STRESS CORROSION TESTING
A symposium
presented at the
Sixty-ninth Annual Meeting
AMERICAN SOCIETY FOR
TESTING AND MATERIALS
Atlantic City, N. 1, 26 June-1 July, 1966
ASTM SPECIAL TECHNICAL PUBLICATION NO. 425
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Foreword
The Symposium on Stress Corrosion Testing was presented in four
sessions during the 69th Annual Meeting of the Society, in Atlantic City,
N. J., 26 June-1 July, 1966. The symposium was sponsored by Committee G-l on Corrosion of Metals. The symposium chairman was H.
Lee Craig, Jr., Reynolds Metals Co. Presiding at the four sessions were
M. A. Streicher, E. I. du Pont de Nemours & Co., Inc.; M. G. Fontana,
The Ohio State University; G. J. Danek, Jr., U. S. Naval Marine Engineering Laboratory; and W. F. Gerhold, National Bureau of Standards.
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Related
ASTM Publications
Stress-Strain-Time-Temperature Relationships in Materials, STP 325 (1962), $5.25
Stress-Corrosion Cracking of Titanium, STP 397 (1966),
$14.00
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Contents
Introduction
Stress Corrosion Testing Methods
Influence of Stress and Temperature on Short-Transverse Stress Corrosion Cracking of an Al-4.2Zn-2.5Mg Alloy—W. J. HELFRICH . .
A Comparison of Three Precracked Specimens for Evaluating the Susceptibility of High-Strength Steel to Stress Corrosion Cracking—
C. D. BEACHEM AND B. F. BROWN
Application of an Accelerated Stress Corrosion Test to Alloy Development—E. E. DENHARD, JR., AND R. R. GAUGH
Discussion
A Rapid Stress Corrosion Test for Aluminum-Magnesium Alloys—
H. L. CRAIG, JR., AND H. B. ROMANS
Discussion
Influence of Environment on Crack Propagation Characteristics of
High-Strength Aluminum Alloys—J. H. MULHERIN
Discussion
Stress Corrosion Cracking of High-Strength Bolting—C. S. LIN, J. J.
LAURILLIARD, AND A. C. HOOD
Discussion
Stress Corrosion of High-Strength Steel Alloys—Environmental
Factors—A. GALLACCIO AND M. A. PELENSKY
Stress Corrosion of Magnesium Alloys—Environmental Factors—
21
31
41
50
51
65
66
81
84
98
99
107
M. A. PELENSKY AND A. GALLACCIO
Discussion
Resistance of Ferritic Stainless Steels to Stress Corrosion Cracking—
A. P. BOND, J. D. MARSHALL, AND H. J. DUNDAS
Discussion
Some Techniques Used in the Study of Stress Corrosion Cracking—
H. L. LOGAN
Discussion
A Proposed Mechanism for the Stress Corrosion Fracture of a CopperBeryllium Alloy—W. D. SYLWESTROWICZ
A Quantitative Stress Corrosion Test for Al-Zn-Mg Alloy Plate—
Stress
1
3
115
116
125
127
142
145
H. ROSENTHAL AND H. R. PRITCHARD .
165
ROMANS
182
Corrosion Test Environments and Test Durations—H. B.
Reactions Contributing to the Formation of Susceptible Paths for
Stress Corrosion Cracking—D. A. VAUGHAN AND D. I. PHALEN .
209
Discussion
227
Critical Species in Stress Corrosion Phenomena—E. N. PUGH AND A.
R. c. WESTWOOD
228
Circulating Autoclave System for Stress Corrosion Cracking Studies—
R. W. STAEHLE
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Stress
Corrosion Cracking
Rates of a Nickel-Brass Alloy Under
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CONTENTS
Discussion
291
Reporting and Evaluating Stress Corrosion Data—D. O. SPROWLS .... 292
Evaluation of Various Techniques for Stress Corrosion Testing Welded
Aluminum Alloys—M. B. SHUMAKER, R. A. KELSEY, D. O.
SPROWLS, AND J. G. WILLIAMSON
Stress Corrosion Testing of 7079-T6 Aluminum Alloy in Various
Environments—B. W. LIFKA AND D. O. SPROWLS
Environmental Factors Affecting the Stress Corrosion Cracking Behavior of an Aluminum-Zinc-Magnesium Alloy—H. B. ROMANS
AND H. L. CRAIG, JR
342
363
Discussion
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STRESS CORROSION TESTING
Introduction
A recent survey by ASTM Committee G-l revealed that nearly
one hundred laboratories are engaged in stress corrosion testing. This
survey was part of a task group effort to determine the state-of-the-art
concerning stress corrosion testing, organized into three areas: specimen
design; test environments and duration; and data evaluation and reporting.
The committee decided to provide a forum at which current practices
would be presented and discussed by those active in the field. Papers
were solicited from engineers and scientists who use these tests for specification purposes, for material selection guides, and for design and use
criteria. Theoretical papers were included where the theory was shown to
influence test methodology.
The theme of the conference was the development of reproducible,
standardized test methods. Nearly all structural metals were covered:
high strength steels, stainless steels, and alloys of aluminum, copper, magnesium, and zirconium. Some new methods were fully discussed, including the precracked cantilever beam and tests on joints, such as bolts
and weldments.
The breadth of coverage shows the wide interest in this subject. Despite
its longevity as a metallurgical phenomenon, stress corrosion cracking
mechanisms are still being investigated. Three previous symposia have
concentrated on this problem: in Philadelphia (1944), in Boston (1954),
and in Pittsburgh (1959). Sufficient time has elapsed since the last one
for those working in the field to realize that the complex nature of the
problem has not yet yielded to the ever-growing amount of scientific
and technical investigations into the causes.
The designer, for. one, is faced with materials selection problems. Yet
he must rely upon stress corrosion tests which are highly empirical and
only poorly correlated with service performance.
To select an example with which the writer is familiar, take highstrength aircraft alloys. Originally, these alloys were fabricated in sheet
form. Stress corrosion tests, among others, were used to develop optimum
alloys and tempers for the strength levels desired. Adequate service performance has been obtained from these materials. When designers and
builders went to different forms of construction, using thick sections,
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2
STRESS CORROSION TESTING
a few cases, it was not achieved, and problems with stress corrosion
cracking grew in number with each new model aircraft. An unsuspected
sensitivity to stress corrosion cracking in the short-transverse direction
was discovered. Stress corrosion cracking was not well enough understood
to foresee these problems. In addition, test methodology was such that
development programs required months of testing. Designers are now
able to deal with the shortcomings of these alloys hi a satisfactory fashion.
This symposium reveals many such problems. The amount of work
which is suggested on the basis of current investigations is staggering, if
the desire to bring rigor into stress corrosion testing is to be fulfilled. The
urgent and critical needs of industry, the defense effort, and the space
program compel us to undertake this goal. From this point, or state-ofthe-art, the committee plans to launch its effort in fulfilling these needs.
The task group for the symposium would like to acknowledge the efforts of the many persons who prepared papers, those who reviewed them,
and the contribution of the ASTM staff for their help in making the
symposium a significant step forward in fighting the problem of stress
corrosion cracking.
H. Lee Craig, Jr.
Research supervisor, Department of Applied Chemistry and
Mathematics, Reynolds Metals Co., Richmond, Va.;
Chairman, Subcommittee VI on Stress Corrosion Testing
and Corrosion Fatigue of Committee G-l on
Corrosion of Metals
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STRESS CORROSION TESTING METHODS
REPORT OF TASK GROUP 1 ON STRESS CORROSION TESTING METHODS
OF SUBCOMMITTEE X ON STRESS CORROSION OF COMMITTEE B-3 ON
CORROSION OF NONFERROUS METALS AND ALLOYS*
Stress corrosion tests are conducted
for a variety of reasons, the reason frequently determining the type of test.
Some of the reasons are:
(1) Evaluate a metal or alloy, or various heat treatments of one alloy for susceptibility to stress corrosion in certain
environments.
(2) Compare stress corrosion susceptibilities of various alloys.
(3) Evaluate environments which
might accelerate stress corrosion cracking in various alloys.
(4) Evaluate a specific service requirement with regard to the possibility
of stress corrosion.
(5) Evaluate the effectiveness of coatings or other protective measures for
reducing stress corrosion of susceptible
metals and alloys.
cur, but complete fracture from stress
corrosion also may not occur because of
stress relaxation during the earlier stages
of cracking.
For (1) and (2), both types of tests
should be conducted in order to obtain
a complete picture of the susceptibility
of an alloy to stress corrosion, preferably
using specimens oriented in more than
one direction with respect to the rolling
direction. Behavior in the short-transverse direction is especially important
for study. For (3) and (5), the most important considerations should be selection of a specimen type and a stressing
method which provide reproducibility
of stress pattern and stress level from
one specimen to another, and selection
of a material known to be susceptible to
stress corrosion cracking. In materials
that are notch sensitive, care should be
Stress corrosion tests may be con- taken to eliminate notches, scratches,
ducted under conditions of either con- etc., which might act as stress raisers.
stant deflection or of constant load. An
For (4), the part should be tested unimportant difference between them is der the type of stress it will encounter in
that under constant load general corro- service. Table 1 lists some types of stress
sion may cause a decrease in cross-sec- that prevail in service with some recomtional area with subsequent increase in mendations as to whether constant destress per unit area. This can lead to flection or constant load would more
mechanical rather than stress corrosion nearly duplicate the stress condition.
failure. Under constant deflection, the For (3), (4), and (5), specimens taken
stress per unit area tends either to stay out in the short-transverse direction are
constant or decrease. Therefore, failure preferable, since this is the direction
by stress rupture is not as likely to oc- which shows the greatest susceptibility.
The type of specimen to be used is
* Reprinted from Proceedings, Am. Soc.
Testing Mats., Vol. 65, 1965, pp. 182-197. frequently determined by the form in
Committee B-3 has been replaced by Commit- which the metal is available (sheet, plate,
tee G-l on Corrosion of Metals. Subcommittee
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STRESS CORROSION TESTING
TABLE 1
Source of Sustained
Tension in Service
Residual stresses:
Quenching!
Forming /
Misalignment
Interference Bushings:
Rigid
Flexible
Flareless fittings \
Clamps
/
Stressing Method Most
Applicable
constant deflection
constant deflection
constant deflection
constant load
either, but constant
defection probably
better
Hydraulic pressure. . . constant load
Dead weight
constant load
Faying surface corrosion
constant load
TABLE 2
Form in Which Material
Is Available
Possible Types of
Specimens
Sheet
. bent beam, preform,
tension specimen, Ubend
Plate (less than 2 in.
thick)
C-ring, bent beam, tuning fork
Plate (more than 2 in.
thick)
short-transverse tension
specimen
Bar (depending on
thickness or diameter)
C-ring, tuning fork
transverse
tension
specimen
Tubing
C-ring, (Battelle system
of internal pressure)
Wire
tension specimen, loop
SPECIMENS FOR SIMULATING SPECIFIC
CONDITIONS
1. Preformed specimens for simulating residual
stresses
2. Welded assemblies
3. Interference rings for simulating pressed in
bushings, fasteners, etc.
specimen types to material form. Some of
the most common types are described
below.
Bent Beam Specimen (Constant Deflection) :
point loading; or by holding the ends in
a rigid jig (snap-in bent beam). All three
methods require jigs.
The snap-in bent beam is the most
commonly used because it requires the
simplest fixture. The stress pattern and
calculation are, however, the most
complicated. If the broad assumption is
made that the specimen makes a truly
circular arc, then the theoretical relationship between stress on the outer
fibers and specimen length can be calculated by Method 1-A described later.
The calculated stress is always lower
than the actual stress, and, if an accurate
value is desired, strain gages should be
used for stress measurement.
A formula is also available which takes
into account the fact that the arc inscribed by the specimen may not be
truly circular. This is given in Method
1-B.
A detailed stress analysis of the bent
beam specimen has been published elsewhere.1 A photograph of the snap-in bent
beam is shown in Fig. 1. Specimens that
are stressed at low levels may be too
short for easy insertion into the stressing jig, and a stressing device may be
useful. Such a device is also shown in
Fig. 1.
A fixture employing four-point loading
provides uniform distribution of stress
between the two inner supports. The
formula for calculating the deflection is
given in Method 1-C.
With three-point loading, a marked
localization of the maximum stress
exists at the central support. The formula for this is given in Method 1-D.
This type of loading has been employed
in an electrolytic test to determine
rapidly the susceptibility to stress corrosion cracking of aluminum-magnesium
This type of specimen may be stressed
1
G. Wed
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W. Longinow,
Corrosion,
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REPORT OF COMMITTEE B-3
5
FIG. 1—Stressing device and jig for bent beam specimen.
alloys. Details of this test are given in stant-deflection specimen useful for
multiple testing has also been described.2
Method 1-E.
Three- and four-point loading may also Tension Specimens (Constant-Deflection
be of the constant-load type if the deor Constant-Load, Depending on the
flection is accomplished by application
Stressing Method}:
of a dead load.
Either flat or round tension specimens
Precautions should be taken to insure
that no galvanic corrosion can take place are suitable for stress corrosion testing.
between the specimens and the supports. Direct loading may be accomplished by
2
G. J. Heimerl and D. N. Braski, "A Stress
Also, with bent beams it is important
Corrosion
Test for Structural Sheet Materials,"
that the edges be broken.
Materials Research
& 16
Standards,
Vol.EST
5, No.
1,
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A self-loading, bent-beam type of con- January, 1965,
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STRESS CORROSION TESTING
use of compression springs, lever arms,
or the tension rings shown in Figure 3.
As pointed out by Godard and Harwood,3
however, dead loading by means of
levers is difficult and requires proper
placement and alignment of pivots to
ensure that the desired load is being applied.
With this type of test, uniaxial stress
is increased by crack nucleation and
propagation due to reduction in crosssectional area. Therefore, the rate of
crack propagation should increase in a
tension specimen. The increase will continue under constant load, while under
constant strain the presence of small
FIG. 2—Specimen holder for electrolytic test.
The load to be applied is measured by
calibrating the compression springs. The
tension rings can either be calibrated for
deflection under load, or a strain gage
can be placed on the bolts and the bolt
calibrated for strain.
Constant deflection can be obtained
by means of the fixture shown in Fig. 4.
The | in. diameter tension specimen
shown in Fig. 4 is frequently used for
testing susceptibility of aluminum plate,
forgings, and castings to stress corrosion
in the short-transverse direction. The
strain to be applied is best measured by
a strain gage or an extensometer on the
specimen.
cracks may tend to cause relaxation in
the specimen.
C-Ring Specimen:
The C-ring is useful for making transverse tests of a wide variety of products.
If possible, the C-rings should be oriented
in such a way that the direction of maxialso
mum tensile stress will be parallel to the
short-transverse direction of the piece.
The recommended orientation and a
drawing of this specimen are shown in
Fig. 5. The formula to be used for stressing the C-rings is given in Method 2-A.
A specimen similar to a C-ring but requiring less machining is shown in Fig.
3
6a. As the legs are drawn together by
H. P.
Godard and
J. Harwood,
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REPORT OF COMMITTEE B-3
with a strain gage mounted in the center
of the surface being placed in tension.
Two other types of "tuning fork"
specimens are shown in Figure 6b. Formulas that can be used for stressing these
specimens are given in Method 2-B.
7
posed, the face or root sides of both
strips are stressed in tension, otherwise
the face and root sides of the weld are
stressed against each other (Fig. 8). The
applied stress does not take into account
Sheet-Type Preform Specimen:
Sometimes methods of fabrication or
assembly induce residual stresses in the
metal. The effects of these stresses on
corrosion can be determined by exposing
the fabricated or assembled parts to a
suitable corroding medium. Stresses induced by fabrication or assembly are
likely to be of greater magnitude than
the design stresses.
One type of specimen that may be used
to evaluate the effects of stresses induced
by cold work is termed the "preformed"
specimen (Fig. 7). A high elastic bending
stress is superimposed upon residual
stresses introduced by forming, giving a
total sustained stress that is relatively
high but not readily determined.
In this specimen, a depression is formed
in the center of a strip of metal 9 in. long
machined to a width of 0.700 ± 0.005 in.
The formed specimen can then be exposed
without further treatment, or else the
ends can be forced together and sprung
into a jig or stressing plate with a 7f in.
span. This latter condition is generally
used and is a particularly severe test
since it involves both residual and external stressing.
Welded Beam Assemblies:
FIG. 3—Tension rings for loading tension
specimens.
A bent-beam specimen can be used for
testing butt welds. For such investigathe stress concentrations accompanying
tions, strips full thickness by 1 in. wide
by 8f ft long are sawed and machined the variations around the weld bead.
The formula for calculating the defrom large welded panels.
flection
is given in Method 3.
The specimens are stressed in pairs
against each other using H-type saddles Interference Ring Specimen:
and half-rounds and drawing up the
The interference ring type of specimen
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STRESS CORROSION TESTING
desired to develop a hoop stress in a particular part and where other more economical specimens are not suitable. This
ring is particularly advantageous, for
example, in circular die forged parts
where the most critical grain structure
tails of the method and the formula for
stressing are given in Method 4.
Tubing Specimen:
A specimen design and jig for stress
corrosion testing of tubing under con-
FIG. 4—Tension specimen and stressing fixture.
exists at the surface of the part. Other stant load is shown in Fig. 10. This
advantages are that a relatively large method can be adapted to study a variety
surface area of metal is placed under a of alloys under either aqueous or gaseous
uniform tensile stress, and the assembled environments. Details of the method and
unit simulates practical situations on the formula for calculating the hoop
structures containing various types of stress are given in Method 5.
interference-fit components.
U-Bend Specimens:
The ring is stressed by pushing it onto
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FIG. 5—C-ring specimen.
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6—Tuning
fork specimen.
REPORT OF COMMITTEE B-3
U-bends are formed by plastic deformation from sheet stock. Tensile stresses are
applied by closing the ends of the U with
a nut and bolt. The U-bend specimen is
highly sensitive to stress corrosion crack-
11
bend specimen. A detailed discussion of
the U-bend has been published elsewhere.4
With very high-strength alloys it is
sometimes impossible to form a U-bend
FIG. 7—Preformed sheet specimen.
FIG. 8—Welded beam stress corrosion assembly.
ing, but the test is mainly considered to without fracturing the specimen. Such
be qualitative since it combines a rela- materials can be bent to 150 deg in the
tively unknown stress condition with a annealed condition and then heat treated
relatively unknown condition of work before drawing the legs parallel. Care
hardening. These specimens frequently
4
N. Wed
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Corrosion
Stainless
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STRESS CORROSION TESTING
PRESSURE AND CORRODENT INSIDE TUBE
FIG. 9—Stressing tools for interference ringtype specimens.
should be taken to avoid springback before closing the ends with a bolt. A stressing device to prevent springback has
FIG. 10—Jig and specimen for stress corrosion testing of tubing under internal pressure.
been described elsewhere.5 It is also important that edges be broken for this
type of specimen.
FORMULAS FOR STRESSING BENT BEAMS
METHOD 1-A
Stress-Strain Relationship:
where:
e = outer fiber strain,
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REPORT OF COMMITTEE B-3
a = outer fiber stress, psi, and
E = modulus of elasticity, psi.
Strain on Outer Fibers When Flat Specimens Are Bent into Simple Arcs:
where:
t = material thickness, in., and
r = radius of curvature, in.
13
h = distance between outer fiber and
neutral axis, and
rm = minimum radius of curvature of
the bent beam.
where:
t = thickness of specimen and
y = maximum deflection of the bent
beam.
FIG. 12—Four-point beam loading.
FIG. 11—Diagram for calculation of specimen length.
• Calculation of Specimen Length (see Fig.
FIG. 13—Three-point beam loading.
When y is much greater than t:
11):
and
where:
S = specimen length, in.,
/ = jig length, in., and
9 = central angle, radians.
METHOD 1-B
The basic equation for determining the
stress, (r, in the outer fiber of a bent beam
specimen is:
The curve of a beam, bent by restricting the ends, closely assumes a parabolic
shape. The semi-latus rectum of a parabola is the minimum radius of the curve.
The length of the arc of a parabola which
closely approximates the length, /, of
the specimen may be calculated from the
equation:
where:
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STRESS CORROSION TESTING
FIG. 14—Correction factor for stresses in C-ring specimen.
where:
d = jig span and
rm = Et/2(T has been substituted into
the equation.
METHOD 1-D
Three-Point Loading Stress Calculations
for Bent Beams (Outer Fibers) (See
Fig. 13):
METHOD 1-C
Four-Point Loading Stress Calculations
for Bent Beams (Outer Fibers) (see Fig,
12):
Symbols and units are the same as for
Method 1-C.
METHOD 1-E
This test involves stressing a 3-in.
where:
strip cut from heavy sheet or plate as a
y = deflection, in.,
simple beam in the holder shown in Fig.
2. A high strain value is used. The speciF = stress in outer fibers, psi,
L = beam length, in.,
men is made the anode in an electrolytic
A = distance between inner and outer cell.
load points, in.,
The deflection of the middle of the
E = modulus
of
elasticity,
psi,
and
specimen
is calculated
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D = specimen
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REPORT OF COMMITTEE B-3
15
FIG, IS—Dimensions of tuning fork specimen.
where;
y — deflection, in.,
« = strain in outer fibers,
L — distance between end supports, in.,
and
t = thickness of specimen, in.
FORMULA FOR STRESSING C-RING SPECIMENS
METHOD 2-A
Procedure:
1. Measure with a micrometer to the
nearest 0.001 in.:
(a) The outside diameter parallel to
the stressing screw (averaging the two
ends of the ring).
(b) The wall thickness.
2. Set up a table to calculate the final
diameter (0Z?/) required to give the desired stress using the following equations:
where:
A
= change of OD giving desired
stress, in.,
/
= desired stress, psi,
OD = outside diameter, in.,
t
D
= wall thickness, in.,
= mean diameter (OD ~ t), in.,
= modulus of elasticity,
= a constant (function of ring
D/f), see Fig. 14, and
ODf = final outside diameter of stressed
C-ring.
To simplify calculations, certain terms
in the above equation may be combined
into a constant that will be applicable
for a group of rings of the same alloy and
size. Let 4EZ/ir ~ K, a constant. Then
A - fiy/Kt. If the alloy or the size of
the ring is changed, a different K must
be calculated.
E
Z
NOTE—The main source of error 10 the stress
determination lies in the measurements of the
C-ring dimensions. If in a typical example of a
0.750 in. O.D. by 0.060 in. wall thickness C-ring
the measurements are determined accurately to
the nearest 0.001 in., the total error should not
exceed about 6 per cent.
FORMULA FOR STRESSING TUNING FORK SPECIMENS (JIG. 15)
METHOD 2-B6
where:
total amount of closure at tine
ends, in.,
length of tines, in. (see Fig. 15),
* Derived from
deflection formula tor cantileverCopyright
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16
STRESS CORROSION TESTING
/ = thickness of each tine, in.,
E = modulus of elasticity, psi, and
a = maximum stress in outer fiber of
either tine, psi.
STRESSING OF WELDED BEAM ASSEMBLY
METHOD 3
The deflection used in stressing the
welded beam assembly is obtained by the
use of the following formula:
where:
da = change in the distance between the
/
E
i
L
a
plates along the longitudinal axis
of the bolts, in.,
= the required stress, psi,
= modulus of elasticity, psi,
= thickness, in.,
= length (center to center of bolts),
in., and
= distance from center of bolts to
nearest flange of "H" section, in.
STRESSING OF INTERFERENCE RING SPECIMENS
METHOD 4
Interference Ring Specimen:
The nominal dimensions of this specimen can be varied to suit the part being
tested. The following limitations should
be kept in mind when specifying dimensions for the specimen:
1. The width of the ring should be at
least four times the wall thickness to
insure maximum uniformity of stress
across the ring. (Note—There is no
known maximum limitation to the width
of the ring and probably this would be
limited only by the ability to insert the
plug into the ring.)
2. The stress varies through the thickness of the ring, being highest at the inside surface.
stress on ID
3. The width of the plug should be two
times the width of the ring + f in.
4. The tolerance in the plug diameter
is 0.0005 in. This is adequate for interferences of 0.0080 in. or more, resulting
in a maximum possible error of 6 per cent.
For smaller interferences, it may be desirable to tighten this tolerance so as to
maintain the maximum possible error at
6 per cent.
Stressing:
The following formula is used to calculate the interference required for stressing the ring.
where:
/ = interference (on the diameter) between ring and plug, in.,
E = modulus of elasticity, psi,
ID = inside diameter, in.,
OD = outside diameter, in., and
F = circumferential stress desired at
the OD, psi.
For example, if the wall thickness of the
ring is 0.042 times the OD, the maximum
The ID of the ring must be measured
and minimum stress in the ring will differ by about 10 per cent. For greatest as accurately as possible (usually to the
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uniformity,
the wall thickness should be nearest half mil) so that the diameter of
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REPORT OF COMMITTEE B-3
give the required interference as accurately as possible.
A ssembly:
The ring is stressed by pushing it onto
the plug, using a lubricant to prevent
galling. Liquid butyl stearate or petroleum jelly are good lubricants that are
clean and easy to remove from the exposed surfaces after assembly of the
17
units. Figure 9 shows an assembled ring
and plug and the stressing tools employed.
In cases where it is not practicable to
make the plug of the same material as
the ring, the entire exposed surface of the
plug and the fillet at the edges of the
ring should be effectively coated to prevent galvanic action between the ring
and plug. A suitable wax or paint may be
used for this purpose.
FIG. 16—Barlow correction factor.
STRESS CORROSION TEST FOR TUBING (CONSTANT LOAD)
METHOD 5
The hoop stress, which is uniform
across the tubing thickness, can be calculated as follows :
The test specimen is inserted into the
jig and welded at each end; therefore, the
material used for the jig should possess
welding compatibility with the tubing
material. The wall thickness of the tubing is reduced in the center as shown in where :
Figure 10. The corrodent is introduced a = hoop stress,
into the
tubing
there
P Wed
= internal
pressure,EST 2015
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/ = wall
thickness.
18
STRESS CORROSION TESTING
Correct calculated stress by CF from
the ratio of D0/D\ which can be obtained
from curve in Fig. 16 (D0 = outside
diameter).
SUMMARY
This report represents an attempt to
provide the novice in stress corrosion
testing with some idea of the principles
and practical aspects that must be considered in planning a meaningful stress
corrosion testing program. Some reasons
for conducting stress corrosion tests are
outlined.
Differences between the two types of
stressing, constant-deflection and constant-load, are pointed out, and some
basis is provided for selecting one or the
other. Some of the most common types
of specimens are described in detail.
Emphasis is placed on specimens that
could be used for studying more than one
type of problem, or that could be used
for specific situations of more or less
universal concern.
The Subcommittee is indebted to the
following task groups for developing and
preparing this report:
Sara J. Ketcham; Task Group Chairman.
W. E. Berry
W. W. Binger
J. F. Eckel
O. B. Ellis
H. B. Romans
R. S. Shane
D. H. Thompson
A. C. Willhelm
Respectfully submitted on behalf of
the Subcommittee,
H. Lee Craig, Jr.,
Chairman
ADDENDUM
A more complete picture of the stress
corrosion characteristics of an alloy is
obtained if the effect of stress raisers is
also taken into consideration. Stress
raisers such as screw threads, sharp fillets,
fatigue, and weld cracks may be built
into a part, while others such as pits may
naturally develop during exposure to a
corrosive environment. Notched or precracked specimens are useful for this
purpose.
by changing the depth or the radius of
the notch or both. Related values for
these parameters are available from
Ref 2.
Precracked Specimens:
The philosophy behind the use of precracked specimens, detailed test procedure, and treatment of the stress
aspect by fracture mechanics have been
set forth in several publications [J, 4\.
Briefly, the purpose of this test procedure
Notched Specimens:
is to permit characterization of a metal
Notched tensile bars or C-rings such in terms of the size of a preexisting flaw,
as those shown in Fig. 17a and b can be which would be expected to initiate a
used to investigate the effect of stress stress corrosion crack in that metal under
raisers on susceptibility of an alloy to a specific stress in a particular environstress corrosion cracking. A detailed ment.
description of the use of the notched CThe most common type of specimen is
ring is available [7]7. Higher or lower stress shown in Fig. 17c. A notch is first maconcentration factors can be obtained chined into a rectangular bar, then the
' The italic numbers in brackets refer to the specimen is held against an eccentric
bearing
in 16
a machine
lathe2015
and fatigued
list of Copyright
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at the
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FIG.(all17—Notched
and
precracked
specimens.
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20
STRESS CORROSION TESTING
until the machined notch is extended.
The specimen is then stressed as a cantilever beam.
Methods for Wire, Rope, and Cable Products:
The types of specimens most commonly used for these products are tension specimens, either with or without
reduced sections. These can be dead
loaded by means of calibrated springs or
levers. The fixtures shown in Figs. 3 and
4 could be modified for this purpose.
A free loop is used to stress wire into
the plastic region. The test consists of
looping wires around mandrels three
times their diameters, permitting them
to spring back. After exposure to a corrosive media, damage is detected by
failure in unlooping. Cracking occurs on
the inside of the loops, where the surface
layers are in tension after springback.
Use of such a specimen and the approximate residual stresses have been described [5, 6].
given in the report are accurate enough
for most test work. However, for precise
work, the experimenter can only be sure
of the stress levels by determining the
strain with strain gages. This particularly
applies to specimens such as those shown
in Figs. 6 and 8.
Consideration should be given to the
possibility of crevice corrosion developing
at areas where specimens are in contact
with a jig. In some cases, this can provide
galvanic protection to the stressed area.
In other cases, it may give rise to hydrogen embrittlement by the cathodic corrosion product, hydrogen. The areas where
the jig and the specimen are in contact
should then be either masked off or,
where possible, the jigs suspended in
such a way as to keep the contact area
out of the corroding medium.
Respectfully submitted on behalf of
Task Group I,
Sara J. Ketcham
Naval Air Engineering Center,
Aeronautical Materials Laboratory,
Philadelphia, Pa.;
Chairman
Precautionary Notes:
Stresses calculated from the formulas
SUGGESTED REFERENCES
F. A. Champion, Corrosion Testing Procedures,
John Wiley, New York, 1952.
TJ. R. Evans, The Corrosion and Oxidation of
Metals: Scientific Principles and Practical
Applications, chapter XVI, St. Martin's
Press, New York, 1960.
H. H. Uhlig, Corrosion and Corrosion Control,
chapter 7, John Wiley, New York, 1963.
Stress Corrosion Cracking and Embrittlement, W.
D. Robertson, editor, John Wiley, New York,
1956.
Symposium on Stress Corrosion Cracking in
Metals, ASTM-AIME, 1944.
Physical Metallurgy of Stress Corrosion Fracture,
T. H. Rhodin, editor, Interscience, New York,
1959.
CITED REFERENCES
[1] ¥. S. Williams, W. Beck, and E. Jankowsky,
"A Notched Ring Specimen for Hydrogen
Embrittlement Studies," Proceedings, Am.
Soc. Testing Mats., Vol. 60, 1960, p. 1192.
[2] R. E. Peterson, Stress Concentration Design
Factors, John Wiley, New York, 1953.
[3] B. F. Brown and C. D. Beachem, Corrosion
Science, Vol. 5, 1965, p. 745.
[4] B. F. Brown, "A New Stress-Corrosion
Cracking Test for High-Strength Alloys,"
Materials Research & Standards, Vol. 6,
No. 3, March, 1966, p. 129.
[5] G. T. Spare, Wire and Wire Products, Vol.
29, No. 12, 1954, p. 1421.
[6] G. Brewer and H. C. Ihsen, Metal Progress,
April, 1945, p. 707.
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W. J. Helfrich1
Influence of Stress and Temperature on
Short-Transverse Stress Corrosion Cracking
of an AU.2Zn-2.5Mg Alloy
REFERENCE: W. J. Helfrich, "Influence of Stress and Temperature on
Short-Transverse Stress Corrosion Cracking of an Al-4.2Zn-2.5Mg Alloy,"
Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967,
p. 21.
ABSTRACT: Stress corrosion cracking of an age-hardened Al-Zn-Mg alloy (7039) subject to continuous immersion in a 1 N NaCl solution was
studied at stresses of 5 to 55 ksi and temperatures of 30 to 100 C and
found to be an activated process. The rate, r, of stress corrosion (proportional to the inverse failure time) can be expressed as:
r = /-o exp{ - [Q*(0) - sV*}/RT\
where Q* (O) is the activation energy in the absence of stress, s, and V*
is the activation volume. Q* (O) has a value of 20 ± 0.8 kg-cal/mole, and
V* is approximately 28 to 34 cmVmole for applied stresses of 25 ksi or
higher. It is suggested that the rate-determining step in the stress corrosion
of 7039 involves anodic dissolution of MgZn2 at grain boundaries. The
mechanism of stress corrosion cracking is discussed with reference to
previously derived models.
KEY WORDS: aluminum alloy, stress corrosion, cracking, activation,
electrochemical dissolution, pitting (corrosion), corrosion
The effect of temperature on the kinetics of stress corrosion cracking
of metals has not been studied in great detail. Yet, for a process in
which electrochemical dissolution may determine the time dependence
for failure, temperature should be a powerful rate-determining factor.
Temperature effects undoubtedly account, in part, for the effectiveness
of hot concentrated nitrates and alkalines in causing stress corrosion of
mild steels, boiling 42 per cent magnesium chloride in cracking of austenitic stainless steels, and hot saline solutions on the reduced time-tofailure in certain aluminum alloys. In this latter connection, Gruhl [I]2
1
Research engineer, Department of Metallurgical Research, Kaiser Aluminum
& Chemical Corp., Spokane, Wash.
2
The italic numbers in brackets refer to the list of references appended to this
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22
STRESS CORROSION TESTING
has shown that stress corrosion cracking in an Al-Zn-Mg alloy was dependent upon temperature to the extent that the observed failure times
were almost halved for each 10 C rise in testing temperature in the
range of 25 to 70 C. However, not all aluminum alloys are similarly
affected. The stress corrosion failure times of an age-hardened 2024
aluminum alloy are not measurably dependent upon the testing temperature in the range of 3 to 60 C [2]. In a like manner, the rate of intergranular corrosion in unstressed 2024 specimens is independent of testing temperature below 60 C but strongly temperature dependent above
this temperature [3]. Although stress corrosion tests were not conducted
at temperatures above 60 C, Nichols and Rostoker [3] questioned the
compatibility between a process of intergranular penetration that is presumably independent of temperature and theories of stress corrosion
cracking which are based on an intergranular corrosion mechanism.
While it is well known that increasing stress shortens the observed
time-to-failure in aluminum alloys, arguments against a mechanism for
stress-enhanced anodic dissolution are based on the fact that elastic
stress (strain) adds little to the thermodynamic driving force for dissolution [4]. That is, elastic strain energy causes a change in the anodic
corrosion potential of aluminum of only 10~4 to 10~3 v per increment
of 10,000 psi stress, which is negligible compared with normal reversible
corrosion potentials of about 1 v. Thus, it appears that the problem is not
one of whether stress accelerates stress corrosion cracking, but rather
the manner in which the rate of cracking is dependent upon stress. Hillig
and Charles [5] recognized that the rates of some chemical reactions
involving solids are dependent upon the mechanical stress at or near the
reaction boundary. Their theory for stress corrosion of an amorphous
elastic solid fits the known experimental data of Mould and Southwick
[6], Wiederhorn [7], and Charles [8] for static fatigue of soda-lime glass.
The fracture process in glass has been shown to be stress dependent and
thermally activated, consistent with a corrosion process involving transport of sodium ions to the reaction interface.
Since it is the purpose of this paper to show that stress corrosion
cracking in certain aluminum alloys is an activated process, new studies
on the t6mperature and stress dependence for failure in an Al-4.2Zn2.5Mg alloy are presented. These results are discussed with reference to
the model for stress corrosion proposed by Hillig and Charles and hi
terms of the mechanism of stress corrosion cracking.
Experimental
A single lot of commercially available3 1-in. 7039-T64 alloy plate
3
Kaiser Aluminum and Chemical Corp., Oakland, Calif.
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HELFRICH ON INFLUENCE OF STRESS AND TEMPERATURE
23
(66.1 ksi tensile strength, 57.5 ksi yield strength) was obtained for study.
The chemical composition is given below.
Weight per cent.
Zn
Mg
Mn
4.2
2.5
0.24
Cr
0.16
Fe
0.30
Si
Cu
Ti
V
Al
0.14
0.03
0.01
0.01
balance
Specimens for short-transverse stress corrosion testing were machined
into C-rings measuring 0.948 in. in outside diameter, 0.750 in. in width,
and 0.060 in. in thickness, then dressed with 180-grit emery, degreased
in acetone, and stressed. Machining and the method of stressing have
been described elsewhere.4 Prior to testing, all but the upper tension
surface of each specimen was dip-coated in liquid neoprene to eliminate
the possibility of dissimilar alloy corrosion between the C-ring and the
aluminum fastener used in stressing the ring.
Specimens in groups of five C-rings for each variable of stress (5, 10,
15, 25, 35, 45, and 55 ksi) and temperature (30, 40, 60, 80, and 100 C)
were tested by continuous immersion in an aerated 1 N sodium chloride
solution. Solutions were made up with reagent grade salt and distilled
water with a pH of about 6.0 and resistance greater than 500,000
ohm • cm. All tests were conducted in 2000 ml of freshly prepared solution,
in turn, contained in resin reaction kettles refluxed to prevent vapor loss.
Solutions were changed every five to seven days. Temperature control to
±1.0 C was obtained with a mercury-in-glass thermoregulator, by agitating the solution with a magnetic stirrer and by regulating the power
input to a 600-w Pyrex immersion heater.
In general, all specimens were tested to failure, with the exception
of those tested at or near the threshold or endurance stress. The criterion for failure was based on first visual evidence of cracking at five
diameters magnification. Since it was found that the failure times were
log-normally distributed, the geometric mean failure time was selected
as representative of the stress corrosion performance for each group of
test specimens [9].
Discussion
General Approach
Stress corrosion cracking is assumed to be a thermally activated
process in which the rate of activation is dependent upon the local state
of stress (driving force) at or near the reaction boundary. According to
4
See p. 3.
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24
STRESS CORROSION TESTING
Hillig and Charles [5] the activation free energy, Q*(s), for the corrosion reaction can be expressed by a Taylor series expansion at s = O
where higher than first-order power terms of stress, s, are considered
negligible. Q*(O) is the activation free energy for the corrosion reaction
on a stress-free surface. The term dQ*/ds has the dimensions of volume
and is termed the activation volume, V*, from analogy with terms such
FIG. 1—Effect of applied stress at constant temperature on stress corrosion
cracking of 7039-164.
as (3G/dp)T r = V. In the case of tensional forces, it is expected that
dQ*/ds will be negative [5].
The rate of activation, r, for the corrosion reaction in the direction of
the applied force can be given by
where R and T have their usual thermodynamic significance. The preexponential kinetic factor, r 0 , expresses the rate of corrosion of an unstressed planar surface and is exponentially dependent upon the zerostress molar free energy and the surface free energy at the reaction
boundary. This latter term is associated with the surface energy at a
concave
solid-corrosion product interface and represents the surface
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tension
in
retarding
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HELFRICH ON INFLUENCE OF STRESS AND TEMPERATURE
25
interface may become very large, this term can become limiting when
the applied force (stress) is small [5]. Elastic strain energy, which can
also be associated with the driving force for the corrosion reaction, exerts
a negligible thermodynamic effect and is neglected.
Equation 1 describes how stress and temperature influence the rate of
activation for corrosion of a homogeneous brittle solid such as sodalime glass [7]. Hillig and Charles [5] extended these arguments to include stress corrosion cracking in homogeneous plastic materials. It will
be shown that this approach is applicable to stress corrosion cracking
in a nonhomogeneous precipitation-hardened Al-Zn-Mg alloy as well.
Stress Corrosion in Al-4.2Zn-2.5Mg
The logarithms of the median failure times, as a function of the applied
stress at constant testing temperature, are plotted in Fig. 1. It is apparent
that the data obtained for stresses in excess of about 25,000 psi are adeTABLE 1—Activation volume V* for stress corrosion of Al-4.2Zn-2.5Mg at various
temperatures and stresses between 25 and 55 ksi.
Testing Temperature, deg C
100
80
60
40
30
V*, cm'/mole
27 ± 5
26
30
29
28
quately represented by straight lines and, assuming that the rate of reaction is proportional to the inverse of the failure time, satisfy Eq 1. That
is, the log failure times at constant testing temperature are directly proportional to the applied stress. The slopes of the straight line portion of
the curves in Fig. 1 are proportional to the activation volume V* which,
in turn, should be independent of temperature. The data in Table 1
illustrate this point.
The average value of the activation volume obtained from these data
is approximately 28 ± 5 cm3/mole.5 (The error was computed for a
two-sided 95 per cent confidence interval and log standard deviation of
0.12.)
Real departures from linearity exist at applied stresses below about
25 ksi (Fig. 1). The time-to-failure increases with decreasing stress at a
faster than exponential rate, and an approach to a threshold stress is
5
Presumably, during the early stages of stress corrosion, the stress at the reaction interface can be equated to the applied stress. However, stress concentration
undoubtedly occurs with continued chemical attack. Since the acting interfacial
stresses could approach values of several times the applied stress, it is not imposCopyright
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sible
that the
activation
somewhat
smaller
than
the value
quoted here.
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26
STRESS CORROSION TESTING
indicated. It is interesting, however, that for practical purposes, the
threshold stress is sensitive to the testing temperature. Specimens stressed
to 10 ksi and tested at temperatures above 40 C and specimens stressed
at 5 ksi and tested at 100 C failed in a week or less, yet no failures were
observed in specimens stressed at 10 ksi and tested more than 14 weeks
at 30 C. While it was considered instructive to continue these latter tests,
FIG. 2—Temperature dependence on stress corrosion cracking of 7039-T64 at
applied stresses of 5 to 55 ksi.
in that failures may have occurred, prolonging the test period had the
undesirable side effect of causing general corrosion in the form of pitting
attack.
Reduced times-to-failure, with increasing temperature at constant
stress, were noted because stress corrosion is a thermally activated
process. The log of the median failure times for specimens subject to
constant stress, as a function of the inverse of the absolute test temperatures, are graphically presented in Fig. 2. Excepting the results for specimens stressed at 5 and 10 ksi, the data are adequately represented by
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straight
lines, thebyslopes of which are a measure of the activation energy
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HELFRICH ON INFLUENCE OF STRESS AND TEMPERATURE
27
for stress corrosion under constant loading. These results are given in
Table 2.
The approximately uniform decrease in the observed activation energy
of 0.5 to 0.6 kg-cal/mole with increasing stress in increments of 10 ksi
at or above 25 ksi again satisfies Eq 1 and is statistically significant within
a 95 per cent confidence interval. The average change in the activation
energy of 0.56 kg-cal/mole per 10 ksi gives a measure of the activation
energy Q*(O) in the absence of stress, namely, 20 ± 0.8 kg-cal/mole.
The data are not entirely self-consistent, however, in that a decrease of
0.56 kg-cal/mole per 10 ksi yields an activation volume V* of about
34 cmVmole, which is somewhat larger than the value quoted earlier
(28 cmVmole).
Mechanism of Stress Corrosion
An activation energy Q*(O) of about 20 kg-cal/mole and activation
volume V* of 28 to 34 cmVmole point to a rate determining step in the
TABLE 2—Observed activation energy for stress corrosion of Al-4.2Zn-2.5Mg at
temperatures of 30 to 100 C and stresses of 15 to 55 ksi.
Q*(0) - sV*, kg-cal/mole
Applied Stress, ksi
16.9 ± 0.8
17.5
18.1
18.6
20.4
45
35
25
15
55
process of stress corrosion of 7039 alloy which does not involve transport
phenomenon, for example, diffusion of a chemically reactive species,
such as magnesium or zinc, to the reaction boundary. In aluminum, the
latter might generally be characterized by an activation volume in the
neighborhood of 10 cmVmole [10] and an activation energy of 28 to 31
kg-cal/mole [11]. If, in fact, the rate of stress corrosion in Al-Zn-Mg
alloys is controlled by electrochemical dissolution of an anodic phase, it
may be possible to account for the values quoted above.
The principal age-hardening phase in aluminum (3.5 to 6 per cent
zinc, 2 to 4 per cent magnesium) alloys aged at temperatures of less than
200 C has been identified as MgZn2 . During aging, M-MgZn2 forms in
the grain boundaries and, depending upon the degree of aging, GuinierPreston zones and platelets of M'- and M-MgZn2 form in the matrix
[12-15]. Taking the measured lattice spacings [72] and the density [76]
of the M phase, then by calculation the molar volume of MgZn2 is found
to be approximately 30 cmVmole. Since the results of the present study
indicate that V* equals 28 to 34 cmVmole, it is suggested that the rateCopyright byprocess
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controlling
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intergranular
stress
of 7039
alloy is
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28
STRESS CORROSION TESTING
based on stress and thermally activated anodic dissolution of MgZn2 at
grain boundaries. Unfortunately, comparative data relating to the zerostress activation energy for dissolution of Al+MgZn2 in aqueous salt
solutions are not available.
The suggested mechanism and the findings of other studies have yet
to be related to the overall process of stress corrosion cracking in AlZn-Mg alloys. Hillig and Charles [5] suggest that corrosion in an ideally
plastic solid increases the stress concentration near the crack tip until
the flow stress, s f , a i s r e a c h e d . A s l o n g a s t h e a p p l i e d s t r e s s , s , i s l e s s t h a n
st, the maximum stress in the vicinity of the crack tip will not exceed sf.
To develop the stress concentration necessary for mechanical rupture,
the value of smust locally attain larger values than the matrix yield
stress. Probable factors favoring this condition in a nonhomogeneous
(precipitation-hardened) solid are:
1. The presence of a high density of sessile dislocations in the vicinity
of the corroding phase. Recently, Jacobs [13,14] proposed that dislocations associated with MgZn2 precipitates are in some manner responsible
for the high susceptibility of 7075-T6. High dislocation densities are not
observed in 7075-T73, and this product exhibits excellent resistance to
^stress corrosion cracking. As suggested by Hillig and Charles [5] in regions of linear dimensions less than the mean spacing between dislocations, the local stress could exceed sf.
f.
2. The absorption of interstitial hydrogen. Haynie et al [17] have suggested that local cathodic charging of hydrogen at grain boundaries
causes stress corrosion failures in 7079 alloy by an embrittling effect.
Although aluminum typically exhibits a low solubility for hydrogen,
stressed grain boundaries may absorb sufficient hydrogen for embrittlement.
It is conceivable that either, or both, of these factors contribute to
mechanical crack propagation during stress corrosion of aluminum alloys.
However, as an alternative, we will consider the following model in
which the specimen is being stressed in the elastic range and stressenhanced corrosion of an anodic phase (MgZn2 in 7039 alloy) has initiated a pit or notch at the grain boundaries. Again, according to Hillig
and Charles [5], the stress ahead of the notch tip remains substantially
constant up to a distance y from the tip equal to p, the radius of curvature
at the notch tip. As tip sharpening by corrosion occurs, the level of stress
at the notch rises dramatically until, at a distance y — y0 from the tip,
the stress equals the matrix yield stress, s f . With further sharpening,
local yielding would maintain the stress approximately constant at some
distance from y0 to an elastic region where the stress would drop below
the yield strength. Thus, the conditions for brittle rupture can exist, even
when local yielding occurs, if p is less than y0 [5].
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HELFRICH ON INFLUENCE OF STRESS AND TEMPERATURE
29
spacing between MgZn2 precipitates at the grain boundaries. Therefore,
the significant parameters which would favor brittle rupture are (a) a
small particle spacing at or near the grain boundaries and (6) a maximum
value for the matrix flow stress. Clearly, the latter condition is obtained
when the alloy has been aged or otherwise fabricated to produce maximum strength. Overaging or other practices which reduce the matrix flow
stress or increase grain boundary ductility would favor increased resistance to stress corrosion.
A minimum precipitate spacing at the grain boundaries would favor
maximum susceptibility to stress corrosion cracking and, in all probability, would be a sensitive function of the quench rate from the homogenization temperature. Rapid cooling would favor the growth of large numbers of smaller precipitates, effectively reducing the particle spacing and
the resistance to stress corrosion cracking. Of course, a continuous film
of precipitate at the grain boundaries (nil-ductility) would presumably
result in the maximum attainable susceptibility and may also account
for the continuous rates of crack growth noted in 7079 alloy [IS].
The proposed model is entirely consistent with the statistical approach
to stress corrosion cracking in aluminum alloys. As noted previously, the
failure times observed in the stress corrosion of 7039 alloy are log-normally distributed. If, as suggested, the time dependence for failure is a
function of the rate of corrosion, and if this rate is subject to the influence
of a large number of small independent events which operate simultaneously, then a log-normal distribution of failure times would occur
quite naturally [79]. Physically, the log-normal distribution of failure
times describes a process involving the simultaneous development of a
large number of small independent sites of corrosion (pits or notches)
at grain boundary precipitates. These sites of corrosion then join, presumably by mechanical rupture, to form a macroscopic stress corrosion
crack.
Conclusions
As a result of the present studies, the following observations are believed to be applicable to intergranular stress corrosion cracking of 7039
alloy:
1. Stress corrosion is a stress-activated and thermally activated process
exhibiting a zero-stress activation energy of 20 ± 0.8 kg-cal/mole and
an activation volume of 28 to 34 cm3/mole.
2. The rate-controlling, step in stress corrosion presumably involves
anodic dissolution of MgZn2 at grain boundaries, in part, consistent with
an electrochemical mechanism. Factors favoring rapid or easy mechanical
crack propagation are believed to be associated with a minimum particle
spacing and minimum ductility at the grain boundaries and a maximum
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30
STRESS CORROSION TESTING
References
[1] Gruhl, W., 'The Temperature Dependence for Stress Corrosion Cracking
in AlZnMgs," Zeitschrift fuer Metallkunde, Vol 53, No. 10, 1962, pp. 670675.
[2] Nichols, H. and Rostoker, W., "Analogies Between Stress-Corrosion Cracking
and Embrittlement by Liquid Metals," Transactions, American Society for
Metals, Vol 56, No. 3, Sept. 1963, pp. 498-500.
[3] Nichols, H. and Rostoker, W., "Intergranular Corrosion Penetration in an
Age-Hardenable Aluminum Alloy," Journal of the Electrochemical Society,
Vol 112, No. 1, Jan. 1965, pp. 108, 109.
[4] Barnartt, S., "General Concepts of Stress-Corrosion Cracking," Corrosion,
Vol 18, No. 9, Sept. 1962, p. 328t.
[5] Hillig, W. B. and Charles, R. J., "Surfaces, Stress Dependent Reactions, and
Strength," Report No. 64-RL-3756M, Oct. 1964, General Electric Research
Laboratory, Schenectady, N.Y.
[6] Mould, R. E. and Southwick, R. D., "Strength and Static Fatigue of Abraded
Glass Under Controlled Ambient Conditions," Journal of the American Ceramic Society, Vol 42, 1959, pp. 542, 582.
[7] Wiederhorn, S. M., "Effect of Environment on the Fracture of Glass," NBS
Report 8901, June 1965, U. S. Department of Commerce, National Bureau of
Standards, Washington, D.C.
[8] Charles, R. J., "Static Fatigue of Glass: II," Journal of Applied Physics, Vol
29, No. 11, Nov. 1958, pp. 1554-1560.
[9] Booth, F. F., Tucker, G. E. G., and Godard, H. P., "Statistical Distribution
of Stress Corrosion Endurance," Corrosion, Vol 19, No. 11, Nov. 1963, p.
390t.
[10] Butcher, B. M., Hutto, H., and Ruoff, A. L., "Activation Volume and Energy
for Self-Diffusion in Aluminum," Applied Physics Letters, Vol 7, No. 2, July
1965, pp. 34-35.
[11] Stoebe, T. J. et al, "Nuclear Magnetic Resonance Studies of Diffusion of
Al(27) in Aluminum and Aluminum Alloys," Acta Metallurgica, Vol. 13, July
1965, pp. 701-708.
[12] Embury, J. D. and Nicholson, R. B., "The Nucleation of Precipitates: The
System Al-Zn-Mg," Acta Metallurgica, Vol 13, April 1965, pp. 403-417.
[13] Jacobs, A. J., "Electron Microscopy and Stress Corrosion Cracking Studies on
7075 Aluminum Alloy," Research Report No: 65-10, March 1965, Rocketdyne
Division of North American Aviation, Inc., Canoga Park, Calif.
[14] Jacobs, A. J., 'The Role of Dislocations in Stress Corrosion Cracking of
7075 Aluminum Alloy," Transactions, American Society for Metals, Vol 58,
No. 4, Dec. 1965, pp. 579-600.
[15] Thomas, G. and Nutting, J., 'The Aging Characteristics of Aluminum Alloys:
Electron Microscope Studies of Alloys Based on the Aluminum-Zinc-Magnesium System," Journal of the Institute of Metals, Vol 88, 1959-60, pp.
81-90.
[16] Donnay, J. D. H., ed., Crystal Data: Determinative Tables, 2nd ed., American
Crystallographic Assn., April 1963, p. 755.
[17] Haynie, F. H. et al, "A Fundamental Investigation of the Nature of StressCorrosion Cracking in Aluminum Alloys," AFML-TR-65-258, July 1965 and
Oct. 1965, Air Force Materials Laboratory, Research and Technical Div.,
AFSC, Wright-Patterson Air Force Base, Ohio.
[18] Haynie, F. H. et al, Seventh Progress Report, AF 33(615)-1710, 15 December,
1965, Battelle Memorial Institute, Columbus, Ohio.
[19] Booth, F. F. and Tucker, G. E. G., "Statistical Distribution of Endurance in
Electrochemical Stress-Corrosion Tests," Corrosion, Vol 21, No. 5, May 1965,
p. 173.
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C. D. Beachem1 and B. F. Brown1
A Comparison of Three Precracked
Specimens for Evaluating the Susceptibility
of High-Strength Steel to Stress Corrosion
Cracking
REFERENCE: C. D. Beachem and B. F. Brown, "A Comparison of
Three Precracked Specimens for Evaluating the Susceptibility of HighStrength Steel to Stress Corrosion Cracking," Stress Corrosion Testing,
ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 31.
ABSTRACT: The susceptibility of AISI 4340 steel to stress corrosion
cracking in dilute NaCl solution is examined using the center-cracked and
part-through-crack tension specimens and the edge-cracked cantileverbeam specimen. The data are treated in terms of linear elastic fracture
mechanics. It is found that all three types of specimens can be used to determine the value of the plane-strain stress-intensity factor above which
cracking will occur in the presence of this aggressive environment.
KEY WORDS: stress corrosion cracking, high-strength steels, precracked
specimens, cracking (fracturing), corrosion
It has been conclusively shown [7]2 that significant susceptibility of a
material to stress corrosion cracking (SCC) may escape detection unless
precracked specimens are used. If cracks are not intentionally introduced,
the test is usually dependent upon the formation of stress concentrators
such as fatigue cracks or corrosion pits. Crack development at the bottom of a corrosion pit may not occur if the material does not undergo
pitting attack. The use of precracked specimens, on the other hand, more
closely approximates that critical part of a high-strength structure which
contains a crack either built into it or developed during the testing or use
of the structure. It is for this reason that the precracked SCC specimen
was adopted at the Naval Research Laboratory.
1
Head, Micro-Mechanical Metallurgy Section, Physical Metallurgy Branch,
and head, Physical Metallurgy Branch, respectively, Metallurgy Div., Naval Research Laboratory, Washington, D. C. Personal members ASTM.
2
The italic numbers in brackets refer to the list of references appended to this
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STRESS CORROSION TESTING
The precracked specimen must be properly designed, however, or it,
too, may give misleading data [1]. Several steels and titanium alloys
have been tested in different sheet and plate thicknesses, and the data
show that whereas stress corrosion cracks may propagate readily under
plane-strain conditions (in terms of linear elastic fracture mechanics),
they often do not propagate in somewhat thinner sections. If the service
application of the material will be in thick enough sections so that planestrain stress conditions will exist around the leading edge of a crack, the
test specimen must also be designed to provide plane-strain conditions
at the leading edge of its starting crack.
The use of linear elastic fracture mechanics to analyze crack extension
data permits appropriate characterization of relative propensity toward
SCC. In addition, this analysis enables one to predict the cracking behavior of structures in terms of critical flaw sizes and stresses [1,2].
This paper gives the results of SCC experiments in which AISI 4340
steel tempered at 400 F was tested in plane strain, in the form of three
types of crack propagation specimens in order to investigate further the
usefulness of the cantilever-beam specimen.
Material and Test Procedure
AISI 4340 steel (air melted) was chosen for these specimen comparisons due to its wide use and high yield strength, with the latter making
the data more amenable to linear analysis methods. Specimens 9 by 1 l/z
in. were machined from a V^-m.-thick sheet with the long dimension
oriented in the rolling direction. Holes l/z in. in diameter were drilled
through the specimen center lines 1 l/a in. from the specimen ends. Some
of these specimens were then centrally slotted at midlength and tensiontension fatigued to extend the slot to a total length of about 5/s in. in the
RW direction. The general scheme of succinctly and unambiguously
describing crack propagation directions in wrought plate and sheet may
be found in Ref 3: RW indicates the crack propagated perpendicular to
the rolling direction and parallel to the direction of the sheet width; RT
indicates the crack grew perpendicular to the rolling direction and in the
thickness direction. Some of these centrally cracked specimens were then
slit lengthwise to make edge-cracked cantilever-beam specimens about
0.71 in. in depth. Other specimens were fatigued in bending against a
sharp point located at midlength to form an RT surface crack at the
root of a surface notch on the opposite surface. All the specimens were
then austenitized together for l/z hr at 1550 F in argon, oil-quenched, and
tempered at 400 F for 1 hr, giving a hardness of Rc 51.
Containers were slipped over the central (cracked) portions of the
specimens, and the 3.5 per cent sodium chloride (NaCl) solution was
added after the specimens were loaded to the selected stresses. The salt
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FIG. 1—Specimen dimension identification and equations.
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STRESS CORROSION TESTING
Specimens tested in tension were tested in either stress rupture racks or
in a standard tensile machine, with the last being necessary for the higher
stresses. The cantilever-beam specimens were tested in the racks described in Ref 1. Briefly, these racks individually consist of a sturdy
upright support into which one end of a specimen is clamped (with the
cracked edge up), a lever arm clamped to the other end of the specimen,
and a system for adding weights to the end of the lever arm.
Specimens of all three designs were tested at various selected stresses
and the time-to-fracture recorded. If fracture did not occur after an extended period, the specimens were removed from the solution, dried,
heated overnight in air at 400 F, and broken to make the extent of crack
growth visible on the fracture surface.
Specimen dimensions are given in Fig. 1, along with the equations.
The stress-intensity factors from the cantilever-beam specimen tests were
calculated from the Kies [4] equation
where:
KI = plane-strain stress-intensity factor, ksi
m = bending moment, in-lb.,
B = specimen width (sheet thickness), in.,
D = specimen depth, in., and
a = 1 — a/D, where a is the crack depth.
The stress-intensity factors for the surface-cracked (or part-throughcrack) specimen results were calculated from the Irwin equation
where:
K! = plane-strain stress-intensity factor, ksi
a = crack depth, in.,
<r = gross stress, ksi,
(rys = the yield strength of the material, ksi, and
$ = function of the crack shape.
Values of <f> may be obtained from elliptical integral tables [5].
The stress-intensity factors from the center-cracked specimens were
calculated using the equation [6]
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BEACHEM AND BROWN ON SUSCEPTIBILITY OF HIGH-STRENGTH STEEL
35
TABLE 1—Summary of data.
CENTER-CRACKED SPECIMENS
Specimen
No.
57/^- Time, min
ksi v in-
88.5
85.0
37.1
27.2
25.0
19.1
14.6
14.2
18
16
7
8
14
17
11
18
2
4
27
54
84
399
452
464
024
545
Comments
WW, in.
B, in.
2a, in.
Load, Ib
dry
dry
broke
broke
broke
broke
broke
no break
1.473
1.499
1.503
1.493
1.499
1.500
1.508
1.473
0.123
0.125
0.126
0.127
0.126
0.125
0.127
0.123
0.750
0.598
0.587
0.617
0.600
0.617
0.640
0.740
12 150
15 000
6 500
4 630
4 330
3 150
2 460
1 970
SURFACE-CRACKED SPECIMENS
SpecKT
imen . . /'-.—
No. ksiVm.
58.0
58.0
57.0
40.0
37.0
36.0
25.0
18.0
17.2
14.7
13.4
11.2
28
23
21
26
27
30
20
29
22
21
28
19
Time, min
Comments
WW, in.
B, in.
a, in.
1C, in.
Load, Ib
39
15
86
130
451
630
3 042
1 452
26 706
dry
dry
dry
broke
broke
broke
broke
broke
broke
no break
no break
no growth
1.465
1.466
1.467
1.465
1.466
1.465
1.465
1.466
1.465
1.465
1.465
1.465
0.125
0.126
0.125
0.124
0.126
0.126
0.125
0.124
0.125
0.125
0.125
0.129
0.031
0.048
0.034
0.049
0.031
0.045
0.040
0.033
0.034
0.031
0.029
0.020
0.120
0.195
0.125
0.191
0.134
0.196
0.163
0.131
0.154
0.125
0.120
0.115
34 350
27 600
33 550
20 700
24 150
17 250
13 800
12 000
9 960
9 357
8 822
8 490
CANTILEVER-BEAM SPECIMENS
Specimen No. . . /-.—
ksi V m.
1-B.
3-B.
2-B.
4-B.
5-B.
2-A.
3-A.
10-A
1-A.
68.4
50.6
34.3
22.0
17.8
14.6
12.8
12.7
11.6
Time, min
Comments
a, in.
B, in.
D, in.
5
20
36
144
330
1 144
21 600
1 200
dry
broke
broke
broke
broke
broke
no break
no growth
no growth
0.260
0.258
0.272
0.249
0.268
0.288
0.270
0.249
0.248
0.126
0.126
0.125
0.125
0.125
0.124
0.125
0.127
0.125
0.705
0.706
0.713
0.710
0.703
0.713
0.720
0.707
0.714
M, in. -lb •
640
486
320
215
159
124
121
126
100
where a is the gross stress, a is half of the crack length, and / (2a/W)
is obtained from Paris [6].
Results
The object of using the precracked specimens in these SCC tests was
to Copyright
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STRESS CORROSION TESTING
FIG. 2—Stress corrosion cracking tests of AISI 4340 steel, using cantileverbeam specimen in 3.5 per cent Nad solution.
FIG. 3—Stress corrosion cracking tests of AISI 4340 steel, using center-cracked
specimens tested in 3.5 per cent Nad solution.
called KIscc.
SCC • This critical value of Kt is found by loading a series of
specimens at different stress intensities (KJ until the K^ level is found
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BEACHEM AND BROWN ON SUSCEPTIBILITY OF HIGH-STRENGTH STEEL
37
FIG. 4—Stress corrosion cracking tests of A1SI 4340 steel, using surfacecracked specimens in 3.5 per cent NaCl solution.
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FIG. 5—Surface-cracked
specimen results plotted relative to a behavioral
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STRESS CORROSION TESTING
each specimen type are listed in decreasing order of plane-strain stress
intensity, Kt. The Comments column indicates whether or not the crack
grew. The data are plotted in Figs. 2 through 5.
Figure 2 shows the data from the cantilever-beam specimen tests. The
KTc in this test was found to be about 68 ksi \/in. Kle is the critical
plane-strain stress intensity for fracture in the absence of the corrosive
medium—a measure of the fracture toughness of the material. The curve
in Fig. 2 is seen to become horizontal at a KIscc
JSCC of 12.7 ksi \/ni! after
about 1000 min, which means that (for this material and this specimen
design) if a specimen does not break after one day it is likely not to
break at all and may be taken out of the solution, broken, and examined
for possible crack growth.
Figure 3 gives similar data for the center-cracked specimens, only in
this case the Klc is seen to be 88 ksi \/m. The Klscc is seen to be below
14.2 ksi -\/in. Specimen 18 was loaded to this last stress intensity for
more than a month but did not break. However, a little crack growth
was apparent on the fracture surface after the specimen was dried and
broken. The curve in Fig. 3 is not horizontal after more than a month.
Figure 4 shows the data from the surface-cracked specimens. The KIe
is seen to be 58 ksi \/in., and the SCC
KI is seen to be about 12 ksi \/in.
with the curve becoming horizontal after several days or a week.
Discussion of Results
The cantilever-beam test appears to be a much more rapid test of a
material's susceptibility to SCC than either the center-cracked or the
surface-cracked specimens. Moreover, the cantilever-beam test results
enabled accurate predictions of the test results from the other two
specimen types.
For example, the 12.7 ksi -\/in. value obtained from the cantileverbeam tests was used, along with an estimated a/2C value of 0.25 in the
equation for the surface-cracked specimen, to draw a break-no break
curve in Fig. 5. The cantilever-beam test results were thus used to predict
the critical lengths (or depths since the ratio was rather constant) of
cracks for specific gross stresses. No SCC was expected below the curve,
but was predicted above the curve. The surface-cracked specimen data
are plotted on the same graph, and one can see that the times-to-fracture
(number in parentheses near data points) increase as one approaches the
curve from above and that SCC did not occur below the curve. Since
Specimen 19 had a relatively shallow crack, the break-no break curve
was plotted using its a/2C ratio (0.174) and is seen as a dashed line to
the right of the curve for a/2C values of 0.25.
Of the three values of Klc found with these three specimen types, the
value of 58 ksi \/in. is probably the most meaningful for service use.
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BEACHEM AND BROWN ON SUSCEPTIBILITY OF HIGH-STRENGTH STEEL
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fracture uncertainties, which would not influence the Klscc determinations.
The surface-cracked specimen results are seen to be within the limitations [7] of (a) crack depth less than | the specimen thickness and (b)
gross stresses less than the yield strength, as indicated on the ordinate
and abscissa, respectively. It is interesting, however, to consider a limiting, but practical, approximation using the surface crack equation. For a
long shallow crack, where it may be considered infinitely long, Eq 2 may
be rewritten as
At yield point stresses, this reduces to
and the critical crack depth (of a long crack or sharp scratch) to cause
SCC at yield point stresses in this material is seen to be 0.0007 in. The
calculated plastic zone size at the tip of such a crack is approximately
0.001 in. Therefore, the value of 0.0007 in. for the critical flaw size for
SCC does not have the indicated accuracy. The approximate value of
0.001 in. for a critical flaw depth for SCC- in this steel is real and significant, since this kind of flaw might easily be introduced in service and
escape detection.
The lesser fracture times in the cantilever-beam tests and the surfacecracked specimen tests are probably due to the more rapid increase of
/sTj with crack size for these two specimens. Thus, once a crack starts to
grow, the Kt increases rapidly, accelerating crack growth and lessening
the time-to-fracture.
Further consideration of some of the data in Table 1 provides a forceful
argument for the use of stress intensities instead of net section stresses
(uncracked area divided into the load) in the analysis of SCC data. The
surface-cracked Specimen 19 and the center-cracked Specimen 18 both
were stressed to give stress-intensity factors close to 12.7 ksi \/in. and
reacted accordingly to their environments, with the surface-cracked
specimen (slightly below the Klscc) not cracking and the center-cracked
specimen (slightly above the ATISCC) very slowly cracking. The net section
stresses for these two specimens, however, were quite different, with the
center-cracked specimen being stressed at an = 21,800 psi, and the
surface-cracked specimen stressed at <TN = 45,400 psi. Thus, while the
stress intensities of the two specimens were nearly the same and close to
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40
STRESS CORROSION TESTING
Conclusions
The cantilever-beam specimen test for the susceptibility of highstrength alloys to SCC yields the same Klscc as the surface-cracked specimen and probably the same as the center-cracked specimen. The cantilever-beam test not only is more efficient in permitting testing at high
stress intensities but also allows a quicker determination of Klscc than
the other two types of specimens.
The critical crack size for SCC is seen to be about 0.005 in. for a structure containing a surface flaw which has an a/2C value of 0.25 and is
stressed to its yield strength. Similarly, a critical flaw size of 0.017 in. is
predicted for stresses of one half the yield strength.
For a long shallow crack and yield point stress, the critical depth of a
crack for SCC is seen to be about one mil.
A cknowledgments
The authors thank the Bureau of Naval Weapons for sponsoring this
work and G. R. Irwin for his helpful and constructive criticism.
References
[7] Brown, B. F., "A New Stress-Corrosion Cracking Test for High-Strength Alloys," Materials Research & Standards, Vol. 6, No. 3, March 1966, pp. 129133.
[2] Brown, B. F. and Beachem, C. D., "A Study of the Stress Factor in Corrosion
Cracking by Use of the Pre-Cracked Cantilever Beam Specimen," Corrosion
Science, Vol 5, 1965, pp. 745-750.
[3] 'The Slow Growth and Rapid Propagation of Cracks," Second Report of a
Special ASTM Committee, Materials Research & Standards, Vol 1, No. 5,
May 1961, pp. 389-393.
[4] Kies, J. A. et al, "Fracture Testing of Weldments," Fracture Toughness Testing and Its Applications, ASTM STP 381, American Society for Testing and
Materials, Philadelphia, 1965, p. 328.
[5] Irwin, G. R., "Crack-Extension Force for a Part-Through Crack in a Plate,"
Journal of Applied Mechanics; Transactions, American Society of Mechanical
Engineers, Series E, Vol 29, No. 4, Dec. 1962, pp. 651-654.
[6] Paris, P. C. and Sih, G. C., "Stress Analysis of Cracks," Fracture Toughness
Testing and Its Applications, ASTM STP 381, American Society for Testing
and Materials, Philadelphia, 1965, pp. 30-81.
[7] Irwin, G. R., "Relation of Crack Toughness Measurements to Practical Applications," Welding Journal Supplement, Welding Research, Nov. 1962.
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E. E. Denhard, Jr.,1 andR. R. Gaugh1
Application of an Accelerated Stress
Corrosion Test to Alloy Development
REFERENCE: E. E. Denhard, Jr. and R. R. Gaugh, Application of an
Accelerated Stress Corrosion Test to Alloy Development," Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 41.
ABSTRACT: A direct-tensile type of stress corrosion test was used to
develop a corrosion resistant alloy that is relatively lean in nickel content
as compared to other alloys that are immune to cracking in boiling magnesium chloride. This new alloy also is readily hot worked and welded.
The same type of test specimen was used to evaluate the susceptibility
of Type 410 and 17-4 PH stainless steels in several heat-treated conditions
to stress cracking in environments containing hydrogen sulfide. The resistance of these alloys to cracking was found to be improved by tempering or overaging to lower strength levels, confirming earlier findings obtained on other types of tests. When heat treated to equal levels of yield
strength, 17-4 PH was somewhat more resistant to cracking than Type
410.
KEY WORDS: corrosion, stress corrosion, corrosion resistant alloys,
stainless steels, magnesium chloride, hydrogen sulfide, hot working, welding, heat treatment, yield strength
It has been customary to specify the highly alloyed nickel-base materials for equipment that must be immune to chloride stress corrosion
cracking. Typical is the extensive use of Inconel2 (15.8Cr-76Ni) piping,
tubing, fittings, and hardware in heat exchange equipment in the nuclear
industry. Although some of the less costly stainless steels display substantial resistance to stress corrosion cracking, these alloys have not met
with general favor for critical applications. Thus, there appears to be a
family of alloys with intermediate nickel content (35 to 60 per cent) which
has not been given research attention.
Copson [7]3 recognized that alloys containing 16 to 20 per cent chro1
Senior research engineer and research engineer, respectively, Research & Technology, Baltimore Laboratories, Armco Steel Corp., Baltimore, Md. Mr. Denhard
is a personal member of ASTM.
2
Registered trademark of International Nickel Co.
3
The italic numbers in brackets refer to the list of references appended to this
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42
STRESS CORROSION TESTING
mium and 45 to 50 per cent nickel seemed immune to cracking in boiling
42 per cent magnesium chloride (MgCl2). One of the serious shortcomings, however, for fully austenitic alloys of this type is the tendency for
hot cracking or microfissuring to occur upon welding. The development
of a relatively lean austenitic alloy, fully immune to chloride stress corrosion cracking, together with excellent weldability and hot workability,
was the challenge that became the subject of the present paper.
FIG. 1—Stress corrosion test specimen.
FIG. 2—Stress corrosion test equipment.
To meet the above objectives, a rapid, severe, and discriminating
stress corrosion test method was needed to screen large numbers of
experimental alloys. A notched tensile-type stress corrosion test with
boiling 42 per cent magnesium chloride as a corrodent was used for this
purpose.
During the course of the investigation, it became evident that the test
method would be useful for evaluating the effect of heat treatments and
studying
theASTM
behavior
of alloys
in Wed
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mediaEST
other
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DENHARD AND GAUGH ON ALLOY DEVELOPMENT
43
Experimental Procedure
Lean Austenitic Alloy—Resistant to Chloride Cracking
The experimental procedure consisted of making small induction
melts, forging and rolling these to bar stock, and machining stress corrosion specimens. In the case of commercially available materials, stress
corrosion specimens were prepared from small-diameter bar stock. All
specimens were laboratory annealed at 2000 F, unless noted otherwise.
The type of specimen and method of testing has been reported in an
earlier publication [2]. Briefly, the specimens were 1A in. diameter by
FIG. 3—Stress corrosion tests in boiling 42 per cent MgClz.
9l/2 in. long, containing four H-in.-radius circumferential grooves (Kt =
1.3) in the center portion, spaced at Vg-in. intervals (Fig. 1).
The test cell (Fig. 2) consisted simply of a length of glass tubing fitted
with top and bottom rubber stoppers, through which the specimen was
inserted. Heat was supplied by a small coil of resistance wire which was
wound around the bottom portion of the cell. During testing, two of the
circumferential grooves were immersed in boiling 42 per cent MgCl2
solution; the other two grooves were exposed to the vapor. Temperature
was monitored by means of a thermocouple well extending into the solution. The cell was equipped with a water-cooled condenser to maintain a
constant concentration of the test solution.
The test specimens were stressed by applying dead weights through a
lever-arm system. Various levels of stress were applied, starting at just
under the ultimate tensile strength of the alloy under study.
If both by
annealed
annealed-plus-sensitized
specimens
lasted
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44
STRESS CORROSION TESTING
Failure beyond this time invariably was caused by pitting or general
attack due to the severe environment.
Various existing commercial austenitic stainless steels and nickel-base
alloys were tested first. These results are shown graphically in Fig. 3.
The chemical analyses of the materials are shown in Table 1.
Modifications of Type 330 containing substantial manganese and other
small additions were studied next. Linnert's and Larrimore's earlier work
[31 had shown the beneficial effects of manganese in preventing microTABLE 1—Chemical composition and annealing treatment of commercial alloysGrade
Condition
Type 304
Type 310
Type3306
Incoloy 800' . .
Inconel 600C . .
2000
2000
2000
1950
1800
F,
F,
F,
F,
F,
WQ°
WQ
WQ
WQ
WQ
c
Mn
P
s
Si
Cr
Ni
0.048
0.102
0.076
0.046
0.044
1.18
1.78
0.47
0.92
0.26
0.022
0.011
0.022
0.013
0.003
0.011
0.006
0.009
0.008
0.003
0.60
0.40
0.52
0.35
0.30
19.27
27.31
15.11
19.67
15.06
9.26
24.47
35.39
31.51
75.5
0
WQ = water quench.
Laboratory heat.
c
Registered trademark International Nickel Co. ,
b
TABLE 2—Stress corrosion resistance of experimental alloys.
Heat No.
c
Mn
p
s
Time-to-Failure,
hr»
Si
Ni
Cr
Cb
Annealed
R-2759
R-2760
R-4474
R-4475
033120
R-5481
035069
a
0.014 0.49
0.076 0.47
0.045 9.87
0.11 9.80
0.096 10.00
0.022 3.30
0.030 4.88
0.61
0.52
0.26
0.23
0.008 0.008 0.98
0.004 0.008 1.23
0.011 0.005 0.73
14.88
15.11
16.16
15.72
17.82
20.19
19.22
34.31
35.39
35.6
35.7
40.49 0.15
50.32
45.48 0.14
Sensitized
360
>1000
55
>1000
>1000
330
>1000 >1000
>1000 >1000
Tested at 75,000 psi in boiling 42% MgCl2 (154 C).
fissuring in fully austenitic weldments, including alloys of Type 330 composition. This investigation was directed toward reassessing their findings
in conjunction with the stress corrosion studies reported below.
Table 2 contains the analyses and the results of stress corrosion tests
of alloys with intermediate nickel content and manganese additions.
Heats R-2759 and R-2760 represent Type 330 with low- and highcarbon content. The remarkable beneficial effect of higher carbon on
stress corrosion is evident. This is again displayed in alloys R-4474 and
R-4475, in which manganese was added to improve weldability. Up to
this point, all alloys were tested in the annealed condition.
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DENHARD AND GAUGH ON ALLOY DEVELOPMENT
45
in the sensitized condition (1250 F 1 hr air cool), all subsequent alloys
were tested in both the annealed and annealed-plus-sensitized conditions.
To meet this requirement, it was found necessary to lower the carbon and
manganese and raise the nickel contents. Heats R-5481 and 035069
represent compositions with these changes. In the case of Heat 035069,
a small columbium addition was made.
The welding characteristics of Inconel, Type 330, and manganese
modifications of intermediate nickel content alloys were studied using
notched-type specimens of 0.050-in.-thick sheet. For this work, two
coupons 2-in. square and one coupon 2 by 4 in. were arranged and fusion
butt welded (Fig. 4). This procedure provided a fabricated notch to con-
FIG. 4—Weld crack tests.
centrate stresses at the weld bead. Inert-gas tungsten-arc welding was
employed, with no filler metal added. Argon was used at the rate of 20
ft3/hr for shielding gas. The weld speed was 12 in./min using 75 amp
at 9Vz v.
Each alloy was rated on the basis of the resistance of the weld metal
deposit to cracking. The notched area was judged by the severity of
cracking at the root of the notch (heavy—3, medium—2, light—1, and
no cracking—0).
After these specimens were evaluated, longitudinal bend test coupons
were cut as shown by the dotted lines of Fig. 4. These coupons were bent
approximately 90 deg around a ?4 in. radius to emphasize any cracks
present. Evaluation of these specimens was determined by counting the
number of cracks observed. Both types of specimens were examined
under 20-power magnification. The overall crack sensitivity was the
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46
STRESS CORROSION TESTING
Tables 3 and 4 show the analyses of the alloys tested and the results
of the weld tests, respectively. All values are the average of two or more
tests.
The order of mechanical properties of the several alloys is shown hi
Table 5.
TABLE 3—Composition of weld test alloys.
Grade
Type 330
330 Cb
Inconel 600
Experimental
Experimental
Experimental
Heat No.
. . R3524
. . R-3525
R-4475
R-4566
035069
C
Mn
Si
Cr
0.062
0.060
0.032
0.11
0.058
0.030
0.67
0.60
0.22
9.80
9.62
4.88
0.50
0.72
0.25
0.23
0.56
0.73
18.15
18.02
15.77
15.72
16.26
19.22
Ni
Cb
34.44
34.69
75.87
35.7
40.46
45.58
0.16
0 14
TABLE 4^-Weld crack sensitivity.
Grade
Type 330
330 Cb
Inconel 600
Experimental
Experimental
Experimental
Heat No.
Crack
Sensitivity
at Notch
Longitudinal
Bend
Overall Crack
Sensitivity
3
3
2
1
1
1
0
4
1
0
3
7
3
0
1
R-3524
R-3525
R^475
R-4566
035069
1
TABLE 5—Mechanical properties.
Grade
Type 330 .
Incoloy 800. . .
Ultimate Tensile
Strength, psi
80 000 min
75 000 to
100 000
Experimental. . . .
85 000
Inconel 600
. . . 80 000 to
100 000
0.2% Yield
Strength, psi
40
30
50
32
30
50
000 min
000 to
000
500
000 to
000
Elongation in Reduc4 Diameters, tion of
Area, %
%
Hardness,
Rockwell B
30
30 to 50
30
90
49
30 to 50
71
65
78
65 to 85
max
88 max
Other Applications of Accelerated Stress Corrosion Test
Although the stress corrosion test referred to in this investigation was
primarily used for austenitic alloys with boiling magnesium chloride as
a corrodent, it has proven useful for other environments and purposes.
For example, this accelerated test method has been used to evaluate
the resistance of martensitic stainless steels hi various heat-treated conditions to stress cracking4 in simulated sour (that is, hydrogen sulfide
(H2S) containing) oil and gas well environments. For this purpose, the
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DENHARD AND GAUGH ON ALLOY DEVELOPMENT
47
test cell was altered as shown in Fig. 5. This very simple arrangement
uses the same type of specimen as for the MgCl2 test, and it allows testing
to be performed at room temperature in any aqueous solution. Provision
is made to allow charging the test solution with H2S (or any other gas
desired).
For this program, test specimens were machined from 1/4-in.-diameter
bars taken from commercial heats of Type 410 and 17-4 PH.5 The chem-
FIG. 5—Modified stress corrosion test equipment.
TABLE 6—Composition of alloys used for simulated sour well environmental studies.
Alloy
Type 410
17^4 PH
c
Mn
0.50
0.12
0.038 0.28
P
s
Si
Cr
Ni
Cu
Cb+Ta
0.014
0.017
0.013
0.014
0.19
0.68
12.10
16.03
0.22
4.23
3.52
0.24
ical analyses of these materials are shown in Table 6. Larger sizes could
have been used as well, but by starting with */4 -in.-diameter material, the
necessary machining was greatly minimized.
Since the resistance of martensitic stainless steels to stress cracking
in hydrogen sulfide environments has been found to improve with decreasing hardness [4], the subject alloys were evaluated hi the tempered
or overaged conditions. So heat treated, it was hoped that they would
offer good resistance to cracking, while maintaining a moderate level of
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48
STRESS CORROSION TESTING
mechanical strength. Accordingly, the 17-4 PH was heat treated to
Conditions H-1150 and H-1150 M, and the Type 410 was tempered at
1000 and 1100 F. These heat treatments were chosen to produce equal
levels of yield strength (125,000 and 100,000 psi, respectively) for the
two alloys. Details of the heat treatments are shown in Table 7.
TABLE 7—Stress corrosion resistance of Type 410 and 17-4 PH steels.
Alloy
Heat Treatment
0.2% Yield
Strength, psi
Type 410. . . . 1800 F, % hr, oil quench +
1000 F, 4 hr, air cool
125 000
Type 410 . . . 1800 F, }4 hr, oil quench +
1100 F, 4 hr, air cool
100 000
17-4 PH
1925 F ]4 hr air cool +
1150 F, 4 hr, air cool
125 000
17-4 PH
1925 F % hr air cool +
1500 F, 2 hr, air cool +
1150F, 4 hr, air cool
100 000
Applied
Stress, psi
75
50
25
75
50
25
75
50
25
75
50
25
000
000
000
000
000
000
000
000
000
000
000
000
Time-toFailure, hr"
1.3, 1.4
5.7, 6.5
11, 20
6.2, 18
27, 39
195, 845
6.1, 6.1
12, 18
83, 102
16, 22
50, 95
345, 1105
0
Tested at room temperature in 6% NaCl-0.5% acetic acid solution, saturated
with H2S.
FIG. 6—Stress corrosion tests in NaCl-acetic acid H^S solution.
These tests were performed by submerging all four machined grooves
of each specimen in a 6 per cent NaCl-0.5 per cent acetic acid aqueous
solution saturated with H2S and loading in direct tension to several stress
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levels.
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of
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steels
to
cracking
in
sour
well
environments.
The
NaCl
addition
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DENHARD AND GAUGH ON ALLOY DEVELOPMENT
49
is necessary to increase the severity of the test and produce cracking in
reasonably short lengths of time. The results of these tests, all performed
at room temperature, are listed in Table 7 and illustrated graphically in
Fig. 6.
To resolve the question of the mode of failure in this test medium
(whether stress corrosion cracking or hydrogen embrittlement), additional work is currently in progress following the method described by
Brown [5], in which the stressed specimens are polarized either positively
or negatively, with respect to their corrosion potentials, by means of a
potentiostat. This area of endeavor also will provide information on the
susceptibility of martensitic stainless steels to hydrogen embrittlement as
a function of their heat-treated condition.
Conclusions
The accelerated stress corrosion test used in this investigation was an
efficient means of evaluating the resistance of various alloys to cracking in
hot chloride environments. The results obtained on commercial alloys
(Fig. 2) closely parallel the findings of others, and of industrial experience.
An experimental austenitic alloy, nominally containing 0.03C-20O45Ni-5Mn, was developed to meet the objectives of immunity to stress
corrosion cracking and good welding characteristics.
It was noted that Type 330 could be either susceptible or fully resistant to chloride cracking, depending on the carbon and nickel contents.
For instance, at lower nickel levels, immunity to cracking was attained
by increasing the carbon content. However, this approach had the concomitant disadvantage of decreasing the resistance to intergranular attack. These opposing effects were balanced by judicious adjustment of
the carbon and nickel levels in a number of experimental alloys (not all
reported herein), to arrive at the optimum composition.
Manganese, when added in amounts of 3 to 10 per cent, greatly improved the welding characteristics of this alloy. Such additions did not
materially affect the resistance to stress corrosion cracking. During this
phase of the study, a noteworthy correlation between the hot working and
welding characteristics was observed. These alloys exhibited not only
improved weldability, but they also forged and hot rolled without the
tearing or cracking that often occurred without such additions.
Only minor modifications were necessary to adapt this stress corrosion
test for the evaluation of Type 410 and 17-4 PH in aqueous NaCl-acetic
acid-H2S solutions. The resistance to cracking of both of these alloys
was improved by tempering or overaging to lower strength levels. This
confirmed the earlier work of Bloom [4], who performed a similar study
using other types of test specimens. Also, when heat treated to equal
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50
STRESS CORROSION TESTING
A cknowledgments
The authors wish to thank J. F. Kreml, for developing much of the
stress corrosion test data contained in this paper, and R. H. Espy and
J. J. Junod, for carrying out and evaluating the welding tests. We also
wish to thank G. N. Goller, manager, Baltimore Research Laboratories;
M. E. Carruthers, director, Stainless Research; and T. F. Olt, vice president, Research and Technology, for permission to publish this paper.
References
[1] Copson, H. C. and LaQue, F. L., Corrosion Resistance of Metals and Alloys,
2nd ed., Reinhold, New York, 1963, pp. 518 and 519.
[2] Denhard, E. E., Jr., "Effect of Composition and Heat Treatment on the Stress
Corrosion Cracking of Austenitic Stainless Steel," Corrosion, Vol 16, July 1960,
pp. 359t-369t.
[3] U.S. Patent No. 2,894,833, 14 July, 1959.
[4] Bloom, F. K., "Stress Corrosion Cracking of Hardenable Stainless Steels,"
Corrosion, Vol 11, Aug. 1955, pp. 351t-361t.
[5] Brown, B. F., "Stress Corrosion Cracking and Corrosion Fatigue of High
Strength Steels," DMIC Report 210, Problems in the Load Carrying Application
of High-Strength Steels, 26-28 October 1964, pp. 91-102, Defense Metals
Information Center, Battelle Memorial Institute, Columbus, Ohio.
DISCUSSION
R. W. Staehle1 (written discussion)—Regarding the many possible
mechanisms for explaining stress corrosion cracking in a given alloy,
I think that we should reconsider the detailed processes which are operating. Generally, the microprocesses of significance are: anodic reactions,
cathodic reactions, adsorption and absorption processes (for example, of
hydrogen), film formation, bulk transport of interstitial elements, dislocation-surface interactions, and internal dislocation interactions. I think
that in most cases there is not sufficient definitive evidence to ascertain
to what extent and which of these processes are operating. We should
therefore be very careful about reaching conclusions concerning mechanistic interpretations. However, if we are ever to reach our ultimate objective of predicting cracking susceptibility as affected by imposed
environmental conditions and by metal chemistry, we must proceed
open-mindedly and in detail to determine quantitatively the nature of the
microprocesses above and determine then* significance with respect to
the specific stress corrosion process.
1
Assistant professor, Department of Metallurgical Engineering,
Center, The Ohio State University, Columbus, Ohio.
Corrosion
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H. L. Craig, Jr.,1 and H. B. Romans1
A Rapid Stress Corrosion Test for
Aluminum-Magnesium Alloys
REFERENCE: H. L. Craig, Jr., and H. B. Romans, "A Rapid Stress
Corrosion Test for Aluminum-Magnesium AUoys, Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 51.
ABSTRACT: The susceptibility of aluminum-magnesium alloys to stress
corrosion cracking may be determined using a rapid, electrolytic test
method. The stressed specimen, deflected as a beam, is made the anode in
a 3.5 per cent salt solution by impressing a current of 40 ma/in.2 Prior to
testing, a set of specimens may be heated, each one for an increased
period of time over the previous one, to develop susceptibility in a part
of the set of specimens. Test conditions have been chosen to maximize
chances of failing materials with low degrees of susceptibility, while still
not failing materials known to be nonsusceptible.
Results with two commercially produced alloys are presented. Alloy
5083 shows less tendency to develop susceptibility than alloy 5456. The
effect of cold work in accelerating the sensitivity to stress corrosion cracking is also shown.
KEY WORDS: aluminum-magnesium alloys, stress corrosion, corrosion,
aging, heat treatment, electrolysis, salt solutions
During the past ten years, wrought aluminum alloys containing 4 to
5.5 per cent magnesium have become available on a commercial basis
in the United States. Two such alloys are known as 5083 and 5456, the
prime difference between the two being a slightly higher nominal magnesium content in 5456 alloy leading to slightly better mechanical properties in the annealed condition. They were foreshadowed in their appearance here by widespread usage in Great Britain and Europe [I].2 In
addition to a high degree of corrosion resistance, these alloys are easily
welded and possess good mechanical properties. This combination of
factors places them in a position to compete with steel in such structural
uses as the fabrication of railroad cars, barges, and ships. One specific
1
Research supervisor and scientist, respectively, Department of Applied Chemistry and Mathematics, Reynolds Metals Co., Richmond, Va. Personal members
ASTM.
2
The italic numbers in brackets refer to the list of references appended to this
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52
STRESS CORROSION TESTING
example is a railroad gondola with a 271,000-lb gross weight capacity,
made possible by using a 5083 alloy body for the car.
In the literature, there are reports about the possibility of stress corrosion cracking occurring with these alloys when cold-worked, either as
a result of exposure to temperatures in the range of 150 to 400 F, or aging
at room temperature for many years [2]. Most published data have been
developed using laboratory-produced, thin sheet specimens, while the
commercial production mentioned above began with plate.
There is a considerable difference in the fabrication and resultant
metallurgical structure of sheet as compared with plate. The authors'
laboratory has primarily been occupied with the testing of plant produced
metal ranging from 0.125 to 1.25-in. gage heavy sheet and plate. This
program was conducted to verify or disprove the above mentioned reports and interpret their significance in terms of commercial production.
A conventional approach to this testing is to expose suitably prepared
specimens to alternate immersion in 3.5 per cent sodium chloride solution, while they are stressed to 75 per cent of their yield strength. A
device for accomplishing this objective is known as a "stress rack." The
design for this rack was adapted from plans supplied by the National
Bureau of Standards. The preparation of tension-test-type specimens for
this method is comparatively expensive. The procedure also entails the
determination of mechanical properties. When a large number of metallurgical variables are to be investigated, it is not feasible to test more
than two or three specimens per variation. This is an acknowledged
weakness of the test method, as there is a widespread scatter inherent in
stress corrosion cracking data. If two specimens last a specified time in
a given test, there is really no assurance that if a third one were tested,
it would not fail.
Furthermore, the corrosion resistance of the aluminum-magnesium
alloys in salt solution is so high that it requires long-exposure times for
a significant amount of attack to occur. Even then, it is not always clear
whether a failure is due to pitting attack which reduced the cross section
or to true stress corrosion cracking.
To overcome these difficulties, the following criteria were used to develop a rapid laboratory stress corrosion test:
1. In testing closely related materials, it should be sufficient to stress
them all to the same value, as inaccuracies in setting stress levels mask
small differences in mechanical properties.
2. A large number of specimens of a given metal variation should be
tested, including both specimens that fail and specimens that do not fail
in the test.
3. Failure, when it occurs, should happen in a short time, so that
pitting is not a significant factor in assessing the failure.
4. Thebytest
should
berights
ablereserved);
to pickWed
outDec
materials
with
low degree of
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CRAIG AND ROMANS ON ALUMINUM-MAGNESIUM ALLOYS
53
Degree of Susceptibility
In alloy systems that are liable to stress corrosion cracking, various
degrees of susceptibility are found. These degrees may be defined with
respect to three terms of reference: (1) intensity of corrosion, (2) stress
level, and (3) metallurgical factors.
In the absence of corrosion attack, stress corrosion cracking cannot
occur. On the other hand, cracking often occurs when very little general
corrosion has taken place rather than when the metal is severely pitted
or otherwise corroded. In long-term tests, failure of a stressed specimen
may occur due to reduction in cross section by pitting attack. It is difficult
then to decide if the specimen represents a low degree of susceptibility
or not. In the case of early failures, a high degree is easily assessed. For
purposes of this present test, the corrosion attack is made intense so that
materials with "low degrees of susceptibility" as determined in less corrosive environments fail quickly.
The same attitude is adopted toward the stress level: "no stress" should
produce "no failures," and increasing stress levels should result in shorter
and shorter lives. In testing similar materials with but slightly different
mechanical properties, a high stress level above the yielding point tends
to mask these differences, and it also eliminates the need for determining
the yield strength for each specimen.
It is possible to vary the metallurgical condition of some alloys to
produce a range of susceptibilities to stress cracking. This is the procedure
which was adopted for purposes of this test.
Sensitization
In the aluminum-magnesium alloy system, an important variable is
the age of the metal. It is a well-documented fact that freshly prepared
alloys with all the magnesium in supersaturated solution are not susceptible to cracking. However, prolonged aging at ambient temperatures or
shorter times at slightly elevated temperatures (under the solid solubility
curve in the phase diagram) result in precipitation of magnesium in the
form of an intermetallic compound with aluminum. If this precipitate
is distributed unfavorably, by being concentrated mainly in the grain
boundaries, the structure is susceptible to stress corrosion cracking.
This aging process is called sensitizing. This precipitation phenomenon
is not completely understood, nor is the relationship between precipitation and stress corrosion cracking entirely explained on the basis of any
simple mechanism. The role of dislocations has not been studied yet for
this system.
It is easy to demonstrate in a simple binary alloy a process that might
be called the "classical example" of precipitation at grain boundaries.
This condition results in stress corrosion cracking by intergranular attack
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STRESS CORROSION TESTING
across the specimen. However, it is more difficult, if not impossible, to
correlate microstructure of complex alloys with their corrosion behavior. For this reason, it is necessary to test a large set of specimens
that have been subjected to a variety of sensitizing treatments, or aging
practices.
A degree of susceptibility may then be defined as the time required
to produce a susceptible structure from one that is nonsusceptible, by
FIG. 1—Effect of time at temperature on susceptibility to stress corrosion
cracking—NP6-Y4H.
either aging at ambient or slightly elevated temperatures, for example, a
material that fails in a stress corrosion test after being heated for one
week at 212 F is more susceptible than one that does not fail until after
it has been heated for one month at 212 F.
Outline of Test
Thus, a stress corrosion test for the aluminum-magnesium alloys consists of the following steps:
1. Heating a set of specimens at various times in a temperature region
calculated to produce susceptibility.
2. Stressing the specimens to a high tensile stress value.
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3. Exposing the specimens to a corrosive environment.
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CRAIG AND ROMANS ON ALUMINUM-MAGNESIUM ALLOYS
55
Example
Due to the favorable experience reported for British alloys, a lA-m.
gage plate was obtained of a 5 per cent magnesium alloy under the British
Standard Specification 1470 for Wrought Materials of Alloy NP6 - 1A H.
The chemical analysis is:
Si
Fe
Cu
Mn
Mg
Cr
0.11
0.14
0.03
0.23
4.5
0.01
Ni
Zn
Ti
<0.01
0.07
0.01
FIG. 2—Specimen position'in parent plate.
The mechanical properties are:
Tensile
Strength, psi
Yield
Strength, psi
Elongation,
%
43 200
43 700
32 500
31 700
19.7
18.8
Longitudinal
Transverse .
Results of stress corrosion testing using direct loading tension test
specimens, alternately immersed in a 3.5 per cent salt solution, gave the
following results:
Condition
As received
Specimens first heated one week at 212 F. . . .
Specimens first heated one month at 212 F . .
Days to Failure, Duplicate Specimens
942, one specimen not failed
157, 173
2,4
The alternate-immersion cycle was 10 min immersed in each hour.
Specimens of the same plate were taken according to the procedure
outlined in detail below. Figure 1 gives the results of 92 tests, each with
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STRESS CORROSION TESTING
perature. Note that susceptibility to cracking is rather easily induced in
this alloy as revealed by this test through exposure to slightly elevated
temperatures. The test itself lasts but one day, and the results may be
compared with the data given previously, where it took an average of
165 days for the specimens sensitized for one week at 212 F to fail in
the alternate-immersion test. Two other related points should be noted:
whereas the specimen tested "as received" failed in the alternate-immersion test, another specimen taken from the same parent plate passed this
test. Examination indicated that the as-received material failed due to
a reduction of cross-sectional area by pitting. Then also, it should be
noted that an extrapolation of an "imaginary" line between nonfailures
and failures obtained by heating specimens does not follow a straight
line, particularly at the ambient temperatures.
FIG. 3—Specimen holder.
The main conclusion to be drawn from this example is that the test
used causes failures of specimens with very low degrees of susceptibility,
as determined by other tests, but that it does not cause failures in nonsusceptible material.
Experimental Procedure
1. Prepare specimens with machined (milled) transverse surfaces,
3.00 by 0.064 in. by the plate gage. See Fig. 2 for position of specimen
with respect to rolling direction. If the gage exceeds 0.75 in., cut the specimens down to a width of 0.75 in. or less, retaining at least one rolled
surface.
2. Smooth the 3.00-in.-long edges with a fine file or sandpaper.
3. Stencil or otherwise mark specimen on one end to identify it.
4. Sensitize according to the following schedule: (a) as received, (ft)
one
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at Wed
212 Dec
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CRAIG AND ROMANS ON ALUMINUM-MAGNESIUM ALLOYS
57
5. Degrease sensitized specimens with acetone. Handle by ends from
this point on.
6. Insert specimen in holder (Fig. 3) and tighten screw to hold in
place. Do not tighten enough to deflect specimen.
7. Place holder in a clamp; bring dial gage up until the point touches
specimen.
8. Adjust for zero displacement of specimen. Then turn down on
screw until the necessary displacement is obtained. See the Discussion
section below. A value of 8 X 10~3 in./in. for the nominal outer fiber
strain has been used in this work.
9. Place holder in a 600-inl beaker containing 340-ml 3.5 per cent
salt solution. Attach an alligator clip to the end of the specimen. (This
amount of solution in a 600-ml beaker will immerse 2 in. of specimen
in solution. This length is used hi the area calculation.)
10. Insert a piece of metal of the same alloy composition opposite the
specimen to serve as a cathode.
11. Make electrical connection to a dry cell, storage battery, or rectifier through a variable potentiometer.
12. Adjust the current so that a value of 40 ma/in.2 of immersed
specimen surface flows, with the specimen acting as the anode. (A drop
of phenolphthaiein in solution will turn solution around cathode pink.)
NOTE: For specimens wider than 0.250 in., it is necessary to reduce the current density to 25 ma/in.2, due to shielding of the reverse side of the specimen by
the holder. The test is not very sensitive to changes in the current density in
this range.
13. Leave on test for 22 to 24 hr, or until a crack appears in the
specimen, whichever is sooner.
14. After the test period is over, remove the specimen holder and
examine the specimen in place. A failure has occurred if a crack is visible.
If no crack is visible, slowly deflect one end until it touches the opposite
side of holder. In this manner, any incipient or microscopic cracks may
become visible.
15. Report specimens: If the specimen fails as in Step 14, designate
as F. If the specimen does not fail, designate as N.
Discussion
Specimen Design
One of the most difficult problems in stress corrosion testing is sampling
the metal. Especially in heavy plate, the structure may vary through the
thickness. Distinct differences in susceptibility are found depending on
the orientation of the grain structure with respect to the direction of
tensile
stress.
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58
STRESS CORROSION TESTING
form cross-sectional area was selected [3]. It was then determined that
the most sensitive direction with respect to stress corrosion was in the
transverse plane, perpendicular to the direction of working the metal
(Fig. 2). In addition to this feature of maximum sensitivity, machining
specimens is a simple operation. This procedure also has the added benefit of testing the entire cross section, up to 0.75-in. gage plate. Heavier
plate may be milled or sawed to reduce its thickness in half and then
specimens taken. In the case of 3-in. or thicker plate, a short transverse
specimen can be taken. At the other end of the gage scale, sheet as light
as 0.125 in. has been tested successfully. Gages lighter than this are tested
with the rolled surface being stressed. Allowances should be made for
TABLE 1—Variation in response to stress corrosion test
depending on plane of stress.
Stressed Surface*
Rolled Surface
(transverse)
0
Condition
As received
Sensitized at 212 F, days
3.5... .
7
14
28
56
Perpendicular to
Rolled Surface
Specimen 1
Specimen 2
Specimen 1
Specimen 2
N
N
N
N
N
F
F
F
not tested
N
N
N
N
not tested
F
F
F
F
F
N
N
F
F
F
a
Specimens from 5456 alloy sheet, 0.125-in. gage. Specimens 1 and 2 are not
duplicates, but represent different fabricating practices, which were expected to
respond with different degrees of susceptibility.
6
N — no failure; F = failure in test.
reduced sensitivity when a comparison is made between these tests and
those made of a transverse section.
Table 1 gives two sets of data collected to show that the transverse
direction is more sensitive than the rolled surface to stress corrosion
cracking.
Specimen Holder Design
A 1 Vz-in.-diameter rigid poly (vinyl chloride) pipe is used for the
holder body. For purposes of illustration, however, Fig. 3 shows a model
of the holder with a body of clear plastic tubing. After some experimentation with four-point loading and several methods of three-point loading,
the best deflector was found to be a screw machined from Vs-in.-diameter
nylon rod. Studies of stress distribution using the Photostress technique
showed an acceptable ratio of E\, the strain perpendicular to the end
supports, to E2, the strain parallel to the supports. The end supports
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CRAIG AND ROMANS ON ALUMINUM-MAGNESIUM ALLOYS
59
TABLE 2—Stress corrosion failure times in four test media, as a function of strain.
Strain,
in./in.
0.002
0.025
0.003
0.004
0.005
0.006
0.008
Alternate Immersion
3.5% NaCl
F (26 days)
F (13 days)
F (7 days)
Chromic Acid-6
Salt Solution
F
F
F
F
(13 days)
(25 days)
(26 days)
(11 days)
F (1 day)
Constant Immersion-ImpressedCunent0
3.5% NaCl
Chromic AcidSalt Solution
N (15 hr)
F (280 min)
F (90 min)
N (20 hr)
F (14 hr)
F (13 hr)
F (20 min)
F (120 min)
0
A value of 40 ma/in.2 is used.
36 g/1 chromic acid (tech) plus 30 g/1 potassium dichromate (tech) plus 3
g/1 sodium chloride.
6
is drilled in the pipe wall opposite the point of loading for the dial gage
point (not shown in Fig. 3).
This holder is supported in a suitable clamp, the specimen inserted,
and the deflector screw lowered until it just touches the specimen. The
dial is raised until it touches the specimen opposite the deflector screw.
The screw is then lowered until the desired amount of deflection is
reached. This is calculated from the formula for the deflection of a simple beam, d = EL2/6t, where d is the deflection (in.); E, the nominal
outer fiber strain (in./in.), L, the distance between end supports (in.);
and t, the thickness of the specimen (in.). Two hundred holders were
drilled from a jig, ten holders measured, and an average for L calculated
to be 2.48 in. To simplify calculations, the constants in the above equation are collected together into a "deflection factor," which gives the deflection for 0.001 in./in. strain. Using the above figure for L, values of
the deflection factor are:
Thickness, in.
Deflection Factor
0.050
0.055
0.060
0.065
0.070
0.0205
0.0187
0.0171
0.0158
0.0147
0.125
0.0082
Most specimens tested have been 0.064-in. thick, but specimens between 0.050 and 0.125 in. thick have been satisfactorily tested. To find
the desired deflection, the thickness of the specimen is measured, the
corresponding deflection factor is found from the table and multiplied
by the desired strain value. A value of 8 X 10~3 in./in. has been used
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60
STRESS CORROSION TESTING
deflected 0.1264 in. (8 X 0.0158). This is well above the yield point,
and, consequently, the strain value quoted is not accurate as the simple
beam formula does not hold; however, the value is reproducible and convenient for testing of alloys with yield strengths in the range of 30,000
to 40,000 psi.
FIG. 4—Allay 5083 sensitized one week at 297 F, HNO, etch (XlOO).
A comparison was made of constant deflection using this holder and
constant load, using an adaptation of the holder, loading the specimen
by means of lead weights placed on a pan riding on a rod hi place of the
screw. The material tested in this case was an aluminum-7 per cent magnesium sheet, 0.050-in. gage, cold rolled 36 per cent, and aged for three
years at ambient laboratory temperatures before this testing. In this
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CRAIG AND ROMANS ON ALUMINUM-MAGNESIUM ALLOYS
Strain, in./in.
0.001.
0.002.
0.003.
0.005.
Constant Load
N
F
not tested
not tested
61
Constant Deflection
not tested
N
N
F
FIG. 5—Alloy 5083.
As expected, the condition of constant load lowered the strain required for failure.
Corrosion Medium
Two variations in two solutions were tried before selecting the test:
alternate immersion in a solution and constant immersion with an impressed current, using the specimen as the anode [4]. In this instance,
a very susceptible material was used, whose susceptibility developed by
natural aging at ambient temperature. The metal was a binary aluminum7 per cent magnesium alloy sheet, 0.050-in. gage, which had been cold
rolled 22 per cent, and aged two years before these tests. The time-toCopyright
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levelWed
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TableEST
2. A
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62
STRESS CORROSION TESTING
of the material, solution heat treated 30 min at 930 F and quenched in
cold water before the corrosion test, did not fail when strained at a value
of 0.008 in./in. using impressed current in a 3.5 per cent salt solution.
This behavior was due to a resolution of precipitate in addition to the
effect of annealing and consequent loss of properties.
This series of tests shows that an impressed current greatly accelerates
the time-to-failure over an alternate-immersion test at equivalent strain
values. The interesting thing is that the effect of a chromic acid-salt
solution is reversed by the impressed current: in alternate immersion
this latter solution is more aggressive than plain salt, whereas under the
influence of an impressed current, the chromic acid inhibits cracking.
These results indicate that testing with an impressed current is definitive of stress corrosion cracking, in spite of its apparent severity, since
low strain values do not cause failures in susceptible material, nor does
a high strain value cause failure in a nonsusceptible material.
The function of the impressed current is to overcome the inherently
good corrosion resistance of the aluminum-magnesium alloys. The dissolution appears to follow the grain boundaries, which are the site of the
stress corrosion crack. In adopting this means of shortening the test
time, the philosophy is analogous to that proposed by Brown [5] in his
precracked cantilever-beam test. Whereas Brown's test is limited to
materials with relatively high yield strengths, the present test only requires that a properly selective corrosive environment be found for the
material to be tested. There is no limitation on the mechanical properties
of the test alloys.
Type of Failure
Metallographic examination of a typical failure is shown in Fig. 4.
The characteristic intergranular nature of the failure is evident. Only
slight pitting has occurred during this exposure which was 30 min. A
similar specimen left on test for the complete 22 hr was pitted as well
as cracked. In susceptible metal, the cracks appear to form in preference
to pits, while in nonsusceptible material uniform dissolution occurs by
general, shallow pitting, or even fairly uniform etching of grain faces.
This behavior of specimens in the test fulfills the third criterion, which is
to allow susceptibility to stress cracking to be measured without the
interference by reduction in cross-sectional area due to pitting attack.
Results with Commercially Produced Plate
The usefulness of this method can be demonstrated by the results obtained in comparing two alloys, in the form of 0.250-in. gage plate, mill
fabricated. In addition to the alloy variable, the effect of cold work was
studied: four levels—0, 5, 15, and 40 per cent cold reduction.
Figure 5 shows alloy 5083, tested after sensitizing as described above
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at
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CRAIG AND ROMANS ON ALUMINUM-MAGNESIUM ALLOYS
63
after about 1100 days ambient temperature aging. The failed condition is
represented by an "X," the unfailed result by a dot. A few anomalous
results are evident, but the majority of the tests follow a logical pattern,
showing the development of a susceptible condition after a period of
time, depending on the specific temperature.
Figure 6 shows the influence of increasing the magnesium content on
this rate of precipitation. Alloy 5456 has, on the average, 0.75 per cent
more magnesium (5.25 versus 4.5 per cent nominal). The effect of cold
work was changed by the higher magnesium, as a comparison of the two
FIG. 6—Alloy 5456.
patterns of behavior at 15 per cent cold reduction readily shows. The
same trend is evident at 5 and 0 per cent, also.
Conclusions
A rapid stress corrosion test for aluminum-magnesium alloys has been
developed consisting of exposing a deflected-beam specimen, 3 by 0.064in. by the gage of the parent plate, to a 3.5 per cent salt solution under
2
anCopyright
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64
STRESS CORROSION TESTING
specimens is heated at a temperature, 212 F, for periods of time calculated to produce some failures in the test.
This test lasts but one day, in addition to the time required to sensitize
the material to stress corrosion cracking.
The conditions of the test have been chosen to maximize the chances
of failure of materials with low degrees of susceptibility, while still precluding failure of nonsusceptible material.
A cknowledgment
The authors gratefully acknowledge the contributions of the staff of
the Metallurgical Research Division in evolving this procedure. B. A.
Niemeier aided in the development of the specimen holder, and J. S.
Prestley contributed the metallographic studies. R. G. Connell, Jr. determined the effect on constant load.
Numerous persons have kindly discussed these methods with the authors and have aided in their development. They are: S. J. Sansonetti,
director of research technology and applied science, and L. E. Householder, chief metallurgist, both of Reynolds Metals Co.; F. A. Champion,
British Aluminum Co.; Fred Reinhart, National Bureau of Standards,
and especially H. P. Godard, Aluminum Laboratories, Ltd., who provided the initial suggestion of using an impressed current.
References
[1] Champion, F. A., 'The Interactions of Static Stress and Corrosion with
Aluminium Alloys," Journal of the Institute of Metals, Vol. 83, 1954-1955,
p. 390.
[2] Dix, E. H., Jr., Anderson, W. A., and Shumaker, M. B., "Influence of Service
Temperature on the Resistance of Wrought Aluminum-Magnesium Alloys to
Corrosion," Corrosion, Vol. 15, 1959, pp. 55t-62t.
[3] Fraser, J. P., Eldridge, G. G., and Treseder, R. S., "Laboratory and Field Methods for Quantitative Study of Sulfide Corrosion Cracking," Corrosion, Vol. 14,
1958, pp. 517t-523t.
[4] Booth, F. F. and Godard, H. P., "An Anodic Stress-Corrosion Test for Aluminum-Magnesium Alloys," Proceedings, First International Congress on Metallic
Corrosion, London, 1961.
[5] Brown, B. F., "A New Stress-Corrosion Cracking Test Procedure for HighStrength Alloys," Materials Research & Standards, Vol. 6, No. 3, March
1966, pp. 129-133.
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DISCUSSION ON ALUMINUM-MAGNESIUM ALLOYS
65
DISCUSSION
A. R. C. Westwood1 (written discussion)—What is the significance of
the 22-hr test period adopted in this work? Could equivalent data be obtained in, say, a 4-hr test period?
It is possible that the role of copper ions in the environment is to act
as step-poisons, changing corrosion from general to specific and causing
tunnelling or pitting. The copper may be present as simple ions or complexes. A discussion of such possibilities may be found in Corrosion
Science,2 and on p. 228 of this symposium.
H. L. Craig, Jr., and H. B. Romans (authors)—A 22-hr test period
was selected on the basis of obtaining the maximum number of failures
without causing failure of nonsusceptible material. Equivalent data are
not obtained in 4 hr, as some failures occur after that time limit.
We have not investigated the role of copper ions in this test, as the
alloy contains less than 0.1 per cent, and either distilled or deionized
water is used to make up the solutions. Later work with high-purity
aluminum has shown the influence of trace amounts of copper during
pitting corrosion, but its role in stress corrosion cracking has not yet
been defined. See the paper by Romans and Craig on p. 363 of this
symposium.
M. B. Shumaker3 (written discussion)—This type of test has also been
of considerable interest to us at Alcoa. Of particular concern has been
the possibility of the test failing material having a microstructure only
slightly outlined at the grain boundaries with aluminum-magnesium
precipitate and having performed well in service. What evidence do the
authors have as to whether this does or does not occur with their tests?
H. L. Craig, Jr., and H. B. Romans (authors)—The correlation between this test and atmospheric stress corrosion testing is being presented in paper No. 27 at the Symposium on Atmospheric Testing, 25-30
June, 1967. Specimens which fail this test, in 22 hr, have failed after
five years in the atmosphere.
There have been no reports of service failure of 5083 alloy, except
where inadvertent exposure to elevated temperatures has occurred.
1
Associate director, Research Institute for Advanced Studies, Martin Co., Baltimore, Md.
2
Westwood, A. R. C., Corrosion Science, 1966.
3
Research engineer, Chemical Metallurgy Div., Alcoa Research Laboratories,
New Kensington, Pa.
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/. H. Mulherin1
Influence of Environment on Crack
Propagation Characteristics of HighStrength Aluminum Alloys
REFERENCE: J. H. Mulherin, "Influence of Environment on Crack
Propagation Characteristics of High-Strength Aluminum Alloys," Stress
Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 66.
ABSTRACT: The stress corrosion behavior of four alloys of aluminum
has been investigated. Using cantilever-loaded fatigue-cracked specimens,
the crack propagation characteristics of 7178-T6, 2014-T6, 2024-T351,
and 7075-T6 aluminum alloys were determined. The results are interpreted
in terms of fracture mechanics parameters. Where susceptibility to stress
corrosion exists, the rate of crack propagation is dependent upon stress intensity and environment. A method is also presented to obtain threshold
values of stress intensity for stress corrosion attack with minimal experimental effort. The aqueous environments used were immersion in distilled water, immersion in 3Vz per cent sodium chloride, and alternate immersion in salt solution and air. The cantilever technique is evaluated and
the results compared to other testing methods.
KEY WORDS: corrosion, stress corrosion, cracking, crack propagation,
fracture mechanics, aluminum alloys
Nomenclature
a
ag
o0
B
BN
d
K
Klc
K0
L
M
m
Crack depth, in.
Crack extension in stress corrosion environment, in.
Initial value of a, in.
Gross specimen thickness, in.
Net specimen thickness, in.
Gross specimen depth, in.
Stress-intensity factor, psi • in.1/2
Plane-strain fracture toughness, psi-in.1/2
Initial value of K, psi • in.1/2
Half-span length for three-point-loaded notch bend specimens, in.
Applied bending moment, in • Ib
Exponent in the correction factor for the side notches
1
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Pa.
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
67
Force, Ib
Failure time, min
Poisson's ratio
P
//
v
In the usual testing method for the determination of stress corrosion
susceptibility of aluminum alloys, a ring, bent beam, or U-bend specimen
is used as discussed by Sprowls and Brown [I].2 Since these specimens
are carefully fabricated to minimize stress raisers, they can be generally
characterized as "nondefect" specimens. In the experimental procedure,
TABLE 1—Nominal chemical compositions of the aluminum alloys investigated
(weight per cent}.
Alloying Elemc:nt
Cu
Mg
Mn
Si
Zn
Cr
. .
2014
2024
7075
7178
44
4.5
1.5
2.5
2.0
5.5
0.3
6.8
0.3
0.4
1.5
0.6
0.8
0.8
2.7
TABLE 2—Mechanical and fracture properties.
Alloy Designation
Temper
Plate thickness, in
Direction0.
ffu i t psi
(fya , psi
Elongation,6 %
Average Kjc , psi- in.1'2
2014
T-6
y.
LT
70 400
63 300
12 9
24 000
2024
T-351
1%
ST
57 600
46 700
37
22 100C
7075
7075
7178
T-6
T-6
1%
ST
79 700
69 100
4.0
23 200
T-6
y.
LT
84 100
74 100
10 5
20 600
1/4
^
LT
91 700
85 300
11.3
18 500
0
LT denotes the long-transverse direction and ST the short-transverse direction.
6
Two-inch gage length in the %-in. plate and ^-in. in the 1^-in. plate.
c
Lower bound value.
the stressed specimens are exposed to either a natural atmospheric or
laboratory environment. Failure is considered to occur when a crack
apparent to the naked eye has developed. This test method concerns
the period for crack initiation plus the time period for macroscopic
crack development.
In contrast with the conventional approach, this paper is concerned
with the crack propagation characteristics of the material. There are advantages in using this approach. First, it is generally recognized that
with the usual testing technique, crack initiation is sensitive to the surface
condition of the test specimen. By starting with a precracked specimen,
2
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STRESS CORROSION TESTING
FIG. 1—Specimen configuration, nominal dimensions, and three-point loading
locations.
SHORT-TRANSVERSE SPECIMEN
LONG-TRANSVERSE SPECIMEN
FIG. 2—Schematic showing the location of the major specimen notch in relation to the rolling direction.
this area of uncertainty is minimized. Second, it has been demonstrated
that a titanium alloy was immune in a usual stress corrosion test using
nondefect type specimens. However, in the presence of a crack, this
material was very susceptible to stress corrosion cracking [2]. In a real
structure, cracks can originate from a number of phenomena, for example, welding operations or fatigue. Therefore, the behavior hi the presence
ofCopyright
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tion
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Third,
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
69
analysis is available, fracture mechanics provides a well-developed
analytical approach by which the stress situation can be described [3, 4].
Finally, by eliminating the crack initiation phase in the study, possible
additional insight into the mechanism of crack growth may be obtained.
Experimentally, notched and fatigue-cracked bend-bar specimens were
used. These specimens were cantilever loaded as in similar work reported
by Brown [2]. A stress-intensity factor, K, from fracture mechanics is
used to describe the macroscopic stress field at the crack front. The crack
propagation characteristics of 2014, 2024, 7075, and 7178 aluminum
alloys were studied in several laboratory aqueous environments.
Experimental Procedure
Materials
The nominal chemical compositions of the four alloys used in this
investigation, 2014, 2024, 1075, and 7178, are shown in Table 1. The
temper designations, engineering mechanical properties, and fracture
data in the direction investigated are summarized in Table 2.
Test Method
The specimens used were side-notched bend bars as shown in Fig. 1.
Side notches were introduced in the plane of the anticipated crack propagation to approach plane-strain conditions for fracture toughness tests
conducted under ambient laboratory conditions. For the sake of standardization, side notches were used through the entire experimental program. In the short-transverse direction, the L dimension was increased
to 2 in. by the use of adapters for the fracture toughness tests. The orientation of the notches in relation to the plate is shown in Fig. 2.
Fracture toughness data were obtained using three-point loading on a
universal testing machine. Electrical resistance strain gages bonded to
the crack tip area were used as the deformation sensing device to obtain
load-deformation curves on an X-Y recorder. For the environmental
tests, the method of loading was with a cantilever arrangement. These
tests were run with a constant applied load.
The environmental containers consisted of flexible polyethylene cups
of approximately 150 cm3 volume. Although the liquid medium was
not circulated, the entire notched area of the specimen was exposed to
the environment. The environments used were constant immersion in
distilled water, constant immersion in 3l/2 per cent aqueous solution of
sodium chloride (NaCl), and an alternate-immersion cycle consisting of
10 min in 3J/2 per cent NaCl solution and 50 min in air. During the
environmental tests, evaporation losses were restored with distilled water.
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STRESS CORROSION TESTING
the alternate-immersion cycle. The environmental test facility has been
previously described in Ref 5.
Analytical Procedure
In this paper, the stress is described by means of a stress-intensity
parameter, K, from fracture mechanics. K is a parametric description
of the magnitude of the stress field in the vicinity of the crack front.
As a measure of the material response to this stress field, either a KIO
or K0 designation is used. Klc represents the plane-strain fracture toughness obtained in a fast fracture test, and K0 represents the initial stress
intensity applied to environmentally exposed specimens.
The three-point loading equation of Srawley and Brown [6] was
used as the analytical basis for the Klc computation. This equation,
however, does not include the influence of side notches. Freed and
Krafft [7] proposed a correction factor to account for the side notches
through a ratio of the unnotched-to-notched specimen breadths, B and
BN , respectively. The modified equation used as the basis for the calculation including the correction factor is
Since the cantilever-loaded specimens have a shear stress, it appeared
reasonable to use a three-point loading equation rather than a fourpoint loading equation for calculating the K0 values. However, the
results are substantially the same. The equation can be modified to
include a moment term, M, with consideration that M = % PL. Therefore, the equation used for calculating K0 including the side notch
correction factor is
Experimental Results and Discussion
7075-T6 Aluminum Alloy —Short-Transverse Direction
For other materials such as steel and titanium, the relationship
between the initial applied stress intensity, K0 , and the time-to-failure,
// , has been presented in the literature [2,5]. Many of these data are
characterized by extremely short failure times at high K0 levels. At
lower K0 levels an "apparent" endurance limit is reached as indicated
by an abrupt lengthening of the time-to-failure. This type of data was
obtained for 7075 alloy in the solution treated and artificially aged conCopyright
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
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The plane of anticipated crack propagation in these specimens is parallel
to the plate surface at midthickness (Fig. 2).
The variation in the failure time as a function of the initial stress
intensity for the three environments is shown in Fig. 3. For a constant
immersion in 3^ per cent NaCl solution, the failure times gradually
increased with decreasing initial stress intensity. No abrupt change was
FIG. 3 —Delayed failure characteristics of 7075-T6 aluminum alloy in the shorttransverse direction.
found in the relationship. In distilled water, the length of time necessary
for specimen fracture was substantially longer than in the salt solution.
However, over the limited stress-intensity range used in the distilled
water experiments, the curves were of the same general shape. For the
alternate-immersion cycle, in a range of K0 values between 11,500 and
15,700 psi-in.1/2, the failure times closely approximated the data obtained under constant immersion.
Determinations were also made of the crack extensions which occurred
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STRESS CORROSION TESTING
FIG. 4—Crack propagation of 7075-T6 in the short-transverse direction for
several exposure times. Also shown is the growth necessary to cause catastrophic
fracture of the specimen. Arrows indicate the anticipated Ko at which failures of the
specimens occur at 4000, 6000, and 8000 min (from Fig. 3).
FIG. 5—Schematic showing the relationship between K and a for a constant
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
73
periods. The resulting data (Fig. 4) show the relationship between the
three variables, initial stress intensity, crack extensions, and exposure
time.
With an increment of crack extension, an incremental increase will
occur in the stress intensity per Eq 2. The generalized relationship is
shown in Fig. 5. With increases in a, related increases will occur in K
until the catastrophic value of K is attained, Klc. From the schematic
Fig. 5, it can be seen that a change in the initial crack length will cause a
variation in the rate of change of K with a. This variation will also occur
with a change in some other specimen dimensions such as the depth.
Therefore, crack extension data as a function of exposure time and
Ko shown in Fig. 4 are sensitive to the specimen configuration to a greater
degree than the time-to-failure. A test to determine the time-to-failure
FIG. 6—Photograph of fracture surfaces of the 7075-T6 aluminum specimens
oriented in the short-transverse direction showing the crack propagation which occurred during 6000-min exposure in an environment of 3Y2 per cent NaCl
solution. From left to right K = 3200; 3300; 6100; 9200; 9300; 12,000; and 12,000
psi-in.1/2
would cover a much greater range of both a and K and thus would be
less sensitive to initial crack depth.
Obviously, a variation in the rate of change of K as a function of the
initial crack depth can also be substantial among the various types of
specimens with different stress analyses. Beachem and Brown3 have
reported the influence of specimen configuration on the tf parameter.
Also shown on Fig. 4 is the extension of the crack necessary to cause
catastrophic fracture as a function of the original stress intensity. These
data were obtained from the specimens exposed to a constant immersion
of 3^ per cent NaCl reported as tf data on Fig. 3. The arrows along
the catastrophic fracture curve show the K0 level at which fracture of
the specimen would be anticipated to occur after exposure for 4000,
6000, and 8000 min. These values were obtained from Fig. 3. Extrapolation of the exposure curves to these values shows substantial agreement.
The successful use of this crack extension technique is contingent
upon the delineation and measurement of the mechanically induced
fatigue crack, the stress corrosion crack, and the catastrophic fracture
3
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FIG. 7—Composite photomicrograph of a stress corrosion crack in a short-transverse specimen of 7Q75-T6 aluminum alloy showing intergmnular
propagation. Reduced 31 per cent in reproduction. Arrow indicates approximate termination of the fatigue crack. Keller's etch (X/00).
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
75
crack. Immediately after exposure, these specimens were broken by an
impact blow at right angles to the direction of anticipated crack propagation. The fracture surface can then be examined for the different crack
propagation modes. For the 7075-T6 a definitive difference in macro
appearance existed. Figure 6 is a photomacrograph showing the difference in crack extension as a function of stress intensity. A photomicrograph is shown in Fig. 7 of a cross section of an unbroken exposed
FIG. 8—Comparison between the delayed failure characteristics of 2024-T351
and 7075-T6 aluminum alloys in the short-transverse direction in an environment of
3Vz per cent NaCl solution.
specimen to show the transgranular nature of the fatigue crack and the
intergranular nature of the stress corrosion crack.
2024-T351 Aluminum Alloy—Short-Transverse Direction
In the T351 temper (solution treated and stress relieved by stretching
\% to 3 per cent), the susceptibility of 2024 to stress corrosion attack
in conventional testing has been reported to be approximately equivalent to the 7075-T6 material [7].
On the basis that the susceptibility was approximately equivalent
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STRESS CORROSION TESTING
pattern for comparison purposes. The values developed using a maximum exposure time between 25,000 and 30,000 min are shown in Fig. 8.
These data show a substantially less stress corrosion cracking susceptibility when compared to the 7075-T6.
The plane-strain fracture toughness measurements of this material
using both the maximum load and departure from linearity (pop-in)
from ^-in.-thick specimen were as follows: KIe maximum load, 24,300
lc pop-in, 20,300 and 23,900 psi-in.1/2 Based on the pop-in
determinations, an average Klc value of 22,100 is reported in Table 2.
These data indicate that plane-strain conditions were not achieved,
FIG. 9—Crack length of 7075-T6 specimens oriented in the long-transverse
direction after exposure to aqueous environments for 6000 min. No trend which
demonstrates a dependency upon stress intensity or environment is evident in
these data.
due, most likely, to the higher Klc-to-ays ratio for this material. The
Klc reported represents a lower bound value. Also, for the K0 data
shown on Fig. 8, the data above approximately 14,000 psi-in.1'2 represent conservative values of initial stress intensity. These findings are
consistent with the recent recommendation that the specimen thickness
should be 2>£ times (K/ays)2
(see footnote 4).
T
With a higher level of fracture toughness, a lower level of susceptibility
to crack extension might be anticipated, which has been found in the
case of steel [5]. A comparison between the results of the 7075-T6 and
2024-T351 (Fig. 8) demonstrates this decrease in crack propagation
susceptibility.
* Brown, W. F. and Srawley, J. E., communication to Subcommittee I of ASTM
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
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7075-T6 Aluminum Alloy—Long-Transverse Direction
Since initial experimental probing indicated that failure of 7075-T6
specimens oriented in the long-transverse direction would not occur
within 30,000 min testing times, an exposure time approach was used.
The exposure time selected was 6000 min, and the three environments
were used. After exposure the specimens were broken at 90 deg to the
anticipated direction of crack propagation as previously reported. Subsequent examination of the fracture surfaces indicated that the most
FIG. 10—Photomicrographs of the fatigue cracks in 7075-T6 aluminum alloy
specimens oriented in the long-transverse direction after exposure to an alternate-immersion environment for 6000 min. (top) Ko = 17,400 psi-in.1'2 and (bottom) K0 =
15,200 psi-in v* Keller's etch (X700).
reliable crack length determinations were from the specimen edge to
the fast fracture portion. This includes the machined notch depth, the
mechanically induced fatigue crack, and any potential stress corrosion
crack propagation. Note that variation commonly exists in the size of
the fatigue crack from specimen to specimen. Using this technique, the
data obtained are shown in Fig. 9. It is apparent that no trend is present
in these data which would indicate either a stress intensity or environmental dependency present. Metallographic examination did not show
any clear evidence of intergranular crack extension. Photomicrographs
of typical cracks are shown in Fig. 10 to demonstrate the crack characCopyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
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STRESS CORROSION TESTING
FIG. 11—Crack length of 2014-T6 specimens oriented in the long-transverse direction after exposure to aqueous environments for 6000 min.
FIG. 12—Photomicrographs of the cracks in 2014-T6 aluminum alloy specimens
oriented in the long-transverse direction after exposure1/ato an alternate-immersion environment for 6000 min, (top) Ke = 19,300 psi-in. and (bottom) Ko = 14,100
psi-in.1'" Keller's etch (XlOO). Reduced one third for reproduction.
2014-T6 Aluminum Alloy—Long-Transverse Direction
Using the same experimental procedure as was used for the 7075-T6
in the long-transverse direction, the stress corrosion susceptibility of
2014-T6 in the long-transverse direction was investigated. The values obtained for the crack length as a function of KO are presented in Fig. 11.
Examination of these data also does not indicate any dependency upon
stress intensity or environment.
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MULHERIN ON HIGH-STRENGTH ALUMINUM ALLOYS
79
indicated a corrosive mechanism operative in the fatigue cracked area.
This was indicated by a broadening effect in the main crack and possible
crack extension. Since the orientation of the photomicrograph is in a
short-transverse plane, this broadening would indicate propagation in
the short-transverse direction. The photomicrographs showing this condition are presented in Fig. 12.
7178-T6 Aluminum Alloy—Long-Transverse Direction
Similar behavior was observed on 7178-T6 aluminum alloy as was
obtained on the 7075-T6. Therefore, a series was run with exposure
time as variant instead of stress intensity and environment. A series of
specimens was prepared using additional care to minimize any difference
in fatigue crack length. Specimens were exposed to a distilled water enTABLE 3—Influence of time on the crack propagation of 7178-T6 aluminum alloy
in the long-transverse direction.
Exposure Time, min
<1
<1
500
1 000
2
6
6
10
20
20
000.
000
000
000
000
000
a, iin.
Unstressed
0.210
0.210
0.209
0.208
0.208
0.208
0.207
0.209
0.204
0.204
Stressed
0.206
0.209
0.209
0.210
0.207
0.206
Environment
ambient
ambient
distilled water
distilled water
distilled water
distilled water
hydraulic oil
distilled water
distilled water
hydraulic oil
vironment in both the loaded and unloaded condition. For the loaded
specimen, the KQ applied was 13,600 ± 200 psi-in.1/2 Exposure times
used varied from 500 to 20,000 min. Two specimens were broken immediately after fatigue cracking to obtain a zero time crack length. Also,
two specimens were exposed in the unloaded condition to an environment of light hydraulic oil for periods of 6000 and 20,000 min. The results obtained are compiled in Table 3. Examination of these data do not
indicate any pertinent influence of exposure time up to 20,000 min.
Metallographic evidence was similar to that obtained on the 7075-T6
in the long-transverse direction.
Evaluation of Experimental Results
The experimental results are evaluated by comparison to results obtained by a conventional testing method. As a basis of comparison, information is used from the reference by Sprowls and Brown [1]. From
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STRESS CORROSION TESTING
direction can be categorized as having low resistance to stress corrosion
cracking. On the basis of sustained tensile load which did not cause
failure in tension and ring type specimens, 7075-T6 has slightly better
resistance than 2024-T351. In the data reported in this paper, however,
a substantial difference in resistance to crack propagation is indicated
between the two alloys. This observation, naturally, gives rise to a question on the relative importance of the crack initiation period as compared to the crack propagation phase.
In the long-transverse direction, a perceptible difference exists on the
resistance to stress corrosion cracking as shown in Ref 1. The lowest
resistance is in the 2014-T6 alloy with increased resistance in the
7075-T6 and a further increase in resistance in the 7178-T6. In all
cases, the resistance was substantially higher than in the short-transverse
direction. No quantitative distinction was found in the present investigations between the three alloys within the experimental conditions investigated.
Summary of Results
1. Substantial intergranular stress corrosion crack propagation occurred in 7075-T6 aluminum alloy specimens oriented in the shorttransverse direction. Susceptibility decreased with decreasing stress intensity.
2. Within the experimental range investigated, the material behavior
patterns were similar in environments of distilled water, 3|/2 per cent
NaCl solution, and an alternate-immersion cycle in the NaCl solution and
air. Both the alternate and the constant immersion in salt solution were
more aggressive environments than the distilled water.
3. A substantially lower level of susceptibility in the short-transverse
direction was demonstrated in 2024-T351 aluminum alloy as compared
to 7075-T6. A lower level would be anticipated in view of the higher
fracture toughness of this material.
4. Using the crack length after exposure as the susceptibility criterion,
stress corrosion crack propagation on long-transverse specimens from
i4-in.-thick plate was not demonstrated on 7075-T6, 7178-T6, and
2014-T6 aluminum alloys.
A cknowledgment
The author wishes to express appreciation to his colleagues at Frankford Arsenal, especially H. Markus, H. Rosenthal, S. Lipson, and D. F.
Armiento, for many hours of discussion, advice, and encouragement
during the course of this investigation.
References
[1] Sprowls, D. O. and Brown, R. H., "What Every Engineer Should Know About
Stress by
Corrosion
of Aluminum,"
Metal Wed
Progress,
Vol.
81, No.
April and
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DISCUSSION ON HIGH-STRENGTH ALUMINUM ALLOYS
81
[2] Brown, B. F., "A New Stress-Corrosion Cracking Test Procedure for HighStrength Alloys," Materials Research & Standards, Vol. 6, No. 3, March 1966,
pp. 129-133.
[3] Irwin, G. R., "Fracture," Handbuch Der Physik, Vol. VI, Springer, Berlin,
1958.
[4] Paris, P. C. and Sih, G. C. M., "Stress Analysis of Cracks," Fracture Toughness Testing and Its Applications, ASTM STP 381, American Society for
Testing and Materials, Philadelphia, 1965, pp. 30-83.
[5] Mulherin, J. H., "Stress Corrosion Susceptibility of High-Strength Steels in
Relation to Fracture Toughness," Journal of Basic Engineering, Transactions,
American Society of Mechanical Engineers, Vol. 88, Series D, No. 4, pp.
777-782.
[6] Srawley, J. E. and Brown, W. F., Jr., "Fracture Toughness Testing Methods,"
Fracture Toughness Testing and Its Applications, ASTM STP 381, American
Society for Testing and Materials, 1965, pp. 133-198.
[7] Freed, C. N, and Krafft, J. M., "Effect of Side Grooving on Measurements of
Plane-Strain Fracture Toughness," Journal of Materials, Vol. 1, No. 4, Dec.
1966, pp. 770-790.
DISCUSSION
D. O. Sprowls1 (written discussion)—The use of a stress-intensity
factor, KI , to describe the stress that causes stress corrosion cracking
of an alloy is very intriguing. This paper is significant, because it contains the results of some of the first testing done with precracked specimens of high-strength aluminum alloys. It is evident from the apparent
lack of correlation of these limited test results on 7075-T6 and 2024T351 with the published results of conventional stress corrosion tests
of these alloys, however, that additional tests are needed to evaluate
this procedure for use with aluminum alloys.
By way of caution it should be remembered that the susceptibility
to stress corrosion of an alloy depends not upon the metallurgical structure and the attendant mechanical properties alone, nor upon the electrochemical characteristics of this structure alone, but upon these factors
acting together in a specific environment. In the stress corrosion cracking process both the initiation and the propagation of cracking are important. An accelerated test that involves either aspect of the process
to the exclusion of the other may not fully characterize the serviceability
of an alloy from the stress corrosion viewpoint. E. H. Dix, Jr., in the
introduction to the ASTM-AIME Symposium on Stress-Corrosion
Cracking in 1944, said, truly, that "Accelerated corrosion tests, at best,
1
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chief, Int'l
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STRESS CORROSION TESTING
are something of a hazard and accelerated stress-corrosion cracking tests
are a hazard to the nth degree."
It is well known that alloy 2024-T351 has greater notch toughness
and fracture toughness than 7075-T6. Also it has been established that
the fractures of notched and precracked specimens used for these tests
are predominantly transgranular. On the other hand, stress corrosion
cracks in these alloys follow a predominantly intergranular path. Thus,
the susceptibility to propagation of intergranular stress corrosion cracks,
in the opinion of this writer, must depend upon electrochemical reactions
in the path of the crack as well as upon the fracture toughness of the
metal. That such is the case is borne out, in part, by the fact that 2024T851 plate (artificially aged'2024-T351) is highly resistant to stress
corrosion cracking with conventional stress corrosion tests of shorttransverse, as well as long-transverse, specimens, yet its fracture toughness is relatively low, even when compared to that of 7075-T6.
B. W. Lifka2 (written discussion)—One point that is puzzling is the
disparity between results obtained on the two alloys, 7075-T6 failing
rapidly and 2024-T351 not failing.
As mentioned by the author, stress corrosion tests using conventional
specimens show these two alloys to be very similar; if anything, alloy
2024-T351 is the more susceptible of the two in the type of environment
used. In 3.5 per cent sodium chloride by alternate immersion, both 2024T351 and 7075-T6 short-transverse conventional specimens (without
any stress concentrator) will fail within one to two weeks even at very
low stresses, such as 10 ksi (approximately 15 per cent yield strength).
Since the principal purpose of a precracked specimen is to minimize the
incubation period, it does not seem likely that its use should alter the
relative performance of the two alloys so drastically. Another disparity
with conventional stress corrosion tests of 7075-T6 alloy is that shorttransverse specimens fail by stress corrosion cracking considerably
faster when intermittently immersed than when continuously immersed in
3.5 per cent NaCl solution.
One possible explanation might be that the particular 2024 plate used
by the author had higher than normal resistance for the -T351 temper.
It is suggested, therefore, that a number of conventional stress corrosion
tests also ought to be made to establish the quality of the material by
this method. Any differences subsequently observed with use of precracked specimens could then more definitely be attributed to the change
in testing procedure.
It has been stated that use of precracked specimens "eliminates" the
incubation period for stress corrosion cracking. This is undoubtedly an
overstatement. Since the metal forming the walls of the fatigue crack
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DISCUSSION ON HIGH-STRENGTH ALUMINUM ALLOYS
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has not been corroding, some time period, however slight, probably will
be required before the corrosion process starts when the specimen is subsequently exposed. Furthermore, it is known from tests of conventional
specimens and from service experience that stress corrosion cracks very
frequently do not occur at the site of highest stress concentration. Rather
they occur at the site where all factors—stress, environment, electrochemical, and metallurgical—are optimum in combination to generate
a stress corrosion crack. No metal is ideally homogeneous, and this is
true as regards the electrochemical activity of the metal surface. The
likelihood of positioning the fatigue crack at the most active anodic sites
is therefore a matter of chance. In one case, the root of the crack may
be favorably located so that further propagation by stress corrosion
cracking could occur almost immediately. In another case, the crack
may be unfavorably located so that initially only general corrosion would
occur, gradually widening and deepening the fissure.
In tests of smooth specimens, the probability of developing a stress
corrosion fissure will increase with increased area of exposure. Use of
a precracked specimen tends to restrict this area of possible failure to
the walls of the fatigue crack. Therefore, it must be established that
the area and shape of the crack employed are such that the probability
factor is comparable for all items tested.
/. H. Mulherin (author)—The author wishes to express appreciation
to Messrs. Sprowls and Lifka for their very interesting and informative
comments and looks forward to the publication of the efforts of other
investigators in this area.
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C. S. Lin,1 J. J. Laurilliard^ and A. C. Hood1
Stress Corrosion Cracking of
High-Strength Bolting
REFERENCE: C. S. Lin, J. J. Laurilliard, and A. C. Hood, "Stress Corrosion Cracking of High-Strength Bolting," Stress Corrosion Testing, ASTM
STP 425, Am. Soc. Testing Mats., 1967, p. 84.
ABSTRACT: Exposure of high-strength steel bolts (260,000 psi) to the
natural environments of temperature changes, moisture, and salt air
have resulted in a number of stress corrosion cracking failures. Cadmium
plating of the bolts has not prevented these failures. This paper presents
a study of bolts made from H-ll, 4340, and maraging steel heat treated
to several strength levels. The authors have adopted the 3Va per cent NaCl
solution, intermittent exposure test to high-strength bolting and have used
it as a basis for comparison of bolt life. The influence of thread rolling
sequence and applied and residual stresses, as well as electroplated and
zinc primer coatings, were investigated. The results indicate that a H-ll
high-strength steel bolt with threads rolled after heat treatment will fail
in the shank. Cadmium plate with an undercoat of nickel increases the
life by five to seven times. The comparison of steel types shows the
maraging steel (300) to have the best life of the three alloys but not sufficient to overcome the environmental effects.
KEY WORDS: stress corrosion, corrosion, bolts, steels, environmental
testing, salt water, cadmium, nickel, gold, plating, rolling, shot peening
Problem
The needs of aerospace structural designers for higher strength-todensity ratios have resulted in demands placed on the fastener producers
for bolting of higher and higher strengths. Today mechanical fasteners are
being used with strengths in excess of 300,000 psi.2
Exposure of high-strength steels to the natural environment of temperature changes, moisture, and salt air have resulted in a number of service
failures by stress corrosion cracking. The exposure of high-strength bolts
to these same environments, even when the bolts were cadmium plated,
1
Research metallurgist, research engineer, and manager, respectively, SPS
Laboratories, Standard Pressed Steel Co., Jenkintown, Pa. Messrs. Laurilliard and
Hood are personal members ASTM.
2
Hood, A. C. and Sproat, R. L., "Ultrahigh-Strength Steel Fasteners," Structure and Properties of Ultrahigh-Strength Steels, ATSM STP 370, American
Copyright
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Materials,
Philadelphia,
1965, p.EST
208.2015
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LIN ET AL ON HIGH-STRENGTH BOLTING
85
TABLE 1—Chemical composition of bolt materials (weight per cent).
Element
Maraging (300)
c
0.02
0.04
0.005
0.008
0.006
18.95
Mn
P
S
Si
Ni
Cr
Mo
Fe
Co
Al
Ti
B
Zr
Ca
4.96
balance
9.09
0.15
0.66
0.0049
0.010
0.05
4340
H-ll
0.40
0.72
0.013
0.013
0.27
1.69
0.75
0.22
balance
0.40
0.29
0.006
0.012
1.00
5.11
1.33
balance
TABLE 2—Heat treatment and tensile properties of high-strength bolts.
Alloy
Hardening Temperature
and Time
Maraging (300) . . .2050F, 5 min;
1050 F,
salt quench
H-ll
4340
1850 F, 20 min;
atmosphere cool
1550 F, 20 min;
oil quench
Tempering Temperature
and Time
975 F
2 hr +
1050 F
2 hr +
1020 F
2 hr +
1035 F
2 hr +
1080 F
2 hr +
1130 F
2 hr +
500 F
2 hr +
600F
2 hr +
800 F
2 hr +
ProporUltimate
tional
Tensile Load, Limit,
lb (avg)
lb (avg)
2 hr + 2 hr
10 410
8000
2 hr + 2 hr
9 690
7200
2 hr + 2 hr
11 750
8350
2 hr + 2 hr
11 270
7500
2 hr + 2 hr
9 963
7350
2 hr + 2 hr
8 810
6400
2 hr
9 550
6550
2 hr
8 670
6200
2 hr
7 670
6100
resulted in similar failures. This prompted an investigation of the bolts in
a laboratory test designed to simulate some of the worst conditions of
environmental exposure. It was intended to learn what the influences of
bolt material, bolt strength, surface condition, and applied stress would
have on the resistance-to-failure by stress corrosion cracking.
Manufacture of Bolts
The standard
260,000
tensile
boltWed
made
H-ll EST
steel2015
in a size
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86
STRESS CORROSION TESTING
1.
2.
3.
4.
5.
6.
7.
8.
blank cut
head hot forged
headed blank annealed
rough machined
headed blank hardened and tempered
body and thread roll diameters ground
retempered after grinding for stress relief
threads and head-to-shank fillet rolled
FIG. 1—Load extension curve of lA-28 by 2 uncoated, 290,500 psi ultimate
strength H-ll bolts.
The normal practice after thread rolling is to dry-blast clean and
vapor deposit cadmium on the bolt. Since several different coatings were
to be applied, the uncoated bolt was used for a control.
The other alloys, AISI 4340 and maraging steel (300), were also used
in the investigation. They were made in the same fashion as the control
bolt except that the heat treatment corresponding to the alloy was used.
Chemical composition for three alloys is shown in Table 1, and heat
treatment and strength are shown in Table 2.
Proportional limits were established with an overall bolt extensometer
and calculated as shown in Fig. 1.
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LIN ET AL ON HIGH-STRENGTH BOLTING
87
stress calculations, the applied loads are always given in pounds. Bolt
strengths are indicated in pounds per square inch, and the stress is
calculated using the area at 98 per cent of the basic pitch diameter. This
is the conventional method used for this type fastener and is in accordance
with Specification NAS1348. For the V4-28 thread used in these studies,
an area of 0.0388 in.2 was used. Where unit stresses are shown for the
shank of the bolts, an area of 0.0491 in.2 was used.
Material and Bolt Strength
Three steels, H-ll, maraging (300), and 4340, were manufactured
as uncoated bolts in several strength levels to measure the influence of
material and tensile strength on stress corrosion cracking. All threads
and fillets were cold-worked after heat treatment.
TABLE 3—Surface coatings.
Type
Average Coating
Thickness, in.
Remarks
Vapor deposited cadmium (vacuum)
0.0005
Chromate conversion coatings on
electroplated cadmium
MIL-C-8837, no baking after
deposition
Cd 0.0005
Electroplated nickel
0.0005
Douglas proprietary cadmium and chromate coating
AMS 2424, post plate baked
for 3 hr at 630 F
Electroplated nickel and electroplated cadmium (undiff used) . . . .
Electroplated nickel and electroplated cadmium (diffused)
Electroplated gold
Zinc chromate
Zinc chromate (encapsulated dry
film)
Ni 0.0003
Cd 0.0002
Baked after cadmium plating
for 3 hr at 375 F
Ni 0.0003
Cd 0.0002
0.00006
AMS 2416, diffused at 630 F
for 3 hr
MIL-P-8585
National Cash Register Co.
proprietary process
Thread Rolling Sequence
Bolts of H-ll steel were manufactured as the control bolt except that
threads were rolled before heat treatment. These bolts were compared
with H-ll control bolts. Both lots were heat treated to several strength
levels for a comparison of the influence of thread rolling sequence on
resistance to stress corrosion cracking.
Residual Stress
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88
STRESS CORROSION TESTING
sive stresses applied to the shank as well as to the threads of the control
bolt. The control bolt is essentially free of residual stresses in the shank
area since it is retempered after grinding.
One group of bolts was shot peened over the entire bolt surface with
O.Oll-in.-diameter cast steel shot to an Almen strip intensity value of
FIG. 2—Loaded test cylinder used for stress corrosion. Note the slots which
facilitate contact of the salt solution with the bolt shank.
FIG. 3—Test equipment for stress corrosion studies. Note the fans used to
accelerate drying during the 50-min portion of the stress corrosion cycle.
0.008 A. Another group was glass bead shot peened at 80-psi air pressure. The nozzle was held 6 in. from the bolt surface.
Decarburization
One of the groups of control bolts was made with the heat treatment
atmosphere so adjusted that the finished bolt was decarburized. If surface
hardness or carbon content influences the resistance of H-ll to stress
corrosion cracking, then this experiment would measure it. The shank
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LIN ET AL ON HIGH-STRENGTH BOLTING
89
surface was decarburized to a hardness of VHN 506 (Rc 47.5) which
increased to a core hardness of VHN 606 (Rc 53.5) at 0.020 in. below the
surface.
Surface Coatings
Various coatings were applied to the H-ll control lot of bolts as
shown in Table 3. Metallic coatings were applied by both electrodeposition and vapor deposition. Zinc chromate primer was also applied by two
FIG. 4—Effect of applied load of the stress corrosion cracking susceptibility of
/4-28 by 2 uncoated H-ll bolts. The bolts had an ultimate strength of 290,500 psi
with threads rolled after heat treatment. All failures occurred in the bolt shank.
l
methods. Bolts were coated with a 1:1 solution of zinc chromate primer
(Mil-P-8585) and allowed to dry while standing on the thread end. In the
second method, bolts were coated with plastic encapsulated zinc
chromate. The pressure of bolt tightening burst the spheres and applied
the liquid in this fashion. The process of encapsulation is a development
of the National Cash Register Co. and the U.S. Air Force.3
3
Hanny, J. F. and Price, J. E., "Dry Films of Encapsulated Zinc Chromate
Primer," AFML TR-65-54, March 1965, Air Force Materials Laboratory, WrightPatterson Air Force Base, Ohio.
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STRESS CORROSION TESTING
Test Equipment and Methods
The test equipment consisted of turned steel cylinders (Fig. 2) with
slots milled in them to permit free flow of the corrosive media and air.
Bolts were tightened on the cylinders to a predetermined percentage of the
proportional limit. Bolts had previously been tested in a tensile machine
to develop a load-elongation curve (Fig. 1). This was used to measure the
proportional limit. A supermicrometer was used to measure the bolt
elongation while tightening on cylinders was taking place.
The bolted cylinders were hung from a test fixture (Fig. 3). The test
fixture was immersed in a room-temperature solution of 3l/z per cent
sodium chloride (NaCl) in distilled water. The fixture was immersed for
10 min of every hour and withdrawn for 50 min. While the fixture and
cylinders were suspended over the bath, fans were turned on to facilitate
drying with rapidly moving air. Previous investigation indicated that
rapid drying would accelerate failure.
Specimens were examined, and the test solution changed daily. Tests
were continued until bolts failed or a minimum of 1000 hr.
In comparison testing, bolts were loaded to either 90 or 100 per cent
of the proportional limit. The test loads are indicated in the results. In
one experiment, the H-ll control bolt was loaded to varying percentages
of the proportional limit to see if the applied stress had an effect on the
resistance to stress corrosion cracking.
Test Results
Effect
of Applied Stress
The H-ll steel control bolts, uncoated, but with threads and fillets
rolled after heat treatment, were tightened to several applied loads in the
cylinder, and the life was determined. Figure 4 shows the basic stressrupture type curve. The data appear to show that there is a lower level of
applied stress, below which no failure will occur. All failures below the
proportional limit occurred in the bolt shank, which would indicate that
the problem is basically a material deficiency rather than something which
is adversely affected by bolt manufacturing techniques.
Effect
of Material and Bolt Strength
The control bolt of H-l 1 was compared with 4340 and maraging steel
bolts manufactured in a similar fashion and loaded to the proportional
limit for the respective bolts. The results hi Table 4 show that the
maraging steel (300) is affected to a lesser degree than the other two
alloys. The H-l 1 bolts are slightly better than 4340 at the same tensile
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LIN ET AL ON HIGH-STRENGTH BOLTING
91
TABLE 4—Effect of material and strength on susceptibility to stress corrosion
cracking ofuncoated bolts."
Material
H-ll steel
Ultimate Bolt
Strength, psi
Time-to-Failure,6 hr
63C
88'
302 800
290 500
256 800
227 100
Maraging (300)
268 300
249 700
4340 steel
246 100
222 200
197 700
74
106
106
136
1515*
227
1030*
87
196
1002d
136
185
252
103
117
124
140
184
164
1517d
504
1054d
90
220
1291d
167
1610d
603
1127*
353
1000d
1291d
0
Bolts are K-28 by 2 with rolled threads after heat treatment and loaded to
100% of their proportional limit.
6
Failures were shank failures unless noted.
e
Head failure.
d
No failure, test discontinued.
TABLE 5—Effect of bolt strength and rolling threads before and after heat treatmen t
on susceptibility to stress corrosion cracking ofuncoated H-ll steel bolts.°
Thread Rolling Condition
Rolled threads before heat treatment
Rolled threads after heat treatment
Ultimate Bolt
Strength, psi
283 500
251 300
217 800
iPime-to-Failure!, hr
I6
100*
140»
1516
185 600
75P
302 800
63d
88d
74e
290 500
256 800
227 100
106e
106«
136«
1515'
3.56
106*
1526
176*
8926
103e
117«
124e
140*
184'
164"
1517e
3.8*
177*
1343C
1343"
1344"
136e
185«
252*
167e
1610"
0
Bolts are ^-28 by 2 loaded to 100% of their proportional limit.
Thread failure.
c
No failure, test discontinued.
d
Head failure.
• Shank failure.
6
Effect of Thread Rolling Sequence
The bolt life in hours as a function of bolt strength for both sequences
of thread rolling is shown in Table 5. The influence of thread rolling
after heat treatment, which shifts the failure to the bolt shank with a
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92
STRESS CORROSION TESTING
Effect of Residual Stress
The results of shot peening and glass bead peening of H-ll control
bolts are shown in Table 6. When the threads are rolled after heat treatment, further working of the bolt thread surfaces does not affect the results. Apparently the peening of the bolt shank provides insufficient compressive stresses to influence the environmental conditions.
Effect of Coatings
The results of the tests as shown in Table 7 indicate that nickel and
nickel-cadmium coatings provided a considerable increase in the resistTABLE 6—Effect of residual stress on the susceptibility to stress corrosion cracking
of H-28 by 2 uncoated H-ll bolts.
Ultimate
Strength,
psi
Applied
Load,
lba
284 800
6800
138 500 175 300
3
lc
Standard bolt (rolled
threads after heat treatment)
290 500
7500
152 700 192 300
7
Steel shot peened
290 500
7500
152 700 192 300
3
Glass shot peened
290 500
7500
152 700 192 300
5
74
106
106
124
120
123
128
102
114
123
Method of Stressing
Standard bolt (rolled
threads before heat
treatment)
Shank
Stress,
psi
Thread
Stress,
psi
Bolts
Tested
Time-toFailure,6 hr
3.5C
3.8"
140
184
252
133
271
a
100% of proportional limit.
Failures were shank failures unless noted.
c
Thread failure.
5
ance of the standard bolting to stress corrosion failure. Vapor deposited
cadmium offered some improvement over the bare bolt. Electroplated
cadmium with a post plate dichromate more than doubled the resistance
of vacuum cadmium but did not match the nickel-cadmium coatings.
Gold plating, with a thickness of 0.00006 in., showed little improvement
over uncoated bolts. Zinc chromate primer was approximately equivalent
to vapor deposited cadmium whether dipped or applied by the encapsulated process.
Effect
of byDecarburization
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LIN ET AL ON HIGH-STRENGTH BOLTING
Surface Treatment
Neutral
Decarburized
Bolts Tested
Time-to-Failure, hr
4
5
106, 106, 123, 124
74, 112, 114, 128, 138
93
All failures occurred in the bolt shank. Note that no significant difference in bolt life was observed.
TABLE 7—Effect of coatings on the susceptibility to stress corrosion cracking of
V±-28 by 2 H-ll bolts*
Bolts
Tested
Bolt Coating
Bare
8
Vacuum cadmium
Time-to-Failure,6 hr
68
70
52
76
92
92
6
Douglas proprietary electroplated cadmium and chromate conversion coating.
186
242
256
291
329
329
6
Electroplated nickel
5
668
734
1010"
1010'
5093"
743
761
5093°
5093'
784
787
11
1587e
1587"
1587"
1587"
1587"
1587°
2581
Electroplated nickel and electroplated
cadmium (undiffused)
995
1010^
1010"
1252
4
Electroplated gold
Zinc chromate primer (dipped)
3
4
Encapsulated zinc chromate primer
6
767
1512'
274
125d
21 ld
224
241
1512C
1512"
336
223d
268
246
379
Electroplated nickel and
cadmium (diffused)
electroplated
140
140
467
472
480
a
Ultimate bolt strength = 290,500 psi; applied load = 6750 Ib (90% proportional limit); shank stress = 137,500 psi; and thread stress = 174,000 psi.
6
Failures were shank failures unless noted.
c
No failure, test discontinued.
d
Head failure.
Mechanism of Failure
From microscopic appearance, it was found that the location of stress
corrosion cracking originates at the bolt surface (Fig. 5). This was related
to the corrosion pits on the shank surface. Examination of the fracture
surface on bare and plated bolts showed that considerable rust had
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94
STRESS CORROSION TESTING
All primary stress corrosion cracking was found to propagate transgranularly and intergranularly at a direction approximately 90 deg to the
direction of loading.
In the majority of cases, stress corrosion cracking failures were located
in the shank. Occasionally, failure occurred in the thread area, head area,
and lightening-hole area. Head failures occurred only at a bolt strength
FIG. 5—Corrosion pit and crack initiation in the shank of an H-ll steel bolt.
White arrow indicates the direction of applied stress.
of 302,800 psi on H-l 1 bolts with rolled threads after heat treatment
Thread failures occurred only on bolts with threads rolled before heat
treatment. The photomacrographs of thread, shank, and head failures are
shown in Figs. 8 to 10.
The lightening-hole failure (Fig. 11) was found only with plated bolts
that were capable of resisting head, shank, and thread fracture for long
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LIN ET AL ON HIGH-STRENGTH BOLTING
95
FIG. 6—Stress corrosion fracture of H-ll bolt. Note the point source, corroded
area with fracture radiating from this point.
FIG. 7—Shank failure of H-ll steel bolt. Dark zones are patches of rust.
FIG. 8—Thread
failure
of H-ll
steelWed
bolts
threadsEST
rolled
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96
STRESS CORROSION TESTING
stress corroded bolts from the test fixture. Upon application of torque,
the twelve-point serration sheared off. Figure 11 indicates that the crack
path initiates near the base of the lightening hole and curves smoothly
in a slight downward arc to the outside surface of the twelve-point area.
The break is symmetrical around the bolt circumference. An experiment
FIG. 9—Head failures of H-ll steel bolts with threads rolled after heat treatment.
FIG. 10—Shank failures of H-ll steel bolts with threads rolled after heat treatment.
was run to determine retention of bolt load after twelve-point serration
failure. Bolts with strain gages were loaded in cylinders to the prescribed
preload. The serrated section was cut off and the bolt load measured.
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Measurements
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LIN ET AL ON HIGH-STRENGTH BOLTING
97
Discussions and Conclusions
If one considers this investigation in the light of the use of high-strength
steel bolting, then a number of interesting aspects are revealed. It is not
unreasonable to assume that the 3l/2 per cent salt solution immersion is
an accelerated version of what happens to structural bolts used in aircraft in many parts of the world. The loads imparted to the bolts in the
cylinders are typical of what one would expect these bolts to carry in an
aircraft structure.
What appears to be the most significant discovery relates to the
metallurgy of high-strength steels for structures. If the best possible
metallurgical practices were used to produce a 260,000 psi threaded
fastener and yet one lacked the technology of producing threads and fillets
FIG. 11—Stress corrosion cracking of bolt head lightening hole.
by rolling after heat treatment, then the life in our experimental test would
average about 2 hr. Where threads and fillets are rolled after heat treatment, life is raised to at least 74 hr. The addition of vacuum cadmium
plating more than doubles this Me.
This is where standard high-strength steel bolting is today. It would
appear that the most immediate improvement could be obtained by the
addition of a nickel layer under the vacuum cadmium coating. It is
possible that this may be sufficient to prevent stress corrosion cracking of
high-strength bolting within the expected life span of today's aircraft,
since it more than triples the predictable life.
The other approach may be through the use of a bolting material which
in itself has high resistance to stress corrosion cracking. The authors
believe that a material with high resistance is the most foolproof soluCopyright
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98
STRESS CORROSION TESTING
hydrogen embrittlement must be considered, otherwise we solve the
problem of stress corrosion cracking only to create others.2
It is believed that any future evaluation of high-strength bolt materials
must include an assessment of their resistance to stress corrosion cracking. The present test appears to be adequate for bolts of this strength
level.
A cknowledgment
The authors are grateful to R. L. Sproat, director of engineering,
Standard Pressed Steel Co., for continued encouragement and guidance
during the course of this investigation. We also wish to acknowledge the
help and suggestions of W. Boyd and M. Epstein, Battelle Memorial
Institute, provided during the course of a private research program for
Standard Pressed Steel Co.
DISCUSSION
Anthony Gallaccio1 (written discussion)—You report that the nickel
plate and the nickel-cadmium plate on H-ll steel bolts provided
markedly increased resistance of the bolts to stress corrosion cracking.
Can you explain why the nickel plate is so effective, and why the nickel
plate is so much better than the cadmium plate?
C. S. Lin, J. J. Laurilliard, and A. C. Hood (authors)—The authors
appreciate the interest of Mr. Gallaccio in the stress corrosion cracking
of high-strength bolts. The nickel is believed to offer greater resistance
than cadmium to stress corrosion cracking primarily because of its higher
resistance to pitting. We believe that when the pit progresses to the surface of the steel, then one of the several mechanisms advanced for stress
corrosion cracking takes place. However, it takes longer for this to occur
with nickel, and the advantage over cadmium is therefore one of time.
1
Chief, Protection and Preservation Branch, Pitman-Dunn Research Laboratories, Frankford Arsenal, Philadelphia, Pa.
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A. GaUaccio1 and M. A. Pelensky1
Stress Corrosion of High-Strength Steel
Alloys—Environmental Factors
REFERENCE: A. GaUaccio and M. A. Pelensky, "Stress Corrosion of
High-Strength Steel Alloys—Environmental Factors," Stress Corrosion
Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 99.
ABSTRACT: Several years ago it became evident to members of ASTM
Committee B-3, Subcommittee X (now Committee G-l, Subcommittee
VI) on Stress Corrosion that a review of the reported information and,
where available, work in progress on stress corrosion testing was needed.
This was considered essential as a precursor to the proposed development
of standards and criteria for determining the stress corrosion susceptibility
and behavior of various alloys. The subcommittee established three task
groups to accomplish a review of the subject. The task group assignments
were as follows: TG-1, sample selection and method of stress; TG-2,
environments under which stress corrosion occurs and duration of tests;
and TG-3, reporting of stress corrosion results and interpretation of
results. This report is part of the TG-2 assignment. It concerns a review
of reported stress corrosion investigations relating to the effects of various
environments on specific high-strength steels.
KEY WORDS: corrosion, stress corrosion, steels, environmental testing
A survey is presented of selected reported information from investigations performed to determine the effects of various environments or
media on the stress corrosion of a variety of high-strength steels. The
results of the stress corrosion tests were compiled with other pertinent
related information. These included the specific environment, the composition and strength level of the alloy, the type of specimen used, the
stress level employed, and the direction and method of stressing. Informational sources are cited.
Based on the data reviewed and as deemed appropriate, comments
and suggestions are offered. These concern the selection of environments,
sampling, and the various integrated factors bearing on the methods of
testing and the test results.
1
Chief and research chemist, respectively, Protection and Preservation Branch,
Copyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
Pitman-Dunn
Research Laboratories, Frankford Arsenal, Philadelphia, Pa.
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STRESS CORROSION TESTING
Environments and Materials
The conditions of test and the materials covered in this survey are as
follows.
Environment
(Unless otherwise indicated, ambient temperature and total immersion
of the specimen apply.)
Air:
Indoor, 16 to 74% RH, 62 to
92 F
Indoor, 100% RH, 140 F
Indoor, 100% RH, 175 to 205 F
Outdoor, industrial
Outdoor, semi-industrial
Outdoor, marine
Waters:
Distilled
Distilled, 120 F
Distilled, 160 F
Tap
Sea, natural
Sea, synthetic
Solutions:
H2S, saturated
MgCl 2 , 42%, boiling
NaBr, 10.3% (1 Af)
NaC2H4O2, 8.2% (1 Af)
NaCN, 4.9% (1 Af)
NaF, 4.2% (1 M}
NaHCO 3 ,8.4% (1 Af)
NaCl, 3%, sprayed, followed
by exposure to air (dry)
NaCl, 3%, sprayed, followed
by exposure to air, ambient,
indoors
NaCl, 3%, sprayed, followed
by exposure to air, 100% RH
NaCl, 3%
NaCl, 3%, boiling
NaCl, 3^%, alternate immersion
NaCl, 5%, alternate immersion
NaCl, 5.7%
NaClO3, 10.64% (1 Af)
Nal, 15% (1 Af)
NaN0 2 ,6.9% (1 Af)
(NaNO2 + NaNO3), marquench, 1%, elevated temperature
NaNO 3 ,8.5% (1 Af)
NaOH, 4% (1 Af)
NaPO 3 , 10.2% (1 Af)
Na 2 HPO 4 , 14.2% (1 Af)
Na2Cr2O7, 0.25 %
Na2SO3, 10.6% (1 Af)
Na2S04, 14.2% (1 Af)
Na3P04, 16.4% (1 Af)
Alloys:
AFC 77
Cr (5 %); hot-worked die steel
AM 350
D6AC (ladish); martinsite, low
AM
355;
cold-worked
PH
steel
alloy
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GALLACCIO AND PELENSKY ON HIGH-STRENGTH STEEL ALLOYS
101
17-4 PH
300M; silicon modified, 4300
series
4330M
4335M
4340
4340M
Hll
Maraging steel, 18Ni
Maraging steel, 20Ni
Maraging steel, 25Ni
MBMC, No. 1
Ni (9%), Co (4%)
Vascojet 1000
12MoV
Data
The tabulated data compiled in this survey are available through the
authors. Table 1 is a sample of the data.
TABLE 1—Sample of data.
Alloy
Yield
Stress,
ksi
Test
Specimen
Stressed,
%
yield
Direction of
Stress
Stress
Method
No. Specimens,
time-to-failure
Ref
, 3%, AMBIENT TEMPERATURE
D6AC. . .
18Ni maraging
steel
0.23T1
0.52T1.
0.5Ti..
0.5Ti..
0.49T1.
0.4T1..
0.55Ti.
0.62T1.
0.5Ti..
0.62T1.
l.OOTi.
0.5Ti..
l.OOTi.
197.5
197.5
222.5
222.5
235
f
f
f
f
f
wr
wr
wr
wr
wr
75
dm
75
dm
75
dm
dm
dm
dm
dm
bb
ub
bb
ub
bb
235
252
f wr
f wr
75
dm
dm
dm
ub
bb
252
f wr
dm
dm
ub
181.5248.2
249.9
255.4
269.7
278.0
279.1
283
302.5
323
323.3
331
354.4
f
f
f
f
f
f
f
f
f
f
f
f
f
dm
dm
75
dm
dm
75
dm
75
75
75
75
75
75
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
ub
ub
bb
bb
ub
bb
ub
bb
bb
bb
bb
bb
bb
wr
wr
wr
wr
wr
wr
wr
wr
wr
wr
wr
wr
wr
Abbreviations:
f wr = flat wrought
ub =
dm = data missing
m =
bb = bent beam
avg =
nf = nobyfailure
Copyright
ASTM Int'l (all rights
3
2
3
2
3
nf 21 days
nf 198 days
nf 21 days
nf 198 days
nf 21 days, 3 nf
104 days
1 62 days
3 nf 21 days, 3 nf
104 days
1 116 days, 1 19
days
[7]
2 nf 4150 hr
3 1776 hr (m)
3 140 hr (avg)
2 nf 3600 hr
3 1174 hr (m)
2 nf 3600 hr
3 4704 hr (m)
3 36 hr (avg)
2 3200 hr (avg)
3 1580 hr (avg)
2 7 hr (avg)
2 nf 3600 hr
20 20 hr (avg)
[4]
U-bend
mean
average
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102
STRESS CORROSION TESTING
Discussion Concerning the Data
Atmospheres
On the whole, indoor atmospheres are less severe than outdoor atmospheres for the stress corrosion cracking of the high-strength steels.
The stress corrosion susceptibilities of D6AC and 300M alloys are
shown to be greater at the higher temperature and humidity conditions
in air, compared to a normal ambient indoor atmospheric condition [I].2
This observation would be expected to be evident for other alloys in
similar investigations but is not borne out by the data. The independent
effects of moisture or temperature conditions of the atmosphere are not
clearly indicated for clean metals, but some results shown for specimens
sprayed with salt solution and exposed indoors at different humidity
levels indicate the effect of humidity [2].
The following alloys fail in a short time when exposed to marine
environments (that is, Kure Beach 80-ft site):
4340, 218 to 254 ksi yield [2,3]
MBMC No. 1, 218 to 254 ksi yield [2]
Cr (5 per cent) steel alloy, 229 to 237 ksi yield [2]
Kure Beach exposure results for 20Ni maraging and 4340 steels indicate these alloys are more readily stress corrosion cracked by marine
exposure than by the industrial atmosphere at Bayonne [3]. Further, the
Kure Beach atmosphere results for various alloys are observed to be
comparable to those obtained with sea water immersion tests [3].
Waters
Apparently distilled water more readily causes stress corrosion failures
in certain high-strength steels than does NaCl, 3 per cent [4]. From
immersion test data, this observation is applicable in the case of a few
maraging, 18Ni steels, for example, 0.62 Ti (323 ksi yield); 0.5 Ti (249.9
ksi yield). Conversely the salt solution is more active in bringing on the
stress corrosion cracking of other maraging 18Ni steel alloys, for example,
the l.OTi (323.3 and 354.4 ksi yield) [4].
Alloy D6AC is shown to fail more readily in distilled water than in
salt solution [1].
For maraging steels, 18 and 20Ni alloy steels, tested by the bent-beam
method in distilled water, the influence of temperature at the high temperatures is evident. Failure times are shorter at the high temperature
[4]. U-bend tests of the same material under the same environments
were also found to fail more rapidly [4].
2
The italic numbers in brackets refer to the list of references appended to this
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GALLACCIO AND PELENSKY ON HIGH-STRENGTH STEEL ALLOYS
103
Solutions
Alternate immersion in NaCl, 3.5 per cent, using bent-beam specimens
of alloys D6AC, 4340, 4335M, HI 1, AFC 77, and AM 355 caused more
rapid failures than did outdoor exposures in a semi-industrial environment
(Seattle) [5]. However, those materials first to fail during alternate immersion were not necessarily the first to fail in the atmosphere [5]. The
frequency of rainfall or the persistence of condensate will materially
influence the initiation or promotion of corrosion outdoors.
Alloy 12MoV tested by immersion in NaCl, 3 per cent, at pH 1
through pH 11, with bent-beam stressing, reveals an inverse relationship
of pH to stress corrosion cracking; with the increase of pH, stress corrosion
failures decreased. No failures occurred at pH 11.5 or above [6].
In 1 M solutions of NaCl, NaNO3 , NaPO3 with bent-beam or U-bend
stressing, the following order of increasing susceptibility to stress corrosion of the following high-strength alloys is revealed [7]: AM 355,
D6AC, 300M, and Vascojet 1000.
One molar solutions of NaCl, NaNO3 , and Na2SO4 to approximately
the same degree attacked D6AC alloy of different yield strength levels.
But NaPOs resulted in more aggressive attack of the alloy of 206 and
223 ksi strength levels [7]. The same solutions, on the whole, exhibited
somewhat greater attack of the 300M and Vascojet 1000 alloys [7]. The
aggressiveness of the Na2SO4 solution was second to the NaPO 3 , especially for 300M yield strengths 218 ksi and greater. The NaCl was most
aggressive in the case of the Vascojet 1000, and the NaNO3 was least aggressive [7].
Alloys
The susceptibility to stress corrosion failure of maraging steels increases with higher titanium content and with increasing yield strength
[4].
In general, for a single alloy composition, the higher the stress level,
the shorter the time to stress corrosion failure [7,8].
For a given yield strength level, the maraging steels, 18Ni, are less
susceptible to stress corrosion failure than the conventional high-strength
steels [4].
Some of the reported data permit comparisons of the stress corrosion
behavior of maraging steel, 18Ni, to maraging steel, 20Ni, at the same
U-bend stress levels and similar environments. The indication is that
the 18Ni alloy is less prone to stress corrosion failure than the 20Ni
alloy [3,4], but, for the same alloys and environments, the limited data
for the bent-beam tests are less consistent in that respect [4].
Specimens of the 4340 steel drawn at 800 and 900 F have yielded results showing about 25 times greater resistance to stress corrosion crackCopyright
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ing
than those
drawn
at 475
700 F.Wed
These
comparisons
based on
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104
STRESS CORROSION TESTING
alternate-immersion tests in NaCl, 5 per cent, under sustained tensile
loading. Resistance to stress corrosion cracking is poor when the ultimate
tensile strength of the alloy exceeds 200 ksi [9].
Tempering temperatures above 1050 F used for triple drawing of Hll
alloy (ultimate tensile strength near 260 ksi) results in improved resistance
to stress corrosion cracking as determined by immersion in NaCl, 5 per
cent, under sustained tensile loading [9].
Steel alloys of 9 and 18Ni are reported to be less susceptible to failure
from stress corrosion than the 4340 alloy. The 9 per cent alloy is the
better of the two nickel alloys as revealed by alternate immersion in NaCl,
5 per cent, with sustained tensile loading [9].3
Stress Methods
Test results under all environmental exposures show that the U-bend
stressing is more severe than bent-beam stressing [4,7]. Predominantly
these were the test methods employed. Stressing by means of two- or
three-point loading or by tensile loading was employed in some cases
[3,9 JO].
Observations and Suggestions
The data reviewed reveal that a wide variety of environments have been
used for investigating stress corrosion susceptibility of high-strength steels.
Frequently the choice of environment is based on an actual or an expected
condition of service for the metal. Such choices are entirely proper. However, to attain a measure of universality or uniformity of data, at least
at the earlier stages of stress corrosion testing of new or unusual alloys,
the variety of environments—particularly the laboratory test media—
might be selectively limited. Even on the basis of the incomplete and
varying experimental data collected, a few of the environments show
some consistency, specificity, and aggressiveness of stress corrosion attack
of high-strength steels.
Because of the variability of natural atmospheres from day-to-day,
season-to-season, and year-to-year, tests of materials conducted in natural atmospheres should be repeated frequently enough and extended
over sufficiently long periods of time to ensure acquiring adequate and
dependable data. Frequently, test information obtained under laboratory
controlled conditions are used to estimate or predict the service-life
expectancy of a material. Sometimes the performance of materials under
operating or natural exposure conditions fall severely short of such
forecasts.
3
Data supporting the statements citing Ref 9 are not included in the tabular
compilation.
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GALLACCIO AND PELENSKY ON HIGH-STRENGTH STEEL ALLOYS
105
In selecting a controlled environment for stress corrosion testing, as
well as for general corrosion testing, the state or combinations of states
of the environment, the concentrations of constituents in the medium,
pH, density, surface tension factors, and temperature must be taken
into consideration. It might be important to ascertain the role of a single
factor, or of a combination of factors.
Reasonable and functional criteria for corrosion testing methods, reproducibility of test results, and suitable confidence levels of reliability
of the data should be established in order that some measure of standardization might be achieved. In addition to the factors mentioned above
regarding environment, other factors must be acknowledged so that useful and reliable standardization might be realized. Essentially the factors
to be considered regarding the specimen include sampling, the form or
shape, grain orientation or direction of flow, the applied stress relative
to the grain orientation, the level and mode of stressing, the rate of
applied strain, and the number of replicates. It is interesting to note that
specific information concerning these factors, in many cases, is not included in the published reports. In very many cases the data are highly
scattered, and the replicates are few.
The usual wide scatter of results associated with stress corrosion testing may be attributed in part to the inherent variability of the parent
and specimen metal microstructure, composition, processing, and residual
stresses. Sufficiently large numbers of replicates and statistical treatment
of the test data should effectively help to establish an enhanced confidence in stress corrosion testing results. Further such treatment
of the data, with controlled variations of the individual factors associated
with the specimens, environments, and test methods, will expose those
factors of predominant influence.
References
[1] "Stress Corrosion Cracking of High Strength Alloys," 0414-01-5 (Quarterly),
Jan. 1962, Contract DA-04-495-ORD-3069, Aerojet, Asuza, Calif.
[2] Phelps, E. A. and Loginow, A. W., "Stress Corrosion of Steels for Aircraft
and Missiles," Corrosion, Vol. 16, July 1960.
[3] Dean, S. W. and Copson, H. R., "Stress Corrosion Behavior of Maraging
Nickel Steels in Natural Environments," Corrosion, Vol. 21, March 1965.
[4] "Stress Corrosion Cracking of High Strength Alloys," 2914 (Final), Aug.
1964, Contract DA-04-495-ORD-3069, Aerojet, Asuza, Calif.
[5] Dreyer, G. A. and Gallaugher, W. C., "Investigation of the Effects of Stress
Corrosion on High Strength Steel Alloys," Technical Documentary Report
ML-TDR-64-3, Feb. 1964, Air Force Materials Laboratory, Wright-Patterson
Air Force Base, Ohio, AD 605 672.
[6] Batt, H. J. and Phelps, E. H., "Effect of Solution pH on the Mechanism of
Stress Corrosion Cracking of a Martensitic Stainless Steel," Corrosion, Vol.
19, Sept. 1961.
[7] "Stress Corrosion of High Strength Steels and Alloys; Artificial Environment,"
Research Project 389-2 (Final), July 1960 to July 1962, Mellon Institute,
Pittsburgh,
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106
STRESS CORROSION TESTING
[8] Davis, R. A., Dreyer, G. A., and Gallaugher, W. C., "Stress Corrosion Cracking Study of Several High Strength Steels," Corrosion, Vol. 20, March 1964.
[9] Hildebrand, J. F., "Stress Corrosion Cracking of High Strength Nickel Alloys
for Aircraft Applications," Materials Protection, Vol. 3, Sept. 1964.
[10] Hildebrand, J. F., Turns, E. W., and Nordquist, F. C., "Stress Corrosion
Cracking in High Strength Ferrous Alloys," Materials Protection, Vol. 2,
Nov. 1963.
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M. A. Pelensky1 and A. Gallaccio1
Stress Corrosion of Magnesium
Alloys—Environmental Factors
REFERENCE: M. A. Pelensky and A. Gallaccio, "Stress Corrosion of
Magnesium Alloys—Environmental Factors," Stress Corrosion Testing,
ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 107.
ABSTRACT: Several years ago, members of ASTM Committee B-3,
Subcommittee X (now Committee G-l, Subcommittee VI) on Stress Corrosion determined a need for review of work on stress corrosion testing
prior to development of standards and criteria for determining stress
corrosion susceptibility and behavior of various alloys. The subcommittee
established three task groups to accomplish a review of the subject. Task
group assignments were as follows: TG-1, sample selection and method of
stress; TG-2, environments under which stress corrosion occurs and duration of tests; and TG-3, reporting of stress corrosion results and interpretation of results. This report is part of the TG-2 assignment. It concerns a review of reported stress corrosion investigations relating to the
effects of various environments on specific magnesium alloys.
KEY WORDS: corrosion, stress corrosion, magnesium alloys, environmental testing
A survey is presented of reported investigations concerning environmental effects on stress corrosion failure of a variety of magnesium alloys.
Reported test results were tabulated by environments (airs, waters, and
solutions) with alloy compositions, strengths, stress levels, forms, direction, and methods of stressing where this information was reported. Informational sources are cited.
Comments and observations concerning selection of environments,
test methods, sampling, and results are presented. Some suggestions are
offered.
Environments and Materials
The conditions of test and the materials covered in this survey are as
follows:
1
Research chemist and chief, respectively, Protection and Preservation Branch,
Copyright by ASTM
Int'l (all
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Wed DecArsenal,
16 15:53:43
EST 2015 Pa.
Pitman-Dunn
Research
Laboratories,
Frankford
Philadelphia,
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108
STRESS CORROSION TESTING
Environment
(Unless otherwise indicated, ambient temperature and continuous,
total immersion of the specimen! apply.)
Air:
Indoor, ambient RH and temperature
Indoor, 85 to 90% RH
Indoor, 90% RH, ambient temperature; O2 added; SO2 added; CO2
added
Indoor, 95% RH, ambient temperature; O2 added; SO2 added; CO2
added
Indoor, 100% RH, ambient temperature; O2 added; SO2 added; CO2
added
Outdoor, marine
Outdoor, NBS (National Bureau of Standards)
Outdoor, rural
Waters:
Distilled (partial and total immersions)
Sea
Solutions:
NaCl (0.58 g/1) + K2Cr207
CsCl (101 g/1)
(0.15 to 14.9 g/1)
HF (0.20 g/1)
HN03 (0.63 g/1)
NaCl (29 g/1) + K2Cr207 (14.9
H2SO4 (0.010 to 12.25 g/1)
g/1)
KC1 (45 g/1)
NaCl (35 g/1) + K2CrO4 (1 to
KHF2 (5 g/1) (partial immer50 g/1)
sion)
NaCl (35 g/1) + K2Cr04 (20
K2CrO4 (20 g/1) + NaCH3COO
g/1)
(82 g/1)
NaCl (58 g/1) + K2Cr04 (20
K2Cr04 (20 g/1) + Na2C03
g/1)
(53 g/1)
NaCl (40 to 220 g/1) + K2CrO4
K2CrO4 (20 g/1) + NaNO3
(5 g/1)
(85 g/1)
NaCl
(200 g/1) + K2Cr04 (5 to
K2Cr04 (20 g/1) + Na2SO.
200
g/1)
(71 g/1)
Na2CO3 (0.265 to 53 g/1)
NaCH3COO (82 g/1)
NaF (0.42 g/1)
NaBr (62 g/1)
Nal (90 g/1)
NaCl (0.058 to 58 g/1)
NaN03 (85 g/1)
NaCl (0.1 g/1) (intermittent
NaOH (0.40 g/1)
immersion)
(0.71 to
712015
g/1)
NaCl(35g/l)
H(all
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2S0
2 S0
4 (4.9g/l)
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TABLE I—Sample of data.
A»<*
sJeSlsi
Test
Specimen
Stressed,
% yield
]Direction
of Stress
NACL (40 TO 220
AZ61 (AM-C57S-H)
dm
dm
dm
dm
dm
f
f
f
f
f
wr
wr
wr
wr
wr
>100
>100
>100
>100
>100
G/L)
Stress
Method
dm
dm
dm
dm
dm
f
f
f
f
f
wr
wr
wr
wr
wr
100
100
100
100
100
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
Ref
+ K 2 CRO 4 (5 G/L)
NA€L (200 G/L) --1- K2CRO 4
AZ61 (AM-C57S-H)
No. Specimens, time-to-failure
bb
bb
bb
bb
bb
(5 TO
bb
bb
bb
bb
bb
dm 0 g/1 NaCl nf 980 hr
dm 40 g/1 NaCl 17 hr (e)
dm 80 g/1 NaCl 6 hr (e)
dm 160 g/1 NaCl 3 hr (e)
dm 220 g/1 NaCl 5 hr (e)
[in
200 G/L)
dm
dm
dm
dm
dm
5 g/1 K2CrO4 2 hr, 12 hr (e)
20 g/1 K 2 CrO 4 0.07 hr (e)
50 g/1 K 2 CrO 4 0.07 hr (e)
100 g/1 K 2 CrO 4 0.04 hr (e)
200 g/1 K 2 CrO 4 0.02 hr (e)
dm
dm
dm
dm
dm
dm
0.265 g/1 21 hr (e)
0.53 g/1 23 hr (e)
5.3 g/1 22 hr (e)
7.95 g/1 24 hr (e)
15.9 g/1 23 hr (e)
53 g/1 12 hr (e)
(in
NA2C03 (0.265 TO 53 G/L)
AZ61 (MA3)
Abbreviations:
dm = data missing
f wr = flat wrought
bb
= bent beam
Copyright by ASTM Int'l (all
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
nf = no failure
hr = hour(s)
e = estimated from graphs
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
dm
rights reserved); Wed Dec 16 15:53:43 EST 2015
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[9]
110
STRESS CORROSION TESTING
Alloys
The magnesium alloys reported in this review include the following:
Ml
AZ31
AZ31, Ml clad
AZ51
AZ61
AZ80
ZK-60A-T5
Mg (high purity)
Mg + 5A1
MWY1 (5A1, 0.3Mn, O.OBFe)
MWY3 (5A1, 0.3Mn, 0.0019Fe)
Mg + 8A1
Mg + l.SIMn, 0.15 to O.SCe
Mg + 2Mn
MA2
(3 to 4A1' °-2
0.15 to O.SMn)
Mg + 14Li
to
°-8Zn>
Data
The tabulated data compiled in this survey are available through the
authors. Table 1 is a sample of the data.
Discussion Concerning the Data
Air Environments
No failures of ZK-60A-T5 alloy are reported for ambient indoor air
exposures [I].2
In indoor air at ambient temperatures, the "critical" relative humidity
at which AZ61 alloy undergoes cracking, is approximately 98 to 100
per cent [2]. Increasing concentrations of O2 or CO2 in the air decreases
the "critical" relative humidity to 95 per cent, while increasing the concentration of SO2 has no apparent similar effect [2]. When the relative
humidity is close to 100 per cent, the presence of significant amounts of
SO2 and CO2 decrease the tune to cracking [2].
The incidence of cracking rises during periods of rain or of high
humidity and temperature [3]. The seasonal time period of exposure is
an important factor to be considered in attempting to compare results of
specimens exposed outdoors [3].
AZ31, clad and unclad, and AZ61 alloys exposed in marine environments are shown to fail by stress corrosion [3,4,5], Results, however,
do not indicate any failures of the Ml alloy in this environment [3].
The AZ31, AZ51, AZ61, and AZ80 alloys fail, after different periods
of exposure outdoors at the NBS site, Washington, D. C. [4\. Comparison
of AZ31 test results suggests that the outdoor exposure environment at
the NBS site is more severe than the marine site at Hampton Roads, Va.
M2
The italic
numbers
refer toWed
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PELENSKY AND GALLACCIO ON MAGNESIUM ALLOYS
111
Results indicate stress corrosion failures of AZ31 and AZ61 specimens exposed to outdoor rural environment in the vicinity of Midland,
Mich. [3].
Time-to-failure results of AZ61 specimens exposed at Kure Beach
(800 ft from ocean) are shown comparable to those exposed to the
rural-industrial Michigan environment [3].
Waters
ZK 60A-T5 alloys fail in very short periods of time when exposed to
sea water [7]. The sea water exposure is much more severe than exposure
to distilled water [7].
The Mg + 5A1 alloys fail when partially immersed in distilled water,
usually at or just below the water level [6].
Solutions
ZK 60A-T5 alloy stressed at 90 per cent of the yield strength failed
rapidly (in 7 min or less except for one specimen) in 0.6 N solutions of
KC1 (45 g/1), CsCl (101 g/1), NaBr (62 g/1), NaCl (35 g/1), and Nal
(90 g/1) [7].
As reported, the time preceding cracking of AZ61 (MA3) alloy in
NaCl (35 g/1) + K2CrO4 (20 g/1) decreases as temperature is increased
from 41 to approximately 104 F [7]. However, on further increase of
temperature the time preceding cracking increases, and at temperatures
above 104 to 122 F the alloy acquires a relatively high resistance to stress
corrosion cracking [7]. The time-to-cracking of MA2 alloy in NaCl (35
g/1) + H2SO4 (4.9 g/1) is shown to decrease with increasing temperature
from 50 to 158 F [7,8]. This relationship is also shown for AZ80 (MA5)
[7].
In Na2CO3 , susceptibility of AZ61 (MA3) alloy to stress corrosion
failure remains almost constant as the concentration of Na2CO3 increases
from 0.265 to 15.9 g/1 but increases appreciably with concentration increases to 53 g/1 [9].
Results indicate that Mg (high purity) and Mg + 2Mn alloys fail by
stress corrosion when partially immersed in KHF2 (5 g/1) [6].
No failures of AZ61 (MA3) alloy were reported to occur in 260 hr in
the following solutions [70]:
2
KCrO4 (20 g/1) + NaCH3COO (82 g/1)
2Cr04 (20 g/1) + NaC03 (53 g/1)
K
2Cr04
K
(20 g/1) + NaC03 (53 g/1)
A decreasing order of susceptibility of AZ61 (MA3) to stress corrosion
failure in various solutions is indicated by the time-to-failure data as folCopyright
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STRESS CORROSION TESTING
NaCl (58 g/1) + K2Cr04 (20g/l), 4 min
K2CrO4 (20 g/1) + Na2SO4 (71 g/1), 9 min
Na2SO4 (71 g/1), 40 hr
NaN03 (85 g/1), 43 hr
Na2CO3 (53 g/1), 107 hr
NaCl (58 g/1), 265 hr
NaCH3COO (82 g/1), 650 hr
Time-to-stress corrosion failure of AZ61 (AM-C57S-H) is shown to
decrease significantly with increasing concentrations of K2CrO4 (3 to 200
g/1) in NaCl (35 g/1) solution [11]. Similarly, increasing concentrations
of NaCl (40 to 200 g/1) in K2CrO4 (5 g/1) solution result in increasing
susceptibility of AZ61 to stress corrosion failure [11].
Increased susceptibility to stress corrosion failure of AZ61 (MA3) and
Mg + 1.5Mn alloys with increasing concentration of NaCl solutions is
also indicated [72,73].
Stress corrosion failures of various AZ31 and AZ61 alloys exposed to
NaCl (35 g/1) + K2CrO4 (20 g/1) are shown in the tabulated data [3,4,77,
14].
Alloy AZ61 (AM-C57S-H) immersed in NaCl (35 g/1) + K2CrO4 (20
g/1) fails rapidly (8 min or less) by stress corrosion at pH 2 through pH
12; however, at pH 13 no failure is reported after 980 hr exposure [77].
Alloys
The susceptibility to stress corrosion of various magnesium alloys increases with increasing stress [3,4,6,13,14].
Threshold stress levels are suggested for various alloys in certain exposures, for example, AZ61 in NaCl (35 g/1) + K2CrO4 (20 g/1) at
approximately 60 per cent of the yield strength when tested by the bentbeam method [14].
AZ31 sheet when subjected to constant axial stress in a rural environment appears to have a stress corrosion threshold strength in the neighborhood of 20 ksi [3]. However, it is reported that under bending stress
and exposed to marine environment the minimum stress to produce fracture is 25 ksi [3].
AZ61 is shown to be more sensitive to stress corrosion failure than
AZ31. Threshold limits under axial loading appear to be of the order of
12 to 15 ksi [3]. Under constant bending, the threshold limit appears to
be approximately 15 to 20 ksi [3].
At moderate stresses, resistance to cracking of wrought magnesium
alloys (AZ31, AZ51, and AZ61) decreases with increasing aluminum
content
up to approximately 6.5A1 [4].
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PELENSKY AND GALLACCIO ON MAGNESIUM ALLOYS
113
Ml clad AZ31 alloy is shown to be more resistant to stress corrosion
failure in marine environment at high stresses than unclad AZ31 [4]; no
failure occurred after approximately 500 days for Ml clad AZ31 stressed
to 30 ksi compared to failure at approximately 214 days for unclad AZ31
stressed to 20 ksi.
Susceptibility to stress corrosion of Mg + 5A1 alloys in distilled water
appears to increase with increase in iron content [6].
Results indicate that stress corrosion susceptibility of Mg-Li alloys in
humid air is associated with aluminum content and with the condition
resulting from rapid cooling of these alloys from high temperatures [75].
Heating and stabilizing at 300 F appears to result in stress corrosion resistance in these same alloys [75]. Mg-Li alloys with no significant aluminum content withstand this humid air exposure under the various conditions tested [75].
Hardened AZ31 alloys appear less susceptible to stress corrosion failure
than do the annealed alloys exposed to the NBS site [4].
AZ61 (Jlh) alloy annealed at 350 F appears to have greater resistance
to stress corrosion at a higher per cent yield strength in marine environment than do the cold-rolled (J1CR) and hot-rolled (Jlr) AZ61 specimens
and also the AZ31 (FS1) alloys [3].
Stress Methods
The stressing methods reported on include U-bend, bent beam, direct
tensile, spring load, and eccentric loading. Where comparisons can be
made, it appears that an axial tensile stress method may be more severe
at lower stress levels than bent beam, for example, AZ31 (FSlh) and
AZ61 (Jla) alloys exposed in a marine environment [3].
Observations
A number of environments with their effects on a variety of magnesium
alloys have been investigated. The alloys reported include those of different strengths and those which were subjected to different applied stresses
(from approximately 11 to 43 ksi yield stress and stresses applied up to
approximately 135 to 140 per cent of yield).
It may be observed from the data that most of the environments (including airs, waters, and solutions) selected for testing result in stress
corrosion failure of the magnesium alloys. The environments selected in
some instances are intended to duplicate or at least closely represent service conditions. Other selections were made to accelerate testing but were
intended also to aid in predicting actual service life.
Results in some cases lead to specific conclusions regarding stress
corrosion
failure,
other
results
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latter
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114
STRESS CORROSION TESTING
situation may be due to the wide scatter of data in a number of instances,
or the small number of specimens employed in testing, or to both.
The large variations or scatter in some test results could be due to: differences in the alloy (variations in the as-received condition or resulting
from preparation of specimens for testing), differences in environment
(contamination, concentration, or temperature gradients, seasonal time
of exposure, etc.), or the test method employed.
It is difficult to attempt to make reliable predictions of stress corrosion
service life based solely on accelerated testing results. Relative susceptibilities to stress corrosion failure are frequently estimated from laboratory
tests. Caution should be used in attempting to relate accelerated test results with results of outdoor exposure and in predicting service life of materials based on results of exposure to outdoor environments. The latter
reflect differences in test results attributable to daily and seasonal variations of temperature, humidity, and rain conditions.
In setting up any stress corrosion test to evaluate environmental effects,
many factors must be considered. The aims or purpose of the test should
be clearly established prior to initiating the test. The alloy or alloys selected
and their mechanical and metallurigical variables (composition, structure,
form, treatment, strength, level, and types of stress) must be considered.
The rate of strain also may be an important factor to consider. The specific environment or environments to which the material is to be subjected
must be determined, including, as applicable, consideration of temperature, relative humidity, composition and concentration, and pH. Total
or partial, continuous or intermittent immersion, and static or dynamic
conditions may be involved. Finally, it is important that the experiments
be designed to facilitate statistical analysis of the test results.
References
[1] Crossley, F. A., "Research on the Basic Nature of Stress Corrosion for Various
Structural Alloys at Room and Elevated Temperature," ASD-TR-61-713, May
1962, Contract AF 33(616)-7612, Armour Research Foundation for Directorate of Materials and Processes, Aeronautical Systems Div., Air Force Systems Command, Wright-Patterson Air Force Base, Ohio.
[2] Timonova, M. A., "Corrosion Cracking of Magnesium Alloys and Methods
of Protection Against It," Intercrystalline Corrosion and Corrosion of Metals
Under Stress, Levin, I. A., ed., Consultant Bureau Enterprises, Inc., 1962.
[5] Loose, W. S. and Barbian, H. A., "Stress Corrosion Testing of Magnesium
Alloys," ASTM-A1ME Symposium on Stress Corrosion Cracking of Metals,
1944, American Society for Testing and Materials, Philadelphia, 1945.
[4] Logan, H. L. and Hessing, Harold, "Stress Corrosion of Wrought Magnesium
Base Alloys," Journal of Research of the National Bureau of Standards, Vol.
44, March 1950.
[5] Romanov, V. V., "Corrosion of Magnesium," Translation No. 1-8207, 11
April, 1963 (refers to Romanov, V. V., Corrosion of Magnesium, Academy of
Sciences, USSR, 1961).
[6] Ferryman, E. C. W., "Stress Corrosion of Magnesium Alloys," Journal of the
Institute of Metals, Vol. 78, Sept. 1950 to Feb. 1951.
Copyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
[7] Romanov, V. V., "Corrosion of Magnesium," Translation No. 1-8207, 11
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DISCUSSION ON MAGNESIUM ALLOYS
115
April, 1963 (refers to collection of papers, Strength of Metals, IME T,
Academy of Sciences, USSR, published by USSR Academy of Sciences,
Moscow, 1956).
[8] Romanov, V. V., "Effect of Some Factors on the Susceptibility of Magnesium
Alloys to Corrosion Cracking," Inter crystalline Corrosion and Corrosion of
[9]
[10]
[11]
[12]
[13]
[14]
[75]
Metals Under Stress, Levin, I. A., ed., Consultant Bureau Enterprises, Inc.,
1962.
Romanov, V. V., "Corrosion of Magnesium," Translation No. 1-8207, 11
April, 1963 (refers to Zaretsky, E. M., Zhurnal Prikladnoy Khimii, 1951).
Romanov, V. V., "Stress Corrosion Cracking of Metals," Israel Program for
Scientific Translations, 1961, (refers to Romanov, V. V., Sbornik, IME T,
AN SSR, 1958).
Sager, G. F., Brown, R. H., and Mears, R. B., 'Tests for Determining Susceptibility to Stress-Corrosion Cracking," ASTM-AIME Symposium on Stress
Corrosion Cracking of Metals, 1944, American Society for Testing and
Materials, Philadelphia, 1945.
Romanov, V. V., "Corrosion of Magnesium," Translation No. 1-8207, 11
April, 1963 (refers to Zaretsky, E. M., Doklady Akademii Nauk SSSR, 1947).
Tomashov, N. D. and Modestova, V. N., "Effect of Stress on the Corrosion
and Potentials of Alloys of the Magnesium-Manganese System," Intercrystalline Corrosion of Metals Under Stress, Levin, I. A., ed., Consultant Bureau
Enterprises, Inc., 1962.
Mears, R. B., Brown, R. H., and Dix, E. H., Jr., "A Generalized Theory of
Stress Corrosion of Alloys," ASTM-AIME Symposium on Stress Corrosion
Cracking of Metals, 1944, American Society for Testing and Materials, Philadelphia, 1945.
Kiszka, J. C., "Stress Corrosion Tests of Some Wrought Magnesium-Lithium
Base Alloys," M65-1-1, July 1964, Frankford Arsenal, Philadelphia, Pa.;
Materials Protection, Feb. 1965.
DISCUSSION
F. A. Smith1 (written discussion)—I would appreciate your consideration of the possible heat-to-heat variation in chemical composition of
magnesium alloys as a possible way to explain the large variation in timeto-failure for specimens discussed in your paper.
M. A. Pelensky and A. Gallaccio (authors)—It is expected that heatto-heat differences in chemical composition could account for some of the
variation in time-to-failure, since, as reported in this paper, variations in
chemical composition of somewhat different alloys and even different heat
treatments of the same alloy can result in different times-to-failure. However, the degree of influence of heat-to-heat variation in chemical composition is difficult to determine in view of other factors which can result in
variations in time-to-failure. These other factors which must be considered
include variations in physical microstructure, differences in concentration,
contamination, temperature, and pH of the corrosive medium, and also
nonuniformity of stresses, both residual and applied.
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1
Physicist, Argonne
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A. P. Bond,1 J. D. Marshall? and H. J. Dundas*
Resistance of Ferritic Stainless Steels to
Stress Corrosion Cracking
REFERENCE: A. P. Bond, J. D. Marshall, and H. J. Dundas, "Resistance
of Ferritic Stainless Steels to Stress Corrosion Cracking," Stress Corrosion
Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 116.
ABSTRACT: Stress corrosion cracking tests were carried out on 0.020in.-diameter wires under dead-weight loading in boiling solutions of
MgCU , Ca(NO3)2 , and NaOH. The T430 and T434 ferritic stainless steels
were immune to stress corrosion cracking in MgCU under conditions which
caused various austenitic grades to fail in 6 to 46 min. Pitting corrosion
in MgCl2 reduced the cross-sectional area of the ferritic steel specimens
to the extent that they could not support the load, and they failed in a
ductile manner. Stressed specimens of ferritic stainless steels did not
fracture in Ca(NOs)2 or NaOH even after heat treatments which led to
intergranular corrosion. Corrosion potential measurements were made
on the stressed specimens and correlated with various stages in the stresscorrosion cracking process.
KEY WORDS: corrosion, stress corrosion, steels, stainless steels, pitting,
heat treatment
The sensitivity of austenitic stainless steels to stress corrosion cracking
in chloride-containing environments often limits the use of these steels.
Since it has been reported that Type 430 ferritic stainless steel is highly
resistant to cracking even in the presence of chlorides [./,2],4 the present
work was undertaken to ascertain the combined effects of stress and corrosive environments on Types 430 and 4345 ferritic stainless steels. In
addition to boiling magnesium chloride (MgCl2) solutions, boiling solutions of calcium nitrate (Ca(NO3)2) and of sodium hydroxide (NaOH)
1
Research group leader, Climax Molybdenum Company of Michigan, Ann
Arbor, Mich.
2
Formerly with Climax Molybdenum Company of Michigan as research associate; now assistant metallurgist, Metallurgical Dept., W. S. Tyler Co., Cleveland,
Ohio.
3
Laboratory associate, Climax Molybdenum Company of Michigan, Ann Arbor,
Mich.
4
The italic numbers in brackets refer to the list of references appended to this
paper.
B
Type 434 is the proposed designation for stainless steel containing 17 per cent
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chromium
+ 1 per cent molybdenum.
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BOND ET AL ON RESISTANCE OF FERRITIC STAINLESS STEELS
117
were used as test environments, since low-alloy steels are known to undergo stress corrosion cracking in hot nitrates or caustics [3].
Test Procedure
Materials and Heat Treatment
The materials evaluated during this investigation were commercial heats
of stainless steels; their compositions are presented in Table 1. All of the
steels were evaluated in the annealed condition. The annealing treatment
for the austenitic steels consisted of heating the wires at 1950 F for 15
TABLE 1—Composition of alloys (weight per cent).
Alloy Type
c
201 + 1.7 Cu...
301
304
316
430, Heat A . . . .
430, Heat B
434, Heat A
434, Heat B
434, Heat C
0.095
0.045
0.058
0.074
0.043
0.065
0.066
0.059
0.077
0.069
201
Mn
6.32
6.50
0.60
0.75
1.57
0.49
0.40
0.49
0.51
0.42
Si
0.44
0.56
0.43
0.63
0.47
0.34
0.36
0.47
0.38
0.54
S
p
0.007
0.009
0.002
0.005
0.015
0.007
0.011
0.010
0.006
0.007
0.029
0.030
0.017
0.024
0.029
0.015
0.017
0.018
0.022
0.016
Cr
Ni
Mo
N2
Cu
16.53 4.83
0.098
0.039 1.73
16.50 5.40
17.56 7.19 0.23 0.037 0.189
18.38 9.40 0.27
17.52 13.05 ? 16
17.06 0.29 0.07
16.00 0.24 0.20
17.71
1.01
17.37
1.01
0.11
16.52 0.30 0.99
TABLE 2—Strength properties of the 0.020-in.-diameter ferritic
stainless steel wires, psi.
Type 430, Heat B
Condition
Annealed
1 hr, 900 F
15 min, 1800 F
Type 434, Heat A
Proportional
Limit
Tensile
Strength
Proportional
Limit
Tensile
Strength
35 000
40 000
92 000
80 000
80 000
134 000
52 000
45 000
78 000
91 000
90 000
135 000
min; the ferritic steels were annealed by heating at 1500 F for 1 hr. The
ferritic steels were also evaluated after sensitizing treatments of 1 hr at
900 F, or 15 min at 1800 F.
The proportional limit and the tensile strength of the ferritic steels in the
three different conditions are presented in Table 2.
The majority of the tests were carried out on 0.020-in.-diameter wires
of the steels. However, some tests on the Type 434 material were conducted on 0.175-in.-thick strip specimens.
Stress Corrosion Tests
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118
STRESS CORROSION TESTING
for use with Ca(NO3)2 and with MgCl2 environments. A similar cell, constructed of Teflon, was used for the tests with NaOH and for additional
experiments with MgCl2 environments.
The corrosion media were: MgCl2 solution, the concentration of which
was adjusted to yield a boiling point of 284 F (140 C); 55 per cent
FIG. 1—Stress corrosion test cell for wire specimens.
TABLE 3—Failure times for 0.020-in.-diameter wire specimens of the
various steels tested in MgClz solution boiling at 284 F (140 C).
4900 psi Stress
Alloy
Treatment
701
201 + 1.7 Cu..
301
304
316
430, Heat A
430, Heat B
434, Heat A
434, Heat A
434, Heat A
434, Heat B
annealed
annealed
annealed
annealed
annealed
annealed
annealed
annealed
1 hr, 900 F
15 min, 1800 F
15 min, 1800 F
0
6
c
StandAvg Time- No.
ard
of
Devito-Failure.
hr
Tests ation,
hr
0 .35
0 .37
1.53
1.68
2 .92
42 NF"
42 NF°
3
3
5
7
5
1
1
0.048
0.026
0.535
0.588
0.400
53,000 psi Stress
Avg Time- No.
of
to-Failure,
hr
Tests
0. P
0. P
0. 475
0. 436
0. 776
1. 82^
1.98'
45. 6'
60. 5C
327"
297'
3
3
2
63
6
3
1
1
1
1
1
Standard
Deviation,
hr
0.0
0.0
0.106
0.042
0.103
0.31
NF = no failure in indicated exposure time.
Failure by stress corrosion cracking.
Ductile failure by pitting corrosion.
Ca(NO3)2 , boning at 242 F (117 C); and 25 per cent NaOH, boiling at
232 F (111 C). The MgCl2 solution was reused for up to ten tests with the
austenitic stainless steels. The solution was checked from time-to-time by
running a Type 304 specimen at 53,000 psi; a failure time near 26 min
was taken as indicating that the solution was suitable for use. In all tests
of the ferritic steels, a fresh solution was used for each experiment. During
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BOND ET AL ON RESISTANCE OF FERRITIC STAINLESS STEELS
119
some of the tests in MgCl2, potential measurements were made on the
stressed specimens. For these potential measurements, a saturated calomel
electrode was inserted in the side arm shown in Fig. 1. The potential as a
function of time was autographically recorded.
In addition to the runs on wire specimens, a few tests were performed
on Type 434 steel using U-bend specimens made by bending l/z by 2J/2 by
0.175-in. strips over a 3/s-in. mandrel. The legs of the bend were secured
by a steel bolt. The bolt and the adjacent areas of the specimen were
covered with neoprene, and the U-bend was suspended in the solution in
such a way that the coated bolt was not immersed in the liquid.
FIG. 2—Longitudinal cross section of Type 434 Heat A specimen heat treated
15 min at 1800 F and exposed 327 hr in boiling MgClz at a stress of 53,000 psi.
Unetched (X500).
Results and Discussion
The results of stress corrosion tests in boiling MgCl2 are shown in Table
3. Note that no stress corrosion cracking occurred in the ferritic steels,
Types 430 and 434. The failures which did occur were of the ductile type,
resulting from reduction of cross-sectional area due to pitting attack.
Metallographic examination did not reveal any cracks in the failed specimens. The photomicrograph (Fig. 2) shows the appearance of a typical
failure zone in ferritic specimens.
Both ferritic steels were subject to pitting in this aggressive environment, but the molybdenum-containing Type 434 was much more resistant
to pitting, as evidenced by its time-to-ductile failure being greater than that
of the straight chromium steel. The Type 434 wires tested in the glass cells
failed at the point where the wire entered the side arm. This suggests that
a form of crevice attack was accelerating the pitting failure. The tests were
repeated in Teflon cells hi which the wire entered a Teflon plug side arm
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120
STRESS CORROSION TESTING
which fitted the wire tightly, thus minimizing the crevice attack. The Teflon
cell increased the time-to-failure due to pitting. All results shown in Table
3 for the ferritic steels were obtained in the Teflon cell.
The heat treatments for the ferritic steels were selected to give several
degrees of sensitivity to intergranular corrosion. The 1800 F treatment is
known to sensitize ferritic steels to intergranular attack [4]. The 900 F
treatment sensitizes Type 434 to a lesser extent but has little effect on Type
430, according to work in progress at this laboratory. Of course, the
ferritic steels in the annealed condition exhibit the minimum sensitivity
to intergranular corrosion.
The depth of pitting of the ferritic steels was also affected by heat
FIG. 3—Potential versus time for stainless steel specimens stressed at 53,000 psi
while immersed
in
at 284
MgCl
F (140 C).
2 solution boiling
treatment. The 1800 F treatment caused a great increase in the time-tofailure (due to pitting) in MgCl2 . Apparently, this was mainly due to the
initiation of a much larger number of pits on the sensitized Type 434, with
the individual pits then growing more slowly than on the annealed steel as
a result of the reduced cathode-to-anode area ratio. Thus, the large number of pitting sites introduced by heating at 1800 F had the effect of distributing the corrosion more uniformly relative to the annealed steel. Of
course, the greater tensile strength of the specimens heat treated at 1800
F (Table 2) would also serve to increase the time-to-failure by pitting.
However, this effect is probably too small to account for all of the observed
increase in time-to-failure.
Although the failures of the ferritic stainless steels under dead-weight
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loading
appeared to be completely ductile and gave no indication of stress
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TABLE 4—Results of exposing 0.020-in.-diameter specimens of ferritic steels to boiling 55 per cent Ca(NOz)z and
to boiling 25 per cent NaOH at a stress of 53,000 psi.
NaOH
Ca(N08)2
Alloy
430, Heat B
430, Heat B
430, Heat B
430, Heat A
434, Heat A
434, Heat A
434, Heat A
434, Heat B
Mild steel
0
Heat Treatment
Exposure
Time, hr
Remarks
Exposure
Time, hr
annealed
1 hr, 900 F
15 min, 1800 F
15 min, 1800 F
annealed
1 hr, 900 F
15 min, 1800 F
15 min, 1800 F
1500 F
336 NF°
very little corrosion
very little corrosion
intergranular attack
intergranular attack
very little corrosion
very little corrosion
intergranular attack
intergranular attack
failed intergranularly
at 30,000 psi
336 NF"
uniform corrosion
428 NF
intergranular attack
336 NF
355 NF
336 NF
336 NF
uniform corrosion
uniform corrosion
intergranular attack
intergranular attack
336 NF
336 NF
362 NF
336 NF
336 NF
408 NF
500 NF
20.7
NF = no failure in indicated exposure time.
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Remarks
122
STRESS CORROSION TESTING
corrosion cracking, it was desirable to confirm this lack of stress corrosion
cracking under conditions in which ductile failure would not occur. To do
this, U-bend specimens of annealed Type 434, Heat C, were exposed to
boiling MgCl2 solution. In this type of test, the maximum stress is somewhat greater than the yield strength but is not appreciably increased by
FIG. 4—Longitudinal cross section of Type 434 Heat A specimen heat treated
15 min at 1800 F and exposed 336 hr in boiling 25 per cent NaOH at a stress of
53,000 psi. Unetched (X500).
FIG. 5—Longitudinal cross section of Type 434 Heat B specimen heat treated
15 min at 1800 F and exposed 500 hr in boiling 55 per cent Ca(NO^ at a stress
of 53,000 psi. Unetched (X500).
pitting corrosion. Thus, ductile failure due to reduction of cross-sectional
area cannot occur. Metallographic examination of two ferritic stainless
steel specimens stressed as a U-bend and exposed for 75 days hi boiling
MgCl2 revealed no evidence of stress corrosion cracks.
All the austenitic stainless steels investigated were found to be very
susceptible to stress corrosion cracking under the test conditions used. The
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fact
that the austenitic alloys cracked in very short times (see Table 3)
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BOND ET AL ON RESISTANCE OF FERRITIC STAINLESS STEELS
123
emphasizes the importance of the absence of cracking in ferritic stainless
steels evaluated under identical conditions.
Potential-time relationships of the specimens stressed at 53,000 psi are
shown in Fig. 3. The behavior of the austenitic steels is similar to that reported by other workers [5,6]. The initial rise in potential has been associated with partial film repair, while the region of potential drop corresponds to the crack propagation stage [5]. The initial rise in potential was
not observed for the Type 201 steels tested at 53,000 psi, indicating that
crack propagation began almost immediately. At a stress of 4900 psi,
however, the initial potential rise was detected on all the austenitic steels.
The variation of corrosion potential with exposure time was quite different
for the ferritic steels. No peak in the corrosion potential was observed,
although the Type 430 specimen did show a small potential drop as ductile failure occurred.
The ferritic stainless steels were also tested in boiling 25 per cent NaOH
and in boiling 55 per cent Ca(NO3)2 . Both of these media are known to
cause intergranular stress corrosion cracking in plain carbon steels and in
low-alloy steels which have undergone certain heat treatments [7,8], The
results, summarized in Table 4, showed no failure of any of the test specimens. The specimens sensitized at 1800 F were subject to intergranular
corrosion in both NaOH and Ca(NO3)2 , while the annealed specimens
did not undergo intergranular corrosion. The extent of the intergranular
corrosion is illustrated in Figs. 4 and 5. That intergranular corrosion and
not stress corrosion cracking was the corrosion mechanism was confirmed
by exposing unstressed specimens heat treated at 1800 F to the test environments. A very similar type of attack occurred on the unstressed
specimens.
One annealed Type 434 specimen (Heat A) was tested in boiling 50
per cent NaOH. Very rapid uniform corrosion caused ductile failure in
7 hr, so no further tests were made in this medium.
It is not clear to what extent the resistance to stress corrosion cracking
of the ferritic stainless steels, Types 430 and 434, is due to their crystal
structure and to what degree to their chemical compositions. It has been
suggested that the greater number of possible slip planes in the body-centered-cubic lattice as compared to the face-centered-cubic structure prevents stress concentration on any one glide plane reaching as high values
in the body-centered-cubic as in the face-centered-cubic metals [9]. This
cannot be the complete explanation of the resistance of the ferritic stainless steels to stress corrosion cracking, because some body-centered-cubic
metals do undergo stress corrosion cracking under some conditions, for
example, mild steel in nitrates. Another possibility is suggested by the observations of others [10,11] that some austenitic stainless steels containing less than 1 per cent nickel are immune to cracking in MgCl2 (as are
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austenitic
stainless
steels
containing
large
of nickel).
Thus, it may
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STRESS CORROSION TESTING
be that the absence of nickel is responsible for the immunity of the ferritic
steels to stress corrosion cracking. In fact, recent work with some experimental alloys has shown that fully ferritic 18 to 25 per cent chromium
alloys containing as little as 1.1 per cent nickel do stress corrosion crack
in boiling MgCl2 solution.
Conclusions
1. Types 430 and 434 stainless steels do not undergo stress corrosion
cracking even when stressed above their yield points in the following solutions: (a) MgCl2 boiling at 284 F (140 C), (b) Ca(NO3)2 boiling at 242 F
(117 C), or (c) NaOH boiling at 232 F (111 C).
2. Type 430 stainless steel exhibits poor resistance to pitting corrosion
in boiling MgCl2 solution. An addition of 1 per cent molybdenum to this
steel (Type 434) increases the resistance to pitting corrosion.
3. Types 430 arid 434 steel heat treated at 1800 F are subject to intergranular corrosion in boiling Ca(NO3)2 and in boiling NaOH solutions.
References
[1] Scheil, M. A., "Some Observations of Stress-Corrosion Cracking in Austenitic
Stainless Alloys," ASTM-AIME Symposium on Stress Corrosion Cracking of
Metals, 1944, American Society for Testing and Materials, Philadelphia, 1945,
pp. 395-410.
[2] Scheil, M. A. et al, "First Report on Stress Corrosion, Corrosion Cracking
of Stainless Steel in Chloride Solutions," Supplement, Journal of the American
Welding Society, Oct. 1943, pp. 493-S-504-S.
[3] Logan, H. L., "Stress Corrosion Cracking in Low Carbon Steel," Physical
Metallurgy of Stress Corrosion Fracture, Rhodin, T. N. ed., Interscience,
New York, 1959, pp. 295-309.
[4] Baumel, A., "Relationship Between Heat Treatment and Corrosion Behavior
of 17 Cr Stainless Steels in Boiling Concentrated Nitric Acid," Archiv fuer das
Eisenhuttenwesen, Vol. 34, 1963, pp. 135-146.
[5] Hoar, T. P. and Hines, J. G., 'The Electrochemistry of the Corrosion and
the Stress-Corrosion Cracking of 18-8 Chromium-Nickel Steels in Hot
Aqueous Magnesium Chloride Solution," Proceedings, 8th Meeting, International Committee of Electrochemical Thermodynamics and Kinetics, Madrid,
1956, Butterworths, London, 1958, pp. 273-291.
[6] Van Rooyen, D., "Qualitative Mechanism of Stress Corrosion Cracking of
Austenitic Stainless Steels," Corrosion, Vol. 16, 1960, pp. 421t-429t.
[7] Long, L. M. and Uhlig, H. H., "Effect of Carbon and Oxygen in Iron on Stress
Corrosion Cracking in Nitrate Solution," Journal of the Electrochemical Society, Vol. 112, 1965, pp. 964-967.
[8] Uhlig, H. H. and Sava, J., "Effect of Heat Treatment on Stress-Corrosion
Cracking of Iron and Mild Steel," Transactions, American Society for Metals,
Vol. 56, 1963, pp. 361-376.
[9] Denhard, E. E., discussion, Physical Metallurgy of Stress Corrosion Fracture,
Rhodin, T. N., ed., Interscience, New York, 1959, p. 223.
[10] Copson, H. R., "Effect of Composition on Stress Corrosion Cracking of Some
Alloys Containing Nickel," Physical Metallurgy of Stress Corrosion Fracture,
Rhodin, T. N., ed., Interscience, New York, 1959, pp. 247-267.
[11] Riedrich, G. and Kohl, H., "Effect of Various Alloying Elements on the Susceptibility to Stress Corrosion of Austenitic Corrosion Resistant Steels," Bergund Hiittenmdnnische
108, 1963,
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DISCUSSION ON RESISTANCE OF FERRITIC STAINLESS STEELS
125
DISCUSSION
/. F. Boyce1 (written discussion)—Exposure to boiling 42 per cent
MgCl2 solution seems to have become a fad in stress corrosion testing. Is
it not possible to be seriously misled by test results in this very severe environment in predicting behavior of the subject materials in sea water, tap
water, or other mild environments, particularly those not containing
appreciable concentrations of chloride ions?
/. B. Marriott2 (written discussion)—The authors have shown the
superiority of certain ferritic stainless steels over austenitic stainless steels
in a number of test environments. In practice, however, the environments
to which these steels are exposed may be much less corrosive, typical
examples being chloride-contaminated high-pressure hot water and condensing steam. It would be desirable to know whether the results given
here can be used when assessing these or other materials for mildly corrosion service.
The authors also have referred to the effect of corrosion pit formation
and growth on the life of their test specimens. Do they consider that this
variable could be eliminated by the use of a precracked specimen, perhaps
of the type described by Beachem and Brown?
R. W. Staehte3 (written discussion)— Regarding possible objections to
the use of boiling MgCl2 as a test environment, it is obvious that it is not a
normal industrial fluid. However, the results obtained in this experiment
with respect to alloy behavior are in good general agreement with work
which we are conducting in very dilute chloride environments in the 200
to 300 C range. Furthermore, there is an enormous amount of information
available from work in this environment, and I think that there is great
value in having essentially an international standard for comparison
among laboratories.
A. P. Bond, J. D. Marshall, and H. J. Dundas (authors)—In reply to
Mr. Boyce, a very large amount of data on the behavior of austenitic
stainless steels in boiling MgCl2 has been obtained over many years. Thus,
this solution seemed to be a logical one to use in comparing the stress
corrosion cracking behavior of ferritic stainless steels with that of austenitic stainless steels. If chloride-induced stress corrosion cracking does
not occur in this solution, it is usually considered that such cracking will
not occur at lower temperatures or chloride-ion concentrations. Cer1
Metallurgist, Gillette Safety Razor Co., Gillette Park, Boston, Mass.
The English Electric Co., Ltd., Central Metallurgical Laboratories, Cambridge
Road, Whetstone, Nr. Leicester.
3
Assistant professor, Department of Metallurgical Engineering, Corrosion Center,
The Ohio
State University,
Columbus,
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STRESS CORROSION TESTING
tainly the occurrence of stress corrosion cracking in this severe test does
not prove that it would occur under milder conditions, but it does indicate
that it might.
To obtain some information about the behavior of ferritic steels in environments free of chloride, testing was conducted in both boiling 55 per
cent Ca(NO3)2 and in boiling 25 per cent NaOH. These environments were
chosen because stress corrosion cracking of sortie mild steels had been observed in them.
In reply to Mr. Marriott, the results in MgCl2 solution at 140 C do not
guarantee that cracking might not occur at higher temperatures and lower
chloride concentrations, especially under conditions in which the actual
chloride concentration at the metal surface may be much greater than it is
in the bulk environment. However, our results strongly suggest that the
examined ferritic stainless steels would be at least highly resistant to stress
corrosion cracking in the example cited. In this connection, Professor
Staehle's comments are of great interest.
Finally, it must be emphasized most strongly that pitting corrosion of
the ferritic, Types 430 and 434, stainless steels in MgCl2 solution led to
ductile failure when the reduction of the cross-sectional area caused the
constant load to exceed the tensile strength of the materials. No evidence
of stress corrosion cracking was found in this case. Thus the use of precracked specimens would serve no purpose.
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H. L. Logan1
Some Techniques Used in the Study of
Stress Corrosion Cracking
REFERENCE: H. L. Logan, "Some Techniques Used in die Study of
Stress Corrosion Cracking," Stress Corrosion Testing, ASTM STP 425,
Am. Soc. Testing Mats., 1967, p. 127.
ABSTRACT: Techniques used at the National Bureau of Standards in the
study of stress corrosion cracking of metals are described together with
precautions taken in these investigations. Especially designed specimens
of low-carbon and stainless steels and a titanium alloy and supplementary
techniques for obtaining data as to the mechanism of the stress corrosion
process are also described. A specimen and technique recently used to determine whether hydrogen plays a part in the delayed failures of highstrength steels in chloride solutions is described.
KEY WORDS: corrosion, stress corrosion, steels, titanium alloys, hydrogen embrittlement, chlorides, salt solutions, environmental testing, exposure testing
The study of the stress corrosion cracking of metals has been a more or
less continuous project at the National Bureau of Standards (NBS) for the
past 25 years and a continuing project since 1950. This paper describes
techniques that have been employed in the study of both the susceptibility
of certain materials to stress corrosion cracking and the mechanism of
cracking in certain alloy systems. Pitfalls that sometimes have not been
avoided are described and certain interesting results, some of which have
not been reported earlier, are mentioned.
Early work during the war years (1941 to 1945) was devoted to problems such as: (a) determining the effect of elevated temperature aging on
the resistance of the 2024 aluminum alloy to stress corrosion cracking [I],2
(b) determining the resistance of various high-strength aluminum alloys to
stress corrosion cracking [2], and (c) determining the relative resistances
of various magnesium alloys to stress corrosion cracking [3]. Subsequently, most of the investigations have had as their primary objective the
1
Physicist, National Bureau of Standards, Washington, D. C. Personal member
ASTM.
2
The italic
numbers
brackets
refer
the16list
of references
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STRESS CORROSION TESTING
study of the mechanism or mechanisms of stress corrosion cracking. These
objectives have on occasion required special types of specimens. The specimens used for particular investigations have been described in papers giving the results of those investigations, but there has been no general
discussion of specimen types, of the advantages or disadvantages of these
specimens, or of precautions that may be either necessary or desirable in
certain investigations. Similarly, specific methods of studying corrosion
FIG. 1—Specimens in weather exposure racks at the National Bureau of Standards, Washington, D. C. Unstressed specimens were exposed between stressed specimens. Note grip supports for specimens and crossed flexure plates in pull rods.
Common fulcrum for levers is seen on reverse side of rack.
products have been discussed in individual papers but not in a review paper.
Semiconventional Tension Specimens
Because it was desirable to know the initially applied stress, much of the
work has been done with specimens loaded in tension by means of lever
systems and weights.3 Because stresses in bent beams can only be approx3
This stress is superimposed on any residual stresses from heat treatment or cold
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LOGAN ON TESTING TECHNIQUES
129
imated, bent-beam specimens have been used only where practical considerations rule out other types of specimens. (Stresses in U-bend specimens
are unknown and for that reason they have not generally been used in our
investigations.)
Specimens Exposed in the Weather
Ideally each specimen should have its own complete lever system. Such
equipment is expensive, and some of the work at NBS has been done using
racks having one continuous knife edge used by several levers for each of
several specimens. A portion of a rack of this design used on the roof of
the Northwest Building at NBS Washington site is shown in Fig. 1.
Gripping devices for sheet specimens such as those used in tension testing are prohibitively expensive for multiple unit stress corrosion test racks.
Pin devices, however, were found to be reasonably inexpensive and were
satisfactory if properly designed and used (Fig. 1). The grip ends of Vs-hi.
reduced section sheet tension specimens were either 1 in. or for some materials 11A in. wide. Because of the reduced stress concentration, a specimen that would fail in the pin area when supported by a V^-in.-diameter
pin would not fail if a %6 or %-in. pin were used. The grip ends of pin
supported thin sheet specimens (up to possibly l/s in. thickness with some
alloys) must also be laterally supported to prevent failures at the pin holes.
Axial alignment of specimens to prevent twisting or bending is very important. It was found that a slight twisting of magnesium alloy specimens
in the corrodent produced failures either in the shoulders or the grip ends
of the specimens, rather than in the reduced section, and at lower stresses
than if specimens were properly aligned. The use of crossed chains or
crossed flexure plates (Fig. 1) as a part of the pull rods can eliminate
bending, but careful assembling of the specimen-support system, prior to
loading, is required to prevent twisting.
In the weather exposure tests (Fig. 1) unstressed specimens were
placed adjacent to stressed specimens so that losses in tensile properties
due to general corrosion and not to stress corrosion cracking could be
evaluated.
If resistances of different alloys are to be compared in the weather,
either many specimens must be exposed or tests on all specimens should
be initiated on the same day. Apparently weather conditions during the
first 24 hr of exposure have a significant effect on the lives of the specimens. Loose and Barbian [4] reported that specimens placed in weather
exposure racks in April had a much longer life than those initally exposed
during the summer months. Further, it is our opinion that if the relative
susceptibilities of different lots of material to stress corrosion cracking are
to be compared, these materials should be statistically distributed in the
weather exposure racks.
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It was noted in the NBS work with magnesium alloys that many speci-
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STRESS CORROSION TESTING
FIG. 2—Sheet specimen exposed, dead loaded in tension, in a boiling corrodent.
Air condenser and fan prevent evaporation of corrodent. Support system can be
interchanged so as to take a cylindrical specimen with threaded ends.
mens failed after a rain. Some specimens that were severely cracked after a
rain, nevertheless, did not fail until the next rain. The postulated mechanism [5] was that rain water plus corrosion products temporarily destroyed the protective film at certain locations on the specimen surface.
This permitted escape of dislocations with resulting slip and the exposure
of narrow film-free surfaces to the moisture. These areas were anodic to
the
very much
larger
surface
ofDec
the 16
specimen.
If the
resulting
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electrochemical-mechanical
action was more rapid than the film repair,
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LOGAN ON TESTING TECHNIQUES
131
the crack propagated to specimen failure. In some of the specimens, the
film repair process eventually became sufficiently rapid to stop crack
propagation. These specimens did not fail until the next rain.
Specimens Exposed in Laboratory Corrodents
If specimens are to be exposed to a specific environment rather than the
weather, they must be enclosed in suitable cells. Figure 2 shows a typical
cell used at the NBS. This was made from 40-mm Pyrex glass tubing with
the ends drawn down to fit available stoppers. For elevated temperature
work, the cell wall may be insulated with asbestos paper and a heating coil
of resistance wire wound onto the wet paper. This is, in turn, covered with
another thickness or two of asbestos paper. Heating tapes can of course be
used in place of the coils described above. The cells are closed at the bottom and top with slotted rubber stoppers. Evaporation of the solution can
FIG. 3—Four-point loaded specimens that may be exposed in a furnace.
normally be prevented by the use of an air condenser and a small electric
fan.4
If the specimens are to be subject to intermittent immersion, the solution
may be raised into the cells from a central reservoir by means of either a
mechanical pump or compressed air operated on a fixed cycle by means of
a clock system.
It is difficult to adjust the flow from a central reservoir to a number of
cells so that there will be no overflows. For that reason an overflow tube,
of slightly larger diameter than the input tube with its intake at the maximum desired solution level, should lead back through the lower stopper
from each cell to the reservoir.
Note again that if the corrodent produces any general corrosion at all
in the material under test, unstressed specimens should also be exposed
4
The use of small 110-v a-c electric fans such as are available for cooling elecCopyright
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E. H. Phelps
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STRESS CORROSION TESTING
under identical conditions, so that the damage resulting from general corrosion may be evaluated.
Sheet specimens have also been tested as bent beams. Figure 3 shows a
four-point loaded specimen of a salt-covered titanium alloy that had
failed after exposure in air at 750 F for 64 days [6]. The stress in the
outer fiber was computed from the deflection which was measured by the
dial gage device shown in Fig. 4. The loading device also shown in Fig. 4
is a variation of the jig used by many investigators. Over a short length
FIG. 4—Device for measuring deflection of bent-beam specimens in a 2-in. gage
length and jigs for stressing bent-beam specimens. Knife edges rather than points are
used in current models to avoid scratching of specimens. Upper jig designed so that
specimens may be inserted into large mouth flasks.
of arc, the specimen will approach circular curvature, and the strain can be
computed by simple geometrical considerations. We have done little work,
however, with bent-beam specimens except in a marine atmospheric
exposure of stainless steels and the high-temperature exposure of saltcoated titanium alloy sheet material.
Notched Specimens
By changing the grip arrangements, the lever systems (Fig. 1) can, of
course, be used for stressing specimens having circular cross sections.
In the work on stress corrosion cracking of low-carbon steels in nitrate
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solutions,
conventional round tension specimens stressed to their yield
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LOGAN ON TESTING TECHNIQUES
133
strengths failed in 300 to 1800 hr. Notched specimens, however, failed
in 5 to 7 hr [7]. Sixty-degree notches were cut in Vs-m.-diameter rods,
so that the specimen diameter at the root of the notch was 0.25 in. and the
radius of the root of the notch was 0.005 in. This gave a notch concentration factor of approximately 4. To determine the loads to be used with
these specimens, true stress-true strain curves5 were obtained on some of
them, and the remainder were then loaded at a true stress just equivalent
to that which produced a deviation from the linear portion of the true
stress-true strain curve, that is, to a stress that would give a small amount
FIG. 5—Upper chart shows sudden extension of low-carbon steel specimen indicating initiation of mechanical failure prevented from going to completion because
of the work hardening of material at the crack tip. Change in potential results from
exposure of bare material to the corrodent. Specimen was in corrodent at 100 C.
of plastic deformation at the root of the notch. The extension of the specimen in the corrodent during an experiment could be plotted on a strip
chart by using a differential transformer attached to the upper and lower
shoulders of the specimen. The electrochemical solution potential of the
specimens, in terms of a saturated calomel electrode, could be plotted on
a.second channel of the recorder by connecting the electrode to the corrodent through an agar-agar bridge. Figure 5 shows a sudden extension of a
low-carbon steel specimen that was accompanied by a change in potential
in the active direction [7]. The extension of the specimen was believed to
6
Cross-sectional
wererights
calculated
from
diameter-gage
readings
at
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STRESS CORROSION TESTING
FIG. 6—Longitudinal section through root of low-carbon steel specimen. Note
that a pair of cracks had developed (XlOO).
FIG. 7—Longitudinal section through a notch in a stainless steel specimen that
had been subjected to plastic deformation prior to exposure in the corrodent. Note
that some of the cracks (out of the highly deformed region) were approximately
parallel to the tensile axis (X50).
be a mechanical fracture prevented from going to completion by work
hardening and strain aging of the metal at the tip of the crack. This fracture exposed bare film-free metal to the corrodent, hence the change in
potential.
Cracks developed in pairs at the roots of notches in both low-carbon
and
stainless
steelInt'l
specimens.
A longitudinal
section
through
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LOGAN ON TESTING TECHNIQUES
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a low-carbon steel specimen shows that two cracks had developed, one on
either side of the apex of the notch (Fig. 6). Figure 7 shows a longitudinal
section through the root of a notch in a stainless steel specimen that had
been pulled in tension to maximum load, unloaded, and subsequently exposed to a boiling magnesium chloride (MgCl2) solution while stressed
at about one third the previously determined maximum load. Note that
after cracks had penetrated through the severely cold-worked area they
extended almost parallel to the nominal tensile axis [8]. Neither this behavior nor the fact that cracks develop in pairs at the roots of cylindrical
FIG. 8—Stainless steel foil that had been used as an anode in the cathodic protection of stainless steel specimens in a boiling MgClz solution. Note the extensive
stress corrosion cracking resulting from the applied anodic current and the residual
stresses in the foil.
notched specimens has been explained. A number of stress corrosion
cracks penetrated to about the same depth in an annealed stainless steel
specimen having a cylindrical reduced section and then split into pairs.
One can postulate from this behavior that, when the stress concentrations
resulting from the penetration of the individual cracks reached a certain
value, further cracking was in pairs as was the case with the notched specimens.
Plastic tape has been used effectively up to the boiling temperature of
MgCl2 (154 C) to stop off parts of the specimen that might modify the
electrochemical solution potential of that part of the specimen involved
in the cracking reaction. Specifically, all of the area exposed to the solution
exceptby the
reduced
of sheet
round
tension
is
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STRESS CORROSION TESTING
normally masked out. The connections to pull rods are also normally outside the solution. This avoids galvanic couples that might adversely affect
the experiment. Hence, these do not have to be stopped off.
The notched specimens are usually covered with tape up to the edges
of the notch. In one instance, stop-off lacquer was used to cover a whole
magnesium alloy specimen including the notched area. A circumferential
cut was made through the lacquer at the root of the notch, and the specimen was exposed under stress in the corrodent. Its life was several times
that of a specimen exposed subsequently with the whole notched area exposed. Exposing the whole notched area increased the cathodic area many
times and hence the ratio of cathodic area to anodic area.
A stainless steel foil is sometimes placed around the inside of the glass
cell for use as an anode, if the effects of applied currents (anodic or cathodic) on a stainless steel specimen are to be evaluated. Because of the
FIG. 9-—A hollow specimen used for investigating the resistance of materials to
stress corrosion cracking at elevated temperatures and elevated or reduced pressures.
Note that the specimen had failed completely, most probably after an earlier leak
(marked by arrow) had developed.
residual stresses in stainless steel foil, it would fail in a few hours by stress
corrosion cracking when used to cathodically protect an austenitic stainless steel specimen in a boiling MgCl2 solution (Fig. 8). This points up the
fact that anodic currents have also been used to accelerate stress corrosion
cracking in materials such as low-carbon steel [7, 9], stainless steels [10],
and aluminum alloys [77].
To avoid the difficulty of the foil anode failing in the magnesium chloride solution, a cell was constructed from a low-carbon ferritic steel tube.
This cell served very well as an anode to stop stress corrosion cracking by
cathodic protection in the austenitic stainless steel specimens. However,
when the protection was removed, there was a marked difference using
the different anodes. With the stainless steel foil anode, crack propagation
would start again in a few minutes. With the ferritic steel anode, crack
propagation did not start again for several hours. In the first case, a gelatinous magnesium hydroxide or hydrated oxide film formed on the specimen surface, but once protection was removed (current stopped) this was
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LOGAN ON TESTING TECHNIQUES
137
men protected by the ferritic anode could not be identified from its X-ray
diffraction pattern.
The Stainless Steel Specimen That Is an Autoclave
A hollow specimen for elevated temperature work that could in itself
be an autoclave was described at the 1962 annual meeting of ASTM [72].
The specimen is shown in Fig. 9, and a specimen, after it had been split
so that the interior could be examined, in Fig. 10.
Results of an investigation to determine the limits of chloride that could
produce cracking at elevated temperatures in a Type 304 stainless steel
were reported in Materials Research & Standards [13]. When an oxygen
pressure was introduced above the solution, water containing 5-ppm chlo-
FIG. 10—Hollow titanium alloy specimen sectioned for examination of the interior wall. Stress corrosion cracks are indicated by arrows.
ride produced cracking at 575 F. If no oxygen was added, a water solution
containing 20,000-ppm chloride at 575 F did not produce cracking.
The Hollow Ti-8-1-1 Specimen
This same type of specimen has been used in the study of the mechanism of salt stress corrosion cracking in titanium alloys. Some of the
results obtained with the Ti-8Al-lMo-lV alloy were reported at the
Seattle meeting of ASTM in November 1965 [6]. To coat the inner wall
of the specimen with salt, a 10 per cent sodium chloride (NaCl) solution
was poured into and then slowly pipetted out of the specimen maintained
at a temperature >100 C. This left crystals of NaCl distributed, not necessarily uniformly, on the inner wall. The specimen was plugged, connected
to a manifold through the small tube going through the plug, and placed
inCopyright
a creepbyfurnace.
It was stressed in tension through pull rods connected
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STRESS CORROSION TESTING
cavity was evacuated to about 10~3 torr. Wet or dry oxygen, or wet or dry
argon, was permitted to flow into the specimen cavity, and it was then
heated to 750 F and stressed to 90 per cent of its yield strength at that
temperature. If the specimen contained oxygen or water vapor, or if it
had been heavily oxidized prior to the experiment, it would fail under these
conditions of exposure. Figure 10 shows cracks extending into the interior
wall of a specimen.
FIG. 11—Fractured end of a specimen of high-strength 4340 steel showing area
of brittle fracture resulting from hydrogen embrittlement that had initiated the failure. Evidence indicates that the brittle fracture had originated in the interior of the
specimen wall (X12).
Specimens for Investigation of Hydrogen Embrittlement
Low-Carbon Steel Specimen
Hollow specimens have also been machined from steel rod or steel tubing for the purpose of determining whether or not hydrogen was involved
in the failure of the specimen exposed under tensile stress in a corroding
medium.
In one set of experiments, hollow-notched low-carbon steel specimens
were simultaneously evacuated (to approximately 1 X 10~3 torr), exposed
in a boiling aqueous 20 per cent NH4NO3 solution, and subjected to a
tensile stress sufficient to produce a small amount of plastic deformation at
the roots of the notches. When there was evidence that cracks were developing, the residual gases in the specimens were analyzed for hydrogen by
mass spectrographic methods. There was no evidence that hydrogen had
penetrated through the steel in these experiments [14].
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LOGAN ON TESTING TECHNIQUES
139
FIG. 12—Hydrogen bubbles emerging from a stress corrosion crack in a Type
304 stainless steel specimen exposed in a boiling 3Vi per cent NaCl-1 per cent
NHtNOs aqueous solution.
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STRESS CORROSION TESTING
High-Strength Steel Specimens
In experiments having the same objective, specimens were machined
from H-in.-outside diameter AISI 4130 and 4340 tubing. A reduced
section about 1 l/z in. long was machined on each length of tubing leaving
a wall thickness of approximately 0.035 in. The tubes were heat treated
to give tensile strengths of about 180,000 psi for the 4130 and 255,000
psi or more for the 4340 steel.
These specimens were simultaneously evacuated to about 10~3 torr,
exposed to a boiling 3 l/z per cent aqueous NaCl solution, and stressed to
approximately 75 per cent of their yield strengths. Once the specimen was
stressed, the pumping system was cut off and the pressure buildup in the
specimen noted. Periodically, the hydrogen concentration in the residual
gas in the specimen was determined using standard hydrogen analysis
techniques. The results have been reported in some detail elsewhere [75].
The fractured surface of a 4340 specimen is shown in Fig. 11. There is
considerable evidence that the brittle fracture marked out by the arrows
originated in the interior of the metal. Elongation measurements indicated
that the failures occurred after an incubation period of several hours.
Failure was usually complete within 30 sec after there were indications that it had been initiated. This behavior of the specimens and the
fact that appreciable amounts of hydrogen were always measured in the
residual gas prior to specimen failure indicate that failures in these steels
under these conditions were the result of hydrogen embrittlement.
Studies of Corrosion Products
If one is to understand the mechanism of stress corrosion cracking, a
knowledge of corrosion products is helpful. The fact that hydrogen diffused into the interior of the hollow high-strength steel specimens suggested
that the failures resulted from hydrogen embrittlement.
Figure 12 shows bubbles emerging from a developing stress corrosion
crack in an austenitic stainless steel specimen exposed in NaCl, NH4NC>2
solution. Subsequent analysis of gases given off during the corrosion process indicated that these bubbles were hydrogen [16]. Chromium and
nickel were found by chemical analyses in the liquid remaining in the interiors of the stainless steel autoclave specimens, which were removed before cracks penetrated entirely through the specimen walls, that is, before
failure occurred [13]. Microspectroscopic analyses of the solid corrosion
products associated with a crack in an austenitic stainless steel specimen
indicated that iron, chromium, and nickel were present in approximately
the same proportions as in the steel. This indicated that cracking was not
due to the selective attack of the corrodent on a particular component of
the steel.
A gas sampling tube was included in the vacuum system connected to
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LOGAN ON TESTING TECHNIQUES
141
graphic analyses of samples of the gas from the interior of the specimen
did not reveal the presence of chlorine. It had been reported as a corrosion
product by another investigator. At least one as yet unidentified solid
corrosion product was removed from the interior walls of these specimens.
Summary
1. Techniques developed during 25 years of investigating the susceptibility of various alloys to stress corrosion cracking are described. Certain interesting data, some of which have not been previously published,
are also reported.
2. If results obtained on exposure of different materials to the weather
are to be compared, specimens should be exposed at the same time.
3. Damage due to general corrosion should be evaluated in all stress
corrosion testing, using unstressed specimens exposed under identical conditions, and taken into account in assessing damage due to stress corrosion
cracking.
4. Notched specimens having cylindrical cross sections and stressed
so as to produce a small amount of plastic deformation at the root of the
notch may fail after exposure periods two orders of magnitude smaller
than those for unnotched cylindrical specimens. Cracks at the root of the
notch characteristically develop in pairs above and below the apex of the
notch.
5. Hollow specimens have been designed that can contain the corrodent
under either high or reduced pressures of oxygen, air, or an inert gas and
that can be subjected to tensile stresses at elevated temperatures. This type
of specimen has been used to investigate the resistance of stainless steels
and a titanium alloy to special corrosive conditions.
6. Specimens machined from heavy-wall, high-strength alloy steel tubing were used to determine that hydrogen was a factor in the delayed failure of this steel in a hot chloride solution.
7. In efforts to understand the mechanism of stress corrosion cracking,
gaseous and solid corrosion products have been identified by mass spectrographic, spectrographic, X-ray diffraction, and wet chemical analyses.
References
[1] Logan, H. L., Hessing, H., and Francis, H. E., Journal of Research, National
Bureau of Standards, Vol. 38, 1947, p. 465, RP1788.
[2] Logan, H. L., and Hessing, H., Journal of Research, National Bureau of Standards, Vol. 41, 1948, p. 69, RP1905.
[3] Logan, H. L., and Hessing, H., Journal of Research, National Bureau of Standards, Vol. 44, 1950, p. 233, RP2074.
[4] Loose, W. S., and Barbian, H. A., ASTM-AIME Symposium on Stress Corrosion Cracking of Metals, 1944, American Society for Testing and Materials,
Philadelphia, 1945, p. 283.
[5] Logan, H. L., Journal of Research, National Bureau of Standards, Vol. 61,
1958, p. 503,RP2919.
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STRESS CORROSION TESTING
[7] Logan, H. L., Journal of Research, National Bureau of Standards, Vol. 66C,
1962, p. 347.
[8] Logan, H. L., The Stress Corrosion of Metals, Wiley, New York, 1966.
[9] Pearson, E. E., and Parkins, R. N., Welding Research, Vol. 3, 1949, p. 95R.
[10] Hoar, T. P. and Hines, J. G-, Stress Corrosion Cracking and Embrittlement,
Wiley, New York, 1956, p. 122.
[11] Booth, F. F. and Godard, H. P., First International Congress on Metallic Corrosion, Butterworths, London, 1962, p. 703.
[12] Logan, H. L., "A Specimen for Use in Investigating the Stress-Corrosion
Cracking of Metals at Elevated Temperatures," Materials Research & Standards, Vol. 2, No. 2, Feb. 1962, p. 98.
[13] Logan, H. L., McBee, M. J., and Romanoff, M., "Stress-Corrosion Cracking
of Type 304 Stainless Steel at 455 to 615 F," Materials Research & Standards,
Vol. 3, No. 8, Aug., 1963, p. 635.
[14] Logan, H. L. and Yolken, H. T., Second International Congress on Metallic
Corrosion, NACE, Houston, 1966, p. 109.
[15] Logan, H. L. and Wehrung, J. M., Corrosion, Vol. 22,1966, p. 265.
[16] Logan, H. L. and Sherman, R. J., Jr., Welding Journal, Supplement, Welding
Research, Vol. 35, No. 8,1956, p. 389-S.
DISCUSSION
F. A. Smith* (written discussion)—I would appreciate your comments
on laboratory techniques for stress corrosion cracking of materials conducted at temperatures up to 1600 F in a sodium (liquid metal) atmosphere. The need for such testing is related to any liquid metal fast reactor
program. One can project that liquid metal fast reactor technology cannot
progress from laboratory devices to industrial power plants until the stress
corrosion properties of a variety of alloys in sodium is well understood.
Therefore, there exists a need for the development of "standard techniques" for stress corrosion testing in the new environment, sodium.
E. N. Pugh2 (written discussion)—It seems to me that the term stress
corrosion cracking is a generic term and that several different mechanisms
of failure are operative, depending on the system and on the environmental
conditions. I have the impression that you regard the electrochemicalmechanical mechanism, which you describe as a generalized model, applicable to all systems. Do you consider the model to be generalized?
R. W. Staehle3 (written discussion)—The reason for the potential rise
associated with deformation is undoubtedly rupture of the passive film
1
2
Md.
3
Physicist, Argonne National Laboratory, Argonne, 111.
Staff scientist, Research Institute for Advanced Studies, Martin Co., Baltimore,
Assistant
professor,
Department
of Metallurgical
Engineering,
Corrosion CenCopyright
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DISCUSSION ON TESTING TECHNIQUES
143
by mobile dislocations and subsequent exposure of active metal. This
nonprotected metal then reacts rapidly as shown in Fig. 13 (left). However, this process of film rupture and subsequent dissolution also occurs in
alloys which do not crack (that is, pure nickel in chloride) as shown in
Fig. 13 (right). Since this process of film rupture and dissolution occurs
in alloys which crack, what do you consider to be the critical process in
determining inherent susceptibility to cracking?
H. L. Logan (author)—Dr. Smith's question is one that I am sure has
received much thought by many people. One could use a hollow specimen
such as I have described to contain the sodium at temperature. The problem is how to contain the hot sodium when the specimen fails. Two ideas
FIG. 13—Thin foils are shown which have been exposed to boiling MgCh and
subsequently examined in the electron microscope. Light areas are regions of dissolution: (left) Fe-15Ni-20Cr-1.5Si (15 min); (right) pure nickel.
come to mind: (1) Could you use as a heating bath an alloy that will combine with liquid sodium in a manner that can be controlled? (2) Would it
be possible to surround the hollow specimen with a metal bellows (by
welding it to the shoulders), determine its effect on the stress applied to
the specimen, and depend upon it to contain the sodium after specimen
failure?
Both Dr. Pugh's and Dr. Staehle's questions pertain to theory which I
discussed in the presentation but is not included in the paper. My ideas
are best presented elsewhere.4
As is well known, stress corrosion cracking can be prevented or, once
* Logan, by
H.ASTM
L., The
Stress
Corrosion
of Metals,
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New York,
1966, pp. 33Copyright
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STRESS CORROSION TESTING
started, stopped in many systems at least by the application of cathodic
protection. Conversely, stress corrosion cracking can be accelerated by
the application of anodic currents. This would indicate that the process is
in part, at least, electrochemical with small anodes at the crack sites, the
remaining area forming large cathodes. It is a reasonable supposition that
the function of the stresses is to rupture a protective film (visible or invisible) on the metal surface, exposing film-free areas which are more
chemically active than the filmed surfaces, and become the anodes. Electrochemical attack of these anodes can trigger successive mechanical
fractures until failure occurs. I believe that this is the most satisfactory
explanation of the phenomenon of aqueous stress corrosion cracking as
it occurs in many systems and, therefore, can be considered a generalized
model. Yes, I think that we can present a valid generalized model for stress
corrosion cracking.
There are two possible reasons for the fact that the example posed by
Dr. Staehle does not crack. There may not be a sufficient difference in the
electrochemical solution potential of the filmed and film-free surfaces of
nickel in chlorides to propagate a crack by an electrochemical process.
We found that Type 321 stainless steel did not crack in a boiling lithium
chloride (LiCl) solution unless tank oxygen was bubbled through it. The
tank oxygen had the effect of making the electrochemical solution potential of the metal about 90 mv more noble and, hence, increased the potential difference between the film covered and film-free alloy by 90 mv.
The second possibility is that plastic deformation of nickel in the
chloride solution may result in much larger ratio of film-free-to-film
covered areas than in materials susceptible to stress corrosion cracking.
We postulate that if stress corrosion cracks are to propagate by an electrochemical mechanism, we must have small anodes associated with very
large cathodes so that the current density at the anodes will be large.
Specifically, then, if an alloy is to fail by stress corrosion cracking, there
must be sufficient difference in the electrochemical solution potential
between the film-free and film covered surfaces to promote rapid electrochemical attack. Second, plastic deformation must be sufficiently nonhomogeneous so that there will be relatively small regions of high stress
concentration.
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W. D. Sylwestrowicz1
A Proposed Mechanism for the Stress
Corrosion Fracture of a
Copper-Beryllium Alloy
REFERENCE: W. D. Sylwestrowicz, "A Proposed Mechanism for, the
Stress Corrosion Fracture of a Copper-Beryllium Alloy," Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 145.
ABSTRACT: Polycrystalline specimens of a copper-beryllium alloy (1.8
per cent Be) were bent and exposed to an ammonia atmosphere. Two lots
of polycrystalline material of similar chemical composition but of different
grain size and distribution of segregates were investigated. Specimens were
subjected to different values of plastic deformation and humidity conditions. Plastic deformation appears to be an essential factor in stress
corrosion fracture. The importance of humidity conditions was emphasized by the results obtained. Both modes of fracture, transcrystalline
and intercrystalline, were observed. It was found that the distribution
of segregates determines which of these two modes of fracture will occur.
A mechanism of stress corrosion fracture in a copper-beryllium alloy is
proposed.
KEY WORDS: stress corrosion, corrosion, copper-beryllium
ammonia atmosphere, plastic deformation, fracture, humidity
alloys,
In essence the process of stress corrosion fracture is a propagation of
a crack in a nonbrittle material with some features of brittle behavior.
This process occurs when a stressed material is exposed to a specific corrosive environment. The value of stress under which the specimen fractures
is usually much smaller than that for fracture in the usual static test.
In the search for a model to describe stress corrosion fracture, a number of mechanisms have been suggested. They are based on a large volume
of experimental data and can be divided into two general groups: electrochemical theories and mechanical theories. A comprehensive review of
this field was given recently by Parkins [I].2 An attempt is made here to
formulate a mechanism based on experiments with a copper-beryllium
1
Member of technical staff, Bell Telephone Laboratories, Inc., Murray Hill, N. J.
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STRESS CORROSION TESTING
alloy, which takes into account both electrochemical and mechanical effects.
The proposed model requires that the segregates form thin layers around
the pileups of dislocations inside the grains or around the dislocations
piled against the grain boundaries. The thickness of these layers is a critical factor. By an electrochemical process, the material is removed from
these layers, forming narrow cracks with large stress concentrations at
their tips. As a result of the diffusion of the solute and removal of the
material, these layers become locally brittle, and with a large stress con-
FIG. 1—Etched surface of a specimen from the first lot of material. Specimen
was neither strained nor exposed to the corrosive atmosphere. Very small precipitates, probably of /3 phase, are visible (XlOOO).
centration at the tip of the crack, cleavage occurs. But as the bulk material is not brittle, the propagation of the crack stops after proceeding a
short distance with the formation of a number of fresh dislocations in
front of the crack. Then the process repeats. The result is a discontinuous
propagation of the crack.
Material and Experiments
The experiments described in this work were made on pressure cast
copper-beryllium alloy of the following composition: 97.7 per cent
copper, 1.8 per cent beryllium, and 0.3 per cent cobalt. Spectrochemical
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analysis showed:bycobalt, magnesium, and iron from 0.01 to 0.3 per cent;
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FIG. 2—Etched surface of a specimen from the second lot of material. Specimen
was neither strained nor exposed to the corrosive atmosphere. Large precipitates,
probably of p phase, are present (XlOOO).
FIG. 3—Etched surface of a specimen from the second lot of material. Specimen
was neither strained nor exposed to the corrosive atmosphere. Large precipitates,
probably of /3 phase, are present at the grain boundary, also broad segregates are
shown along
grainInt'l
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(XlOOO). Wed Dec 16 15:53:43 EST 2015
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STRESS CORROSION TESTING
aluminum, magnesium, and silicon <0.03 per cent; silver < 0.005 per
cent; copper, nickel, and tin < 0.001 per cent. The alloy was solution
annealed at 800 C, quenched in water and aged at 355 C for 3 hr. From
this material bars were machined 2 in. in length and % 6 by % 6 in. in
cross section. The matrix of these polycrystalline bars consisted of the
f ace-centered-cubic a phase in which was dispersed a small amount of the
body-centered-cubic /? and y phases and cobalt beryllide particles. Two
lots of material were used. Their chemical compositions were nearly
identical, but the grain sizes and distributions of the precipitates and impurities differed. The grain diameter of the first lot was approximately
0.07 mm and of the second 0.3 mm. In the first lot, segregates were dispersed evenly through the grain with a comparatively small amount of
segregation (probably y phase) at the grain boundaries (Fig. 1). In the
FIG. 4—An oblique view of the bending jig.
second lot, the /? phase formed large precipitates inside the grains (Fig.
2) and y-phase segregates at the grain boundaries (Fig. 3). Specimens
from the first lot were used in experiments I and II; specimens from the
second lot in experiments IV and V.
The specimens after cutting from the bulk material were mechanically
polished and then electropolished in 80 per cent phosphoric acid. Specimens were subjected to three-point bending in a specially designed jig (Fig.
4). By sliding the wedges one in relation to the other, the specimens were
bent to the desired deflection. The deflection was measured with a microscope. With a known deflection, as long as a specimen is deformed elastically, maximum tensile stress can be easily computed from the measured
deflection by the equation
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TABLE 1—Experimental data.
Specimen No.
1-3
II-l
II-2
H-3
H-4
IV4
IV-2
IV-3
V-l
V-2
V-3
Lot
1
1
1
1
.1
2
2
2
2
2
2
(
Pf
"max,6 kg/cm"
0.01
0.61
0.09
0.07
0.01
0.74
0.38
0.10
0.76
0.38
0.10
4.850
11.050
9.400
9.100
6.570
11.200
10.650
9.550
11.200
10.650
9.550
Water
Pressure,
mm Hg
Ammonia
Pressure,
mm Hg
20
20
20
20
25
25
25
4.5
4.5
4.5
720
720
720
720
550
550
550
550
550
550
Crack 0.2 mm Long
Crack 1.0 mm Long
dl/dt,d
Crack 3.0 mm Long
dl/dt,d
c
Tex, hr
dl/dt,d
mm/hr
3
0.08
10
0.10
20
0.15
36
40
0.01
100
425
1.1
1.7
8.0
0.015
0.015
2.0
0.6
0.24
240
480
2.0
3.5
15.0
0.025
0.07
1«
1<
1
>60
>60
>60
c ,,_
T
Tex nr
'
a
epi = plastic strain.
omax = maximum tensile stress.
Tex = time of exposure.
d
dl/dt = rate of crack propagation.
6
e
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mm/hr
Tex ,c hr
mm/hr
3.5
1.5
0.7
150
STRESS CORROSION TESTING
specimen, c0 is the half-height, and / is the half-length of the specimen.
For three-point bending, the value of the maximum tensile stress is in the
outside layer in the middle of the specimen. The stress decreases toward
and is zero at the neutral plane of the specimen. When the maximum stress
is larger than the yield point, the calculation is more complex. For comparatively small values of the plastic strains, only a very thin outside layer
of material near the central load point will be plastically deformed. The
bulk of the specimen is deformed elastically. It is assumed, therefore,
that for the computation of the total strain, the curvature of the deflected
specimen is the same as for the specimen deformed elastically only. With
FIG. 5—A relation between the length of the crack and the time exposure is
shown for Specimen (1-3).
this assumption and the measured deflections, total strain can be evaluated,
and with the use of stress-strain curve for the material, the value of plastic
strain in the outside layer can be determined. The values of strain for the
specimens used in these experiments are given in Table 1.
In the first series of tests, specimens were placed in a desiccator above
a crystallizing dish containing ammonium hydroxide. After about 30 min
of exposure, a black film developed on the surface of the specimen. To
observe the propagation of cracks, specimens were taken periodically
from the desiccator, cleaned in 10 per cent hydrochloric acid for a few
seconds to remove corrosion film, and examined under a microscope. The
structure of this black film was studied by an X-ray diffraction method.
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Recorded
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indicated
a structure
related
toEST
cupric
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SYLWESTROWICZ ON FRACTURE OF A COPPER-BERYLLIUM ALLOY
151
This was confirmed by examination with an electron-diffraction technique.
Even after a very short exposure (20 min), observed lines corresponded
to the cupric hydroxide structure with no indications of the presence of
copper oxides. To check if the process of stress corrosion is sensitive to the
humidity of the environment Specimen (I-3)3 was placed in a desiccator
above a crystallizing dish containing ammonium hydroxide. After 60 hr
of exposure, ammonium chloride was added to the dish to lower the vapor
pressure of water in the desiccator with the result that a pronounced decrease in the rate of crack propagation occurred (Fig. 5).
FIG. 6—The surface of Specimen (11-3) after 3 hr of exposure. Trans- and intercrystalline cracks are visible (XlOOO).
In the second series of tests, Specimens (II-1, 2, 3, 4) were placed in an
evacuated container (pressure below 0.1-mm mercury) at 25 C to which
water vapor at 20-mm mercury pressure and ammonia at 720-mm mercury pressure were added. During the tests as some of the water reacted
with the ammonia and copper, the water vapor pressure probably dropped
below the initial value. The fourth series of tests (Specimens (IV-1, 2, 3))
was conducted at saturated vapor pressure of water (25-mm mercury)
at 25 C with additional water supply in the system to keep this pressure
3
Specimen designation: first figure refers to the experiment and second to the
specimen,
for example,
Specimen
(1-3)
refers
to 15:53:43
the thirdEST
specimen
from number one
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STRESS CORROSION TESTING
constant through the test. The ammonia pressure was 550-mm mercury.
At the beginning of the tests in the fifth series (Specimens (V-l, 2, 3)),
the water vapor pressure was 4.5-mm mercury at a test temperature
25 C and ammonia pressure of 550-mm mercury. After 60 hr of exposure even in the most prestrained Specimen (V-l), cracks were not
observed, and the test was then discontinued.
Observations and Discussion
Initiation and Propagation of a Crack
In these series of tests, a number of observations were made related
to the initiation and propagation of the crack. As it was difficult at a
FIG. 7—An electron microscopic replica taken from the surface of Specimen
(1-3) after 3 hr of exposure (X20.000).
very early stage to distinguish between an etched slip plane or an etched
grain boundary and a genuine crack, the time when a crack obtained the
length of 0.2 mm was considered as the time of initiation of the crack.
At this size a crack was already well defined. The main observations
resulting from these tests are summarized in Table 1.
In the process of stress corrosion fracture, two stages will be distinguished: first the initiation of a crack, then its propagation. Prior to the
appearance of initial cracks, an appreciable amount of etching takes
place along the grain boundaries and along the intersection with the
free surface of crystallographic planes inside of the grains. The initial
cracks were either inter- or transcrystalline (Fig. 6). In the first lot of
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153
FIG. 8—A crack in the Specimen (11-4) after 480 hr of exposure.
FIG. 9—The surface of Specimen (II-3) after 3 hr of exposure (XlOOO).
in the second lot, in an intercrystalline one. Transcrystalline cracks
usually develop in the middle of the grain and then spread toward the
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STRESS CORROSION TESTING
ment with the finding of Tromans and Nutting [2]. They found that in
face-centered-cubic alloys transcrystalline cracks are in {111} planes.
It is difficult to measure the radius of curvature of a crack at its tip,
but a rough estimate can be made judging from the photograph of a
surface replica in the electron microscope (Fig. 7). This radius seems to
be smaller than 10~5 cm. A number of initial cracks appear along the
central part of the stressed edge of the specimen. At this stage, the length
of the crack does not depend on the position of the crack. This might
be the result of local stress concentrations, crystallographic orientation
of the grains, and grain boundary conditions.
FIG. 10—The surface of Specimen (11-1) after 3 hr of exposure (XlOOO).
The following observations were made on the propagation of transcrystalline cracks. Initially, cracks appear in the area stressed in tension.
They are comparatively shallow (approximately 10 to 20/x), and the general direction of propagation is hi the plane perpendicular to the bending
axis of the specimen (Fig. 8). A transcrystalline crack conveniently
oriented to the bending axis of the specimen extends in its crystallographic
plane. When the crack reaches a grain boundary, it passes into the
adjacent grain again following crystallographic planes in that grain
(Fig. 9). If the slip planes in the adjacent grain are unsuitably oriented
to the propagation direction of the crack, while the grain boundary is
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SYLWESTROWICZ ON FRACTURE OF A COPPER-BERYLLIUM ALLOY
155
conveniently oriented slip planes. While the overall direction of spreading
of a crack is well defined, locally the crack may deviate quite appreciably
from this direction (Figs. 8 and 10). Once a crack spreads for some
distance on one of the side faces of the specimen, its propagation slows
down or stops until the crack spreads across the front face to the opposite side and penetrates through the whole depth of the specimen. Then
it resumes its normal rate of propagation. This in itself results in changes
of velocity of propagation of a crack measured at the surface of the specimen. If two cracks are formed close together, the stress field between
cracks will be relieved and the crack in the less favorable condition will
stop growing (Fig. 11).
FIG. 11—The surface of Specimen (H-l) after 3 hr of exposure (left) and 7 hr
of exposure (right) (X200). Reduced on the photograph by 2.
Another factor contributing to changes in the rate of propagation of a
crack is related to the proposed mechanism of spreading of a crack in a
nonbrittle material. This mechanism will be discussed in what follows.
It is not surprising, therefore, that a marked variation in the rate of
propagation of the cracks is observed experimentally. Discontinuous
motion of a crack was previously reported by Gilbert and Hadden [3]
and Edeleanu and Forty [4]. Gilbert and Hadden attribute the changes
in the velocity propagation of a crack to the sucessive fracturing and
restoring of the surface film; Edeleanu and Forty to the presence of
deformation bands in the specimens.
Threshold Stress
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It was confirmed in this study, as has been reported by many investiDownloaded/printed by
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156
STRESS CORROSION TESTING
cracks were restricted to the central part of the specimens in the regions
of high stress. In the areas close to the cylindrical supports where the value
of stress is small (zero at the supports), no cracks were observed. Also,
the existence of a threshold stress for the propagation of a crack can
be demonstrated in the following way. The value of the maximum
tensile stress is given by Eq 1. If a stress concentration at the tip of a
crack is taken into consideration the maximum stress is
where L is the length of the crack; P is the radius at the tip of crack; 2c0
and 2c are initial and instantaneous heights of the specimen; and other
FIG. 12—The time of exposure necessary to produce a crack 0.2 mm long as a
function of plastic prestrain or stress in the second series of experiments.
symbols have the same meanings as in Eq 1. With increasing length of
the crack, c decreases; therefore, <rmax decreases also to the zero value
when the crack approaches opposite face (c = 0). Assuming that a minimum value of stress is required to propagate a crack in the conditions
described here, complete fracture should not occur. A crack should stop
before reaching the opposite face unless it develops so large a velocity
that it propagates by its momentum through the area of small stresses.
In agreement with this prediction, complete fracture occurred in only
one test, and this was in the specimen of the highest prestrain. In all
other tests, cracks stopped a fraction of a millimeter before complete
penetration of the specimen.
Effect
of Plastic Prestrain
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SYLWESTROWICZ ON FRACTURE OF A COPPER-BERYLLIUM ALLOY
157
that the threshold stress may correspond to the value of stress necessary
for the initiation of plastic deformation in the grains most favorably
oriented to the applied stress field, as was suggested by Bakish and
Robertson [5].
The rate of crack propagation is affected also by prestrain in the
specimen. The values of time of exposure and of velocity of crack
propagation are shown in Table 1. An exponential dependence was found
for the velocity of crack propagation and plastic prestrain, as can be seen
in Fig. 13 where values of velocity against plastic deformation plotted
on semilogarithmic scale presents a straight line relation. No simple
relation could be established between the time of exposure to fracture
or velocity and value of prestress. This indicates that the main effect of
prestraining a specimen is in the amount of plastic deformation produced
FIG. 13—The velocity of crack propagation as a function of plastic prestrain
for a crack 1 mm long (curve a) and 3 mm long (curve b) in the fourth series
of experiments.
and not in the value of stress. This does not imply that stress itself is not
a factor in the process of stress corrosion. The existence of a threshold
stress clearly indicates that imposed stress is a factor. The explanation
of this seemingly inconsistent behavior may be in the fact that, after
reaching the yield stress, increasing prestrain produces large amounts of
plastic deformation with comparatively small increase of stress. In summary, there are three factors affecting the stress field around a crack:
the shape of the crack, which is related to the amount of plastic deformation as will be discussed in what follows; the orientation of the
crack in relation to the applied stress; and the value of applied stress.
Among these, the shape of the crack is the predominant factor.
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Effect ofby Dislocations
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158
STRESS CORROSION TESTING
formations larger than 1.5 per cent (in 70:30 a-brass alloy) and intergranular cracking with smaller deformations. This dependence of the
type of cracking on the amount of prestrain has not been confirmed for
the copper-beryllium alloy investigated in this study. Specimens plastically
prestrained by only 0.01 per cent developed transcrystalline cracks, and
specimens of 0.7 per cent plastic prestrain showed some intercrystalline
cracking. Therefore, it is concluded that the amount of prestrain is not
related to the mode of formation and propagation of a crack.
It is assumed here that the initiation of a crack is a result of removal
FIG. 14—Etched surface of a specimen from the first lot of material. Specimen
was neither strained nor exposed to the corrosive atmosphere. Very small precipitates with a tendence to align in the crystallographic directions are visible (XlOOO).
of segregated material electrochemically and not by the mechanism proposed by Stroh [7], in which he postulates that dislocations in a pileup
mechanically develop a crack. A dislocation acts normally as a sink for
segregates. Segregates diffusing toward piled-up dislocations in their
slip plane can form a thin plate around them. Through the removal of
segregated material by an electrochemical process, a narrow crack is
formed. The diffusion coefficients, temperature of test, amount of segregates, and disposition of dislocations will be decisive factors in determining the layer thickness of the segregates. With easy cross slip, tangles of dislocations will be formed. This results in broad cracks or rounded pits
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and
comparatively small stress concentrations around them. Likewise, a
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SYLWESTROWICZ ON FRACTURE OF A COPPER-BERYLLIUM ALLOY
159
prolonged heat treatment or large content of solute will result in broad
segregates and produce cracks with small stress concentrations. This can
explain why the susceptibility of some materials to stress corrosion fracture changes [8] when conditions of heat treatment are changed. With
increasing prestrain, the number of dislocations increases, pileups become longer, and this leads to a larger number of long segregation plates.
With a larger number of plates the chance of formation of a crack
increases and the propagation of the crack is faster.
In the copper-beryllium alloy studied, formation of transgranular
cracks was restricted to the interior of the grain away from grain
boundaries where large pileups of dislocation would be expected. This
can be explained as follows: grain boundaries act as sinks for the solute,
therefore, in their vicinity there might be insufficient solute to form a
continuous layer in the slip planes and to form a crack. In some cases when
a large amount of solute is present in the material, a partial migration into
the grain boundaries might create conditions favorable for formation of
thin cracks in the vicinity of boundaries.
This interpretation of the effect of the distribution of segregates in
the material on the form of fracture in the stress corrosion processes is
supported by the results of the experiments described. Specimens from the
first lot, with a very small amount of segregation at the grain boundaries,
have segregations dispersed uniformly throughout the grains with a
tendence to align in the crystallographic directions (Fig. 14). In these
specimens, fracture is mostly transcrystalline. In the specimens from the
second lot with an appreciable amount of segregates at the grain boundaries (Fig. 3), fracture is intercrystalline.
Recently, Tromans and Nutting concluded [2] that change of mode
of fracture is related to electrochemical potential, originating from a segregation of solute elements on dislocations and not to the value of stacking
fault energy as they considered previously [6]. They believe that segregation is an essential condition for stress corrosion fracture. In the present
proposal, it is also accepted that segregation of solute atoms is a fundamental condition for the process but with an important difference. Tromans and Nutting assumed that propagation of a crack occurs through
"dissolution of dislocations," followed by the plastic shearing of material
between dislocations. In the present work, it is postulated that material
becomes locally embrittled as a result of diffusion of solute atoms toward
pileups of dislocations and removal of material from this region. With
applied stress and small enough radius of the crack, a large stress develops at the tip of the crack and cleavage occurs.
It is implied in the presented model that at room temperature diffusion
of solute atoms toward dislocations is possible at least through the distance of a few interatomic spacings. Recent investigations of the effect
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STRESS CORROSION TESTING
by lowering temperature to — 20 C the process was stopped while general
corrosion still occurred. Lowering temperature by 50 C does not embrittle
tested material but strongly affects diffusion to which is attributed the
slowing of the process of stress corrosion cracking.
Effect
of Humidity in a Corrosive Atmosphere
The effectiveness of an electrochemical cell, among other factors,
depends on the presence of a conducting liquid phase at the surface of
the specimen [9]. This can be obtained from the water vapor in the
corrosive atmosphere providing that a hygroscopic film is formed on the
surface of the specimen. The condition of high humidity in the corrosive
environment and formation of the hygroscopic film on the surface of
the specimen are, therefore, essential for the initiation of a crack in the
stress corrosion process.
The effect of humidity is clearly shown in the experiment with a specimen exposed to the corrosive atmosphere in the test chamber with
ammonium hydroxide. By addition of ammonium chloride, which reduces
the vapor pressure of water in the test chamber, a very pronounced decrease of the rate of propagation occurs (Fig. 5). Even more striking are
the results of the IVth and Vth series of experiments (Table 1) with
vapor pressure of water 25 and 4.5-mm mercury, respectively. After 1
hr of exposure of Specimen (IV-1) a crack of 0.8 mm length was observed
in the specimen. After 2 hr of exposure, the crack reached 3 mm length. In
Specimen (V-l) of the same 0.7 per cent prestrain, after 60 hr of exposure, there was no evidence of formation of cracks.
Effect
of Corrosive Film
The effect of humidity on the process of cracking is related to the
presence of a hygroscopic film on the surface of the specimen. This is
shown in the experiment with Specimens (II-2, 3). These two specimens
had similar prestrain and were tested simultaneously. The surface film
of Specimen (II-2) was left untouched while that of Specimen (II-3) was
removed periodically (nine times). This periodic removal of the corrosion
film resulted in the increase of exposure time required to obtain a crack
of 4 mm length from 50 to 300 hr.
In the mechanism of stress corrosion fracture which depends on film
rupture [10-15], it is postulated that a surface film forms a protective
layer over the specimen. The applied stress cracks this film exposing fresh
material susceptible to the corrosive attack. This mechanism is not supported by the observations in the experiments described here. These
show that removal of the surface film delays the process of stress corrosion
(Specimens (II-2, 3)).
In the studied copper-beryllium specimens, the surface film consists
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of
copper hydroxide of orthorhombic structure. It is unlikely therefore
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SYLWESTROWICZ ON FRACTURE OF A COPPER-BERYLLIUM ALLOY
161
that cracks which are in crystallographic orientation in face-centeredcubic matrix coincide with cracks in the film which has an orthorhombic
structure.
To investigate the effect of the corrosion film, Tromans and Nutting
[2] studied a carbon replica of the stress corroded surface of a specimen
of Cu-30Zn alloy. They found a large number of cracks on the surface
film of the grains. This suggests that transcrystalline cracking should be
present in the matrix material if these cracks result from the cracks in
the surface film. In fact, only intercrystalline cracks were observed.
FIG. 15—Debris of the surface of a "dome" which collapsed under the heat of
the illumination light in a microscope (X200).
The cupric hydroxide, observed in this study, is not formed by direct
reaction between copper and water. The reactions leading to formation
of cupric hydroxide are not known fully, but the presence of ammonia
is essential. A specimen exposed for three months to an atmosphere
saturated with water vapor but without ammonia has not developed a
cupric hydroxide film.
A complex reaction probably occurs between copper, ammonia, and
water. The product of this reaction may be a cupro-ammonium complex
formed in the liquid phase from which cupric hydroxide precipitates.
In some tests after longer exposures, a thin crystalline film formed over
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162
STRESS CORROSION TESTING
on the surface of the specimens. These domes appeared to be filled with a
liquid. Due to the heat of the light during microscopic examination the
thin surface of the dome collapsed, and liquid evaporated leaving debris
of the surface film (Fig. 15) and exposing deep pits which had formed
under the domes. The extrusions observed by some authors [2,16] on
the surfaces of stress corroded specimens may in fact be this debris.
It has also been suggested that wedging action of corrosion products
is responsible for the propagation of a crack [76,77]. The fracture delay,
caused by the removal of corrosion layers, could be considered as a
confirmation of this mechanism. However, the observation that below a
level of applied stress cracks do not propagate indicates that a wedging
mechanism cannot be considered as a principal cause of crack spreading
but only as a minor factor in this process. The results of Forty and
Humble [72] showing that in a brass there is no increase of volume of
the corroded material put in doubt the existence of a wedging action. The
conclusion can be made that the importance of the surface film in the
process of the stress corrosion fracture, at least in copper-beryllium alloy,
lies in the hygroscopic properties of the film formed and not in the initiation of a crack in the matrix by a crack first developed in the surface film.
Conclusions
To summarize, the following model of a stress corrosion mechanism
is proposed for copper-beryllium alloy. This model is probably applicable
to a wider range of materials susceptible to stress corrosion fracture. A
material under tensile stress is exposed to a corrosive atmosphere. As a
result of the applied stress, a number of fresh dislocations are generated.
These dislocations located in the slip planes act as a sink for solute atoms,
which preferentially segregate around dislocations. These segregates will
form thin layers in slip planes or at grain boundaries.
At the same time as a result of corrosion, a film is formed on the surface of the specimen. This film is hygroscopic and absorbs water from
the atmosphere to form a conducting layer. In the presence of the liquid,
an electrochemical cell is formed between a layer of segregates and the
matrix of the specimen. The effect of this electrochemical process will
be the removal of the material from the zones around the slip planes or
grain boundaries. With these two processes working simultaneously,
material becomes brittle in these areas, and, in the most favorable location, a crack can be initiated on the surface of the specimen. Transcrystalline cracks lie in crystallographic planes (slip planes).
At the tip of a crack which is roughly perpendicular to the direction
of applied stress, a large stress concentration is present. In this embrittled
area, conditions for cleavage propagation are produced and a crack starts
to spread in the slip plane. But, as the bulk material is not brittle, the
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velocity of crack will decrease as the crack proceeds into the nonbrittle
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SYLWESTROWICZ ON FRACTURE OF A COPPER-BERYLLIUM ALLOY
163
volume, new dislocations in front of the propagating crack will be formed,
and the crack will stop. Then the process repeats. In this way a crack
will spread through the specimen in a jerky motion.
The essential parts of this model are:
1. Segregation of solute atoms to the dislocations in slip planes. The
segregated material differs from the matrix material of the specimen in
electrochemical properties. Also, as the result of diffusion of solute
locally, the material becomes more brittle.
2. Formation of an electrochemical cell. This leads to removal of
material from the areas already rich in solute atoms and to still more
embrittlement of material. For the operation of the electrochemical process, the presence of a liquid phase is necessary. This liquid phase is
produced as a result of the action of the corrosive atmosphere.
3. Formation of a crack with a small enough radius at its tip to produce stress concentration large enough to initiate propagation of a crack
in locally embrittled material.
A cknowledgments
The author would like especially to thank H. C. Theuerer for the
suggestions and helpful discussions on the chemical problems connected
with this work; W. C. Ellis, T. D. Schlabach, and U. B. Thomas for
critical reading of the manuscript; Mrs. A. M. Hunt for performing the
electro-diffraction determination of the corrosion film structure; and Miss
S. E. Koonce for taking electron microscopic photographs.
References
[/] Parkins, R. N., "Stress-Corrosion Cracking," Metallurigical Reviews, Vol. 9,
1964, p. 201.
[2] Tromans, D. and Nutting, J., "Stress-Corrosion Cracking of Face-CenteredCubic Alloys," Corrosion, Vol. 21, 1965, p. 143.
[3] Gilbert, P. T. and Hadden, S. E., "A Theory of the Mechanism of Stress-Corrosion in Aluminium-7% Magnesium Alloy," Journal of the Institute of Metals,
Vol. 77, 1950, p. 237.
[4] Edeleanu, C. and Forty, A. J., "Some Observations on the Stress-Corrosion
Cracking of a-Brass and Similar Alloys," Philosophical Magazine, Vol. 5, 1960,
p. 1029.
[5] Bakish, R. and Robertson, W. D., "Structure-Dependent Chemical Reaction
and Nucleation of Fracture in Cu3 Au Single Crystals," Acta Metallurgica,
Vol. 4, 1956, p. 342.
[6] Tromans, D. and Nutting, J., "Electron Microscope Studies of Stress Corrosion
Cracking," Fracture of Solids, Drucker, D. C. and Oilman, J. J., eds., Interscience, New York, 1962, p. 637.
[7] Stroh, A. N., "Crack Nucleation in Body-Centered-Cubic Metals," Fracture,
Averbach, B. L. et al, eds., Wiley, New York, 1959, p. 117.
[8] Uhlig, H. H. and Sava, J. P., "Origin of Delay Time in Stress Corrosion Cracking of Austenitic Stainless Steel," Proceedings, Second International Congress
on Metallic Corrosion, 1963, National Association of Corrosion Engineers,
1966.
[9] Uhlig,
H., Corrosion
and Corrosion
Control,
New York,
1963, p. 146.
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[10] Mears, R. B., Brown, R. H., and Dix, E. H., Jr., "A Generalized Theory of
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164
[11]
[12]
[13]
[14]
[15]
[16]
[77]
STRESS CORROSION TESTING
Stress Corrosion of Alloys," ASTM-AIME, Symposium on Stress-Corrosion
Cracking of Metals, 1944, American Society for Testing and Materials, Philadelphia, 1945, p. 323.
Logan, H. L., "Film-Rupture Mechanism of Stress Corrosion," Journal of Research of the National Bureau of Standards, Vol. 48, 1952, p. 99.
Forty, A. J. and Humble, P., "The Influence of Surface Tarnish on the StressCorrosion of a-Brass," Philosophical Magazine, Vol. 8,1963, p. 247.
McEvily, A. J., Jr., and Bond, A. P., "On the Initiation and Growth of Stress
Corrosion Cracks in Tarnished Brass," Journal of Electrochemical Society,
Vol. 112, 1965, p. 131.
Forty, A. J., "Surface Films and Stress Corrosion Cracking," Proceedings, Conference on the Environment-Sensitive Mechanical Behavior of Materials, 1965,
to be published.
McEvily, A. J. Jr., and Bond, A. P., "On Film Rupture and Stress Corrosion
Cracking," Proceedings, Conference on the Environment-Sensitive Mechanical
Behavior of Materials, 1965, to be published.
Nielsen, N. A., 'The Role of Corrosion Products in Crack Propagation in Austenitic Stainless Steel Electron Microscopic Studies," Physical Metallurgy of
Stress-Corrosion Fracture, Rhodin, T. N., ed., Interscience, New York, 1959,
p. 121.
Pickering, H. W., Beck, F. H., and Fontana, M. G., "Wedging Action of Solid
Corrosion Product During Stress Corrosion of Austenitic Stainless Steels,"
Corrosion, Vol. 18, 1962, p. 230.
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H. Rosenthal1 and H. R. Pritchard1
A Quantitative Stress Corrosion Test for
AI-Zn-Mg Alloy Plate
REFERENCE: H. Rosenthal and H. R. Pritchard, "A Quantitative Stress
Corrosion Test for Al-Zn-Mg Alloy Plate," Stress Corrosion Testing,
ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 165.
ABSTRACT: The stress corrosion susceptibility of Al-4.25Zn-2.9Mg plate
was determined in the short-transverse direction. C-ring type specimens
were exposed to alternate immersion in a NaCl solution. It was shown that
laboratory temperature and humidity must be controlled to obtain reproducible results. Time-to-failure results for a given lot of material have
a normal distribution when failure times are converted to logarithms.
Tests were also made in a marine environment and in an industrial environment. These outdoor results for various lots follow the same ranking
for susceptibility as determined by the alternate-immersion tests.
KEY WORDS: aluminum alloys, stress corrosion, corrosion, marine exposure, industrial exposure
It is well known that the 7000 series alloys (Al-Zn-Mg-Cu) are susceptible to stress corrosion when heat treated to maximum hardness. The
susceptibility is mainly confined to the short-transverse direction in rolled,
forged, or extruded sections. A number of specifications require testing
such material by subjecting specimens to intermittent immersion in a sodium chloride (NaCl) solution.
Recently, there has been interest in using a copper-free weldable grade
of an Al-Zn-Mg type of alloy plate for army vehicles. The data in this report include some of the information on stress corrosion characteristics
of this alloy developed by Frankford Arsenal in cooperation with the
aluminum industry.
The objectives of the work were: (1) to define the laboratory test conditions such that the test results would be quantitatively reproducible and
(2) to determine if there was a correlation between the laboratory tests
and outdoor exposures.
1
Research advisor, and physical science technician, respectively, Frankford
Arsenal, United States Army, Philadelphia, Pa. Mrs. Pritchard is a personal memCopyright
by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
ber ASTM.
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STRESS CORROSION TESTING
Experimental Procedure
Work described in this report covers alternate-immersion tests conducted under ambient conditions of laboratory temperature and relative
humidity, as well as outdoor tests in marine and industrial environments.
It also covers a later series of tests in which laboratory temperature and
relative humidity were carefully controlled.
Material
Material consisted of aluminum plate conforming to Military Specification MIL-A-46063. Plate thickness ranged from SA to 3% in. Plates
were in the solution-treated, quenched, and artificially aged temper.
The chemical composition limits in weight per cent were as follows:
zinc, 3.5 to 5.0; magnesium, 2.0 to 3.8; manganese, 0.10 to 0.70; chromium, 0.06 to 0.25; iron, 0.40 max; silicon, 0.30 max; zirconium, 0.20
max; copper, 0.10 max; titanium, 0.10 max; others, each 0.05 max, 0.15
total; aluminum, remainder.
Minimum mechanical properties were as shown in Table 1.
TABLE 1—Mechanical properties.
Plate Thickness, in.
Up to 1.500 inclusive
Over 1.500
Tensile Strength,
min, psi
Yield Strength,
0.2% offset,
min, psi
Elongation,
min,
per cent
60 000
57 000
51 000
48 000
9
8
FIG. 1—Orientation of C-ring in relation to plate.
Test Specimens
C-ring type specimens were machined from the plate so that the centerline of the specimen, which is the area of maximum applied stress, represented the centerline of the plate. A sketch of the specimen orientation in
relation to the rolling direction of the plate is shown in Fig. 1.
Two sizes of C-rings were used, % and 11A in. diameter with wall
thicknesses of 0.060 and 0.100 in., respectively. Three-quarter-inch rings
were made from plate !J/4 in. and less in thickness, and lV4-in. rings from
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
167
Stress Corrosion C-Ring Specimen
OD,
in.
t,
in.
Hole Dia (H) Hole Dia (H)
for Bushing, w/o Bushing, Al Screw
Size
in.
in.
0.750 0.060
0.257
0.203
10-32
1.250 0.100
0.257
0.203
10-32
Outside Finish - 63
FIG. 2—C-ring dimensions.
FIG. 3—Two methods of preventing galvanic corrosion between screws and
C-rings, (upper) Micarata bushing (ring partially sectioned) and flower) neoprene
dip.
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STRESS CORROSION TESTING
Figure 2 shows the geometry of the specimens, with two alternate sizes
of holes for the stressing screws. In early tests, a Micarta bushing was
used to insulate the screw from the specimen to prevent galvanic corrosion.
Later, no bushing was used and galvanic corrosion was prevented by
dipping the bolt and nut area of the stressed specimen in liquid neoprene.
Figure 3 shows two rings: one partly cut away to show the bushing and
the other dipped in liquid neoprene.
FIG. 4—Z correction factor used in formula by which C-ring stress is calculated.
Stressing of Specimens
Specimens were stressed by tightening a screw and nut to reduce the
ring diameter. The following formula2 was used to determine the desired
stress in terms of the properties and dimensions of the specimens.
where:
A = change of outside diameter required for desired stress,
/ = desired stress, psi,
t = wall thickness, in.,
D = mean diameter (OD — i), in.,
2
Sprowls,
D. Int'l
O.,(all
inter-laboratory
memorandum,
Alcoa
Research Laboratories,
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
169
E = modulus of elasticity, and
Z = a constant (function of ring D/t, see Fig. 4).
Specimen Preparation
Specimens were degreased after machining. When no bushings were
used, specimens were dipped in liquid neoprene to cover the bolt, nut, and
surrounding specimen area and allowed to dry for 1 l/z hr. All specimens
were wiped with acetone-soaked cotton after stressing and immediately
put in test.
FIG. 5—Schematic of alternate-immersion test equipment.
Equipment
A schematic drawing of the equipment is shown in Fig. 5 and a photograph in Fig. 6.
Alternate immersion of the specimens is accomplished by movement of
the salt solution rather than raising and lowering the specimens themselves.
At the beginning of the immersion cycle, the solution is contained in the
bottom tank. The cycle starts with the timer actuating the solenoid valve
which causes air to enter the tank. The bleeder valve in the tank permits
air to escape, but the air pressure rises, because the rate of leakage is much
less than the rate at which the air is entering the tank. The rising pressure
in the tank causes the liquid to flow into the tray above the tank. The second signal from the timer closes the solenoid valve and stops the flow of
air into the tank. The pressure in the tank now reverts to atmospheric
pressure because of the continued leakage from the bleeder valve. As the
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pressure in the by
tank becomes lower, the salt solution flows back into the
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
171
tank from the tray. The timer is so adjusted that the cycle for the specimens
consists of 10 min immersion in the solution and 50 min drying time. The
trays are tilted slightly to facilitate drainage and also to provide different
depths of solution to cover test specimens of various thicknesses for the
same period of time. Specimens in the trays were covered to a depth of
approximately l/z in.
Test Medium
The test solution was 3.5 per cent (by weight) reagent grade NaCl and
distilled water prepared weekly. The salinity of the solution was checked
five days a week with a hydrometer, and distilled water was added to compensate for evaporation.
No attempt was made to control pH, since it was 7.0 when the solution
was made and 7.3 to 7.4 when the solution was discarded after a week of
use.
Outdoor Test Conditions
Specimens prepared in the same manner as for alternate immersion
were also exposed to marine and industrial outdoor environments.
The marine site was located in southern New Jersey, approximately 500
ft from mean high tide. Specimens were supported on racks at an angle of
45 deg facing the ocean.
The inland site was located on the roof of a Frankford Arsenal building
(Philadelphia, Pa.). The atmosphere in the vicinity of Frankford Arsenal
can be considered an industrial type atmosphere, since it is located in an
area surrounded by heavy industry. Specimens here were also supported
on racks at a 45 deg angle facing south.
Examination for Cracks
The presence of cracks was determined by visual inspection of the
specimens. If doubt existed, visual examination was supplemented by
examination with a binocular microscope at a magnification of X7. If
doubt still remained, the area was noted, and subsequent examinations
were directed to this area of the specimen to see if crack growth had occurred.
The laboratory tests in alternate immersion were kept under observation
except over the weekends. The observation period on weekdays was between the working hours of 7:30 a.m. and 4:30 p.m. The number of
examinations carried out daily varied from two to four.
Examination of specimens at the marine site was made weekly and at
the industrial site every weekday.
Tests Under Ambient Laboratory Conditions
A series of alternate-immersion tests was made in which ambient lab-
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STRESS CORROSION TESTING
TABLE 2—Results of tests in alternate immersion at different seasons of the year.
Lot
A
B
Stre ss
. psi
21
21
39
39
000
000
000
000
Date Started
9 March
12 June
25 February
27 April
Avg
Temperature,
degF
Failure Time,
days
82
36, 42, 56
4, 9, 10
6, 80, 80
7, 14, 14
87
83
85
FIG. 7—Time-to-failure at 35,000 psi of 14 lots of plate in three environments
arranged in order of increasing resistance to alternate immersion. Tests started
during summer months.
heating months, the air temperature ranged from 77 to 84 F and the
relative humidity from 30 to 50 per cent. In the summer, the air temperature varied from 77 to 101 F and the relative humidity from 30 to 90 per
cent. The temperature of the test solution was from 4 to 7 deg less than the
air temperature.
Table 2 lists results of two lots of material tested in alternate immersion
at different times of the year. These data show the effect on failure time of
the temperatures prevailing during periods of winter heating and summer
ambient conditions. Tests were run in triplicate, and individual failure
times for specimens are shown in the table. It can be seen that failure time
decreases with increasing temperature.
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
173
Comparative Summer Tests
As was previously indicated, the variation of temperature and humidity
was in a much narrower range in the winter than in the summer. It is therefore interesting to compare, in general, results of tests carried out in the
summer months with those conducted during the winter.
Figure 7 shows data obtained from tests started in the spring and summer, when laboratory ambient conditions showed their maximum variation. Data on stress versus time-to-failure were obtained on 14 lots of
FIG. 8—Time-to-failure at 35,000 psi of 11 lots of plate in three environments
arranged in order of increasing resistance to alternate immersion. Tests started
during winter months.
plate with at least 3 specimens per stress level. A curve of stress versus
median time-to-failure was plotted for each lot. From these curves, a
median time-to-failure at 35,000 psi was determined for each lot. These
are arranged in Fig. 7 in order of increasing resistance to stress corrosion
in the alternate-immersion test. Corresponding data developed simultaneously for the same lots in the marine and industrial environments are
also shown.
The ranking determined by the laboratory tests was considerably different from the ranking for the outdoor exposure tests (Fig. 7).
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STRESS CORROSION TESTING
Comparative Winter Tests
Figure 8 shows data for eleven lots which were started during the winter
heating months. These data are also arranged in order of increasing resistance to stress corrosion in the alternate-immersion test. The plots show
good correlation between the alternate-immersion and the outdoor results
except near the terminal part of the graph, where the median failure time
in alternate immersion is in excess of twelve days.
Data from Fig. 8 have been replotted in Fig. 9 in the form of per cent
survival versus failure time. In this graph, the shape of the curve of the
marine exposure is similar to that of the alternate immersion; in contrast,
FIG. 9—Per cent survival versus time-to-failure at 35,000 psi in three environments. Specimens represent the 11 lots of plate on which tests were started during
winter months.
the industrial curve is nearly a straight line. Other investigators have noted
that an industrial atmosphere is often extremely aggressive.
Tests on Micarta Bushings
In addition to the effects of temperature and relative humidity on
reproducibility of results, it was found that the Micarta bushings used as
insulators affected results in the alternate-immersion test, particularly at
high stress levels. It was suspected that these bushings were deforming
under stress and thus permitting a relaxation of stress in the specimens.
An experiment was conducted to determine the extent of this effect on
alternate-immersion test specimens and on two groups of control specimens. One control group was exposed to alternate immersion without
bushings and the other group, with bushings, was exposed to air only. The
air temperature was 80 ± 2 F and the solution temperature 76 ± 2 F.
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Specimens were % in. diameter and were stressed at several levels. Diam-
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
175
eters were measured over a period of time to determine whether relaxation had occurred.
It was found that the maximum change occurred after five days in the
specimens with bushings exposed to alternate immersion and that the
amount of relaxation increased with increasing applied stress. No change
in diameter occurred in controls exposed to alternate immersion without
bushings nor in controls exposed only to air at any stress level.
Results for the alternate-immersion test are shown in Table 3.
Apparently the Micarta absorbed moisture and became softened, permitting the specimens to expand and relax the applied stress. For this
reason, the use of bushings was discontinued. Galvanic corrosion was
thereafter avoided by dipping specimens in liquid neoprene to cover the
screw, nut, and specimen contact areas.
TABLE 3—Change in diameter of %-in. rings exposed to alternate immersion.
Applied Stress,
psi
20 000
30 000
40 000
No. of
Specimens
Increase in Diameter
in 5 Days, in.
Stress
Relaxation, psi
2
2
4
0.001, 0.002
0.002, 0.003
0.002, 0.004, 0.004, 0.004
1600, 3200
3200, 4800
3200, 6400
Tests Under Controlled Conditions
On the basis of the results cited, it was decided that constant temperature and relative humidity must be maintained in the alternate-immersion
test to obtain significant and reproducible results.
Therefore, the equipment previously described was installed in a room
in which the air temperature was controlled to 80 ± 2 F and the relative
humidity to 45 ± 6 per cent by the use of an air-conditioner, dehumidifier,
and electric heater. With this combination of temperature and relative
humidity, the solution had an average temperature of 76 ± 2 F. The temperature of the specimens during the 50-min drying cycle increased above
the solution temperature by 1 to 2 F, as determined by thermocouples
placed on specimens.
At the relative humidity maintained, the specimens apparently dried
completely between immersions.
To determine whether quantitative and reproducible results could be
obtained with the improved facilities, an experiment was designed in which
two series of replicate specimens were tested in alternate immersion. All
specimens were stressed to 35,000 psi. Details of these tests follow.
Material
Material was Al-Zn-Mg plate from two commercial lots in the solutiontreated,
quenched,
artificially
aged
One lotEST
was
a nominal
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176
STRESS CORROSION TESTING
thickness of 1V4 in. and the other, 2 in. They were identified as Lots 1 and
2, respectively. These lots were selected because of their known low resistance to stress corrosion. Their composition is shown in Table 4.
The mechanical properties in the long-transverse direction are shown
in Table 5.
Specimens
Three-quarter-inch-diameter specimens were machined from Lot 1 and
1 V^-in.-diameter specimens from Lot 2. A total of 72 specimens was made
TABLE 4—Chemical composition, weight per cent.
Lot 1
Element
Lot 2
0.1 to 0.2
0.1 to 0.3
<0.05
0.1 to 0.3
2.88
0.1 to 0.3
3.75
<0.05
none
remainder
Si
Fe
Cu
Mn
Mg
Cr
Zn
Ti
Zr
Al
0.1 to 0.2
0.1 to 0.3
<0.05
0.2 to 0.5
2.97
0.1 to 0.3
3.57
<0.05
none
remainder
TABLE 5—Mechanical properties.
T ,
Tensile Strength,
psi
1
2
64 800
64200
Yield Strength,
0.2% offset, psi
57 000
55 900
Elongation,
%
13.0
13.0
from each lot and tested in eight groups of nine specimens each. Specimens
in each group were tested under identical conditions, with the start of the
test for each group constituting a sequence with three- to seven-day intervals between starting times.
Specimens were stressed, prepared, and tested according to methods
previously described, except that no bushings were used as insulators.
Instead, specimens were dipped in liquid neoprene as shown in the lower
part of Fig. 3.
Determination of Cracking Time
Since examinations for cracks were not made during nights or weekends,
there was necessarily a large uncertainty as to when cracks actually occurred in those instances when they were found after the weekend or night
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TABLE 6—Failure time data on C-rings stressed to 35,000 psi and tested in alternate immersion under controlled conditions, days.
Group No.
Elapsed Time
Averaged Time
LOT No. 1—1J4-IN. PLATE, %-IN. C-RiNos
1
2.7, 3.7, 3.7, 3.7, 3.7, 3.7, 3.7,
3.7, 3.7, 4.7, 4.7, 5.0, 5.0, 5.7,
2.7, 3.0, 3.7, 4.0, 5.0, 5.0, 5.7,
4.0,4.7,4.7,4.7,4.7,4.7,4.7,
2.7, 3.0, 3.0, 3.0, 3.0, 3.0, 3.7,
2.7, 3.0, 3.0, 3.7, 3.7, 3.7, 4.0,
2.7, 2.7, 2.7, 2.7, 3.7, 3.7, 3.7,
2.7, 2.7, 2.7, 3.0, 4.0, 5.6, 5.6,
2
3
4
5
6
7
8
4.7,
5.7,
5.7,
7.0,
3.7,
6.7,
4.6,
6.7,
5.7
6.7
5.7
9.7
3.7
6.7
5.6
...«
2.7, 3.4, 3.4, 3.4, 3.4, 3.4, 3.4, 4.4,
3.4, 3.4, 4.4, 4.4, 5.0, 5.0, 5.4, 5.4,
2.7, 3.0, 3.4, 4.0, 5.0, 5.0, 5.4, 5.4,
4.0,4.4,4.4,4.4,4.4,4.4,4.4, 7.0,
2.7, 3.0, 3.0, 3.0, 3.0, 3.0, 3.4, 3.4,
2.7, 3.0, 3.0, 3.4, 3.4, 3.4, 4.0, 6.4,
2.7, 2.7, 2.7, 2.7, 3.4, 3.4, 3.4, 4.3,
2.7,2.7,2.7,3.0,4.0,5.3,5.3,
5.4
6.4
5.4
8.4
3.4
6.4
5.3
6.4,
LOT No. 2—2-iN. PLATE, I^-IN. C-RiNGS
2.7, 3.0, 3.7, 4.0, 4.7, 5.7, 6.7,
0.8, 2.8, 4.0, 4.0, 4.0, 6.7, 6.7,
3.7, 4.0, 4.0, 4.7, 4.7, 4.7, 4.7,
2.7, 3.0, 3.7, 4.7, 4.7, 5.7, 9.7,
4.0, 4.7, 4.7, 4.9, 4.9, 5.7, 9.7,
2.7, 3.7, 4.0, 4.7, 5.7, 6.7, 6.7,
2.7, 2.7, 3.7, 4.6, 4.6, 5.6, 7.0,
2.7, 3.0, 3.7, 5.0, 5.6, 6.7, 9.6,
1
2
3
4
5
6
7
8
0
6
12.0,
7.0,
11.0,
9.7,
9.7,
7.1,
9.6,
11.0,
24.0
14.0
14.0
11.0
9.7
11.0
9.6
13.0
2.7, 3.0, 3.4, 4.0, 4.4, 5.4, 6.4,
. . . " 2.4, 4.0, 4.0, 4.0, 5.4, 5.4,
3.4, 4.0, 4.0, 4.4, 4.4, 4.4, 4.4,
2.7, 3.0, 3.4, 4.4, 4.4, 5.4, 8.4,
4.0, 4.4, 4.4, 4.9, 4.9, 5.4, 8.4,
2.7, 3.4, 4.0, 4.4, 5.4, 6.4, 6.4,
2.7, 2.7, 3.4, 4.3, 4.3, 5.3, 7.0,
2.7, 3.0, 3.4, 5.0, 5.3, 6.4, 8.3,
Specimen lost.
Omitted from analysis.
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12.0,
7.0,
11.0,
8.4,
8.4,
7.1,
8.3,
11.0,
24.0
14.0
14.0
11.0
8.4
11.0
8.3
13.0
178
STRESS CORROSION TESTING
period. For the overnight cases, the true cracking time was assumed to be
midway between the morning of discovery and the last previous examination of the preceding day. If the test was started on a Friday afternoon, as
happened in many instances, the midpoint averaging method for cracks
found the following Monday morning would be biased, because there is
FIG. 10—Log™ of days-to-failure of 3A-in. C-rings, Lot 1, versus per cent
failure plotted on normal probability paper. All specimens stressed at 35,000 psi.
FIG. 11—Logio of days-to-failure of llA-in. C-rings, Lot 2, versus per cent
failure plotted on normal probability paper. All specimens stressed at 35,000 psi.
less likelihood of cracking during the early stage of the test. In these cases,
the Monday morning cracking time was used. However, these values are
identified hi graphs containing them.
Table 6 shows the actual elapsed times when cracks were found and also
the averaged times for the two lots of plate.
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
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Tests of Normality
Figures 10 and 11 contain the data for Lots 1 and 2 on normal probability paper using the logarithms of failure times. Booth et al3 have shown
that failure times are normally distributed when the times are converted to
logarithms. The technique of plotting the points is that of Lewis.4
A straight line was fitted to the points visually, since the fit is quite
simple. Because of this straight line relationship, the normality of the population is established.
It is possible to determine from Figs. 10 and 11 various statistical quantities. Since the distribution is normal, the mean and the median can be
considered to be identical. Other statistical data derived from these figures
are discussed in the following section.
TABLE 7— Statist ical data for Lots 1 and 2.
Lot
No.
1
2
a
Median Failure Time
Expected Median Min-to-Max Range0
Log 10 ,
days
Days
s Log 10 ,
days
» = 7,
days
n = 9,
days
0.590
0.720
3.9
5.2
0.130
0.220
3.0 to 5.1
3.3 to 8.2
3.1 to 4.9
3.6 to 7.8
» = is,
days
3.3 to 4.6
4.0 to 6.9
95% confidence level.
Determining Number of Specimens To Be Tested
In developing a specification for stress corrosion susceptibility, the
number of specimens to be tested must be kept as small as possible because of the costs of machining and handling. However, this desire must
be balanced against the fact that the median failure time is more accurately
determined with a larger group of specimens. The relevant relationship is:
expected range of median failure times5 = median ±
ts
—
where:
s = estimated standard deviation,
n = number of specimens, and
t = a quantity dependent on n and the confidence level desired for
the range of failure times. For 95 per cent confidence level, t =
2.36 (for n = 7), t = 2.30 (for n = 9), and t = 2.13 (for n =
15).
'Booth, F. F., Tucker, G. E., and Godard, H. P., "Statistical Distribution of
Stress Corrosion Testing," Corrosion, Vol. 19, 1963, pp. 390T-395T.
* Lewis, C. F., "Statistics—A Useful Tool for the Examination of Corrosion
Data," Corrosion, Vol. 9, 1953, pp. 38-43.
5
Dieter, G. E., Jr., Mechanical Metallurgy, McGraw-Hill, New York, 1961.
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180
STRESS CORROSION TESTING
In Table 7, the estimated standard deviation and median failure times
for Lots 1 and 2 are shown as determined graphically from Figs. 10 and
11.
Also shown in the table are the calculated expected median failure times
for 7, 9, and 15-specimen groups for a 95 per cent confidence level. The
choice of a 9-specimen group appears reasonable since there is a worthwhile improvement in the calculated spread over a 7-specimen group. The
added precision obtained with a 15-specimen group does not appear to
FIG. 12—An analysis of data relating to Lots 1 and 2 and a number of other
lots tested at 35,000 psi. The number of specimens tested for each lot is shown
between the data points. Plotted on the ordinate scale is the range of median failure
times calculated on the basis of nine specimens at 95 per cent confidence level;
corresponding points on the abscissa are the median failure times.
have good economic justification. With a group of nine specimens, the
minimum median value for Lot 1 is approximately 20 per cent less than
the experimental median; the corresponding value for Lot 2 is 30 per cent.
The plots of experimental median versus 9-specimen minimum and
maximum median are shown for Lots 1 and 2 in Fig. 12. Also plotted are
a number of points corresponding to other lots tested under similar conditions. These lots represent a lesser number of specimens, but the calculations are based on a hypothetical 9-specimen group.
The meaning of this graph can be further illustrated as follows: Assume
an unknown lot of material is to be tested and that a group of nine specimens willbybeASTM
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ROSENTHAL AND PRITCHARD ON TEST FOR AL-ZN-MG ALLOY PLATE
181
mum of 5-day median failure time with a confidence level of 95 per cent.
The 5-day level is taken from the ordinate, and the corresponding minimum value on the abscissa must experimentally show a median of ll/z days
to have less than 1 chance in 20 that the median of a group of 9 specimens
falls below 5 days.
When the consumer, on the other hand, accepts material having a median failure time of 5 days, there is some risk that he will accept material
actually having a median time of 4 days.
Conclusions
When the laboratory temperature was controlled to 80 ± 2 F and the
relative humidity to 45 ± 6 per cent (the conditions of these tests), results were reproducible.
Using the controlled laboratory conditions for two lots of Al-Zn-Mg
plate, the times-to-failure follow a normal distribution if failure times are
converted to logarithms.
Outdoor exposures in a marine atmosphere and in an industrial atmosphere gave similar results (on the basis of ranking) to the laboratory alternate-immersion tests carried out in the winter months.
A cknowledgment
The authors express appreciation to SP/7 B. Collins (U.S. Army) for
preparation of the line drawings in this report; to the various staff members of the research laboratories of Alcoa, Kaiser, Reynolds, and Dow
aluminum companies; and to General Motors personnel operating the
U.S. Army Tank Automotive Plant (Cleveland).
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H. B. Romans1
Stress Corrosion Test Environments
and Test Durations*
REFERENCE: H. B. Romans, "Stress Corrosion Test Environments and
Test Durations," Stress Corrosion Testing, ASTM STP 425, Am. Soc.
Testing Mats., 1967, p. 182.
ABSTRACT: The most frequent environments used for stress corrosion
testing of aluminum and titanium are presented. The general effects of
the various environmental factors which may affect the tests are described.
These include temperature, humidity, pH, etc. The data on a definite test
period are presented in a quite limited form due to the .scarcity of data in
published papers on this subject. The results of a questionnaire which was
submitted to all the members of Committee G-l concerning their testing
methods for stress corrosion testing are presented.
KEY WORDS: corrosion, stress corrosion, environmental testing, aluminum alloys, titanium alloys, temperature, humidity, pH, salt solutions,
oxygen
The types of environment in which stress corrosion cracking is best
produced are those in which highly localized corrosion occurs with the
absence of general surface corrosion. This makes the choice of the environment fairly specific for each of the metal families. Since no universal
test environment can be adopted, it is necessary to discuss each of the
metals separately.
The test period beyond which a metal can be judged safe from stress
corrosion cracking must depend on the end use being studied as well as
the metal. At present, the test periods may range from one day to three
years or until the test metal fails. In most cases, both of these considerations are left to the judgment of the persons conducting the test.
Subcommittee VI of ASTM Committee G-l on Corrosion of Metals
was formed for the purpose of setting up standard procedures for conducting stress corrosion tests.
1
Scientist, Department of Applied Chemistry and Mathematics, Reynolds Metals
Co., Richmond, Va. Chariman, Task Group 2, Subcommittee VI, ASTM Committee G 1. Personal member ASTM.
* Report of Task Group 2 of Subcommittee VI of ASTM Committe G-l on
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
183
Task Group 2 of Subcommittee VI was formed for the purpose of compiling:
1. Data on test environments which are in general use for stress corrosion cracking tests.
2. The length of test required to give reasonable ensurance that a metal
will perform satisfactorily in a given environment.
The data presented in this report were obtained from three major
sources. These were: (1) a literature survey by the chairman, (2) information supplied by the members of Task Group 2, and (3) a questionnaire
submitted to all the members of Committee G-l.
FIG. 1—Effect
of stress on the time-to-stress corrosion failure.
The report is divided into three general parts: I, General Philosophy
of Testing; II, Stress Corrosion of Nonferrous Alloys; and III, Stress
Corrosion Questionnaire on Ferrous Alloys. The results of the stress
corrosion questionnaire are entered as part of the particular metal concerned.
At the February 1966 Committee Week meeting in Washington, it was
decided by the members of Subcommittee VI to present for publication
only those metals which have been well covered. These include aluminum
and titanium. The other metals will be added as supplements to the original report as they are completed.
General Comments
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STRESS CORROSION TESTING
applying the environment to produce stress corrosion. It is also considered
that the publication of large amounts of data, test results, and test correlations would serve no useful purpose since it would be a repeat of published
work. Therefore, the direction of this report has been aimed at discussing
the environments used and the general methods of application. Titanium
is presented in a more detailed form which does include test methods and
periods. The work in this field is relatively new and the published work
scarce; this makes a compilation useful.
There has been no effort to describe specimen geometry or methods of
stress as this was covered very well by Task Group 1. Interpretation of
FIG. 2—Influence of the season of the year on rate of stress corrosion of brass.
data and test results will be covered by Task Group 3 and is not necessary
in this report.
Comments on Test Periods
Very little is mentioned in the literature as to a definite test duration.
Most workers in the stress corrosion field use a test period which is based
on the type test, test solution, material, convenience or space, and end use
for the data.
A number of practical considerations must enter into the selection of a
test period other than those above. For instance, the stress level with respect to the yield strength of the material must be considered. As can be
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small
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
185
differences in stress at low stress levels. At very high stress levels, the
stress per unit area is not a very large factor [7].2
If outdoor exposures are made, the time of year the test is started must
be considered in selecting the test period. Figure 2 shows the variation in
time-to-failure which may be obtained by exposing at different months of
the year.
The corrosivity of the environment is another important factor in
choosing a test period. Extremely corrosive environments may produce
stress cracks at a slower rate than a mild environment.
Careful consideration and judgment with regard to the general application of the metal may dictate the test period to be used. A consideration
which may set the period of test is quality control. For instance, one quality control test requires that specimens pass a 5-day median test period.
Another requires a 30-day test.
PART I. GENERAL PHILOSOPHY OF TESTING
Laboratory Testing for Stress Corrosion Susceptibility
Stress corrosion tests conducted in the laboratory are generally under
artificial environments conducted in such a way as to accelerate the corrosion and duplicate the type of failure experienced in service environments.
Laboratory tests are advantageous since they can be performed under
carefully controlled conditions. These conditions may include controlled
atmospheres such as temperature, humidity, and air pollution. These are
obtained by intermittent immersion, partial immersion, total immersion,
wick and drip feed, and spraying. Only if the conditions are controlled to
a high degree can accurate data comparisons be made.
The test conditions must be specific to the alloy system under study
since the stress corrosion of each metal is associated with a definite environment. Quite often an environment which develops stress corrosion hi
one metal, inhibits the cracking of another due to severe general attack.
Care should be taken to ensure that laboratory tests are not so severe
that failures are erroneously assumed to be stress corrosion cracking when
it is actually mechanical failure due to the reduction in cross-sectional
area. This danger is minimized by careful selection of the corrodent,
specimen size, and method of loading.
Accelerated stress corrosion tests are most useful if the results correlate
with service experience or natural environments. If the tests do not correlate with either of the above, they may still be useful for screening purposes.
Variations in a given environment can cause a marked change in the
2
The italic numbers in brackets refer to the list of references appended to this
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STRESS CORROSION TESTING
mode and type of cracking. For instance, in certain magnesium alloys,
a change in pH of the corrodent (NaCl-K2CrO4) will cause the failure to
change from transgranular, in high pH solutions, to intergranular, in low
pH solutions [2],
'Various Environmental Factors Which Affect Stress Corrosion Test Results
Physical Factors
Temperature—The temperature of the test solution is probably the
most important single factor in the role of stress corrosion cracking for
some metals and alloys such as the aluminum-zinc-magnesium alloy. The
rate of cracking increases with increase in temperature.3 A possible explanation is that the increase in temperature increases the conductivity
of the solution. Also, diffusion at elevated temperatures is faster which
allows increased activity in the corrosion cells.
Humidity—The effect of humidity on the results of stress corrosion
tests applies especially where the method of alternate immersion is used.
There is general agreement that the specimens never become completely
dry due to water absorption in the oxide layer. However, there is probably
a critical amount of moisture which must be removed to get reproducible
results.
During periods of high humidity, the specimens can stay dripping wet
for a long period and often do not dry between immersions. Some of this
deficiency may be overcome if the air is circulated to dry the specimens.
However, the test configuration often does not allow forced air drying.
The literature indicates that the effect of humidity is maximum for
aluminum at about 85 per cent and gives a marked decrease in the time-tofailure3 [3].
The stress corrosion cracking of all metals is affected to some extent by
the relative humidity. The effect of this variable can and should be eliminated by controlling the condition of the test area.
Atmospheric Environments
For the most part, stress corrosion testing in the atmosphere is used to
correlate laboratory tests and to determine how the metal will behave in
a natural environment. Unfortunately, long periods of exposure are
needed to produce failures in all but the most susceptible alloys, or when
the superimposed load is very near the yield strength of the material.
Moisture, temperature, and various atmospheric impurities, that is,
seacoast, industrial, rural, etc., influence differently the failures of different metals.
8
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Aluminum specimens may fail faster in an industrial atmosphere depending on the alloys3 [4].
Magnesium alloys subjected to a marine and a rural atmosphere failed
in about the same time period. Stainless steels tested at Kure Beach, N.C.,
and Middletown, Ohio, showed failures occurring within a few weeks at
Kure Beach, while the same alloys did not fail in two years at Middletown.4
Note that the rate of stress corrosion cracking under atmospheric
conditions increases at high humidity and high temperatures. This means
that exposures should be made at the same season of the year if accurate
data comparisons are to be made. The time-to-failure can vary as much
as a factor of eight by varying the month of exposure [5].
Chemical Factors
Neutral Salt Solutions—The stress corrosion cracking of metals in
neutral solutions of various salts, particularly sodium chloride, has been
widely investigated.
Many of the metal alloys are susceptible to stress corrosion cracking
in distilled water at elevated and ambient temperatures. However, cracking occurs much more slowly in distilled water than in water containing
a few parts per million of chlorides.
An increase in sodium chloride (NaCl) concentration has been shown
to cause a continual increase in the rate of stress corrosion cracking and
in the rate of total corrosion of magnesium and some aluminum alloys.
However, this does not hold true for all metal alloys.3
The addition of oxidizers (K2Cr2O7 or K2CrO4) to solutions of NaCl
inhibits total corrosion and increases the rate of stress corrosion in some
of the nonferrous alloys. This may be explained in part if we consider
that the formation of a large number of pits on a metal surface hinders
the development of isolated stress raisers into stress corrosion cracks.
The Influence of Oxygen—Oxygen plays an essential role in the process
of corrosion of unstressed metals. It probably has an even greater significance in the process of stress corrosion cracking. For example, an
experiment was performed with Al-5Mg and Al-7Mg alloys in aerated
and deaerated 3 per cent solutions of NaCl [6\. In aerated solutions, the
specimens cracked after several hours and in deaerated solutions, no
cracking was observed in 94 days.
The influence of oxygen from the air was also confirmed under conditions of full immersion and spraying with a 3 per cent solution of NaCl.
The immersion test produced failures in 365 days, while spraying produced failures in 55 days [7],
It has been noted that in alternate immersion testing, specimens of
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STRESS CORROSION TESTING
aluminum-zinc-magnesium alloys at a shallow depth in the test tank
failed faster than comparable specimens tested at a greater depth.5
Other work [7] states that for stainless steel, "salt spray stress corrosion
testing produced failures while alternate immersion proved inadequate
to cause stress corrosion failures in the materials tested."
The Influence of pH—The rate of stress corrosion cracking of most
metals generally decreases with increase in pH, especially in the alkaline
range [8,9]. The pH of most tests is a function of the environment used
unless an effort is made to control it. Most test environments will be
neutral or very weakly acid.
PART II. NONFERROUS ALLOYS
Aluminum
Laboratory Tests
Laboratory tests are advantageous in that they can be performed in a
convenient location under conditions that can be controlled. An ideal
laboratory stress corrosion cracking test must be simple to control, rapid,
and correlatable with service conditions.
The Alternate-Immersion Test—This is a popular test for aluminum
alloys, including aluminum-copper, aluminum-magnesium, and aluminum-zinc-magnesium, aluminum-zinc-magnesium-copper and aluminummagnesium-silicon types. Data obtained in this test may be correlated
with seacoast atmosphere and seawater tests [4,10]. This test is more
severe than inland industrial atmospheres for most alloys, the exception
being those of the aluminum-zinc-magnesium family that contain little
or no copper. In the case of the latter alloys, the 3.5 per cent NaCl
alternate-immersion test is less effective in some cases than exposure to
an industrial atmosphere. This test is fairly easy to control, but is not
rapid enough for some purposes. The cycle generally used is 10 min in
solution and 50 min out of solution.
Continuous Immersion—Continuous-immersion tests are relatively
simple to operate and control, but for some alloys they are not so effective
as alternate-immersion types. More aggressive electrolytes have been
used to make this procedure more effective. Immersion in boiling 1 N
sodium chloride solution is an effective test for aluminum-zinc-magnesium
alloys,3 but not so reliable for aluminum-zinc-magnesium-copper and
other types of alloys as the 3.5 per cent NaCl alternate-immersion test.
Acidified salt chromate solutions are effective in producing stress corrosion cracking, but they also produce severe localized intergranular or
pitting attack in unstressed specimens. The corrosive attack is so drastic
that "corrosion" failures will occur in relatively short exposure periods,
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
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and the test data obtained will be very unrealistic unless pains are taken
to separate the failures that are not the result of stress corrosion cracking.
Salt Spray Test—This is reported to be an effective test environment
and is reported to give results that are comparable to alternate-immersion
tests on most alloys. In general, the spray type of test is not so. easy to
control as an intermittent-immersion test. A salt spray test procedure is
outlined by Chadwick and Grainger. It is also reported that an acidifie^
(pH 3) 5 per cent intermittent spray is effective with an exposure period
of two weeks being sufficient for many aluminum alloys [11,12].
Electrolytic Tests—Electrolytic tests are highly accelerated stress corrosion cracking tests. They are conducted by impressing a current between
the specimen as the anode and a suitable cathode in a sodium chloride
solution. This type of test has been reported to be most successful for
aluminum-magnesium type alloys. The method has been shown to place
aluminum-magnesium alloys in the same order of stress corrosion susceptibility as the slower immersion test of stressed specimens in salt
peroxide solution and in the same order as obtained by immersion in
the sea and exposure in a seacoast atmosphere6 [13].
Highly accelerated tests such as this usually are most useful as screening tests for alloy development and quality control purposes.
Environmental Factors Which May Affect the Stress Corrosion Test Results of Aluminum
Humidity—An environment that will support the stress corrosion of
aluminum must provide an electrolyte in which the electrode relationships of the readily corrodible paths, such as the grain boundaries, are
anodic to the rest of the metal. It has been established that only very
small amounts of moisture are required on the metal surface to provide
the electrolyte3 [14]. Thus, stress corrosion cracking of susceptible aluminum alloys may occur in mildly corrosive environments such as the
atmosphere and in distilled water. The moisture films which develop
on a metal surface exposed to the atmosphere are acidic, especially in
seacoast and industrial locations. The cracking susceptibility of aluminum alloys reaches a maximum at about 85 per cent relative humidity.
Atmospheric Environments—Not all alloy types react the same to
different environments. For example, aluminum-copper, aluminummagnesium, and aluminum-zinc-magnesium-copper alloys are more
prone to stress corrosion cracking in seacoast atmospheres than in inland
industrial or rural atmospheres. An aluminum 5 per cent magnesium
alloy fabricated to be very susceptible to stress corrosion was exposed
at a tropical marine site in Aruba, Netherlands Antilles. It failed in one
day. At a marine site in Kure Beach, N.C., failure occurred in 32 days
and at a rural atmosphere in Richmond, Va., in 258 days [75]. But
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STRESS CORROSION TESTING
aluminum-zinc-magnesium alloys with relatively low copper (7079) and
no copper (7039) are just as susceptible in a less corrosive inland atmosphere as in the seacoast locations7 (also see Refs 4 and 16).
Temperature—This is the most critical single factor affecting the rate
of stress corrosion cracking for some aluminum alloys and should be
carefully controlled. An increase in temperature decreases the time-tofailure. If the log of the life is plotted against the reciprocal of the absolute
temperature, a straight line function is obtained. The straight line holds
to about 10 C at which point a change in slope indicates that the time-tofailure at 0 C would be infinitely long.3
Work indicates that the effect of temperature on time-to-failure of
aluminum alloys depends greatly on the alloy. An aluminum-zinc-magnesium alloy showed the time-to-failure was increased 12 min for each
degree centigrade. In a less susceptible aluminum-zinc-magnesium alloy
the time-to-failure was increased to 629 min per degree centigrade.3
Chlorides—The presence of certain anions, notably chlorides, tend to
accelerate the stress corrosion of aluminum. For this reason, and because
sodium chloride is so prevalent in natural environments, most stress
corrosion tests are carried out in sodium chloride solutions. Constant
deflection "loop tests" of sensitized Al-7Mg alloy sheet were conducted
by immersing specimens hi sodium chloride solutions of different concentrations ranging from 1.5 to 25 per cent [6]. As the concentration
was increased, the specimen life decreased from about 26 to 3 hr at 9
per cent NaCl after which the life remained constant. An experiment
was performed by one research laboratory using a sensitized Al-5Mg
alloy exposed by alternate immersion in sodium chloride solutions ranging
in concentration from 3.5 down to 0.005 per cent (50 ppm). Tension
specimens stressed to 75 per cent of the yield strength were subjected
to bending strain and exposed. Stress corrosion cracking occurred in
even the most dilute solution although the specimen life increased from
20 hr at 3.5 per cent NaCl to about 180 hr at 0.005 per cent [10]. In
another laboratory an aluminum-zinc-magnesium alloy showed no difference in failure time when specimens were tested at 60 per cent of the
yield strength in 0.01, 0.1, 1, and 2 N sodium chloride solution at 180 F.
But the time-to-failure was increased five times when tested in distilled
water.3
Hydrogen Ion Concentration (pH)—A change in the pH or other ions
present in the corrosive medium may alter the electrochemical' relationship of the metal constitutents and thereby affect the type of corrosion,
which in turn could influence the stress corrosion resistance. Work indicates that a change in mechanism occurs at about pH 3.6 (Fig. 1 from
footnote 3). In a laboratory test of an Al-lOMg alloy stress corrosion
cracking occurred in a neutral 53 g/liter NaCl + 3 g/liter H2O2 solution
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
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in 15 min. When the pH was raised by the addition of 10 g/liter NaOH,
no cracking occurred because general corrosion predominated rather than
intergranular attack [77]. In another experiment an Al-7Mg alloy failed
in 5 min in 3 per cent NaCl solution at pH 1 and hi 60 days at pH 10.7;
no failures occurred above this pH value.
Oxygen—Oxygen plays an essential role in the stress corrosion cracking of aluminum alloys. Tests have been conducted hi aereated and
deaereated 3 per cent solutions of NaCl, using Al-5Mg and Al-7Mg
alloys. In aereated solutions, the specimens cracked after several hours,
and in deaereated solutions, no cracking was observed in 94 days [6].
The influence of oxygen from the air was also shown by a comparison of
full immersion versus spraying with 3 per cent NaCl solution. Spraying
produced failures in 55 days hi contrast to 365 days for the immersion
test.
Results of Questionnaire on Stress Corrosion Testing of Aluminum
The survey by Task Group 2 of the Committee G-l members turned
up 37 investigators conducting stress corrosion tests on aluminum.
The tests were conducted for one or more of the following purposes:
1. Quality control
18
2. Development research
31
3. Applied research
35
4. Pure research
20
All the investigators are involved in more than one of the above programs. There is no exclusive pure research being conducted on aluminum
except through government-sponsored research. The groups doing this
work were not contacted by this survey.
The three general methods being used for testing are as follows:
1. Laboratory tests
32
2. Field tests
27
3. Service tests
16
It may be noted that most people conducting laboratory tests also
conduct field tests.
The method of exposing the specimens is by total or alternate immersion and high humidity. The number using each test is:
1. Total immersion
26
2. Alternate immersion
20
3. High humidity
14
The test environments are not so varied as in some of the other metals.
They are divided up as follows:
1. Sodium chloride
24
2. Salt spray
5
3.
Magnesium
chloride
4
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4. Seawater by
4
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STRESS CORROSION TESTING
5. Metal cleaning
2
Sodium chloride is certainly the most popular solution in use.
The water used for the test solution shows that distilled water is preferred. The tabulated list reads as follows:
1. Distilled
24
2. Deionized
10
3. Tap
6
4. Seawater
2
5. Unknown
2
The survey showed that about half of the people use a written procedure. The exact figures are 18 using a written procedure, 17 not using
a procedure, and 2 with "no comment."
Most testing is conducted under controlled temperature conditions.
The number reporting temperature control of the test environments are
27 and uncontrolled temperature of the test environments are 10.
The grade chemicals used may be tabulated as follows:
1. Industrial
9
2. U.S. pure
12
3. Chemical pure
21
4. Reagent
1
5. Unknown
3
The test environments are changed as shown below:
1. Continuously
2
2. Daily
4
3. Weekly
8
4. Monthly
2
5. Each two days
1
6. Varies
8
7. After each test
1
8. Unknown
5
9. Other
5
There appears to be more disagreement about the period of solution
change than any other factor.
Most members were noncommittal on the questions of factors other
than the above which affect stress corrosion. Some comments were, surface preparation, stress, specimen configuration, solution flow rate, and
pH.
Copper Alloys Questionnaire Results
The survey results from 19 investigators on copper and copper alloy
stress corrosion cracking are tabulated below.
Tests are conducted for one or more of the following reasons:
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1. Quality control
11
2. Development research
14
3. Applied research
15
4. Pure research
8
The tests are conducted in:
1. Laboratory tests
16
2. Field tests
11
3. Service tests
6
The specimens are exposed to the environment by:
1. Total immersion
14
2. Alternate immersion
7
3. High humidity
5
The test environments are:
1. Sodium chloride
7
2. Magnesium chloride
7
3. Mercury compounds
5
4. Ammonia vapor
5
5. Molten bismuth
2
6. Mercury metal
1
7. Mattsson's solution
1
8. Seawater
2
9. Other
2
Eleven investigators use a written procedure while eight do not. Seven
have a definite test period and twelve do not.
The water used for solution make up is:
1. Deionized
3
2. Distilled
10
3. Tap
2
4. Seawater
2
The grade chemicals are:
1. Industrial
3
2. U.S. pure
5
3. Chemically pure
11
4. Unknown
2
Fourteen control the test temperature and five do not. The solution is
changed:
1. Daily
1
2. Weekly
1
3. Constantly
4
4. Varies
4
5. Unknown
4
6. Other
4
7. As necessary
2
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STRESS CORROSION TESTING
Magnesium
Results of the Questionnaire on Magnesium Alloys
The survey results from three investigators on magnesium stress corrosion cracking are as follows:
Tests are conducted for one or more of the following reasons:
1. Quality control
2
2. Development research
2
3. Applied research
3
4. Pure research
3
The tests are conducted in:
1. Laboratory tests
3
2. Field tests
3
3. Service tests
2
One investigator has a written procedure while two do not. None has
a definite test period. The water used for solution makeup is:
3
1. Deionized
2. Distilled
3
The grade chemicals are:
1. U. S. pure
1
2. Chemically pure
2
Two control the test temperature and one does not. The solution is
changed:
1. Daily
1
2. Weekly
1
3. Varies
1
4. Unknown
1
Nickel
Results of the Questionnaire on Nickel Alloys
The survey results from eleven investigators on nickel and nickel alloy
stress corrosion cracking are tabulated below.
Tests are conducted for one or more of the following reasons:
1. Quality control
7
2. Development research
10
3. Applied research
8
4. Pure research
5
The tests are conducted in:
1. Laboratory tests
9
2. Field tests
7
3. Service tests
4
The
specimens
are
exposed
to
the
environment
by:
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1. Total immersion
10
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
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2. Alternate immersion
5
3. High humidity
3
The test environments are:
1. Sodium chloride
5
2. Magnesium chloride
5
3. Seawater
2
4. Unknown
1
Five investigators use a written procedure while six do not. Three have
a definite test period and eight do not.
The water used for solution make up is:
1. Deionized
2
2. Distilled
5
3. Tap
1
4. Unknown
3
The grade chemicals are:
1. U.S. pure
2
2. Industrial
3
3. Chemically pure
6
4. Unknown
3
Eight control the test temperature and two do not. The solution is
changed:
1. Constantly
1
2. Varies
7
3. Unknown
3
Titanium8
Service failures associated with stress corrosion cracking of titanium
alloys have been attributed to hot salt, hot chlorinated hydrocarbons,
hydrochloric acid, fuming red nitric acid, N2O4 , and methanol. Liquid
metal embrittlement of titanium has been attributed to molten cadmium
from overheated cadmium-plated fasteners.
Fairly reproducible laboratory tests for detecting stress corrosion cracking of titanium alloys have been devised for all of the above environments
except perhaps methanol. The cracking problem with methanol has arisen
only recently during the pressure testing of a Ti-6Al-4V Apollo fuel tank.
This failure is so recent that insufficient laboratory tests have been made
to define the conditions.
On the other hand, a laboratory test has been developed which causes
stress corrosion cracking in salt (NaCl) solutions, although there have
been no service failures of titanium in this environment. The test consists of propagating a crack which has been initiated in a bent-beam
specimen. The details of the test are described in the section on 3.5 per
cent NaCl (precracked specimens).
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STRESS CORROSION TESTING
A summary of the conditions of each of the laboratory tests is presented in the following sections along with a listing of the sensitive alloys
and a bibliography. The bibliography does not contain all the available
references on the subject but includes selective references which are
representative of those describing the particular test procedure.
Hydrochloric Acid Tests
The tests in hydrochloric acid have been conducted with bent beams
and U-bends. Susceptible alloys include: Ti-6.8Al-2.2Sn, Ti-8Al-lMoIV, Ti-12Zr-7Al, and unalloyed titanium. The test is described in Table
1.
TABLE 1 — Hydrochloric acid tests.
Test solution
Test temperature
Specimen type
Failure time
Type of cracking
Alloys which cracked
Alloys which did not crack
5 to 10% HC1 in water
95 F
4-point loading, bent beam and U-bend
instantaneous to >12 days
intergranular
(1) A-110AT (6.8A1, 2.2Sn)
(2) A-55 (commercial purity)
(3) A-70
(4) Ti-12Zr-7Al
(5) Ti-8Al-lMo-lV (transgranular)
C-110M (7.3 Mn) and Ti-3.36Mn.
TABLE 2— Red fuming nitric acid.
Test solution
Test temperature
Specimen type
Failure time
Type of cracking
Alloys which cracked
red fuming nitric acid (RFNA), 2.5 to 20% NO22
room temperature
U-bends and Erichsen cups
3 hr to 1 week (168 hr)
intergranular
(1) 75A (commercial purity) in anhydrous
RFNA containing 7% or more NO2 and in
RFNA -20% NO2 containing <0.6% water.
(2) Ti-8Mn in anhydrous RFNA containing
2.5% or more NO2 and in RFNA-20% NO2
containing <0.7% water.
Alloys which did not crack : . . .. none listed
Tests in Red Fuming Nitric Acid
Stress corrosion cracking in red fuming nitric acid depends upon the
NC>2 and moisture contents of the acid. Titanium 75A (commercial purity)
cracks in anhydrous acid when the NO2 content is greater than 7 per cent,
while Ti-8Mn cracks at NO2 greater than 2.5 per cent. Both resist cracking
in acid containing 20 per cent NO2 if the moisture content exceeds 0.6
to 0.7 per cent. Additions of 1 per cent NaBr to red fuming nitric acid
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also inhibits
stress
cracking
The
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197
Tests in Chlorinated Hydrocarbons
Cracking of titanium in chlorinated hydrocarbons has been observed
during stress relieving treatments at about 1150 F and hydrostatic testing at 600 to 700 F. The corrodent in the. stress relief failures was
believed to be residual trichloroethylene from vapor degreasing, while
that in the hydrostatic test was a proprietary high-temperature fluid.
Laboratory tests with chlorinated hydrocarbons have been conducted
in these two temperature ranges. Ti-5Al-2.5Sn was found to be susceptible to cracking at both temperatures. The tests are described in
Table 3.
TABLE 3 — Tests in chlorinated hydrocarbons.
Trichloroethylene:
Test solution
Test temperature . . . .
Specimen type
Failure time
Type of cracking
Alloys which cracked
Chlorinated Diphenyl:
Test solution
Test temperature
Specimen type
Failure time
Type of cracking
Alloys which cracked
Alloys which did not crack
as-received trichloroethylene
1150 to 1500 F
U-bend and circle-cross patch welds
1 to 16 hr
not given (probably intergranular)
A-110AT (5Al-2.5Sn)
as-received chlorinated diphenyl
600 to 700 F
welds and circular patch welds with 0 to 70,000
psi applied stress
V± to 3 hr
not given (probably intergranular)
RC A-110AT (5Al-2.5Sn)
. none listed
Hot Salt Cracking Tests
The hot salt cracking problem has received considerable attention
because the skin temperatures of the titanium wings in supersonic aircraft are expected to approach the 550 F temperature where cracking
can occur and because of the possibility of salt contamination during
heat treatment at higher temperatures (1100 to 1300 F).
Laboratory tests have been devised using NaCl and sea salt (natural
and synthetic) in the 550 to 950 F range and NaCl in the 1100 to 1300 F
range. Recent tests indicate that NaCl produces cracking more readily
than sea salt. The salt is applied by several cycles of dipping in a water
solution or paste followed by drying prior to exposure at high temperatures. The temperature for drying the salt solution ranges from room
temperature to 300 F.
Some tension tests are run for long periods of time at applied loads of
greater than 25 to 30 per cent of the 0.2 per cent offset yield strength.
The specimens are then pulled in tension to determine whether incipient
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STRESS CORROSION TESTING
TABLE 4— Tests in hot NaCl (550 to 950 F).
Starting test solution
Test temperature
Specimen type
Test time
Type of cracking
Alloys which cracked
Alloys which did not crack
. . 3% NaCl to supersaturated boiling NaCl solution dried on specimen at room temperature or
up to 300 F
. . 550 to 950 F
. . tension, bent beam, self-stressed spot welded
beams, right angle bends
. . 20 to 7000 hr
. . intergranular
. . Ti-8Al-lMo-lV, Ti-6AMV, Ti-12Zr-7Al, 2.5A1lMo-10Sn-5Zr, Ti-12V-llCr-3Al
. . TMAl-3Mo-lV (stressed to 100,000 psi and exposed 7000 hr at 500 F)
TABLE 5—Tests in sea salt (550 to 800 F).
Starting test solutions
Test temperature
Specimen Type
Test time
Type of cracking
Alloys which cracked
Alloys which did not crack . .
. . 6 parts NaCl-1 part MgCl
2 or 7 parts NaCl-1 part
MgCl2 (synthetic sea salt solutions); natural
sea salt slurry
. . 550 to 800 F
. . tension, precracked tension, self-stressed welded
beams, tube-type tension, tension specimens
with side hooks, notched and unnotched cantilever beams, bent beams
. . 50 to 20,000 hr
. . intergranular
. . Ti-8Al-lMo-lV, Ti-6AMV, Ti-5Al-2.5Sn, Ti4A1-3MO-1V, Ti-5Al-2.75Cr-l.25Fe, Ti-13VHCr-3Al, Ti-6Al-6V-2Sn
. . Ti-5Al-2.5Sn, TMAl-3Mo-lV, Ti-5Al-2.75Cr1.25Fe (precracked specimens exposed 50 hr
at 600, 700, 800 F at 50 to 80 per cent of 0.2 per
cent 100-hr creep strength), 6A1-4V (notched
and unnotched cantilever beams exposed
20,000 hr at 650 F at 30 to 100 per cent of yield
strength)
TABLE 6— Tests in hot NaCl (1100 to 1300 F).
Starting test solution
Test temperature
Specimen type
Test time
Type of cracking
Alloys which cracked
Alloys which did not crack . .
. 100 ppm solution to salt slurry (NaCl)
. 1100 to 1300 F
. bent beam, U-bend, tension
. . y% to 150 hr
. . intergranular
. . Ti-5Al-2.5Sn, Ti-5Al-5Sn-5Zr, Ti-12Zr-7Al,
TiSAl-lMo-lV
. . none listed
IMo-l V, and Ti-5Al-5Sn-5Zr appear to be the most susceptible to hot salt
cracking. The alpha-beta alloys are less susceptible, and the degree of
susceptibility appears to increase with increasing aluminum content. However, the Ti-8Mn alloy, which contains no aluminum, is also susceptible
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
199
Among the alloys with intermediate resistance to cracking are: Ti-5Al5Sn-5Zr-lMo-lV, Ti-6Al-4V, Ti-6Al-6V-2Sn, Ti-5Al-2.75Cr-l.25Fe,
and Ti-13V-llCr-3Al. Alloys exhibiting the most resistance to hot salt
cracking are: Ti-4Al-3Mo-lV, Ti-10Sn-5Zr-2.5Al-lMo, and Ti-4Mo4Zr-2Al. The laboratory tests are described in Tables 4, 5, and 6.
TABLE 7— Tests in 3.5% Nad (precracked specimens).
3.5% NaCl solution pH 8 or synthetic sea salt0
ambient
cantilever beam, notched, and with and without
crack initiated at base of notch by fatigue
less than 1 hr at critical load
Failure time
Type of cracking
intergranular and transgranular
unalloyed RS-70 (annealed)
Alloys which are sensitive
Ti-7Al-2Cb-lTa (annealed)
Ti-7Al-3Cb (annealed)
Ti-6Al-2.5Sn (annealed)
Ti-5Al-2.5Sn (annealed)
Ti-6Al-3Cb-2Sn (annealed)
Ti-7Al-3Cb-2Sn (annealed)
Ti-8Al-3Cb-3Sn (annealed)
Ti-8Mn (annealed)
Ti-8Al-lMo-lV (slightly) (annealed)
Ti-6AMV (very slightly) (annealed)
Ti-6.5Al-5Zr-lV (aged at 1100 F)
Ti-6Al-4V-lSn (aged 1100 F)
Ti-6Al-6V-2.5Sn (aged 900 F)
Ti-6Al-2Mo (aged 1100 F)
Ti-7Al-3Mo (annealed)
Ti-13V-llCr-3Al (annealed)
Alloys which are not sensitive.. Ti-65A (annealed)
Ti-6Al-4V (annealed and annealed and aged)
Ti-7Al-2.5Mo (annealed)
Ti-6Al-2Mo (annealed)
Ti-6Al-2Sn-lMo-lV (annealed)
Ti-6.5Al-5Zr-lV (annealed)
Ti-6Al-2Sn-lMo-3V (annealed)
Ti-5Al-2Sn-2Mo-2V (annealed)
Ti-6Al-2Cb-lTa-0.8Mo (annealed)
Ti-4Al-3Mo-lV (age hardened)
Ti-13V-llCr-3Al (age hardened)
Test solution
Test temperature
Specimen type
0
ASTM Specifications for Substitute Ocean Water (D 1141 - 52).
Tests in 3.5 Per Cent NaCl Solution or Seawater
Titanium alloys have exhibited an exceptional resistance to stress
corrosion cracking in salt solutions or seawater when exposed as bent
beams, U-bends, or tension specimens. However, in recent months it
has been shown that certain alloys show a marked tendency toward
crack propagation in salt water if they are precracked prior to exposure.
The test usually consists of machining a notch across the short edge of a
rectangular
cross-section beam. The beam is then fatigued until a small
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crack
is initiated. A
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200
STRESS CORROSION TESTING
the crack. The precracked specimen is then stressed as a cantilever beam.
Sucessively larger weights are added until the crack propagates. A stress
intensity factor is then calculated:
where:
m = moment at the notch (in-lb),
B = horizontal thickness of the specimen (in.),
D = vertical depth of specimen (in.), and
a = 1 — (a/D) where a = total initial depth of notch plus fatigue
crack (in.).
TABLE 8— Testing in NZO± .
Test solution
Test temperature
Specimen test
Failure time
Type of cracking
Alloys which cracked
Alloys which did not crack
liquid N2O4 containing <0.18% H2O and
<0.06% NO
85 to 165 F
bent beams or pressurized tanks > 40 ,000 psi
stress
1 to 20 days
intergranular
. Ti-6AMV; Ti-75A (commercial purity)
. . none listed
Similar tests are conducted in air (no salt solution) and the corresponding intensity factor is calculated. If the intensity factor for exposure to
salt water is less than that for air, the alloy is judged to be sensitive to
stress corrosion cracking.
Evaluations of the implications of this test are now under way. Initial
tests have indicated that a number of alloys are sensitive to stress corrosion cracking, but many are not (Table 7). In general, it appears that
sensitivity to cracking in seawater is associated with aluminum content in
the alloy, isomorphous beta stabilizers (columbium, molybdenum, and
vanadium) and aging heat treatments (900 to 1300 F).
Tests in N2O4
Service failures of Ti-6Al-4V tanks have occurred in cylindrical vessels
(0.020-in. wall) which were pressurized to 250 psi at 105 F with propellant
grade N2O4 (< 0,20 weight per cent moisture). Subsequent tests with
pressurized tanks and bent-beam specimens have shown that cracking
can occur at moderately high stresses in N2O4 containing less than 0.18
per cent H2O and < 0.06 per cent NO. Cracking is inhibited by additions
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
201
to the N2O4 of > 0.18 per cent H2O, > 0.06 per cent NO, or 7 per cent
HNO3 . The laboratory tests are described in Table 8.
Results of the Questionnaire on Titanium
The survey results from eight investigators on titanium stress corrosion
cracking are tabulated below:
Tests are conducted for one or more of the following reasons:
1. Quality control
2
2. Development research
8
3. Applied research
7
4. Pure research
4
The tests are conducted in:
1. Laboratory tests
6
2. Field tests
7
3. Service tests
8
The specimens are exposed to the environment by:
1. Total immersion
5
2. Alternate immersion
4
3. High humidity
3
The test environments are:
1. Sodium chloride
4
2. Magnesium chloride
1
3. Solid salt
1
4. Seawater
2
Six investigators use a written procedure while two do not.
Two have a definite test period and six do not. The water used for
solution makeup is:
1. Deionized
1
2. Distilled
5
3. Seawater
2
The grade chemicals are:
1. U.S. pure
1
2. Industrial
1
3. Chemically pure
5
4. Unknown
1
5. ASTM sea salt
1
Four control the test temperature and six do not. The solution is
changed:
1. Daily
2
2. Weekly
1
3. Constantly
1
4. Varies
3
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202
STRESS CORROSION TESTING
PART III. FERROUS ALLOYS
Mild Steel
Results of the Questionnaire on Mild Steel
The survey results from 14 investigators of mild steel stress corrosion
are tabulated below.
Tests are conducted for one or more of the following reasons:
1. Quality control
7
2. Development research
13
3. Applied research
12
4. Pure research
8
The tests are conducted in:
1. Laboratory tests
13
2. Field tests
10
3. Service tests
6
The specimens are exposed to the environment by:
1. Total immersion
11
2. Alternate immersion
7
3. High humidity
5
The test environments used are:
1. Sodium chloride
10
2. Magnesium chloride
5
3. Other
1
Four investigators have a written procedure and ten do not. Two have
a definite test period and twelve do not. The solution makeup water is:
1. Deionized
5
2. Distilled
6
3. Tap
2
4. Seawater
2
5. Unknown
1
The grade chemicals used are:
1. Industrial
1
2. U.S. pure
2
3. Chemically pure
6
4. ASTM sea salt
1
5. Unknown
3
Ten control the temperature of the test environment and four do not.
The solution is changed:
1. Weekly
2
2. Continuously
1
3. Varies
6
4. Unknown
5
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
203
High-Strength Steels
Results o] the Questionnaire on High-Strength Steel
The survey results from eight investigators on high-strength steel
stress corrosion cracking are tabulated below.
Tests are conducted for one or more of the following reasons:
1. Quality control
1
2. Development research
6
3. Applied research
7
4. Pure research
4
The tests are conducted in :
1. Laboratory tests
7
2. Field tests
6
3. Service tests
1
The specimens are exposed to the environment by:
1. Total immersion
6
2. Alternate immersion
4
3. High humidity
3
The test environments are:
1. Sodium chloride
5
2. Magnesium chloride
2
3. Salt spray
1
4. Other
4
Four investigators use a written procedure while four do not. Three
have a definite test period and five do not. The water used for solution
makeup is:
1. Deionized
2
2. Distilled
6
3. Tap
1
4. Seawater
2
The grade chemicals are:
1. U.S. pure
2
2. Chemically pure
5
3. Unknown
1
Three control the test temperature and five do not. The solution is
changed:
1. Daily
3
2. Weekly
3
3. Constantly
3
4. Varies
1
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STRESS CORROSION TESTING
Stainless Steels
Results of the Questionnaire on Stainless Steels
The survey by Task Group 2 of the members of Committee G-l turned
up 51 investigators conducting stress corrosion tests on stainless steels.
The tests are conducted for one or more of the following purposes:
1. Quality control
18
2. Development research
41
3. Applied research
42
4. Pure research
20
With five exceptions, each investigator includes more than one of the
above reasons for conducting tests. Two conduct only development research, one conducts only applied research, and two pure research.
The tests are conducted under three general headings:
1. Laboratory tests
47
2. Field tests
26
3. Service tests
23
Thirty-two people conduct two or three of the listed tests and seventeen use only one of the tests.
The specimens are exposed to the test environment by either partial,
total, alternate immersion, or high humidity. The number using each
test are:
1. Partial immersion
3
2. Total immersion
39
3. Alternate immersion
20
4. High humidity
18
The greatest number of variables are the test environments. While
sodium chloride and magnesium chloride appear to be the favorites,
many others are used as can be noted below:
1. Sodium chloride
25
2. Magnesium chloride
30
3. Distilled water
1
4. Sodium sulfate and acetic acid
1
5. Selenium oxide and hydrochloric acid
1
6. Acetic acid and hydrogen sulfide and sodium chloride
3
7. Acetic acid and hydrogen sulfide
1
8. Hydrogen sulfide and sodium chloride.... ,
1
9. Water and potassium hydroxide 650 F
1
10. Water and ferric chloride 650 F
1
11. Magnesium chloride and hydrogen peroxide 200 F
1
12. Salt spray
1
13. High purity water
1
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1
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
205
15. Process mixtures
2
16. Solid sodium chloride 1100-1400 F
1
17. Cathodic current in H2SO4
1
18. Seawater
5
19. Artificial seawater
1
20. Sodium hydroxide
1
With all the environments involved, it would seem that there would be
wide ranging disagreements on the degree of susceptibility of various
stainless steel alloys.
The survey revealed that less than half the investigators use a written
procedure. To be more specific, 20 use a written procedure, while 32 do
not.
There were also 38 of the polled who do not have a definite test period
and 13 who do.
The test solution makeup water shows that deionized and distilled
water are the most popular. The tabulated list is given below:
1. Deionized
20
2. Distilled
24
3. Tap
6
4. Unknown
2
Most tests have temperature controls possibly due to the fact that a
major portion are conducted at elevated temperatures. A total of 37
control the temperature while 16 do not.
The grade chemicals used show that certified pure is most widely
used. The tabulation is shown below:
1. Reagent
2
2. Certified pure
34
3. U.S. pure
13
4. Technical
1
5. Industrial
7
6. Unknown
4
The test environment was changed at intervals as shown below:
1. Constantly
5
2. Daily
8
3. Weekly
9
4. Monthly
2
5. Varies
18
6. Each specimen
2
7. Unknown
7
Most members were noncommittal on factors other than those above
which affect stress corrosion tests.
Some of the comments given were: sample configuration, stress application, surface preparation, specimen selection, solution flow rate,
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STRESS CORROSION TESTING
References
[1] Harwood, J. J., "The Influence of Mechanical Factors on Stress Corrosion,"
University of Tennessee Corrosion Conference, 1-3 March 1955, University
of Tennessee Press, Knoxville, Tenn.
[2] ASTM-AIME Symposium on Stress Corrosion Cracking of Metals, 1944,
American Society for Testing and Materials, Philadelphia, 1945.
[3] Godard, H. P., "The Corrosion Behavior of Aluminum," Corrosion, Vol. 11,
No. 12, Dec. 1955, pp. 542T-552T.
[4] Sprowls, D. O. and Brown, R. H., "What Every Engineer Should Know About
Stress Corrosion of Aluminum," Metal Progress, Vol. 81, No. 4, April and
May 1962.
[5] Bobylev, A. V., "Stress Corrosion Cracking of Brass," Metallurgizdat, 1955.
[6]Ferryman, E. C. W. and Hadden, S. E, "Stress Corrosion of Al-7Mg Alloy,"
Journal of the Institute of Metals, Vol. 77, 1950, p. 207.
[7] Croucher, T. R., "Northrop Norair ARTC Project 10-59."
[5] Gilbert, P. T. and Hadden, S. E., "A Theory of the Mechanism of Stress
Corrosion in Al-7Mg Alloy," Journal of the Institute of Metals, Vol. 77,
1950, p. 237.
[9] Romanov, V. V., "Stress Corrosion Cracking of Metals, A Bibliography,"
The National Science Foundation, Washington, D.C.
[10] Dix, E. H., Jr., Anderson, W. A., Jr., and Shumaker, M. B., "Development of
Wrought Aluminum-Magnesium Alloys," Alcoa Technical Paper 14, Alcoa
Research Laboratories, New Kensington, Pa., 1958.
[11] Lifka, B. W. and Sprowls, D. O., "An Improved Exfoliation Test for Aluminum Alloys," Corrosion, Vol. 22, No. 1, Jan. 1966.
[12] Chadwick, R. and Grainger, A. B., "Stress Corrosion of Wrought Ternary
and Complex Alloys of the Al-Zn-Mg System," Journal of the Institue of
Metals, Vol. 85, 1956-1957, p. 161.
[13] Booth, F. F. and Godard, H. P., "An Anodic Stress Corrosion Test for Aluminum-Magnesium Alloys," International Congress on Metallic Corrosion, London, 1961, p. 8.
[14] Dix, E. H., Jr., "Acceleration of the Rate of Corrosion by High Constant
Stresses," Transactions, Institute of Metals Div., American Institute of Mining
and Metallurgical Engineers, Vol. 137, 1940, p. 11.
[15] Ailor, W. H., "World Wide Atmospheric Test Program," Corrosion Technology, Nov. 1965.
[16] Aeronautical Systems Division Contract No. AF 33 (657)-8543, Douglas Aircraft Report 31421, 1 April, 1963.
[17]Mears, R. B., Brown, R. H., and Dix, E. H., Jr., "A Generalized Theory of the
Stress Corrosion of Alloys," ASTM-AIME Symposium on Stress Corrosion
Cracking of Metals, 1944, American Society for Testing and Materials, Philadelphia, 1945, pp. 323-339.
Bibliography on Titanium
[1] Fontana, M. G., "Stress Corrosion in Titanium and Its Alloys," Industrial &
Engineering Chemistry, Vol. 48, No. 9, 1956, pp. 59A and 60A.
[2] Meredith, Russell and Arter, W. L., "Stress Corrosion of Titanium Weldments,"
Welding Journal Supplement, Welding Research, Vol. 36, 1957, pp. 415s—418s.
[3] Kochka, E. L. and Peterson, V. C., "The Salt Corrosion of Titanium Alloys at
Elevated Temperatures," Final Technical Report, 15 January, 1961, Crucible
Steel Company of America.
[4] Rittenhouse, J. B., "The Corrosion, Pyrophoricity, and Stress-Corrosion Cracking of Titanium Alloys in Fuming Nitric Acid," Transactions, American Society for Metals, Vol. 51, 1959, pp. 871-895.
[5] Kiefer, G. C. and Harple, W. W., "Stress-Corrosion Cracking of Commercially
Pure Titanium," Metal Progress, Vol. 63, No. 2, 1953, pp. 74-76.
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[6] Rittenhouse,
J. Int'l
B., "The
Corrosion
andWed
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Nitric
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ROMANS ON TEST ENVIRONMENTS AND TEST DURATIONS
207
Acid," paper presented at the Second Pacific Area National Meeting of ASTM
in Los Angeles, Calif., 17-21 September, 1956.
[7] O'Connor, Eugene, "Investigation to Determine the Reaction of A-110AT Titanium Alloy to Various Manufacturing Process Materials as Contaminants
in an Air Atmosphere at Elevated Temperatures," Factory Laboratory Report F-4-3-59, 23 April, 1959, Solar Aircraft Co.
[8]Brown, Hiram, "Stress-Corrosion of Ti-5Al-2.5Sn," Memorandum 60, 4 August, 1960, Defense Metals Information Center, Columbus, Ohio.
[9] Braski, D. N. and Heimerl, G. J., 'The Relative Susceptibility of Four Commercial Titanium Alloys to Stress Corrosion at 550 F," NASA-Tn D-2011,
National Aeronautics and Space Administration.
[10]Covington, L. C. and Early, F. R., "Methods of Protecting Titanium Against
Hot Salt Stress Corrosion," Progress Report 21, Aug. 1964, Titanium Metals
Corporation of America, New York.
[11] Pride, R. A. and Woodward, J. M., "Salt-Stress-Corrosion Cracking of Residually Stressed Ti-8Al-lMo-lV Brake-Formed Sheet at 550 F (561 K)," NASATM X-1082, National Aeronautics and Space Administration, April 1965.
[12] Braski, D. N., "Preliminary Investigation of Effect of Environmental Factors
on Salt-Stress-Corrosion Cracking of Ti-lAl-lMo-lV at Elevated' Temperatures," NASA-TM X-1048, National Aeronautics and Space Administration,
Dec. 1964.
[13] Heimerl, G. J. et al, "Salt Stress Corrosion of Ti-8Al-lMl-lV Alloy Sheet at
Elevated Temperatures," paper presented at the Fifth Pacific Area National
Meeting of the American Society for Testing and Materials, Seattle, Wash., 31
October-5 November, 1965.
[14] Piper, D. E. and Fager, D. N., "The Relative Stress-Corrosion Susceptibility of
Titanium Alloys in the Presence of Hot Salt," paper presented at the Fifth
Pacific Area National Meeting of ASTM, Seattle, Wash., 31 October-5 November, 1965.
[15] Donachie, M. J., Danesi, W. P., and Pinkowish, A. A., "Effects of Salt Atmosphere on Crack Sensitivity of Commercial Titanium Alloys at 600-900 F,"
Pratt and Whitney Aircraft, East Hartford, Conn.
[16\ Rideout, S. P., Louthan, M. R., and Selby, C. L., "Basic Mechanisms of Stress
Corrosion Cracking of Titanium," paper presented at the Fifth Pacific Area
National Meeting of ASTM, Seattle, Wash., 31 October-5 November, 1965.
[17] Boyd, W. K. and Fink, F. W., 'The Phenomenon of Hot Salt Stress-Corrosion
Cracking of Titanium Alloys," NASA CR-117, National Aeronautics and
Space Administration, Oct. 1964.
[18] Avery, C. H. and Turley, R. V., "Chloride Stress Corrosion Susceptibility of
High Strength Stainless Steel, Titanium Alloy, and Superalloy Sheet," MLTDR-64-44 Vol. II, May 1964.
[79] Crossley, F. A., "Research on the Basic Nature of Stress Corrosion for Various
Structural Alloys at Room and Elevated Temperature," ASD-TR-61-713, May
1962.
[20] Logan, H. L. et al, "The Mechanism of Stress Corrosion of Titanium Alloys
Exposed to Sodium Chloride at Elevated Temperatures," Report 8690, May 4,
1964, National Bureau of Standards, Washington, D. C.
[21] Kirchner, R. L. and Ripling, E. J., "The Diffusion of Corrosion Products in
Hot Salt Stress-Corrosion Cracking of Titanium," Materials Research Laboratory, Inc., Oct. 1965.
[22] Martin, George, "Investigation of Long-Term Exposure Effects Under Stress
on Supersonic Transport Structural Alloys," paper presented at the Fifth
Pacific Area National Meeting of ASTM, Seattle, Wash., 31 October-5
November, 1965.
[23] Newcomer, R., Tourkakis, H. C., and Turner, H. C., "Elevated Temperature
Stress Corrosion Resistance of Titanium Alloys," Corrosion, Vol. 21, No. 10,
Oct. 1965, pp. 307-315.
[24]
Braski,by
D. ASTM
N., "Preliminary
Investigation
Effect
Environmental
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STRESS CORROSION TESTING
tures," NASA-TM X-1048, National Aeronautics and Space Administration,
Dec. 1964.
[25] Heimerl, G. J. et al, "Salt Stress Corrosion of Ti-8Al-lMo-lV Alloy Sheet at
Elevated Temperatures," paper presented at the Fifth Pacific Area National
Meeting of ASTM, Seattle, Wash. 31 October-5 November, 1965.
[26]Faulkenberry, B. and lannucci, A., "Effects of Sodium Chloride on Stress Corrosion Cracking of Titanium Alloy During Stress Relieving," Convair Report
MP 59-053, 19 May, 1959.
[27] Myers, D. E., "DOD High Strength Titanium Alloy Sheet Research Corrosion
Program," Report NA57H-527-18, 1 August, 1962, North American Aviation,
Inc., Columbus, Ohio.
[25] Murphy, J. F., and Carpenter, S. R., "Final Report On Alpha Titanium Manufacturing Development Sheet Program," RTD-TDR-63-4010, Oct. 1963.
[29] Brown, B. F., "A New Stress-Corrosion Cracking Testing Procedure for HighStrength Alloys," paper presented at the 68th Annual Meeting of ASTM
Lafayette, Ind. 13-18 June, 1965.
[30] Lane, I. R., Cavallaro, J. L., and Morton, A. G. S., "Seawater Enbrittlement
of Titanium," paper presented at the Fifth Pacific Area National Meeting of
ASTM, Seattle, Wash., 31 October-5 November, 1965.
[31] Brown, B. F. et al, "Marine Corrosion Studies, Third Interim Report of Progress," NRL Memorandum Report 1634, July 1965, Naval Research Laboratory.
[32] Hatch, A. J., Rosenberg, H. W., and Erbin, E. F., "Effects of Environment on
Cracking in Titanium Alloys," paper presented at the Fifth Pacific Area National Meeting of ASTM, Seattle, Wash., 31 October-5 November, 1965.
[33]Dohogne, C. L. et al, "A Study of the Stress-Corrosion Cracking of Titanium
Alloys in Seawater with Emphasis on the Ti-6Al-4V and Ti-8Al-lMo-lV
Alloys," Research Report R471, Oct. 18, 1965, Reactive Metals, Inc., Niles,
Ohio.
[34] "Nitrogen Tetroxide/Titanium Alloy Stress Corrosion Investigation," Report
8271-928060, Vol. I, Bell Aerosystems Co., Buffalo, N. Y. Contract NAS 9-150,
DMIC No. 65045.
[35] "Nitrogen Tetroxide/Titanium Alloy Stress Corrosion Investigation," Report
8271-928060, Vol. II, Bell Aerosystems Co., Buffalo, N. Y., Contract NAS
9-150, DMIC No. 65046.
[36] "Nitrogen Tetroxide/Titanium Alloy Stress Corrosion Investigation," Report
8271-928060, Vol. Ill, Bell Aerosystems Co., Buffalo, N. Y., Contract NAS
9-150, DMIC No. 65047.
[37] "Nitrogen Tetroxide/Titanium Alloy Stress Corrosion Investigation," Report
8271-928060, Vol. IV, Bell Aerosystems Co., Buffalo, N. Y., Contract NAS
9-150, DMIC No. 65048.
[38]Berry, W. E., White, E. L., and English, J. J., "Review of Recent Developments," Corrosion and Compatibility, Defense Metals Information Center, 30
November, 1966.
[39] Mori, Kensi, Takamura, Akira, and Shimose, Takaaki, "Stress Corrosion
Cracking of Ti and Zr in HCl-Methanol Solutions," Corrosion, Vol. 22, No. 2,
Feb. 1966, pp. 29-31.
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D. A. Vaughan1 and D. I. Phalen1
Reactions Contributing to the Formation
of Susceptible Paths for Stress
Corrosion Cracking
REFERENCE: D. A. Vaughan and D. I. Phalen, "Reactions Contributing
to the Formation of Susceptible Paths for Stress Corrosion Cracking,"
Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967,
p. 209.
ABSTRACT: Reactions between austenitic steels, martensitic steels, and
ternary aluminum alloys have been investigated by optical and electron
microscopy and by X-ray diffraction to evaluate metal-atomic hydrogen
reactions as potential methods of generating susceptible paths for stress
corrosion cracking of these materials. Significant structural changes are
observed to result from cathodic charging treatments which have been
compared with those that occur during stress corrosion. However, the
reaction products are shown to be highly anodic to the uncharged metal
and thus are very infrequently, if ever, observed in stress corrosion
cracks. The mechanisms by which susceptible paths may be generated in
stress corrosion cracking tests of these metals are discussed in terms of
the observed metal-hydrogen reaction. Structural characteristics of these
metals are shown to be contributing factors in the formation of these
paths.
KEY WORDS: corrosion, stress corrosion, cracking, steels, aluminum
alloys, stainless steels
The paths through which metals fail during stress corrosion have been
difficult, if not impossible, to predict on the basis of preexisting metallurgical or structural characteristics. Therefore, the mechanism by which
susceptible paths form prior to cracking in stress corrosion environments has been of considerable interest to the present authors for several
years. Through an understanding of the phenomena for various alloys, it is anticipated that methods for preventing stress corrosion cracking will be derived and fuller utilization of the physical properties will
result. Although alloy development or selection has been successful in
minimizing overall external corrosion reactions, there are insufficient
1
Associate chief and metallurgist, respectively, Structural Physics Div., Bat-
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STRESS CORROSION TESTING
data on the nature of internal reactions to identify specific characteristics
which lead to failure in stress corrosion environments. There are at least
two structural characteristics which could provide susceptible paths: (1)
lattice defects such as dislocation or stacking faults within the grains or
at grain boundaries and (2) second-phase material. Either of these structural properties will undoubtedly form connecting paths through which
failure may occur during stress corrosion. The density of defects is so
very high in most commercial alloys that if they are the sources of susceptible paths, a corresponding high density of cracks might be expected
to form during the initial exposure to stress corrosion environments.
Even though this correlation does not appear to exist, the defect structure cannot be ignored, and some excellent work is being carried out by
Pickering and Swann and others on the orientation of defects in material
with different susceptibility to stress corrosion cracking. Alternatively,
there are, in general, due to impurities or precipitation reactions, minor
phases present in most commercial alloys, which provide sites for initiation of stress corrosion. A direct correlation between stress corrosion
cracking and preexisting phases has not been attained. Thus, the studies
to be reported here suggest possible mechanisms by which susceptible
paths form during stress corrosion cracking tests for various alloys. The
specific reaction investigated is that which occurs at cathodic sites,
namely, the generation of atomic hydrogen on the metal surface.
Experimental Approach
Owing to the poor correlation between preexisting minor phases and
the susceptible paths for stress corrosion cracking of most alloys, the
approach taken by the authors has been directed toward understanding
reactions which could result in the formation of susceptible paths during
stress corrosion tests. The reactions of particular interest were those involving atomic hydrogen and the metal as would be expected to occur
during aqueous corrosion. It has been established by Nielsen and others
that hydrogen is produced during stress corrosion and that a rather large
quantity of hydrogen is released from cracks. The hydrogen that escapes
is molecular and would not be expected to react with the metal, but corrosion-produced hydrogen is initially in the atomic state, and reaction
with the metal is quite probable. Although only a small amount of hydrogen pickup has been reported to result during stress corrosion cracking
of metals, local concentration of hydrogen may be sufficient to produce
a second phase or to alter the mechanical properties of the matrix metal.
To increase the hydrogen concentration in bulk quantities of the metal,
electrolysis of aqueous solutions has been employed to increase the
atomic hydrogen concentration at the metal surface. This treatment increases the magnitude of the cathodic reaction that occurs at metal surCopyright
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in corrosion
processes.
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VAUGHAN AND PHALEN ON FORMATION OF SUSCEPTIBLE PATHS
211
cathode, that is the metal-hydrogen reaction products, can in some cases
be retained on the metal for analysis by metallographic and X-ray diffraction methods. Through an analysis of these bulk reaction products, a
better understanding of the potential cathodic reactions resulting from
metal corrosion may be possible.
Several types of austenitic stainless steel plus a high-strength (martensitic) steel and a high-strength aluminum alloy (Table 1) have been
reacted with cathodically generated hydrogen, as described above, and
the resulting products have been examined by optical and electron miTABLE 1 — Composition of alloys investigated (weight per cent).
AUSTENITIC STAINLESS STEEL
Alloy
Commercial
304°
Ab
B°
e'-
er
Ni
Mn
18.8
17.7
17 6
17 3
9.4
8.0
17.8
17.9
1.6
1.5
15
2.5
N
0.023
0.006
0 036
0 010
C
0.06
0.075
0 017
0 055
Si
P
S
Fe
0.58
0.50
0 49
0 24
0.04
0.028
0 023
0 007
0.016
0.012
0 021
0 Oil
balance
balance
balance
balance
P
S
Fe
HIGH-STRENGTH STEEL ALLOY
Cr
Commercial
4340°
0.80
Ni
Mn
Mo
1.82 0.67 0.21
C
0.41
Si
0.27 0.008 0.016 balance
HIGH-STRENGTH ALUMINUM ALLOY
Commercial
7079°
D°
0
6
Zn
Mg
Cu
Mn
4.36
4.05
3.14 0.65 0.23
3.01 ...
Fe
Cr
Si
0.16
0.01
0.13 0.07
... 0.01
Ti
Al
0.02 balance
balance
Susceptible to cracking.
Resistant to cracking.
croscopy and by X-ray diffraction. Throughout these studies, a comparison was made of structural characteristics of the metals before charging
with those obtained after charging. Furthermore, the comparison was extended to include structural characteristics of these materials after exposure to stress corrosion cracking environments.
Results and Discussion
Austenitic Stainless Steel
The results on cathodic charging of austenitic stainless steels have been
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STRESS CORROSION TESTING
FIG. 1—Photomicrograph of austenitic stainless steel specimens after cathodic
charging for Vi hr; (a) susceptible and (b) resistant to stress corrosion cracking
(X250).
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VAUGHAN AND PHALEN ON FORMATION OF SUSCEPTIBLE PATHS
213
both susceptible and resistant to stress corrosion cracking, were reacted
with cathodically generated hydrogen and examined by the methods previously described. Distinct differences in the reaction products were observed between the two types of material, namely, the hydrogen formed
a metastable solution in the austenite lattice for susceptible alloys but
formed hydride phases immediately when reacted with the resistant alloys. This difference is revealed metallographically in Figs, la and b
which show the surfaces of susceptible and resistant specimens after cathodically charging for l/2 hr. X-ray diffraction data are presented in Fig.
2 showing changes in lattice structure of the austenite due to cathodic
FIG. 2—Graphical representation of X-ray diffraction
charged stainless steel.
results on hydrogen-
charging of these materials. Plots A, D, F, and G of Fig. 2 represent single phases, which are present in varying amounts in plots B, C, and E.
The hydrogen martensite indicated in plot B, so designated because of
the broad diffraction lines and its formation on charging, is unstable and
reverts back to austenite upon aging 8 to 16 hr at room temperature.
This phase has a very small crystallite size as the breadth of the diffraction lines would suggest. Furthermore, there is no indication of its presence in the photomicrograph shown in Fig. la.
The above reactions could be expected to occur to a limited extent in
corrosion. In stress corrosion cracking, however, the applied stress is essential and is suspected as a contributing factor in the formation of the
susceptible paths. Therefore, the cathodically charged specimens were
deformed
TheWed
resulting
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STRESS CORROSION TESTING
FIG. 3—Photomicrograph of cathodically charged stainless steel specimen after
plastic deformation; (a) susceptible and (b) resistant to stress corrosion cracking
(X250).
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VAUGHAN AND PHALEN ON FORMATION OF SUSCEPTIBLE PATHS
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FIG. 4—Transmission electron micrographs of A1S1 steel quenched from
1600 F and tempered at temperatures indicated; (a) as quenched and (b) 400 F
(X14875).
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STRESS CORROSION TESTING
FIG. 4 (cent.)—Transmission electron micrographs of AISI steel quenched
from 1600 F and tempered at temperatures indicated; (c) 800 F and (d) 1300 F
(X14875).
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VAUGHAN AND PHALEN ON FORMATION OF SUSCEPTIBLE PATHS
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are incorporated in Fig. 3, which show that transgranular bands form
perpendicular to the tensile direction in the susceptible alloy (Fig. 3d),
but only grain boundary separation occurs in the resistant alloy (Fig. 3b).
Thus, a potential mechanism for the formation of susceptible paths during
stress corrosion cracking of stainless steel consists of corrosion produced
hydrogen dissolving in the austenite lattice and stress-induced diffusion
of this hydrogen to concentration sites, which are probably local regions
of maximum tensile stress. Confirmation of the susceptibility of the transgranular bands, so generated, was established by exposing the charged
and deformed steel to boiling 42 per cent magnesium chloride (MgCl2)
FIG. 5—Change in microstress and crystalline domain size with tempering temperature.
in which the bands were attacked very rapidly. Also, potential measurements revealed cathodically charged specimens to be 200-mv anodic to
an uncharged specimen of the steel when the two are coupled in the
MgCl2 solution.
To evaluate the above proposed mechanism for stress corrosion cracking of the austenitic stainless steels and to design alloys resistant to failure by stress corrosion, it is recommended that factors controlling the
type of reaction between atomic hydrogen and the metal need to be understood in more detail. For example, the effects of structural defects,
of electron structure resulting from compositional changes, and of impurities in the metal on hydrogen solubility in the austenite lattice have
not been investigated. Also, the hydrogen transport mechanism under
applied stress has not been resolved.
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STRESS CORROSION TESTING
FIG. 6—Transmission electron micrographs of AISI 4340 steel cathodically
charged after tempering at indicated temperatures; (a) 400 F and (b) 1300 F
(X14875).
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Martensitic High-Strength Steel
In the case of high-strength steel (AISI 4340), the reaction with cathodically generated hydrogen is more subtle in that no second phases
are produced or detected either by X-ray diffraction or by optical metallographic methods. The inability to detect a second phase by the latter
method is not surprising in view of the fact that the martensitic grains
are difficult to resolve by optical methods. X-ray diffraction analyses of
the quenched and the 400 F tempered specimens show the domain size
to be approximately 200 A units. This size was confirmed by transmission electron microscopic methods. The change in microstructure as a
function of tempering temperature is shown in Fig. 4. Although the defect structure present in the as-quenched specimen (Fig. 40) appears to
be partially removed by tempering at 400 F (Fig. 46), X-ray diffraction
studies show little or no change in domain size. Structural changes become quite apparent by electron microscopic methods after tempering at
higher temperature (Figs. 4c and d), which coincide with an increase in
domain size on tempering at 600 and 800 F (Fig. 5). Included in Fig. 5
are the results of microstress analysis as a function of tempering temperature. The microstress decreases with tempering temperature, the largest
change occurring in the temperature range 600 to 800 F as was observed
for the domain size.
The specimens described in Fig. 4 were cathodically charged and reexamined to determine structural changes resulting from the reaction
between the steel and atomic hydrogen. Representative electron micrographs of the 400 and 1300 F tempered specimens after hydrogen charging are presented in Figs. 6a and b, respectively. The most interesting
result in this study is the substructure produced in the 400 F tempered
steel by cathodic charging. Little or no substructure is produced upon
cathodic charging the 1300 F tempered steel, but the ferrite grain boundaries are widened to some extent. A comparison of Fig. 6a with Figs.
4a and b reveals that the structural defects of Fig. 4a are also present in
Fig. 6a but not in Fig. 4b. This sequential investigation indicates that the
defect structure, which appears to be partially removed upon tempering
at 400 and 600 F, is regenerated by cathodic charging. Thus, hydrogen
appears to enter defects in the martensite plates of this high-strength (284
ksi) steel, but it does not appear to enter the grains of the steel tempered
at 1300 F, which is of considerably lower strength (120 ksi).
As this change in defect structure is not observed upon charging the
steel specimens tempered at 1300 F, compare Fig. 4d with Fig. 6b; the
reaction with atomic hydrogen appears to differ from that of the higherstrength (low-temperature temper) material. In particular, the ferrite
grain boundaries (Fig. 6b) are broadened by hydrogen charging. An examination of the metal surfaces," after charging with hydrogen, also reCopyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
vealed differences in reaction behavior for the steels tempered at low
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STRESS CORROSION TESTING
FIG. 7—Electron micrographs of the surface of AISI 4340 steel cathodically
charged after tempering at the temperatures indicated; (a) 400 F and (b) 1300 F
(X8750).
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VAUGHAN AND PHALEN ON FORMATION OF SUSCEPTIBLE PATHS
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and high temperature. The difference in surface reaction is illustrated
in Fig. 7. For low-temperature tempered steel (Fig. Id), the surface structure remained unchanged, which is undoubtedly due to the easy entry of
hydrogen through the defect structure in this specimen. In the case of
higher-temperature tempered material, cathodic charging produced blisters on the metal surface (Fig. 76). This indicates that a lower void-space
volume is present in the grain boundaries of the latter specimen than is
available in the domain boundaries of the higher-strength steel.
These structural changes resulting from metal-hydrogen reactions suggest that the cathodic portion of a stress corrosion reaction may be
effective in altering the microstructure of this steel. These structural
changes are probably restricted to the surface layers which would be in
the tensile zone of stress corrosion test specimens, where the morphology
of stress corrosion crack exhibits a transgranular path, namely, through
the martensite plates of the high-strength steel. This path changes to an
intergranular type fracture in the neutral and compressive zones of stress
corrosion cracked specimens. In these latter zones, the fracture morphology indicates that the fracture path is the prior-austenite grain
boundaries. In view of the above discussion on the probable effect of
hydrogen in the tensile zone, it is likely that the intergranular failure in
the other zones results from hydrogen diffusing to and concentrating in
the prior-austenite boundaries. This is consistent with the mechanism
proposed by Fetch and Stables on the lowering of grain-boundary energy
by its adsorption of hydrogen.
The structural modifications induced by hydrogen charging of the
lower-strength (high-temperature temper) steel specimens are limited to
ferrite grain boundaries, which in some respects are similar to those described above, even though failures do not occur as frequently in the
lower-strength steel. This decrease in incidence of failure may be a result of partial elimination of the prior-austenite grain boundaries so that
the neutral and compressive zones do not have rapid diffusion paths for
hydrogen. The lack of a continuous grain boundary would not necessarily eliminate delayed failures by the metal-hydrogen reaction. However, at higher tempering temperatures, the structural defects within the
martensite and ferrite grains are reduced or eliminated, so that hydrogen
diffusion would be expected to be much slower. The studies of hydrogen
reaction with 4340 steel specimens tempered at 800 and 1300 F indicate
that hydrogen enters the ferrite boundaries formed by these heat treatments. As these grains are much smaller than the prior-austenite grains,
the diffusion path would be appreciably longer. The hydrogen reaction
with the surface of these specimens was found to differ from that observed for the low-temperature tempered specimens in that blisters
formed on the former but not on the latter specimens. This difference was
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STRESS CORROSION TESTING
FIG. 8—Electron micrograph of (a) cathodically charged surface
stress corrosion fracture of Al-4Zn-3Mg alloy (X8750).
and (b)
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VAUGHAN AND PHAUEN ON FORMATION OF SUSCEPTIBLE PATHS
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slower rate, and the higher tempering temperature has reduced the defect concentration and thus the void space for hydrogen to enter the lattice.
Improvement in the resistance of the high-strength 4340 steel to stress
corrosion fracture may be attained (1) by preventing diffusion of hydrogen picked up during corrosion through eliminating the prior-austenite
grain boundaries and (2) by reducing or changing the character of the
defect structure within the martensite grains.
FIG. 9—Electron micrograph of thin section of Al-4Zn-3Mg alloy; (a) as
thinned (X12250).
High-Strength Aluminum Alloy
The fracture path for stress corrosion cracking of Type 7079 aluminum alloys has been established as the grain boundaries. Since the details of the failure mechanism are not completely understood, it was of
interest to determine how the cathodic portion of the corrosion reaction
might contribute toward understanding acceleration of the attack when
stress is present. For this investigation, a special heat of Al-4Zn-3Mg
ternary alloy was prepared which had failure times very nearly the same
as those of commercial Type 7079 aluminum. As an initial analysis of the
mechanism, the morphological characteristics of cathodically charged
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specimens
were compared with those in the fracture surface of stress
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STRESS CORROSION TESTING
FIG. 9 (cent.)—Electron micrographs of thin section of Al-4Zn-3Mg alloy;
(b) cathodically charged and (c) exposed to 3.5 per cent Nad solution without
applied potential (X12250).
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VAUGHAN AND PHAUEN ON FORMATION OF SUSCEPTIBLE PATHS
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corrosion cracks. Figure 8 shows that the surface structures are very
similar. While the structure on the surface of the charged specimen was
essentially uniform from area to area, the corresponding pit structure
(black dots) was observed only near the apex of the crack in the stress
corrosion specimen. The remaining surfaces of the crack were relatively
pit free and were similar to those observed when bulk specimens were
made anodic in the test solution which contained 3.5 per cent sodium
chloride (NaCl). The pits observed in the electron micrographs of Fig.
8 are believed to result from a two-step process: (1) reaction of the metal
with atomic hydrogen and (2) preferential dissolution of the reaction
product by the test solution. Thus, in aluminum alloys as in the case of
austenitic stainless steel, the metal-hydrogen reaction products are attacked rapidly by the stress corrosion testing solutions. However, in view
of the results on the steel-hydrogen reactions, it appeared that other
structural changes could occur upon cathodically charging the aluminum
in electrolytes other than the NaCl solution.
Thin sections of the ternary alloy were prepared and 'examined (1) as
thinned, (2) after cathodic charging in an arsenic saturated 5 per cent
H2SO4 electrolyte, and (3) after exposure to 3.5 per cent NaCl solution
in the absence of an applied voltage. Electron micrographs of thin sections of the ternary alloy in the above conditions are shown in Fig. 9.
Although this alloy was given the T6 aging treatment, the grains exhibit
little or no second phase, and only narrow strain lines are seen traversing
the grains in the as-thinned condition. There are very small precipitates
in the grain boundaries. The narrow width of the boundaries suggests that
these adjacent grains have nearly the same crystallographic orientation.
It would be of extreme interest to determine whether stress corrosion
cracking paths propagate along high- or low-angle boundaries.
After cathodic charging a thinned section of this alloy (Fig. 9Z>) was
obtained. Here the pits, light spots on the print, are distributed rather
uniformly over the grain except in a zone, approximately 2500 A wide,
on either side of the grain boundary where the density of pits decreases.
The structure within this zone differs from the matrix grain after charging in that the fine strain lines or subgrain structures that formed in this
narrow zone appear to be oriented perpendicular to the grain boundary.
As no stress was intentionally applied during charging, it is evident that
these structural characteristics are inherent in the metal and when reacted with atomic hydrogen become detectable. This zone may result
from depletion of the alloy during formation of the grain-boundary precipitate. It is quite likely that the physical and chemical properties are
thus altered so that this narrow zone becomes the susceptible path for
stress corrosion cracking.
Figure 9c shows an electron micrograph of a thin section of this alloy
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STRESS CORROSION TESTING
occurring as a result of this treatment is a broadening of the strain contour lines even though no stress was applied. This was also observed
upon cathodic charging (Fig. 96). In addition, Fig. 9c shows that a portion of the grain boundary has been attacked. This attack appears to extend into the grain, particularly in the upper right corner of Fig. 9c,
which would indicate an accelerated reaction in a narrow zone adjacent
to the grain boundary. This attack switches from one side to the other
along the boundary. Although the previously mentioned depleted zone is
not as pronounced in Fig. 9c, there are some indications of its presence
by the type of attack and by the shape of some strain contour lines as
they approach the grain boundary.
In this aluminum alloy, it would appear that the susceptible path may
be present, but that the cathodic portion of the corrosion reaction may
provide additional structural changes and thus increase the sensitivity of
the grain-boundary zone to attack by the corrosion medium or to the
propagation of cracks by an applied stress.
Conclusions
From the observations described in this research, it must be concluded
that the type of metal-hydrogen reaction depends, to a large extent, upon
the prior metal structure. The similarity between structural changes that
occur by stress corrosion and by cathodic charging suggests that the cathodic portion of corrosion reaction may contribute significantly toward
the generation of susceptible paths during stress corrosion cracking tests.
In all cases, the cathodic reaction is limited to generation of the susceptible path, but an anodic reaction is necessary to produce a stress corrosion crack. It must be pointed out that the high-strength steel and,
possibly, the aluminum alloy may fracture under cathodic charging alone;
however, this type of failure is usually defined as hydrogen embrittlement. These alloys do not form hydride phases as does the austenitic
stainless steel. The latter steel does not become embrittled, which is undoubtedly due to the short-range diffusion of hydrogen before the precipitate occurs.
Acknowledgments
The authors wish to acknowledge United States Steel Corp.; U.S.
Navy, Bureau of Weapons; and U.S. Air Force, Materials Laboratory,
Wright-Patterson Air Force Base; who contributed the support for this
research.
Bibliography
[1] Swann, P. R. and Nutting, J., Journal of the Institute of Metals, Vol. 88, 1960,
p. 478.
[2]
Swann, by
P. R.,
Corrosion,
Vol.
19, No.
3, 1963,
p. 102t.
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VAUGHAN AND PHALEN ON FORMATION OF SUSCEPTIBLE PATHS
227
[5] Pickering, H. W. and Swann, P. R., Corrosion, Vol. 19, No. 11, 1963, p. 373t.
[4] Nielsen, N. A., Corrosion, Vol. 20, No. 3, 1964, p. 104t.
[5] Vaughan, D. A. et al, Corrosion, Vol. 19, No. 9, 1963, p. 315t.
[6]Vaughan, D. A. and Phalen, D. L, Metals Engineering Quarterly, Vol. 5, No.
3, 1965, p. 39.
[7] Haynie, F. H. et al, 66-267, 1966, Air Force Materials Laboratory.
[8] Fetch, N. J. and Stables, P., Nature, Vol. 169, 1952, p. 842.
DISCUSSION
Daniel van Rooyeri^ (written discussion)—It was interesting to hear of
an 18Cr-8Ni stainless steel which is immune to stress corrosion cracking. Can the authors give some details of the tests in which the alloy was
immune and also any metallurgical or other reason for this resistance to
cracking in an austenitic 18Cr-8Ni alloy? Also, details of the 18Cr-18Ni
alloy, which was mentioned, would be welcome.
Was the electrochemical potential, used for charging specimens with
hydrogen, in the range which may be expected at local cathodes during
stress corrosion cracking of austenitic stainless steels?
D. A. Vaughan and D. I. Phalen (authors)—The compositions of the
stainless steel alloys used in this investigation, as given in Table 1, show
the major difference to be in nitrogen content which is significantly higher
for susceptible than for resistant material. The susceptibility to stress
corrosion cracking was established during exposure to boiling 42 per cent
MgCl2 solution for times up to 500 hr with some tests extending to 8000
hr. Specimens of these materials were used to determine the structural
details of the metal-hydrogen reaction. The electrochemical potential
used to generate hydrogen at the metal surface was of the order of 5 v,
which may have exceeded that at local cathodes during stress corrosion
as the absolute value is not well known. However, the present study was
intended to produce sufficient reaction product to be detected by X-ray
diffraction rather than just enough to initiate one or a few susceptible
paths.
Section supervisor, The International Nickel Co., Inc., Paul D. Merica Research Laboratory, Sterling Forest, Suffern, N. Y.
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E. N. Pugh1 andA.R.C. Westwood1
Critical Species in Stress Corrosion
Phenomena
REFERENCE: E. N. Pugh, and A. R. C. Westwood, "Critical Species in
Stress Corrosion Phenomena," Stress Corrosion Testing, ASTM STP 425,
Am. Soc. Testing Mats., 1967, p. 228.
ABSTRACT: Consideration has been given to the identification of the
critical species in several stress corrosion systems. It is shown that in the
a-brass/aqueous ammonia system, cupric complex ions of the type
Cu(NH3)n2+ play a controlling role in the cracking process. Complex ions
are also found to constitute the critical species in the embrittlement of
silver chloride in certain aqueous environments. In the case of materials
such as stainless steels and magnesium and aluminum alloys, which undergo stress corrosion cracking in chloride environments, the critical species may be the chloride ion itself or metal-chloride complexes.
Attention is given to both the role of the critical species in the mechanisms of failure and the practical significance of these findings to stress
corrosion testing. It is suggested that more attention to the chemistry of
environments which cause stress corrosion cracking, with particular regard
to the identification of the critical species, could be of significant practical
value.
KEY WORDS: corrosion, stress corrosion, cracking, brass, stainless
steels, magnesium alloys, aluminum alloys
It is generally accepted that, for any given material, stress corrosion
phenomena occur only in certain specific environments. Thus, it is somewhat surprising that relatively little attention has been given to the chemistry of these environments, particularly to the identification of the particular species responsible for failure. Nevertheless, it is readily apparent
that if the nature of the critical species and its role in the cracking process
can be established, then the possibility of being able to suggest means of
preventing or inhibiting stress corrosion cracking is much improved. This
paper describes the results of recent studies which have led to the identification, and in one instance control, of the critical species in certain
systems, and then discusses the implications of this work with regard to
1
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PUGH AND WESTWOOD ON CRITICAL SPECIES
229
the mechanisms of stress corrosion cracking and to the general problems
of stress corrosion testing.
For the purposes of this paper, the term stress corrosion cracking is
considered, quite arbitrarily, to encompass the embrittlement of crystalline solids in aqueous environments. The term is also used in a generic
sense, for it is becoming increasingly evident that no single unified mechanism of failure exists. Indeed, more than one mechanism may operate
FIG. 1—Stress corrosion data for 70-30 brass specimens stressed in fresh or
preconcentrated 15 N aqueous NHs. The copper exists in the preconcentrated
solutions as C^NHs)^ ions [5].
in any given material/environment system. For example, it will be seen
in the following section that in the classical a-brass/ammonia system at
least two mechanisms are thought to operate, the determining factor being the chemical composition of the environment [I].2
The a-Brass/Aqueous Ammonia System
An important contribution to our understanding of the role of the
ammoniacal environment in the stress corrosion cracking of a-brass was
2
The italic numbers in brackets refer to the list of references appended to this
paper.
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STRESS CORROSION TESTING
made by Althof [2] in 1944, when he demonstrated that cracking did not
occur until the testing solution turned blue and that stress corrosion life
was markedly reduced when the test solutions were preconcentrated
with copper before the test commenced. The significance of these observations has become apparent following their confirmation in more
recent studies [1,3-6], for example, Fig. 1. It is known that in aqueous,
oxygenated ammoniacal solutions, copper exists as cupric complex ions
of the type Cu(NH3)n2+, the number of ammonia ligands, n, varying
from 1 to 5 depending on the ammonia concentration of the solution
FIG. 2—Effect of varying the volume of the testing solution on the time-tofailure of 70-30 brass specimens stressed in fresh or preconcentrated 15 N aqueous
NH3 [5].
[7]. The presence of these complex ions gives rise to the blue color of the
solution. If it is then assumed that such complex ions play an important
role in the stress corrosion process, the differences in time-to-failure of
specimens tested under a given stress in either "fresh" or preconcentrated
solutions (Fig. 1) may be attributed to the fact that time is required in
fresh solutions for the production, via dissolution of the specimen, of a
sufficient concentration of these complex ions. In preconcentrated solutions, of course, this critical ion species is present at the beginning of the
test. Data from experiments in which the volume of the testing solution
was varied (Fig. 2) also may be explained on the basis of this hypothesis.
In fresh solutions, the time required to build up a sufficient concentration
ot complex ions to cause failure increases with increasing volume; in
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preconcentrated
solutions, time-to-failure is independent of volume [5].
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PUGH AND WESTWOOD ON CRITICAL SPECIES
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Some differences of opinion appear to exist concerning the exact
complex ion which is responsible for failure. Complex ions containing
1.5 to 3.0 [8] and 4 [2-4] ammonia ligands each have been claimed to
play major roles in the stress corrosion process, but these claims have not
been substantiated by experimental evidence. Recently, Pugh et al [5]
FIG 3
- —Effect of copper content of 15 N aqueous NH* on (a) time-to-failure
and (b) rate of weight loss of 70-30 brass. In the stress corrosion data, the points
represent the average of at least four tests and the bars indicate the highest and
lowest values [1].
have combined studies of stress corrosion cracking of a 70-30 brass with
spectrophotometric studies of the ammoniacal testing solutions. Experiments conducted in 15 N and 1 N solutions preconcentrated with copper
established that rapid cracking occurred in each solution. Comparison
of the absorption spectra of the test solutions with published spectroCopyright by data
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232
STRESS CORROSION TESTING
were Cu(NH3)52+ and Cu(NH3)42+, respectively. Work is in progress to
determine whether cracking can occur when complexes containing fewer
ligands predominate.
To investigate the role of the complex ions in the cracking process,
Pugh and Westwood have studied the effect of varying the complex-ion
concentration on the behavior of 70-30 brass tested in oxygenated 15 N
aqueous ammonia [1]. The time-to-failure, tF, rate of weight loss, and
surface condition were each found to depend on the complex-ion concentration, in this case Cu(NH3)52+. Figure 3a illustrates the relationship
between tF for specimens tested under a constant load and copper content of the solution which, in these experiments, was directly proportional
to the concentration of Cu(NH3)52+ ions. It can be seen that tF decreased
with increasing copper content, as expected from the earlier results, for
example, Fig. 1, but that a well-defined inflection occurred at a critical
concentration. This inflection coincided with a change in surface condition. Specimens tested in solutions of copper content exceeding the
critical value were coated with the characteristic black oxide coating
commonly termed the tarnish, while specimens tested in solutions of
lower concentrations were apparently free from this coating. The relationship between rate of weight loss and copper content of the solution
also exhibited a maximum at this critical concentration (Fig. 3b).
These and other observations have led to the conclusion that two
mechanisms of stress corrosion cracking are operative in this system, one
in the presence of the tarnish and the other in the absence of this layer
[1]. The inflection in the stress corrosion data (Fig. 3a) is considered to
correspond to the transition between these mechanisms.
The mechanism of stress corrosion cracking in solutions which do not
cause tarnishing is not fully understood, but it is currently considered to
occur by a dissolution-dependent mechanism involving the following
autocatalytic reaction between the cupric complexes and copper atoms
at the brass surface.3
Since this reaction does not involve zinc, then it might be expected
that cracking also should occur in pure copper. Earlier work suggested
that the pure metal is immune to this type of failure [10, 11], but recent
studies have demonstrated that cracking does in fact occur provided that
the ammoniacal solutions contain certain critical concentrations of
Cu(NH3)52+ ions [12].
3
This reaction accounts for the initial increase in the rate of weight loss with
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PUGH AND WESTWOOD ON CRITICAL SPECIES
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In tarnishing solutions, stress corrosion cracking is thought to occur
by the tarnish-rupture mechanism first proposed by Forty and Humble
[13]. According to this mechanism, cracking is confined to the tarnish
and proceeds by the repeated formation and rupture of this brittle layer
(Fig. 4). Evidence for this mechanism is convincing. For example, McEvily and Bond [14] have shown that the fracture surfaces of specimens
FIG. 4—Schematic representation of the tarnish-rupture mechanism, (a)-(f),
and resulting fracture surface (g).
stress corroded in tarnishing solutions exhibit striations perpendicular to
the direction of crack propagation. This establishes the discontinuous
nature of the fracture process and is thus fully consistent with the tarnish-rupture model (Fig. 4g). Further, this mechanism predicts that
cracking should also be produced by repeated cycles of immersion of an
unstressed specimen in a tarnishing environment followed by stressing
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STRESS CORROSION TESTING
standard stress corrosion tests, can in fact be produced by this procedure.
This observation would appear to invalidate the claim that stress corrosion cracking requires the simultaneous action of stress and corrosive attack [16].
The cupric complex ions are considered to play a controlling role in
the growth of the tarnish [7]. This layer, which consists largely of cuprous oxide [6,13], has been reported to exhibit a parabolic rate of
growth [14]. Accepting the general view that the oxide is cation deficient [17], then it would appear that the rate of tarnish growth is controlled by the diffusion of cuprous ions across the oxide film, which is in
turn dependent on the concentration of cation vacancies. The presence of
zinc in the solid has been found to be a prerequisite for tarnishing [7].
It is probable that zinc exists in the oxide lattice as the divalent ion, occupying cation sites. To maintain electrical neutrality, each divalent zinc
ion requires the presence in the lattice of a cation vacancy, so that it
might be argued that the role of zinc in promoting tarnishing is simply
to introduce these vacancies. However, if this were the case, then tarnishing in brass would be expected to occur in many oxygenated aqueous
solutions, whereas, in fact, tarnishing has been reported only in ammoniacal solutions containing a sufficient concentration of cupric complex
ions (Fig. 3). Accordingly, it has been proposed that the role of the complex ions is to react preferentially with zinc ions at the oxide-solution
interface, and that the preferential removal of zinc in effect results in the
injection of vacancies into the oxide [7]. The presence of these vacancies
is considered to be primarily responsible for the high rate of cation diffusion necessary for rapid tarnish growth. Further work is in progress
to determine the details of this model, but it is interesting to note that it
is supported by recent studies using an electron microprobe analyzer,
which established that the tarnish is severely depleted with respect to
zinc [18].
The conclusion that two mechanisms of failure are operative in the
a-brass/ammonia system, both dependent on the concentration of the
cupric complex ions, raises the question of which is responsible for commonly observed service failures. While it is possible that both may occur
under different conditions, it appears probable that failures in moist industrial atmospheres, that is, season cracking, proceed by the tarnishrupture mechanism.4 Under these conditions, shallow layers of adsorbed
water formed on the surfaces readily pick up oxygen and ammonia from
the atmosphere. The small volume of solution leads to the formation of
large concentrations of cupric complex ions which, in the presence of a
tensile stress, cause rapid cracking (Fig. 2). It is probable that the complex-ion concentrations produced in these circumstances would be suffi* Other copper-base
alloys,
not become
tarnished,
would EST
be expected
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PUGH AND WESTWOOD ON CRITICAL SPECIES
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cient to cause tarnishing. This view is supported by the observation that
when specimens are immersed in fresh 15 N aqueous ammonia, removed, and then stressed in air while the surfaces are still wet (that is,
corresponding to tests in small volume), they rapidly become tarnished,
and cracking is observed [5].
If it is accepted that season cracking occurs by the tarnish-rupture
mechanism, then it is apparent that tests for determining the susceptibility of brasses to this failure should employ conditions which lead to
tarnishing. The long standing practice of simulating practical conditions
FIG. 5—Effect of applied stress on time-to-failure of poly crystalline AgCl tested
in aqueous NaCl solutions at room temperature [21].
by carrying out tests in air containing controlled partial pressures of water vapor and ammonia [11,19] probably satisfies these conditions provided that the adsorbed aqueous film does not become excessively thick.
However, such tests can be carried out more simply and reproducibly by
totally immersing the stressed specimens in tarnishing ammoniacal solutions. The studies by the authors suggest that oxygenated 15 N aqueous
ammonia preconcentrated with >3 g/liter copper is suitable (note the
reproducibility of data in the tarnishing range in Fig. 3d). Alternatively,
the tarninshing solutions described by Mattsson [20], containing 0.05
g-atom/liter copper, 1 g-mole/liter ammonia, and of pH about 7, may
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STRESS CORROSION TESTING
FIG. 6—Polycrystalline AgCl deformed in 6 N aqueous NaCl presaturated
with AgCh3' ions, demonstrating initiation and subsequent growth of intercrystalline cracks. Transmitted light [23].
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PUGH AND WESTWOOD ON CRITICAL SPECIES
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Silver Chloride/Aqueous Chloride Systems
The nonmetal silver chloride has been observed to undergo stress corrosion cracking in certain aqueous solutions [21-26]. For example, Fig.
5 illustrates the relationship between time-to-failure, tF , and engineering
stress for polycrystalline specimens tested under constant tensile load
either in air or in aqueous sodium chloride solutions of various concentrations [27]. Fracture in air and in solutions of concentration <3 N was
ductile and transcrystalline, while brittle intercrystalline failure occurred
in solutions of concentration >3 N. It can be seen from Fig. 5 that, as
in the stress corrosion cracking of many metals [75] (compare Fig. 9), tF
increased with decreasing stress and that a limiting stress existed below
which cracking did not occur in finite times. The limiting stress was dependent on the chloride-ion concentration of the solution.
The solubility of silver chloride in aqueous chloride environments is
dependent on the concentration of chloride ions in solution. In water at
room temperature, the solubility is about 10~5 N, whereas in 6 N sodium
chloride solution it is about 10~2 N', the increase in solubility with increasing chloride-ion concentration is due to the formation of highly
soluble complex ions, such as AgQ2~, AgCl32~, and AgQ43~ [27].
The importance of the most highly charged of these to the stress corrosion process can again be demonstrated by preconcentration experiments. For example, Fig. 5 illustrates the marked reduction in tF at a
given stress caused by presaturating 6 N sodium chloride solutions with
AgCl43~ ions. Conversely, increasing the volume of fresh 6 N sodium
chloride solutions results in a significant increase in tv at a given stress
(Fig. 5) [27].
Metallographic studies of specimens stressed in 6 TV sodium chloride
solution presaturated with AgCl43~ ions have indicated that cracking is
initiated where slip bands are arrested at grain boundaries of large misorientation [23]. For example, in Fig. 6a, intercrystalline cracks (arrowed), which appear dark in transmitted light, have been initiated at
each of the blocked slip bands A, B, and C. Further cracks are formed
at other blocked bands during subsequent stressing (Figs. 6b to e).
Cracking is not observed when the stress field associated with the arrested slip bands is relieved by deformation in the neighboring grain, for
example, Fig. 6a at S. Once initiated, an intercrystalline crack propagates
in a relatively brittle fashion provided that the grain boundary containing
it is approximately perpendicular to the tensile axis (note in Fig. 6 that
crack propagation occurs more rapidly at L to K, where the grain boundary is approximately perpendicular to the tensile axis, T.A.) and that
the embrittling solution is present at the crack tip.
Tests on monocrystals have established that unnotched specimens are
essentially
immune to cracking in sodium chloride solutions [27,22],
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STRESS CORROSION TESTING
Thus it may be concluded that embrittlement is not a consequence of
some inherent property of the grain boundary, but that the boundaries
play an important role by acting as barriers to glide dislocations, hence
causing stress concentrations and facilitating crack initiation.
Embrittlement of polycrystalline silver chloride has also been observed in other complex-forming solutions [21,25,26], The complex ions
formed in these solutions can be positively or negatively charged, but the
magnitude of the charge must be > 1 + or >2—; solutions in which complex ions of lower charge are produced do not cause embrittlement (Table 1). The degree of embrittlement has been found to increase with concentration of the critical species in the environment and with the charge
on the complex ion (Table 1) and to be a function of the distribution of
TABLE 1 —Relationship between charge on complex ions and their ability to embrittle polycrystalline silver chloride [21,25,26].
Solution
10 N NH4OH
<0.1 N AgNO3
17 N AgNO3
0.1 AT NaCl
1 N NaCl
11 NCsCl
5.8 N KC1
6 N NaCl
11.8 NHCl
11.8 Af HC1, presaturated with
CuCl2
18.8 ATLiCl
8 ATNaBr
20 AT LiBr
17 ATNaSCN
6 N NazSjOa
Complex Ion
AgCNIW
AgsCl*
Ag,Cls+
AgCl2AgCla2(AgCl43-)/nCs»+
(AgCl4*-)mK»+
AgCI4*AgCtfCuCls3AgCl4*AgBr4»AgBr4*~
Ag(SCN)4»AgC&O.,),*-
Charge
Embrittlement
1+
1+
2+
12<3<333-
no
no
yes
no
no
no
no
yes
yes
3
33335-
yes
yes
yes
yes
yes
yes
the charge on the complex ion [25]. It is of interest to determine whether
a "common" ion need be present in an adsorbed complex species to
cause embrittlement of a given ionic-covalent solid. Work in progress
suggests that this is not necessary, for adsorbed AgGU3" complexes have
been shown to cause a marked reduction in the surface microhardness
and, therefore, in the strength of magnesium oxide monocrystals [27].
The practical implication of this observation is that complex species produced by the dissolution of one structural component in a given environment might be responsible for the subsequent embrittlement of another
material exposed to the same environment.
It has been proposed that the embrittlement of silver chloride is associated with the adsorption of highly charged complex ions [21,25]. In
earlier papers, the authors suggested that the presence of the adsorbed
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PUGH AND WESTWOOD ON CRITICAL SPECIES
239
complex ions in the vicinity of strained surface bonds, for example, at
sites where glide dislocations are arrested, or at crack tips, might perturb
the distribution of bonding electrons, reducing bond strength and hence
allowing rupture at reduced stress levels [21,25]. However, current work
FIG. 7—Relationship between time-to-failure and temperature, T K, for polycrystalline AgCl tested at various stresses in aqueous solutions containing (a) positive and (b) negative complex ions [28].
favors an alternative mechanism involving the formation of adsorptioninduced charge double layers [25]. The predominant charge carriers hi
silver chloride are Frenkel defects, that is, interstitial silver ions and
silver-ion vacancies [29]. Thus the adsorption of negatively charged complex ions might be expected to induce a compensating positively charged
surface layer in the solid, containing an excess concentration of niter-
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STRESS CORROSION TESTING
stitial silver ions, and a more diffuse, negatively charged layer, containing a greater than equilibrium concentration of silver-ion vacancies, extending into the bulk. Adsorption of positively charged complex ions
would be expected to produce a double layer of opposite sense.
FIG. 8—Inhibition of stress-corrosion cracking in poly crystalline AgCl in 6 N
aqueous chloride environments by replacement of Na* ions by K+ or Cs+ ions.
Applied stress 600 g/mm2 [25].
The presence of large concentrations of point defects in the surface
layers might be expected to cause significant surface hardening, possibly
leading to the formation of a brittle surface layer. Then, by analogy with
the tarnish-rupture mechanism (Fig. 4), it might be imagined that cracking would proceed by the formation and rupture of these defect-hardened
layers. Such a process might be expected to lead to striated fracture surfaces. Fractographic studies, using the optical microscope, have established that the fracture surfaces of embrittled monocrystals do in fact
exhibit striations which are perpendicular to the direction of crack propagation [23]. Similar markings have been observed in intercrystalline
fracture surfaces, but in this instance they were confined to certain regions [28]. Failure to detect striations in all regions of the fracture sur-
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PUGH AND WESTWOOD ON CRITICAL SPECIES
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faces may be due simply to the limited depth of focus and limited resolution of the optical microscope. Electron-microscope (replica) studies are
being undertaken to clarify this point.
Recent investigations of the temperature dependence of the fracture
process provide strong support for this "film-rupture" model. According
to this mechanism, the rate of cracking should be controlled by the rate
of formation of the embrittled double layer ahead of the crack and hence
by the rates of diffusion of the point defects. Studies of the effect of temperature on tF at a given stress have established that the relationship between log tF and l/T (note that tF is inversely proportional to the rate
of cracking) is linear for specimens tested in solutions containing either
negative or positive complex ions (Fig. 7). Calculations of activation energies from these data yield values of 0.08 and 0.39 ev, respectively, in
reasonable agreement with reported values of the activation energies for
the motion of interstitial silver ions, 0.05 to 0.15 ev, and silver-ion vacancies, 0.27 to 0.37 ev [29-37]. In addition, evidence for the postulated
increase in surface hardness has been obtained from indentation studies.
The identification of highly charged complex ions such as AgCl43~ as
the critical species in the embrittlement of silver chloride allows for the
possibility of inhibiting this phenomenon. Clearly, from Table 1, any
addition to the environment which causes an effective reduction in the
charge of embrittling complexes is likely to inhibit cracking. Thus the
successive replacement of sodium ions in 6 N chloride solutions by either
potassium or cesium ions has been found to be an effective means of preventing embrittlement [25] (Fig. 8). Such inhibition is believed to result
from the formation of mixed complexes of lower charge (Table 1). Additions of Group III B cations Zn2+, Cd2+, or Hg2+ to 6 N sodium chloride solutions also inhibit embrittlement. In this case, inhibition was attributed to competition of these ions for chloride ions, again resulting
in the formation of complexes of charge <3 — [25], for example.
Such observations suggest the possibility of "built-in" inhibitors for
systems such as these. Specific alloying elements could be incorporated
in the solid which, following release to the environment during dissolution, would associate with or cause breakdown of the potentially embrittling species, thus preventing embrittlement.
Other Systems
In the systems discussed in the preceding sections, the critical species
responsible for cracking are produced by dissolution of the material in
the environment. Thus, in each case, preconcentration of the environment
with critical species before the start of the test causes significant reductions
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STRESS CORROSION TESTING
in stress corrosion life. To investigate the generality of this effect, preconcentration experiments have been extended to several commercial alloys
tested in aqueous chloride containing environments. The solutions were
presaturated by stirring with filings taken from the alloy under test:
AZ31 Magnesium (3Al-lZri)—Tension specimens were tested over
a range of engineering stresses in solutions containing 35 g/liter sodium
chloride and 30 g/liter sodium chromate. Times-to-failure for fresh and
presaturated solutions were not significantly different (Fig. 9).
2024 Aluminum (4.5Cu-1.5Mg-0.6Mri)—Age-hardened tension spec-
FIG. 9—Effect of applied stress on time-to-failure of AZ31 magnesium alloy
stressed in fresh and presaturated NaCl-NazCrO*.
imens were tested at an engineering stress of 17 kg/mm2 in solutions
containing 53 g/liter sodium chloride, 50 g/liter sodium chromate, and
hydrochloric acid (pH about 0.2). Specimens tested in fresh solutions
failed after an average of about 160 sec compared to about 2100 sec
for those tested in presaturated solutions.
304 Stainless Steel (18Cr-8Ni)5—Specimens were tested at a stress of
42 kg/mm2 in boiling 42 per cent magnesium chloride. Failure occurred
after an average of 1200 sec in both fresh and presaturated solutions.
The fact that presaturation of the test environments did not produce
detectable reductions in tp for these systems indicates either that a sufficient concentration of the critical species is produced rapidly, or that
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PUGH AND WESTWOOD ON CRITICAL SPECIES
243
the critical species is not generated by specimen dissolution but exists in
fresh solutions. In the latter event, the chloride ion itself would appear
to be the most probable species. While the mechanism or mechanisms
of stress corrosion cracking are not understood in these systems, it is
nevertheless interesting to speculate on the possible roles of this ion.
Chloride environments are known to cause pitting corrosion in each
of these alloy systems [52]. Moreover, recent studies suggest that pitting
in aluminum results from the incorporation of the chloride ion in the protective oxide film, reducing the resistance of this film to the diffusion of
ions [33]. By analogy with the case of the a-brass/ammonia system discussed above, it might be proposed that stress corrosion cracking in these
systems proceeds by a similar oxide-rupture mechanism (Fig. 4) and that
FIG. 10—Schematic representation of the mechanism of transcrystalline stress
corrosion cracking proposed by Pickering and Swann [47] and Swann and Embury
[48].
the role of the chloride ion is to modify the defect structure of the surface oxide, permitting its growth to brittle dimensions. Recent fractographic evidence might be considered to support this view, since stress
corrosion fracture surfaces for each of these alloy systems exhibit striations similar to those observed in the case of a-brass and silver chloride
[34-37]. It is evident, however, that further study is necessary to confirm
that the oxide-rupture mechanism is operative in these systems.
The observation that chloride ions adsorb strongly on stainless steel
[35] suggests an alternative mechanism. It has been proposed that stress
corrosion cracking in metals may occur by an adsorption-dependent
mechanism [39-41], similar to that proposed for liquid-metal embrittlement of metals [42,43], but at the present time there appears to be no
unambiguous evidence to support this view. On the other hand, the presence of strongly adsorbing species, acting as step poisons, may be releCopyright
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244
STRESS CORROSION TESTING
dissolution behavior [44-46]. To illustrate, in the mechanism for transcrystalline stress corrosion cracking proposed by Pickering and Swann
[47] and Swann and Embury [48], it is postulated that cracking occurs
by the ductile rupture of a slot weakened by many corrosion tunnels
(Fig. 10). Such tunnels have been observed in Type 301 stainless steel
exposed to 42 per cent magnesium chloride at 140 C [47]; magnesium-7
per cent aluminum exposed to sodium chloride-potassium chromate solutions [47], aluminum exposed to aqueous sodium chloride [49], and
also in copper-25 per cent gold exposed to 10 per cent ferric chloride
solution [47].
Little attention has been given to the role of the chemistry of the environment in the formation of the observed corrosion tunnels. However,
similar tunnels have been observed in lithium fluoride following immersion in a slightly corrosive aqueous environment containing strongly adsorbing step poisons, for example, fatty acid molecules [46] or ferric
fluoride complexes [45]. Tunnelling did not occur in the absence of such
poisons. Westwood and Rubin [46] have suggested that tunnels form
and grow because the dissolution process is less efficiently inhibited at
the bottom of a pit or growing tunnel than at the external surface. This
is a consequence of differences in the concentration of the step poison at
the two sites.
Westwood [50] has suggested that tunnel corrosion in metals may also
be associated with the presence of strongly adsorbing step poisons. The
adsorbing species may be simply the chloride ion itself, or, alternatively,
a metal-chloride complex ion. The latter could be produced by dissolution of the material. Since concentrations of only 10~6 N of a sufficiently
active step poison can markedly affect the dissolution behavior of a solid
in a slightly corrosive environment, then the formation of the necessary
concentration of metal-halide complexes would occur rapidly, thus rationalizing the absence of a detectable difference in tv in the tests carried
out in fresh and presaturated solutions. For the same reason, the metal
ion in the complex need not necessarily be one of the major alloying
constituents of the alloy. It could be provided, for example, by dissolving
inclusions. In this instance, tunnelling or pitting would be expected to be
first observed in the vicinity of such impurity particles, as is often the
case [48]. It would also follow that for a material cracking by a tunnelling mechanism, minor alloying additions could significantly influence
stress corrosion life. Such an effect has been observed for a high-purity
stainless steel containing 20 per cent nickel and 20 per cent chromium
[51]. In this instance, the addition of 1 to 2 per cent molybdenum to the
steel decreased the life in boiling 42 per cent magnesium chloride solution at a given stress from 56 to 9 hr. It should be possible, in terms of
this approach, to reduce the degree of susceptibility to cracking by
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PUGH AND WESTWOOD ON CRITICAL SPECIES
245
breakdown of the "embrittling" complex species in solution (the "builtin" inhibitor approach). Thus it is interesting to note that additions of 2
per cent copper to the high-purity stainless steel increased the life to 139
hr [57].6
Conclusions
It may be concluded from the above discussion that, in certain systems, the presence of critical ion species in the testing environment is
primarily responsible for stress corrosion failure. In these systems, the
critical species were produced by dissolution of the specimen during testing. From the standpoint of stress corrosion testing, realization of this
point is most important, for in such systems variable test results will be
obtained if the specimens are exposed to the environment for different
times before stressing, or if testing solutions are reused. The volume of
the testing solution used also will be a factor (Fig. 2). On the other hand,
if the critical species can be identified, and the testing environments preconcentrated with this species, then extremely reproducible test data can
be expected, as was illustrated in Fig. 3a.
At the present time, however, knowledge of the chemical species existing in solution in most of the stress corrosion systems of technological
interest is extremely limited. It is suggested that increased attention to
this aspect of stress corrosion could be of immediate practical value to
both stress corrosion testing and the more important problem of preventing stress corrosion cracking.
A cknowledgment
The authors are pleased to acknowledge financial support received
from the U. S. Army Research Office (Durham) and the Office of Naval
Research.
References
[1] Pugh, E. N. and Westwood, A. R., "Complex Ions and Stress-Corrosion
Cracking in a-Brass," Philosophical Magazine, Vol. 13, 1966, pp. 167-183.
[2] Althof, F. C., "Inter- and Intracrystalline Corrosion and Its Causes," Zeitschrift fur Metallkunde, Vol. 36, 1944, pp. 177-186.
[3] Graf, L. and Richter, W., 'The Problem of Stress Corrosion of Homogeneous
Solid Solutions," Zeitschrift fur Metallkunde, Vol. 52, 1961, p. 833.
[4] Pugh, E. N. and Westwood, A. R., "Stress-Corrosion Cracking in Brass,"
High-Strength Materials, Wiley, New York, 1965, pp. 701-704.
[5] Pugh, E. N., Montague, W. G., and Westwood, A. R., "On the Role of Complex Ions in the Season-Cracking of Alpha-Brass," Transactions, American
Society for Metals, Vol. 58, 1965, pp. 665-671.
[6] Hoar, T. P. and Booker, C. J., "The Electrochemistry of the Stress-Corrosion
Cracking of Alpha-Brass," Corrosion Science, Vol. 5, 1965, pp. 821-840.
8
Note that
in terms
oxide-rupture
mechanism,
effects such
as these may
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be ascribed to changes in the defect structure of the oxide.
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246
STRESS CORROSION TESTING
[7] Cotton, F. A. and Wilkinson, G., Advanced Inorganic Chemistry, Interscience, New York, 1962, p. 756.
[8] Johnson, H. E. and Leja, J., "On the Potential/pH Diagrams of the Cu-NHrHsO and Zn-NHy-H2O Systems," Journal of the Electrochemical Society,
Vol. 112, 1965, pp. 63 8-641.
[9] Jorgensen, C. K., Absorption Spectra and Chemical Bonding in Complexes,
Pergamon, London, 1962, p. 286.
[10] Edmunds, G., "Season Cracking of Brass," ASTM-AIME Symposium on
Stress-Corrosion Cracking of Metals, 1944, American Society for Testing
and Materials, Philadelphia, 1945, pp. 67-89.
[11] Thompson, D. H. and Tracy, A. W., "Influence of Composition on the StressCorrosion Cracking of Some Copper-Base Alloys," Transactions, American
Institute of Mining, Metallurgical and Petroleum Engineers, Vol. 185, 1949,
pp. 100-109.
[12] Pugh, E. N., Montague, W. G., and Westwood, A. R., "Stress-Corrosion
Cracking of Copper," Corrosion Science, Vol. 6, 1966, pp. 345 and 346.
[13] Forty, A. J. and Humble, P., 'The Influence of Surface Tarnish on the StressCorrosion of a-Brass," Philosophical Magazine, Vol. 8, 1963, pp. 247-264.
[14] McEvily, A. J., Jr., and Bond, A. P., "On the Initiation and Growth of StressCorrosion Cracks in Tarnished Brass," Journal of the Electrochemical Society,
Vol. 112, 1965, pp. 131-138.
[15] Pugh, E. N., Montague, W. G., and Westwood, A. R., "On the Mechanism(s)
of Stress-Corrosion Cracking," Environment-Sensitive Mechanical Behavior,
Gordon and Breach, New York, 1966, pp. 351-402.
[16] Sutton, H. et al, Journal of the Institute of Metals, Vol. 71, 1945, p. xvii.
[17] Kubaschewski, O. and Hopkins, B. E., Oxidation of Metals and Alloys, Butterworths, London, 1953, p. 30.
[18] Forty, A. J. and Humble, P., "Surface Films and Stress-Corrosion Cracking,"
Environment-Sensitive Mechanical Behavior, Gordon and Breach, New York,
1966, pp. 403-420.
[79] Jamieson, A. L. and Rosenthal, H., "Aqua Ammonia Test," ASTM-AIME
Symposium on Stress-Corrosion Cracking of Metals, 1944, American Society
for Testing and Materials, Philadelphia, 1945, pp. 36-46.
[20] Mattsson, E., "Stress Corrosion in Brass Considered Against the Background
of Potential/pH Diagrams," Electrochimica Acta, Vol. 3, 1961, pp. 279-291.
[21] Westwood, A. R., Goldheim, D. L., and Pugh, E. N., "Embrittlement of
Polycrystalline Silver Chloride," Discussions of the Faraday Society, No. 38,
1964, pp. 147-156.
[22] Levine, E., Solomon, H., and Cadoff, I., "Fracture Characteristics of Polycrystalline AgCl Wet with Aqueous Solutions," Acta Metallurgica, Vol. 12,
1964, pp. 1119-1124.
[25] Westwood, A. R., Goldheim, D. L., and Pugh, E. N., "Fracture Behavior of
AgCl in Aqueous NaCl," Acta Metallurgica, Vol. 13, 1965, pp. 695-700.
[24] Levine, E. and Cadoff, I., "The Embrittlement of Single Crystals and Bicrystals of AgCl in Aqueous Salt Solutions," Acta Metallurgica, Vol. 13, 1965,
pp. 875-880.
[25] Westwood, A. R., Goldheim, D. L., and Pugh, E. N., "Complex-Ion Embrittlement of Silver Chloride," Materials Science Research, Vol. 3, Plenum,
New York, 1966, pp. 553-576.
[26] Westwood, A. R. and Goldheim, D. L., "Embrittlement of Silver Chloride by
Copper Chlorocomplexes," Proceedings, First International Conference on
Fracture, Sendai, Japan, Vol. 2, 1965, pp. 1999-2014.
[27] Westwood, A. R., Goldheim, D. L., and Lye, R. G., "Rebinder Effects in
MgO," Philosophical Magazine, 1967, in press.
[28] Westwood, A. R., Goldheim, D. L., and Pugh, E. N., "A Double-Layer Mechanism for the Complex-Ion Embrittlement of Silver Chloride," Philosophical
Magazine,Vol. 15, 1967, pp. 105-120.
[29]
Abbink,
C. and
D. S.,reserved);
"Ionic Conductivity
Silver Chloride
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PUGH AND WESTWOOD ON CRITICAL SPECIES
[30]
[31]
[32]
[33]
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[50]
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Vol. 27, 1966, pp. 205-215.
Ebert, I. and Teltow, J., "Zur lonenleitung und Fehlordnung von Silberchlorid mit Zusatzen," Annalen der Physik, Vol. 15, 1955, pp. 268-278.
Abey, A. E. and Tomizuka, C. T., 'The Effect of Hydrostatic Pressure on
the Ionic Conductivity of Silver Chloride," Journal of Physics and Chemistry
of Solids, 1966, Vol. 27, pp. 1149-1159.
Metals Handbook, 8th ed., American Society for Metals, Metals Park, Ohio,
Vol. 1, 1961.
Heine, M. A., Keir, D. S., and Pryor, M. J., 'The Specific Effects of Chloride
and Sulphate Ions on Oxide Covered Aluminum," Journal of the Electrochemical Society, Vol. 112, 1965, pp. 24-32.
Fairman, L. and West, J. M., "Stress-Corrosion Cracking of a Magnesium
Aluminum Alloy," Corrosion Science, Vol. 5, 1965, pp. 711-716.
McEvily, A. J., Jr., and Bond, A. P., "On Film Rupture and Stress-Corrosion
Cracking," Environment-Sensitive Mechanical Behavior, Gordon and Breach,
New York, 1966, pp. 421^53.
Nielsen, N. A., 'The Role of Corrosion Products in Crack Propagation in
Austenitic Stainless Steel. Electron Microscopic Studies," Physical Metallurgy
of Stress-Corrosion Fracture, Interscience, New York, 1959, pp. 121-143.
Logan, H. L., McBee, M. J., and Kahan, D. J., "Evidence for an ElectroChemical-Mechanical Stress Corrosion Fracture in a Stainless Steel," Corrosion Science, Vol. 5, 1965, pp. 729 and 730.
Staicopolus, D. N., "Electrocapillary Studies on Solid Metals," Journal of the
Electrochemical Society, Vol. 108, 1961, pp. 900-904.
Uhlig, H. H., "New Perspectives in the Stress-Corrosion Problem," Physical
Metallurgy of Stress-Corrosion Fracture, Interscience, New York, 1959, pp.
1-17.
Coleman, E. G., Weinstein, D., and Rostoker, W., "On a Surface Energy
Mechanism for Stress-Corrosion Cracking," Acta Metallurgica, Vol. 9, 1961,
pp. 491-496.
Nichols, H. and Rostoker, W., "Analogies Between Stress-Corrosion Cracking and Embrittlement by Liquid Metals," Transactions, American Society
for Metals, Vol. 56, 1963, pp. 494-507.
Westwood, A. R. and Kamdar, M. H., "Concerning Liquid Metal Embrittlement, Particularly of Zinc Monocrystals by Mercury," Philosophical Magazine, Vol. 8, 1963. pp. 787-804.
Stoloff, N. S. and Johnston, T. L., "Crack Propagation in a Liquid Metal Enivronment," Acta Metallurgica, Vol. 11, 1963, pp. 251-256.
Sears, G. W., Growth and Perfection of Crystals, Wiley, New York, 1958, p.
441.
Sears, G. W., "Dislocation Etchings," Journal of Chemical Physics, Vol. 32,
1960, pp. 1317-1322.
Westwood, A. R. and Rubin, H., "Etch-Tunnels in Lithium Fluoride," Journal of Applied Physics, Vol. 33, 1962, pp. 2001-2007.
Pickering, H. W. and Swann, P. R., "Electron Metallography of Chemical
Attack Upon Some Alloys Susceptible to Stress-Corrosion Cracking," Corrosion, Vol. 19, 1963, pp. 373-3891.
Swann, P. R. and Embury, J. D., "Microstructural Aspects of Stress-Corrosion Failure," High-Strength Materials, Wiley, New York, 1965, pp. 327-359.
Edeleanu, C., 'The Propagation of Corrosion Pits in Metals," Journal of the
Institute of Metals, Vol. 89, 1960, pp. 90-94.
Westwood, A. R., "Concerning the Possible Role of Adsorbed Step-Poisons
in Stress-Corrosion Cracking," Corrosion Science, Vol. 6, 1966, pp. 381-384.
Saxena, M. N. and Dodd, R. A., 'Transgranular Stress-Corrosion Cracking
Mechanisms in High Purity Austenitic Stainless Steels," Environment-Sensitive Mechanical Behavior, Gordon and Breach, New York, 1966, pp. 455—479.
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R. W. Staehle1
Circulating Autoclave System for Stress
Corrosion Cracking Studies
REFERENCE: R. W. Staehle, "Circulating Autoclave System for Stress
Corrosion Cracking Studies," Stress Corrosion Testing, ASTM STP 425,
Am. Soc. Testing Mats., 1967, p. 248.
ABSTRACT: A circulating autoclave system has been constructed and
operated in which time-to-breaking can be determined exactly at temperatures to 700 F, pressures to 5000 psi, a wide range of stresses, various
concentrations of dissolved gases and ionic species, and for a large number of specimens exposed simultaneously. Preliminary experiments have
studied commercial and specially melted iron-nickel-chromium alloys. Results have confirmed already established trends for the effect of environment and alloy composition at 500 F.
KEY WORDS: corrosion, stress corrosion, cracking (fracturing), autoclaves, steels, stainless steels, Inconel, iron-chromium-nickel alloys
There is a need to obtain reliable information on time-to-breaking
due to stress corrosion cracking as a function of alloy composition, environmental composition, temperature, and stress. This information is
necessary for both direct use in engineering applications and to furnish
boundary conditions for the development of theories of stress corrosion
cracking. In the past, it has been difficult to obtain quantitative information on time-to-breaking at temperatures above the solution boiling temperature or for any range of dissolved gas concentrations. Control of
applied stress at the elevated temperatures and at pressures above atmospheric has also been difficult. Information available at the higher
temperatures has generally been obtained by periodic examination of
specimens at the end of specified test periods.
Described herein is a circulating autoclave system capable of measuring time-to-breaking for up to 200 specimens at temperatures to 700 F,
pressures to 5000 psi, dissolved oxygen concentrations to 1000 ppm, and
stresses to any desired level. The system is capable of operating in the
once-through or recirculating modes.
At the outset it was expected that the oxygen and chloride concentra1
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ing,
The Ohio State University, Columbus, Ohio. Personal member ASTM.
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
249
tions required for cracking would be relatively high, that .is, several
hundred parts per million. For these ranges of concentration, it was adequate to operate the system in the recirculating mode. Thus, a slight contamination due to recirculating the solution would exert little influence
compared to the relatively high chloride concentration. However, it was
found that the cracking was occurring in very short times at concentrations of chloride and oxygen less than 10 ppm. Therefore, it was decided
that it would be necessary to change to a once-through mode of circulation in which the solution would be furnished from a distilling unit and
dumped to drain after passage through the test region. The experimental
results described herein are from preliminary work performed while the
system was operating in the recirculating mode. As of the writing of this
paper, the necessary equipment for the once-through system had been
installed and preliminary shakedown runs begun.
Significant components, procedures, and preliminary experimental results will be described. Emphasis is placed on features which are unique
to measuring time-to-breaking and to controlling solution composition.
Description of Circulating System and Components2
Design Objectives
Circulation of the medium is required to prevent stagnation of the
solution, replenish depleted reactants, remove tramp ions resulting from
corrosion of the system, provide continuous monitoring of significant
solution properties, and provide a means for rapidly heating the solution
to test temperature.
Since dissolved gases, especially oxygen, appear to exert a strong
influence on the stress corrosion cracking phenomenon, it is necessary
to introduce dissolved gases in quantities desired and according to whatever schedule is required to meet experimental objectives. The maximum
quantity of gas introduced will depend on the solubility at temperature,
requirements of the experiment, and maximum pressure capability of
the system.
The system was designed to operate up to 5000 psi with no component operating with a maximum internal applied stress of 10,000 psi.
Materials exposed to the corrosive medium were selected to provide
maximum resistance to corrosion and subsequent malfunction or contamination. Inconel 600 was selected as the main material of construction. A maximum temperature of 700 F was based on the temperature
of liquid water being 705 F at the critical point. To obtain interpretable
data, it was necessary to keep the corrosive medium single phase; twophase (boiling or gas bubble) environments were considered to produce
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2
All high pressurebycomponents for this system were fabricated by Pressure ProdDownloaded/printed
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FIG. 1—Schematic arrangement of circulating autoclave systems used for stress corrosion cracking studies.
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FIG. 2—Schematic arrangement of electrical power for circulating autoclave system.
252
STRESS CORROSION TESTING
more uncertainties than desirable. These requirements provided the incentive for the capability of operating at pressures above the equilibrium
vapor pressure. (The vapor pressure of steam in equilibrium with water
at 705 F is 3200 psi.)
To obtain data of maximum usefulness, it was necessary to provide
a procedure for indicating time-to-breaking of any specimen under
FIG. 3—External view of test autoclave. Total height approximately 72 in.
any condition of temperature and pressure without shutting down the
circulating system.
The components and system arrangements described in subsequent
sections were designed specifically to meet the above objectives.
Circuating System
A schematic arrangement for the circulating system is shown in Fig. 1,
and a schematic of the electrical power arrangement is shown in Fig. 2.
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The
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on a once-through
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
253
it is introduced at valve 2. The water coming to valve 2 is produced continuously in a 100 gal/hr still and prior to entering at valve 2 is mixed
with the required chloride or other specie additions using positive displacement pumps. Dissolved gases, if required, are introduced at the main
pump. The pump raises the liquid pressure from ambient to the desired
operating pressure, usually about 4000 psi. The solution moves to the preheater in which it is heated to test temperature. Passing from the pre-
FIG. A—Cross section of test vessel used for stress corrosion cracking tests.
heater, the solution enters the test vessel from which it passes through a
cooler and back pressure regulator. The cooler is necessary to prevent
damage to the back pressure regulator valve and to prevent a large
volume of hot steam from entering the drain system. The back pressure
regulator together with the positive displacement pump acts to maintain
solution pressure.
Main Te,sf Fe-s.se/
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The main
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and 164. 15:53:43
The primary
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254
STRESS CORROSION TESTING
mens at the same time to identical environmental conditions, thereby enhancing the self-consistency of results. A shell-liner construction was
used primarily to keep the stresses low on the inside surfaces of the liner.
By shrink fitting the liner into the shell, the calculated tangentical stress
on the inside surface is less than 1000 (negative) psi when the autoclave
operates at 700 F and 5000 psi internal pressure. The liner was centrifugally cast Inconel 600, because this method of fabrication permitted some reduction in cost, and the shell was 4340 steel. A double-
FIG. 5—Top of autoclave head showing electrical lead fittings in place. Plugs
are placed in openings not used for electrical leads.
end construction of the vessel was used to permit cleaning and for
possible future use of the bottom head for bringing in electrical leads.
The top head contains 15 openings through which electrical or other
leads can be brought into the autoclave. Figure 5 shows the top head
with both seals and electrical lead fittings. Figure 6 shows a schematic
cross section of an electrical lead-through fitting. Each of these fittings
permits up to 16 electrical leads to be brought through the pressure
boundary to be used for determining time-to-breaking of wire specimens.
These fittings are capable of withstanding up to 5000 psi without losing
their pressure seal. The Teflon gasket is a standard Conax design, and
the use of cooling fins maintains the temperature sufficiently low so that
the Teflon does not flow. Housings were of Inconel 600, the insulators of
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alumina,
and the electrical leads of Teflon-coated Inconel 600. These
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
255
FIG. 6—Schematic cross section of fitting for bringing leads through autoclave
head.
insulated electrical lead wires have been satisfactory to 600 F and have
shown no evidence of significant deterioration below this temperature.
Preheater
The preheater consists of an Inconel 600 tube (%6 by %
ap6 in.)
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130 ft by
long and is heated by using the tube itself as a reDownloaded/printed
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256
STRESS CORROSION TESTING
FlCr. 7—-Relationship among power, tube length, and voltage for preheater.
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FIG. 8—Assembled preheater showing insulation in place. One side removed.
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
257
sistor. The relationship between power, tube length, and voltage is shown
in Fig. 7. Self-heating was chosen because it was the most economical
and involved only the cost of the transformer. The assembled and insulated preheater is shown in Fig. 8. Figure 2 shows the electrical arrangement for providing power to the preheater. The length of the preheater tubing was chosen to keep the thermal stresses to a minimum,
with the maximum thermal stress at full power being less than 3000 psi
in tension at the inside surface.
Pump and Gas Injection
A duplex positive displacement pump is used to circulate the solution
and ha^a maximum flow capacity of 40 gal/hr against a pressure of
5000/psi. One end of the pump is shown in Fig. 9.
FIG. 9—One end of duplex hydraulic pump used to circulate solution. Pump
fluid system driven by 5-hp motor.
Any dissolved gases desired are introduced at the pump. This procedure eliminates uncertainties which would result from direct injection
of gas under pressure. Figure 10 shows the schematic arrangement for
introducing the dissolved gas. The gas is trapped between the solenoid
valves (8 and 9 on Fig. 1 and shown on Fig. 10) at a pressure higher
than system pressure and in a known volume. The entrapped gas is released on each stroke of the positive displacement pump. Thus, on each
half cycle of the pump, solution containing dissolved gas of a measured
and required concentration is injected into the system. The pump cam
and switches for activating the solenoid valves are shown in Fig. 11.
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FIG. 10—Schematic arrangement for injecting gas and for the hydraulic system.
STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
259
ture at a rate of 40 gal/hr. Tap water is used as a coolant. The tubing
in the cooler, Inconel 600, is about 120 ft long and is spirally wound.
An exterior view of the cooler is shown in Fig. 12.
Specimen Stressing and Break Indication
Important objectives in devising techniques for stressing the specimens
are: There should be a positive indication when the specimen breaks;
the applied load should be adjustable and controllable; the specimen
should not be subject to selective chemical effects due to edges, corners,
or crevices; a maximum number of specimens should be exposed in a
single experiment; and the final result of time-to-breaking should be
indicative of the early stages of the cracking process.
FIG. 11—Cam on pump used for activating microswitches which control oxygen
additives.
The optimum result from considering the above objectives was to
utilize a compression spring and 0.015-in.-diameter wire specimen.
Essential features of the design are shown in Fig. 13. The wire specimens
are first attached to the upper clamp, and a weight is attached to the
other end. The magnitude of the weight is determined from the load
required on the wire and the temperature affected change in the shear
modulus. Since the spring material is Inconel X heat treated for 10 hr
at 1350 F, there is essentially no relaxation of the spring material up
to the 700 F maximum operating temperature.3 The adjustment for
effect of temperature on shear modulus is included in the initial load.4
When the prescribed load is applied, the lower clamp is tightened. The
deflection of the spring due to the weight is maintained when the
3
"Superalloy
Wire,"
Sept.Wed
1963,
Standard
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260
STRESS CORROSION TESTING
clamp is tightened; thus, the spring exerts the same tensile force as the
weight.
Important consideration in designing the springs are: minimizing
spring volume to maximize the number of specimens; minimizing stress
in the springs to reduce relaxation and chemical attack; maximizing
FIG. 12—Cooler used to cool solution from test temperature to ambient.
spring deflection rate to reduce errors in stress level; restrict length of
springs to keep below the critical buckling length; and minimize solid
height of the spring to improve circulation around the specimen. The
foregoing requirements conflict and compromises are required. Since
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there
is no obvious way to derive an analytical expression for optimizing
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
261
FIG. 13—Design of jig used for stressing specimens.
spring size, it was arbitrarily decided that the autoclave should contain
approximately 100 specimens. Using the standard equations for spring
deflection and spring surface stresses as plotted in Fig. 14, together with
Wahl's expression5 for the buckling criterion, the spring parameters of
wire diameter (d), number of coils («), length (L), and average diameter
(D) were selected.
Figure 15 shows a close-up of the actual springs and electrical lead
B
Wahl, A.
Mechanical
Springs,
McGraw-Hill,
York, 1963;
Cheronis,
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by and Application, McGraw-Hill, New York, 1961.
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262
STRESS CORROSION TESTING
wires, and Fig. 16 shows an experiment ready to be placed in the autoclave. The middle zone contains no specimens in Fig. 16. All materials
in the specimen support and testing jigs are Inconel 600. In Fig. 16 it
should be noted that the autoclave head rotates, but the specimen rack
does not.
Recording Time-to-Cracking
Times-to-cracking are recorded in an Esterline Angus event recorder
as shown in Fig. 17. When a specimen breaks (Fig. 18), a circuit is
closed and a pen in the event recorder moves to indicate that the circuit
FIG. 14—Spring design curves relating deflection, load, stress in spring, spring
wire diameter, number of coils, and stress in test wire for springs of 1.25 in. mean
diameter.
is closed. The event recorder has modules of 20 pens each, but the unit
used contains two modules per cabinet so that 40 pens are available. In
the present system, each pen can monitor several specimens simultaneously if a cyclic timing system is utilized. Thus, for one second out of
a minute, circuit a would check all specimens on the a circuit, and during a subsequent interval the b circuit is checked. The individual pens
may be used for double or triple duty. The circuit for multiple pen use
is shown in Fig. 19. One of the difficulties associated with the use of the
"no-contact and subsequent contact" system is the conductivity of the
environment.
If theInt'l
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is too conducting,
there isEST
little
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
263
ination between the nonbroken and broken situations. This difficulty is
circumvented by reducing the available area of the contacting electrodes
or by reducing the conductivity of the solution. The latter approach is
generally the most applicable, since critical studies of cracking are
generally conducted at low ionic concentrations.
FIG. 15—Close-up of assembled load of stressing jigs ready to be inserted into
autoclave.
Protective Circuits
The nature of the circulating system is such that certain accidents are
possible such as loss of pressure or overheating. Since it is obviously
desirable to protect both the experiment and system, a safety circuit
system (Fig. 20) is utilized. This system will shut off any of the major
components on signals from either the pressure monitor or the multipoint temperature recorder. The multipoint temperature recorder contains six alarms which can be connected to any of the 24 available temperature positions.
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byEST
a barricade
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264
STRESS CORROSION TESTING
FIG. 16—Specimen stressing jigs attached to autoclave head and ready to be inserted into autoclave.
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
265
FIG. 17—Dual event recorder for recording time-to-cracking.
FIG. 18—Schematic showing configuration of stressing jigs with broken and unbroken wire specimens.
(Fig. 21). This barricade has two thicknesses of %-in. plywood and one
thickness of ^-in. boiler plate. All valves are mounted on the inside
with their valve stems extending through the wall. A monorail host is
used to transport the experiments over the safety barricade into the
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autoclave.
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266
STRESS CORROSION TESTING
Miscellaneous Components
All valves have Inconel 600 bodies. Valves for higher temperature
portions of the system have their packings extended about 5 in. from
the solution. The temperature drop to the packing is about 300 F at an
operating temperature of 500 F.
The system is protected by gold-covered stainless steel ruptive disks.
Gold covering of the ruptive disks is essential to prevent their premature
failure.
FIG. 19—Wiring arrangement for multiple use of event pens.
Joints are of the mechanical cone type using a collar and gland nut.
These joints frequently lose their seal when there is appreciable temperature cycling. This requires the use of welded joints in critical applications such as the preheater.
Source of Distilled Water
The new distilled water source is a Barnstead Model SS-100 still having a capacity of 100 gal/hr. Two positive displacement pumps are
used to mix the distilled water and the chloride. The water is pumped
with a Clark Cooper c-p 2 positive displacement pump, and the chloride
with
a Beckman solution metering pump No. 746.
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
267
FIG. 20—Schematic wiring diagram for protective circuits.
Experimental Results and Discussion
Experiments performed to date have been conducted using the system
in a recirculating mode. This initial work has been directed primarily
toward determining the general behavior of the circulating system and
the time-to-breaking circuitry. Such factors as oxygen concentration,
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268
STRESS CORROSION TESTING
alloy composition, initial stress, metallography, and specimen mounting
procedures have been investigated.
For the recirculating experiments double distilled water was used and
sodium chloride (NaCl) additions were made as required. Specimens
were stressed as described previously.
All specimens in these experiments were wires of 0.015 in. diameter.
Results from the first series of experiments are tabulated in Table 1.
Important features of the data in Table 1 are as follows:
1. With increasing nickel content, the time-to-breaking increases.
FIG. 21—Safety barricade for circulating autoclave system showing instrumentation.
2. As the concentration of reactants in the environment decreases,
there is a significant increase in time-to-breaking.
3. At high concentration of environmental reactants, the effect of
stress is relatively small, but as the concentration decreases, there is an
increasing difference between the high- and low-stress levels.
Table 2 summarizes results from a preliminary experiment which
was conducted to determine the general significance of alloy composition. The experiment involved 93 hr and 30 min of operation at 50 ppm
Nad and 8 ppm oxygen, after which the oxygen was changed to 158
ppm and the experiment continued. The data in Table 2 have been
arranged
that
portions
of 16
the15:53:43
experiment
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treated
individually.
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Type 304
Type 309
Type 310
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Alloy
6
TABLE I—Time-to-breaking for three stainless steels at 500 F, min."
1000-158d
1000-158
21 000
36 800
21 000
23, 27, 28
22,27
26, 25, 27
24,25
25,26
25, 25, 25
36 800
21,25
24,24
91,20, 35
17, 17
35
NB (365)
NB (365)
27,23
21 000
42,43
36 800
41,30
57, 35
NB (235)
33,30
NB (365)
3 specimens
26, 147
c
Stress, psi
1000-8
10-8
50-8
50-158"
3581, 376, 634
186, 20
NB (4350)
1870
903
-6
4501
NB (4350)
3 specimens
NB (4350)
2 specimens
30, 90, 685
43, 42, 58
NB (5610)
2 specimens
58, 53
NB (5610)
2 specimens
74, 59
" Time-to-cracking is given to nearest minute. The designation NB indicates that the specimen did not break for the duration of the
experiment with the number in parentheses giving the length of the experiment. The flow for all experiments was 33.5 gal/hr, and the
operating pressure was 3250 psi. Time-to-cracking is measured from a time 15 min after starting oxygen additions. The oxygen additions
were started when the temperature of the preheater outlet was 330 F. Specimens broke both in the grips and between grips. Times-tobreaking were found to be generally the same for breaking in either location. Metallographic examination showed no differences in the
morphology of the cracks.
6
Specimens in the as-received condition.
c
The stresses are corrected for change in shear modulus. No check was made later to determine the relaxation of the individual specimens which in many cases were broken and thus did not permit such a measurement.
d
The first number indicates amount of chloride in parts per million; the second the amount of oxygen. In cases where oxygen was listed
as 8 ppm, it was assumed that the water was saturated with air prior to the experiment. The 8 ppm value is the approximate amount present in water in contact with air at 25 C. In experiments showing 158 ppm oxygen, the oxygen was injected as a measured amount on each
stroke of the positive displacement pump.
» Specimens previously exposed to 50 ppm NaCl and 8 ppm oxygen before oxygen was raised to 158 ppm. Because of the nature of the
results, it was considered legitimate to divide the experiment into two parts.
270
STRESS CORROSION TESTING
TABLE 2—Results from stress corrosion studies of iron-chromium-nickel base
alloys at 500 F.a
Breaking Times with 50 ppm
NaCl and 8 ppm Oi
Alloy*
304
30
310
54
54C
54Mo
54N
54P
54Pd
54P
54Re
54Si
54W
58
Total Individual Time,
hrrmin
0:30, 1:30, 11:25
Breaking Times with 50 ppm NaCI
and 158 ppm O:
Average
Total Individual Time,
hr:min
Total
Average
Average
after Os
Increase
4:30
94:11, 94:12, 94:28
94:28, 94:23
94:44, 94:29
94:36, 94:12, 94:44
94:25,94:21,94:18
94:18
94:03
93:57
93:53, 93:50
94:17
94:26
94:37
94:31
94:21
94:18
94:33
93:57
93:52
0:47
0:56
1:07
1:01
0:51
0:48
0:33
0:27
0:22
94:33
94:31
94:34
97:46
1:03
1:01
1:04
4:16
0:58, 0:58
0:43
81:40,76:20
0:58
0:43
79:00
2:50, 57:49
2:23
30:05
2:23
94:18,94:47
94:31
94:34
96:39, 96:21
101:58,96:06
0
All specimens loaded to 21,000 psi stress. Notes from Table 1 apply.
Alloy 54 = Fe-15Ni-20Cr. Fourth component additions are made at 1.5 atomic
per cent for metallic elements and 0.1 atomic per cent for nonmetallic elements.
Alloy 58 = Fe-45Ni-20Cr.
6
TABLE 3—Time-to-breaking (tf)° as affected by surface preparation for Type 310
stainless steel at 500 F and 90 per cent of yield stress (0.2 per cent offset}.
No.
Specimens
Tested
No.
Failed
As-received6
Vacuum annealede
Electropolished**
11
11
0
12
2
1
Mechanically polished*
12
6
Condition
Times- to-B reaking,
hr:min
0:29, 0:47
0:28
0:40,0:44
0:48, 0:49
18:00, 0:42
Avg,
hr:min
0:38
0:28
0:44'
0
NaCl content: 0.0050 weight per cent, O2 content: 0.0010 weight per cent
(approximately). All specimens broke at ends and inside Teflon sleeves. Cracking
at this location has been determined previously to occur at about same time as
outside sleeve. Also, failure occurs by cracking and not general oxidation.
6
Specimen held approximately 10 sec at 1100 C in strand anneal.
« Vacuum annealed approximately 2 hr at 1090 C.
d
Vacuum annealed as in (c) and electropolished.
• Vacuum annealed as in (c) and mechanically polished with 600-grit paper.
(Note that questions of importance of local mechanical work, roughness; and bulk
deformation are being checked in subsequent experiments.)
f
Did not include 18:00 hr breaking time in average.
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FIG. 22—Effect of surface preparations on time-to-breaking of Type 310 stainless steel.
272
STRESS CORROSION TESTING
most pronounced effect of composition is the effect of nickel. The
best comparison is between alloys 54 and 58 (15 versus 45 per cent
nickel, both with 20 per cent chromium). The breaking time for the
alloys is 61 versus 256 min after the increase in oxygen concentration.
There are other obvious trends with various fourth component alloy additions, but the significance of trends based on these observations is not
presently apparent.
Table 3 summarizes a preliminary experiment to compare the effect
of surface preparation on time-to-breaking. Clearly, there is a significant
FIG. 23—Effect of chloride and oxygen concentration on time-to-breaking of
Type 304 stainless steel.
difference among the preparations from the standpoint of total number
of failures. Again the significance of this experiment is not presently
obvious except to note that surface preparation is indeed important.
These data should be compared with the results from similar experiments6 in boiling magnesium chloride (MgCl2) shown in Fig. 22. These
results show the same general trends, that is, that the mechanically
abraded surfaces tend to break earlier and tend to exhibit a tighter spread
of cracking time.
Results from experiments described above and summarized in Tables
1 to 3 are very similar to results from many previous experiments by
a
Cochran,byR.,
unpublished
results,
Corrosion
Department
of MetallurCopyright
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gical
Engineering, Theby
Ohio State University, Columbus, Ohio.
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STAEHLE ON CIRCULATING AUTOCLAVE SYSTEM
273
other investigators. For example, the beneficial effect of nickel, the relatively rapid cracking of Type 304 stainless steel, and the effects of chloride and oxygen concentration have been described many times. However, the unique aspect of the work described herein is the capability
for measuring times-to-breaking exactly for any combination of significant environmental parameters. Subsequent experiments will study in detail the effect of alloy, temperature, stress, and environmental chemistry
on time-to-breaking.
Figure 23 summarizes time-to-breaking information for Type 304
stainless steel. Points plotted are averages of data from Table 1. The
strong effects of chloride and oxygen concentrations are evident.
Conclusions
1. A circulating autoclave system has been developed in which exact
time-to-breaking can be determined for specimens stressed at designated
stresses, various dissolved gas concentrations, various ionic species, temperatures to 700 F, and pressures to 5000 psi.
2. Preliminary stress corrosion cracking experiments show that the
system performs as intended. These experiments show for the ironchromium-nickel alloys tested that time-to-breaking is increased by reducing the stress, reducing chloride and oxygen, and increasing nickel
concentration.
3. From an engineering point of view, note that cracking of Type
304 stainless steel will occur in about 6 hr at the annealed yield stress
for chloride contents of 10 ppm and oxygen concentrations of 8 ppm.
For this same chemistry, cracking can occur in 20 min at 36,800 psi.
A cknowledgments
This work was sponsored by the U.S. Atomic Energy Commission,
Nuclear Technology Branch, under the supervision of J. M. Simmons
and A. E. Van Echo. Experimental work described herein was performed by P. J. Simmons and J. Frey. It is a pleasure to acknowledge
the very considerable contributions to the design and construction of
the high pressure equipment by Pressure Products Industries, Hatboro,
Pa. Robert Wolf and Anthony Ostrowski of that organization were especially helpful.
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Nancy McKinney1 and H. W. Hermance1
Stress Corrosion Cracking Rates of a
Nickel-Brass Alloy Under Applied
Potential
REFERENCE: Nancy McKinney and H. W. Hermance, "Stress Corrosion Cracking Rates of a Nickel-Brass Alloy Under Applied Potential,"
Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967,
p. 274.
ABSTRACT: A method was developed to test the susceptibility of nickelbrass alloys to stress corrosion cracking in the presence of nitrates or other
salts. A wire of the alloy under test was stressed, and an electrical leakage
path was provided to a cathode by a filter paper impregnated with the electrolyte being studied. Stresses, potentials, salt concentrations, relative humidities, and time were varied to evaluate these factors. In the case of
ASTM Grade D nickel-brass alloy, using ammonium nitrate as the electrolyte, the rate of crack penetration increased with increases in the salt
concentration in the leakage path, the applied stress, relative humidity,
and temperature. There was no direct correlation with the applied potential. Other hygroscopic nitrates such as zinc, and to a lesser degree calcium
and copper, also caused stress corrosion cracking. Chlorides and sulfates
were ineffective under conditions in which nitrates produced cracking.
From the data obtained, the expected failure times of nickel-brass parts at
average ambient temperature and humidity could be estimated.
KEY WORDS: copper alloys, nickel alloys, zinc alloys, corrosion, stress
corrosion, electric potential, nitrates, humidity, sulfates, chlorides, salt
water
Failures of stressed nickel-brass parts of telephone equipment have
been observed in central offices in the Los Angeles area. The parts that
failed, wires of 0.023 in. diameter, were normally under about 6-g load
and had a positive potential. An inspection showed that failure was due
to a form of stress corrosion cracking. These failures often occurred
within as short a time as two years. Large amounts of particulate matter
had accumulated on these metal surfaces near areas that had cracked.
This particulate material was an accumulation of air-borne dusts. It
seemed possible that corrosive materials in this dust during periods of
high humidity could have caused the cracking.
Chemist and chemist (retired), respectively, Chemical Dept, Bell Telephone
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Laboratories,
Inc., Holmdel, N. J.
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McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
275
The Los Angeles dusts are hydrophilic, containing large amounts of
adobe dusts, which are fine claylike materials, as well as containing oxidized organic compounds and inorganic ions. About half of the solids
obtained from an evaporation of a water extract of Los Angeles dust
deliquesced at 55 per cent relative humidity. Table 1 shows that these
air-borne particulates are composed of sulfates and nitrates; about 8.2
per cent by weight is sulfate; 8 per cent by weight is nitrate. Generally
TABLE 1—Analysis of Los Angeles air-borne dust specimen (weight per cent).0
Material
4-g Specimen
Benzene soluble
Water soluble
Sulfate
Nitrate
Sodium
Ammonium
Chloride
Calcium
Potassium
Magnesium
Insoluble-combustible
Insoluble-noncombustible
a
18
28
8.2
8.0
3.0
2.6
1.9
1.5
1.0
0.7
46
8
Specimen collected by high volume sampler from 18 February to 19 July, 1960.
TABLE 2—Relative humidity over saturated solutions.
Salt
Calcium chloride
Calcium nitrate
Ammonium nitrate
Sodium nitrate
Sodium chloride
Ammonium chloride
Ammonium sulfate
Sodium sulfate
Calcium sulfate
RH, %
Temperature
Reference, pp.
28.99
50.7
61.8
73.93
75.93
78.4
80.0
81
98
25 C
77.7 F
25 C
25 C
25 C
71 F
25 C
25 C
20 C
521-528°
519-520
507-517
521-528
521-528
519-520
507-517
507-517
675
a
Wexler, A., ed., "Humidity and Moisture Measurement and Control in Science
and Industry,". Vol. 3, Reinhold, New York, 1965.
6
International Critical Tables, 1st ed., Vol. 1, McGraw-Hill, New York, 1926.
the nitrate salts are more hygroscopic than the sulfate salts as is shown
in Table 2.
Preliminary qualitative experiments were made to determine if either
nitrate or sulfate salts could cause cracking of nickel brass. It was
found that the nitrate salts such as ammonium and calcium caused rapid
cracking at 75 per cent relative humidity but that ammonium sulfate did
not.
A laboratory
was Wed
thenDec
begun
to study
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effect
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and
other
anions
as
well
as
to
study
the effect of apDownloaded/printed by
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276
STRESS CORROSION TESTING
plied potential, relative humidity, temperature, and stress on the rate of
stress corrosion cracking of nickel-brass wire. In addition, it was hoped
that a method for testing the susceptibility of various alloys to this type
of stress corrosion cracking could be developed.
Experimental Details
The unit shown in Fig. 1 was used. It permitted simultaneous testing
of six wires so that average effects could be obtained. The frame, 2, was
FIG. 1—Test unit.
made of Plexiglas and contained holes through which the test wires, 1,
were inserted. Clamping screws, 3, were used to hold the individual
wires in place. Two metal plates of gold-plated brass that were fastened
to the frame were used as electrodes. The plate, 4, served as the cathode
and the plate, 6, as the anode. The openings in the cathode plate were
symmetrical about the test wire. The anode contained small holes
through which the wires were inserted, so that the wires became anodic
during the testing. The vertical position of the anode could be adjusted
to give the desired constant deformation on the test wires. Filter paper
disks (0.005 in. thick, 0.06 in.2 area), 5, containing a desired amount
ofCopyright
electrolyte,
wereInt'l
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about
each Wed
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McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
277
punching out the papers to a larger diameter than that of the cathode
rim and using a central hole in the paper disk of a smaller diameter than
the wire.
The alloy studied was ASTM Grade D 65-12-23 copper-nickel-zinc
alloy, spring tempered, 0.025 in. in diameter (Specification for CopperNickel-Zinc Alloy (Nickel Silver) Wire, B 206 - 56). The wire was
straightened, cut to 4 to 5-in. specimens, and degreased before testing.
The initial loads on the wire were set with the anode plate at either
0, 5, 10, or 15 g corresponding to cantilever stresses of 0, 19,000, 39,000,
or 56,000 psi.2 All these test loads are under the proportional limit for
this wire which is about 91,200 psi. Unless otherwise stated, the testing
was done at 15-g load.
Ammonium nitrate was generally used as the electrolyte, since it was
the most active of the electrolytes we had tested in producing stress
corrosion cracking. Moreover, these ions are present in large amounts in
the Los Angeles dusts. The amounts of nitrate used per disk ranged from
360 to 0.4 p,g as nitrate ion. When no potential was applied, at least 60/*g nitrate per disk were used. To prepare these disks, a calculated amount
of nitrate was incorporated into Whatman No. 1 filter paper by dipping
papers into a nitrate solution and air drying them. The exact amount of
nitrate per square inch of paper was then determined. Disks were punched
out of these sheets and stored at low humidity until used.
The assembled units were tested either at 43 or 75 per cent relative
humidity. These humidities were chosen to be above and below 62 per
cent relative humidity, which is the equilibrium humidity for a saturated
solution of ammonium nitrate (Table 2). Two saturated salts were used to
obtain these test humidities. A saturated solution of sodium chloride gives
75 per cent relative humidity at room temperature;3 a saturated solution of
potassium carbonate gives 43 per cent relative humidity at 24.5 C.4 As
many as eight units at one time could be placed in a relative humidity
chamber containing the appropriate saturated solution for simultaneous
testing. Before starting the tests, the air in the chamber was circulated by
a blower to ensure a uniform relative humidity in the chamber.
Except when otherwise indicated, a potential of 45 v connected
through a protective 1-megohm resistor to the unit's terminals was used.
Since the measured current was always much less than 45 ju,a, it was
obviously limited by the resistance of the filter paper disks, which was
2
Stresses were calculated using the following equation: S = 3EdD/2l2, where
S = stress, psi; E = Young's modulus, psi; d — diameter of wire, in.; D — deflection
due to load, in.; and / = length of wire, in.
3
Wexler, A. and Hasegawa, S., "Relative Humidity-Temperature Relationships
of Some Saturated Salt Solutions in the Temperature Range 0 C to 50 C," Journal
of Research of the National Bureau of Standards, Vol. 53, No. 1, July 1954, pp.
19-26.
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278
STRESS CORROSION TESTING
FIG. 2—Photomicrograph of a diametrical section of a test wire. This wire was
exposed to 29-ng nitrate, 75 per cent relative humidity, 15-g load, for 5 days. The
photo shows the transcrystalline nature of the crack.
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McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
279
governed by the relative humidity of the test chamber and the nitrate
concentration.
Discussion
Crack Development
Within a few hours after starting a test run, a small band of dark
corrosion products appeared on the surfaces of the wires directly under
FIG. 3—Initial corrosion (points 1 and 2).
the filter paper about the wire's maximum stressed area. As the test
continued, the corrosion began to penetrate into the wire only within the
area of the wire which was under tensile stress and formed small transcrystalline cracks, which may have followed lattice imperfections in the
metal. A photomicrograph of a section of a test wire is shown in Fig. 2.
These cracks are not as yet deep enough to cause wire yawning. With
the passage of time, the corrosion continued to deepen and widen. When
it reached about three fourths of the way through the wire, visible surface
cracks appeared.
To determine just how far corrosion did penetrate into each of the test
wires,
they had to be broken open through their corroded area by bendCopyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
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280
STRESS CORROSION TESTING
ing them in the same direction as the applied stress. Sometimes the corroded areas in the wire appeared only slightly discolored but usually
were dark brown-black. Figure 3 shows the beginning of visible wire
penetration. This wire, after bending, contained two large pits on its
surface with discolored edges. It also had other very small pits, above
point 1 and a little to the left, which cannot be seen in this figure. Figure
FIG. 4—Cross sections of test wires showing progressive stages of penetration.
4 shows increasingly deep penetrations. Estimates of damage to the interior of these wires are listed on each illustration. These estimates were
based on the per cent of total cross-sectional area that appeared to be
covered by the corrosion product. The accuracy of this estimate was about
±5 per cent.
Assuming that a typical corrosion penetration had progressed into
the test wire from three main corrosion points or deep pits, Fig. 5 illustrates probable successive stages of corrosion through its cross section
Copyright
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15:53:43 EST 2015
based
on observations
variousWed
stages
The three
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McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
281
hypothetical corroded areas eventually join into a larger more continuous
area of corrosion which then continues to deepen until wire yawning
occurs. This type figure has been used by Hines5 to illustrate the depths
of corrosion reached in austenitic chromium-nickel steel wires exposed
to boiling magnesium chloride under varying stresses.
Penetration Rates at Fixed NO3-,
RelativeHumidity
Humidity
and
and
Load
Load
CondiCon
3~, Relative
tions
The rate of corrosion penetration into wires under various conditions,
such as nitrate concentration, relative humidity, and load, was studied.
This was done by placing up to eight test units under equal loads and
nitrate concentrations in the same relative humidity chamber. Each unit
FIG. 5—Progressive stages of penetration starting from three areas.
was removed after a different time interval, and the depth of corrosion
penetration into the test wires was determined. The average depth as
well as the maximum and minimum depths of penetration reached for
each of these time intervals was recorded. From these data, plots of
average penetration versus time and maximum penetration versus time
were made.
Average and Maximum Rates—Figure 6 shows the rate curves based
on average values obtained at high humidity. The range from the highest
to the lowest penetration obtained with the six test wires is indicated by
the straight lines through the average points. The maximum penetration
rates are also shown here by the dotted lines. The rate of penetration, as
is clearly shown in this figure, is directly dependent upon the amount of
nitrate that was in the paper disks.
6
Hines, J.by
G.ASTM
and Hugill,
R. W.,
Physical
Metallurgy
of Stress-Corrosion
FracCopyright
Int'l (all
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ture, Rhodin, T. N., ed., Interscience, New York, 1959, pp. 193-223.
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282
STRESS CORROSION TESTING
FIG. 6—Effect of nitrate concentration on penetration rates at 75 per cent relative humidity.
FIG. 7—Effect of nitrate concentration on penetration rates at 43 per cent relative humidity.
These rate curves are parabolic. Since the 5, 11, and 29-^.g rate curves
do not appreciably slow down, it could be assumed that these concentrations could probably cause penetrations into greater depths of metal
than is afforded by the size that we tested. The fastest penetration rate in
terms of area obtained in this work was about 0.4 /*2/sec calculated from
the maximum rate curves shown for 29-ju.g nitrate. No penetration occurred after a month in wires exposed to paper disks without nitrate.
Figure 7 shows the average and maximum rate curves obtained at 43
per cent relative humidity. The rates are much slower at this humidity
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but
are still dependent upon nitrate concentration. It was surprising that
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occur (University
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nitrate No
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McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
283
not appreciably hygroscopic until the relative humidity is 62 per cent
(Table 2). A possible explanation might be that the small amount of
moisture adsorbed by this salt plus the amount adsorbed by the filter
paper was sufficient to permit a small amount of electrolysis and migration of ions to occur. Some of the products formed might have been an
extremely deliquescent salt such as zinc nitrate. (A saturated solution of
this salt at room temperature gives an equilibrium humidity of 42 per
cent relative humidity.5) These could then provide the necessary moisture
to continue the corrosion penetration reactions.
Figure 8 shows the average rate curves for wires tested at different
loads. All the wires were exposed to the same nitrate concentration at
either the high or low humidity. No visible penetration occurred in wires
FIG. 8 —Effect of load on penetration rate at high and low humidity, with 29-m
nitrate.
tested under no load after 20 days. The rates, as the figure illustrates,
are dependent upon the load on the wire. The wires with 10- and 15-g
loads were penetrated to about the same depths; the wires under 5-g
loads were penetrated at a much slower rate. Other nitrate concentrations were also tested at 5- and 10-g loads at both humidities. A similar
effect of load on the penetration rate was noticed.
An empirical rate equation was derived based on the parabolic curves
of Figs. 6 to 8. It is expressed as
dx/dt = ci/(2jc + c2)
where t = time in days, and x — per cent cross-sectional penetration. The
constant c^ appears to depend upon the concentration of nitrate in the paper. The constant c2 is apparently dependent upon relative humidity and
load. The rate equation therefore indicates that the rate of penetration is
directly dependent upon the concentration of nitrate in the paper disk and
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inversely
dependent upon the area already covered by corrosion.
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Iso-Penetration
Levels—
To anticipate
the fastestpursuant
rate at which
penetraUniversity of Washington
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284
STRESS CORROSION TESTING
tion moves through the wires under the fixed test conditions, various points
were selected from the maximum rate curves (Figs. 6 and 7) and used to
plot iso-penetration curves.
Figure 9 shows such a plot for penetration levels of 5, 25, 60, and 80
per cent at high humidity and indicates the minimum time that the wires
FIG. 9—Minimum time to reach given depths of penetration at 75 per cent relative humidity.
FIG. 10—Minimum time to reach given depths of penetration at 43 per cent
relative humidity.
took to reach a given depth of penetration. The lines are almost parallel
and apparently level off. The plots indicate that for various penetration
depths to be reached within 100 days of laboratory testing certain definite
amounts of nitrate would be required. For example, it took at least 2-ju.g
nitrate to cause a penetration of 80 per cent and 0.4-/xg nitrate to cause a
penetration
of 25 per cent in this time. Since 80 per cent penetration is
Copyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
deep
enough for surface
cracking to occur, amounts of nitrate in the paper
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285
McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
disks as small as 2 /xg could cause permanent wire damage at high humidity. By extrapolation of points on the figure, the minimum amounts of nitrate that would be needed to cause penetrations such as 5 and 60 per cent
can be estimated.
Figure 10 shows a plot for penetration levels of 5, 25, and 60 per cent at
the low humidity. The 60 per cent line was obtained by drawing a line
through the available point (29 /xg, 101 days) parallel to the others. A
rough estimate of the shortest time that a wire exposed to 5-/ig nitrate
would take at low humidity to reach 60 per cent penetration using this line
would be about 5l/z years. Contrasting 5M> years with the time required
by 5-^g nitrate to reach 60 per cent penetration at high humidity which is
about 16 days, the importance of reducing humidity as a means of control
is readily apparent.
Penetration Rates at Other Fixed Test Conditions
Temperature—The rate of penetration is dependent upon temperature.
Wires that were tested at temperatures near 0 C showed less penetration
than wires tested under the same conditions of stress and humidity at room
temperature. Wires tested at higher temperatures showed greater penetration than wires tested under the same conditions at room temperature. The
results are listed in the following table:
Concentration NOs~ , Mg
29.
29
Test Period,
days
3
5
Average Penetration at 75% RH, %
3C
Room
Temperature
50 C
24
37
98
6
Applied Potential (relationship to leakage current, nitrate consumption)
—Wires were tested in the units at high humidity without applying potential. Penetration did occur but only when disks contained very large concentrations of ammonium nitrate. This is shown in the following table:
Concentration NOa~ , jig
360.
150.
60.
Test Period, days
Average Penetration
(two or more
wires), %
4
14
40
48
30
12
No potential apparently is needed if the concentrations of nitrate in the
disks are very large. But to obtain penetrations with the smaller nitrate
concentrations
within
a rights
reasonable
anEST
anodic
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286
STRESS CORROSION TESTING
Tests were also made to determine if varying the applied voltage would
vary the rate of penetration. The wires were exposed to the same nitrate
concentration, 29 /ng, at high humidity. Table 3 shows that, after runs of 9
days, about the same average depth of penetration was obtained with
voltages between 6 to 135 v.
TABLE 3—Effect of applied potential on depth of corrosion penetration
(after nine days at 75% RH, 29-v-g NOr).
Potential, v
0
6
45
90
135
Avg Penetration, %
none
57
60
57
65
FIG. "11—Electrolysis current through assembly containing 6 nickel-brass wires
at various voltages at 75 per cent relative humidity, 11-fi.g nitrate.
Figure 11 shows the amount of leakage current obtained at various applied voltages, when wires were tested with 1 l-/xg nitrate at high humidity.
The initial current is highest with the highest voltage, but within 2 to 3 hr
the current for all the voltages tested dropped to about the same steady
value of 0.1 /u.a.
Approximately the same number of coulombs passed in each case.
Figure 12 shows the currents obtained with various nitrate concentrations
at a fixed voltage of 45 v. The current, after an initial transient peak is apCopyright dependent
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16 15:53:43
EST 2015
parently
on (all
the rights
amount
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present
in the paper
but not
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(as
was seen in Fig.
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McKINNEY AND HERMANCE ON NICKEL-BRASS AUOY
287
FIG. 12—Electrolysis current through assembly containing six nickel-brass wires
at various nitrate concentrations at 75 per cent relative humidity, 45v.
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FIG. 13—Rate of nitrate removal from filter paper originally containing either
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5, or 2-fig nitrate.
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288
STRESS CORROSION TESTING
Figure 13 shows that the nitrate in the filter paper disks is only slowly
removed with a 45-v potential at high humidity. This indicates that the ion
migration under these conditions is slow. At low humidity, this rate is still
slower. For example, after 19 days of test, no measurable amount of nitrate was removed from the disks containing 29-ju,g nitrate, even though
the test wires showed an average cross-sectional penetration of 11 per cent.
TABLE 4—Effect of electrolytes on rate of corrosion penetration
(75%RH,29ngN03~).
3~).
Salt
Zinc nitrate
Ammonium nitrate
Calcium nitrate
Copper nitrate
Ammonium sulfate0
Ammonium chloride
0
Test Period, days
Avg Penetration, %
5
5
5
5
30
30
39
38
18
5
0
0
This salt, after 50 days at 96% RH, causes 15% average penetration.
FIG. 14—Effect of ammonium nitrate and ammonium sulfate on penetration of
wires at 75 per cent relative humidity.
After 40 days, 1.2-ju.g nitrate were removed from these disks, and the wires
showed a 33 per cent penetration.
Electrolytes (individual, mixed)—Nitrate salts other than ammonium
nitrate, such as zinc, calcium, and copper, also cause a similar type of
stress corrosion cracking. These data are shown in Table 4. Ammonium
sulfate and ammonium chloride did not cause any penetrations after one
month's testing. The test wires were heavily corroded however. At over 96
per cent relative humidity, sulfate salts can cause a similar type corrosion
cracking. This would be similar to nickel-brass cracking in sulfate soluCopyright
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McKINNEY AND HERMANCE ON NICKEL-BRASS ALLOY
289
Wires were tested using a mixed electrolyte of nitrate and sulfate in the
paper disks. The wires were examined after 18 and 42 days. Figure 14
shows the results obtained from this study. The top line shows the rate of
penetration with 5-//,g nitrate alone. As the amount of sulfate is increased,
the rate of penetration is decreased. However, when this test was repeated
with a nitrate concentration of 11 ^.g, the retarding effect of the sulfate ions
was not as great.
Wires were tested with paper disks containing a known amount of an
aqueous extract of the Los Angeles dust specimen (Table 1). The concentration of nitrate in the disks was 10.4 j«,g, the ratio of nitrate to sulfate in
this dust extract is about 1:1. It was found that the Los Angeles dust did
cause a slower penetration rate than 11-ju.g nitrate as ammonium nitrate.
Test Period, da
14
41
£j!S3?
NH4NOs ,
11 ng NOs-
66
100
32
70
Here it would appear that the sulfate and possibly other salts in the dust
extract reduced the rate.
Relationship of Laboratory Rates to Field Breakage Rates
It was found that the rates generally correlated with field failures, if the
differences in stress were taken into account. Normally, where the nitrate
accumulations on equipment were the greatest, the failure rates were
greatest under the same relative humidities. In some instances, where the
relative humidity was lower, breakage was noticeably less, as would be
expected.
Evaluation of Other Alloys
Other alloys were tested with 29-/*g nitrate at high humidity for 5 days.
The results are shown below. The 80-20 alloy did not show corrosion penetration under the conditions of test, while the same type of alloy with active metal additives did exhibit evidence of corrosion.
Alloy Composition
80Cu-20Ni
80Cu-20Ni with Al, Mg additives.
No. Wires Tested
5
23
No. Wires with
Penetration
0
1
Conclusions
1. A test apparatus was developed which was useful hi obtaining the
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of various
on the
rates
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penetration
into 16 15:53:
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290
STRESS CORROSION TESTING
2. Of all the salts tested, only those containing nitrates were capable of
causing cracking or serious corrosion penetrations into the test wires at 75
per cent or lower relative humidity.
3. Large concentrations of nitrate can cause cracking of the test wire
even without an applied potential. With small amounts of nitrate, noticeable corrosion penetrations occurred only when the wire was held at a
positive potential.
4. The rate of corrosion penetration was found to be directly dependent upon nitrate concentration, stress, relative humidity, and temperature.
Increases in any of these variables increased the penetration rate. The
rate of corrosion penetration appears to depend on the sign but not the
magnitude of the applied potential.
5. Large amounts of Los Angeles air-borne dusts under periods of high
humidity are capable of stress corrosion cracking nickel-brass parts. At
low humidity, serious cracking would not be expected to occur since the
penetration rates are extremely slow.
A cknowledgment
The authors would like to thank D. W. Long, for his help in constructing the equipment and his aid in obtaining experimental data; Miss B.
Russiello, for the quantitative determination of the various anions; and
W. A. Lawrence, Jr., for his assistance in the mathematical interpretation
of the data.
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DISCUSSION ON NICKEL-BRASS ALLOY
291
DISCUSSION
W. C. Harding1 (written discussion)—The authors are to be congratulated on an extremely interesting and thorough analysis of a phenomenon
that is puzzling to us also. We have experienced this same stress corrosion
failure of spring temper 12 per cent nickel silver relay springs. However,
our failures occurred in Texas and New Jersey. The authors mention Los
Angeles only. Have any other failures been noted in other locations in the
United States? Have any tests been conducted on plated or coated nickel
silver? Have the authors conducted corrosion tests on 18 per cent nickel
silver to determine whether this composition would be superior in resistance to stress corrosion to the 12 per cent grade?
Nancy McKinney and H. W. Hermance (authors)—Other areas of the
United States have not reported failures of this type. A recent study was
made of telephone central office dust deposits in industrial areas to determine if conditions for causing nitrate cracking existed outside Los Angeles.
No cities in Texas were included in this survey. Dusts containing the largest
amounts of nitrates were found in cities along the eastern seaboard with
New York City, Bayonne, N. J., Baltimore, Md., and Washington D. C.
leading. But these eastern seaboard dusts and dust found in other coal and
oil burning industrial cities contain large amounts of sulfate and are heavy
in tarry, oily, carbonaceous material which have a lesser tendency to deposit as fine wettable material. Unless the dust contamination on nickel
silver contains a very large amount of nitrate, enough to overcome the inhibiting effects of its sulfate content and nonwettable nature, and unless
the humidity conditions have been met, cracking would not be expected to
occur. For example, in New York City after about 15 years, some stress
cracking has occurred in nickel silver component parts which were within
4 ft of the windows and were not under an applied potential. The dust and
corrosion on these surfaces contained 360-yu.g nitrate/in2, 2280-/j,g sulfate/in2. Nickel silver components away from these same windows did not
become stress cracked. Their dust deposit contained only 9-jug nitrate/in2.
We have not tested plated or coated nickel silver.
An 18 per cent nickel silver wire alloy (55Cu-18Ni-27Zn) was tested by
us under the conditions described in this paper. It appears to be even more
susceptible to cracking than the 12 per cent nickel silver possibly due to its
greater zinc content.
1
Materials and process engineer, Westinghouse Electric Corp., Jersey City, N. J.
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D. O. Sprowls1
Reporting and Evaluating Stress
Corrosion Data*
REFERENCE: D. O. Sprowls, "Reporting and Evaluating Stress Corrosion Data," Stress Corrosion Testing, ASTM STP 425, Am. Soc. Testing
Mats, 1967, p. 292.
ABSTRACT: Susceptibility of a metal to stress corrosion implies a greater
deterioration of its mechanical properties through the simultaneous action of a static stress and exposure to a corrosive environment than
would be produced by the combined effects of each factor operating
separately. In recent years this point of view on stress corrosion and
stress corrosion cracking (the ultimate in stress corrosion) has become
widely recognized. However, confusion still exists, with the result that
some work is published that has little value because of the poor design
of testing procedure. It is still sometimes erroneously assumed that failure of a specimen under stress corrosion test conditions suffices to demonstrate susceptibility to stress corrosion cracking. This conclusion is
unjustified, for the specimen may have become so weakened by corrosion,
even without any acceleration by stress, as to fail under the applied load.
Thus, it is most essential to select techniques (type of specimen and
method of loading, test medium and period of exposure) that produce
failure that is purely the result of stress corrosion cracking.
Because of the marked effect that test procedure can have upon ordinary criteria, such as specimen life, percentage of specimen survival,
threshold stress, etc., investigators should report details of procedure and
technique as well as detailed data along with their analysis of the results.
This is necessary to enable the reviewer to adapt the information to his
need and to make possible a reanalysis of the data by some method that
may be developed in the future.
KEY WORDS: stress corrosion, cracking (fracturing), corrosion, data,
evaluation, analysis
Just to define "resistance to stress corrosion cracking" is a difficult
and controversial task. Although it is relatively easy to stress corrosion
crack a susceptible alloy in a suitable environment, a considerable
* Report of Task Group 3 of Subcommittee VI of ASTM Committee G-l on
Corrosion of Metals. Task Group members are D. O. Sprowls, chairman; E. G.
Haney; J. F. Hildebrand; D. S. Neill; H. R. Pritchard; W. G. Renshaw; H. B.
Romans; T. J. Summerson; and D. H. Thompson.
1
Assistant chief, Chemical Metallurgy Div., Alcoa Research Laboratories, New
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Kensington,
Pa.
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
293
amount of testing is necessary to determine whether or not the degree of
susceptibility constitutes a practical limitation upon the use of the alloy.
The latter task generally involves testing at several stress levels under
conditions of loading and corrosive environment that have an established
relationship to the intended service of the alloy.
As increasing numbers of alloys and tempers are developed, and as
applications become more sophisticated, complex, and demanding, the
need for information on stress corrosion behavior of metals has become
increasingly greater. Comparisons of stress corrosion test data obtained
by different investigators, however, are impossible unless the test procedures and the methods of reporting the results are similar. This situation
has led to the need, not only for standardized test procedures, but also
for accepted methods of analyzing stress corrosion test data.
Scope
The considerations of Task Group 3 have been directed along the following lines:
1. Critical review of the physical and mathematical significance of
various methods of reporting and analyzing data of the types required of
the various test purposes.
2. Sample evaluations of stress corrosion performance using selected
criteria or "representative values."
3. Compilation of a pertinent bibliography.
Definitions
Dix [7]2 referred to stress corrosion cracking as: "Spontaneous failure
by cracking of a metal under the combined action of high stress and corrosion." Most workers in the field of stress corrosion have agreed that
stress corrosion cracking results from tensile stresses at the metal surfaces
acting for prolonged periods of time. Champion [2] has expanded this
definition and commented as follows: "Susceptibility of a metal to stress
corrosion implies a greater deterioration in the mechanical properties of
the material through the simultaneous action of a static stress and exposure to a corrosive environment than would occur by the separate but
additive action of those agencies."
In recent years these views on stress corrosion and stress corrosion
cracking (the ultimate in stress corrosion) have become widely recognized. However, confusion still exists, with the result that some work is
published which has little value because of the poor design of testing
procedure. It is still sometimes erroneously assumed that failure of a
specimen under stress corrosion test conditions suffices to demonstrate
2
The italic numbers in brackets refer to the list of references appended to this
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294
STRESS CORROSION TESTING
susceptibility to stress corrosion cracking. This conclusion is unjustified,
for the specimen may have become so weakened by corrosion, even
without any acceleration by stress, as to fail under the applied load. Thus,
it is most essential to select techniques (type of specimen and method of
loading, test medium and period of exposure) that produce failure that
is purely the result of stress corrosion cracking.
Verification of Stress Corrosion
Stressed Versus Unstressed Specimens
It follows from the definitions presented above that resistance to stress
corrosion of an alloy with unknown susceptibility can only be established
FIG. 1—Distinguishing stress corrosion failure from mechanical failure [3].
by exposing both unstressed and stressed specimens to corrosive conditions. Various investigators have emphasized the importance, especially
with constant load tests, of distinguishing between failures resulting from
stress corrosion cracking and failures resulting from a reduction in load
supporting area by corrosion [3].
The difference between these two situations is illustrated by the diagram in Fig. 1. Three identical tension specimens may be assumed as
examples. Specimen A is the unexposed, unstressed blank; its breaking
load is represented by the height of Block A. Specimen B is exposed to a
corrosive environment in the unstressed condition. Specimen C, exposed
to the same corrosive environment, is stressed by a constant load equal,
for example, to three quarters of the breaking load of specimen A. DurCopyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
ing
the exposure, specimen C breaks under the combined influence of
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
295
stress and corrosion, and at that time specimen B is tension tested. The
closer the breaking load for specimen B is to that of specimen A, the
greater is the evidence of stress corrosion of specimen C and the certainty
that it failed as a result of stress corrosion cracking. On the other hand,
if the breaking load for B is sufficiently close to the load on specimen C
to be within experimental error, then the failure should not be attributed
to stress corrosion cracking, but rather to the probability that the specimen merely had corroded to the point where it could no longer support
the load.
Under test conditions involving prolonged exposures at ambient
temperatures or elevated temperatures that could result in weakening of
the alloy, failure resulting from stress rupture could be encountered; in
this situation, additional unstressed and stressed specimens should be
subjected to the same conditions of temperature and time but without
exposure to the corrosive medium, so that the loss in strength due to
corrosion can be separated from the loss due to the simultaneous metallurgical change in the alloy.
As a general rule, test procedures are designed to serve best the purpose of the investigation. For some purposes, of course, tests are performed on materials of known susceptibility to stress corrosion cracking,
and there is no need to expose accompanying unstressed specimens.
Examples are tests to evaluate protective coatings or specific structural
assemblies and quality control tests. For the purpose of quality control,
a rapid test is preferred, and the standards adopted frequently have no relationship to the serviceability of the alloy.
Stress Corrosion Index
Jones [4] has developed a "percentage loss of strength at failure due to
stress corrosion" by relating the breaking loads in Fig. 1 as follows,
(B — C)/(A — C). For practical purposes this relationship may be expressed in terms of stress, and Booth et al have termed this the "stress
corrosion index" [5].
where:
TSU = tensile strength (based on original cross-sectional area) of
unstressed specimen exposed for the time-to-failure of the
stressed specimen,
TS = tensile strength of unexposed specimen, and
<r
= nominal stress applied to stressed specimen resulting in its
failure.
Higher index values indicate a greater tendency toward stress corrosion
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STRESS CORROSION TESTING
It is evident from an examination of the formula that the corrosion
loss of the unstressed specimen has a great deal of influence upon the
stress corrosion index and, therefore, should be determined carefully.
It is advantageous to test unstressed specimens after several periods of
exposure and plot a TSu versus time curve, from which TSu values can
be taken for individual stressed specimen lives (or suitable representative
specimen lives). It may be readily appreciated also that the shape and
distribution of sites of localized corrosion (stress concentrators) can
have an appreciable effect on the load required to break a stress corroding
or a corroded unstressed specimen. Thus, the stress corrosion index for
an alloy can be expected to vary with testing conditions and is not very
useful for ranking alloys except under specific test conditions. An exTABLE 1—Comparison of stress corrosion indexes for several aluminum alloys
exposed at different stress levels and in different environments.11
3.5% NaCl Alternate Immersion
Industrial Atmosphere
Tensile
Strength
Median
Specimen
Life,
days
Tensile
Strength,
Unstressed,
psi
SCI
Median Tensile
Speci- Strength,
men
UnSCI
Life,
stressed,
days
psi
45
64
5
66
0.84
30
25
20
15
10
10
48
40
32
24
16
16
2
2.5
4.5
12
73
1
59
48
44
40
36
27
0.88
0.61
0.56
0.52
0.49
0.32
Applied Stress
Tensile
Strength,
ksi
ksi
2014-T65
70
2024-T35
63
Alloy
%
623
62
0.60
0 Source: Unpublished data, Alcoa Research Laboratories. Test specimens
were 0.125-in.-diameter by 2-in.-long tensile bars machined transversely from
rolled rod. Stressed specimens were exposed in triplicate.
ample of such variation is shown in Table 1, where the SCI varies markedly for a given alloy depending upon the test environment and upon
the magnitude of the applied stress.
Metallographic Examination
One of the most useful methods of verification of stress-corrosion cracking in a specimen corroded under stress is the microscopic examination
of sections taken parallel to the direction of the applied stress and including the fracture or major crack and an appreciable length of the adjacent
tensile surface. Several important factors can be established: (1) the
path of the crack, that is, intergranular or transgranular, which for many
alloys indicates whether or not stress corrosion cracking has occurred;
(2) identification of questionable surface indications of cracks, directional
corrosion,
etc.;Int'land
thereserved);
prevalent
corrosionESTand
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presence
of stress corrosion
stringers (directional corrosion perpendicular
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
297
to direction of tensile stress, not present in unstressed specimen). It frequently is necessary to examine also a similarly corroded unstressed
specimen to establish the evidence of stress corrosion when cracks are
not detected.
In addition, electron microscopic examination of replicas of fracture
surface can be helpful in establishing evidence of stress corrosion cracking. The consideration of the specific characteristics of stress corrosion
cracks in various alloys and environments is considered beyond the scope
of this report.
FIG. 2—Effect of applied current on cracking time, USS 12MoV stainless
steel in aerated 3 per cent NaCl solution [6].
Anodic and Cathodic Polarization
For alloys that are susceptible both to stress corrosion cracking and to
hydrogen embrittlement, polarization experiments can be used to distinguish between the two mechanisms of cracking. Phelps and Loginow
[6] have described the procedure and discussed the interpretation of the
results in the case of high yield strength steels. In an anodic-cathodic
polarization experiment with USS 12MoV steel, anodic polarization
accelerated failure (Fig. 2). Cathodic polarization at low current density
greatly extended the time-to-failure. On the other hand, cathodic polarization at relatively high current densities again accelerated failure. These
trends
indicate that (1) with no current, (2) with anodic polarization,
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and
(3)
with cathodic
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STRESS CORROSION TESTING
cracking was caused by corrosion along an active path through the steel,
that is, stress corrosion cracking; with cathodic polarization at relatively
high current densities, hydrogen was evolved and hydrogen embrittlement was the probable cause.
Reporting Stress Corrosion Data
Stress corrosion testing technique can have a marked effect upon
ordinary criteria of stress corrosion resistance; hence, it is desirable for
investigators to describe the material evaluated and the test procedure
in detail. Information should be given on: (1) test material, including
composition, manufactured form and dimensions, particulars of heat
treatment or other tempering practice, and method of sampling; (2) type
of test specimen and its orientation relative to grain structure; (3) method
of applying stress (or strain); (4) possible residual stress in the test specimen; and (5) test medium, giving such details as concentration, temperature, period of exposure, etc.
It is desirable also that the raw test data be reported along with the
interpretation of the investigator. The basic data are valuable to readers
wanting to make other comparisons.
Analysis of Stress Corrosion Test Data
An adequate analysis of stress corrosion test data is of very great
importance. The first problem is to determine the most useful measure
of stress corrosion performance. Then, in order to compare one material
with another, representative values must be selected. The most promising
approaches to this problem found in the literature or gained from the
experience of members of this task group are discussed in the sections
that follow.
Specimen Life
Frequently, stress corrosion testing involves determining the lives of
the specimens under specific test conditions. It is well known that considerable scatter in results occurs, and it is often found that the majority
of specimens in a test fail rapidly, leaving a few which fail at much
longer times, or even do not fail at all before the test is discontinued.
This behavior raises considerable difficulties, both theoretical and practical, in deciding when to terminate a test, choosing a satisfactory "representative value," and in comparing such values.
Criteria—Various criteria of the time-to-failure, or specimen life,
have been used, depending upon the test procedure. Under direct-tensile
loading with weights or springs, the initiation of a stress corrosion crack
generally results in rapid fracture of the specimen, and the specimen life
is easily determined. There is the necessity, mentioned above, of verifyCopyright by ASTM Int'l (all rights reserved); Wed Dec 16 15:53:43 EST 2015
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
299
Under direct-tensile constant-deformation type loading to a relatively high stress, the initiation of a stress corrosion crack also generally
causes rapid fracture of the specimen. Because of the stiffness of such
stressing frames,3 however, the problem of the tensile failure of a corrosion-weakened specimen is greatly reduced and occurs only under
conditions when severe localized corrosion has developed. With relatively low applied stress the time required to initiate a stress corrosion
crack will be longer, and under constant-deformation load, a specimen
of an alloy that is susceptible to acceleration of general corrosion by
stress may develop many stress corrosion stringers without resulting in
actual cracking of the specimen. In this case, comparative tension tests
with corroded unstressed specimens and microscopic examination must
be relied upon to reveal the tendency for stress corrosion.
Constant-deformation loaded bent specimens of various types, including beams, loops, U-bends, C-rings, etc. of high-strength alloys that
are appreciably susceptible to stress corrosion cracking provide fractures
that are sharply defined and easy to detect. Frequently, however, with
lower-strength alloys, or more resistant alloys, or with relatively low
applied stress, cracking will be initiated slowly and is difficult to detect.
Cracks may initiate at multiple sites, and a real problem arises in deciding
when to consider a specimen "failed" and when to terminate the test.
Inasmuch as these specimens do not always fracture, it is preferable
to report the first crack as the criterion of failure. It is common practice
to make this inspection with the naked eye or a low-power (X10 to X15)
magnifying glass. If there are indications noted that cannot be established
definitely as a crack by this type of examination, the investigator should
either (1) note the date of this first suspicion of cracking and continue the
exposure of the specimen and watch for further growth that will confirm
the first indication as the failure date, or (2) discontinue exposure of the
specimen and perform a microscopic examination of a cross section taken
through the suspected crack area to verify the crack. Because of this
difficulty, some investigators have resorted to other criteria, such as a
fixed amount of "sag," or permanent set, to denote failure.
Representative Values—The arithmetic mean specimen life is widely
used, because it can be manipulated algebraically and can be used in
most standard statistical tests of significance. It should be remembered,
however, that extremely large or extremely small values may seriously
affect the mean and render it atypical of the average distribution. Moreover, in using the mean it is assumed that the population is normally or
very nearly normally distributed. The median, on the other hand, has the
advantage that it is influenced less by extreme values and has the important property of requiring no assumption about the population dis-
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300
STRESS CORROSION TESTING
tribution. When practicable, the population distribution should be determined.
Schikorr and Wasserman [7] have suggested that specimen lives do
not follow a normal distribution and that the median is a more reproducible figure. These authors also point out that median values may be
obtained much faster than arithmetic mean values because only about
half the number of specimens exposed need to be tested to failure. The
TABLE 2—Comparison of several representative values for ranking protective
coatings for preventing stress-corrosion cracking.0'
Coating 7, ZnCrO4
primer
Specimen Life, L
Specimen Life, L
Days
Log L
Days
43
43
43
91
641
1.63347
1.63347
1.63347
1.95904
2.80686
Coating 5, Alumilite 205
14
43
56
56
63
Log L
1.14613
1.63347
1.74819
1.74819
1.79934
Mean
172
Range
43 to 641
86
Geometric mean . . .
Median
43
Coating 11, polyurethane
43
63
63
91
117
Mean
46
1.61506
Range
14 to 63
41
Geometric mean
Median
56
Treatment 2, shot
peened
63
1.79934
1.63347
2.04922
112
1.79934
2.04922
112
1.79934
2.38021
240
1 .95904
2.38021
240
2.06819
Mean
Range
Geometric mean
Median
1.85188
75
43 to 117
71
63
1.93326
Mean
Range
Geometric mean
Median
2.13164
153
63 to 240
135
112
0 Source: Ref 11. Interference ring specimens of 2014-T6 alloy stressed at 75%
of the yield strength exposed in quintuplicate to the seacoast atmosphere at
Point Judith, R. I.
median is used in a German specification [8], which provides additionally
for the use of the geometric mean if the number of specimens is small.
Booth et al [5] and Booth and Tucker [9] have investigated the statistical distribution of specimen lives (endurances) in accelerated laboratory tests of aluminum-magnesium alloys in which all of the highly
stressed specimens failed. These authors observed that raw specimen
lives were not satisfactory for use in statistical analysis of experiments
using most standard tests of significance for two reasons: (1) the variance
of the specimen lives was not independent of their arithmetic mean, and
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Both of these objectives are overcome if the logarithms of the endurances
are taken as the experimental data. Besides producing a symmetrical distribution, taking logarithms produced a distribution that appeared to be normal. These considerations apply equally well even if no statistical analysis is
contemplated, and they demonstrate that the arithmetic mean of raw endurance is not a very meaningful parameter. On the other hand, the geometric
mean (which is the antilog of the arithmetic mean of the logarithm) is meaningful.
FIG. 3—Comparison of various rankings of protective coatings for prevention
of stress corrosion cracking (Table 2).
Further investigation of the frequency distribution of specimen lives
for other alloys and other test conditions is needed. It should be established, for example, whether or not the population distribution of
specimen life will be the same for specimens that are highly resistant as
for those that are less resistant under a given set of test conditions.
Further examination of stress corrosion data may reveal an analogy to
fatigue testing experience which has shown that, in the finite-life range
of the S-N curve, normal distributions generally result from transformation to log cycle-life; however, at stresses near the fatigue limit, where
runouts are observed, normal distributions are not obtained even after
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STRESS CORROSION TESTING
The problem of choosing a "representative value" for comparing one
item with another is illustrated by the data tabulated in Table 2 and
summarized graphically in Fig. 3. The one extremely long life of a specimen with Coating 7 badly distorts the ranking using the arithmetic
mean. Including either the range or the time for the first failure for each
coating helps to show a more complete picture, and it appears in this
FIG. 4a—Stress/cracking-time curves for a mild steel immersed in boiling
LiNOa solutions [12].
FIG. 4b—Stress-cracking-time curves of steels of different compositions [13].
example that the combination of geometric mean and first failure time
gives the most representative summary. Another way of presenting these
types of data is discussed under section on per cent survival curve.
Stress-Life Curve
More information about the resistance to stress corrosion cracking of
a material can be obtained by testing specimens over a range of applied
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stresses.
Such data usually are presented graphically with stress plotted
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
303
against specimen life. The primary interest is generally in the long-life
portion of the curve (stress corrosion "threshold"); therefore, more
tests should be made in this vicinity. If only a relatively small number of
specimens or stressing frames are available, single specimens may be
tested at each stress level. When more tests can be performed simultaneously, information can be gained sooner by exposing groups of
specimens at several different stress levels.
In the latter procedure, each group should consist of at least four
specimens if it is desired to estimate the variability of the data. (Ten or
more specimens are preferable to obtain some indication as to the shape
of the distribution of specimen-life values.) Furthermore, to obtain
approximately an equal degree of precision throughout the range of
FIG. 5—Comparison of stress corrosion tests of Al-Zn-Mg 3 under constant
tensile and bending load [14].
stress levels, more specimens should be tested in the long-life than in the
short-life range. Smaller groups, of course, can be tested if the specimenlife distribution shape is known or if statistical comparisons are not desired. Additional guidance may be found in ASTM Recommended
Practice for Choice of Sample Size to Estimate the Average Quality of
a Lot or Process (E 122 — 58).
Stress-specimen-life curves generally are of a hyperbolic type as shown
for mild steel [12,13] in Fig. 4a. Similar curves have been observed for
other metal alloy systems, including those of aluminum [14,15], magnesium [16], austenitic stainless steel [77], and copper [18]. When the
specimen lives are plotted on a log scale (Figs. 4b and 5), the curves consist of two straight line branches with one being parallel to the time
axis
and designating
a "threshold
stress" for
theDec
particular
test EST
conditions.
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304
STRESS CORROSION TESTING
aluminum-zinc-magnesium alloy, Brenner and Gruhl [14], and Gruhl
[79] obtained curves such as those shown in Fig. 5 and derived the
following equation for the higher stress level branch
where:
L =
a =
c =
K =
specimen life,
applied stress,
constant relating to the degree of stress dependency, and
constant which depends upon temperature in accordance with
the Arrhenius equation.
Azhogin [20] has proposed an equation for the entire curve.
where:
<r = applied elastic stress
<Tcr = critical stress (threshold)
t = specimen life
K = constant
The curve is a hyperbola displaced along with the o-axis by the value of
<r cr . When the applied stress is equal to or less than <r cr , the alloy is not
susceptible to stress corrosion cracking. When (a — o-C7.) was plotted
against / on log-log paper for experimental results on tests of highstrength steels, brass, and magnesium alloy, straight lines were obtained, thus satisfying the equation for the curve.
To evaluate a commercial product it can be useful to plot the stresslife points for a large amount of data for which upper and lower limit
curves can be drawn (Fig. 6). The lower limit represents the threshold
stress. Another alloy can then be compared with the commercial alloy
by superimposing data points on the performance band for the commercial alloy (Fig. 7).
Per Cent Survival Curve
A curve can be constructed using the values of per cent survival of a
group of specimens after a given period of exposure that are observed
for several (at least three) values of applied stress. This method of
analyzing stress corrosion data is especially useful when some of the
specimens "run out," that is, survive the duration of the test. An example
is given in Fig. 8. Similarly, per cent survival can be plotted against
time of exposure (Fig. 9). These curves are applicable particularly when
relatively large size groups of specimens have been tested.
Curves of
type,
which
can bereserved);
plotted Wed
eitherDec
as per
cent survival
or
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per
cent failure, provide
one of the most significant comparisons of
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FIG. by
6—Specimen
and section
size
can16affect
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structures are shown
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taken
in theWed
short-transverse
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minum alloy
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
307
FIG. 8—Effect of environment on the resistance to stress corrosion cracking
of short-transverse specimens from aluminum alloy 7079-T6 forgings (see footnote 3).
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FIG. 9—Effectiveness of various types of protective coatings in delaying stress
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cracking of 2014-16
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308
STRESS CORROSION TESTING
materials that have different degrees of susceptibility to stress corrosion
cracking. While the curve can be drawn on any type of graph paper, the
per cent survival values often will lie along a straight line when the data
are plotted on normal probability paper (Fig. 10) [21}. Confidence
limits can be computed and also plotted on the graph.
FIG. 10—Distribution of stress corrosion results for USS 12MoV stainless
steel [221.
Stress Corrosion Thresholds
Suss [22] has pointed out the many pertinent factors that can influence
the corrosion process and the actual tensile stress within a member.
Some of these are: (1) variations in test specimens (size, shape, stress
concentrators, internal stress, and method of load application), (2)
variations in environments (chemistry, temperature, galvanic cells, etc.),
and (3) variations in materials nominally within the same specification.
In consideration of the many variables and of possible interactions
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among
it would
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to attempt
to establish
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FIG. 11—Relative resistance of extrusions in several high-strength aluminum alloys to stress corrosion cracking. The "highest sustained tension
stress that did not cause Jailure" was obtained from the bottom limit of a band drawn in the same manner as those in Fig. 6. Arrows indicate no
stress corrosion failures at the highest stress employed [15].
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310
STRESS CORROSION TESTING
parameters below or above which stress corrosion will not occur in
service. Nevertheless, for the purpose of comparing alloys, various
thresholds can be used advantageously.
Critical Stress—Stress corrosion thresholds in terms of the sustained
tensile stress (more accurately defined as the initially applied strain or
load, with allowance for residual or internal stress in the metal) determined under specific test conditions can be useful as guide lines for
comparing alloys and manufactured products [75]. A graph such as
that shown in Fig. 6, which represents a summary of a great many tests,
is of special significance. The threshold stresses so obtained for various
materials can be compared as shown in Fig. 11. It should be emphasized,
however, that so-called threshold stresses determined in laboratory or
in field tests should not be regarded by designers as limiting stresses,
TABLE 3—Effect of test solution composition on resistance to cracking of steel.0
Solution
Atmosphere
pH
Specimens
Critical
train, Sc
9 N
distilled water
0.01% acetic acid
0.01% acetic acid
0.01% acetic acid
H2S
H2S
H2S
H2S
4.02
3.64
3.16
2.93
12
8
12
31
2.9
1.6
1.2
0.7
N-80
0.1% acetic acid
0.5% acetic acid
H2S
H2S
3.16
2.93
12
42
3.8
2.8
Alloy
0
b
Data excerpted from Table 2 of Ref 24.
Standard error is about 0.1 X 10~3 in./in.
in the same way that fatigue or endurance limits are used for dynamically
loaded members.
Azhogin and Pavlov [23] have used the "critical stresses," <r cr , to
compare various high-strength steel alloys and to predict their stress
corrosion performance in outdoor atmospheres based on laboratory
stress corrosion tests.
Critical Strain—Fraser et al [24] have developed a quantitative procedure using the probit method for studying sulfide corrosion cracking
of steel, by determining the degree of loading for a specific alloy at which
the probability of cracking failure with the test period is one-half. The
number obtained is called the "critical strain," and it is a function not
only of the alloy tested but also of the test environment and procedure.
Very susceptible steels have low critical strain (Se) values, whereas
nonsusceptible steels have high Sc values. Using alloys of varying Sc
values, all tested at the same high strain, the relative severity of any
given test environment can be measured. Analysis of the data from field
tests yields a severity rating, Rs, which is the critical strain of an alloy
which
would
be expected
to give
50 per
cent
in test.
High Rs
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
311
values are associated with severe test environments and low Rs values
with mild environments.
Determining the stress (or strain) at which the probability of failure
is one half is a more simple approach statistically and is just as useful
for comparing the relative susceptibility of two or more alloys as determining the threshold stress. The authors preferred to use strain rather
than stress because the loading necessary to produce failure exceeded the
yield strength in many instances. The test procedure consists of exposing
a number of specimens, usually about twelve, at a variety of different
strain levels for a definite time period. The test strains are chosen as
close as possible to the estimated critical strain so as to obtain maximum
FIG. 12—Effect of initial stress intensity on time-to-fracture, AIS1 4340 steel
heat treated to about 150 kg/mm3 yield strength [27].
information from each of the limited number of test specimens. The
critical strain is then calculated by the statistical technique of the probit
analysis (described in detail by the authors) [24]. The probit method also
is described in detail as used for making fatigue tests in Ref 10. A sample
of the authors' data illustrating the effect of test solution composition
on the resistance to cracking is given in Table 3.
Critical Stress Intensity—Tiffany [25] and Johnson and Willner [26]
have concluded the apparent existence of a stress-intensity threshold
value required to initiate stress corrosion cracking in a high-strength
steel. Studies by Brown and Beachem [27] at the U.S. Naval Research
Laboratory on a large variety of high-strength steels and titanium alloys
have reinforced this conclusion. By inserting a fatigue crack in the
specimen
commencing
the reserved);
test, stress
crackingEST
can2015
be
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STRESS CORROSION TESTING
caused to initiate immediately upon application of sufficient stress. The
stress at the root of the precrack can be described quantitatively by the
stress intensity parameter K ± , and a threshold value of this required to
initiate stress corrosion cracking is defined as ^Tlscc. An example is
shown in Fig. 12.
Percentage Stress Corrosion Susceptibility
By combining the stress corrosion index and a stress specimen life
curve, Jones [4] has arrived at another parameter, percentage stress corrosion susceptibility. This is obtained by plotting the index against the
applied stress expressed as per cent of the tensile strength. The crosshatched areas in the graphs in Fig. 13 were obtained by extrapolation of
FIG. 13—Effect
alloys [4].
of stress on corrosion of strain-aged aluminum magnesium
the plotted curves EF to E', the tensile strength, and to G, the threshold
stress, for these test conditions. The percentage stress corrosion susceptibility, then, is an expression of the cross-hatched area as a percentage
of the total area bounded by the axes. This parameter, while it has the
advantage of combining a considerable amount of test data into a single
number, has some of the same limitations discussed above for the stress
corrosion index.
Change in Mechanical Properties
Another way to evaluate quantitatively the degree of susceptibility or
resistance to stress corrosion cracking is to determine the effect of cracking upon some mechanical property or physical characteristic of the
test specimen. Such tests are meaningful, however, only if it is established
first
that the
in mechanical
property
the16result
of stress
corroCopyright
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sion
cracking and not
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
313
Relaxation—An example of this technique is a simple test described
by Thompson [25] for determining the resistance of copper alloys to
stress corrosion cracking in moist ammoniacal atmospheres and in natural
environments. In this procedure the resistance to stress corrosion cracking
is represented by the time to produce 50 per cent relaxation (permanent
set) of loop specimens. This may be found from relaxation versus time
data by interpolation. The data given in Fig. 14 show how a series of
copper alloys was rated by this method.
FIG. 14—Stress-corrosion cracking of brasses, and brasses containing a third
element in moist ammoniacal atmosphere [28].
Ductility—The relative susceptibility of titanium alloy "self-stressed"
beam specimens was assessed by Braski and Heimerl [29] using the loss
of bend ductility shown by a compression test. An example of the author's
data is given in Fig. 15, in which the presence of and the relative degree
of stress corrosion cracking is represented by the amount of shortening of
the specimen at fracture.
Summary
The selection
a criterion
of the degree
of 16
susceptibility
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314
STRESS CORROSION TESTING
the test. It should be recognized that there is a great deal of inherent
variability in the results of stress corrosion tests, and it is hazardous to
base comparisons upon small differences in the criteria. Statistical
analyses have not been used to any appreciable extent because usually
too few specimens are tested and also because of the difficult problem
of choosing suitable representative values.
Because of the marked effect that test procedures can have upon
ordinary criteria, such as specimen life, percentage of specimen survival,
threshold stress, etc., it is desirable for investigators to report details of
FIG. 15—Results of compression tests for Ti-8Al-lMo-lV specimens stressed
at 50 ksi [29]. (Shortening at fracture as a function of exposure time at 550 F.)
procedure and technique as well as detailed data along with their
analyses of the results. This is necessary to enable a reviewer to adapt
the information to his need and to make possible a reanalysis of the data
by some method that may be developed in the future.
Except in simple tests for which a crack-no-crack type of answer is
adequate, it is advantageous to test a material at several different levels
of applied stress, including unstressed specimens. This procedure enables
the proper identification of stress corrosion failures, and permits the use
of the most significant comparison methods, such as stress-per cent survival curves and stress-specimen-life curves including estimates of the
critical stress or strain (thresholds). Comparisons of specimen lives are
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SPROWLS ON REPORTING AND EVALUATING STRESS CORROSION DATA
315
Selecting representative values for specimen lives should be done with
caution: It appears that the geometric mean or the median may be more
meaningful than the arithmetic mean, but in any case, it is advisable to
cite also the range of lives or time until the first failure. Additional research is needed to establish the population distribution of specimen
lives under various testing conditions. The percentage stress corrosion
susceptibility and changes in mechanical properties are valuable under
special conditions.
The burden of reporting and interpreting comparative stress corrosion
test results can be lightened only as additional research leads the way to
fuller understanding of the general acceptance of standardized tests or
recommended practices.
References
[1] Dix, E. H., Jr., "Acceleration of the Rate of Corrosion by High Constant
Stresses," Transactions, American Institute of Mining, Metallurgical, and
Petroleum Engineers, Vol. 137, 1940, p. 11.
[2] Champion, F. A., Corrosion Testing Procedures, 2nd ed., Wiley, 1965, p. 133.
[3] 'The Stress Corrosion of Metals," Metals Handbook, American Society for
Metals, 1948, pp. 227, 228.
[4] Jones, E. L., "Stress Corrosion of Aluminum Magnesium Alloys II—Methods for Expressing Stress Corrosion Susceptibility on a Comparative Basis,"
Journal of Applied Chemistry, Jan. 1954, pp. 7-10.
[5] Booth, F. F., Tucker, G. E., and Godard, H. P., "Statistical Distribution of
Stress Corrosion Endurance," Corrosion, Vol. 10, No. 19, Nov. 1963, pp.
390t-395t.
[6] Phelps, E. H. and Loginow, A. W., "Stress Corrosion of Steels for Aircraft
and Missiles," Corrosion, Vol. 16, 1960, pp. 325t-335t.
[7] Schikorr, G. and Wasserman, G., Zeitschrift fuer Metallkunde, Vol. 40, 1949,
p. 201.
[8] Deutsche Normen, DIN 50908, "Prufung von Leichtmetallen Spannungskorrosion Versuche."
[9] Booth, F. F. and Tucker, G. E., "Statistical Distribution of Endurance in
Electrochemical Stress Corrosion Tests," Corrosion, Vol. 21, No. 5, May
1965, pp. 173-177.
[10] A Guide for Fatigue Testing and Statistical Analysis of Fatigue Data,
ASTM STP 91-A, 2nd ed., American Society for Testing and Materials,
Philadelphia, 1964.
[H] Sprowls, D. O. et al, "Investigation of the Stress-Corrosion Cracking of
High-Strength Aluminum Alloys," Eleventh Quarterly Report, 20 Jan., 1966,
Contract NAS 8-5340, George C. Marshall Space Flight Center, Huntsville,
Ala.
[12] Parkins, R. N. and Usher, R., "The Effect of Nitrate Solutions in Producing
Stress-Corrosion Cracking in Mild Steel," First International Congress on
Metallic Corrosion, Butterworth's, London, April 1961, p. 289.
[13] Parkins, R. N., 'The Stress-Corrosion Cracking of Mild Steels in Nitrate
Solution," Journal of the Iron and Steel Institute, Vol. 172, Oct. 1952, pp.
149-161.
[14] Brenner, Paul and Gruhl, Wolfgang, "Stress-Corrosion Cracking Tests of
Al-Zn-Mg 3 Under Constant Tensile and Bending Strain," Zeitschrift fuer
Metallkunde, Vol. 52, No. 10, 1961, pp. 599-607.
[15] Sprowls, D. O. and Brown, R. H., "Resistance of Wrought High Strength
Aluminum
Alloys
toreserved);
Stress Wed
Corrosion,"
Technical
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Laboratories, New Kensington, Pa. 1962. Also, published under title, "What
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316
STRESS CORROSION TESTING
Every Engineer Should Know About Stress Corrosion of Aluminum," Metal
Progress, Vol. 81, No. 4, April and May 1962.
[16] Loose, W. S. and Barbian, H. A., "Stress-Corrosion Testing of Magnesium
Alloys," ASTM-A1ME Symposium on Stress-Corrosion Cracking of Metals,
1944, American Society for Testing and Materials, Philadelphia, 1945, pp.
273-292.
[17] Denhard, E. E., Jr., "Effect of Composition and Heat Treatment on the
Stress-Corrosion Cracking of Austenitic Stainless Steels," Corrosion, Vol. 16,
1960, pp. 359t-369t.
[18] Bulow, C. L., "Stress-Corrosion Testing of Copper-Base Alloys," ASTMAIME Symposium on Stress-Corrosion Cracking of Metals, 1944, American
Society for Testing and Materials, Philadelphia, 1945, pp. 19-35.
[19] Gruhl, Wolfgang, 'The Temperature Dependence of the Stress-Corrosion
Cracking of Al-Zn-Mg 3," Zeitschrift fuer Metallkunde, Vol. 53, 1962, pp.
670-675.
[20] Azhogin, F. F. /'Corrosion Cracking of High Strength Steels," Inter crystalline
Corrosion and Corrosion of Metals Under Stress, Consultants Bureau, New
York, 1962.
[21] Loginow, A. W., "Specimens Used in Stress-Corrosion Testing of Alloys,"
Twenty-first Annual Conference of the National Association of Corrosion
Engineers, St. Louis, Mo., 15-19 March, 1965. To be published in Corrosion.
[22] Suss, Henry, "Practicality of Establishing Threshold Values to Eliminate
Stress Corrosion Failures in Metals and Alloys," Corrosion, Vol. 17, No. 2,
Feb. 1964, pp. 83-88.
[23] Azhogin, F. F. and Pavlov, Yu. L., 'Tendency of Steel Toward Corrosion
Cracking in Various Media," Korroziya i Zashchita Metallov, Sbornik Statey,
Moscow, Oborongiz, 1962, pp. 112-117. (Translation FTD-TT 64-643,
Foreign Technology Div., Air Force Systems Command, Wright-Patterson
Air Force Base, Dayton, Ohio.)
[24] Fraser, J. P., Eldredge, G. G., and Treseder, R. S., "Laboratory and Field
Methods for Quantitative Study of Sulfide Corrosion Cracking," Corrosion,
Vol. 14, No. 11, 1958, pp. 517t-523t.
[25] Tiffany, C. F., "Progress in Measuring Fracture Toughness and Using Fracture Mechanics," Fifth Report of a Special ASTM Committee, Materials
Research & Standards, Vol. 4, No. 3, March, 1964, p. 107.
[26] Johnson, H. H. and Willner, A. M., "Moisture and Stable Crack Growth in a
High Strength Steel," Applied Materials Research, Vol. 4, No. 1, Jan. 1965,
p. 34.
[27] Brown, B. F. and Beachem, C. D., "A Study of the Stress Factor in Corrosion Cracking by Use of the Precracked Cantilever Beam Specimen,"
Corrosion Science, Vol. 5, 1965, pp. 745-750.
[28] Thompson, D. H., "A Simple Stress-Corrosion-Cracking Test for Copper
Alloys," Materials Research & Standards, Vol. 1, No. 2, Feb. 1961, p. 108.
[29] Braski, D. N. and Heimerl, G. J., 'The Relative Susceptibility of Four Commercial Titanium Alloys to Salt Stress Corrosion at 550°F," NASA Technical
Note D-2011, National Aeronautics and Space Administration, Dec. 1963.
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Af. B. Shumaker,1 R. A. Kelsey,1 D. O. Sprowls,1 and
J. G. Williamson'i
Evaluation of Various Techniques for
Stress Corrosion Testing Welded
Aluminum Alloys
REFERENCE: M. B. Shumaker, R. A. Kelsey, D. O. Sprowls, and J. G.
Williamson, "Evaluation of Various Techniques for Stress Corrosion
Testing Welded Aluminum Alloys," Stress Corrosion Testing, ASTM STP
425, Am. Soc. Testing Mats., 1967, p. 317.
ABSTRACT: In the development of high-strength, weldable aluminum alloys it is necessary to determine the resistance to stress corrosion cracking
of experimental combinations of parent plate and filler alloys. Several
types of specimen and methods of loading have been studied to evaluate
techniques suitable for rapid screening of alloys and for demonstrating
the serviceability of alloys. Stress corrosion tests have been conducted in
3.5 per cent sodium chloride by alternate immersion and in seacoast
and industrial atmospheres, comparing beam and tension specimens. The
beam specimens were loaded by constant deformation, and tension specimens were loaded both by constant deformation and by constant load.
Welded specimens designed to investigate the effect of residual welding
stresses upon both butt welds and fillet welds also are included.
KEY WORDS: corrosion, stress corrosion, weldments, aluminum alloys,
salt solutions, chlorides, exposure testing, residual stresses
Stress corrosion has not been a problem with weldments of commercial
aluminum alloys. Many years of satisfactory service can be cited for
welded structures of various alloys of the aluminum-manganese, aluminum-magnesium, aluminum-magnesium-silicon, and aluminum-copper
types. The search for ever stronger aluminum alloys as materials of construction, however, leads investigators into complex alloy systems for
which the resistance to stress corrosion is a necessary consideration. Also,
the trend toward larger and more sophisticated structures in the transpor1
Research engineer, Chemical Metallurgy Div.; research engineer, Engineering
Design Div.; and assistant chief, Chemical Metallurgy Div., respectively, Alcoa
Research Laboratories, New Kensington, Pa.
2
Head, Corrosion and Finishes Sections, Materials Div., Propulsion and VeCopyright
by ASTM
Int'l (all rights
reserved);
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16 15:53:43
EST 2015 Ala.
hicle
Engineering
Laboratory,
George
C. Marshall
Flight
Center, Huntsville,
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STRESS CORROSION TESTING
FIG. 1—The extent of metallurgical changes in various aluminum alloys caused
by ASTM
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15:53:43 EST 2015
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heat of welding
as indicated
by potential
andWed
hardness
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
319
tation and aerospace industries has resulted in some problems with
residual and locked-in assembly stresses. Hence, there is a need for
reliable methods for testing the resistance to stress corrosion of weldments
of new alloys and tempers.
Stress corrosion testing a weldment presents a number of complications that are not encountered in the testing of the parent metal. Fusion
welding not only introduces a region of metal with a cast grain structure,
but also creates a number of metallurgical and mechanical property
Specimen dimensions for vdriOuS
plate thicknesses , inches
Formula for stressing
Where Ad = deflection, inch
f = nominal stress , psi
£ = Young's modulus, psi
J_
_o_
b_
1/8
1/4
4
4
2
2
3/8
1/2
3/4
1
43/4
43/4
31/2
31/2
5'/2
H/2
6'/2
6
8
12
6
s.
.k
10
10
13
13
17
12
12
15
15
19
20
25
22
27
FIG. 2—Beam stress corrosion assembly.
changes of the parent metal in the heat-affected zone (HAZ). These
changes have variable effects on the hardness and upon the electrochemical properties of different aluminum alloy systems (Fig. 1). Such
changes are especially important in the case of heat treated alloys
and can have a bearing on the choice of the most suitable test procedure.
It is the purpose of this paper to compare the merits of several test
procedures that have been used for evaluating aluminum alloys at the
Alcoa Research Laboratories. This program was designed to provide
information useful not only to the producers and the users of aluminum
alloys but it is anticipated that this comparison of test methods will be of
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also to the producers and the fabricators of other metals.
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320
STRESS CORROSION TESTING
Test Specimens and Methods of Loading
Four techniques have been evaluated at the Alcoa Research Laboratories for stress corrosion testing welded aluminum alloys. A description
of the types of specimens, methods of stressing, and a discussion of the
advantages and disadvantages of each follows.
Simple Beam Specimen
The beam specimen (Fig. 2) has been in use for over 20 years and has
been found satisfactory for all types of aluminum alloys. This specimen
is simple to make from any thickness of plate and is easy to stress in a
quantitative manner. Beams may be conveniently stressed in pairs (Fig.
FIG. 3—Calibration tests of welded beam-type stress corrosion specimen.
2), or they may be stressed individually and grouped in racks as has
been done by some other investigators [I].3
Beam deflections required to develop the intended tensile stress are
calculated with the formula given in Fig. 2 and are then applied by bolting
together the ends of the beams. The deflections are measured with a dial
gage to within ±0.0005 in. Thus the error in stress application, if the
beams were of homogeneous material arid the cross section were uniform,
is within 2 per cent; the precision of the deflection measurement is within
0.5 per cent; arid the error in determining Youftg's modulus of elasticity
(ASTM Method for Determination of Young's Modulus at Room
Temperature, E 111-61) is within 1 per cent.
The extent to which measured stresses in welded beam specimens agree
with the nominal stress Was determined from tests on welded l-in.-thick
plate. Strains were measured immediately adjacent to the weld bead arid
3
The italic numbers in brackets refer to the list of references appended to this
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
321
in the parent material outside the HAZ by means of foil type electrical
resistance strain gages, located as shown in Fig. 3. Calibration tests under
tensile loading showed the average measured strains in the parent material
to be in close agreement with the strain calculated by dividing the applied
load by the cross-sectional area and Young's modulus.
The results obtained in the bend tests on X7106-T6351 are shown in
Fig. 3. In these tests the beam specimens were deflected by an amount
calculated to stress the parent metal to 28.7 ksi (75 per cent of the 10-
FIG. 4—Welded U-bend stress corrosion specimen.
in. gage length yield strength of the weldment). For this case, involving
essentially elastic strains, the stress in the parent material outside the
HAZ was within 5 per cent of the nominal stress; while in the HAZ
adjacent to the weld, the stress, concentrated by the close proximity of
the weld bead fillet, was about 20 per cent greater than the nominal
stress.
The results of the bend tests of 2219-T8 type alloy beam specimens
are shown in Fig. 3. In these tests it was desired to stress the parent
metal to 30 ksi (75 per cent of the 10-in. gage length yield strength) for
which the calculated beam deflection was 0.46 in. The calibration showed
that a deflection of 0.46 in. resulted in an elastic strain in the center span,
outside
HAZ,Int'l
of(all
2800
to a stress
of 29 ksi (also
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STRESS CORROSION TESTING
within 5 per cent of the nominal stress). In this instance, the material immediately adjacent to the weld bead underwent plastic deformation, the
measured strain being 4500 /iin./in., equivalent to a stress of only 22
ksi. The latter stress was determined from a stress-strain curve obtained
by a tension test of a full section welded specimen, on which strains were
measured in the HAZ by electrical strain gages.
These data show that the stresses outside the HAZ are in close agreement (within 5 per cent) with the nominal values, regardless of the strains
in the HAZ. Although in tensile loading the stress in the HAZ will be at
least equal (or higher, due to stress concentration) to that in the parent
material, regardless of the amount of local yielding in the HAZ; it must
be realized that in bending, local yielding may result in a lower stress in
FIG. 5—Welded sheet tension specimen.
the HAZ than in the parent metal. Thus, to make direct comparisons between bending and tension tests, where local yielding occurs, it is desirable
to develop the same strain in the HAZ, accepting the fact that the stress
in the parent material will be greater in the bent specimen.
V-Bend Specimen
The U-bend specimen (Fig. 4) is a qualitative, highly stressed specimen
that has been widely used for various metals with and without welds. It
is of limited value and is realistic only for certain applications that
involve severe deformation of the welds. At the Alcoa Research Laboratories it has been used chiefly as a rapid screening of those welds that
are sensitized by plastic deformation [2].
The U-bend consists of a rectangular strip, its length depending upon
the desired bend radius, bent 180 deg around a mandrel. The bent specimen is allowed to spring back elastically to a stable position and then
stressed by bringing it back to the 180 deg position and holding it in
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
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FIG. 6—Constant-deformation type of stressing frame used for 0.125-in.-thick
sheet tension specimens in direct tension.
FIG. 7—Constant-load type of stressing frame for 0.125-in.-thick sheet tension
specimens in direct tension.
place by means of a bolt. This produces the same stress state that was in
the specimen hi the original bent condition.
Tension Specimen
Tension specimens of the general type shown in Fig. 5, like the Ubend, have been used extensively for testing various materials. A quantitative stress can be applied by a variety of methods ranging from constant
load to constant deformation. The stressing frames (Figs. 6 and 7) were
designed and built at Alcoa Research Laboratories to use in a study of
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324
STRESS CORROSION TESTING
weld testing methods [3], Practical considerations limit these methods of
stressing to relatively thin specimens. The constant-deformation type
stressing frame (Fig. 6) is a modification of the one used by the Rocketdyne
Division of North American Aviation, Inc.4 that, in turn, uses the
principle used much earlier for round tension specimens at the Alcoa
Research Laboratories [4]. To load the specimen in this stressing frame,
the side struts are simultaneously forced inward toward the specimen until
the desired strain in the specimen is achieved. A special loading device
is used to obtain steady and equal movement of the side struts to assure
uniaxial stressing of the specimen.
FIG. 8—Application of load to tension specimen in the constant-load stressing
frame. In both stressing frames (constant-deformation and constant-load), the
strain equivalent to the desired stress was measured over a Vi-in. gage length to
±1 pin. by means of an electrical strain gage placed adjacent to the weld heat-affected zone.
With both types of stressing frames, the strain, equivalent to the intended stress (according to Hooke's law), is measured by means of an
electrical extensometer attached to the specimen just outside of the heataffected zone on one side of the weld. Because of warpage during welding
the specimens may not always be perfectly flat and some bending stress
may be developed during loading. This effect, usually very small, is
averaged out by taking strain readings on both surfaces of the specimen
(Fig. 8). Because of the high precision of the strain measuring technique,
the error of stress application is governed by the error inherent in Young's
modulus (less than 1 per cent).
4
Private communication,
J. W., Rocketdyne
North American
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
325
An investigation was performed to determine the extent to which the
constant-deformation and constant-load fixtures (Figs. 6 and 7) approach
the intended conditions of loading. First the relationship between load
and fixture stiffness was experimentally determined as indicated in Figs.
9a and b. Load-deflection data were obtained for specimens slotted at
their midlength to give reduction in area of 0, 20, 40, 60, and 80 per cent
(Figs. 9a and c). The stress corresponding to a given slot (crack) depth
was found by trial and error. Let P° be the force required to stress a
FIG. 9—Procedure used for determination of fixture stiffness
cracking on specimen load-deflection relationship.
and effect
of
smooth specimen to the initial stress <T° and P' the force accompanying a
given crack depth. Then to satisfy equilbrium requirements
and
where:
s = specimen, and
/ = fixture.
Let AL°g be the initial change in specimen length produced by P°s and
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AL'S
the change in length corresponding to P's. The change in specimen
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326
STRESS CORROSION TESTING
FIG. 10—Effect of stressing frame stiffness (K) and increasing depth of cracking
on average net section stress in sheet-type tension specimen.
length, J(AL) S , may be found as shown in Fig. 9c. To maintain equilibrium, the incremental change in specimen length must equal the incremental deformation of the fixture or
where
the values of K° and K' being obtained from Fig. 9b.
From the value of P' which satisfies the above requirements, the average net area stress is
where A' is the net area for the given slot (crack) depth.
The variation of net section tensile stress with the per cent reduction
of net section through a propagating stress corrosion crack under various
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in Fig.Wed
10.Dec
Two
limitingESTcurves
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
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eluded, one for dead-weight loading when the stiffness, K = 0, and the
other for an infinitely stiff stressing frame for which K = oo. It can be
seen from the curves that the average tensile stress on the net section
increases most rapidly with stressing frames having the lowest stiffness.
The curve for the constant-load stressing frame, as expected, almost
FIG. 11—Typical H-plate specimen for stress corrosion tests of weldments.
Berry mechanical strain gage is in position for measurement just outside the heataffected zone of center section.
duplicates the dead-load curve. The curve for the constant-deformation
stressing frame, however, instead of duplicating the curve for an infinitely
stiff stressing frame, tends to approach the curve for the constant-load
stressing frame. This is because there is a certain amount of elasticity in
the frame components. The similarity of the curves for the two stressing
frames also is associated with the assumption of localized cracking, which
is the usual case for susceptible weldments. In testing materials other
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than
weldments, however, generalized cracking or intergranular attack
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328
STRESS CORROSION TESTING
may occur, and the similarity of the curves would be decreased. Calculations have shown5 that the curve for a similar constant-deformation
stressing frame in the case of generalized cracking approaches closely
the limiting curve for an infinitely stiff fixture.
Residual Tensile Stress
Butt Welds—The H-plate specimen (Fig. 11) was designed as a selfstressed specimen for testing thick plate in direct tension by utilizing the
residual stresses developed in plates welded under constraint. Tensile
stresses are developed across the weld as a result of the thermal mismatch
and plastic deformation accompanying welding of the center section. The
H-plate is a modification of a specimen developed at the Alcoa Research
FIG. 12—Sandwich specimen simulating rigid structure. Note stress corrosion
cracking in edge of center plate of this specimen immersed for 2 hr in 6 per cent
boiling sodium chloride solution.
Laboratories for the measurement of residual stresses developed in weldments. The advantages of this type of specimen are that special stressing
frames are not required, any thickness of plate can be used, and the
stresses are representative of those produced by constraint welding and
can be measured. Moreover, the stresses can be adjusted by reducing the
width of the center or side struts after welding. A disadvantage is that
residual stresses are inherently variable and the reproducibility of the
specimen is not as high as that of the previously described specimens.
The magnitude of the tensile stress developed in the test section was
determined from strain measurements made before welding and after
final sizing. Figure 11 shows the Berry mechanical strain gage in position for measurement, just outside the heat-affected zone of the center
section.
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See p. 342.
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
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Fillet Welds—Most of the weldable aluminum alloys have a good
resistance to stress corrosion cracking in the short-transverse direction;
therefore, residual stress developed across the edges of the plate during
fillet welding do not introduce a stress corrosion hazard. Care must be
taken, however, to avoid exposure of the short-transverse grain structure
when stressed in tension for two notable exceptions; namely, certain heat
treated aluminum-zinc-magnesium alloys and improperly processed strain
hardened aluminum-magnesium alloys containing more than 4 per cent
magnesium.
FIG. 13—Sections from sandwich specimens of 1.5-in.-thick 7039-T6 plate after
exposure to boiling 6 per cent sodium chloride solution for 96 hr. Failure was observed after only 2 hr in test for specimen at left with no protection. The specimen
at right which had the edges of the plate overlayed with 5356 filler alloy was free
from cracking.
To simulate rigid assemblies involving fillet or butt welds situated
close to the edge of a plate, a sandwich type of specimen (Fig. 12) has
been devised. This type of specimen consists of a center panel with
machined edges, with slightly smaller outer panels welded on each side.
Another type of specimen used for this purpose has been obtained by
removing slices from a cruciform weld-cracking specimen [5]. Critical
stub lengths (Fig. 13) for both this and the sandwich specimen range up
to about 1.5 times the plate thickness.
As in the case of the H-plate specimen, the fillet weld specimens are
realistic in that they contain actual welding stresses, and elaborate testing
fixtures are not required. A disadvantage, however, is that the residual
stresses cannot be controlled readily, and the specimens do not provide a
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STRESS CORROSION TESTING
Test Environments
Test environments used by the Alcoa Research Laboratories include
the 3.5 per cent sodium chloride (NaCl) alternate immersion (1-hr cycle
with 10-min immersion followed by 50-min drying in air at ambient room
temperature), the seacoast atmosphere at Point Judith, R.I., and the
industrial atmosphere at New Kensington, Pa. In the atmospheric exposure, specimens were exposed at 45 deg angle on racks facing south.
Test Materials
A description of the composition and properties of alloys for which
stress corrosion data are cited is given in Table 1.
TABLE 1—Nominal chemical composition of aluminum alloys discussed in this paper
(weight per cent).
Alloy
5456
5556°
2014
40432219
2319°
7039
X7139
X7106
X5180-.
51830
Cu
4.5
6.3
Si
0.8
5.25
Mn
Mg
0.8
0.8
5.25
5.25
05
0.8
Zn
Cr
0.10
0.10
0.3
0.25
0.25
0.25
0.45
0.75
2.8
2.8
2.2
4.0
4.8
Zr
4.0
4.0
4.2
2.0
0.20
0.10
0.10
0.15
Ti
V
0.10
0.15
0.15
0.15
0.10
0.10
0.15
0.15
0.10
0.10
Filler alloy only.
Experimental Results
Data from the Alcoa Research Laboratories' files were chosen to
illustrate the utility and the engineering significance of the various techniques described above for stress corrosion testing welded aluminum
alloys.
Specimens Stressed in Bending
Simple Beam Versus U-Bend—Simple beam and U-bend specimens
have been used extensively in the development of aluminum-magnesium
alloys. The simple beam specimen is used as a practical specimen to
demonstrate the serviceability of plate filler combinations. The U-bend
specimen is used as a screening test to investigate such effects as plastic
deformation (strain), magnesium content, and long-time natural aging
upon the aluminum-magnesium alloys.
Plastic deformation (strain) will accelerate the precipitation of aluminum-magnesium constituent during natural aging of alloys of relatively
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content
them to
become
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331
SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
corrosion cracking [2]. An illustration of the effect of plastic deformation of welds is given in Table 2 for 5456-H321 alloy welds. Fortunately,
the specific sequence of forming, heating, and stressing that developed
susceptibility to stress corrosion cracking in this test are not likely to
occur in industrial uses of welded aluminum alloys. Under the more
realistic sequence of heating, forming, and stressing, using both U-bend
and simple beam specimens, excellent resistance to stress corrosion
TABLE 2—Effect of plastic deformation on the resistance to stress corrosion
cracking of Y^-in.-thick 5456-H321 alloy welded plate.°
Beam Specimens Stressed 75% of Weldment Yield Strength (4-in. gage length)
U-Bend Specimens (4< radius)
Formed, Stressed, Formed, Heated
1 Week
Formed and Stressed and Heated 61 Week
at 212 Ffr, and
Stressed
212 F
F/N C
Time
in Test
F/NC
Time
in Test
F/NC
Time in Test
As Welded
1 Week 212 F6
Time
in Test
F/N C
Time
in Test
4 years
0/8
4 years
F/NC
3.5% NaCl ALTERNATE IMMERSION
0/4
4 years
0/4
4 years
4/4
5, 5, 33, 63 0/4
days
SEACOAST ATMOSPHERE—POINT JUDITH, R.I.
0/2
9 years
0/2
9 years
2/2
212, 436
days
0/4
0/8
11A
years
1Y2
years
INDUSTRIAL ATMOSPHERE— NEW KENSINGTON, PA.
0/2
9 years
0/2
9 years
2/2
859, 859
days
0 Butt-welded in three passes with parent filler by the consumable electrode
process, direct current straight polarity-metal inert gas (DCSP-MIG). Tensile
properties of weldment—tensile strength, 50.8 ksi; yield strength (0.2% offset in
4-in. gage length), 27.9 ksi; % elongation in 4-in., 15.7.
6 Heating 1 week at 212 F used to simulate long-time natural aging at room temperature.
c F/N = number of stress corrosion failures over number of specimens exposed.
cracking has been demonstrated. It is significant that aluminum-magnesium alloys containing 4 to 5.5 per cent magnesium have been used extensively for welded constructions of various types with eminently satisfactory results.
Specimens Stressed in Tension
Constant Deformation Versus Constant Load—The data in Table 3
obtained
on aluminum-zinc-magnesium and aluminum-copper alloys
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that the constant-deformation
and the constant-load stressing
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TABLE 3—Resistance to stress corrosion cracking of DCSP-TIG welded %-in.-thick sheet of several heat treated aluminum alloys loaded
by different methods.
Original Properties
°hcct \lloy
2014-T6 (as welded)
2219-T87 (welded as
2219-T37; aged 24
hr at 325 F)
Mloy
4043
2319
Condition
bead on
Tensile
Strength, ksi
Yield
Strength, ksi
Elongation in
2 in., %
52.0
36.4
3
Loading
Method0
Industrial
Atmosphere
3.5% NaCl Alternate
Immersion
F/N5
Days 0
F/N &
BCD
TCD
TCL
0/4
1/2
0/2
84
(1) 84
0/4
0/2
84
bead off
49.6
35.1
4
BCD
TCD
TCL
0/2
0/2
0/2
84
84
84
bead on
45.4
40.7
1
BCD
TCD
TCL
0/4
0/2
1/2
84
84
(43) 84
bead off
49.1
41.2
2
BCD
TCD
TCL
0/4
0/2
0/2
84
84
84
0/2
0/4
0/2
0/3
484
498
498
835
815
X5183
bead on
55.1
38.1
5
BCD
TCD
0/4
0/3
180
180
0/4
0/3
7039-T6 (welded as
7039-T6; aged 8 hr
at 225 F + 16 hr at
300 F)
X5183
bead on
61 .4
50.1
6
BCD
TCD
0/4
2/3
(137), (151)
180
0/4
3/3
BCD
TCD
0/4
0/3
(143)" 180
180
0/4
2/3
TCL
0/2
(123)* 180
3/3
X7139-T6 (as welded) X5180
bead on
55.1
38.1
5
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484
498
498
7039-T6 (as welded)
180
Days"
812
(153), (421),
(545)
835
(526), (625),
815
(481), (483),
(530)
61.4
50.1
X7139-T6 (welded as X5 180
X7139-T6; aged 8
hr at 225 F + 16 hr
at 300 F)
bead on
X7106-T6 (as welded)
bead on
53.0
37.7
bead on
59.8
50.0
X7106-T6 (welded as
X7106-T6; aged 8
hr at 225 F + 16 hr
at 300 F
X5180
bead off
51.0
45.4
6
BCD
TCD
1/3
3/3
(6) 180
(2), (2), (2)
2/2
3/3
TCL
3/3
(D, (2), (2)
3/3
5
BCD
TCD
TCL
0/4
0/3
0/3
180
180
(129)* 180
0/4
0/3
1/3
835
815
(584) 815
6
BCD
TCD
1/3
3/3
(13) 180
(2), (3), (5)
2/2
3/3
TCL
3/3
(4), (4), (5)
3/3
(83), (150)
(45), (58),
(58)
(69), (73),
(73)
BCD
TCD
0/4
0/3
180
180
0/4
2/3
TCL
0/3
180
3/3
3
(87), (164)
(21), (34),
(35)
(47), (47),
(54)
469
(113), (113),
478
(116), (165),
(197)
* Loading Methods: BCD = bending, constant deflection (Fig. 2); TCD = tension, constant deflection (Fig. 6); and TCL = tension,
constant load (Fig. 7). In all cases parent metal outside of HAZ stressed to 75 per cent yield strength in 2-in. gage length.
6
F/N = number of stress corrosion failures over number of specimens exposed.
c
Number
in parentheses is the time required for failure of specimen(s) that failed.
d
These failures, unlike those in the post weld aged aluminum-zinc-magnesium alloys which occurred along the fusion line, occurred in
the heat-affected zone of the parent metal. At this location, severe undermining corrosion occurred causing concentration of the stress and
initiation of small intergranular stress corrosion cracks extending parallel to the surface of the sheet.
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334
STRESS CORROSION TESTING
frames were equally effective, and there was no tendency for specimen
lives to be shorter under constant load, as might be inferred from a comparison of the curves in Fig. 10. Thus, the simpler and less expensive
constant-deformation stressing frame is preferred for testing weldments
when the application of direct tensile stress is desired.
Residual Tension (butt welds)—The procedure used in making the Hplate specimens represented an extreme effort to develop maximum
residual tensile stress across the 2-in. weld test section. The actual tensile
stresses measured in as-welded H-plates of X7106-T6 alloy averaged
about 15 ksi for tungsten inert gas (TIG) welded Vs-in. sheet, 20 ksi for
TIG or metal inert gas (MIG) welded %-in. plate and 26 ksi for MIG
welded 1-in.-thick plate. Post weld aging the specimens (8 hr at 225 F
+ 16 hr at 300 F) tended to reduce the above values only by about 2 to 3
ksi. Similar stresses were also obtained in Vs-in. and %-in.-thick H-plates
of other aluminum-zinc-magnesium alloys and of aluminum-copper
alloys 2014-T6 and 2219-T87.
Stress corrosion failures have occurred, which indicate that a surface
residual stress transverse to the weld of sufficient magnitude to cause
stress corrosion cracking of post weld aged XV106 and X7139 alloys can
be developed in H-plate specimens. Only a few H-plates have failed to
date, suggesting that the stresses are close to stress corrosion thresholds
for these alloys.
This is the first investigation in which the H-plate has been used, and
the information available to date indicates that this type of specimen
is worthy of further trial, especially for relatively thick sections, about
% in. and above in thickness.
Residual Tension (fillet welds)—In sandwich specimens (Fig. 12) with
test plate of 1.5-in.-thick 7039-T6, extensive stress corrosion cracking
occurred in the plate edges after only 2-hr immersion in boiling 6 per cent
NaCl solution. Using this specimen in an investigation of various protective measures to prevent this cracking, it was found that one of the most
effective procedures was to "butter" the edges, that is, to lay down on
the edges a protective coating of filler alloy (Fig. 13).
Bending Versus Tension
The first stress corrosion tests of welded joints providing a direct comparison of tension versus bending were performed on welded 2014-T6
sheet. Two tension specimens (Fig. 5) were machined from a butt welded
panel 0.090 in. thick. One specimen was stressed as a beam, and the other
was stressed in tension in a constant-deformation fixture similar to that
shown in Fig. 6. The tension specimen was loaded to 25 ksi as measured
in the parent metal outside the HAZ by means of a mechanical strain
gage. The strain developed in the HAZ adjacent to the weld bead was
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also
measured
using
VHn.
strain
The beam
specimen
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335
FIG. 14—Representative of failures of TIG welded X7106-T6 and X7139-T6
sheet (VI in. thick) employing X5180 filler wire (post -weld aged). Etch-.Keller's.
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336
STRESS CORROSION TESTING
having a similarly located electrical strain gage was then loaded to the
same local strain in the HAZ as determined for the tension specimen.
When placed in the 3.5 per cent NaCl alternate-immersion test, the
tension specimen failed in the HAZ adjacent to the fusion line after 2
days, but the beam specimen was still intact after exposure for 144
days. Metallographic examination of both specimens established the evidence of stress corrosion cracking in the tension specimen and the absence
of cracking in the beam specimen.
Welded Vs-in.-thick sheet of several aluminum-copper and aluminumzinc-magnesium alloys was evaluated using beam specimens and directtension specimens. In all cases the specimens were loaded by constant
deformation to develop a tensile stress in the parent metal outside the
HAZ equal to 75 per cent of the 2-in. gage length yield strength of the
weldment. The test data are summarized in Table 3. Of the items that
showed some susceptibility to stress corrosion cracking, the percentage of
failures was higher and the specimen lives were shorter for the tension
test (TCD) than for the beam test (BCD). A recap of the failures is as
follows:
No. Failures/No. Specimens Exposed
Alternate
Immersion
2Q14-T6
7039 -T6 PWA°
X7139-T6, PWA
X7106-T6, PWA
Total
0
Tension
Bending
Alloy
0/4
0/4
1/3
1/3
Atmosphere
0/4
2/2
2/2
2/14
4/8
6/22 ( 27%)
Alternate
Immersion
1/2
2/3
3/3
3/3
Atmosphere
2/3
3/3
3/3
8/9
9/11
17/20 (85%)
PWA = post weld aged.
Thus, the tension test was shown to be more discriminating, particularly
when testing an alloy in the vicinity of its stress corrosion threshold. It
would be expected, therefore, that stress corrosion cracking threshold
stresses indicated by tensile loading would tend to be slightly lower than
those indicated by beam specimens. A typical stress corrosion failure of
the post weld aged aluminum-zinc-magnesium alloys when loaded either
by bending or tension is shown in Fig. 14. These results are in general
agreement with the findings of Brenner and Gruhl [6] in tests on aluminum-zinc-magnesium alloy sheet (not welded).
It is of interest to consider possible reasons for the difference in results
of the bending and the tension tests. Although it is difficult to compare
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the
state of stress at the crack tip in bending and tension specimens, some
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
337
FIG. 15a—Distribution of yield strength and residual stresses in a longitudinally
welded 5456-H321 plate 36 in. wide and Vi in. thick.
FIG. 156—Residual stresses produced transverse to weld by multipass welding.
(ll/4-in. 7039 aluminum alloy plate MIG welded with ^LQ-m.-diameter X5180 electrode).
information may be gained by comparing the energy stored in the two
types of specimens. Since no external energy is applied to the system during stress corrosion cracking, the energy required to propagate the crack
must be supplied by the release of stored elastic-strain energy in the
specimen and the stressing frame. There are at least two factors which
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to greater stored energy for specimens loaded in tension as
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338
STRESS CORROSION TESTING
compared to those loaded in bending. First, with increasing crack depth,
the potential energy corresponding to the tensile stress field remains
greater in a tension specimen than in a beam specimen. Second, for tensile
stressing frames of the types shown in Figs. 6 and 7, considerable energy
is stored in the frames, whereas the only stored energy in the bending
fixture shown in Fig. 2, is that stored in the supports at midspan and in
the bolts. This is relatively small compared to that in the tensile fixtures.
There are two other factors which could contribute to greater crack
sensitivity in tension tests than in bending tests. First, both surfaces (and
the edges) of the tension specimen are stressed in tension and hence are
vulnerable to the initiation of stress corrosion cracking, whereas only the
convex side of the beam specimen is stressed in tension and likely to
initiate cracks. Second, in a direct-tension test, the specimen will fracture
when the crack depth is sufficient to produce an average stress on the net
section equal to the fracture stress of the material (under the existing
stress state); however, in bending, only the extreme fibers in the vicinity
of the crack tip are stressed to the breaking strength, so that failure will
tend to occur progressively rather than suddenly. (This appears to be
borne out in tension tests of notched bars; sharply notched tension
specimens generally fail suddenly, whereas sharply notched beam specimens fail gradually by pulling the sections apart.)
General Discussion
The stress corrosion performance of a welded structure is influenced by
(1) the compositions of the parent plate and the filler alloy; (2) the
metallurgical conditions of the parent plate, the heat-affected zone, and
the filler alloy; (3) the magnitude of the sustained tensile stress, which is
determined by the algebraic sum of the stresses introduced on the exposed surface; and (4) the physical and chemical nature of the corrosive
environment.
In the evaluation of the stress corrosion performance of weldments,
factors (1) and (2) can be markedly influenced by variations in welding
procedure. The heat from welding produces metallurgical changes in the
parent metal and reduces the strength in part of this heat-affected zone
(HAZ). Examples of such effects are shown for both strain-hardened
and heat treated alloys in Figs. 1 and 15.
The yield strength across butt welds made in aluminum alloys, as
determined at 0.2 per cent permanent set, depends upon the gage length
used, the yield strength increasing with gage length. Structural designers
recognize from past experience with riveted and bolted joints that a small
amount of local yielding is characteristic and will not significantly affect
the performance of a structure [7]. Thus, the yield strength value corresponding to 0.2 per cent offset on a 10-in. gage length is considered to
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SHUMAKER ET AL ON WELDED ALUMINUM ALLOYS
339
same magnitude commonly encountered in riveted or bolted joints at
loads that cause an average net section stress equal to the yield strength
of the plate. For weldments of most alloys, the proportional limit based
on 10-in. gage length tests will be 50 to 70 per cent of the 10-in. gage
length yield stress.
At the Alcoa Research Laboratories, the practice of loading stress
corrosion specimens to 75 per cent of the 10-in. gage length tensile yield
stress has been adopted. This stress level will result in some inelastic
deformation of the metal in the reduced strength zone. The justification
for this is that locked-in assembly stresses developed in structural members welded under constraint may be as high as the yield strength of the
weakest metal in the heat-affected zone. When the structure is subjected
to external loads, the load stresses add to the locked-in stresses and
generally result in some inelastic deformation of the metal in the reducedstrength zone.
In addition to the assembly stresses developed in weldments made
under constraint, it is possible for local residual stresses resulting from
thermal mismatch and plastic deformation during welding to be developed
in the immediate vicinity of the weld. These stresses may be developed
parallel to the weld [9] (Fig. I5d) or transverse to the weld (Fig. 15b).
The former stresses will be relieved when cross weld specimens are cut
from the welded panel. However, the latter type stresses, which generally
occur only in thick plate joined by multipass welds, will not be relieved
when full-thickness cross weld specimens are cut from the panel. It is
evident from the data shown in Fig. I5b that such stresses can be appreciable and variable and should be recognized when it is attempted to
develop a specific stress in the test specimen.
It is obvious that in attempting to make a quantitative stress corrosion
test of a weldment, the mechanics of developing a known stress in the test
specimen is complicated by the presence of the weld bead (or beads).
Application of a load could be simplified in the case of light gage material
by machining the weld beads flush with the plate or in the case of multipass welded plate by machining uniform thin (0.125-in.-thick) specimens from the weldment. Data in Table 3 for 0.125-in.-thick sheet tested
"bead-off" versus "bead-on" both as beams and as tension specimens
show that tests of bead-off specimens may fail to reveal an appreciable
susceptibility to stress corrosion cracking. Although testing bead-off
specimens may be advantageous for some purposes, these data raise a
doubt as to whether or not such results would be representative of the
majority of industrial weldments.
Summary
Several types of test specimens and methods of loading have been
described,
theInt'l
investigator
may choose
a 16
test15:53:43
procedure
that best
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340
STRESS CORROSION TESTING
associated with the testing of weldments be understood so that a procedure can be selected that will not give misleading results. Although more
information is needed on this subject, the following observations are
significant.
Beam specimens with the weld bead(s) intact provide a versatile,
reliable test for determining the serviceability of a weldment in varied
environments. The test is reasonably quantitative provided that the
specimens are not overstrained during loading. In testing beams of relatively thick plate containing multipass welds, consideration must be given
to the presence of residual welding stress transverse to the weld.
For laboratory studies that in some cases may require the most rapid
and discriminating test, the direct-tension specimen is advantageous.
Stressing frames of the constant-deformation and spring-loaded types are
equally effective, and more practicable, especially the constant-deformation frame, than dead-weight loading equipment. Often there is a
tendency for selective attack in the HAZ causing a reduction of the
cross-sectional area, which in the case of the direct-tension test may result
in failure because of tensile overload that could be confused with a stress
corrosion cracking failure.
To develop the maximum tensile stress that is tolerable in a welded
structure, the specimen should be loaded to the point of developing a
small amount of inelastic deformation in the HAZ. As a general rule, this
can be accomplished by stressing the test specimen to 75 per cent of the
10-in. gage length yield strength of the weldment.
The U-bend specimen is a qualitative, highly stressed specimen of
limited value, because it is realistic only for certain applications that
involve severe deformation of the welds. It can be useful for rapid screening of weldments that are sensitized by plastic deformation.
Testing specimens with the weld beads machined off will permit a
closer control of applied stress but may not provide as severe or as
realistic a test as testing bead-on specimens.
For testing butt welds in thick plate, the self-stressed H-plate specimen permits the combination of extreme assembly stress due to welding
under constraint and local residual welding stress and is worthy of further
evaluation.
A cknowledgment
A portion of the work was supported by the George C. Marshall
Space Flight Center of the National Aeronautics and Space Administration (Contract NAS 8-5340) to whom the authors express their appreciation. The contract work was administered under the technical direction
of the Propulsion and Vehicle Engineering Laboratory, Materials Div.,
George C. Marshall Space Flight Center, with J. G. Williamson acting
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References
[7] Baysinger, F. R., "Investigation of Welding and Fabricability of Kaiser Experimental Alloy MR39A," Contract NAS 8-5065, Request PT2-82464, 21
May, 1964.
[2] Dix, E. H., Jr., Anderson, W. A., and Shumaker, M. B., "Development of
Wrought Aluminum-Magnesium Alloys," Technical Paper 14, Alcoa Research
Laboratories, New Kensington, Pa., 1958.
[3] Sprowls, D. O. et al, "Investigation of the Stress-Corrosion Cracking of High
Strength Aluminum Alloys," Contract NAS 8-5340, summary report, 1 August,
1965, sponsored by the George C. Marshall Space Flight Center, Huntsville,
Ala.
[4] Sager, G. F., Brown, R. H., and Mears, R. B., "Tests for Determining Susceptibility to Stress-Corrosion Cracking," ASTM-AIME Symposium on StressCorrosion Cracking of Metals, 1944, American Society for Testing Materials,
Philadelphia, 1945, p. 255.
[5] Weiss, S., Ramsey, J. N., and Udin, H., "Evaluation of Weld-Cracking Tests
on Armor Steel," Welding Journal, Vol. 35, No. 7, 1956, p. 348S.
[6] Brenner, Paul and Gruhl, Wolfgang, "Stress-Corrosion Cracking Tests of AlZn-Mg 3 Under Constant Tensile and Bending Strain," Zeitschrift fuer Metallkunde, Vol. 52, No. 10, 1961, pp. 599-607.
[7] Hill, H. N., Clark, J. W., and Brungraber, R. J., "Design of Welded Aluminum
Structures," Transactions, American Society of Civil Engineers, Vol. 127,
Part II, 1962, pp. 102-126.
[5] Task Committee on Lightweight Alloys, "Suggested Specifications for Structures
of Aluminum Alloys 6061-T6 and 6062-T6," Paper 3341, Proceedings, American Society of Civil Engineers, Journal of the Structural Division, Dec. 1962, p.
1; Alcoa Handbook of Design Stresses for Aluminum.
[9] Hill, H. N., "Residual Welding Stress in Aluminum Alloys," Metal Progress,
Aug. 1961.
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B. W. Lifka1 andD. O. Sprowls1
Stress Corrosion Testing of 7079-T6
Aluminum Alloy in Various Environments
REFERENCE: B. W. Lifka and D. O. Sprowls, "Stress Corrosion Testing
of 7079-T6 Aluminum Alloy in Various Environments," Stress Corrosion
Testing, ASTM STP 425, Am. Soc. Testing Mats., 1967, p. 342.
ABSTRACT: Stress corrosion cracking of certain 7079-T6 aluminum alloy
aircraft structures involving sustained tensile stress acting in an unfavorable grain direction has occurred sometimes in surprisingly mild environments. The resistance of short-transverse specimens of 7079-T6
products has been evaluated at different levels of applied stress in several
different seacoast and inland industrial atmospheres. It was found that the
atmosphere was more critical than the generally accepted alternate-immersion test using 3.5 per cent sodium chloride. The results of the outdoor
tests are correlated with eight accelerated exposures; and consideration is
given to the most suitable laboratory stress corrosion test environment for
this alloy. A detailed analysis of the constant deformation methods of
stressing is included.
KEY WORDS: corrosion, stress corrosion, aluminum alloys, environmental testing, exposure testing, chlorides, salt solution, salt spray test,
anisotropy
Aluminum alloy 7079-T6 is attractive for aerospace and hydrospace
applications because of its desirable mechanical properties in thick sections. In fact, 7079-T6 has become the alloy most widely used in the
United States for large high-strength aluminum forgings. Excellent service
over a number of years has been demonstrated in a wide variety of parts
involving large tonnages of metal. This generally satisfactory service
record is the result of the cooperation of materials engineers and designers. There is a strong incentive for such cooperation when it is realized
that an understanding of the capabilities and the limitations of materials
combined with necessary precautions in design and assembly [I]2 will
permit the use of stronger alloys and tempers.
1
Engineer and assistant chief, respectively, Chemical Metallurgy Div., Alcoa Research Laboratories, New Kensington, Pa.
2
The italic numbers in brackets refer to the list of references appended to this
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
343
Service failures resulting from stress corrosion cracking occasionally
have occurred in parts of 7079-T6 and other high-strength aluminum
alloys. Invariably these failures have resulted from residual or assembly
surface tensile stress acting continuously in the short-transverse direction (transverse in the case of round and square sections) relative to the
grain structure. Longitudinal tensile stresses, on the other hand, have
rarely caused stress corrosion problems. This record is consistent with the
results of laboratory tests [2], which have shown that all of the commercial high-strength aluminum alloys are highly resistant to stress
TABLE 1—Resistance to stress-corrosion cracking of 2-in.-thick 7079-T651 plate.
Number of triplicate specimens failing and days-to-failure.a
Test Environment
Test
Period,
days
Industrial atmosphere (New
Kensington, Pa.). 1460
Seacoast atmosphere
(Point Judith,
1460
R-i.)
3.5% NaCl alternate
immersion
84
Acidified 5% NaCl
intermittent
14
spray
Acidified 5% NaCl
intermittent
30
spray
4.55% NaCl + 1.16%
CrO3 alternate
immersion
First run
30
Second run
30
0
Short-Transverse
Stress, 10 ksi
No.
Days
Long-Transverse
Stress, 53 ksi
No.
Longitudinal
Stress, 56 ksi
No.
Days
3 448, 644, 1202 0
0
3 126, 126, 716
3 716, 716, 826
0
0
0
0
3 6, 7, 10
0
0
? 10, 17
3 10, 15, 17
2
19, 24
0
3
3
7, q, Q
7, 7, 8
2
2
Days
23, 28
17, 24
Type of specimen: % -in. -diameter tensile bar.
corrosion cracking when stressed in the longitudinal or the long-transverse
directions; but that many alloys, including 7079-T6, have a relatively
low resistance when stressed in the short-transverse or the transverse
(round or square sections) directions relative to the grain structure. A
sample of data illustrating this directional behavior for 7079-T6 plate is
given in Table 1.
Purpose of Stress Corrosion Testing
One of the limitations associated with all high-strength metals is stress
corrosion cracking. "While it is relatively easy to determine if a product
is 'susceptible to stress-corrosion cracking,' it is far more difficult to
determine
possesses
a 'degree
hamper Wed
its
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STRESS CORROSION TESTING
{**;
(a) As heat treated 7079-T6 landing gear. Bored hole extends to point C, section
AB shown below.
(b) Section AB showing cracks on ID.
(c) Shows depth of crack in parting plane, Flick's etch.
(d) Tip of stress corrosion crack, Keller's etch (X100).
(e) Shows interfragmentary nature of crack, Keller's etch (X500).
FIG. 1—Example of a service failure.
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
345
general usefulness" [3]. The mechanism of stress corrosion of aluminum
alloys and their resistance to stress corrosion cracking have been under
continuous investigation at the Alcoa Research Laboratories for over 35
years. The objectives of this research have been to: (1) determine the
factors which must be controlled in order that aluminum alloy products
of superior strength may be successfully employed and (2) develop stress
corrosion resistant compositions and tempers. A number of publications
have been issued, some of which are identified in Refs 2 through 8.
Use of Accelerated Tests
It is readily recognizable that the most reliable data are those obtained
in environments representative of the conditions the final assembly will
encounter in use. For aerospace applications, typical environments are
industrial or marine atmospheres; while total, partial, and intermittent
immersion in fresh and sea waters are representative of hydrospace applications. Unfortunately, the prolonged exposure periods required for tests
in these environments to be conclusive are often prohibitive and accelerated laboratory tests must be used. It is essential, however, that an
accelerated test reproduce the same mode of failure as occurs in service
environments. It should rate the percentage failures and the times-tofailure of various thermal treatments and anisotropic behaviors in the
same order of merit as do natural environments.
The accelerated stress corrosion test most widely accepted by the
aluminum industry is the 3.5 per cent sodium chloride (NaCl) alternateimmersion test (1-hr cycle: 10 min immersion plus 50 min drying in air).
Extensive evaluation over many years at Alcoa Research Laboratories
[2,7] has established a good correlation between this procedure and extended seacoast and industrial atmospheric exposures for alloys and
tempers of the aluminum-copper and aluminum-magnesium systems
and for most aluminum-zinc-magnesium-copper alloys such as 7075 and
7178. Susceptible specimens of these alloys are just as likely, and frequently more prone, to stress corrosion crack in 3.5 per cent NaCl
alternate immersion than in seacoast or inland industrial atmospheres.
This information has enabled the industry to reliably test these alloy
types in the alternate-immersion test without the continual necessity of
back-up tests in natural environments.
Deviation of 7079-T6 Alloy
These same extensive tests have shown that alloy 7079-T6 deviates
from the general pattern in that at low stress levels transverse and shorttransverse specimens are less likely to stress corrosion crack in the alternate-immersion test than in atmospheric tests.
An example of this is illustrated by the case history of the 7079-T6
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STRESS CORROSION TESTING
FIG. 2a—Specimen used to simulate various levels of residual quench tensile
stress at the parting plane of the forging shown in Fig. 1.
FIG. 2b—Resultant data for specimen shown in Fig. 2a.
the rough machined forging unavoidably resulted in circumferential tensile
stress on the finish-machined inside wall of the cylinder. Cracking was
revealed in the inside wall of a number of the parts by a penetrant
inspection. Microscopic examination of the failed parts (Fig. 1) established that the cracks occurred in the parting plane of the forging where
the
residual
stress
a short-transverse
grain
structure.
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
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The cracks were intergranular (and interfragmentary) in nature, and
appeared as typical stress corrosion cracks. Measurements by layerremoval techniques on four parts that did not fail showed residual tensile
stresses in the circumferential direction of about 10 ksi at most locations,
with a maximum of 18 ksi. However, it is possible that higher stresses
than these might have existed in»the parts that actually failed.
A laboratory test was undertaken to duplicate the service failures using
the C-ring specimen shown in Fig. 2a. Rings were machined from the
forging, so that the parting plane was at the middle of the C. The wall
FIG. 3—Comparison of the probability of survival of short-transverse specimens of 7079-T6 alloy with alloys 7075-T6 and 7039-T63 in two environments
(21 to 30 specimens representing 7 to 10 lots of rolled plate as a basis for each
point).
thickness of the C-rings was machined to l/s in. by removing only the
outer surface so that the test surface was the same as in the original part.
The C was then spread to impose tensile stress on the inner surface, the
level of stress being determined from electrical strain gage measurements.
Exposures were made in 3.5 per cent NaCl alternate immersion and hi
seacoast and industrial atmospheres. The resultant data are presented
graphically in Fig. 2b.
In the alternate-immersion test, failures occurred at 40 and 30 ksi
but not at the 20 and 10 ksi stress levels. On the other hand, failures did
occur at the 20 and 10 ksi levels with prolonged exposures to the
atmosphere, indicating a need for an accelerated test correlating better
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348
STRESS CORROSION TESTING
firm the residual stress determinations of 10 to 20 ksi, although the timesto-failure were relatively long (11A to 3 years).
Investigation of the stress corrosion characteristics of 7079-T6 alloy
and of various experimental aluminum-zinc-magnesium alloys has shown
that this deviation of 7079-T6 alloy is related to its difference in composition from aluminum-zinc-magnesium-copper alloys such as 7075-T6, the
most significant difference being the decrease in copper content. Data
plotted in Fig. 3 for the 3.5 per qent NaCl alternate-immersion test show
FIG. 4a—Type of die forging used for comparison of the effect of environment.
FIG. 4b—Macroetched cross section shows parting plane structure and specimen location.
that at relatively low stresses, such as 15 or 25 per cent yield strength,
the percentage survival of 7079-T6 and 7039-T63 was far higher than
that of 7075-T6. However, in the case of the atmospheric exposure data
shown in Fig. 3, the percentage survival of 7079-T6 and 7039-T63 at
stresses of 15, 25, or 50 per cent yield strength has reversed and is now
far lower than that of 7075-T6.
Investgation of Stress Corrosion Cracking of 7079-T6 in Various Environments
Investigation was made to study the susceptibility of 7079-T6 to stress
corrosion cracking in various types of atmospheric environments and to
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FIG. 5—Shows the Vs-in.-diameter tension specimen, the various parts of the
stressing frame, and the final stressed assembly.
search for a more definitive laboratory accelerated test for 7079-T6
than the 3.5 per cent NaCl alternate-immersion test.
Procedure
Test Materials •
The major portion of the investigation was conducted on three 7079T6 die forgings of the type shown in Fig. 4a. The principal reason for
using this forging was that its uniform cross section permitted a large
number of replicate specimens to be taken across the short-transverse
parting plane grain structure as shown in the macroetched section in Fig.
4b. Tests were also made on transverse specimens from IVa-in.-diameter
7079-T651 rolled rod and in all three directions of 2-in.-thick 7079-T651
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350
STRESS CORROSION TESTING
FIG. 6—ARL's synchronous loading device used to stress specimens. A stressed
assembly and one assembled fingertight for stressing are shown to the left.
plate. All items had composition within the specified limits and had been
produced using commercial fabricating and heat-treating facilities. Aluminum Association composition limits for alloy 7079 are:
Alloy
Si
Fe
Cu
Mn
Mg
Zn
Cr
Ti
7079
0.3
max
0.4
max
0.4 to
0.8
0.1 to
0.3
2.9 to
3.7
3.8 to
4.8
0.1 to
0.25
0.1
max
Test Specimen and Method of Loading
The data to be discussed hi this paper were obtained with Vs-in.diameter by 2-in.-long tension specimens, stressed hi the wedge-type
frame shown in Fig. 5. Components of the frame are assembled fingertight, and the specimen is then stressed with ARL's synchronous loading
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device
shown in Fig. 6. The inward movement of the wedge-shaped side
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
351
pieces induces uniaxial tensile stress in the specimen. Applied stresses
usually are kept below the proportional limit, or about 75 per cent of the
0.2 per cent yield strength. Thus, only the straight line portion of the
stress-strain curve for the test specimen is needed, since the applied stress
is considered to be related directly to the measured strain by Hooke's law
and Young's modulus. During the stressing operation, the strain is
FIG. 7—Effect of loading method and extent of cracking on average net
section stress, local cracking.
measured with a Type F Huggenberger tensometer with a 0.5-in. gage
length and a magnification of about 2000.
The error of stress application at the point of measurement is probably
less than about 2 per cent, since the precision of the strain measurement
is within 0.5 per cent, and the error in determining Young's modulus is
within 1 per cent (Method for Determination of Young's Modulus at
Room Temperature, E 111-61). Because there is a certain amount of
inherent eccentricity in mating threads, it is impossible to completely
avoid some slight bending of the specimen. The likelihood of developing
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352
STRESS CORROSION TESTING
of the specimens and the synchronous loading device that ensures equal
and simultaneous movements of the side pieces of the stressing frame.
Tests have shown that the uniformity in stress around the circumference
is of the same order as is achieved in tension tests employing this same
type of specimen.
FIG. 8—Effect of loading method and extent of cracking on average net
section stress, general cracking, or corrosion.
A comparison of this method of loading with other techniques, such
as loading with dead weight (constant load) or with an infinitely stiff fixture (constant deformation), is given in Figs. 7 and 8. The change in the
average tensile stress on the net section caused by the initiation and
propagation of a stress corrosion crack or by general cracking has been
calculated and checked by a compliance experiment (Appendix I). With
the initiation of localized cracking, the average tensile stress on the net
section increases rapidly, as in the case of dead-weight loading, until the
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
353
fracture strength is reached and the specimen breaks. The average stress
does not increase quite as rapidly as with dead-weight loading, so the
specimen life can be expected to be longer for specimens loaded in this
frame than for dead-weight loaded specimens. In the case of general
cracking or severe general corrosion, however, there is a considerable
FIG. 9—Effect of environment on the resistance to stress corrosion cracking of
7079-T6 die forgings stressed in the short-transverse direction.
difference in the effects of the loading method on the change in stress:
whereas with dead-weight loading, a reduction in area of about 40 per
cent will result in tensile failure, a specimen loaded in the stressing frame
can be corroded completely away without actually breaking. Thus, this
method of loading is similar in effect to dead-weight loading when
localized cracks are initiated, and it has the advantage over dead-weight
loading when generalized corrosion occurs in that fractures from tensile
overload are not so likely to occur.
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TABLE 2—Individual specimen lives, in days, of the data shown graphically in Fig. 9. Triplicate, short-transverse, %-in. -diameter'by2'-in. -long
tensile bars from 7079-T6 alloy die forging.
Forging
No. 1
No. 2
No. 3
atSOF
New Kensington, Pa.
at 120 F
No. 3
Point Judith, R.I.
lOksi
20 ksi
30 ksi
10 ksi
20 ksi
30 ksi
10 ksi
20 ksi
30 ksi
10 ksi
20 ksi
30 ksi
10 ksi
20 ksi
30 ksi
10 ksi
20 ksi
30 ksi
730°
730"
730°
187
200
730°
180
200
302
730°
730°
730"
59
391
535
81
300
339
173
200
730°
730°
730°
730°
303
303
730°
3
39
66
20
44
50
44
135
140
94
94
99
559
633
643
1460
1460
1460°
123
123
123
198
201
205
262
293
303
23
24
25
20
77
80
38
150
155
20
23
38
598
598
836
1460
1460°
1460°
221
221
305
221
221
221
221
221
221
221
221
221
221
221
221
221
221
221
221
221
589
920
1322
232
1460°
1460°
232
232
232
232
180
730°
730°
140
140
151
249
296
421
94
94
94
77
730°
730°
730°
296
339
339
1141
1460°
1460°
140
151
339
232
232
145
232
232
145
145
232
67
145
145
145
145
232
145
145
145
Forging
No. 2
Point Comfort, Tex.
Cleveland, Ohio
Alternate Immersion — 12 Weeks
No. 1
Seacoast Atmosphere—4 Years
Industrial Atmosphere—4 Years
100% Relative Humidity
3.5% NaCl + 0.1%
NHiHCCh
3.5% NaCl
10 ksi
20 ksi
30 ksi
10 ksi
20 ksi
84°
84'
29
29
42
35
48
64
4
4
6
11
11
84°
84°
15
21
84°
84°
20
28
80
82
28
32
84°
84"
84=
84"
11
84°
84"
13
6
8
10
4
4
11
30 ksi
Total Immersion
Synthetic Sea Water*"
10 ksi
20 ksi
84°
84°
60
84°
84°
35
42
84°
10
18
35
84°
84°
84°
84°
84°
84°
84°
3.5% CrOs + 0.3%
NaCl + 2.9% KzCnOr—
12 Weeks
30 ksi
10 ksi
20 ksi
13
15
18
21
61
84°
61
84°
84°
84°
84°
84°
22
26
26
84°
84°
7
84°
84°
11
Boiling 6% NaCl 4 Days
30 ksi
10 ksi
20 ksi
30 ksi
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
1
3
7
2
2
2
4
21
28
4°
4°
4°
4°
4°
4°
4c
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c
Specimen did not fail.
Acidified 5% NaCl
Intermittent Spray at
120 F— 2 Weeks
10 ksi
7
11
11
7
14°
14°
14°
14°
14°
20 ksi
1
1
3
4
5
8
7
7
14°
30 ksi
1
1
3
2
3
3
3
3
4
LIFKA AND SPROWLS ON ALUMINUM ALLOY
355
Results and Discussion
Survey of Various Environments
The three die forgings were first tested in four different outdoor atmospheric environments: two inland industrial sites, New Kensington, Pa.,
and Cleveland, Ohio; a seacoast location in the northeastern United
States, Point Judith, R.I.; and a seacoast location in a warm climate,
Point Comfort, Tex. A summary of the test results (Fig. 9)3 shows similar
performance in all four locations, with a markedly higher percentage of
failures at the lowest stress level (10 ksi) than in the 3.5 per cent sodium
chloride alternate-immersion test. It is noteworthy that failure times
(Table 2) of specimens exposed to the seacoast atmosphere at Point Comfort, Tex., were not appreciably different from the lives of specimens exposed at Point Judith, R.I., where the average temperature is about 20 F
lower.
Exposures to atmosphere saturated with water vapor (100 per cent
relative humidity) at 80 or 125 F caused failures of all specimens stressed
to 30 ksi as did the exposures to the natural atmospheres, but in water
vapor there were fewer failures at 20 ksi and no failures in over two years
of testing at 10 ksi.4 In these tests in saturated atmospheres, a slightly
greater number of failures was encountered at the higher temperature.
Five laboratory test media were investigated in a search for an improvement over the 3.5 per cent sodium chloride alternate-immersion test.
Synthetic sea water (Specifications for Substitute Ocean Water, D 114152) was even less discriminating than 3.5 per cent sodium chloride
solution. The addition of 0.1 per cent ammonium bicarbonate (based
on work of Farmery and Evans [9]) to the 3.5 per cent sodium chloride
(Fig. 9) gave only a slight improvement. Continuous immersion in an
acidified-salt-dichromate solution suggested by Sager et al [5] was not
so effective as the 3.5 per cent sodium chloride alternate immersion.
Continuous immersion in boiling 6 per cent sodium chloride solution
[5] for four days was no more effective in producing failures at the low
stress levels than the alternate-immersion test but had the advantage of
being of shorter duration. A recently developed acidified 5 per cent
sodium chloride intermittent spray at 120 F [10] (Fig. 9) appears to be
the most promising, although this test still did not produce quite so
high a percentage of failures at the 10 ksi stress levels as did the atmospheric exposures.
Modifications o] the 3.5 per cent Nad Alternate Immersion
Tests utilizing the 2-in. plate were made to study the effect of purity
3
Tabular data on which Fig. 9 is based are given in Table 2.
*
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356
STRESS CORROSION TESTING
of the salt and water and to evaluate alternate-immersion tests using
various salt plus chromate solutions. These showed that the tendency
to cause cracking increased slightly with increasing purity of both the
salt and water. They also showed that a combination of sodium chloride
and chromic acid would cause more and quicker failures than would
salt by itself. However, tests of specimens taken in all three grain directions revealed a characteristic of the chromate-containing solutions that
seriously detracts from their usefulness.
It is not enough for an accelerated test to cause rapid failure of susceptible specimens. To be most useful it should correlate well with
natural environments and service experience in all respects and should
not be so severe as to cause failures that would not occur in normal use.
FIG. 10—Relative resistance to stress corrosion cracking of three items of 2Vzin.-diameter aluminum alloy rolled rod. The stress levels employed corresponded to
75, 50, and 25 per cent of the respective transverse yield strengths.
Alloy 7079-T6 shows appreciable susceptibility to stress corrosion
cracking only to stresses acting in the short-transverse direction. Like
most high-strength alloys, it is relatively resistant to stresses acting in
the long-transverse and longitudinal directions. In fact, only a few isolated instances have ever been recorded of service failures due to stress
in the latter directions.
Tests on the 2-in. plate showed that when a chromate-bearing solution was aggressive enough to cause 7079-T6 short-transverse specimens
to fail at low levels of stress, it also causes fractures in specimens from
the long-transverse and longitudinal directions. An example of this is
given in Table 1 for a 4.5 per cent NaCl +1.2 per cent CrO3 solution,
which was one of the solutions most able to cause fractures in shorttransverse specimens at 10 ksi. Data obtained in the acidified 5 per cent
sodium chloride intermittent spray test are also included to show that
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
357
short-transverse direction failed at a low stress (10 ksi—15 per cent yield
strength) in all environments, excepting the standard alternate-immersion
test. For highly stressed (53 ksi—75 per cent yield strength) long-transverse specimens, the industrial atmosphere (four years), the 3.5 per cent
NaCl alternate immersion (twelve weeks), and the acidified intermittent
spray (extending over two weeks) did not cause failure. Failures did occur,
however, in the northern seacoast atmosphere and in the intermittent
spray but only after relatively long exposures. On the other hand, the
salt-chromic-acid solution fractured the long-transverse specimens in
shorter time than the short-transverse specimens. Finally, none of the
environments failed the longitudinal specimens except the salt-chromicacid solution. Consequently, use of acidic chromate solutions is not recommended because of their unrealistic aggressiveness.
Comparison of Alloys
The data obtained on the rolled rod [11] are shown in Fig. 10 and
provide another comparison of alloy 7079 with aluminum-zinc-magnesium alloys of higher and lower copper content. Again the alloy with
the higher copper content, 7178 (1.6 to 2.4 per cent copper), shows
good agreement between results obtained in the alternate-immersion test
and in the atmospheres; while the newer, copper-free weldable alloy,
X70065 (0.1 per cent copper max), shows a disparity between results obtained in alternate immersion and the atmospheres, following a pattern
similar to that of 7079. As was the case with the tests of forged material,
the acidified spray test shows promise of being able to rapidly produce
results similar to those obtained in atmospheres.
General Considerations for a Laboratory Test Medium
The search for a more satisfactory accelerated test has not as yet
produced a method that is completely satisfactory. Thus far, the accelerated test showing the best correlation for 7079-T6 with atmospheric
exposure, without being too severe, has been the acidified 5 per cent
sodium chloride intermittent spray test. However, there are other factors
that should be considered in the selection of a new test method.
Applicability to Other Alloys
It obviously would be advantageous to be able to test several alloys
and alloy systems in the same test medium. Not only would this reduce
the need for a multiplicity of test methods, but, more significantly, it
permits data obtained on one alloy to be placed in their proper perspective to other alloys. Tests on nine alloys of aluminum-zinc-magnesium
5
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STRESS CORROSION TESTING
(7000 series), aluminum-magnesium (5000 series), and aluminum-copper
(2000 series) systems showed the acidified intermittent spray gave good
correlation with atmospheric exposure for all alloys with the exception
of copper-free aluminum-zinc-magnesium alloys, such as 7039-T63, at
stresses below 25 per cent yield strength. This is unfortunate, because this
is the only other major alloy group for which the standard alternate-immersion test is not effective.
Note that the acidified intermittent salt spray test has a disadvantage in
that it is relatively corrosive to most aluminum alloys. The general surface pitting may impair visual detection of fine cracks in specimens.
Also, it can result in failure in prolonged tests by mechanical rupture due
to reduction of the cross section. Because of this, a two-week test period
is recommended as the maximum period suitable for all alloys and sizes
of specimens. However, longer periods of three or four weeks may be
employed for alloys with high resistance to general corrosion or for
specimens for which the load-carrying area is large.
The boiling 6 per cent NaCl solution test is the most effective accelerated test thus far developed for the copper-free aluminum-zinc-magnesium alloys. While it is not so effective on 7079-T6 at very low stresses
as is the intermittent spray test, it is somewhat better than the standard
alternate-immersion test. Unfortunately, the major drawback to this
method is that it is not effective on aluminum-zinc-magnesium alloys
with a copper content of 1 per cent or more, nor is it effective on the
aluminum-copper and aluminum-magnesium alloy systems.
The amount of general corrosion that occurs in the boiling 6 per cent
NaCl solution test is very slight. This simplifies detection of cracks and
makes the possibility of mechanical rupture extremely remote. Some relaxation of applied stress due to creep has been noted, but, in general,
the amount has not been so large as to invalidate the test method. Because
of this relaxation, some investigators favor use of total immersion in 6
per cent NaCl at 70 F rather than at 212 F. Obviously, additional effort
is needed to develop an accelerated test that is completely satisfactory for
copper-free aluminum-zinc-magnesium alloys and also suitable for other
alloy systems.
Method of Exposure
The procedures for both the acidified intermittent spray and the boiling 6 per cent NaCl solution tests are outlined in Appendix II. The procedures for continuous spray tests are definitely established and well
known. Use of an acidified solution and intermittent operation do not
complicate these in any way. In all such spray tests, however, corrosion
is not uniform on all surfaces of a specimen; corrosion is maximum on
surfaces facing upward on which the dispersed mist can settle out. This
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
359
does not present any real problem for specimens where all surfaces are
equally susceptible. It could be of major significance on complex shapes
or assemblies if a particularly susceptible area is shielded from the mist.
The simplicity of the boiling 6 per cent NaCl solution test is one of its
principal advantages and results in a high degree of reproducibility. All
that needs to be done is to suspend the specimens in such a manner that
they do not touch the vessel or one another.
Availability and Cost of Equipment
The intermittent spray test is operated in equipment designed to comply with ASTM specification for continuous spray tests. Such equipment
is commercially available in a wide range of sizes and can be readily and
inexpensively modified for automatic cyclic operation.
The boiling 6 per cent NaCl solution test requires nothing special and
can be run with ordinary laboratory equipment. Other than a heat supply
and any inert vessel, all that is required is a sealed lid equipped with a
condenser to maintain the solution at a constant concentration. This is
fortunate because it permits both tests to be run for maximum correlation
with very little additional expenditure.
Summary
It has been established that alloy 7079-T6 rarely fails by stress corrosion cracking in 3.5 per cent NaCl alternate immersion at stresses below
20 ksi, whereas the alloy has failed by this mechanism in natural atmospheres at stresses as low as 7 ksi [2].
Various accelerated test media have been investigated in search of a
laboratory test more closely paralleling results obtained in service-type
environments represented by seacoast and industrial atmospheres. The
best correlation for 7079-T6 was achieved with an acidified 5 per cent
NaCl intermittent spray test. This test has the advantage of being short
(two weeks) and applicable to most other aluminum alloys. The exception
is that this method represents only marginal improvement over 3.5 per
cent NaCl alternate immersion for the copper-free aluminum-zinc-magnesium alloys, such as 7039-T63. It has the disadvantage of causing
severe generalized corrosion that could result in "corrosion failures" due
to excessive attack if the period of exposure is too long.
Because of its simplicity and low cost, the boiling 6 per cent NaCl
solution test is recommended as a back-up test on 7079-T6 and to
provide better correlation with copper-free aluminum-zinc-magnesium
alloys.
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STRESS CORROSION TESTING
APPENDIX I
Procedure Used for Determination of Fixture Stiffness and Effect of Cracking on Specimen Deformation
The change in average net area tensile stress due to cracking (Figs. 7 and 8)
was determined as follows:
The relationship between load and fixture stiffness was experimentally determined as indicated in Figs. \\a and b. Load deflection data were then obtained
for specimens slotted at their midlength to give reductions in area of 0, 20, 40,
60, and 80 per cent (Figs, lie and d). The stress corresponding to a given crack
depth was found by trial and error. Let P° be the force required to stress a plain
specimen to the initial stress <r° and P' the force on the specimen for a given crack
depth. Then to satisfy equilibrium requirements
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LIFKA AND SPROWLS ON ALUMINUM ALLOY
361
Let AL° be the initial change in specimen length produced by P° and AZ/ the
change in length corresponding to P'. The change in specimen length, d(AL)s,
can be found from the specimen load-deflection data (Fig. lid). To satisfy compatibility, the incremental change in specimen length must equal the incremental
deformation of the fixture or
where
From the value of P' which satisfies the above requirements, the average net
area stress is
where A' is the net area stress for the given crack depth.
APPENDIX II
Test Procedure for Acidified Intermittent Spray and Boiling 6 per cent NaCl
Solution Tests
Acidified Intermittent Spray Test [10]
A 5 per cent NaCl solution is made with salt of 99.7 per cent purity and distilled water and acidified with acetic acid to pH 3. Specimens are exposed in
cabinets designed to meet the requirements of ASTM Method of Acetic AcidSalt Spray (Fog) Testing (B 287 - 62). The recommended length of exposure is
two weeks, and specimens are inspected daily for failure. Test conditions are the
same as those required by ASTM except for the following variations:
1. Operating temperature may be increased from 95 to 120 F.
2. Specimens are intermittently sprayed in 6-hr repetitive cycles, consisting of
M-hr spray (ASTM Method B 287), 2 hr of dry-air purge, and S^-hr soak at
high relative humidity.
3. Due to the small percentage of spray time, periodic measurements of pH
and specific gravity are made only on the reservoir solution and not on condensate
specimens.
Boiling 6 per cent NaCl Solution Test
This test employs a 6 per cent NaCl solution made up from salt of 99.7 per cent
purity and distilled water. The solution is brought to a rapid boil, and the specimens are then freely suspended in the boiling solution. The usual exposure period
is 96 hr (4 days) which is occasionally extended to 7 days. The majority of failures
occur during the first few hours of test, consequently, specimens are inspected
for failure quite frequently during the initial portion of the test. A typical inspection schedule is:
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STRESS CORROSION TESTING
(c) Every hour for the next 4 hr.
(d) Three times a day thereafter
References
[1] "Metallic Materials and Elements For Flight Vehicle Structures," MIL-HDBK5, Section 3.1.1.3, Precautionary Notes on Aluminum (first appended in Aug.
1962).
[2] Sprowls, D. O. and Brown, R. H., "What Every Engineer Should Know
About Stress Corrosion of Aluminum," Metal Progress, Vol. 81, No. 4, 1962,
pp. 79-85; Vol. 81, No. 5, 1962, pp. 77-83.
[3] Dix, E. H., Jr., "Prevention of Stress-Corrosion Cracking in Service," Metal
Progress, Vol. 56, Dec. 1949, pp. 803-806.
[4] Mears, R. B., Brown, R. H., and Dix, E. H., Jr., "A Generalized Theory of the
Stress-Corrosion of Alloys," ASTM-AIME Symposium on Stress-Corrosion
Cracking of Metals, 1944, American Society for Testing and Materials, Philadelphia, 1945, pp. 323-339.
[5] Sager, G. F., Brown, R. H., and Mears, R. B., 'Tests for Determining Susceptibility to Stress-Corrosion Cracking," ASTM-AIME Symposium on
Stress-Corrosion Cracking of Metals, 1944, American Society for Testing and
Materials, Philadelphia, 1945, pp. 255-272.
[6] Nock, J. A., Jr., 'Today's Aluminum Aircraft Alloys," Transactions, Society
of Automotive Engineers, Vol. 61, 1953, pp. 209-220.
[7] Dix, E. H., Jr., Anderson, W. A., and Shumaker, M. B., "Influence of Service
Temperature on the Resistance of Wrought Al-Mg Alloys to Corrosion,"
Corrosion, Vol. 15, No. 2, 1959, pp. 55-62.
[8] Sprowls, D. O. and Rutemiller, H. C., "Susceptibility of Aluminum Alloys to
Stress Corrosion," Materials Protection, Vol. 2, No. 6, June 1963, pp. 62-65.
[9] Farmery, H. K. and Evans, U. R., 'The Stress Corrosion of Certain Aluminum Alloys," Journal of the Institute of Metals, Vol. 84, 1955-56, pp. 413422.
[10] Lifka, B. W. and Sprowls, D. O., "An Improved Exfoliation Test for Aluminum Alloys," Corrosion, Vol. 22, No. 2, 1966, pp. 7-15.
[11] Lifka, B. W. et al, "Investigation of the Stress-Corrosion Cracking of High
Strength Aluminum Alloys," Contract NAS 8-5340, two-year summary report,
1965.
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H. B. Romans1 and H. L. Craig, Jr.1
Environmental Factors Affecting the Stress
Corrosion Cracking Behavior of an
Aluminum-Zinc-Magnesium Alloy
REFERENCE: H. B. Romans and H. L. Craig, Jr., "Environmental
Factors Affecting the Stress Corrosion Cracking Behavior of an Aluminum-Zinc-Magnesium Alloy," Stress Corrosion Testing, ASTM STP 425,
Am. Soc. Testing Mats., 1967, p. 363.
ABSTRACT: The stress corrosion cracking phenomenon in an aluminumzinc-magnesium alloy is affected by the environmental factors in a different fashion than the way these factors affect other corrosion processes.
The presence of traces of moisture is sufficient for stress corrosion cracking to proceed; the absence of moisture prevents cracking. The time required to initiate stress corrosion cracking is affected mostly by the
temperature of the environment. Wide variations in sodium chloride concentration have no effect on the time-to-failure. Changes in solution pH
and contamination with dissolved copper alter the time required to produce failure, so that stress corrosion cracking is not a continuous function
of either variable. Evidence is presented which shows that the manner of
stressing a specimen and specimen configuration both affect failure time.
These anomalies may be used to understand the lack of reproducibility
of stress corrosion testing and the difficulties encountered in correlation
studies, either among different tests or the same test carried out by different
laboratories.
KEY WORDS: corrosion, stress corrosion, environmental testing, cracking
(fracturing), aluminum alloys, zinc alloys, magnesium alloys, sodium
chloride, pH, copper ions, temperature, humidity, immersion tests (corrosion)
Aluminum alloys subjected to stress corrosion tests are usually judged
on the degree of stress corrosion susceptibility by the time required
for a specimen to fail. In the past, tests have not been very reproducible
and small differences in alloy behavior were usually unnoticed due to the
larger scatter in data. There has been very little concern for this behavior
since most people were interested with gross differences in alloy systems.
However, isolated cases of service failures, combined with a large
1
Scientist and research supervisor, respectively, Department of Applied Chemistry and Mathematics, Reynolds Metals Co., Richmond, Va. Personal members
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STRESS CORROSION TESTING
amount of publicity, have brought about an ever-increasing trend toward
the writing of stress corrosion test specifications by some of the aircraft
and military agencies of the government. One such specification designated as MIL-H-6088D for alloy 7075-T73 states:
The 7075-T73 alloy shall be capable of passing the following tests without any stress corrosion cracking. Thirty-days' exposure by alternate immersion in 31A% NaCl solution (meeting the purity and pH requirements
of method 811 of Federal Test Method Standard No. 151) at room temperature while stressed in any direction to 75% of the yield strength. The
exposure cycle shall consist of 10 minute immersion in the solution and 50
minutes out of solution.
A definite period of time is set as the criteria for an acceptable product.
This means that reproducible tests are now a must in order for the specifications to be meaningful. Note that there is no specified test temperature
nor is there any mention of humidity or purity of water used to make
up the solution. In fact, most of the tests conducted today ignore these
factors.2
To determine if these and other, environmental factors affected the
time-to-failure, a series of tests was conducted on the effect of humidity,
temperature, solution contaminants, variations in stress, and specimen
configuration. Tests were also conducted between two different laboratories at controlled and uncontrolled temperature and humidity to determine the correlation of test results. Temperature has probably the
greatest single effect on the time-to-failure, but it is seldom controlled.
Contaminants from the air or heavy metal ions in solution probably have
a large effect, although no numerical value has yet been placed on them.
The effect of humidity is usually ignored, but actually this is the only
factor needed to produce stress cracking in certain alloys.
The purpose of this paper is to discuss the test environment and to
show some of the effects of testing conditions on the stress corrosion
cracking of a single alloy in the aluminum-zinc-magnesium system.
Materials
An aluminum-zinc-magnesium alloy with the following nominal composition (in per cent) was used for the experiments:
Zn
4
Mg
Fe
Si
Mn
Cr
Ti
2.8
0.4
0.3
0.3
0.2
0.1
Cu
Others
Each
0.1 max 0.05 max
Total
0.15 max
All specimens were from plate 1 to 3 in. thick.
2
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ROMANS AND CRAIG ON ALUMINUM-ZINC-MAGNESIUM ALLOY
365
Two tempers for experimental purposes were prepared with several
lots representing each temper. One temper, designated as Temper R, was
prepared to provide good resistance to stress corrosion cracking. The
other, designated as Temper S, was prepared to provide low resistance
to cracking. The typical longitudinal mechanical properties were: tensile
strength, 70 ksi; yield strength, 60 ksi; and per cent elongation in 2 in.,
10 per cent.
Procedure and Results
The method of specimen preparation prior to testing, the test solutions,
test equipment, and other details which apply to each test are given in
the procedural details in the Appendix.
Distilled Water Tests
C-ring specimens of a sample in Temper S were stressed to 35 ksi for
testing in total immersion in distilled water at 100 C. The following
steps were taken to ensure that no contamination would be present in the
test chamber:
1. The bolts used for stressing the specimens were fabricated from the
parent sample to ensure the absence of dissimilar metal contacts and to
eliminate the need to use foreign material, such as the wax usually used
for insulation.
2. The specimens and bolts were cleaned as specified in the procedural details section.
3. An all-glass reaction flask was cleaned in sulfuric acid cleaning solution, rinsed in distilled water, dipped in nitric acid, and again rinsed.
4. After cleaning, everything in the test was handled with rubber gloves
to prevent contamination from the hands.
The specimens all failed from stress corrosion cracking within 15 hr
(overnight). These failures occurred in the absence of any contamination
other than that from the glassware and dissolved metal ions from the
specimens.
The distilled water was 8 X 105 ohm-cm prior to the test, and at the
end of the test it was 2 X 105 ohm-cm. This shows a relatively high
pickup of ions of which the major portion is assumed to be aluminum.
Humidity
To determine the effect of humidity, tests were conducted on C-ring
type specimens from a sample in the S temper stressed at 35 ksi. Three
to five specimens were exposed to each of the following environments:
1. A dynamic vacuum at an average of 30 C and 0 per cent relative
humidity.
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STRESS CORROSION TESTING
3. An air-conditioned laboratory atmosphere at an average of 23 C
and 50 per cent relative humidity.
4. A normal laboratory atmosphere at an average of 30 C and 85
per cent relative humidity.
5. A controlled humidity test chamber at 38 C and 95 per cent relative humidity.
Humidity alone may play a major role in the time required to produce
stress corrosion cracking as proved by the following results. Note that
in Table 1 failures occurred in all environments except the dynamic
vacuum.
The fact that failures occurred in the desiccated atmosphere may be
explained by the ability of the oxide film to adsorb and hold moisture
TABLE 1 —Effect of humidity on time-to-failure.
Approximate
Relative
Humidity,
%
Average
Temperature,
degC
0
30
dynamic vacuum
0
30
desiccated atmosphere
50
23
air-conditioned atmosphere
85
30
laboratory atmosphere
95
38
controlled humidity test
chamber
Environment
Days-to-Failure
5 specimens did not fail in 50
days
3 specimens failed within 12
days
3 specimens failed within 4
days
3 specimens failed within 3
days
4 specimens failed within 1
day
in preference to the desiccant. This moisture can be removed under
vacuum. Much more rapid failures occur in environments where the
moisture is readily available. There appears to be a close relationship
between the per cent humidity and the time-to-failure. There is a strong
possibility that other contaminants in the atmosphere also affect time-tofailure.
Effect of Solution Concentration
Forty C-ring specimens from Temper S (Lot 1) were stressed at 35
ksi for testing in total immersion at four concentrations of sodium chloride
at 100 and 80 C. Five specimens were tested to failure at each concentration of 0.01, 0.1, 1.0, and 2.0 M at each temperature. The time-tofailure at stress of 35 ksi was not affected by the sodium chloride content
in the concentrations tested. All specimens failed within 20 min at 100 C
and within 40 min at 80 C. This is the normal failure time for this
temper for a 3.5 per cent sodium chloride solution.
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Temperature
Tests to determine the effect of temperature on stress corrosion cracking were made on four different types of specimens fabricated from a
sample of Temper S (Lot 1). The specimen description is listed below in
the procedure.
Procedure for C-Ring Jesto—Thirty C-ring specimens, similar to the
one shown in Fig. 1, from Temper S (Lot 1), as well as 20 C-ring speci-
, FIG. 1—C-ring specimen.
FIG. 2—Subsize constant strain specimen and stressing frame.
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STRESS CORROSION TESTING
mens from Temper S (Lot 2) were stressed to 35 ksi. Five C-rings at
each temperature of 100, 77, 50, 36, 8, and 3 C were tested, except
Lot 2, which was not tested at 8 and 3 C. The specimens were totally immersed for the entire period of testing.
The bolt and nut used for stressing was insulated from the specimen
with a plastic dip coating to prevent galvanic effects.
FIG. 3—Deflected-beam specimen and stressing fixture.
FIG. 4—Constant-load specimen.
Procedure for Subsize Constant Strain Tests—Thirty specimens similar to the one shown in Fig. 2 were stressed to 35 ksi. Five specimens
were tested in total immersion at each of 100, 84, 80, 40, 20, and 3 C.
The stressing frame was covered with a plastic dip coat to prevent contact with the test solution and prevent galvanic contact. Tests with other
coatings demonstrated that this dip coat did not affect the results.
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strength. (This stress is used because previous experience has shown
that this type of specimen gives erratic results if stressed below the yield
strength.)
Three specimens were tested at each temperature of 100, 70, 47, and
22 C. The specimens were tested using a standard deflected-beam holder3
as shown in Fig. 3. However, the holder for the elevated temperatures
was fabricated from aluminum pipe so it would not be distorted. The
specimens were electrically insulated from this holder with glass sleeves
over the aluminum bolts.
Procedure for Constant-Load Tension Specimens—Twelve specimens
similar to the one shown in Fig. 4 were stressed to 35 ksi. They were
FIG. 5—The effect of temper and temperature on the time-to-failure. C-rings
stressed at 35 ksi.
threaded into flexible grips and loaded by means of a lever system. Calibrated lead weights were added to obtain the desired stress. Three each
were tested at 88, 70, 50, and 35 C; all specimens were tested to failure
except those at 35 C. These were removed after 9 hr, since previous work
has shown that in this test the life is considerably extended at temperatures
below 50 C.
The results of the above tests show that the temperature is the greatest
single environmental factor governing the time-to-failure by stress corrosion. A linear function is normally obtained if the log of the time-tofailure is plotted against the reciprocal of the temperature in degrees
Kelvin. This linear function holds true down to temperatures of around
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STRESS CORROSION TESTING
8 C. An exception to this is discussed later. There are indications, as
shown in Fig. 5, that, in the vicinity of 0 C, the time-to-failure would be
infinitely long. Variations in heat treatment of the alloy change the slope
of the curve.
The data from the C-ring tests are plotted in Fig. 5. An analysis of
variance was made on the data using the procedure set forth by Ref 1.
The analysis shows that the variance at each temperature is linear when
plotted on a log scale. This is represented by the broken lines in Fig.
5. The slope of the more resistant lot of Temper R is presented to show
the differences obtained. The results from these tests also manifest some
FIG. 6—Effect of specimen configuration and method of applying stress on the
time-to-failure.
of the effects of the method of stress application and specimen configuration.
The data from the subsize constant-strain tests, the deflected-beam
test, and the constant-load tension test are plotted in Fig. 6. The C-ring
test data in Fig. 5 are repeated for comparison. All specimens except
the constant-load tension type produce straight line functions when the
log of the time-to-failure is plotted against the reciprocal of the temperature in degrees Kelvin.
Note from the plot of the data, a different slope is obtained depending
on the shape of the specimens and manner of applying stress. This
prevents direct comparison of failure times among the different specimens.
Effect of Stress Level on the Time-to-Failure
Fifteen subsize constant-strain tension specimens (Fig. 2) were
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mens from each temper were stressed at each stress level of 10, 15, 25,
35, and 45 ksi and tested by total immersion at 100 C. A Huggenberger
Model 892 extensometer was used to set the stress.
A major factor affecting the time-to-failure by stress corrosion is the
stress level at which the specimen is stressed. As would be expected, the
higher the stress, the more rapid the failure. If the log of time-to-failure is
plotted against the stress, the resultant data can be represented by two
straight lines with different slopes (Fig. 7).
At high stresses, there is little difference in time-to-failure due to differences in stress. However, at some point which may be called the critical
FIG. 7—Effect of stress on the time-to-failure at 100 C constant-strain tension
specimens.
stress, a sharp break is noted and small differences in stress cause large
differences hi time-to-failure.
Effect
of pH
Five C-ring specimens from Temper S, Lot 1, were stressed at 35 ksi
and tested at each pH level of 1.0, 2.0, 3.0, 3.6 (ten specimens were
tested at this pH), 4.0, 4.7, 5.0, 6.3, 8.2, and 10. The tests were conducted
in total immersion at 55 C. The pH was adjusted with either hydrochloric
acid or sodium hydroxide and measured with a Beckman zeromatic pH
meter.
Solution potential measurements were made at each pH level except
pH
10, using a recording potentiometer. A 0.1 N calomel cell with eight
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STRESS CORROSION TESTING
one time. The reference electrode was at room temperature; the specimens
were at 55 C.
The effect of pH on the time-to-failure varies with the pH value. A
definite area of prolonged life occurs at a pH of 3.6 (Fig. 8). On either
side of this value, the time-to-failure is much shorter. At a very alkaline
pH of 10, the life is considerably extended. The solution potentials do
not explain the stress corrosion behavior as a function of pH. The
potential at pH 1 was —0.99 v and the remainder up to pH 8 were —1.03
to-1.04v.
FIG. 8—Time-to-failure as a function of pH at 80 C. C-ring specimens stressed
at 35 ksi.
Solution Contaminants
Copper Ion Additions—Eighteen C-ring specimens of Temper S (Lot
1) were stressed at 35 ksi for testing in total immersion at 55 C at five
levels of copper. Three specimens were tested at each of 0.1, 0.5, 1.0,
3.0, 10, and 100-ppm copper, based on dilution from a stock solution
prepared by a weighed addition of copper chloride. Three specimens were
tested with no copper added to the solution. The solution was agitated
during the test to maintain a flow over the specimens and thereby reduce
the
effect by
of depletion
therights
diffusion
layer.
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ROMANS AND CRAIG ON ALUMINUM-ZINC-MAGNESIUM ALLOY
373
In short-term tests, the effect of copper ion additions is to accelerate
the rate of cracking. For instance, with no added copper in solution,
the specimens fail in an average time of 33 min, while with 0.1 to 10 ppm,
the failures occur in an average of 20 min. If amounts are present in the
range of 100 ppm, then cracking is retarded with failures occurring in
about 40 min. However, in this range, there is noticeable plating out of
copper on the surface; this apparently delays the reaction.
The alternate-immersion test solution used for routine testing was
monitored for copper ion contamination to determine the exact amount
of the contamination and its variation as a function of time.
The solutions were monitored in two tanks. One tank contained only
copper-free alloys (Tank 1); the other tank contained only copper-containing alloys (Tank 2) (Table 2).
Aluminum Ion Additions
Twelve C-ring specimens of Temper S (Lot 1) were stressed at 35 ksi
for testing in total immersion at 55 C at four levels of aluminum ion
TABLE 2— Typical copper contamination in alternate-immersion testing
(precision of measurement : ±0.02 ppm).
Solution Age
New
4 days old
Tank 1, Copper-Free Alloys
0.07
0.04
Tank 2, Copper-Bearing Alloys
0.05
0.18
concentration. Three specimens were tested at each of 1, 10, 50, and
100-ppm A1+ + + ion based on dilution from stock solution. The solution was agitated during the test to maintain a flow over the specimens.
The results of the tests were inconclusive. However, it was established
that if the aluminum ion does affect the time-to-failure, it was not measurable in the test.
Environmental Comparisons
Data Correlation Between Two Laboratories
Uncontrolled temperature and humidity alternate-immersion tests were
conducted in two different laboratories on four lots of Temper R
and two lots of Temper S stressed at 35 ksi. Nine specimens of each
lot were tested in 3.5 per cent sodium chloride alternate-immersion
tests at each laboratory. The average temperature in these tests was 32 C
with an average humidity of 80 per cent. At the time that these tests were
conducted, both laboratories conducted tests in an open laboratory area
where all types of contaminants were in the air. There was a large spread
inCopyright
the seasonal
temperature
and reserved);
humidity Wed
sinceDec
neither
laboratory
was
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374
STRESS CORROSION TESTING
air-conditioned. However, the tests were conducted to conform closely
with Federal Test Method Standard No. 151.
Controlled temperature (38 C) and humidity (45 per cent) tests were
conducted between the same two laboratories on three lots of Temper
R and three lots of Temper S with nine specimens of each lot tested
in 3.5 per cent sodium chloride alternate immersion at each laboratory.
These tests were conducted in a small insulated room of very similar construction at both laboratories. The rooms were air-conditioned and the
humidity controlled with a dehumidifier.
The main difference between the two tests was the water for solution
makeup and surface preparation of the specimens. Laboratory A used
deionized water and etched the specimens prior to testing as outlined in
TABLE 3—Data comparisons between laboratories median failure time of nine
specimens of each lot tested in 3.5% NaCl alternate immersion, days.
Temper and Lot No.
Temper R:
1
2
3
4
Temper S:
1
2
3
Laboratory B c
Laboratory A
51°
43
27
29
II 6
12
40
10"
15
<6
7
86
6
13
19
10
6
7
7
<7
<6
4
4
4
0
Uncontrolled ambient temperature and humidity, approximately 90 F and
80% RH.
6
Temperature controlled at 38 C and humidity at 45%.
c
Surface preparation of specimens in Laboratory B consisted of degreasing in
acetone only. They were not cleaned as outlined in the Appendix.
the Appendix. Laboratory B used distilled water and degreased the specimens in acetone prior to testing.
There is some correlation in the order of the ranking of lot numbers
based on time-to-failure between alternate-immersion tests conducted on
the same material in separate laboratories at uncontrolled temperature
and humidity. However, the failures in one laboratory occur three to five
times as fast as in the other laboratory. The correlation in the order of
ranking as well as the time-to-failure is greatly improved when temperature and humidity are controlled. However, somewhat faster failures are
still in Laboratory B (Table 3).
Discussion of Results
The number of factors which are shown to affect the time required to
produce stress corrosion cracking should dictate more careful controls on
future testing. The fact that moisture is the only requirement to initiate
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cracking
shows the sensitivity of this corrosion process to its environment.
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ROMANS AND CRAIG ON ALUMINUM-ZINC-MAGNESIUM ALLOY
375
Ansbacher [2]4 observed that the oxide film is capable of adsorbing
water vapor as well as atmospheric contaminants. It is known from the
experimental data and other sources [3] that the per cent moisture affects,
directly, the time-to-failure which decreases with increasing humidity. It
is reasonable to assume that air contaminants would also accelerate the
cracking process.
It has not been possible to place an exact numerical value on the
effect of humidity due mainly to the interaction between it and temperature. It is shown that temperature is one of the biggest factors and should
be the one most closely controlled. A linear function is obtained when
the log of the time-to-failure is plotted against the reciprocal of the
temperature in degrees Kelvin. This holds true for all the specimens used
except those stressed by loading with a dead weight.
The stress level at which a specimen is tested determines to a large
degree the expected life. High stresses produce rapid failures with little
differences in the time required. However, once a critical level is reached,
the life is greatly extended by small decreases in stress.
A plot of the stress versus life can be represented by two straight lines
with different slopes. The critical stress is referred to as being at the
intersection of the two lines. Brenner and Gruhl [4] state that, "Below
the kink in the stress/life curve the life is practically unlimited." However, our data shows that, for the alloy tested, failures do occur below the
break in the curve. There is a possibility of a second break below which
failures would not occur even with this alloy. The data are sometimes
represented by a curve in which the sharp break is eliminated [5].
It has been reported in the literature [6] that the failure time of
aluminum-magnesium alloys is shortened by increasing the sodium
chloride concentration. This is not the case with the aluminum-zincmagnesium alloy system which is insensitive to the effect of concentration. The pH affects the time-to-failure in a somewhat radical manner. A
maximum life occurs at a pH of 3.6 in the range from pH 1 to pH 8.
Investigations are underway to determine the reason behind this phenomenon.
Mattsson [7] shows a similar relationship with brass but as a mirror
image of the plot in Fig. 1. He shows a maximum life at a pH of 7.8 in
the range 11.2 to 4.7.
Copper ions in elevated temperature tests reduce the time required to
obtain cracking. It is postulated by the authors that their role is to increase
the potential difference between the cathodic and anodic sites which at
elevated temperatures promotes cracking instead of pitting. Bell and
Campbell [8] and Godard [9] found that copper in amounts as low as
0.02 ppm have a large effect in stimulating pitting corrosion.
4
The italic
in (all
brackets
to theWed
list Dec
of references
appended
to this
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376
STRESS CORROSION TESTING
The test results show that correlation of tests between laboratories is
greatly improved by controlling temperature and humidity. It is predicted
that good correlation can be obtained if the tests are controlled to a high
degree of accuracy.
Conclusions
Carefully prepared tests conducted in distilled water and in a desiccated
atmosphere show that the only requirement for the initiation and propagation of stress cracks is the presence of moisture. The moisture does not
have to be added intentionally. The oxide film is capable of entrapping
or adsorbing sufficient moisture to allow stress corrosion cracking to
occur. The surface moisture can be removed under a dynamic vacuum.
This is evidenced by the fact that stress corrosion failures do not occur
under this condition.
The experimental data show that the per cent moisture affects the
time-to-failure which decreases with increasing humidity. It was not
possible to place an exact numerical value on the effect of humidity due
to its interaction with temperature.
Temperature is the biggest factor controlling the time required for
stress corrosion cracking to occur. Extremely rapid stress corrosion
failures occurring in minutes can be obtained at 100 C, while at 20 C the
time required may be several days. There are indications that at 0 C,
the time-to-failure would be infinitely long. Between 100 and 30 C,
the log of the time-to-failure is linear if plotted against the reciprocal of
the absolute temperature. This holds true for all specimens and configurations tested except the constant load which has a linear function only
between 100 and 50 C. The fact that this test method does not produce
linear data makes it undesirable from the standpoint of extrapolating to
lower temperatures.
The other three specimen configurations produce linear data but with
different slopes. This will prevent direct comparisons between the
different types of specimens. If the metallurgical condition of the alloy
is changed, it will also change the slope of the curve. It is postulated that
the reason for this behavior is that the fabrication practice can form a
grain boundary precipitate which is much more active at elevated temperature than ambient temperature tests would indicate.
The stress level at which a specimen is tested determines to a large
degree the expected life. A stress-life curve can be represented by two
intersecting straight lines. At high stresses, rapid failures are obtained
and there is very little difference in the time required to produce stress
corrosion failure as the unit stress is lowered. However, once a critical
stress level is reached, there is a sharp break in the curve giving a large
increase in life with small decreases in stress.
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ROMANS AND CRAIG ON ALUMINUM-ZINC-MAGNESIUM ALLOY
377
sensitive to changes in the sodium chloride concentration in amounts of
0.01 to 2 M. The pH of the sodium chloride test solution affects the timeto-failure in a somewhat radical manner. There is an increase in life
at a pH of 3.6 in the range from pH 1 to pH 8. At the present time, there
is no plausible explanation for this behavior.
Copper ions in elevated temperature tests accelerate failures. It is
postulated that their role is to increase the potential difference between
the cathodic and anodic sites, which at elevated temperatures promotes
cracking instead of pitting.
The reproducibility of tests on the same materials between laboratories
is poor. This indicates the need for better controls on environmental
factors than is presently being used. The fact that the correlation of data
between two laboratories was greatly improved by controlling only the
temperature and humidity indicates that additional controls on water
purity, grade salt, pH, and other factors should be beneficial.
A cknowledgment
The assistance of F. E. Loftin and Z. L. Vance in obtaining the data
for this presentation is gratefully acknowledged. The authors gratefully
acknowledge the permission of Reynolds Metals Co. to publish this paper.
APPENDIX
Procedural Details
All of the specimens were prepared for testing according to the following
procedure:
1. The specimens were machined from the sample in such a way as to
orient the short-transverse direction of the plate parallel to the applied stress.5
The four types may be listed as the C-ring 0.750 in. outside diameter by
0.060-in. wall (Fig. 1); the constant-strain subsize tension specimen 1.875 in.
long by 0.125-in. reduced section (Fig. 2); the deflected-beam specimen 3 in.
long by 0.064 in. thick by plate gage (Fig. 3); and the constant-load tension
specimen 3 in. long by 0.125-in. reduced section (Fig. 4). The largest portion
of this work used a C-ring specimen.
Exceptional care was taken to ensure the smoothest possible machined
surface. No abrasive was ever used due to the danger of smearing the surface.
2. The specimens were degreased in acetone to remove residual oil and
then etched for 30 to 40 sec in 5 per cent sodium hydroxide solution at 77 C.
This was followed by a dip in concentrated nitric acid at 25 C to remove the
smut.
5
Prestley, J. S., "The Effects of Specimen Orientation on Resistance to Stress
Corrosion of Aluminum Alloys—A Mathematical Model," presented at the sixtyninth ASTM Annual Meeting, Atlantic City, N. J., 26 June-1 July, 1966, unCopyrightwork.
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STRESS CORROSION TESTING
After the specimens were stressed to the desired level, all portions of the
fixture were waxed to prevent any galvanic effect or solution contamination
from the fixture. The specimens were always cleaned and stressed directly
before testing to minimize surface contamination by laboratory atmospheres.
Test Solutions
All the test solutions used for the experiments reported in this paper
were prepared from deionized water with an average resistance of 200,000
ohm-cm, unless otherwise noted. The sodium chloride was USP grade
added to obtain a concentration of 3.5 per cent. The solution was aged
for seven days prior to use to allow stabilization of pH. The solution is about
pH 4 when first prepared and requires about five days for the CO2 content
to reach equilibrium with atmospheric conditions. After five days, the pH will
adjust to 6 to 6.5. The solution then meets the requirements of Federal Test
Method Standard No. 151.
Test Equipment
Unless otherwise stated, the alternate-immersion tests reported have
the temperature controlled at 28 ± 1 C. The humidity was controlled at 45 ± 2 per cent. The solution containers were 50-gal glass aquaria
with Plexiglas immersion racks. The specimens were exposed in a single
layer to prevent differences in drying rates or drainage onto a lower layer.
All-plastic immersion pumps were used to circulate the solution to insure a
homogeneous condition.
When other alloys were tested, the ones containing deliberate additions of
copper were placed in a separate container from the copper-free alloys.
The total immersion tests were conducted in glass containers immersed in
a water bath in which the temperature was controlled to ±1 C. A hot plate
was used for the total immersion specimens tested at boiling temperature.
References
[1] Acton, F. S., Analysis of Straight Line Data, Wiley, New York.
[2] Ansbacher, F., "The Effects of Water Vapor on the Electrical Properties of
Aluminum," Nature, Vol. 171, 24 January, 1953.
[3] Champion, F. A., 'The Interactions of Static Stress and Corrosion with
Aluminum Alloys," Journal of the Institute of Metals, Vol. 83, 1954-55.
[4] Brenner, P. and Gruhl, W., "Stress Corrosion Testing of Al-Zn-Mg Under Constant Tension and Bending," Zeitschrift fuer Metallkunde, Vol. 52, No. 10,
1961, pp. 599-607.
[5] Symposium on Corrosion Fundamentals, University of Tennessee Corrosion
Conference, University of Tennessee Press, March 1965.
[6] Ferryman, E. C. and Hadden, S. E., Journal of the Institute of Metals, Vol.
77, 1950, p. 207.
[7] Mattsson, E., "Stress Corrosion in Brass Considered Against the Background
of Potential/pH Diagrams," Electrochemical Acta, Vol. 3, 1961.
[8] Bell, W. A. and Campbell, H. S., "Aluminum in Fresh Waters," British NonFerrous Metals Research Association Laboratories, London N.W. 1, 8 April,
1961.
[9] Godard, H. P., "The Corrosion Behavior of Aluminum," Corrosion, Vol. 11,
No. 12, Dec. 1955, pp. 542T-552T.
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DISCUSSION ON ALUMINUM-ZINC-MAGNESIUM ALLOY
379
DISCUSSION
D. O. Sprowls1 (written discussion)—It is generally recognized that the
specimen configuration and the manner of stressing affect both the
probability of developing a stress corrosion crack and the rate of crack
propagation in susceptible alloys. The data given in Fig. 6, however,
provide an unusual opportunity of making a direct comparison of three
different procedures, namely: (a) direct tension, constant load; (b) direct
tension, "constant strain;" and (c) bending, constant strain (C-ring). (The
deflected-beam specimen is excluded from this comparison because the
beams were stressed at a much higher level.) One would expect from
the distribution of stress in the specimens that the tension specimens,
especially when placed under constant load, would fail more rapidly than
the C-ring specimen. Brenner and Gruhl2 have, in fact, shown that
aluminum-zinc-magnesium alloy sheet tension specimens stressed in
tension failed more rapidly and at a lower stress level than similar specimens stressed in bending. This same comparison was shown for weldments by Shumaker et al.3
At the Alcoa Research Laboratories we regularly use large numbers
of C-rings and 0.125-in.-diameter by 2-in.-long tensile bars stressed by
"constant strain".4 Usually one specimen or the other is used, the choice
depending upon the dimensions of the product to be tested. However, in
several tests of experimental heat treatments for -T6 type 7039 core,
rolled plates in the 2 to 4-in. range of thickness were stress corrosion
tested in the short-transverse direction using both types of specimen. As
a general rule, the tensile bars cracked sooner than the C-rings, and in
cases of borderline susceptibility, the percentage of failures was higher
for the tensile bars. A sample comparison is shown in Fig. 9.
The comparisons shown by the authors in Fig. 6 are surprising in
that: (1) the C-rings failed more rapidly than the tension specimens
stressed in tension by "constant strain," and (2) at the lower temperatures
the tension specimens stressed by constant load presumably would have
failed after longer exposures than the C-rings. The expression "constant
strain" is used with quotation marks in the case of the tensile bar, because with this particular tensile bar and stressing frame the specimen is
1
Assistant chief, Chemical Metallurgy Div., Alcoa Research Laboratories, New
Kensington, Pa.
2
Brenner, Paul and Gruhl, Wolfgang, "Stress-Corrosion Cracking Tests of AlZn-Mg 3 Under Constant Tensile and Bending Strain," Zeitschrift fuer Metallkunde,
Vol. 52, No. 10, 1961, pp. 599-607.
3
See p. 317.
4
Sprowls, D. O. and Brown, R. H., "What Every Engineer Should Know About
Stress Corrosion of Aluminum," Metal Progress, Vol. 81, April and May 1962.
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380
STRESS CORROSION TESTING
not loaded purely in constant strain. An analysis of the loading characteristics of this system5 has shown that the strain energy developed in the
specimen and the stressing frame during loading is sufficient that when
localized cracking of the specimen occurs, the average tensile stress on
the net section through the crack increases in almost the same manner
as in the case of dead-weight loading. Therefore, our experience at Alcoa
Research Laboratories indicates that the behavior of this "constant
strain" tensile bar and the constant load tension specimen generally are
FIG. 9—Comparison of stress corrosion test specimens, short-transverse tests
of 7039 alloy plate.
about the same and both constitute a slightly more severe test than the
C-ring specimen.
It should be pointed out that the "critical stress" identified by the
authors in Fig. 7 should not be confused with the stress corrosion "thresholds" referred to by many investigators. The term "threshold stress" is
generally used to denote the highest sustained tensile stress that did not
cause failure by stress corrosion cracking under the conditions of the
test, and this stress is not necessarily related to the shape of the stress
specimen life curve. The threshold stress so determined will, of course, be
related to the specific test conditions. Threshold stress and critical
stresses as used by various investigators are discussed in more detail in
6
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DISCUSSION ON ALUMINUM-ZINC-MAGNESIUM
ALLOY
381
the report of Task Group 3 of ASTM Subcommittee VI of Committee
G-l on p. 292 of this volume.
H. B. Romans and H. L. Craig, Jr. (authors)—In general, the authors
can agree with the contention that the subsize constant-strain specimen
provides faster stress corrosion failures than the C-ring. However, we
have shown that this does not hold true for all materials under all test
conditions. We believe this behavior should show that making a flat
statement about anything in stress corrosion testing is really hard to
justify.
A. R. C. Westwood6 (written discussion)—Would you comment on
the variation in slope of the log tp versus (1/T) curve with temper for
a given alloy composition? If an activated process is controlling, this suggests that some change in mechanism is occurring.
Messrs. Romans and Craig—The explanation for the slope variations
may be explained by considering the potential differences which can exist
between the grain boundary and the grain. If the volume of the precipitate
particles in the matrix were equal to the volume in the grain boundary,
then the potential difference would be small. However, if no matrix
precipitate were present and there were a large volume of grain boundary
precipitate, then there would be a very large potential difference. This
would result in a highly anodic grain boundary.
The specimens we were working with were between these two extreme
cases. At the high temperatures the grain boundary of all of the specimens
was attacked very rapidly. The temperature factor was great enough to
overcome some of the potential difference due to precipitate distribution.
This resulted in a narrow spread among specimens in time-to-failure. At
low temperatures the potential difference between matrix and boundary
was the controlling factor. This gave a wider spread in stress corrosion
cracking results due to small differences in precipitate distribution resulting in different slopes for each specimen.
9
Associate director, Research Institute for Advanced Studies, Martin Co., Baltimore, Md.
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in connection with its work of promoting knowledge
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specifications and tests for materials. Much of the data
result from the voluntary contributions of many of the
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ASTM technical committees, or compilations of data
developed in special Society groups with many organizations cooperating. A list of ASTM publications and
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