Surface residual stresses due to turning operations in AISI 304 type stainless steel were studied as a function of machining speed, feed rate, depth of cut, and tool geometry and coating. Residual stress tensors were determined using X-ray diffraction technique. The effects of turning conditions and tool on the residual stresses were discussed in terms of mechanically and thermally induced non-homogeneous plastic deformation of the surface layers of the workpiece.

1. WEAR
ELSEVIER Wear 194 (1996) 168-173
Surface residual stresses in machined austenitic stainless steel
D.Y. Jang a,*, T.R. Watkins b, K.J. Kozaczek b, CR. Hubbard b, O.B. Cavin b
’Mecharzicul and Aemspme Engineering Depurtmerrt. University o~Missorrri-Colronbicc, Cofrmbiu. MO 652 ll. USA
’High Ternperuture Muterial Laboratory, Ouk Ridge Nationul l*rborcrtory. Oak Ridge. TN 37831-6064. USA
Keceived 3 February 1995; accepted 19October1995
Abstract
Surface residual stresses due to turning operations in AISI 304 type stainless steel were studied as a function of machining speed, feed rate,
depth of cut, and tool geometry and coating. Residual stress tensors were determined using X-ray diffraction technique. The effects of turning
conditions and tool on the residual stresses were discussed in terms of mechanically and thermally induced non-homogeneous plastic
deformation of the surface layers of the workpiece.
Keywords: Stainlesssteel;Machining;Residualstress;X-raydiffraction;Surface;Machinability
1. Introduction
Stainless stee1 is one of the principal materials being util-
ized in manufacturing critical parts for modern power and
chemical plants because of the combination of appropriate
mechanica1 properties and high resistance to corrosion. How-
ever, these useful properties can be reduced during fabrication
such as machining. Stainless steel has been characterized as
gummy during machining, showing a tendency to produce
long and stringy chips, which seize or form a built-up edge
on the cutting tool. This degrades the surface finish of a
machined product and results in reduced tooi life. In order to
utilize the useful properties of stainless steel by improving
fabrication quality, it is necessary to study surface quality of
machined stainless steels.
Surface integrity is a measure of the quality of a machined
surface and is interpreted as elements which describe the
actual structure of both surface and subsurface. Surface integ-
rity is generally defined by its mechanical, metallurgical,
chemical and topological states of surface properties such as
surface roughness, hardness variation, structural changes and
residual stress, etc. The demand for high quality production
focuses attention on the surface condition of products. espe-
cially residual stresses on the machined surface because of
their effects on component performance, longevity and reli-
ability. Severe failures produced by fatigue, creep and stress
corrosion cracking invariably start at the surface of compo-
nents and depend to a great extent on the quality of surface.
* Corresponding author.
0043-1648/96/$15.00 0 1996 ElsevierSclence S.A. All rigbts reserved
SSD10043-1648(95)06838-4
It is therefore necessary to quantify surface integrity proper-
ties as a function of machining conditions. This paper deals
with the residual stress aspect of surface integrity in terms of
machining speed, feed rate, depth of cut, and tooi geometry
and coating.
1.1. Residual stresses in machining
Residual stress in a machined surface is one of the crucial
factors in determining surface quality. Residual stress is
defined as the stress that exists in an elastic body after al1 the
external loads are removed. Machining generally involves a
large amount of plastic deformation with extremely high
strain and strain rate. It imparts residual stresses in the surface
layer of a workpiece, which are undesirable but unavoidable
in the cutting process. The cause and mechanism of residual
stress in machining have been studied by several researchers
both analytically and experimentally. Henriksen [ l] pre-
sented empirical data on the residual stress distribution in
metal cutting and suggested that residual stresses are gener-
ated by the mechanica1 rather than the thermal effects of the
cutting process when light cuts were taken on low carbon
steels at the low and medium cutting speeds. Liu and Barash
[ 21 found experimentally that the length of the shear plane
in chip formation correlated with subsurface deformation and
residual stress formation. Hence, the complicated and unclear
chip-tool interface boundary conditions were eliminated for
the finite element analysis. Leskovar and Peklenik [ 31 inves-
tigated the iníluence of cutting process parameters on surface
and subsurface conditions in turning, and showed that tensile

2. D. Y. Jang et ~1./ Wear I94 (1996) X68-1 73 169
Table 1
Cutting conditions for residual stress measurement
Condition no. Speed ( rpm ) Depth of cut (mm) Feed rate (mm rev- ’) Radius (mm) Coated
1 310 0.762 0.084 0.397 NO
2 310 0.762 0.104 0.397 NO
3 310 0.762 0.132 0.397 NO
4 310 0.508 0.104 0.397 NO
5 310 1.016 0.104 0.397 NO
6 180 0.762 0.104 0.397 NO
7 580 0.762 0.104 0.397 NO
8 310 0.762 0.084 0 NO
9 310 0.762 0.104 0 NO
10 310 0.762 0.132 0 NO
11 180 0.762 0.104 0.397 Yes
12 310 0.762 0.104 0.397 Yes
13 580 0.762 0.104 0.397 Yes
stresses are dominant after turning and that higher speed tends
to produce higher residual stresses. Matsumoto et al. [4]
discussed the effect of hardness of the workpiece material on
residual stresses in orthogonal machining by measuring the
residual stresses in the machined surfaces of workpieces with
different hardness. Jang and Seireg [51 presented a compre-
hensive computer-based simulation model to predict the
residual stress on the machined surface using a model con-
sidering cutting mechanics and thermal effects between tool
and workpiece. An excellent review of previous works on the
residual stresses has been given by Brinksmeier et al. [ 61.
According to the published literature, it has been generally
accepted that for a given material, the nature of the residual
stress distribution in the surface region of a machined work-
piece depends on the cutting speed, feed rate, depth of cut,
tool geometry, and whether or not a lubricant is used. The
residual stresses are generally produced by inhomogeneous
plastic deformation induced by the mechanica1 and thermal
events associated with the process of chip formation and
interaction between the tool nose region and the freshly
machined workpiece surface [ 5-71.
Residual stress formation during machining of stainless
steel is not wel1 understood. Austenitic 304-type stainless
steel is one of the most troublesome materials to machine
because of its tendenties to work-harden, high temperature
and adhesion [ 81. There have been only a few attempts to
calculate and measure the residual stresses on machined stain-
Cutting direction
Axial q,
Fig. 1. Cutting direction and stress tensor coordinate relative to the
workpiece.
less steel components [ 9-111. However, in order to analyze
the mechanism of residual stress formation in the machined
stainless steel correctly and have the practica1 machinability
data of a turning process, it is necessary to determine the
influence of cutting conditions on the residual stress. The
current study investigated effects of cutting conditions on
cutting speed, feed rate and depth of cut, including tool geom-
etry and tool coating intensively. X-ray diffraction from
Cr KP source, which has a relatively lower penetration depth,
was used to measure residual stresses in the layer of compo-
nents close to the machined surface.
2. Experiments
AIS1 304 austenite stainless steel bars of 38 cm in length
and 4.9 cm in diameter were supplied by Oak Ridge National
Laboratory. The chemical composition of this alloy in weight
percent was reported as 69.493 Fe, 0.059 C, 1.26 Mn, 0.44
Si, 18.6 Cr, 9.5 Ni, 0.033 P, 0.015 S, 0.35 Mo and 0.25 Cu.
Each bar was premachined and then annealed at 1100 “C for
3 h and slowly cooled for 36 h to remove any pre-existing
residual stresses and to avoid the formation of residual stress
due to rapid cooling, respectively.
Machining was performed on a 7.5 HP lathe with maxi-
mum speed of about 1500 rpm without cutting fluid. The
cutting conditions for residual stress measurements are given
in Table 1 and the longitudinal cutting distance for each cut-
ting condition was about 1.6 cm. Fig. 1 shows the cutting
edge and cutting direction relative to the workpiece shape,
including stress tensor coordinate system. A mechanica1 chip
breaker and a specially designed tool dynamometer using a
tool holder and a three component force transducer were used.
Carbide inserts of P30 ISO class were used. Sharp tool inserts
and inserts with 0.397 mm edge radius were used. The shape
of inserts was 80”-diamond with 11” relief angle. A brand
new insert was used for machining each part of workpiece in
order to insure the same initial tool conditions. The cutting
operations were monitored to avoid chatter using a Tektronix

3. D.Y.Jangetal./Wear194(1996)168-173
oscilloscope to display cutting force signals from the tool
dynamometer. A computer with an IEEE 488 standard A/D
board was used for data acquisition.
The machined bar was cut using a vertical band-saw with
14 teeth/in carbon hard-back blade and divided into sectional
specimens for the residual stress measurement. The length of
a sectional specimen, which includes three different cutting
conditions, was 5.1 cm long. In the sawing operation, the
cutting speed was 60.96 cm min- ’and a sulphur-base cutting
oil was used as a coolant to avoid changing the residual
stresses generated by the machining process.
The stress measurements for the triaxial with shear analysis
were performed for the samples containing cutting conditions
from 1 to 7 at $= O”, 4=45”, and $= 90”. For the biaxial
stress analysis, stress measurements on the samples of cutting
conditions 8-13 were conducted at 4 = 0” and 4 = 90”. Dif-
fraction data for $= 0” and $ = 90” from the triaxial with
shear stress analysis can be used for the biaxial stress analysis.
The (3 11) peaks were scanned using a 0.05” 2 0 increment
from 145” to 153” 20 (0 is the Bragg angle) [ 131 at each
tilt angle, iy ( - 55”, -42”, - 28.2”, O”,28.2”, 42” and 55”).
Count times of 15-20 s were used for accurate peak shapes
and good counting statistics.
The residual stresses were measured in the High Temper-
ature Materials Laboratory at Oak Ridge National Labora-
tory. A four-axis ($, x, 0, 20) goniometer with a 360 mm
radius was employed for the stress measurements using the
psi-goniometer geometry (i.e. the psi, q, or tilt axis was ,Y,
not 0) [ 121. Data acquisition and peak profile analysis were
controlled by the DMS software, which was installed in a
Micro-VAX computer. The cylindrical samples were
mounted in the cradle type’sample holder employing elastic
bands to hold the sample in place. Sample alignment was
accomplished using a dia1 gauge probe (accurate to
+ 10 p_m) and a telescope. Fig. 2 shows the setup of a sample
in the sample holder. The goniometer alignment was ensured
by examining a plate containing Si powder in epoxy using q
tilting in the range of - 55”< 91 +55”. The maximum
observed peak shift for the (533) reflection of Si at 136.8”
20 was less than 0.03” 20.
The measured strain distributions, i.e. the plot of d spacing,
(d,, -d,,) ld, VS sin’ q along the cutting direction, were
linearly fitted using the generalized least squares fitting
method [ 12,151. The terms dv+ and do are the measured and
the unstressed {311) interplanar spacings, respectively. X-
ray elastic constants for the austenitic phase for the stress
calculation from the measured lattice strains of (3 11) planes
of 304 stainless steel were S, = - 1.87 X 10-h MPaä’ and
S2/2 = 7.6 X 10e6 MPa-’ (E= 172 GPa, v = 0.29) [ 141
and do was 1.0824 A. The average strain-free interplanar
spacing, do, was determined from the annealed filings at
1010 “C for 3 min in vacuum followed by rapid cooling in
air of the stainless steel stock [ 161. The stresses were cal-
culated using the Dolle-Hauk method [ 121.
3. Results and discussion
In the stress measurements, a high-angle diffraction peak
with sufficient diffracted intensity is needed for accuracy
and precision. Therefore, Cr KB radiation (wavelength A=
Table 2 from the triaxial with shear stress analyses showed
negligible shear and radial stresses ( ur3, ,= 1,2,3= 0). Hoop
stresses ( gZ2) in the cutting direction were the largest resid-
ual stresses and tensile up to 600 MPa. The axial components
in the feed direction ( CJ,,) were predominantly compressive.
Table 2
Residual stress components from measurements of triaxial stress with shear analysis
2.08487 A) was used from a rotating anode generator oper-
ating at 9 kW (30 kV and 300 mA). The effective depth of
X-ray penetration was about 5 km for 50% of the total
integrated intensity [ 13,141. A double pinhole collimator
with a 1.5 mm opening was used with 3 and 0.3 mm receiving
slits. Smal1 X-ray beam size and low divergente, combined
with the relatively large radius of curvature of the sample,
minimized the experimental errors due to the uncertainty
regarding the true 0 angle on a curved specimen.
1 -67+81 330+91 25*58 -29k43 - 12_t26 -22+21
2 -246k 177 271 k238 -132+136 681102 -5k73 14+42
3 -79f93 4811114 24+69 -21 k53 -0.4f33 - 15+-23
4 29191 552*110 29+67 -89+46 5432 -20*23
5 -941101 462 k 95 78+69 -72+81 -2+-24 -30+27
6 - 139k87 255 15108 -14+64 -87k50 -4+31 5*21
7 -53&84 629 I99 14561 -3149 -2&28 - 17rt21

4. D.Y. Jang et ~1./ Weur 194 (1996) 168-173 171
Table 3
Residual stress components from measurements of biaxial stress analysis
Condition no.
1
2
3
4
5
6
1
8
9
10
11
12
13
-92+34 234 rt 68 7*50
- 114k68 479f 135 31*118
- 102537 346 f 74 35+59
1137 507 * 74 -81*54
- 172+45 317*90 -39*91
- 1251135 346 f 69 - 125*54
-68f35 517*69 46+55
-365*66 -55f 132 -27f90
-389f84 242 k 168 14* 138
186*63 749 f 124 -452t 100
-218f58 273f 116 19+96
- 172*58 -372* 115 -105*114
-59+80 492+ 160 -108f122
Al1 units are MPa.
I
-200
I I . I , I , , . , I , , . . I , , , , I , *, ,
0.06 0.09 0.1 0.11 0.12 0.13 0.14
Feed Rate (mmlrev.)
Fig. 3. Residual stress variation w.r.t. feed rate.
As the tool moves along the hoop direction of a cylindrical
workpiece in turning, three components of cutting forces are
generated in the cutting zone between tool and workpiece.
The cutting force in the hoop direction is usually the largest
and has the most significant effect on the residual stresses in
a turned surface. Therefore, it is plausible to expect the largest
stress ( azz) in the cutting direction. The feeding component
( u1 1) shows tension or compression depending on the cutting
conditions [ 9-111. Because measurements were conducted
on the free surface, residual stress state is plane stress with
the principal axes directions close to the hoop and axial direc-
tions, and ui3 (i = 1,2,3) can be neglected in the turned com-
ponents [9,10]. Tabie 3 summarizes residual stresses from
the biaxial stress analysis for 13 different cutting conditions
to consider plane stress state in turned components consisting
of hoop and axial stresses.
Figs. 3-7 show the effects of cutting conditions including
tool edge radius and tool coating on the hoop ((T~*), axial
(<Ti1) and shear stresses ( g12). The hoop stresses (c2?) were
shown to peak at a feed rate of 0.104 mm rev ~ ’(see Fig. 3 )
Also, the hoop stresses were shown to increase and decrease
a
400 -
z
200 -
s
?!
z o-
-200
I
$-_--_-_l 47 -- - __
---*--4
l II 8, I, 1, c 1 * * c * 1, t 8 < 1 j,. 1 - -, 8 1, 5 8 I /
0.4 0.5 0.6 0.7 0.6 0.6 1 1.1
Depth of Cut (mm)
Fig. 4. Residual stress variation w.r.t. depth of cut.
a
400 -
z
t
200 -
E
G
o-
-200 -
l..’...-‘.-
-.4....
f’iI __ _____- _ _ ____- - 4
100 200 300 400 500 600
Speed (RPM)
Fig. 5. Residual stress variation w.r.t. speed
2 600-
z 600 - __--
3
!? 400
200 ;
_---
__--
__._.------
G
_____----
____--
a 300
1 4b __--
_____.----
x 0
-1;
JI
100 -
0 ~11~',~~~',,~~',~~~'~~~~
100 200 300 400 600 600
Speed (RPM)
Fig. 6. Tool coating effect on hoop stress ( mzz)
with speed and depth of cut, respectively. However, u,, and
(T,~were effectively constant over the ranges of cutting speed,
depth of cut and feed rate examined (see Figs. 3-5). Fig. 6
showed negligible effect of tool coating on hoop residual
stress (<Tak), but Fig. 7 showed that tool sharpness has a large
effect, which was stronger than the effects of the other cutting
conditions.

5. 172 D.Y. JUIIRet trl. / Weírr 194 (1996) 16X-173
G 600
g 600
fL 400
ùi
0
300
g 200
100
1’
0”““““““““““““““’
0.06 0.09 0.1 0.11 '0.12 0.13 0.14
Feed Rate (mmlrev.)
Fig. 7. Tool edge radius effect on hoop stress (CT?~).
A sharp tool introduced the greatest tensile stress of rrZZat
high feed rate as wel1 as the greatest compressive residual
stress of (T,, However, at feed rates of 0.084 and 0.104
mm revv’ , rounded tools of 0.397 mm edge radius generated
higher tensile hoop stress and lower compressive g,, This
effect can be explained due to higher contact friction between
tool and workpiece.
Tensile residual stresses in the turned stainless steel com-
ponents arc typically caused by plastic deformation in front
of the advancing tool and the localized thermal effect. The
large contact area between tool and workpiece formed by a
rounded edge of a tool is expected to generate more frictional
heat in the cutting zone. Because of low thermal conductivity
of stainless steel, thermal effect in the cutting zone is greater
than the other regular steels. Therefore, stainless steel
machining with a round edge tool generates higher tensile
stress on the machined surface [9-111. Although the
mechanica1 effect generally creates compressive residual
stresses on the machined surface, strong work-hardening and
a considerable increase in the microstructural defects close to
the machined surface, which causes greater elastic relaxation
upon unloading of the thin surface layer as compared with
the underlying workpiece, result in tensile residual stress
[ 10,l 1] From significant computer simulations of residual
stress generation in machined stainless steel components
[ 9,101, a large compressive plastic zone was formed in front
of the advancing tool. Thc wake of plastically compressed
material left in the surface layer receives plastic deformation
constrained by the undeformed bulk of the workpiece, which
causes tensile residual stresses on the machined surfacc.
In conventional turning, the heat absorbed in the workpiece
is relatively smal1 ( 10-15 % of the total energy in the primary
shear zone) 15, 101, and the penetration depth of the heat is
much smaller compared to grinding. However, since thermal
conductivity of 304 stainless steel is low compared to other
common steels, the absorbed heat wil1 be retained longer in
the smal1 cutting zone. During dry cutting, the material ele-
ment near the machined surface is heated and elongated more
than the bulkmaterial. After cutting, the surfaceelement stays
hotter than the bulk of the workpiece, because the majority
of the cooling occurs from inside and the air is a poor con-
ductor of heat. Therefore, the surface element expands and
experiences a compressive stress from the bulk material
below. This can lead to compressive plastic deformation of
the surface layer facilitated by the low yield strength at high
temperature and can result in tensile residual stresses.
4. Conclusions
Residual stresses were measured experimentally as a func-
tion of cutting speed, depth of cut and feed rate in order to
characterize surface integrity in the machined 304 stainless
steel components. Triaxial with shear stress analysis showed
that hoop stresses (crZ2) in the cutting direction were the
largest and tensile up to 600 MPa, with the radial components
( (T,~,,= ,,2,3)close to zero because measurements were made
on free surface. The axial components in the feed direction
( u, , ) were predominantly compressive. Therefore, the
machined surfaces of the cylinders were in a state of plane
stress with the principle axes directions close to the hoop and
axial dircctions [ 9.101. The axial components of stress in the
feed direction were predominant compressive and insensitive
to the aforementioned cutting conditions. The shear stresses
also behaved in this manner. The hoop stresses were predom-
inantly tensile, increasing with cutting speed, decreasing with
depth of cut, and peaking with feed rate. TOOI sharpness was
found to have the largest effect on residual stress. The data
suggest that a sharp tool, a low feed rate (0.08 mm rev- ‘),
a deep depth of cut ( 1 mm), and a slow cutting speed
( < 180 rpm) wil1 result in a minima1 amount of residual
stress.
Acknowledgements
The authors would like to thank reviewers for their sincere
and valuable comments on the paper. Author (D.Y. Jang)
would like to thank Oak Ridge Institute for Science and
Education (ORISE) for their support of travel expense under
the contract no. S-3506. This research was sponsored by the
U.S. Department of Energy, Assistant Secretary for Energy
Efficiency and Renewable Energy, Transportation Technol-
ogies, as part of the High Temperature Material Laboratory
User Program under contract DE-AC05840R2 1400, man-
aged by Martin Marietta Energy Systems, Inc. This research
was supported in part by an appointment to the Oak Ridge
National Laboratory Postdoctoral Research Associates Pro-
gram administered jointly by the Oak Ridge National Labo-
ratory and ORISE.
References
[ 11 E.K. Henriksen, Residual stresses in machined surfaces, J. Appl. Mech.
A.W.5 73 ( 195 1) 69-76.

6. D.Y. Jatrg et al. / Wear 194 (1996) 168-173 173
[2] C.R. Liu and M.M. Barash, The mechanical state of the sublayer of a
surface generated by chip removal process, J. Eng. Ind.. ASME, 98
(1979) 1192-1208.
[3] P. Leskovar and J. Peklenik, Influence affecting surface integrity in
the cutting process, Ann. CIRP (1981) 245-248.
[4] Y. Matsumoto, M. Barash, and C.R. Liu, Effect of hardness on the
surface integrity of AIS1 4340 steel, J. Eng. Ind., ASME. 108 (1986)
169-175.
[S] D.Y. Jang and A. Seireg, A model for predicting residual stresses in
metal cutting, Proc. Jup. Int. Tribolqy Conf: Nag~q~u, 1990, pp. 439-
444.
[6] E. Brinksmeier. J.J. Cammett, P. Leskovar, J. Peters and H.K.
Tonshoff, Residual stresses-measurements and causes in machining
processes, Ann. CIRP. 31 (1982) 491-510.
[7] H.K. Tonshoff and E. Brimksmeier, Determination of mechanica1 and
thermal influences on machined surfaces by microhardness and
residual stress analysis, Ann. CIRP, 29 ( 1980) 519-530.
[81 G. Boothroyed, FundrrrnentalsofMeralMuchinir~~andMachine Tools,
McGraw-Hill, Washington, DC, 1975.
[9] D.Y. Jang and L. Wang, Predicting stress distribution in workpiece
using F.E.M., Mo. Acad. Sci. Conf., 1992.
[ 101 J. Liou and D.Y. Jang, Study of residual stress distribution in the
machined stainless steel, J. Marh. Model. Sci. Comput., in press.
[ I 11C. Wiesner. Residual stress after orthogonal machining of AIS1 304;
numerical calculation of the thermal component and comparison with
experimental result, Mefull. Trans. A.. 23-a ( 1992) 989-996.
[ 121 I.C. Noyan and J.B. Cohen, Residual Stress Determinarion by
DifJrucrion-Mensurernenl und Interpretution, Spinger, New York,
1987.
[ 131 Residual Stress Measurement by X-Ray Diflructiort, SAE J784a, SAE
Handbook, 197 1.
[ 141 H.M. Ledbetter, Monocrystal-polycrystal elastic constants of stainless
steel, Phys. Srur. Sol., 85 ( 1984) 89-96.
[ IS] T. Inmura, S. Weissman and J.J. Sladem, Jr. A study of age-hardening
of Al-3.85% Cu by the divergent X-ray beam method. Acta Cq~sf., 15
( 1962) 786-793.
[ 161 I.C. Noyan, Determination of the unstressed lattice parameter “a«”
for (triaxial) residual stress determination by X-rays,Adv, X-ruyAnal.,
28 (1985)281-288.
Biographies
Dong Young Jang: is Assistant Professor at Mechanica1 and
Aerospace Engineering Department of the University of Mis-
souri-Columbia. He won his BSc (1978) and MSc (1981)
from Seoul National University and his Ph.D. from the Uni-
versity of Florida ( 1990). His research expertise is in design,
manufacturing and tribology, and he has published over 30
technical papers. He is amember of ASME, STLE and KSEA.
Thomas R. Watkins: is a Research Staff Member at the High
Temperature Materials Laboratory in the Metals and Ceram-
ics Division of Oak Ridge National Laboratory. His BSc
(Ceramic Engineering) is from Alfred University (1985),
and his MSc and Ph.D. (Ceramic Science) are from Penn-
sylvania State University ( 1988 and 1992). Research inter-
ests include X-ray diffraction, residual stresses and
mechanica1 properties of materials. He is a member of the
American Ceramic Society.
Krzysztof J. Kozaczek: is a Development Staff Member at
the High Temperature Materials Laboratory in the Metals and
Ceramics Division of Oak Ridge National Laboratory. His
MSc (Mechanica1 Engineering) is from the Technical Uni-
versity of Warsaw (1978), and his Ph.D. is from the Penn-
sylvania State University ( 1991). His research interests
include nondestructive characterization of materials (X-ray,
ultrasonic, magnetic), experimental stress analysis (mostly
X-ray diffraction), crystallographic texture analysis (X-ray
diffraction), and modeling of elasticity and plasticity of poly-
crystals. He has published over 40 technical papers and is a
member of the ASME and the Society of Experimental
Mechanics.
Camden R. Hubbard: is leader of the Diffraction and Physical
Properties Group at the High Temperature Materials Labo-
ratory in the Metals and Ceramics Division of Oak Ridge
National Laboratory. He won his BSc (Chemistry) from the
University of California at Berkeley ( 1966), and his Ph.D.
(Physical Chemistry/Crystallography) from Iowa State Uni-
versity ( 197 1). His research interests include X-ray and neu-
tron diffraction and thermophysical properties of materials.
He is a member of the American Ceramic Society, American
Chemical Society, American Crystallographic Association
(member, former officer, Materials Science Special Interest
Group) , International Centre for Diffraction Data (member,
former Vice Chairman and Board of Directers member),
Materials Research Society, Neutron Scattering Society of
America.
0. Burl Cavin: is a retired research member at the High
Temperature Materials and Ceramics Division of Oak Ridge
National Laboratory. His BSc (Mathematics, Physics and
Chemistry) is from Lincoln Memorial University (1953),
and his Msc is from the University of Tennessee (1960). His
research interests include phase identification at room and
elevated temperatures, polycrystalline and single crystal stud-
ies of texture and orientation.