soc.,
8, 11 (1965).

NUCLEAR APPLICATIONS & TECHNOLOGY VOL. 8

Scott and Eatherly

GRAPHITE AND XENON BEHAVIOR

7. G. M. WATSON, R. B. EVANS, III, W. R. GRIMES,
and N. V. SMITH, ‘‘Solubility of Noble Gases in Molten
Fluorides,”’ J. Chem. Eng. Data, 1, 2, 285 (1962).

8. F. N. PEEBLES, ‘“Removal of Xenon-135 from
Circulating Fuel Salt of the MSBR by Mass Transfer to
Helium Bubbles,”” ORNL-TM-2245, Oak Ridge National
Laboratory (1968).

9. R. J. KEDL, ‘‘Bubble Generator,”” MSR Program
Semiannual Progress Report, February 28, 1969, ORNL-
4396, Oak Ridge National Laboratory (1969).

10. R. H. CHAPMAN, ‘‘HRE-2 Design Manual,”” ORNL-
TM-348, p. 112, Oak Ridge National Laboratory (1964).

11. R. J. KEDL and A. HOUTZEEL, ‘‘Development of a
Model for Computing ¥Xe Migration in the MSRE
Graphite,”’ ORNL-TM-4069, Oak Ridge National Labora-
tory (1967).

12. R. J. KEDL, “A Model for Computing the Migration
of Very Short-Lived Noble Gases into MSRE Graphite,”’
ORNL-TM-1810, Oak Ridge National Laboratory (1967).

FEBRUARY 1970 189