ORNL-2599.txt

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ORNL-2599
C-85 — Reactors-Aircraft Nuclear
AEC RESEARCH AND DEVELOPMENT REPORT Propulsion Systems 7as

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3 445k 0251081 5

M-3679 (22nd ed.)

L

AIRCRAFT NUCLEAR PROPULSION PROJECT
SEMIANNUAL PROGRESS REPORT

A
FOR PERIOD ENDING SEP TEMBER 30, 1958 : &é

REpT Feetrrn
CATION Crance: ' R R
AEC 9-g-¢ < 7g
OAK RIDGE NATIONAL LABORATORY
operated by

UNION CARBIDE CORPORATION
for the

U.S. ATOMIC ENERGY COMMISSION

LEGAL NOTICE

This report was prepared as an account of Government sponsored work. Neither the United States,
nor the Commission, nor any person acting on behalf of the Commission:

A. Mokes any warranty or representation, express or implied, with respect to the accuracy,
ion contained in this report, or that the use of
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any information, apperatus, method, or process disclosed in this report may not
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any information, apparatus, method, or process disclosed in this report.
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with the Commission.

ORNL-2599
C-85 —~ Reactors-Aircraft Nuclear
Propulsion Systems

M-3679 (22nd ed.)

This document cansists of 232 pages.

Copy ?/ of 234 copies. Series A,

Contract No. W-7405-eng-26

AIRCRAFT NUCLEAR PROPULSION PROJECT
SEMIANNUAL PROGRESS REPORT

For Period Ending September 30, 1958

A. J. Miller, Project Coordinator

DATE ISSUED

JAN 22 1359

OAK RIDGE NATIONAL LABORATORY
Oak Ridge, Tennessee
operated by
UNION CARBIDE CORPORATION
for the
U. S. ATOMIC ENERGY COMMISSION

MARTIN MARIETTA ENERGY SYSTEMS LIBRA

- i

3 445& 0251081 5

FOREWORD

The ORNL-ANP program primarily provides research and development support in
shielding, reactor materials, and reactor engineering to organizations engaged in the
development of air-cooled and liquid-metal-cooled reactors for aircraft propulsion. Most
of the work described here is basic to or in direct support of investigations under way
at General Electric Company, Aircraft Nuclear Propulsion Department, and Pratt &
Whitney Aircraft Division, United Aircraft Corporation. This report is divided into rour

major parts: 1. Metallurgy, 2. Chemistry and Radiation Damage, 3. Engineering, and
4, Shielding.
'

ANP PROJECT SEMIANNUAL PROGRESS REPORT

SUMMARY

PART 1. METALLURGY
1.1. Fabrication

Studies of the effect of heat treatment on the

mechanical properties at room temperature of arc-
" cast columbium prepared by the electron-beam
melting technique have indicated that specimens
heated at intermediate temperatures in the range
700 to 1000°C are more ductile than those heated
at the extremes of the range of treatment tempera-
tures. There is evidence, however, that the in-
creased ductility resulting from intermediate tem-
perature anneals of welds would not be retained at
higher temperatures and that stabilizing elements
will be required to neutralize the effects of inter-
stitial elements,

Electron-beam-melted columbium billets were
found to extrude readily if they were protected from
atmospheric contamination during heating, Arc-
melted material was found to be more difficult to
extrude, possibly because of a higher impurity
content,

Cracks found in drawn duplex and unclad colum-
bium tubing have indicated insufficient annealing
before tube reduction, and therefore a high-tempera-
ture, high-vacuum annealing furnace was designed
and is being built,

Since it appears that the production of satis-
factory columbium bodies will require either high-
purity metal or special alloys, purification and
alloying investigations are under way. Resistance
heating of columbium wire in vacuum at tempera-
tures up to its melting point has shown significant
reductions in nitrogen and oxygen content. Melting
stock has been obtained from which special alloys
will be prepared. Beryllium, zirconium, and cerium
will be used to neutralize residual oxygen. High-
purity columbium specimens contaminated with
known quantities of oxygen or nitrogen are being
prepared for corrosion testing in lithium,

Methods for obtaining highly pure yttrium metal
and hydrogen are being studied in order to de-
termine the effects of impurities on the hydriding
of yttrium. It is hoped that purification will alle-
viate the tendency of yttrium metal to crack during
hydriding and during thermal cycling of the hydrided
metal, The electron-beam melting process was

shown to be useful for removing fluorides and
traces of magnesium from yttrium metal, but it was
ineffective in removing traces of oxygen or nitrogen
or farge quantities of magnesium. Equipment for
the production of yttrium metal from YF 5 has been
expanded to pilot-plant scale,

1.2. Corrosion

Studies of the various methods of purifying lithium
of oxygen and nitrogen have continved. As part of
these investigations, samples were taken of the
lithium shipments received from vendors and ana-
lyzed for oxygen by several methods in order to
compare various analytical techniques. The analy-
ses showed that it will probably be necessary to
purify all lithium received.

Experiments are being conducted to determine
the solubility of lithium nitride and lithium oxide in
tithium, An average value of approximately 1300
pPpm oxygen was obtained recently that is consider-
ably lower than the value obtained previously.

The ability of yttrium to getter impurities from
lithium is being investigated. Hardness increases
and vacuum-fusion analysis results indicate that
yttrium getters both oxygen and nitrogen from
lithium, but no reductions in the oxygen or nitrogen
contents of the lithium bath have been found.

Static tests of several grades of arc-cast colum-
bium tubing have been conducted. Preliminary
results indicate that surface contamination and
subsequent diffusion of the impurities into the
cofumbium during annealing may account for some
of the previously observed attack of columbium by
lithium,

Two joining techniques for columbium, braze
welding and fusion welding, and three types of
gaseous environment were investigated to determine
their effects on the hardness of columbium and its
corrosion resistance to lithium. A marked reduc-
tion in depth and severity of attack was noted as
the purity of the gaseous welding environment was
improved. Appreciable decreases in the hardness
of the as-welded columbium specimens as a result
of exposure to lithium again indicated gettering of
oxygen from the columbium by the lithium.

A series of seesaw-furnace tests of molybdenum
in contact with lithium was conducted. The results
indicate, in general, that molybdenum has excellent
resistance to attack and mass transfer at hot-zone
temperatures as high as 1900°F, Dissimilar-metal
mass transfer of Inconel to the molybdenum was
observed in the Inconel-sodium-molybdenum section
of the test capsule. No molybdenum failures either
in the base material or in weld zones have been
encountered in the tests conducted thus far,

Similar seesaw-furnace tests of columbium-lithium
systems have been initiated. In the one test (hot
zone, 1600°F; cold zone, 1100°F; test duration,
300 hr) completed to date, columbium exhibited
excellent resistance to attack and mass transfer,

Static corrosion tests of beryllium specimens in
lithium in iron capsules and in beryllium capsules
at 1500 and at 1830°F showed beryllium to be
quite resistant to attack by static lithium in an
all-beryltium system. It was very heavily attacked,
however, when tested in the iron container,

Refractory-metal-base brazing alloys are being
corrosion tested in high-temperature lithium, since
commercially available brazing alloys have very
limited corrosion resistance to lithium,

A method was developed (under subcontract at
NDA) for measuring the rates at which container
metals dissolve in liquid lithium, The solution rate
measurements confirmed the deleterious effects of
nitrogen and oxygen and the beneficial effect of
aluminum on mass transfer in lithium as noted in
thermal-convection loop tests.

1.3. Welding and Brazing

A survey of phase diagrams of refractory-metal-
base binary systems has revealed several promising
brazing alloys. Vacuum-brazing equipment has
been constructed for studying these systems, and
flowability and corrosion tests are under way.
Studies of the diffusion of the brazing alloy into
columbium are also being initiated. If diffusion or
alloying occurs it may increase the remelt tempera-
ture of the brazing alloy and thus increase the
maximum service temperature,

A small,glass dry box containing a stationary
torch and a movable table was assembled for
welding studies on columbium and molybdenum.
This equipment will be used in investigating the
weldability of these materials under highly con-
trolled degrees of atmospheric purity. Several butt
welds made on arc-cast columbium sheet in a puri-
fied argon atmosphere were found to be ductile,

vi

1.4. Mechanical Properties

Equipment has been developed for tensile and
creep tests of materials in controlled environments
at temperatures up to 2500°F. Several successful
tests have been completed on molybdenum at
2000°F.

Creep and rupture properties of Inconel have been
investigated in tubes with both internal pressure
and axial stress, The creep data obtained compare
favorably with the theoretical curves derived from
Soderberg’s analysis and the von Mises and Tresca
flow rules. Fracture cannot be predicted by the
maximum principal stress criterion,

Dynamic load tests have been conducted on
Inconel rods and tubes that were programed on the
basis of plastic strain rather than stress. A simple
relationship, in the form N = K, exists between
the number of cycles to failure and the plastic
strain absorbed in each cycle, The constants a
and K are found to vary as a function of the ma-
terial and test conditions. On the other hand,
geometry, temperature variations in the range 1300
to 1600°F, and small variations in chemical compo-
sition do not significantly alter the cycle life of a
material, Strains induced thermally show excellent
correlation on the basis of equivalent amounts of
plastic strain per cycle with mechanically induced
strains in temperature ranges where the material is
metallurgically stable. The frequency at which the
material is strained is observed to infiuence the
number of cycles that can be survived. The slow
frequencies, particularly at low amplitudes, appear
to produce the most deleterious results.

1.5. Ceramics

Small batches of zirconium boride were synthe-
sized for incorporation into BeO bodies. The ma-
terial prepared by an improved process was found
to be >97% pure. Several additives were tested as
means for increasing the density of BeO. Of the
various combinations tested, it was found that a
body with 94 wt % BeQ, 5 wt % MgO, and 1 wt %
B4C had the greatest density.

Screening tests were run to determine the oxida-
tion resistance of TaB,, TiB,, CrB2, HfBz, BN,
and ZrB, when added to high-density BeO. The
ZrBz-BeO and HfBz-BeO combinations were found
to be superior to the other mixtures. Attempts to
fabricate large hot-pressed blocks of the ZrB,-BeO
mixture for which a commercial grade of ZrB_ was
used revealed the need for pure ZrB.,. The commer-
cial grade of ZrB2 was found to contain about 10%
Fe. Uniform blocks were obtained when pure ZrB,
was used,

1.6, Nondestructive Testing

Studies of the use of x-ray sensitive Vidicon in a
closed-circuit television system for remote viewing
of x-ray images were continued. It was found that
the sensitivity of the selenium photoconductor to
x-rays could be improved by admitting visible light
to the photoconductor surface, Contrast sensitivity
was increased by as much as a factor of 3. Thick-
ness changes of about 2,5% were detected in E}B-in.-
thick aluminum,

Investigations are being made of the applicability
of existing ultrasonic techniques to the inspection
of duplex tubing, It appears that the resonance
ultrasonic technique can be developed as a means
for evaluating the quality of the bond between the
materials,

1.7. Metallography

A 20-tube semicircular heat exchanger fabricated
of Inconel in which a fused salt flowing around the
tubes exchanged heat to NaK flowing in the tubes
was examined metallographically to determine the
location and effects of the leak that caused termi-
nation of operation. The heat exchanger, which
had operated 1139 hr at 1200°F and above, had
experienced 194 thermal cycles between isothermal
operation at 1200°F and operation at design condi-
tions (fused salt inlet and outlet temperatures of
1600 and 1250°F and NaK inlet and outlet tempera-
tures of 1070 and 1500°F.)

The leak of NaK to fused salt was found to have
been the result of fracture of three tubes in the
high-temperature header area. The cracks were
radial and propagated from the tension side of the
tubes; they were in the tubes with the shortest
tube bend-to-tube sheet length (]'/8 to 3/4in.). The
maximum depth of corrosion attack on the fused-
salt side of the tubes was 0,006 in., and, on the
NaK side, it was 0.002 in.

PART 2. CHEMISTRY AND RADIATION DAMAGE
2.1. Materials Chemistry

Sufficient fluoride mixture for the first charge
of the equipment in the yttrium-metal pilot plant was
prepared. The YF, needed was prepared by direct
fluorination of dry Y,0,. A mean conversion
efficiency of 99% and suitably low oxygen contents
were obtained. Magnesium fluoride for the process

was prepared by dry hydrofluorination of MgO; the
mean conversion efficiency was 97.7%.

Sufficient information with which to construct a
portion of the phase diagram of the LiF-YF3 system
was obtained, A single eutectic occurs in the
system at 19 mole % YF,; mp, 682°C.,

Further experimental information was obtained
regarding a tentative procedure for purifying lithium
metal by extracting oxide, nitride, and carbide
impurities with a eutectic mixture of lithium halide
salts. The second extraction of each series of
extractions showed considerable reduction in the
concentration of impurities.

2.2, Analytical Chemistry

The method for the determination of oxygen in
fluoride salts by fluorination with KBrF , was
successfully applied, The coefficient of variation
of the method, as based on duplicate analyses of
seven samples, was 6% in measurements of the
salt LiF-Mng-YFB. The recovery of oxygen added
as MgO and Y203 exceeded 98%, with a precision
of 2%.

A rapid method was adapted for the determination
of magnesium and yttrium in LiF-Mng-YF3 by
titration with a standard solution of ethylene-
diaminetetraacetic acid (EDTA). Amalgamated
gold serves as the indicator electrode. Both ele-
ments are titrated in the same solution by adjust-
ment of the pH of the solution. No interferences
were encountered in the application of the method
to LiF-MgF ,-YF,. The precision of the method is
about 1%.

2,3. Radiation Damage

Experimental apparatus is being prepared with
which to test the stability of yttrium hydride and
beryllium oxide thermally cycled at high tempera-
tures in a high gamma-ray heating region of the
Engineering Test Reactor. Mockup tests have
been run on the heat removal system and on the
thermocouple design. The heat removal system
will provide a temperature control method in that
it will be possible to vary the heat conductivity of
the cooling gas by varying the concentration of
argon and helium in the gas mixture. Stainless steel
was selected as the test capsule material as a re-
sult of compatibility tests. Examinations and meas-
urements of the samples will be made before irradi-
ation and repeated after irradiation.

Stress-rupture experiments in air on tube-burst
specimens of Inconel were completed in the MTR.

vii
Preparations are being made for similar tube-burst
tests and creep tests in the fast flux of the ORR,

In conjunction with studies of the effects of
radiation on semiconductor barriers, a cryostat was
constructed that has a temperature range of —200°C
to +350°C, Requirements, other than the tempera-
ture range, were long term reliability and “‘fail

_safe’’ operation. A cooling water failure or an
electrical failure will shut the cryostat down with-
out harmful effects to any of the system components,
Temperatures in the sample chamber can be main-
tained to within ]/4°C, with a gradient along the
length of the copper section of the sample chamber
of ]/4°C.

A grown-junction silicon diode has been irradiated
in a fast-neutron flux facility in the ORNL Graphite
Reactor to an integrated dosage of 1.8 x 101>
neutrons/cm?. Changes in forward and reverse
current at l-volt bias were recorded. A change in
reverse current on the order of a factor of 100 in-
crease was observed that is to be compared with an
increase of 200% for a comparable irradiation of an
alloy-junction diode and an increase of 30% in a
point-contact diode.

Attempts were made to alloy indium and germanium
at low temperatures, and crystal growths were pro-
duced that had not been noted before. Penetration
of indium into the germanium conformed to precise
geometric configurations, and it is thought that
such penetration may be indicative of the early
steps of the alloying process.

PART 3. ENGINEERING

3.1, Component Development and Testing

The oil-lubricated pump rotary element that is
being operated in a gamma-radiation field in the
MTR canal will have accumulated a total dosage
of 1 x 10'? by September 30. Operation of the seal
continues to be comparable to that of a seal
operated without irradiation for a similar period of
time. Analyses of the bulk and leakage oil have
indicated little change.

A Fulton-Sylphon seal on a sump-type centrifugal
pump that is circulating NaK at 1200°F, 1200 rpm,
and 1200 gpm has operated satisfactorily during
more than 7200 hr. A similar pump that had oper-
ated in the cavitation region with NaK as the
pumped fluid for 1096 hr was stopped because of
a drain valve leak. No damage was evident upon
disassembly of the pump, The data are being

viii

analyzed. A similar pump that is operating with a
fused salt has logged over 10,300 hr of operation. -
A third thin-shell mode| that was tested under
high-temperature thermal cycling conditions is
being disassembled for examination. A test loop
failure caused termination of the test after 602
thermal cycles; that is, 302 more than the scheduled
300 thermal cycles.

3.2. Heat Transfer Studies

Experimental studies of the effect of thermal-
stress cycling on structural materials were con-
tinued with the pulse-pump system, The dependence
of surface failure (cracking of Inconel) on cyclic
frequency was investigated in tests at 0.1, 0.4, and
1.0 cps. Maximum damage occurred at 0.1 cps,
with cracks as deep as 0.207 in. being observed.

At the highest frequency, no cracks were found at

the end of a 100-hr (360,000 cycles) exposure. An

attempt was made to correlate the results with

mechanical fatigue data, and, in general, it appears

that, for a given strain, failure occurs earlier (fewer

cycles) with surface thermal cycling than with

mechanical cycling. v

Analyses of data obtained from a mercury system
in the study of heat transfer in a liquid metal with
internal heat generation were completed. On the
average, the data indicate surface-to-fluid mean
temperature differences 49% higher than those
predicted by the analogy. Since the experiment was
designed so that gaseous films or scale at the wall
could not influence the experimental results, it is
concluded that the diffusivity ratio is less than
unity for mercury.

Studies of heat transfer between a heated tube
surface and air in high-velocity vortex motion were
continued. Data with both forced- and free-vortex
flow indicate values of the heat-transfer coefficient
significantly above those obtained with linear flow
and equal flow power dissipation in identical
geometries.

3.3. Instrumentation and Controls

An initial check of the data acquisition system
purchased from the G. M. Giannini Company was
made and minor troubles and design inadequacies
were noted, The system is being modified and -
should be back in operation soon.

Life tests of three level probes operating in NaK
were continued. Similar tests in sodium were
initiated. Overrange tests designed to accelerate

-

the effects of aging on pressure transmitters were
continued. Slight bulging of the Inconel bodies
was observed after 5100 hr at 1400°F.

Life tests of a series of Inconel-sheathed thermo-
couples exposed to a fused salt were terminated
after about 11,000 hr at 1500°F. The thermocouples
were performing well when the test was stopped,
but the errors were increasing. Similar thermo-
couples have been under test in a sodium environ-
ment for about 16,000 hr. Many of these thermo-
couples now show considerable error.

Studies of the fundamental behavior of thermo-
couples in the range 300 to 1100°C are under way
under subcontract at the University of Tennessee.
Work on the improvement of the Oracle program for
thermocouple-data handling is nearing completion,

3.4, Applied Mechanics

Basic studies are under way of the elastic be-
havior of five structural configurations commonly
used in reactor design. The configurations being
analyzed are a flat circular plate, a tapered circular
plate, a cylindrical shell, a conical shell, and a
spherical shell. The stress and displacement
formulas and tables of numerical values for all the
functions involved are being obtained. The study
of the conical shetl, which has been completed,
provides for analyses of a conical shell with a
combination of axisymmetrical loads. The loads
may include membrane forces and uniform pressure,
in addition to edge moments and shear forces.

3.5. Advanced Power Plant Design

As part of a program for examining the feasi-
bility of the vortex reactor concept, experimental
studies were made of the vortex strengths ob-
tainable in a gas under a variety of conditions.

A survey was made of various types of auxiliary
power units suitable for use in satellites, A pre-
liminary evaluation was made of reactor cycles
employing either aluminum chloride or rubidium
vapor as the heat transfer medium.

PART 4. SHIELDING

4.1, Shielding Theory

A calculation was performed to obtain predicted
pulse-height spectra of capture and inelastic-
scattering gamma rays which could be compared
with the experimental spectra obtained in an ex-
periment at the Tower Shielding Facility. While

experimental cross sections were available for the
calculation of the nitrogen capture gamma-ray
spectrum, many of the cross sections used to pre-
dict the spectrum of inelastic-scattering gamma
rays were theoretical, In spite of this, the shapes
of the calculated and experimental pulse-height
spectra were in agreement,

It has previously been reported that an Oracle
Monte Carlo calculation of the penetration of mono-
energetic, monodirectional gamma rays in a lead
and water shield has included a total of 512 prob-
lems. Some typical plots of the heating results
from these calculations are presented as the per-
centage of the total energy incident upon the slab
absorbed in a specified region. The results are
compared with an empirical formula,

The Oracle Monte Carlo code was used to calcu-
late the primary gamma-ray heating in a stab shield
similar to that being used in the GE-BSF experi-
ment. |t was assumed that the slab thicknesses
were the same as those in the experiment and that
they were infinite in the other two directions. Most
of the assumptions made in the calculation caused
the results to be too high, possibly as much as a
factor of 1.5. This is corroborated by GE calcu-
lations of the heating which were based on an
analysis of the thermocouple measurements and
which were a factor of 2 lower than the results of
this Monte Carlo calculation,

A Monte Carlo code is being developed for the
IBM-704 machine to calculate the angular and
energy distributions, at a point detector, of gamma
rays emitted from a monoenergetic, point isotropic,
or point monodirectional source embedded in an
infinite homogeneous isotropic medium, The code
is now in the ‘‘debugging’’ stage.

Neutron dose-rate distributions beyond water
slabs 1, 3, 4, and 6 mfp thick were calculated for
plane monodirectional, monoenergetic sources
incident on the slabs at angles of 0, 30, 60, and
75deg. The source energies considered were
0.55, 1.2, 2, 4, 6, and 8 Mev. Dose-rate buildup
factors calculated for the various sources are pre-
sented, as well as the dose rates at the rear of the
slabs resulting from neutrons multiply scattered
within the slabs,

4.2. Lid Tank Shielding Facility

The final measurements taken in the L TSF ex-
periment designed by GE-ANP for the study of the
production of secondary gamma rays in configura-
tions containing advanced shielding materials are
presented. |n most of these last configurations,

all of which were tested in oil, the effectiveness
of various thickness of stainless steel with and
without boral was investigated, One configuration
that consisted only of berylium and lithium hydride
in oil was also tested,

A study of the production of secondary gamma
rays in lead was made with lead thicknesses that
varied from 1 to 9 in. The lead was followed by
either an oil medium or a borated-water medium in
which gamma-ray dose-rate measurements were
made. From these tests it appears that the first
3 in. of lead attenvated most of the primary gamma
rays from the LTSF source plate; further, the total
dose rutes at fixed distances from the source were
not affected when the lead thickness was increased
beyond 3 in, The measurements in the borated
water were a factor of 3 lower than the dose rates
in the oii in the region close to the lead and were
a factor of 13 lower approximately 120 cm beyond
the lead.

A corrected curve of thermal-neutron flux meas-
urements in oil at the LTSF is presented. It is
16% higher than the curve presented previously.

Shields consisting of randomly distributed ab-
sorbing chunks in a relatively transparent matrix:
must be from a few per cent to several hundred per
cent greater in mass than homogeneous shields
which give the same amount of attenuation, This
is the result of radiation ‘‘channeling’’ in the
spaces between the absorbing chunks. A method
for calculating the transmission of radiation through
heterogeneous shields has been developed, and a
typical calculation for boral has been performed.
The computed results are compared with experi-
mental results.

4.3. Bulk Shielding Facility

The planned series of measurements for the GE-
BSF study of the production of heat in radiation
shields was completed, but the results are of
questionable value since it has been discovered
that all of the heating samples leaked and oil
entered the air region surrounding the samples.
The series, which consists of configurations of
beryllium and lithium hydride separated by a
gamma-ray shielding section (iron, lead, or Mallory
1000), is being repeated with new samples.

4.4. Tower Shielding Focility

The experiment performed at the Tower Shielding
Facility in cooperation with Convair, Fort Worth,
to obtain information complementary to that obtained
from the Nuclear Test Airplane (NTA) program has
been completed. During the NTA program of ex-
periments, gamma-ray and fast-neutron dose rates
and thermal-neutron fluxes were measured both
inside and outside the airframe containing the
reactor while in flight and on the ground. Measure-
ments were also made on the ground in the absence
of the airplane structure. This last set of measure-
ments was duplicated in the TSF experiments, and,
in addition, measurements were made as a function
of altitude in the absence of the airplane structure.
With these additional measurements the influence
of the air, the ground, and the aircraft structure
on the various measurements can be determined.
The TSF experiment also included angular mappings
of the radiation around the various reactor shield
configurations, near the ground and at altitude,
to obtain data which will yield more accurate source
terms than were previously available. Some gamma-
ray and neutron energy spectra were also determined.

Predictions of the gamma-ray and fast-neutron
dose rates at large distances from the Tower
Shielding Reactor |l (TSR-Il) have been made for
two operating conditions. For the first case the
reactor was assumed to be bare and to be operating
at a 5-Mw power at an altitude of 200 ft, a condition
which would result in the highest dose rates that
could be achieved with the reactor, The gamma-ray
dose-rate calculations for the bare reactor were
based on measurements taken at large distances
from the TSR-|, while the fast-neutron dose-rate
calculations were based both on the TSR-| data
and on other measurements made at large distances
from the Aircraft Shield Test Reactor (ASTR). For
the second case the reactor was assumed to be
encased in a “‘beam’’ shield which would allow a
highly collimated beam of radiation to be emitted
from the reactor. This calculation, which was
based on measurements made at large distances
from the TSR-l in a simulated beam shield, is con-
sidered to be more representative of a typical
operation. The results for both cases are plotted
as a function of distance from the reactor. As part
of this investigation, a calculation was made of
the uncollided neutron flux 4200 ft from the TSR-I

to obtain a fast-neutron buildup factor for the case

Fad
in which an air-filled collimator extending through
the reactor shield is pointed at the detector,

4.5, Tower Shielding Reactor Il

The latest design of the TSR-Il, with its asso-
ciated controls and 5-Mw water cooling system, and

several studies supporting the design are pre-
sented. The critical mass has been set at 8.1 kg
with 1,6% excess reactivity. Control will be
achieved by moving six umbrella-shaped grids of
Inconel-clad cadmium in the internal water re-
flector, A description of the first shield is also
presented,
CONTENTS
SUMMARY ...ttt ettt e et et ee e e e e e e e es o2 s ee ettt ee s e et e
PART 1. METALLURGY

LTe FABRICATION ..ottt et e et ee et seee e e e es e ses e eee oo,
Columbium INVeSH GAHONS ..cvee ettt ettt ee e e et e e s e s s ee s et sese e
Effect of Heat Treatment on Mechanical Properties w.oo.uooooioooeeeeeeeee oo
EXTPUSION STUAIS ettt et e oo et e e e s e s ne e o
PUrification EXPErimMeNts ....ooiuiuieiiiiiiieceeeeeeeee e ee et e e e e e e e oot

SPECIal Alloy Preparations ...t e eeee et s et e eeseese s e s s e e reseseeseanes

Effects of Oxygen and Nitrogen Contamination on Corrosion by Lithium.....ccocvueeeuemrervernnn,

Yttrium and Yttrium Hydride Investigation S ... iieeciieoe oo eeeee e e e v s es s eses s see s ereras e e
Effect of Metal Purity on Hydriding .cooovvvoee oot s se e
Dissociation Pressure STUdies ...t ettt ettt er e eee e aeeeesaeas
Deformation Mechanisms of YHIIUM wo.ociiiiccc ettt et es et seet e e reas
Electron-Beam Melting of Yttrium and Yttrium-Magnesiom Alloy......ocoocoieeieeieeeeeeee s
Expansion of Ytrium Preparation ProCess ..o et e e ee e eseeeeve e e e e,

1.2, CORROSION L.ttt ee et et e et ese s ees eeeseasesetessasasen s ssnsessesessseasas
Lithium PuUFHiCation ..ot et ne st ee ettt s s essaeeeeemessssemeesraenas
Columbium Exposed 10 Lithium ..ottt e et e s e e s e s e sres e s s sene e
Arc-Cast Columbium Tubing Tested in Static Lithium .......cccoeiviiiioecieece e

Effect of Welding Procedures on Corrosion ca.. ittt et eva e tee et e
Comparison of Methods for Determining the Nitrogen Content of Columbium .........o.ocovvvvuerieernneee.
Molybdenum and Columbium Tested in Lithium in Seesaw-Furnace Apparatus .cccoceeeveereerveenennee.
Berylium Exposed 10 Lithium .o et eeee e eee s et e e eae e et areaeesnaeas
Refractory-Metal-Base Brazing Alloys Exposed to Lithium .....occooiiiiiiiiiieiieeeee e,
Determination of the Solution Rate of Metals in Lithium ....o...occcooiiiiiiieciceeee e

1.3, WELDING AND BRAZING ..ot ettt ire st e sre e ba b ans v e sae st eatabssassbasssbssesnsaaensarens
Development of Brazing Alloys for Lithium Service........cocooiiiioiiiiei e,

1.4.

1.5.

Welding Studies of Columbium
MECHANICAL PROPERTIES

and Molybdenum ..o

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

Development of Test Equipment for High-Temperature Investigations ...........cccccocoovrieviccivieciiecenaa.

Multiaxial Creep Studies..........
Creep Analysis ................
Rupture Analysis................

Strain-Fatigue Studies..............
Temperature Dependence ..
Frequency of Dynamic Loa
Design Factors ....ccoeueneee.

..................................................................................................................
------------------------------------------------------------------------------------------------------------------

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

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

..................................................................................................................
dS ............................................................................................................

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

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

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

VOO NN O WWW W

xiii

1.6.

1.7.

2.1

2.2.

2.3.

3.1

3.2,

xiv

Densification of Beryllium Oxide ..ottt

Oxidation Resistance of Boron-Containing Beryllium Oxide Bodies ...c.ccocimiiiiiciiciiiiiiiiiininnne,

NONDESTRUCTIVE TESTING oottt et s seere s s e a et e e sarene san e snb bbb b ns
Remote XeRay VieWing ....ccooiiiiiiiiiiriniis ettt e s st st b ees s aaa b e n e e ae b e s abesinness
DUPLEX TUBING e bbb s e bbb e b s
Metal [dentification METEr .. ...cocuueiiiiiiiete et eeeetes bttt e ssees s aans e eeses s besbra s e s bt ae e s e rbaras s sanean treneaeares

METALLOGRAPHY Lottt es e et s ans sha s ek bbb e e s s s s e a st st e

Results of Metallographic Examination of Small Semicircular Fused-Salt Fuel-to-NaK
Heat Exchanger Operated at High Temperatures ..o

PART 2. CHEMISTRY AND RADIATION DAMAGE
MATERIALS CHEMIS TRY ittt e et st st ere e e ea e e se bbb b s b ibe s e e eas s

Preparation of Charge Material for Reduction to YHIUM .cccoooiiiiiiiiieiie e
Conversion of Y0, 10 YF 4o
Preparation of MgF , ..o
Preparation of the LiF-MgF,-YF, Mixture ..o

Phase Equilibria in the System LiF-YF ..o s

Extraction of Lithium Metal Impurities with Molten Salts ..o

ANALY TICAL CHEMISTRY .ottt ettt sttt re et sbe st s srs et e sasasess e s e e e nseasnesmsnbessssnsesias
Determination of Oxygen in Fluoride Salts ... s

Determination of Yttrium and Magnesium in LiF-Mng-YF3 ................................................................

RADIATION DAMAGE ...ttt ettt e e te s te st e s e e e s e st e e eabe st e e eaeeneesebeesesenbnansaara s
ETR Irradiation of Moderator Materials for Use at High Temperatures ...,
Creep and Stress Rupture Tests Under lrradiation ..o

Radiation Effects in Electronic Components ... it et te s e e meanse st aeesessssas
Wide-Range Multipurpose Cryostat.. ..ottt e
Grown-Junction Silicon Diode lrradiation ........cooieiieiieieee et s e
Low-Temperature Indium-Germanium AllOy .......ccooiiviiiiiiic e

PART 3. ENGINEERING

COMPONENT DEVELOPMENT AND TESTING ..ccoiiieieeciitceiee et snen s s
Irradiation Test of Oil-Lubricated Pump Rotary Element ..o,
Test of NaK Pump with Fulton-Sylphon Seal.......cccooiiiiiiircr s

Cavitation Tests of Centrifugal PUmps ..ot
Thermal Stability Tests of Metal Shells. ...

HEAT TRANSFER STUDIES w..oooovoc e SO USSR
Studies of the Effect of Thermal-Stress Cycling on Structural Materials ...
Liquid-Metal Heat Transfer Experiment ..o,

Heat Transfer with Vortes Flow .ot tetaee s e e st sen e senneaeans

3.3.

3.4.

3.5.

4.1.

4.2.

INSTRUMENTATION AND CONTROLS 94
Data ACQUISITION Sy STEM ..ottt e 94
Liquid-Metal-Level TranSducers .......cooovuomoooeeeoeoeeoeeoeee e 94
Pressure Transmitter TestS ..ot ee oo e 94
ThermoCOUPIE TeSES ........oiiiuieriirn ettt ettt ee et ee et eoee e e oo e oo eseee e 95
Thermocouple Development STUdIES ............cooemiivviieeoeeeeeeseeeeeeree oo oo oo 95
Behavior of Thermocouples in the Temperature Range 300 t6 1100°C weeveveveoeoooeeoeeooo 95
Oxidation and Cold Work STUdies .....oiuueieeeeceeeeeeeese e eeeeee e 96
Calibration SHUAIES ..ot ettt ee e et eeee e 9%
Oracle Program for Thermocouple-Data Handling ..o eeeeeoeevoeooeeoeoeeoeeeoeeoeeoeeeeeeeee 96
APPLIED MECHANICS. . ..ottt es e e 98
Basic Problems in EIastiity oiiuiiriiiiocececeeceeee e et 28
NUMEFICAl ANGIY SIS ©oiviuitiiteicecececet ettt 99
ADVANCED POWER PLANT DESIGN .. ..ot 101
Vortex Reactor Experiments.. ...cooocoviiuiiiiiiieec oo e e et 101
Survey of Design Problems of Auxiliary Power Units for Satellites ..oooomommoooooeoeoooo, 110
Cycle Performance Considerations ..o oo oo 110
PLANNE PROGIAM ..ottt e s et e e e e et 113
PART 4. SHIELDING
SHIELDING THEORY ..ottt st ee e aes s ee s e e e s ae sttt s s eee oo eeoe e 119
Analysis of Neutron Physics Division Experimental Study of Gamma Rays
Produced by Neutron Interaction i AQr ......o.ooeeeoeee oo e 119
Spectra of Gamma Rays Incident on the Outside of the Collimator .....cocooveeveeeeereveseereereorrorans 121
Spectrum After Attenuation and Buildup in the Water Collimator.o.ovimmmeeoooooeooooo, 122
Calculation of Pulse-Height Distributions from the Detector ... imreoneeeeroseeeeesreseesress oo 123
Monte Carlo Calculation of the Deposition of Gamma-Ray Heating in Stratified
Lead and Water SIabs ..o ettt e ettt 123
A Monte Carlo Calculation of the Gamma-Ray Heating in Several Shielding
Configurations Adjacent to the Bulk Shielding REQCIOr ...oc.oveiiereeereeeeeeeeeeeee e e ees e e 128
MEthod Of Caleulation ..ocoiiiiiiiii et v e eee e et e est e s eeesenee s seeons 128
CONCIUSTONS .ttt oo es e e e ee e e te et e e s s s e s e oot ee s 131
A Monte Carlo Code for the Calculation of Deep Penetrations of Gamma Rays ......cccooeevveevrennan., 133
IMPOrtanCe SAMPIING ...ttt enesesese e ne s rasesans 134
Double Systematic SAMPling .....ciiiieiiceeeeee e ettt ee s 134
Statistical EStimation ..ottt et e e 135
Weighted ISotropic SCattering ..ottt et et e s oo 135
Splitting and Russian RouUlEtte .. ...ttt st enan e 135
DUBPUL <ttt bt ettt et tetae et s st ettt ettt et se et ret e ettt ee s s ar s e ee s s e aenen 135
T eSTING PrOC@AUIE oot ettt et e et ee e tee et e ren e e e s e e s eneeeas 136
A Monte Carlo Calculation of the Neutron Penetration of Finite Water Slabs oo.voovovoeveeveveea. 136
LID TANK SHIELDING FACILITY oo ettt eee e ee e e e e s, 141
Study of Advanced Shielding Materials (GE Series) .u..iiiiiiimiicice e eeeeeee e e e e oo een e ee e 141
Study of Secondary Gamma-Ray Production in Lead.. ..o, 145

4.3.

4.4,

4,5,

xvi

Thermal-Neutron Fluxes Measured in Qil at the LTSF: An Erratum..ceiiiiiiireeiee e 148
Radiation Transmission Through Boral and Similar Heterogeneous Materials
Consisting of Randomly Distributed Absorbing Chunks . ..o 148
BUL K SHIE L DING F A C I T Y it e e e ettt teas tesaasssestbasss e beassasssarn s eraranansenes 154
The GE-BSF Study of Heating in Shields ... s 154
TOWER SHIE L DING F A CILITY oottt ettt e e e ee e aresssat v tes s e e s eeeaeentesessaensnssaeare abees 158
Aircraft Shield Test Reactor Experiment at the Tower Shielding Facility ...ccoovviiinieniciicinecc, 158
ASTR Shield Configurations ......ooiiciiiiieiee e e ettt s r e n s 158
DT OIS e e e e ettt e et e e e e e e e et ee et e et e e eeettttretetetes e ieneiatabatteattatrestran e —rtaretae et asestieraratnreseaataans 158
Radiation Mappings in the Reactor Horizontal Midplane.......ccocooiiiiiiiiiiie 158
Measurements in Crew Shield Mockups ...t 166
Gamma-Ray and Neutron Spectral Measurements .........ccoooveiiiiiiniiiiiii 168
MiSCEllaNEOUS MEASUIEMENTS .. oottt et ettt tr et teae et nsmsassesbatsnsbassanans 170
Prediction of Radiation Intensities at Large Distances from the TSR-H ... 174
Measured Dose Rates 4000 £t from TSR=l oot rnas 174
Predicted Dose Rates 4200 ft from TOR=Il oot et es s e e rara s 175
Calculation of Uncollided Fast-Neutron Flux 4200 ft from TSR-l ..., 181
TOWER SHIELDING REACTOR bl oottt e e e e et se et ae st et e nbe s s svnssaaeassrnnnes 184
Mechanical Design.......cooiiiiiiiiiiiiiiec ettt s r et et s et eb e s st s e b e e et ers 184
COMIEO| SYSTEMS ettt ettt te et e st et s e bt e st ae et ceab e e cb bt te e sae s s e ebe e nesans s e s esane s s s e et nn e 186
CONTrO ] MO O AN SN oot e ettt et et e r et e e s s e ra s e s e ss rarasa s e raaaeeas 186
Electrohydraulic Transducer for Control of Water Pressure in the
CoNtrol M AR SM oottt e et rae s e et tetes it seat s astnbe s aeraaesaeeaeaneeasnnantnsnsraeenananans 187
TSR Block Diagram.. ..o...oocoieiiiiiiiieiiitie ettt et st et st ere e s b e e e s eas e e n e 188
ReeaCtor Controls Ui e .ot e e et e ae s e e e b s raes 191
Heat Removal EQUIPMENT ...ttt et e s et ee 191
NUCTEAE CalUl G ON S oo e ettt ettt e e ee e e e s at e eeteass s aanan semsnnenns senananans 191
Comparison of Calculations and Critical Experiments ... 192
Reactivity CoeffiCients ...ttt et s r e e 193
Contro] SRl Effeetivemes s oo ettt rte e e e eee e e e e ee e enn aeeaeaeeene et eaaeebaes 194
Contro] Grid EffeCti VEMES S oo oeeieteee e eeeeee ettt et eeseeeeeee s tete e s s abae e e tatasnss abressasnenseseseseassmrnsaanees sasnbrannnnnnesans 194
Selection of Critical Mass For T ORIl oottt ettt e eeee e e e e e eeee e st etsat et nenessararnnn 195
Calculation of Heating in the Fuel Plates.......coiiiiiiii et 196
F oW DS tribUITION STUAIES woneeiiitiie e ieee ettt e e e ee e eeeee s eeeetee s e ar st s baesrn i ee baasbans e saranassissssansrnnnnnnnns 202
Flow Studies of Central Fuel Elements oot ettt it et i s eerere e re e eeee e rees 202
Flow Studies of Annular Fuel Elements ..ot tie st sit e er e e tr e s e s e rar e e e e eaes 202
ReACtOr Core Kimetic StUGI@S o oiivirii ettt ee e et e e ttee st teeee e e et aae e e e e e e e s tesesabesesns s e ansn e snaieeeeanasaians 206
Safety Sy Stem MEASUFEMENTS ....cociiiieiieticeceticee ettt et e st e ses b eabesteateetseeresbensesaer e ssane e atassenanseamenaes 206
Shi@ld DESTGNS et ettt e et e s e b e e s 210
Investigation of Stresses in the Tower Structure from Water and Electrical Lines ..o 211

-

Part 1
METALLURGY

W. D. Manly

Metallurgy Division

T. Hikido
U.S. Air Force
].].

FABRICATION

J. H. Coobs

COLUMBIUM INVESTIGATIONS

Effect of Heat Treatment on
Mechanical Properties

D. O. Hobson

Studies were continued in an investigation of the
effect of heat treatment on the room-temperature
properties of a cast columbium ingot prepared by the

H. Inouye

electron-beam melting technique. The results of
these experiments will provide information needed
for determining the feasibility of forging castings
at room temperature to prevent contamination and
for assessing room temperature evaluation of the
probable performance of fusion welds.

Notched-bend specimens machined from the ingot
were heat treated at 50°C intervals from 700 to
1000°C in a dynamic vacuum of 2 x 10~¢ mm Hg
for periods ranging from 168 hr at the lower temper-
atures to 72 hr at the higher temperatures. The
effects of these heat treatments were evaluated
from the standpoints of fracture mode, microstruc-
ture, and ductility as determined from an integrated
plot of the load-deflection curves. In all cases, the
fracture occurred transgranularly and showed no
trend with respect to the heat-treatment temper-
ature. Microstructurally, no differences were noted,
and no second phases were detected when the
samples were examined at a magnification of 1500
times. The results of triplicate slow-bend tests of
specimens heat treated at a particular temperature
exhibited considerable scatter, however. The
fractures did not always occur at the base of the
notch, and it is thought that the variations in the
location of fracture may have affected the test re-
sults. The data presented in Table 1.1.1 indicate,
however, that the specimens treated at intermediate
temperatures were more ductile than those heated at
the extremes of the range of treatment temperatures.
The values presented in Table 1.1.1 are the lowest
of the three values obtained in the triplicate tests.
The material from which the specimens were ma-
chined had the following impurities: 0.01% C,
0.12% 0,, 0.0013% H,, and 0.032% N ,.

The data of Table 1.1.1 show some correlation
with results obtained in previous cold rolling and
bend tests of similar heat-treated specimens from
this ingot. |t was found previously that an inter-
mediate-temperature anneal improved the fabri-
cability. The instability exhibited by the metal

Table 1,1, 1.

Room«Temperature Ductility of Cast Columbium

The Effect of Heat Treatment on the

Specimen dimensions: 0,125 x 0.340 x 1.00 in.

Y notch: 45 deg, 0.005-in. root radius, 0,040 in.
deep

Strain rate: 0.05 in./min

Energy Absorbed

in Room-Temperature

Specimen Heat Treatment

Time Temperature Notched-Bend Test
(hr) (°C) (in.-1b)
Asereceived 10.03
168 700 0.96
132 750 12.04
120 800 15.35
96 850 32.55
96 900 26.36
72 950 18.10
72 1000 25.20

indicates that heat-treatment effects on welds wili
be only temporary and that stabilizing elements
will be required to neutralize the effects of inter-
stitial elements.

Extrusion Studies
D. O. Hobson

Fifteen columbium billets were extruded during

H. Inouye

the period covered by this report. Nine tube ex-
trusions and two rod extrusions were successful,
and four billets failed to extrude satisfactorily,
The data obtained in these experiments are pre-
sented in Table 1.1.2. No difficulties were ex-
perienced in extruding electron-beam-melted ma-
terial so long as it was protected from atmospheric
contamination during heating. The protection was
provided by copper plating the billets, covering
them with powdered graphite, and heating them in
an inert-atmosphere furnace. Arc-melted material
tends to be more difficult to extrude, possibly be-
cause of its higher impurity content. Efforts have
been made to extrude columbium bare, that is, with-

out the copper plate, but this leads to surface con-
tamination and in the case of tube blanks results

ANP PROJECT PROGRESS REPORT

Table 1,1.2. Extrusion Data on Columbium Billets

w
Extrusion - TExfrusuon Reduction Plating Quality of
Designation* ype emp:roture Ratio Material Extruded Material
Q) -
wWC-1 Tube 815 7:1 Copper Good
wWC-2 Tube 815 7:1 Copper Good
WC-3 Tube 815 7:1 Copper Good
WC-4 Tube 815 7:1 Copper Good
WC.-5 Tube 815 7:1 Copper Good
WC-6 Tube 815 7:1 None Poor
EM-1 Tube 815 7:1 None Bad, stalled press
K-1 Tube 815 7:1 None Bad, stalled press
K-2 Tube 815 7:1 Copper Bad, stalled press
WC.7 Tube 815 7:1 Copper Good
wC-8 Tube 815 7:1 Copper Good
WC-9 Red 815 8.75:1 Copper Good
WC-11 Tube 815 7:1 Copper Good
wC-12 Tube 815 7:1 Copper Good .
K-4 Rod 1200 9.75:1 None Good
*WC: Wah Chang Corporation; electron-beam melted material. -
EM: Electro Metallurgical Corporation; arc-cast material.
K: Kennametal; arc-cast material,
in cracking and peeling of the surface. After the signed to operate at a maximum temperature of
extrusions were completed, all the good tube 1400°C at a pressure of ~ 1 x 1075 mm Hg. |t will
blanks were sent to the Superior Tube Company for be able to handle extrusions up to 36 in. long, and
drawing into tubing. it will be provided with equipment for moving the
The present invenfory of columbium fubing con- extrusions into and out Of the hOf zone. Most O{
sists of ~ 100 ft of 0.620-in.-dia, 0.090-in.-wail, the machine work has been completed, and the in-
draw-clad duplex tubing consisting of 0.030 in. of duction coil and outer furnace shell have been
columbium clad with 0.060 in. of type 446 stainless received.
steel, and ~ 200 ft of 0.500-in.-1D, 0.030-in.-wall An extrusion billet was designed for the ex-
unclad columbium. All the tubing received has trusion of duplex tubing consisting of a stainless
been submitted to the Nondestructive Testing steel outer shell and a refractory metal inner shell.
Group for inspection. A preliminary report on the The design (Fig. 1.1.2) will permit extrusion of an
inspections indicates that the unclad tubing has evacuated billet without the use of a protective
shallow (0.003 to 0.005 in. deep) cracks on both core material, The vacuum is not broken until the
the inner and outer surfaces. Since such cracks actual moment of extrusion. One type-310 stainless-
are probably due to insufficient annealing before steel-clad vanadium billet was extruded at a 16:1
tube reduction, a high-temperature high-vacuum reduction in area, and a smooth, uniform extrusion -

annealing furnace has been designed and is being
built. The fumace, shown in Fig. 1.1.1, is de-

In future extrusions columbium and
molybdenum are to be used as the inner layer.

was obtained.

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
ORNL-LR-DWG 33724

ROTARY VACUUM
SEAL WITH PULLEY~—--—.O

-

FEET

4-in. SCHED-40 TYPE 347
STAINLESS STEEL PiPE

2-in. SCHED-40 TYPE 347
HIGH-VACUUM VALVE STAINLESS STEEL PIPE

4-in, DOUBLE -~ TOUGH
+~ PYREX PIPE

S~ >_in.-ID MULLITE TUBE

2-in, VACUUM

IORE Rt L

E

nnnoonnéoo
w

s
b
P
b
g = |INDUCTION COIL
4]

T
C
D
2
o
O
m

o
o
D
D
D
P
D

4-in, DIFFUSION PUMP -
D e - QUARTZ TUBE

L |

SRR

( QQOOOOOOOOODOO‘OODOO [

";\—GRAPHiTE SUSCEPTOR
J o)
FORE PUMP

\\ — b

5~—— EXTRUSION BILLET
]
7 b

. L/ - BRASS ENDPLATE

AN
/[ A\

Fig. 1.1.1. Layout of High-Temperature High-Vacuum Annealing Furnace.
ANP PROJECT PROGRESS REPORT

ORNL-LR-DWG 33725

COLUMBIUM OR MOLYBDENUM)

V.

WELD \

S5
e

Fig. 1.1.2. Billet Design for Extrusion of Stainless-
Steel-Clad Refractory-Metal Tubing.

WELD

)
7

N

Purification Experiments
D. O. Hobson

The possibility of purifying columbium by anneal-
ing at high temperatures in vacuum is being investi-
gated. A vacuum system is being used in which a
columbium wire can be resistance heated to temper-
atures up to its melting point, 2415°C. The results
of one series of tests for 30-min periods at various
temperatures and pressures are presented in Fig.
1.1.3. In future experiments, the effect of varying

H. Inouye

the heating time at constant temperatures will be
studied.

Special Alloy Preparations
T. K. Roche

Special columbium alloy specimens are being pre-
pared for evaluation in molten lithium. The avail-
able evidence indicates that the mechanism of
corrosion of columbium by lithium is related to the
presence of oxygen in the “‘pure’’ metal. The
addition of alloying materials that are potentially
capable of neutralizing the residual oxygen should
therefore be beneficial, The alloy additions pres-
ently being considered are beryllium, zirconium, and
cerium, Columbium melting stock has been pre-
pared from sheet material received from DuPont,
which has been found, by chemical analysis, to
contain 0.032 wt % C, 0.0003 wt % H,, 0.0084 wt %
N,, and 0.053 wt % O,. Prior to the preparation of
the alloys, the fabricability of the unalloyed ma-
terial melted by normal arc-melting procedures will
be evaluated.

UNCLASSIFIED
ORNL-LR-DWG 33726

PRESSURE {mm Hg)
(x10°3) 15 15 24 2.4 28 a a5
400 :

IMPURITY CONCENTRATION (ppm)}

of S —— —

ROOM TEMP 1400 1500 1600 1700 1800 1300 2000
(AS RECEIVED)

ANNEALING TEMPERATURE (°C)

Fig. 1.1.3. Effect on Impurities af Heating Columbium
Wire in Yacuum for 30 min at Yarious Temperatures and

Pressures. (@OWiamyahmisianaayetiom

Effects of Oxygen and Nitrogen Contamination
on Corrosion by Lithium

J. E. Spruiell H. Inouye

Experimental studies of the absorption of oxygen
and nitrogen by pure columbium in the temperature
range of 800 to 1600°C at pressures ranging from
3x 104 to 1 x 10~ mm Hg were initiated. The
preliminary tests of nitrogen pickup consisted of
(1) sealing specimens of columbium and hot-pressed
boron nitride or hot-pressed silicon nitride in
quartz capsules and heating for 20, 50, and 75 hr at
temperatures of 900, 1000, and 1100°C; and (2) in-
troducing purified nitrogen of a given partial pres-
sure into a previously evacuated furnace which
contained columbium.

The results of the tests in which the columbium
specimens and boron nitride or silicon nitride were
sealed in quartz capsules were evaluated in terms
of weight change and change in chemical analysis.

VSummer employee.

The weight-change measurements indicated that
the columbium specimens had absorbed some gas,
but chemical analysis proved the absorbed material
to be primarily oxygen rather than nitrogen. The
data for one columbium specimen sealed with boron
nitride under a pressure of 3 x 10~ mm Hg in
quartz are presented in Tables 1.1.3 and 1.1.4,
Since the change in weight indicated a total im-
purity pickup of only 0.09%, it appears that the
analyses are somewhat incorrect, but they do indi-
cate the relatively higher pickup of oxygen than of
nitrogen. Eleven specimens were given similar
tests, and all the results were of the same order of
magnitude. The hot-pressed boron and silicon
nitrides apparently contained decomposable oxide
impurities.

In a second series of tests, columbium speci-
mens were heated in a purified nitrogen atmosphere
in a furnace at temperatures up to 1300°C for 1 hr.
No significant weight changes were observed. In
these tests nitrogen pressures from 0.5 to 3 x 10~3
were used. The specimens were bright after the
tests, but they reflected light in @ way that re-
vealed thin transparent films on their surfaces.
The effect of the pressure of the nitrogen on film
formation has not been studied completely, but it is

Table 1,1,3, Weight Change Data for Pure Columbium
Exposed to Boron Nitride in an Evacuated Quartz
Capsule for 72 hr at 1000°C

Columbium Boron

Specimen Nitride
Original weight (g) 2,8825 0.7263
Weight after test (g) 2,8853 0.7134
Weight change (g) +0.0028 ~0.0129

Table 1.1.4. Results of Impurity Analyses of Columbium
Before and After Exposure to Boron Nitride in an
Evacuated Quartz Capsule for 72 hr at 1000°C

Columbium Impurities (wt %)

Specimen 02 N2 H2
As-received 0.0061 0.0046 0.0005
After test 0.12 0.017 0.0046

PERIOD ENDING SEPTEMBER 30, 1958

believed to be small. There may be a critical pres-
sure below which absorption occurs without the
formation of the nitride film. None of these speci-
mens were analyzed because the weight-gain data
indicated no measurable absorption of any gas.
Furnaces and equipment capable of operating at
much higher temperatures are being fabricated so
that the range of study may be increased.

Oxygen-contaminated specimens were made by
introducing purified oxygen into a previously evac-
uated furnace containing columbium specimens.
The gas was introduced in such a manner as to
maintain a constant partial pressure of oxygen.
Various concentrations of oxygen were obtained by
leaving the specimens at temperature in the oxygen
atmosphere for various times. By this method,
specimens containing from 61 to 5000 ppm oxygen
were prepared by heating at 850°C under a partial
pressure of 2 x 10~ mm Hg O, for times ranging
up to 16 hr, These specimens are being tested to
determine the effects of oxygen contamination on
corrosion of columbium by lithium.

YTTRIUM AND YTTRIUM HYDRIDE
INVESTIGATIONS

W. J. Werner T. Hikido

Effect of Metal Purity on Hydriding
It has been reported by GE-ANPD that con-

siderable difficulty has been encountered with re-
spect to cracking of hydrided pieces of yttrium.
The cracks form both during hydriding and upon
thermal cycling of the hydrided piece. There has
been some evidence that the cracking is associated
with nonmetallic inclusions which may be observed
in the microstructure of the metal, and it is hoped
that improvement of the purity of the yttrium will
alleviate the cracking problem. Conclusive proof
of the effect of impurities is still unavailable, how-
ever, because of the lack of high-purity metal. At
the present time, studies are under way on the zone
refining of yttrium by the floating-zone procedure
developed by Kneip and Betterton 2 for the zone
refining of zirconium, Work has also been started
on the evaluation of various procedures for the
purification of hydrogen. With pure metal and pure

2G. D. Kneip, Jr,,and J, O, Betterton, Jr., J. Electro-
chem, Soc. 103, 684 (1955).

ANP PROJECT PROGRESS REPORT

hydrogen, it will be possible to compare the prop-
erties of a high-purity hydride with the properties
of hydrides having various impurity contents,

Dissociation Pressure Studies

No accurate set of dissociation pressure curves
for the YH system is, as yet, available. A dis-
sociation pressure rig is therefore being constructed,
and measurements will be made of as pure a hy-
dride as is obtainable. The effects of impurities
on the dissociation pressure will also be determined.

Deformation Mechanisms of Yttrium

The growth of single crystals of yttrium is being
attempted in connection with a program to study the
deformation mechanisms of high-purity yttrium. The
first method to be tried for the growth of these
single crystals is that of growing the crystals from
the melt (zone refining). A strain-anneal technique
will also be tried.

Electron-Beam Melting of Yttrium and
Yttrium=Magnesium Alloy

W. J. Werner T. Hikido

The Temescal Metallurgical Corporation, Rich-
mond, California, has developed a technique for
melting metals under a high vacuum in which the
energy of a focused electron beam is used. The
furnace consists essentially of a vacuum pumping
system, an electron **qun’’ and power supply, a
melting-stock feeding device, and a water-cooled
copper crucible with an ingot extraction mechanism.
The electron gun has a circular tungsten filament
electron source, focusing devices, and a power
source which supplies a high-potential electric
field through which the electrons are accelerated,
The electron gun is designed to concentrate part of
the electron beam on the end of the melting stock
as it is lowered into the furnace and the balance
on the surface of melted metal ingot in the crucible.
The metal melted from the stock drips into the
crucible where a molten pool is maintained and the
ingot is continuously extracted. The major ad-
vantages of the Temescal process are (1) the high
vacuum (104 mm Hg) maintained during melting and
(2) the use of water-cooled copper to minimize
crucible contamination,

The electron-beam melting technique was used to
investigate the possibility of purifying yttrium metal
with respect to O, N,, Li, Mg, and C content, and
the possibility of consolidating yttrium-magnesium

alloy directly into a pure metal ingot without going
through the vacuum-distillation step, Calculations
by A. R. Miller® showed that it should be thermo-
dynamically possible to purify yttrium by melting it
under a high vacuum. The consolidation of the
yttrium-magnesium alloy was also thought to be
possible if the melting were done by using a slow
feed rate and a low power input to the electron gun.

For the experimental investigation, one bar of
yttrium metal made up of six finger castings Heliarc
welded together and two bars of yttrium-magnesium
alloy ~ 1Y% in. in diameter and 10 in. long were sent
to the Temescal Metallurgical Corporation. The bar
of yttrium metal was double melted. The first meit
was accomplished at a feed rate of ~ ]/3 Ib of metal
per hour with maximum power input to the gun.

This melt was characterized by a heavy purple
plasma which surrounded the electron gun, metal
stock, and the melt, The density of this plasma,
which was composed chiefly of lithium and mag-
nesium, controlled the melting rate for a given power
input. At too fast a melting rate, the plasma be-
came dense enough to short out the filament to the
focusing device. The plasma was considerably less
dense during the second melt, and a faster melting
rate, was, therefore, possible.

The original analyses of the six finger castings
and the analysis of samples taken from the bottom
of the first melt and from the top, middle, and
bottom of the second melt are presented in Table
1.1.5. The results indicate that the electron-beam
melting did not purify the metal with respect to
oxygen and nitrogen and that the melting may have
actually resulted in some additional contamination.
There did seem to be some degree of purification
between the first and second melts, however, and
the method should be very effective in removing
magnesium, lithium, and any volatile fluorides.
Therefore, consideration will be given to further
experiments with electron~beam melting when larger
quantities of metal become available., Even if re-
moval of the oxygen and nitrogen does not prove to
be possible, it will be of considerable interest to
minimize the other impurities in order to determine
the true effect of the gases. The fluorides and
oxyfluorides are currently suspected of contributing
heavily to the difficulties encountered in working
yttrium metal and with cracking in hydrided yttrium.

3A. R. Miller, A Thermodynamic Study of Metals and
Ceramics, report prepared for Stauffer Chemical Company,

PERIOD ENDING SEPTEMBER 30, 1958

Table 1.1.5. Chemical Analyses of Yttrium Stock and Melts Prepared by Electron«Beam
Melting Techniques at the Temescal Metallurgical Corporation

Impurities (ppm)

Casting No.
0, N, Mg Li C

Yttrium Stock

L.9B-1 1300 1100 27 50 66

L-11A-3 990 48 25 16

L-T1A-4 970 30 23 16

L-10A-2 980 170 29 410

L.10A.7 980 100 100 <10 140

L-10A-8 850 110 120 <10 150
Melt

T-1B (bottom of first melt) 1600 330 <10

T-2T (top of second melt) 1500 210 <10

T-2B (bottom of second melt) 1300 170 <10

T+2M (middle of second melt) 1100 220 <10

Attempts to process the 80 wt % Y~20 wt % Mg
alloy directly to the metal ingot by the electron-
beam technique were unsuccessful. It was thought
that if a low melting rate were maintained and the
power input ke pt low, the magnesium could be
driven off and the metal melted. However, when the
beam of electrons was foacused on the-alloys the++
magnesium boiled out violently with considerable
splattering. Some of the magnesium deposited on
the filament and focusing shield and caused short
circuiting of the filament. It was not possible to
control the splattering and shorting, and therefore
the melting attempts had to be abandoned.

Expansion of YHrium Preparation Process
T. Hikido
In March 1958, the Oak Ridge National Laboratery

was requested to consider scaling up the yttrium
preparation process, which was described in the
previous report.* Official approval of the proposed
ORNL program was received at the end of April,
and the preparation of facilities was begun. The

objective set forth for the program was the prep-
aration of ytirium metal containing less than 1000-
ppm O, in 20-Ib ingots.

The equipment for this program, which is installed
in Building 9201-2, Y-12 area, consists of two
parallel 50-1b fluoride purification systems that
supply YF, mixtures to a National Research Corp-
oration 150-lb (rated in steel) vacuum-induction-
melting furnace. A system for transferring into the
furnace the lithium used as the reductant is included
in the installation.

The control panel and furnaces for one of the
fluoride purification systems may be seen in Fig.
1.1.4 beside the vacuum-reduction furnace, which
is in the center of the photograph. The lithium and
fluoride loading ports on the side of the vacuum
furnace may be seen. Another view of the reduction
furnace which shows the control panel for the 100-kw
motor-generator reduction power supply is shown in

Fig. 1.1.5,

47, Hikido and W. J, Werner, ANP Quar, Prog. Rep.
March 31, 1958, ORNL-2517, p 4.

ANP PROJECT PROGRESS REPORT

Fig. 1.1.4. Yttrium Preparation Facility. ekttt

10

PERIOD ENDING SEPTEMBER 30, 1958

Fig. 1.1.5. Yttrium Preparation Facility and Control Panel for Power Supply. (Queesiemsasaiaiesstie
-

ANP PROJECT PROGRESS REPORT

1.2. CORROSION
E. E. Hoffman

LITHIUM PURIFICATION
E. E. Hoffman

Various methods have been investigated for the
purification of lithium.! Although the effect of
nitrogen and oxygen contamination of the lithium on
the corrosion resistance of columbium to lithium at
elevated temperatures is not yet well understood, it
is generally felt that oxygen and nitrogen concen-
trations should be reduced to a practical minimum to
eliminate the purity of the lithium as a test vari-
able. A method of reducing the oxygen contami-
nation is also needed in order to produce the low-
oxygen-content lithium required in the production of
low-oxygen-content yttrium (see Chap. 1.1, this
report).

Analyses of lithium from three shipments supplied
by the vendor in gas-tight stainless steel containers
(under an inert gas) are presented in Table 1.2.1.
The specification of less than 100 ppm nitrogen
was, as may be seen, not met by the vendor. Ef-
forts will be made to obtain improved as-received
material, but it is felt that it will probably be nec-
essary to purify all lithium used for tests.

Three different methods have been used for de-
termining the oxygen content of lithium. In order to
provide a comparison of the results obtained by the
three methods, dip samples were taken from approxi-
mately 25 different batches of lithium. These dip
sampdes weresthenssuwbmimtedalor analysis by each

TE. E. Hoffman, W. H. Cook, and D. H. Jansen, ANP
Quat, Prog, Rep. March 31, 1958, ORNL-2517, p 9.

Table 1.2.1. Nitrogen and Oxygen Analyses of Recent

Lithium Shipments

Impurity Content

Shipment Quantity (ppm)*
No. {Ib)
Nitrogen Oxygen
1 20 605 390
2 46 1940 415
3 46 3310 1980

*Each value is an average of four analyses; sample
were taken for analysis at approximately 300°C.

12

method: (1) an activation analysis method de-
veloped in the ORNL Analytical Chemistry Di-
vision, (2) a butyl iodide-iodine method also de-
veloped in the ORNL Analytical Chemistry Di-
vision, and (3) a methyl alcohol-Karl Fisher reagent
method developed by the Nuclear Development Cor-
poration of America. Although the agreement of the
results obtained with the three methods was not
good, the best agreement was obtained between the
butyl iodide-iodine method and the activation
analysis method. Improved sampling and analytical
techniques will be employed in the future to further
determine the reliability of these two methods for
determining oxygen in lithium.

Results of preliminary determinations of the solu-
bility of lithium nitride and lithium oxide in lithium
were presented previously,] and one additional
oxide solubility determination has been made for
which an improved sampling technique was used. A
total of five analyses made by two methods by the
Analytical Chemistry Division on samples taken at
250°C showed good agreement. An averaged value
of approximately 1300-ppm oxygen was measured.
This value is considerably below that obtained in
one previous test, and additional work will be re-
quired to resolve the difference.

The purification experiments, described previ-
ously,! in which titanium and zirconium metals
were used as gettering agents for nitrogen and
oxygen in lithium, indicated that although fhesg
metals are effective in removing nitrogen, they do
not getter oxygen from the lithium. In several
tests the oxygen content of the titanium getter foil
actually decreased, as would be expected, inas-
much as lithium oxide is more stable than titanium
oxide. Yttrium oxide is much more stable from a
free energy of formation standpoint than either ti-
tanium or zirconium oxides, and, therefore, tests
are being performed to determine the effectiveness
of yttrium metal as a getter for oxygen in lithium.

A gettering test was conducted in which 88.5 g
of yttrium turnings was added to 628 g of lithium,
The stainless steel container, which had a tan-
talum liner, was held at 1500°F for 72 hr. Yttrium-
metal cylinders (0.3 in. in diameter and 2 in, in
fength} introduced into the purification vessel at
24-hr intervals were analyzed for oxygen and nitro-
gen at the conclusion of the test. The results of

PERIOD ENDING SEPTEMBER 30, 1958

the analyses, which were made by the vacuum- and further tests will be conducted to check the

fusion method, are presented in Table 1.2.2. usefulness of yttrium for this application.
Hardness measurements were made on the yttrium

metal specimens tested in lithium and on similar COLUMBIUM EXPOSED TO LITHIUM

specimens given the same heat treatment in an E. E. Hoffman

evacuated capsule. The purpose of these measure-

ments was to determine the extent of gettering of Arc-Cast Columbium Tubing Tested in Static

oxygen and nitrogen from the lithium by the yttrium, Lithium
Pickup of these elements results in increases in Static corrosion tests of the various grades of
bardness in yttrium. The hardness increases which  columbium stock received from vendors are being
occurred in one of the yttrium cylinders are given performed. Specimens are exposed to lithium for
in Table 1.2.3 as a function of depth from the sur- 100 hr at 1500°F and at 1800°F in an effort to
face. It may be seen that an appreciable hardness correlate metal history and analysis with any
increase occurred to depths of 65 to 105 mils. observed corrosion. Only a few samples have been
Specimens No. 2 (Y-6) and No, 3 (Y-8) showed tested thus far, and there are, as yet, no clear cut
similar hardness increases near the surface but to quantitative indications of the effects of the nitro-
lesser depths, as would be anticipated, since they gen and oxygen content of the metal on corrosion.
were exposed to the lithium for shorter times. The tests have shown, however, that the presence
Unfortunately, these promising results were not of oxygen is detrimental to corrosion resistance,
accompanied by reductions of the oxygen content In general, specimens of welded tubing have been
of the lithium, Lithium samples taken at various heavily attacked in the weld and heat-affected
time intervals showed a continuing increase in zones, and there has been little or no attack of the
nitrogen and oxygen as the test run proceeded. |t base material,
is felt, however, that the yttrium metal analyses The effect of surface contamination introduced
are less subject to error than the lithium analyses; during processing is itlustrated in Fig. 1.2.1, which

Table 1.2,2, Analyses of Yttrium Metal Getter Specimens Before and After Exposure to Lithium at 1500°F

Specimen Time Period of Exposure Oxygen Content (ppm) Nitrogen Content (ppm)

No. (hr) Before Test After Test Change Before Test After Test Change
1(Y-3) 0-72 1300 3800 +2500 450 1100 + 650
2 (Y-6) 24-72 1400 3900 +2500 570 850 + 280
3(Y-8) 48-72 240 1800 +860 110 170 +60

x . . ]

Table 1.2,3. Diamond Pyramid Hardness Values for an Yttrium Cylinder Specimen (Y-3) Used in a LithiumYttrium
Gettering Experiment as a Function of Distance from Surface Exposed to Lithium Compared with Values

for a Specimen Given a Similar Heat Treatment in Vacuum

Distance from Surface (mils)

Specimen Treatment 5 (Edge) 25 65 105 145 (Center)

Diamond Pyramid Hardness (500-g load)

Exposed to lithium for 72 hr ot 1500°F 117 95 71 66 59

Heat treated as above but in vacuum 59 60 58 * 59 59

rather than in fithium

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED] OUTER

SURFACE
OF TUBE
WALL
£
©
=3
s}
S

UNCLASSIFIED
Y-26263

0.015 in.

~=—— INNER
SURFACE
OF TUBE

WALL

Fig. 1.2.1. Outer and Inner Surfaces of Tube Wall of DuPont Arc-Cast Columbium Specimen Following Exposure
to Static Lithium for 100 hr ot 1500°F. Both sides were exposed to lithium. Note attack on outer surface. Etchant:

HF-HN03~HZSO‘-H20 (parts: 25-10-10-55). 250X. ((umeeemmming )

shows the outer and inner surfaces of a specimen Table 1.2.4, Diamond Pyramid Hardness of DuPont
of DuPont arc-cast columbium tubing that was ex- Arc-Cast Columbium Tubing Before and After Exposure
posed to lithium for 100 hr at 1500°F. Since both to Lithium for 100 hr at 1500°F

surfaces were exposed to a common lithium environ- S A
ment, the lithium purity was eliminated as a vari- oloes BNy low

able. It appears that some contamination was intro- Near Outer  Middle of  Near Inner
duced during tubing fabrication and was diffused Surface Wall Surface
into the outer surface during a subsequent anneal.

Hardness measurements also indicated the possi- Before test 145 121 133
bility of outer surface contamination. The Diamond Aot et 109 % 18
Pyramid hardness values for this arc-cast tubing

before and after the test are given in Table 1.2.4. *Values are averages for four determinations.

The decrease in hardness may have been due to
leaching of impurities by the lithium,

Effect of Welding Procedures on Corrosion

A series of tests was conducted to study the ef-
fect of various Heliarc welding procedures on the
corrosion of arc-cast columbium in lithium. Two
joining techniques and three types of welding en-
vironment were used in weld sample preparation.
The arc-cast columbium stock used was 0,035 in.
in thickness, and it contained, as determined by
chemical analysis, 360-ppm 0,, 20-ppm N,,, and
140-ppm C.

The joints were made by using a fusion welding
technique in which no filler metal is applied or
with a filler metal consisting of an 85% Zr-15% Cb
alloy. Sheet material for the fusion welds was ob-
tained by splitting and flattening 0.5-in.-OD seam-
less tubing. For the welds made with filler metal,
the split tubing was not flattened. Three joints of
each type were made that were prepared in three
different environments. One was made in air with
good, conventional inert-gas (argon) coverage; one
was made in an inert-atmosphere chamber containing
argon gas contaminated with 1600 ppm N, + CO and
400 ppm 02; and the third was made in an inert-
atmosphere chamber containing high-purity argen,
which analyzed 67 ppm N, + COand 3 ppm 0,. All
the specimens were exposed to lithium simultane-
ously in a single columbium container for 100 hr at
1500°F.

The results of these tests are presented in Table
1.2.5. The specimens joined in argon plus air
showed weight losses and those joined in the inert
atmosphere chamber showed weight gains, The
specimens joined with the filler metal showed es-
sentially no attack of the weld metal. The effects
of environment on the heat-affected zone and the
weld metal are illustrated in Fig. 1.2.2. As may be
seen, the heat-affected zone of the columbium ad-
jacent to the braze weld of the specimen joined in
a contaminated environment (air plus argon) was
heavily attacked, but the same zone showed very
little attack when the specimen was joined in a
high-purity argon environment. The base material
was attacked in both cases but to a lesser degree
in the case of the specimen joined in high-purity
argon,

2'

PERIOD ENDING SEPTEMBER 30, 1958

The specimens joined by fusion welding were
heavily attacked in both the weld zone and the
heat-affected zone. The attack in these areas de-
creased in both depth and concentration as the
purity of the cover gas increased. The weld zones
are shown in Fig. 1.2.3 after exposure to lithium at
1500°F for 100 hr, The attack was extremely heavy
on the specimen welded in argon plus air. It is
surprising, however, that even with the purest weld-
ing environment the specimen was subsequently at-
tacked by lithium along the grain boundaries. The
introduction of impurities during welding should
have been quite low. The results indicate that
unalloyed columbium must not only be welded under
very pure, inert atmospheres, but that the parent
metal must be quite low in nitrogen and oxygen.

An enlarged photograph of a crack in the weld zone
of the specimen welded in argon plus air is shown
in Fig, 1.2.4. The grain-boundary phase shown in
Fig. 1.2.4 has not yet been identified; it was not
visible in the as-welded specimens prior to the test
in lithium.

As was expected, the hardness values for speci-
mens in the as-welded condition were guite high
when the joining operation was performed in impure
argon because of pickup of interstitial elements
such as oxygen and nitrogen. All test specimens
decreased in hardness as a result of exposure to
lithium, and thus it appears that the lithium gettered
the contaminants responsible for the hardness in-
crease,

The free energies of formation indicate that co-
lumbium nitride is more stable than lithium nitride ?
whereas, lithium oxide is considerably more stable
than columbium oxide. Therefore, lithium metal
would probably reduce any columbium oxide present
as a precipitate in the grain boundaries of colum-
bium welds, but it should not attack columbium
nitride. Specimens selectively contaminated with
nitrogen and oxygen are being tested that were
prepared as described in Chap. 1.1 in order to
identify the contaminant responsible for the cor-
rosion of columbium welds,

2E, E, Hoffman and W. D. Manly, Lithium as a Nuclear
Reactor Coolant, ORNL-2538 (August 8, 1958), p 31.

ANP PROJECT PROGRESS REPORT

Table 1,2,5. Results of Exposure of Columbium Weld Samples to Lithium for 100 hr at 1500° F

Diamond Pyramid

Specimen Hardness
Joining Welding Weight Results of Metallographic (500-0z load)
Technique Environment Change Examination of Weld Sample
. Before After
(mg/in.
Test Test
Braze Welding Argon + air -0.7 Braze weld: 1 mil of attack 361 343
(85% Zr=15%
Cb Filler alloy
Heat-affected zone: complete grain- 360 94
boundary penetration of specimen
Base material*: 4 mils of subsurface 129 112
voids
Argon +1.0 Braze weld: no attack 285 246
+1600-ppm N2 Heat-affected zone: 15 mils of attack 114 98
+400-ppm O2
Base material: 4 mils of subsurface 120 104
voids
Argon +0.7 Braze weld: no attack 289 242
+67-ppm N2 Heat-affected zone: 6 mils of attack 122 98
+3-ppm O2
Base material: 3 mils of subsurface 113 105
voids
Fusion Welding Argon + air -0.4 Weld: complete and very heavy attack 356 128
of entire weld
Heat-affected zone: 15 mils of inter- 127 90
granular attack
Base material: 3 mils of attack 127 97
Argon +0.6 Weld: complete attack of grain bound- 155 105
+1600- N aries; attack not as concentrated as
ppm Ny in fusion weld described above
+400-ppm 02
Heat-affected zone: 9 mils of inter- 116 93
granular attock
Base material: 3 mils of attack 128 107
Argon +0.2 Weld: complete attack of grain bound- 131 95
4 67-ppm N aries in a few scattered areas; less
PP 2 attack than in either of the fusion
+3-ppm 02 welds described above
Heat-affected zone: 7 mils of inter- 118 92
granular attack
Base material: 1 mil of attack 127 93

*The as-received arc-cast columbium contained 360-ppm 02, 20-ppm N2, and 140-ppm C, and it had a Diamond

Pyramid hardness of 137.

16
PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
Y-26784

WELD AREAS

85% 2r =15 % Cb

FILLER METAL ‘\‘

0500 -in.-0D,
0.035-in.—
WALL TUBING

\COLUMBIUM

BRAZE WELDED
IN ARGON +AIR

BRAZE WELDED IN
ARGON + 67ppm Ny + 3ppm O,

HEAT-AFFECTED AREAS

Fig. 1.2.2. Effect of Welding Environment on the Corrosion Resistance of Columbium Braze Welded with an
85% Zr—15% Cb Alloy and Exposed to Lithium at 1500°F for 100 hr. Unetched. (Swmmgt with caption)

UNCLASSIFIED
(@) (6) (e) Vezeris

0.03 in.

Fig. 1.2.3. Effect of Welding Environment on the Corrosion Resistance of Columbium Fusion Welded and Ex-
posed to Li

ium at 1500°F for 100 hr. (a) Weld zone of specimen welded in argon plus air. (b) Weld zone of speci-
men welded in argon plus 1600-ppm N, plus 400-ppm O,. (c) Weld zone of specimen welded in argon plus 67-ppm N,
plus 3-ppm O,. Etchant: HF-HNO-H,S0,-H,0. Wwmmpt with caption)

17

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED |

Y-26663

0.010in.

Fig 1.2.4. Enlorgement of Crack Shown in Fig. 1.2.3a

COMPARISON OF METHODS FOR DETERMINING
THE NITROGEN CONTENT OF COLUMBIUM

E. E. Hoffman

The validity of the vacuum-fusion method for de-
termining the nitrogen content of columbium has
been evaluated by checking ¥he results against
those obtained with the micro-Kjeldahl method,
Several columbium samples were analyzed by the
Special Analysis Laboratory of the ORNL Ana-
lytical Chemistry Division by both methods. The
results obtained by the two methods were in agree-
ment, as may be seen in Table 1.2.6.

The micro-Kjeldah| method consists of dissolving
a sample of approximately 0.1 g in concentrated
H,S0,, transferring the solution to the Kjeldahl
upparmus, adding NaOH, and distilling the ammonia.
The distilled ammonia was analyzed colorimetri-
cally with the use of Nessler's reagent.

In the vacuum-fusion method, a platinum bath in
a graphite crucible is used for sample dissolution.
The technique used was standard except that pieces
of platinum weighingatotal of about 10g are dropped
into the bath after the samples have been added in
order to thoroughly mix the bath material. In addi-
tion, the columbium metal samples are wrapped in
platinum foil so that when dropped they will sink

18

Table 1.2.6.
of Columbium by Vacuum-Fusion and the
Micro-Kieldahl Methods

Determination of the Nitrogen Content

Nitrogen Content (ppm)

Columbium
Samplo Code  BY Vacuum-Fusion By Micro-Kjeldahl
Method Method
Cb Wire 35 35
No. 11
Cb Wire 3500 3500
No. 12
3400
Cb Wire 34 33
No. 13
WC-1A 160 150
160 160
160
130

deep into the bath and not float on its surface. The
evolved gases are passed over hot copper oxide,
and the CO, which forms is taken out in a cold
trap. The H is converted to water and removed in
a magnesium peychlora'e absorption tube. The re-
maining gas is N,, which is defermined in a McLeod

gage.

MOLYBDENUM AND COLUMBIUM TESTED IN
LITHIUM IN SEESAW-FURNACE
APPARATUS

E. E. Hoffman

A series of dynamic corrosion tests of molybdenum
in contact with lithium were conducted with the use
of a seesaw-furnace apparatus. The test assembly
is shown in Fig. 1.2.5. In the tests, a 15-in.-long
molybdenum pipe (0.80 in. OD x 0.10 in, wall) was
partially filled with lithium, and the molybdenum
pipe was placed inside an Inconel pipe (1.31 in.
OD x 0.07 in. wall). The annular space between
the molybdenum and the Inconel was partially filled
with sodium, The Inconel pipe protected the mo-
lybdenum from oxidation, while the sodium acted as
the heat-transfer bond between the two containers.
The test assembly was then placed in the seesaw
funace, where it was rocked up and down (% cpm)
with the desired temperature gradient between the
ends of the capsule. Hot-zone and cold-zone
-~
UNCLASSIFIED
ORNL— LR-—-DWG 30678

SPECIMEN
1.0x0.60x0.065 in.

INCONEL TC WELL

Fig. 1.2.5. Seesow-Furnace Test Assembly for Evalu-

ating the Corrosion Resistance of Molybdenum in Lithium.

Tflble ]n207.

PERIOD ENDING SEPTEMBER 30, 1958

specimens were suspended in the ends of the mo-
lybdenum tube, so that weight changes and the
attack in each zone could be determined.

The extruded molybdenum pipes used for the six
tests conducted thus far were made from unalloyed
vacuum-arc-cast material, The specimens sus-
pended in the hot and cold zones were made from
arc-cast molybdenum-titanium alloy sheet, The
unalloyed stock was found by analysis to contain
0.02 wt % C, 0.0012 wt % 0,, and 0,0005 wt % N,,.
The alloy stock contained, in addition to the 0.4§
wt % titanivm, 0.02 wt % C, 0.0031 wt % G,, and
0.0005 wt % N,. A summary of the results obtained
to date in the experiments is given in Table 1,2.7,
The results show that molybdenum has very good
resistance to lithium at quite high temperatures.
The results of test S5-514 were not in agreement,
however, with those obtained in the other five
tests. The cold-zone specimen showed a weight
gain and metal-crystal deposition. Spectrographic

Results of Seesaw-Furnace Tests of Molybdenum Exposed to Lithium

T Specimen
emperature .
o Duration - : Weight Change
("F) Lithium Analysis .2 )
Test No. of Test (mg/in.) Results of Metallographic
Hoet  Cold (hr) N2 (ppm) 02 {ppm) Hot Cold Examination of Specimens
Zone Zone
Zone Zone
§$S-514* 1700 1300 500 284 210 —25.8 +9.0 Hot zone: surface pits to a depth of
PRSCE 7L R 0.4 mil
Cold zone: metal crystal deposition to
a depth of 0.3 mil
55-525* 1700 1000 500 4770 1450 -0.1 +0.4 Not yet examined
55-523** 1800 1000 100 284 910 —-0.3 —0.7 Hot zone: surface roughened slightly,
less than 0,2 mil of attack
Cold zone: no attack or deposition
$S-5246** 1800 1050 500 4770 1450 ~0.1 +2.2 Not yet examined
§S-522** 1900 1000 100 284 910 —0.4 —0.8 Hot zone: surface roughened slightly,
less than 0.2 mil of attack
Cold zone: no attack
§5-527** 1900 1100 150 4770 1450 -0,3 +0.4 Not yet examined

*The tested specimens were hand polished for metallographic examination; the polishing removed ™~ 0,25 mil of

surface.

**The tested specimens were surface ground for metallegraphic examination; the grinding removed ™~5 mils of
p 9 R Lo g p g g

surface,

19

ANP PROJECT PROGRESS REPORT

analysis of the specimen surface did not reveal any
differences between as-tested and before-test speci-
mens. Test $5-525 was a recheck of test $5-514,
and the weight-change data were in good agreement
with results obtained in the tests at higher temper-
atures. At the present time the anomalous results
of test 55-514 are not understood, and further work
is planned in an effort to explain the previous re-
sults. The ends of the molybdenum capsule and
the hot- and cold-zone specimens from test $5-526
are shown in Fig. 1.2.6, and an enlargement of the
hot zone of the molybdenum capsule is shown in
Fig. 1.2.7. No molybdenum capsule failures have
occurred to date. Test $5-527 was terminated by a
sodium leak through the Inconel protective capsule
wall, and, although the molybdenum capsule was
attacked in the region of the leak, it did not fail.
That the molybdenum capsule did not fail may have
been due to gettering of the available oxygen by
the sodium during the time required for the furnace
to cool. Hot- and cold-zone specimens from test
55-522 are shown in Fig. 1.2.8. Very little, if any,
attack may be detected even at very high magnifi-
cation. The hot-zone specimen partially recrystal-
lized during the test. The cold-zone specimen sur-
face shown in Fig. 1.2.8 is quite similar to that of
the metal before test.

In all these seesaw-furnace tests, some dis-
similar-metal mass transfer from the Inconel pipe
wall to the outer surface of the molybdenum
capsule walls in the hot zone was observed. Mass
transferred crystals were also found on the cold-
zone wall of the Inconel pipe exposed to sodium.
Thermal-convection loop tests of molybdenum-
lithium systems are planned in order to check the
corrosion resistance of molybdenum in a system
with unidirectional flow.

A series of columbium-lithium seesaw-furnace
tests is under way for which a test configuration
similar to that shown in Fig. 1.2.5 is being used.
In the columbium tests, the outer protective cap-
sule is type 316 stainless steel rather than Inconel
as in the molybdenum tests. The columbium tubes
are 15 in. in length, 1.25 in. in outside diameter,
and 0.032 in. in wall thickness. The only test per-
formed to date was terminated after 300 hr by a
sodium leak in the stainless steel pipe. The co-
lumbium tube did not fail as a result of the failure
in the protective pipe. The hot- and cold-zone
specimens were at 1600 and 1100°F, respectively.

20

The hot-zone specimen showed a slight weight
gain (0.1 mg/in.2), which may be attributed to
gettering of residual nitrogen from the lithium,
while the cold-zone specimen showed no weight
change. Further tests are planned at higher tem-
peratures and for longer test periods.

=

HOT ZONE, 1800°F

WEIGHT CHANGE,
—-0.4mg/in2

weLassiEs
Vi

COLD ZONE, 1050°F

WEIGHT CHANGE,
+2.2 mg/in2

pialiiin
o 0.5 10
INCHES

Fig. 1.2.6. Hot- and Cold-Zone Molybdenum Capsule
Sections ond Specimens from Test 55-526. Lithium cir-
culated inside the seesaw capsule for 500 hr under the
temperature conditions indicated. Sodium flowed over the
outside of the molybdenum capsule (see Fig. 1.2.5).
(@ww————! with caption)

wepssnes

——oesin——f

Fig. 1 Enlargement of Hot-Zone End of Molyb-
denum Capsule From Test $5-526. Welding operations
were performed in an inert-atmosphere chamber. This
section of the test capsule was at 1800°F for 500 hr.

(Swmtlm—! vith caption)

UNCLASSIFIED
Y+26675

INCHES

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
Y-26676

“.'

(@) HOT-ZONE, 1900°F

1500

st

1000°F

Fig. 1.2.8. Molybdenum Specimens from Seesaw-Furnace Test §5-522. Etchant: 50% NH,OH-50% H,0,.
Specimens nickel #Y8F8® to preserve edges during metallographic polishing. (UMl with caption)

BERYLLIUM EXPOSED TO LITHIUM
E. E. Hoffman

In previous corrosion tests? a beryllium speci-
men in an iron capsule exposed to lithium at 1830°F
for 400 hr showed extensive attack and large weight
losses. Tests have recently been conducted in
iron capsules and in all beryllium systems, and the
results are compared in Table 1.2,8. In the recent
tests, as in the earlier test, the specimen tested in
an iron capsule ot 1830°F suffered extensive solu-
tion attack as a result of dissimilar metal mass
transfer of beryllium from the specimen to the cap-
sule wall. The beryllium in an all beryllium system
showed very good corrosion resistance to static
lithium at elevated temperatures.

The serious effect of dissimilar metal mass trans-
fer on the corrosion resistance of the beryllium is

3A. deS. Brasunas, Interim Report on Static Liquid-
Metal Corrosion, ORNL-1647 (May 11, 1954).

o

illustrated in Fig. 1.2.9. The rectangular shape of
the beryllium specimen was altered appreciably in
the test conducted in an iron capsule. The slight
weight gains of the beryllium specimens fested in
beryllium containers are probably due to gettering
of impurities such as oxygen and nitrogen from the
lithium. The surfaces of the beryllium specimens
tested recently at 1830°F are shown in Fig. 1.2.10.
Considerably more porosity and surface roughening
is apparent on the specimen tested in an iron con-
tainer than on the specimen tested in a beryllium
container.

Some indication of the extent of dissimilar metal
mass transfer is illustrated in Fig. 1.2.11. The
Be,Fe intermetallic compound on the surface was
identified by x-ray analysis. This phase has a
hexagonal structure, and its anisotropy is apparent
in the photomicrograph taken with polarized light
(Fig. 1.2.118).

21
ANP PROJECT PROGRESS REPORT

Table 1.2.8, Corrosion of Beryllium Specimens by Lithium in Static Tests
Attack (mils)

Temperature Time o Specimen Weight Change -

(°F) () Commotun (mg/em?) Intergranular Solution

1830% 400 ron ~101.8 127 30 P

1500 100 Iron -3.9 4 0

1500 100 Beryllium +0.3 2 0

1830 100 Iron —66.1 5 2

1830 100 Beryllium +03 3 [)
*Previous test, ref 3.

UNCLASSIFIED
Y-26285
[¢] 0.5 10
INCHES

BEFORE TEST TESTED IN Fe CAPSULE TESTED IN Be CAPSULE
WEIGHT CHANGE, WEIGHT CHANGE, p
—66.4 mg /cm? +0.3 mg/cm?

Fig. 1.2.9. Effect of the,Expasure.of Beryllium to Static Lithium for 100 hr ar 1830°F. (q@g@mwith caption)

UNCLASSIFIED UNCLASSIFIED
Y-26413 Y-26409

Fig. 1.2.10. Surfaces of Beryllium Specimens Shown in Fig. 1.2.9 Which Had Been Exposed to Static Lithium for e
100 hr ot 1830°F. (a) Tested in an iron capsule. (b) Tested in a beryllium capsule. As-polished. 500X. Wlwmme
with caption)

22

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
Y-26416

Fig. 1.2.11. Wall of Armco Iron Capsule in Which a Beryllium Specimen Was Tested in Contact with Static
Lithium for 100 hr at 1830°F. Note dissimilar metal mass transfer of beryllium to capsule wall. As-polished. (a)

Bright*fi@ldeM|umination. (b) Polarized light illumination.

REFRACTORY-METAL-BASE BRAZING ALLOYS
EXPOSED TO LITHIUM

D. H. Jansen

Most commercially available brazing alloys
contain constituents used to alter the melting point,
lower the flow point, increase flowability, or im-
prove the ductility of the alloy. These constituents
include the precious metals, copper, and manganese,
to mention a few. The additives are beneficial for
their specific purposes, but they may produce alloys
that possess limited corrosion resistance to the
liquid metals, especially lithium. A typical ex-
ample, a 60% Mn—40% Ni brazing alloy which failed
completely when exposed to lithium for 100 hr at
1500°F, is shown in Fig. 1.2.12. Analyses of
microdrillings from the remaining alloy fillet indi-
cate that the manganese was attacked to the same
extent as the nickel. Another example, an 86%
Fe-5% Si—5% Cu-4% B brazing alloy on a type 347
stainless steel tube-to-header joint which was
tested in lithium in a seesaw-furnace apparatus for
100 hr at 1500°F, is shown in Fig. 1.2.13. The
copper concentration of the lithium bath after this
test was very high.

As a result of the poor corrosion resistance of
the alloys described above, refractory-metal-base

®wmy with caption)

brazing alloys are being investigated for use in
high-temperature lithium systems, Columbium
tubing (‘/2 in. OD) has been used for capsules in
tests in which small (5 g) brazing alloy samples
(buttons) have been corrosion tested in static
lithium at 1700°F, Conventional methods for
loading and testing the brazing alloys were used.

The corrosion testing program on these refractory-
metal-base alloys has involved, initially, static
tests of zirconium- and titanium-base binary and
ternary alloy buttons. The alloys which appear
promising with respect to corrosion resistance will
be used as a basis for developing other alloys with
lower flow points or better flow characteristics.
(Melting-point and flow-point determinations and
flow-characteristic studies are being made by the
Welding and Brazing Group, see Chap. 1.3.) The
more promising alloys will then be used for brazing
molybdenum or columbium T-joints which will be
given more severe corrosion tests (seesaw-furnace
tests or tests in thermal-convection loops with
brazed sections in the hot legs).

The results for all the refractory-metal-base
alloys corrosion tested in lithium to date are
presented in Table 1.2.9. The alloy button con-
taining vanadium (70% Zr-30% V) exhibited the

23

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED 094
Y-7776 005

BASE MATERIAL

BASE MATERIAL Lo

Fig. 1.2.12. A 60% Mn—40% Ni Brazing Alloy on a Type 316 Stainless Steel T-Joint After Exposure to Static
Lithium for 100 hr at 1500°F. Complete penetrationgafethie brazing alloy is evident. Unetched. 150X. (iiiimes!

with caption)

EEE]

INCHES.
i

i

fuone
foor|
- Lo
fooef
@ foef
o |
§ Lo

[t

Fig. 1.2.13. (o) An 86% Fe-5% Cu—5% Si—4% B Brazing Alloy on Type 347 Stainless Steel Tube-to-Header Joint
After 100 hr of Exposure to Lithium in a Seesaw-Furnace Apparatus at a Hot-Zone Temperature of 1500°F. (b) En-
largement of Area Enclosed by Dotted Lines in (o). Unetched. 50X, 250X. Reduced 28%. Mmmiissmmb with

caption)

24
PERIOD ENDING SEPTEMBER 30, 1958

Toble 1.2.9. Results of Static Corrosion Tests on Refractory-Metal-Base Arc-Cast Brazing Alloy
Buttons Exposed to Lithium in Columbium Capsules at 1500 or 1700°F for 500 hr

Braze Material Test
Composition Temperature Weight Change* Metallegraphic Results and General
(wt %) (°F) (%) Condition of Button After Test

95 Zr-—-5 Be 1700 +0.005 Subsurface void formation to a depth of less than

84 Zr-16 Fe 1500 -0.573 No attack; numerous small cracks in specimen
-0.62

82 Zr.18 Cr 1700 -1.81 No attack observed

75 Zr-25 Cb 1700 -0.227 No attack cbserved
-~0.11

75 Zr—15 Cr=10 Fe 1500 —-2.5 Edge of specimen attacked nonuniformly toa depth
-1.7 of 1 mil

70 Zr-30 V 1700 —-10.87 Second phase leached to a depth of 9 mils
-2004

69 Zr-31 Mo 1700 -1.34 Neo attack observed
- 1030

80 Co-20 Cb 1700 -2.6 Specimen uniformly depleted to a depth of 0.5 mil;

spotty attack found that varied from 0 to 10 mils

55 Ti—45 Zr 1700 -0.063 No attack observed
+0.079

72 Ti~28 Co on Cb Spotty attack to less than 1 mil in depth

T-jcint PP

*Each weight change is the result of one test; the buttons weighed approximately 5 g.

most weight change during the test. The micro-
structure of the edge of the alloy before and after
the test is shown in Fig. 1.2.14, Both vanadium
and zirconium had previously exhibited good cor-
rosion resistance to lithium,**> and a microspark
traverse on the as-tested 70% Zr-30% V alloy
button showed essentially no alteration of compo-
sition from the edge into the interior of the alloy.
It appears therefore that the zirconium and va-
nadium may both have gone to the columbium cap-
sule by a process of dissimilar metal mass transfer,
This particular test is being repeated to determine
whether the large weight loss is due to mass trans-
fer.

4D. H. Jansen and E. E, Hoffman, ANP Quar. Prog.
Rep. Sept. 30, 1957, ORNL-2387, p 215.

SNuclear Pyppulsion Pyogram, Engineering Progress
Report, April 1, 1958—]June 30, 1958, PWAC.582,

The only alloys that showed severe corrosion at-
tack were the 75% Zr-15% Cr—10% Fe and the 80%
Co-20% Cb (Fig. 1.2.15) alloys. Attack on these
alloys was quite spotty and nonuniform, The in-
homogeneity of the alloy buttons, which may be the
cause of nonuniform attack, is being studied.

Previous static tests on pure cobalt in lithium
under slightly more severe conditions (1832°F for
447 hr) showed intergranular attack and a weight
loss of 7% (0.7 g/in.2). Consequently, the attack
seen on the 80% Co-20% Cb alloy (Fig. 1.2.15) is
understandable, in view of the high cobalt content
of the button.

Titanium-base brazing alloys in button form and
in the form of braze fillets on columbium T-joints
were tested and the results are being evaluated at
present, One titanium-base alloy (72% Ti-28% Co)
on a columbium T-joint tested in 1700°F lithium for
100 hr showed approximately 1 mil of nonuniform
attack on the exposed surface (Fig. 1.2.16),
ANP PROJECT PROGRESS REPORT

Fig. 1.2.14. Alloy Button of 70% Zr—30% V, () Before and (b) After a Static Corrosion Test in Lithium at
1700°F for 500 hr. Note depletion of the second phase fo o depth of 10 mils. 250X. (@mmissssswmb with caption)

2 Ni PLATE ; UNCLASSIFIED

EEE

T
INC‘!ES

T
f L

EEEEREER

Fig. 1.2.15. An 80% Co-20% Cb Alloy Button After Exposure to Static Lithium for 500 hr at 1700°F. The attack
shown at the exposed edge was nonuniform and varied from 0 to 10 mils. Unetched. 250X. (Gamimimmbind with

caption)

26

PERIOD ENDING SEPTEMBER 30, 1958

Fig. 1.2.16. A 72% Ti-28% Co Brazing Alloy () Before and (b) After Testing in Static Lithium at 1700°F for
100 hr. Dark areas at the braze and base material interface are microcracks. Corrosion of the exposed edge is

li
(dSmmiinimtionh vith caption)

No attack of the type observed on the 80% Co-20%
Cb alloy button was seen on this 28% Co alloy on
a T-joint. Attack may be seen at only a few, scat-
tered places along the surface of the fillet, The
results of this test, as compared with those for the
alloy button containing 80% Co, indicate that cor-
rosion resistance is decreased when large per-
centages (80%) of cobalt are used.

DETERMINATION OF THE SOLUTION RATE
OF METALS IN LITHIUM®

E. E. Hoffman

A method was developed at NDA for measuring
the rate at which container metals dissolve in
liquid lithium under controlled nonequilibrium con-
ditions.” The method consists essentially of im-
mersing a thin test specimen into a comparatively
large volume of liquid metal held at the desired
temperature in an inert container. With suitable
choices of surface area and immersion time, the

SSubcontract with Nuclear Development Corporation of
America, July 1, 1957 to July 31, 1958,

7B. Minushkin, Determination of the Solution Rate of
Metals in Lithium, NDA-44 (June 30, 1958).

ited fo intermittent attack to a minimum depth of 1 mil. Etchant: 46 H,0, 46 HNO,, 8 HF. 100X. Reduced 21%.

test specimen can be made to lose about 1 mg of
weight in a test period of less than one day. This
weight loss does not significantly increase the
concentration of test specimen material in the
liquid metal, but the loss can be determined with
ample accuracy and precision on a semimicro
balance.

The solution rates determined in this manner are
proportional to the specific solution rate constant,
. Based on experimental evidence from thermal-
convection loop tests, it is believed that inhibitors
and impurities exert their effect on mass transfer
rates by altering the specific solutionrate constant.
Therefore, this method is suitable for preliminary
studies of the effects of additives and impurities
on mass transfer.

The data from fests on stainless steel indicate
that variables can be detected which affect the
initial solution rate by as much as +20%. The re-
sults indicate that nickel-bearing stainless steel
dissolves in lithium at 1600°F at an initially high
rate of 3.0 mg/in.Zshr because of the preferential
leaching of nickel. Within a relatively few hours,
depletion of nickel in the specimen surfaces re-
sults in a rapid decrease in the solution rate to a

27

ANP PROJECT PROGRESS REPORT

value of about 0.3 mg/in.2+hr at solute concen-
trations of a few hundred parts per million. It was
shown that this solution rate accounts for the mass
transfer rates observed in thermal-convection loop
tests. The solution rate tests confirm the dele-
terious effects of nitrogen and oxygen and the
beneficial effects of aluminum additions on mass
transfer rates that were observed in thermal-con-

28

vection loop tests. In addition, it appears that
Misch metal and tantalum additions may have
beneficial effects.

Solution rate studies of molybdenum in lithium at
1600°F indicate that this material is not signifi-
cantly attacked by lithium. Tests on niobium
yielded conflicting results which have not been
resolved.

W o &
PERIOD ENDING SEPTEMBER 30, 1958

1.3. WELDING AND BRAZING

P. Patriarca

DEVELOPMENT OF BRAZING ALLOYS
FOR LITHIUM SERVICE

R. G. Gilliland G. M. Slaughter

Refractory-metal-base brazing alloys are being
developed for application in high-temperature re-
actors that use lithium as a coolant. As a starting
point, a literature survey was made of the binary
systems of columbium, zirconium, titanium,
tantalum, and molybdenum in order to find eutectics
or minimums in the approximate temperature range
of interest. The data obtained for columbium and
zirconium were presented previously, ! and data for
titanium, molybdenum, and tantalum are summarized
in Tables 1.3.1, 1.3.2, and 1.3.3. Small buttons of
approximately fifty binary or ternary alloys of in-
terest have been prepared for testing by arc melt-
ing, and results of tests of a number of these alloys
in lithium are presented in Chap. 1.2, Corrosion.

A vacuum-tube fumace that has been modified
for ultrashigh-temperature work is being used for
preparing brazed joint specimens for testing. The
furnace and the quartz muffle for use at temper-
atures up to 1200°C are shown in Fig. 1.3.1. A
ceramic muffle made primarily of zirconium oxide
is available for use at 1350°C. Vacuums of less
than 1 u are consistently obtained.

A number of the arc-melted alloys have been
tested for flowability at various temperatures by
using the T-joint sample design with both
columbium and molybdenum as base metals., Flow
points were obtained from these tests, and metal-
lographic examinations of the resulting joints were
conducted. The flow points and general brazing
characteristics of the several alloys tested thus far
are presented in Table 1.3.4.

The brittleness of the brazing alloy in the as-
brazed condition is an important problem in this
study, as may be seen in Fig. 1.3.2, which shows a
brazed molybdenum T-joint that contains numerous
fillet cracks. The brazing alloy used was 84 wt %
Zr and 16 wt % Fe. In comparison, the molybdenum
T-joint brazed with an alloy containing 70 wt % Zr
and 30 wt % V, shown in Fig. 1.3.3, is considered
to be satisfactory,

16, M. Slaughter, ANP Quar, Prog. Rep. March 31,
1958, ORNL-2517, p 35.

A study of the diffusion of brazing alloys on re-
fractory metals has been initiated. Lap joints
made from columbium, and possibly molybdenum,
brazed with the different altoys will be used in the
study, and particular attention will be paid to
brazing alloys containing refractory metals, The
joints will be aged just befow, at, and just above

Table 1,3.1. Titanium Binary Systems of Potential
Interest As Brazing Alloys

Eutectics or Minimums*

System Composition Melting Point
(wt %) (°F)
Ti 3035
Ti-Ag 96 Ag, eutectic 1650
Ti-Au 15 Au, eutectic 1530
82 Au, eutectic 2345
Ti-Be 3 Be, eutectic 1750
90 Be, eutectic 2300
Ti«Cr 47 Cr, minimum 2550
Ti-Co 28 Co, eutectic 1870
81 Co, eutectic 2075
Ti«Cu 50 Cu, eutectic 1740
70 Cu, eutectic 1600
Ti-Ge 19 Ge, eutectic 2570
Ti-Fe 31 Fe, eutectic 1985
85 Fe, eutectic 2415
Ti=Pb 26 Pb, eutectic 2250
TieMn 42 Mn, eutectic 2150
Ti=Ni 28 Ni, eutectic 1750
65 Ni, eutectic 2030
83.8 Ni, eutectic 2350
Ti=5i 9 Si, eutectic 2430
51 Si, eutectic 2700
78 Si, eutectic 2430
Ti-Y 30 V, minimum 2950
Ti=Zr 50 Zr, minimum 2900
Ti-Th 88 Th, eutectic 2200

*A. D. McQuillan and M. K. McQuillan, Titanium,
New York Academic Press, New York, 1956.

ANP PROJECT PROGRESS REPORT

Table 1.3,2. Molybdenum Binary Systems of Potential Table 13,3, Tantalum Binary Systems of Potential
Interest As Brazing Alloys Interest As Brazing Alloys
Eutectics or Minimums* Eutectics or Minimums*
System Composition Melting Point System Composition Melting Point
(wt %) °F (wt %) (°F)

Mo 4800 Ta 5425
Mo-Ni 47 Ni, eutectic 2400 To-Co 67.6 Co, eutectic 2325
Mo-Co 62 Co, eutectic 2444 Ta-Si 94 Si, eutectic 2525
Mo-Si 12 Si, eutectic 2600 To-Fe 80 Fe, eutectic 2570
Mo-Al 22 Al, eutectic 4200 TaeNi 62 Ni, eutectic 2480
Mo-B 1B, eutectic 3540 39 Ni, eutectic 2552
Ve D0y Anctlc 2624 *M. Hansen, Constitution of Binary Alloys, McGraws
Mo-Mn 99 Mn, eutectic 2191 Hill, New York, 1958.
Mo-Cb 53 Cb, eutectic 2400

*M. Hansen, Constitution of Binary Alloys, McGraw-
Hill, New York, 1958.

UNCLASSIFIED
PHOTO 44634

Fig. 1.3.1. Vacuum Tube Furnace and Quartz Muffle for Refractory-Metal Brazing Alloy Studies. The furnace is
mounted on a track fo permit rapid heating and cooling of the test sample.

30
PERIOD ENDING SEPTEMBER 30, 1958

Table 1,3.4. Data Obtained in Refractory-Metal«Base Brazing Alloy Study

Metallographic Observations

. Flowability
Brazing Alloy Flow of Joints
Composition Point Brazed on Brazed on
(wt %) ) Ch Mo Brazed on Brazed on
Cb Mo

84 Zr—16 Fe 9347 Good? Good Severe cracks Severe cracks

65 Zr~25 V-10 Fe >1300

75 Zr~15 Cr-10 Fe >1300 Good Poor® Severe cracks

82 Zr18 Cr 13007 Good Goad Severe cracks No cracks

65 Zr~25 V-10 Ge ~ 1300 Fair? Good Severe cracks No cracks

67 Zr—29 V4 Cr >1300

63 Zr=-27 V-10 Cr >1200

60 Zr—-26 V14 Cr >1200

57 Zr~24 V-19 Cy >1200

70 Zr~-30 V 1230% .Good Good No cracks No cracks

67 Zr—29 V-4 Fe 1300 Good Good

65 Zr--28 V-7 Fe >1200

60 Zr-26 V-14 Fe > 1200

L S B 4 .

63 Zr-27 V-10 Fe > 1300

50 Zr—-21 V=29 Mo >1300

57 Zr—-24 V-19 Mo >1200

63 Zr~27 VY-10 Mo >1200

60 Zr-25 V-15 Cb 1300 Goed Good No cracks No cracks

69 Ti-31 Fe 10857 Good Good Slight cracks Severe cracks

72 Ti~28 Co 10254 Goad Gaad Na cracks No cracks

M. Hansen, Constitution of Binary Alloys, McGraw-Hill, New York, 1958.

bGood indicates continuous filleting and extensive spreading en joint.

“Poor indicates wetting only at contact points.

dFair indicates intermittent filleting and little spreading on joint.
the brazing temperature. After 8- and 100-hr aging WELDING STUDIES OF COLUMBIUM
periods the specimens will be examined to determine AND MOLYBDENUM
the extent of diffusion of the brazing alloy elements R. L. Heestand
into the base metal or alloying of the base metal
with the brazing alloy. If such diffusion or alloy- A small, glass dry box containing a stationary
ing occurs, it is hoped that it will increase the re- torch and a movable table was assembled, as shown
melt temperature of the brazing alloy and thus in- in Fig. 1.3.4, with vacuum equipment capable of
crease the maximum service temperature attainable. maintaining vacuums as low as 4 x 105 mm Hg, for

LY I -4

31

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED

Fig. 1.3.2. As-Brazed Molybdenum T-Joint Brazed with an Alloy Containing 84 wt % Zr and 16 wt % Fe. The

joint contains numerous fillet cracks. As-polished. 150X. (Gumdigssm®l with caption)

. Molybdenum T-Joint Brazed with an Alloy Containing 70 wt % Zr and 30 wt % V.
®Swmtigiauin| with caption)

32

UNCLASSIFIED

Y-26820

As-polished.

kR

INCHE S

ETTTTRITTTRITITEIL:

o
1s0x

50X.
PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
PHOTO 43496

Fig. 1.3.4. Welding Dry Box with Stationary Torch and Movable Table.

welding studies of columbium and molybdenum.
The equipment was designed to make both fusion
welds (no filler metal addition) and welds with
filler metal, and it can be used for the evaluation
of the weldability of refractory metals and their
alloys in controlled atmospheres. The setup for
preparing edge-fusion and butt joints in the con-
trolled atmosphere chamber is shown in Fig. 1.
Several typical butt welds made on arc-cast
columbium sheet in the automatic welding dry box
in a purified argon atmosphere are shown in Fig.
1.3.6. Free-bend tests at room temperature showed

samples of all these welds to be ductile. Metal-
lographic studies of the samples are under way.

Five different heats of arc-cast molybdenum sheet
for use in welding studies were analyzed for carbon,
nitrogen, and oxygen. Samples of the 0.040-in.-
thick sheet were then welded in the dry box de-
scribed above and machined into bend specimens.
The bend specimens are being tested to determine
the effect of the contaminants on the ductile-
brittle transition temperature of material from each
heat. All specimens tested thus far have been
found to be brittle at room temperature.

33

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
PHOTO 43497

Fig. 1.3.5. Close-Up of Dry Box Equipment Showing Setup for Preparing Edge-Fusion and Butt Joints.

UNCLASSIFIED
Y-26115

INCHES

Fig. 1.3.6. Butt Welds in Columbium Sheet Made with (a) the Addition of 85 wt % Zr-15 wt % Cb Filler Wire, (b) the
Addition of 82 wt % Zr—15 wt % Cb=3 wt % Mo Filler Wire, (c) the Addition of 100 wt % Cb Filler Wire, and (d) Without

Filler Wire. (@uWniiggmsim with caption)
The selection of structural materials capable of
withstanding service loads at the high temperatures
at which reactors for the propulsion of aircraft will
operate has necessitated determinations of the
basic strength properties of new materials and the
development of accurate and precise analytical pro-
cedures for the evaluation of materials. Equipment
is now available which permits the rapid testing of
materials in controlled environments at temperatures
up to 2500°F. Means for obtaining quantitative de-
sign information on the behavior of metals under
complex states of stress and strain have been
studied, and it has been found that experimental
creep data are in substantial agreement with the
results of calculations based on accepted theories
of creep and fracture.

DEVELOPMENT OF TEST EQUIPMENT
FOR HIGH-TEMPERATURE INVESTIGATIONS

D. A. Douglas

In the development of equipment for studying the
mechanical properties of both metals and nonmetals
in the temperature range from 1800 to 3000°F, it
was necessary to provide control of the test en-
vironment, which would be gas or a vacuum, to
provide control of the temperature gradient and the
temperature fluctuation comparable to that provided

UNCLASSIFIED.
ORNL=LR-OWG 33728

U-CUP SEALS
FOR PULL ROD.

~PULL ROD

 WATER-COOLED FLANGE

|— CHAMBER (10in. Dia
X 12in. LONG)

FURNACE (THREE

SPECIMEN WITH PLATINUM ELEMENTS

P1-Rh THERMO- SEPARATELY CONTROLLED)
COUPLES —— 1
__TO VARIAC AND
ALuNDUM | | [CONTROLLER 110 v
EMFFLE === | O VARIAC AND
i [[CONTROLLER 1 v
U E===—rocroumo

TO VARIAC AND

CONTROLLER 110 v

Fig. 1.4.1. Diagram of an Apparatus for Creep and
Tensile Tests in Controlled Environments at 2000°F and
Above.

PERIOD ENDING SEPTEMBER 30, 1958

1.4, MECHANICAL PROPERTIES
D. A. Douglas

for conventional creep and tensile tests, to assure
accurate strain and load measurements, and to keep
the design simple and compact so that specimens
could be tested quickly. A schematic diagram of
the apparatus designed within this framework is
shown in Fig. 1.4.1, and a photograph of the parts
ready for assembly is shown in Fig. 1.4.2. The
use of U-cup O-rings as sliding seals simplifies
the application of the load and accommodates the
usual displacement in the tensile and creep tests.

y Lo
h Vason

)

Fig- 1.42. Specimen, Pull Rods, and Grips Used in
Apparatus Shown in Fig. Ld.1.

35

ANP PROJECT PROGRESS REPORT

The furnace element currently being used in the
apparatus is platinum, but molybdenum or tungsten
could be used for operation at higher temperatures.
The three elements of the furnace are separately
controlled by Variacs, and temperature deviations
along the gage length can be held to +5°F,
High-purity argon has been used as the test en-
vironment for all of the tests conducted to date.
The apparatus is amenable to evacuation, but the
increased time and trouble involved in vacuum
tests make high-purity inert gases more desirable
as test environments, The reliability of the ap-
paratus has been checked in three tests on mo-
lybdenum specimens at 2000°F, and the results
indicate that a suitable device has been achieved.

MULTIAXIAL CREEP STUDIES
C. R. Kennedy

Analytical studies have been made of creep data
obtained for Inconel tubular specimens in tests in
which internal pressure, as well as axial loading,
was applied in the apparatus described previously.’
The internal pressure required to produce a given
tangential stress was calculated by using the thin-
wall formula. The axial stress was calculated
conventionally, with the axial stress produced by
the internal pressure being taken into consideration.
The average radial stress was obtained by dividing
the internal pressure by 2, After-test measurements
of the tube dimensions were made to determine the
three principal strains at rupture. In those cases
where end effects or bulging occurred, tangential
strains measured at regions of highly localized de-
formation could not be compared directly with the
over-all axial rupture strain in connection with
subsequent data analysis. |t was necessary, there-
fore, to use average tangential strains for this
purpose. These were obtained by averaging five
measurements of the tangential strain taken at
equidistant positions along the 21/2-in.-goge length,
An observation that the sum of the three measured
average strains often was greater than zero sug-
gested an increase in volume during the test, Since
a volume change is inconsistent with the assump-
tions used in the mathematical analysis of creep
flow and fracture, this matter was investigated by

c. R Kennedy and D, A, Douglos, ANP Quar. Prog.
Rep. Sept. 30, 1957, ORNL-2387, p 185.

36

measuring the specific gravities of after-test speci-
mens by using a Jolly balance. In these tests, ob-
served density changes were found to be less than
the sensitivity of the balance, and thus the apparent
volume change based on after-test strain measure-
ments was not real. The error was believed to be
associated primarily with the radial strain measure-
ments and was attributable to the inaccuracies in-
herent in measuring small changes in the sizes of
small sections with curved surfaces. In the data
analysis which follows, the ‘‘measured’’ radial
strains at fracture were determined from axial and
tangential strains based on constancy of volume,

Creep Analysis

A workable formulation for steady-state creep is
that of Soderberg,? which is based upon the follow-
ing assumptions: (1) the material is originally
isotropic and remains isotropic during the creep
process so that the stress effect on creep rates
canbe expressed in terms of the principal stresses;
(2) the principal stresses coincide with the principal
strain directions; (3) the volume remains constant
so that hydrostatic pressure has no influence and
the stress effect must therefore be a function only
of the difference of the principal stresses (that is,
shear stresses); (4) the stress state remains constant
with time; and (5) the effective creep stress, 7,
and the effective creep strain rate, €, for all stress
states at a given temperature are related through
the material constants K and » by the relation
(1) E=KG" .

According to this formulation, the principal steady-
state creep rates are

2¢. R, Soderberg, Trans, Am, Soc. Mech, Engrs, 63,
737 (1941),

where the subscripts Z, 6, and R denote axial, tan-
gential, and radial stress or strain rates, respec-
tively. Based on the assumptions stated, only the
principal stresses and the relation between creep
stress and creep strain for simple tension are re-
quired in order to compute multiaxial creep rates
according to Eq. 2, The value of & depends on the
flow criterion selected to correlate the data. Based
on the von Mises (distortion energy) criterion, it is
given by

1
3) © =———{(az-—08)2+ (ag-oR)2+

V2

and, by the Tresca (maximum shear stress)
criterion, it is given by

(4) g zamax —Omln .

The effective creep strain rate, €, is the creep
strain rate associated with & for a value of uniaxial
stress equal to the calculated effective stress.

The design criterion based on creep rate in many
applications is the time to reach a limiting strain.
For the purpose of comparing experimental results
with those predicted by Eq. 2, the average creep
rate to 2% strain was used. In those cases where
rupture occurred before 2% axial strain had oc-
curred, the average creep rate to one-half of the
total strain was used,

Comparisons of the measured axial creep rates
with those predicted from Eq. 2 by using both the
von Mises and the Tresca criteria are shown in
Figs. 1.4.3, 1.4.4, and 1.4.5. In general, the corre-
lation of the experimental data obtained by using
the von Mises criterion is slightly better than that
obtained by using the Tresca criterion. In most
design situations, however, the stresses are known
only approximately, and, in view of the many inde-
terminate variables which influence creep, the
simpler and more conservative Tresca criterion is
clearly adequate for engineering purposes, The
advantages of using the Tresca criterion for multi-
axial creep have been demonstrated by Wah!? for
the case of rotating disks. According to Wahl, the
use of the von Mises criterion predicted much lower
creep rates than those obtained experimentally.

3A. M. Wahl, J. Appl. Mechanics 23, 231 (1956).

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
ORNL-LR-DWG 31530

-1

10 N S ,717211’ - = ’4177 pa— :‘ -
e 2 0, 6000 psi — - |

~ B L

= 4 —F r

E e o //* i;[ o7y =5000psi ~ ]

o e .

S B

g, e A !

é TO .. — /Il 4

o S S—— i 4000 psi ]

fl — = T |

5 [ A

= ,’

3 . 2 s, =3000psi
wn o, = psi
——— T Pt 4

e e e LI
= "~ "
_I ~— 1V
g
>
< ——-I7 = TRESCA CRITERION
~——[¥ = von MISES CRITERICN .
104 }
@ 4 2 1.33 1 0.8 0.67 0.57

STRESS RATIO (a7 /)

Fig. 1.4.3. Axial Tension Creep Rates vs Stress

Ratios for Four Principal Axial Stresses.

UNCLASSIFIED
ORNL-LR-DWG 34534R

10 % 1
5
2 rad
L 0= 6000 psi
77 4
4 ‘l I,
> 7
5 7 7
AR
// v Wa
2 ,/ - [5=4000psi ]
7/ - ®
107! LA s
77 > 7
r i d P
5 r/ /" rd —~
/ 7 2V 1 7 9,=3000 psi
7 r P
> /// // Py ,1/ /
L] 7/ c/ ,/
1g72 4 A
'I‘II, ,ll I,I
5 17— 7
p 7 Pl
/ )
7
) * 1/ / v
7
1672 ’
5 7 ' —
4 ~=—— T =TRESCA CRITERION _|
4 V'=VON MISES CRITERION
2 H
T
G0.25 o] -0.25 -0.5 -075 -4 -14.33

STRESS RATIC (O}/Oé)

Fig. 1.4.4. Axial Compression Creep Rates vs Stress

Ratios for Three Tangential Stresses.

37

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL-LR-DWG 34532

10 [ I T
i ! 1
I ] ]
1 T 1
5L == 7=TRESCA CRITERION
| em—— =VON MISES CRITERION
T2
~
B2
o 5g=2000psi
'_
<1 —
(0 /I
[a
w a4
0.5
g 7 /
2 /S
o / /
)
Q 0.2 7 74
o / /@
= /
o oA /
© ) 77—
- A 4
< YA 4
% 0.05 //
0.02
0.01
4 -2 -3 -4 -5

STRESS RATIO (g /)

Fig. 1.4.5. Axial Compression Creep Rates vs Stress
Ratios for Stress Ratios of Less than —1 with a Constant
2000-psi Tangential Stress.

Rupture Analysis

Rupture data are presented in Fig. 1.4.6 in which
the stress ratios calculated by the thin-wall formula
are plotted against the rupture life. The rupture
data are also presented in the form of 100-, 500-,
and 2000-hr isochronous fracture envelopes in Fig.
1.4.7. The envelopes were obtained from cross
plots of the data shown in Fig. 1.4.6. The 100-hr
fracture envelope predicted by the maximum princi-
pal tensile stress criterion is also plotted in Fig.
1.4.7.

It may be seen that the results calculated and
plotted in the manner described show deviations
from the simple maximum principal stress criterion.
There are, however, at least two important factors
which should be considered in a realistic analysis
of the data, Both of these factors involve in some

38

UNCLASSIFIED
ORNL LR- DWG 31533

© . TTT
T A
6000 - psi MAXIMUM
2.00 |- PRINCIPAL STRESS
— “_#_ \__H
b@
& 100 —
(@]
5 050 - %—1
a o
(5]
2 ]_T \
@ Q — - - +
= = [
m . . H ' H H
L - f£3000- psi MAXIMUM
| 1/ | [T PRINCIPAL STRESS
—-0.50 — —A 1
B J H o W il
—1.00 L L] H\ 1N [ J_J___
20 1000 10,000

TIME TO RUPTURE (hr}

Fig. 1.4.6. Rupture Times vs Stress Ratios for Three

Average Maximum Principal Stresses,

UNCLASSIFIED
ORNL-LR-DWG 31611

8000
.-\c\ i
—— |
6000 [— I—— |
GI\
."'"‘--.‘ [
~— |
4000 T3
'--...______-.‘.-‘ ‘
.\_T > P 1
! |
2000 } | '
.’(7‘) ™Y ® |
e ’ » I
o s ’
a J |
o M |
= , 1 !
w
g I ; |
% 2000hr § I H |
< -2000 .' 7 / .
P / |
[ ] ®
-4000
100 hr 1
- 6000 100-hr FRACTURE ENVELOPE
PREDICTED BY MAXIMUM /}
PRINCIPAL STRESS CRITERION |
-8000 | |
0 2000 4000 6000 8000 10,000

TANGENTIAL STRESS {psi}

Fig. 1.4.7. Rupture Data Presented as 100., 500-, and

2000-ht Isochronous Fracture Envelopes.

manner the changes in stress state associated with
deformation of axially loaded pressurized tubes.
The first factor is concerned with the stress-state
changes in uniformly deformed tubular specimens.
For an equivalent amount of strain under these con-
ditions, the increase in the axial stress is less
than the increase in the tangential stress over the
entire range of stress ratios studied. For example,
after 10% axial strain in a specimen tested in
simple tension the '‘true’’ axial stress becomes 1.1
times as great as the original axial stress; however,
after 10% outer-surface tangential strain of a speci-
men subjected to tangential stress only, the aver-
age tangential stress becomes 1.18 times as great
as the original tangential stress, For this example,
the increase in the tangential stress is greater than
that of the axial stress by a factor of 1.07. This is
to be compared with the data plotted in Fig. 1.4.7,
where it is shown that the axial stress required to
produce rupture in 100, 500, or 2000 hr is greater
than the corresponding tangential stress by a factor
of about 1.25. This discrepancy can be rationalized,
however, by noting that the effect being considered
is intensified by the second important factor which
is associated with localized deformation or bulging,
particularly in tension-compression states,

The relative amounts of bulging are demonstrated
in Fig., 1.4.8, which is a plot of the difference in
the maximum and the average tangential strain at
rupture versus the stress ratios. It is shown that
the bulging between stress ratios of « to ‘/2 is
insignificant; however, the extent of bulging in-
creases substantially as the stress ratio drops
from ‘/2 to ~1. The result of this bulging, if only
the increase in diameter of the specimen were con-
sidered, would be to increase the tangential stress,
which is the maximum principal stress in this
range. The relation between the degree of bulging
and the deviation of results from the maximum
principal stress criterion may be seen by comparing
Figs. 1.4.7 and 1.4.8. For example, in Fig. 1.4.8
the degree of bulging obtained by testing at 3000
psi is shown to be less than that obtained by
testing at 4000 and at 6000 psi. This is consistent
with the rupture results presented in Fig. 1.4.7,
which show that the deviation of the 2000-hr frac-
ture envelope from a vertical line representing a
maximum principal stress criterion is much less
than that for the 500- or 100-hr curves (vertical
reference lines not shown). Accordingly, the re-
sults of this study appear to support the maximum

PERIOD ENDING SEPTEMBER 30, 1958

principal stress criterion for time-dependent frac-
ture under multiaxial stress conditions. It has
been shown, however, that in applying this cri-
terion it is important to compensate for the re-
duction in rupture life caused by changes in the
stress state associated with deformation,

The role of the maximum principal stress is dem-
onstrated by crack patterns (Fig. 1.4.9) on the out-
side surfaces of specimens. A close inspection of
these patterns reveals that the cracks propagate in
a stair-step manner, and, in all cases, the general
direction of the intergranular cracks is normal to
the maximum principal stress, |t was observed that
all cracks were intergranular and that none of the
specimens exhibited necking prior to failure,

The effect of multiaxial stress states on the
rupture elongation was also investigated. Fracture
strains based on after-test measurements of speci-
mens tested with a 4000-psi maximum principal
stress are shown in Fig. 1.4.10, in which the aver-
age axial, tangential, and radial strains are plotted
versus the stress ratio. The strain at failure, €/,
for tests with the same maximum principal stress
appears to be related to the deviatoric stress in the
following way:

f ' 06+0R
ALY

f OZ+09
(o5

A comparison of these relationships with the data
in Fig. 1,4.10 indicates good agreement, The value
of B is constant for all the stress states tested
with the same maximum principal stress, but it
varies with the maximum principal stress, as shown
in Fig. 1.4.11. The rupture strains from all tests
are compared with Eq. 5 in Fig. 1.4.12, which is a
plot of the ratio of rupture strains to the axial strain
for simple tension vs the stress ratio, Again, there
appears to be good agreement with Eq. 5.

The ratio of the shear strain at rupture, as deter-
mined from

yf: Ef --Ef

max min 7

to the shear strain at rupture for specimens tested

39

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL-LR-DWG 31535

y2
/20

¥ /

6000 - psi MAXIMUM
PRINCIPAL STRESS

o //

] //

! >
4000 - psi MAXIMUM/

PRINCIPAL /

52
z
a
o
|_
0
J
g
=
z
w
>
< 6
= STRESS
w
T}
&
w 5 -
>
a
o, /
= ®2000-psi MAXIMUM 5
s PRINCIPAL STRESS |
=
- L
z ( Sl
w
L (0] /
L
a 0 /

-1

-2 J

-3

© q 2 £33 { 0.75 0.5 0.25 0 -0.25 -0.5 -075 -{

STRESS RATIO (0'2/0'8)

Fig. 1.4.8. Difference in Average and Maximum Tangential Strains at Rupture for 3000-, 4000., 6000-psi

Maximum Principal Stresses at Various Stress Ratios.

40

184

UNCLASSIFIED
Y-26205

6000,/4000 6000,/6000

}

TANGENTIAL STRESS AXIAL STRESS
s

1500,/6000

Fig. 1.4.9. Crack Patterns of Outside Surfaces of Specimen Tested with a Maximum Stress of 6000 psi at 1500°F in Argon. Dye penetrant applied
after test.

8561 ‘06 ¥39W3Ld3S INIONI dOI¥3d
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL—-LR-DWG 31586

14 A
2 5
//
10
o, + O [
f_ 8 i /< o,+ o,
e, =80, - —) F_ Z' R

@)
- 7
7 (? 0
b@ 4
2 / AN
é 2 y. N
}_
J
g o . :. -1
T /
W / \‘
g _o A N //
/T \‘ -~
Kl Xefi:g(%— >
[ ]
-6 - )
\i I i/ \‘
s N\
[ ] \‘
—10 \
® AXIAL STRAIN \
-4 O TANGENTIAL STRAIN N
B RADIAL STRAIN €,=—(€, +€,) \?
L
© 4 p 1.33 { 075 05 025 O -025 -05 -075 -1

STRESS RATIO (g, /0)

Fig. 1.4.10. Strain at Rupture vs Stress Ratio for Specimens Tested with a Maximum Principal Stress of 4000 psi.

42
UNCLASSIFIED
ORNL-LR-DWG 31587

£ 10000

wy

wy

w -

e

5 —

W 5000 //

5 -

z /'

W

& - —

= o

<

a

2 2000

a

a

=

2

=

% 1000

3 0,015 0.020 0.025 0.030
8

Fig. 1.4.11. Variation of B {Constant of Equation Re-
lating Strain at Failure and Deviatoric Stress) with
Maximum Principal Tensile Stress for Specimens Tested
at 1500°F in Argon.

under simple tension is also plotted against stress
ratio in Fig. 1.4,12, This shows that specimens
which obey Eq. 5 with the same maximum stress in
the tension—~tension stress state fail with the same
shear strain regardless of rupture life, Thus, for
specimens of this type, the rupture strain in the
tension—tension state is a function of the maximum
principal stress only, It should be realized that al-
though the total elongation in a particular direction
may be less under combined tension—tension
stresses, the shear strain or total deformation at
rupture is the same.

Specimens tested in the tension—compression
stress state cpparenfly can sustain a much greater
shear strain before fracture than those in the
tension—tension state. As shown in Fig. 1.4.12,
specimens tested under pure shear sustained twice
the shear strain at fracture as the specimens tested
in simple tension,

STRAIN-FATIGUE STUDIES
R. W. Swindeman

Many of the failures encountered in engineering
devices result from dynamic loads which are im-
posed on the structure either mechanically or
thermally. Most often, fatigue failures are con-
sidered to be the result of rapidly fluctuating
stresses which introduce damage in the material
on a very microscopic scale. Thus in low-tempera-
ture fatigue, no bulk plastic straining of the metal
can be discerned. However, at elevated tempera-
tures, where relaxation of stress can occur, meas-
vrable amounts of plastic strain may be induced

PERIOD ENDING SEPTEMBER 30, 1958

during each stress reversal. This is particularly
true in the case of restrained structures when they
are subjected to large thermal fluctuations, Quite
often, under such conditions, the thermally induced
stresses occur over relatively long time cycles and
are large enough to exceed the elastic limit of the
metal. Thus, in studying the behavior of materials
loaded in this manner it is more meaningful to think
of the metal as experiencing a number of strain
reversals which ultimately lead to failure, rather
than to attempt to base such failures on the inde-
terminate stress state,

Recent studies have revealed that the plastic
strain history of a metal under dynamic load con-
ditions provides extremely useful data for calcu-
lating the metal life consumed and the service life
remaining under expected operating conditions.
This idea was conceived by Manson? and then dem-
onstrated experimentally by Coffin® by thermally
cycling stainless steel under conditions of restraint
until failure occurred. The relationship between
plastic strain and cycles to failure was found to be
of the form N%€_ = K, where N is the cycles to
failure, € _is the strain per cycle, and @ and K are
constants which depend on the material and test
conditions.

Most of the subsequent work has been within the
temperature range where the rate of work hardening
was appreciably greater than the recovery rate.
The work described below concerns metals at tem-
peratures where creep and relaxation are the domi-
nant factors. The fact that creep and relaxation
are dominant in the temperature range of interest
made it both feasible and attractive to substitute
mechanical loads for thermal fluctuations as a
means of producing strain cycles. The apparatus
used is described in detail in a separate report.’

Most of the investigation was conducted with
Inconel as the test material; but, a few tests were
made with Hastelloy B and with beryllium, Thus,
representative data can be compared for metals of
quite widely different degrees of strength, duc-
tility, and fabrication history. Conventional plots
of the plastic strain per cycle, €p1 VS N, the number

4s, s, Manson, Bebavior of Materials Under Conditions
of Thermal Stress, NACA-TN<2933 (July 1953),

3. F. Coffin, Jr,, Trans, Am. Soc. Mech. Engrs. 76,
931 (1954),

6C. R. Kennedy and D, A, Douglas, Plastic Strain
Absorption as a Critetion for High Temperature Design,
ORNL-2360 (April 17, 1958).

43

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL—LR—DWG 31536R

2.0 .
Z
&
’(7) S i
o %S :
el | |
= 1.5  — L — |
© 9 | |
| = ’ :
xr o y! e: eg '
P~ - — =( — = — ) FOR o, =CONSTANT
a Yo €7, @ 2 @
G I I N N
1.0 J J
T
1.5 S S R SR R
1 1 :
1 i ! ; ‘(
i Tr + 0, '
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o = >£
1.0 ‘ | L. il -
0.5

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o 4 2 1.33 { 075 0.5 0.25 0 —-025 -0.5 —075 —{°

STRESS RATIO (92 /0p)

Fig. 1.4.12. Plots of Rupture Strain Ratio and Rupture Shear Stress Ratio vs Stress Ratio.

of cycles to failure are shown in Fig. 1.4.13, The
co-ordinates were plotted on logarithmic scales so
that if Manson’s relationship N®¢, = K were satis-
fied, the data points would define a straight line,
It is apparent that this condition is satisfied for
each material, but it is also evident that there is a
variation in the slope of the line from one metal to
another, These slopes are approximately —0.58 for
Hastelloy B, —0.76 for Inconel, and —0.81 for be-
ryllium. Such values are in contrast to the findings
of Coffin’ who shows ato be very nearly ~0.5 for
several different metals tested at temperatures be-
low 1200°F,

It has also been noted that for any given material,
K can be varied by changes in the metallurgical
structure. A plot of the strain-cycling properties of
Inconel rod specimens machined from the same heat
but differing in the annealing treatment given before
the test is presented in Fig. 1.4.14. The specimens
annealed at 1650°F had a finer grain size and were
much more resistant to fracture than the specimens
annealed at 2050°F, Comparative data for rod ma-
terial at 1300, 1500, and 1600°F, which was re-
ported previously,® indicate that grain size effects
are most pronounced at 1500 and 1600°F; however,
at 1300°F coarse- and fine-grained material have
nearly identical strain-cycling properties, Inconel
is a solid-solution type of alloy that is not subject
to phase changes or aging reactions, and therefore
the difference in behavior seems to be clearly de-
pendent on the relative grain size, It is also note-

L. F. Cotfin, Strain Cycling and Thermal Stress
Fatigue, paper presented at the Fourth Sagamcre Con-
ference, Raquette Lake, N. Y., Aug. 21-23, 1957.

UNCLASSIFIED
ORNL-LR-DWG 31443

L-© HASTELLOY B TESTED AT 4650°F| Ml
=4 BERYLLIUM TESTED AT 1250°F | [Nl 4

€p, TRUE PLASTIC STRAIN PER CYCLE (%)

05 - oy Ll
S l’iHflfl e 1 A S A1
0.2 L1 NN I ‘ L L
041 1 10 100 1000 10,000

A, CYCLES TO FAILURE

Fig. 1.4.13. Comparison of the Strain-Cycling Char-

acteristics of Inconel, Hastelloy B, and Beryllium.

PERIOD ENDING SEPTEMBER 30, 1958

worthy that the two lines are parallel, and thus the
a value is consistent for both types of specimens,

The question of whether specimen geometry seri-
ously affects the strain-cycling properties also
merits consideration. The strain-cycling properties
at 1500°F of rod and tube specimens are compared
in Fig. 1.4,15. Tube data fall slightly below the
rod data, but this may be attributed to two factors.
First, the methods of sensing failure are ditferent;
the tubes are considered to fail when the first
crack propagates through the wall, whereas rods
are tested to complete failure. Further, the rod
specimens have a slightly finer grain size than that
of the tube specimens. The tube data include test
results from two different heats of metal, and the
agreement indicates that the cycling properties are
not particularly sensitive to small variations in
alloy composition.

UNCLASSIFIED
ORNL-LR-DWG 31445

: ‘ T

Nl

L@ FINE-GRAINED RODS

1 ==

e Ll

A, CYCLES TO FAILURE

€5, TRUE PLASTIC STRAIN PER CYCLE (%)
3}

Fig. 1.4.14. Effect of Grain Size on the Strain Cycling
Properties of Inconel at 1500°F,

UNCLASSIFIED
CRNL-LR-DWG 31444

200 ‘ : T ‘ | |
|
100 it M l i1
w - ,,,,,,g\ i R D Y D 1
d 50 ™
C ; ™
E 20 [ i i Ny J i : I ;
=z L L | NN ! __ L ‘i
Z {0 N +
< [T T " AN - T
E s T 1 eSS T R R—
” [ ; Hd e ‘ LI
© st I R SR \\\ ]
= IR i ! . NG L
2 : N
q * RODS ; : \\\ y
& rlfatuees eat Ay Jr L LUL | TNEIN |
w =& TUBES (HEAT B) e i RIS
z 05 fi o TUBES TO COMPLETE RUPTURE i}~ I B —
= SCATTERBAND I Ll { T P S ™
. ! N |
‘Uq Q.2 | ‘ ’{ ‘ | 1T —t ‘: Hb ; T J HER ‘\ \\A
S I TN O A A RS
0.2 3 10 100 1000 10.000

A, CYCLES TO FAILURE

Fig. 1.4.15. Strain Cycling Properties of Inconel
Tubes and Rods at 1500°F.

45
ANP PROJECT PROGRESS REPORT

Temperature Dependence

One of the most important considerations in regard
to test variables is the influence of temperature on
the number of cycles to failure. This is particularly
important where it is desired to apply isothermal
mechanical strain-cycling data to design problems
where thermal cycling is involved. The results for
tests at 1300, 1500, and 1600°F that were con-
ducted to determine the effect of temperature are
presented in Fig. 1.4.16. The band represents the
scatter of points obtained in an extensive program
of tests of tubes and rods at 1500°F, Nearly all
the data for rods at 1300°F and for tubes at 1600°F
fall inside this band. The rod data obtained at
1600°F fall consistently above the band because of
the grain size effect mentioned previously. It ap-
pears, therefore, that within this particular range
there is no major change in fracture behavior as a
function of temperature. The fact that Inconel is
metallurgically stable, that no grain growth occurs,
and that the deformation mechanism is uniform
throughout this span are important points to bear
in mind when considering these results.

Since temperature appears to have no major influ-
ence on the number of cycles to fracture in the
range 1300 to 1600°F, it would appear reasonable
to expect mechanical strain cycle results under iso-
thermal conditions to correlate with those obtained
by thermally cycling about a corresponding mean
temperature. Data illustrating the validity of the
assumption are presented in Fig. 1.4.17. The
thermal excursions about the 1300°F mean covered
the range from 1000 to 1600°F. At the 1500°F

mean the temperature extremes were 1400 and

UNCLASSIFIED
ORNL-LR-DWG 31442

,\:400 B R ESE: ‘F TR TTTF {fl TF
: [ I T e
w T ™ ; T .
C 20 -t -1 : j - Ll
«
g 10 = 1 S SEEREL
z ot v menan)
4 5 - e
a T
7 [ & 1300°FRODS ~ iy : y i
) 2 — & 1600°FRODS P B R : e M ]
E o {600°F TUBES o MNTB
< 1 1500°F SCATTERBAND, ~ 5= I S R RN
a - RODS AND TUBES (¢~ ' T i
w OS5 — 17 7rrmn NN N
2 BATHI I LTSNS
}-_ 0.2 u - ‘ ! ’ ; I T R ) + _\_\;..
ot g | ! ‘ |

of] : L il ‘ L

X 1 10 100 1000 10,000

A, CYCLES TO FAILURE

Effect of Temperature on the Strain

Fig. 1.4.16.

Cycling Properties of Inconel.

46

UNCLASSIFIED
ORNL-LR-DWG 31444

10 l ‘ ] T 1 1T
g SEsHimliREss T i ERil
wl E } — ‘1 l‘ i -
é 2 1 \\ : ‘
(8} } ‘\.\l\h
G ok M AL
a - — - o H
z 05 " . S \\ H
- S S i N
.g_ oL L N |
5 ol L LI A AN
2 7 ® THERMAL, 1300°F MEAN at RN
o4 |- © THERMAL; {500°F MEAN Q‘&, 4 b
< F——MECHANICAL; SCATTERBAND e N
& 005 | - | WHICH COVERS 1300, 1500, N |
) " | _AND1600°F DATA -
F o002 |
3 |

0.0 i
10 100 1000 10,000 100,000 1,000,000

A, CYCLES TO FAILURE

Fig. 1.4.17. Compatison of the Thermal and Mechanical

Strain Cycling Properties of Inconel.

1600°F, It is evident that in the absence of metal-
lurgical instabilities a metal responds to the me-
chanically produced strain in the same manner as
to a strain thermally induced.

Frequency of Dynamic Loads

A very important variable to be considered in
studying the effect of dynamic loads on the be-
havior of metals at elevated temperatures is that of
frequency. It is well known that the test frequency
can be changed over a wide range without preducing
a marked effect on the plot of stress vs number of
cycles to failure (S-N diagram) in low-temperature
fatigue tests. On the other hand, several investi-
gators have found that frequency does exert an in-
fluence on the cycles to rupture at high tempera-
tures. This is especially true at lower frequencies
where creep or relaxation may occur. Hyler® has
conducted fatigue studies on Inconel at 1200 and
1600°F at load frequencies of 60 and 600 cpm. He
observed essentially no difference in the S-N dia-
gram at 1200°F, but he found an appreciable fre-
quency effect at 1600°F. Specimens cycled at 600
cpm endured 10 times more cycles than those cycled
at 60 cpm. Fatigue studies on lead have yielded
similar results. Dolan? reports data for reversed
flexure fatigue tests on lead at room temperature for
frequencies of 1, 6, 44, and 248 cpm. Curves re-
lating the strain per cycle and the number of cycles
to rupture were determined for each frequency. A
comparison of the curves revealed that the cycles

BW. S. Hyler, Battelle Memorial Institute, unpublished
work,

9. J. Dolan, Metal Progr., 61 (3), 55 (1952).

to rupture increase with increasing frequency. Un-
fortunately, the strain values ranged only from 0.1
to 1% and the reported strains include both elastic
and plastic components. Since plastic strain per
cycle was not determined in any of these studies,
the direct use of these data for studying the fre-
quency effect on the strain-cycling behavior is not
possible. The results raise the question, however,
of whether or not the difference in the plastic strain
absorbed could explain the variation in properties
with frequency. Since plastic strain is generally
considered a reliable measure of damage to a
structure, it would appear that specimens strained
within the same limits of plastic deformation should
survive an equivalent number of cycles before
fracture, regardless of the rate at which the strain
is introduced. Design problems involving tempet-
ature fluctuations are, however, often concerned
with slow thermal cycles, and therefore a series of
tests to investigate frequency effects was initiated.

The majority of the data previously discussed was
obtained by using a constant load; thus the creep of
the specimen continued until the limiting strain
condition was reached in the desired time interval.
This type of test is quite easy to regulate for short
time cycles, but the prescribed load required to
produce specific strains in 10 or 30 min is diffi-
cult to obtain, Thus a slight modification of the
equipment was made to provide adjustable, but
rigid, reference limits.® In test, the specimen is
strained to the selected limit very rapidly and held
at this strain for the selected time interval, During
this period the elastic strain is gradually converted
into plastic strain by the relaxation process, It is
therefore quite simple to control the frequency and
to vary it from times as short as l/2 min per cycle
up to whatever time cycle seems practical, A com-
parison of results obtained by the creep method and
by the relaxation method is shown in Fig. 1.4.18,
Since creep and relaxation are manifestations of
the same deformation phenomena, the excellent
correlation shown between the results of the two
types of tests was expected.

The data from a series of tests at 1500°F in
which the standard 2-min/cycle period was used
are compared in Fig. 1.4.19 with data from tests in
which a 30-min/cycle period was used. At high
percentages of plastic strain per cycle, good corre-
lation exists, but a significant deviation is shown
for the low strain values. Data obtained at 1650°F
with a 30 min/cycle period follow a trend similar
to that found at 1500°F, but at 1300°F, as shown

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED

—~ 100 ORNL-LR-DWG 31439
e e e i s o = 5 1 e e S
— I —r TR ST HE - 8 - I
== e ==
g l i
5 . TN T fi T
" | A |
10— NGNHE Ll
z — N EEs

| i [
x 51— 1 B Al N
= | LT | | ] i
= SN T
© 2| ORELAXATION CYCLE T NN I
= CREEP CYCLE; SCATTERBAND N L ]
% S , NN |

T 1 15N | 0 1 S— ~ il ‘
= o - N
w 0.5 o o Lo o fi
z 1 i ' SN
i Ii I S
. 0.2 ™
S0, Il | J

0.4 f 10 100 1000 10,000

A, CYCLES TO FAILURE

Fig. 1.4.18. Comparison of the Strain Cycling Proper-
ties of Inconel Subjected to Creep and to Relaxation
Cycles at 1500°F.

UNCLASSIFIED
ORNL-LR-DWG 31438

100 — -
5 RS SRS naa: R e m i e
< 50 ™ :
5 [ 1‘ ™ P J TTTH I
O | N
s 20 }‘ R 1 ST HII
& 10 | e NGl H-

T . s NN

pt ‘ % i ™17 N t

o - o H } T B A, R ATy, N (|

5 i ) Al ‘ NN

o2 i b [ : ' TP TN

54 7' ‘ “ L [ “ L o fo] INIIN L
< I~ 0 1500°F TUBES; 30 MIN PER CYCLE i — e Nt
& 0.5 ——1500°F SCATTERBAND; 2MIN PER CYCLE | H
g 7 ] ITHHJ_ ] H%[F I TSI
E oo S .
[ TETRAT T

O8] 1 10 100 1000 10,000
A, CYCLE TO FAILURE

Fig. 1.4.19. Effect of Frequency on the Strain Cyeling
Properties of Inconel at 1500°F.

in Fig. 1.4.20, the only test point which falls out-
side of the 2 min/cycle data occurs below 1% strain,
and, even here, the deviation is not appreciable.

There is, of course, considerable engineering
interest in the fact that for the same amount of
repeated plastic deformation, a metal subjected to
slowly fluctuating loads will survive for only a
fraction of the cycles of a specimen cycled more
rapidly. This is especially important since the
frequency appears to exert the greatest influence
for values of strain below 2%, which are also equi-
valent to the thermal strains most often encountered
during service, There is no evidence to indicate
that the longer thermal exposure encountered in the
low-frequency tests is detrimental, so it must be
concluded that the mechanisms controlling fracture
at high temperatures are strongly influenced by time
as well as plastic flow.

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL-LR-DWG 31440

I ® 1650°F TUBES; 30MIN PER CYCLE ' |#

| |——SCATTERBAND WHICH COVERS 1300, ||

E 1500 and 1600°F TUBE DATA; 2 MIN
05 PERCYCLE i o [y pff

ol L | o
i {0 100 1000
N, CYCLES TO FAILURE

T
|
T

= Ik — I e jas =S ss
D — T T T TR e
g © R S B T T
o SR N o N AN .. ™ .'.T . Lo L i %
& 20 ] Lfi.‘\\ '| [ Lf
z A T QU : :
& 10 b ;‘ : SSENR
Z t— 1 l e e IR
==,
— A . L U S tal
5 © {300°F TUBES; 30MIN PER CYCLE !
Q 2 i -
=
»
<«
]
i
w
5
o
-

.Ii

Fig. 1.4.20. Effect of Temperature on the Strain
Cycling Properties of Inconel Tested at a 30 min/cycle
Period.

Design Factors

This investigation of the high-temperature be-
havior of metals under dynamic loads reveals the
following significant points.

1. The relationship, NOLE!J = K, holds for several
metals at elevated temperatures provided the cycle
time is relatively short. The exponent, o, is de-
pendent on the material and varies from —0.58 to
—0.81. Since the strain-cycling curves of many

48

metals have exhibited a slope of —0.5 in low-tem-
perature tests, the greater slopes at high tempera-
tures indicate a dependence on temperature or
deformation mechanism rather than directly on ma-
terial composition. Since there is no significant
variation in a for Inconel over a 300°F range, it
appears that the deformation mechanism is the
major factor that affects the slope of the strain-
cycling curve for Inconel. Thus, there may be a
break in the strain-cycling curve similar to the one
observed for the equicohesive temperature in creep.
The K factor is changed significantly by changes
in grain size, and thus it is believed that the dis-
placement of the rupture curve can be attributed

to metallurgical changes.

2. Metals respond in the same manner to plastic
strain whether it is introduced by thermal changes
or mechanically imposed loads. This statement
must be qualified to the extent that the deformation
mechanism and metallurgical structure must not be
altered by the temperature fluctuations.

3. The rate at which strains are reversed has a
marked effect on the number of equivalent strain
cycles a metal can survive. This phenomenon,
which is apparently confined to the high-temperature
region, establishes the fact that plastic deformation
is not the sole mechanism controlling the fracture
life of a metal.

PERIOD ENDING SEPTEMBER 30, 1958

. 1.5. CERAMICS

SYNTHESIS OF ZIRCONIUM BORIDE
R. A. Potter

Small batches (approximately 200 g each) of zir-
conium boride continved to be synthesized in the
laboratory for additions to beryllium oxide control
samples. Some improvement has been made in the
synthesis process' by adding kerosene to the
batch. The zirconium silicate-boron carbide-
graphite reaction was expedited and carried to
near completion by the presence of the hydro-

carbons resulting from the ignition of the kerosene.

Several batches of ZrB, were semiquantitatively
analyzed by spectrographic methods, A typical
batch is >97% pure, with silicon being the pre-
dominant impurity.

DENSIFICATION OF BERYLLIUM OXIDE
R. A. Potter

Seven different compositions, given in Table
1.5.1, were mixed in a jar on rolls for 18 hr, and
the material was then pressed at 20,000 psi in
stee| dies, broken up, granulated by passing
through a 30-mesh screen, and remixed for 2 hr,
Right-cylinder shapes approximately 3‘/4 X 3/4 in.

]L. M. Doney, R. L. Hamner, and R. A. Potter, ANP
Quar. Prog, Rep. March 31, 1958, ORNL-2517, p 45.

Table 1.5.1. Compositions of Densified BeQ Bodies

Composition (Parts by Weight)

Composition

L. M. Doney

were made by pressing the materials at 15,000 psi.
These shapes were then fired to six different tem-
peratures (1650 to 1900°C in increments of 50°C).
The furnace consisted of an induction-heated
graphite tube with a positive flow-through of
helium. In each case a fast firing schedule of
approximately 2 hr was used.

Screening tests indicated that composition 2 was
superior to the others with respect to density. The
densities of this body at the six test temperatures
are compared with those of the 100% BeO control
sample in Table 1,5.2.

Magnesium oxide alone, as in composition 3,
acted as a densifier, but at about 1700°C it began
to volatilize rapidly and caused distortion of the
shape of the cylinder. At 1650°C the density of
composition 3 was found to be 2.9 g/cm?; how-
ever, there was a slight distortion in the shape.

For a further test, composition 2 was pressed
into right cylinder shapes approximately ]3/8 x 1 in.
in height. Forming pressures on these pieces were
6,000 and 10,000 psi. The slugs were fired in an
oxidizing atmosphere to a temperature of 1500°C.
Densities of these samples were in the range
2.67 to 2.75 g/cm>. Density values for similar

100% BeO control samples were 2,02 to 2.04 g/cm3,

Specimens in the shape of bars approximately
5x 'll/4 X 3’/4 in. have been fabricated for additional
testing.

Table 1.5.2. Comparison of Densities at Various
Temperatures of a 100% BeO Body and a Body
=W%ith 94 wt % BeO, 5 wt % MgO, and 1 wt % B,C

No. BeO*  Mg0  Fe,0, B,C
1 100
2 94.0 5.0 1.0
3 94.0 5.0
4 94.0 1.0
5 98.0 1.0 1.0
6 98.0 1.0
7 98.0 1.0

*G. C. grade, Brush Beryllium Company.

Sintering Density (g/cm3)
Temperature
©C) 100% BeO 94% BeQ-5% MgO-1% B4C
1650 2.71 2.91
1700 2.79 2.88
1750 2.83 2.89
1800 2.84 2.88
1850 2.85 2.90
1900 2.85 2.86

ANP PROJECT PROGRESS REPORT

OXIDATION RESISTANCE OF
BORON-CONTAINING BERYLLIUM
OXIDE BODIES

R. L. Hamner

Screening tests were also run to determine the
oxidation resistance of hot-pressed boron-contain-
ing BeO bodies being considered for GE-ANPD
applications. The oxidation characteristics of
the mixtures tested are compared in Fig. 1.5.1.
The additives TaB, (9.35 wt %) and BN (2.29 wt %)
not only have poor oxidation resistance above
1000°C but also apparently have very volatile
oxidation products, Ta, O, and 8,0, which ac-
count for the comparatively large weight losses.

It was expected that the BeO-CrB, body would
show an over-all weight loss upon oxidizing, be-
cause the oxides of chromium, Cr203 and Cr03,
tend to be volatile. A green deposit adjacent to
the specimen indicated that a chromium oxidation
product slowly volatilized along with the B,0;
an over-all weight loss was detectable after about
125 hr at 1300°C.

The borides of Ti (2.31 wt %), Zr (5.22 wt %),
and Hf (9.25 wt %), had good oxidation resistance
when incorporated in dense BeQ, as evidenced
by appearance and over-all weight changes, with
the ZrB,-BeO and HfB,-BeO combinations being
slightly better than the TiB,-BeO combination. It
is not known why the HfB,-BeO compact lost
weight despite the known stability of the oxida-
tion product, HfO,.

The thermal stability of ZrB, in air during long-
term oxidation testing at 1300°C is illustrated in
Fig. 1.5.2. A specimen of pure BeO of comparable

ORNL-LR-DWG 33729

Dl [ TiB2tBe0 -
o
o s e S -y 18,1+ Be0
R~ . — CrB,+ Be0
[ © HiBy+BeO
2

5 -05 . : ‘ !
& [ ' i
v . o
g o | i o
5 o ;
= -1.5 ""‘—l‘b: N S - —_

| !
5 1000°C| & I|300°c\ [\L
Y oo L NG BN + BeO s

! ‘ ! T ———y
-25 | - s [ - R—
by ‘ 1
f
30 pt—t \JoB; * 220 ; N

0 25 50 75 10C 125 150 175 200 225 250 275 300
TIME (hr)

Fig. 1.5.1. Results of Oxidation Tests of Boron-Con-
taining BeO Bodies.

50

o) ORNL—;_R-DWG 33730
0.20 i
5 /”lo—’
&
= o6 L ZrB, +BeD 1
Z o P T :
g ¢ Pl o ]
3 P
5 008 7
I '
S .04 //‘ _
" i 1L
o ¢ . re—— [, BeO
e g
-004 ,
~0.08 [
o) 100 200 300 400 500 600 700
TIME (hr)

Fig. 1.5.Z Results of Oxidation Tests of BeO and
BeO + Zr82 in Air at 1300°C.

density was tested with the ZrB,-BeO compact to
give a more complete picture of weight changes
involved, since BeO is somewhat volatile in the
presence of water vaporat 1300°C, X-ray and
chemical analyses showed the boron to be present
after the long-term oxidation test. There was not
sufficient ZrO, present as an oxidation product to
be detectable by x-ray analysis,

Information was received during this reporting
period that the Brush Beryllium Company had dis-
continued the manufacture? of ““Luckey S. P.”
grade BeO, which has been used aimost exclusively
in this laboratory for BeQ development work. |t
was, therefore, necessary to establish fabrication
characteristics of a new material representative
of future supplies; Brush's G. C. grade BeO, which
is derived from the sulphate and calcined at ap-
proximately 1100°C, was received for this purpose.

Specimens of G. C. grade BeO and BeO contain-
ing 5.22 wt % of laboratory synthesized ZrB, were
hot pressed at temperatures of 1550, 1650, and
1750°C under a pressure of 2000 psi for 15 min,
The G. C. grade of BeQ was found to be much
more reactive than the Luckey S. P. grade; notice-
able compaction began at approximately 250 to
300°C lower (at about 1250°C). This difference
in reactivity of the two materials can be explained
by the difference in grain size, as shown by elec-
tron photomicrographs, Figs, 1.5.3 (S, P. grade)
and 1,5.4 (G. C. grade).

Densities of pure BeQ obtained at 1550, 1650,
and 1750°C were 92, 95, and 98%, respectively.
Densities obtained for the BeO-ZrB,, mixture were
approximately the same for all the temperatures,

2 . . .
Private communication to R, L. Hamner.
Fig. 1.5.3. Brush Beryllium Company's S. P. Grade
BeO. 23,000X. Reduced 57%.

being 96.7, 96.4, and 96.4%, respectively; thus
ZrB,, appears to be somewhat effective as a densi-
fier at the lower temperature. Above 1750°C a
rather vigorous reaction occurred between the
ZrB,-BeO mixture and the graphite die, the attack
being severe enough to prevent removal of the
piece from the die.

Attempts were made to fabricate large hot-
pressed blocks of the ZrB-BeO mixture, 3 x 3 x
1 in., for physical and mechanical properties meas-
urements. A commercial grade of ZrB,, was used
because of the large quantities of boride involved.
Upon cutting the blocks, which were pressed at
1550°C and 2000 psi, an irregularity in color was
noted that was roughly in the shape of a “bow-
tie’’ and which approximated the temperature profile
of the piece in the graphite die. The top and center

PERIOD ENDING SEPTEMBER 30, 1958

W

s b
1

Fig. 1.5.4. Brush Beryllium Company's G. C. Grade
BeO.

specimens shown in Fig. 1.5.5 are typical of this
defect. It was not eliminated by time-temperature
variations, although it was less pronounced at
lower temperatures.

Analysis of the ZrB, revealed a large quantity
of iron (approximately 10%). When iron-free ZrB,
was used, either laboratory synthesized or an
acid-treated commercial grade, the appearance was
uniform, as illustrated by the bottom specimen
shown in Fig. 1.5.5. Six ZrB,-BeO blocks are
being fabricated at ORNL for testing at GE-ANPD.
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
PHOTO 31988

ARananaml T \\\-\‘\w\3
2

° \NCHES

Fig. 1.5.5. Hot-Pressed BeO-ZrB, Blocks. Top block and center blocks made with commercial ZrB,. Bottom
block made with iron-free laboratory synthesized ZrB,. SulSl with caption)

52

PERIOD ENDING SEPTEMBER 30, 1958

1.6. NONDESTRUCTIVE TESTING
J. W. Allen

REMOTE X-RAY VIEWING
J. W. Allen R. W. McClung

Further studies were made of the use of x-ray
sensitive Vidicon' in a closed-circuit television
system for remote viewing of x-ray images. As
reported previously, the sensitivity of the present
selenium photoconductor Vidicon system is con-
siderably less than that which can be obtained
with existing film techniques. The primary ad-
vantages of this system (in contrast to those con-
sisting of closed-circuit television applied to
fluoroscopy) are the inherent high resolution,
which approaches the resolution attainable with
a fine-grained radiographic film, and the inherent
magnification of images.

Recently it was discovered that the sensitivity
of the selenium photoconductor to x-radiation could
be improved in many situations by admitting visible
light to the photoconductor surface. Although the
phenomenon is not well understood, it appears that
the sensitivity is increased merely by raising the
level of total incident radiation, including both
x-radiation and visible light. This effect is readily
observable and has been used to increase the con-
trast sensitivity by as much as a factor of 3. The
wave length of the enhancing light does not seem
to be critical, as evidenced by the use of both
white fluorescent lamps and incandescent lamps
as the light source.

By using this light-enhancing technique, thick-
ness changes of about 2,5% can be detected in
s/a-in.-fhick aluminum. This indicates that practical
inspections could be accomplished at contrast
sensitivities of 5 to 6%. Although this figure is
considerably lower than that attainable with film
techniques, it compares favorably with those pub-
lished for other television systems which use
fluorescent screens.

Further investigation of the light-sensitizing
technique is planned. In addition the use of photo-
conductor materials other than selenium is being
studied.

DUPLEX TUBING
R. W. McClung

The increased interest in the fabrication and use
of duplex tubing has created a need for adequate

inspection techniques for this configuration. The
problems which are common to the inspection of
conventional tubing, such as dimensional gaging
and the detection of cracks, seams, laps, and other
discontinuities, are also present in duplex tubing.
In addition, an evaluation must be made of the

bond quality between the layers, and measurements

should be made of the individual layer thicknesses.

To detect the more common discontinuities, the
conventional techniques? of encircling-coil eddy-
currents, pulse-echo vltrasonics, radiography,
fluorescent penetrants, and visual inspection
should be adequate for complete evaluation. Also
some of these techniques may provide methods for
determining some of the properties peculiar to
duplex tubing. The use of the pulse-echo ultra-
sonic technique has detected the presence of lami-
nations in welded and drawn tubing.3 The detection
of such laminations is probably dependent upon
the size and location of the lamination, the ultra-
sonic frequency, and the dimensions of the tube.
At present, not enough is known about the mecha-
nism of detection of these laminations to assure
a high level of confidence in their detection under
all conditions. Many of the areas of nonbonding
will resemble simple laminations, and, under
proper conditions, may be located by the pulse-
echo ultrasonic technique.

If a gross separation exists between layers,
radiography may offer a means for detection, either
as a function of the decreased metal thickness due
to the voids when passing the radiation perpen-
dicularly through the lack-of-bond, or by observing
the poor bond in profile with the radiation passing
tangentially through the separation between layers.
However, as the separation between layers de-
creases, the confidence level of such an examina-
tion decreases rapidly.

The resonance ultrasonic technique is being
evaluated as a means for the detection of non-
bonded areas. This technique has been used in

'R. B. Oliver and J. W. Allen, ANP Quar. Prog. Rep.

2R, B. Oliver et al., ANP Quar. Prog. Rep. Dec. 31,
1956, ORNL-2221, p 260.

3R. B. Oliver, R. W. McClung, and J. K. White, Am.
Soc. Testing Materials, Spec. Tech, Publ. 223, 62-79
(1957).

53
ANP PROJECT PROGRESS REPORT

the past to measure wall thickness. On duplex
tubing with a well-bonded interface, the total wall
thickness of both layers would be indicated as a
function of a fundamental or harmonic resonance of
an ultrasonic frequency. Under proper conditions,
if a lack of bonding exists, this indicated thick-
ness would be that of the clad or outside layer
only. However, to achieve this detection of non-
bonding, it is necessary to make a careful selec-
tion of frequency range and calibration to avoid
confusion of indications.

A few small batches of duplex tubing have been
examined to determine the bond quality. These
include approximately 14 ft of 0.504 x 0.042 in.
tubing and 24 ft of 0.375 x 0.036 in. tubing. These
lots were fabricated by cladding about 0.025 in. of
low-carbon steel on types 304 and 347 stainless
steel, respectively. Radiography, resonance ultra-
sonics, and pulse-echo ultrasonic tests revealed
the presence of gross discontinuities in the
0.504-in.-dia tubing. Metallographic sectioning of
several typical defects revealed considerable lack
of bonding, with gross voids and separation pre-
valent. The worst bond condition detected is
shown in Fig. 1.6.1. The transverse cracks seen

in Fig. 1.6.2 are probably responsible for the

many indications noted in the pulse-echo ultra-
sonic examination. The same ultrasonic and
radiographic techniques as those used on the
0.504-in.-dia tubing, when applied on the 0.375-in.-
dia tubing indicated the presence of a good bond,
similar to that shown in Fig. 1.6.3. The resonance
ultrasonic technique seems to provide an evalua-
tion for bond quality which is the best and most
consistent of any of the current techniques. Addi-
tional development effort is needed and will be
made to provide a high level of confidence for the
complete evaluation of duplex tubing.

METAL IDENTIFICATION METER
R. A. Nance J. W. Allen
A transistorized model of the Metal Identification
Meter? (MIM) has been developed. Although the
internal operation has been altered, the new instru-
ment also utilizes eddy-current methods for identi-
fication and retains the simple external operation

“R. B. Oliver, J. W. Allen, and R. A. Nance, ANP
Quar. Prog. Rep. June 30, 1957, ORNL-2340, p 256.

UNCLASSIFIED
26826

Fig. 1.6.1. Very Poor Bond Between Type 1030 Steel Cladding and Type 304 Stainless Steel in }-in.-OD, 0,042-
in.-Wall Tube.

54
PERIOD ENDING SEPTEMBER 30, 1958

Fig. 1.6.2. Transverse Cracks at Bond Interface Between Type 1030 Steel Cladding and Type 304 Stainless Steel
in ‘/Z-rn.-on, 0.042-in.-Wall Tube.

4 S

INCHES

Fig. 1.6.3. Good Bond Between Type 1008 Steel Cladding and Type 347 Stainless Steel in %-in.-OD, 0.036sin.-
Wall Tube.

55
ANP PROJECT PROGRESS REPORT

of the vacuum-tube model. The MIM-Mark 1I, shown
in Fig. 1.6.4, is 7% x 3% x 2% in., and it weighs
approximately 20 Ib, representing an 80% reduction
in size and weight.

In developing the transistorized model, several
changes were made which greatly increased its
versatility. These are: (1) the instrument has
been made truly portable by the incorporation of a
battery power supply, (2) the range of metals identi-
fiable has been expanded to include all ferromag-
netic and nonferromagnetic metals, (3) the effec-
tive depth of penetration of the eddy-currents has
been decreased by increasing the operating fre-
quency and thereby reducing the minimum metal
thickness necessary for accurate identification,
and (4) the diameter of the probe coil has been de-
creased to reduce the error produced when a flat
reference standard is used in conjunction with the
identification of a specimen having a curved sur-
face.

A block diagram illustrating the internal opera-
tion of this instrument is shown in Fig. 1.6.5. The
frequency of the probe oscillator is determined by

the inductance of a probe coil, which, in tum, is
determined by the conductivity and permeability of
the specimen. The frequency of the tuning oscilla-
tor is controlled by the dial on the front of the
instrument. The signals from these two oscillators
are heterodyned to produce a signai having a fre-
quency equal to the difference of the frequencies

UNCLASSIFIED
ORNL=LR-DWG 302938

REGIENCY i, L, ThansiTons
N ]
TUNING
ETERODYNE | (/=%)| AMPLIFIER
ILLATOR
\| ,'-0355(:.:“5 wl & MIXER |AND DETECTOR]
20k

12140
TRANSISTOR

PROBE
OSCILLATOR PROBE COIL

=200 320 ke 774 specmen

|
Dflmwmm

METER

Fig. 1.6.5. Block Diagrom Describing Operation of
Metal Identification Meter-Mark IL.

UNCLAsSIFIED
Y220

Fig. 1.6.4. Metal Identification Meter-Mark Il

of the two original signals, The designs of the
amplifier and detector are such that only signals
of a narrow band of frequencies, centered about
120 kc, will be amplified, detected, and presented
on the meter,

Since the eddy-current flow in a specimen is a
function of its conductivity, magnetic permeability,
and geometry, it is possible by controlling geom-
etry factors to utilize the conductivity or con-
ductivity and permeability characteristics of a
metal as identifying features. This is accomplished
by tuning the instrument for @ maximum meter de-
flection with the probe on a reference standard.
The difference frequency between the two oscil-
lators is then 120 kc. If the probe is then placed
on a specimen having a different conductivity or
permeability, the frequency of oscillation of the
probe oscillator will shift so that the difference
frequency between the two signals is no longer
120 kc and the meter will no longer indicate a
maximum. |f the difference between the metals
is great enough, the meter will cease to indicate
altogether,

The response of this instrument to changes in
conductivity in nonferromagnetic materials is
illustrated in Fig. 1.6.6. The dial settings 565,
545, and 495 represent the tuning dial settings
that were necessary to obtain maximum meter
deflection when the probe was placed on samples
of Hastelloy B, Inconel, and type 316 stainless
steel, respectively. It is evident that, if the
instrument is tuned with the probe on the Inconel
sample (dial setting 545), the instrument will not
respond when the probe is placed on the other
two samples. [t can also be seen in Fig. 1.6.6
that the slight variations of conductivity normally
found among samples of the same alloy will not ap-
preciably decrease the meterreading from a maximum.

UNCLASSIFIED
ORNL-LR-DWG 30634R

10 T T T | T T T
DIAL SETTING 565 ' '
A A LT
| ;
275 =—CIAL SETTING 545 "~ [ f—F -
2 [ ‘ / |
w DIAL SETTING 495 - -+t '
[1d 5 ‘ - Y U
o {
w
= !
w i
225 - / — \ L Y g
LN

1.2 1.4 16 1.8 2.0 2.2 2.4 2.6 2.8 3.0
CONDUCTIVITY {7 IACS)
Figo
Mark
Tuning Dial Setting.

1.6.6.
Il to Conductivity Changes as a Function of

Response of Metal Identification Metere

PERIOD ENDING SEPTEMBER 30, 1958

It is important to note that this instrument cannot
be used to separate two nonferromagnetic metals of
differing composition which have the same conduc-
tivity or overlapping ranges of conductivities. Ac-
curate identification can only be made when it is
known that none of the metals in the group to be
identified have the same or overlapping conductivi-
ties. A similar problem is encountered in identify-
ing ferromagnetic materials, since different com-
binations of conductivity and permeability can
produce the same instrument response.

Variations in the geometry of a specimen can
adversely affect its identification when a thick,
flat reference standard is used. The extent of the
error produced by the curvature variations on the
surface of a specimen is illustrated in Fig. 1.6.7.
For all practical purposes an item may be con-
sidered to present a flat surface to the probe if
its radius of curvature is greater than 1.5 in.
Variations in the thickness of the specimen will
produce an error in identification only if the speci-
men is thinner than the effective depth of the eddy-
current penetration, which varies with the con-
ductivity and permeability of the part and the
operating frequency of the probe oscillator.

The variations in meter reading caused by thick-
ness variations of a specimen can be used as a
measure of thickness. This has been satisfactorily
accomplished in the gaging of 3/4 x 0.05 in. type
304 stainless steel tubing with an accuracy of 2'/2%.

This instrument can also be used to separate
specimens from the same heat of an alloy which
have been subjected to different mechanical and
thermal treatments, since they produce changes in
the conductivity and in some cases the magnetic
permeability of the part. [n addition, this instru-
ment can be used to locate cracks normal to the
inspected surface, since they interrupt the eddy-
current flow and cause an apparent conductivity
change in the part.

UNCLASSIFIED
ORNL-LR-DWG 30295

50 Y
\
g a0} \ —‘ .
g £ 30 |
I
% T 820 . e - _
bal \ }
[ 10 — . ™ — ———t
N
O \
0 0.5 1.0 1.5 2.0 2.5 3.0 @
CYLINDER RADIUS {in.)
Fig. 1.6.7. Effect of Cylinder Radius on Error of

Metal Identification Meter-Mark Il.

57

ANP PROJECT PROGRESS REPORT

1.7. METALLOGRAPHY

R. J. Gray

RESULTS OF METALLOGRAPHIC EXAMINATION
OF SMALL SEMICIRCULAR FUSED-SALT
FUEL-TO-NaK HEAT EXCHANGER
OPERATED AT HIGH TEMPERATURES

R. J. Gray J. E. VanCleve

A metallographic examination has been com-
pleted of a 20-tube semicircular heat exchanger
fabricated of Inconel in which a fused salt flow-
ing around the tubes exchanged heat to NaK flow-
ing in the tubes. Details of the construction and
operation of this Inconel heat exchanger were pre-
sented previously.'*? Testing of this heat ex-
changer was terminated after 1139 hr of high-
temperature operation at 1200°F and above because
of a leak of NaK into the fused salt. At the time
of termination the heat exchanger had operated at
design conditions for 258 hr with a temperature
differential from fused-salt inlet to outlet of 350°F
(1600 to 1250°F) and a temperature differential
from the NaK inlet to outlet of 430°F (1070 to
1500°F). There had been 142 hr of operation under
transitional conditions and 739 hr of isothermal
operation. The heat exchanger had experienced
194 thermal cycles between isothermal operation
at 1200°F and operation at design conditions.

The general area of the failure was readily
located after the shell was removed because of a
change in the color of the fluoride salt from green
to gray in the NaK outlet header region. The gray
material ignited during sectioning of the header,
and the presence of NaK mixed with the fluoride
salt was thus evident.

Two of three fractured Inconel tubes found in
the high-temperature header area are shown in
Fig. 1.7.1. The cracks were radial and were in
the so-called ““tension” side of the tubes. The
cracks were found in the tubes with shortest tube
befld-'o-?ube sheet length. This length ranged
from 1) % to 3/ in. The side of the tube that cracked
was in fms:on because of the temperature differ-
ence between the tube and the shell.

1). C. Amos et al., ANP Quar. Prog. Rep. March 31,
1957, GRNL- 22740 o 43,

2y C. Amos, R. L. Senn, and D. R. Ward, ANP Quar.
Prog. Rep. Szpl 30, 1957, ORNL-2387, p 38 and esp
Fig. 1.2.7,

58

The typical condition of the tube wall in all
three ruptured tubes is shown in Fig. 1.7.2. The
tube wall shifted laterally 0.002 in. after the
largest crack completely penetrated the wall. The
other large crack completely penetrated the tube
wall but it did not open as did the first. A third
crack may be seen that propagated from the braze
alloy surface. This is the only instance of a small
crack penetrating the material, and it suggests the

UNCLASSIFIED
3 823

0.101n./DIV
| |

Fig. 1.7.1. Cracks in Two of Three Cracked Inconel
Tubes Found in the High-Temperature Header Area of o
Fused

Cracks are in the tension side of the tubes.

Semicircular Salt-tosNaK Heat Exchonger.

Fig. 1.7.2 Typical Cracks Found in Tension Side of a Tube. Etchant:

possibility that a mechanism other than the progres-
sion of grain-boundary voids had operated to aid in
the fracture of the tubes.

The largest of the cracks shown in Fig. 1.7.2 is
shown at a higher magnification in Fig. 1.7.3. The
surface of the tube which was exposed to the fused
salt was generally corroded and leached to a depth
of 0.007 in. The grain-boundary voids which were
found in large numbers in similar areas of other
heat exchangers of this type were not present in
this unit,

The smaller of the large cracks shown in Fig.

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
Y-26166

10% oxalic acid. 100X.

1.7.2 is shown in Fig. 1.7.4 ot a higher magnifica-
tion. The crack follows the grain boundaries, and
the greater width of the crack at the inner, or NaK
contacted, surface indicates that it originated at
the inner surface. A few grain-boundary voids may
be seen that extend from the penetration. The
grain-boundary voids do not seem to be as numer-
ous or large as those found in previously examined
heat exchanger tubes. The maximum general cor-
rosion found in the heat exchanger is shown in
Fig. 1.7.5. On the fused-salt side the attack is

to a depth of 0.006 in. and on the NaK side the
attack is to a depth of 0.002 in.

59

ANP PROJECT PROGRESS REPORT

-{\: UNCLASSIFIED
SN Y-26165

Fig. 1.7.3. The Larger of the Cracks Shown in Fig. 1.7.1 at a Higher Magnification. General corrosion and
leaching to a depth of 0.007 in. may be seen on the surface that was exposed to the fluoride salt. Etchant: 10% =
oxalic acid. 200X.

UNCLASSIFIED |2 :

Fig. 1.7.4. The Smaller of the Large Cracks Shown in Fig. 1.7.2 at o Higher Magnification. Complete pene- 3
tration of the tube wall may be seen. Small grain-boundary voids subtend the large crack. Etchant: 10% oxalic acid.

200X.

60
Fig. 1.7.5. Maximum General Corrosion:
* oxalic acid. 200X.

0.006 in. on Salt Side and 0,002 in.

on NaK Side. Etchant: 10%
Part 2
CHEMISTRY AND RADIATION DAMAGE
2.1. MATERIALS CHEMISTRY

W. R. Grimes
Chemistry Division

PREPARATION OF CHARGE MATERIAL FOR
REDUCTION TO YTTRIUM

G. J. Nessle

A pilot plant has been constructed that will
provide for the monthly production of about 200 Ib
of massive ytrium metal. The Chemistry Division
is responsible for the preparation of the suitable
fluoride mixture, either LiF-Mng-YF3 or LiF-YF,,
and the Metallurgy Division will carry out the re-
duction, distillation, and sparge-melting steps that
will yield the final product. The pilot plant was
completed during August, and the first batch of
fluoride-salt mixture (LiF-YF,) has been prepared.
Phase equilibrium studies of the LiF-YF, system,
as reported below, have indicated a single eutectic
with 19 mole % YF , that has a melting point of
682°C. Therefore an attempt will be made to
prepare low-oxide-content yttrium from this binary
mixture and thus eliminate the Mg-Y alloying step
and subsequent distillation of Mg from the alloy
that would be required if the LiF-MgF ,-YF_ mixture
were used, If this attempt is unsuccessful, it will
be necessary to revert to the ternary system. The
metal reduction operation is described in Chap. 1.1
of this report.

Conversion of Y,0, to YF,
J. Truitt
About 400 Ib of YF, was produced by direct

hydrotluorination of dry Y,0, in a large-scale
(250-1b capacity) converter. In addition, about

100 Ib of YF, was produced in a small, pilot-scale
unit (10 Ib per batch). These units are operated at
approximately 1100°F, and the raw material is
treated for about 10 hr with HF. Analyses of the
batches produced in the pilot-scale unit are given

in Table 2. 1.1,

Preparation of MgF,
J. Truitt

No satisfactory grade of MgF . was available
commercially, and therefore six 10-lb batches of
MgF , were prepared by dry hydrofluorination of
MgO. Previous results had shown that by extended
treatment at 600°C a product of better than 99.5%

could be obtained. However, extended hydro-
fluorination is not economically desirable if, as

was later shown, 95% assay material is satisfactory
for charging to the final molten salt purification
step. The range of product analysis and conversion
efficiency of the six batches is given in Table 2,1.2.
It is now possible to purchase a satisfactory grade
of MgF, and sufficient material has been ordered

for several months of operation.

Preparation of the LiF-MgF,-YF ; Mixture
J. Truitt J. E. Eorgan

Several small-scale batches of the ternary mixture
LiF-MgF ,-YF, were purified for yttrium-metal pro-
duction. Although it was known that the raw
materials contained oxides to some extent, the
standard 3-hr hydrofluorination time used in other

Table 2.1.1. Analyses of Batches of YF3 Obtained
by the Conversion of Y203 with HF in a
Pilot-Scale Unit

Minimum Maximum Mean
(%) (%) (%)
Analytical results*
Yfl'rium 59.7 6]-3 60-7
Fluorine 38.1 39.0 38.7
Oxygen 0.09 1.5 0.7
Conversion efficiency* 97.2 100.0 99.0

*Individual analyses for Y and F and conversion effi-
ciencies are not necessarily from the same batch.

Table 2.1.2. Analyses and Conversion Efficiencies
Obtained in the Batch Conversion of Mg0 to MgF2

Minimum Maximum Mean
(%) (%) (%)
Analytical results*
Magnesium 39.0 40.0 39.5
Fluol’ine 59.4 60.7 60.3
Conversion efficiency* 95.0 99.6 97.7

*Individual analyses for Mg and F and conversion effi-
ciencies are not necessarily from the same batch,

65
ANP PROJECT PROGRESS REPORT

salt purifications was used on the ternary mixture
in order to test the new method of oxide deter-
mination in fluoride salts and to determine the
reproducibility of oxide removal.

The analytical results obtained for ten batches
are presented in Table 2.1.3. Petrographic and
x-ray diffraction examination of these mixtures
did not detect oxides. As can be seen from the
data, however, the chemical analyses indicated
approximately 1000-ppm oxygen in the salt, There
is some reason to believe that the oxide analyses
gave high results because of sampling techniques,
but, on the basis of these analyses, the hydro-
fluorination time will be increased gradually until
a product purity of 500-ppm oxygen or less can be
assured.

Table 2.1.3. Analyses of Ternory Mixtures of
LiF-Mu_:ng-YF3

Analytical Results*

Constituent Minimum Moximum

Found Found Mean
Y 30.6% 33.5% 32.3%
Mg 8.70% 9.23% 8.95%
Ni 40 ppm 145 ppm 75 ppm
Cr 30 ppm 200 ppm 100 ppm
Fe 80 ppm 200 ppm 150 ppm
0, 1000 ppm 1500 ppm 1100 ppm

*Individual analyses are not necessarily from the

same batch.

Extending the hydrofluorination will have a
marked effect on the processing time per batch of
mixed salt because there will be increased dis-
solution of the reactor vessel, usually nickel, and
of the accumulated impurities left from previous
batches. Therefore, care must be taken to use only
the HF needed in order to prevent the processing
time (or cost per batch) from becoming unrea-
sonable, if possible,

PHASE EQUILIBRIA IN THE SYSTEM LiF-YF3
R. E. Thoma

Determinations of the character of the phase
equilibria occurring in the system LiF-YF have
been deterred by the presence of considerable

66

amounts of a phase which is probably an oxyfluoride
of yttrium, This phase appears as a biaxial
negative material with an optic angle of 10 deg,

and it has refractive indices just above and below

1.700.

All samples of YF, used previously had been

prepared by solid-state hydrofluorination of Y,0,.

R

ecently a LiF-YF, batch was hydrofluorinated in

the liquid state in order to minimize the content of
the oxyfluoride., This batch was subsequently
analyzed chemically and used as the beginning

m

aterial for several thermal-gradient quenching

experiments, The quenched samples from these
experiments were apparently free of oxygen-
containing phases. The quench results are thus
believed to be more nearly representative of

L

iF-YF, equilibria than any results previously

available, A portion of the phase diagram of the
system has been constructed from these results
and is shown in Fig. 2.1.1. A single eutectic
occurs in the system at 19 mole % YF,; mp,
682°C. This melting temperature is lower than
any of the values previously reported, ! and it

is believed to be the most accurate value yet
attained.,

'R. E. Thoma, ANP Quar. Prog. Rep. March 31, 1958,

ORNL-2517, p 54.

TEMPERATURE (°C)

UNCLASSIFIED

ORNL-LR-DWG 33937
900

LIQUID

800 \
LiF + LIQUID
YLIF-XYFy
+
LIQUID
700 3
o ;[ o -
LiF + YLiF - XYFy
600
0 10 20 30

YF3 (mole %)

Fig. 2.1.1. A Portion of the System LiF-YF, (0~30

Mole % YF3)0
EXTRACTION OF LITHIUM METAL IMPURITIES
WITH MOLTEN SALTS

G. M. Watson J. H. Shaffer

A tentative procedure for purifying lithium metal
by extracting oxide, nitride, and carbide impurities
with a eutectic mixture of lithium halide salts, such
as LiF-LiBr (12-88 mole %; mp, 453°C) or LiF-LiCl
(20-80 mole %; mp, 485°C), was described pre-
viously.? The effectiveness of the method depends
on the differences in solubilities of impurities in
the salt and metal phases. Indirect measurements
in this laboratory indicate that the solubility of
Li,O is relatively high (over 1 mole %) in molten
salts. The solubility of the oxide in liquid lithium
is not known, however, because of difficulties en-
countered in sampling and analysis, even though
some progress has been recently reported? in the
determination of oxygen in lithium,

It is expected but not necessarily assured, how-
ever, that the differences in the solubilities of the
impurities in the salt and metal phases are
favorable toward their accumulation in the salt
phase and that a suitable extraction method for
lithium purification may be developed.

Five experiments were performed on two different
batches of lithium in attempts to gain familiarity
with the procedure and to determine whether a cor-
relation could be obtained between the number of
extractions on a given batch of lithium and the
amounts of impurities extracted. The results of
these experiments are summarized in Table 2.1.4.

No particularly good correlation has been obtained
to date between the number of extractions performed
on a given batch of lithium and the concentration of

2G. M. Watson and J. H. Shaffer, ANP Quar. Prog.
Rep., Marchk 31, 1958, ORNL-2517, p 55.

3N. I. Sax and H. Steinmetz, Determination of Oxygen
in Lithium Metal, ORNL-2570 (Oct. 15, 1958).

Table 2,1.4.

PERIOD ENDING SEPTEMBER 30, 1958

impurities extracted. |t is true that the second ex-
traction in each of the series shows a considerable
lowering in concentration of the impurities extracted
as compared with the initial concentration; however,
the concentration of 50 ppm shown is probably in
error, since in this experiment a considerable
amount of bismuth-lithium alloy was trapped in

the salt. Substantial corrections determined by
measuring the hydrogen evolved from the salt when
treated with water had to be applied. Thus the
value of 50 ppm was obtained as the net result

of the difference between two relatively large
numbers. Accordingly this result can be dis-
counted for the present. The result of the third
extraction of batch number 1 appears to be

totally due to nitrogen impurities, as given by
chemical analysis of the salt. It is not immediately
apparent why no oxygen was found (by difference)
unless the analytical results for the nitrogen
determination are erroneous.

For future experiments greater care will be
exercised in minimizing contamination in the
experimental assembly. Also the lithium metal
will be heavily contaminated with known amounts
of oxide, introduced as CuQ, in order to attempt to
determine on a much firmer basis whether the
partition coefficient for the oxide extraction is
favorable for the accumulation of impurities in
the salt phase.

Recent improvements in analytical techniques
have made it feasible to analyze lithium samples
before and after the salt extraction, and it is no
longer necessary to depend entirely on the complex
procedures required for analysis of the salt phase.
The liquid lithium samples will be obtained as
emulsions of lithium in paraffin.

4

4A, S. Meyer, Jr., and R. E. Feathers, ANP Quar. Prog.
Rep. March 31, 1958, ORNL-2517, p 57.

E xtraction of Impurities from Liquid Lithium by Molten LiF-LiX Eutectic Mixtures

Total Impurities

Batch Extraction Salt Extracted (Calculated Nitrides
No. Number Mixture as ppm 0) {as ppm N)
1 1 LiF-LiBr 400 Not determined
2 LiF-LiBr 50 <50
3 LiF-LiBr 484 500
2 1 LiF-LiCl 2860 <20
2 LiF-LiCl 467 <20

ANP PROJECT PROGRESS REPORT

2.2. ANALYTICAL CHEMISTRY

J. C. White
Analytical Chemistry Division

DETERMINATION OF OXYGEN IN
FLUORIDE SALTS

A. S. Meyer, Jr. G. Goldberg

The method for the determination of oxygen in
fluoride salts by fluorination with KBrF ,, as de-
scribed in a previous report, was applied success-
fully to determinations of oxygen in the mixed salt
LiF-MgF -YF ;. Seven salt mixtures were analyzed
during this period. The results for duplicate
samples are presented below:

Sample No. Oxygen Content (%)
Al 0.11, 0.11
A2 0.21, 0.22
A3 0.11, 0.10
A4 0.11, 0.10
AS 0.14, 0.15
Aé 0.24, 0.26
A7 0.11, 0.13

The coefficient of variation of the results is 6%.
Since Sheft, Martin, and Katz? had reported that
KBrF, was too basic to react with sufficient satis-

faction with an oxide as basic as MgO and had
recommended the use of an acidic addition com-
pound, SbF BrF2, for such oxides, the validity of
the results ?or oxygen in LiF-MgF ,-YF , was cor-
roborated by analyzing standard samples of MgO and
Y,0,. The recovery of oxygen from these oxides
exceeded 98% with excellent precision (coefficient
of variation, 2%). Highly refractory MgO, however,
did not react appreciably with KBrF .. The method
described is now being used for routine determi-
nations of oxygen in the mixed salt and also to
determine the extent of conversion of oxides of
magnesium and yttrium to the fluorides.

TA. S, Meyer, Jr., and G. Goldberg, ANP Quar. Prog.
Rep. Sept, 30, 1957, ORNL-2387, p 150.

2|, Sheft, A. F. Martin, and J. J. Katz, J. Am. Chem.
Soc. 78, 1557 (1956).

68

DETERMINATION OF YTTRIUM AND
MAGNESIUM IN LiF-Mng-YF3

J. P. Young R. F. Apple

A potentiometric method for the determination of
both yttrium and magnesium in LiF-MgF ,-YF . was
developed and evaluated. Yttrium and magnesium
form stable complexes with ethylenediaminetetra-
acetic acid (EDTA) and can be determined by direct
titration by applying the technique of Reilley and
Schmid,® who use a modified mercury electrode. The
electrode system is essentially an amalgamated
gold wire as the indicator electrode and calomel as
the reference electrode.

Both yttrium and magnesium are determined in the
same sample because of the difference of the
stability constants of their respective EDTA com-
plexes. Yttrium, the more stable complex, is
determined at a pH of 4.0 £ 0.5. Then the pH of the
sample solution is increased to 9.7 £ 0.3, and
magnesium, the less stable complex, is determined.
An example of the resultant titration curve is shown
in Fig. 2.2.1. When titrating yttrium and magnesium
with 0.01 M EDTA, at least 5 mg of yttrium and at
least 1 mg of magnesium should be present in order
to obtain a satisfactory potential break at the end
point. Although macro amounts of iron, nickel, and
chromium would definitely interfere with these
determinations, the method is quite insensitive to
the micro amounts of these corrosion products that
would be expected in routine samples of LiF-

MgF -YF .

Up to approximately 400 ug of fluoride ion can be
tolerated in the sample solution without affecting
either determination. The titrations can be per-
formed in either a nitrate or sulfate medium; large
concentrations of chloride ion interfere with the
response of the mercury electrode. Typical results
for the determination of yttrium and magnesium in
synthetic standard sample solutions which con-
tained nitric or sulfuric acid are given in Table 2.2.1.

3C. N. Reilley and R. W. Schmid, Anal. Chem. 30,
947 (1958).

The precision of the method is about 1%. The
results of these determinations are accurate to with-
in 2%. This method is considerably less time-
consuming than the gravimetric methods that were
previously used for the determination of these
metals. The procedure has also been applied to
the determination of yttrium in yttrium oxide and in
yttrium fluoride.

UNCLASSIFIED
ORNL-LR-DWG 33938

R C T T 1T

Mg CONTENT: 2.18 mg ,/
pH: 9.7 £ 0.3

[

-Sva I
i [

|

POTENTIAL (WITH REFERENCE TO
A SATURATED CALOMEL ELECTRODE)

6 8 10 12 19 16 18 20 22 24 26
0.01# EDTA (mi)

Fig. 22,1, Results of Potentiometric Titration of
Yttrium and Magnesium with Ethylenediominetetraacetic

Acid.

PERIOD ENDING SEPTEMBER 30, 19

Table 2.2.1. Typical Results for the Determination
of Yttrium and Magnesium in Solutions Which
Contain Yttrium, Magnesium, and Lithium by

Medns of a Potentiometric Titration with

Ethylenediaminetetraacetic Acid

58

Yttrium (mg) Magnesium (mg)

Present Found Present Found

With 1 M Sulfuric Acid Present

7.24 7.35 1.45 1.44
7.30 1.44
10.9 10.9 2.18 2.16
11.1 2.17
14.4 14.6 2.90 2.84
14.4 2.89

With 0.2 M Nitric Acid Present

7.50 7.58 1.45 1.49
7.58 1.49
11.3 11.4 2.18 2,15
11.3 2.16
15.0 15.3 2.90 2.84
15.2 2.87

69
ANP PROJECT PROGRESS REPORT

2.3. RADIATION DAMAGE

G. W. Keilholtz
Solid State Division

ETR IRRADIATION OF MODERATOR MATERIALS
FOR USE AT HIGH TEMPERATURES

W. E. Browning, Jr.
R. P. Shields J. E. Lee, Jr.

Experimental apparatus is being prepared with
which to test the stability of yttrium hydride and
beryllium oxide at temperatures up to 1800 in
the high-level irradiation environment of the Engi-
neering Test Reactor (ETR), where the gamma-ray
heating is approximately 25 w/g. Temperature
gradients will be induced in the specimens, and the
reactor programing will provide thermal cycling.

The yttrium hydride samples will be checked for
changes in crystal structure, as well as for the
effects of stresses resulting from temperature
gradients.

Before the samples of the moderator materials are
irradiated, x-ray examinations will be made of the
yttrium hydride samples in order to determine the
crystal structures, the modulus of elasticity and
shear modulus of each sample will be measured,
marked regions on the samples will be photographed,
and exact measurements of the pieces will be made.
The radioactivities of the samples should be low
enough for the examinations and measurements to be
repeated after irradiation.

The heat generated in the irradiated capsules will
be conducted across a small annulus by a helium-
argon or a helium-nitrogen mixture which will be
essentially static as far as heat transfer is con-
cerned. There will be just enough flow to permit
changing the mixture in order to maintain a constant
temperature at the outside surface of a capsule.
Helium has about four times the heat transfer
capacity of argon and will make up the largest part
of the gas mixture.

Since the intensity of gamma-ray heating in the
reactor is uncertain it is necessary to provide means
for raising the experimental equipment to a region of
lower gamma-ray intensity if the heating is found to
be too great. In order to accommodate the necessary
raising devices, the apparatus for the initial experi-
ment will include only one yttrium hydride and one
beryllium oxide capsule.

The beryllium oxide sample will contain three
right cylinders 0.636 in. in diameter and 1 in. in

70

height. These will be placed end to end in a type
410 stainless steel capsule. The yttrium hydride
sample is similar, but the diameter of each cylinder
is 0.800 in. and the cylinders will be enclosed in
type 430 stainless steel. The yttrium hydride
cylinders and the steel capsule prior to assembly
are shown in Fig. 2.3.1. As may be seen, one
cylinder, which will be at the top of the assembly,
has a weil for a thermocouple with which to monitor
the central temperature. The top beryllium oxide
cylinder temperature will be monitored similarly.

The assembled capsules and the tubes which pro-
vide the gas annuli are shown in Fig. 2.3.2.
Thermocouples will be attached to the sides of the
capsules midway between the ends. In order to pass
these thermocouples through the thin gas annulus,
sheathed, swaged magnesia thermocouple leads will
be silver soldered into the small metal plugs shown
in Fig. 2.3.2. The thermocouples will be spot-
welded to the capsule. Flexibility in the thermo-
couple wires where they pass through the gas
annulus will be provided by helical coils.

After the outer tube assembly is completed, it will
be encased in an aluminum spacer that will provide
a ]/] -in. water annulus and will also anchor the
sheathed thermocouple leads and gas tubes. The
spacer unit will then be fitted into an aluminum tube
which will be attached to a 1-in.-dia stainless steel
pipe. This pipe, which will extend to a point about
10 ft above the ETR lattice, will support the experi-
mental apparatus. All thermocouple leads and gas
lines will be inside the pipe and will leave the
reactor cell through a 1-in.-dia flexible steel tube
that will connect with the support tube and the
flange on the reactor. The steel cable attached to
the support pipe and a drum near the reactor flange
will provide the means for raising the apparatus to
a position of lower gamma-ray flux if the termperature
becomes too high.

The design work on the thermocouple installation
was the major developmental effort required for this
experiment. Each thermocouple must be capable of
operating at a high thermal stress in the gas annulus.
Also, the sheath which encloses the leads and is
sealed at the junction with the tube that forms the
gas annulus must be capable of withstanding the
high pressure of the reactor cooling water.
PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
PHOTO 44085

Fig. 2.3.1. Yttrium Hydride Cylinders and Steel Capsule Ready for Assembly and lrradiation in the ETR at High

Temperatures. (Secret with caption)

UNCLASSIFIED
PHOTO 44558

Fig. 2.3.2. Capsule Irradiation Assemblies Showing Tubes Which Form Gas Annuli.

A series of mockup devices was used in designing
a thermocouple arrangement and developing the
technique for installing the thermocouples, and the
operating conditions were simulated as nearly as
possible in the mockup tests. The mockup used to
test the thermocouples is shown in Fig. 2.3.3. The
heating element (bottom of Fig. 2.3.3) is a 10-mil-
wall Inconel tube attached to copper electrodes. At

900°F, approximately 2 kw of heat can be conducted
from this element across the 10-mil annulus to a
water jacket. The thermocouple wires are spot-
welded to the heating element, and the Inconel
sheath containing the wires extends through a small
cap which anchors the sheath to the water jacket.
Cycling the heating element from room temperature
to 900°C flexes the thermocouple wires; the life of

71

ANP PROJECT PROGRESS REPORT

Fig. 2
for High-Temperature Irradiation of Moderator Material s.

the thermocouple is a function of the number of such
cycles. It was found that annealing the wires after
removing the sheath and before bending them greatly
increased the life.

This mockup device also served to develop a
temperature control method. With argon flowing
through the annulus at a very low rate, a Moore
controller, activated by a thermocouple, introduces
helium into the argon in varying amounts to increase
or decrease the heat conductivity of the gas.
Varying the conductivity of the gas mixture allows
a constant temperature to be maintained on the out-
side of the capsule although considerable changes
in gamma-ray flux may take place.

72

UNCLASSIFIED
PHOTO 44814

. Mockup Device Used in Designing and Testing the Thermocouple Installation Used in the Apparatus

The compatibility of yttrium hydride and beryllium
oxide with various probable encapsulating materials
was also investigated. The materials tested were
Inconel, nickel, and types 446, 430, and 410 stai
less steel. An Inconel capsule containing yttrium
hydride that was heated for 200 hr at 900°C showed
no attack. However, when a second capsule was
tested at 1000°C, the Inconel reacted with the
hydride, and o large opening was formed in the
capsule in a matter of 10 to 15 hr. The stainless
steels mentioned were all satisfactory and the final

selection was made on the basis of most favorable
machining, welding, and heat properties. At the
present time the apparatus for the experiment and a

reactivity mockup required for the ETR critical
facility are being assembled.

CREEP AND STRESS RUPTURE TESTS
UNDER IRRADIATION

J. C. Wilson
W. E. Brundage W. W. Davis
N. E. Hinkle J. C. Zukas

The first MTR tube burst tests ' were carried out
in a helium atmosphere. The short times to rupture
were attributed to thermocouple errors arising from
contamination of the helium atmosphere by thermal
insulation. A second series of tests have now been
completed in an air atmosphere in the MTR, and the
resultant data are being evaluated. Equipment is
being constructed for operation of tube burst tests
in the lattice of the ORR and for carrying out creep
rate experiments on the pool side.

RADIATION EFFECTS IN ELECTRONIC
COMPONENTS

Wide-Range Multipurpose Cryostat
0. E. Schow

In conjunction with studies of the effects of radi-
ation on semiconductor barriers, a cryostat has been
constructed with a temperature range of ~200°C to
+350°C. Requirements, other than the temperature
range, were long-term reliability and ‘‘fail-safe’’
operation, that is, cooling water failure or electrical
failure should shut the cryostat down without
harmful effects to any of the system components.

The cryostat consists of concentrically nested
metal ‘*cans,’’ suspended from the top, as shown in
Fig. 2.3.4. The innermost container, the sample
chamber, is in two sections. The part which
actually holds the sample is a 4-in.-ID copper
cylinder having %-in. walls. The bottom of the
cylinder is c|ose3 with Y -in. copper plate recessed
2 in. from the end. This recessed bottom is neces-
sary to reduce the heat loss through the bottom of
the cylinder. The over-all length of the copper
section is 12 in., of which 10 in. is available for
use. The open end of this section is silver-soldered
to a 4-in.-ID stainless steel cylinder, 14 in. long,
having 0.019-in. walls. The stainless steel

11, C. Wilson et al., ANP Quat Prog. Rep. Dec. 31,
1957, ORNL-2440, p 207.

PERIOD ENDING SEPTEMBER 30, 1958

cylinder is, in turn, attached to a brass plug which
bolts to the suspension plate. The entire sample
chamber, along with its associated thermocouples
and heaters, is completely removable for servicing.
Around the sample chamber and separated from it by
a ]/2-in. heat-exchange-medium chamber is the liquid-
nitrogen cooling jacket. This jacket is filled
automatically and provides the heat sink for the
sample chamber. Around the cooling jacket is a
]]/2-in.-high vacuum chamber to isolate the system
from the ambient temperature. The brass suspension
plate is supported by the outermost container, a
I/B-in. wall, stainless steel cylinder, 30 in. high.

Each chamber, with the exception of the liquid-
nitrogen cooling jacket, is connected by means of
holes in the suspension plate to a gas-and-vacuum
manifold panel, from which it is possible to intro-
duce vacuum, any desired gas, or air into each
chamber. These operations may be performed
independently or simultaneously. The system is
monitored by gages with which it is possible to
determine pressures in the chambers from 2 atm to
10~% mm Hg. Also connected through this panel is
a mechanical roughing pump to reduce the system
pressure sufficiently to prevent damage to the
diffusion pump.

The diffusion pump for the system is a water-
cooled 300-liter/sec fractionating pump. Silicone
oil is used to prevent pump damage if the system
goes to atmospheric pressure while the pump is on.
This pump is suspended below the cryostat and
connected through the outer cylinder by means of a
4-in. ‘‘O’'-ring-sealed flanged joint.

The diffusion-pump cooling coil and the boiler coil
are connected to a monitoring and control panel.
Loss of water pressure, reduction of water flow, or
failure of electrical power will automatically initiate
an emergency shutdown. The pump cannot be started
again until corrective measures have been taken.

Temperature control of the cryostat is effected by
means of a Speed-O-Max ‘“‘PAT’’ temperature con-
troller. This controller uses, as reference, one of
seven thermocouples which have been embedded in
the copper section of the sample chamber and held
there by copper amalgam. Heat is applied to the
sample chamber by means of five heaters wound
around the chamber. The total electrical input to
this heater system is controlled by the controller,
but separate adjustments can be made to determine
what portion of the total electrical power goes to
each heater. Temperatures in the sample chamber

73

|74

UNCLASSIFIED
ORNL-LR-DWG 33939

r LIQUID NITRCGEN

-~m—HIGH VACUUM

\> STAINLESS STEEL

HIGH VACUUM TEST CHAMBER HEAT EXCHANGER
yd
l L)
I —< AIR DC HEATER SAMPLE CHAMBER
POWER
S N
NITROGEN TEMPERATURE
MANIFOLD PANEL CONTROLLER = :””
—— HELIUM
| [ ‘
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1
|
MECHANICAL | [
PUMP —
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/ 1 | I
| E
THERMO - | o =~
COUPLE DISCHARGE 1 | SR
VACUUM VACUUM [ | R
GAGE GAGE B —3 S
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COOLING PANEL - L %
ION GAGE | : : ,Ljf : R
. 2 I
L | £ i
o | ; §
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DIFFUSION AND
MECHANICAL PUMP
CONTROL PANEL

WATE RJ

e

SUPPLY

COPPER

POWER
INPUT

EMERGENCY
DIFFUSION PUMP
SHUTDOWRN
CONTROL PANEL

< COMPRESSED AIR

g

——=—+= TO DRAIN

)

7

Fig. 2.3.4. Wide-Range Multipurpose Cryostat.

MECHANICAL PUMP

LY0dI Y S53490¥d LIITFr0dd NV
can be maintained to within ]/4°C with a gradient

along the length of the copper section of the sample
chamber of ]/4°C

Grown-Junction Silicon Diode Irradiation

J. C. Pigg C. C. Robinson
0. E. Schow

An experiment has been performed to determine the
effect of fast-neutron radiation on a grown-junction
silicon diode. The facility used for the irradiation
was the same as that used for bombardment of a
point-contact silicon diode and an alloy-junction
silicon diode so that comparisons of the three types
of diodes would be possible. The sample studied in
this irradiation was 1.6 ¢m long, 5 mm wide, and
5 mm thick. The junction was about 8 mm from the
end of the p side. The sample was ground and
carefully etched. Nickel was plated to the ends,
and lead wires were attached. It was then placed
in a glass container which was evacuated, flushed
with helium, and filled to a pressure of ]/2 atm with
helium. The leads and thermocouple wires were
brought out through Kovar seals.

The irradiation was performed in hole 51 of the
ORNL Graphite Reactor. This hole is a fast-flux
facility with a fast flux of about 8.3 x 10!
neutrons/cm?.sec when the reactor is at 3500 kw.
The flux is proportional to the power level. The
facility is water-cooled to maintain a temperature of
about 23°C. Because the sample was not in
intimate contact with the side of the facility, its
temperature during the irradiation rose to 33°C. The
irradiation was begun with a reactor power of 100 kw,
which was raised to a maximum of 2100 kw. At this
power the fast flux was approximately 5 x 101!
neutrons/cm2.sec. To further assure bombardment
by fast neutrons only, the sample was wrapped in
several layers of 4.5-mil cadmium. A bias of 1 v
was applied in the forward and reverse directions
alternately, and the current was measured.

A rapid increase in forward current and a small
decrease in reverse current which occurred during
the very early part of the irradiation is illustrated
in Fig. 2.3.5, that is, up to about 5 x 1013
neutrons/cm?. Although not shown conclusively,
experience with previous irradiations indicates that
this may possibly be attributed to photo-emf's pro-
duced in both the sample and in the lead wire
insulation by the gamma-ray flux of the reactor.

From the dosage of about 1 x 10" to about
1 %103 neutrons/cm?, the changes in the reverse

PERIOD ENDING SEPTEMBER 30, 1958

current may be attributed to the increase in temper-
ature. |nasmuch as the temperature levels off after
1x10'3 neutrons/cm?, while the reverse current
continues to increase, this change in reverse current
is believed to be the result of radiation. The
irradiation was terminated at 1.8 x 10'® neutrons/cm?
because the ratio of forward to reverse current had
reached a magnitude which suggested other studies
of the sample.

A preliminary comparison of the results of this
irradiation with those of previous irradiations? show
a change in this sample on the order of a factor of
100 increase in reverse current as compared with an
increase of 200% for a comparable irradiation of an
alloy-junction diode and an increase of 30% in a
point-contact diode.

Low-Temperature Indium-Germanium Alloy

J. C. Pigg C. C. Robinson
0. E. Schow

During some preliminary work on the problem of
alloying indium into germanium for the purpose of
making electrical barriers, some heretofore un-
familiar phenomena were noted. This portion of the
work dealt with attempts to alloy at temperatures
which were not greater than 280°C. On a phase
diagram this temperature represents a concentration
of indium of about 5%. Alloying times were on the
order of 15 min to 1 hr, with additional 10-min
periods to bring the samples to temperature and
10-min periods to cool the samples to room temper-
ature. Inasmuch as no continuous time-temperature
recordings were taken, it would be difficult to
estimate the cooling rate of the sample as it passed
through various significant temperatures, that is,
the eutectic or the freezing point of the indium. |t
is believed, however, that the thermal inertia of the
heater would prevent extreme, 50°C/min or faster,
temperature changes. The alloying was carried out
in a helium atmosphere.

After the alloying cycle, the major part of the
indium was removed by careful cutting, and the
remainder of the sample was etched in nitric acid
only. This served to remove the indium without
significantly attacking the germanium. Microscopic
examination of the surface showed no major amount
of penetration of the indium into the germanium;

2J. C. Pigg and C. C. Robinson, Solid State Semiann.
Prog. Rep. Aug. 30, 1954, ORNL-1762, p 78.

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL-LR-DWG 33940

3 | 34
SAMPLE TEMPERATURE (°C)—|
el
—— » ] .—i_. a .—_.%
—al ® S—— — 32
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3 p
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1 ®
21— S
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a o
£ / "
S o—0 A o
= o
W 5
% _ — 26 E:_j
S [a
o so—o-0.0-* A m
~ -
wees A
®
R - 24
100 kw /
fl Va
z 500 1000 1400 2100 kw
> kw kw kw
Q v - REACTOR POWER LEVEL -
o1 /‘ —] 22
oJ
WA W /. AT {-vbias (x10 % amp)
A [, AT {-vbias (x10™% amp)
A ® SAMPLE TEMPERATURE (°C)
T‘:“ — 20
oA
0 1x10" 5x 10" 1x10'"° 1.5 x 102 2x10%

FAST-NEUTRON DOSE (neutrons/¢m?)

Fig. 2.3.5. Irradiation Effects on a Grown=Junction Silicon Diode.

indeed, most of the surface appeared to be un-
touched. (Microscopic examination of germanium
etched with nitric acid alone does show some attack,
but the action is slow and does not compare with

the action of nitric acid on the indium and is not
believed to detract from the phenomena observed.)

In the areas where penetration was noted, it was
seen to be, for the most part, penetration which
conformed to rather precise geometric configurations.
Some of these may be seen in Fig. 2.3.6. Geometric
crystal growths on the surface of the germanium may
also be seen. Other examples are shown in Fig.
2.3.7. The sample shown in Fig. 2.3.7 is not the
same as the one shown in Fig. 2.3.6. On still
another sample, well-defined triangular *‘screw
growth’’ was seen. In one instance, a pyramidal
crystal was growing from the bottom of a hexagonally
sided pit. The pyramidal and triangular shapes of

76

the growths and pits might be attributed to the
crystal structure of indium, but it is suspected that
such growths would have been removed by nitric
acid.

Hexagonal pits and growths might be attributed to
the germanium. If so, a pit in germanium conforming
to the shape of the germanium crystal would not be
expected unless the germanium had been dissolved
by the indium and the resulting mixture removed by
the nitric acid. This leads to speculation about the
action of nitric acid on different mixtures of germa-
nium and indium. Triangular pits which seem to be
progressing towards a hexagonal shape can also be
seen in the photographs. The spotted or ‘‘dusty’’
surface of the germanium is thought to be a result
of the previously mentioned action of nitric acid on
germanium.

Fi

Penetration.

35X.

It is well known that indium-germanium alloy
junctions do not result in an even mixing or diffusion
of one metal into the other. Recrystallization of
single crystal germanium on the indium side of the
barrier has been seen, and the irregularity or
“raggedness’’ of the junction is still a problem to
those interested in junction preparation.

Examination of germanium alloy transistors by the
above technique, that is, removal of the indium by
nitric acid, shows a number of indentations in the
germanium surface, but they are not of a geometric
configuration. Other samples show no significant
pits, but the surface is smoothly curved. The pits
of these commercial units may be due to entrapped
gas, and, if so, it could be said that the slight
curvature of the surface is a result of more complete
melting than in samples alloyed at low temperatures.
Since a considerably higher temperature is involved

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED |
PHOTO 45703

2.3.6. Photomicrograph of Germanium Surface Showing Crystal Growth (Indicated by Arrows) and Indium

in the preparation of the commercial units, it may
be that samples alloyed at low temperatures indicate
the beginning of the alloying process, that is,
penetration takes place at some particular spot on
the germanium surface in a disordered region of
some sort such as, perhaps, a dislocation or an
etch pit. Since etching rates along crystal faults
are different than on a normal surface, the prefer-
ential attack might be limited to a crystal fault
inherent in the crystal or introduced by cutting and
not removed by etching. Such a conclusion does
not appear to be completely unfeasible, since it is
known that ““screw growths”’ occur at crystalline
faults. If such is the case, it would lead to specu-
lation about alloying indium into germanium which
has been ground and lapped but not etched. Further
speculation would suggest that recrystallization of
indium-germanium alloy junctions begins from these
growths.
Pl UNCLASSIFIED
PHOTO 45704 |

Fig. 2.3.7. Dark-Field llluminated Photomicrograph Showing Germanium Surface Crystal Growth. 250X.

The germanium used in this work was n-type, and
it had a resistivity of approximately 2 ohm/cm,
which is not extremely different from that of com-
mercial units. Before alloying it was etched in
CP-4, washed in distilled water, and kept in ethyl
alcohol until used. The indium was lightly etched
in nitric acid long enough to produce a shiny
surface, washed in distilled water, and stored in
ethyl alcohol until used.

78

Examinations of the samples after exposure to air
for a period of 1to 2 weeks showed that the pits
and growths had undergone some change. They were
still on or in the germanium, but the edges had been
rounded as if they had been lightly etched. Another
sample kept in ethyl alcohol for a period of two
weeks showed no changes in the shape or sharpness
of the pits and growths.
Part 3
ENGINEERING

A. P. Fraas A. L. Boch

Reactor Projects Division
3.1. COMPONENT DEVELOPMENT AND TESTING

H. W. Savage

E. Storto

Reactor Projects Division

IRRADIATION TEST OF OIL-LUBRICATED
PUMP ROTARY ELEMENT

D. L. Gray

The device which consists of the rotary element
of an oil-lubricated pump and which is installed
in a gamma-irradiation facility in the MTR canal,
as previously described, ! has continued to operate
with downtime limited to 20 min as a result of a
power failure and 1 hr as a result of a low=lube-oil-
flow alarm. As of August 19, the lower bearing
and seal area had received a total gamma-ray
dosage of 8.9 x 107 r, and, by September 30, a total
gamma-ray dosage of 1 x 1010 ¢ is expected, A
large oil reservoir is used with this type of rotary
element and thus only a small portion (about 2%)
of the lubricant (Gulfcrest 34) is exposed to radi-
ation at any time. The average gamma-ray dosage
to the total quantity of oil in the lubricating system
is therefore expected to be about 2.2 x 108 ¢ by
September 30.

Insofar as can be ascertained, the operation of
the seal has been comparable to that of a seal op-
erated without irradiation for a similar period of
time. The leakage of oil from the seal into the
catch basin has usually been at a rate of 5to 7
ml/day, although higher rates have been observed
for short periods from time to time. Results of

lD. L. Gray and W. K. Stair, ANP Quar. Prog. Rep.
Sept. 30, 1957, ORNL.-2387, p 32.

analyses of samples of oil taken from the test equip-
ment at periodic intervals are presented in Table
3.1.1.

The viscosities of both the bulk and the leakage
oil have increased but not deleteriously so, as
indicated by operation of the apparatus, The in-
crease in the bromine number indicates that irradi-
ation has displaced some hydrogen atoms from the
hydrocarbon molecules, also without noticeable
effect on operation. The acidity number indicates
very little contamination of the system.

TEST OF NoK PUMP WITH
FULTON-SYLPHON SEAL

D. L. Gray

A sump-type centrifugal pump designed to operate
at a shaft speed of 1200 rpm and to circulate NaK at
1200°F at a rate of 1200 gpm was modified to in-
clude a Fulton-Sylphon seal and placed in test
operation as described previously.? During more
than 7500 hr of operation, the oil leakage past the
test seal, which separates the lubricating oil from
the pump tank helium cover gas, has been of the
order of 10 cm? per week, which is exceptionally
small for a seal of this type. The similar upper
seal, which separates the lubricating oil from the
atmosphere, has been equally satisfactory.

2p. L. Gray, ANP Quar. Prog. Rep. March 31, 1958,
ORNL.-2517, p 67.

Table 3,1.1. Results of Analyses of Oil Samples from MTR Test Apparatus

Bulk Circulating Oil

Lower Seal Oil Leakage

Sample Sample Sample
Control Sample
Accumulated Taken Accumulated
Sample of 7-31-58
3-13 to 3-27-58 4-28-58 524 to 6-21-58
Yiscosity, SUS 90.2 118.5 108.0 112.6 140.2
at 250°C
Bromine No., mg of 0.87 5.6 3.4 4,1 7.3
Br/100 g of oil
Acidify, mg of 0 0.013 0 0 0

KOH/g of oil

81

ANP PROJECT PROGRESS REPORT

CAVITATION TESTS OF
CENTRIFUGAL PUMPS

P. G. Smith W. E. Thomas

One of two pumps being tested, as described
previously,? to study the effect of operation in the
cavitation region was stopped because of leakage
of NaK from the system drain valve, This sump-
type centrifugal pump had circulated NaK in the
temperature range of 1200 to 1450°F for 2136 hr
at a flow rate of 645 gpm, a pump speed of 2700 rpm,
and a pump tank cover gas pressure of 2.5 psig. At
the time of termination of the test, the pump had
been operating under cavitation conditions for
1096 hr. The pump appeared to be in good con-
dition after the test; there was no evidence of
cavitation damage, The data accumulated during
this period are being analyzed.

A similar pump that is being operated with a fused
salt as the pumped fluid has logged more than
10,300 hr of continuous operation at 1200°F, in-
cluding about 8200 hr under cavitation conditions.
The oil leakage from the lower seal that separates
the lubricating oil from the pump tank cover gas has
averaged 10 cm>/day during the past six months.
Seal leakage from the upper seal that separates

3P, G. Smith and W, E. Thomas, ANP Quar. Prog.
Rep. March 31, 1958, ORNL-2517, p 69.

82

the lubricating oil from the atmosphere has averaged
5 cm3/day.

THERMAL STABILITY TESTS
OF METAL SHELLS

J. C. Amos R. S. Senn

A third thin-shell model was tested under high-
temperature thermal cycling conditions similar to
those previously imposed on two other thin shell
models.?*> This shell had the same geometry as
those used previously, but it was fabricated by
welding over-size sections and reducing the shell
to final dimensions by machining the weldment.

Testing of the third shell was continued beyond
the scheduled 300 thermal cycles to 602 thermal
cycles. At this point the test was terminated be-
cause of a test loop failure. There was no indication
that the loop failure was in any way associated with
the test shell. The shell is currently being examined
for dimensional changes, and it will be metal-
lurgicaily examined for effects of thermal cycling on
the welds,

45, C. Amos and L. H. Devlin, ANP Quar. Prog. Rep.
March 31, 1957, ORNL-2274, p 50.

3), C. Amos and R. L. Senn, ANP Quar. Prog. Rep.
Sept. 30, 1957, ORNL-2387, p 45.

PERIOD ENDING SEPTEMBER 30, 1958

3.2. HEAT TRANSFER STUDIES

H. W. Hoffman
Reactor Projects Division

STUDIES OF THE EFFECT OF THERMAL-STRESS
CYCLING ON STRUCTURAL MATERIALS

J. J. Keyes A. |. Krakoviak

Experimental studies of the effect on Inconel of
surface thermal-stress cycling in a fused.salt en-
vironment (NaF-ZrF -UF ,, 56-39-5 mole %) were
continued with the use of the high-frequency pulse-
pump loop.! In order to obtain data for comparison
with data obtained in the University of Alabama
bulk thermalcycling studies? and to establish the
frequency dependence of cycles to crack initiation,
tests 9 through 11 were conducted at a pulse fre-
quency of 0.1 ¢cps and test 12 at 1.0 cps. Tables

3.2.1, 3.2.2, and 3.2.3 present summaries of the
operating conditions and results for these runs. The
wall thicknesses of the Inconel specimens used for
these tests were selected to give just perceptible
temperature fluctuations at the outside surface.
Further, as seen from Table 2.3.3, tests 9 through

11 constitute a comparison of the effect of exposure
time (or total number of cycles) for nearly equiv-
alent stress conditions at a given frequency.

]J. J, Keyes, A. |, Krakoviak, and J. E. Mott, ANP
Quar, Prog. Rep., Dec. 31, 1957, ORNL-2440, p 54.

2J. F. Gonee, Jr,, and W. D. Jordan, Thermal Fatigue
Tests II, MR7, Bureau of Engineering Research,
University of Alabama, June 1957,

Table 3,2, 1. Operating Conditions of High~Frequency Thermal-Stress-Cycling Test of Inconel Pipes

Test Number

Test Variables

9 10 11 12
Average fluid temperature, °F 1407 1412 1411 1415
Maximum temperature difference between 550 569 557 575

hot and cold streams, °F

Frequency of temperature oscillations, cps 0.1 0.1 1.0
Total pulsing time, hr 200 75 25 100
Total cycles 72,000 27,000 9,000 360,000
Total fluid flow,rate, gpm 3.6 3.6 3.7 3.8

Table 3,22, Results of Metallographic Examination of Test Pieces from High«Frequency
Thermal«Stress«Cycling Tests of Inconel Pipe

Test No. Metallographic Results

9 Severe intergranular cracking found throughout test section;
depths of cracks increased from 0.128 in. at top of test
section to 0,207 in. in center section

10 Intergranular cracking found throughout; depths of cracks
increased from 0,012 to 0.023 to 0.103 in. in top, center,
and bottom of test section, respectively

11 No evidence of cracking

12 No evidence of cracking

83
ANP PROJECT PROGRESS REPORT

Table 3,2,3. Summary of Results of High«Frequency Thermal-Stress<Cycling Tests of Incone! Pipe

Mean temperature:

1405 + 10°F

Inside diameter of test section:

0.485 in.

Estimated Measured Wall Calculated
T Calcul ated Total Numb
est Inside Wall Temperature Thickness Stress on ) otal Number
N i ) Total Strain f Cyel
O Temperature Amplitude on of Test Inside Wall per 1/ Cycle ot Lycles
Amplitude* QOutside Wall Section Fibers ?7)
0
(°F) (°F) (in.) (psi)
5 113 t7 0.250 19,400 0.095 194,400
6 1121 Not measured 0.415 23,400 0.114 32,000
7 169 0 0.415 13,400 0.065 224,000
8 170 Not meosured 0.415 13,500 0.066 508,600
9 +120 18.6 0.445 21,700 0.106 72,000
10 127 19.5 0.445 22,800 0.111 27,000
11 122 *+8.8 0.445 22,100 0.108 9,000
12 176 t5.5 0.147 12,000 0.058 360,000

*90% efficiency assumed; Jakob's equation used for calculating 7; entrance effects neglected, and flow de-

termined by heat bolance.

The photomacrographs of Fig. 3.2.1 picture
typical cracks observed at three locations along
the length of the test section used in test 9:

(a) shows the conical transition piece (diameter
reduction from 1 in, to 0.485 in.) at the entrance to
the test section, (£) shows the entrance region of
the test section, and (c) shows the central region
approximately 4 in. downstream from the testesection
entrance. A photomicrograph showing one of the
cracks in the central region of the test section of
test 9 is presented in Fig. 3.2.2, Finally, Fig.
3.2.3 presents a comparison of photomacrographs of
equivalent fransverse sections (near the exit) from
tests 9 and 10. The effect of the total number of
cycles of exposure (72,000 cycles in test 9 and
27,000 cycles in test 10) on crack depth is apparent
in Fig. 3.2.3. The deepest crack resulting from
test 9 was 0,207 in, deep, and there were many
cracks as deep as 0,120 in, {n test 10, however,
with only slightly more than ]/3 the exposure, the
maximum crack formed was only about 0.100 in.
deep, and most of the cracks were 0.012 to 0.020 in.
deep. A further decrease to a total exposure of
9,000 cycles (test 11) resulted in no cracking.

84

Test 12, the first of a series extending the fre-
quency range of these studies, was terminated at the
end of 100 hr (360,000 cycles). Post-operation ex-
amination showed no cracking of the test section.

A cross comparison of the available lower-frequency
data indicates a 0.012- to 0.020-in. average (0.100-

in. maximum) crack depth at 0.1 ¢ps and only 0.003

in. at 0.4 cps for the same strain and approximately

the same number of cycles of exposure.

Estimates of the stress on the inside wall fibers
and the corresponding strain per one-fourth cycle,
€, are presented in Table 3.2.3. In these esti-
mates, allowance was made for tube-wall curvature,
but a linear temperature gradient through the wall
was assumed. The quantity, €, is plotted as a
function of the cycles of test duration in Fig, 3.2.4
in an effort to summarize the results of this investi-
gation; the experimental data are differentiated as
to cyclic frequency. The point of incipient ma-
terial failure was chosen as the appearance of the
first discernible cracks on the basis that the ma-
terial had been weakened and that application of a
design load could lead to rupture. An arrow attached
to the data point and pointing to the right indicates

Fig. 3.2.1. Effect of High-Frequency Thermal Cycling
on Inconel in @ Fused-Salt (NaF-ZrF -UF ,, 56-39-5 mole
%) Environment — Test 9. Photomacrographs showing

cracks [(a) and (5) entrance region and (c) central region]
found after 72,000 cycles at 0.1 cps and a strain level of

0.11% per one-fourth cycle. 3X. Reduced 24.5%.

PERIOD ENDING SEPTEMBER 30, 1958

that no cracks were observed and that failure would
be expected to occur only after longer exposure.
Conversely, an arrow pointing to the left indicates
that the observed crack depth was such that initial
cracking occurred earlier in the exposure. Thus, the
incipient failure point occurs, for a given strain,
somewhere along a horizontal line between the
oppositely directed arrows. In only one case was a
run (test 6) terminated at the condition of incipient
failure. At the lowest strain (0.058% at 1 cps), no
cracks were observed for N = 360,000; a longer test
is in progress. For comparison, Fig. 3.2.4 also
shows experimental mechanical fatigue data ob-
tained at ORNL by Douglas and Swindeman at a
cyclic rate of 1.0 cps. The limits on their data
represent the maximum and minimum strain to which
the samples were subjected during a cycle. It is
recognized that a similar range applies to the data
of this study; however, uncertainties in temperature
make it impossible to accurately estimate this
range. In general, although the data are limited,
there is indication that, for a given strain per cycle,
failure occurs at fewer cycles with surface thermal
cycling than with mechanical cycling.

LIQUID-METAL HEAT TRANSFER EXPERIMENT
G. L. Muller® H. W. Hoffman

The initial phase of the studies of liquid-metal
(mercury) forced-circulation heat transfer with in-
ternal heat generation and no heat transfer through
the walls of the tubing has been completed. The
purpose of this investigation has been to provide
not only experimental information useful in the de-
sign of systems containing internally heated circu-
lating liquid metals but also to test the heat-
momentum transfer analogy used in the theoretical
treatment of this problem. Further, since heat was
not transferred through the walls, the results were
unaffected by interfacial resistances, such as
gaseous films or scale, and a direct check could be
made of the “‘interface thermal resistance’ theory
postulated to explain the reduced values of the ex-
perimentally measured Nusselt modulus when heat is
transferred through a pipe wall.

3Work performed while on assignment from Pratt &
Whitney Aircraft.

85
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
T-15497 4

Fig. 3.2.2. Effect of High-Frequency Thermal Cycling on Inconel in a Fused-Salt Environment — Test 9. Photo-
micrograph of crack in central region of test section.. 100X.

UNCLASSIFIED]
15565

Fig. 3.2.3. Effect of High-Frequency Thermal Cy

Test 10. Comparison of photomacrographs of test-section exits. 3X.

g on Inconel in a Fused-Salt Environment — (o) Test 9, (5)

86

The experimental apparatus used was described
previously,* and is illustrated in Figs. 3.2.5 and
3.2.6. The heat-transfer section is shown in Fig.
3.2.7; 36-gage copper-constantan tube-wall thermo-

UNCLASSIFIED
ORNL-LR-DWG 33944

: 7T Rl
& |1
o 02l - e
EJ_; ' i i H
> | ;
3 eI |

o i ! :
W ooz T L Fu S U S S 1[ , I
[ - e
= : LY it = AR (I
| =1 el
o i ] o de= -P ]
5 0.05 Il Pl 1 P
a ° 0.4 cps, MEDIUM-GRAINED INCONEL ROD
oA 1| * 0.t cps, MEDIUM-GRAINED INCONEL ROD
E —4—4 1.0 cps, MEDIUM-GRAINED INCONEL ROD +—
E [ ORNL METALLURGY DIVISION FATIGUE ‘ J
7002 F— DATA, MAXIMUM AND MINIMUM ¢, FOR — BERR
g COARSE-GRAINED 1.0 cps INCONEL ROD '
= TESTED AT {400°F AND 1.G cps l ‘

0.01 | T |

104 2 5 103 2 5 108
N, NUMBER OF CYCLES

Fig. 3.2.4. Effect of High-Frequency Thermal Cycling
on Inconel in a Fused-Salt Environment. A comparison

with mechanical fatigue data in terms of total strain.

PERIOD ENDING SEPTEMBER 30, 1958

couples were located at the positions B in this
figure. Power was supplied for electrical resistance
heating of the mercury at 5 v and 1900 amp, maxi-
mum. The 60-cycle frequency of the power supply
ensured a uniform current distribution with radius
(uniform radial volume heat generation). Axial
variation of the heat generation as a result of the
temperature dependence of the mercury electrical
resistivity was neglected, since the variation about
the mean was estimated to be only +2%, Heat
losses from the heated section were minimized by
guard heating and insulation.

The experimental results are presented in Table
3.2.4. The Reynolds modulus covered a range from
29,000 to 165,000, and the volumetric heat gener-
ation rate varied from 1.4 x 107 t0 8.3 x 107
Btu/hr-ft3. The radial temperature differences
given in the final column of the table are based on

4G. L. Muller, ANP Quar. Prog. Rep. Sept. 30, 1957,
ORNL -2387, p 106; ASME preprint No.58-HT-17, Second
Annual ASME-AIChE Joint Heat Transfer Conference,
August 1958,

Table 3.2,4. Experimental Results for Mercury Forced-Circulation Heat Transfer

with Internal Heat Generation

Velumetric . Radial
Heat Removed Axial
Heat Reynolds Prandtl Electrical Temperature
Run from Test in Temperature
Generation Number,  Modulus, Power Input, . ) Difference,
No. R P Bto/h Section, 90t 7 Rise
ate, g 3 NRe NPr 9in (Btu/hr) (Bru/hr) out (°F) two- t
(Btu/hr-ft ) ( F)
x 107
1 3.48 73,300 0.0200 13,880 13,040 0.939 104.47 3.46
2 1.65 29,000 0.0200 6,700 6,060 0.905 127.05 3.90
3 2.55 45,200 0.0200 9,940 10,280 0.969 125,51 4.90
4 7.53 147,000 0.0192 29,500 27,400 0.929 113.35 4.08
5 5.84 102,900 0.0192 22,670 20,530 0.907 125.41 3.25
6 8.30 164,500 0.0191 32,700 29,900 0.915 113.30 5.87
7 5.10 95,700 0.0200 20,600 18,100 0.878 119.03 4.57
8 2.68 88,500 0.0225 10,870 9,710 0.894 63.71 2,23
9 3.20 58,400 0.0205 13,060 11,860 0.908 122,42 3.33
10 2.25 40,000 0.0204 9,060 8,510 0.939 124, 11 2,83
11 1.44 39,200 0.0224 5,660 5,200 0.919 75.49 1.81
12 6.51 124,000 0.0207 26,450 24,100 0.91 116.46 3.78

87

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORNL—LLR—DWG 25908R

_—MIXING CHAMBER AND
//THERMOCOUPLE WELL

/

T
]
uT _‘
aNt :
o0 208000 A
COOLER//
POWER } N POWERSTAT
TRANSFORMER
/ — THIN-PLATE ORIFICE
IN
v = ol«“NT
/ Al OO
TROL\" )y
W o0 Q
0
[
SUMP

SURGE
Z CRAMBER .

N
AUXILIARY 7!
c0
e
SN <Y
\_ PUMP DRIVE MOTOR AND
') .O SPEED SELECTOR

POSITIVE DISPLACEMENT MOYNGC PUMP

MERCURY
FLOW DIRECTION

Fig. 3.2.5. Schematic Diagram of Experimental Apparatus for Study of Forced-Circulation Heat Transfer in

Mercury with Volumetric Heating.

88

68

UNCLASSIFIED|
PHOTO 30644/

Fig. 3.2.6. Photograph of Experimental Apparatus for Study of Forced-Circulation Heat Transfer in Mercury with Volumetric Heating.

8561 ‘0€ ¥IFWILJIS INIANI d0I¥3d
ANP PROJECT PROGRESS REPORT

l«— THERMOWELL

.
.

GLASS
N TEST
N SECTION

s
//”

.

,: N’

AN AN’ COPPER
VOLTMETER E N N’ SHELL

Ty MICARTA Srlhi ] T CLAMSHELL
HERMOCOUPLE 1 SRS P S
CLAMP | e HEATER

Pl Lae L
AR BRRRRRRY

MERCURY FLOW

DL Do RS ST SO
o, e . DA W0
A R A RN~
L e P LA,

p; v A T
B R A .

L, , .

. ’ BN B
00 PR

r; j——r_J

INSULATION

1~ MIXING CHAMBER

COPPER BUS BAR

v Vv

VvV

UNCLASSIFIED
ORNL—~LR—DWG 33942

GLASS TEST SECTION

CLAMSHELL GUARD HEATER

COPPER SHELL

A TTZZLE
P
S

AR

INSULATION

THERMOCOUPLE

CURRENT TRANSFORMER, 1200:5

DETAIL OF
THERMOCOQUPLE
AMMETER INSTALLATION
THERMOCOUPLE
CONTROL
THERMOCOUPLE
AMMETER DETAIL OF CONTROL

AND MEASUREMENT
THERMOCOUPLES

CURRENT TRANSFORMER, 4200:5

~{™SCREW-TYPE COUPLING

S~ STAINLESS STEEL TUBE

Fig. 3.2.7. Details of Test Section for Mercury Heat-Transfer Experiment.

90

average values obtained from the six central thermo-
couples of the upstream half of the test section.
The results are thus relatively free of end effects
resulting from the massive electrodes and they
represent the fully established thermal boundary-
layer region. A typical set of fluid- and wall-
temperature measurements is given in Fig, 3.2.8.
The results are presented graphically in Fig, 3.2.9
in terms of T, a dimensionless temperature dif-
ference.® On the average, the data indicate temper-
ature differences 49% higher than those of the
analogy prediction; and a line drawn through the
data at this level and parallel to the lower T-curve
of Fig. 3.2.9 has the empirical equation:

I
T =

- 0.92
53.0 +0.152 N

SA discussion of the factor T, including a comparison
with previous analyses is to be found in ANP Quar,
Prog, Rep. Dec, 31, 1957, ORNL-2440, p 59.

UNCLASSIFIED
ORNL-LR-DWG 33943

190
Nge=173,300
| Wp,=0.020 74
180 Pr 4
g" =3.48x10" Btushr- 13 b4
= N ” . @
170 |—# =546 8tusnrt-oF v
27y =0.2510n. 7/
p— — /
60 |—"" " mlprea=270°F L
WALL TEMPERATURE —~| 7% .
e GUARD HEATER \\:’fl/
50 TEMPERATURE 7
7
140 _//
/s
4

130 pex
¥C,

/

/A

120
A/ MIXED - MEAN FLUID TEMPERATURE
/ fo—Tm= 3.46°F

110
/

A .
100 ° "‘{/

TEMPERATURE (°F)

T i

0 2 4 6 8 10 12 14 16 18
CISTANCE FROM ¢ OF DOWNSTREAM ELECTRODE (in.)

Fig. 3.2.8. Typical Experimental Temperature Profiles
in Mercury Heat=-Transfer Experiment Test Section — Run

1.

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
4 ORNL-LR-DWG 33944

0 I H T !
fl:bl»#%—fi ] !
o N ) ]
5 7=
— 65+0.24 ».fgf’s Tt
1
=
2 79+O.22€w‘v":,('l'92
_2 L
10 P ik
‘:’@ o —DATA AVERAGE
X s N 1
3
" N \.\
2 3T
Oy - N
Il ™
. N
-3 \\
10 ™
My
N
\\
5 \\
L N
N
2 — I
\\\
10-4 N
102 2 5 10° 2 5 109 2 5 10°
Du_ye
= Tm7%
NF'e - I

Fig. 3.2.9. Comparison of Theoretical and Experimental
Temperature Differences for Mercury Heat Transfer with
Internal Heat Generation and No Wall Heat Transfer.

The maximum deviation of the data from this corre-
lation is approximately +35%. An error analysis
indicates that most of the observed deviation occurs
in the measurements of the small radial temperature
differences.

This investigation thus shows differences between
theory (based on the Martinelli heat~momentum-
transfer analogy in which a, the ratio of the dif-
fusivity of momentum to the diffusivity of heat, is
assumed to be unity) and experiment similar to
those observed in wall heat-transfer studies with
liquid metals. Since anomolous wall effects®
could not influence the experimental results, it is
concluded that the diffusivity ratio is somewhat
smaller than 1 for mercury.

HEAT TRANSFER WITH VORTEX FLOW
W. R, Gambill N. D. Greene

The study of heat transfer with air in high-velocity
vortex motion within electrical-resistance heated
tubes was continued.” Details of the two types of

8 .
Gaseous films or scales.

Details reported in A Preliminary Investigation of Air
Film Heat-Transfer Coefficients for Free- and Forced-
Vortex Flow Within Tubes, N, D, Greene and W. R.
Gambill, ORNL CF-58-5-67 (May 23, 1958),

91
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
CRNL—LR-DWG 29642

FLUID [N

SPIRAL RAMP

ORIFICE (0.104-in. DIA)
HEAT TRANSFER SURFACE
{ BRASS)

// F /,O ' VORTEX ACCELERATOR CONE

= 756%

(o) THE FORCED VORTEX GENERATOR

HEAT TRANSFER SURFACE

0.260-in.1D
{COPPER)

0.300-in.0D

EXTERIOR WALL INSULATED

TANGENTIAL HOLES, 0.0135-in. DIA
(50 HOLES /in.)

At = 4049

{$#) THE FREE VORTEX GENERATOR

Fig. 3.2.10. Vortex Flow Generators.

92

vortex generator used are shown in Fig. 3.2.10. In
the forced-vortex® generator (3.2.102) the air enters
through a small orifice and is given a rotary motion
by the spiral ramp. The vortex is further accelerated
by a converging cone |located immediately before the
test section. In the free-vortex generator (Fig.
3.2.106) fluid is introduced through numerous, small
tangential inlets distributed along the tube length.
With equal and constant pressure drop across each
inlet, the incoming fluid cannot be displaced by
fluid from an upstream inlet. Thus, in proceeding
downstream, the fluid is forced to rotate through a
progressively decreasing radius with progressively
increasing tangential velocity.

The result of experiments with these two generators
are given in Figs, 3.2.11 and 3.2.12. In these plots,
a parameter related to the Colburn j-factor is pre-
sented as a function of flow power dissipation (the

81n a forced vortex the tangential velocity is directly
proportional to the radius; in a free vortex the tangential
velocity is inversely proportional to the radius,

UNCLASSIFIED

3 ORNL-LR-DWG 29613R
10 . I ) g —
I i il 2 L
! | " | e | :
5 | ! RS MU N ’/‘(/ .
- ; C\ e ——
j F\iO\N \CE\:’//// . : ol
: LT
\N | et iy V 1
\ A;L‘L F\’O -t ‘ ‘ | R
== —en e e
& s ‘ _CONDITIONS: T ]
T 35 CUTIINL T TuBE: L/p=20, £,=0.378in. | |[]
= | . {SEE Fig. 3.2.10) T
! ”*‘ FLUID: AIR J ‘ H '
Z ‘ . i ‘ . L.{-.‘-_.. J
N R e 1 T
ot 2 5 f 2 5 10 2 5 408

FLOW POWER DiSSIPATION (hp/flz)

Inside Air Film Coefficients in a Tube for

Fig. 3.2.11.

Linear and Forced-VYortex Flow.

PERIOD ENDING SEPTEMBER 30, 1958

energy required to pump the fluid per square foot of
tube heated surface). Since the physical property
factor, Nzr /c, was found to be nearly constant
within the limitations of the experiment, this is
essentially a plot of the heat-transfer coefficient,
h. A comparison is made in each case with linear
flow heat transfer through tubes of equivalent
length and diameter. From Fig. 3.2.11 it may be
seen that forced-vortex flow is superior to linear
flow only for power dissipations in excess of

17 hp/ft2. However, more extensive boiling-water
vortex-flow data’ indicate that increasing the ratio
of inlet area to tube cross-sectional area and de-
creasing the ratio of tube length to tube diameter
will result in an upward shift of the vortex flow
curve. The free-vortex results given in Fig. 3.2.12
indicate that, for the specific conditions of the test,
area ratio = 0.4 and L/D = 11,5, free vortex flow is
much superior to linear flow,

"W. R. Gambill and N. D. Greene, A Preliminary Study
of Boiling Burnout Heat Fluxes for Water in Vortex Flow,
ORNL CF-58-4-56 (April 12, 1958).

UNCLASSIFIED
CRNL-LR~-DWG 296414R

4
to T T —- T -1
— CONDITIONS" H‘ EfT [ IIHT - J] g
5 % TUBE: £/D=417, D,;=0.260 in,(SEE Fig.3.210)[ ]
. . FLUID AR o R H-'—— - 4 ‘”1
. i R o1
‘. [N /;-
H [y !
VORTEX FLOW s -
DG Ll Tel - ; g ;
5 |— AR i,f,,,,,, I
= 4/' B [J/ |
Ny s T
—T" ILINEAR FLOW(CALC)
T N H
‘ ‘ [ S ! i
i . \
40?—/Kmll | L 1\1\|
o4 0.2 05 2 5 10 20 50 100

FLOW POWER DISSIPATION (hp/11%)

Fig. 3.2.12. Inside Air Film Coefficients in a Tube for

Linear and Free-Vortex Flow.

93
ANP PROJECT PROGRESS REPORT

3.3. INSTRUMENTATION AND CONTROLS
H. J. Metz

instrumentation and Controls Division

DATA ACQUISITION SYSTEM
G. H. Burger

The data acquisition system purchased from the
G. M. Giannini Company, which was briefly de-
scribed in the previous report, ! was installed and
placed into operation. For the initial check of the
system, three temperature recorders and one
pressure system, consisting of a 12-point pressure-
scanning system feeding a strain gage and a
Gilmore strain gage circuit, were fed into the in-
stallation. The data were recorded on the standard
log sheets for the 20-in.-carriage typewriter. In
addition, the data were logged on the punch tape in
Oracle code. The code was verified by the Oracle
tape readers. Since the initial check of the system,
a total of seven recorders giving 84 separate inputs
has been fed into the system, the system has been
operated to test all aspects of system operation,
that is, the typewriter, tape punch, tape reader,
tape code, logging format, and operational char-
acteristics, such as typewriter operational format,
carriage return, carriage return and reset, etc,,
have been checked. In addition, all data operational
functions of the system have been checked, in-
cluding the circuitry that provides for operations on
the data such as multiplication, division, and
addition,

Since the unit began operation, it has run a total
of 612 hr, There have been some operational
troubles with the system, mostly in the logging
unit, the typewriter, and typewriter linefinder,
Other minor troubles occurred that were a result
of incomplete checking at the factory and initial
design inadequacies. The faults attributable to
the vendor are now being corrected by the vendor,
and it is expected that the system will again be
in operation within a month. The system was
shipped from the fractory without a final system
check in order to meet the delivery schedule.

It is expected that with the completion of the
corrections and modifications now being made on
the system, it will operate in a satisfactory manner,

16. H. Burger, ANP Quar. Prog. Rep. March 31, 1958,
ORNL-2517, p 81.

94

It is planned to operate the system with all the
maximum inputs possible (120) for a short period
to again check the system, After this period, it

is expected that the system will be used to record
and handle data from an actual operating test in
order to determine the capabilities or shortcomings
of the machine and to establish a list of criteria
for or characteristics desirable in future machines,

LIQUID-METAL-LEVEL TRANSDUCERS
G. H. Burger

Life tests of the three level probes operating in
NaK at 1200°F and above were continued. Two of
the units have now operated more than 9000 hr and
the other has operated more than 4600 hr.

Analyses of data obtained previously in wetting
tests of Inconel probes in NaK verified that, al-
though wetting is rather irregular and unpredictable,
at temperatures above 600°F it is complete within
a very short time (almost instantly). At tempera-
tures below 600°F, no definite time vs temperature
factor could be established.

The iron-plated [nconel probe tested as described
previously ! gave unsatisfactory results in that
there was poor wetting of the iron plate. Since
only one iron-plated probe was tested, this ob-
servation is not conclusive, but no further tests
on iron-plated probes are contemplated.

Wetting and calibration tests of four [nconel
probes in sodium have indicated that wetting in
the sodium will be similar, if not superior, to that
in NaK. Tests of these probes in sodium are con-
tinuing.

PRESSURE TRANSMITTER TESTS
C. M. Burton

Two Taylor Instrument Company pressure trans-
mitters designed for use at high temperatures in
the pressure range 0 to 200 psig are undergoing
overrange (300 psig) tests, as described pre-
viously,? in order to accelerate the effects of
aging at 1400°F. Micrometer measurements made

2C. M. Burton, ANP Quar. Prog. Rep. March 31, 1958,
ORNL-2517, p 82.

after 5100 hr showed slight bulging of the Inconel
bodies, but the calibrations have remained within
the error limits. These tests are continuing,

THERMOCOQOUPLE TESTS
C. M. Burton

Life tests of a series of thermocouples were
terminated after about 11,000 hr of operation at
1500°F. Thirteen l/d-in.-OD, Inconel-sheathed,
magnesium-oxide-insulated, Chromel-Alumel, dual
thermocouples, and three similar platinum—
platinum-rhodium thermocouples were tested.

The end closures were Heliarc-welded caps. The
thermocouples were immersed in a fused-salt
mixture (NaF-ZrF ,-UF ,, 50-46-4 mole %). The
normal operating temperature was 1500°F, except
when calibrations were made at 1100 and 1300°F
against a platinum—platinum-rhodium standard
thermocouple.

On the whole, the thermocouples were performing
well when the test was stopped, although the
errors were increasing with age. Two Chromel-
Alumel thermocouples failed before the end of
the test. The test was stopped because all the
helium purge and vent lines had become plugged
with ZrF , vapor deposits.

After shtudown, it was found that a weld of a
spark-plug-level riser to the vessel had failed,
but the buildup of ZrF , was sufficient to prevent
significant helium leakage at the 2- to 3-psiqg
operating pressure, .

The Inconel sheaths submerged in the salt were
bright and in good condition, but the sheaths ex-
posed to the blanket gas were considerably cor-
roded. All the risers were plugged with ZrF ,, and
it would probably have been impossible to have
removed and replaced a thermocouple.

A similar life test in a sodium environment has
now run about 16,000 hr. Many of the thermocouples
show considerable error,

THERMOCOUPLE DEVELOPMENT STUDIES
J. F. Potts

Behavior of Thermocouples in the Temperature

Range 300 to 1100°C

Means are being sought for improving thermo-
electric measurements by obtaining a better under-
standing of the fundamental behavior of thermo-
couples in the range 300 to 1100°C. The major

PERIOD ENDING SEPTEMBER 30, 1958

portion of the study is being conducted under a
subcontract at the University of Tennessee En-
gineering Experiment Station by D. L. McElroy,
The progress of the first year of this work was
covered by McElroy in a previous ORNL report.3

In the past year, a comparative evaluation of
thermocouples constructed of a number of com-
mercially available nickel-base alloys has been
made in order to establish relations describing
the effects of time and temperature on stability.
At a sacrifice of thermoelectric power and linearity
of the temperature-vs-emf curve, it was demon-
strated that alloys of nickel and chromium used
as positive legs demonstrated greatest stability
at 800 to 1100°C in air when the chromium content
was increased to about 20%. Chromel-P contains
roughly 9% chromium, With the same conditions
and objections as those given above, it was
demonstrated that nickel-silicon alloys containing
about 3% silicon evidenced greater stability than
the nickel-silicon-aluminum-manganese alloys,
such as Alumel. |t was further demonstrated that
thermocouple life and stability could be improved
by the use of carefully fabricated swaged sheaths.

Attempts to improve the performance of Chromel-
Alumel thermocouples have led to the conclusion
that isothermal heat treatment of Chromel-Alumel
thermocouples at 565 and 840°C causes shifts in
the original calibration with time at temperature,
Further, isothermal heat treatment at 900°C in
oxygen has indicated that for short exposures
Chromel-P dominates the thermocouple changes
and for long exposures Alumel dominates the
drift. Pickling of oxidized wire was found to be
a beneficial treatment; and partial immersion heat
treatments were found to be conducive to small
drifts of calibration in the temperature range 385
to 830°C. Further refinements of the techniques
for using the unilateral gradient furnace described
in the previous report? have led to its considerable
value as an analytical tool for quantitatively cor-
relating thermocouple performance with various
inhomogeneity-producing phenomena, such as
oxidation and recovery from cold work in tem-
perature gradients.

3p. L. McElroy, Progress Report I, Thermocouple Re-
search Report for the Period November 1, 1956 to
October 31, 1957, ORNL-2467.

95

ANP PROJECT PROGRESS REPORT

Oxidation and Cold Work Studies?

Although much of what is known about base-metal
performance at temperatures up to 1100°C can be
extended in order to develop improved thermo-
couples for use at temperatures in excess of 1100°C,
major gaps exist in the fundamental information. A
survey of the present knowledge of thermoelectric
measurements has revealed that information is
needed on relations between oxidation and thermal
emf and between cold work and thermal emf. Tests
are under way over the range 300 to 1000°C on
variations of nickel-chromium alloys as the positive
thermal elements and nickel-silicon alloys as the
negative elements, including Hoskins Chromel-P
and Alumel and other alloys available from the
Hoskins Co., Driver-Harris Co., and Kanthal Corp.
For comparison, samples of thermocouple-grade
iron, constantan, and copper will also be studied.

Data have been obtained on oxidation (measured
by weight gain) in air at 700, 900, and 1000°C.
Tests are now in progress at 800°C. The data
indicate that greater oxidation of Chromel occurs
at 700 to 800°C than at 1000°C.

Tests of oxidation in a temperature gradient
(maximum temperature of 1000°C) of base-metal
samples in various sizes of quartz and Alundum
tubes were initiated recently. Microstructure
examinations have shown intergranular attack of all
the alloys tested above 850°C, whereas surface
attack predominates in alloys tested at tem-
peratures below 850°C. Depth of immersion tests
are also under way on Geminal-P and Geminal-N
(Driver-Harris), Kanthal-P and Kanthal-N (Kanthal
Corp.), Chromel-P plus columbium and Alumel
(Hoskins Co.) in quartz and Alundum tubes with
and without Inconel, stainless steel, and titanium
shavings or fines placed in the bottom of the
tubes. Preliminary tests indicate that the results
are greatly influenced by the choice of thermo-
couple well material. Drift tests in an atmosphere
of oxygen at a pressure of 30 psig indicate that
the thermocouple drift rates are not markedly dif-
ferent than those measured in air.

The cold work studies have progressed to the
point where the relations between rolling and
drawing reduction, respectively, and percentage
of cold work have been established for wire

4Work under subcontract at University of Tennessee
Experiment Station.

96

samples that were initially round. For example,
wire flattened to 10% of its original diameter
receives only about 45% of the cold work it would
receive if it were die-drawn to 10% of its original
diameter, The reason for establishing this relation
is that it is experimentally more convenient to
flatten wire between rolls than to draw it through
dies, particularly when it is desired to impose
varying amounts of cold work on one continuous
sample. A series of experiments is being con-
ducted to establish the relation between tem-
perature, time, and percentage of cold work and
the ranges for recrystallization and recovery.
Hardness measurements will be made to establish
recrystallization temperatures, and electrical re-
sistance measurements will be made to establish
recovery temperatures. After these data are ob-
tained, the cold work vs thermal emf relations
will be established with the use of the unilateral
gradient furnace mentioned above. These cold
work studies will establish the practical limits
of precision of available base-metal thermocouple
alloys.

Calibration Studies?

Temperature-standardization techniques are being
developed through a series of closely controlled
intercomparisons of the results of melting and
freezing experiments with water, tin, lead, zinc,
antimony, aluminum, and silver. The temperature
measurements are being made with various platinum
resistance thermometers and platinum—platinum-10%
rhodium thermocouples certified by the National
Bureau of Standards.

Oracle Program for Thermocouple-Data Handling
J. W. Reynolds

Work on the improvement of the Oracle program for
thermocouple-data handling is nearing completion
at ORNL. The format for processing of raw thermo-
couple millivoltage data is being changed to include
the ‘‘standard’’ temperature only once on the final
picture, whereas previously the standard tem-
perature had been repeated for each thermocouple
test point. The coding of the thermocouple point
identification is to be changed to include the
location of the test facility and the type of test.

In addition to the columns for error, test millivolts,
and standard millivolts, a column is to be included

for the resistance or other desired parameter. For
aid in statistical analysis of a particular type of
thermocouple, the new format is to include the
summation of errors, the summation of squares of
the errors, the standard deviation, the average error
plus and minus three standard deviations, and the

PERIOD ENDING SEPTEMBER 30, 1958

degrees of freedom. As an additional aid to the
reduction of the data to comprehensible form, a
program for presentation of the accumulated data

in analog form after it has been statistically treated
is being included.

97

ANP PROJECT PROGRESS REPORT

3.4. APPLIED MECHANICS

B. L. Greenstreet
Reactor Projects Division

BASIC PROBLEMS IN ELASTICITY
F. J. Stanek

Fundamental studies of conical, cylindrical, and
spherical shells were continued in which classical
small-deflection theory is being used to determine
elastic properties. The basic objective of each
study is to list the required stress and displace-
ment formulas and to provide tables of numerical
values of all required functions.

The conical-shell sfudy] is nearing completion
and a preliminary report? giving the results of the
work has been issued. The results provide a means
whereby complete analyses can be made for shells
subjected to a combination of axially symmetric
loads. These loads include edge forces and mo-
ments, membrane forces, and uniform pressure, as
shown in Fig. 3.4.1.

Equations are given for displacements in both the
meridional direction and nermal to the shell wall,
as well as for the stresses; The values of all the
necessary functions are tabulated for the dimension-
less coordinate x from 0 through 40, The dimension-
less variable is given by

1/2

ctn a
x-=2(3.3o45 - y) ,

where 2a is the vertex angle of the cone, 4 is the
wall thickness, and y is the coordinate measured
from the vertex along the surface of the cone. A

tabulation of the functions for larger values of x

will be included in a later report.

Comparisons have been made of the stress dis-
tributions obtained through analytical and experi-
mental means for @ tube-and-spherical-shell con-
figuration loaded as shown in Fig. 3.4.2.%4 Work
done in evaluating and tabulating the functions ob-
tained in the solution of the spherical-shell

R, . Meghreblian, ANP Quar. Prog. Rep. June 30,
1957, ORNL-2340, p 63.

2g, J, Stanek, Stress Analysis of Conical Shells,
ORNL CF-58-6-52 (Aug. 28, 1958).

3. 1. Stanek, Spherical Segment with Circular Hole
at Vertex Loaded Axisymmetrically Along the Edges,
ORNL-2207 (Dec. 19, 1956).

4R, V. Meghreblian, ANP Quar. Prog. Rep. Dec. 31,
1956, ORNL-2221, p 6.

98

UNCLASSIFIED
ORNL—LR—DWG 33845

Fig. 3.4.1. General Loading for Conical-Shell Study.

UNCLASSIFIED
ORNL—LR—DWG 33946

AXIAL FCRCE, F

TEST ANALYTICAL
MODEL MODEL
a 3.4 deg 4.0 deg
B 69.9 deg 68.0 deg
a 13.440 t2.675
A 0.507 0.507
% 26.51 25.00
! 0.500 0.500
r 0.7995 0.8842
EDGE
CONDITION FIXED FIXED
AT ¢=f

Fig. 3.4.2. Loading for Tube-and-Spherical-Shell Study.
equations resulted in an analytical model that nearly
approximated the actual configuration used in ob-
taining the experimental data. The geometric
quantities for the test shell and the analytical
mode| are given in Fig. 3.4.2

The meridional stresses at the inner and outer
surfaces of the spherical shell are plotted in Figs.
3.4.3 and 3.4.4. The circumferential stresses are
shown in Figs. 3.4.5 and 3.4.6. Excellent agree-
ment exists for both the meridional and circum-
ferential stresses.

NUMERICAL ANALYSIS
F. J. Witt

Second-order, partial differential equations with
variable coefficients are encountered in analyses
pertaining to transient neutron diffusion and heat
transfer. Therefore, the finite difference equations
required for solving an equation of the form

Flx,t) ) Inlx)
ot Ox 2
+ [(g/x) + B(x)] in%l + C(x)n(x,t) + D(xrt) '

with zero initial condition, were derived. Forward

UNCLASSIFIED
ORNL-LR-DWG 33947

6000
T TICA
5000 ) HECRETICAL |
® EXPERIMENTAL
|
4000
3000

STRESS (psi}

\
2000 \
\

l’@‘-

1000
\ |
0 ) O 5-‘ o
®
~1000 1
0 10 20 30 40 50 60 70

¢, COLATITUDE ANGLE
MEASURED FROM CENTER OF HOLE (deg)

Fig. 3.4.3. Meridional Stress at Outer Surface for an
Axial Force of 1 Ibh.

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
ORNL-LR-DWG 33948

2000
@
®
1000 u
10)
L \‘Q
O I l 7 hd
%‘
a
<A -1000 -
L
o
(-
e B
—-2000 |— .
-3000 ~=— THEORETICAL
® EXPERIMENTAL
-4000
o 10 20 30 40 50 60 70

¢, COLATITUDE ANGLE
MEASURED FROM CENTER OF HOLEF {deg)

Fig. 3.4.4. Meridional Stress at Inner Surface for an
Axial Force of 1 Ib.

UNCLASSIFIED
ORNL-LR-DWG 33949

4000

3000 L
THEORETICAL
@ EXPERIMENTAL

2000 \

STRESS (psi}

\

it
? p\ PR e g
g 1
-1000
O 40 20 30 40 50 60 70

¢, COLATITUDE ANGLE
MEASURED FROM CENTER OF HOLE (deg)

Fig. 3.4.5. Circumferential Stress at Outer Surface for

an Axial Force of 1 Ib.

99
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED

ORNL-LR-DWG 33950

0 -
@ ]

o & " e

\

STRESS (psi)
oy
o
O

THEORETICAL

® EXPERIMENTAL
1400 - e

0 10 20 30 40 50 60 70

¢, COLATITUDE ANGLE
MEASURED FROM CENTER OF HOLE (deg)

Fig. 3.4.6. Circumferential Stress at Inner Surface for

an Axial Force of 1 Ib.

100

and central difference approximations were used.
The terms 82q/8x2 and dn/dx were approximated to
a fourth-order correction factor in Ax by using a
Taylor series in each case. The term dn/0 was
approximated to at least a second-order correction
factor in At

The difference equations were then used in
writing a code for the Oracle. Through the use of
this code, a solution to the partial differential
equation can be obtained when D(x,t) = D(x). The
coefficients of the equation may be of any form
provided they are continuous functions. The ac-
curacies associated with the difference approxi-
mations employed allow the use of a relatively
large increment in t. This is very useful where
solutions are required for many values of that
variable,

The code may also be used to solve the differential
equation

d*n dn
A(x) — + B{x)—+ C(x)p + D(x) =0 .
dx2 dx

3.5. ADVANCED POWER PLANT DESIGN

VORTEX REACTOR EXPERIMENTS
J. J. Keyes

Reactor Projects Division

The feasibility of the vortex propulsion reactor
concept which combines the principles of gaseous
fission heating and separation of the fissionable
material! depends strongly upon the nature and
strength of the vortex field which can be generated
at mass flow rates permitted by a binary diffusion
process. |t is shown in ref 1 that the desired
vortex field is one in which the velocity varies
inversely with radius — the so-called ideal *‘free-
vortex'' —~ as well as one in which the peripheral
Mach number is high. Allowable mass velocities
of the low-molecular-weight propellant are in the
neighborhood of 0.01 Ib/sec-ft. These require-
ments could be fulfilled, it was felt, provided the
flow remained laminar, although the influence of
viscosity was only approximated in the analysis.

An experimental study was therefore initiated in
an attempt to measure vortex strengths in a real
gas and to ascertain the effects of viscosity and
turbulence. The study to date has consisted
mainly of two phases: (1) measurements on solid-
wall tubes with all of the mass flow leaving at the
center of one end of the vortex tube, as assumed

in the analytical study; and (2) measurements with
a fraction of the total input flow bled off at some
radial position other than the tube center. In some
of the latter experiments, an attempt was made to
make the boundary layer laminar by bleeding
through a uniformly porous wall which formed the
outer boundary of the vortex tube.?

In both phases of the study, nitrogen was used
as the working fluid at room temperature. Mass
flow rates and pressures were chosen to bracket
estimated operating conditions in the actual
device by assuming similarity of the characteristic
Mach and Reynolds moduli and of the mass flow
parameter, m]/277,u, which appears in the approxi-
mate solution of the Navier-Stokes equations! and
in which m, is the mass flow per unit of tube
length in Ib/ftesec and p is the viscosity in the
same units as m,,

The apparatus for carrying out phase 1-A (Table
3.5.1) of the experiment is shown in Figs, 3.5.1

1J. L. Kerrebrock and R. V. Meghreblian, An Analysis
of Vortex Tubes for Combined Gas-Phase Fission-
Heating and Separation of the Fissionable Material,
ORNL CF-57-11-3, Rev. 1 (April 11, 1958),

2}, L. Kerrebrock and P. G. Lafyatis, Aralytical
Study of Some Aspects of Vortex Tubes for Gas-Phase
Fission Heating, ORNL CF-58-7-4 (July 21, 1958),

Table 3.5.1, Configurations and Operating Conditions of Vortex Tube Experiments

Working gas: N, at 80°F

Experimental Description of Inlet Nozzle Upstream Tube Wall Mass Flow, r, (Ib/sec-h) Ble.ed Exit Figure*
Phase Vortex Tube Geometry Header F’r‘essure, Pressure, Inlet,  Exit Tubs Bleed, Ratia, Con figuration M References
by (psia) by (psia) ! Center, E 8 B/E
T-A Plastic tube No. 1, Twelve nozzles 215 45 0.007 Same as ! 0 0 0.280-in.-0D 3.5.7, 3.5.9,
2in. ID, 12 in. long 0.0135 in. in als 85 0.011 annular nozzle, and 3,5,10
diameter at 0.187-in.-dia
0.92-in. radius 440 135 0.015 drill hole
565 185 0.020
Plastic tube No. 2, Twelve nozzies 315 65 0.0045 Same as { 0 0 0,280-in.-0D 3.5.8, 3.5.9,
2in, ID, 12 in. long 0.010 in. in 440 115 0.0066 annular nozzle and 3.5.11
diometer at
0.84-in. radius 245 165 0.0082
840 0.0097
1.B Plastic tube No. 3, Eleven holes 0.020-in. 386 52. 0.028 0.028 0 0 0.280-in. nozzls 1,95 3.5.12
1in.ID, 12in, long, in diameter, in line
with o plastic insert with nozzles at
tube 0.45-in. radius
2 Porous metal tube, Tweilve nozzles 715 68 0.028 0.028 0 0 0.280-in. nozzle 2,18 3.5.12 and
2in, ID, 12 in, long 0.0135 in. in 3.5.12
diameter at
0.92-in, radius
715 95 0.028 0.0074 0.0 2.8 0.125in. drill 1.98 3.5.13
{uniform hole
wall bleed})
740 81 0.029 0.0062 0.023 3.5 0.22-in. drill hele 2.7 3.5.14
{radial at at center, twelve
7= 0.25) 0.055-in. drill
holes at 0.25-in.
radius

101
ANP PROJECT PROGRESS REPORT

ORNL-LR-DWG 32566

B

4in
PROBE TAP (4 TOTAL)

- - AXIAL PROBE — «-00i3-in-dio NOZZLE SPACED in.,
\ PR(EéSTUORTiLTf‘P T _ 90° ROTATION (12 TOTAL)

\\ —

1 a

- [

- © 2.000-in~dia BORE

i .

SRRy '!

|
GAS SUPPLY CHANNEL
‘ (4 TOTAL)

e 12 in, —

Fig. 3.5.1. Sketch of Experimental Vortex Tube.

and 3.5.2. The geometry selected was a 1-ft
section of a 2-in.-|D vortex cavity bored into a
4-in. Lucite block. Twelve brass feed nozzles
were located in the wall of the block to form a
broken spiral of entrance jets spaced 1-in. apart
axially with a 90-deg rotation between adjacent
nozzles throughout the length of the block. The
inside surfaces of the nozzles were honed flush
with the inner surface of the tubular cavity. The
feed nozzles were supplied with gas from four
headers drilled into the plastic block so that

each header fed three nozzles in line. The arrange-
ment of feed headers and nozzles may be seen in
Fig. 3.5.2. Gas was exhausted at one end through
an adjustable annular orifice directly to the atmos-
phere. The geometrical configurations used for
the various experiments and the experimental con-
ditions are described in Table 3.5.1.

102

In phase 1-B, the vortex tube consisted of a
1-in.-ID, 12-in.-long plastic tube inserted into the
2-in.-ID tube for operation in the same test setup
as that used in phase 1-A. The tube was fed by
means of eleven symmetrically spaced, 0.020-in.
drilled holes in line with nozzles located at a
radius of 0.45 in. In order to measure the radial
static pressure distribution, a number of pressure
taps were located in the feed plate.

An over-al| flow and instrumentation diagram of
the test system is presented in Fig. 3.5.3. Helium
or nitrogen gas was supplied through pressure
requlators and appropriate metering devices to
each of the four supply headers at pressures up to
600 psig. Valves were used to control the indi-
vidual header flows. The pressure in the vortex
tube was regulated by the exit valve up to a maxi-
mum of 150 psig.
PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED)
PHOTO 31173

Fig. 3.5.2. Plastic Test Block — Vortex Tube No. 1.

103

ANP PROJECT PROGRESS REPORT

2209

EXIT
NOZZLE

ORNL-L!

\ AxiaL PRESSURE PROBE

WALL PRESSURE OR PROBE TaP

YORTEX TUBE

RADIAL PRESSURE TAPS

i i Y e

API
MANOMETER
BOARD
0-100 IN Hg
0-100 IN Hz0

AT A7, R

DESUR

PRESSURE INDICATOR
API DIFFERENTIAL PRESSURE INDICATOR

PR PRESSURE REGULATOR

FI

FLOW INDICATOR
THERMOCOUPLE

PL
TC

TRACE GAS GENERATOR

>
L
N

T
Ter
FI
F1
1L
LI
@0-600 psig
TC

"QUICK-CONNECT” COUPLINGS —&—

@0-600 psig

LAB. TEST GAGE
0-150psig

£

|

=600 psig
<

TC

0 in.
o}
g

=10
H

(5)°
I
17
FI

600 psig

—<t+—T"<

PR-2
TO He
SUPPLY

~1500 psig

1000 psig

@ G-600psig
3
l:%ZPRA
3
TO N,
SUPPLY
~1500 psig

Pl

104

TC—

Fig. 3.5.3. Over-All Flow and Instrument Diagram of Vortex Tube Test System.
PERIOD ENDING SEPTEMBER 30, 1958

The apparatus for carrying out the second phase
of the experiment (Fig. 3.5.4) was similar to that
employed in the first phase with one significant
addition: the wall of the vortex tube was made of
uniformly porous metal (sintered nickel) and it
was surrounded by a pressure jacket so that a
controllable fraction of the total throughput could
be bled off through the wall and metered. The

metal vortex tube and jacket are shown sche-
moticolly in-Fig. 3.5.5, in which the/ariaaemsnt
is similar to that of Fig. 3.5.1. The tube installed
in its jacket and in the test rig is shown in Fig.
3.5.6.

The objective of the experiments in phase 1-A
was to ascertain vortex strengths in the 2-in.-ID
tube under conditions wherein all mass flow was

UNCLASSIFIED
PHOTO 32913

AR R AR RRAGARRRRERVARRE

=1
ORNL-LR-DWG 34775

-PROBE TAP BLEED FLOW
(6 TOTAL) EXHAUST
POROUS NICKEL TUBE BACK-PRESSURE | ADJUSTABLE
2-in. 10 x Ya-in. WALL JACKET ’ EXIT NOZZLE
=
r
3
L]
e e — = |
=
AN \li

PRESSURE TAP (20 TOTAL)
POSITIONED RADIALLY

LGAS SUPPLY| HEADER
(4 TOTAL)

oorssin-om sverr nozae
(12 TOTAL)

4
FT %5.5. Sketch of Porous Metal Vortex Tube and Jacket,

105

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED)|
PHOTO 32915

Fig. 3.5.6. Metal Vortex Tube Installed in Experimental Facility, WSssm® with caption)

expelled radially at the tube center, for which case

- the flow had to be maintained sufficiently low to
conform to the limitations imposed by the binary
diffusion process in the device. The results are

, depicted graphicaily in Figs. 3.5.7 through 3.5.11,
in which the measure of vortex strength is the
ratio M /M].,where M, is the peripheral tangential
Mach number, and M. is the inlet jet Mach number.
The mass flow, m,, was varied by changing the
nozzle inlet pressure and the diameter of the jets.
The value of M, is determined by fitting the meas-
ured pressure d't;sfribution data to the derived
relationship:

p (v -1y

b, 2 LAY
P
where p is the absolute pressure at position r“=
r/ro, p,, is the pressure at the tube wall, and y is
the heat capacity ratio. Since the equation

assumes a free vortex, the values of Mp are only
approximate.
. S
ORNL-LR-DWG 32295
0.07
L
0.06 =
° f
o
0.05
a
o
0.04 s .
My a
Mf
0.03
TUBE NO. !
WALL PRESSURE, psia
o 65
0.02 - H5
a 165
0.0
0
0 0.004  0.008 0.012 0.016 0.020
7, (Ib/sec-ft)

Fig. 3.5.7. VYariation of Ratio of Yortex Tangential
Mach Number and Jet Exit Mach Number with Mass Flow
¢ Rate per Unit of Tube Length and Tube Pressure for
Tube No. 1.

PERIOD ENDING SEPTEMBER 30, 1958

dtiien oy
ORNL-LR-DWG 32294
0.04
b ©
®
g 2 -
0.03 > 5
4
TUBE NO. 2
Mp WALL PRESSURE, psia
— Q.02 —
M o 65
[ ] 115
A 165
0.0
o}
0 0.004 0.008 0.0t2 0.0t6 0.020
7 (Ib/sec-ft}

Fig. 3.5,8. Variation of Ratio of Yortex Tangential
Mach Number and Jet Exit Mach Number with Mass Flow
Rate per Unit of Tube Length and Tube Pressure for
Tube No. 2.

ORNL—LR—DEG 34571

0.8

1

|
I
1
07 -flm‘%—* —"— TUBE NC.1, 777,= 0.02 Ib/sec-ft
|

£,= 585 psia; p,= 85 psia
MJ’,-=1,9O

08 &= Tuse NO. 2, M, = 0.0097 Ib/sec-ft  —
p,= 640 psio;, p, = 65psia

M/-32.16

0.5

DASHED CURVES ARE FOR IDEAL
"FREE VORTEX"

M/A/Jlf. 04

0.3

0.2

o 0.2 0.4 0.6 0.8 1.
RADIAL POSITION, r'

Fig. 3.5.9. Mach Number Profiles for 2«in, Plastic
Vortex Tubes.

107

ANP PROJECT PROGRESS REPORT

i ORNL-LR-DWG 32293 The local Mach number, M, at r’, is obtained
' directly by differentiation of the pressure data as
applied to a relationship for circular flow of an
. ideal gas that neglects the small contributions
0.8 . .
° ° due to a radial velocity term:
a [w)
L ]
a 2 1 r’ dp rd
0.6 y p’ dr’
e where p”is the ratio p/p_ .
In phase 1-B of the experiments, the diameter of
0.4 TUBE NO 1 the tube was decreased to 1 in., and the experi-
waLL PREE?URE (psia) mental Mach profile shown in Fig. 3.5.12 was
o 15 obtained. Also shown is a result from a phase 2
0.2 o 165 7 experiment with the 2-in. metal tube under nearly
the same conditions.
0
0 0.004 0.008 0.012 0.016 0.020
m1 (Ib/sec - ft) 16 ORNL-LR-DWG 34572
. . . . +e
Fig. 3.5.10. Variation of € in the Expression v oc r |
|
with Mags Flow Rate per Unit of Tube Length for Tube 4 ]
No. ]o ll
iR
12 i
Y + \\
L \
10 ORNL-LR-DWG 32292 \ | —o— PLASTIC TUBE No.3, fin. 1D
' \ M= 0028 Ib/sec-ft
‘ 1
1.0 \ PO = 386 psio; £,= 52 psio
\ My =196
h\
0.8 o \\ —e— METAL TUBE, 2in. ID
MM, 0.8 X 71, = 0028 Ibfsec-ft
o P . \\ A = 745 psie, £, = 68 psic
\  M-2m
0.6 ‘ \\ (REPLOT FROM Fig. 3.5.13)
E o o 06 A
e \
) N
o N " "
04 \ \\,\,.FREE VORTEX
0.4 b NS
TUBE NO. 2 N
WALL PRESSURE (psic) N
(o] 65 e
o 15 SN
o 165 0.2 ‘\\"‘"---_(_ —
o \/h-________"
0 0.004 0008  0.012 0.016 0.020 .
77?1 (Ib/sac - ft) 0 02 0.4 06 0.8 1.0
RADIAL POSITION,
+
Fig. 3.5.11. Variation of € in the Expressiony o r €
with Mass Flow Rate per Unit of Tube Length for Tube
No. 2,

Fig. 3.5.12. Comparison of Mach Number Profiles for
1=in.=ID and 2-in,-ID Vortex Tubes with no Bleed Flow.

108

The objective of the phase 2 experiments was to
investigate the effect of two methods of bleeding
off a fraction of the inlet flow at some radial
position away from the tube center. In the first
method, the bleed was from the tube periphery
through a uniformly porous wall, in the hope that
sufficient boundary layer stabilization could be
achieved to significantly lower the eddy losses
induced by shear forces at the wall and thus in-
crease the vortex strength. The Mach profile
obtained with a uniform wall bieed ratio of 2.8 is
compared in Fig. 3.5.13 with that obtained with no
bleed at the same inlet mass flow. Results are
presented in Fig. 3.5.14 for an alternate method
of bleed in which the excess flow is removed
radially at a position 7 = 0.25 so that the exit

ORNL - LR-DWG 34573
0.9
0.8 4
a =0 M, = M, = 0.028 Ib/sec- ft

0.7 |

& |

\
1
l \
06 il |

|I
|‘\ :—:2: m] = 0.028 |b/SeC « ft

ALL FLOW RADIAL AT TUBE CENTER
By = 715 psia; A, = 68 psia

M = 248

‘\ Mg = 0.0074 lo/sec - ft
)
0.5 \‘l UNIFORM WALL BLEED RATIO = 2.8 ]
M/M, \ -
/M, \ M, = 1.98
‘\
04 \
\
) DASHED CURVES ARE FOR IDEAL
\\ "FREE VORTEX"
0.3
0.2
/
O % ‘M
' —&< ;
0

0 02 0.4 0.6 0.8 10
RADIAL POSITION, r/

Fig. 3.5.13. Mach Number Profiles for o 2-in.-ID Metol
Vortex Tube Showing Effect of Uniform Wall Bleed.

PERIOD ENDING SEPTEMBER 30, 1958

ORNL-LR-DWG 34574
0.5
0.4 | /%
7= 0029 Ib/sec ft
M= 0.0062 Ib/secft

Po=T40 psic;pw= 61 psia
M;=2.27

\
0.3 \

MM, \

c.z2 \

.

o \_____ .AP.\

0 0.2 04 0.6 0.8 1.0
RADIAL POSITION, r'

Fig. 3.5.14. Mach Number Profile for o 2«in.«ID Metol
Vortex Tube with Radial Bleed 0,25 in. from Center,

flow at r” = 0 is held within allowable limits.
The main effect here is that of increasing the
inlet mass flow parameter m /27y over a large
fraction of the radius.

The foilowing general observations from the
data are noteworthy:

1. The vortex strength, M_/M ., for a 2-in.-diam-
eter increases from about 0.63 for plastic tube
No. 2 with m, = 0.0045 Ib/sec-ft (Fig. 3.5.8) to
about 0.08 for the metal tube with m, = 0.028
Ib/sec-ft (Fig. 3.5.13). Thus, the vortices which
were generated at low mass flows appear to be
significantly weaker than would be required in
vortex reactor applications. The plots of Fig.
3.5.9 for the 2-in. plastic tubes show that the
maximum value of M/M. which can be achieved at
low mass flows is severely limited. The small
peak in the profiles which occurs near the wall
for all the 2-in. tubes is not understood but may
be related to the nature of intfroduction of the
gaseous |et.

2. As indicated in Figs. 3.5.10 and 3.5.11,
the value of the exponent € in the relationship
v« r*€, where v is the tangential velocity, in-
creases from —0.4 at m, = 0.0045 Ib/sec ft to
about -0.8 at m, = 0.011 Ib/sec-ft; a further in-
crease in m, appears fo produce little change
in €, Since deviation of € from —1 is a measure
of the deviation from the ideal ‘‘free vortex,’’ it is

109
ANP PROJECT PROGRESS REPORT

concludgd that although the vortex strength is low
the velocity distributions roughly approximate free
vortex behavior at the higher mass flows.

3. The effect of decreasing the tube diameter
from 2 to 1 in., as illustrated in Fig, 3.5.12, is
to increase significantly the velocity at a given
radial position, 7”. [n the 1-in. tube a value of
M/M . of 1.25 was achieved (M = 2.45) at "= 0.085;
whilé at the corresponding absolute radius in the
2-in, tube at 7 = 0.043 in., M/M.= 0.6 (M =1.31).
At a radius of 0.25 in., the 1-in. tube produced
M/M. = 0.25, while for the 2-in. tube M/M.=0.27.

], 1.
Thus it appears that by decreasing the radius
the performance of a given vortex tube is not
impaired, and the advantage is gained of enabling
more tubes to be included in a given volume of
reactor,

4. When a uniform wall bleed is applied at a
given inlet mass flow, as illustrated in Fig.
3.5.13, the effect is to increase slightly the
velocity near the periphery of the tube at the
expense of that nearer the center. The rapid rise
at r* <0.06 for the profile with bleed is a conse-
quence of the small exit hole employed in this
run as contrasted with the larger annular opening
in the run with no wall bleed (see Table 3.5.1).
This effect is discussed in a forthcoming report.3
It is felt that the failure of wall bleeding to
achieve a significant improvement may be due
to the high degree of turbulence induced by the
jets in the vicinity of the wall so that a laminar
boundary layer cannot be achieved with practical
bleed ratios.

5. As an alternate method of bleed-off, a run
was made in which the excess mass flow was
removed at a radial position 0.25 in. from the
center of the 2-in. metal tube; no wall bleed was
employed. The results are depicted in Fig.
3.5.14, which indicates when compared with
Fig. 3.5.13, that for 7 > 0.1 the velocity profile
agrees well with that obtained when all the flow is
removed at the tube center. This technique shows
promise for increasing velocities and vortex
strength at the expense of lowering exit mass flow
rates.

In addition to the basic experiments discussed
here, numerous runs have been carried out to

3). L. Kerrebrock and J. J. Keyes, Jr., An Experi-
mental Study of Vortex Tubes for Gas Phase Fission
Heating — Part I, ORNL CF-58-7-5 (July 5, 1958).

110

study the effects of number, size, and location of
the inlet nozzles, tube length, methods of bleed-
off, gas properties, and pressures. Some pre-
liminary separation experiments have also been
made using He-Br, and He-C Flé gas pairs, ds
discussed in ref 3. Since these results are not
conclusive, it is intended to continue the separa-
tion studies.

SURVEY OF DESIGN PROBLEMS OF AUXILIARY
POWER UNITS FOR SATELLITES

A. P. Fraas

Reactor Projects Division

The performance of various types of auxiliary
power unit that might be used in satellites has
been reviewed during the past several months.
Informgtion from the SNAP (Systems for Nuclear
Auxiliary Power) program indicates that the weight
of either solar cells or thermoelectric (or thermo-
ionic) power units employing radicactive isotopes
as a heat source for a reconnaissance satellite
will be about 500 to 1000 Ib/kw. While lighter
weights (350 Ib/kw) might be achieved with solar
cells continuously oriented toward the sun,
satellite missions which involve spending almost
half of the time in the earth’s shadow immediately
reduce the efficiency of such solar cells by
roughly 50%. Further, it is difficult to effect the
orientation of the solar cells toward the sun when
mission considerations entail orientation of other
elements of the satellite toward the earth. In
addition, the specific weight of either the solar
cells or the radioisotope power sources is not
reduced as the power output of the unit is in-
creased, and the voltage obtainable from such
units is low. From preliminary estimates it
appears that the use of a reactor as a heat source
might make possible auxiliary power units of
only 20 to 40 |b/kw for power outputs in the range
of 20 to 200 kw net electrical output.

Cycle Performance Considerations

One SNAP || power plant proposal suggests
three fluid circuits: a sodium circuit cools the
reactor, mercury serves as the thermodynamic
cycle working fiuid, and terphenyl is used to
transfer heat from the condenser to the radiator.
This approach entails three heat exchangers,
three pumps, three expansion tanks, and three
sets of instruments and controls. A very large
improwament in reliability and a substantial
saving in weight should be possible through the
use of a simpler system employing a single fluid.
In the mercury-vapor cycle used in both the
SNAP | (radioisotope source) and SNAP ||
(reactor) power plants, there will probably be a
substantial holdup of droplets of mercury on the
condenser surfaces, and therefore the weight of
the mercury in the condenser will probably be
greater than the dry weight of the condenser. It
would also be desirable to increase the operating
temperature level in the condenser to save con-
denser weight, and this is not practicable be-
cause it would imply too high a pressure level
for the high-temperature portion of the mercury-
vapor cycle. Further, the low specific volume
of the mercury vapor leads to so small a turbine
that the shaft speeds tend to be high, and the
aerodynamic efficiency of the turbine is low.
Cycles employing either aluminum chloride or
rubidium vapor should have several important
advantages relative to mercury for a light-weight
power plant, In both cases, a single fluid can
be used to cool the reactor and carry out the
thermodynamic cycle. Flow diagrams for power
plants employing these cycles are presented in
Figs. 3.5.15 and 3.5.16. The thermodynamic
properties of these fluids have been estimated
and reports are being prepared to present the

UNCLASSIFIED
ORNL—LR—DWG 34575

EVAPORATING -
DRUM / TURBINE

I X
Y
GENERATOR
REACTOR
!
COMNDENSER
‘ FEED PUMP
CIRCULATING
PUMP

Fig. 3.5.15. Flow Diagram of Nuclear Power Plant
Employing o Rubidium-Vapor Cycle. Wimany with caption)

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
ORNL—LR—DWG 34576

GENERATOR

S
// TURBINE

COOLER

REACTOR

| COMPRESSOR

Fig. 3,5.16. Flow Diagrom of Nuclear Power Plont
Employing an Aluminum Chloride Gos Cycle, (MDY

with caption)

resulting data. Some preliminary calculations for
some typical cycles have been made and are in-
cluded below.

The rubidium vapor cycle would operate much
like a boiling-water reactor with a pump circu-
lating a weight flow through the reactor of
roughly ten times the weight of rubidium vapor-
ized per pass. Of course the vapor bubbles
formed in the reactor will pose serious reactivity
and control problems, but it seems likely that
the reactor can be designed to reduce these
effects to tolerable levels.

A gas-turbine cycle utilizing aluminum chloride
has some unusual characteristics. Aluminum
chloride dissociates from Al,Cl, to AICI, in the
temperature range between 700 and 1600°F, with
the bulk of the dissociation taking place in a
400°F range that increases with pressure. Pre-
liminary calculations indicate that the cycle
can be designed so that the gas will be mostly in
the form of Al ,Cl, during compression, while
during the expansion process it will be mostly
AICIl,. This, in effect, will cut the compression
work in half and thus produce a marked improve-
ment in cycle efficiency. The nature of this
effect can be visualized by examining the P-V
diagrams of Fig. 3.5.17, which compare similar
ideal gas turbine cycles for helium and aluminum
chloride. The work involved in each compression

11
ANP PROJECT PROGRESS REPORT

RANKINE CYCLE (RUBIDIUM)

BRAYTON CYCLE (Al,Clg—AICIg)

ORNL— LR-DWG 34577

BRAYTON CYCLE (HELIUM)

° ] C] \
g 2 g N\
w @ & \
3 2 @
§ IDEAL WORK (NPUT = 0.45 Btu/Ib ? IDEAL WORK INPUT =25.1 Btu/Ib & IDEAL WORK INPUT = 477 Biu/Ib
o o 1l
a a a
Y
| ! | | [ I i T T [ y

IDEAL WORK OUTPUT = 144 Blu/Ib IDEAL WORK OUTPUT = 627 Btu/Ib \
3 3 2
g \ 2|\ g N
N 0 \ "
2 \ 2 ?
& \ & % IDEAL WORK OUTPUT = 604 Btu/Ib
5 N & \\ &

\_{-—_.__ \1
P
[ [ ! [ | I ! [ I |

IDEAL NET WORK =113.8 Btu/Ib IDEAL NET WORK = 37.6 Btu/Ib \
3 3 2
oA e : O
& i \ W \
> \ D =)
[72] w (72}
& \ & \ \ @ | IDEAL NET WORK =127 Btu/lb

\_ \\‘-—_____

VOLUME (ft3/1b)

VOLUME (ft3/1b)

VOLUME (#t3/1b)

Fig. 3.5.17. P-V Diagrams for Typical ldeal Thermodynamic Cycles (Data from Tables 3.5.2, 3.5.3, and 3.5.4).

or expansion process is directly proportional to
the area of the P-V diagram, and the net work is
proportional to the net area for the cycle. A
Rankine cycle utilizing rubidium vapor is also
included in Fig. 3.5.17 to show that the proposed
aluminum chloride cycle is roughly midway be-
tween a gas turbine (or Brayton) cycle and a
Rankine cycle in its requirements for work input
during the compression process.

The diagrams of Fig. 3.5.17 were prepared for
ideal cycles with no ailowances for losses. The
most important losses are associated with the
efficiencies of the compressor and the turbine,
which are likely to be of the order of 80%. This
means that with an 80% efficient compressor the
ideal work input will be 80% of the actual work
input, while the actual work output of the turbine
will be only 80% of the ideal. In addition, pres-
sure drops between the compressor and the

112

turbine will cause losses in net output from the
cycle. The nature of these effects can be seen
readily in Fig. 3.5.18.

If allowances are made for these losses to
obtain the actual net outputs for the cycles of
Fig. 3.5.17, the diagrams of Fig. 3.5.19 result,
The relatively large work input required for the
compression process of the Brayton cycle makes
it very sensitive to compressor inlet temperature,
because the compression work increases rapidly
with the initial temperature. As a result, the net
work output and over-all thermal efficiency of
the Brayton cycle drop off so rapidly with in-
creasing temperature at the compressor inlet that
for most plants the compressor inlet temperature
must be held below 200°F. At the same time, the
turbine inlet temperature must be at least 1200°F,
and preferably should be above 1400°F. For the
helium cych_gifig. 3.5.19, even with a 1540°F
- a4 UNCLASSIFIED
ORNL-LR-DWG 34727
! I I T I
80 PRESSURE LOSS THROUGH HEATER
TURBINE INLET TEMPERATURE =

1500°F

60 ! t
r—NET EITFECTIVE AREA
INE

40
20 [EDDY LOSSESIN A

COMPRESSOR —

‘ Tl-llROUGH CIIOOLER

I l I
'PRESSURE LOSS THROUGH HEATER

|

1
TURBINE INLET TEMPERATURE =

w
(@]

PRESSURE {psi)
(o]

| 1200°F
60 | i
NET EFFECTIVE AREA
40 EDDY LOSSES IN TURBINE

EDOY LOSSES IN
20 I~ COMPRESSOR “1

PRESSURE LLOSS
THROUGH COOLER

0 a4 8 {2 16 20 24 28
SPECIFIC VOLUME (ft¥1b)

Fig. 3.5.18. P-V Diagrams for ldeal Gas-Turbine Air
Cycles with Cross-Hatched Areas to Indicate the Magni-

tude of the Principal Losses.

turbine inlet temperature, the net work from the
cycle is negative when the compressor inlet tem-
perature is raised to the level required to give a
reasonable radiator specific weight.

The aluminum chloride possesses another
important advantage over conventional gases for
use in a gas turbine; the heat transfer coeffi-
cient during the heating and cooling processes
should be exceptionally high. The large amounts
of energy involved in dissociation appear to give
effects analogous to those responsible for the
high heat transfer coefficients characteristic of

PERIOD ENDING SEPTEMBER 30, 1958

boiling and condensing heat transfer conditions,
This should reduce the heat fransfer area re-
quired in comparison with the area required for
gases such as helium or nitrogen.

A further advantage of aluminum chloride is that
it could be used in the vapor phase throughout the
cycle so that there would be no free liquid surface
at any point in the system, and hence no expansion
tank would be required. While it is possible that
a free liquid surface could be stabilized by surface
tension forces or other factors, it is also true that
a great deal of difficulty has been experienced at
ORNL with stabilization of free liquid surfaces
in pumped systems, To date no one has yet sug-
gested a means of carrying out laboratory tests
which would approximate free-flight conditions.
Thus there is a major advantage to the use of a
working fluid which would avoid any difficulties
with liquid surface stabilization.

Summaries of the calculations on which Figs.
3.5.17 and 3.5.19 were based are presented in
Tables 3.5.2, 3.5.3, and 3.5.4.

Planned Program

The next phase of this work will entail a com-
prehensive series of analytical studies to de-
termine the performance obtainable from rubidium
and aluminum chloride as working fluids as com-
pared with more conventional fluids such as
mercury, sodium, helium, air, and water, The
studies will also disclose the major problem areas
most deserving of attention. If this phase of the
work indicates a substantial superiority for either
rubidium or aluminum chloride as a working fluid
in an auxiliary power plant unit it will be fol-
lowed by experimental determination of heat
transfer characteristics of the working fluid
chosen, investigation of corrosion problems
associated with the fluid, and more complete de-
sign studies,

113
ANP PROJECT PROGRESS REPORT

RANKINE CYCLE (RUBIDIUM)

BRAYTON CYCLE {Al»Clg-AICl3)

e rnd
ORNL-LR -DWG 34578
BRAYTON CYCLE (HELIUM)

ACTUAL WORK INPUT = 0.3 Btu/1b ACTUAL WORK INPUT = 31.2 Btu/Ib
% @ @
] ] a
w w w
@ 1 [1ef
=} 2 pn }
@ & @ ‘ |
w v w
w w W
14 @ [1 4
a a %@h Q@ — ACTUAL WORK INPUT =597 Btu/Ib ——
R—
[ |
1 T | | | I l
ACTUAL WORK QUTPUT =79.8 Btu/Ib ACTUAL WORK OUTPUT=50.2 8tu/Ib
2 2 z
w w E
5 5 z
w ()] w
2 Q @ ] ‘
w l&J w
T a : ,‘% &l ACTUALWORK QUTPUT= 493 8tu/ib —|
1 H l
% ACTUAL NET WORK=79.5 Btu/Ib B ACTUAL NET WORK=19Btu/Ib ACTUAL NET WORK=-114 Btu/Ib
— -—_ | —
‘@ ‘@ ‘@
s & 2
& % & &
o 2 =}
w w w
(73] w w
& %\ 2 L
Q. Q o
%%h—.

VOLUME {ft/1b)

VOLUME (£t ¥1b)

VOLUME (ft3/ib)

Fig. 3.5.19. P-V Diagrams for Typical Ideal Thermodynamic Cycles with CrosssHatched Areas to Represent the
Losses Entailed by Compressor and Turbine Efficiencies of 80%.

Table 3.5.2. Rubidium Vapor Cycl

-]

Temperature Pressure, Specificm Enthalpy, b Quality Change in
Condition (°R) b (psia) Volgme, v (Bty/1b) (%) Enthalpy, Ab
(ft~/1b) (Btu/Ib)
Yapor to turbine 2060 66.33 3.892 498 100 417
Yapor to condenser 1460 2.66 384 114
(isentropic)
Yapor to condenser 1460 2,66 60.5 418 88.1 79.8
(70% etficiency)
Liquid to feed pump 1450 2.46 0.01275 81 0
79.8
Cycle efficiency = —= 19.1%
417

114

PERIOD ENDING SEPTEMBER 30, 1958

Table 3.5.3. Aluminum Chloride Cycle

(isentropic)

Cond Temperature  Enthalpy, Enthalpy, Pressure, VSpIeCIflc- DProportior; Eha:gle "
ondition o o \ olume, v issociated, nthalpy,
("R) h (Btu/Ib) S (Btu/"F) (psia)
(Bro P (£13/1b) x Ab (Btu/Ib)
Compressor inlet 1200 203.48 0.0828 5 10.346 0.0726
Compressor outlet 1356 228.6 0.0828 60 0.967 25.1
(isentropic)
Compressor outlet 1378 234.9 0.0873 60 0.992 0.0889 31.2
(80% efficiency)
Turbine inlet 2000 490.2 0.2388 60 2.516 0.8782 255.3
Turbine outlet 1653 427.5 0.2388 5 24.46 0.8410 62.7
(isentropic)
Turbine outlet 1685 440.0 0.2458 5 25.40 0.8755 50,2
(80% efficiency)
37.6
ldeal Cycle Efficiency = —— = 14.37%
261.6
19
Actual Cycle Efficiency = —— = 7.43%
255.3
Table 3.5.4. Helium Cycle
Temperature Pressure, Specific
Condition (°R) b (psia) Yolume, v
(#13/1b)
Compressor inlet 1200 100 7.2
Compressor outlet 1582 200 4,75
(isentropic)
Turbine inlet 2000 200 6
Turbine outlet 1517 100 9.1

483 — 382 101
ldeal Cycle Efficiency = ———— = — = 24.2%
2000 - 1582 418

386 - 478 =92
Actual Cycle Efficiency = =
418 - 96 322

115

Part 4
SHIELDING

E. P. Blizard

Neutron Physics Division
4.1, SHIELDING THEORY

ANALYSIS OF NEUTRON PHYSICS DIVISION
EXPERIMENTAL STUDY OF GAMMA RAYS
PRODUCED BY NEUTRON
INTERACTIONS IN AIR

F. L. Keller 0. S. Merrill!

The gamma-ray dose rate which results from
secondary gamma rays produced by neutron inter-
actions with the air can be neglected in most
shield design problems because it is small com-
pared with the dose rate which results from air-
scattered source gamma rays. Situations may
arise, however, in which it is desirable to allow
a large amount of neutron leakage from the reactor
shield and place part of the neutron shielding
material around individual regions which must be
further shielded. In many of these situations the
air-scattered gamma rays, which are usually of
rather low energy (that is, <l Mev), are attenuated
very rapidly by the shielding material around the
individual region, whereas the neutron-induced
gamma rays, which are isotropically emitted with
rather high energies (up to approximately 10.8 Mev),
penetrate the shielding material much more readily.
Therefore, as the thickness of the shield around
the individual region is increased, the relative
importance of the neutron-induced gamma rays is
increased and it is possible, in some instances,
for these gamma rays to be the major contributors
to the dose rate.

The two important methods of secondary gamma-
ray production in air are neutron capture in N4 and
inelastic neutron scattering with both nitrogen and
oxygen. In order to accurately calculate the dose-
rate contributions from these two processes, it is
necessary to know the cross sections involved and
the spectra of the gamma rays produced. The re-
sults of experimental measurements of the thermal-
neutron radiative capture cross section of N'4 and
the associated capture gamma-ray spectrum have
been reported by Bartholomew and Campion.? Very
little experimental information is available on
either the cross sections or the associated gamma-
ray spectra for inelastic neutron scattering with
nitrogen and oxygen, but Lustig, Goldstein, and

IOn assignment from Convair, San Diego, Calif.

2. A, Bartholomew and P, J, Campion, Can. J. Phys.
35, 1347 (1957).

Kalos 34 have published theoretical values of the
inelastic neutron scattering cross sections of
nitrogen and oxygen from which it is possible to
estimate the spectrum of inelastic scattering gamma
rays produced from a given fast-neutron distribu-
tion, These cross-section calculations are, how-
ever, based on a mode| which assumes that a
statistical model of the nucleus may be applied to
the compound nucleus, and the results are question-
able for nuclei as light as those of oxygen and
nitrogen, Thus, it was considered important to
make a comparison between the shape of a meas-
ured inelastic scattering gamma-ray spectrum and
that of one calculated with the use of the theoreti-
cal cross sections for at least one situation.

An experiment was therefore performed at the
Tower Shielding Facility (TSF) that was designed
to separate out and measure as accurately as
possible the spectra of gamma rays produced in air
from neutron captures in N'4 and inelastic neutron
scattering.® This experiment, in which a 3- by
3-in. sodium iodide crystal spectrometer was used,
yielded two pulse-height distributions, one of which
was attributed to capture gamma rays and the other
to inelastic scattering gamma rays. These experi-
mental pulse-height distribution curves are shown
in Fig. 4.1.1. In order to eliminate background
counts from the measurements of neutron interac-
tions in the crystal, it was necessary to make the
gamma-ray measurements inside a thick shield
through which a water-filled collimator was inserted,
Therefore the gamma rays which were detected had
penetrated through the thick (4 ft) water colli-
mator, and the measured pulse-height distributions
had been considerably altered from the desired ones
by the water attenuation and buildup.

As may be seen, it would be an almost impossible
task to take the measured pulse-height distribu-
tions and attempt to work backwards to obtain the

3H. Lustig, H. Goldstein, and M. H. Kalos, The
Neutron Cross Sections of Nitrogen, NDA-86-1 (June
30, 1957).

4, Lustig, H. Goldstein, and M. H. Kalos, An
Interim Report on the Neutron Cross Sections of

Oxygen, NDA-086-2 (1958). :

C. E. Clifford, V. R. Cain, and F. J. Muckenthaler,
ANP Quar. Prog. Rep. Dec. 31, 1957, ORNL-2440,
p 285; see also V. R. Cain, F. J. Muckenthaler, and
C. E. Clifford, Study of Gamma Rays Produced by
Neutron Interactions in Air, ORNL CF-58-4-77 (May
29, 1958).

119
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED

2—01—056—-22—596
20 <l)
or o]
10
0
o
*® [e]
5 S -—
*ol i i _
% - 1 CTTT T ) T
| 1 b
0 .: o %%090 ‘ j | o ; : : e T
] i } : |
. Oooto i |t NO BORAL ON REACTOR COLLIMATOR
. °r°°¢o o%é% v MINUS BORAL ON REACTOR COLLIMATOR
. . i AV SRR B T T T
e I B . L | |
. - } B S S DR R N ——
- : A I
0.5 -ib Ot . N
° o ‘ !
Lo J - o i T ol {
[ ] %O
o 1
0.2 A L 'y I U N oo, o GDQOJ 96,,4 S
L i I 00
. ! | 070) e oo° I
.".O.. ; : i ° % !
- lb... ! [s] .
g o4 - . O e
c "o ] ]
g '_..
(&3 [ ] B T
§ . 005 . lr t T B o ]
2 3 / . [EUR S S (L . PO
'; ———— - J PR PR N R SR > L
= INELAST!IC-SCATTERING GAMMA RAYS / . o
g -. PR JE—
o c
0.0? D . T e
° l : i
| o
®
0.0 .. 'Q. _ . ]
- 8 i s b e ]
0.005 |—+ P .
S |
| _ o ]
0.002 -t :
! *
* |
0.004 - e SO S SRS s s SR —
T , S S ——— RS
0.0005 S - ; ; AN S - f T—
‘ . i s . | ]
0.0002 -
| |
0.0004 L
0] 100 200 300 400 500 600 700 800 200 1000 1100 1200

PULSE HEIGHT

Fig. 4,1.1, Pulse-Height Spectra of Capture and Inelastic Scattering Gamma-Rays Resulting from Neutron Inter-

actions in Air.

120

spectra of gamma rays incident on the outside of
the collimator. Therefore, it was considered ad-
visable to make the comparison by first calculating
the spectra incident on the outside of the water
collimator from both capture and inelastic scat-
tering gamma rays, then determining the attenuation
and buildup associated with the passage of the
radiation through the water collimator, and, finally,
taking into account detector response functions to
obtain calculated pulse-height distributions which
could be compared directly with the experimental
results presented in Fig. 4.1.1.

Spectra of Gamma Rays Incident on the Outside
of the Collimator

A rather accurate estimate of the complete spec-
trum of gamma rays from neutron captures in N4
was obtained by making use of the direct experi-
mental measurements of Bartholomew and Campion,
The report by Lustig, Goldstein, and Kalos on
neutron cross sections of nitrogen3 supplied the
needed theoretical values of the total inelastic
scattering cross sections for neutrons with energies
up to 18 Mev and also the results of Hauser-
Feshbach calculations of the cross sections for
exciting the first four individual nuclear levels
with various energy neutrons up to 6 Mev, Since
these individual level cross sections are relatively
flat in the region of 6 Mev, they could be extrapo-
lated to somewhat higher neutron energies. Then,
for neutron energies above 6 Mev, the difference
between the total inelastic cross section and the
sum of the individual level cross sections repre-
sents the cross section for exciting some level
whose energy is greater than 6 Mev. Branching
ratios obtained from a review article by Ajzenberg
and Lauritsen® were used with the individual level
cross-section curves to obtain cross-section curves

2

for producing gamma rays with discrete energies
less than 6 Mev from these levels. The branching
ratios also indicated that most of the levels in the
region from 6 to 9 Mev decay directly to the ground
state, Since these levels are rather closely spaced
and since they decay predominantly to the ground
state, it was assumed in calculating the gamma-ray
spectrum that the difference cross section men-
tioned above could be treated as a cross section
for producing an inelastic scattering gamma ray

6F. Ajzenberg and T. Lauritsen, Revs., Modem Phys.
27, 77 (1955).

PERIOD ENDING SEPTEMBER 30, 1958

with an energy approximately equal to the initial
energy of the neutron. In the case of oxygen the
first, and most important, level is at 6.09 Mev, and
it was assumed that the entire inelastic oxygen
cross section could be taken as the cross section
for producing a 6.09-Mev gamma ray.

With these cross sections and an assumed neu-
tron distribution it was then possible to estimate
the spectrum of the inelastic scattering gamma rays
produced. The spectrum of fast neutrons above 3
to 4 Mev (which are the only neutrons of importance
for inelastic scattering in air) was assumed to have
a fission shape in these calculations. The spec-
trum of inelastic scattering gamma rays obtained in
this manner is shown in Fig. 4.1.2, where the
discrete energy gamma rays have been represented

( 10—26) UNCLASSIFIED

X 10 ORNL-LR-DWG 31793
5
2 +—nH

s a s S S SN

l

O e S

o
[ 1]
U
S~
Y -7 e [ ]
§ / i
I 1 - 1
3 ‘4 ’
2 3 | H 1 iy !
S |
o 4.4 /-
I 1 04 ’ 7z
x ZAZ A 7
<?
z NN 7 é ’
] ?
4 m A 1
& ! ’ /2 ; ’ Z
= T I/
= e E s B
~ /582050007 7, B/
2 THTAW A b Zmn;
S FHddIE 7y
o AR R ZRL
Rl /2 an /R Y
N b B A 7 fi"
5 - d4laplp a1
g el Jhn
’ ‘A’ 1 1] R
T o siizi/Bl/7 7y TR, ]
HEULTVE LB\
. HIBA U ¢ 1MH >
s | duagd AN
iKY [ \
D|0'2 .7 M?L’w_.;_fifi_.fié_fi’ AN k
7750 . — - ANNAN ANSANN:Y
- 7E 2R - — - NN AN -
_._‘fifi_ _/__?‘__fifi%_j ; \\\\\Q NN \\
--EZE700/ 7 . RN DRRNAY
IZEN707007007 7 202 I SN\ ANWNY
S THHUIH B 0 AR N
I+ HH41—A />_ -
7 1 1 ‘N Ty
o qadidg 7 a0 N\ N
WU EBRZEEINN N
AT EINN
2 I AN % g NN
‘U’ BRZ R
HUNRBRZRY
AR B A GH
A il ’ YA HH
1 2 3 4 5 6 7 8 9 10 #H 12 13

£y, GAMMA RAY ENERGY (Mev]

Fig. 4.1.2, Calculated Inelasfi'ic Scattering Gamma-
Ray Number Flux Spectrum at the Detector (No Water

in Collimator),

121
ANP PROJECT PROGRESS REPORT

in histogram form with an energy width of 0.2 Mev.
The fact that the calculated distribution above

6 Mev is continuous rather than discrete, is, of
course, due to the lack of detailed knowledge con-
cerning the excitation of individual levels in this
region. Fortunately the number of gamma rays in
the continuous portion of the curve is small com-
pared with the number of gamma rays in the dis-
crete lines, and hence the approximations used in
the calculations concerning these gamma rays are
probably sufficiently accurate.

Spectrum After Attenuation and Buildup
in the Woter Collimator

Since all gamma rays which reached the detector
without making a collision in the water collimator
travelled through approximately the same water
thickness (™4 ft), the uncollided spectra at the
detector were determined by simply multiplying
the incident spectra by e ““(E)T gt each energy
point, where u(E) is the total linear absorption
coefficient of water at energy E and T is the water
thickness (~4 ft), Next, it was necessary to
determine the distribution of scattered, or buildup,
gamma rays at the detector for each case. Since,
as mentioned above, all the uncollided gamma rays
which reached the detector penetrated essentially
the same water thickness, it was assumed that
the distribution at the outside of the water coili-
mator could be artificially represented by a single
point isotropic source whose strength was nor-
malized to give the same uncollided flux at the
detector as that calculated above.

In the case of gamma rays it appears that even
for point isotropic sources in infinite media most
of the scattered gamma rays which reach a detector
are gamma rays which leave the source in almost
the right direction to reach the detector without
scattering. Goldstein and Wilkins? have calcu-
lated the spectra of buildup gamma rays at a de-
tector for cases of point isotropic sources in
infinite media, and, since the assumed artificial
point source was normalized so that the correct
number of gamma rays left the source in the direc-
tion toward the detector (although an incorrect
number left at large angles to the source-detector

7H, Goldstein and J. E. Wilkins, Jr., Calculations of
the Penetrations of Gamma Rays. Final Report, NYO-
3075 (June 30, 1954).

122

axis), it was further assumed that the Goldstein-
Wilkins results could be applied directly to the
artificial point source to determine the buildup
spectrum at the detector in the actual problems.

Plots of the total unnormalized gamma-ray
number flux distributions at the detector (that is,
collided, or buildup, plus uncollided) for capture
and inelastic scattering gamma rays, which were
obtained in the manner described above, are shown
in Figs. 4.1.3 and 4.1.4, respectively. In these
figures the histograms representing the uncollided
discrete-energy gamma rays are superimposed on
continuous distributions which are almost entirely
due to scattered, or buildup, gamma rays.

UNCLASSIFIED

-2 ORNL-LR-DWG 34795

10
3
S B R
},_
O
w
b s
=)
a
N \ -
¥ -
£ 9
w -
i‘_) 2
: |
= -
-
g e pa

-3
S 10
@
o}
~
[aN)
£
[#]
~
4 5 =
<
[r -
L= ot
=
=
<
U)
x 2
>
’—
a
S L
L _4 N
O 10
o
W
o
= e S
)
=4
O 5
w
~N
3
<
=
or
s
1 2
=z
2
E
RS 40—5
1 2 3 4q 5 6 T 8 9 10 M 12

£, GAMMA RAY ENERGY (Mev)

Fig. 4.1.3. Calculated Total Capture Gamma-Ray
Number Flux Spectrum at the Detector (Water in
Collimator).
UNCLASSIFIED

-26
(X140 ) ORNL-LR-DWG 31796

B |

S
y &
M —

GAMMA RAYS /cm?/sec/ Mev WHICH STRIKE A DETECTOR)
3
4

Tie (E'MUN-NORMALIZED NUMBER OF INELASTIC SCATTERING

1+ 2 3 4 5 6 7 8 9 10 M 42
57’,, GAMMA RAY ENERGY {Mev)

Fig, 4.1.4, Calculated Total Inelastic Scattering
Gamma-Ray Number Flux Spectrum at the Detector

(Water in Collimator),

Calculation of Pulse-Height Distributions
from the Detector

Finally, it was necessary to include the detector
response so as to obtain pulse-height distribution
curves whose shapes could then be compared
directly with the shapes of the curves obtained from
the experimental results, Experimental measure-
ments of response functions of a 3- by 3-in. sodium
iodide crystal are available for various mono-
energetic gamma-ray sources, These response
functions were replotted in histogram form, where
the averaging was performed over 1-Mev intervals,
and interpolations were then made to obtain esti-
mates of the response functions at 0,2-Mev intervals
over the entire energy range of interest. These

PERIOD ENDING SEPTEMBER 30, 1958

response functions were then applied to the calcu-
lated number flux distributions at the detector to
give calculated pulse-height distributions from the
detector for both capture and inelastic scattering
gamma rays. A comparison of the shapes of the
experimental and calculated pulse-height distri-
butions from both capture and inelastic scattering
gamma rays is shown in Fig. 4.1.5,

The shapes of the measured and calculated
curves are in excellent agreement, Since the
spectrum of capture gamma rays which was taken to
be incident on the outside of the water collimator
was known to be good because most of it was ob-
tained from the results of direct experimental
measurements, the excellent agreement between the
shapes of the calculated and measured pulse-height
distributions from capture gamma rays indicated
that the methods used to determine the effects of
water buildup and detector response could be con-
sidered to be reliable. As a corollary, the ex-
cellent agreement which was also obtained between
the shapes of the calculated and measured pulse-
height distributions from inelastic scattering gamma
rays gives confidence in calculations based on a
spectrum of inelastic scattering gamma rays ob-
tained in the manner outlined, This analysis has
been published in more detail in a topical report.®

MONTE CARLO CALCULATION OF THE
DEPOSITION OF GAMMA-RAY HEATING IN
STRATIFIED LEAD AND WATER SLABS

L. A. Bowman® D. K. Trubey

It was reported 10 previously that an Oracle Monte
Carlo calculation of the penetration of monoener-
getic, monodirectional gamma rays in a lead and
water shield had been made that included a total
of 512 problems. The problems resulted from all
combinations of eight different lead and water con-
figurations (see Fig. 4.1.6), four total slab thick-
nesses (1, 2, 4, and 6 mfp), four energies for the
incident gamma rays (1, 3, 6, and 10 Mev), and
four angles of incidence (0, 60, 70.5, and 75,5 deg).
The results obtained included the dose rate and
energy flux throughout the slab and at the rear of the

8F. L. Keller and O. S. Merrill, Analysis of the
Recent TSF Secondary Gamma Ray Experzmenz ORNL-
2586 (Avug. 25, 1958).

%0n assignment from Wright Air Development Center,

195, Auslender and A. T. Futterer, ANP Quar, Prog.
Rep. March 31, 1957, ORNL-2274, p 286.

123
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2—01—056—-22—596R{
20— | r ]
L9 ‘ ‘ L ' ’ ‘
L i : \ i Lo
¢ | N SN
100} 1+ — : P
[ W H - ! — — - i i
I ¢ . e _ _ i :
e O ;
ol — ‘ T s
S e e B S R
e S NS EEEE .
— - B’ J ( — J ’ RO ' : R
l ., i ‘ ‘ i
— | RS R —_— ] " 1 | 1
i | . -
.. 0% o } ! : ;
ZITe e, T 7 | =T DIFFERENCE CURVE:
. %, . | i -7 NO BORON PLEXIGLAS ON REACTOR
. o°B o & . afe | | COLLIMATOR MINUS BORON PLEXIGLAS
‘ | A L ON REACTOR COLLIMATOR |
e I e R I L ' o ——
I l 1 SR !
i L ] | H
R A~ o SR S SN TV SR | | — -
' . : ‘ ; ‘ i o
i e A?..A . ‘OO,Q,_QARi,,, ,
| ‘ (Y : ooo} i
i —_— e e S Lo -
: ..'. ; A &bo
02}~ v | o e ———
! Co 4
l | ! %, i A—f'\f ‘% !
_ | L e, e o 1
& ol s, [PcAlLcULATED PONTS  ©
5 Yt S S S ‘,/ (NORMALIZED) L T % -
f’ - — ; FAYNNES S S e - ‘ R
= L D , B A I A o
o i L ! : (o]
& i ST e ; ' o |
Noos i | oy 1 W T
= i ! ! P
c | H . + i e L - | [ .
3 INELASTIC-SCATTERING GAMMA RAYS |  ® o " ‘
© L -y N ; R R e oo 77777(..7777"71 PR - P e S
T ‘ |
i 3 ! o : '° |
oo2 + -+ -t ! : : : B e A S : I — - s
| C ‘ ’ 1
| . | o
I o
. ! _
.01 || o ‘ S
~ : S e R R
0.005 j:_ ; S — . - 4 . . . _1.__ .. . . R o
i - . - i
- 1
0.002 | 4 | et - S |
| | .
: [ ]
0.004 - S e . : . N e b
o RN R :
0.0005 e B ; : - e b S -
_ 4 i S __5 _— ? — e
! i |
0.0002 +—- e ‘ % ‘ - e et
‘ : ]
‘ | e
0.0004 —— i 1 i ‘ R [ :
0 100 200 300 400 500 600 700 8OO 900 4000 1100 1200

PULSE HEIGHT

Fig. 4.1.5, Comparison of Experimental and Calculated Pulse-Height Spectra of Capture ond Inelastic Scottering

Gamma Rays Resulting from Neutron Interactions in Air,

124

PERIOD ENDING SEPTEMBER 30, 1958

slab; dose-rate buildup factors; the heat deposited of the heating results with an empirical formula
x throughout the slab; and the energy and angular is described here.
distribution of the gamma rays reflected from and
transmitted back into the slab, A comparison of
. the dose-rate buildup factors for normal incidence
obtained from this calculation with the factors
obtained from an empirical formula was also given
previously,! 1 and a recently completed comparison

The heating results are given as the percentage
of the total energy incident upon the slab that is
absorbed in a specified region in the slab. Some
typical plots of the results are shown in Figs.
4,1.7 through 4.1.10, which compare the Monte

Carlo results averaged over a region of four in-

V1L, A. Bowman ond D. K. Trubey, ANP Quar. Prog. tervals with values obtained by using the following
Rep. Sept. 30, 1957, ORNL-2387, p 320. empirical formula:
pa(EO,Mafx) X, +x,
J (E ,0,x,Mat )= [sec 0 ————— exp = | ~——
X PRI X u,(E 5, Mat ) cos 0
] X, X, . X, ] X+ %, . ] X,
exp - + _ - exp —
“1\ecos § © 2\cos @ O cos @ %2\ cos 9" 0 cos 6
,ua(Eo,Mafl) X+ %, 2

exp —<4cos (1l —cos ) |1~ ’
,(E g Mat,) (xy +x)4+ 1 VE,
where

J (Ey,0,x,Mat ) = percentage of total energy incident upon the slab that is absorbed in the slab at
point x per mean free path,

x, = number of mean free paths of the first material,
x , = number of mean free paths of the second material,

energy of the incident gamma ray,

Il

0
¢ = angle between the direction of the incident gamma ray and the normal to the slab,
'“a(EO'me) energy absorption coefficient

,uz(E'o,Maf ) " total absorption coefficient

(COS

) = NDA point isotropic energy absorption buildup factor for the first material,!’
B, ( 5 > NDA point isotropic energy absorption buildup factor for the second material,
cos

Xy +x,
exp - ( 3 )— exponential attenuation to point heating is calculated, and
cos

,ua(Eo,Maf]) X, +x

# (g, Mat,) (x, +x2)4 +1] VE,

is the empirical short-circuiting correction,

. exp~<4cos (1l —cos @) |1~

125
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2-01-059-238A

p} = NORMAL THICKNESS IN MEAN FREE PATHS AT
INITIAL ENERGY

{ / LEAD
pd
2 / LEAD WATER
e S /
@ 7 LEAD
b3S A WATER
2 - Hen ;s
D
=
z WAT
o 4 LEAD - AfER L .
= HTh,0
a
=]
o
[V
Z 5 WATER
o
Q
(s3]
[=4
a /
w6 WATER LEAD
7 WATER / LEAD
8 WATER LEAD
r/4v'+fir/4t+— ’74—"’4‘ 7/g ——w=
7, TOTAL THICKNESS (mean free paths) ————m=—

Fig. 4.1.6. Lead and Water Slab Configurations Used

in Monte Carlo Calculations,

The first bracketed factor represents the fraction
of incident energy expected to be deposited per
mean free path if the scattered gamma rays are
neglected. The second bracketed term is the build-
up factor. Near the boundary (x , small), where the
spectrum is largely determined by the first material,
the buildup is given by the first term. This term
damps out as x, gets large, and the buildup factor
is characteristic of the second material, The
buildup factors used in the formula were the results
of the weli-known NDA moments method calcula-
tion. 12 The energy absorption buildup factors used
were for a point isotropic source, since these were
the only buildup factors presented in ref 12,

24, Goldstein and J. E. Wilkins, Jr., Calculations
of the Penetrations of Gamma Rays. Final Report,
Y0-3075 (June 30, 1954).

126

UNCLASSIFIED
2-04-059- 354

2.00 —
C 3412
| REFLECTEC: ©.759

‘ --i——— TRANSMITTES: G148
1.80 ~—— ABSORBED: 99.03

T — Eo= 3 Mev O FORMULA

5 1 | 9= 60deg ® MONTE CARLO

" ( I 11 P |

& 7 11 N

= — — — :

3 | R
.- [

= | L .._L_L__

B:J q ) ‘ ' 1 [ T

a 140 —t —

o j o |

w 1 | [ 1

2 - —r i

w 1 T,_.._,fi _

2 120 r— W | |

IS ‘ -

E:" \ l 7 ! 1

> \ |

w S —

Z 1.00 1

T .

s

= \

<] I S S

© \

~ 0.80

g

e R

Q — T

= \

2 \

I 080

O .

et

w

O

g \

2 0.40

a % -

© A

=

0.20 L
Y
iz 1
0 e B B O
0 { 2 3 4

Hor, NORMAL THICKNESS OF COMPOSITE SLA8 IN MEAN FREE PATHS
AT INITIAL ENERGY

Fig. 4.1.7. Gamma«Ray Energy Absorption in a Lead
Shield as @ Function of the Shield Thickness,

The third bracketed factor is the ‘‘short-circuit-
ing’’ factor. An attempt was made to separate the
effects of the various parameters in the exponent.
The factors which depend on the angle reach a
peak at 60 deg. It seemed reasonable that a peak
might occur at about 60 deg when the combination
of a decreasing path length and a decreasing cross
section, as well as the final energy of a scattered
gamma ray as the angle of scattering increases,
were taken into account. The effect of distance
from the initial boundary also shows a peak (near
1 mfp). There is little short circuiting at short
distances, since the heating is due largely to first
collisions. The short circuiting damps out at
large distances since the buildup factor adequately
accounts for the scattered gamma rays far from
UNCLASSIFIED
2-04-059- 359

t 6431
T
; REFLECTED: 0,019
| TRANSMITTED: 3.21
1.80 ABSORBED: 96.8
= L E;= 6 Mev O FORMULA
a | 8= O deg @ MONTE CARLOD
w
1 : -t J
“ 4160 LEAD % WATER
g U B T
3 e _
: BN :
@ 440 ‘
] ] 1
@ I
z : -
] + i [ [
< 420 1 - -
> | I E |
&) |
14
w —t
z
w . | N N— 1
I 100
T i
a [ I ]
: T N
Q
<I I —
° 080 \ |
= - T
z ! A
8 A
z \
I 060
- o
2 )
LcL) 1 I !
L oot
z b f
o 040
o |
2 N\, |
w .\ |
0.20 R,
"-.a} I |
1o~ [ | ||
S el » ol
0
o) 4 2 3 4

e, NORMAL THICKNESS OF COMPOSITE SLAB IN MEAN FREE PATHS
AT INITIAL ENERGY

Fig. 4.1.8. Gomma-Ray Energy Absotption in o Lead-
Water Shield as a Function of the Shield Thickness.

boundaries. The factor

,ua(Eo,Mot])

i, (Eg,Mat)

is usually taken to be that of the first material,
since, in general, the short-circuiting effect is due
to scattering near the initial boundary. This pro-
cedure seems adequate if the first material is 0.25
mean free path thick but probably is not adequate
if the thickness of the first material is less than
this. The variation with energy seems to break
down with low energy and as a result the formula
can be low by as much as 20% for the 1-Mev case.
It should be emphasized that the data obtained
with this empirical formula have only been com-
pared with data from the Monte Carlo calculation

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
2-01-059- 3684

.00 2454
r
REFLECTED: 1,64
TRANSMITTED: 0,00t

L 380 ABSORBED: 98.3
= Eg= {0 Mev O FORMULA
a 8=755deg @ MONTE CARLC
v} 0
o T 1
L 320 WATER
= i} -
o
= >
14
w
o
o 2.80
w
[14]
@
(o]
wn
m
<
> 2.40
o
x
w
3
> e
<I
& 200
<I
=
>
<
(L]
> 160

18
2] \
g o
¥ {
5 1.20 \
-
w
© —
g \
":;’ 0.80 \\
[0
w
=

\
040
’ \
o l \'.-'0—--.. -
0 1 2 3 4

Haor, NORMAL THICKNESS OF COMPOSITE SLAB IN MEAN FREE PATHS
AT INITIAL ENERGY

Fig. 4.1.9. Gomma-Ray Energy Absorption in a Water
Shield as a Function of the Shield Thickness.

described above, which had a very limited number
of parameter values, and therefore it is possible
that the fit would not be so good for other values of
the parameters, particularly those outside the energy
range examined. The worst fits were obtained for
low energies, especially with lead following water,
but even in these cases the error was less than

20%. In nearly all the cases examined, the error
was less than 5%,

This work has been described in more detail in a
topical report. '3

13, A. Bowman and D. K. Trubey, Deposition of
Gamma-Ray Heating in Stratified Lead and Water
Slabs, ORNL CF-58-7-99 (July 28, 1958).

127
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2-01-059 - 3686

2.00 T

3472
T
] REFLECTED: 2.48
TRANSMITTED: 0.298
1.80 ABSORBED: 97.2
= Eo= 3 Mev O FORMULA
g 8=60deq @ MONTE CARLO
w
: | A
L 160 — WATER — LEAD —
2 — ] 1 =
<1
s
x ; .
Lt |
T 140 ’:
2
x } :
I | ] |
[4p]
2 20 | 3 |
& [ i
€ \ | |
g y a i
w 1
> oo L
a4 1.00
:
z 1
<l
© \
< 0.80 \
5
= \
- \
-
E 0.60 < j
-
i |
= h.
g 0.40
5 N
<
s
fl- [ T
0.20 "
\cx\ i[
-9
\)\‘.‘\1 1 i
0 N."-qn [ ] [ Y
o) i 2 3 4

Hof+ NORMAL THICKNESS OF COMPOSITE SLAS IN MEAN FREE PATHS
AT INITIAL ENERGY

Figs 4.1,10, Gamma-Ray Energy Absorption in a
Water-Lead Shield as a Function of the Shield

Thickness.

A MONTE CARLO CALCULATION OF THE

GAMMA-RAY HEATING IN SEVERAL SHIELDING

CONFIGURATIONS ADJACENT TO THE
BULK SHIELDING REACTOR

L. A. Bowman'?
W. W. Dunn'?

R. H. Lessig'?
D. K. Trubey

Some measurements of the heating produced in
several specific shielding configurations at the

Bulk Shielding Facility were presented previously,!

140n loan from Wright Air Development Center.
]50n loan from U.S. Air Force.
wK. M. Henry, ANP Quar. Prog. Rep. March 31,

1958, ORNL-2517, p 113,

128

]

and more recent measurements are described in
Chap. 4.3 of this report. For a comparison with the
experimentally determined heating, a Monte Carlo
calculation of the primary gamma-ray heating for
slab geometry was performed on the Oracle, and
the results are presented here.

Three different shielding configurations were
considered. In the experimental arrangement, each
configuration consisted of slabs of shielding ma-
terial immersed completely in a 3/8-in.-l‘hick alumi-
num tank filled with insulating oil. The material
of each configuration was the same, with the ex-
ception of the gamma-ray shielding section. The
slab that made up the gamma-ray shield was a
different material in each configuration; in one it
was 3 in, of lead, in another it was 3 in. of iron,
and in the third it was 2 in. of Mallory 1000 (90 wt %
tungsten, 4 wt % copper, and 6 wt % nickel, den-
sity = 16.9 g/em3). In all configurations, two slabs
of lithium hydride were placed behind the gamma-
ray shield, one being 4 in. thick and the other
being 1 ft thick. Also, in all cases, a 4-in.-thick
slab of beryllium was placed in front of the gamma-
ray shielding material. The experimental arrange-
ment is shown in Fig. 4.1.11, and the details of
the metallic heating samples, containers, and
thermocouple locations are shown in Fig. 4.1.12,

The experimental conditions were approximated
in the calculational procedures by assuming that
all the materials had the thicknesses mentioned
above and that they were infinite in the other
dimensions, These assumptions were necessary
since only slabs of infinite extent can be considered
with the Oracle Monte Carlo code.!” The gamma-
ray energy and angular distribution of the BSF re-
actor used in this calculation were the same as
those reported previously,!®

Method of Calculation

The Oracle Monte Carlo code yields gamma-ray
heating in a region of specified thickness within a
shield as a function of initial energy and angle of
incidence of the gamma rays. Initial gamma-ray
energies, E, of 0.5, 1, 2, 4, 6, and 8 Mev and in-
cident angles, 6, of 0, 30, 50, 65, and 80 deg were

175, Auslender and A. T. Futterer, ANP Quar. Prog.
Rep. March 31, 1958, ORNL-2274, p 286.

]8L. A, Bowman et al., ANP Quar. Prog. Rep. March
31, 1958, ORNL-2517, p B5.
s
2-01-058~-0-391
™——O0IL TANK; 3g~in.-THICK
ALUMINUM WALLS
WATER SLABS 4 ft SQUARE
% |+ 5 ft SQUARE
31/2“'1.“‘-{ T g
=6
2|2
Jdl =l I
15| 3] O
ol g
| =
o=
b —INSTRUMENT
4T HOLESY, 2,83
AT ' WATER
x:\;.:\
—~HEATING
SAMPLES
BULK
SHIELDING
REACTOR = £ =
(LOADING NO. 33) || |3 2|5 -
Yo in,—= |-—
GULF 10-C
INSULATING OIL

GAMMA SHIELDING MATERIALS
3in. OF IRON
3in.OF LEAD
2 in. OF MALLORY 1C00

Fig. 4.1.11. Plan View of the BSF Shield Heating
Experiment.

considered in this calculation, with 9 being meas-
ured from the normal to the shield. Heating esti-
mates integrated over all energies and angles were
made in the shield only at the points where thermo-
couples were located, as shown in Fig. 4.1.11,
although the Oracle calculation yielded heating
and dose-rate information throughout the shield.
It was assumed that the thickness of the shield
preceding the region being considered was meas-
ured from the front of the slab to the front surface
of the air annulus. This resulted in a slight over-
estimate of the heating.

The Oracle results, given in terms of a function
] .(6,E), were plotted for each energy E at each
position; J (6,E) is the percentage of the incident
energy deposited in a region x. A typical plot of
J .(6,E = 4 Mev) for the front thermocouple location
in the lead gamma-ray shielding material, that is,
the thermocouple location near the beryllium-lead
interface, is given in Fig. 4,1,13. With the use of

PERIOD ENDING SEPTEMBER 30, 1958

Rt L UNCLASSIFIED
2-01—058-0- 392

0.037-in.-0D CERAMIC
SUPPORT TUBES, 0.009-in.
THICK WALLS —

1.25-in.
A DIAMETER
EXTERNAL VIEW SECTION A-A
FROM EDGE
2 in.

Lin —CERAMIC FEED-THROUGH INSULATOR

/ FOR THERMOCOUPLE LEADS (SEALED

T /N WITH "DEVCON B8")
AT
7 ® LOCATION OF THERMOCOUPLES
SAMPLES
- MATERIAL  THICKNESS
, L& x ‘cr IRON 0.2 in.
LA P et e LEAD 0.4 in,
Y G T f|w MALLORY 4000 0.1 in.
//c v Nz o2 BERYLLIUM 0.4 1in,
/‘;/I/, " // '5 5
| 1

-— TWO HALVES OF CONTAINER CEMENTED
SECTION B-B TOGETHER WITH "DEVCON B PLASTIC STEEL"

Fig. 4.1.12. Details of Metallic Heating Samples, Con-
tainers, and Thermocouple Locations for BSF Heating

Experiment.

the energy and angular distribution of the gamma
rays from the BSF reactor as previously estimated '8
and curves such as that of Fig. 4.1.13, the heating
in the desired layer within the shield was com-
puted for each energy and angle of incidence to the
shield, The heating in a region due to gamma rays
from an infinite plane source is:

(1) H_=K IQ Y] (6,E) ¢.(0,4,E) cos 0dQ
E

Il

277Kf E]x(G,E) ng(G,E) cos @ d(cos 6) ,
E

where
K = factor to convert from Mev/sec to watts,

qBE(G,E) = gamma-ray energy flux in group E leav-
ing the BSF reactor at angle ¢ averaged
over ¢ (MevicmZosec ™ 'osteradian='.
kw™ 1,

129

ANP PROJECT PROGRESS REPORT

b«
2-01-059-382

0.04 ’

(@]
(@]
L8]

T

i

i

|

I

o
2
|
i
i

Jy 8, £), FRACTICN OF INCIDENT ENERGY ABSORBED IN REGION X
(@]
O
r
i
{
1
!
|

o \

0 10 20 30 40 50 60 7O B8O 90
8, ANGLE OF INCIDENCE (deg)

Fig. 4.1.13, Calculated Heating in the Front Lead
Sample Produced by 4-Mev Gamma Rays as a Function

of the Angle of Incidence.

cos O = term to convert from flux to current,

d(l = differential solid angle = sin 0 d0 d¢ =
d(cos 0) do.

In order to correct for the finite area of the source,
a Hurwitz correction was applied to the heating
equation, Since the BSF reactor is not a disk,
a radius @ was assumed which would give the same
disk area as the area of the north face of the
reactor. This correction, G(E), can be expressed
as the ratio of exponential integrals if it is assumed
that the correction for the total flux is the same as
that for the uncollided flux:

2) G(E) = =
El(z-uz'tz')—El E.'uz'tz'
E(Lut,)
where

H(z,©) = attenuation function for the uncollided
flux from an infinite source plane a

130

¥+ - distance z from the point at which the
heating is being calculated,

H(z,a) = attenuation function for the uncollided
flux from a finite disk (of radius a) a
distance z from the point at which the

heating is being calculated,
t .= thickness of ith layer of shield,

p, = total cross section for ith layer of shield,

[¢3) e—y
E (=) = dy .
LA s

Incorporating this correction in the heating equation
gives

(3) H (watts-cm =2 w1

- Zan ¥ 7 (0,E) G(E) $(6,E)] cos 6 d(cos 6).
E

The integrand of this equation is plotted in Figs.
4.1.14 through 4.1.16 for four thermocouple locations
in the shield configuration made up of beryllium,
Mallory 1000, and lithium hydride. The effect of

the shield on the angular distribution is easily

seen. In the case of the Mallory 1000 sample the
heating results were normalized to a sample 1 ecm

thick.

2-01-059-383

N O I e R |4

Hy, HEATING (w-cm™ 2. kw™ 1)

1.0 09 08 o7 06 0.5 04 03 0.2 04 0]
cos &

Fig., 4.1.14, Calculated Heating in the Front Beryllium
Sample Produced by the BSF Gamma-Ray Energy Spectrum
as a Function of cos 0 (Beryllium—Mallory 1000—~Lithium
Hydride Configuration).

-4
(x107%) 2-01-059-384

5.0 \
4.0 \

\,fFRONT SAMPLE

AN

Hy  HEATING (w-cm™2-kw™ ")

~ SAMPLE \
1.0 \ AN

\\
] T —
0 I c— T~
{0 09 08 07 06 05 04 03 0.2 o] 0
cos 8

Fig. 4.1.15. Calculated Heating in the Front and Rear
Mallory 1000 Samples Produced by the BSF Gamma-Ray
Energy Spectrum as a Function of cos  (Beryllium-
Mallory 1000-Lithium Hydride Configuration),

(x10° 7 2-01-059-385%
3.6

}

H,, HEATING (w-cm 2 kw

0 I~
1.0 09 0.8 o7 06 0.5 0.4 0.3 0.2 [eX ] o]
cos &

Fig. 4.1.16. Calculated Heating in the Front Lithium
Hydride Sample Produced by the BSF Gamma-Ray Energy
Spectrum as a Function of cos € (Beryllium—Mallory
1000~Lithium Hydride Configuration),

PERIOD ENDING SEPTEMBER 30, 1958

Equation 3 was numerically integrated from plots
similar to those of Figs. 4.1.14 through 4.1.16 for
each thermocouple location in the three different
shield configurations, Since the sample thick-
nesses varied, the integration was normalized to
heating per unit volume by dividing by the thickness
of the sample. These values are given in histogram

form in Figs. 4.1.17 through 4.1.19,

Conclusions

The results for heating in the beryllium of the
three configurations are in good agreement, differing
only by statistical variation. |t is interesting to
note that the heating is nearly the same in the
gamma-ray shield materials, since the attenuation
of the radiation in the shielding in front of the
sample and the absorption in the sample tend to be
compensating effects,

Most of the assumptions made in the calculation
cause the result to be too high. These are:

1. All the flux at the reactor surface is outwardly
directed.

2. The heating rate is that of a sample located
at the beginning of the void.

3. The source strength and angular distribution
are the same over the entire reactor face.
The last assumption probably gives the worst effect.
It might be possible to apply some correction fac-
tor to approximately take this into account, but
this has not been done because of the complexities
involved and the additional uncertainties which
would be introduced. A crude estimate of the
average source strength (based on the thermal-
neutron flux distribution near the center of the
reactor) indicates that the results are possibly a
factor of 1.5 too high. This estimate is corroborated
by the heating results based on an analysis of the
thermocouple measurements which were a factor of
2 lower than these calculated results,'?

19 . G. Carver et al., Proceedings of the Fifth Semi-
annual ANP Shielding Information Meeting, Atlanta, Ga.,
May 14~15, 1958, C/25801, vol. |, paper 10.

131
ANP PROJECT PROGRESS REPORT

(“0—4) 2-01-059-386
1.80 a®

d=22.5cm

7. Fe

(FRONT SAMPLE)

7 ‘ ¢ = DISTANCE BETWEEN REACTOR
L e —— AND SAMPLE ' I
d=12.4 cm |
1

’T'“ 1.20 1——— - T ——

z i

i ‘

E 1.00 - - - e —— — e ]

2 ! :

z : ;

— : |

o ! !

g 0.80 | S O I E—

s d=17.5cm

T a’=25,1(im

™ 0.60 o . ST T
e

(REAR SAMPLE)

(FROQTBE‘»AMPLE% RS / oy
0.40 % / B
/ d=29.5¢cm |
0.20 . g=359cm--— .
// / / AOQ/TLQLMPLE’// (REA/RU:AMPé/)
] 0
1 2 3 4

SAMPLE THICKNESS {cm)
-,
Fig. 4.1.17. Calecuvlated Total Heating Produced by the BSF Gamma-Ray Energy Spectrum in All Samples in the

Beryllium=lron—Lithium Hydride Configuration.

] 6

—4
tx10™% 2-01-059-387
1.80 [
|
d=22.5cm
1.60 e . S e
MPLE) o = DISTANCE BETWEEN REACTOR
AND SAMPLE
1.40 —p - -]
g=12.4 cm
’/
_. 120 / —
|
_x
v /
E 1. / .
S 00 7 Be //
z /(FRONT SAMPLE)Y]
3
2 os8o0 ]
<[
w
I
™ 0.60
e
0.40 #=25.4cm T —
Pb ;
{REAR i
0.20 SAMPLE} — i | o
d=29.5cm d=359cm
LiH LiH
|
° FRONT SAMPLE —+—"—REAR SAMPLE 1
o) 1 2 3 4 5 6

SAMPLE THICKNESS (cm)

Fige 4,1,18, Calculated Total Heating Produced by the BSF Gamma-Ray Energy Spectrum in All Samples in the
Beryllium—Lead~Lithium Hydride Configuration,

132

-3

PERIOD ENDING SEPTEMBER 30, 1958

(x10 M)
2-01-059-3
1.80 T ° >
d=225cm
1.60 MALLORY 1000
(FRONT SAMPLE) 7 = DISTANCE BETWEEN REACTOR
AND SAMPLE
1.40

1.20 //

d=4%7.5¢cm

M, HEATING {w-em™3 kw1)
MMM

0.60

\

LALLM

0.40 /. Be // 7 Be

(FRONT SAMPLE ) 71/ (REAR SAMPLE)

o
_

0.20

_

MALLORY 1000 4 | - T
(REAR SAMPLE)

g=27.0cm gd=33.4cm

N

2

/////// LiH LiH
{FRONT SAMPLE )——— {REAR SAMPLE)
3 4 5 6

SAMPLE THICKNESS (cm)

Fig. 4.1.19. Calculated Total Heating Produced by the BSF Gammo-Ray Energy Specttum in All Samples in the

Beryllium—Mallory 1000—Lithium Hydride Configuration,

A MONTE CARLO CODE FOR THE CALCULATION

OF DEEP PENETRATIONS OF GAMMA RAYS
S. K. Penny

A Monte Carlo code is being developed to calcu-
late at a point detector the angular and energy dis-
tributions of gamma rays emitted from a monoener-
getic, point isotropic or point monodirectional
source embedded in an infinite homogeneous iso-
tropic medium of constant density. The word
“‘monodirectional’’ is used here in a loose sense;
the emission directions are actually in a half-cone
specified by a polar angle, where the source-
detector axis is the polar axis, and they are uni-
formly distributed in the azimuthal angle. The
code is now in the ‘‘debugging’’ stage on the
IBM-704 electronic data processing machine. It is
hoped that by using special techniques it will be
useful for penetrations as deep as 20 mean free
paths, In its final form the code can be used for
neutron penetration also, provided a differential
cross section for scattering is given,

The purpose of this code is twofold, First, it is
hoped that it will shed some light on the discrep-
ancy between calculated and experimental energy
distributions from a Co%® source in water, the
chief difficulty being the magnitude of the distri-
butions. The calculated values were obtained by
the moments method,2%~22 and the experimental
results were obtained by Peelle, Maienschein, and
Love.?4 Second, there is merit in experimenting

20L. V. Spencer and U, Fano, Phys. Rev. 81, 464
(1951); see also J. Research Natl. Bur. Standards 46,
446 (1951).

2.lL. V. Spencer and F. Stinson, Phys. Rev. 85,
662 (1952).

22, Fano, J. Research Natl. Bur. Standards 51,
95 (1953).

23y, Goldstein and J. E. Wilkins, Jr., Calculations
of the Penetrations of Gamma Rays. Final Report,
NDA-15C-41 or NY0-3075 (June 30, 1954).

24R, W. Peelle, F. C. Maienschein, and T. A, Love,
Energy and Angular Distribution of Gamma Radiation

from a Co%0 Source After Diffusion Through Many
Mean Free Paths of Water, ORNL-2196 (Aug. 12, 1957).

133
ANP PROJECT PROGRESS REPORT

with specid®¥onte Carlo techniques in order to see
when und where they can be applied.

The special techniques referred to above are
importance sampling on the first-collision distribu-
tion coupled with double systematic sampling,
statistical estimation, splitting and Russian
roulette, and a special form of output, Other
features in the code are an energy cutoff, a weight
cutoff, and a cutoff for distance from the detector.

Importance Sampling

The importance sampling is discussed here from
the viewpoint of a point isotropic source. The
technique for the monodirectional source is simpler
and straightforward.

In the code, an attempt is made to estimate

z= [dx fdy fdw ... fdt bx,yw,et) 2(x,y,w,0t)

and it is assumed that the probability distribution
function can be written as

b(xfylwl"‘lt) = f(xf)’) g(w,...,t) .
By definition

z(:x,y) = fdw... fdt (W, eee,t) 2(%,¥,00,00e,t)
22(:x,y) = fdw... fdt glw, vee,t) zz(x,y,w,...,t)

and f(x,y) is the probability distribution function for
the first collision, Further, x is the cosine of the
angle between the direction of emission and the
source-detector axis, and y is the number of mean
free paths to the first collision. If it is desired to
sample from f*(x,y) instead of from f(x,y) (impor-
tance sampling), then

— [(x,y)
dx Jdy f*(x,y) z(ix,y) ——— =% ;
fax fdy f*(x,y) z(:x,y) *y)

however, the variance is not usually the same when
importance sampling is employed as when the
normal sampling is used. The probability distri-
bution function for the first collision that gives the
minimum variance is of the form:

*(x,y) = f(e,y) V22ex, ) N,

where N is the normalizing constant. Kahn?® shows
this for importance sampling with one variable, and

25y, Kahn, Applications of Monte Carlo, AECU-3259
(1954) (rev ed. April 27, 1956).

134

his argument can easily be extended to two vari-
ables, The code will begin with the task of esti-

mating z2(:x,y)by performing an auxiliary Monte
Carlo calculation, |nstead of sampling from a
first-coliision distribution, x and y will be fixed

and z2(:x,y) will be estimated, This will be

done for a number of points {x,y). This will then
provide estimates for the functions f*(x,y), G(y:x),
and F(x), which will be placed in tabular form in
the machine. The functions G(y:x) and F(x) are
defined as

f_);o dn [*(x,n)

Glyx) =
S dn e

and

Flx) = f_: d«ff_: dn *(€m) .

Once the auxiliary Monte Carlo problem has been
completed and the importance distribution functions
have been calculated, the main Monte Carlo routine
is started.

Double Systematic Sampling

The importance distribution for the first collision
is sampled with the use of a double systematic

sampling technique, 2% as follows:
1+ ]/2
F(x1)= i=0,],2,o..N_]
i+ %
G()/l:xz) = ] = O, ]’ 2, see N - ] .

in these expressions, N = 2%= number of histories;
a is a positive integer 20. The point (x ,y )} is then
found by interpolation in the tables. The 's must
be randomized, since they should not be correlated
to i, This can be done by using the periodicity of
the pseudorandom numbers generated by the con-
gruence method. Each binary bit in a pseudorandom
number has a certain period, as illustrated below:

26 2524 23 22 21 20 20 erind
x x x 0 1 bit

eee X X X

26peyised by S. K. Penny and C, D, Zerby, ORNL.
If it is desired to obtain all the even integers, 0,
2, . (2%%1 = 2), logical computer operations are
used to obtain from a pseudorandom number the fol-
lowing number:

20 20~1 232271790
000...0 x x

period

we x x x 0 bit

Then, generating 2% random numbers in succession
will provide the 2% integers desired in random
order, and the random correspondence i €>7 can be
obtained. Of course, a separate random number
generator must be provided which has the same
basic random number but an initial random number
deep in the series. When a first-collision point is
chosen by the sampling technique, the initial
weight must be modified by the factor

flx oy V(2 0y) o

Statistical Estimation

The technique of statistical estimation will be
used at every collision; that is, the probability of
the radiation reaching the detector after the colli-
sion will be calculated and used to estimate the
flux at the detector,

Weighted Isotropic Scattering

It is planned to sample the direction after
collision from an isotropic distribution, rather
than the Klein-Nishina differential cross section,
and “‘weight’’ it with the Klein-Nishina distribu-
tion, This procedure was derived primarily for
simplicity and to save in calculational cost, [t
is certainly biased sampling, but for small sepa-
ration distances it is probably the best method of
sampling. For large separation distances the
variance of the estimated flux will increase, but
the increase may be compensated for intuitively
by the use of importance sampling and splitting.

Splitting and Russian Roulette

The techniques of splitting and Russian roulette
are to be used, with the splitting occurring at
the collision point rather than at a splitting
boundary. In other words, if the particle crosses
b boundaries with the boundaries successively
closer to the detector, it will be split into 28
particles at the final collision point. If a particle
crosses b boundaries with the boundaries suc-
cessively farther away from the detector, it then

PERIOD ENDING SEPTEMBER 30, 1958

has a probability of 279 of surviving., Hf it sur-
vives, its weight is increased by 2%, The split-
ting boundaries are concentric spheres with the
detector as their center. Again the reason for this
technique of splitting at the collision point is
chiefly simplicity. In all cases the statistical
estimation is performed before either the splitting
or the Russian roulette techniques are used.
Although it may be argued that splitting at the
collision point will introduce biasing, it remains
to be shown whether the increased cost associ-
ated with an unbiased calculation would be
warranted.

Output

The output will be of a special form; there will
be no histograms. The usual process, which
provides histograms, consists of estimating the
flux at the detector and then determining the
angular and energy distributions from the total
flux arriving in a sector. The new code makes
use of the expansion:

k+127+1
2 2

Pk(w) Pl(T)

4,0 = [ do [ a7 Py(@) Py 91D,0,7)

E-E
c
T=2——" -]
E, - E,
E = energy of radiation,

!

E = cutoff energy,

m
I

initial energy,

cosine of angle between source de-
tector axis and direction of radiation,

€
Il

D = separation distance,
P, P ;= Legendre polynomials,

é(D,w,T) = flux at D per unit w per unit 7.

At each collision statistical estimates are made

of the A, (D) term?7 rather than simply the

A . (D) term, which is easily seen to be the total

flux. Single scattering is treated separately from
multiple scattering. The code allows &,/ < 15 for

2755 sugyested by R, R, Coveyou, ORNL.

135
ANP PROJECT PROGRESS REPORT

single scattering and &,/ < 7 for multiple scatter-
ing. Of course, the coefficients for single scat-
tering and multiple scattering are completely
additive,

A choice may be made of any combination of the
following values in the output: number flux,
energy flux, energy deposition in the medium
(ergs/g), and energy deposition in carbon (ergs/g)
at the detector, carbon being the standard material
for dose-rate calculations or measurements, Also,
the variance of the estimated coefficients (4, ,)
will be automatically calculated.

Testing Procedure

The code will be tested against three standards.
First, results for no energy degradation and iso-
tropic scattering will be checked against the exact
solutions given by Case, deHoffman, and Placzek. 28
Second, results for a medium other than water will
be checked against the energy distributions of
energy flux fumished by the moments method,2%~23
Third, results for water as a medium will be checked
against the experimental work of Peelle, Maien-
schein, and Love?? and values obtained by the
moments method, 2923

A MONTE CARLO CALCULATION OF THE
NEUTRON PENETRATION OF FINITE
WATER SLABS

J. B. Hilgeman??
C. D. Zerby

Monte Carlo calculations? of the neutron dose-
rate distribution beyond finite water slabs were
performed for various parameter combinations, In
the idealized probiem a plane monodirectional,
monoenergetic beam of neutrons was assumed to be
incident on a water slab of infinite area and finite
thickness. The angle between the incident neutron
direction and the normal to the slab was denoted

F. L. Keller

as 6, and the number of neutrons which penetrated
the siab was assumed to be recorded by a spherical

28y M. Case, F. de Hoffman, and G, Placzek, In-
troduction to the Theory of Neutron Diffusion, Los
Alamos Scientific Laboratory, Los Alamos, 1953,

29On assignment from the U.S. Air Force.

30The Monte Carlo machine program which was used

in this study is a revision of the neutron air-scattering
program which has been described by C, D, Zerby, A
Monte Carlo Calculation of Air-Scattered Neutrons,
ORNL-2277 (April 23, 1957). This program employs

the method of statistical estimation.

136

detector. The space around the detector was
divided into a number of solid angle intervals with
the apex at the detector point, and the number of
neutrons which entered into each of these solid
angle intervals was recorded. The energy spec-
trum of the radiation which entered each of the
solid angle intervals was determined by dividing
the energy range from some cutoff energy, E_
the source energy, E . into an arbitrary number of
equal intervals and recording each contribution in

to

its appropriate energy interval, This energy spec-
trum was then used to determine the tissue dose-
rate contribution from each energy and solid angle
interval,

The scattering probability was assumed to be
isotropic in the center-of-mass system. This is
essentially exact for the scattering by hydrogen
(the major scatterer) and is fairly satisfactory for
the scattering by oxygen over most of the energy
range of interest in these calculations. The density
of water was taken to be 1 g/cm3, All the results
were normalized to one incident neutron per second
per square centimeter of slab surface.

A parameter study was carried out in which plane
monodirectional beams of neutrons with energies,
Eqy of 0.55, 1.2, 2, 4, 6, and 8 Mev were incident
on the water slabs at angles, Oy of 0, 30, 60, and
75 deg. The thicknesses of the slabs ranged from
I to 8 mfp (mean free paths). The number of case
histories used for a particular problem varied from
5,000 to 10,000, depending upon the slab thickness
and the angle of incidence. The solid angle inter-
vals at the detector were defined by 90-deg
azimuthal angle intervals and 15-deg polar angle
intervals with respect to a polar axis which was
normal to the slab surface., A cutoff energy, E_
of 0.1 Mev was used. This value was chosen to
approximate the low-energy cutoff of most of the
present dosimeters. Fifteen equal energy intervals
were used to determine the spectrum,

Figures 4.1.20 through 4.1.23 show a set of
representative curves which were generated from
data obtained from this study. In these figures
dose-rate buildup factors, B , are plotted as a func-
tion of the finite slab thicknesses for various
energy neutrons incident at angles, 60 of 0, 30,

60, and 75 deg. The dose-rate buildup factor may

be defined as the ratio of the total dose rate at the
detector to the dose rate which would result if every
collision were equivaient to an absorption, It should
be noted that the results are plotted as a function of
UNCLASSIFIED
2-04-059-259

1000
500
ol
'
200 o
100 1
/
% 50 /l
/
“ 20 /
© [ o2
o
10 ] //
/
{
/ /
; / /
/ ‘/
N/ e
/ / // I 8o=0 deg |
2 %/ P i
1

0 i 2 3 4 5 6 7 8
NORMAL THICKNESS IN MEAN FREE PATHS

Fig. 4.1,20, Dose-Rate Buildup Factors for 0.55-Mev
Neutrons Incident at Yarious Angles on Finite Water

Slabs.

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
2—01—059-260

1000

500

*— 8,= 75 deg

~~——

200

100

50

""'--....____

BUILD-UP FACTOR
\

20 /

o

Ob
//Q’/

R0

£,

] // L~

2 %l/
Q { 2 3 4 5 6 7 8
NORMAL THICKNESS iIN MEAN FREE PATHS

Fig. 4.1.21. Dose-Rate Buildup Factors for 2-Mev
Neutrons Incident at Yarious Angles on Finite Water

SIUbS.

137
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED UNCLASSIFIED
2—-01-052-261 2-01-059-262
1000 1000
.
500 Jo P ‘T R I 500
T /90= 75 deg [
o
9]
M
200 f=— S R 200 y
QO
100 R RS s +— —
__ o 100
. / _ f
50 — P
S S 50 /
el 5 o] o [
a i
o a
= N ? f
3 / 3
= =
@ o
o 20 , R — - 2o /
/ ® [
bQJQ/
e
10 l // 90= 60 dEQ QO//
/ 7 ]l //
—
/ /
/ / /
5 /
[ / ] 5 / /
/ // 8q= 30deg V %
// | / o 8y = Odeg
/_,-——" 8,=0deg / %/ —]
> V . /
v | LA =
| 4
4 |
0 1 2 3 4 5 6 7 8 f
NORMAL THICKNESS IN MEAN FREE PATHS 0 1 2 3 4 5 8 7 8

NORMAL THICKNESS IN MEAN FREE PATHS

Fig. 4.1.22, Dose-Rate Buildup Factors for 4-Mev
Neutrons Incident at Various Angles on Finite Water Fig. 4.1.23. Dose-Rate Buildup Factors for 8-Mev
Slabs. Neutrons Incident at Various Angles on Finite Water

Slub5|

138
the normal thickness of the slab in mean free paths.
When plotted in this manner it is obvious that the
buildup factor for a given energy and slab thickness
should increase with increasing values of 4. be-
cause of the ‘‘short-circuiting’’ effect, These re-
sults were compared with the results of a similar
calculation which was performed by QObenshain,
Eddy, and Kuehn,3! and the buildup factors from
the present calculation were found to be consider-
ably smaller, in general, than their values, The
cause of this apparent discrepancy is not yet known.
Figures 4.1.24 through 4,1.27 show the angular
distribution of the scattered neutron dose rate at

UNCLASSIFIED
2-01-059-255

2 N | \DETECTOR'
40-7 \ \ | | JO S |
- \ AN [ i
N\ 3-mfp-THICK SLAB
: \\ N, _
- |® 5
| N ol
HE N
.:‘:’ [ \ \ \
5|8
e 2 N
aly
et 4 mfp
(=]
-._,m 10—3 \
N )
w T
=y
a
& 5
Q
(]
=
Q
o
-
@
=4 2 a
A
(92
S 9 .
10” ~
B
\\
5 T~
—
2 S
i
:
10710 E
-4 )

Cos a

Fig. 4.1.24, Fast-Neutron Dose Rates at Rear of
Finite Water Slabs Resulting from Multiply Scattered
Neutrons (EO = 0,55 Mev),

PERIOD ENDING SEPTEMBER 30, 1958

the detector for neutron beams of various energies
normally incident (6, = 0) on water slabs of various
thicknesses, Since there is azimuthal symmetry at
the detector for these cases, the curves are plotted
as dose rate per steradian versus cos @, where a

is the polar angle at the detector. These plots were

31, Obenshain, A, Eddy, and H. Kuehn, Polyphemus.
A Monte Carlo Study of Neutron Penetrations Through
Finite Water Slabs, WAPD-TM-54 (1957).

UNCLASSIFIED
2—01—059-2564A

-6 \
10 N
.

5 \\

) \ \\{fp_THICK sLaB
/—T: . \ \
- | %10 \
5|
HE N\
s s
T Tg, \ AN
o g \
— 2
2
o
wl 468 \\ \imfp
W
b AN e
(o] 5 \ .
= N
o
z AN
1 \
w
a 2 N

w
15° %

- % DETECTOR

Cos a

Fig. 4.1.25. Fast-Neutron Dose Rates at Rear of
Finite Water Slabs Resulting from Multiply Scattered
Neutrons (E'0 = 2 Mev).

139
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED

6 2—01—-059-257
10 : : ‘
i !
5 o0 S I
Z \"'kDETECTOR ]

——
21
] S e N i —
3§
| 34g7 | o .
AR Z-mfp-THICK SLAB |
:J 5 A ek ——— “. P |

3
S 5
(1)
—
<
@
w
w
O
) 2
=
(@]
[ia
5
o 108
T
-
w
a
= 5

2 |-
107° ' :

Cos a

Fig. 4.1.26, Fast-Neutron Dose Rates at Rear of
Finite Water Slabs Resulting from Multiply Scattered
Neutrons (EO = 4 Mev),

generated by drawing smooth curves through the
histogram output of the machine calculations.

For cases of normal incidence (65 = 0), a very
large fraction of the total scattered dose rate at
the detector is contributed by neutrons which have
undergone only one scattering event in the slab,
The fraction of the total scattered dose rate which
was contributed by singly scattered neutrons for
each of these cases is given in Table 4.1.1. From
this table it is seen that single scattering calcu-
lations may be expected to yield fairly accurate
results for dose rates from neutrons which are
normally incident on thick water slabs. The higher
orders of scattering become more important, how-
ever, as the angle of incidence, 90, is increased.

A detailed description of ail of the results of this
parameter study will be given in a separate report.3?

32, B, Hilgeman, F. L. Keller, and C. D. Zerby,
Neutron Penetration of Finite Water Slabs, ORNL-2463
(unpublished).

140

UNCLASSIFIED
2-01-059-258

SLAB

0 —

SN DETECTOR __ |

\ | ‘
f-mfp-THICK SLAB
107° I ~ - _ 1

—1

rep-hr ' steredian

source neutron-cm—2.

FAST-NEUTRON DOSE RATE (

CcOs @

Fig. 4.1,27, Fast-Neutron Dose Rates at Rear of
Finite Water Slabs Resulting from Multiply Scattered
Neutrons (E0 = 8 Mev),

Table 4.1.1. Fraction of Total Scattered Dose Rate at
the Detector Contributed by Neutrons Which Have
Undergone Only One Scattering Event (90 = ()

Fraction of Dose Contributed by

E,(Mev) Singly Scattered Neutrons
1 mfp 3 mfp 4 mfp 6 mfp
0.55 0.827 0.864 0.902
2.0 0.727 0.86 0.846
4.0 0.791 0.832 0.877
8.0 0.707 0.829 0.762

PERIOD ENDING SEPTEMBER 30, 1958

4.2. LID TANK SHIELDING FACILITY

* W. Zobel
STUDY OF ADVANCED SHIEL DING The configurations for which the fast-neutron and
" MATERIALS (GE SERIES) gamma-ray measurements are reported here are 4-J,
L. Jung 4-K, and 4-L, in which the thickness of a stainless

steel layer adjacent to the source plate was in-

creased in steps from 1 to 3in. In Fig. 4.2.1 the
gamma-ray dose rates beyond these configurations
are compared with the gamma-ray dose rates in oil
only. For purposes of comparison, the dose rates
beyond configuration 4-A, which consisted of 4 in.

The experiment designed by GE-ANP for studying
the production of secondary gamma rays in con-
figurations containing advanced shielding materials
has been completed. Most of the results of the fast-
neutron and gamma-ray dose rate measurements
bey'ond the various conflgurc'mons lisgd in this ex- of stainless steel in oil, are also presented.
periment were reported previously. < The remain- Corresponding fast-neutron dose rates are shown in
ing data on thermal-neutron flux measurements and Fig. 4.2.2.
on fast-neutron and gamma-ray dose rate measure-
ments are presented here, The configurations for
which the data were obtained are described in Table . .

4,2,1. The dimensions and compositions of the ma- Dec?‘;;"‘ 5:9“5‘]7’,' ?grffi.%;fl;',”;'éflfvp Quar. Prog. Rep.
terials used were the same as those given in Table 2|, Jung, ANP Quar. Prog. Rep. March 31, 1958,
4,2.2 of ref 2. ORNL-2517, p 90.

Table 4.2 1. Description of Latest Configurations Tested in GE Series”

. Confi " zgr Location of Unavoidable Oil Gap
nfi
° N gut: ren Description of Configuration Probe” Within Configuration®
umbe
) e (cm) (em)
4-A 4 in. of stainless steel + oil 16,2 2.8
4-) 1 in. of stainless steel + oil 7.0 1.3
4-K 2 in. of stainfess steel + oil 10.0 1.8
4-L 3 in. of stainless steel + oil 12.8 2.0
4-N 4 in. of stainless steel + ]/2 in. of boral + oil 17.6 3.0
4-Q ]/2 in., of boral + 4 in. of stainless steel + ]/2 19.6 3.6
in, of boral + oil
4-T 4 in. of stainless steel + ]4 in. of boral 4 oil 17.6 3.7
5-A 4 in, of Be + 12 in. of LiH + ail 45.8 1.1
41-A 4 in. of stainless steel (dry) + oil in an Al tank 15.4 1.9
(Jrine-thick woll) (oir gap)
Oil Qil only in canfiguratian tenk 3.2d

%Includes some configurations described in ANP Quar. Prog. Rep, Dec. 31, 1957, ORNL-2440, p 263, ond in
ANP Quar. Prog. Rep. March 31, 1958, ORNL-2517, p 90.

Actual distance from source plate assembly to back of salid configuration.
“Includes 1-cm recession of the aluminum windaw in the configuration tank.

Pasition of inside surface of the aluminum window in the configuration tonk,

141

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2-01-057-71-435

104 T T LT T T4 T s
RN =R
SERSERE

: o Ee I B A |

s 7_,,_T50:~cc IPN C?AMBER_” _jr__ —_ ,—) ‘

]

GAMMA-RAY TISSUE DOSE RATE (ergs-g ' - hr !

S0 . ~ANTHRACENE ‘-
} SCINTILLATION ]
P ‘ DOSIMETER

2 ‘ -

! ‘
o
b CONFIGURATION STAINLESS |
| NUMBER STEEL (in)
5 - g ——4-J 1
a - 4-K 2
v 4-L 3

2 1 -0 4-A 4
X OIL ONLY

L S PO S
2c 40 60 80 100 120 140 160

75, DISTANCE FROM SOURCE PLATE {em)

Fig., 4,2.1. Gammao=-Roy Tissve Dose Rates in Oil and
Beyond Configurations 4-A, 4-J, 4-K, and 4-L: Effect
of Increasing the Thickness of a Stainless Steel Configu-

ration,

The thermal-neutron fluxes beyond these same
configurations are presented in Fig. 4.2,3. Intro-
ducing 1 in, of stainless steel to the oil medium re-
duced the flux by 11.5%. Stepwise additions of 2,
3, and 4 in, reduced the flux by 24.6%, 32.2%, and
42.1%, respectively, at a distance of 100 ¢m from
the source, compared with the flux measured in oil
alone,

The curves in Fig. 4.2.4 indicate that placing
]4 or I/2 in. of boral behind the 4 in, of stainless
steel had only a negligible effect on the flux 100 cm
from the source. It also appears that ]/2 in. of boral
in front of the stainless steel in addition to the

142

UNCLASSIFIED
2-01-0587 ~71-436

R - "

R I N
CONFIGURATION NO. STAINLESS STEEL (in}
— : 1

| —

a-
a-
4.
a4- .
oieonty o : ,,,Z[
H : |

t
)

W

FAST NEUTRON TISSUE DQOSE RATE (en;s-g_“hr_I

o] 20 30 40 50 60 70 80 20 100 Ho
Z,, DISTANCE FROM SOURCE PLATE (cm}

Fig. 4.2.2, Fast-Neuvtron Tissue Dose Rate in Oil
and Beyond Configurations 4-A, 4-J, 4-K, and 4-L:
Effect of Increosing the Thickness of o Stainless Steel

Configuration,

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
2-01-057-74-432

Sx 107 — ] T Il
2 4 (.‘;
100 3
LY N - -
W T
5 LAY

S LAY - —
- I\

eacd A\, ¥
5 \

6 \\\\\ CONFIGURATION NO. STAINLESS STEEL (in.)
40 e . - JE— 1 [— p—— - ——

" SecC

THERMAL NEUTRON FLUX (neutrons-cm™ 2
n

-2
10
0 10 20 30 40 50 60 70 80 20 100 10 120 130 140 150
2y, DISTANCE FROM SOURCE PLATE (cm)

Fig. 4.2.3. Thermal-Neutron Fluxes in Oil and Beyond Configurations 4-A, 4-J, 4-K, and 4-L: Effect of Increasing

the Thickness of a Stainless Steel Configuration.

143
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2--057-71-433

.
|
i
e

— e e e ]
CONFIGURATION — BORAL  STAINLESS  BORAL )
NUMBER (in) STEEL (in) (in)

0 4 0 +0IL
0 4 14 +0IL
0 4 o +0IL
VE 4 1o +0IL —
0 4 (DRY) +OIL IN Al TANK |

THERMAL NEUTRON FLUX (neutrons-cm 2-sec”w™!)

0 10 20 30 40 50 60 70 80 20 100 110 i20 130 14Q 150
Zey DISTANCE FROM SOURCE PLATE {cm!}

Fig. 4.2,4, Thermal-Neutron Fluxes Beyond Configurations 4-A, 4-N, 4-Q, 4-T, and 41-A: Effect of Inserting

Boral Within Configurations Containing Stainless Steel.

144

2 in. of hgral behind the stainless steel did not
affect the flux at 100 cm appreciably. The boral
did, however, cause a decrease in the thermal-
neutron fluxes at distances greater than 120 c¢m
from the source, possibly as a result of the re-
duction in the number of photoneutrons arising from
high-energy gamma rays. High-energy gamma rays
originate in the stainless steel following the capture
of thermal neutrons, and the boral reduces the
number of captures by absorbing most of the neu-
trons which are thermalized in oil and then re-
flected back into the stainless steel, 374

Thermal-neutron flux measurements beyond con-
figuration 41-A are also plotted in Fig. 4.2.4 for
comparison with the curve for configuration 4-A.

At 100 ¢m from the source the flux beyond con-
figuration 41-A was 16.5% higher than that beyond
configuration 4-A. The two configurations differed
only in that configuration 41-A was dry; that is, the

1 cm of oil between the aluminum window of the con-
taining tank and the first slab of stainless steel, as
well as the oil between the slabs, was removed.

As described previously,? several types of gamma-
ray shields were used in conjunction with 4 in. of
beryllium and 12 in. of lithium hydride. Measure-
ments beyond a configuration consisting of only
the beryllium and the lithium hydride (configuration
5-A) were made to assist in evaluating the data. In
order to complete the set of measurements, thermal-
neutron fluxes beyond this configuration are shown

in Fig. 4.2.5.

STUDY OF SECONDARY GAMMA-RAY
PRODUCTION IN LEAD

J. M. Miller

A comprehensive program for measurement of the
production of secondary gamma rays in lead was
initiated at the Lid Tank Shielding Facility (L. TSF)
for comparison with the calculation being made by
the Nuclear Development Corporation of America
(NDA). In the tests run thus far, thicknesses of
lead varying in l/z-in. steps from 1 to 6 in. and
additional thicknesses of 7]/2 and 9 in. have been

3p. K. Trubey, An Estimation of Photoneutrons from
Carbon-13 in an Oil Shield, ORNL-2200 (Aug. 13, 1958).

4. T. Chapman et al., Measurement of an Effective
Neutron Removal Cross Section of Lithium at the Lid
Tank Shielding Facility, ORNL CF-54-11-3, p 3 (Nov.
2, 1954).

PERIOD ENDING SEPTEMBER 30, 1958

utilized. In each case the lead was kept dry by
placing it in a steel tank (5/ in.~thick walls)
posmoned against the source plate. An aluminum
tank (Y-in.-thick walls) filled either with oil or with
borc:fec‘3 water was always placed immediately behind
the lead. The steel tank which holds the entire con-
figuration has a ¥-in.-thick aluminum window on

the source side, and the recession in the tank wall
between the window and the first slab of lead in-
troduces a l-cm-thick air gap at this poeint in each
configuration. Gamma-ray tissve dose-rate measure-
ments were made in the oil or borated water, and in
some cases fast-neutron tissue-dose-rate and
thermal-neutron flux measurements were made in

the oil.

The gamma-ray dose-rate measurements in oil and
borated water beyond the various thicknesses of
lead are plotted as a function of the distance from
the source plate in Fig. 4.2.6. The measurements
in borated water were made primarily to determine
the effect of suppressing the thermal-neutron flux
beyond the lead. The gamma-ray dose rates in the
borated water were a factor of 3 lower than the dose
rates in oil in the region close to the lead and a
factor of 13 lower approximately 120 cm beyond the
lead.

Cross plots of the data from Fig. 4.2.6 are shown
in Figs. 4.2.7 and 4.2.8. The plots in Fig. 4.2.7
present the gamma-ray dose rates at points 100 cm
beyond the configurations, corrected for the inverse
72 attenuation, and show that most of the primary
gamma rays are attenuated by the first 3 in. of lead.
The gamma rays observed beyond greater thick-
nesses of lead are practically all secondary gamma
rays. The plots in Fig. 4.2.8 present the gamma-
ray dose rates beyond the various configurations at
distances of 80 and 100 cm from the source. This
figure shows that a lead thickness of 3 in. gives
the maximum effectiveness in reducing the gamma-
ray dose rate at a fixed distance from the source
and that increasing the thickness beyond 3 in. does
not result in a further reduction.

The thermal- and fast-neutron measurements in oil
beyond the various configurations are plotted as
functions of the distance from the source in Figs.
4.2.9 and 4.2.10, respectively. This study will be
continued with configurations in which beryllium and
lithium hydride will be used in combination with
the lead.,

145
ANP PROJECT PROGRESS REPORT

a 2-01-057-0-74-437

n 2
T
','U CONFIGURATION 5-A: 4 in. OF Be +1{2 in. OF LiH + OIL
o2
10 N
N\
AN
. N\

d
|

M

THERMAL-NEUTRON FLUX (neutrons-cm™ 2

N

5 - \\

— — \‘k\ —
2 \
1 \\\

AN
. N\
N\
N
N

{0
40 50 60 70 80 90 100 {10 120 130 t40
z,, DISTANCE FROM SOURCE PLATE (cm)
Fig. 4.2.5. Thermal-Neutron Fluxes Beyond Configuration 5-A,
146

PERIOD ENDING SEPTEMBER 30, 1958

UNCL ASSIFIED
2-04-057-73-428

5 x 10°
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CONFIGURATION NO. COUNTER TYPE LEAD (in.
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10 20 30 40 50 60 70 80 90 100 1o 120 130 140 {50 160

Zy, DISTANCE FROM SOURCE PLATE (cm)

Fig. 4.2,6. Gamma-Ray Dose Rates in Oil and Borated Water Beyond Various Thicknesses of Lead as a Function

of the Distance from the Source Plate.

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED

2 2-01-057-73-427
{10
5
Iy
\ IN OlL BEHIND LEAD
I; ‘ \
,_:" \ U\Kfi 5
s [\
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2
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LEAD THICKNESS (in.}

Fig. 4.2.7. Gamma-Ray Dose Rates in Oil and Borated
Water 100 cm Beyond Various Thicknesses of Lead

{Measurements Corrected for Inverse r2 Attenuation),

THERMAL-NEUTRON FLUXES MEASURED IN OIL
AT THE LTSF: AN ERRATUM

L. Jung

Results of measurements of the thermal-neutron
fluxes in oil in the usual LTSF configuration tank
were presented in an earlier report.> During a
periodic check of measurements, a computational
error was discovered in the normalization of the
instruments to gold-foil measurements in the oil
which had caused the reported curve to be 16% too
low; Fig. 4.2.11 shows the corrected curve.
Similarly, all other thermal-neutron measurements

148

UNCLASSIFIED
2 2-01-057-73-429

//,.\fi T
|
|

_BORATED WATER
y 1 —]

BORATED WATER
WA
R S i

GAMMA-RAY DOSE RATE (ergs-g~'-hr™l-w™!)

© 80c¢m FROM SOURCE PLATE
2 A 100 cm FROM SOURCE PLATE

0 2 4 6 8 10 12
LEAD THICKNESS (in.}

Fig. 4.2.8. Gamma-Ray Dose Rotes in Oil and Borated
Water at Points 80 and 100 cm from the Source Plate
as a Function of the Thickness of Lead Adjacent to

the Source,

in oil which have been reported in two of the pre-
vious progress reports 145 should be increased by

16%.

RADIATION TRANSMISSION THROUGH BORAL
AND SIMILAR HETEROGENEOUS MATERIALS
CONSISTING OF RANDOMLY DISTRIBUTED
ABSORBING CHUNKS

W. R. Burrus®

One material commonly used as a thermal-neutron
suppressor is boral, a heterogeneous mixture of

5D. W. Cady and E., A, Warman, ANP Quar. Prog. Rep.
Sept. 30, 1957, ORNL-2387, p 297.

*Now at QOhio State University, Department of Physics,
Columbus, Ohio.

PERIOD ENDING SEPTEMBER 30, 1958

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Fig. 4.2,9. Thermal-Neutron Fluxes in Oil Beyond Various Thicknesses of Lead as a Function of Distance

from the Source Plate.

149

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2—-01-057-73-43

103 - T -
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Distance from the Source Plate.

150

Z, DISTANCE FROM SOURCE PLATE {cm)

Fast-Neutron Dose Rates in Oil, Beyond Various Thicknesses of Lead as A Function of the

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
2-01-057-0-71-434

5X107
2
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10 \-
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z\
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THERMAL NEUTRON FLUX (neutrons - ¢m 2 -sec—1-w™1)

2 ,l
© GOLD FOILS
0 f——id——1 v Y—in FISSION CHAMBER
A
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v

3—in. FISSION CHAMBER N
12%—in. BF; PROPORTIONAL COUNTER X

0 10 20 30 40 50 60 70 80 90 100 Ho 120 130 140 150 160
z, DISTANCE FROM SOURCE PLATE (cm}

Fig. 4.2,11, Thermal-Neutron Flux in Oil in the LTSF Configuration Tank.

151
ANP PROJECT PROGRESS REPORT

commercial-grade boron carbide and aluminum

sandwiched between aluminum sheets. The sand-
wich is usually rolled to a thickness of 7 or 1/8 in.

Since the mixture is nonuniform, considerable space
exists between the chunks of boron carbide attenuat-
ing material, and, consequently, radiation can pene-
trate the shield by passing unattenuated between
the chunks. Because of this “‘channeling effect,”’
which is a statistical effect caused by the increased
transmission along paths which pass through less
than the average material thickness, the masses of
boral and similar heterogeneous shields must be
from a few per cent to several hundred per cent
greater than the masses of homogeneous shields
which would give the same amount of attenuation.

R. R. Coveyou’ has suggested a mode! with which
to calculate the approximate transmission of radi-
ation through materials that consist of randomly dis-
tributed chunks. The material is considered to be
divided into layers which have a thickness charac-
teristic of the size of the chunks. The holes in the
layers are assumed to be located in a manner
statistically independent of holes in adjacent layers,
so that the over-all transmission is the product of
the transmissions of all the layers, As the chunks
are made more attenuative, the radiation passing
through the holes between the chunks becomes
more important,

A method which is based on the Coveyou model
and which has been extended to include a distri-
bution of various chunk sizes and shapes has been
developed for calculating the fransmission of radi-
ation through heterogeneous shields. The method
has been used to compute the transmission of neu-
trons through boral. For the calculation it was
assumed that the boral sandwich was rolled to a
thickness of ]/8 in. and that the thickness of the
B,C-Al mixture was 0.085 in. with 40 vol % boron
carbide. This resulted in an over-all volume
fraction of approximately 25% for the absorbing
chunks, which were assumed to be spherical in
shape. The chunks were first considered to be of
11 different sizes between 20 and 100 mesh; how-
ever, it was found that assuming only four sizes
gave approximately the same results, and only four
groups were used thereafter,

The transmission calculated by this method for
normally incident 2200-meter/sec (0.0253-ev) neu-
trons through ]/s—in.-'rhick boral was 0.076. This
is to be compared with a transmission of 0.0015

70 . . .
Private communication,

152

calculated for normally incident 2200-meter/sec
neutrons by the homogeneous approximation.
However, the homogeneous approximation is an
inappropriate first approximation for this type of
shield.

The transmission of nomally incident neutrons
through a ]/B-in.-thick boral shield as a function of
energy is shown in Fig. 4,2,12, along with the limit
as the chunks become opaque (low energies). The
average transmission over the neutron distribution
shown (Maxwell-Boltzmann distribution at room
temperature) is 0.096 for a constant efficiency de-
tector and 0.084 for o 1/v detector.

The transmission of isotropically incident neutrons
through %-in.-thick boral as a function of energy is
shown in Fig. 4.2,13, For this case the average
transmissions are 0.024 for a constant efficiency
flux detector, 0.021 for a 1/v flux detector, 0.041
for a constant efficiency current detector, and 0.034
for a 1/v current detector.

The calculated results can be compared with the
results of two experiments which have been per-
formed at ORNL to determine the transmission
through '/B-in. thicknesses of boral as measured by
1/v detectors. In the first experiment® the radi-
ation consisted of thermal neutrons escaping from a
thermal column on top of the ORNL Graphite Reactor
with an angular distribution of the (1 + /3 cos 6)
type,? which is more forwardly peaked than an iso-
tropic flux. Consequently, the experimental values
should be between the computed values for normal
incidence and those for isotropic incidence. The
transmission obtained for a Brooks and Perkins
boral sample was 0.070, while the transmission for
a Carbide sample was 0.094.
periment '? the radiation was a collimated beam of
normally incident neutrons from a beam hole at the
ORNL Graphite Reactor. The transmissions ob-
tained for two different Alcoa samples were 0.065
and 0.070, respectively.

This method will be described in detail in a
separate report, 1

In the second ex-

BR. O. Maak, B, E. Prince, and P. C, Rekemeyer,
Boral Radiation Attenuation Characteristics, MIT
Engineering Practice School, KT-251 {(Nov. 27, 1956).

9R. F. Christy, Lecture Series in Nuclear Physics,
MDDC-1175, p 115 (Dec. 1947).

]OG. deSaussure, private communication.

W. R. Burrus, Neutron Transmission Through Boral
and Similar Heterogeneous Materials Consisting of
Randomly Distributed Absorbing Chunks, ORNL-2528
{to be published).

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
ORNL-LR-DWG 25040

1.0
0.5
/’/
v
//
0.2 2
P e
L1 /
z 7 /
(.T’.n / /l
@
z 0 /
E: / /l
= -~ g
// A
e /
— HOMOGENEOUS
0.05 P / APPROXIMATION
\ LIMIT AS £ —= 0O
{OPAQUE CHUNKS)
002
474 =0.0253 ev
TYPICAL 20-100 MESH B,C DISTRIBUTION
007  o0d 0.2 0.5 1.0 2.0 50 10 20 40
E/kTy

Fig. 4.2.12. Neutron Transmission Through l/a-in.-Thick Boral as a Function of Energy; Normally Incident Flux,.

UNCLASSIFIED
ORNL-LR-DWG 25044

1.0
#75=0.0253 ev
0.5 | TYPICAL 20-100 MESH B,C DISTRIBUTION
"/
A
NORMAL INCIDENCE>/ P
-
02 —~ //
/ \\ ‘//' /
o
z » v /
3 THERMAL NEUTRON 7 -
0 SPECTRUM (PER UNIT LETHARGY) | il
Z O FOR 4T =0.0253ev  / — A\ —7
z - 3 LA~ P
P 4 L \ |~ e
[ L ”
— (SOTROPIC =Sy
0.05 N / CURRENT DETECTOR —-am, \ L+ ISOTROPIC
- p— / ¥ FLUX DETECTOR
[ ,/ \
/ /,/ //
/ 1] /
/r // /
0.02 7 /‘/ / \\
e
/ // ‘
faarmaeent"] L1 \
/
AT \
O'O-‘ /
007 0d 0.2 05 1.0 20 50 10 20 40
VAT,

Fig. 4.2.13. Neutron Transmission Through 1/8-in.-TI1i¢:k Boral as a Function of Energy; Isotropically incident Flux.

153

ANP PROJECT PROGRESS REPORT

4,3, BULK SHIELDING FACILITY
F. C. Maienschein

Neutron Physics Division

THE GE<«BSF STUDY OF HEATING IN SHIELDS
K. M. Henry

Most shielding research programs, as was pointed
out previously, ' are devoted to optimizing shields
with respect to radiation attenuvation and to minimiz-
ing the size and weight of the shield. This effort
tends to increase the power densities, energy
deposition rates, and temperatures of the shield.

If energy deposition rates could be determined, then
temperatures could be calculated and cooling
systems designed accordingly. Errors in the size
and location of the cooling channels which result
either in less than maximum power operation or in a
shield that is not optimized in weight or size thus
could be avoided.

Until the study of radiation heating in shields’
was initiated recently at the BSF in cooperation
with the General Electric Company, energy
deposition (heating) tests had been performed only
on a small scale with samples in geometries that
neglected scattering and interface conditions. For
example, the energy absorption in small samples of
lead, iren, and aluminum, with the calorimeter com-
pletely surrounded by water, were reported by
Binford.2 In the experiment now in progress an
attempt is being made to compare the energy
deposition rates found in calorimetric samples in
relatively large laminated shields with calculated
results. If the calculations can be verified ex-
perimentally, then shield designs can be optimized
with respect to heating as well as attenuation.

The Laboratory participation in the experiment
includes providing the facility for the tests and
personnel to perform the experiment and to tabulate
the raw data. In addition, the BSF is supplying
neufron foils and counting channels for gamma-ray
and fast-neutron dosimetry. The General Electric
Company is furnishing all the shield slabs, the
heating samples, and the containing tank with its
associated parts. They are also supplying the
U233 fission chamber and its electronic counter-
parts. In addition, GE personnel will be responsible

‘K. M. Henry, ANP Quar, Prog. Rep. March 31, 1958,
ORNL-2517, p 113,

2, T. Binford, Nucleonics 15(3), 93 (1957).

154

for all data analysis, point kemel calculations, data
correlation, and preparation of the final report. The
necessary reactor flux and power calibration data

for the BSR loading being used, as well as the
radiation attenuation in water at various distances
from the reactor, have been supplied to GE by ORNL.

The shield mockup being investigated consists
of slabs of beryllium and lithium hydride separated
by a gamma-ray shielding section which consists
of one of three materials: iron, lead, or Mallory
1000. Measurements of the radiation intensities
and temperatures at the slab interfaces in the Be-
Fe-LiH configuration were reported previously.’
Subsequently, the entire series of measurements for
all three of the configurations has been completed,
but the results, including those published pre-
viously, ! are of questionable value since it has
been discovered that all the heating samples leaked
and oil entered the air region surrounding the
sample. Consequently, further publication of
temperature measurements will be postponed until
the series of measurements is repeated with new
samples. A detailed description of the experiment
and the results of some measurements of radiation
intensities are given below.

The BSR loading being used in the experiment
consists of an array of 28 elements, including two
shim-safety rod elements and one regulating rod
element. The face adjacent to the shield is 15 in.
wide (five elements wide), and 24 in. high; the
thickness of the core is either 15 or 18 in., depend-
ing on whether the row of elements contains five
or six elements. More detailed information on the
structure was presented in a previous reporT,3
along with the power and flux distributions, Data
on the attenuation of the flux and dose rates from
the face of the core through various thicknesses of
water were also published in previous reports. 43

3E. B, Johnson, Power Calibration for BSR Loading 33,
ORNL CF-57-11-30 (Nov. 28, 1957).

4E. C. Maienschein et al., Attenuation by Water of
Radiations from a Swimming Pool Type Reactor, ORNL-
1891 (Sept. 7, 1955).

SAttenuation in Water of Radiation from the Bulk
Shielding Reactor: Measurements of the Gamma-Ray
Dose Rate, Fast-Neutron Dose Rate, and Thermal-
Neutron Flux, ORNL.2518 (July 8, 1958).

A plan view of the reactor, the shield, and the con-

falnmg tank was presented in the previous repor'r
and is repeated in this report as Fig. 4,1.11 in
Chap. 4.1. The shielding slabs are submerged in

oil in a containing tank which is positioned 4.75 in.

from the reactor face. (In Fig. 4.1.11 the tank is
shown to be only 3.5 in. from the reactor. It had
been hoped that it could be positioned this close
to the reactor, but for the tests reported here, as
well as those reported previously, it was not
possible. The recent installation of a new reactor
support will allow the closer positioning for future
runs.)

The containing tank, which was designed to
withstand a 5-ft differential water-cil head, was
fabricated from ¥-in.-thick aluminum plates rein-
forced with 0.25 by 3 by 3 in. channels. It is 4.5
ft wide, 8.5 ft long, and 29 ft deep. After it was
positioned in the pool, it was filled with concrete
blocks to a height of 5.5 ft. An aluminum platform,
which rests on the concrete and is bolted to the
tank walls, provides locating pins with which a
removable aluminum tray holding the shielding con-
figuration can be positioned.

Each shielding configuration consists of a 4-in.-
thick slab of beryllium, followed by a gamma-ray
shield and 16 in. of aluminum-encased lithium
hydride. The lithium hydride section consists of
two slcbs of 4 and 12 in., respectively, separated
by a /8 -in.-thick aluminum spacer which is fitted
to the smaller slab in an attempt to reduce the
thickness of the ** in the assembled shield.
The larger lithium hydride section was supplied by
the LTSF and is used in this experiment only to
provide an effectively infinite lithium hydride
medium behind the smaller section. The gamma-ray
shield consists of either 3 in. of lead, 3 in. of iron,
or 2 in, of Mallory 1000. A slab of oil extends be-
yond the lithium hydride to the tank wall. In ad-
dition there is a %-in.-thick layer of oil between
the tank wall and the beryllium. (The desired
thickness of this oil layer was ¥% in., as indicated
in Fig. 4.1.11 in Chap. 4.1). All shields are as-
sembled on the pool room floor and installed in the
tank as a unit. This simplifies the positioning of
the slabs and provides an opportunity for measuring
the thickness of cracks between the slabs. The
cracks do not tend to change size with handling, as
indicated by measurements made before and after
the tests,

The beryllium slab and each of the gamma-ray
shielding slabs have stepped holes along the

crack’’

PERIOD ENDING SEPTEMBER 30, 1958

horizontal shield axis in which plugs holding the
heating samples are inserted. These slabs alse
have l/z-in. deep grooves on one side that extend
from the horizontal holes upward to the tops of the
slabs. When matched, these grooves form instru-
ment wells for vertical flux and dose-rate traverses
at the interface of the beryllium and the gamma-ray
shield, The small lithium hydride slab has a
straight-through hole along its axis, and two instru-
ment wells, one near each side, extending from the
hole upward to the top of the slab. When vertical
traverses are not being made, the instrument wells
are filled with plugs of the same material as the
slab, the metallic plugs being bare half cylinders
and the lithium hydride plugs being full cylinders
encased in aluminum.

As has been described previously, ' the containers
that hold the beryllium, iren, and lead heating
samples are each 3 in. in diameter and 2 in. long
with an intemal void 2 in. in diameter and 1 in.
thick. The design of the sample cases, which were
fabricated from the same materials as the samples
they hold, is shown in Fig. 4.1.12 of Chap. 4.1. The
samples themselves are 1Y in. in diameter and of
various thicknesses (the lead is 0.1 in. thick, the
iron 0.2 in. thick, and the beryllium 0.4 in. thick);
they are suspended by three small insulating support
tubes in the center of the void in the container, The
case for the Mallory 1000 sample dlffers in thot it
is stepped (3 i |n. in diameter for 1Y% in., then 2/ in.
in diameter for / in.). The Mallory 1000 sample
thickness is 0.1 in. Because of fabrication diffi-
culties, the case for the lithium hydride sample was
made from 0.25-in.«thick aluminum with external
dimensions measuring 2.96 in. in diameter and 1Y,
in, in thickness, The lithium hydride sample is
1.25 in. in diameter and 0.5 in, thick. Calibrated
iron-constantan thermocouples are attached to the
center of each side of the samples and to each in-
side face of the cases. The leads are brought
through holes in the cases (sealed with Devcon B
plastic steel} out to the surface of the oil.

Neutron foil activation measurements were made
at the interfaces of the Be-Pb-LiH configuration.
Cobalt foils were positioned 12 in. apart along the
horizontal centerline and 8 in. apart along other
radii that extended from the center on 45-deg angles.
All the foils were placed in recesses in an aluminum
foil holder 4 ft by 4 ft by 0.09 in. thick. The results
for both bare and cadmium-covered foils are shown

in Figs. 4.3.1 and 4.3,2.

155

ANP PROJECT PROGRESS REPORT

ol
2-01-058-0-444
793 x 10?
“ 2.02 x 108 (Ca)
s 4
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177 x 108 (ca) 8.63 X 107
7.24 % 10°? 2.38 x 108 (ca)
w
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326 x 10'° Y
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9,53 x 10% (ca) ™ 3.28 % 10'°
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L LINE OI!-_" SHIELD 130 x 102 (cd)
10
8 542 X 10 10
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1( 153 X 10 (N 5.78 x 108 (cq)
AT s
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POWER: 49.3 kw OF SHIELD
(100 kw NOMINAL POWER)

Fig. 4'3.10
ments at the Be«Pb Interface (Facing Reactor).

Cobalt Foil Neutron Activation Measures
The

2- o NN,
1
\
|
|
. | 7
" + 713 X 107 (Cd)
S 1.34 x 10° y
' Ve
1.49 x 10° 792 X107 Cd)
7.38 X 407 (Cd) ‘ 158 X 10°
e g o
/ 4.94 X 10
/\\ . 2.67 x 10% {Ca)
/ s,
/ e ,2.54 X 10° (Ca) , 634X 102
/_,- ) 599 x 10° sz 108 (e
e
/ G,,;\ /ON PLUG .
2, NS e 1.61 X 108 [Cd)
327 %107 e / 4.08 X +0°
—— P a — - —— e e
) \ 1.88 X 10° (Cd) “
HORIZONTAL CENTER \‘\
LINE OF SHIELD -
~,
2.34 X 10% (Cd) 9.66 X 10°
9.75 x 10° 4.49 x 108 (Cd)
3.80 X 16° (Ca) )
1.32 x 10'0
|
2.23 x10° ; 140 x 108 (Cd)
144 x 108 (cd ' 279 X 102 =
ye
. 426 x 0%
1,50 X 109 (Cd)
‘ ee— VERTI CENTER LI
POWER: 49.3 kw - oF SH::E'T_"D ENTER LINE
{100 kw NOMINAL POWER) 1

Fig. 4.3.2, Cobalt Foil Neutron Activation Meosure-

top number refers
first run, and the
run. The two are

of the uncertainty

to the observed neutron flux in the
lower number to that in the second
not necessarily comparable because

in positioning between the two runs

The
in the

ments at the PbeLiH Interface (Facing Reactor).

top number refers to the observed neutron flux

first run, and the lower number to that in the second

run, The two are not necessarily comparable because

and the steep flux gradient at each interface.

Vertical thermal-neutron traverses were made in
each instrument well in the Be-Pb-LiH configuration
with a 3/4-in.-0D U235 fission chamber. Each

L . -
counter was attached to a %-in.-dia by 3-ft-long tube

which formed part of a four-tube preamplifier housing.

A conventional amplifier, scaler, and power supply
completed the electronic setup. Difficulty was en-
countered during these measurements as a result of
the photoneutron background produced by reactor
fission-product decay gamma rays. |n addition, the
reactor had to be operated at such a low power (to
prevent overloading of the counting channel) that

156

of the uncertainty

in positioning between the two runs

and the steep flux gradient at each interface.

these fission-product gamma rays made a linear
power response versus demand difficult to achieve.
The total error introduced in the results, which are
plotted in Fig. 4.3.3, is probably £10%.
Fast-neutron measurements were made in the
vertical instrument wells with a *‘scaled-up”’ 7gmin.-
OD version of the Hurst-type phantom fast-neutron
dosimeter. The detector was attached to its pre-
amplifier in a manner similar to that of the thermal-
neutron detector. Binary-type integrating scalers
were used as the summing device. Some pulse-
height distributions were measured simultaneously.

6 2-01-058-0-443
10 T T T I
— |
&
O
O
[¢]
0.
T_ 5 )
T 10
3 S HOLE
2 —(Be-Pb INTERFACE)—
g Ve v
[ = ..
-’ o]
» - HOLE 2 —--}—e
bt (Pb-LiH INTERFACE) ® o
4
S (o) .
£ 5
(.J L
c .
= [e]
E L
£
5
3
T 10 %
2 -
)
o
'_
o]
W
s
I fo o
= 102 Co HOLE 3
] —O {LiH~-LiH INTERFACE)
T O O
|- )
~ B @]
0
o a
L
10 I Q9
0 4 8 12 16 20 24 28
DISTANCE ABOVE SHIELD AXIS (in.)
Fige 4.3.3. Thermal-Neutron Fluxes at the Shield

Interfaces as a Function of Distance Above the Shield
Axis (GE-BSF Shield Heating Experiment).

The effects caused by the fission-product decay
gamma rays are believed to have influenced the fast-
neutron measurements 5%, These measurements
are plotted in Fig. 4.3.4.

The ionization chamber used for the gamma-ray
vertical traverses was a ‘'scaled-down’’ version of
the conventional graphite-walled CO-filled
ionization chamber. lts output was read on a
direct-reading micromicroammeter. The results are
presented in Fig. 4.3.5.

Additional information concerning this experiment
can be found in several reports published by GE.6=9

6). G. Carver et al., Proceedings of the Fifth Semi-
annual ANP Shielding Information Meeting, Atlanta, Ga.,
May 14~15, 1958, C/25801, vol. |, paper 10.

7). G. Carver, Description and Preanalysis of Proposed
BSF Nuclear Heating Measurements, XDC-58-1-51 (Dec.
23, 1957).

8R. 4. Maier, Report of Data from First Nuclear Heating
Measurements in Bulk Shielding Facility, XDC-58-8-61
{July 28, 1958).

YA, W, Casper, Point Kernel Calculations of Core

Gamma Heating in the BSF Nuclear Heating Experiment,
XDC-58-8-234 (1958).

PERIOD ENDING SEPTEMBER 30, 1958

2-04-058-0-444

10
° HOLE 1
o (Be-Pb INTERFACE)
5 |
£ i
T 10
; U
El
R=
E ,
=]
c W o
T HOLE. 2
= (Pb-LiH INTERFACE) $ i
(=8
Q
= ]
w - d T Q
g 1 e ¥ 2
® ' HOLE 3 H
@ [ (LiH-LiH INTERFACE) 1
- ?
=z
o
14
—
jun] J ]
3 ,
b 407
7] CONFIRMING POINTS =
E ‘
o
1073
0 4 8 12 16 20 24 28

DISTANCE ABOVE SHIELD AX!S (in.)

Fig. 4.3.4. Fast-Neutron Dose Rates at the Shield
Interfaces as o Function of Distance Above the Shield
Axis (GE<BSF Heating Experiment).

2-01-058-0-445

HOLE f

o ¥ (Be-Pb INTERFACE)
[ ]

[ ]

o
o
®

o
N
-

GAMMA-RAY DOSE RATE {r - hr— ' nominol watt ")

_FHOLE 2
O F~ (Pb-LiH INTERFACE )
g =
0.05 — . .
o ‘ ‘ A . B A
F 1 - r
0.02 ’ &
HOLE 3
0.01 | (LIH-LHH INTERIFACE)//‘ :l r
) 4 8 12 16 20 24 28
DISTANCE ABOVE SHIELD AXIS (in.)
Fig. 4,3.5. GammasRay Dose Rates at the Shield

Interfaces as a Function of Distance Abeve the Shield
Axis (GE«BSF Heating Experiment),

157
ANP PROJECT PROGRESS REPORT

4.4. TOWER SHIELDING FACILITY

C. E. Clifford

Neutron Physics Division

AIRCRAFT SHIELD TEST REACTOR
EXPERIMENT AT THE TOWER
SHIELDING FACILITY

Y. R. Cain S. C. Dominey1
F. Malone'

The experiments performed at the Tower Shield-
ing Facility in cooperation with Convair, Fort
Worth, 2 to obtain information complementary to
that obtained from the Nuclear Test Airplane (NTA)
program of experiments has been completed. During
the NTA experiments gamma-ray and fast-neutron
dose rates and thermal-neutron fluxes were meas-
ured both inside and outside the airframe contain-
ing the reactor while in flight and on the ground.
Measurements were also made on the ground in the
absence of the airplane structure. This last set of
measurements was duplicated in the TSF experi-
ments, and, in addition, measurements were made
as a function of altitude in the absence of the air-
plane structure. With these additional measure-
ments the influence of the air, the ground, and the
aircraft structure on the various measurements can
be determined. The TSF experiments also included
angular mappings of the radiation around the vari-
ous reactor shield configurations near the ground
and at various altitudes in order to obtain data
which will yield more accurate source terms than
were previously available. Some gamma-ray and
neutron energy spectra were also determined.

During the TSF experiments a large volume of
data was collected, most of which has not yet been
properly evaluated and correlated. For this reason,
and because power level corrections have not yet
been made, this report cannot be a complete ac-
count of the experiment. Rather it is intended only
to indicate the scope of the measurements and to
present samples of typical data, Detailed reports
will be published by Convair.

ASTR Shield Configurations

The Aircraft Shield Test Reactor (ASTR) consists
of an approximately cylindrical array of MTR-type
fuel elements mounted horizontally between two

]Convoir, Fort Worth,

2¢, E. Clifford et al., ANP Quar, Prog, Rep. March 31,
1958, ORNL-2517, p 117,

158

grid plates; demineralized water is used as the
moderator, reflector, and coolant. The core and
moderator section is surrounded by a layer of lead
and tanks to contain the water shield, as shown in
Fig. 4.4.1. The lead thicknesses, 3 in. on the side
and 6 in. on the front, were not altered during the
experiment, but various configurations of plain and
borated water were used in the two 7-in.-thick side
tanks. The primary configurations are shown in
Fig. 4.4.2. For most of the measurements a ]/a-in.-
thick boral plate was mounted on the outside of
the reactor shield, but for two configurations,
which were otherwise identical with configura-
tions 3 and 5, the boral plate was removed. These
two configurations are referred to as 3NB and 5NB.
In another configuration, identified as configura-
tion 5.5, only the lower halves of the two side
tanks were filled with plain water to effectively
give a ‘277"’ shield.

The support from which the ASTR was suspended
allowed the reactor to be rotated 180 deg. For all
ASTR configurations a 6-in.-thick lead shadow
shield was suspended in front of the reactor and
was rotated with it.

Detectors

Most of the dose-rate and flux measurements
taken during this experiment were obtained with
three types of instruments: (1) a Hurst-type single-
barrel fast-neutron dosimeter; (2) an anthracene
scintillation dosimeter for gamma-ray dose-rate
measurements; and (3) a BF,; proportional counter,
in which the boron is enriched to 96% in the B'°
isotope, for thermal-neutron flux measurements.
Both bare and cadmium-covered BF , counters were
used at each location. When the detectors were
not positioned in or on a shielded crew-shield
mockup, they were suspended from a special sup-
port truss. In addition to these instruments, foil
detectors were used for special measurements.

Radiation Mappings in the Reactor
Horizontal Midplane
Total Doses and Fluxes. — During the NTA ex-
periments, measurements of radiation were made
with unshielded detectors around the ASTR in the
absence of the airframe to obtain data which could

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
ORNL -LR-DWG 28942

GAMMA MONITOR

AFT GRID

PRESSURE VESSEL LEAD CASE
CONTROL RCD

MAIN SHIELD TANKS

FUEL ELEMENTS
FORWARD GRID

LEAD RINGS AND DISKS
FORWARD SHIELD TANKS
ION CHAMBER

CoPNE

Fig. 4.4.1. Cutaway View of the Aircraft Shield Test Reactor (ASTR).

be used to determine a source term. All these
measurements were made in the horizontal mid-
plane of the reactor at an altitude of 12,5 ft. Since
the ground would be expected to have considerable
effect on the measurements at this altitude, the
measurements were repeated and expanded in the
TSF experiments at an altitude of 186 ft (height of
horizontal axis of the reactor above the ground) and
were compared with similar measurements taken at
the TSF at an altitude of 12,5 ft,

A bottom view of the experimental arrangement
used for these radiation mappings is shown in
Fig. 4.4.3. The reactor was rotated around its

vertical axis, through an angle 8, measured clock-
wise from the reactor-detector vector to the re-
actor’s forward-drawn axis, and radiation measure-
ments were made with unshielded detectors at three
detector stations suspended from the support truss.
For these measurements the detector stations were
staggered to minimize direct-beam shielding of one
group of detectors by another.

Typical gamma-ray dose rates taken around the
ASTR at a constant reactor-to-detector separation
distance of 59 ft and an altitude of 186 ft with and
without a 14-in.-thick water side shield on the re-
actor are shown in Fig. 4.4.4 (ASTR configurations

159
091

,,,,,,,,,

Z9

7
‘.-;:‘é‘.-‘:-‘.

S

o //
= ; .
2

'--"--"'J ,'_:/:'.{ XKL

\\\\\\\\\\\\\\ T

/ ////// /

fflf:;.
\

x.\\"\‘\'.\\‘\"\:'\&.'\'.\'\'\_\j\‘\':'-’*

Nk

b, S f.ézi

\\\..\.\\ NN

\

/
7
87
N7
:':'y
Z
%
27

o
2-01-056~-23-792

"‘-4////////////7

]
47t

.

SIS, a0

{;:\‘%{//////////////fl--

CONFIGURATION 3%

*IN CONFIGURATION 3NB THE
BORAL PLATING ON THE OUT-
SIDE SURFACE OF THE SHIELD
WAS REMOVED

*%* N CONFIGURATION S5NB THE
BORAL PLATING ON THE OUT-
SIDE SURFACE OF THE SHIELD
WAS REMOVED

¥ IN CONFIGURATION 5.5 ONLY
THE LOWER HALVES OF THE
SIDE TANKS WERE FILLED
WITH WATER

Fig. 4.4.2. Shielding Configurations Used on the Aircraft Shield Test Reactor (ASTR).

e
SIS A

B ‘\\\\.\X..\.\.\_\\\.\\_\\.\.\;::;._-::_..
\\\\ ..

B

CONFIGURATION 15

CONFIGURATION 4

CONFIGURATION 5%* *

aééu

REACTOR CORE

PLAIN WATER

9l
\\\\\\\X\\\\\

LEAD

\\:--‘:-':-': 0N
~<////////////// /“

R (’

ANAAANARRNARNN
OO NN AN

BORATED WATER

SOMNONNNNNNNANN

CONFIGURATION 3B

AIR

Ld0dIY SS3A400¥d LDOIr0dd dNV
PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED

- 2-01-056-23-775

100 ft —

- 94 ft -

|(= 59 fi
3 H
!;. 53 ft

; 3 > ]

& [ ] |

® BF; COUNTER
+ FAST-NEUTRON DOSIMETER 4
4 ANTHRACENE SCINTILLATION DOSIMETER

+
+
.8

27 ft

:

33 ft

BOTTOM VIEW
wtill

Fig. 4.4.3. Experimental Arrangement for Total Radiation Measurements in Horizontal Midplane Around the ASTR.

2—01-056-23—-1-758 R{

2 x 101 : : [ :
_—ASTR SIDE TANKS EMPTY
; 10~ A~ KL (CONF 5) N
yd V4 Y
- y, NG -
- |; 5 / \-"\ /’ \
T / / ~ e AN N\
J.: "4 —” N
: / N\
c 10_2 y d N
E / ASTR SIDE TANKS FILLED N\
W . / WITH PLAIN WATER (CONF 3) \
= |
ae \
L
2 L/ \
5 / UNSHIELDED DETECTOR KEPT AT CONSTANT \
& 1073 |4 DISTANCE OF 59 ft FROM ASTR \
é 4 ALTITUDE = 186 ft ~
= 5 6= 0 deg ALONG REACTOR AXIS
(@]
2
1074

0 30 60 90 120 150 180 210 240 270 300 330 360
8, REACTOR ORIENTATION ANGLE (degq)

Fig. 4.4.4. Total Gammo-Ray Dose Rates in Horizontal Midplane Around the ASTR.

161

ANP PROJECT PROGRESS REPORT

3 and 5). (The reactor-to-detector separation dis- Fig. 4.4.3, the fast-neutron measurements were

tance is measured from the center of the reactor to also made at a point 59 ft from the reactor, but *
the center of detection of the counter.) The dose the thermal-neutron flux measurements were made

rates at @ = 90 and 270 deg increased approxi- at a point 53 ft away,

mately a factor of 4 when the water side shield The angular radiation mappings were supple- .
was removed. This increase is only slightly mented by foil exposures, most of which were made

greater than would be expected on the basis of on the reactor shield surface. However, foils were

the thickness of water removed; however, at 6 =0 also exposed at various positions in space, some

deg the dose rate increased a factor of 10. This as far away as 33 ft. The foils used were gold

indicates the importance of the side neutron shield-  (bare, gold-covered, and cadmium-covered), indium

ing in reducing the gamma rays at this point. Cor- {bare and cadmium-covered), copper, aluminum,

responding fast-neutron dose rates (Fig. 4.4.5) and sulfur, and magnesium. [n addition, long pieces

thermal-neutron fluxes (Fig. 4.4.6) consistently in- of copper wire were exposed near the reactor shield.

creased by a factor of 100 at all points when the The results of three such exposures are shown in

water side shield was removed. As shown in Fig. 4.4.7. The wires were positioned parallel to

FAST-NEUTRON DOSE RATE (mrem-hr-"'-w™)

162

2-01-056-23-776

#
ASTR SIDE TANKS EMPTY (CONF 5)——/ N
! t y 4 N
7 AN
J N
\ -
AN
/ N\

| UNSHIELDED DETECTOR KEPT AT CONSTANT
o'l .. DISTANCE OF 59 ft FROM ASTR

ALTITUDE =186 ft

8=0deg ALONG REACTOR AXIS

|

T

ASTR SIDE TANKS FILLED WITH PLAIN WATER (CONF 3)
L

)

2 i & /fl

yd NG 7 A
7 NG y 4 N\
/ N pd N
,/ \\\_ / AN
// \\\

L N
103 y

0 30 60 90 120 150 180 210 240 270 300 330 360

8, REACTOR ORIENTATION ANGLE (deg)

Fig. 4.4.5. Total Fast-Neutron Dose Rates in Horizontal Midpliane Around the ASTR.

PERIOD ENDING SEPTEMBER 30, 1958

oy
3 2-01-056—23-777
10
5
[~ ASTR SIDE TANKS EMPTY (CONF 5} .
2 -\
102 \
T; 5 UNSHIELDED DETECTOR KEPT AT CONSTANT
T DISTANCE OF 53 ft FROM ASTR
E ALTITUDE = 186 ft
» 8 = O deg ALONG REACTOR AXIS
[
a 2
£
e
o
o 10
= .
g ASTR SIDE TANKS FILLED WITH PLAIN WATER (CONF 3)
5 5 7 Py
=z
é / \Z""--. \\
3
1
1 R I
5 A,
2 -
16"
0 30 60 90 120 150 180 210 240 270 300 330 360

8, REACTOR ORIENTATION ANGLE (deg)

Fig. 4.4.6. Total Thermal-Neutron Fluxes in Horizontal Midplane Around the ASTR.

the reactor centerline, in the horizontal plane con-
taining the reactor axis, at distances of 0, 20, and
40 cm from the shield surface. The results plotted
in Fig. 4.4.7 are given as a function of the dis-
tance from the front end of the shield. For these
exposures the two ASTR side shield tanks were
filled with water {(configuration 3).

Direct-Beam Doses and Fluxes. — In addition to
the total dose-rate and flux measurements made
around the reactor, direct-beam neutron measure-
ments were made as a function of the angle 6. For
the fast-neutron measurements, the dosimeter was
placed in a cylindrical water collimator located
12.5 ft above the ground. In this position the
center of detection of the dosimeter was in the

same horizontal plane as the reactor axis and 33 ft
from the center of the reactor. The collimator hole
was 3'/4 in. in diameter and approximately 24 in.
long and was surrounded by an 18-in. minimum
thickness of water. Typical plots of the direct-
beam fast-neutron measurements taken while the
ASTR was rotated around its vertical axis are
shown in Fig. 4.4.8. These measurements were
token both with and without the water side shield
on the reactor (ASTR configurations 3 and 5).
Direct-beam thermal-neutron flux measurements
were made at a reactor-detector separation dis-
tance of 15 ft and an altitude of 186 ft. For these
measurements the BF3 counter was positioned in a
cylindrical, paraffin collimator which was lined

163
ANP P

5
10 - : e e _
«tn | : : \ r o
; » : ‘ | J
5 9 I
_
c : 4
g ; - 20 cm FROM
w2 ON SURFACE SURFACE |
5
Q
L
= 10?4
> i : : ;
i ~ N Ny
= : - .
8 5 -/- 1 \.\5\ H
'_ - . : .\\\ - —
2 40 ¢m FROM ,
z / SURFACE :
R — ) | ) . 1\,_, .
= /. | \ N\
= ‘ ; ! i AN,
I 3 i ' : i
= 10 . i ! ! \
L — = 4 1 . | | \
- A7 I |
L. y : ! 1
fi 5 - ',/" | i i
[¥E]
@ - ASTR SIDE TANKS FILLED WITH PLAIN |
WATER; BORAL PLATE REMOVED FROM ----
SHIELLD SURFACE (CONF. 3NB) |
! ’
, ]
{0 I ‘ |
0 12 24 36 48 60 72 84 96

Fig. 4.4.7. Relative Thermal-Neutron Fluxes Alang
Side of ASTR Shield: Copper Wire Activation Measure-

ments on Shield Surface and 20 and 40 c¢cm from Surface.

ROJECT PROGRESS REPORT

"
2-0t-056-23-778

with cadmium. This collimator hole was 2 in. in
diameter and was surrounded by a 6-in. minimum
thickness of paraffin; the distance from the out-
side end of the collimator hole to the center of
detection was approximately 12 in. The collimator
was suspended from the support truss, and a re-
motely positioned cadmium shutter allowed both
bare and cadmium-covered measurements to be

made. Examples of these measurements are given

in Fig. 4.4.9 for ASTR configurations 3 and 5.

Scattered Doses. — Measurements of the scattered
fast-neutron and gamma-ray dose rates were also
made as a function of 8 for various ASTR shield
configurations at distances of 33, 59, and 100 ft,
although none are reported in this paper. For these
measurements, which were made both at an altitude
of 12.5 ft and at an altitude of 186 ft, the detectors

DISTANCE FROM FRONT END OF ASTR SHIELD (in.)

were suspended from the support truss in a line
which, when 6 = 0 deg, was coincident with the
reactor axis. The arrangement is shown in Fig.
4.4.10. The details of the direct-beam shields are
shown in Fig. 4.4.11, These shields were remotely
positioned to vary the direct-beam shield angle a
from 8 deg to approximately 30 deg and thus to
eliminate various solid angles of radiation while

et
2-01-056-23-1-779

2
it \\ — — . t — —
-~ Y 7 4 N - _
- -~ b 1
5 P il A / N/ U —
/ \‘ ‘ l RS,
L AN L ol
- “—F— ASTR SIDE TANKS \ / \
s o EMPTY (CONF 5) — N = / fiomne Ao
. | - ¥ ) T . T
T A F" \ i - ‘\ I T
< I S I \ ] L R
; — \ / <\
E / \
~ o \ } N \ I
b i — 7 NN
'<_[ ~ Y 4 l[ J N B .S D
o i /< / 4 _ 7~7\7Wi T A\fi
" . p—
8 / / / % h o T
(] 2 - 4 ; r \ i ]
2 / K ‘ ;
S 41073 |- , .1 ™ASTR SIDE TANKS FILLED J \
= 7 WITH PLAIN WATER (CONF 3) h Y
o yd - A Y
w3 7 I I I l N
il . | l \ | AN
g 2 ~{ DETECTOR IN COLLIMATOR KEPT AT CONSTANT CISTANCE OF 33 ft FROM ASTR N\
104 / ALTITUDE = 12.5 ft \\
———1  § =0 deg ALONG REACTOR AXIS - —
. —
2 — —
1073
0 20 40 60 80O 100 120 140 160 180 200 220 240 260 280 300 320 340 360

164

8, REACTOR ORIENTATION ANGLE {(deq)

Fig. 4.4.8. Direct-Beam Fast-Neutron Dose Rates in the Horizontal Midplane Around the ASTR.

PERIOD ENDING SEPTEMBER 30, 1958

2—O|—056—2!-|-—757

T -
I -
I

|
yd S~ ASTR SIDE TANKS EMPTY ___/ N\
N (CONF 5)

\ / ,,a

]
T
..F

L Y

A

\ Y .
\ / \
\

\ / \

o/ A \/ \

THERMAL-NEUTRON FLUX (counts-min~t-w™1)

7 i A\ \
v 4 7 N \ L |
y i 7 N\ | _——ASTR SIDE TANKS FILLED \ \
5 LA V4 " WITH PLAIN WATER CONF 3} A \ A
// C 7 \ N\
, / \\ 1
1 1/ \v/\ /\/ \ —
¥ 4 \fv N
7 LY
y 4 AY
5 P N—
DETECTOR IN COLLIMATOR LI S
/ KEPT AT CONSTANT DISTANCE \
/ OF 15 ft FROM REACTOR
» ALTITUDE = 186 ft —
/ 6 = O deg ALONG REACTOR AXIS
| | | |
0 30 60 30 120 150 180 210 240 270 300 330 360

8, REACTOR ORIENTATION ANGLE {deg)

Fig. 4.4.9. Direct-Beam Thermal-Neutron Fluxes in the Horizontal Midplane Around the ASTR.

UNCLASSIFIED
2-01-056-23-783

DIRECT BEAM SHIELDS

ANTHRACENE SCINTILLATION DOSIMETER
FAST-NEUTRON DOSIMETER

BFy COUNTER

o+t H

A4 m9!>
{

Fig. 4.4.10. Experimental Arrangement for Scattered Radiation Measurements in Horizontal Midplane Around the
ASTR.

165
ANP PROJECT PROGRESS REPORT

the reactor-to-deteator separation distances re-
mained fixed. By taking many measurements in
this manner, it was possible to extrapolate the
plots of dose rate versus a to obtain the scattered
dose rates for a = 0. By subtracting the resulting
values from the total dose rates obtained at the
same separation distances, a second method for
determining direct-beam dose rates was available,

UNCLASSIFIED
2-01-056-23-784

+ FAST-NEUTRON DOSIMETER
A ANTHRACENE SCINTILLATION DOSIMETER

DISTANCE
VARIABLE
DISTANCE |
FIXED
a
&) Pb 8in.
f

‘\ BORATED RUBBER

Fig. 4.4.11. Direct-Beam Shield Used Between the
ASTR and Fast-Neutron and Gamma-Ray Detectors.

Measurements in Crew Shield Mockups
In the NTA experiments the ASTR and its shadow

shield were located approximately midway between
the nose and the tail of the airplane. A cylindrical
crew shield mockup was located 33 ft in front of
the reactor in a position that was midway between
the reactor and the crew compartment in the nose
of the airplane. The total distance between the
reactor and the crew compartment was 65 ft,

In the TSF experiments, measurements were made
both in and around the crew compartment and the
crew shield mockup, both of which were suspended
from the support truss, as shown in Fig. 4.4.12;
however, the typical data presented in this paper
for this set of measurements were all obtained in
the crew shield mockup. The walls of the crew
shield mockup were fabricated from borated rubber;
the thickness at the rear was 21 in., and the side
and front thicknesses were varied from 0 to 11 in.
(see Fig. 4.4.13). The 30-in.-dia by 70-in.-long
cavity in the mockup was lined with 2.5 in. of lead
on the rear and with 0.093 in. of lead on the side
and front.

In addition to the measurements in the Convair
shields, measurements were also made in an ORNL

UNCLASSIFIED
2-04-056-23-790

@® LOCATIONS FOR DETECTORS QUTSIDE
CREW SHIELDS

{(ONE FAST-NEUTRON DOSIMETER, ONE
ANTHRACENE SCINTILLATION DOSIMETER,
AND TWO BFy DETECTORS)

CONVAIR CREW COMPARTMENT i
OR ORNL-TSF COMPART-
MENTALIZED TANK

CONVAIR CREW
SHIELD MOCKUP

|l
bl i |
il o JiE
l ’lH“ »“:F ’1[‘ ‘ -
i7" N-REACTOR
‘ ROTATOR
LEAD J
SHADOW SHIELD\L .
™asTR

r 33 ft =‘

Fig. 4.4.12. Configuration of ASTR and Crew Shields.

166

comparLTIe‘Q!t’c:Iized tank that had been used for
earlier experiments at the TSF. The compartments
in this tank, shown in Fig. 4.4.14, were filled
either with plain or with borated water. The rear
shield was 37 in, thick, and the side shield was
varied from O to 20 in. When this tank was used

it replaced the Convair crew compartment.

2-01?6—23-789
INSIDE DIMENSIONS 30-in. DIA x 70-in.LONG
V/A BORATED RUBBER (PERMANENT)}, 21-in. THICK
BORATED RUBBER, VARIABLE THICKNESS FROM O TO il in.

B Lc20 (PERMANENT), 0.093-in. THICK ON SIDE AND FRONT, 2.5 in. ON REAR

DETECTOR STATIONS 251, 252

DETECTOR STATIONS 243A, 2438
DETECTOR STATIONS 24¢, 242

FAST-NEUTRON DOSIMETER
ANTHRACENE SCINTILLATION DOSIMETER
CADMIUM-COVERED BF, COUNTER
BARE BF ; COUNTER

oepp

Fig. 4.4.13. Convair Cylindrical Crew Shield Mockup.

PERIOD ENDING SEPTEMBER 30, 1958

Effect of_Reiecior Side Shield Thickness. — The
oo

measurements of radiation in and around the
Convair crew shield mockup 33 ft from the ASTR
were taken as a function of the thickness of the
borated rubber side shield on the crew shield (de-
tector) for most of the ASTR shield configurations
at altitudes of 12.5 and 186 ft. Typical gamma-ray
dose rates, fast-neutron dose rates, and thermal-
neutron fluxes at the 186-ft altitude are shown in
Figs. 4.4.15, 4.4.16, and 4.4.17, respectively, for
configurations 3 and 5.

The measurements in the ORNL compartmental-
ized tank 65 ft from the ASTR as a function of the
water side shield thickness were made only at the
186-ft altitude. Three reactor shield=detector
shield combinations were used: (1) Plain water
shields both on the reactor (configuration 3) and
on the detector tank; (2} a plain water shield on
the reactor and a borated water shield on the de-
tector tank; and (3) borated water shields both on
the reactor (configuration 3B) and on the detector
tank. Typical measurements in this tank are also

plotted in Figs. 4,4,15 through 4.4,17.

Effect of the Ground. - In order to determine the
effect of the ground on the measurements made in
and around the Convair crew compartment and

UNCLASSIFIED
2—-01-056—15—P-490

e - 131.60 in. OUTSIDE e
1
~—-—37.05 in. QUTSIDE —-—‘ o~ 73.68in OUTSIDE—- -~ -——— —=Ji= 20,49 in OUTSIDE =
i
I
12.0in. (TYP)
g 1] 8k .
t = = |2.|2f|"l. { |

=~ £ - ML 2.0 in T F

g-— < [ ot o = e -73.48 in: e o @ (f\(p).

t m l - | 8 " | }

py Lol e L R _

= ) 1.90 in. -

= = 10 413211 36.0in.DIA (TYP) —s le S D D D D

oo ‘ NI f

o~ = S— * ?

N 5 } J ’ { 17.74 in. DIA—

o 6 : 18,0 in. OD
7 2.0 in. (Typ)-A
1" 8 —={ 11.90in. =—
9
-Hz‘sean‘—-—t —= 24.06in. (TYP) |- - 7.96 in.—=]
- ) '

0.375in.

Fig. 4.4.14. ORNL-TSF Cylindrical Compartmentalized Tank. (@immgt with caption)

167

ANP PROJECT PROGRESS REPORT

—1 2-01-056~23-1-756

10 e e
R R
ALTITUDE =186 ft B —

" | CONVAIR MOCKUP: 33 ft FROM REACTOR |
ORNL TANK: 65 ft FROM REACTOR

—2—~ASTR SIDE TANKS EMPTY (CONF 5)
| L N
N N A S S —

10 ' A\ i +
— | =
- ’IN CONVAIR MOCKUP WITH — ]

7/‘ BORATED-RUBBER SIDE SHIELD
] |

_ASTR SIDE TANKS |
7 FILLED WITH PLAIN 3
|- WATER (CONF 3} ]

GAMMA-RAY DOSE RATE (mrem-hr™!-w

\Q
" IN ORNL TANK WITH | §§\
107 E=——— P AIN WATER SIDE SHIELD =~ +— S
"IN ORNL TANK WITH —— i =
BORATED-WATER SIDE SHIELD=""T T * ]
| ASTR SIDE TANKS FILLED | A
WITH BORATED-WATER (CONF 3B)
0 4 8 12 16 20

DETECTOR SHIELD SIDE THICKNESS (in.)

Fig. 4.4.15. Gamma-Ray Dose Rates in the Convair
Crew Shield Mockup and in the ORNL-TSF Compartmen-
talized Tank as a Function of the Detector Shield Side
Thickness: ASTR Experiment.

crew-shield mockup, many of the measurements
were repeated as a function of altitude. Typical
gamma-ray measurements made in air and in the
crew-shield mockup 33 ft from the reactor are
shown in Fig. 4.4.18 for two ASTR shield con-
figurations: configuration 5, in which there was
no side shielding on the reactor, and configuration
5.5, in which only the lower halves of the two
ASTR side tanks were filled, Corresponding fast-
neutron measurements are given in Figs. 4.4.19
and 4.4.20. Thermal-neutron flux measurements
made in air for ASTR configuration 5 and three
reactor-to-detector separation distances, 27, 53,
and 94 ft, are presented in Fig. 4,4.21.

e EAEY

168

T
2-01-056-23-1-755R4

=1
2 X140 1 T [ T T T T 1
40_‘ \\ 9=Ode9 —
e ALTITUDE =186 ft —
S CONVAIR MOCKUP 33 ft FROM REACTOR—
\\ ORNL TANK 65 ft FROM REACTOR ]
N N S N SR B
" \\ASTR SIDE TANKS EMPTY (CONE 5)
10 =
k. N i
T N
o =
£ 103 N { \
[}} N 1
£ S +;
= =N CONVAIR MOCKUP WITH
= \< BORATED RUBBER SIDE SHIELD
o A | | 1 i | l
w HEEEE
w404
9 X ASTR SIDE TANKS FILLED  ——
- N WITH PLAIN WATER (CONF. 3) —]
S AN
= N
D s N
Z 10
|_
[75]
& AN
™ IN ORNL TANK WITH N
~ BORATED WATER SIDE AN
156 | SHIELD: SAME CURVE WAS N
= OBTAINED WHEN ASTR SIDE TANKS = —
=~ WERE FILLED WITH PLAIN WATER N
~ (CONF. 3) OR WITH BORATED WATER {CONF. 3B)
g7
0 4 8 12 16 20 24

DETECTOR SHIELD SIDE THICKNESS {in.)

Fig. 4.4.16. Fast-Neutron Dose Rates in the Convair
Crew Shield Mockup and in the ORNL-TSF Compartmen-
talized Tank as o Function of the Detector Shield Side
Thickness: ASTR Experiment.

Gamma-Ray and Neutron Spectral Measurements

In addition to the measurements discussed above,
which essentially repeated those taken in the NTA
experiments at Convair, the TSF experiments in-
cluded a series of measurements at an altitude of
186 ft to determine the gamma-ray and neutron
energy spectra at various locations. Gamma-ray
spectral data were obtained in the Convair crew
shield mockup for the 11-in.-thick side shield and
in the ORNL compartmentalized tank for several
thicknesses of borated water on the side. For
these measurements two sodium iodide crystals
were used with an RCL 256-channel pulse-height

2-01-056-23-1-7

10 o
| I 1 | I 1 | 1 Al
8=0deg p—
ALTITUDE =186 ft —
CONVAIR MOCKUP 33 ft FROM REACTOR |
ORNL TANK 65 ff FROM REACTOR
2 ¥\
10
A\
LY
\
\
= \ASTR SIDE TANKS EMPTY (CONF. 5)
i 10 \
* N\
'c A
E
@ X
2 1 r\ A\
RE] 3 2N
LY 7 LY
§ /1' A Y
| \ /
- \ ( N
=
e \ |\ IN CONVAIR MOCKUP WITH
5 10 \——= BORATED RUBBER SIDE SHIELD -—
w N ¢ : ’ —
W \ ASTR SIDE TANKS
N X7 FILLED WITH PLAIN
g N\_| || WATER (CONF 3)
£, \\ \\ IN ORNL TANK WITH
T 10 \. PLAIN WATER SIDE =
. N S SHIELD -
™ ~
™ IN ORNL TANK WITH N
"~ BORATED WATER SIDE AN
10~ | SHIELD: SAME CURVE
= WAS OBTAINED WHEN =
= ASTR SIDE TANKS WERE —X
" FILLED WITH PLAIN WATER
_ (CONF. 3) OR WITH BORATED \\
_4 | WATER (CONF. 38)
40 4 I 1 | 1 1 1 N
0 4 8 12 16 20 24

DETECTOR SHIELD SIDE THICKNESS (in.)

Fig. 4.4.17. Thermal-Neutron Fluxes in the Convair
Crew Shield Mockup and in the ORNL-TSF Compartmen-
talized Tank as o Function of the Detector Shield Side
Thickness: ASTR Experiment.

analyzer. Typical pulse-height spectra for four
side shield thicknesses on the crew shield (de-
tector) and ASTR configuration 5 (reactor side
tanks empty) are shown in Fig. 4.4.22. A pulse-
height spectrum for a 20-in.-thick detector side
shield is shown in Fig. 4.4.23 for ASTR configura-
tion 3 (both reactor side tanks filled with plain
water), A spectrum for the same detector shield
thickness but with only one of the two ASTR side
shields filled with water (configuration 15) is pre-

sented in Fig. 4.4.24.

Neutron spectral measurements were made in a
6-ft-dia, 6-ft-long cylindrical water tank which was
suspended from the support truss 65 ft from the
reactor and had eight conically shaped collimator

PERIOD ENDING SEPTEMBER 30, 1958

- 2-04-056-23-1-780

10
5 oy
/‘\\
\ T
\\dEASUREMENTS IN AIR
2 e
—
£ \ —
®
' 162 < ¥
. Ty | ]
£ I MEASUREMENTS IN CONVAIR ]
s i MOCKUP WITH NO SIDE SHIELD.]
m > [ ASTRSIDE TANKS/ —~—
> — EMPTY (CONF. 5)
3 A
i
‘—-.N
S ~ )
> ~~ MEASUREMENTS IN CONVAIR
P MOCKUP WITH 11—in.~THICK
! BORATED RUBBER SIDE
<L _
s 163 >\\ SHIELD
= 7 Ry /
g / 7/ —
© [
5 |_“LOWER HALVES OF
| ASTR SIDE TANKS FILLED WITH
PLAIN WATER (CONF. 5.5)
2
REACTOR-DETECTOR SEPARATION DISTANGE = 33 fi
8= 0deg
0 25 50 75 100 125 150 {75

ALTITUDE (ft)

Fig. 4.4.18. Gamma-Ray Dose Rates in Air and in the

Convair Crew Shield Mockup as a Function of Altitude:
ASTR Experiment.

ot
2-01-056-23-4-7864

i

MEASUREMENTS IN AIR

—

ASTR SIDE TANKS EMPTY
(CONF. 5)
; I

MEASUREMENTS IN CONVAIR
/NOCKUF’ WITH NO SIDE SHIELD

\

<

FAST-NEUTRCN DOSE RATE (mrem-hr '-w

'~ REACTOR-DETECTOR SEPARATION
" DISTANCE =33 ft

T 8=0deg
I}

5X40 2 ' ! ‘

0 20 60 S0 120 150 180
ALTITUDE (ft)

Fig. 4.4.19. Fast-Neutron Dose Rates in the Convair

Crew Shield Mockup with No Side Shielding as a Function
of Altitude: ASTR Experiment.

169

ANP PROJECT PROGRESS REPORT

2 —
8=0deg

ASTR-CREW SHIELD
SEPARATION DISTANCE = 33 ¢

FYZRE
L]
-3 2-01-056-23—-1-754Ri
10 T I
. / ASTR SIDE TANKS EMPTY {CONF 5}
'z / \
T ]/ \
£
] 5
E
)
2
o
Ll "
8 --_-—'_'—-_~
[m]
=z LOWER HALVES OF ASTR SIDE TANKS
8 FILLED WITH PLAIN WATER {CONF 5.5)
5
V8]
T
—
w
&

0 30 60 90 120 150 180
ALTITUDE (f1)

Fig. 4.4.20. Fast-Neutron Dose Rate in the Convair
Crew Shield Mockup with 11ein.-Thick Borated Rubber
Side Shield as a Function of Altitude: ASTR Experiment.

holes radiating from near its center. The collima-
tor axes were spaced at 45-deg angles, each col-
limator half-angle covering 5 deg. Neutron-sensi-
tive plgtes were placed at the inner ends of the
collimator holes, which were sealed off from one
another. The results of all of these measurements
will be reported by Convair,

Miscellaneous Measurements

During the course of the experiment it became
apparent that several measurements were needed
which were not originally scheduled. These in-
cluded resonance and threshold detector measure-
ments as a function of distance from the reactor
and sulfur-foil measurements on the ASTR shield

170

2-01-056-23-1-782

10" | !
\ 6=0 deg ]
= N\ ASTR SIDE TANKS EMPTY
= N\ (CONF. 5) ,
g T !
ES
T > \
c
c AN
» 27 ft FROM ASTR
£ \ N
g \ T
>
=
|
w 2 ™,
=
S ~ Ns ft FROM ASTR
i—
-
,_éJ \ ]
J 402 N\94 ftFROMASTR | |
<J
=
E:J R
T T —
= —
q
5X10
0 30 60 90 120 150 180
ALTITUDE (f1)

Fig. 4.4.21, Total Thermal-Neutron Fluxes in Air at
Various Distances from the ASTR as a Function of Al-
titude,

surface. For the sulfur-foil exposures, ASTR con-
figuration 3 was used, and the foils were placed
around the reactor shield at 5-in. intervals in the
vertical center-plane of the reactor. The addi-
tional measurements also included thermal-neutron
flux measurements from 2 ft below the reactor water
pool surface to 10 ft above the surface, while the
reactor was held at altitude, and gold-foil measure-
ments of the thermal-neutron flux from 5 in. below
the surface of the ground to 5 in. above the surface,
also while the reactor was held at altitude. In
addition, thermal-neutron flux traverses were made
between the ground level and the reactor at an
altitude of 70 ft with a bare BF , counter and with

a BF; counter having a 277 cadmium shield on the
bottom or on the top. The data obtained in these
traverses are plotted in Fig. 4.4.25.

1
W

—1

-channel

-1

Counts-sec

ENERGY {Mev)

PERIOD ENDING SEPTEMBER 30, 1958

2-01-056~23-1- 785

B Y

“+—— ©
o

THICKNESS OF BORATED WATER

ON SIDE OF TSF TANK {(in.)
o 12
« 16
a 18
g 20

30in. OF BORATED WATER CN REAR

QF CREW SHIELD MOCKUP
ALTITUDE = 186 ft

a1
] éjs‘

F T s s )Y

i e

[ —

ah T
&.‘.a ‘@:Qfl - ._.o‘?°°°n°°0 ° .,
229552 o

60 80 100 120

CHANNEL NUMBER

Fig. 4.4.22. Pulse-Height Spectra of Gamma Rays Observed in the ORNL-TSF Compartmentalized Tank 65 ft from
the ASTR: ASTR Side Tanks Empty (Configuration 5).

171

ANP PROJECT PROGRESS REPORT

LY L]
2-01-056-23-1-786
CHANNEL NUMBER
-2 50 100 150 200
10 ] i I T
L] T T T
S
0
o}
o
B o
0 a
o}
2.25 Mev
[o)
® o |1
0
- % %
1
z
T °F %
©
: s R,
e}
L
__(.J 10‘4 &:Dfi
#U +r
3 o
g o
o
o (0]
B
A
b
o
ALTITUDE = 186 ft o
10_5 20 in. OF BORATED WATER ON SIDE OF TSF TANK; 30 in. ON REAR ool
oU |9
O O
0@90—03
o o Ralks lele)
QO J
o ‘o—“O“‘D“——
o
9 o
10'6 i I 1 ?O
e} 1.0 2.0 3.0 4.0 5.0 8.0 7.0 8.0 9.0
ENERGY (Mev}

10.0

Fig. 4.4.23. Pulse-Height Spectrum of Gamma Rays Observed in the ORNL-TSF Compartmentalized Tank 65 ft
from the ASTR: One ASTR Side Tank Filled with Plain Water (Configuration 15).

172

PERIOD ENDING SEPTEMBER 30, 1958

]
ENEAN 2=~ 0{—056—23—4-787
CHANNEL NUMBER
20 40 60 80 100 120 140 160 180 200
1 4 ] T i I i I ] I

ALTITUDE = 186 ft
20in. OF BORATED WATER ON SIDE OF
TSF TANK; 30in. ON REAR

gopP

o o
s  _3
5 10 o
T
Q
2 %g_xm
w 5 gy
£
g
[+)
2 %
[« 1)
08 %
L)
o
qu’
5
Yo
[
&0 0.
Yo
o Cg¥|o® 00%
2
0
(o]
%
o5 | | ! | | | 1 l | ! ®
0 1 2 3 4 5 6 7 8 9 10
ENERGY (Mev)

Fig. 4.4.24. Pulse-Height Spectrum of Gamma Rays Observed in the ORNL-TSF Compartmentalized Tank 65 ft
from the ASTR: Both ASTR Side Tanks Filled with Plain Water (Configuration 3).

173
o
2-04-056-23-1-788

2X140° [ ‘ |
ASTR ALTITUDE =70 ft

ASTR SIDE TANKS FILLED

10 |— WITH PLAIN WATER {CONF. 3) f

1 y 4 I

g

w
wn
N

///

7
BARE DETECTC% /

2 7 CADMIUM COVER ON
/ BOTTOM OF DETECTOR

|

NEUTRON COUNT RATE (counts-min~*-w ")

-_.._...._/
10 A
7
’/

_/

5 -~ CADMIUM SHIELD ON |
o~ ] TOP OF DETECTOR
2
CENTER LINE OF ASTR ——=
, |
0 10 20 30 40 50 60 70

DETECTOR ALTITUDE (f1)

Fig. 4.4.25. Neutron Count Rate Measured with a BF ,
Counter Below the ASTR as a Function of Detector Al-
titude.

PREDICTION OF RADIATION INTENSITIES AT
LARGE DISTANCES FROM THE TSR-ll

Y. R. Cain

The proposed construction of the Melton Hill Dam
on the Clinch River near the TSF area has made it
necessary to determine what the gamma-ray and
fast-neutron dose rates will be at large distances
from the Tower Shielding Reactor-1| (TSR-1}); there-
fore, calculations were performed to estimate the
dose rates for two operating conditions. For one
case the reactor was assumed to be ‘‘bare,’’ that
is, to have no external shield, It is to be noted,
however, that even the “‘bare’’ reactor will always
be surrounded by small thicknesses of water, alu-
minum, and lead (see Fig. 4.5.1 in Chap. 4.5 of
this report). This calculation was for an operating
power of 5 Mw, the maximum design power, at a
200-ft altitude, a condition which would result in
the highest dose rates that could be achieved with
the TSR-Il. For the other case the reactor was
assumed to be encased in a ‘‘beam’’ shield (see
Fig. 4.5.2 of section 4.5) which would allow a
highly collimated beam of radiation to be emitted

174

from the reactor. The results of this latter calcu-
lation, which was also for a 5-Mw operating power
at a 200-ft altitude, are expected to be more repre-
sentative of a typical operation. As part of this
investigation, a calculation was made of the un-
collided neutron flux 4200 ft from the TSR-| to
obtain a fast-neutron buildup factor for the case
in which an air-filled collimator extending through
the reactor shield is pointed at the detector.

Measured Dose Rates 4000 ft from TSR-I

Gamma-ray and fast-neutron dose rates at dis-
tances up to 4000 ft from the TSR-| have been
measured® for a reactor water shield thickness of
15 em. The data obtained, given in Figs. 4.4.26

3F. N. Watson, ANP Quar, Prog. Rep. March 10, 1956,
ORNL-2061, p 5; it should be noted that the quoted 15-
cm thickness is the distance from the center of the re-
actor face to the wall of the 12-ft-dia reactor tank,

UNCLASSIFIED
5 2-01-056-10—-2-267R!
t0
(R {_ I _+._......_:r____.{ B . I —}—
] ! T | | |
i | i T
—— ® 390C-cm3 ION CHAMBER, REACTOR VISIBLE — ]
" O PORTABLE ELECTROSCOPE; REACTOR VISIBLE
A PORTABLE ELECTROSCOPE; REACTOR NOT VISIBLE |

w

z

o~ S
|

&

TL

— = iflAfi

E 5 A AN

; A4

L.h N

Q

10! |
0] 800 1600 2400 3200 4000 4800 5600
r, DISTANCE FROM REACTOR TO DETECTOR (ft)}

Fig. 4.4.26. Measured Gamma-Ray Dose Rates as a
Function af the Distance from the TSRl with a 15-em-
Thick Water Shield.

and 4.4,27, were fitted with expressions of the
type

C e-—f/A

(n Prspa=—>2—,

where

D = gamma-ray or fast-neutron dose rate
(mrehr=tow=1 or mrep-hr"]-w"),
C = arbitrary constant,
r = reactor-detector separation distance (ft),
A =relaxation length
= 785 ft for gamma rays

= 630 ft for fast neutrons.

UNCLASSIFIED
2-01-056-10-3-268R{

10 I ) T )
| I T
! ] I
— | | |
& —— a2 HURST-TYPE DOSIMETER; REACTOR
% |L VISIBLE
| PORTABLE DOSIMETER (DIRECTIONAL); |
2 4 REACTOR VISIBLE _l,_¥
‘A PORTABLE DOSIMETER {DIRECTIONAL);
3 ’ REACTOR NOT VISIBLE
10 A - i _—
P N Y I
N i |
L N }
5 T f
7 N\ A=630 ft
N
T e
“ i
T A2 ;
= 10 T ] \\\ ]
. i
o N -
E S N
N\Q A\
S N
2 LN
\.\
I N\
I _ AN
AN
5 \\
2
q

0 800 1600 2400 3200 4000 4800
r, DISTANCE FROM REACTOR TO DETECTOR (ft)

Fig. 4.4.27. Measured Fast-Neutron Dose Rates as a
Function of the Distance from the TSR-l with a 15-cm-
Thick Water Shield.

PERIOD ENDING SEPTEMBER 30, 1958

Predicted Dose Rates 4200 ft+ from the TSR-ii

Gamma-Ray Dose Rates for Bare TSR-Il. — The
gamma-ray data presented in Fig. 4,4.26 were used
to predict the gamma-ray dose rates from the bare
TSR-{l simply by accounting for the differences
between the thicknesses of the shielding materials
on the two reactors along a straight line from the
reactor to the detector. While the TSR-| had a
15-cm-thick water shield, the TSR-Il shield will
consist of 2.5 cm of water, 3.5 cm of aluminum,
and 5.1 cm of lead. The modified expression is
therefore

(2 Doy(TsR-1t,bare) =

-1/ A
k4 ”(EH2O"H2O+2AI"AI+EP b*P b)

C - P
e e
1 r2
—r/A
(3) ¢,
-2 T2 f
where
D., = gamma-ray dose rate (mr-hr™1),

C,=172x 104 me-hr= 't 20w =1 (a constant
taken from ref 3),
P = reactor power level (w),
x = thickness of a shield material on the

TSR-I! minus the thickness of the same
material on the TSR-1{,

=2.5¢cm = 15 em = =12.5 cm,
xH2O cm cm cm

Xpp =35 cm,
xp = 5.1 cm,
2 = macroscopic absorption cross section for
gamma rays in the specified material,

-1
E,0 =0.0396 cm™",

2, =0.0953 cm=!,
26y, =0.467 cm™ .

With these values of x and 3 substituted in Eq. 2
and with the assumed power of 5 Mw, the value of
C, in Eq. 3 becomes 9.35 x 10° mrehe=1-ft2. The
complete expression for the gamma-ray dose rates
at various distances from the bare TSR-|l operating
at a power of 5 Mw.is thus

-r/785
9 e
() DoirsR-it,bare) = 935 % 10" —

r

mrehr=1

This expression is plotted as D,yr2 versus rin

Fig. 4.4,.28

175

ANP PROJECT PROGRESS REPORT

Fast-Neutron Dose Rates for Bare TSR-11 (Based UNCLASSIFIED
on ASTR Data). — Rather than attempt to predict 10° £ p1-0s6-2o- A TeR
fast-neutron dose rates for the bare TSR-Il from 5
the available TSR-| data, it was considered pre-
ferable to obtain measurements from a reactor \
shield that more nearly resembled that to be used R . \
on the bare TSR-1l. For this reason, measurements " 0 \\
were taken at a distance of 4200 ft from the Convair -,T \\ — —
Aircraft Shield Test Reactor (ASTR) (see preceding £ 3 \\ L
section of this chapter) with its thinnest shield 5 N ]
configuration (see configuration § in Fig. 4.4.2). £ 2 \(’FAST NEUTRONS
This shield consisted of 15 ¢m of water, 4.45 cm qu‘o -
of aluminum, and 7.3 ¢m of lead. It should be : e e e
pointed out, however, that the ASTR is approxi- 2 > \ AN
mately cylindrical, and the shield thickness given ; \\\ \\
is in the perpendicular plane bisecting the longi- 2 2 \ \ ’
tudinal axis of the reactor, X 10 . \
= N\ AN

The point at which the fast-neutron dose rate B N,

was measured was 4200 ft from the side of the " S \\\ \\\

ASTR along a line perpendicular to the longitudinal 2 N
axis of the reactor and originating at the reactor N N

center. The angle between this line to the detec- ;» 108 <

tor and the reactor axis is identified in the discus- B

sion below as the angle 6. With the reactor operat- >  GAMMA RAYS - \\\

ing at a power of 1 Mw, the measured dose rate was AN

0.78 mrem/hr (ref 4). As a check on this measure- e N

ment, a comparison was made using data taken at o N
various distances from the Nuclear Test Airplane 0 {000 2000 3000 4000 5000 6000
while it was in flight at altitudes greater than r, DISTANCE FROM REACTOR TO DETECTOR (ft)
10,000 ft (ref 5). An extrapolation of these data

to ground level resulted in a fast-neutron dose rate Fig, 4.4.28. Predicted Maximum Values of D, r2 and
of 0.77 mrem-hr"l, which seemed to confirm the Dn r2 as a Function of the Distance from the Bare TSR-II
measurement at the TSF. Operating at a Power of 5 Mw.

The TSF measurement 4200 ft from the ASTR was then used to calculate the dose rate from the bare
TSR-1l by normalizing Eq. 1 to the measured value, accounting for the differences in the shielding on the
two reactors, accounting for the differences in the neutron fluxes on the surfaces of the two reactors, and
assuming that the fast-neutron dose rates 4200 ft from the reactors would be proportional to their total fast-
neutron leakages. The expression for the dose rate is thus
- (EHzotzo"'zAlel"'EP b*P b)

G
22 p, ,
Gl

®
©

(3) D (TSR-11,bare) = Ca 3 3
1

where

D, = fast-neutron dose rate (mremehr=1),

C3 = a constant with units of mrem-hr=1ft2.w™1,

At the same point the gamma-ray dose rate was 0.55 mr/hr.

55. C. Dominey and J. B. Moore, Airborne Radiation Mapping Data, NARF-58-14T (March 14, 1958).

176

PERIOD ENDING SEPTEMBER 30, 1958

(,‘152/(;51 =ratio of the neutron flux at the surface of the TSR-II to that at the surface of the ASTR,

tzo =2,5em ~15em =~12.5 ¢m,
%, =3.5¢cm -4.45cm =-0.95 cm,

Xpp =91 ecm =73 ecm =-2.2 cm,
2 = macroscopic cross section for neutrons in the specified material,
Iy o =0.1538 em™!,
EAI = 0,079 em™T,
S5, =0117 em!,
G, = geometric factor for ASTR, which is the ratio of the average fast-neutron dose rate at a given

distance to that at 6 = 90 deg for the same distance

D(6)

" D(6 =90 deg)

] .[;" D(6) sin 0 6
 D(6 = 90 deg) -

f sin 0 d0
0

] n

- D(6) sin 6 dO,
2D(6 = 90 deg) -l; () sin

D(6) = dose rate at a constant distance from the reactor as a function of the angle 8 (taken from the
top curve in Fig. 4.4.5 of this chapter),
G, = geometry factor for the TSR-I, which, because of the spherical symmetry of the reactor, is
equal to unity,
The final value of Gz/G] = 1/0.599.
If it is assumed that the power distributions in the ASTR and the TSR-Il are flat, the flux is inversely
proportional to the volumes of the cores and the term v, /V2 can be substituted for (,'152/(;5], where vV and
V, represent the volumes of the ASTR and the TSR-Il, respectively. Equation 5 then becomes

—7/A
e n v, G, ‘(szoxH20+EAIxAI+ZPbeb)
(6) Dn(TSR-lI,bare) = C3 T2 _V_ G_ Pe
2 Uy
—r/)\n
(=4
(7) =0y

Assuming a reactor power of 5Mw and reducingEq.6 to Eq. 7 gives a value of 6,40 x 10'! mremhr= 112
for the constant C4+ The complete expression for the fast-neutron dose rate from the bare TSR-II operating

at a power of 5 Mw is thus
~7/630

1 -1
(8) Do (TSR-11,barey = 6-40 < 10 —z mrem-hr

177
ANP PROJECT PROGRESS REPORT

This is plotted as Dflr2 versus r in Fig. 4.4.28, When the value of ris 4200 ft, D_ is 46 mremshr~ ',

FastsNeutron and GammasRay Dose Rates for TSRall in a Beam Shield. — No experimental information
was available on which to base the calculation for the highly collimated beam; therefore, it became neces-
sary to perform an experiment with the TSR-1 in which a 15-in.-dia air-filled collimator was placed adja-
cent to the reactor in its 12-ft-dia water tank, Measurements of fast-neutron and gamma-ray dose rates and
thermal-neutron fluxes were made at a distance of 4200 ft as a function of the beam orientation angle, 6,
with the detectors positioned at § = 0 deg (6 deg west of due south). These measurements are plotted in
Fig. 4.4.29. The fast-neutron and gamma-ray measurements had a very high background, so the statistical
accuracy is poor. At the best point, 6 = 0 deg, the probable errors total 32% for the gamma-ray measure-
ments and 8% for the fast-neutron measurements. The statistical accuracy of the thermal-neutron data is
much better, however; the probable errors range from 7.4% at 6 = 130 deg to 2.4% at 0 = 0 deg. As a result,
the thermal-neutron data were used to obtain an assumed shape for the other measurements,

The measurements obtained for fast-neutron and gamma-ray dose rates at 0 = 0 deg were used to predict
dose rates from the TSR-ll with a similar collimator by accounting for the differences in the shielding ma-
terials on the two reactors along the beam axes. The shield thicknesses used for the TSR-1l were the same
as those given above for the bare reactor. The shield thicknesses on the TSR-l were reduced to 1.9 cm of
water and 3.0 cm of aluminum, The fast-neutron dose rate at 4200 ft from the TSR-Il operating at 5 Mw in a
beam shield thus becomes
o) 5 _p , e'(EHZOxH20+EAIxAI+EPbeb)

n{TSR-1l,beam) n(TSR-l,beam)
where, for this case,
xH2O =25em-19cm=0.6cm,
xpp=35¢em - 3.0ecm=0.5¢cm,
Xp) = 5.1 em,

and the values of X are the same as those given following Eq. 5. The resulting value of D (TSR-11, beam)
4200 ft from the reactor for a power of 5 Mw is 1,07 mrem+hr=1.

A similar procedure gave a value of 0.23 mrshr=! for the gamma-ray dose rate, Do (1sR-11, beam) The
total dose rate (neutrons and gamma rays) from the highly collimated beam shield was thus 1.3 mrem«hr—7,
82% of which was due to fast neutrons.

Figure 4.4.30 shows these predicted dose rates as a function of distance from the reactor. These dose
rates apply only for a point directly in line with the beam, and they can be reduced by about a factor of 6
by turning the beam 90 deg.

FasteNeutron Dose Rates for the Bare TSR«ll (Based on TSRe«l Data). — The fast-neutron collimator
dose rates presented in Fig. 4.4.29 were used to make an additional prediction of the fast-neutron dose
rates at large distances from the bare TSR-Il. The calculational procedure was as follows. If ¢, represents
the neutron flux on the surface of the TSR-| and it is assumed that the flux has a cosine current distribu-
tion, the net number of neutrons crossing 1 cm? of the TSR-| face per second per steradian in the direction

a{a being the angle from the normal) is /277) cos a neutronsscm™2.sec” ' .steradian~'. If it is further
g 1

178

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED UNCLASSIFIED
_ 2-04-056-25-1-752 04 - —26-5-794
9o 2X|O3 2-04-056-25

2x10?
AR o \X 10?
s AR RN W

. ,__,_/, A N\
N\ 5 X\ L> FAST NEUTRONS 5

/ N ‘.\‘
\ \
d \\—D—
<\THERMAL-NEUTRON FLUX 2 2
A\

1
~n
\
]
N

-4 i

GAMMA-RAY DOSE RATE (mr-hr w '
sec w

o ~6 P 3
'g 10 - - 10 L » 07
A 7 _ = \ P
; G - = \ £
3 s 0 E 3 \ > €
< [ f/f"\\ w \ E
ST T Y N B /AR 5 \ -
g 1 ! ® 2 2 W
R e 7N\ . \ s
£ 5 / 9 4o \ oF 3
£ = > A 3
- % = —T < Q
£ 2 /| \ £ i\ 3
£ 30’ y i e ! A\Y z
g & - +— e D W a 5 \L="CAMMA RAYS 5 &
w g a7 AN 2 @
2 = 7 X = A\N 5
= oo . | - . i \ )
: s hd o \\ E:
g /\ 2 2
— i
= 7 FAST-NEUTRON DOSE RATE \ \ =
W . / Lo | \ b
: . { AN\ o &
5 // | | A
b 2 2 15+in.~dic COLLIMATOR ADJACENT TO REACTOR —— \\ ~
IN TANK
8 = HORIZONTAL ANGLE BETWEEN COLLIMATOR 5 \\ \\ 5
AXIS AND REACTOR-DETECTOR AXIS N
-8 | | | i | \
1)
-150 -120 -30 =60 -30 0 30 6C \
8, BEAM ORIENTATION ANGLE {deg) 2 N \\ 2
\ \
10! Ay
Fig. 4.4.29. Measured Fast-Neutron Dose Rates, 0 1000 2000 3000 4000 5000
Gamma-Roy Dose Rates, and Thermal-Neutron Fluxes DISTANCE FROM REACTOR (ft)
4200 ft from the TSR-1 as o Function of the Beam Angle
& Beam Emitted Through Air-Filled Collimator Pene- Fig. 4.4.30. Predicted Gamma-Ray and Fast-Neutron
trating Reactar Shield. Dose Rates as a Function of the Distance from the TSR-II

Operating at a Power af 5 Mw in a Beam Shield.

assumed that there is no wall scattering, then the total number of neutrons leaving the collimator of length

L and radius a is

(10) ? cosal A AQ = (;;cos a) (7a?)

where
A = area of the collimator,
AQ) = solid angle subtended by the end of the collimator, as seen from the reactor,

cos a =approximately 1, since a is small for L >> a.

The dose rate at an angle 6 from the collimator due to 1 neutron per second emitted from the collimator is:

2712

" P

179

ANP PROJECT PROGRESS REPORT

where D(60) is the measured dose rate 4200 ft from the TSR-l as a function of the angle 6. Then the total

dose rate from a unit isotropic source (a source emitting 1/47 neutrons+sec™!.steradian=) is:

D(6) / 2mL? L2 -
(12) K == ) 27 sin 6 d6 = — D(6) sin 6 40, y
411\ (17a”) ma'p, J8=0

where 277 sin 6 40 is the differential solid angle. It ¢, represents the neutron flux on the surface of the

TSR-Il and it is assumed that it also has a cosine current distribution, the total number of neutrons emitted

per second from the TSR-I| would be

v

13) wr [ (22 2 si aniig, [ i z
( T Tcosa (27 sin a da) = 77”]9/)2 fa:() Cos a sin ada=277r]q.‘>2 '

a=0

where 7, is the outer radius of the core; the other terms have been defined previously.
{t is now possible to calculate the dose rate at 4200 ft from the bare TSR-II by multiplying Eq. 12 by

Eq. 13 and introducing an attenuation term:

-(ZH,0%H,0" A1 A" EP L P
e

L2 T
(14) D, 1l bare) = 2t {—— f D(6) sin 0 46
(TSR-11l,bare) 172 7Ta4q')] o

2.2 ~ 5

2L /Y, m ‘(\ZHzotzOJ"EAIxAIJrEPbeb) .

(15) S (s f D(6) sin 048 | e :
\V2 6=0

where the values of £ and x are the same as those given following Eq. 9, and, as in Eq. 6, it is assumed
that the flux is inversely proportional to the volumes of the cores. (In this case, however, V, is equal to
the volume of the TSR-l.) The resulting dose rate for a 5-Mw operating power is

= (szoxH20+zAlel+zP b*P b)
e

e b 2(48)2(14.75)2 (6,480)
n(TSR-1l,bare) — (75)4 (10,640)

f'" D(6) sin 0 d6
F==0

=193 [(1.79 x 10-7 mrem-hr‘l-w"])(S x 108 w)] ]—
2.07
= 83 mrem-hr ™!

This value compares reasonably well with the value of 46 mrem:hr=! calculated above.

As a check on the validity of this type of calculation, a similar procedure was used to predict the fast-
neutron dose rates 4200 ft from the ASTR which could be compared with the measured dose rate reported
above for the same reactor shield configuration. The calculated result was 2.78 mrem+hr=! and differed
from the measured value by a factor of 3.56, the latter value being 0.78 mrem+hr=1, As a further check, a
calculation of the fast-neutron dose rate 4200 ft from the TSR-I in its 12-ft-dia tank gave a predicted value .
which was a factor of 1.8 higher than the measurement. These two calculations indicate that the value of
83 mremshr~! given above is probably a factor of 2 or 3 too high. The discrepancies probably arise partly -

because the number of neutrons which leave the collimator is greater than Eq. 9 predicts and partly because

180

PERIOD ENDING SEPTEMBER 30, 1958

the beam results for fast neutrons may not vary with angle in exactly the same manner as the results for
thermal neutrons (which were used to give the curve shape for these calculations). In any event, it appears
that the dose rate at 4200 ft which was calculated by this method should be conservative in the sense that

the calculated dose rate should be larger than the actual dose rate.

Calculation of Uncollided Fast-Neutron Flux 4200 ft from the TSR-|

A calculation of the uncollided neutron flux at a distance of 4200 ft from the TSR-| was performed to
obtain a fast-neutron buildup factor for the case in which a beam is pointed at the detector. For this
calculation the measured neutron spectrum, N(E), at the face of the Bulk Shielding Reactor® was used
for neutron energies up to 2.7 Mev and a normalized Watt's spectrum was used for higher energies, where

N(E) is given in units of neutrons.cm™%.sec™'.w™='.Mev~.steradian='.

If a cosine distribution on the
face of the reactor is assumed, the uncollided fast-neutron dose rate at a distance ! is:

o ~2.(E)I
17 N(E) cos a C(E) AQ e

(17) D, U(TSR-LI) = f y dA_ dE ,
E=0 AS

det

where
A_ = area of the source (the reactor face),
dA_ = differential area on the reactor face,
A, ,, = area of detector,
ET(E) = macroscopic total cross section for air,
C(E) = flux-to-dose rate conversion factor,
AQ) = solid angle subtended by the detector at dA_,
! = distance from dA_ to the detector.
In order to base the calculation of D, u(tsr.y ©n @ measured fast-neutron dose rate at the surface

of the reactor, it was necessary to calculate the dose rate at the surface of the reactor as well as at

the distance /. For this case, the distance ¢, defined in Fig. 4.4.31, approaches zero and e—ET = L

(18) D Im | f N(E) <t>C(E) <—A"e*> 2mx dx | dE
TSR-1,surface) ~ 7
n(TSR-lsurface) E=0 A_ Ager \! 12

4] [ (¢ 0] tx dx

=2 I N(E) C(E) f -
E=0 | Yx= !

2 Im N(E) C(E) — fm | e

= 27T —
o | Jo (x2 + t2)3/2
e 4] —t 03]

= 27 N(E) C(E dE
o (E) C( (x2 + t2)]/2

=27 f N(E) C(E) 4dE ,
E=0

®R. G. Cochran et al., ANP Quar. Prog. Rep. Dec. 10, 1953, ORNL-1649, p 117.

181
ANP PROJECT PROGRESS REPORT

where x is the distance from the horizontal centerline of the reactor to the area dA.
When the detector is at large distances from the reactor, / >> x, cos ¢ is equal to 1, and the integral

over the source simply gives the area of the collimator:

I

Ao [ ~Z (E) ! ‘(2H2o"H20+2A|"A|)
(19) D uctsra,n = 5 N(E) C(E) e dE e
[ E=0

&Y
-(“'H2o"H20+EA1"A|)
where the exponential e accounts for the attenuation of the 1.9 cm of water and 3.0
cm of aluminum between the reactor and collimator.

The buildup factor, B, is:

2 fE zo N(E) C(E) dE

A ® “3. (E)
coll f N(E) C(E) e T dE
12 E=0

(20) B = Dmecs,l=4200 ft

meas,surface

by <
‘(“Hzo"H2oT~A|"A|)

e

(442 x 10-7)  (27)(3.65 x 10°) ]
(2,05 x 10%)  (6.83 x 10~8)(16.62) (0.59)

=7.38

<
—2..(E)I
Figure 4.4.32 is a plot of N(E) C(E) e T as a function of the neutron energy E for two separate

calculations. In the first calculation the averaged cross sections were taken every 0.5-Mev energy interval.
These cross sections are given in Table 4.4.1. For the second calculation the cross sections used were
taken from BNL-325 (ref 7) at 20-kev intervals below 2 Mev, at 100-kev intervals from 2.0 to 6.0 Mev, at
200-kev intervals from 6.0 to 8.0 Mev, and at 0.5-Mev intervals from 8.0 to 12.0 Mev. The resulting inte-

grations under these two curves differed by less than 10%.

’D. 4 Hughes and J. A, Harvey, Neuiron Cross Sections, BNL-325 (July 1, 1955).

UNCLASSIFIED
2-01-056-25-D-795

-
|

DETECTOR

- # -
|

Fige 4.4.31. Geometry Used for the Calculation of the
Uncollided Flux 4200 ft from the TSR-l,

182
UNCLASSIFIED
2-01-056-25-A-753R

2
o' _a—— USING AVERAGE TOTAL CROSS ____|
] SECTIONS AT O.f-Mev INTERVALS ——]
I
® \
1
_ i '\
w
5, % \ N
o r ‘
Q
s 1 \
2 0 }
8 40 i Wi/ N\
o I X f \\
W
_L 5 l \
W 1 Ay A\
»E | IS
N | v \\
S, ] .
g ! ; NN\
= | \ \
-
10 I —t \\ LY
I | LU
T v N AY
Y
5 AW
USING AVERAGE TOTAL CROSS LN
| SECTIONS AT 0.5-Mev INTERVALS/ \
, )\
10
0 2 4 6 8 10 i2
ENERGY (Mev)

Fig. 4.4.32. Calculated Uncollided Fast-Neutron Spec-
trum 4200 ft from the TSR-l for the Case in Which an Air

Filled Collimator Penetrates the Reactor Shield and

Points at the Detector,

Table 4.4.1. Average Cross Sections Used for

Calculation of Uncollided Flux

PERIOD ENDING SEPTEMBER 30, 1958

E, Neutron S Air E, Neutron S Air
Energy T'_] Energy T'_]
(Mev) (em™") (Mev) {em™")

0.5 13.7 x 10~3 6.5 6.59 x 10~
1.0 10.6 x 10~3 7.0 6.79 x 1073
1.5 9.23 x 103 7.5 6.79 x 10~3
2.0 8.19 x 10~ 8.0 6.59 x 10~
2.5 7.18 x 10~2 8.5 6.59 x 10~
3.0 8.06 x 1073 9.0 6.59 x 10>
3.5 8.50 x 10™° 9.5 6.59 x 10>
4.0 8.50 x 1077 10.0 6.59 x 10~
4.5 7.09 x 10> 10.5 6.59 x 10~°
5.0 7.04x 1077 11.0 6.59 x 10~°
5.5 7.24 x 1073 1.5 6.59 x 10~°
6.0 6.74 x 1073 12.0 6.59 x 107>

183
ANP PROJECT PROGRESS REPORT

4,5. TOWER SHIELDING REACTOR Il

C. E. Clifford

Design work on the Tower Shielding Reactor ||
(TSR-Il) is now essentially complete. The several
design changes which have been incorporated since
the previous report! and recent studies of operating
conditions and experimental facilities are described
below. The two shields which will be used initially
with the TSR-|| are also discussed.

MECHANICAL DESIGN?

The latest concept of the TSR-Il is shown in
Fig. 4.5.1. The design and much of the fabrication
of the major components inside the reactor tank have
been completed, although the fuel elements will be
modified following full-scale flow tests, The
present design differs from the one last reported
in several respects. The inside dioameter of the
upper portion of the reactor tank has been increased
from 36.75 to 39.75 in. to facilitate removal and in-
sertion of the fuel elements and the shielding above
the core. In addition, the ionization chambers have
been moved outward in the upper region of the inner
cylinder to provide more room for the control
mechanism positioning and actuating devices which
will be housed in a turret atop the central cylinder
region,

Another modification in the design has been the
elimination of one of the boral hemispherical shells
just outside the lead region. This will ensure
adequate clearances in the annular fuel region.

A 2-ft-thick shield of lead shot and water has been
added above the central fuel elements to reduce
the gamma-ray leakage through the central cylinder.
This will be especially important when the reactor
is enclosed in a shield mockup. This lead-and-
water shield, which can be seen in Fig. 4.5.1,

]C. E. Clifford and L. B. Holland, ANP Quar. Prog,
Rep. Dec. 31, 1957, ORNL-2440, p 275.

2The mechanical design work on the TSR-ll is being
done by the Engineering Department,

184

L. B. Holland

will be penetrated by approximately 120 helical
ducts through which the cooling water will flow.
The gamma-ray leakage through this shield will
never exceed the leakage through any shield used
with the TSR-ll. The shield shown above the core
in the region between the reactor tank and the inner
cylinder will be removable.

The design of the TSR-Il is such that it can be
encased in various shields. The first shield with
which the reactor will be used will be a spherical
shield having a beam hole both fore and aft (see
discussion of shield designs in subsequent section
of this chapter). The entire assembly will be sus-
pended from a platform, as shown in Fig. 4.5.2, and
it can be rotated around the vertical axis so that
the beam holes can sweep a horizontal plane. When
the reactor is suspended in this manner, the cooling
water will pass through a coaxial swivel joint be-
fore it enters and after it leaves the reactor tank.

The reactor and beam shield may also be sus-
pended so that the beam holes will sweep a vertical
plane, For this case, two 12-ft girder-like members
will be attached to the side of the shield through an
axle and bearings, and the cooling water will enter
through a swivel fitting and the hollow axle of one
support member and leave through a similar arrange-
ment on the other support member. By alternately
using the two beam holes, a beam can sweep the
whole 360 deg in the vertical plane without tilting
the reactor more than 90 deg from the vertical
position,

Calculations have indicated that, with a 30-deg
conical section removed from the shield, a 6775
ft-lb torque will be required to rotate the reactor
and shield so that the removed section will sweep a
vertical arc. A much smaller torque will be re-
quired to sweep the horizontal arc. A Char-Lynn
torque booster working through a worm and 180-
tooth gear will be used to rotate the reactor as-
sembly. A torque booster under full load has per-
formed satisfactorily in laboratory tests,

CONTROL TURRET ~—a

DRIVE GEAR
N
ROTATION BEARING ‘i‘ 7
o

s I!
7 NE
CENTRAL CYLINDER #i§
ASSEMBLY 95}
a1
oI
SHIELD SUPPORT A
SLEEVE 0

FISSION CHAMBER —-3

HELICAL

COOLING-WATER TUBES —&f{

INCHES

LEAD-WATER SHIELDING —

WATER
INLET

: - I_"_ nl 1 ——
N IRt A
- ]

-

5% 4
Pt

i

Zfi

N

— o o
o ©

-

il

ou. e? © a% o
g0 2

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED

ORNL-LR-DWG 32708R1{

WATER
OUTLET

PLATFORM

o=

7 L
5 s s,

SN

|

I,
¥— IONIZATION 8f1 4in.
M CHAMBER

6ft 7in.

y e~ ALUMINUM
REACTOR
TANK

" |8 PERMANENT
N LEAD-WATER
SHIELDING

CORE
REGION

TP T TT LT

- ,.
[ 22777700

CONTROL MECHANISM AND
SUPPORT ASSEMBLY

Fig. 4.5.1. TSR-ll Design (Vertical Section).

185

ANP PROJECT PROGRESS REPORT

o i
v TR N
A ‘ ’W _, !
{ i [ i,

CONTROL AND INSTRUMENT LINES

DRIVE MECHANISM AND SAFETY
MECHANISM OPERATORS

BEAM HOLE

- COAXIAL SWIVEL JOINT

UNCLASSIFIED
2-01-060-368

# ——— SUPPORTING PLATFORM FOR
HORIZONTAL ROTATION

SUPPORTING FLANGE FOR
VERTICAL RCTATION

Fig. 4.5.2. TSR-il in Beam Shield.

CONTROL SYSTEMS?

Control Mechanism

The experimental mockup of the control mechanism
for the TSR-II, which was described previously, '
was tested and found to have a short life owing to
excessive wear of the traveling nut of the position-
ing device, Several modifications of the design
failed to increase the life. Design changes were

186

then mode so that the pressure on the traveling nut
was reduced by nearly a factor of 10. This was
accomplished by reducing the size of the seal that
restricts the flow out of the chamber and increasing
the area on which pressure is applied to seal the
control plate assembly against the traveling nut.

3The controls system for the TSR-ll is being develop-
ed by the Reactor Controls Group.

The force on the traveling nut was reduced to 10 |b,
the difference between the spring force and the hy-
draulic force. Proper guidance and longer life were
ensured by using stellite bearings instead of minia-
ture roller bearings on the control plate assembly
and by using piston rings to provide a good seal
between the assembly and the cylinder wall. With
these changes, a model of the control mechanism
(see Fig. 4,5.3) was still operating successfully
when a test was terminated after 5000 cycles. The
model is shown in the test stand in Fig. 4.5.4. The
model will be tested later under flow conditions.

A total of five control mechanisms will be used to
operate neutron absorbing grids for shim-safety pur-
poses in the TSR-Il. A sixth grid, which does not
scram, will be used for fine control. The neutron-
absorbing umbrella-shaped grids will be made of
Inconel-clad cadmium ribbons. It had previously
been planned? to use perforated boron-loaded
plates as the neutron absorbers byt the fabrication
of the Inconel-clad cadmium was more feasible.

4C. E. Clifford and L. B. Holland, ANP Quar. Prog.
Rep. June 30, 1957, ORNL-2340, p 323.

%“ CONTROL PLATE

PERIOD ENDING SEPTEMBER 30, 1958

Electrohydraulic Transducer for Control
of Water Pressure in the
Control Mechanism

J. E. Marks?

An electrohydraulic transducer will be used to
actuate the control mechanism. The purpose of the
transducer is to allow electrical control of the water
pressure in the TSR-Il control mechanism with the
standard ORNL magnet circuitry. The transducer
is a bypass pressure regulator with electrical rather
than mechanical adjustment. As a force-balancing
device it compares the control mechanism water
pressure acting on a diaphram with the thrust of a
solenoid and applies the difference between the
two forces to a spring. Any differential force
causes the spring to deflect proportionately, and
the deflection is used to displace a valve plug away
from or toward its seat, depending upon whether
the water pressure effect is greater or less than the
thrust of the solenoid. This valve is in a shunt
relation to the control mechanism.

Slnstrumenfuficm and Controls Division.

UNCLASSIFIED
2-01-060-378

TSN SRR RSN SIRSRNLX
7777777%é%/// 7T 77l

// | 00 WATER FLOW
PN g oo

] T
\ S N S

\ \ 1IN OVERTRAVEL

58

LTI LL VAV AP

AN NN

RN

N

CLEANING RING/

\DRIVE NUT

Fig. 4.5.3. TSR.ll Control Mechanism.

187
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
Bl P0T0 43086

Fig. 4.5.4. TSR-1l Control Mechanism in Test Stand.

Between the source of water and the control
mechanism~transducer load, there is a restriction
of known magnitude. The load pressure is princi-
pally a sum of the source pressure less the pressure
drop across this restriction, and the pressure drop
is in turn proportional to the square of the total
flow through both devices. Hence the transducer
can affect the pressure seen by the control mech-
anism.

188

The electrical input to the solenoid is from the
circuitry of the standard magnet amplifier and sigma
bus. One design model of the transducer, shown in
Fig. 4.5.5, has been built and tested. With an in-
put of 100 to 0 ma, the control range was from the
operating pressure P to (P — 20) psig. This
range was sufficient to regulate the load pressure
between the scram condition and the stable operat-
ing condition of all the control mechanisms tried.
The load pressure followed a scram signal input
after a delay of 20 msec. This was adjudged as
satisfactory; however, a later design incorporating
several improvements is to be tested soon.

The advantages of the transducer over an “‘on-
off’" type of device, such as a solenoid valve, are
the same as those of magnet amplifiers in com-
parison with relays in conventional safety magnet
confrol circuits. Furthermore, the fransducer will
supply a regulated source of water at any required
pressure to improve stability and reliability of the
control mechanism operation.

TSR-Il Block Diagram
L. C. Oakes®

The logic of the TSR-II block diagram (shown in
Fig. 4.5.6) falls into the general pattern of that of
other ORNL reactors and therefore it is not dis-
cussed in detail here. However, many small dif-
ferences brought about by special features of the
TSR-1I have altered the specific modes of operation,
and these differences are described below.

Grid Motion. ~ Since the TSR-I| will have one
drive unit common to all five of the shim-safety
grids, it will not be possible to move one grid in-
dependently of the others. Thus any operation of
the drive is analogous to ‘‘group operation” in the
conventional reactor. For orderly operation, a means
for producing more modest changes in reactivity
must be available to the operator. For this
operation, a slow insertion and slow withdrawal
mode is provided wherein the drive runs at approxi-
mately 20% of full speed. All grid withdrawal is
stopped by 10-sec periods, and the grid drive
switching is so arranged that a 25-sec period will
stop fast withdrawal and automatically limit shim
motion to the slow-speed mode.

Avutomatic Shim Insertion. ~ Although automatic
shim insertion has been used previously, the
scheme presented here is aslight departure from
the normal concept. As shown in the block diagram,
when the servo is on and the regulating grid reaches

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED

2-01-060-38
FEEDBACK
ZERO ADJUSTMENT, DIAPHRAGM
SPRING AND —
CENTERING DIAPHRAGM SOLENQID ORIFICE
A A A
4 N N 4 N
N
000G N
N %
. A — D)ISCHARGE
% \'§
AN
)
3 in 1} — A Wi HHl” —— INLET
N
AN
- /
i__ —
Fig. 4.5.5. Electrohydraulic Transducer for Control of Water Pressure in TSR-Il Control Mechanism.

the lower limit, the shims will be inserted at the
slower speed until the regulating grid withdraws to
the intermediate limit switch. This action takes
place only if the period is more positive than

-100 sec.

There are two normal conditions which will
cause the regulating grid to drive to the lower limit
during operation. One will occur when, for some
reason, such as the reduction of the water inlet
temperature, the reactivity is increasing and the
regulating grid drives to the lower limit while
holding a constant power level. Under this con-
dition, since the reactor period is infinite, the
shims would be expected to drive in to re-establish
the regulating grid in the center of its stroke. The
other condition will occur when a demand for re-
duced power causes the regulating grid to drive to
the lower {imit. In the latter case, however, the
period is likely to be shorter than ~100 sec, and
automatic shim insertion will be inhibited. In this
condition a substantial shim insertion would drive
the reactor far subcritical, and upon reaching the
desired lower power level the regulating grid would
be driven to the upper limit with the reactor still
subcritical. The 100-sec period interlock eliminates
the need for reshiming after a reduction in power
level.

Fission Chamber Drive. —~ An automatic mode of
operation is provided for the fission chamber which,
at request, will hold the counting rate between 2
counts/sec and 10° counts/sec. This mode of
operation relieves the operator of the task of keep-
ing the counting channel on scale and contributes
to an orderly operating system. |n addition, since
the pulse preamplifier is contained in the same can
as that used to contain the fission chamber and sees
the same flux, the automatic mode of operation en-
sures that the electronic components will not be
unduly irradiated. When the automatic mode of
operation has not been requested, an annunciator
warns the operator when the counting rate has ex-
ceeded 10° counts/sec.

Test and Automatic Rundown Inhibit (TARDI). -
TARDI permits periodic testing of shim drive
mechanism performance without the encumberance
of certain operating interlocks. |t may be used only
so long as all seat switches indicate that the grids
are seated, the shim pump pressure is zero, and the
key switch is off.

Door Interlocks. = At the request of the operator,
the front door interlock may be temporarily blocked
to permit personnel to enter and leave the control
house at shift change without shutting down the
reactor. Permission to bypass the front door in-
terlock constitutes a mode of operation and is

189
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
ORHL-LR-QOWG 32707R{

| |

INSERT
- ’{ OFF }‘ ‘F|SSIONCH.I

CONTROL POWER

WITHORAW
FISSION CH.

AUTOMATIC
|

MAIN PUMP FLOW WATER FLOW
SEAL SEAL

>200 gpm BLOCK OUT
>100 cps l < 100 ¢ps

NO REVERSE
log # OR CRM
CONFIDENCE

MAIN PUMP FLOW J WATER FLOW

SERVO "ON"
TARD! "OFF"

T MORE POS.
THAN =100 sec

TARDI "ON"

REG. GRID IN OR [PS:"D
/| ABOVE INT. LIMIT
// >200gpm BLOCK OUT
rd
SEAL l_—.,

Lnu. GRIDS SEATED | | "NO" SAFETY TROUBLE i

START-UP PERMIT
STOP SEAL

PR R
> AST | RTJ Wl RTJ REG. GRID
" TARDI "ON | “F ST INSE | [ SLOW INSE IN LOWER LIMIT

_PERIOD > 10 sec
| DOCRS ANDO HOIST SEAL

NO FRONT DOOR BY-PASS | | ND INSERT NO WITHDRAW

INTERLOCK PERMIT
I FRONT DOOR MON!TRON—‘ REAR ODOOR MONITRON }
ORIVE NOT AT DRIVE NOT AT
UPPER LimIT LOWER LtMIT

I INSERT
NO CLUTCH WATER FLOW
Pee20ib BLOCK OUT
FRONT DOOR BY-PASS
SAFETY TROUBLE
INTERLOCK PERMIT PERIOD < 5 sec I ’ 2 CHANNELS FAILED
SAFETY LEVEL] To>180°F *

< {Omr/hr < {0 mr/hr
FISSION CHAMBER
NOT MOVING

{ SERVO ERROR | M WDR. REG. GRID | | I
WOR. GRID

|

PUMP FLOW

l LOW EMERGENCY

I SERVO ON J

log &/ LEVEL

REVERSE ]

>{0 X SERVO DEMAND >4.2 Wy
{ PERIOD > 25 sec | I PERIOD < 25 sec_l SCRAM RESET
s NO WATER FLOW
log # > 10 BLOCK OUT ISERVO 0FF1 | SERVO ON }
’ L NO INSERT] I NO INSERT j GRID DRIVE NOT
AT UPRER LIMIT

|

GRID DRIVE NOT
AT UPPER LIMIT

GRID DRIVE NOT

GRID DRIVE NOT
AT UPPER LIMIT

AT LOWER LIMIT

S

GRIO DRIVE NOT
AT LOWER LIMIT

REG.
GRID
INSERT

GRID DRIVE NOT
AT LOWER LIMIT

FISSION FISSION

CHAMBER

REG.
GRID
WITHDRAW

SLOW
GRID
WITHDRAW

FAST
GRID
WITHDRAW

CHAMBER
INSERT

WOR.

INSERT INSERT

HOIST PCWER

WATER FLOW
BLOGK OUT
CONTINUOUS
BELL Sw. rcHANGEOVER l

L PUMP DELAY

s

EXP. OR OPER.
DISABLE FRONT BY-PASS TARDI = TEST AND AUTOMATIC RUN DOWN INHIBITOR
HOIST MOTION REQUEST DOOR BY-PASS "ON"
REACTOR REACTOR
I 5-sec DELAY }-—— IN POCL OUT OF POOL

DEADMAN SW.

— 3-min DELAY log # <107

ANY GRID
OUT OF SEAT

|
; ‘—i 5-sec DELAY

START-UP DELAY

FRONT

NC
FRONT DOOR

REACTOR OCOR
HOIST CONTINUOUS MOTION START-UP BvePASS BY-PASS
B
PERMIT BELL PERMIT PERMIT INTERLOCK INTERLOCK

PERMIT PERMIT

Fig. 4-506- Block Dingram of fhe TSR'IIO
190

described as follows: When the log N indicates
that the power level is below 500 w the operator
may request that the front door interlock be by-
passed. While the bypass is in effect the following
conditions prevail:

1. Personnel may enter and leave the front door
without producing a scram.

2. No shim grid withdrawal is permissible.

3. The servo will remain in operation to counter-
act reactivity changes resulting from such things as
reduction in the temperature of the cooling water.

4. Any rise in power above 500 w will produce a
reverse.

At the request of the operator, normal operating
conditions may be restored. Before permission is
granted to withdraw shims, however, the horn which
is always sounded prior to reactor operation at the
TSF must again be sounded for 3 min to warn per-
sonnel who may have been left outside the building.

Water System Interlocks. — There are two modes
of operation of the reactor from the standpoint of
the water system: the power mode and the test
mode. With the power mode the reactor is operated
at or near full power, and the flow is varied between
200 and 1000 gpm over the power range. With the
test mode the reactor is operated at a power level
sufficiently low to prevent overheating with no
water flow. This is brought about by a test switch
designated throughout the block diagram as ‘‘water
test.”’ Prior to a startup for a power run the main
pump must be operated at the full flow of 1000 gpm
for 3 min in order to produce sufficient mixing of
the water to remove any severe temperature gradients
from the system. This pump delay may run con-
currently with the startup delay. Failure of the
main pump will produce a scram rather than a re-
verse, The more drastic measure is preferred in
order to lessen the afterheat probiem,

The small volume of water in the reactor vessel
and the exposure of the vessel and interconnecting
hoses to the cold weather bring about special re-
quirements with respect to afterheat and winter
freezes. A highly reliable turbine pump which will
deliver between 10 and 40 gpm to the reactor through
the main pump lines will be operated 24 hr a day.
Since this emergency pump must remove the after-
heat in the event of main pump failure, it will be
operated from an a-¢ to d-¢ converter and a bank of
60 station batteries capable of running the pump at
full delivery for 8 hr or more. If a failure of the
emergency pump should occur, the reactor will be

PERIOD ENDING SEPTEMBER 30, 1958

reversed to shutdown level. An electric heater in
the emergency pump line will maintain a safe temper-
ature when the system is unattended in the winter.

Reactor Controls Cubicle

The reactor controls cubicle has been delivered
and set up at the Tower Shielding Facility. The
cubicle, which is 14 ft long, 7 ft 5 in. high, and
7 ft deep, is shown in Fig. 4.5.7. It is designed
for unattended operation and is patterned after the
cubicles for the Bulk Shielding Reactor and the Pool
Critical Facility. The cubicle will be partially wired
and then moved into its final position when the con-
trol console and vertical board for the Tower
Shielding Reactor | are removed.

HEAT REMOVAL EQUIPMENT

The portion of the 5-Mw heat removal system which
is being constructed by an outside contractor should
be completed in October., The only holdup at present
is the delivery of one pump to the contractor. The
system will be subjected to a hydrostatic test with
the 6-in. lines to the tower legs capped off. A flow
test of the system will be made by short circuiting
the 6-in. pipes to the tower legs with a 6-in. hose.

In order to provide for the event that the studies
by Gambill and Hoffman (see below) indicate that
a Reynolds number of 7000 is necessary, an in-
vestigation is being made of the cost of raising the
cooling water flow rate or raising the system static
pressure or both to achieve the desired Reynolds
number for turbulent flow.

NUCLEAR CALCULATIONS
M. E. LaVerne

Nuclear calculations which were performed for the
TSR-1l with the 3G3R Oracle code were reported
previously. '+6=7 A multigroup, multiregion reactor
code® has now become available, and an adaptation
of this code to a Goertzel-Selengut type has been
used to perform a number of further calculations for

the TSR-|l, as discussed below.

8C. E. Clifford and L. B. Holland, ANP Quar, Prog,
Rep. March 31, 1957, ORNL-2274, p 294.

7C. E. Clifford and L. B. Holland, ANP Quar. Prog.
Rep. Sept, 30, 1957, ORNL.-2387, p 304.

Bw. E. Kinney, R. R, Coveyou, and J. G. Sullivan,
Neutron Phbys, Ann. Prog. Rep. Sept, 1, 1958, ORNL.-
2609, p 84,

191
ANP PROJECT PROGRESS REPORT

T cLassiFieD |
PHOTO 44sds
2 STSS

Fig. 4.5.7. TSR-1l Reactor Controls Cubicle.

Comparison of Calculations and
Critical Experiments

In order to confirm the validity of this calcu-
lational method and also to determine the effect of
a lead-boral layer outside the fuel region, calcu-
lations were first performed for three critical ex-
periments which were set up in finite slab geometry
as shown in Fig. 4.5.8 (the figure shows only one-
half of each configuration). Each fuel region con-
sisted of a 4 by 6 array of standard BSR fuel ele-
ments, and boral sheets simulated the control
plates. Clean, or nearly clean, experiments® were
obtained for configurations A and B by heating the
water in the assembly tank and simultaneously

9These critical experiments were performed by D. F.
Cronin, J. K. Fox, and L. W. Gilley.

192

withdrawing a cadmium rod to maintain the assembly
critical. On configuration B, a portion of the rod
(2.75¢ worth) was still inserted when the rod travel
was used up.

Configuration C was set up to evaluate the effect
on the reactivity of removing the lead and boral
layers. Because estimates of excess reactivity at
room temperature for both configurations B and C
were available from rod calibrations, configuration
C was not heated to attain the ‘‘clean’’ state.

A comparison between the results of experiments
and the calculations is presented in Table 4.5.1.
The error for the three cases averages about 1%.
The experimental & is the multiplication constant
that would have been observed with the rod com-
pletely withdrawn, i.e., in the “‘clean’’ condition.
Previously computed temperature coefficients were

PERIOD ENDING SEPTEMBER 30, 1958

Table 4.5.1. Comparison Between Calculated and Experimental Values of Multiplication Constant

Water

Correction to

T Multiplication Constant, k Calculated k& Corrected Error
Confi ti emperature
nrguration o Experimental Calculated . Thermal Calculated & (%)
("F) Density
Base
A 108 1.0000 0.9897 0.0014 0.9911 0.9
B 118.9 1.0002 0.9872 0.0018 0.989%90 1.1
C 73.4 1.0081 0.9921 0.0035 0.0002 0.9958 1.2
soamre o ez
-10 | k
= I J
zm™ ! o
[ Jwarer w E 15 N &
L3 Q| y. I 8
b - =20 8[ Y T w
8 g al \ / I _1
| 7 o% -es| &l — E
0 q w | o
> | I
CONFIGURATION A 30 | |

CONFIGURATION B

i

CONFIGURATION C

[ ] . 1 ; | L 1 1 ] L | J
0 10 20 30 40 50 60
{cm)

Fig. 4.5.8. Configuration for TSR-1l Critical Experi-

ments in Slab Geometry.

used to correct the calculated & for differences be-
tween the experimental and the calculational thermal
base and dens ity.

Reactivity Coefficients

Void coefficients for the core were computed from
the change in the multiplication constant produced
by deleting the water (but not the aluminum or
uranium) from thin spherical annuli in the core. The
results are shown as the points in Fig. 4.5.9. A
“teasonable’’ curve was drawn through the calcu-
lated points and volume-averaged to obtain a mean
void coefficient of ~2.15 x 1074 Ak/k (in per cent)

20 25 30 35 40
REACTOR RADIUS (cm)

Fig. 4.5.9. Core Void Coefficient as a Function of the

Reactor Radius.

per cubic centimeter of void, This is to be com-
pared with a value of =3.5 x 10=# which was reported
for SPERT |.19

Temperature coefficients were calculated both for
density and for thermal base changes. The net
temperature coefficient was =0,77 x 102 Ak/k
(in per cent) per °C, which can be compared with a

valve of -0.9 x 10~ 2 for SPERT 1.0

For various reasons (for example, cumulative
tolerances in fabrication), core boundaries may de-
viate from the desired locations, An estimate of
the effect of such deviations on the reactivity was
made by deleting uranium from a thin spherical
annulus at the outer edge of the core while the
total amount of uranium was kept constant. The
radial coefficient was found to be 0.23 Ak/k (in per
cent) per centimeter,

These reactivity coefficients are summarized in
Table 4.5.2, in which a mass coefficient is also
given.

10F . Schroeder et al., Nuclear Sci, and Eng, 2, 96
(1957).

193

ANP PROJECT PROGRESS REPORT

Table 4.5.2, Summary of Computed Reactivity
Coefficients for the TSR-!l

Average void coefficient for core, -2.15 % 10-4
Ak/k (in per cent) per cubic

centimeter of vaid

Temperature coefficients, Ak/k
(in per cent) per °C

~1.39 x 10=2

Density coefficient

+0.62 x 10=2

Thermal base coefficient

Net coefficient -0.77 x 10-2
Radial coefficient, Ak/k (in per +0.23
cent) per centimeter
Mass coefficient, (Ak/k)/(Am/m) 0.30

Control *‘Shell’’ Effectiveness

The control grid effectiveness was recalculated
as a function of the separation distance from the
fuel, with the contral grids being treated simply as
thin regions or shells. The results are shown in
Fig. 4.5.10. The shape of the curve is in good
agreement with the results of the earlier 3G3R
Oracle calculations, ! although it is a factor of
5.2 higher. The apparent discrepancy is due
primarily to the difference in absorption cross
sections used for the shells. The experimental
data shown in the figure are from the Bulk Shielding
Reactor test with eight solid stainless steel plates,
which was described previously.!! The BSR data
are normalized to the calculated values at the first
experimental point.

CONTROL GRID EFFECTIVENESS

The nuclear model used for the calculations de-
scribed above is a complete shell which is assumed
to move uniformly away from the fuel region, and
therefore it does not duplicate the control grids
which will be used in the TSR-ll. The actual con-
trol and safety grids will cover less than one-half
of the core-reflector interface, and, because of the
fixed radius of the plates, the edges of the plates

e E. Clifford and L. B. Holland, ANP Quar. Prog,
Rep. Junme 30, 1957, ORNL-2340, Fig. 5.3.6, p 327.

194

UNCLASSIFIED

100 2-01-060-54
] ] I ]
———
| ——= ORACLE CALCULATIONS:
46-cm-dio INTERNAL WATER REFLECTOR, —]
50 14 - ¢cm - thick CORE,
2-cm-~thick LEAD,
L 0635-cm-thick BORAL,
20 - cm-thick EXTERNAL|WATER REFLECTOR
| |
T
e I B
_~ EIGHT 77-mil-thick STAINLESS STEEL PLATES
X 20 IN LOADING 628, FOUR CENTER ELEMENTS,
x VALUES NORMALIZED AT FIRST EXPERIMENTAL
< \ POINT.
. .
> 4
[ \
o>
i 7o) "
o b N
< *o\
g ]
[ — . ”"
L T
5 .
\\
N
2
0 q 2 3 4 5 6 7

SEPARATION DISTANCE (cm)

Fig. 4.5.10, Calculated Effect of a Boral Shell in the
Internal Water Reflector on the Reactivity as a Function
of the Distance Between the Shell and the Core: Com-
parison with BSR Data.

will move away from the core at a slower rate than
will the center of the plate.

In order to determine the worth of an actual grid,
the grid area was divided into rings with their
common center at the center of the grid, and the
worth of each ring was determined as a function of
its normal distance from the core. The total worth
of the grid was then defermined by summing the
worth of all areas of each grid for a given separation
distance between the core and the center of the
plate. The results are presented in Fig. 4.5.11.
The Inconel-clad cadmium ribbons in the regulating
grid and in the lower shim-safety grid will be
spaced ]’46 in. apart to permit water to flow through
the internal reflector region; therefore, these plates
are worth only one-half as much in reactivity as the
remaining shim-safety grids.

After the control grid worth was determined, the
thermal-neufron flux distribution throughout the
reactor (see Fig. 4.5.12) and the source distribution
in the core (see Fig. 4.5.13) were determined from
the multigroup calculations with the control grid po-
sitioned for operation at 5 Mw.

UNCLASSIFIED
2-01-060-50
5 ; I [ I

TEMPERATURE 66 °F~COLD CLEAN

4 \\ ATION IN SEPARATION DISTANCE

REACTIVITY CORRECTED FOR VARI-

x OVER PLATE AREA,
S |
2 3 P :
g \\ALL SAFETY GRIDS
>
E
E 2 \\
O
<I
&J \\
— CONTROL GRID
o I B
o} 1 2 3 4 5 6 7

SEPARATION (cm)

Figs. 4.5.11. Reactivity Worth of the TSR.ll Regulating

and Shim-Safety Grids as a Function of Distance from the

PERIOD ENDING SEPTEMBER 30, 1958

SELECTION OF CRITICAL MASS FOR TSR-ll

Nuclear calculations were performed previously
with the 3G3R code to aid in establishing the
critical mass of the TSR-Il. When the effects of
the control shells, the external shield, xenon build-
up, fuel burnup, and possible errors in fuel loading
were considered, the calculated critical mass for
the reactor was 8.2 kg of U235 (see ref 1). Since
the calculations and critical experiments reported
above have indicated that the TSR-Il lead-boral
shield does not measurably change the critical
mass, additional critical mass calculations were
performed with the multigroup, multiregion code for
a clean, cold, water-reflected, spherical annulus
with the TSR-|| core dimensions and volume
fractions. Sufficient U?3° was then added to in-
crease the reactivity by 6.8 Ak/k (in per cent).

The results gave a loading of 8.1 kg of U?3% for the

Fuel. TSR-Il core, and this was set as the critical mass.
UNCLASSIFIED
2-01-060- 45
s BORAL Al Al—H,0 BORAL
(x10°°) H,0 \ H,0 / CORE —\ Pb [ H,0
7 ‘.0.’.0.00‘
S
/

N‘-fi

: \

[®]

: \

w

[y

e {6

5

@

£

>

>

c {2

—

=

o s

1

-

)

=

4 0.8 \

I

=

x

ul

I

-

0]
0 {0 20 30 40 50 60 70

RADIUS {cm)

Fig. 4.5.12, Thermal-Neutron Flux in the TSR-1! as a Function of Radius: 5-Mw Power Level.

195

ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2-0{-060-44

150 — Coe e e e

SQURCE INTEGRAL
4T

N - —]

= 150.326 neutrons /sec

140

SOURCE INTEGRAL AT 5 Mw, 3.84 x 10'7

130

120

"o ———- [

SOURCE DISTRIBUTION (neutronsscm 7sec)

{00

90

| : !

80 i ! | 1 i

23 25 27 29 L1 33 35 37
RADIUS IN FUEL REGION (cm)

Fig. 4.5.13, Source Distribution in the TSR-Il Fuel
Region as a Function of Radius: Boral Shell 4 cm from
the Fuel.

Of the 6.8% increase in the reactivity, 1.5% is to
cover the worth of the control grids in the fully
withdrawn position, 0.3% is allowed for structural
material in the core region, 1.4% is to cover the
worth of fixed neutron absorber plates in the in-
ternal reflector region, 2.0% is to allow for an error
in the calculations, and 1.6% is the amount that is
controlled by the control grids.

The 1.6% that is controlled by the grids represents
approximately one-half the reactivity worth of the
shim-safety grids and covers the following: 0.6%
to counteract the negative temperature coefficient
of reactivity from 40 to 180°F, 0.3% excess re-
activity necessary to permit servo control, 0.6% to
counteract xenon buildup for 8 hr of operation at 5
Mw, and 0.1% for short-time burnup of U233,

The 0.3% allowed for the structural material in the
core region was determined by calculating the re-
activity change caused by deleting the U235 from
shells in the core region. An integral value was
obtained by summing, as a function of radius, the
areas of these shells which were intersected by
the cylindrical support members.

196

The 1.4% is to cover the fixed neutron absorbing
plates which will be mounted in the internal re-
flector region near areas of the core-reflector in-
terface not covered by the movable plates. These
plates may be adjusted when the reactor is unloaded «
so that some shimming may be done in the initial
loading, as long as sufficient excess reactivity is
left to provide 0.8% for U233 burnup.

A 2.0% excess was allowed for calculational
errors because the comparison of the calculation
with critical experiments showed an average error
of this amount,

The TSR-I| elements (see Fig. 4.5.14) have been
fabricated to contain the 8.1 kg of U2 established
as reported above, However, since the geometry of
the core permits only a limited adjustment of excess
reactivity by control plates and no means of adding
excess reactivity by the insertion of additional
fuel elements, the fuel elements have been assembled
in a temporary fashion so that the core loading can
be altered by changing a limited number of fuel
plates if a critical experiment should indicate that
it is necessary. Such a critical experiment will be .
performed at the Critical Experiments Facility.

When the fuel elements are assembled in their final
form the worth of the control plates will be deteri-
mined and some of the reactivity coefficients will be
measured.

CALCULATION OF HEATING IN
THE FUEL PLATES

A hand calculation was performed to estimate the
heating per fuel plate in each element. The source
distribution given in Fig. 4.5.13 was used to de-
termine the power distribution in each fuel plate
as a function of the length. The power generation
for some of the plates in the central fuel elements
is shown in Fig. 4.5.15 and for the annular elements
in Fig. 4.5.16. There are twelve annular elements
and eight central elements. The total power per
fuel plate was found by integrating the area under
the distribution plots for each fuel plate. The re-
sults are plotted as a function of the fuel plate
number for the central fuel elements in Fig. 4.5.17 .
and for the annular fuel elements in Fig. 4.5.18.
These values will be used in the flow distribution
studies to determine the required flow per fuel .
channel (see below).

PERIOD ENDING SEPTEMBER 30, 1958

|UNCLASSIFIED
PHOTO 45083

Fig. 45.14. Photo of TSR-Il Fuel Elements.

197
861

UNCLASSIFIED
2-0i-060-46

i \ |
76 -~ |-&— EDGE DISTANCE (VARIES FROM 0.250 in. FOR PLATE NO.
\\ TO 0.486 in. FOR PLATE NO. 40).

72 i’/_‘*“‘_—"__'—_"'_‘"\{ 1 ]

| |

| (

.
68 e /T FUEL REGION
\ \ FUEL PLATE
64 \ \ PLATE AVERAGE POWER GENERATION —---rd
NO. (Btu /hr/t2)
PLATE NO.

40 4 53,980
™ \ 10 54,270 ]
N 20 54,650
\ 30 30 55,340

40 56,980

56 SN20 N U

52 10’ s \\
N\
N

60

]

POWER GENERATION (BTu/hr/ffz)

48

N
A\ NN

40

10 12 14 16 18 20 22 24 26
FUEL LENGTH {cm)

O
no
D
2}
@

Fig. 4.5.15. Power Generation in the Fuel Plates of a Central Fuel Element as a Function of the Distance from the Inside Edge of the Fuel.

LId0d3IY §SSIH00dd LI23F0¥d ANV
661

POWER GENERATION (Btu/hr/ft2)

UN

CLASSIFIED

2-01-060-47

80 I '
76 \\ \ | HORIZONTAL MIDPL ANE
; PLATE AVERAGE POWER GENERATION
S M \ u/hrsfte
7 \\ q {_““—I ______ )i--— FUEL REGION Z?' (8;91:6;” )
! 42 58,190
\ | 44 55,520
68 \ N FUEL PLATE :S 5;‘31:?323
\ \ 53 51,350
ca N~ . 56 50,010
\ PLATE NO. 58 48,510
T d o
60 \\44 — 7 41,350
\\\:wEE\
56 e . =
\ 53 ‘\ Q
o '\_‘\5856 \\\s\\\
44 I— 66 \‘\\ X ™~ \ N
74 \ \\ \I
w0 — 7 T 1T
o) 2 4 6 8 10 12 14 16 18 20 2 24 26 28 30

FUEL LENGTH {(cm)

Fig. 4.5.16. Power Generation in the Fuel Plates of an Annular Fuel Element as a Function of the Distance from the Horizontal Midplane to

the Edge of the Fuel.

8561 ‘0 ¥IEWILJ3IS ONIONT a0l¥3d
00C

POWER PER FUEL PLATE (kw)

10

UNCLASSIFIED

2-01-060-42

INNER RADIUS OF PLATE{4=1.600 in.
PLATE THICKNESS = 60 mils A
CENTER-TO-CENTER SPACING OF FUEL PLATES =180 mils
NUMBER OF CENTRAL FUEL ELEMENTS = 8

(4 ABOVE MIDPLANE, 4 BELOW)

///
5 10 15 20 25 30 35 40

CENTRAL FUEL ELEMENT PLATE NUMBER

Fig. 4.5.17, Power Generation in the Fuel Plates of a Central Fuel Element as a Function of the Fuel Plate Number.

1Y0d3IY SSIYO0¥d LD3rodd dNV

102

POWER PER FUEL PLATE (kw)

UNCLASSIFIED
2-01-060-58

14
2
10 \_\\\

8 ™.

5 N

INNER EDGE OF PLATE 41 IS AT 8.945 in.

4 ————— PLATE THICKNESS = 60 mils \
CENTER-TO-CENTER SPACING OF FUEL PLATES = {80 mils

NUMBER OF ANNULAR FUEL ELEMENTS = 42

40 45 50 55 60 65 70
ANNULAR FUEL PLATE NUMBER

Fig. 4.5.18. Power Generation in the Fuel Plates of an Annular Fuel Element as a Function of the Fuel Plate Number.

75

8561 ‘0 YIHWILHIS ONIANT QOiy¥3d
ANP PROJECT PROGRESS REPORT

FLOW DISTRIBUTION STUDIES

Flow Studies of Central Fuel Elements

The hydraulic flow test stand which has been as-
sembled to test the confrol mechanisms under 1000-
gpm flow conditions is shown in Fig. 4.5.19, and
the complete piping arrangement and holdup tank is
shown in Fig. 4.5.20. This test stand will also be
used to study the flow distribution in the central
fuel elements and in the internal reflector region.
The stand will be extended to accommodate not only
the central fuel elements but also the lead-and-
water shield above the core region.

Flow Studies of Annular Fuel Elements
W. R. Gambill'? H. W. Hoffman'?2

The full-scale, quarter-sphere model constructed
for study of the flow features of the annular fuel
elements in the TSR-I| core is shown in two stages
of assembly in Figs. 4.5.21 and 4.5.22. Figure
4.5,21 shows the containment fixture, the ends of
an inner fuel plate assembly, and the separator
plate. In Fig. 4,5.22 the second-pass *‘‘orange

' or outside fuel plates, have been placed in

slices,’
position and covered with a Lucite shell and back-
up cage. The groove in the head-end plate (shown
in the left lower corner of Fig. 4.5.22) receives a
pass rib which separates inlet and exit water fiows.
A movable mockup of the control plate assembly is
located in the first pass cavity separating the two
The entire model is
The separation between

inner fuel element sections.
fabricated from aluminum.
the inner surface of the Lucite shell and the outer
circumference of the outside fuel plates is approxi-
mately 0.5 in.

The static pressure drop across the full length of
every other flow channel is measured with manom-
eters connected to pressure taps located at channel
inlets and exits. Taps for the outer (second pass)
flow channels are visible in Fig. 4.5.21. The
manometer fluid is a CCl -1, solution.

The flow model, horizontal for the present tests,
will later be placed in a vertical position. The
flow rate is measured with a calibrated orifice and
a differential pressure (DP) cell, the nozzle inlet
and exit pressure with Bourdon-type gages, and
the water temperature with a standard Weston pipe-
line thermometer. The watertemperature may be

12Reactor Projects Division.

202

varied from 58 to 102°F with a steam-heated heat
exchanger. Tests to date have been made at a con-
stant flow rate of 250 gpm.

The measured static pressure drops for the model
in the ‘“‘as-received’’ state are shown in Fig. 4.5.23.
Variation of control plate position had very little
effect on the pressure drops for the second-pass
flow channels, as was expected; since all the first-
pass flow must reverse through the 0.5-in. gap
around the separator plate.

Extreme radial asymmetry of flow is evident. A
very large portion of the first-pass pressure drop
is apparently due to flow complications in the
central control cavity. It is probable that some sort
of coupled vortex flow pattern, induced by the
geometry and the control assembly mockup, exists
in the central cavity.

In the outer region of the core it was found that
most of the water entered the first seven channels
and that flow in the remaining outersregion channels
was either negligible or slightly reversed. Visual
observation of air bubbles, injected upstream of the
first pass, confirmed the right portion of Fig. 4.5.23.
Thus, it was observed that the bubbles streamed
through the first few channels but were in reverse
flow in the outermost channels.

Velocities have been calculated from the pressure-
drop data of Fig. 4.5,23. These are relatively un-
informative for the first pass because of the large
extraneous flow resistance associated with the
central cavity. For the uninterrupted channels of
the second pass, however, the results indicating
velocities from —2 to + 11 ft/sec are considered to
be valid. Calculations were based on the recent
parallel-plate relations given by Rothfus er al., 13
which indicate a Reynolds number of ~ 7000 as the
upper extension of the transition region or lowest
level of fully established turbulence.

Preliminary calculations indicate that the core
hydrodynamics will be more of a problem than the
heat transfer as such. |f velocities are sufficient
for full turbulence to exist, fuel element surface
temperatures will not exceed saturation temperatures
even at low static pressure levels, To maintain
full turbulence in all channels, however, a larger
total flow rate may be required. For a Reynolds
number of 7000 in each first-pass flow channel, a

]3R. R. Rothfus et al., J. Am. Inst, Chem, Eng. 3,
208 (1957).

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED
PHOTO 44878

Fig 4.5.19. Hydraulic Flow Test Unit.

203

ANP PROJECT PROGRESS REPORT

Fig. 4.5.21. Full-Scale Quarter-Sphere Model for Flow Test Studies of Annular Fuel Elements (Before Instal-
lation of Outer Pass).

204
AP(psi)

PERIOD ENDING SEPTEMBER 30, 1958

Fig. 4.5.22. Full-Scale Quarter-Sphere Model for Flow Test Studies of Annular Fuel Elements (Outer Pass and
Plexiglas Shell Installed).

o8

06

oz

UNCLASSIFIED
2-01-060-48

INSIDE ZONE (FIRST PASS)

T
OUTSIDE ZONE (SECOND PASS)
! |

e
I

|
T T
| |

CONTROL PLATE AT MAXIMUM RADIUS
[ e 5 0| \/]
‘ |

1| FLOW RATE: 250 gom

e —

WATER TEMPERATURE: 58°F

| MODEL POSITION: HORIZONTAL

o

| | |
CONTROL PLATE HALF-EXTENDED \‘[;//
1 f /

CONTROL PLATE AT MINIMUM RADIUS —/ |

T
A
/ T

RS SN

~— SEPARATOR PLATE

|
|

6 20 24 28 32 36 40 44 48 52 56 60

FLOW CHANNEL NUMBER

Fig. 4.5.23. Radial Variation of Pressure Drop in the Fuel Element Flow Test.

64

&8

205
ANP PROJECT PROGRESS REPORT

total flow of 910 gpm is necessary; for the outer
region, a flow of 1380 gpm is required. These

tlow rates are based on an equivalent diameter

that is equal to twice the water gap between the
plates and an evaluation of viscesity at mean pass
temperatures (131°F for the first pass and 148°F
for the second pass). The fuel element surface-to-
centerline temperature rise appears to be negligible
for the worst operating condition.

The next test will be made with hot water and with
sealing strips around the equator of the outside re-
gion to stop a small bypass stream observed earlier,
Later tests will involve placing flow resistances
(perforated plates or screens) in the inlet ends of
the smaller-radius outer-region flow channels. It is
hoped that a redistribution of flow-satisfying heat-
transfer requirements can be attained in this fashion.
Flow redistribution in the first pass is potentially
a considerably more difficult problem, and studies
of first-pass flow have barely begun.

A flow test in which cold water will pass through
a single channel of an outside fuel-element section
is being prepared. The flow rate and static pressure
drop will be measured, and a calculated friction
factor vs Reynolds number plot should reveal the
extent of the transition region for the particular
entry conditions and curved fuel plates of the

TSR-]| reactor.

REACTOR CORE KINETIC STUDIES

The results of the initial kinetic studies of the
reactor core, which were made on the ORNL Re-
actor Controls Analog Facility, 4 are presented
here without analysis. The first two cases were
run to determine how well the control mechanism
could protect the reactor without the inherent safety
features of a heterogeneous reactor. Three other
cases were run to partially investigate the self-
limiting features of the reactor based on the temper-
ature coefficient of reactivity.

With the reactor system simulator operating at an
initial power of 5 Mw and in a prompt critical con-
dition, excess reactivity was added and the simu-
lator delayed neutron circuits were closed. The re-
sulting power excursion was recorded and the safety
rods were allowed to drop when a level of 7.5 Mw
was reached. Concurrent with the reactor power

M4 This investigation was made by R, K. Adams, F. P,
Green, and E. R. Mann of the Instrumentation and Con-
trols Division.

206

excursion the mean temperature of the exit fuel
section was recorded. Both the power excursion
data and the mean temperature are plotted on Fig.
4,5.24a. The net negative temperature coefficient
of the reactor was disconnected; hence no limiting
effects of temperature coefficient were present.
For purposes of comparison, the time vs reactivity
profile of the safety grids is plotted on the same
figure (see Fig. 4.5,246). The same type of in-
formation for an initial power of 5 Mw and a level
trip at 6 Mw is presented in Fig. 4.5.25. The
safety grid profile shows that the insertion time for
the grids was longer in this case.

With the reactor operating at an initial power of
0.5 Mw and with the net temperature coefficient of
the reactor “‘active,’’ the water temperature leaving
the air cooler was reduced from 145.8 to 32°F.
(There is an approximately 130-sec delay time be-
fore the water reaches the reactor.) The resultant
power excursion and system temperature changes
are plotted as functions of time in Fig. 4.5.26.

The value which was used for the temperature
coefficient of reactivity (Ak/k) was ~6.5%. Later
calculations and critical experiments show the
value to be near that of the BSF (see above).
These cases will be re-run. Figure 4.5.27 pre-
sents the same type of data for an initial power of
S Mw,

With the reactor operating at 5 Mw, an abrupt stop
in cooling water flow was also simulated, The re-
sulting reactor power and system temperature changes

are plotted in Fig. 4.5.28.

SAFETY SYSTEM MEASUREMENTS
J. E. Marks

The evaluation of the TSR-I[ Safety System was
based upon a value of delay time and a time function
of control plate position that were determined ex-
perimentally with a test assembly incorporating the
best designs of the proposed system components.
The measuring techniques described below were
considered capable of yielding the accuracy re-
quired for a reliable reactor analysis.

The test assembly included a simulated sigma bus
circuit, a standard magnet amplifier, an electro-
hydraulic transducer, and a control mechanism with
connections between these components similar to
those expected in the final reactor assembly. A
differential transformer type of linear motion trans-
ducer was mechanically connected to the control
plate to translate the position of the plate into a

1000

900 —

800

700

600

500

REACTOR POWER (Mw)

400

300

200

100

|
N d

REACTIVITY (% Ak/k)

i
W

Fig. 4.5.24. (a) Reactor Power Level and Fuel Temperature Rise as a Function of Time After Insertion of

PERIOD ENDING SEPTEMBER 30, 1958

UNCLASSIFIED

2-01-060-57
N [
POWER EXCURSION - POWER EXCURSION
| PERIOD = 070 ¢ 102 1 |
970 x 10° Mw 8.6 msec 2 300 1O x40 w‘""\/-
| w
o« 2
|~ 9.0msec g s 8.70 x (0™ Mw
= g
8.70 x10% Mw o E \
w ¥ 200
B s ’
=
o d 5.55 x 10° Mw
Z 100 f
/ z 2.7 %102 Mw
w
=
0
[ 0 40 80 100
5.55 x {0° Mw TIME (msec)
{0.5 msec
2.7 x10% Mw
%\ 1.0 msec
g
INITIAL / \
POWER
5 Mw é
| S
LEVEL TRIP AT \
7.5 Mw \
SAFETY GRID PROFILE
(TIME AFTER TRIP)
\\
]
\
0 20 40 60 80 100 120 140 160 180
TIME (msec)

200

Various Amounts of Positive Excess Reactivity with Power Overshoot Being Limited Only by Safety Grid Insertion
After . TSR-Il Reaches 7.5 Mev. (b) Rate of Reoctivity Insertion by Safety Grids as a Function of Time After Scram.

207
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED

2-01-060-52
1000
POWER EXCURSION
900 : b —m 3
PERIOD 0.9 x 10
16.5 msec /
800 — e e
[
w
3
Q_ TOO — - [ o —_— . - — [ I E—
.|
= 0.7x10°
=
Z
o 600 [t — -
}—.
.|
l
2 3
Y 0.5x10
500 _
[
w
=
(@]
a
& 400 - —
=
(&)
g
w
[0}
© 300
—
L«
& 3
0.25 x 10
22.3 msec
200
100
INITIAL
POWER
-0 \
o &_
0 , ——
LEVEL TRIP AT 1.2 x I—
INITIAL POWER \
X - \ —
3 SAFETY GRID PROFILE \
a (TIME AFTER TRIP)
: \
>
}—.
a9
L
@x -3 \_\
-4
0 20 40 60 80 100 120 140 160 180 200

TIME (msec)

Fig. 4.5.25. (a} Ratio of Power Level to Initial Power as a Function of Time After Insertion of Various
Amounts of Positive Excess Reactivity with Power Overshoot Being Limited Only by Safety Grid Insertion After
TSR-H Reaches 6 Mw. (b) Rate of Reactivity Insertion by Safety Grids as a Function of Time After Scram.

208

PERIOD ENDING SEPTEMRER 30, 1958

UNCLASSIFIED
2-01-060-53
, 50

500 T ‘ 1 ‘ 1
i ot ’/
| /V ’/{//
400 ‘ . ’
‘ ‘ / - | Z
| EXIT MEAN ‘ T #’
FUEL TEMPERATURE
-— //
L 300 — ;, , : ~—REACTOR POWER .
[} | 1 /»' y
x = ‘ o
P | v
BOILING POINT—="1"
E L -~ J o r%
: /
H 200 : - v s
EXIT MEAN ‘ / / R
FUEL TEMPERATURE — o
F | A /< _74.7
! i —— -
EXIT SECTION MEAN /\\:\__ —] <UEL EXIT
WATER TEMPERATURE WATER TEMPERATURE
! ‘ 1 ‘ ‘
| | /‘iREACTOR POWER e
INLET WATER TEMPERATURE-"=] 7 —
——
o A
0 05 1.0 15 2.0 25 3.0 3.5
TIME (sec)

40

30

20

REACTOR POWER {Mw)

Fig. 4.5.26. Reactor Power and Fuel Temperature as a Function of Time After the Initial Reduction of the Inlet

Water Temperature from 145.8 to 32°F. Reactor power initially at 0.5 Mw.

UNCLASSIFIED

2-01- 060 -54
500 —— —
S
// <
// //
// s /
400 - it 7
1 reactor ' J#
POWER
EXIT MEAN d )‘-’/ 7
FUEL TEMPERATUREv/ _— s
& 300 — |- e — >
& / I P
5 BOILING POINT —__ y .
« S\ v
! — N - RS e e (e
o . S 7
E / //- ‘ ,//» //
5 / o e -
— 200 T //
—— / ,//
I
- 4 - P ‘-——-——-_-_-
FUEL EXIT — ey .
WATER TEMPERATURE S~ EXIT SECTION MEAN s /
I WATER TEMPERATURE Sy
100 » -] ced - .
S
/ ! A
L INLET WATER | , e
TEMPERATURE ~— | — e : :
o ,
0 0.5 L0 5 2.0 2.5 3.0 35
TIME {sec)

40

30

20

REACTOR POWER {Mw)

Fig. 4.5.27. Reactor Power and Fuel Temperature as a Function of Time After the Initial Reduction of the Inlet

Water Temperature from 145.8 to 32°F. Reactor power initially at 5 Mw.

209
ANP PROJECT PROGRESS REPORT

UNCLASSIFIED
2-01-060-55

300 T ——
| EXIT MEAN FUEL SRR IR _—
; TEMPERATURE\\ R e s
250 : Lol s 4
BOILING POINT ~—_ L ——/7"""_# 4
e
! o S R
; | g ) /'/’ A
= ~—— EXIT SECTION s s ~
e 200 MEAN WATER /'/ S 7 3z
S S e - o /
l TEMPERATURE SN S s S x
2 o 3
< &
& . - x
s S BRI
TR N S {1, 4
\ — ~ = Wi
/, / e y ’, ) m
\/REACTOR POWER o
_ I ] | \\
\ s
// S
100 — Y
.///
50 RA PR I
0 0.5 1.0 15 2.0 35
TIME (sec)

Fig. 4.5.28. Reactor Power ond Exit Section Mean Fuel and Mean Water Temperoture os a Function of Time

After Loss of Coolant Woter Flow.

3000-cps differential voltage. This voltage was
applied to a cathode-ray oscilloscope with a
200-msec sweep period initiated by the same change
in sigma bus voltage as that which initiated the
scram. The plot of position vs time that resulted
was calibrated in tenths of an inch between the
limits of travel of the control plate. The plate was
considered to have started and stopped moving when
the slope of the curve left and retumed to zero.

The delay time was measured by using the change
of sigma bus potential that initiated the scram to

simultaneously trigger the sweep of the oscilloscope.

The length of sweep from the point it began to the
first discontinuity was taken as an indication of the
*‘dead’’ time between initiation of a scram signal
and the first perceptible motion of the control plate.

SHIELD DESIGNS

The design criteria for the beam shield which will
be used initially with the TSR-ll were that (1) on a

210

Reactor initially at 5 Mw.

rem basis and with an rbe of 10 for fast neutrons,
there was to be an equal number of gamma rays

and fast neutrons at the reactor shield surface, and
(2) the background dose at the crew compartment
from the beam shield 64 ft away was to be a factor
of 50 to 100 less than that produced from a beam at
right angles to the reactor-crew compartment axis.
With these criteria and a gamma-ray shield of 50-50
volume per cent lead and water, the following shield
dimensions were obtained.'® The lead-and-water
gamma-ray shield should be 43 ¢m thick with an
inner radius of 48 cm. The water neutron shield
should be 81 cm thick, which gives an outer shield
radius of 172 cm, Thus, the shield will contain
33,500 Ib of lead and 42,900 |b of water. The
possibility of using 5/us-in. raschig rings with a
’/a-in.-dicu central opening in the lead-and-water

lsThis work was performed by Lloyd Byrnes of GE-
ANP, Cincinnati, Ohio.

gamma-ray shield is being investigated. The
beam holes will be stepped halfway through the
shield, with the outer cylinder 15 in. in diameter
and the inner cylinder 10 in. in diameter., The
Laboratory is proceeding with the design of this
shield with the additional criterion that it must be
self-supporting when it is placed on the ground.

A second high-performance shield for the TSR-I
is being designed by the Y-12 Engineering Depart-
ment according to specifications set by Pratt &
Whitney Aircraft. The shield will utilize a de-
pleted uranium gamma-ray shadow shield and a
lithium hydride neutron shield. [t will be fabricated
in the Y-12 shops and is scheduled to be completed
by May 1, 1959. The estimated cost of the shield is
$330,000.

INVESTIGATION OF STRESSES IN THE TOWER
STRUCTURE FROM WATER AND
ELECTRICAL LINES

An investigation has been made !¢ to determine
whether water hoses and electrical cables can be
suspended from the reactor support cables and
tower legs without exceeding the allowable stresses
in the tower structure.

The investigation was carried out for the following
conditions:

1. weight of reactor = 55 tons,

2. weight of crew compartment = 30 tons,

3. 80-mph wind (20 psf),

4, two water hoses, each 170 ft long and weighing
25 Ib/ft,

5. four electrical cables, each 170 ft long and
weighing 2.75 |b/ft.

The reactor and crew compartment were considered

to be in the plane of normal operation, that is, in the

16, A, McCarthy, Report of Investigation for the
TSF-1I, McPherson Co., July 1958,

SO PERIOD ENDING SEPTEMBER 30, 1958

vertical plane bisecting the longitudinal axis of the
Tower Shielding Facility (east and west) and ex-
tending 200 ft from the ground level.

The water hoses and electrical cables were as-
sumed to be symmetrically positioned from tower
legs | and Il. One end of each hose and each cable
was considered to be attached at a distance of
123 ft 6 in. above the base of the leg and the other
end was considered to be attached to the reactor
support cables at a point 33 ft from the reactor.

The calculations were compared to a similar cal-
culation made in connection with the original de-
sign of Knappen-Tippetts-Abbott-McCarthy (KTAM). 7
The calculation by KTAM did not take into account
items 4 and 5 above and showed the allowable
stresses in the tower legs to be:

1. as given by KTAM, 19.90 ksi, and

2. allowed by AISC code, 19.55 ksi.

The stresses in the tower legs determined in this

calculation are:

1. before addition of water hoses and electrical
cables, 18,072 ksi, and

2. after addition of water hoses and electrical
cables, 19.545 ksi.

In the calculation it was assumed that the electrical

cables and water lines were rigidly fastened to the

tower legs. In the present design all four electrical

cables will be suspended from Leg | and one water

hose will be suspended from Leg | and the other

from Leg Il. However, the suspension arrangement,

which is shown in Fig. 4.5.29, eliminates the greater

portion of the horizontal load in the tower legs by

shifting it to the deadmen which now anchor the

tower legs.

”Loadz'ng Criteria and Analysis for Tower Shielding
Facility, Knappen-Tippetts-AbbotteMcCarthy Co,, New
York, 1953,

UNCLASSIFIED
2-01-060-56

-s— TOWER LEG

MESSENGER CABLE
FOR WATER HOSE

GUY ANCHOR

]
' 1

HOSE TO REACTO)/7/

HOSE FROM STANDPIPE)

Fig. 4.5.29. Tower Suspension Arrangement of Reactor Cooling Water Lines.

212

VONO U AN~

A A

@EMMO =

CMEEPIMZAOMPETPREI>COCCOMECODROPMAM

. Allen

. Billington

. Blankenship

. Blizard

. Boudreau

. Boyd

. Breeding

. Briggs

. Bruce

. Callihan

. Center (K-25)

. Charpie

. Clifford
Coobs

. Cottrell

. Culler

. Cuneo
DeVan

. Doney

. Douglas

. Emlet (K-25)

. Fraas

Frye

. Furgerson

Gray

. Greenstreet

. Grimes

. Grindell

Guth

. S. Harrill

. Hikido

R. Hill

E. Hoffman

W. Hoffman

. Hollaender

. H. Jordan

. W. Keilholtz

. L. Keller

OAr—“—AToorzTaroImMmrprmoxTtmemmMUmMw =X

. J. Keyes

. S. Livingston
. N. Lyon

INTERNAL DISTRIBUTION

42.
43.
44,
45.
46.
47.
48.
49,
50.
51.
52.
53.
54.
55.
56.
57.
58.
59.
60.
61.
62.
63.
64.
65.
66.
67.
68.
69.
70.
/1.
72,
73.
74,
/5.
76.
77.
/78-80.

81-87.
88.
89-91.

ORNL -2599
C-85 —~ Reactors-Aircraft Nuclear
Propulsion Systems

M-3679 (22nd ed.)

. G. MacPherson

. C. Maienschein

. D. Manly

. R. Mann

. J. Miiler

. Z. Morgan

. J. Murphy

P. Murray (Y-12)

L. Nelson

. J. Nessle

. G. Overholser

. Patriarca

. K. Penny

. M. Perry

. M. Reyling

. W. Savage

. W. Savolainen

. D. Schultheiss

L. Scott

. D. Shipley

. Simon

. J. Skinner

. H. Snell

. Storto

. Susano

. Swartout

. Trauger

. Trubey

. Watson

. Weinberg
White

. Wigner {consultant)

. Williams

. Wilson

. Winters

W. Zobel

ORNL - Y-12 Technical Library,
Document Reference Section

Laboratory Records Department

Laboratory Records, ORNL R.C,

Central Research Library

TrOIXT-MA>METNT

OCOME> 000N M>TPM-TP> T VP>

mONoNDzzxmPO

213

EXTERNAL DISTRIBUTION

92-95. Air Force Ballistic Missile Division
96. AFPR, Boeing, Seattle
97. AFPR, Boeing, Wichita
98. AFPR, Douglas, Long Beach
99-101. AFPR, Douglas, Santa Monica
102-103. AFPR, Lockheed, Marietta
104. AFPR, North American, Canoga Park
105. AFPR, North American, Downey
106. AFPR, North American, Los Angeles
107-108. Air Force Special Weapons Center
109-110. Air Research and Development Command (RDZN)
111, Air Technical Intelligence Center
112, Air University Library
113-115. ANP Project Office, Convair, Fort Worth
116. Albuquerque Operations Office
117. Argonne National Laboratory
118. Armed Forces Special Weapons Project, Sandia
119. Armed Forces Special Weapons Project, Washington
120-121. Army Ballistic Missile Agency
122, Army Rocket and Guided Missile Agency
123. Assistant Secretary of Defense, R&D (WSEG)
124. Assistant Secretary of the Air Force, R&D
125-130. Atomic Energy Commission, Washington
131-133. Bettis Plant (WAPD)
134. Brookhaven National Laboratory
135. Bureau of Aeronautics
136. Bureau of Aeronautics General Representative
137. BAR, Aerojet-General, Azusa
138. BAR, Chance Vought, Dallas
139. BAR, Convair, San Diego
140. BAR, Goodyear Aircraft, Akron
141. BAR, Grumman Aircraft, Bethpage
142. BAR, Martin, Baltimore
143. Bureau of Ships
144. Bureau of Yards and Docks
145-146. Chicago Operations Office
147. Chicago Patent Group
148. Curtiss-Wright, Quehanna
149, Director of Naval Intelligence
150. duPont Company, Aiken
151-158. General Electric Company (ANPD)
159-160. General Electric Company, Richland
161. Hartford Aircraft Reactors Area Office
162. {daho Test Division (LARQO)
163. Jet Propulsion Laboratory
164. Knolls Atomic Power Laboratory
165. Lockland Aircraft Reactors Operations Office
166-167. Los Alamos Scientific Laboratory
168. Marquardt Aircraft Company
169. National Advisory Committee for Aeronautics, Cleveland

170.
171.
172.
173.
174.
175.
176.
177.
178.
179.
180-181.
182-185.
186.
187.
188-189.
190.
191-192.
193.
194-195.
196-208.
209-233.
234,

‘

National Advisory Committee for Aeronautics, Washington

Naval Air Development Center

Naval Air Material Center

Naval Air Turbine Test Station

Naval Research Laboratory

New York Operations Office

Oak Ridge Operations Office

Office of Naval Research

Office of the Chief of Naval Operations

Patent Branch, Washington

Phillips Petroleum Company (NRTS)

Pratt and Whitney Aircraft Division

Sandia Corporation

Sandia Corporation, Livermore

School of Aviation Medicine

USAF Headquarters

USAF Project RAND

U. S. Naval Radiological Defense Laboratory

University of California Radiation Laboratory, Livermore

Wright Air Development Center

Technical Information Service Extension

Division of Research and Development, Atomic Energy Commission,
QOak Ridge Operations

216

Reports previously issued in this series are as follows:

ORNL-528
ORNL-629
ORNL-768
ORNL-858
ORNL-919
ANP-60
ANP-65
ORNL-1154
ORNL-1170
ORNL-1227
ORNL-1294
ORNL-1375
ORNL-1439
ORNL-1515
ORNL-1556
ORNL-1609
ORNL-1649
ORNL-1692
ORNL-1729
ORNL-1771
ORNL-1816
ORNL-1864
ORNL-1896
ORNL-1947
ORNL-2012
ORNL-2061
ORNL-2106
ORNL-2157
ORNL-2221
ORNL-2274
ORNL-2340
ORNL-2387
ORNL-2440
ORNL-2517

Period Ending November 30, 1949
Period Ending February 28, 1950
Period Ending May 31, 1950
Period Ending August 31, 1950
Period Ending December 10, 1950
Period Ending March 10, 1951
Period Ending June 10, 1951
Period Ending September 10, 1951
Period Ending December 10, 1951
Period Ending March 10, 1952
Period Ending June 10, 1952
Period Ending September 10, 1952
Period Ending December 10, 1952
Period Ending March 10, 1953
Period Ending June 10, 1953
Period Ending September 10, 1953
Period Ending December 10, 1953
Period Ending March 10, 1954
Period Ending June 10, 1954
Period Ending September 10, 1954
Period Ending December 10, 1954
Period Ending March 10, 1955
Period Ending June 10, 1955
Period Ending September 10, 1955
Period Ending December 10, 1955
Period Ending March 10, 1956
Period Ending June 10, 1956
Period Ending September 10, 1956
Period Ending December 31, 1956
Period Ending March 31, 1957
Period Ending June 30, 1957
Period Ending September 30, 1957
Period Ending December 31, 1957
Period Ending March 31, 1958