ORNL-1721.txt

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ORNL-1721
This document consists of 107 pages.

Copy 7,,,2,::4 231 copies. Series A.

Contract No. W<7405-eng-26

AIRCRAFT REACTOR ENGINEERING DIVISION

ORNL AIRCRAFT NUCLEAR POWER PLANT DESIGNS

A. P. Fraas
W

. Savolainen

A,

May 1954

DATE ISSUED

ROV 10 1854

OAK RIDGE NATIONAL LABORATORY
Operated by
CARBIDE AND CARBON CHEMICALS COMPANY
A Division of Union Carbide and Carben Corporation
Post Office Box P
Oak Ridge, Tennessee

 

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FOREWORD

Formal Air Force interest in nuclear propulsion for aircraft dates from October 1944,
when the head of the Power Plant Loboratory (WPAFB), Col. D. J. Keirn, upproached
Dr. Vannevar Bush on the subject. Subsequent to that and other discussions, the NEPA
group was formed in 1946, The NEPA group moved to Qak Ridge in 1947, and by 1948,
ORNL had begun to provide assistance in research and testing. The ORNL effort gradu-
ally expanded, and the ORNL-ANP General Design Group was formed in the spring of
1950 to help guide the program and to evaluate and make use of the information being
obtained.

Four years of work at ORNL on the design of aircraft nuclear power plants have dis-
closed much of interest. In a project so complex and so varied it is inevitable that many
of these points should escape the attention of nearly all but those immediately concerned
or be forgotten in the welter of information produced., Some of this material is buried in
ANP quarterly reports, and much has never been formally reported. '

Many reactor designs have been prepdared, but each design has represented an isclated
design study, and the issues have been much confused by variations in the assumptions
made in the course of each reactor design. This report is intended to provide a critical
evaluation of the more promising reactors on the basis of ¢ common, reasonable set of
design conditions and assumptions,

 
 

CONTENTS
FOREWORD & v evvvveeenennnenss, e BT v
SUMMARY i it ittt i e et n e st s e s e s e s e e e e |
PART I, DESIGN CONSIDERATIONS
MILITARY REQUIREMENT S & i it ittt it s tr e s ssn o s e oo neosanononsasos 2
PROPULSION SYSTEM CHARACTERISTICS & . i vttt i i i it s aastannsnssnsssansss . 2
Vapor-Cycle Compressor-Jet . o v v i i i st it onavsonaesnnonesnsesensssessss 3
Gas-Cycle Compressor-Jet . . i v it o it ittt et it in oo onsetnonesenosoasonssans 3
T U O e v vt v it it i e e e e e e e e 3
Specific Thrust and Specific Heat Consumption .o v v it it iv v vunas b s es e 4
Chemical Fuel as a Supplementary Heat Source ... ..... S e e e e ae i e 4
REACTOR TYPES .. ittt ittt et st osannsnssnan C et e e et 4
AIRCRAFT PERFORMANCE ............ ettt e e e e e e e 6
FfHects of Reactor Design on Aircraft Gross Weight ... .. o i ittt i it it enennsn 6
Effect of Chemical Fuel Augmentation .............. f et et e e n et e s H
SHIELDING ..ttt i i i ittt i isnenn e e n et e e e e e e e e e 16
Units of Radiation Dose Measurements . v o v v vt vt t ot nnnsansnnssesssonssonssnsas 16
Permissible Dose Rate for Crew . . . o it i it it it it it i s st s ot s asmo s nonononon s 17
Radiation Damage to Organic Materials and Activation of Structure . .o o4 L. e 17
Ground-Handling and Maintenance Problems . .o v v v v v s i i ittt s s i e i8
Shield Weight « v v it v it et et it e e s st e ottt saaneaears 21
NUCLEAR PROPERTIES & it it ettt sttt asassnssetsnsastossoenssasesess 33
Moderating and Reflecting Materials + . v o v v i it i i i it i i e 35
Effect of Moderating Material onDesign . . s i i vt i it o s ettt nesnnsas 36
Reflector-Moderated Reactor .. v v v i v ot o s v ot v s et s sososvososnsnassossosasssas 38
REACTOR CONTROL 4t v ittt i it i et snasssoaasananssanssosstoisanosnarssss 39
MATERIALS o v v vt vt et es v snonnncsosn et e e e e e e e ey 40
SHUCTURE & « 4+ s e o s o s v v s o o oo s o e o ssosooeonnsstnsasasanorsansassseesnsseass 41
Solid Fuel Elements o v v v e s o v ot o v o v am o s o s 0 s o s asosonusnsesncensnsosscnsaas 42
High-Temperature Liquid Coolants and Fuel Carriers . .. ..o v v i i i oo iin s, 48
HEAT REMOVAL i et it e e et e e 51
TEMPERATURE GRADIENTS AND THERMAL STRESSES . . o vt it i s i e ittt e e 55
TEMPERATURE DISTRIBUTION IN CIRCULATING-FUEL REACTORS ..o vt v it it iien s 59
PART Il. REACTOR STUDIES
COMPARISON OF REACTOR AND CYCLE TYPES ... .o it ittt et i oo 63
REACTOR, HEAT EXCHANGER, AND SHIELD ARRANGEMENTS . .. . ittt enneenn 64
Shield and Heat Exchanger Designs . . . o v v i i it i vt i ien s i s s i et o ancsan 67
Reactor Core Configurations . v v v v v vt s e v s n v s s et sasaccsnsnsas S e e s 76

Vil

 
DETAILED DESIGNS OF REACTORS

 

--------------- R a8 4 8 & B B F s 8 6 a s s SN e oo 79

Sodium-Cooled Solid-Fuel-Element Reactor . .. v ii it i i it in it it et te v in e ensennn 79
Circulating-Fuel Aircraft Reactor Experiment . ., .,..... S et s e et et e e e 83
Fluid-Moderated Circulating-Fuel Reactor .. i i ittt i ittt ittt it ettt ensanans 83
Reflector-Moderated Circulating-Fuel Reactor .. ., i i e i i i ii it i ii it i it e eanaans 87
SECONDARY FLUID SYSTEM 1. i ittt ittt ettt it ettt onssansnnsaatonens 93
MAJOR DEVELOPMENT PROBLEMS L. it it it ittt it e sttt aosttanansens 96

vili

 
 

ORNL AIRCRAFT NUCLEAR POWER PLANT DESIGNS

A, P, Fraas

A, W. Savolainen

SUMMARY

The detailed design of an aircraft nuclear power
plant poses an extraordinarily ditficult set of
problems. V*2:3 |t will be found implicit in this
report that the problems are so intimately inter-
related that no one problem can be considered
independently of the others; yet each problem is
sufficiently complex in itself to be confusing. In
an effort to correlate the work that has been done,
a tentative sef of military requirements for nuclear-
powered aircraft is presented first and accepted
as axiomatic, . The types of propulsion system that
might be used are discussed next, and the turbojet
engine is shown to be the most promising. Aircraft
performance considerations are then presented on
the basis of a representative power plant, and the
shield data used are validated in a section on
shielding. It is shown in these sections that the
reactor should be capable of a power density in
the reactor core of at least 1 kw/em® and, prefer-
ably, 5 kw/ecm®, and it should operate at a suf-
ficiently high temperature to provide a turbine ajr
inlet temperature of at least 1140°F for the turbojet
engines. The effects of nuclear considerations

 

M The Lexington Project, Nuclear-Powered Fiight,
LEXP-1 (Sept. 30, 1948). fi

2R\epor’r of the Technical Advisory Board to the Tech-
nical Committee of the ANP Program, ANP-52 (Aug. 4,
1950).

37. A. Sims, Final Status Report of the Fairchild
NEPA Project, NEPA-1830 (no date).

 

on the size, shape, and composition of the reactor
core are presented, and in the light of the preceding
presentation, possible combinations of materials
and the limitations on the materials are discussed.
The effects of the physical properties of several
representative coolants on the maximum power
density obtainable from a given solid-fuel-element
structure is determined on the basis of a consistent
set of assumptions. Design limitations imposed by
temperature - distribution and thermal stress are
also examined. |

From the data presented in the section on air-
craft performance and in the sections on nuclear
materials and heat removal considerations, it is
shown that the reactor types having the most
promising development potential and the greatest
adaptability to meet the wide variety of military
requirements are those in which a liquid removes
heat from the reactor core at temperatures of
1500°F or higher. Designs fer several high-temper-
ature reactors are presented, and their advantages
and disadvantages are discussed.

The problems involved are too complex to permit
anything approaching an Aristotelian proof to
support a choice of reactor type, but it is hoped
that this report will convey something more than
an oppreciation for the various decisions and
compromises that led first to the circulating-
Hluoride-fuel reactor and then to the design of the
reflector-moderated reactor type recently chosen
as the main line of development at ORNL. '
PART I. DESIGN CONSIDERATIONS

MILITARY REGQUIREMENTS

The potential applications of nuclear-powered
gircraft to the several types of Air Force mission
are quite varied. Robot aircraft, ram-jet and rocket
missiles, and unmanned large nuclear-powered tugs
towing small manned craft have been suggested
as a means of avoiding the shielding problem in-
volved in the use of nuclear power. As will be
shown later, it is probable that even in missiles
some shielding would be required because of
difficulties thot would otherwise arise from radi-
ation damage and rodiation heating.? Furthermore,
from the information available, it appears that
these applications, while possibly important, either
would not justify the large development expense
of the nuclear power plant required or, in the case
of nuclear rockets, would represent such an ex-
trapolation of existing experience as to be very
long-range projects, A number of different missions
for monned aircraft with shielded reactors are,
however, of such crucial importance as to more
than justify the development cost of the nuclear
power plant, All these missions involve strategic
bombing, Studies by Air Force contractors have
indicated that the aircraft should be capable of
operation (1) at sea level and o speed of approxi-
mately Mach 0.9, or (2} at 45,000 ft at Mach 1.5,
or (3) at 65,000 ft at about Mach 0.9. A plane of
vnlimited range that could fly any one or, even
better, two or three of these missions promises to
be extremely valuable if available by 1965, |In
addition to the strategic-bombing application, there
are important requirements for lower speed (Mach
0.5 to 0.6), manned aircraft, such as radar picket
ships and patrol bombers. The problems associ-
ated with supplying a beach head a substantial
distance from the nearest advance base indicate
that a logistics-carrier airplane of unlimited range
would also be of considerable value.

In re-examining these requirements, it is seen
that a nuclear power plant of sufficiently high
performance to satisfy the most difficult of the
design conditions, namely, manned aircraft flight
at Mach 1.5 and 45,000 ft, would be able to take
care of any of the other requirements, except those
involving rocket missiles., Because of the rapid

 

4R, W. Bussard, Reactor Sci. Technol., TID-2011, 79-
170 (1953).

rate of advance of ceronautical technology and
because of the inherently long period of time re-
quired to develop a novel power plant of such
exceptional performance, it appears that develop-
menta! efforts should, if at all possible, be centered
on a power plant of sufficiently promising develop-
mental potential fo meet the design condition of
Mach 1.5 at 45,000 ft either with or without the
use of chemical fuel for thrust augmentation under
take-off and high-speed flight conditions. [t has
heen on this premise that work at ORNL has
proceeded since the summer of 1950,

PROPULSION SYSTEM CHARACTERISTICS

Several types of propulsion system well svited
for use with manned aircraft are adaptable to the
use of nuclear power as a heat source for the
thermodynamic cycle on which they operate. One
of these is the turbopropeller system in which a
steam or gas turbine is employed to drive a con-
ventional aircraft propeller, with heat being added
to the thermodynamic cycle between the compressor
and the turbine. A second is the compressor-jet
system, a binary cycle in which a steam or gas
turbine is employed to drive a low-pressure-ratio
air compressor. The air from the compressor is
heated in the condenser or cooler by the turbine
working fluid and then expanded through a nozzle
to produce thrust. A third system, the turbojet,
employs a gas-turbine cycle. Inthis system enough
energy is removed from the air passing through the
The balance of
the expansion of the air is allowed to take place

turbine to drive the compressor.

through o nozzle to produce a relatively large
thrust per pound of air handled. A fourth system,
the ram-jet, will work well only at flight speeds
above Mach 2.0, because it depends upon the ram
effect of the air entering the engine air inlet duct;
the ram effect provides the compression portion of
the thermodynamic cycle. Heot is added after
compression and the air is allowed to expand
through a jet nozzle to produce thrust. Because
it eliminates the relatively heavy and complicated
parts associaoted with the compressor and turbine,
the ram-jet system appears, on the surface, to be
much the simplest mechanically, but in practice,
serious complications arise because any given
unit will work well only in the very narrow range
of flight speeds for which it was designed. |t
should be nofed that each of these four systems
operates on a thermodynamic cycle that involves
an adiabatic compression, followed by addition of
heat at constant pressure, and then an adicbatic
expansion.

Qf the four types of propulsion system cited,
only the compressor-jet and the turbojet look prom-
ising for the applications envisioned. The turbo-
propeller system is handicapped by the poor aero-
dynomic performance of propellers obove high
subsonic speeds and by the very serious problems
associated with the high blade stresses inherent
in such designs. The ram-jet power plant is use-
fess for take-off and landing ond is so sensitive
to speed and altitude that it does not look prom-
ising for manned aircraft.

Vapor-Cycle Compressor-Jet

The wide-spread use of vapor cycles has di-
rected attention to water as a working fluid for the
thermodynomic cycle of an aircraft power plant.
The principal difficulty associated with such
a power plant is the size, weight, and drog as-
sociated with the condenser. in attempting to
establish the proportions of such o power plant,
it soon became evident that only by going te high
temperatures and pressures and by using the cycle
in conjunction with a compressor-jet engine to give
a binary cycle could o reasonably promising set of
performance characteristics be obtained.® By
superimposing the water-vapor cycle on a com-
pres sor-jet cycle, the power generated in the steam
turbine could be used to drive the air compressor,
while the condenser that would serve as the heat
dump for the steam cycle could also serve to heat
the air of the compressor-jet cycle. With this
arrangement, the air pressure drop across the
condenser could be kept from imposing an intoler-
able drag penalty on the airplane.

Vapor cycles essentially similar to the water-vapor
cycle have been proposed which use mercury,®
sodium,” or rubidium as the working fluid. These
fluids make possible much lower operating pres-

 

SA. P. Fraas and G. Cohen, Basic Performomce Char-
acteristics of the Steam Turbine-Compressor-Jdet Ajreraft

Propulsion Cycle, ORNL-1255 {(May 14, 1952).

6A. Dean and 5. Naokazate, Invesfigotion of a Mercury
Yapor Power Plant for Nuclear Propulsion of Aircraft,

NAA-SR-110 (Mar. 21, 1951).

Ty, Schwartz, [nvestigation of a Sodium Vopor Com
pressor Jet for Nuclear Propulsion of Ajrcraft, NAA-SR-
134 {(June 25, 1953).

sures than could be used with water af any particu-
lar temperature level. Unfortunately, the weight
of the mercury required per unit of power output
for the mercury-vapor system appears to be too
high,® while the sodium-vapor system must be
operated at a temperature well above that feasible
for iron-chrome-nickel alloys.”

Gas-Cycle Compressoar- et

A somewhat similar system has aolso been con-
sidered which would use helium as the working
fluid with a closed-cycle gas turbine.® Helium
could be compressed, passed through the raactor,
expanded through o turbine, directed through a heat
exchonger to reject its heat to the gir stream of
the compressor jet, ond returned to the helium
The exiro power obtained from the
helium turbine, over und above that required to

COMPIressor..

drive the helium compressor, would be employed
to drive the air compressor of the compressorsjet
cycle. This system would have the advantage of
using helium to cool the reactor and thus would
avoid any form of corrosion of materials in the
reactor. '

Turbajet

Several cycles thot use air us the thermodynamic
working fluid have been proposed. The first of
these would employ the redctor to heat the air
directly by diverting it from the compressor through
the reactor before directing it to the turbine of the
turbojet engine.® With this arrangement the only
large heat exchanger in the system would be the
reactor core, because, with an open cycle, no
bulky condenser or cooler would be required.

A versatile variant of the turbojet system. is
based on o high-temperature liquid-cooled reactor
that could serve as the heat source for not enly a
turbojet but for any of the other propulsion systems
menticned, that is, turbopropeller, compressor-jet,
or ram-jet. Versatility would be obtained by com-
pletely separating the air that would serve as the
working fluid of the thermodynamic eycle from the
reactor and by using a good heat transfer fluid to
corry the heat from the reactor to a heat exchanger
placed ot a convenient position in the propulsion
system. While heat exchangers would be required
with systems of this type, they could be kept

 

B4 Schwartz, An Anolysis of Ineri Gas Cooled Re-
actors for Applicotion to Supersonic Nuclear Aircrofs,

NAA-SR-111 (Sept. 8, 1952).
relatively small because they would operate ot a
high temperature with superior heat
mediums.

transfer

Specific Thrust and Specific Heat Consumption

In evaluating the merits of any particular pro-
pulsion system, it is convenient to work in ferms
of specific thrust and specific heat consumption
because the size and the weight of the power plant
depend on these two parameters, The higher the
specific thrust in pounds per pound of air handled,
and the lower the specific heat consumption in
Btu per pound of thrust, the smaller and lighter the
power plant will be, The most important factor
that offects these two parameters is the peak
temperature of the working fluid in the thermo-
dynamic cycle.?:? in the binary cycles, such as
the supercritical-water and helium cycles, the peak
temperature in the air portion of the cycle is also
a very important factor. A comprehensive pres-
entation of the effect of temperature on specific
thrust and heat consumption can be found in the
report of the Technical Advisory Board,? which
shows that the specific thrust is dependent mainly
on the peak temperature of the thermodynamic cycle,
irrespective of whether a compressor-jet or a
turbojet is employed. This is a very important
conclusion, since it indicates that compressor-jets
and turbojets give substantially the same perform-
ance for the same design conditions, except insofar
as the weight and drag of the machinery required
is concerned.

Chemical Fuel as a Supplementary Heat Source

The use of chemical fuel as a supplementary
heat source has important implications. The
foremost among these is that the chemical fuel
could be used to sustain flight in the event of a
nondestructive reactor failure. Another very im-
portant application would be the use of chemical
fuel for warmup and check-out work when operation
of the reactor would present radiation hazards to
ground personnel. Yet another important possi-
bility would be the use of chemical fuel for inter-
burning to raise the air temperature just ahead of
the turbine in the turbojet engine or for afterburning
following the turbine. Either arrangement could be
used to obtain increases in thrust of as much as

 

%A. P. Fraas, Effects of Major Parameters on the
Il:’ée;—f{c;rmnce of Turbojet Engines, ANP-57 (Jan. 24,
51).

100% with little increase in the weight of the
machinery required. Such arrangements would be
most attractive to meet take-off and landing or
high-speed requirements. The use of interburning
or afterburning would not be practical with the
vapor or helium cycles because the low pressure
ratio of a compressor-jet engine makes it inherently
insensitive to the addition of extra heat from a
chemical-fuel burner. Similarly, the large pressure
drop through the direct-air-cycle reactor would
make the air cycle less responsive to the addition
of heat from a chemical-fuel burner than a high-
temperature-liquid  turbojet system would be.
While separate engines operating on chemical fuel
only might be employed, a lighter power plant and
a lower drog installation should be obtainable by
the addition of burner equipment to the nuclear
engines.

REACTOR TYPES

Each of the various types of propulsion system
described in the previous section could be coupled
to one or more of a wide variety of reactor types.
The most promising of the reactor types can be
classified, as in Table 1, on the basis of the form
of the fuel, the manner in which the moderator is
introduced, and the type of fluid passing through
the reactor core. The materials considered for
each design are also given in Table 1, together
with the type of propulsion system to which the
design is best adapted. References to the studies
of these reactor types are given. The only reactor
types for which studies have not been made have
been the boiling homogeneous reactor and the
stationary-fuel-element liquid-fuel reactor cooled
by either a boiling liquid or a gas. Studies were
not made of these types because, at present, there
are no known combinations of materials that would
give good performance in these reactors.

Many factors influence the selection of a reactor
type because many different requirements must be
satisfied. The various limitations imposed on the
reactor design by aircraft requirements, nuclear
and heat transfer considerations, materials prob-
lems, etc., are discussed in the following sections.
The information brought out in this way is then
applied to a critical examination of detailed de-
signs for reactors representative of the more
promising types.
TABLE 1. AIRCRAFT REACTOR TYPES

 

 

FLUHD FLOWING

PREFERRED TYPE

 

FFERE E
REACTOR TYPE FORM OF MODERATOR THROUGH REACTOR | OF PROPULSION SYSTEM REFERENCES
Stationary Solid fuel (sintered UO2 and Circulating H,O Supercriticalewater— 10
fuel stainless steel in a stainless compressorsjet
steel-clad compact, graphite- NaOH High-temperature liquid— 11
UOZ' SiC-UOz, cermets) turbojet
Stationary Ligquid coolant (Liz High-temperature liquid— 1,2,12
{Be, BeC, C, Be, C) Na, Pb, Bi, fused turbojet
: fluorides)
Boiling coolant Sedium=vdapor—compressor- 7
{Na) jet
Gas coclant Helium, gas turbine, 8
(air, helium) compressor-jet
Direct-air-cycle turbojet 1,2,3
Liguid fuel (static fluorides Circulating NaOH High~temperoture ligquid—~ Not reported
in tubes) turbojet
Stationary Liguid coolant High«temperoture liquid— 13,14
(Be, BeO, C, BezC) {Na, Pb, Bi) turbojet
Boiling coolant Not studied
Gos coolant Not studied
Circulating Homogeneous {fuel dispersed or Boiling Neot studied
fuel dissolved in liquid moderator) Nonboiling NaOH-U02 slurry High-temperature liquid— 15
turbojet
Li7GH-N00H°U02 High-temperature liquid—~ 15,16
solution turbojet
Separate moderator Solid (Be, BeO, ) Fused fluorides Highstemperature liquid— 17,18
turbojet
U-Bi High-temperature liquid— 19
turbojet
Liguid {HZO, MNaOH, Fused fluorides High-temperature liquid— 17,18

 

 

NaOD, Li’0D)

 

 

turbojet

 

 
AIRCRAFT PERFORMANCE

Quite a number of different approaches have been
made to the problem of determining the feasibility
of nuclear aircraft. Most of the NEPA studies
were devoted to fairly detailed designs for a few
particular aircraft to meet certain specified con-
ditions. Both the Lexington Committee and the
Technical Advisory Board did some parametric
survey work, but, because of the limited time and
information available, there were many questions
left unanswered. North American Aviation, Inc.,
followed the same general approach as that used
by the Technical Advisory Boord, but again, be-
cause of the limited information available, their
survey was The Boeing Airplane
Company has done a fair amount of parameiric
survey work, but the bulk of that published has
been devoted to the supercritical-water cycle,

The design gross weight of an airplane is a good
indication of its feasibility partly because a high

incomplete,

gross weight with a low payload indicates a
marginal aircraft, and partly because it is doubtful
whether o craft of more than 500,000-1b gross
weight would be tactically useful if it could carry
only a small payload. Further, the costs of con-
struction, operation, and maintenance of aircraft
are directly proportional to gross weight.

Any difficulty that required for its solution a
small increase in component weight over the value
assumed for design purposes would require a large
compensatory increase in gross weight, Therefore
it is important to know the effects on aircraft
gross weight of the key reactor design conditions,

 

]ONuclear Development Associates, Inc., The Super~

critical Water Reactor, ORNL-1177 (Feb. 1, 1952).

”K. Cohen, Circulating Moderator-Coolant Reactor
for Subsonic Aircroft, HKF-112 (Aug. 29, 1951).

120 B, Eilis (ed.), Preliminary Feasibility Report
for the ARE Experiment, Y-F5-15 (Aug. 1950).

IsR. W, Schroeder, ANP Quar. Prog. Rep. Mar, 10,
1951, ANP-60, p. 28.

YR, C. Briant et al,, ANP Quar. Prog. Rep. Dec. 10,
1950, ORNL-919, p. 22!

15¢, Cohen, Momogeneous Reactor for Subsonic Aire

craft, HKF-109 (Dec. 15, 1950).

Yw. B. Cotirell and C. B. Mills, Regarding Homogene-
ous Aircraft Reactors, Y-F26-29 (Jan., 29, 1952).

7y, B, Cottrell, Reactor Program of the Aircroft
Nuclear Prapulsion Project, ORNL-1234 (June 2, 1952).

18A. P. Fraas, C. B. Mills, and A. D, Callihan, ANP
Quar. Prog. Rep. Mar. 10, 1953, ORNL-1515, p. 41,

WK. Cohen, Circulating Fue! Reactor for Subsonic
Adrcraft, HKF-111 (June 1, 1951).

namely, temperature, powser density, and radiation
doses inside and outside the crew compartment.

Effects of Reactor Design on Aircralt Gross Weight

A parametric survey29 of airplane gross weight
was carried out by using the quite complete set of
shield-wesight data prepared in the course of the
1953 Summer Shielding Session2! and the turbojet-
engine performance and weight data given in a
recent Wright Aeronautical Corporation report.22
The shield-weight charts are reprinted here as
Figs. 11 to 15 in the section on ‘‘Shielding.”
These charts constitute the only consistent set of
shield-weight data availoble for a wide range of
reactor powers and degrees of shield division.
The degree of shield division is a function of the
location of the shield material. The more divided
the shield, the heavier is the crew shield and the
lighter the reactor shield. The shield-weight data
are for shields made up primarily of layers of lead
and water. The reactor shields of Figs. 11 to 15
were ‘‘engineered’’ for reflector-moderated circu-
lating-fuel reactors to include weight allowances
for reactor, heat exchanger, pressure shell, struc-
ture, headers, ducts, and pumps. As will be shown
in the latter part of the section on ‘‘Shielding,”’
the total shield weights given are representative
for most reactor types, except air- or gas-cooled
reactors, for which the large veids infroduced in
the shields by ducts and heoders would cause
major increcses in shield weight in comparison
with the values given. The Wright data for turbojet-
engine weight are representative of the propulsion
machinery weight required for the most promising
types of propulsion system.

A set of tables was prepared from the reactor
design dato to focilitate solution of the basic
equation for aircraft gross weight. Studies have
shown that aver-cll power plant performance is not
too sensitive to either the compressor pressure
ratio or the pressure drop from the compressor to
the turbine provided the pressure drop does not

 

20A. P. Fraas and B. M. Wilnher, Effecis of Aircroft
Reactor Design Conditions on Aircraft Gross Weight,
ORNL CF.54-2-185 {May 21, 1954),

212 b, Blizard and H. Goldstein (eds.), Report of

the 1953 Summer Shielding Session, ORNL-1575 (June
14, 1954).

22R. A. Looas, H. Reese, Jr,, and W, C. Sturtevant,
Nuclear Propulsion System Design Anclysis Incorpo-

rating @ Circulating Fue! Reacter, WAD-1800, Parts |
and 11 {(Jan, 1954).
exceed 10% of the obsolute pressure at the com-
pressor outlet.?® Hence the engine compression
ratic was taken as &:1 aond the pressure drop from
the compressor to the turbine was taken as 10%
of the absolute pressure at the compressor outlet,
with one-half of this considered as chargeable to

the radiators. The turbojet-engine data were taken

largely from the Wright report. The specific thrust
and the specific heat consumption were taken from

Figs. 1X-1 through 1X-12,22 the engine, compressor, -

and turbine weight were taken from Fig. 1-19, and
the engine air flow from Fig. {18, Engine nacelle
drag was taken from Fig. 67 of ANP-57,° except
that 50% submergence of the nacelles in the
fuselage was assumed, The weight of the engine
tailpipe, cowling, and support structure was taken
as 25% of the compressor and turbine weight. The
total weight of the NaK pumps, lines, and pump-
drive equipment was calculated from the estimafes
given in ORNL-1515"8 1o be 38 Ib/Mw. The radi-
ator cores were designed to give a turbine air inlet
temperature of 1140°F with a 1500°F peak NaK tem-
perature and an air pressure drop across the radi-
ator core equal to 5% of the compressor outlet
pressure, The radiator size and specific weight
were determined by extrapolation of the experi-
mental curves in ORNL-150923 for a tube-and-fin
core employing 15 nickel fins per inch. These data
were combined with the turbojet-engine data to
obtain the propulsion machinery weight, and then
the installed weight of the propulsion machinery
and the reactor power output as functions of thrust
for various flight conditions were determined. The
results of these calculations are presented in
Tables 2 and 3.

The basic equation used to relate aircraft gross
weight to the weight of the aircraft structure, the
useful load, the shield weight, and the weight of
the propulsion machinery was the same as that
used by the Technical Advisory Board, North
American Aviation, and Boeing:

Wg = Wst + UL + Wsh + me ,
where
Wg = gross weight, Ib,
W_, = structural weight (including landing
gear), Ib,
UL = useful load, Ib,

 

23w, S, Farmer ef al., Preliminary Design and Per-
furmance) of Sodium=to-Air Radiators, ORNL-1509 {Aug.
26, 1953).

W., = shield weight (reactor shield ond crew
shield), Ib,
Wom = propulsion machinery weight, Ib,

The weight of the structure was taken as 30% of
the gross weight. While the value would probably
be closer to 25% for subsonic aircraft (except for
gircraft using power plants with low specific
thrust, such as the supercritical-water cycle), the
value used seemed representative ond adequate
for the purposes of this analysis,

The solution for aircraft gross weight was ob-
tained graphically by preparing charts such as
Fig. 1. The weight of the propulsion machinery
plus reactor and shield that could be carried by
an airplane after providing for structural weight
and useful load was plotted agoinst gross weight
to give o family of steeply sloping paraliel straight
fines., The weight of the shield and the propulsion
machinery required for each of a series of gross
weights was then plotted on the same coordinates,
the aircraft gross weight being taken os the product
of the thrust and the lift-drag ratio. The solution
for the gross weight is defined by the intersection
of the curve for the total power plant weight re-
quired with the line defining the power plant
weight that could be carried with a porticular
vseful load,

The liftedrag (L/D) ratio estimated for each
flight design condition would not be the optimum
lift-drag ratio obtainable with the airplane becouse
take-off, landing, ond climb requirements would
necessitate wing loadings lower than those for
minimum drag. The L/D values used are given in
Table 4, These L/D ratios are for the airplane
configuration without nacelles, an allowance for
nacelle drag having been deducted from the specific
thrust given in Table 2. Thus the L/D ratio with
nocelles would be lower than that indicated, par-
ticularly at high Mach numbers.

The useful load was considered as including the
crew, radar equipment, armoment, bomb load, and
other such items. Since the shield weights used
were for a dose rate of 1 ¢/hr in the crew compart-
ment, the useful lood can alsc be construed to
include any extra crew shielding required to reduce
the crew dose to less than 1 r/hr, For the purposes
of the study, a useful load of 30,000 1b was
selected as typical.

The aircraft gross weights obtained were then
plotted against dose rate at 50 ft from the center
of the reacter (af locations cther than in line with
TABLE 2, CALCULATIONS FOR POWER PLANT SPECIFIC QUTPUT

Compressor Pressure Ratio — 6:1

Ratic of Radiator OQutlet Pressure to Inlet Pressure = 0,90

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

(3600°\d h ‘ .
d b c d e { g=f — |- h i=— i R=1i+]
\3413/e ¢
] Specific e . . . Propulsion
Turbine Specific Turbajet Engine NaK System
Mach Altitude Inlet Specific Thrusi Less Heat Consumption Installed Weight Weight Machinery
Thrust Nacelle Weight
No. (1) Temperature (Ib-sec/Ib) Drag Btu/sec-lb kw /b Ib-sec/Ib ib /b (“;‘/]b of (1b/1b of
CF) (Ib-sec/1b) of thrust of thrust of air of thrust thrust) thrust)
0.6 Sea level 1140 25.7 25.2 6.22 6.69 15.25 0.605 0.494 1.099
1240 30.7 30.2 6.07 6.5 15.0 0.496 0.481 0.977
1340 35.5 35.0 6.04 6.46 14.81 0.424 0.478 0.902
0.4 35,000 1140 40.8 40.3 5.35 571 56.6 1.405 0.532 1.937
1240 45 44,5 5.54 5.91 55.9 1.255 0.550 1.805
1340 48.3 47.8 5.6 5.97 55.1 1.154 0.555 1.709
0.9 Sea level 1140 19.6 18.6 6.86 7.62 12,05 0.648 0.549 ¥.197
1240 24.8 23.8 6.73 7.40 11.89 0.500 0.533 1.033
1349 26.3 28.3 6.55 7.15 11.73 c.414 0.515 0.929
1540 37.5 36.5 6.35 6.88 11.40 0,313 0.496 0.809
0.9 35,6000 1140 35.8 34.8 5.63 6.1 43.5 1.250 0.555 1.805
12490 49 39 5.75 6.22 43,0 1.103 0.5¢5 1.668
13490 43,5 42,5 5.8 5.26 42.4 0.998 0.570 1.568
1540 50.8 49.8 5.85 6.29 41.4 0.831 (.572 1.403
1.5 35,000 1140 24,5 20.0 6.30 8.14 24,6 1.231 0.635 1.866
1240 28.5 24.0 6.26 7.34 24.3 1.010 0.612 1.622
1340 33 28.5 6,20 7.57 24.0 0.843 0.5¢90 1.433
1540 40.5 36.0 6.16 7.32 23,6 0.656 0.571% 1.227
1.5 45,000 1140 24.5 20.0 6.30 g.14 28,6 1,979 0.748 2.727
1240 28.5 24.0 6.26 7.84 39.0 1.625 0,720 2.345
1340 33 28.5 6.20 7.57 38.6 1.353 0.696 2.049
1540 40.5 36.0 6,15 7.32 38.0 1.055 0.673 1.728

 

 

 

*j = £{0.038 NaK piumbing weight + specific radiator weight).
TABLE 3. PROPULSION MACHINERY WEIGHT AND REACTOR OUTPUT FOR VARIOUS THRUST REQUIREMENTS

Ratio of Radiator Qutiet Pressure to Inlet Pressure = 0,90

Compressor Pressure Ratio = 6:1

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

TURBINE THRUST (ib)
MACH | ALTITUDE INLET 10,000 15,000 20,000 25,000 30,000 40,000 50,000 60,000
NOG. {ft) TEMPERATURE
o W o » prx 1 W r W P W P W P W F W P W P
{“F) pm pm pm pm pm pm pm pm
0.6 Saa level 1140 10,99 § 66.9 | 16.48 | 100.4| 21.95] 133.8 27.45 | 167.21 32.951 200,71} 43,95 267.6| 54.95 [334.5| 65.90 | 401.4
1240 9.77 1 65,1 | 14,65 ] 97.61 19.53] 130.2 24.40 1 162.8} 29,30} 195.3 1 39.10| 260.4| 48.85 {325.5] 58,60 390.6
1340 9.02 | 64.6 | 13.52 ] 96,9 18.03; 129,21} 22,521 161,5| 27.05] 193.8 | 36.G5| 258.4{ 45,05 {323.0| 54,05} 387.6
0.6 35,000 1140 19.37 § 57.1 | 29,05 | 85.6] 38.70; 114.2| 48.40 | 142.8] 58.05] 171.3} 77.45| 228.4| 96.80 [ 285.5| 116.2| 342.6
1240 18,05¢ 59.1 { 27.05 | 88.6| 36.10] 118.2 | 45.10 | 147.B| 54.15| 177.3 | 72.20 | 236.4| 90.2 [295.5| 108.3} 354.6
1340 17.09 1 59.7 | 25,65 | 89.6) 34,20 119.41 42,75 | ¥49.2} 51,36 179.1] 68.40| 238,88} 85.50 [298.5| 102.5! 358.,2
0.9 Sea level 1140 11.97 1 76,21 17.95 | 114,31 23,95} 152.4] 29.95{ 190.5] 35.90} 228.6% 47,90 | 304.8} 59,90 [381.0{ 71.80 457.2
1240 10,331 74.0 |} 15.50{ 111.0} 20.65] 148.0 25.801 185.0} 31.00{ 222,01 41.30} 298,01 51.70 {370.,0] 62.00} 444.0
1340 9.29 0 7151 13,951 107.2| 18.60] 143.0| 23.20 178.8| 27.90; 214.5 | 37.15| 286.0| 46.45 {357.5| 55.75 429.0
1540 8.09 ] 68.8 | 12,131 103.2| 16.20| 137.6| 20.20| 172.0| 24.30| 206.4 ! 32.35] 275.2} 40,40 }344,0] 48,50} 412.8
0.9 35,000 1140 18.05! 61.Y } 27.05 ] 9L.6] 36.10| 122.2] 45,10 152.8] 54.15] 183.,3{ 72.20| 244.4] 90,2 |305.5{108.3 366.6
1240 16,681 62.2 1 25,00 93.3] 33,35) 1244 41.70] 155,5] 50,00 186.6] 466,70 248.8] 83.40 ;311.04100.0 | 373.2
1340 15,681 62,61 23.50{ 93.9) 31.35| 125.2{ 39.20} 156.5| 47.00{ 187.81 62.70| 250.4| 78.30 }313.0! 94.00} 375.6
1540 14.3 62,9 | 21.05{ 94.3} 28.10| 125.8] 35.10{ 157.2| 42.20) 188.6| 56.20| 251.6{ 70.25 {314.4} B84.30} 377.2
1.5 35,000 1140 18.66 | 81.4 1 28,00 122,1% 37.30| 162.8| 46.70| 203.5| 56.00] 244.2| 74.70| 325.61 93.30 | 407.01104,5 | 488.4
1240 16,22 78.4 1 24,35 117.6} 32,50 156.8] 40.60| 196.0{ 48.70) 235.2%{ 64.90] 313.6] 81.20 {392.0{ 97.30| 470.4
1340 14,331 75.7 | 21.501 113,6{ 28.70] 151.4{ 35.85] 189.2| 43.00 227.1} 57.30] 302.8] 71.70 {378.5| 84.00] 454.2
1540 12,27 | 73,2 18,40 109.81 24.55{ 146.4{ 30.70] 183,0| 36.801 219.6] 49.10} 292.8| 61.30 | 366.0] 73.6¢| 439.2
1.5 45,000 11406 27.27 81,41 40,851 122,11 54,50 162.8] 68,201 203.5] 81.80) 244.2|109.0 | 325.6{1356.3 407.0f 163.5} 48B.4
1240 23,45| 78.4 ) 35.20 | 117,86} 46,25 156.8]| 58.70¢ 196.0} 70.45] 235.2| 93.90] 313.6(117.30 §392,0] 141.,0} 470.4
1340 20,491 75,7 1 30.75 ) 113.6| 41,00 ¥51.4) 51.30) 189.2| 61.50| 227.1| 82.10} 302.8]|102.5 1378.5{ 123.0| 454.2
1540 17.28 | 73.2 | 25.90 | 109.8| 34.55{ 146,4| 43.201 183.0| 51.80| 219.5| 69.15] 292.8| 86.40 | 366.0} 103.6| 439.2
W = Propulsion machinery weight, 10_3 ib.

om

**P = Reactor power, megowaits,

cheMeeg
T S
ORNL-LR-DWG 69
370

 

 

 

    
      
    
 
 

 

NN

STRUCTURAL WEIGHT =30%, OF GROSS WEIGHT
POWER DENSITY =5.0 kw/em?®

320 | fi J‘ e

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

| o
b
570 te/hr, 1240: F -
1r/hr, 1340 F -
1r/hr, 1540° F \
o 10 r/hr, 12407 F -
% 10 r/hr, 1340° F
2 100 r/hr, 1240 F =3
$ | \
S //
220 S/ e S
;S
,,,,,,,,,, // / / /. / /
// ////
/ // ]
L/ / / k
o b N S 1600 v/hr A540°F
; S 100 r/hr, 1540°F
________ 4 % 1000 r/hr, 1340°F
1000 r/hr, 1240°F |
/ : /// N~ 400 r/hr, 434C°F
--------- ‘ %7/ RO /b, 1540°F T
/I/ 7 % V/ SO0 R B R
/42/2/ . ////4 ] ‘_
NS NN
200 300 500

, (Ibx 1073)

Fig. 1, Chart for Determining Aircraft Gross Weight at Mach 1.5 and 35,000 §t for Yarious Turbine

Inlet Temperatures.

10
 

TABLE 4, L/D RATIOS FOR YARIQUS FLIGHT CONDITIONS

 

 

 

MACH NUMBER ALTITUDE L/D
(f1) {without nacelles)
0.6 Sea level 15
0.6 35,000 15
0.9 Sea level 10
0.9 35,000 12
1.5 35,000 6
1.5 45,000 6

 

 

 

the crew shield) to show the effects of various
degrees of shield division in relation to the im-
portant design conditions (Figs. 2 through 5),
dose rate expressed here, for simplicity, in r/hrf‘!
is actually the personnel exposure dose rate
(rem/hr) from radiation made up of seven-enghfhs
gamma rays and one-eighth neutrons, assuming u
relative biological effectiveness of 10.

A number of important conclusions can be de-

duced from Figs. 2 through 5. Perhaps the most™
important is that the gross weight of the airplane:
is very sensitive to reactor power density and the
operating temperature, except under the subsonic
design conditions with power densities greater |

than 1 kw/em®. For a power density of about 1
kw/ecm® and a turbine air inlet temperature of aboit
1200°F, an increase in reactor temperature level
of 100°F is more beneficial than a factor-of-2 in-
crease in power density. The turbine cir inlet
temperature will be lower than the peak fuel tem-
perature by roughly 400°F, depending on the heat
exchanger proportions, and thus a turbine air inlet
temperature of 1140°F might correspond to a peak
fuel temperature of about 1540°F., Since it is
doubtful whether reactor structural materials will
be available that will permit reactor operating
temperatures of much above 1650°F, it is likely
that, to achieve turbine air inlet temperatures of
much above 1200°F, it will be necessary to provide
for interburning of chemical fuel between the
radiator and the turbine,

For reactor power densities of more than 1
kw/ecm®, the aircraft gross weight is not very
sensitive 1o the degree of division of the shield,
except in the range of reactor shield design dose
rates below 10 r/hr at 50 ft. This effect occurs

The .;,;

-

partly because the incremental weight of a given
radial thickness of shielding material increases
at o progressively more rapid rate as a unit shield
is approached and partly because, for the particular

. » . series of shields used, the secondary gamma rays

“Iproduced in the outer lead layer become of about

the same importance as the prompt gamma rays
from the core if the lead thickness is more than
about & in, The secondary gamma rays make it
necessary to add disproportionctely large amounts
of lead to reduce the dose rote from the reactor

shield to below about 10 r/hr at 50 f,

Effect of Chemical Fuel Augmentation

It is possible to use the same basic techniques
for investigating coses in which chemical fuel is
burned between the radiators and turbines to obtain
extra thrust for take-off, landing, and high-speed
flight. If the power required (in Mw) is multiplied
by 3413 Btu/kw-hr and divided by the lower heating
value of the fuel (about 18,000 Btu/Ib), the equiva-
lent rate of chemical fuel consumption is obtained

in Ib/hr, that is,
(Fuel consumption, Ib/hr)

3413 103
= (Power, Mw) ~——1---—>-<-----~-—-—-
18,000

= 190 (Power, Mw) .

The weight of the burners and the related equip-
ment required for chemical augmentation of nuciear-
powered turbojets should be roughly 25% of the
installed weight of the basic engine without radi-
ators. Thus the extra weight of the equipment for
interburning may be readily calculaoted by multi-
plying column ;i of Table 2 by 25% and by the total
thrust required. It is assumed that the weight of

il
450 oo w———I———?—-?__Tm?T"?_W_ .........

  
   
  
  
 
  

i DENSITY (kw/cm?)

 

 

 

 

 

 

 

\
\ | \ i
| , .
s00 o Ao i |
‘ ; | | REACTOR POWER

 

ORNL—-LR-DWG 195

 

 

 

 

 

 

 

 

 

 

" I I i : l [
! ! : | i : P
Q i | ]
> : [ / J | ( (
2 - | | ’ :
o ‘ \ : L
(1_,:} 300 : \ - - 4‘ ! o 4o ! 230
£ L | | ] |
0 TURBINE AIRINLET . [ ‘
& TEMPERATURE (OF ) ‘[ |
& | ! ‘ 1
| ‘
250 S } LA 1s0 g
L=t : =
@ | ‘ =
g ‘ ‘ x
- L)
<t =
g
200 14 180 g
o
<
lad
o
150 o 10
wootH——"-+—>—" 4L L. P L
L 1000
DOSE AT 50§t {r/hr)
Fig. 2. EfHects of Shield Division, Power Density, and Tuibine Air Inlet Temperature on Aircraft

Gross Weight at Sea Level and Mach 0.9.

the fuel tanks and lines can be offset by savings
in structural weight that can be effected by re-
lieving the wing bending and torsional loads
judicious location of the f{fuel storage
system. Therefore, the fuel tonk system is treated

through

as if the entire weight were made up of fuel. If
additional turbojet engines are required for use
with chemical fuel exclusively, their approximate
weight in pounds per pound of thiust can be ob-
tained from column i of Table 2 as functions of
altitude and Mach number by multiplying by a
factor of 1.25 to account for the burner equipment.

12

The performance of an aircraft with chemically
augmented nuclear power is illustrated in Fig. 6
to show the effect of sprint range on gross weight
for various reactor design conditions for a sprint
condition of Mach 1.5 and 45,000 ft. A comparison
with Fig. 5 shows that, for sprint ranges of 1000
to 1500 miles, the chemically augmented nuclear-
powered airplane is lighter than ths allenuclear-
powered airplane and the reactor power is much
lower, especially for the lower reactor operating
temperafures.
landing would be reduced, and the chemical fuel

Furthermore, the gross weight for
 

GAML-LR-DWG 693

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

500 250
450
DOSE IN 5-ft DIA, 10-ft LONG
CREW COMPARTMENT: 4 /hr
STRUCTURAL WEIGHT: 30%, OF GRDSS WEIG
30,000 1b \
400 ] ‘ TR 200
350
r?h
£ 300 150
”»
0
l.._
I
Q@
wl
=
»n 250
o3
Qo
x
O
e
<
o
2
£ 200 400
<L
150 e
| |
TURBINE AIR
INLET TEMPERATURE (°F) -
100 1140 50
1140
1140
1440
1240
1340
50 1540
0 0
{ 10 100 1000

DOSE AT %0 ft (¢/hr)

TOR POWER { Mw)

RELC

Fig. 3. Effects of Shield Division, Power Density, ond Turbine Air Inlet Temperature on Aircraft

Gross Weight at 35,000 f+ and Mach 0.9.

13
ORNL-LR- OWG 196

 

 

    
   

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

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1| | STRUGTURAL WEIGHT: 30% OF GROSS WEIGHT | ‘ L] !
| | | | USEFUL LOAD- 30,0001b '; o
| | 1 P | ‘ \ |
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‘ ‘ | N : % !
] il s
100 Lo L1 L] Ll
1 2 5 10 20 50 100 200 500 1000

DOSE AT 5C ft (r/br]

Fig. 4. Effects of Shield Division, Power Density, ond Turbine Air Inlet Temperature on Aircraft
Gross Weight at 35,000 f+ and Mach 1.,5.

14
ARCRAFT GROSS WEIGHT (10 ° Ip)

1100

1000

900

800

700

500

400

300

200

100

DRANL-LR~0OWG 197

 

L

DOSE IN 5-ft-DIA, 10-ft-LONG CREW COMPARTMENT: {r/hr
7 STRUCTURAL WEIGHT: 30% OF GROSS WEIGHT

 

|

T |

|

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

1000

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

800

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

600
\ ~—— :
freeee : e — edentid 400
\\.«__‘ \""fi-“\ :
{ \M‘*«-. 1540 \-—-‘.fi
e T r— f
T
f
\
oo e i 200
|
2 5 10 20 50 100 200 500 1000

DOSE AT 50 ft (r/hr)

REACTOR POWER (megawatts)

Fig. 5. Effects of Shield Division, Power Density, and Turbine Air Inlet Temperoture on Aircraft Gross

Weight at 45,000 ft and Mach 1.5.

15
 

 

 

 

 

 

 

 

 

 

    
      
 
 

ORNL-L_R-DWG 525

 

 

 

 

 

T mmmqymeemeees ‘—-----—v—"'—“—“—"——*r ------------------ —fi_“—'fi“’”—!—”””
A SO U BN e ]
| | |
! . |
~ ] RADIATOR OUTLET - . L - b .
INTERBURNER INLET TEMPERATURE {°F) i ; | T
| |
400 |m S e 140 - 4 ! 200
' Hag - 1
L L 424D Jo e ]
! ; 1140 : %
i { 1140 O R
| |
T ) 7
2 |
- e e
o ;
= | —_
i z
- - | 150 =
o = -
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2 | i g
o ! I
« | e ,{ . r
1 i o
b | | =
1 | ! =
d i — - w
S : ; x
o :
X et
20Q b-odo- —— e CRUISE FLIGHT CONDITICNS: Mach C.9-350C0 ft - - — 100
: SPRINT FLIGHT CONDITIONS: Mach £.3-45000 ft ‘
ot ——— - TURBINE INLET TEMPERATURE=1640"F b
‘ i | E ‘ |
__ 4 e e i S sy +
i i | : i ‘
| ; . : 1 i ! i
- — w : A R | |
L P | ! |
i i ! | i ! i |
\ p | . | L ; ‘ .
| ! 1 i | ' ! i | * |
‘ , | ! | i ‘ ‘ !
wo b—~ el ‘ I -‘ ------- e L ‘ ! : ‘ e b 50
0 500 1000 1500 2000 2500

SPRINT RANGE (miles)

Fig. 6. Performance of Airciaft with Chemically Augmented Nuclear Power,

be available for stand-by power in
the event of a reactor failure, Also, as the chemi-
cal fuel is burned, the weight of the aircraft will
decrease; therefore maneuverability, ceiling, ond
speed would be continucusly improving throughout
the sprint relative to the initial Mach 1.5 and
45,000 ft design condition,

system would

SHIELDING

The shield weight is the key to the problem of a
The weight de-
pends partly on the tolerable radiation dose for the
crew and partly on the degree to which the shield
is divided between the reactor and the crew com-
partment.

manned nuclesr-powered airplone.

Division of the shield introduces prob-
lems of radiation damage to organic materials in
the airplane, radio and radar reception and de-
tection, light emission from ionization of air around
the airplane, and the shielding of ground-handling
and maintenonce personnel.

16

Units of Rodiation Dose Measurements

Gamma radiation doses have been expressed in
roentgens (r) for many years, 1 r being the gamma
radiation dose giving on energy deposition of
83.8 erg/g of air. On the same basis, the “‘rep”’
(reentgen equivalent physical}) was devised to
serve as a measure of both neutron and gamma
radiation doses. Thus 1 rep in gommas is equal
to 1 r, and 1 rep in neutrons is a dose giving an
energy deposition of 83.8 erg/g of tissve. The
rep value gives a good measure of radiation damage
to organic materials; that is 1 rep of gamma rays
causes roughly the same damage as 1 rep of fast
neutrons (thermal neutrons cause very little dam-
age).
be very sensitive to fast neutrons, and the unit
“rem’’

Living tissue, however, has been found to

(roentgen equivalent man) was devised to
correlate neutron and gamma radiation damage to
The neutron dose in rem is obtained by
multiplying the neutron dose inrep by the “‘relative

man.
biological effectiveness’’ (RBE). While the RBE
varies from 2:to 50 for various parts of the body, o
factor of 10 is ordinarily used; hence 1 rep of fast
neutrons is usually taken as being equal to 10

rem, and T rep of gammas is equal to 1 rem.
Permissible Dose Rate for Crew

In attempting to establish o permissible dose
rate for military operations, it is instructive to
examine the standard laboratory radiation dose
tolerances in current use. These permit dose rates
of about 15 rem/yr through the normal working life-
time of laboratory personnel. It has been found
that radiclogists ordinarily get an average of seven
times this laboratory tolerance dose, largely through
carelessness. The principal ill effects appear o
to be an incidence of leukemia among radiologists
of approximately twice that of the population as a
whole. The small amount of information available
on humans indicates that genetic effects may be-
gin to appear in the form of mutation rates double
the normal value if the total radiation dose reaches
something like 75 to 100 rem. The threshold for
cataract formation in the eyes is about 200 rem of
fast neutrons, While the dose-rate problem is very
complex, many aspects of it are debatable, and
the data are inadequate, a total dose of 100 to 200
rem, of which about one-eighth could be in neutrons,
should be admissible, particularly if the personnel
were carefully selected so that the probability of
their having children would be low; that is, a mini-
mum age limit of 30 or 35 years might be imposed.
The genetic effects, which would be recessive and
would affect subsequent generations, would be
undesirable even though only a small percentage
of the offsprings would be affected. if a relatively
jow value of 100 rem were arbitrarily specified as
a permissible total dose for the crew, shielding
designed to expose the crew to 1 rem/hr would
permit any individual a total of 100 hr of flying
time in nuclear-powered aircraft, a limited but
possibly acceptable period. If the crew design
dose rate were 0.1 rem/hr, the flying time for on
individual in nuclear-powered aircraft would be
extended to 1000 hr, o period that would seem to
be entirely adequate. On this basis the analysis
presented in the following section has covered
only the range of crew dose rates from 1 to 0.}
rem/hr, of which one-eighth could be in neutrons.

 

2flF. E. Farris, A Compendium of Rodiation Effects
on Selids, Vol. I, NAA-SR-241 (Nov, 2, 1953).

Rodiation Damage 1o Grganic Materials
and Activation of Structure

The amount of maintenance work required will
depend to a large extent on the reliability and
service life of the equipment in the airplane. Since
materials deteriorate
radiation fields, radiation damage to organic ma-

organic in high-intensity
terials seems to impose an upper limit on the degree
of division of the shielding. The results of ex-
tensive experiments are available on rodiation
damage to rubber G-rings, gasket materials, elec-
trical  insulation, lubriconts, hydraulic fluids,
etc,24:25:26  Some representative data have been
organized in o separate reporf27 to put them in a
convenient form for engineering purposes, and
Fig. 7 is an applicable illustration taken from that
report. Briefly, it appeors that, after 300 hr of
exposure at full power, the best rubber hose and
C-ring materials tested to date would be seriously
damaged if they received a dose rate in excess of
30,000 rep/hr, while the poorest would be damaged
by one-tenth of that dose. Greases are similarly
affected, and the best petroleum oils can withstond
doses as much as ten times higher than the greases.
If o nuclear-powered airplane is to become truly
operational, it wili be highly desirakle — if not
essential — to limit the radiotion dose from the
reactor to a value such that elastomers and greases
would have a life of at least 300 hr if located 10
ft from the reactor. To satisfy this condifion the
reactor shield should be designed to give a dose
of mot more than 1000 rep/hr ot o distance of 50
ft from the center of the reactor.

The structural members of the airplane or of the
engines might present serious sources of radiation
if they were activated by absorption of neutrons.
This problem is discussed in considerable detail
in a G-E report,?® which indicotes that neutron
activation of the engines constitutes the more
serious problem and becomes important if the full-
power neutron dose at the engines is greater than
100 rem/hr. Thus, if the reactor shield is de-

signed to give one-eighth of the radiation dose in

 

25(:. D. Bopp and 0. Sisman, Rediotion Stability of
Plastics and Elastomers, ORNL-1373 (July 23, 1953),

zéGeflErOi Electric Company, Aircraft Nuclear Pro-
pulsion Project, Engineering Progress Report Mo. 7,
APEX-7 (March 1953). '

274 4. Stumpf ond B. M, Wilner, Radiation Damage to
Elastomers, lLubricants, Fabrics, and Plastics for Use
in Muclear-Powered Ajrcraft, ORNL CF-54-4-221 (April
15, 1954), :

17
 

 

 

 

10" ——————— e ————— e — —
5 ”|>
!
1
MATERIALS TESTED }
2 1 NATURAL RUBBER ({32a43) 1
2 GR-S30 (32847} ;
3 BUTYL RUBBER GR-150(32546)
108 4 NEQPRENE W (32n44)
- 5 HYCAR OR-15 (32845) |
5 HYCAR Poa-24 (32848) i
5 7 SILASTIC 7170 ‘!
8 THIOKOL-ST (3000-5T)
9 VULCOLLAN
_ 10 HYCAR OR-15 (PISIC2)
S 2 11 HYCAR 0S5-10 (PISIC3)
L’ 12 THIOKOL (Pisic 7
8 13 HYPALON-S2 (PISICAN
2 q0® | - ®
. ! 1]
: |
=
O i
e
5 ‘ 7 ]
| |
I J ‘
2 ‘ \
7 216 i3 7 4 8 5 ‘
10 S T e © G 2 . o -
i
: f
5
2 , |
: 5 4 6 ‘
. *» 0
o7 e28l83e]
108 S e - i

 

 

NO DAMAGE SLIGHT DAMAGE

MODERATE DAMAGE

ORNL-LR-DWG 920

 

 

 

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F- |
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‘ R 9! 10 13 1
e 8 o e
j e !.11
l . » - . e | .- o & —
i T2 g 6] -5 43 ]
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. ,,,013 i
*® a2 » e —
i 12 11 10| |
11012 13— f -
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- * @& & - ® -8 @ e _— _
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T
} I
[ P .
1 -
| |
| |
!
| j
i o e o

 

 

 

 

SEVERE DAMAGE COMPLETE FAILURE

RELATIVE DAMAGE

Fig. 7. Radiation Damage to Representative Elastomers lrradiated in the ORNL Graphite Reactor.

neutrons and the engines are located 15 ft from the
reactor, the shield may be divided to the point
where the total dose at full power 50 ft from the
reactor can be as high as 1000 rem/hr without
introducing any serious maintenance problems re-
sulting from neutron activation of equipment outside
the reactor shield.

Ground-Handling and Maintenance Preblems

it is instructive to examine ground-handling and
maintenance problems on the premise that, because
of radiation damage to orgariic materials and acti-
vation of structural materials, division of the
shielding is limited so that the full-power dose at
50 ft from the reattor may not exceed 1000 rem/hr.
Three major types of work that would require
people to be within 50 ft of the reactor shield are

involved. The first is the regular maintenance

18

work that could be scheduled for a period during
which auxiliary ground shielding might be arranged.
The second type of work is ground-handling or
maintenance immediately prior to take-off or after
landing in the course of which the use of auxiliary
shielding would be very awkward and expensive.
The third type of work includes unscheduled acti-
vity required by an emergency such as a fire or a
crash immediately preceding take-off or following
a landing when auxiliary ground shielding would
probably not be cvailable. The shielding require-
ments for these three major types of ground-handling
and maintenance work differ considerably because
of differences in exposure time and dose.

The first type of work, which might be carried
out with auxiliary shielding in place, covers the

bulk of the regular maintenance operations. It is
of interest to"note that experience in the B-36
flight-test program indicates that over 2000 man-
hours of work of this sort must be carried out per
test flight and that there is an average of one
flight per week. While the dose rate to be ex-
pected in the vicinity of the reactor after shut-
down will vary with the amount of gamma shielding
around the reactor, within minutes of the shutdown
it will generally drop by a factor of at least 20
from the dose rate for full-power operation. If
much of the full-power dose is from secondary
gammas generated in the shield, the reduction upon
shutdown will be correspondingly greater. Decay
of the short-lived fission products will effect a
further reduction in dose rate by a factor of about
Z in the first day and, again, by a factor of 3 in
the next three days; after that the dose rate falls
off very slowly. These effects are shown more
explicitly in Fig. 8.

[f the ground personnel worked 40 hr/wk, it should
be possible, with little inconvenience, to arrange
that they spend only one-third of their time in the
vicinity of the airplane, while the rest of the time
couvid be spent an appreciable distance awaoy.
However, because of the character of the mainte-
nance work fthat would have to be carried out, «
disproportionally large amount of time would have
to be spent in the vicinity of the airplone immedi-
ately after shutdown, A fair assumption might be
that ground personnel would have to take one-half
their total weekly dose during the first 8 hr
following a landing.

Thus if a man is to receive not more than (.35
rem/wk and he receives 0.18 rem during the first
8 hr following shutdown and spends 2 hr of that
time at an average distance of 15 ft from the center
of the reactor, the permissible dose rate would be
0.09 rem/hr at 15 ft. This dose rate ot 15 fi would
give a dose rate of about 0.01 rem/hr at 50 ft from
the center of the reactor. If no auxiliary ground
shielding were used, this weould require that the
shield be designed to give 0.2 rem/hr at 50 ft for
full-power operation; only a unit shield would meet
this requirement. Auxiliary ground shielding for a
divided shield could probably be arranged most
conveniently by draining part or all of the fuel or
water from the outer hydrogenous region of the
shield and replacing it with zinc bromide, mercury,
or an oil-metal shot mixture. The shield structure
would, of course, have to be strong enough to carry
the resulting loads. Since a load factor only «

little greater than 1 is required when the aircrafi
is af rest on the ground because of the absence of
dynamic leads, it should not be difficult to handle
the structural problem.

The second type of activity for which tolerabie
dose levels must be set involves much shorter
periods of exposure to radiation. This category
covers the ground-handling work that will be re-
guired prior to the installation of auxiliary shieid-
ing immediately after a landing or immediately
prior to a take-off after the auxiliary shielding has
been removed. This work would include towing the
airplane into position, last-minute funeup, checking
or repair operations, ond the installation or re-
moval of the ouxiliary shielding. While this work
might be carried ouwt with highly specialized
equipment, the cost and time involved could be
cut tremendously if the dose level in the vicinity
of the airplane could be kept sufficiently low so
that personnel could carry cut the necessary opera-
tions without special protection. If the dose rate
were | rem/hr at 50 ft ond if no appreciable amount
of werk within a 50-ft radius were required, the
personnel might be permitted to get the bulk of
their weekly dose, say 0.25 rem, in ¢ 15-min period.
This indicates that, to meet the requirements for
this second type of work, it would be desirable to
design the shield to give not more than | rem/hr
at 50 ft after shutdown, which would meuan about
20 rem/hr at 50 ft at full power.

The third type of ground-hondiing work that
should be considered in establishing shield speci-
fications is that associated with emergencies.
While it is very difficult to predict the character
of the work and the time required to cope with the
emergencies that might arise in connection with u
fire ot a crash, a total dose of 25 rem is permifted
under such circumstances by AEC regulations.
If the time of exposure to radiation were 15 min, a
dose rate of 100 rem/hr could be telerated on this
basis. Since emergency operations might have to
be carried out up to 15 ft from the center of the
reactor, it oppears that the shield should not be
divided beyond the point where it would give 100
rem/hr at 15 ft immediately after shutdown, that is,
200 rem/hr at 50 ft at full power. If the crash were
so violent aos to strew the surrounding area with
fission products, the shield would be ineffective
and the dose would not be a function of shield
design; hence such a case does not pose o shield
design problem,.

9
20

IN PER CENT OF FULL-POWER DOSE

DOSE AFTER SHUT-DOWN

ORNL-LR-DWG 1135

 

 

 

 

 

 

 
   

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

_ { -
I o ]
|
T=OPERATION TIME (hr)
Q
. ,ENT':'L, ,,,,,,, _ |
——
0 100 200 300 400 500
SHUT DOWN TIME (hr)

Fig. 8. Dose ot 50 ft from Reactor as a Function of Time After Shutdown,

 

600
The results of the above discussion have been
summarized in Fig. 9, The left column gives the
dose at full power 50 ft from the center of the
reactor as an index of the degree of division of the
shielding. The chart is applicable to any shield
design. The first four columns simply give dose
rates for representative conditions. (Note that the
characteristics of any particular shield design fall
along a horizontal line.) Typical radiation damage
limits are given in the fifth column., The sixth
shows the amount of auxiliary shielding required
to reduce the dose after shutdown to a level that
will permit @ man to work within 15 ft of the center
of the reactor for 2 hr shortly after shutdown and
for a total of an additional 10 hr during the suc-
ceeding week (assuming one flight per week). The
last two columns show the effects of neutron
activation of turbojet-engine parts.

In re-examining Fig. 9 and the preceding dis-
cussion, it appears that the reactor shield may be
divided to the point where at full power it would
give 100 rem/hr at 50 ft from the center of the
reactor without imposing exceptionally difficult
limitations on ground-handling and maintenance
operations.
point where at full power it would give 1000 rep/hr
50 ft+ from the reactor, radiation damage to elas-
tomers and greases would be serious. Also, ground
operations would be severely restricted and time-
consuming, and much expensive, specialized
remote-handling equipment would be required.

Shield Weight

The weight of a carefully designed redctor shieid
depends primarily on the reactor power, the power
density, and the specified full-power radiation dose
level at a given distance from the reactor, usually
50 feet. It is also heavily dependent on the dis-
position of equipment such as pumps and heat ex-
changers inside the shield and the presence of
voids such as ducts and headers. Another factor
is, of course, the kind of shielding material used.
A diligent search for superior shielding materials
has failed to disclose any that are markedly su-
perior fo a combination of lead ond woter {the lead
for gamma-ray attenuation and the water for neutron
attenuation). While the investigation of materials
is not complete, the most promising combination of
shield materials found thus far is vranium, bismuth,
and lithium hydride. A shield of these materials
might moke possible a shield weight saving of as

If the shield were divided beyond the

much as 15% in comparison with the more conven-
tional shield of lead and water. The shield weight
also depends on the weights of the shield structure
and of the cooling equipment required to dissipa}e
the energy of the radiation absorbed in the shield.
These items have been responsible for increases
in shield weight of as much as 20% in some de-
signs, and they may increase the weights for
shields of special materials more than they in-
crease the weights of lead-water shields. It should
be mentioned that jet fuel is as effective as water
as o neutron shield on a volumetric basis. The
lower density of the jet fuel gives a small saving
in neutron shield weight that is largely offset by
the additional lead required for gamma shielding.

A general idea of good shield design practice
can be gained from a highly simplified approach.
Roughly, one fast neutron and one hard gamma-ray
(over 1.5 Mev in either case) escape from the
reactor core per fission. This radiation can be
attenuated by a factor of 2.72 by a thickness of
approximately 3 in. of lead or water for the neutrons
or thicknesses of 1 in. of lead or 10 in. of water
for the gamma rays, The neutron flux from a 200-Mw
reactor must be attenuated through the shielding
material by a factor of about 100,000,000 if the
resulting neutron radiation dose is to be reduced
to 0.125 rem/hr (that is, one-eighth of the total
dose), and the gamma flux must be attenuated by a
factor of about 1,000,000 if the resulting gamma
dose is to be reduced to 0.875 rem/hr (seven-eighths
of the total dose) to give a total dose of 1 rem/hr,

About 20 attenuation lengths will be required for
the neutrons and about 15 for gamma rays, Since
the fast-neutron attenuation lengths in lead and
water are about the same, this means that about
60 in. of shielding material must be interposed
between the reactor and personnel 50 ft away to cut
their neutron dose rate to 0.125 rem/hr. Since .60
in. of water represents only six attenuation lengths
for gammas, about 10 in. of the shielding material
would have to be lead instead of water to cut the
total dose rate to 1.0 rem/hr.

The situation is complicated by the generation
of hard, secondary gammas from inelastic scatter-
ing of fast neutrons in lead or structural materials
and from neutron captures in hydrogen, lead, or
structural materials such as steel or aluminom,
The production of secondary goamma rays can be
inhibited by introducing boron or lithium to absorb
the neutrons as soon as they are slowed down by

21
cC

DOSE AT 50 ft FOR FULL -POWER OPERATION {rem/hr)

ORML-LR-OWG 2009

 

 

 

 

6 7 5 RADIATION EFFECTS CAUSED BY
— {0 — 10 — 1.8 sec — 5% 10 —
[ o
NEUTRON ACTIVATION OF AN ENGINE
= 15 it FROM THE REACTOR CENTER
~
E AFTER 400 hr AT FULL POWER
« = (/g OF DOSE iN NEUTRONS)
A
w % r ~
= RE DAMAGE T = - 2
L 40° L o° O | 18 sec = sx0f SEVERE DAMAGE TO =l 12in. OF LEAD (OR Er— ax10°f —
o o T PETROLEUNM _UBRICANTS O FQUIVALENT —~
o 4 — 0w w O £
D O 2
£ o r o = T £
£ = i © o - .
“’ < L - il 3 z
g s < > SEVERE DAMAGE TO TIRES, o o ez
= > K W
L 10* = 10° S 3 min B sx10” _ [ RUSBER HOSE, O-RiNG, AND - K- 10in OF LEAD - 4xi0 W G0
Oy
5 = <t} GASKET MATERIALS s 2 =5
L) - E - o w w O =
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o Wl W0 Q L - o
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o5 -
3 g 4 = s 2 oo w W< 2
L 10 0 + - 30 min o 5x10 = & Si— 8in. OF LEAD b a o
— + < 54 L O W
_ Q = O w = O
D uy = + oo g =
L — O + o« b =
o <1 O O Ui = w
= - O T O =z
7 z > = = - W &0
. 3 5 § Sy 2 3
2 * 3 = o o © ‘5 —1 i E 3
L 10 O 0 = 5hr T 540 2 L_“I_J#%—Sin.OFLEAD Z1— 4x 10 W g 10
- Ll
- W - = & = o
<1 =~ o < - = or £
— O > l w
Ll @ c r = - &g o
wn o = = < T O Il
Q Pt E o = 2 0 0
o O = < = T T w
w ~ < x =3 = 2 =
= -~ - D — o =
© w? E | son 5 < T Ol 4inor Lean = L
4 r o & - in. L Z e
— o w
<1 READ HORIZONTLLLY 4CROSS i %
- = - 5
[4p]
& M
a o
n - ) 8
| SETS L 500 o L 5% 10 .~ REPLACE WATER al— -

 

 

 

 

 

 

BY ZnBry,

 

 

Fig, 9. Approximate Effects on Ground-Handling and Maintenance of Degree of Division of the Shielding.
el A

the water shielding. (B'® and Li® are the only
materials that do not give off hard capture gamma
rays.) The production of secondary gammas in the
lead can be kept to an unimportant level by dis-
tributing the lead through the shield in such a way
that the neutron flux in the lead af any given point
is low enough so that the secondary gamma flux
produced there is below the local level of the
This concept
of the ‘‘matched’’ shield?® has proved invaluable.
It is equally applicable to the disposition of

primary gamma flux from the core.

structural materials.

The above concepts can be best illustrated by
examining their application te a ftypical shield
design.  The arrangement of reactor, heat ex-
changer, pressure shell, and shield assembly
shown in Fig. 10 was evolved in an effort to get
the lightest possible over-all assembly consistent
with reactor physics, heat transfer, and other re-
quirements for a circulating-fluoride-fuel reac-
tor. 1821 The arrangement is such that, except
for the pumps at the top, the varicus regions are
enclosed by surfaces of revelution about the verti-
cal axis of the reactor. It was found that the re-
flector around the reactor core should be at least
12 in. thick and should be followed by a layer of
about 0.13 in. of B0 if the Inconel pressure shell
were to be kept from becoming a more important
gamma source than the core insefuras gammas leck-
ing from the shield surface are concerned. Similar
reflector and boron-layer dimensions were found to
minimize the activation of the secondary fluid
in the heat exchanger by neutrons from the core.
It was also found that decay gammas from the fuel
in the heat exchanger would make the heat ex-
changer a gamma source of about the same im-
portance as the core, and therefore little would be
gained by placing gamma shielding inside of the
heat exchanger. It was found that attenvation of
the fast-neutron flux by the 12-in.-thick reflector
would be sufficiently great that o lead layer of up
to & in. in thickness could be placed just outside
the pressure shell without creating a seriously high
level of secondory gamma-ray productien in the
outer lead layers of the shield.

The only fairly complete set of shield weights
available to show the effects of reactor power,

 

281_“ Tonks and H. Hurwitz, The Economical Disitri-
bution of Gamma-Ray Absorbing Maferial in a Spherical

Pile Shield, KAPL-76 (June 8, 1948).

power density, und shield division is a set com-
puted for the basic arrangement shown in Fig. 10,
This set is presented in Figs. 11 to 15, The shield
weight increases with power at much less than a
ltnear rate, and ot o given power, it is not very
sensitive to reactor core diameter for core diameters
of less than about 24 in.; however, it becomes
progressively more sensitive for larger cores.

The data for seven representative cases have
been crossplotted in Figs, 16 and 17 to show that
the total reactor und crew shield weight is not
very sensitive fo the degree of division of the
shield, except for the nearly unit shields in which
the lead thickness exceeds & in. and hence dis-
proportionally large amounts of lead must be added
to take care of secondary gammas preduced in the
outer lead layers, A considerable saving in weight
might be effected by distributing part of the thick
lead region throughout the water; however, such a
step cannot be taken effectively until experimental
test dota ore available. The lorger crew compart-
ment of Fig. 17 gives less incentive for dividing
the shield thon does the smaller compartment of
Fig. 16.

An effort was made to estimate the shield weight
for a set of “‘ideal’” lead-water shields for the de-
sign conditions used for the engineered shields
described chove?® so that some insight into the
effects of the heot exchuanger, pumps, struciure,
etc. could be obtained. The enly daote available 1o
indicate the proper distribution of the lead to give
an  ideal matched shield cre from experiments
carried out by Clifford®® to establish the lead dis-
position for minimum weight in a lead-water shield
for a 36-in.-dia reactor. Perturbations were applied
to Clifford’s data te obtain estimated fotal shield
weights for a power density of 2.5 kw/cm® and a
dose 50 ft from the reactor of 10 rem/hr with o
crew shield designed to give 1 rem/hr inside of @
crew compartment 5 {4 in diameter and 10 1 in
length, The results are plotted in Fig. 18. To
facilitate comparison, similar data for enginsered
shields (taken from the calculations made for Fig.
12) have been plotted on the same coordinafes,

 

29R, M. Spencer and H. J. Stumpf, The Effeci on the

RMR Shield Weight of Varying Neutron and Gommoa Dosse
Components Taken by the Crew and Comparison of the
RMR Shield ‘Weight 1o That for an ldeafized Shield,
ORNL CF-54-7-1 {to be published).

30¢ E. Clifford et al., ANP Quar. Prog. Rep. May
10, 1950, ORNI_-768, p. 36.

23
DWG. 18743
MAKEUF FUEL TANK-- L~ PUMP QUILIL SHART

      
  
   
 
  
  
   
  
   

.~ CONTROL ROD ACTUATOR
’ »~ PUMP QUILL. SHAF'T

THERMAL INSULATION -, THZRMAL
Y / INSULATION

Nok DUCT ™~ ...
ot

REACTOR PRESSURE SHELL - /.

THERMAL NSULATICA #{f e
/

]
LEAD SHIELDING -
il
4
4

COOLING PASSAGE —

I

BCRATED WATER ——-

DRAIN VALVE ——

FLEL DRAIN L NE -—" N - SEIF - SEALING

  

RUBBER TANK

VERTICA. SECTION

LEAD SHIELDING - -~ REACTOR PRESSURE SHELI
.

   
  
   
 
   

-~ THRERMAL INSULATION

 

 

 

THERMAL
INSULATIGN~.

 

 

 

“._WEB OF CANT.LEVER BEAM
SUPPORTING REACTOR AND SHIELD
TJBE BLNDLE

- e ,
SELF—SFEALING RUBBER TANK — - BORATED WATER

Q & 12 24
HOR!ZONTAL. SECTION e ]
SCALE N INCHES

Fig. 10. Sections Through a Lead-Water Type of Shield.

24
ORNL - LR-!QE 682

 

180

 

160

 

 

140

REACTOR AND SHIELD ASSEMBLY WEIGHT (ib X 40 )

 

 

 

120

 

 

 

 

100

 

 

8C

 

60

 

 

 

 

 

 

 

 

 

 

 

40 l :
O {00 200 300 400 500 600 700 80C
REACTOR POWER (Mw)

 

Fig. 11. Effects of Power Output and Core Bicmeter on the Weight of the Reactor and Reactor Shield Assembly (No Crew Shield) for
Dose 50 ft from Recctor of 1 rem/hr (i/ Neutrons, / Gemmas).
9

: ORN -DWG 683

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

o . o
- ,/ ~
] 7 ,/ .:"":,/
~
CORE DIAMETER, 45.3 in. P / e
R AN L 7 2
POWER DENSITY, 1.25 kw/cm3—\ / e //:/
N f =~
1 \ 4
m CORE DIAMETER, 36.0 in. ')/ -~ q/'/ /ff
€ o0 \ // P .
® /
- AT
}-Iu / \ (// / 7‘4\( \\ l
x ’ //S'/ N \ PO\.«;ER DENSITY, 5.0 kw/om3
< : - : — , B cm
> A7 P /))( N\ | | |
g GO // ‘/ /; ~ A, N 1 ! . !
= /s /fi — POWER DENSITY, 2.5 kw/ecm3
o Z / \ | ! |
5 7 = \ < ]
o /// ya //‘ \ — CORE DIAMETER, 28.5 in.
L
= s 5 | |
T 80 ¥ ! t !
; ,/ s — CORE DIAMETER, 22.7 in.
< 2 M\
: 7/ \
5 /) //
= . L
/é\ CORE DIAMETER, 8.0 in.
N,
7
7
40
¥
\— CORE DIAMETER, 14.3 in.
|
|
|
20 ? :
o 100 200 300 400 500 600 700 800 900

REACTOR POWER ( Mw)

Fig. 12. Effects of Power Qutput and Core Diameter on the Weight of the Reactor und Reactor Shield Assembly (No Crew Shield) for
Dose 50 ¢t from Reactor of 10 rem/hr (1’8 Neutrons, ?’7;3 Gammas).
ORNL-LI!-I DWG 6B4

 

140

 

 

120 |

 

 

 

1CO

 

 

 

30

 

 

60

 

 

 

40

REACTOR AND SHIELD ASSEMBLY WEIGHT (Ib X10 %)

 

 

20

 

 

 

 

 

 

 

 

0 100 200 300 400 500 600 700 800 300
REACTOR POWER (Mw) ’

Fig. 13, Effects of Power Qutput and Core Di«zmeter or the Weight of the Reactor and Reactor Shield Assembly {No Crew Shneid) for
Dose 50 f¢ from Reactor of 100 rem/he (/5 Neutrans, 7 % Gammas},
8C

ORNL - LR- DWG 685

 

 

 

 

 

12 O . l! E ;
5 ; i
F a
t i i
* x 5 5\«;%
, E : R’ OF’N % L
100 @0‘*“25 Wi 10%/
| T A

 

 

 

 

 

 

 

 

 

REACTOR AND SHIELD ASSEMBLY WEIGHT (ib X 70“3)

 

 

 

 

 

 

| J

0 100 200 300 400 500 600 700 800 900
REACTOR POWER {Mw)

 

 

Fig. 14. Etfects of Power Output and Core Diameter on the Weight of the Reactor and Reactor Shield Assembly {No Crew Shield) for
Dose 50 {1 from Reactor of 1000 rem/hr (3‘8 Mautrons, 1?8 Gammas).
60

W
o

B
o

TOTAL WEIGHT OF CREW AND SHADOW SHIELDS (b X 10 °)
n O
o o

O

Lo

6¢

ORNL-LR-DWG 6B

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

L N I !;f
| | |
Pl
HEa
i A
/f
. K | |
E. Q;l"\q, : i
/
/ |
} I 1 L
p _
! *\L///,
2O
>d

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

n
Pt
\%
5

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Pl
' !
: L [ _t

 

20 50 100 200 500 GO0 2000 5GC0 10,000 50,000
DESIGN ATTENUATION OF DOSE FROM REACTOR SHIELD (r/hr)

Fig. 15. Effects of Design Attenuation on Crew Shield and Shadow Shield Weights for Equal Attenuation of Neutrons and !G:o'mmu's.
0t

ORNL—~LR—CWG 457

 

1.4 X 10

 

1.2 X10°

 

 

 

 

 

 

 

 

 

8.0 X 104

 

TOTAL WEIGHT (Ib)

 

6.0 x 10%

 

 

 

 

4.0 % 10?

 

 

 

 

 

 

 

 

 

 

2.0 x 10

 

N —
|8}
N
O

50 100 200 500 1000 2000
DESIGN DOSE 50 ft FROM REACTOR SHIELD {r/hr)

Fig. 16. Effect of Design Dose from Reactor Shield on Total Reactor and Crew Shield Weight for o Dose of 1 rem/hr in a Crew Com.
pariment 5 ft in Diameter and 12 #t in Length, '
DWG. 21!'5’1

 

  
 
 

 

 

 

 

 

 

 

 

 

 

  

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

120
O 285'” dth
= 1 ORE 200)\;1 '
£ 80 _22.7-in- c‘“GCORE 200 - M |
= - )
5 3 L
% 18.0-in.-dig CORE 200 Mw P
= 18.0-in-dia CORE; 100- MweJ‘
= .
o — —48.0- in-dia CORE; 50 - Mw 1]
180m dmCORE 25 - Mw |
- 1
B o
40 ; ( L
-
| | | |
| ‘
< | ; |
: Do ; ! -
| ; i J | | :
L | o
] AR
] ; : . : i i i : : i t | !
5 by bbb ISR
1 2 5 {0 2 5 100 2 5 1000 2

CESIGM DOSE 50ft FROM REACTOR SHIELD {r/hr)

Fig. 17, Effect of Design Dose from Reactor Shield on Total Reactor and Crew Shield Weight for a Dose of 1 rem/hr in a Crew Come
partment § ft in Diameter and 12 ft in Length.
¢t

 

 

 

 

 

 
  

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 
 
  
    

 

 

 

 

 

189 | | | [ | | | | : |
| | | | | | | | |
| ~ i | | SOURCE OF WEIGHT INCREMENT
S I R I ]
oo} —— g -
‘ TOTAL REACTOR, HEAT EXCHANGER, } DUC“‘fS, PUMPS, lflND STRUQTURE
‘ AND REACTOR SHIELD WEIGHT FOR R¥R ~ % ExTRA LEAD FOR DECAY
140 ; | 5 \ | GAMMAS FKOM HEAT EXCHANGER
| o | i
| INSERTION OF HEAT EXCHANGER,
: : PRESSURE SHELL, AND LUMPING OF
120 ! : T LEAD OUTSIDE OF PRESSURE SHELL |
: 1 ‘
o L |
~ 400 F— —— — —
__(2 .
g |
hu ! 3
9 ]
[’ - e [ -l ]
z 80 \ | | |
REACTOR AND SHIELD WEIGHT FOR
IDEALIZED MATCHED LEAD-WATER SHIELD
! 1 ;
o — ; e e e e
: | | | | I 5
| 5 DOSE: 10 r/hr AT 50 ft FROM REACTOR = ‘
‘ (/s gammas, g neutrons) ! : |
40 1 -t —_— - - i - -
g [ POWER DENSITY: 2.5 kw/em? ‘ 7 | |
i \ | i -3 | |
| | | | |
i
20 | = bt — WL—ffl—ffl S —
! 3 ‘ j ‘ ‘ : : :
L N
, 5 | | ! ‘ l } |
0 | | : . 1 | i | \ ‘;
0 100 200 300 400 500 600 700 800 900 1000 1100 1200
REACTOR POWER (Mw)
-

Fig. 18. Comparison of Reflector-Moderated Reactor Shield Weight with Weight of Corresponding ldealized Matched Lead-Water Shield.
The major factors that cause the weight of the

engineered shield to be greater than that of the

idealized shield are the insertion of the heat ex-
changer and the pressure shell and the fumping of
the lead into a single region immediately outside the
Because of the lumping of the
lead, however, there is little increase in shield
weight required to take care of fission-product
decay gammas from the fuel in the heat exchanger.
The ducts through the shield increase the weight,
but this increase is not very serious because the
ducts are filled with liquid and their cross- sectional
area is small. The very much larger ducts required
for an air- or helium-cooled reactor would be re-
sponsible for far more serious increases in shield
weight. ‘

In re-examining Fig. 18, it is interesting to note
that, in the reactor power range of greatest current
interest (200 Mw), the engineered shield is about
one-third heavier thon the idealized
matched shield. While this is a rough estimate
since all the shield weight estimates are subject
to errors totaling about 10%, the character of the
calculations was such that the values for the
various conditions should be comparable.

Since the shield weight data given in Figs. 11
to 15 were used in the parametric aircraft perform-
ance study presented earlier, it is pertinent tfo
consider the applicability of this data to other
types of reactor. Since most other reactor types
make use of cores in the form of right circular
cylinders, an allowance for their less favorable
geometry must usually be made. 1t has been shown
that the diameter of o sphere equivalent to a right
circular cylinder from the shielding standpoint is
about 25% greater than that of the cylinder.? Since
the cube of 1.25 is about 2.0, it can also be said
that the effective power density for the cylinder
shouyld be considered as roughly twice its average
actual power density to be equivalent to that for a
spherical core of the same diameter.

QOther items that might have substantially dif-
ferent effects in other reactor types would be the
heat exchanger and the ducts. From Fig. 18, it
appears that elimination of the heat exchanger
might make possible a weight saving of about 15%.
Most of the other items considered in Fig. 18, ex-
cept the ducts, seem to hove small effects for other
reactor types. Insertion of the large voids required
for the ducts and headers of a gas-cooled reactor
would be responsible for major increases in weight,

pressure shell,

fead-water

particularly for the more nearly unit shields. In
fact, since the shield thickness increases little
with reactor power, while the duct flow-possage
areg must increase in direct proporfion to reactor
power, and since the radiation Jeakage through a
duct increases more rapidly than its cross-sectional
area, many shielding experts feel that the extra
weight required for the ducts for a near-unit shield
for a high-power gas-cooled reactor represents an
extremely large weight increment. The problem is
clearly quite formidable, as is indicated by the
fact that to date no shield weights for gas-cooled
reactors have been published except those for
highly divided shields that give 80,000 rem/hr or
more at 50 ft from the reactor.

While some weight savings might be effected by
using shielding materials other than lead and water,
no substantially lighter engineered design based
on such materials has been prepared to date. In
any event, a weight saving of more than 10% through
the use of special materials seems uniikely. After
reviewing all the above-mentioned factors, it ap-
pears that there is little likelihood of getting an
operational reactor and shield assembly that will
weigh less than perhaps 85% of the values given

in Figs. 11 to 15.

NUCLEAR PROPERTIES

Design proposals for high-powered reactors have
ranged from those for the near-thermal water-moder-
ated reactor of the supercritical-water cycle to
those for fast reactors, as can be seen from the
values given in Table 5 for median energy for
fission. The important aircraft reactor design
proposals are compared in Tobles 5 and 6 with
other representative reactors. In general, it has
appeared that the higher the median energy for
fission, the greater is the crifical mass. This is
particularly true for solid-fuel-element reactors,
because relatively lorge core volumes of at least
2.0 #% are required to satisfy heat transfer and
fluid flow requirements (to be discussed in a later
section), Further, because of the shorter neutron
lifetime inherent in the faster reactors, it has
been felt that they would present markedly more
serious control problems. Because of these
factors, all the reactor designs that have looked
promising enough to receive considerable attention
for aircraft application have been thermal or
epithermal, that is, have hod a medion energy for
fission of between 0.025 and 1 ev. As is evident

33
ve

TABLE 5. REACTOR PHYSICS DATA FOR REPRESENTATIVE REACTORS

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

1 | |
. MTR STR sCw SIR EBR I ARE AC-100 |  RMR
Critical Mass i | |
Clean, kg of U233 2.3 9.5 18 52 48.2 { 8to15 | 24.5 18
Poison allowance 500 g | 8.7 kg 5 kg/1000 hr | None (a) (@)
at 62.5 Mw '
Burnup allowance 38 g/day 2.0 kg 2 kg/cycle | 87.2 g/day 1 g/day | 1.5 g/day 0.833 g/hr ' 2.1 g/hr
‘ at 62.5 Mw |
!
Power, Mw 30 70 700 62.5 1.4 1.5 20 f 60
Hours at full power 600 600 144 900 > 3000 1000 | 100 1000
Per cent thermal fissions ~95 94 ~40 0.7 0 65 |76 35
|
Mean neutron lifetime, ~15 8 ~1 ~2 1 14 4.214 P40
sec X 10—5 ‘ f
|
l | f
Temperature cosfficient of | ~15x107% 3.9 10 15| 5 x 1075 | 1p-5 255x1075 | 2.8x 1075 | 3.6 10-40) | 7 10-5
reactivity, per °C i x 1079 E !
Peak-to-average power density \ 1.8 3.9 ~1.5 2,28 1.25 ; ~2.0 l 1.3 1.7
| {longitudinal)
Fuel heat capacity, cal/°C 22,9 x10%(€) | 37 10% | ~5 x10¢ | 4.9 x 105 1.6 x 10° 3.8 x 194 26.3 x 10° 410 6 x 104
Median energy for fission Thermal Thermal | ~0.1 ev 52 ev Fast Thermal Thermai P V0.2 ev
i |
Free-flow ratio in core 0.63 F ~0,7 ~0,3 0.3 0.273 0.007 0.428 ! 0.75
Power density, kw/cm> ]! 0.3 .10 1.0 0.218 {avg.) 0.167 .03 0.05 ] 1.0

 

 

 

(

Preponderantly a slow

©ar 127°C.

a)FueE to be added as required.

temperature coefficient associated with the water moderator,
TABLE 6. DIMENSIONS AND COMPOSITIONS OF REPRESENTATIVE REACTORS

 

 

 

REFLECTOR | REFLECTOR
I1ZE COR M 0
REACTOR CORE S E COMPOSITION {vol %) THICKNESS COMPOSITION
MTR 73 % 23.4 X 60 cm 0.25, U3, 36.66, Al; 63.09, H,0 | 12 in. Be + 44 in. C | Be, C
STR 36.4 in. dia x 43 in. high 0.15, U235; 57.8, Zr; 42.0, H,0 2 in. HZO
SCW 2.5 ft long % 2.5 §t dia 0.3673, U233, 17.36, stainless 2Y in, H.,0
steel; 82.272, H20
SIR Hexogonal prism (eff.) 55.2, Be; 31.1, Na; 7.9, stainless | 8 in. Re
27.25 in. high x 27.25 steel; 2.2, MgO; 1.8, UO:2 (93%
in., dia enriched}); 2.0 void
EBR Hexagonal eylinder 7.5 in, | 53, uranium; 13.5, stainless 10 in, U238 i u?38
high, 7.5 in. across flats steel; 33.5, NaK breeding blanket
ARE 33.0 in, dia x 35])3 in, high | 82.6, BeO; 7.6, fuel; 4.8, Na; 2.2, 7!‘{2 in. BeO
inconel; 2.8, void for rods coolant: Na
AC-100 Hexcgonal eylinder 30 in. | 40, HEO; 4.7, Al; 11, insulation; ~3 ft H20
high, 29.5 in, across 5.2, fuel, including UOZ; 39.1,
flats, and 33.4 in. across void
diagonal
RMR Sphere, 23.75 cm radius 40.98, F; 39, Na; 17.47, Zr; 31 em Be
2.55, U3

 

 

 

 

 

in Table 6, this has necessitated that the concen-
tration of iron-chrome-nicke! alloy structural ma-
terial be kept to less than 18 vol % and that the
microscopic neutron absorption cross section of
the coolant be less than 1 barn. In fact, the only
reactor listed in Tables 5 and 6 that does not meet
these conditions is the EBR, a nonmobile reactor,

Moderating and Reflecting Materials

At first glance there appear to be several ma-
terials to choose from for the moderator, Beryllium,
beryllium oxide, graphite, water and heavy water,
sodium hydroxide, sodium deuteroxide, lithium,
copper, lead, and bismuth all might be used as
either moderating or reflecting materials. Beryl-
lium, beryllium oxide, D,0, and graphite have such
low capture cross sections that they may be used
in very thick sections without serious loss to the
neutron economy. Thermal stress considerafions
make beryllium oxide of doubtful value for reactors
having core power densities of greater than 0.5
kw/cm?®, even though beryllium oxide is one of the

best of the ceramics from the standpoint of thermal-
shock resistance. Beryllium and grophite appear
to be satisfactory from the thermal stress stand-
point, although they present other problems, The
cost of fabricated beryllium must be expected to
be from $75 to $300 per pound, while the cost of
reactor-grade graphite is only about $0.15 per
pound, However, its much higher atomic density
and its better high-energy scattering cross section
make beryllium much superior fo graphite on a
volumetric basis. The use of beryllium gives amuch
more compact reactor and hence a much lighter
shield, Normal water has such o short diffusion
length that it may not be used in sections thicker
than 1 in. without excessive loss of neutrons to
captures in the water, Partly because of this and
partly because of its predominantly forward scat-
tering, normal water is much less effective as o
reflector than beryllium, beryllium oxide, or graph-
ite. For the same reasons, much the same can be
said for NaOH, NaOD, and Li’OD. The properties
of the principal moderating materials are shown

in Table 7,

35
TABLE 7. PROPERTIES OF PRINCIPAL MODERATING MATERIALS

 

 

 

 

 

 

 

 

 

 

 

H,0 BE BeO D,0 | NaOH | Li’OH C
Density, g/cm> 1.0 1.84 2.84 1.1 1.8 | 1.4 1.6
Age-to-thermal, cm? 33.0 | 98.0 105.0 120.0 | 120.0 350.0
Thermal diffusion length, cm 2.88 | 23.6 28.5 100.0 5.0 | 5.9 50.0
Thermal conductivity, 0.35 | 48.0 15.0 0.35 0.7 72.0
Btu/hr-f?2-(oF/ft)
Therma! expansion coefficient, 10.0 x TO“é 5.5 x ]0m6 1.1 % 10'—6
in./in.-°F
Modulus of elasticity, psi 40.0 x 10° 42.0 x 108 1.5 % 108
Effect of Moderating Materiol on Design SECRET
DWS. 14428

Any detail design is heavily dependent on the
materials used, and many different materials
combinations appear interesting at first glance. If
moderating material is distributed throughout the
core, it displaces fuel and coolant and makes the
core larger for a given power than would be re-
quired by heat transfer and fluid flow consider-
ations. Normal water can constitute as little as
25 vol % of a recctor core for which the fuel in-
vestment is kept to within tolerable limits., |f
beryllium is used, at least 50 vol % of the core
should be occupied by moderator, unless the
principle of reflector moderation is employed, in
which case a fairly uniform power distribution can
be obtained with as little as 25 vol % of beryllium,
Much the same relations hold for D, 0, NaOH, and
Li’OH as for beryllium. The relationships on
which these observations are based were discussed
in earlier reports3'+32 from which Figs. 19 to 22
were taken to show these effects, [f allowances
are made for the volume required for structure,
control rods, etc,, the ratio of the flow passage
area for the reactor coolant to the cross-sectional
area of the reactor can hardly be better than indi.
cated in the free-flow-ratio entry in Table 5 for
representative reactors,

Hydrogen is such an obvious choice as a moder-
ator that some further remarks about its limitations
must be made. The forms in which hydrogen could
be used in a reactor are limited, namely, water, a

 

31w, K. Ergen, ANP Quar. Prog. Rep. Mar, 10, 1952,
ORNL-1227, p. 48.

32c. B. Mills, ANP Quar. Prog. Rep. Dec. 10, 1951,
ORNL-1170, p. 14.

36

 

 

 

 

j | CRITICAL MASS DETERMINED BY

i r BARE REACTOR MAETHOD, USING A

| | REFLECTOR SAVINGS VALUE OF £
g (THE DIFFUSION LENGTH)

(10)

MASS

CRIT.CA:

 

URANIUM

 

 

 

CORE DIAMETER (ft)

Fig. 19, Critical Mass vs Core Diameter for
Hydroxide Moderated Reactors with Thick Hydrox-
ide Reflectors.

hydroxide, an organic compound, or a metal hy-
dride. |f water were used, either it would have to
be kept at a pressure of around 5000 psi, which
would pose exceedingly difficult structural and
pump seal problems, or it would have to be ther-
mally insulated from the hot zone of the reactor,
a measure that would be wasteful of core volume

and would probably introduce poisons. An even
 

KOH

 

. NaOH

N i

 

 

 

 

URAMIUM REGUIRED FOR CRITICAL MASS {(wt %)

 

 

 

 

 

CORE DIAMETER (ft)

(o)
Fig. 20.

Digmeter for Yarious Hydroxides.
Fuaction of Temperature.

DM M I5658

 

-

ORNL—-LR—DWG 11328

 

 

URANIUM SOLUBLE {wi %]

 

 

 

 

 

 

 

500 550 600 650

TEMPERATURE (°C)
(5}

(a) Weight Per Cent of Uranium Required to Achieve Criticality as a Function of Core
(b) Weight Per Cent of Uranium Soluble in Varicus Hydroxides as a

 

 

 

 

 

 

 

ORNL. - A==

 

SOWG 1140

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

- [0 [ ]
EXTERNAL SYSTEM VOLUME
100 | A IS ASSUMED TO BE 16 fi® 177
b L
. J A 3-ft DIA CORE TOTAL
2 3Y,—ft DIA CORE IES‘F":'T%";Y
ft
2 i 4= 11 DIA CORE
£ 60 | bt i ------ SR .
= | ;
g Ll Ll o
2 L]
= 40 e e oo
5 4-f1 DIA CORE | URANIUM
: /j, 3%-ft DIA CORE N
50 | bl L] 3-f1 DIA CORE CORE - |
=T
,,,,,, | s eedheee e
O ] -
0O oM 020 030 040 030 060 070
VOLUME FRACTION FUEL~COOLANT IN CORE
Fig. 21. Total Uranivm Inventory and Uranium

in Core as a Function of the Volume Fraction of
Fuel-Loolant in the Core for BeO-Moderafed Circu~
lating-Fuel Reactors with 3, 3/- and 4.-ft-dia
Cores and NaF-UF , Fuel-Coolant.

 

 

 

 

 

 

 

 

240 Dt
\ 177 ] ‘ T 1
_FUEL CODLANT Lif - ’\lal KF f‘Jr, T _____ [ B T
VOLUME OF THE SYSTEM OUTSILE ] | | 1 '
200 THE CORE: 150 fi¥ | ) )' T
WATER OENSITY: 09567 arem® | L i' |
i
= |
£ - 3%, ‘H UA CORE | !
4 ¥ 4 22
4 (a2l ftDia coge > 5 5 2
= "n\ B i
5 3-fr-DIa CORE |
T
> - . ——eeers foooeom freneene
L w 5/? fH-D1A CORE 5’# 7
- 3-ft~DIA CORE Z 50
L
ke 2V Fr-DIA CORE | 2 p
..‘[__.__%_.___ b i -
ol b oL | L
020 030 G40 030 B0 . Q70 030 030 040
VOLUME FRACTION FUEL ~COOLANT IN CORE
Fig. 22, Total Uranium Inventory and Uranium

in Caore as a Function of the Volume Fraction
of Fuel- Coo!nn‘l in the Core for H,0- Moderuted
CirculatingsFuel Reactors with 2/- 3-, and 3/’-
ft-dia Cores and LiF-NaF-. KF«UF4 Fuei-'Cooiani'.
more important factor if the water were thermally
insulated would be that between & and 15% of the
reactor output would go into heating the water;
thus not only would heat be wasted, but the wasted
heat would have to be dumped through o radiator
at low temperature, and the radiator would impose
a weight-and-drag penalty equivalent to a further
loss in power-plant output of ot least 10%, An
over-all performance penalty of 15 to 20% seems
to be a stiff price to pay for the privilege of using
water as the moderator, The design compromises
that would be necessary to cope with problems of
distortion and differential
pansion would probably entail still further penal-
ties,

thermal thermal ex-

Various organic compounds of hydrogen have
been suggested as moderators, for example, di-
phenyl oxide, cyanides, etc. However, radiation
damage tests on organic compounds indicate that
this is not a promising course because the gamma
flux in the moderator would be about 10'° gam-
mas/cm?.sec for a reactor core power density of
1 kw/em®. All orgonic liquids tested to date
have shown severe radiation damage after an inte-
grated gamma flux of, at most, 1078 gammas/cm?2.
This would give an operating life of only 20 min
for the moderator material, Not only would radi-
ation decomposition of the moderator fluid present
a problem, but it seems likely that deposits of
carbon and sludge on heat transfer surfaces would
tend to render them ineffective. Metal hydrides
might prove sufficiently stable under radiation,
but none with truly satisfactory physical properties
hos been developed to date. Hydrogen gas is too
diffuse for use as a moderator, and liquid hydrogen
would present cooling problems inconsistent with
high-temperature aircraft reactor design.

The lowest estimated critical masses for the
various configurations considered are for some of
the hydrogen-moderated cores. However, there is
less chance to get a low critical mass in o high-
power reactor through the use of hydrogenous
moderators than appears ot first glance, because
the data of Figs., 19 through 22 do not include
allowances for temperature effects, coatrol rods,
burnup, ond fission-product poisons. The low-
critical-mass reactors are highly sensitive to these
poisons and require much larger allowances to take
care of them. This can be deduced from the first
two lines of Table 5. The same data show that
the lower the critical mass for the clean, cold

38

condition the more sensitive is the reactor fo the
accumulation of fission-product poisons and to fuel
burnup.

Reflector-Moderated Reactor

The reflector-moderated reactor presents a num-
ber of important advantages. By removing most of
the moderator from the core to the reflector, the
effective power density in the core can be nearly
doubled for a given average power density in the
fuel region. By heavily lumping the fuel, it is
possible to eliminate much of the parasitic struc-
tural material ordinarily required to separate the
moderator and fue! regions, |f beryllium is em-
ployed as the reflecter-moderotor, a substantial
proportion of the neutrons are reflected back into
the fuel region at epithermal energies so that they
penetrate even fairly thick layers of fuel and keep
the ratio of the peok-to-average fission density
from exceeding something of the order of 1.5 to
2.0,

Many factors influence the critical mass of the
reflector-moderated reactor.  Perhaps the most
important is the poison concentration in the re-
Other factors include the core radius and
annulus thickness., Figure 23 shows
critical mass plotted against these last two factors

sl

ORNL-LR-DWG., 2017

flector,
the fuel

EXTRAPOLATED REFLECTOR

@ 49
THICKNESS = 30 CM NV ~
. |
= |
Y 30 |
§ 0 ’
g
] EQQ
o -
R;
> 1@0\
FS
R
0|

 

EXTERNAL FUFL VOLUME =0 FT2

Fig. 23. Effect of Reactor Dimensions on Con-
centration of Uranium in o NaZrF, Fuel for g
Reflector-Moderated Circulating-Fuel Reactor,
for a 30-cm-thick reflector containing an amount
of poison representative of that which would be
involved if canning of the beryllium with inconel
should prove necessary. If the caonning is not
required, the critical mass will be reduced by
about 30%. A quite complete set of multigroup
calculations (from which Fig. 23 was taken) is

being made to determine these effects,33

REACTOR CONTROL3*

The problem of reactor control is essentially one
of matching the power of the reactor to the Joad.3?
This uswally amounts to keeping the fuel elements
at a prescribed temperature and moking absolutely
certain that they do not go above a maximum
temperature considered to be the threshold for
damage., The ease with which this control can be
accomplished is associated with the temperature
coefficient of reactivity, It has been demonstrated
that o reactor with a large negative temperature
coefficient in the fuel does not even require a
control rod; slow-acting shim rods for shufting
down the reactor, for compensating for long-term
drifts in reactivity, or for changing the operating
temperature may be incorporated in some instances.

A large negative femperature coefficient in the
fuel has been achieved only in liquid-fuel reactors,
such as the “water-boiler’’ and the Homogeneous
Reactor Experiment. The Aircraft Reactor Experi-
ment and the reflector-moderated reactor should
also exhibit the demonstrated stability of other
liquid-fuel reactors. Consequently, the control of
the ARE oand the proposed reactor should be
simple. |t moy well be that no nuclear instrumen-
tation will be required for the circulating-fuel
aircraft reactor; the proposed Homogeneous Test
Reactor (HRT) is not to have mechanical control
rods. ' ' |

Reactors with solid fuel elements do not exhibit
a large negative temperature coefficient in the
fuel, although they may have o small over-all nega-
tive temperature coefficient as a result of ex-

pansion of the moderator or the coolant. |t is not

 

33(:. S5, Burtnette, M. E. LaVerne, and C. B. Mills,
Reflector-ModeratedrReactor Design Parometer Study:
Part [. Effects of Reacfor Proportions, ORNL CF-54-7-5
{(to be issued).

#4This material was prepared with the assistance of

W. H. Jordan, E. 5. Bettis, and E. R. Mann,

351nferim Repart of the ANFP Controi Board for the
Aircroft Nuclear Propulsion Program, ANP-54 (Nov.
1950). '

implied that such recctors cannof be controlled or
that they are even inordinately difficult to control;
nevertheless, they do involve control problems
that do not occur in the circulating-fuel reactors.
The problems are summarized in the following
statements,

1. A solid-fuel reactor must have a large number
of shim rods to override xenon and to compensate
for fuel depletion. This entails much mechanical
gadgetry, as well as distortion of the flux pattern.
Distortion of the flux pattern, in turn, makes the
already difficult problem of hot spots much worse,
By contrast, the liquid-fuel reactor does not have
these problems because fuel can be added to give
a uniformly higher fuel cencentration to take core
of depletion, and xenon may be removed as it is
formed. _

2. While flux-sensing elements may be desirable
in a liguid-fuel reactor, they are so vital in o
solid-fuel reactor that they must be compounded,

3. Most of the proposed control systems for
solid-fuel-efement reactors include a fast-acting
servo-confrolled rod to compensate for quick
changes in reactivity, Such o rod is a hozard in
itself, since it might introduce a sharp increase in
reactivity. Probably the only satisfactory sclution
to this problem is to try to design the reactor so
that abrupt increases in reactivity cannot occur,

Even though step changes in reactivity are not
anticipated, they afford a useful basis for analysis
because a step change of the proper magnitude can
be introduced into an analogous system to simulate
most perturbations of practical interest. Thus, the
controllability of a reactor can be deduced from
the rise in fuel temperature that would result from
a step change in reactivity, This is particulorly
important in aircraft reactors where the operating
temperature is made as close as possible to that
likely to damage the reactor. When the response
of a reactor with a negative temperature coefficient
of reactivity (a) in the fuel is considered, it can
be readily shown that the maximum temperature
rise in the fuel (AT) as a result of a step change
in reactivity 5&/k is given by

Ok
2 —
k

AT =

 

o

Thus, if a =5 x 107%/°C and Sk/k = 3 x 1073,
AT will be 120°C or 216°F. |

39
The temperature rise to be expected in a solid-
fuel reactor can be approximated if it is ossumed
(1) that the heat capacity of the fuel element is
small, (2) that an increase in power produces a
corresponding increase in the film drop between
the fuel element and the cooclant, and (3) that the
increase in power caused by a step change in
reactivity is the transient term only, further in-
creases being stopped by a servo control, In this
case the power increase is given by 8k/&(3. Then

, Bk
(Qf ~ 0) = (1 +—16E> (9}, -~ 6) ,
where
0, = fuel temperature before the step change,
9;' = fuel temperature after the step change,
G = coolant temperature,

delayed neutron fraction.

It can be seen that the increase in fuel tempera-
ture over coolant temperature depends upon the
original difference between fif and 0 _; for example,

=
I

the ratio
07 ~ 0, Sk
- 0= +—0= 1.4
6/ — 6.: kP

for a Sk/k of 3 x 1073,

The power and temperature perturbations for a
sodium-cooled solid-fuel-element reactor were cal-
culated on the ORNL reactor simulator according
to the following conditions:

1. The volumetric heat capacity of the solid fue!
elements was 1.0 cal/cm?® °C.

2. The fuel-region power density at design point
was 5.7 kw/cm3,

3. The coolant was a liquid with thermal proper-
ties comparable to those of liquid sodium.

4, The design-point power was 200 Mw and the
coolant system was designed to extract power at
this rate.

5. The step perturbation, Ak/k, was 0.305%.

A servo system of reasonable proportions was
simulated, and it was presumed that the power
would be controlled from an error signal propoition-
al to (p — p,), where p was the power at any time
t and p, was the design-point power. The transient
responses in power and in fuel temperature are
shown in Fig, 24, where it is clear that the temper-
ature rise depends on the original difference in
temperature between fuel element and coolant,
Thus in a loosely coupled system, such as an air-

40

UNCLASSIFIED
ORNL-LR-DWG 2241

 

 

 

 

 

 

 

  

2600 : : ,
2400
— 2200
e
w
[ve
o
% 2000
o
Ly
o
=
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L
=
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L
=
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1200 b wmreeeeee e
g
z 300 ; | ! ,
o |
wl i
; !
§ 200 | — f
| i
| L ,|

 

 

 

 

0 5 10 15 20 25
TIME (sec)

Fig. 24. Power and Temperature Overshocts for
Step Changes in Reactivity of 0.305 Ak/k for a
Sodivm-Cooled Solid-Fuel-Element Reactor with

an Average Power Density in the Fuel of 5.7
kw/em>.

cooled reactor in which the fuel element tempera-
ture must be much higher than the coolant tempera-
ture, the temperature rise would be much more
severe than in the sodium-cooled reactor.

MATERIALS

The various moderoting materials that might be
employed were discussed in the previous section
because nuclear considerctions are dominant in
their selection. This section covers structural,
fuel element, and coolant materials, the selection
of which is usually based mainly on engineering
considerations.

Structure .

A key factor in the design of a reactor is the
structural material of which it is to be built, An
indication of the structural materials that might
be employed in a high-temperature nuclear power
plant may be gained from an examination of the
program carried on during the past 15 vears for
the development of superior materials for gas-
turbine buckets, The most frequently used re-
fractory alloys have been those of iron, chromium,
and nickel, particularly the 18-8 stainless steels
and Inconel. A group of alloys that give even
better high-temperature performunce are cobalt-
base alloys containing various amounts of iron,
chromium, nickel, molybdenum, and tungsten. Un-
fortunately, cobalt has o high neutron-absorption
cress section and becomes an exceptionaily bad
source of gammas it exposed fo thermal neutrons.
if even trace amounts of cobalt were carried out-
side the shield in a fluid circuit they would be
serious sources of radiation. . Both ceramic ma-
terials and cermets have also been employed, but
their brittleness haos led to difficulties; they have
yet to be developed to the point where they are
capable of withstonding the severe thermal stresses
imposed in turbojet engines.

All the materials mentioned above were con-
sidered because of their oxidation resistance,
However, in certgin types of reactor it would be
possible to employ refractory materials such as
molybdenum, columbium, and graphite in an ombient
completely free of oxygen, for example, a molten
metal. Further, it is conceivable that a completely
new refractory alloy might be developed from such
high-melting-point materials as molybdenum, tung-
sten, columbium, zirconium, chromium, and vana-
dium. The recent development of iron-aluminum-
molybdenum alloys, such as Theromatfor, lends
credence fo this possibility.

A particular system must be examined in order
to evaluate the relative merits of the various
structural materials, but, in general, the structural
metal should have both high creep strength at high
temperatures and ductility throughout the operating
temperature range of at least 2 or 3% so that high
focal thermal stresses will be relieved by plastic
flow without cracking. It further seems necessary

in most instances that the structural metal be
highly impermeable and weldable, with ductility in
the weld zone of at least 2 or 3% throughout the
temperature range from the melting point to room
temperature. Some of the materials that have been
considered for use in aircraft nuclear power plants
are listed in Table 8, together with their significant
properties for this application., The availability
of the material is o most important consideration
in the conduct of o development program, because
a good assortment of bar stock, tubing, and sheet
is essenfial to the fabrication of test rigs, It has
been primarily the availability consideration that
has led to the use of iron-chrome-nicke! alloys in
most of the development work to date. It is hoped,
however, that better materials will be available
for future, more advanced reactors.

The effects of temperature on the stress-rupture
properties and the creep rates of some typical
metals and alloys are shown in Figs. 25 through

UNCLASSIFIED
ORNL-LR~DWS 2010

 

    
  
  

100 e Sty
T

. f ‘
T 8% af 1800°F
90 o o ] | f . i
80 . e e ol e

INCONEL -~

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o

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ELONGATION

 

- HASTELLOY © (AS CAST)

f  HASTELLOY B
J N '

{ \ * |
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i R
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RO e

 

 

 

 

 

 

 

 

1200

1200 1400
TEMPERATURE  (°F)
Fig, 25. Effect of Temperature on Elongation

of Yorious Materials.
28.3%  Unfortunately, these curves do not tell the
whole story., Al the iron-chrome-nicke! alloys
over-age at temperatures above 1650°F, because
the hardening constituents, such as the carbides,
tend to migrate to the grain boundaries. Annealing
and grain growth inevitably accompany over-aging.
Intergranular corrosion would be likely to follow
and would probably cause trouble in thin sections
where a grain might extend all the way through a

0.010- to 0.20-in.-thick sheet or tube wall.

Solid Fue! Eiements

While the bulk of the ORNL-ANP effort since the
fall of 1951 has been directed toward the develop-
ment of a circulating-fluoride-fuel reactor, the

 

3‘SJ. M. Woods, Mechanical Properties of Metals and
Alloys at High Temperatures, ORNL-1754 (to be issued).

UNCLASSIFIED
ORNL-LR-DWG 2011

 

 

20

w

 

RUPTURE STRENGTH {psi x 10"3) FOR 1000 nr
3

      

L e g oo b e e e e e o e e o o T

WIRE-SPAGED PLATE-TYPE FUEL ELEMENT
N T S L
1200 1400 1600 1800 2000
TEMPERATURE (°F)

 

Fig. 26. Effect of Temperature on Rupture
Strength of Yaricus Materials.

42

major effort prior to that time was on the develop-
ment of reactors utilizing stationary fuel elements,
The work on solid fuel elements is continuing, but
on a very limited basis, so that another avenue of
approach to the high-temperature aircraft reactor
may be kept open.

The fissionable material for a high-power reactor
with stationary fuel elements may, in general, take
the form of uranium metal, uranium metal alloy,
uo,, UC,, or, possibly, other uranium compounds.
However, at high temperatures serious difficulties
are encountered because of the low melting point
of uronium metal and most of its alloys.?7 The
melting point of pure uronium is about 2066°F, that
is, not much above the proposed fue! surface
temperature for the reactor, so that the metal would
be so weak at operating temperature as to require
additional support, The support might be provided

 

37R. W. Bussard and H. E. Cleaves, Journal of Metal-
lurgy and Ceramics, Yol. 1, No, 1 (1948),

UNCLASSIFIED
ORNL-LR -DWG 2012
25 :_"'"""“—h”’ ““‘“*T—_"""" "*—_N——‘”'I
: | ! i |
! | I
\ v 66 AT 1200°F |
i i i

     
   

il

 
  
    
 

< MOLYBDENUM (10Q hr)

 

| | | N ‘
~+-TYPE 316 ‘ N !
10 | - \-— STAINLESS STEEL | N
i {1000 hr) ! *X«“

 
  

STRESS TO PRODUCE 1.0% CREEP {psi X 107%)

 

WIRE ~SPACED

LEZTE— TYPE FUEL ELEMENT
|

0 Lb——d | |

{

200 1300 1400 1500 1600 1700 1800 1300
TEMPERATURE (°F)

Fig, 27. EHect of Temperature on Siress to
Produce 1.0% Creep for Yarious Materiols,
TABLE 8, PROPERTIES OF REFRACTORY METALS AND ALLOYS

 

 

THERMAL.NEUTRON

COEFFICIENT OF

 

MELTING |  DENSITY ABSORPTION MODULUS OF 1 e CMAL EXPANSION | SPECIFIC HEAT THERMAL cosT
POINT NEAR 20°C CROSS SECTION ELASTICITY bER OF (cal/a/%0) CONDUCTIVITY AT 70°F {WELDABILITY | AVAILABILITY 5716}
2 0 ;
o} (g/ cm> i callg Btu/hr 12 (°F /1)
( ¢/cm) {barns/ atom) (psi) (in./in.) o

Tungsten 6170 19.3 19.2 52 x 108 2.4 x 10~9 0.032 96 Poor Poor 10.43 (ingot)
Tantalum 5425 16.6 21.3 27 %108 3.6 x 1079 0.036 31 Good Fair 39 (sheet)
Molybdenum 4760 i0.2 2.4 48 % 10° 2.7 x 10~ 0.061 85 Poor Poor 4 {pressed ingot)
Niobium 4380 8.57 1.1 18 x 108 4.0 x 10~°% 0.065 Good Fair 75 {powder)
Vanadium 3150 5.1 4.7 21.5 x 106 4.3 % 10_6 0.15 17 No data Poor 30
Zirconjum 3200 6.5 G.18 11 x ]06 3.0 x 10‘-6 0.08 14 Fair Fair 35
Titanium 3300 4.54 5.6 15 x10° 4.7 x 108 0.13 100 Fair Fair 15 (sheet)
Chromium 3430 7.19 2.9 3.4 x 1076 011 39 Bad Difficult 3.60 (electrolytic)
Iron 2802 7.87 2.43 29 108 6.5 % 10~9 0.11 36 Good Good 0.11 to 1.48
Cobalt 2723 8.9 34.8 30 x 10° 6.8 x10~% 0.099 40 Poor Difficult 2.60
Nicke! 2650 3.90 4.5 30 x10° 7.4 x10~% 0.105 34 Good Good 0.865 (sheet)
Nichrome V 2550 8.4 30 x 108 9.8 x 10~ 0.107 7.8 Good Fair 1.00
Inconel 2600 8.51 4.0 31 x 108 6.4 % 108 0.11 8.7 Good Good 0.925 (sheet)
Incone! X 2600 8.3 4.0 31 x 108 16.0 x 10~% 0.13 8.5 Fair Good 2.75
Type 316 stainless steel 2550 8.02 2.9 28 x 106 9.7 x 10~ 0 G.12 9 Good Good 0.645 (sheet)
Hastelloy B 2900 9.24 3.9 30.7 % 10° 5.6 x 1079 0.091 6.5 Fair Fair 2.50
Zircaioy-2 (1.44% Sn; 6.55 0.25 13.8 x 108 6.5 x10~% 0.08 8.2 Foir Fair 35

0.05 Ni; 0.12 Fe; 0.11 Cr}

 

 

 

 

 

 

 

 

 

 

 

43
 
UNCLASSIFIED
ORNL -LR-DWG 2C13

 

 

 

 

T e

 

 

FATIGUE STREMGTH {psi x 1073

 

 

 

 

 

 

 

 

 

 

 

10 Jrmeremesmnen —
o
1200 1400 1600 1300
TEMPERATURE (°F)
Fig. 28.  Effect of Temperature on Fatigue

Strength of Various Materials,

by using rods of solid moderator (such as BeQ)
cooted with uranium metal, but the uranium coating
would be very thin and would be very likely to
break up and spall off because of thermal stresses.
Even worse, .the metallic uranium would migrate
by diffusion and mass transfer in the coolant to
the walls of the pressure shell, heat exchanger
tubes, ete., where it would tend to diffuse into the
base metal and form o low-melting-point eutectic
in the grain boundaries. The eutectic of iron and

vranium melts at 1337°F, and the eutectics of

uranium with nickel and chromium melt at tempera-
atures well below 1800°F.
alloy mefts at 2345°F over most of the composition

Uranium-molybdenum

range, and thus the presence of uranium metal would
seriously molybdenum,
About the only metal that does not form a low-
melting-point eutectic with uranium is columbium,
but this material is expensive and difficult to

reduce the strength of

procure and fabricate. Cladding or canning metal-
lic uranium or its alloys would serve to reduce
the diffusion rate but would not reduce it suffi-
ciently ot the operating temperatures involved,

The difficulties associated with the low melting
points of uranium metal and uranium alloys, can
be avoided by introducing the uranium as uao,,
which is a chemically stable material with a very
high melting point, 3949°F, Uranium carbide might
also be used, but it is less stable chemicaolly, and
it would react with most moderators, coclonts, or
canning materials at the temperatures considered
here. Therefore, UC, seems quite inferior to UQ,,
except, perhaps, on the basis of thermal conduc-
tivity and resistance to thermal shock. Other
uranium compounds have been considered, but none
appears to be superior to UG, ;

Uranium oxide can be fabricated into fuel eles
ments in a number of ways. For an air- or helium-
cooled reoctor it might be contoined in o matrix
of chromium and Al;0,, in the form of o ceramel,
or in a matrix of silicon carbide, in the form of a
ceramic. Since o large surface area is essential,
the fuel elements could be in the form of thin flat
plates or tubes or a pebble bed, The support of
such fuel elements would be difficult if they were
to be used at high temperatures. [ thin plates or
tubes of o ceramic or a ceramel were held rigidly,
they would be virtually certain to crack under
thermal stress; if they were supported loosely,
they would flutter in the high-velocity gas stream
and fail as a result of abrasion of the contact
surfaces, If o pebble bed were used, the same
difficulties would be encountered, In any case the
support structure would have to be metai, While
the metal could be cooled, it is hard to see how
hot spots could be prevented if the ceramic or the
ceramel were af operating temperatures much above
the temperature of the metal,

The UO, might be poured Ioosely into long
siender metal tubes or pins, but at high power
densities the temperature at the cenfer of even
0.080-in.-1D pins would exceed the melting point

45
of the UO, and cause fusion. Hence, there would
probably be objectionable concentrations of UO,
that would create hot spots at indetferminate regions
in the pin. A better arrangement would appear to
be to place a thin layer of UO, on the inside of
the tube wall, as proposed in the KAPL-SIR de-
sign.>® The problem of supporting and accurately
spacing these pins or tubes is a most serious one,
however, as has been clearly shown by experience
at KAPL.

Probably the most promising way to fabricate
U0, into a fuel element is to clad with stainless
steel o sintered compact of UO, and stainless
steel3? in which the U0, may constitute as much
as one-third of the volume. Sandwiches of this
type can be rolled to give plates with minimum
thicknesses of 0.006 in. of cladding and 0.008 in.
of UQ. compact in the core. One obvious way to
use such fuel plates would be to stack alternate
flat and corrugated plates to give the grrgngement
shown in Fig. 29. The coolant would flow between
the corrugations. This arrangement has the dis-
advantage, particularly when vsed with low-thermal-
conductivity coolants, of giving hot spots in the
low-velocity regions in the vicinity of the points
of contact between the flat and corrugated plates,
Short spacers containing no fuel can be placed
between the corrugated and flot shests to avoid
this, as in the arrangement shown in Fig. 30. In
a third arrangement, shown in Fig. 31, wire spacers
are passed perpendicularly through flat plates at
intervals sufficiently close to maintain good
spacing in spite of tendencies toward thermal
distortion. A fourth arrangement, shown in Fig.
32, is based on the demonstrated procticality of
fabricating the UO, stainless compact in the form
of tubes. This arrangement gives o fue! element
that is very resistont to warping and thermal
distortion, Yet another arrangement, shown in
Fig. 33, depends on the use of UQ, packed into
small-diameter tubes which can be drawn or swaged
to give wires as small as 0.020 in. in diameter,
Ancther arrangement, shown in Fig. 34, employs

 

38K nolls  Atomic Power Laboratory, Reactor Engi-
neering Progress Report July, August, September, 1951,
KAPL-614, p. 13,

39G. M. Adamson, ANP Quar. Prog. Rep. June 10,
1951, ANP-65, p. 181; E. 5. Bomar and J. H. Coabs,
ANP Quar, Prog. Rep. Sepf. 10, 7957, ORNL-1154, p.
147; E. S. Bomar and J. H. Coobs, ANP Quar,.Prog.
Rep, Dec, 10, 1951, ORNL.-1170, p. 128; E. 5. Bomar,
J. H. Coobs, and H. Inouye, Met. Div. Semiann. Apr.
10, 1953, ORNL-1551, p. 58.

46

e ORN_-LR-CW6 1441

 

 

 

 

COOLANT FLOW

Fig, 29. Corrugated Plate Type of Fuel Eiement.

   
   
 

NON FUEL - BEARING
SPACER BARS

  

N Tl = ‘Jr; T
/'4,;,\\’\__/;?_ S o o
B T e, ,,". B At j ]
B g gy

COOLANT FLOW

Fig. 30. Corrugated Plate Type of Fuel Element
with Nonfuel-Bearing Spacer Bars,

sintered blocks of UQ, and stainless steel compact
in which a closely spaced hole pattern would
provide coolont flow possages and heat transfer
surface area, If erosion or spalling should prove
a problem with this arrangement, the holes might
be lined with thin-walled tubes and the gap be-
tween the blocks and the tube walls might be filled
with a molten metal, such as sodium, to provide
@ good thermal bond.

Some idea of the amount of core wvolume that
must be devoted to the fuel elements can be gained
from an illustrative example. If the critical mass
for a reactor were 50 Ib of UZ3° and a sintered
stainless steel matrix containing 33 vol % UO,
were employed, the ceramel matrix volume would
have to be about 0.25 ft3. The volume of cladding
HPNL—I!'!WG a3 . m.’

ORNL -LR- Biv5 1145

     

—7 WIRE SPACERS

e
. 71\ STAINLERS SYEFL CLADDING — -

   

\SINTEHED LG, AND STAINLESS STEEL CORE
Fig. 33. Wire Type of Fuel Element,

SECRET
~ Ol ~[ R -[CWG {148

e

i :

AT

S . P o .
COOLANT FLOW \:\\W P o “

Fig. 31. Wire-Spaced Plate Type of Fuel Element,

     

~.

SELRET
ORNI - LR~DW5 1144

   
 

< STAINLESS STEEL

%

%.CLADDING G
e

 

Fig. 34. Sintered UO, and Stainless Steel Block
Type of Fuel Element, 3

fn summary, a good detoil design for a solid-fuei-
element system for an aircraft reactor should pro-
vide the following:

<
SINTERED UG, AND STAINLESS STEEL CORE

_ . ' : 1. an adequate volume of UD, to insure criticality,
Fig. 32. Sandwich-Tube Type of Fuel Element. 2. adequate surface area to meet heat transfer

requirements,

material required to provide adequate surface arec 3. a surface that would not give trouble with cor-
to meet heat trensfer requirements usually proves rosion, mass transfer, erosion, or spalling, or
to be about the same as the matrix volume, and have a tendency to pick up surface films that
therefore the volume of material in the fuel ele- would impede heat transfer,

ments would be about 0.5 ft3, This would consti- 4. a geometry that would give a fairly uniform
tute 12% of ‘the volume of a 2-ft-dia spherical temperature distribution throughout the fuel
reactor core. . element and avoid both excessive temperatures

47
in the interior and thermal stresses that would
induce cracking or warping under power- and
temperature-cycling conditions,

5. adequate strength and stiffness to insure struc-
tural integrity and the surface spacing required
by heat transfer considerations so that hot
spots could be avoided,

6. a fuel element that could be consistently fabri-
cated with the requisite quality at reasonable
cost,

7. structural material in an amount consistent with
a reasonable critical mass requirement,

High-Temperature Liquid Coclants and
Fuel-Carriers

A thorough survey of materials that appear prom-
ising as heat transfer fluids for high-temperature
aircroft reactors was presented in ORNL-360.40
The first requirement is that the fluid must be
liquid and thermally stable over the temperature
range from 1000 to 1800°F. A melting point con-
siderably below 1000°F would be preferable for
ease in handling, while a substance that would be
liquid at room temperature would be even better,
Other desirable characteristics are low neutron
absorption, high volumetric specific heat, and high
thermal conductivity. Above all, it must be pos-
sible to contain the liquid in a good structural
material at high temperatures without serious
corrosion or mass transfer of the structural ma-
terial, The principal substances so far suggested
that show much promise of satisfying these re-
quirements are listed in Table 9, together with
Of these ma-

terials, sodium hydroxide, lead, and bismuth are

some of their physical properties.

considered to be only marginally useful because
of their corrosion and mass transfer characteristics,
The promising liquid metals can be separated into
a light group, lithium and sodium, and a heavy
group, lead and bismuth, As will be discussed in
the next section, the light metals are highly
preferred because of their superior heat transfer
properties and corrosion and mass transfer charac-
teristics.

Liquids intended to serve as vehicles for uranium
in circulating-fuel reactors that operate at high
temperatures are subject to the same criteria as
those that serve as coolants. In addition, the

 

4OA. S. Kitzes, A Discussion of lLiquid Metals as
Pile Coolants, ORNL=-360 (Aug. 10, 1949).

48

solubility of uranium and its effect on the physical
properties of the fluid must be considered. All
the fluids in Table 9, except the fluorides, can be
shown to be unsuitable, for one reason or another,
as vehicles for uranium. Fortunately there are
many different fluorides that can be used, 4! The
NafF-ZrF, melt (NaZrf;) was chosen for the ARE
because the materials were readily available,
nontoxic, and not too expensive. Unfortunately,
the physical properties of this fluoride mixture,
particularly the melting point, vapor pressure, and
viscosity, leave much to be desired. Both BeF,
and LiF can be used to reduce the melting point,
but BeF, is toxic and LiF would require Li7, a
material that is not avoilable, although it could
be obtained at a price that should not be unreason-
able.  Other promising components are KF and
RbLF; however, KF has a neutron absorption cross
section that is higher than is desirable (Tabie 7),
while RbF

commercial demand for it,

is expencive because there is no
It has been determined
that ample stocks of rubidium-contcining ore are
available, and the price of Rbi should not be un-
reascnable if substantial amounts are ordered.

The terms corrosion and mass transfer need some
clurification. Corrosion implies the remaval of
surface material from the container by a chemical
reaction with the liquid or by simple solution in
the liquid, as is the case with liquid metals, Cor-
rosion damage to a solid material caused by contact
with a fluid results in a loss in strength of the
solid material, Mass transfer in liquid metals is
a phenomenon that involves removal of container
material from the hotter portion and deposition in
the cooler zane of a closed circuit with a tempera-
ture gradient in which the liquid is being circu-
lated.
variations in solubility as a function of tempera-

The removal and deposition result from

However, when the circulating fluid is a
fused salt, the container material is transported

ture,

from the hotter to the cooler zone of the circuit
because of variations in the equilibrium constanis
of the chemical reactions as functions of tempera-
ture. tor example, in an Inconel system circu-
lating a fused-fluoride-salt fuel, the differences
in chemical equilibria at the two temperature zones

may cause mass transfer of chromium according to

41w, R, Grimes and D. G. Hitl, High-Temperoture Fuel
Systems, a Literuiure Survey, Y-657 (July 20, 1950);
The Reactor Handbook, Vol. 2, Sec. 6, p. 9215 (1953).
TABLE 9. PROPERTIES OF REPRESENTATIVE REACTOR COOLANTS

 

 

MACROSCOPIC

 

BOILING
MELTING POINT, THERMAL VISCOSITY SPECIFIC HEAT DENSITY VOLUMETRIC THERMAL-NEUTRCN PREFERRED
COOLANT POINT 7 CONDUCTIVITY o /o3 HEAT CAPACITY ABSORPTION CONTAINING REMARKS
5 60 mm 2 o (cp) {cal/g/"C) (g/em™) o 3
(°F) (°F) [Btu/hr-#t°-(OF /)] (cal/°Crem®) CROSS SECTION MATERIAL
(em™ )
Li’ 354.0 2403 25.0 0.4 1.0 0.46 0.46 £.00131 Type 430 stainless Severe mass transfer above 1150°F
steel
Na 208.0 1621 34.5 0.2 0.30 0.78 0.23 0.0092 Type 316 stoinless Virtuatly no corrosien or mass transfer up to 1600°F
steel or Incone!
NaK (56% Na, 66.2 1518 16.7 0.161 0.253 0.742 0.188 0.0183 Type 316 stainless Virtually no corrosion or mass transfer up to 1600°F
44% K) steel or Inconel
Pb 621.0 3159 8.6 1.2 0.037 10.0 0.37 0.00592 Type 430 stainless Severe corrosion* and mass transfer above 1150°F
‘ steel
Bi 520.0 2691 9.0 1.0 0.039 9.4 0.367 6.000406 Type 430 stainless Severe corrosion and mass transfer above 1150°F
steel
NaQOH 0.7 1.0 0.49 1.7 0.83 0.021 Nickel Severe corrosion and mass transfer above 1150°F
NerF5 950.0 2.0 7.5 0.29 3.0 0.87 0.00367 Inconel No corrosion or mass transfer up to 1500°F when
used as a vehicle for UF3
NaoF-KF-LiF 851.0 2.5 2.5 0.40 1.9 0.76 tnconel Prabably no corrosion or mass transfer up to 1500°F
when used as a vehicle for UF3
HZO {100°F) 32.0 212 0.35 0.7 1.0 1.0 1.0 0.048 Type 347 stainless Severe corrosion above 1650°F**
steel
Air (sea level) 0.04 0.04 0.26 0.00032 0.00008 0.00002 Type 310 stainless Severe corrosion above 1800°F

 

 

 

 

 

 

 

 

 

steel or Nichrome V

 

 

*The term '‘severe corrosion’’ is used where the attack exceeds a depth of 0.010 in. after 500 hr of testing, because in most reactors the thickness of many structurcl elements must be less than 0.025 in,

**G, H, Hawkins et al., Trans. Am. Soc. Mech., Engrs. 65, 301 {1943).

49
 

 
the reaction

Cr + 2UF, =2UF, + CrF, .

Another set of reactions, known as dissimilar metal
transfer, makes it desirable that complex plumbing
systems be fdbricated entirely from one metal or
alloy. Dissimilar metal transfer involves the re-
moval of one or more of the constituents of one
alloy and its transport through the liquid to another
alloy where the deposited material diffuses into
the base metal. The transport driving force in this
case is a difference in chemical potential; the
chemical potential of a constituent of a complex
alloy is lower than that of a pure metal or of a
simple-solution alloy.  Examples of dissimilar
metal transfer have included the plugging of nickel
heat exchanger tubing by iron which was trans-
ported to the nickel surface through the liguid
medium from a stainless steel pump chamber and
a stainless steel expansion tank.

HEAT REMOVAL

The power density in the reactor core is limited
by the rate at which heat can be removed by the
fluid passing through the core.
rate depends, in turn, on the permissible fluid

The heat removal

velocity and the temperature rise through the re-
actor and on the density and specific heat of the
heat transfer fluid, The optimum coolant tempera-
ture rise depends upon the characteristics and
proportions of the over-all power plant, While the
relations -are quite complex and depend in large
measure upon the characteristics of the various
components of the system, for most aircraft reactor
types the optimum temperature rise for the fluid
passing through the reactor core appears to be of
the order of 400°F.42 In fact, a temperature rise
greater than 600°F has been proposed for only one
of the detailed major cycle proposals made to
date — the air cycle. For the air cycle the allow-
able temperature rise will be the difference be-
tween the maximum reactor air outlet temperature
obtainable and the turbojet compressor outlet tem-
perature. The resulting temperature rise is likely
to be of the order of 600°F, depending on the com-
pression ratio.

Once a permissible femperature rise is estab-
lished and a coolant is chosen, a major limiting

 

425 M. Walley, W. K. Moran, and W. Graff, Off Design
Turbojet Engine Performance of a Nuclear Powered

Aircraft, ORNL CF.53-9-80 (Aug. 1953).

factor for solid-fuel-element reactors is the per-
missible pressure drop across the reactor core,
While there is some variation in the pressure drop
associated with different types of fuel element and
different reactor core arrangements, it appears, in
general, that the pressure drop across the core
should be kept to something of the order of 30 to
50 psi because of limitations imposed by pumping
power and fuel-element stress considerations.

A third important facter associated with heat
removal from the core of a solid-fuel-element re-
actor is the difference in temperature between the
fuel element and the coolant.  The higher the heat
transfer coefficient obtainable, the lower this
temperature difference becomes.  In attempting the
detailed design of any particular reactor, it soon
becomes evident that, regardless of how desirable
an increased amount of heat transfer surface area
may be, the problems associated with the fabri-
cation of the fuel elements become progressively
greater as the amount of heat transfer surface area
per unit of volume is increased and the structure
becomes progressively more delicate and ““lacey.”’
In almost every instance the inclination is to de-
crease the hydraviic radius of the coolant passage
to a value as low as possible consistent with
problems of fabricating the fuel-eiement surfaces
and with stress considerations associated both
with the fluid pressure drop and the thermal
stresses that would produce thermal distortion.

A third factor affecting the power density ob-
tainable from a reactor core is the free-flow ratio,
that is, the ratio of the effective flow-passage area
to the total cross-sectional arec of the reactor
core. For most reactors the moximum practical
value for this parameter appears tc be about 0.40,
but for the reflector-moderated high-temperature-
liquid type it appears to be closer to 0,60, and for
the circulating-moderator type it appears to be of
the corder of 0.85. Water-moderated reactors in
which water is not the prime heat transfer medium
can be designed for free-flow ratios as high as
0.50 because water is so potent a moderator,

{f an attempt is made to get the maximum power
density from a given core matrix geometry with a
given coolant, it soon becomes evident thot any
actual reactor must be expected to differ consider-
ably from the commonly assumed ideal reactor in
which there is ¢ perfectly uniform fiuid fiow distri-
bution and the core matrix is stressed by a perfectly

vniferm loading.  Careful consideration of the

51
usual perversities of velocity distribution under
turbulent flow conditions disclosed marked devi-
ations from ideal conditions. Also, the ideal con-
ditions are clearly unreasonable if allowances are
made for ordinary amounts of thermal distortion.
Experience in brazing radiator core matrices, for
excmple, has shown that even the relatively slow
rates of temperature change associated with the
furnace brazing operations produce variations in
passage thickness of as much as 30%. Thus it is
felt that even if great care is taken in the design
to minimize the cumulative effects of thermal
distortion and fabrication tolerances, variations in
effective thickness of the coolant passage of the
order of at least 20% will occur. Heat transfer
analyses show that variations in coolant passage
thickness would lead to the formation of hot spots,
and thus in some regions the local temperature
difference between the fuel-element surface and
the coolant would be greater than the design value,

Substantial variations in power density through-
out the core matrix can be expected to result from
nonuniform fission densities, since even in the
ideal reactor, there would be variations in fission
density because of the effects of geometry on the
neutron flux. Allowances must be made in any
actual reactor for additional irregularities caused
by the presence of control rods and by the non-
uniform distribution of the fission-product poisons
that will accumulate, |t therefore appears that
local power densities at least 50% greater than
the mean power density must be expected. |If
further allowance is made for vagaries in flow
distribution and for irregularities in channel shape
as a result of thermal distortion, it would seem
that in a realistic design, local temperature dif-
ferences hetween the fuel element and the coolant
of at least twice the mean should be anticipated.

A careful examination of the siress anclysis
problem for any core matrix shows that the same
basic reasoning must be applied as that applied to
thermal distortion. Heat removal reguirements for
the maximum available flow passage area and the
maximum possible heat transfer area, coupled with
nuclear requirements to minimize neutron absorption
in structural material, have led to u relatively
complex, finely divided structure in every design
proposed to date. Since a complex structure is
inherent in o solid-fuel-element reactor, the
stresses induced in the fuel elements and their
supports by the pressure drop acioss thes core

52

matrix always constitute o problem. The most
probable cause of failure would be the fotigue
stresses arising from the pressure fluctuations
associated with the turbulent flow of the coolant
through the core matrix, Just as in the blades in
turhojet engines, the stresses induced would
probably be two or three times the direct stresses
indicated by the average pressure drop across the
core matrix,

As was pointed out in the previous section, the
iron-chrome-nicks! alloys are the only structural
materials from which it seems reasonable at this
time to expect to fabricate fuel elements. The
data available on the high-temperature strength of
these alloys indicate that even if the siresses are
kept low, the permissible operating temperature
can scarcely exceed 1800°F. The strength proper-
ties of Inconel and other possible structural ma-
terials are presented in Figs. 25 through 28 as
functions of temperature, Dotted lines on Figs. 26
and 27 show the stresses unticipated in one of the
most favorable fuel element matrices devised to
date, that is, the wire-spaced plate-type fue!l ele-
ment shown in Fig. 31. A consideration of the
safety factor that would be acceptable for so vital
a structure as a reactor core indicates that the
maximum allowable operating temperature of a
solid fuel element would probably be between
1600 and 1800°F,

Since the wire-spaced plate-type fuel element
(Fig. 31) is representative of the possible solid
fue! elements, it was used as a basis for com-
paring the characteristics of various potential re-
actor coolaonts. For this study it was assumed that
the plates were 0.020 in. thick and spaced on
0.120-in. centers and that the wire spacers ob-
structed 5% of the effective flow passage area,
The operating conditions assumed were a fluid
temperature rise of 400°F, a fluid pressure drop of
50 psi, and a limiting fuel-element-metal tempera-
ture of 1700°F. The limiting fluid outlet tempera-
ture was to be determined by hot-spot consider-
ations; that is, the maximum permissikble tempera-
ture differential between the fluid and the fuel
If a higher
limiting fuel element temperature were wused, a
lower fluid pressure drop would sesm to be neces-
sary. Table 10 shows the results of a set of
calculations based on these assumptions.

element was to be twice the mean.

Air, as a coolant, was ireated as a special case.
The limiting flow velocity for the air was de-
TABLE 10. LIMITING POWER DENSITIES FOR VARIOUS REACTOR COOLANTS

 

 

HEAT-TRANSFER-LIMITED

 

ANT P VELOCITY HEAT REMOVED PER COOLANT-FLOW-LIMITED | HEAT TRANSFE EAN TEMP E | POWER DENSITY FOR A
REACTOR COOLL:L A:SAGE FLUID DENSITY FOR SPECIFIC HEAT | UNIT OF PASSAGE | FREE-FLOW PowTE-; DEN';;YT D CO-II:Z:FICTE;TR M DllFEREiéEUR ocar TEMP;RAFTURE
COOLANT G7 (g/cms) 50-psi Ap (col/g/OC) FLOW AREA RATIO 3 2 o o
(in) 2 o (kw/ em3) (Btu/hr-#12:F) (°F) DIFFERENCE OF 100°F
{ft/sec) {Btu/sec-ft°.°F) 3
{kw/em”)
L7 20 0.46 68 1.0 1960 0.6 10.4 55,000 128 8.1
Na 20 0.78 53.5 0.30 782 0.6 4.17 39,300 71.6 5.7
NaK 20 0.742 55 0.25 636 0.6 3.4 24,100 95 3.6
Py 20 10.0 14.9 0.037 343 0.6 1.83 12,800 96.5 1.9
Bi 20 9.4 15.2 0.039 347 0.6 1.85 13,000 95.5 1.9
NaOH 30 1.7 29.5 0.49 1530 0.85 1.6 6,100 601 1.9
NaZrFg 30 3.0 22.2 0.29 1206 0.6 4.27 8,000 362 1.18
NaF-KF-LiF 30 1.87 28.7 0.4 1340 0.6 4.76 16,000 201 2.4
H,0 (100°F, no boiling) 30 1 38.8 1 2420 0.85 18.3 1,200 4840 0.38
Supercritical water™ 30 0.4 0.70 200* 0.70
Air (sea level) 40 0.0059 350+ 0.26 33.4 0.5 0.088 360 198 0.088
Air (45,000 f) 40 0.00134 322% 0.26 7.0 0.5 0.027 18 187 0.027

 

 

 

 

 

 

 

 

 

 

 

*Supercritical water calculations were based on heat transfer and pressure-drop data given in Pratt & Whitney Aircraft Div., Nuclear Propulsion Program Engineering Progress Reports, No, 9, PWAC-75 and No. 10, PWAC-83, The limiting temperaturs difference

of 200°F was the temperature difference in the outlet region.

**Mach .20 at inlet.

53
termined by a compressibility loss consideration;
that is, the Mach number was controlling. In all
other cases the fluid velocity was computed to
give an ideal pressure drop across the fuel element
of 30 psi, and an additional 20 psi was assigned
to the spacers and supports for the fuel plates to
give an over=all pressure drop of 50 psi. The
temperature rise in the air also required special
trectment. |t was taken as being equal to the
difference between the compressor outlet tempera-
ture given in APEX-943 and 1700°F minus twice
the mean temperature difference between the fuel
element surface and the air,

It is evident from columns 9 and 10 of Table 10
that, except for air and the liquid metals, the
principal limitation on reactor power density is
the temperature differential between the fuel ele-
menf and the coolant rather than the rate at which
coolant can be forced through the fuel element
matrix,  Therefore the mean local temperature
differential between the metal surface and the
coolant was specified as 100°F so that the peak
fuel-element-surface temperature would be 1700°F,
the average fuel-element femperature would be
1600°F at the coolant-outlet face, and the average
coolant outlet temperature would be 1500°F. The
resulting heat-transfer-limited power densities are
given in the last column, It can be seen that on a
heat removal basis for a consistent set of con-
ditions lithium is clearly the best reactor coolant
and that sodium is a close second, while air is
the poorest in that it requires a reactor core
volume 50 times greater than that required by
sodium for a given power output. A remarkable
point is that the molten salts are actually superior
to the heavy liquid metals as heat transfer me-
diums. H. F. Poppendiek and M. W. Rosenthal are
preparing a report covering o more sophisticated
and complete analysis than that given in Table 10.
In their work they also varied the hydraulic radius
of the heat transfer passages; however, their work
leads to essentially the same conclusions as those
presented here,

An important point for which no allowance was
made in the above analysis is that the local heat
transfer coefficient is much less sensitive to
vagaries in the local fluid velocity for molten
metals than for the other coolonts. This makes

 

435 eneral Electric Co., Aircraft Nuclear Propulsion,
Department of Engineering, Progress Report No. 9,
APE X9 (Sept. 1953).

the molten metals definitely more desirable be-
cause with them the likelihood of hot spots and
thermal distortion would be reduced. Unfortunately,
however, of the good heat fransfer mediums only
sodium, NaK, and the molten fluorides can be used
for periods of 100 hr or more at temperatures of
around 1500°F in any structural material currently
available and fabricable,

TEMPERATURE GRADIENTS AND
THERMAL STRESSES

Thermal stresses have been referred to o number
of times in previous sections. These stresses may
be induced by a temperature difference between
two fluid streams, as in the tube walls of a heat
exchanger, or by a temperature difference between
the surface and the core of a solid in which heat
is being generated.?? FExamples of the latter are
fission heating in solid fuel elements and gamma
and neutron heating in solid moderator materials,
A calculated thermal stress gives a good indication
of the behavior of a brittle material; that is,
cracking is likely to occur if the calculated thermal
stress exceeds the normal tensile or shear strength
of the matericl. Only a small amount of yielding
is necessary in a ductile material, however, to
relieve the thermal stress. Therefore the calcu-
lated thermal stresses for ductile materials are
significant only in that they indicote that if the
elastic limit of the material is exceeded, plostic
flow and, possibly, distortion will result. Progres-
sively greater distortion may result from thermal
cycling, This might lead, for example, to partial
blocking of a flow passage between adjacent
plates in a solid-fuel-element assembly. Thus
thermal stresses in o fuel plate might lead to a
hot spot and hence to burn-out of a fuel element,

For most purposes, thermal stresses can be
approximated by considering one of two ideal
configurations, namely, flat slabs and thick-walled
cylinders with uniformly distributed volume heat
sources, Charts for the simpler flat-slab configu-
ration are presented in Figs. 35 and 36 to show
both the temperature ditference and the thermal
stress between the surface and the core for the
materials of greatest interest. The values given
are for a uniformly distributed heat source giving

 

‘MF. A. Field, Temperature Gradients and Thermai
Stresses in Heat-Generating Bodies, ORNL CF-54-5-196
(May 21, 1954).

55
ORNL~LR--DWG 1147

10

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0020 0050 010 0.20 050 08
PLATE THICKNESS (in.)

Fig. 35, Moximum Temperature Ditfereatial in Flat Plates with Uniformly Distributed Volume Heot
Scurces.

56
STRESS {psi)

Fig. 36.

<

ORNL —LR-DWG H4B

 

 

005 04 0.2 05 08
' PLATE THICKNESS (in)

Thermal Stress in Flat Plotes with Uniformly Distributed Yolume Heat Sources.

57
a power density of 1 watt/em®, Both the tempera-
ture difference and the thermal stress are directly
proportional to the power density. Many other
geometries can eusily be reduced to the flateslab
configuration. A flat plate with heat generated in
a plane at the center will have twice the tempera-
ture differential and one and one-half times the
stress of a plate of equal thickness with uniform
heat generation, A cylindrical rod with uniform
heat generation will have one-half the temperature
differential and three-eighths the stress of a flat
plate with uniform heat generation and a thickness
equal to the diameter of the rod.

It is instructive to apply Figs. 35 and 36 to some
typical structures, for example, the fuel plates for
a sodium-cooled reactor with stainless-steel-clad
UQ, and stainless steel fuel elements. By taking
the solid fuel element design shown in Fig. 31, on
which Table 10 was bosed, and a reactor core
power density of 4,2 kw/cm3, the power density
in the fuel element will be 35 kw/cm® because it
constitutes only 12% of the total core volume., It
can be seen from the curve for Inconel (Fig. 36),
which has properties about the some as those for
stainless steel, that the thermal stress for 0.020-
in.-thick fue! plates would be 0.4 psi for 1
watt/cm?3, or 14,000 psi for 35 kw/cm®, if the fuel
is uniformly distributed throughout the plates, I,
instead, the cladding constitutes one-half the total
thickness, it can be shown that the temperature
differential and the thermal stresses are approxi-
mately half again as great, and thus there would be
a 120°F temperature difference between the center
and the surface and a thermal stress of about
20,000 psi. By referring to Fig. 27, it can be seen
that this therma! stress is mony times the creep
strength of the stainless steel at a temperature of
1700°F ; therefore severe thermal distortion would
be likely to result. Thus the actual power density
might have to be substantially less than the 4.2
kw/cm® permitted by heat transfer considerations,

The more complex geometry of the thick-walled
cylinder requires a more complex representation.4?
For the purposes of this report, the typical set of
curves shown in Fig. 37 will suffice to show the
basic relotionships, These curves apply to an
important particular case; namely, a reactor moder-
ator region cooled by equilaterally spaced circular
passages, It can be shown that the temperature
and the thermal stress distribution in a rigid block
cooled by equilaterally spaced parallel circular

58

5 ORNL-LR-DWG 49
Z ~ T ] T —-

 

  
    
 

VOL. % iN HOLES

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

J e — CCOLING FOLE DIA = 0125in

 

 

 

 

 

 

108 A c _ 3
"’}#‘T:_:_;;;‘: POWER DENSITY = 200 watts/cm
T m.:lll__:' Sl T I R ]
| {»— e
o
T
o — ,J T o T

 

 

. ! : i )
- Be YELD POINT AT 1200°F ' T
L] oTH o

   

    

 

 

 

HOLE SPACING (in}

Fig. 37. Effects of Heole Spacing and Diameter
on Temperature Distribution and Thermal Stress in
a Rigid Block Having a Usiformly Distributed
Volume Heat Source and Equilaterally Spaced
Circular Cecling Possages.,
passages can be closely approximated by con-
sidering the block to be a stack of thick-walled
cylinders having hole diameters the same as those
in the block and an outside diameter equal to 105%
of the hole spacing in the block. For the case
shown, the power density in the reactor core was
taken as approximately 4 kw/cm3, which gives
gamma- and neutron-heating density in the moder-
ator of about 200 watt/em® (that is, 5% of the
power density in the core). The chart is equally
applicable to a reactor core geometry similar to
that of the ARE or to the regions in the reflector
or the island immediately adjocent to the fuel
region of the reflector-moderated reactor. It is
quite evident that BeO, because of its brittleness,
could be considered for use only if pierced with
many closely-spaced cooling passages; however,
such o structure would be flimsy and easily
damaged.

TEMPERATURE DISTRIBUTION IN
CIRCULATING-FUEL REACTORS

The circulating-fuel regctor poses some special
temperature distribution problems that have not
demanded attention in other fields of technology.
These problems arise beccuse the temperature of
any given element of fluid in the reactor core at
any given instant is a complex function of the time
that it has spent in the fissioning region, the power
density, the amount of heat that it has gained from
or lost to the rest of the fluid through conduction
or turbulent mixing, and its own heat capacity.
As a consequence, there are two major sets of
problems that may arise in any circulating-fuel
reactor, The first of these is the formation of
severe local hot spots as a result of flow sepa-
ration. If the hot spots caused local boiling in a
reactor having a high power density, there would
be erratic fluctuations in power and, possibly,
instability of the reactor. Therefore it seems
essential that the fuel flow passages be carefully
proporfioned o ovoid flow separation. The second
problem, boundary-layer heating, arises because

fissioning in the nearly stagnant fluid at the fuel- °

channel surface makes the temperature there tend
to be much higher than that of the free stream. A
rigorous and comprehensive study of the boundary-
layer phenomenon has been 'in process since
1952.4%:46 A few curves based on that study are
presented here to show some of the more important
relationships. '

A good insight into the problem can be gained
from examination of an important typical case ~
that presented by an ARE type of right-circular-
cylinder reactor core containing parallel circular
passages proportioned so that 50 vol % of the core
is filled with fuel while the remainder of the core
is moderator and structural material, If the pas-
sage wall between the moderator and the fuel were
not cooled, the wall temperature would exceed the
local
amount, as can be seen in Fig. 38. At a given fuel
velocity the temperature difference between the
wall and the fuel is directly proportional to power

mean fuel temperature by a substantial

density, but, if the reasoning of the previous
section is followed and the fluid temperature rise
is kept constant at 400°F for a given reactor core,
the fuel velocity becomes directly propottional to
the power density. Surprisingly encugh, the re-
duction in boundary-layer thickness associated

 

45H. F. Poppendiek and L. D. Palmer, Forced Cone
vecfion Heat Transfer Between Porallel Plates and in
Annuli with Volume Heat Sources Within the Fluids,
ORML-T701 {May 11, 1954). '

464, F. Poppendiek and L. D. Palmer, Forced Cons
vection Heat Transfer in Pipes with Volume Heof

Sources Within the Fluids, ORNL-1395 (Dec. 2, 1952).

ORNL-LR-DWG 2014

‘ {
i
‘ i
500 x’ PASSAGE DIAMETER, 8in. — 1
/ ‘ |

 

400 |- T

N

 

 

200

 

100

 

UNCOOLED WwaALL TEMPERATURE LESS
MIXED-MEAN FUEL TEMPERATURE (°F)

 

 

0.5 1.0 1.5 2.0 25 30 35 40
FOWER DENSITY IN FUEL (kw/ecm®)

Fig. 38. Effect of Passage Diameter and Power
Density on the Difference Between the Uncooled
Wall Temperature and the Mixed-Mean Fuel Tem-
perature for an ARE Type of Reactor Core, A 21-
in.~dia right circular-cylinder reactor core con-
taining 50 vol % of NaF-ZrF4—UF4 as the fuel
was dssumed.,

59
with the increase in the Reynolds number causes
the temperature difference between the fuel and
the wall to diop somewhat with an increase in
power density, This effect is shown in Fig. 38,
together with the effects of variations in passage
diameter. The smaller passoges give markedly
reduced wall temperatures. Unfortunately, reducing
the diameter of the passage also increases the
amount of structural material in the reactor and,
hence, increases the critical mass. A similar set
of data is presented in Fig. 39 for a 2l-in.-dia
fuel onnulus that is typical of reflector-moderated
recctors. The effects of variations in fuel physi-
cal properties are indicated by curves for two
different fluoride welts having respectively about
as good and as poor sets of heat transfer proper-
ties as are likely to prove of practical interest,

It is clear from Figs. 38 and 39 that, since the
temperature of the structural metal wall is the
limiting temperature in the system, there is a
sirong incentive to cool the walls, The temperature
distribution through the fuel stream, the wall, and
the wa!l coolant for several conditions is shown
in Fig. 40 for a 4-in.~-dia fuel tube, a ]/é-in.-'rhick
Incone! wall, ond a 1/3-in.«dic wall coolant channel,
Sodivm was assumed as the wall coolant, and the
fuel assumed had physical properties similar to
those of NaK-KF-LiF-UF, (10,9-43,5-44.5-1.1 mole

%). Similar curves are given in Fig. 41 for a fuel

 

1800 —— - ————————

LESS

TEMPERATURE
TEMPERATURE (°F)

UNCOOLED WALL
MIXED-MEAN FUEL

 

 

 

 

 

05 1.0 15 20 25 30 35 4.0
POWER DENSITY IN THE FUEL fkw/cm?)

Fig. 39. Effects of Power Density in Twe Dif-
ferent Fuels on the Difference Between the Un-
cooled Wall Tempercture and the Mixed-Mecan Fuel
21.in,~dia Core

Temperature for o Reflector-

Moderatad Reactor.

60

having the same physical properties except vis-
cosity, which was assumed to be ten times greater
than that of the fuel assumed for Fig. 40. The
difference in temperaiure betwesen the center of
the stream and the peak fuel temperature is nearly
twice as great as that for the lower viscosity fuel,

The sensitivity of the system to velocity distri-
bution in either the fue! or the wall coolant fluid
streams is shown in Figs. 42 and 43. The curves
in Fig. 42 are for sodium-cccled walls, while those
presented in [ig, 43 are for the same system
except that NaOH is used as the wall coolant, An
examination of these two sets of curves shows
that moderate variations in fuel velocity have
relatively little effect on the temperature distri-
bution. Further, from the temperature distribution
standpoint, the NaQH coolant is inferior to the
sodium because it gives fairly wide variations in
wall temperature for variations in the wall-coolant

ORNL-LR-DWG 2046

 

 

  
  
   

 

 

 

 

 

 

 

 

 

 

400 e
o - UNCOOLED WALL
N s
w
<=
=
g —COCLANT TAKING 3.3% &
<200 OF FUEL HEAT GENERATION
o
=
u
L
O
o
s
o
4ol
z -~ COOLANT TAKING 4.9%
W OF FUEL HEAT GENERATION
a
g g
= SODILM VELGCTITY = 10 1ps =
W FUEL—POWER DENSITY = { kw/cm> o
; Re = 100,000 4
Pr =10 z

@ 3
M 5
ul _d
r lad
E o
o [T
i L =S0DIIV
o R, e
= INCONEL WALL \
2 r—FUEL 30UNCARY LAYER }
O /’
S ,

4—in. DlA

 

 

Fig. 40. Effect of Wall Cooling on the Temper-
ature Distribution Through the Fuel Stream, the
Wall, and the Wall Coonlant (Sodium).
CRNL~-LR-DWG 2017

 

 

 

 

   
    
  

80C
-
e,
- .
W 800 - UNCOOLED WALL -
‘q‘:
T
© -COOLANT TAKING 34% OF
o Py ! FUEL HEAT GENERATION T -
7 /
I
)
x 200 - - N PN e L
Ll
L
=
12
2

O

DOLANT TAKING 45%

 

 

 

 

 

 

 

 

 

 

=
L
x
2 b
g OF FUEL HEAT GENERATION w
T -200 s e e 2
3 4
<=
ot SODIUM VELOCITY = 10 fps o
» FUEL-POWER DENSITY = 1 kw/cm? z
o400 e T -
fj Re = 10,000 !
w Pr= 10.0 v
o .
E L
< -500 Ay ULy SRSy NPy Ay - j._.
e L SODIUM -
£, S
=
a1 A e INCONEL wWALL
S TBO0 g e —
o FUEL BOUNDARY LAYER ~——w
- lpinflyin] 4-in. CIA :
~1000 Lo arent
Fig. 41, Ef#fect of Fuel Viscosity on Temper-

ature Distribution Through the Fuel Stream, the
Wall, and the Wall Coolant (Sedium). Fuel vis-
cosity assumed to be ten times greater than that
of fuel assumed for Fig, 40.

velocity. The wall temperature variations would
be likely to lead to thermal distortion and warping
or buckling of the wall.

The use of a hydroxide as both moderator and
wall coolant for an ARE type of core has some
attractive possibilities. As can be deduced from
Figs. 38 and 39, the use of perhaps 50 fuel tubes
about 2,0 in. in diameter instead of g thick annulus
of fuel would give lower uncooled wall tempera-
tures, Figure 43 gives some idea of the possi-
bilities of such a design. Closely fitted baffles
would be required fo direct the hydroxide flow over
the tube walls at a uniformly high velocity, Ir-
regularities in wall temperature would tend to give
progressive thermal distortion and deterioration in
the hydroxide velocity distribution. The cumulative
effects of this process might lead to a hot spot
and severe corrosion of the tube wall. _

While quantitative data are not now available,
some comments on the fuel boundary-layer heating

ORNL-LA-DWS 2018

{eF}

200

 

FUEL CHANNEL

 

 

 

 

 

 

 

 

 

 

O o= e — e I rrn i - - - - - -
Lo
o
o w—— FUEL e = 100,000
2 o0 f oA —== FUELL Re=87000
1
[ Pr=10
= POWER DENSITY =1 kw/em3
w o
& -2R0 T OSOGIUM FLOWING AT 10 fps axnp 15
T REMOVING 4.9% OF FUEL <EAT Ll_lz
[n
5_1 GENERATED., REDUCTIGN OF i
= 300 - SODIUM VELOGITY BY 50% PRODUCED - }35..
= AN EFFECT TOC SMALL TO BE T
o - - - 03
o DISCERNIBLE. .
- LLI
w 400 b3 - e ] E_
£
5 /-" 3ODIUM
woo .S e INCONE( WALL
% '500 77777 e e e R
’_u_J. / / ~FUEL BOUNDARY LAYER
4 T £
< Yand minl 4-in DIA
& 500 Ll
-d
Fig. 42. Effect of Fuel and Coolant {Sodium)

Velocities on Temperature Distribution Through
the Fuel Stream, the Wall, and the Wall Coolant,

ORML - R- DWG 2019

 

 

 

200 bkttt as s s s
-~

s

2 100

=

I

<«

 

 

RATURE (ESS TEMPERATURE IN CENTER OF FULL

FUEL Re = 100,000
NeOH VELGOITY =40 ps
<100 b b i e FUEL Re= 406,000 ...
NaQH VELOGITY = 5 fps
o aon avis e e ELEL R = 67, 000
NaOH VELGEITY =10 fos
FUEL BOWER DENSITY =1 swiomd |
Pr=1.0 €

-200

-300 e

 

=400 = NOOH v T

   

INCONEL 0
800 e e s I -~

}
L

 

 

 

 

 

 

 

 

< ~ FUEL BOUNDARY LAYER
;f_J
. —~50O0 e
T
(&) o—
3 Yiol'ind 4-in. DI
4 g AL in ~in. 4

—rop LELE

Fig. 43. Effect of Fuel and Coolant {(NaOH)

Velocities on Temperature Distribution Through
the Fuel Stream, the Wall, and the Wall Coolont,

61
problems of the reflector-moderated circulating fuel
reactor can be made. The curves given in Figs.
38 to 43 were all obtained from derivations based
on flow in infinitely long passages with fully
developed boundary layers, This approximation is
good for the exit ends of ARE type cores having
fairly large tube length-to-diameter ratios. For
reflector-moderated reactors, however, the entire
passage through the core will be subject to en-
trance effects that will markedly alleviate the
boundary-layer heating problem.  The high-in-
tensity, fine-grain turbulence induced by the pumps
coupled with the mixing effects of the turbulator

62

vanes af the core inlet 47 should serve both to
increase the eddy diffusivity by a large factor and
to inhibit boundary layer thickening in the diffuser
region between the core inlet and the mid-plane.
In the highest temperature region from the mid-
plane to the outlet the fuel passage converges and
should give a marked reduction in boundary layer
thickness and hence in the heating problem,

 

7p. E. Ball, Investigotion of the Fluid Flow Pattern
in .[crghéc;de! of the “Fireball’' Reactor, Y-F15-11 (Sept.
4, 1952).
PART ll. REACTOR STUDIES

COMPARISON OF REACTOR AND
CYCLE TYPES

The report of the TAB provides what is probably
both the most authoritative and the most compre-
hensive comparison of the principal aircraft re-
actors ond propulsion systems that have been
proposed. The key conclusions of the TAB were
set forth in the form of a list of the various cycles
proposed in the order of their promise as aircraft
power plants, The list is almost as pertinent today
as it was four years ago and is as follows:

1. sodium-cooled stationary-fuel-element reactor,

2. reactor (fused-fluoride fuel),

3. homogeneous reactor (fused-hydroxide fuel),

4. circulating-moderator reactor (fused hydroxide
moderator, solid fuel elements),

5. supercritical-water reactor,

6. helium-cooled solid-fuel-element reactor,

circulating-fuel

7. air-cooled solid-fuel-element reactor.

This list was the first one that included the
circulating-fuel reactor as a promising type for
propulsion of aircraft. [t had been feared that the
loss of delayed neutrons through circulation of the
fuel would make such reactors unstable, and it
was not until the TAB deliberations that the
inherent stability of circulating-fuel reactors was
appreciated. This realization constituted one of
the major advances in the program,

The TAB conclusions con' be justified more
effectively now than was possible four years ago
by using the information presented in the first
portion of this report. Key data have been com-
piled in Table 11 to indicate the major charac-
teristics of the most promising of the reactor types
that have been considered.

One of the best measures of the performance of
a system comprising redactor, shield, and propulsion
machinery is the weight of the system in pounds
per pound of effective thrust, that is, the net
thrust minus the drag chargeable to the engine
installation. As was shown in the section on
“Aircraft Requirements,”’ this weight can be
conveniently split info two parts, namely, the
weight of the reactor and shield assembly and the
weight of the propulsion machinery. For any given
set of reactor temperature ond flight conditions,
the propulsion machinery weight per pound of thrust
is essentially independent of power. The reactor

and shield assembly weight per pound of thrust is
a complex function of the reactor power, but under
all circumstances it decreases rapidly as the
power density in the reactor core is increased.
Since the shield weight is probably the most im-
portani single item and is largely determined by
power density, the first line in Table 11 gives the
limiting power density for each type of reactor as
established by heat removal considerations as
summarized in Table 10, The limiting temperatures
both in the reactor and in the jet engine air stream
have an important influence on the weight and drag
of the power plant installation. These data, along
with the weight of the propulsion machinery per
pound of thrust and the specific thrust, are pre-
sented in the next four lines. The temperature
coefficient of reactivity for fast transients, to-
gether with remarks on the controllability, are
presented next, along with estimates of the fuel
investment required per airplane and the cost of
fabricating and reprocessing the fuel. The last
two lines are devoted to remorks on the efficacy
of chemical fuel augmentation and to the hazards
associated with each type of cycle. |

The data for the helium and mercury-vapor cycles
were taken from studies made by North American
Aviation Corporation under an ORNL subcontract,
Even by going to reactor temperatures substantially
higher than those assumed for the other cycles
listed, it was not found possible to reduce the
weight of the propulsion machinery for the helium
and for the mercury-vapor cycles to an acceptable
level; hence, these cycles are clearly not of
further interest, This approach in which weight
considerations are considered paramount can be
justified by considering that for supersonic aircraft
for which the lift-drag ratio will be 5 or 6, roughly
one-third of the aircraft gross weight must go for
structure and equipment, one-third is available for
the reactor and the shield, and one-third can be
used for the propulsion machinery, Thus the weight
of the propulsion machinery should not exceed
2.0 1b/1b of thrust.

The reactor power densities in Table 11 for the
direct air and the supercritical water cycles were
taken from Table 10. They agree with the power
density estimates published by Air Force con-
tractors except for differences in assumptions,

63
Both the General Electric Company and the Pratt &
Whitney Aircraft Division in some of their designs
have assumed more complex and finely divided fuel
elements and hence greater heat transfer areas per
unit of volume. The General Electric Company has
also assumed higher metal temperatures (peak
temperatures within 150°F of the melting point of
the structural material, for example) and thus
further increased design power densities, If the
higher metal temperature should ever prove practi-
cable for the air cycle, it should be equally appli-
cable to the other cycles. The combined effects
of these more optimistic assumptions yield power
densities for the G-E design that are approximately
twice as high as those given in Table 11.48 The
propulsion machinery weight for the air cycle was
estimated from the same data as that used for the
high-temperature-liquid cycles. The weight esti-
mate agrees well with G-E data if allowances are
made for the differences in turbine air inlet temper-
ature. The propulsion machinery weight for the
supercritical-water cycle was taken from Pratt &
Whitney reports.

In comparing the data in Table 11 for the various
cycles it is evident that the performance of the air
cycle is seriously handicapped by a reactor power
density that is inherently only 1 to 10% of that for
the high-temperature liquid-cooled cycles. This,
coupled with the large air ducts required in the
shield, leads to a high shield weight unless the
shield is very heavily divided, The supercritical-
water cycle has the disadvantage of being o low-
temperature and, hence, a low-specific-impulse
system so that it inherently gives a heavy, bulky,
high-frontal-area power plant with virtually no
promise of thrust augmentation through interburning
or afterburning. The hydroxides are afflicted with
such severe corrosion and mass transfer problems
that even after five years of research there is still
no known method of containing them at temperatures
above 1000°F, Thus performonce would be so
limited as to rule out the circulating-moderator and
the homogenesous reactors,

From the above discussion it follows that during
the past few years the only cyeles giving promise
of high performance with the materials available
have been those employing high-temperature liquid-
cooled reactors coupled to turbojet engines. On

 

48 Gencral Electric Co., Aircraft Nucleor Propulsion,
Department of FEngineering, Progress Resport No. 9,
p. 29, APEX-9 (Sept. 1953),

64

the bosis of materials considerations, the field
was narrowed to the sodium-cooled solid-fuel-
element reactor and the circulating-fue! reactor,
In both these reactors an intermediate heat transfer
fluid is required because the fluid that passes
through the reactor is rendered far too radioactive
to be circulated outside the shield.

REACTOR, HEAT EXCHANGER, AND
SHIELD ARRANGEMENTS

A historical survey of ORNL aircraft reactor
design work provides a further approach to a criti-
cal comporison of high-temperature-liquid reactor-
types. Some work on the ANP Project was started
at ORNL in 1948 to provide experimental data for
NEPA. As the effort directed toward fundamental
problems, such as shielding and materials, was
expanded, the need for supporting work on power
plant design became evident, The General Design
Group was set up in March 1950, and an intensive
study of reactor types and cycles was initiated.
By June 1950, it had been concluded that o high-
temperature liquid-cooled reactor coupled to turbo-
jet engines evinced markedly greater promise than
any other arrangement, The program formulated on
this basis was given great impetus by the TAB
recommendations in August 1950,

In examining the problems associated with the
sodium-cooled solid-fuel-element and circulating-
fuel reactors it was felt that the latter should have
a substantially higher performance potential so far
as upper temperature limit and power density are
concerned. Further, the use of a circulating fuel
would greatly simplify problems of preparing and
repraocessing the fuel, and would give an almost
assuredly simple reactor control system, because
the negative temperature coefficient ossociated
with expansion of the fuel would be entirely ade-
quate to take care of any fast transient pertur-
bations. Problems associated with a solid-fuel-
element reactor include xenon override and limi-
on the operating life imposed by the
tolerable burnup in the fuel elements and by the

tations

degree to which provision can be made to compen-
sate for the reactivity losses associated with
burnup through the use of such devices as the
addition of poisons that would burn out duting the
course of operotion., Insertion of a large number
of control rods would seriously impair the heat
transfer characteristics of the core and require

The high

much complex actuating equipment.
TABLE 11. COMPARISON OF KEY DATA FOR THE MORE PROMISING TYPES OF AIRCRAFT REACTOR SYSTEMS

Reactor Power Density Conditions given in Table 10, 1700°F Peak Metal Temperature (Except for Helium and Mercury Vapor Cycles), Flight at Mach 1.5 and 45,000 #

 

 

 

SODIUM-COOLED

A{R-COOLED

HELIuM-coor epld

 

CIRCULATING-FUEL | HOMOGENEOUS IRCULATING-MODERATO! SUPERCRITICAL-WATE MERC - SOLID-FUEL.ELEME
SOLID-FUEL-ELEMENT REATc:TOR REACTORU S ;Er:c*roa TR RREACTOR(‘” TER | souin-FuEL-ELEMENT CYCLEURR!EYAZ$;2&) ;:A!ETOLR =
{2000 psi heljum)
Hect-removal-limited power 4,2 >10 >10 1.9 0.70 0.027 4,2 0.66
density, kw/cmB
Limiting reactor fluid outlet 1550 1700 1200 1200 1300 1280 1730 3100
temperature, °F
Limiting air temperature, °F 1240 1350 990 990 455 1280 1160 1230
Specific thrust {less nacelie drag), 22.4 29.0 14 14 19.4 26.0 28.8 38.5
1b/1b of air/sec
Weight of propulsion machinery at 2.5 2.0 4.0 4.0 3.3 1.9 4.3 2.7
Mach 1.5 and 45,000 §1, 1b/1b of
thrust
Probable temperature coefficient +1073 -5 % 1073 ~5% 10"% ~10~3 +10™2 +10~7 +10-% +10~5
for fast transients, Ak/&-°C
Remarks on controllability Difficult and Simple and Simple and Startup Difficuit and Difficult and Difficult and Diffieult and
complex inherently inherently procedure complex complex complex complex
reliable reliable difficult
Total fuel investment (U2%) in 60 120 60 25 30 100 60 100
reactor, Ib
Cost of fabrication and 16.00 1.50 1.50 16.00 16.00 16.00 16,00 16.00
reprocessing per gram of U235
dollars'®
Efficacy of chemical fuel Very good Very good Very good Very good Poor Good Poor Poor
dugmentation
Major hozards Reactor runaway and Fuel spill; Fuel spill; NaOH spill; Burst of some Reactor runaway and Reactor runaway Reactor runaway and
melt-down; sodium NaK fire NaK fire NaK fire part of 5000-psi melt-down; burst of melt-down; burst of

 

fire

 

 

 

 

system

 

300-psi system

 

 

3000-psi system

 

(O)qu calculated from Pratt & Whitney Aircraft Div., Nuclear Propuision Program Engineering Progress Reporf, No, 9, PWAC-75, p. 28.

(b)General Electric Co., Aircraft Nuclear Propulsion, Department of Engineering, Progress Report No. 9, APEX-9 {Sept. 195‘3).
(C)Dm‘q calculated from work of A. Dean and . Nakazato, Investigation of Mercury Vapor Power Plant for Nuclear Propulsion of Aircraft, NAA-5R-110 (Mar. 21, 1951).

(d)

Data caleulated from work of H. Schwartz, An Analysis of Inert Gas Cooled Reactors for Application to Supersonic Nuclear Aircraft, NAA-SR-111 (Sept. 8, 1952},

(e)Dcm caleulated from memorandum from C, E., Larson to G. Beardsley, Preliminary Comporison for Reprocessing Fuels from an SCWR and a CFR, ORNL CF-53-12-11 (Dee, 1, 1953).

65
temperatures required of an aircraft reactor coupled
with leaktightness requirements and shield weight
and residual radiation considerations make it seem
unlikely that o sodiumecooled solid-fuel-element
reactor could be reloaded readily. On the other
hand, littie information was available in 1950 on
fluids that might serve as vehicles for fuel. Hy-
droxides had been considered by NEPA,4? but it
was felt that serious corrosion problems would be
inherent in their use because the oxygen in them
is not bound tightly enough to give assurance of
their remaining inactive at high temperatures. R,
C. Briant, in April 1950, pointed out that, on the
basis of chemical thermodynamics, the alkali fluo-
rides should be inherently stable relative to the
iron~-chrome-nickel alloys even at the high temper-
atures required and advocated their use as circu-
lating fuels. It was recognized, however, that the
use of the fused fluorides as o circulating fuel
would mean the opening of o whole new field of
reactor technology that would be filled with un-
knowns and that there would be no guarantee of
success, |herefore it was decided in September
1950 that the major emphasis should be placed on
the sodium-cosled solid-fuel-element reactor, while
a substantial research program should be directed
toward the solution of the corrosion problems
associated with the hydroxides and the fluorides.

Shield and Heat Exchonger Designs

To implement the design effort on the sodium-
cooled solid-fuel-element reactor, an intensive
study was made of the shieilding problem by o joint
ORNL-NEPA committee in the fall of 1950. A
number of the reactor and shield designs included
in the committee report®® are of interest. Figure
44 shows the first design prepared, which followed
the NEPA practice of using conventional tube-
and-shell heat exchangers disposed relotive to the
reactor and to the pumps in a quite conventional
fashion with the shield simply wrapped around the
resulting assembly. The estimated shield weight

for this assembly was over 230,000 1b, The lay-

outs shown in Figs. 45 and 46 make use of an

 

ONEPA Project Quarterly Progress Report for Pericd
April T—June 30, 1950, NEPA-1484.

SOReport of the Shielding Board for the Aircraft
Nuclear Propulsion Program, ANP-33 {Oct. 16, 1950).

unconventional heat exchanger in which the reactor
coolant flows axially through the interstices be-
tween small-diameter, closely spaced tubes, while
the secondory circuit fluid passes through the
tubes to give a virtually pure counterflow system,
The shield weight for the tandem heat exchanger
arrangement of Fig., 45 was estimated to be about
160,000 1b, while that for the annular heat ex-
changer arrangement of Fig. 46 was 122,000 ib.
This was close to the weight of the ideal maiched
lead-water shield, the weight of which was esti-
mated to be 116,000 Ib. A fourth configuration,
which made use of lead as the reactor coelant, is
shown in Fig. 47, This arrangement, the weight
of which was estimated to be about 120,00C Ib,
was designed to employ the lead reactor coolant
as shielding material by placing the heat exchanger
at the same region in the shield as would normally
be occupied by the gamma-ray shielding material.
Differential thermal exponsion appeared to pose
some rather difficult structural problems in con-
nection with the fairly large volume, low-temper-
ature shield region inside the high-temperature
heat exchanger shell, The design also had the
disadvantage that the lead-corrosion
showed no promise of solution.

Concurrently with the 1950 Shielding Beard
investigation, a second joint ORNL-NEPA group
carried out an intensive study of the reactor and
engine control problem.®® This group reluctantly
reached the conclusion that a solid-fuel-element,
high~temperature, high-power-density reactor might
be unstable ond that if at all possible an effort
should be made to obtain a reactor with a negative
temperature coefficient, even if it meant compro-
mising the reactor design, While there was liftle
doubt that the sodium-cooled solid-fuei-element
reactor cowld be controlled, it appeared that on

problem

unusually complex control system would be re-
quired which, when coupled to the very complex
contrel system required for the turbojet engines,
would probably seriously impair the reliability of
the power plant, In view of this serious develop-
ment, the situation wos reappraised. Mt was de-
cided that the matertials research work was still not
sufficiently far along to permit shifting the major
emphasis to a circulating-fuel type of reactor and
therefore development would have to continue on
a stationary-fuel-element reactor, It was felf,
however, that it would be possible to use a design

67
89

ORNL—-LR - DW5. 1150 ORNL—L.R~DWG 1151

 
  

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BOTTOM

 

 

 

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Fig. 44. Shield Design for o Sodium-Cooled Solid-Fuel-Element Reactor with Coaventional Tube-and-Shell Heat Exchengers,
ORNL ~LR-DWG 4152

         

-—-—- PRESSURE SHELL

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

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SECONDARY SODIUM OUTLET _\fl\\‘ —- fy N . JlENy 7 e PRIMARY COOLANT PUMPS
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Fig. 45. Design of a Sodium-Cooled Solid-Fuel-Element Reactor with o Tandem Heat Exchanger
Arrangement,

69
PRIMARY COOLANT PUMP -,
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SPACE FOR THERMAL

 

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DIAPHRAGM TO PERM!IT
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=

 

Fig. 46. Design of a Solid-Fuel-Element Reacter with an Annular Heot Exchanger Arrangement.

70
ORNL-LR~- DWG {154

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INTERMEDIATE HEAT EX-
CHANGER WALLS

  

- BOTH SIDES OF THI5 PLATE TO BE
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Fig. 47. Design of a Lead-Cooled Solid-Fuel-Eiement Reactor with Heat Exchangers Aranged to Act

as Gamma-Ray Shielding.

similar to that of the KAPL Submarine Intermediate
Reactor but with a molten fluoride salt containing
uranium in solution instead of solid UO, in the
fuel pins. Thermal expansion would push some of
this fluid fuel out of the reactor core into an
expansion tank to give the desired negative tem-
perature coefficient of reactivity., Work on this
design proceeded for nearly o year until, in the
fall of 1951, it developed that the GE-ANP project
had dropped its plan to base development on high-
temperature liquid-cooled reactors and had instead
returned to the air cycle. Since this change eased
the pressure for ORNL to get an experimental re-
actor into operation at the earliest possible date,
the entire nuclear-powered aircraft situation was
reappraised.

No structurally satisfactory design had been
evolved for a high-power-density reactor core em-
ploying the molten-saltfilled fuel pins, and two
problems associated with the fuel pins
seemed well-nigh insuperable. First, there was,
inherently, a temperature drop of over 1000°F
between the center and the outside of the fuel pin.
While this would not have been serious in a low
power redactor of the type to be used for the Air-
craft Reactor Experiment, it probably would have
been quite serious in a full-scale aircraft reactor,

major

Second, as in any solid-fuei-eiement reactor, ade-
quate
spaucing of the fuel elements would have been
exceedingly difficult tc arrange. Thus, on the
basis of structural and heat transfer considerations,

support and satisfactory maintenance of

71
attention was turned to the circuloting-fluoride-
fuel reactor with its then more difficult materials
problems. Considerable progress had been made
in the investigations of the chemistry of fluoride
fuels, and it was believed that in the long run the
problems associated with the use of circulating
fuels would prove easier to solve than the less
obvious, but nonetheless vital, problems inherent
in fixed-fuel-element reactors,

initiated in
October 1951 for examining the problems associ-
ated with a full-scale aircraft nuclear power plant
employing a circulating fluoride fuel. In one of
the first studies, an examination was made of the
possibility of piping the fluoride fuel directly to

An intensive design effort was

heat exchangers in the turbojet engines and thus
eliminating the complications associated with an
intermediate heat transfer circuit., The design
study of this proposal is covered in ORNL-1287.31
Even by going to the exceptionally large reactor-
crew separation distance of 120 ft and a crew-
engine separation distance of 135 ft, which badly
compromised the airplone design, the arrangement
led to a shield weight actually greater than that
obtainable intermediate heat transfer
circuit, Also, the radiation dose level of about
6 x 10% r/hr at 50 ft from the reactor that would
result from this arrangement would be completely
intolerable.

with an

Shielding of the engine radiators ap-
peared to be out of the question because of their
large size. Ground-handling and maintenance prob-
lems seemed to many people to be insuperable,
Thus this arrangement was dismissed and attention
was directed to reactor systems employing an
intermediate heat transfer fluid.

A careful study of arrangements of the reactor,
intermediate heat exchanger, and shield was made
in an effort to determine the effect of configuration
on shield weight,>? Activation of the secondary
coolant threatened to be a much more severe prob-
lem in these arrangements than in the arrangements
for use with solid-fuel-element reactors (Figs. 44
through 47) because delayed neutrons from the
circulating fuel would be released in the heat

exchanger. It was found that this key problem

 

3TR. W. Schroeder and B. Lubarsky, A Design Study
of a Nuclear-Powered Airplane in Which Circulating Fuel
is Piped Directly to the Engine Air Radiators, ORNL-
1287 (Apr. 16, 1953).

52A. P. Fraas, Three Recctor-Heat Exchanger-Shield

Arrangements for Use with Fused Fluoride Circulating
Fuel, Y-F15-10 (June 30, 1952).

72

could be handled by keeping moderating material
out of the heat exchanger so that most of the
neutrons would escape before they would slow
down. By filling 5 to 10% of the heat exchanger
volume with boron carbide, most of the neutrons
that did not escape would be captured in boron
rather than in the secondary coolant. This solution
to the problem of neutron activation of the second-
ary coolant makes the circulating-fuel reactor
superior to o homogeneous reactor. In homogeneous
reactors, moderation of the deloyed neutrons by
moderator-fue! would greatly aggravate the prob-
lem.

The first circulating-fluoride-fuel reactor, inter-
mediate heat exchanger, and shield arrangement
studied — an arrangement in which the reactor and
heat exchanger were placed in tandem — is shown
in Fig., 48. To keep the octivation of the sodium in
the seconduary circuit to a tolerable level, it was
found that it would be necessary to separate the
heat exchanger from the reactor core by at least
12 in. of good moderating material followed by a
1 in. thick layer of boron carbide (or, if B'® were
used instead of natural boron, a thickness of 0.2
in.). A carefu!l analysis of this arrangement dis-
closed also that the pressure shell should be
separated from the reactor core by a layer of boron
carbide of similar thickness to keep the pressure
shell from becoming ¢ more important source of
gammas than the core. |t also became clear that
activation of the secondary coolant by delayed
neutrons emitted from the fue! in the heat ex-
changer could be markedly reduced by spreading
the heat exchanger out in a thin layer and thereby
increasing the neutron escape probability, Further,
it was observed that with the tandem arrangement
(Fig. 48), the lead shielding required just for the
heat exchanger constituted ¢ major portion of the
total shield weight,

The annular heat exchonger arrangement shown
in Fig. 49 was evolved to place the heat ex-
changers around the reactor within the primary
reactor shield and thus eliminate the extra lead
shizlding of the heat exchangers.
ment gave an estimuted shield weight of 128,000
Ib, as compared with 156,000 Ib for the tandem
arrarigement,
led to the conclusion that an additional weight

This arrange-

Careful examination of this design

saving could be realized by changing the gecmetry
of the design to make it more nearly spherical,
The spherical arrangement shown in Fig. 50 was
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Fig. 49. Schematic Longitudinal Cross Section Through Hydroxide-Meoderated Circulating-Fuel Reactor with Annular Heat Exchengers,
\ e

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TUBE BUNDLE

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Fig. 50. Arrangement of Circulating-Fuel Reflector-Moderated Reactor with Spherical Heat Exchanger.

then designed, and it was found to give an esti-
mated shield weight of 120,000 b, The shield
weights given here were estimated for o quasi unit-
shield design condition, namely, 7 r/hr at 50 fi
from the reactor, and 1 r/hr inside the crew com-
partment. Various degrees of shield division were

also considered!” in an effort to reduce shield

weights for the various designs, and in each
instance the spherical-shell heat exchanger ar-
rangement shown in Fig. 50 was found to be
superior. In all cases, the reactor output design
condition was 400 Mw, and the reactor core di-
ameter was about 30 in,

75
Reactor Core Configurations

The reactor cores used in the shield design
studies of Figs. 48 and 49 were conventional in
that the moderator was distributed throughout the
active fuel region with relatively little lumping.
However, the reoctor core design shown in Fig. 50
was evolved on a quite different basis. A brief
account of the reasoning that led to this reflector-
moderated reactor design may be of interest. A
number of people had felt that a small (perhaps
18-in.-dia) fast reactor might be built to utilize
one of the uranium-bearing fluoride-salt fuels.
Rough calculations made by T, A, Welton indicated
that the high concentration of uranium atoms re-
quired to oachieve criticality with this type of
reactor would be difficult to obtain in any fluoride
salt ‘melt likely to have desirable physical proper-
ties. Others felt that the concentrations of Li”
and Be in the fluoride melt might be increased to
the point where their moderating effect would be
sufficient to make possible a homogeneous fused-
fluocride reactor., A minimum critical mass of the
order of 150 kg was indicated by one- and two-
group calculations for such a reactor. Since the
shield design studies had clearly shown the
desirability of a thick reflector, it was felt that
it might be possible to capitalize on this thick
reflector and effect a major reduction in critical
mass with a quasi-homogeneous flucride fuel. (At
the time, a fairly high uvranivm concentration in the
fluoride mixture was not considered to be too
serious,)

Multigroup caleulations indicated that a beryl-
lium reflector could be made so effective that the
critical mass could be cut to something of the
order of 15 kg.>® This prediction was later con-

18 |t has also been

firmed by critical experiments,
found that the good high-energy neutron-scattering
cross section of the fluorine in the fuel is more
important for o reactor of this type than the moder-
ating effects of Be or Li’. In fact, the neutron-
scattering cross section is so much more important
that heavier elements such as No and Rb may be
used in the fluorides instead of Be or Li’7 with
fitile effect on critical mass.

The heavily lumped fuel region of the reflector-
moderated reactor has a number of major advan-
tages. The removal of all structural material from

 

33C. B. Mills, The Fireball, A Reflector-Moderated
Circulating=Fuel Reactor, Y-F10-104 {June 20, 1952),

76

the core except the core shells reduces parasitic
neutron capture in structural material to a minimum
and hence reduces the critical mass. The place-
ment of most of the moderating moterial in the
reflector gives a smaller diameter core for a given
power density in the fuel and hence a lighter
shield.

Many circulating-fuel reflector-moderated reactor
arrangements have been proposed to take advantage
of the spherical-shell heat exchanger and shield
arrangement shown in Fig. 50, In general, it
appears that there are eight basic types of con-
struction that might be employed. The simplest
type, shown in Fig. 51, comprises a thick, spheri-
cal shell of moderator surrounding o spherical
chamber containing liquid fuel. Ducts at the top
and bottom of the shell direct cold fuel into the
reactor core and carry off high-temperature fuel.
Such an arrangement has two major disadvantages.
First, the well-moderated neutrons reflected to the
fuel region from the reflector tend to be absorbed
near the fuel-reflector interface so that the power
density falls off rapidly from that interface to a
relatively low value at the center. Second, the
flow pattern through such a core is indeterminate,
and large regions of flow separation and probable
stagnation would be likely to occur in a highly
irregular, unpredictable fashion, although vanes
or screens at the inlet might be effective in slowing
down and distributing the flow, The arrangement
shown in Fig. 52, which makes use of a central
*island,”’ appeared to be more promising. The
central] islond has the advantage that it reduces
the critical mass and yields a more uniform power
distribution.  Thus the extra complexity of a
cooling system for the island appears to be more
than offset by the reduced critical mass, improved
power distribution, and much superior hydrodynamic
characteristics.

The most serious problem associated with the
arrangement of Fig. 52 appeors to be that of
cooling the moderator,®4 a problem common to all
high-power density reactors. If beryllium is used
as the reflector-moderator material, closely spaced
cooling passages must be employed in those
portions close to the fuel region to remove the
heat generated by gamma-absorption and the

 

S4R. W. Bussard ot al., The Moderator Cooling System
for the Reflector-Moderated Reactor, ORNL-1517 {Sept.
1953).
ORNL-LR-DWG 283%

 

e
REFLECTOR- MODERATOR

Fig. 51. Simple Two-Region Reactor Core with
Thick, Spherical Shell of Moderator Surrounding a
Spherical Chamber Containing Liquid Fuel. '

neutron-slowing-down processes. Other arrange-
ments have been considered; for example, the high
femperature gradients and thermal stresses induced
in the beryllium in this fashion might be avoided
if a layer of a liquid such as lead or bismuth
could be interposed between the fuel and the
refiector-moderator regions, as indicated in Fig.
53. This liquid could be circulated to carry off
the heat and the beryllium coeling problem would
be markedly relieved. |

It appears feasible to use graphite in direct
contact with fluoride fuels without damage to the
graphite or contamination of the fuel. Therefore
a possible design (Fig. 54) comprises a block of
graphite drilled to give a large number of parallel
passages through which the fuel might flow, This
design, in effect, gives a very nearly homogeneous
mixture of fuel and graphite in the reactor core.
A variation of this design is shown in Fig. 55.
Several concentric shells of grophite might be
placed in such a way that they would serve to
guide the fuel flow and at the same time oct as
moderating material, From the hydrodynamic stand-
point, either of these arrangements oppears to be
preferable to the screens or vanes placed in the

"

ORMNL-[.R-DWG 2836

 

 

  

L SO
2 OO
5. A

-MODERATOR
REFLECTOR -MODERATOR

Fig, 52. Three-Region Reactor Core with Central
Island of Moderating Materials.

fuel inlet mentioned in connection with Fig, 51.
While these arrangements appear to have the
advantage of simplifying the core design and
dispersing moderator through the fuel region,
calculations indicate that the fluoride fuel com-
pares favorably with graphite as o moderating
material, and therefore the arrangements of Figs.
54 and 55 are little better from the nuclear stand-
point than that shown in Fig. 51,

A number of different types of fluid-moderated
reactors has been considered. One variant is
shown in Fig. 56. A set of coiled tubes through
which sodium hydroxide could be pumped might
be placed in the reactor core, These could be
made to serve both to improve the fuel velocity
distribution and to moderate fast neutrons in the
reactor core, The principal disadvantage associ-
ated with such an arrangement is that it would be
difficult to avoid local hot spots in the liquid fuel
in zones where flow separation and stagnation
might occur. Alse, the relatively large amount of
structural material in the tube walls would capture

77
ORNL-LR-DWG 2837 X
ORNL -LR-DWS 2528

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MODERATOR
REFLECTOR-MODERATOR

Fig. 54. Reactor Core with Fuel Channels in

Fig. 53. Five-Region Reactor Core with Pro- Graphite Block.

vision for Ceoling Reflector-Moderator Regions.

ORNL-LR-DWG 2334

ORNL-LR-DW6 2833 .
- HYOROXIDE

GRAPHITE SHELLS -

Fig. 55. Reactor Core with Grophite Shells in Fig. 56. Fluid-Moderated Reactor Cere with
Fuel Channel. Coiled Tubes for Circulating the Moderator.

78
CHNL-LR-BWE 284G

 

/ .
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Fig. 57. Fluid-Moderated Reactor Core with
Straight-Tube Fuel Possages and Provision for

Circulating Moderator Around Fuel Tubes,

ORNL--L'!'DWE} 2841

 

 

 

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- HYDROXIDE

Fig. 58, Fluid-Moderated Reoactor Core with
Spheriodized Fuel Passages and Provision for
Circuloting Moderator Around Fuel Passages. '

a substantial percentage of the neutrons, and
therefore the critical mass would be increased.
The arrangement of Fig. 57 also presumes the
use of a fused hydroxide as a fluid moderater, The
fluoride fuel would circulate through the circular
passages and pass down through the reactor core,
while the hydroxide moderator would circulate
through the spaces between the fuel passoges.
Because of the fuel boundary-layer heating problem
(cf., section on “‘Temperature Distribution in
Circulating-Fuel Reactors’’), boffles would have
to be provided arocund the fuel tubes so that the
hydroxide could be circulated at a high velocity
over the tube wail with good velocity distribution
to prevent hot spots. The arrangement of Fig, 58
is similar to that of Fig. 57, except that the tubes
are specially shaped to reduce the volume of the
header regions and fo give a more nearly spherical
core and hence a lower shield weight. '

DETAILED DESIGHNS OF REACTORS
Sodium=-Cooled Solid-Fuel-Element Reactor

The first detailed design studies of reactors
were based on sodium-cooled solid-fuel-zlement
reactor cores, ond several types of fuel element
were examined. The pin type used in the SIR
core?? appeared to be attractive, but the problems
of supporting the pins and maintaining wniform
spacing between them were exasperatingly difficult,
particularly . for high-power-density cores.  An
arrangement that promised to give a much higher
power density potential incorporated stainless-
steel-clad sandwich fuel plates having a sintered
UQ, and stainless steel core, os described in the
previcus section on "*Materials.”” The most highly
developed design of this character is that shown
in Fig. 59, which was prepared in the summer of
1952 by A. S. Thompson, This core was designed
to employ a fue!l element of the type shown in
Fig. 60, but it is equally well adopted to the use
of sandwich tube fuel elements of the type shown
in Fig. 61. The core design was based upon the
flow of sodium downward through the annular fuel
element matrix and the reflector and then radially
outward and upward through the heat exchanger in
the annulus between the reflector and the pressure
shell. Pumps at the top of the pressure sheil
were designed to toke the sodium os it left the
heat exchanger and deliver it back to the core
inlet passage.

79
CRNL-LR-DWG 528

  

 

 

 

 

 

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ium-Cooled Solid-Fuel-Element 100-Mw Reacter.

59. Sod

Fig.

80
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ORNL-LR-DWG 527

 

 

 

 

   
    

 

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LI
Il
NOTE : DIMENSIONS IN INCHES

 

 

 

Fig. 61. Reactor Core Employing Stairless-Steel-Clad Sandwich Fuel Tubes.

82
A careful examination of the over-all charac-
teristics of an aircraft power plant using al-
ternatively sodium-cooled solid-fuel-element and
circulating-fuel reflector-moderated reactors led
to the conclusion that the circulating-fuel type
gave the greater promise. The presence of fuel
in the heat exchanger seemed to increase the NaK
activation by about a factor of 10 and gave a
shield weight increase of 3000 to 5000 |b for the
circulating-fuel These disadvantages
appeared to be more than offset by the higher NaK
temperature obtainable. That is, since the limiting

reactor.

reactor temperature appears to be the peak allow-

able metal temperature and since the temperature
drop through the solid-fuel-element core and
cladding and from the solid-fuel-element surface
to the sodium must be 100°F or more (with allow-
ances for hot spots), the NaK temperature leaving
the heat exchanger could be at least 100°F higher
for the circulating-fuel reactor than for the sodium-
cooled solid-fuel-element type. Further, for the
fluid-fuel type, the reactor and power plant should
be muchk easier to control, the fuel reprocessing
costs should be very much lower, and the fuel
loading and unloading operations much simpler.
Thus ORNL effort was concentrated on the circu-
lating-fuel reactor with the thought that most of
the work would be almost equally applicable to
either type and that in the event something quite
unforseen arose to handicap the fluoride fuel it
should prove fairly easy to shift to the sodium-
cooled solid fuel element.

Circulating-Fuel Aircraft Reactor Experiment

Many different designs have been prepared for
circulating-fluoride-fuel ARE is
representative of an important class of these. As
shown in Figs. 62 and 63 the design of the ARE
was based on the use of passages approximately
]]’3 in, in diameter spaced on approximately 3]/2-in.
centers in a BeQO matrix. The BeO matrix was
prepared in the form of hexagonal blocks approxi-
mately 6 in. long and 3]/2 in. across the flats, While
the ARE design was prepared for a reactor power of
only 3000 kw, it was intended to simulate a reactor
capable of developing @ much higher power out-
put.’!  For the high-power case, it was intended

reactors; the

that simple tube-to-header sheets be used at the
top and bottom in place of the complicated return
bend arrangement shown in Fig. 62. The tube-to-
header sheets could not be employed for the ARE
simply because, at the low powers for which the
ARE was designed, the fuel flow through the
reactor would have been so low that laminar flow
would have prevailed and the tube wall tempera-
ture would have been about 500°F higher than the
mean fuel temperature in the core., By using return
bends and connecting eleven of the passages in
series to give five parallel groups it was possible
to increase the flow velocity sufficiently to ensure
turbulent flow and thus reduce the temperature
difference between the tube wall and the mean
fluid temperature to about 50°F,

It was mentioned that the ARE was intended to
simulate in a rough fashion a reactor core poten-
tially capable of power outputs of at least 200
megawatfs. It was recognized, however, at the
time the ARE was designed that there were many
features that would have to be chonged to permit
the higher power output. Of most importance would
be a change to some better moderator arrangement
than the hexagonal BeO blocks, which as shown
in the. earlier section on “Temperature Gradients
and Thermal Stresses,”’ would break up into rather
small pieces under the action of the thermal
stresses thot would be induced by gamma heating
in a high-power reactor. The ARE design also has
the disadvantage that the average power density
in the reactor core is very much lower than the
power density in the fuel. It alse requires a
relatively high concentration of uranium fluoride
in the fuel melt, which is considered undesirable
because the relatively high uranium concentration
yields a fluoride that has inferior physical proper-
ties and less than optimum corrosion character-
istics, '

Fluid-Moderated Circulating-Fuel Reactor

An arrangement somewhat similar to that for the
ARE was worked out for the design shown in Fig.
64. In this reactor, water or hydroxide was to
serve as a liquid moderator that would fill the
between the tubes that carried the
To avoid

interstices
fluoride fuel through the reactor core.

83
REGULATING RCD
ASSEMBLY -~

 
 
  
  
  
 
  
 
  
  
 
  
  
    
  
  

—TUBE EXTENSION
SAFETY ROD -

GUIDE SLEEVE. -
. _— THERMAL SHIEL

P
SAFETY ROD ASSEMBLY / ——FUEL INLET
;/ .
s

  
 
 

THERMCCOUPLE
LAYOUT

CORE ASSEMBLY ——{:=

REFLLECTOR COOLANT
TUSES —— -

FUEL TUBES

PRESSURE SHELL -

 

 

 

 

e

 

>

 

 

BOTTON HEADER -

THZRMAL SHIELD CAP

   

N

  
   

o

  

DweG. 18336

e THERMAL SHIZILD CAP

D TOP

MANIFOLD
1

____TOP HEADER

-—-BeC MODERATOR
AND RZFLECTOR

 

i

‘I/’THERMAL SHIELD

ASSEMBLY

TICM TUBE
SHEET

—— STUD

-———SUPPORT ASSEMBLY

~~| - FUEL OUTLET

MAN:FOLD

"~ THERMAL. SHIZLD
BOTTOM

- 80% in. REF. -

2395 in REF —— —»—

 

Y \ "
“SUPPORT ASSEMBLY HELIUM MANIFOLD

Fig. 62. Vertical Section of the ARE.

84

02 4686 8 12 18
SCAILLE IN INCHES
i ) OWG. 15847

 
 
  
 
 
 
 
 
 
 

—INSTRUMENT TUBE
- (INGCMEL)
TUBE GOl
(INCONEL )

- REFLECTOR DGE BLOCK
- (BERYLLIUM DXICE)
(NGONEL )~/

—-— REFLECTOR BLOCK
(BERYL_IUM OXICE)

 

e REFLECTOR COOLANT
TUBES {INCOMNEL)

/. e / T ; ) ,
o A 0 . el o3 jpem=e e AN )
, , A A ONGONEL)

  
 
 

         
  
  
 

 

N
i
N
(XD
W \‘
e GORE SLEEVE
INCONEL

TTTSGARETY ROD GUIDE
SLEEVE (INCONEL)

SCALE N INCHES

 
 

 

Fig. 63. Cross Section of the ARE.

85
DWG Y-F7-17'4

  
  
  
 
 
 
 
  
 
  
   
  

- FUEL TUBE SPACER AND CONTRCL
ROD TRACK SUPPORT BAFFLE

STEEL ROLLER — o
CADMIUM CYLINDER . | |7 f i)
STEEL SHAFT -

ENDLESS TRACK FOR SPROCKET-
DRIVEN CONTRCL RODS .

Tlem— 10 N
(RCTATED 20°)

 

 

 

 

 

 

 

 

 

- TCP TUBE SHEET

L
FUEL EXPANSION fb Py )
TANK ASSEMSLY - — -~ //\ T

 

COOLANT “

 

  
  
  

COOLANT QUTLET ,,_J He= - COOLANT INLET

 

   

e
=
; 4
< oy
‘ l 2
e : |
i . (] . [
: i T gt C
i |_ il | [ e i
r fii T- i
< [ |7~ HEAT EXCHANGER CENTRAL
= “; ISLAND SPACERS
O "
; o Il

Jlj " HEAT EXCHANGER RING SPACER

 

 

 

 

 

 

  

“~FUEL PUMP VOLUTE

FUEL PUMP.. |/
i CASING

Fig. 64. Reactor Design for Use with Liquid Moderator (Water or Hydroxide) ond Circulating Flvoride
Fuel,

86
freezing the fuel at zero or low power if water
were used, it was planned that a thin layer of
insulation would be placed between the walls of
the double-walled fuel tubes. It would thus be
possible to operate the reactor with the water at
a substantially lower temperature than that of the
fluoride fuel. Unfortunately, the thermal insulation
would preclude cooling of the fuel-tube walls
during high-power operation, and hence the allow-
able flucride fuel temperature would be perhaps
300°F lower thaon might otherwise be possible.
The temperature differential between the water
and the fuel could be reduced, of course, by pro-
viding o heavy pressure shell and operating the
reactor with the water at high temperature and
pressure. However, a major disadvantage of this
arrangement would be that to keep the stresses in
the fuel tube walls to within reasonable values it
would be necessary for the fuel system to operate
at high pressures. In tum, there would be dif-
ficulty with the pump-shaft seals, and the pressure
shell would be excessively heavy,

Two variants of this design were prepared. In
the first, the reactor was designed fo generate
steam for a supercritical-water cycle in which the
moderator region of the reactor would serve as the
feedwater heater. In the alternate arrangement,
the heat added to a hydroxide moderator could be
dumped at a high temperature, while the bulk of
the heat would be transmitted from the fuel to NaK
in the heat exchanger and the NaK would, in turn,
be directed to turbojet engines. A major innovation
in heat exchanger design was introduced which
involved the use of a fairly large number of smali
tube bundles with the tubes terminating in small,
circular-disk headers. This arrangement had the
advantage that the heat exchariger could be fabri-
cated in elements, and each element could be
carefully inspected and pressure tested. The
efements or tube bundles could then be welded
into the pressure shell with a relatively simple,
rugged joint. By breaking the heat exchanger up
in this fashion it was believed that the ultimate
cost could be markediy reduced and the reliability
substantially increased. The principal uncertainty
associated with this alternate arrangement was
that it was difficult to see how a sufficiently uni-
form hydroxide flow distribution could be main-
tained over the outside of the fuel tubes through
the core. If the flow were not uniform hot spots
might form ond rapid corrosion of the tube wall

by the hydroxide would result. As discussed
previously, both designs gave a high shield weight
because of the unfavorable geometric effects
associated with the tandem reactor—heat exchanger
arrangement,

The problems associated with the reactor core
arrangement designed for use with annular heat
exchangers (Fig. 49) are in direct contrast to thase
of the tandem heat exchanger arrangement. Al-
though the hydroxide flow through the moderator
tubes in the core could prebably be kept at a uni-
formly high velocity and hence the hydroxide tube
wall would be cooled effectively, the turbulence
pattern in the fluoride fuel flowing across the
moderator tube coils would probably be erratic and
unpredictable and local hot spots in the fuel would
be likely to occur. The hot spots in the fuel might
cause local boeiling and, possibly, instability from
the reactor control standpoint.

Reflector-Moderated Circulating«Fuel Reactor

The design shown in Fig. 65 is representative of
a series of reflector-
moderated reactors employing sodium-cooled beryl-
lium as the moderator and reflector material, A
fairly complete set of dota for these reactors is
given in Tables 12 and 13. The designs for these
have been the most carefully worked out of any
full scale ORNL-ANP reactor designs prepared to
date and hence merit special attention, particularly
since the problems dealt with are common to most
high-temperature liquid-cooled reactors.

The cross section (Fig. 65) through the reactor
core, moderator, and heat exchanger shows a series
of concentric shells, each of which is a surface of
revolution about the vertical axis. The two inner
shells surround the fuel region at the center (that
is, the core of the reactor) and separate it from
the beryllium island and the outer beryllium re-
flector. The fuel circulates downward through the
bulbous region where the fissioning takes place
and then downward and outward to the entrance of
the spherical-shell heat exchanger that lies he-
tween the moderator outer shell and the main
pressure shell, The fuel flows upward between
the tubes in the heat exchanger into two mixed-~
flow fuel pumps at the top.

circulating-fluoride-fuel

From the pumps it is
discharged inward to the top of the annular passage
leading back to the reactor core. The fuel pumps
are sump-type pumps located in the expansion tank
at the top, A horizontal section through this

B7
 

 

 

  
  
  
   
 
 
 
 

   

 

 

 
 
 
 
 

 
 

'/ 7/ 7 HEAT EXCHANGER -

// HEADER FOR Na TO

 
 

' (STRES$ =400 psi)

_ . INNER GORE SHELL —
Ly« (STRESS =200 psi)

 
 

TUBE WALL

SN

' REFLECTOR Na

  

{ / REFLECTOR--— -

OUTER CORE SHELL — |

CUTLET HEADER--

| Be ISLAND 17 1

 

(STRESS =50 psi) -
. L ~

   
  

e LT
. < S i

 

 

 

  

<,

¥

 
 

-
b

 

 

 

 

 

<

   

 
 
   

 

   

 

DWG. 22342

~—IMPELLER
(STRESS=500 psi)

* - UPPER PUMP DECK

.~ LOWER PUMP DECK

 

TN
——TUBE HEADER SHEET
N (STRESS =1250 psi)

. TUBE WALL
A (STRESS = 45C psi)

- - REFLECTOR
SHELL

8'C LAYER

INCONEL JACKET

  
 

  
  
 
   

71 COOLING TUBE
CONNEGTER

v

wdj
g 4 TUBE BUNDLE
vy SPACERS

JF  ghl/ " TUBE BUNDLE
i /7 JACKET FOR 8'°
/ / LAYER

     
  

- B0 LAYER

- PRESSURE SHELL
NER

T Na PASSAGE
/ " PRESSURE SHELL
< {STRESS = 625 psi)

T ({STRESS = 1250 psi)

- TUBE HEADER SHEET
(STRESS = 600 psi )

Fig. 65. Reflector-Moderated 100-Mw Reactor, Stresses for key structural elements are given,
TABLE 12. PRINCIPAL DIMENSIONS OF A SERIES OF REFLECTOR-MODERATED

CIRCULATING-FUEL REACTORS

Power, megawatts
Core diameter, .in.
Power density in fuel, kw/emS

Pressure shell outside diameter, in,

Fuel System
Fuel volume in core, f3
Core inlet cutside diameter, in.
Core inlet inside diameter, in.
Core inlet areq, in.2
Fuel volume in inlet and outlet ducts, f3
Fuel volume in heat exchanger, ;3
Fuel volume in pump and plenum, 3
Total fuel volume circulating, 13

Fuel expansion tank volume, ftz (8% of system volume)

Fuel Pumps
Fuel pump impeller diameter, in..
Fuel pump impeller inlet diameter, in.
Fuel pump impeller discharge height, in,
Fuel pump shaft center line to center line spacing, in.
Plenum chamber width, in.
Plenum and volute chamber length, in.
Plenum end volute chamber height, in.
Impeller rpm ‘
E stimated impeller weight, b
Impeller shatt diameter, in,
Impeller overhang, in.

Critical speed, rpm

Sedium Pump
Na pump impeller diameter, in.
Na pump impeller inlet diameter, in.
Na pump impeller discharge height, in.
Ha exponsion: tank volume, 13 (]D% of system volume)
Na in Be paséoges, 3
Na in pressure shell, fi

Ma in pump and heat exchanger, f13

Fuel-te-NaK Heat Exchanger
Heat exchanger thickness, in.
Heat exchonger inside diameter, in.
Heat exchanger outside diameter, in.
Heat exchanger volume, £
Angle between tubes and equatorial plane, deg
Number of tubes
Tube diameter, in.
Tube spacing, in.
Number of tube bundles

Tube arrangement in each bundle

50
18
1.35
48.5

1.3
10

40
0.4
1.25
0.3
3.25
0.26

075
3.5
1.1
20
14.5
30
2.0
2700

1.5
12
6000

3.4

2.4
0.75
0.08
0.43
0.15
0.20

1.7
42
45,4

27
2304
0.1875
0.2097
12

B x 24

100
18
2.7
50.6

1.3
10

40
0.4
2.5
0.3
4.5

0.36

4.5
1.5
21
15
31
2.0
2700
12
1.75
13
6000

4.1
2.9
0.9
0.09
0.47
0.15
0.25

2.75
42
47.5
10

27
3744
0.1875
0.2097
12

13x 24

200
20
3.9
56.4

11
7.7
49
0.5

0.5
7.8
0.62

8.5
5.5
1.8
22,5
15.5
33
2.4
2500
17

14
5200

5.0
3.5
1A

4.65
44

53.3
20

27
6600
0.1875
0.2119
12

22 % 25

300
23
3.9
62.0

2.7
12.8

67
0.7
7.5
1.0

11.9
0.95

10
6.75
3.2
27
17.5
37
3.0
2300
24
2.25
15
5000

5.9
4,2
]l2

5.9

.47
58.8
30

27
9072
0.1375
0.2094
12

28 x 27

89
TABLE 12 {continued)

Moderator Region
Volume of Be plus fuel, £13
Volume of Be only, 3
No. of coolant holes in reflector
No. of coolant hales in island
Na coolant tube inside diameter, in.
Na coolant tube wall thickness, in.

Na pressure shell annulus thickness, in.

region is shown in Fig. 66. A pump of the type
proposed recently completed 1600 hr of operation
in a fluoride system with pump inlet temperatures
ranging from 1000 to 1500°F.

The moderator is cooled by sodium which flows
downward through passages in the beryllium and
back upward through the annular space between
the beryllium and the enclosing shells, Two
centrifugal pumps at the top circulate the sodium
first through the moderator and then through the
small toroidal sodium-to-NaK heat exchangers
around the outer periphery of the pump-expansion-
tank region. Two sodium pumps and two sodium-
to-NaK heat exchangers are provided so that failure
of one pump will not completely disable the re-
actor. Two fuel pumps were provided for the same
reason.

The design of Fig. 65 presumes that canning of
the beryllium will be required to protect it from
the sodium, but that trace leaks of sodium through
the Inconel can connections can be tolerated. As
a result, the Incone!l canning tubes that would be
fitted into the rifle-drilled holes in the reflector
were designed to be driven into tapered hores in
the fittings shown at the equator, while the outer
ends of these same tubes would be rolled into
their respective header sheets at the top and
bottom, The tube-connecting fittings at the equator
would also serve as dowels to locate the two
beryllium hemispheres relative to each other.
Corrosion tests on the berylliumesodium-lnconel
system are under way, and preliminary tests indi-
cate that there is good reason to hope that it will be
possible to allow the sodium to flow directly over
the berylliuvm; if so, the rather complex canning
operation would be unnecessary. Of even more
importance, however, elimination of the Inconel
canning would remove poison from the reflector and
reduce both the critical mass and the production
of capture gammas {and hence the shield weight).

90

22.4 22.4 25.8 31.5
21.1 21.1 24,0 28.8
208 208 554 554
86 86 210 210
0.155 0.187 0.187 0.218
0.016 0.016 0.016 0.016
0.125 0.125 0.187 0.200

The spherical-shell heat exchanger that makes
possible the compact layout of the reactor—heat
exchanger assembly is based on the use of tube
bundles curved in such a way that the tube spacing
is uniform irrespective of “latitude.”>® The indi-
vidual in headers that
resemble shower heads before the tubes are welded

tube bundles terminate

This arrangement facilitates assembly
because it is much easier to get a large number of
small tube-to-header assemblies leaktight than one
large unit,

in place.

frurther, these tube bundles give o
rugged, flexible construction {resembling steel
cable) that is admirably adapted to service in
which large amounts of differential thermal expan-
sion must be expected. This basic tube bundle
and spacer construction was used in a small NaK-
to-NaK heat exchanger that operated for 3000 hr
with a NaK inlet temperature of 1500°F3% and in a
fluoride-to-NaK  heat exchanger that
successfully for over 1600 he,37

The allowable power density in the fuel region
may be limited by radiation~-damage, control, moder-
ator-cooling, or

operated

considerations,
While the experimental results obtained to date
are difficult to interpret, no clearly defined radi-
ation-damage limit to the power density has been
established, and it is entirely conceivable that
radiation-damage considerations will prove to be
less important than other factors in establishing a
limit on power density, The kinetics of reactor
control are very complex, Work carried out to date
indicates that conirol considerations are likely

hydrodynamic

 

SSA. P. Fraas and M, E. LaVerns, Heat Exchanger
Design Charts, ORNL-1330 (Dec. 7, 1952).

6G. H. Cohen, A, P, Fraas, and M. E. LaVerne, Heat
Transfer and Pressure Loss in Tube Bundles for High-
Performance Heat Exchangers and Fuel Elements,

ORNL-1215 (Aug. 12, 1952).

578.. Wilner and H. Stumpf, Intermediate Heat Ex-
changer Test Results, ORNL C¥F-54-1-155 (Jan. 29,
1954),
TABLE 13. HEAT TRANSFER SYSTEM CHARACTERISTICS FOR A SERIES OF
REFLECTOR-MODERATED CIRCULATING-FUEL REACTORS

REACTOR POWER, megawatts

Fuel-to-Nak Heat Exchanger and Related Systems

Fuel temperature drop, °F

NaK temperature tise, °F

Fuel AP in heat exchanger, psi

NaK AP in heat exchanger, psi

Fuel flow rate, |b/sec

NaK flow rate, Ib/sec

Fuel flow rate, cfs

NaK flow rate, cfs

Fuel velocity in heat exchanger, fps

Fuel flow Reynolds number in heat exchanger
NaK velocity in heat exchanger, fps
Over-all heat transfer coefficient, Bfu/hr'ffz'OF

Fuel-NaK temperature difference, °F

Sodiumeto-Nak Heat Exchanger and Related Systems

Na temperature drop in heat exchanger, °F
NoK temperature rise in heat exchanger, °F
Na AP in heat exchanger, psi

NoK AP in heat exchanger, psi

Power generated in islond, kw

Power generated in reflector, kw

Power generated in pressure shell, kw

Na flow rate in reflector, |b/sec

Na flow rate in island and pressure shell, Ib/sec
Total Na flow rate, |b/sec

No temperatfure rise in pressure shell, °F
Na AP in pressure shell, psi

Na tempercture rise in island, °F

Ne AP inisland, psi

Na temperature rise in reflector, °F

Na AP in reflector, psi

Na-Nak temperature difference, °F

Shield Cooling System

Power generated in 6-in, lead layer, kw

Power generated in 24-in. HZO layer, kw

to limit the power density in the reactor to a value
such that the temperature rise in the circulating
Hluoride fuel will not exceed something like 1000
to 2000°F /sec. A 2000°F /sec temperature rise in
the fuel would imply a power density of approxi-
mately 4 kw/cm®. The difficulties associated with

50 100 200 300
400 400 400 400
400 400 400 400
35 51 61 75
40 58 69 85
263 527 1,053 1,580
474 948 1,896 2,844
2.1 4.2 8.4 12,6
10.5 21.0 42,0 63.0
8.0 9.9 1.2 12,2
4,600 5,700 6,700 7,000
36.4 44.9 50.9 55.4
3,150 3,500 3,700 3,850
95 100 110 115
100 150 150 150
100 150 150 150
7 7 7 7
7 7 7 7
500 1,000 2,000 3,000
1,700 3,400 7,500 11,200
190 350 500 620
53 72 154 234
22 28 51 76
75 100 205 310
28 39 30 26
4 6 6 6
72 m 120 124
32 21 12 12
100 150 150 150
36 27 18 18

43

110 210 300 350
<3 <6 <12 <18

cooling the moderator and with the hydrodynamics
of fuel flow through the core increase with power
density, as shown in the section on ““Temperature
Distribution in Circulating-Fuel Reactors.”” The
results of that work alse indicate that it would be
desirable to keep the average power density in the

91
¢6

   

DWS E-179864

   
  
   

~—FUEL PASSAGE TO CORE

_—Na PASSAGE IN ISLAND

——- —Na T0 REFLECTOR

 

 

TUheoNoK INLET

 

- Na- TO - NaX HEAT EXCHANGER

-————  NoK INLET

Na RETURN FROM ISLAND
AND PRESSURE SHELL —— -

Fig, 66. Horizontal Section Through Pump Region of Reflector-Moderated Reactor.
fuel to about 4 kw/em®, This power density would
mecn that a core diameter of 21 in, would be re-
quired for a 200-Mw reactor, :

Two major tenets of the design philosophy have
been that the pressures throughout the systems
should be kept fow, particularly in the hot zones,
and that all structure should be cooled to a temper-
ature approximately equal to or below that of the
secondary coolant leaving the heat exchanger.
Great care was exercised in establishing the pro-
portions of the designs presented in Table 12 to
satisfy these conditions. The temperature, pres-
sure, and stress values calculated for the various
stations in a typical design are indicated in Fig.
65. The stresses in the structural parts have been
kept to @ minimum and the ability of the structure
to withstand these stresses has been made as
great as practicable. Thermal stresses are not
indicated on Fig. 65, because it is felt that they
will anneal out at operating temperatures and, at
worst, will cause a little distortion which should
not be serious. Examination of Figs. 35, 36, and
37 discloses that the pressure stresses in the
major structural elements of Fig. 65 are quite
modest.

SECONDARY FLUID SYSTEM
The ORNL effort has been devoted almost wholly

to the reactor and shield, but a small amount of
preliminary design and developmental work has
been done on the rest of the system. This has
been necessary portly because the feasibility of
the power plant as a whole depends to o consider-
able degree on the components outside the shield
and partly because only by doing work of this
character has it been possible to evaluate the
incentives toward higher temperatures and power
densities and such factors as the penalties at»
tached to low<temperature moderator systems. Other
factors that could also influence reactor and shield
design are items such as the size and the shape
of ducts through the shield dnd over-all system
control.

Figure 67 shows a schematic diagram of a typi-
cal complete power plant system based on a circu-
lating-flucride~-fuel reactor, The major part of the
heat generated in the reactor would be transferred
directly from the fuel to the NaK in the intermedi-
ate heat exchanger. About 4 to 5% of the heat
developed would appear in the mederator-cooling
system, which would operote at a somewhat lower

This
heat could be removed by possing perhaps 20% of
the NaK returning from the turbojet engines through
a heat exchanger that would serve to preheat the
NaK before it passed to the main heat exchanger
and, at the same time, would cool the sodium in
the moderator circuit. it might facilitate system
temperature control if the moderator sodium system
were cooled by a separate NaK circuit. This
circuit might be used to heat compressed air for
air turbines to drive the reactor pumps. The air
might be bled off the turbojet compressors or it
might be supplied by a separate compressor. In
either case, the heat would be employed te good
advantage and would not be simply wasted.

temperature than that of the fuel system.

Quite a number of different coolants have been
considered for use in the secondary system. In
addition to the molten metals and fused salts
included in Table 9, it might be possible to use
some other fused salts with less favorable nuclear
properties but more favorable physical properties,
in particular, lower melting points,
number of such salts might be substituted for NaK
in the secondary system, but to date none having
a melting point below 500°F has been suggested.
[t has been felt that the advantages associated
with the essentially roomstemperature melting
point of NoK more than offset the fire hazard
Table 14 lists some of the
fluids that have been proposed for the secondary
system, together with some of the measures of
their desirability, It is clear from this table that
lead, representative of the heavy metals, gives
system weights that are quite out of the question,
Lithium appears to be the most promising from the
weight standpoint, but it cannot be used, at least

Any one of a

inherent in its use.

for the present, because it gives severe mass
transfer at temperatures above 1200%. Also, its
high melting point would probably be a serious
handicap in service. The fused salts are less
effective heat transfer fluids and would therefore
require larger intermediate heat exchangers; also,
they have high melting points. The small weight
advantage of sodivm in comparison with NaK
seems to be more than offset by the 200°F melting
point of sodium which, while not very high, would
present a greater service problem than the 56%F
melting point of the parficular NaK alloy assumed.
(Eutectic NaK melts at 159, but has a somewhat
lower specific heat.)

93
MANIFOLD

FROM
OTHER
ENGINES

 

 

/ FUEL-PUMP DRIVE TURBINES,

 

AlR

 

¥

 

CHECK

TURBINE

 

 

 

TWO REQUIRED

—_———P

o
Dgg 2157

Na-TO-NaK REFILECTOR COOLANT HEAT
EXCHANGER, TWO REQUIRED

 

VAIVE

 

 

 

 

 

 

   

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

AIR —#=

DIFFUSER —

94

  

 

 

 

 

 

 

 

 

 

T FUEL-TO-NaK INTERMEDIATE
HEAT EXCHANGER, FOUR

REQUIRED
4
T Y
FOUR REQUIRED
AIR NaK -
TURBINE PUMP

J > T

- \\ . A - BLEED AIR
CONTROL VALVE f\ 3
NaK -TO-AIR — g
RADIATOR
\ /{
COMPRESSOR TURBINE NOZZILE

Fig. 67. Diogram of Reflector-Moderated Reactor Power System,
TABLE 14, EFFECTS OF SECONDARY CIRCUIT FLUID ON THE WEIGHT OF MAJOR SYSTEM COMPONENTS

 

 

 

WEIGHT WEIGHT WEIGHT WEIGHT OF SHIELD WEIGHT TOTAL WEIGHT
OF OF OF : PUMPS AND INCREMENT INCREMENT
LIQUID PIPES RADIATORS PUMP DRIVES RELLATIVE TO RELATIVE TO
{ib) {ib) (Ib) {ib} MakK (Ib) NaK (Ib)
Nak 2,600 3,300 6,000 1,600 ¢ 0
Lithium 900 2,300 2,700 1,600 ~1,500 ~5,100
Sodium 2,400 3,100 5,850 1,400 =400 ~1,350
Potassium 3,300 3,900 6,500 2,100 z,000 4,100
lead 57,700 5,500 2,000 1,400 2,000 60,300
NaF-KF-LiF 1,150 1,45C 3,900 650 2,000 -1,5950

 

 

 

 

 

 

 

These comparisons were made on the basis of a
200-megawatt system. In order to keep stresses
in high-temperature metal walls to conservative
values, the peak pressure in the system was limited
to 100 psi. Some attempts to optimize line size
have been made, which indicate that the 100-psi
value gives close to a minimum system weight if
allowances are made for the extra weight of pumps
and pump drive equipment and the thicker pipe
walls required for the higher pressures.

A number of methods of system conirol have been
considered, If the circulating-fuel reactor performs
as expected it will serve as an essentially con-
stant-temperature heot source., Hf the pumps are
operated at a constant rpm, the temperature rise
in the NoK passing through the intermediate heat
exchanger will be directly proportional to the
power output, but the mean temperature of the fuel
system wiil remain consfant.- Unfortunately, o
substantial amount of power is required to drive
the pumps both for the reactor and for the secondary
system, and it probably will not prove practicable
to keep the pumps running ot full speed if the

turbojet engines are idling, This will probably be
frue irrespective or whether the pumps are driven
by air turbines or by electric or hydraulic motors.
Actually, the turbojet engine characteristics are
such that it would be more desirable 1o allow the
pump speed to vary with engine speed; in fact, it
seems likely that after the turbojet sngines have
been started, the speed of the pumps could be
allowed to reach equilibrium under any conditions
from idling to full power and still give a reason-
able set of flow rotes ond temperoture rises. Since
the torque output of o turbine wheel falls off with
rpm and since the torque required to drive the
pump impeller increases as the sguare of the rpm,
it is evident that the turbine-pump system would
be exceedingly stable, Preliminary estimates
indicate that the pumpedrive turbine speed would
follow the turbojetengine speed very closely
during an acceleration of the turbojet, It is clear
that a full-scale power plant will be a quite com-
plex system and that possibilities of instability
and oscillation exist, but it also appears that the
components con be proportioned so that o stable
system will result.

95
Priority

A-1

96

MAJOR DEVELOPMENT PROBLEMS

The following outline of the key design problems of the circulating-fuel reflector-
moderated reactor and the status of these problems at this time serves as a summary of
the work that has been covered by this report and of the work that remains to be done to
provide a sound basis for the design of a full-scale aircraft power plant, The principal
reports that cover the work that has been done and an indication of the current priority
of the remaining problems is included to give some idea of the progress that has been
made and of the magnitude of the task that remains,

OUTLINE GF MAJOR RMR DEVELOPMENT PROBLEMS

Development Problem Status May 1254 Reports
Fuel Chemistry and Corrosion

Corrosion

Harp tests and simple thermal-convection
loops
High-temperature-differential, high-velocity
loops
Radiation Damage and Corrosion
In-pile capsule tests

In-pile loop tests

Physical Properties
NaF-KF-LiF, NaF-BeF,, NaZrFyg, ete.
NaF-RbF-LiF

Other fuels and fuel carriers

Selubility of UF4 and UF3
Methods of Preparation
Xenon Removal

Reprocessing Techniques

High-Performance High-Temperature Heat
Exchangers

NaK-to-NakK
Pressure losses for flattened-wire tube-
spacer arrangement

Heat transfer and endurance test

NaK-to-Air
Fabricability, perfoermance, and
endurance tests (including study of

character of failure)

Fluoride-to-NaK
Tube-to-hecder welding, endurance and
performance tests
Effects of trace leaks, and fabricability
of spherical shell type

Much favorable data

No data

Some favorable data
Neo data, equipment

being assembled

Adequate data

Data expected soon

Data expected by
Dec. 30, 1954

Some data

Considerable experience

Little data

Seme favorable data

Adequate data

More tests needed

Mare tests needed

More tests needed

Little data available

ORNL-1515, -1609,
-1649, -1692

ORNL-1649, -1692
ORNL-1692

ORNL. CF-53-3-261

ORNL-1215

ORNL-1330

ORNL-1509, -1692

ORNL. CF-54-1-155
OUTLINE OF MAJOR RMR DEVELOPMENT PROBLEMS (continued)

Priority Development Problem Status May 1954 Reports

Shielding

Preliminary Designs

Lid Tank Tests of Basiec Configurations
Effects of thickness of reflector,
pressure shell, lead, and boron layers
Estimated Full«Scale Shield Weights
Effects of power, power density,
degree of division
Activation of Secondary Coolant
Estimated
Measurements for neutrons from core
Measurements for neutrons from heat

exchanger

Measurement of Short-Half-Lived Decay

Gammas

Refined Lid Tank Tests

Experiments on Air Scattering

Static Physics

Multigroup Calculation—~Effects of
Moderator Materials
Effects of core diameter, fuel-region
thickness, reflector thickness,
reflector poisons, and special

materials

Critical Experiments

Critical maoss with various fuel
regions—Na, fluoride, fluoride-
graphite

Control red effects (rough)

End duct leakage

Danger coefficients for Pb, Bi, Rb,
Li’, Na, Ni, etc.

. Check on Multigroup Calculation
Twao-region
Three-region
Core shell effects

EHects of end ducts

Many designs available ANP-53, Y-F15-10,
ORNL-1575
Adeguate dota for ORNL-1615

preliminary design

Adequate data for ORNL-1575

preliminary design

Data adequate , ORNL. 1575
Data adequate ORNL-1616
Data needed

Tests in progress

Tests planned for late

1954

Tests in progress

Adequate data expected ORNL-1515
by Sept. 1, 1954

Preliminary tests ORNL-1515

promising

Some test data available
Some test data available

Some test data available

Some data expected by
Sept. 1, 1954

Some data expected by
Sept. 1, 1954

Some data expected by
Sept. 1, 1954

Some data expected by
Oct, 1954

Some data expected by
Dec. 1954

97
OUTLINE OF MAJOR RMR DEVELOPMENT PROBLEMS (continued)

Priority Development Problem

Danger coefficient

Control rod effects

Moderator Cooling

Estimation of Heat Source Distribution

Be-Na-lncone!l Corrosion Tests

Static capsule tests

Harp tests
Al High-~temperature-differential, high-velocity
loop
A-1 Thermal Stress and Distortion Test with High

Power Density

Effects of Temperature, AT, Surface Volume

Ratio, etc,

1 Creep-Rupture Properties of Inconel Under

Severe Thermal Cycling

Pumps
Shakedown of Pumps with Face-Type Gas
Seals
1 Model Tests of Full-Scale Pump
1 Endurance Tests of Full-Scale Pump
1 Fabricability of Full-Scale Pump Impeller

Power Plant System

Preliminary Designs

Performance and Weight Estimates

Effects of Temperature, Power Density,

Shield Division, etc,

Reactor Kinetics

1 Theoretical Analyses

ARE Temperature Coefficient Measurements

1 Xenon Effects

98

Status May 1954

Some data expected by

Dec. 1954

Some data expected by

Dec. 1954

Good estimates made

Some favorable data
Some data

Some data

Test nearly ready to run

Tests planned, data
badly needed

Tests planned, data
badly needed

Adequate data for design

Tests being run
Tests planned

Tests planned

Adequate data

Adequate data for

preliminary design

Adequate data for

preliminary design

Preliminary analysis

completed

Tests planned

Data badly needed

Reports

ORNL.-1517

ORNL.-1692
ORNL.-16%92

ANP-57, ORNL. -
1255, -1215,
-1330, -1509,
-1515, -1609,
-1648

ANP-57, ORNL-
1255, -1215,
-1330, -1509,
-1515, -1609,
-1648

ORNL CF-54-2-185

ORNL CF.53-3-231
 

OUTLINE OF MAJOR RMR DEVELOPMENT PROBLEMS (continued)

Priority Development Problem

Hydrodynamic Tests
1 Flow Separation at Core Inlet

Effects of Heat Genaration in the
Boundary Lavyer

Fill and Drain System
Preliminary Design

Water and High-Temperature Tests

High-{emperature Test with Radiocactive

Material

Full-Scale Reactor Tests
Control

1 Temperature coefficient

Xenon effects
Stability

Performance
Heat exchanger, pumps, etc.

Temperature distribution

Endurance Tests

 

Status Moy 1954 Reports

Y-F15-11, ORNL.
1692

ORNL-1701

Some data available

Theoretical analyses com-

pleted for ideal case

Design looks promising

Tests planned for fall,
1954

Tests might be run in 1955

Some information expected
from ARE
Information badly needed

Information badly needed
Tests being planned

Tests being planned

Tests being planned

99