ORNL-CF-58-2-46.txt

UNCLASSIFIED

OAK RIDGE NATIONAL LABORATORY

QOperated By

UNION CARBIDE NUCLEAR COMPANY 0 R N L

(44
POST OFFICE BOX X (ENTRAL F"-ES NUMBER

OAK RIDGE, TENNESSEE

X822

 

........

 

58-2-46

6S

 

 

 

 

DATE: February 5, 1958 COPY NO.

SUBJECT: A MOLTEN SALT NATURAL CONVECTION EXTERNAL TRANSMITTAL
REACTOR SYSTEM AUTHORIZED

TO: Iisted Distribution

FROM: F. B. Romie, AMERTCAN-STANDARD,

and B. W. Kinyon

£
Q

. NOTICE

This document contains information of a preliminary

nature and was prepared primarily for internal use

G at the Oak Ridge MNationel L.aboratory. It is subject

e : tc revision or correction and thersfore does not
‘ represent a final report.

UNCLASSIFIED w5 oot

 
 

 

 

 

 

comtractor of the Commissicn to the extent that such employee or contractor grepares, handles

 

 

This report was prepared os an acceunt of Gevernment sponsored work, Neither the United States,

nor the Cbmmission,. oy any perébfl acting on behalt of the Commissiont

A, Makes any worransy or representation, express or implied, with respect to the acecuracy,
completeness, o usefulness of the informotion contained in this report, or thet the use of
eny information, opperatus, methed, or process disclosed in this report may not infringe
privately owned rights; or

B. Assumes any [iabilities with respeci to the use of, or for domages resulting from the use of
6ny informetion, eppurates, method, or process disclosed in this report.

As used in the above, “‘person acting or beholf of the Commission® inciwdes any employee or

or distributes, or provides cccess to, any infeemation pursuent to his employmant or coniract

with the Cemmission. i "

 

 

 
UNCLASSIFIED
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o Abstract

Fuel-salt volumes external to the core of a molten-salt reactor are calculated
for a system in which the fuel salt circulstes through the core and primary exchanger
by free convection. In the calculation of these volumes, the exchanger heights above
the core top range from 5 to 20 ft. Coolants considered for the primary exchanger
are g second molten salt and helium. External fuel holdup is found tc be the same
with either cooclant. Two sets of terminal temperatures are selected for the helium.
The first combination permits steam generation at 850 psia, 90001?e The second set
is selected for a closed gas turbine cycle with an llOOOF turbine inlet temperature.
Specific power (thermal kw/kg 235) is found to be about 900 kw/kg, based on initial,
clean conditions and a 60 Mw (thermal) output. A specific power of 1275 kw/kg is

estimated for a forced convection system of the same rating.

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UNCLASSIFIED
1.
A MOLTEN SALT NATURAL CONVECTION REACTOR SYSTEM

F. E. Romie¥

B. W. Kinyon
Introduction

One of the problems of a circulating-ligquid-fuel reactor is the provision of

reliable, long-lived fuel circulating pumps. This problem is eliminated for a
gsystem in which the fuel is circulated through the primary exchanger and reactor
core by natural convection. The advantages of omitting the circulating pump and
its attendant problems of meintenance and replacement are purchased at the price
of ificreased fuel-salt volume in the primary exchanger and in the convection
risers. There are applications for a resctor system in which the premium placed
on reliability and ease of maintenance could meke the convection system attractive.
The purpose of this report is to investigate a free-convection reactor system in

crder to afford a basis for assessing its merits.

Seleefiion of Reactor Conditions

A schematic representation of the reactor and primary heat exchanger is shown
in Figure (1). The fuel salt returned to the core after cooling in the heat exchanger
enters the bottom of the reactor sphere and leaves through the top. The temperature
of the fuel salt entering the heat exchanger is specified to be 1225°F, & temperature
which available corrosion information indicates is consistent with long-term life

for the system. With this temperature and a thermal output of 60 Mw, it is possible

to consider selection of steam turbine conditions of conventionsl plants. Thus tem-

perature conditions in the fluids passing through the primary exchanger have been

 

* On loan from AMERICAN-STANDARD, Atomic Energy Division, February 1957 to
February 1958.

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selected to permit generation of 850 psia, QOOOF stesm. For a thermal output of 60 Mw
these steam conditions would give a generator output of about 22 Mw (37% efficiency).

The IQESOF maximum sglt temperature also allows consideration of a closed gas
turbine cyecle using helium at a turbine inlet temperature of 1100°F. For this tem-
perature and the postulated cycle parameters summsrized in Table I# the turbine
output would be about 19 Mw (30.8% efficiency).

The importance of desirable nuclear properties for the fuel salt takes precedence
over the thermal properties in the selection of & fuel salt. For this reason a mixture
of lithium and beryllium fluoride salts is selected. For such mixtures the'viscosity
increases and the melting point decreases with increasing beryllium content. Mixture
130 (63% TiF, 36% BeFé;:vl% UF) on a mole % basis), with a melting point of 850°F,
has been selected as giving a reasonable combination of these two properties. The

pertinent physicel properties of this salt are given in Table II.

Table I.

ASSUMED GAS TURBINE CYCLE PARAMETERS

Turbine inlet temperature 11000F
Compressor inlet temperature (both stages) 1000F
Compressor adisbatic efficiency 874
Turbine adiabatic efficiency 89%
Regenerator effectiveness 884
Compressor pressure ratio (1.52/stage) 2.3
Pressure losses, I AP %
P

Calculated Performance - Helium

Thermal efficiency 30.8%
Helium temperature at salt exchanger inlet 6760F
Regenerator NTU T-33

Cycle output decreases 2.1% for each
percentage point increase in AP
P

 

__________ 4 * Attainment of the helium compressor and turbine efficiencies postulsted in
hid Table I has yet to be demonstrated.

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The secondary fluid in the primary exchanger is limited by compatibility con-

(13

siderations

to either a gas or a molten salt.

Both possibilities are treated.

In the case of the molten salt cooling the heat from the reactor would be trans-

ferred from the fuel salt, to the coolant salt, tc sodium, to water. A similar

system is described ir Reference 1.

Good compatibility with the fuel salt and low

fusion temperature {6SOOF) are the principal reasons for selection of Mixture 84

(35% 1iF, 27% NaF, 38% BeF,) as the coolant salt.

Table TIT.

PHYSICAL PROPERTIES OF SALT MIXTURES 130 AND 84

Mixture 130 Mixture 8h
Unit heat capacity, Btu/Ib°F 0.62 C.59
Thermal conductivity, Btu/hr-ft-CF 3.5 3,2
Density, Ib/ft° (t - OF) 136.4-0.0121t  1%9-0.0142t
Viscosity, Ib/hr-ft, 1000°F: 36.6 29,1
1200°F: 19.2 1k,2
Fusion temperature, °p 850 6L0

 

The use of a gas as a cocolant offers the advantage that, in the case of a
steam cyele, it would be the only fluid intermediate between the fuel salt and
the steam. In the case of the gas turbine éycle there Would be no fluid inter-
mediate between the working fluid and fuel salt. The use of gas rather than salt
also eliminates heaters for the coolant salt and sodium circuits and decreases the
number of drain tanks required for the system. Gases sultable as a coolant are
hydrogen; helium and nitrogen. Helium is specifically considered in the following
because of its good heat transfer properties and nuclear and chemical inertness.

Hydrogen would be preferable from the point of view of heat transfer and pumping

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considerations, while nitrogen, due to its higher molecular weight, would be attrac-
tive for use with a gas turbine cycle. The low moclecular weights of hydrogen and
helium require the use of a large number of compressor and turbine stages. Nitrogen

would offer the advantage of using presently-developed compressor and turbine equipment.

Optimum Riser Conditions

The hydrostatic differefntisl hesd causing the fueli-sglt flow ig proportional
to the product of the temperature change of the fuel salt and the height of the
exchanger above the reactor. The pressure drop available to the exchanger is thg
hydrostatic head minus frictional losses oecurring in the riser pipes. These fric-
tional losses decresse rapidly with increasing riser diameter, but at the expense
of an increased salt hcldup volume in the riser. If the fuel-salt mass rate and
temperature change are both fixed, then, for the attainment cf a specified pressure
drop available to the exchanger, there is one combination of riser diameter and
height of exchanger for which the salt volume in the riser system will be a minimum.
The exiStefice of this minimum is illustrated in Figure (2), in which the variation
of salt volume in the riser and of riser diameter are shown as a function of exchanger
height. Tablé ITT summarizes 8 set of optimm riser conditions used in this reporta
Flgure (2) and Table IV are based on the riser friction data given in Table III.

The frictional losses in the riser are found to be determined primarily by the
expansion and contraction losses and are insensitive to wall friction. Thus replace-
ment of a single riser by two sete of risers having an equal height and total cross
sectional ares will make essentially the same pressure drop available to the exchanger
at the same riser holdup volume. This fact can be of use in decreasing the mechanical
rigidity of the piping and indicates that the use of two exchangers with separate

risers in place of one will not require an increased fuel holdup volume in the risers.

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G Table TII.
FRICTION IOSSES IN THE RISERS

Expansion plus contraction loss

on entering and leaving core L.5 ft

on entering and leaving exchanger headers 1.5 £t
Equivalent length of pipe added for bend~loss allowance 8 ft
Flow friction factor .02
Tength of piping 2H + 17 ft

Priction loss asscciated with risers
=(o.02[8+17+23)+3 we
D (%)2 D Z2gp

Hydrostatic head for flow

= Ofl@f (fi + %)

Salt volume in riser

= xpF (17 + 2H
i

 

 

Part Load Operation

 

For a given core-riser-exchanger system the flow rate of the fuel salt is
determined in terms of the difference in fuel salt temperature in the two legs
of the riser and the flow frietion characteristic of the system. The part load
operation of a given system is shown in Figure (3). For the molten salt reactor
the average of the core inlet and exit temperatures is effectively a constant in-
dependent of heat output. Thus, if at rated conditions the salt temperature
entering the exchanger (leaving the core) were 12250F, the salt temperature leav-

g ing the core at half-load would be, based on Figure (3}, ll870F, and the temperature

of the salt returned to the core would be 1059OF compared to lOOOOF at full-losad.

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500

% LASSIFIED

Temperature change of fuel

salt, OF

Exchanger height, ft
Salt volume in risers, ft5
Diameter of riser, Tt

Pressure drop across
exchanger, lb/ £t2

Salt velocity in riser, ft/sec

200

65
1.78

1.5

OPTTMUM RISER CONDITIONS

200

10

19.5
1.656
13.4

1.74

Q = 60 Mw
a = 0.,0121

200

105.5

1.537
27.8

2.02

Table IV.

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2

 

r£=-Op

225

55
1.602

Tl

106)"'

225

10
68
1.525%

15.6

1.79

225

20

89
1.418
31.8

2.10

250

L8

1.497
8.8

1.69

250

10
58.5
1.41

17.8

1.90

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Salt-Cooled Heat Exchanger

 

S The términal temperatures selected for the coolant salt are 8750F at inlet

 

to the primary exchanger and 10250F at exit. These two temperatures remain un-

changed for all salt-cocled exchanger designs considered. The BTSOF entrance

temperature is 25°F above the fusion temperature of the fuel galt and QBSOF above
the fusion temperature of the coolant salt. Heat transfer data used in the cal-

culations are given in Table V.

Table V.
PRIMARY HEAT TRANSFER AND FRICTION DATA
Fuel Salt
Nusselt modulus for fuel salt* k.0
Friction factor 64/Re
Entrance and exit losses for salt, velocity heads 1.5
Thermal resistance per unit tube length 1
NurziR8 = 0.0227
Coolant Salt
Thermal resistance per unit tube length¥* 0.0032
Exchanger Tubes
Thermal conductivity, Btu/hr-ft-°F 1.5
Wall thickness, in. 0.059
Thermal resistance per unit length:
0.42 in. ID tubes 0.00272
0.634 in. ID tubes 0.00189
¥ (QGraetz modulus is less than 5¢0 for all designs.

*¥% This value is estimated attainable without excessive pressure loss by
appropriate shell-side baffling. (Doubling this resistance would in-
crease the over-all resistance by only 11%.) '

Figures (4) and (5) summarize the results of exchanger calculations. In-
spection of these figures shows that increasing the height of t@e exchanger above
the reactor has no effect on the total length of exphangef tubing required and

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thus no effect on the fuel-salt volume within the exchanger tubes; however, in-
creésing the exchanger height leads to an increased length of the exchanger and
a decrease in the number of tubes in the exchanger. Thus the tube bundle diameter
is decreased and, consequently; the volume of fuel salt contained in the exchanger
header. However, the net result on salt volume external to the reactor is a slight
increase with increasing exchanger height. This increase is due to the increased
sglt volume in the risers. If the number of exchanger tubes were the constant
parametef rather than the tube diameter, then an increase in exchanger height
would lead to a decrease in total holdup volume and a smaller diameter tube.

Decreasing the internal diameter of the exchanger tubes from 0.63L4 to 0.42 in.
produces, for a 10 ft height of the exchanger, about a 23% decrease in external
holdup volume, but at the expense of a 2.3-fold increase in the number of tubes.
Variation of the fuel-salt temperature drop over the range from 200 to 250°F PTro-
duces a relatively minor effect on the exchanger designs.

The low thermal resistance of the coolant salt relative to that of the
laminarly-flowing fuel salt results in a small temperature difference hetween
the exchanger shell and tubes. Thus thermal stresses due to differential expan-
sion of t;bes and shell should prove small enough to permit & straight tube design

despite the rather large temperature difference between the two salts.

Gas-Cocled Exchangers for Steam Cycle

For helium cooling of the primary exchanger, the terminal temperatures of
the helium were fixed at BSOOF for inlet to the exchanger and lOZSOF for outlet.
The drop in fuel-salt temperature wae held at 225°F. The 850°F inlet temperature
insures that the fuel salt will not freeze ifi the tubes, and the 1025°F exit tem-

perature is consistent with the generation of 900°F steam. The exchanger designs,

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which are summarized in Table VI, are based on the use of 0.634 in. ID tubing
with copper fins. Dimensions of the finned tubes (Table VII) were scaled-up from
descriptions in Reference 2, from which the flow fractiéfi and heat transfer data
wére also obtained.

With a single-pass cross flow exchanger, the salt flowing in the row of tubes
first contacted by the entering helium would leave the exchanger at a'considerably
lower temperature than that in the last row of tubes. With laminar flow, the flow
resistance is directly proportional to the viscosity. Thus the salt velocity in
the first row would be less than in the last row, and the mean temperature of the
first row conseqfiently lower than would be calculated cn the basis of uniform flow
in all tubes. The effects leading to non-uniform salt flow are mutually aggravating,
with the result that a single-pass cross flow exchanger is unacceptable. A multi-
pass exchanger is thus required if a cross flow exchanger design is to be used. A
counterflow exchanger has the optimum configuration, both from the point of view
of obtaining uniform flow distribution and minimizing the required over-all conduc-
tance of the exchanger. |

The over-all dimensions given for the helium-cooled exchanger gorrespond to
a three-helium-pass cross flow exchanger. The column labeled "depth" refers to the
distance traveled by the helium in one pass (i.e., distance from front to rear of
tube array}. The "width" refers to the transverse dimension of the exchangers.

The fractional pressure loss, AP/P, listed in the table is based on & helium
pressure level at the exchanger of 100 psia. The fractional pressure loss is in-
versely proporticnal to the square of the pressure level. Thus, for example, if
the pressure of the helium were increased to 200 psia, the AP/P values shown ifi

Table VI would be quartered (without change in any other entries in the table).

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......... B Table VI.
GAS-COOIED EXCHANGER - STEAM CYCIE
0.634k ID Finned Tubing

Total Iength AP Tube Salt Frontal Area
Helium of tubing P Depth Volume per pass
Re No. £L (100 psia He)  ft £t £2
500 k8,300 .0000758 264 106 1060
1000 43,000 000438 469 95.1 530
2000 39,200 .0026 .855 86 265
LOCo 36,200 0177 1.582 79.3 132

_ Iength per Header Total Salt
Helium Number of tube Volume Yolume Width
Re No. Tubes £t £43 ££3 £t
5 ft high: riser vol - 55 cu ft, AP = 7.k psf

500 5520 8.75 2l 185 410
1000 5230 8.22 23 173 194
2000 5000 7.8k o2 163 102
LO00 4800 7.5h 21 155 5k

10 £t high: riser vol - 68 cu ft, APsalt = 15.6 psf
500 3800 12.7 16.7 191 250
1000 3600 12.0 15.8 179 133
2000 3440 11.4 15.1 169 69.4
4000 3310 10,9 14.5 162 36.2

20 £t high: riser vol - 89 cu ft, AP 1y = 31.8 psf
500 2660 18.2 11.7 207 175
1000 2520 17.1 11.1 195 93,2
2000 2k10 16.3 10.6 186 48.8
4000 2320 15.6 10.2 178 25.4

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Table VII.

DESCRIPTION OF FINNED TUBE*

 

Internal tube diasmeter, in. 0.63h
Wall thickness, in. 0.059
Tube material INOR-8
Fin material Copper
Fin area/total area 0.876
Fin thickness, in. 0.03%9
Tube spacing in plane normal to gas flow, in. | 1.745
Tube spacing parallel to flow, in. | 1.4%0
Heat transfer area/unit volume, 1/ft 76.2

* Geometrically similar to surface CF-8.72 (c) in Reference 2.

 

The fraction, s, of the plant output used in pumping the helium through the

exchanger is given by the equation,

in which 7y 1s the specific heat rate, T +the mean gas temperature (OR), AT
the temperature change of the gas, Mo the blower efficiency, and np the plant

efficiency. For a blower efficiency of 80% and plant efficiency of 37%, one obtains,

AP
S - 1097“"?"‘

Thus if 2% (8 = .02) of the generator output is used to blow helium through the
exchanger, the value of AP/P would have to be 0.00187.

In order to facilitate discussion of the calculated results summarized in
Table VI, it is convenient to define the reference heat exchanger described in

Table VIII.

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Table VIII.

e
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REFERENCE GAS~COOLED HEAT EXCHANGER

AP/P 0.00187
Pregsure level, psia 100
Exchanger height, ft 10

Tube length, ft 11.5
Number of tubes 3240
Iength of tubes, ft 8
Total salt volume, ft3 172
Depth, in. 9 %
Helium Reynolds Modulus 1750

 

If the pressure level of the helium were increased from 100 to 200 péia and
the exchanger altered in order to maintain AP/P constant, then the exchanger length
would be decreased from 78 ft to 48 ft and the depth increassed from 0.77 to 1.2 ft.
Accompanying changes in the number of tubes, tube length and salt holdup volume
would be relatively negligible.

If the power expended in pumping the 100 psia helium through the exchanger
were incressed from 2% to 4% of generator output, (i.e., %? = 0.00%74), the depth
and length of the exchanger would change to 0.93 ft and 61 ft, respectively. Corre-
sponding changes in salt holdup volume, number of tubes and tube length would be
relatively negligible.

For a 10 ft height of the exchanger above the reactor, the calculated salt
fiolume external to the core is 172 cu ft for the reference exchanger. This can
be compared to & fuel-salt volume of 160 cu ft for the salt-cooled exchanger of

the same height, using exchanger tubes of the same internal diameter. If the

reactor core diameter is taken as 8 ft, corresponding to a fuel-salt volume of

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268 cu ft, then these figures indicate that gas cooling can be used with a fuel~

salt volume increase of 12 cu ft in a total volume (exclusive of volume in expan-
sion tank) of 440 cu ft.

The major uncertainty in comparing the externsl fuel-salt volumes for the
gas-céoled and salt-cooled exchangers lies in the uncertainty in the salt volume
which should be attributed to the exchanger headers. TFor the present calculations
the header volume hae been assumed to be 0.00438 £t per 0.634 in. ID tube for both
the salt-cooled and gas-cocled exchangers. A more detailed design study is required
to fix the header volume more accurately.

Use of a smaller tube diameter for the gas-cooled exchanger would result in
a8 decreased fuel-sglt holdup volume at the expense of an increased number of
shorter tubes. The effect of using a 0.42 in. ID tube instead of the 0.63L4 in.
tube used in the calcuiations shcould parallel the behavior previously shown for
the salt-cooled exchanger.

A possible configuration for the three-pass exchanger is shown in Figure (6).
The configuratién permits attachment of the riser pipes to the headers with suffi-
cient flexibility to minimize thermael stresses and provides a cylindrical containment
for the pressurized helium. Use of two such cylinders to contain the reference
exchanger would result in twe exchangers, each 19 1/2 ft long. If the helium pres-

sure were ipcreased to 200 psia, the length of each would be reduced to 12 ft.

Gas-Cooled Exchangers for Gas Turbine Cycle

The '‘major problem encountered in designs of exchangers for use with a gas
turbine cycle is the low temperature of the helium entering the helium~-fuel-salt
exchanger. For the cycle parameters given in Table I, this temperature is 676°F,
& temperature lThOF below the fusion temperature of the salt. Exploratory calcu-

lations indicate that this low temperature would freeze the salt if an exchanger

uging the finned surface described in Table VII were used.

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For minimum fuel-sglt holdup in the exchanger tubes, the optimum solution

...............

to the problem of meintaining a wall surface temperature above the freezing point
is found in the use of a counterflow exchanger, in which the gas-side thermal
resistance decreases in the direction of gas flow .in such a msnner that the wall
temperature does not drop below a prescribed value. In the csse of a longitudi-
nelly finned tube, the gas-side thermal resistance variation required could be
effected by increasing the fin height along the tube until the full fin height
is gttained. For the remainder of the tube length, the fin height would be un-
changed and the tube wall temperature would increase above the prescribed minimum.
The exchanger calculations presented for the gas turbine cycle presume the use of
such an exchanger with a minimum wall-fuel-salt interfacial temperature of 9OOOF.
Results of the exchanger calculations are presented in Figuref?}-e The
abscissa, ¢ , is the ratio of the sum of the gas and wall thermal resistances
per unit tube length to the fuel-salt thermal resistance per unit length,
= Rw + Rg . The gas-side thermal resistance is defined as that which is
R

obtainedswith the unaitered finned tubing. Using the thermal resistance data

¢

presented in Table V, it is easily shown that the gas-side thermal resistances

OF-ft-hr
Btu

8 l/z-fold increase in gggwside registance corresponds to an increase of approx=-

corresponding to ¢ values of 0.25 and 1.5 are 0.003785 and 0.0321 . This
imately 40% in external fuel-salt holdup volume, number of tubes and length per tube.
No data are presently available defining the hesat transfer and flow friction
characteristics of tube bundles using longitudinally finned tubes. Thus the
exchanger designs are incomplete. Hafiever, the number of tubes, tube length and
total external salt volume can be estimated with reasonsble accuracy if it is
assumed, as seems likely, that the same value of ¢ can be realized with longitu-

dinally finned tubes as with circumferentially finned tubes. For the steam cycle

R

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UNCLASSIFIED
15.

helium-cocled reference exchanger the value of ¢ 1s C,47. Using this value of ¢
in Figure (7} gives, for an exchanger 10 ft above the reactor, a total salt holdup
volume of 160 £t°, s total length of exchanger tubing of 36,200 ft, with 3300 tubes

11 ft long. A possible configuration for the exchanger is shown in Figure (8).

Comparison of Cooling Means

Teble IX gives a compariscn of the three cooling systems considered for the
conditions listed at the head of the table. The difference in external fuel-salt
holdup volume among the three systems is not large and is therefore not a determin-
ing factor in the selection of the coolant medium for the reactor.

In genersl, the helium-cooled exchangers have much larger over-all volumes
than the salt-cooled exchangers. This i;creased bulk is caused by the larger
spacing required by the finned tubes and also by the large volume reguired by the
helium headers. An increase in helium pressure will decrease the helium header
volune and also should decresse the fuel-salt header volume. The fipper limit on
helium pressure is prcbably set by consideration of the effects of a tube rupture
on the fuel-salt system. It is of interest to note that other gas-cooled reactor
stfidiesf apparently not limited by similsr considerations; have recommended gas
pressures as high as 1000 psia. The 100 psia pressure specificslly considered
in these calculsations is probably conservstively low.

Table IX.

COMPARISON OF COOLING MEANS
Based on 60 Mw, 10 ft high exchanger

 

 

- OF
0.634 ID tubes, AT, o7 = 225°F
Method of Cooling No. of Tubes Total Tube Iength, £t Vol Outside Core, ftE
Salt-cooled for steam 3150 36,000 161
generation
Helium~-cooled for 3240 37,200 172
- steam generation
e Helium-cooled for gas 3300 36,000 16%
turbine cycle
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Comparison with Forced Convecticon System

 

........

Design of a forced convection resctor system (1) has led to an estimation
of 0.56 cu ft of fuel salt externsl to the reactor core per thermal megswatt.
For the free convection system, using C.634% ID tubes in the primary exchanger,
the externsl volume is about 160 cu ft for 650 Mw, giving a specific volume of
2.67 cu ft/Mw. Using these numbers and initial-clean critical mass data (5),
the fuel inventory for 60 Mw minimizes, for both the free and foreed convection
systems, at a core diameter of about 8 ft. At.this diameter the specific powvers
{kw thermal/kg 235) are 895 and 1275 kw/kg for the free and forced convection
systems, respectively. The free convection system is thus estimated to require
Lo% more inventory than the forced convection system.

Addition of a small amount of thorium to the core salt would produce some
breeding and a longer time intervsl between fuel additions. One-quarter percent

thorium requires sbout an 87% increase in fuel inventory at a core diameter of & ft.

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References

1. Kinyon, B. W., and Romie, F. E., "Two Power Qeneration Systems for
s Molten Fluoride Reactor"”, To be presented at the Nuclear Engineer-
ing and Science Conference of the 1958 Nuclear Congress (March),
Chicago, Illinois

2. Kays, W. M., and london, A. L., "Compact Heat Exchangers", The
National Press, Palo Altc, California (1955)

3. Personal communiecaticn from J. T. Roberts, ORNL

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UNCLASSIFIED
ORNL~LR-DWG 27943

EXPANSION DOME

    

HEAT EXCHANGER
1COOLANT 1
DOWNCOMER

Fig.1. Schematic of Natural Convection Reactor.

18 el 020
UNC-LR-DWG 27944

 

 

 

19 et
UNC-LR-DWG. 27945

 

 

LR £LERER
20 68 ok r{j p:i Fa
FIGURE 4

UNC-LR-DWG. 27946

 

.........

e

2 I0 12 20
HEIGHT OF EXCHANGER —FT iy poo
21 W Ued
 

50

40

30

20

FIGURE 5

UNC-LR-DWG. 27947

 

TEMP CHANGE OF FUEL SALT

 

250 °F

 

 

225

 

200

 

 

 

 

TOTAL LENGTH OF TUBING x 103 FT
(BOTH 0.42 & 0.634 1.D. TUBING)

 

250

634 1.D, ,/
//
/

s
?zoo

 

;/
=

\

 

EXCHANGER TUBE
LENGTH,FT _ 250

- 4225
- ~200

 

 

 

 

 

 

 

 

0

20

EXCHANGER HEIGHT,FT
¢e

R

  

Figure 6

SALT HEADER

 

Three-Pass Exchanger

UNCLASSIFIED
GRNL-LR- DWG 27948

~PRESSURE SHELL
FIGURE 7

 

UNC-LR-DWG. 27949

 

-
"
|
mm

 

  
UNCLASSIFIED
ORNL-LR-DWG 27950

ac

A K‘“ji =
ifii ) éj

 

Fig. 8. Schematic of Counterflow Salt to Helium Exchanger.