ASM Metals HandBook Vol. 8 - Mechanical Testing and Evaluation
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*
The section “Creep Properties” was adapted from Ref 1
with additional content by R.W. Hayes, Metals
Technology, Inc. The remaining sections of this article were adapted from “Creep, Stress-Rupture, and Stress-
Relaxation Testing” in Volume 8 of the 9th Edition Metals Handbook. The section “Constant-Stress Testing”
was updated by N.L. Carroll, Applied Test Systems Inc.
References cited in this section
36. E.N. Da C. Andrade, On the Viscous Flow in Metals and Allied Phenomena, Proc. R. Soc. (London)
Vol 84, Series A, 1910–11, p 1
37. E.N. Da C. Andrade and G. Chalmers, The Resistivity of Polycrystalline Wires in Relation to Plastic
Deformation and the Mechanism of Plastic Flow, Proc. R. Soc. (London) Vol 138, Series A, 1932, p
348
38. F. Garofalo, O. Richmond, and W.F. Domis, Design of Apparatus for Constant-Stress or Constant-Load,
J. Basic Eng., Vol 84, 1962, p 287–293
39. M.F. Holmes and P.J. Wray, A Constant Stress Apparatus Suitable for Use with Small Forces, J. Phys.
E, Sci. Instrum., Vol 3, 1970
40. L.M.T. Hopkin, A Simple Constant Stress Apparatus for Creep Testing, Proc. Phys. Soc. B, Vol 63,
1950, p 346–349
41. C.H Carter, Jr., C.A. Stone, and R.F. Davis, High-Temperature Multi-atmosphere Constant Stress
Compression Creep Apparatus, Rev. Sci. Instrum., Vol 51 (No. 10), Oct 1980, p 1352–1357
42. R.L. Fullman, R.P. Carreker, and J.C. Fisher, Simple Devices for Approximating Constant Stress
During Tensile Creep Tests, J. Met., Vol 5, 1953, p 657
Creep and Creep-Rupture Testing*
Computer-Controlled Constant-Stress Testing
As previously noted, precision-machined cam-lever systems are designed to mechanically adjust the load with
respect to the change in length of a specimen with a fixed initial gage length and overall length. Although the
disadvantage of this type of testing system is the physical constraint on the initial geometry of the specimens to
be tested, the lever arm lends a significant mechanical advantage to the testing machine. Moreover, by using the
increased compliance, reduced force, and increased travel on the weight pan end of the lever system (e.g., with
a ratio of 10:1 or 20:1), a significant increase in control parameters and sensitivity results in higher precision
testing.
The use of a computer control system for constant-stress testing can overcome the constraints on specimen
geometry imposed by the cam-lever and similar systems. The computer control system can be programmed to
accommodate a wide variety of specimens as well as provide control over test parameters such as rate of load
application and removal, step-loading, cyclic loading, test time, furnace or oven temperature, and so on. The
computer control system can also be used for data acquisition and analysis and for emergency test situations.
Dynamically Adjusted Constant-Stress Creep Test Machine. A novel computer-controlled creep testing system,
developed at the University of California, Davis, is called the dynamically adjusted constant-stress creep
machine (DACSCM) (Ref 43). This system integrates the positive and beneficial aspects of the lever-arm test
system with a computer-control system that modulates load control apparatus to perform constant-stress creep
experiments. The testing system, at the time of writing, accumulated well over 1200 experimental hours of
usage by various investigators. The longest test lasted for more than 274 h at a single constant stress rate.
The specific algorithm that relates load and tensile elongation to cross-sectional area is based on the assumption
of volume constancy within the deforming gage length. The following expression can be written:
(Eq 15)
where for the i
th
data point, σ
0
is the desired stress, A
0
is the initial cross-sectional area, l
0
is the initial gage
length, and x
i
is the total accumulated specimen displacement. The expression will be assumed to be valid
throughout the entire deformation process until the onset of necking (i.e., localized straining) within the
specimen takes place. Also, since a significant amount of the experimental creep work is carried out at strain
levels that are much less than those typically required to initiate necking, this assumption can be considered
reasonable and acceptable.
Typical Computer-Controlled Constant-Stress Testing System. A typical constant-stress testing system with a
lever arm and computer control and data acquisition is shown in Fig. 18. With proper programming and
selection of lever arm ratios, a system of this type provides precision control and also data acquisition and
analysis. The versatility of these systems makes it feasible to perform various test functions, such as
creep/stress rupture testing, stress relaxation testing, constant-stress testing, and cyclic load or cyclic strain
testing. This type of system can be easily converted from one test mode to any of the other modes. In addition,
compression testing can be performed in any of the test modes by using an appropriate compression test fixture.
Testing systems such as the one shown in Fig. 18, are available with dual ratio lever arms (e.g., 20:1 and 3:1),
which extend the range and provide optimum control and sensitivity.
Fig. 18 Typical computer-controlled constant-stress test system with 90-kN (20,000 lbf) load capacity
and furnace system.
In order to achieve precise test results, measurement is required on the specimen gage length and sometimes on
the specimen diameter(s) as well. A typical extensometer is an averaging extensometer with dual high-precision
capacitance transducers. Extensometers are also available to gage specimen diameter(s) change during testing,
using similar transducers.
Footnote
*
The section “Creep Properties” was adapted from Ref 1
with additional content by R.W. Hayes, Metals
Technology, Inc. The remaining sections of this article were adapted from “Creep, Stress-Rupture, and Stress-
Relaxation Testing” in Volume 8 of the 9th Edition Metals Handbook. The section “Constant-Stress Testing”
was updated by N.L. Carroll, Applied Test Systems Inc.
Reference cited in this section
43. R.B. Grishaber, R.H. Lu, D.M. Farkas, and A.K. Mukherjee, A Novel Computer Controlled Constant
Stress Lever Arm Creep Testing Machine, Rev. Sci. Instrum., Vol 68 (No. 7), July 1997, p 2812–2817
Creep and Creep-Rupture Testing*
Stress Relaxation
Traditional creep testing to develop descriptions of strain rate, stress, and temperature behavior can be time
intensive and expensive, involving many creep test stands, specimens, and thousands of hours of testing. Stress
relaxation offers the potential to eliminate this difficulty by producing strain rate/stress data over a wide range
of rates from a single specimen (see the article “Stress Relaxation Testing” in this Volume). This information is
developed when the elastic strain of a specimen extended (or compressed) to a certain constant length is
converted to plastic strain. As the specimen slowly deforms, the load required to maintain the constant length is
reduced, hence the term stress relaxation test. In addition, this type of experiment simulates the real
engineering problem of the long-term loosening of tightened bolts and other fasteners.
In its simplest form, a stress relaxation test involves loading (straining) a specimen to some predetermined load
(strain), fixing the position of the specimen (halting the crosshead motion in a universal test machine, for
example), and measuring the load as a function of time. With knowledge of the elastic modulus of the specimen
and the stiffness of the testing machine, the load-time data can be converted to stress/strain-rate data. This
information can then be used to determine the stress components and activation energies of creep for
deformation.
In stress relaxation testing, a material is loaded with an initial stress level (σ
0
, in the idealized graphic
representation in Fig. 19). Depending on the application, the initial stress may be below, at, or above the yield
stress on loading. Whether or not there is permanent plastic strain on loading, there is an elastic strain that, in
linear elastic materials, is proportional to the stress through the appropriate temperature-dependent modulus.
Most reported data concern stress relaxation in materials initially stressed below their yield strength.
Conversely, if initial yielding occurs, an initial permanent or plastic strain develops (Fig. 20). The elastic strain
remaining, EC, is directly proportional to the remaining stress, DC in Fig. 20, so that in effect, after the initial
yielding, stress relaxation is a time-dependent conversion of elastic to plastic irrecoverable strain. As time
progresses, line ED in Fig. 20 shifts to the right and the relaxation of stress (Δσ) is defined as equivalent to the
stress required to produce an equal amount of elastic deformation, where (assuming that the modulus of
elasticity is constant and that the change in cross-sectional area is negligible during the test):
Δσ = σ
0
- σ
t
= Eε
p
where σ
t
is the stress at time t, E is the elastic modulus, and ε
p
is the plastic strain at time t. The quantity ε
p
is
equal to the residual strain after load removal.
Fig. 19 Idealized concept of stress relaxation. (a) Constant strain loading in tension. (b) Stress-relaxation
curves in tension
Fig. 20 Stress-strain diagram for determining relaxation in stress
One major drawback of many stress relaxation tests is the number of demands placed on the experimental
equipment. The load-measuring system must be capable of making accurate measurements of very small
changes as a function of time. The effect of the loading rate on the relaxation rate should be evident. In
addition, room temperature, as well as specimen temperature, must be precisely controlled throughout the
experiment. This is critical, because even small fluctuations in temperature will produce thermal expansion
effects that can mask changes due to relaxation. Also, the calculated stress exponents and activation energies
may not be the same as those determined from creep testing. The processes that produce plastic flow could be
different for these two situations. Only comparison of the results from both types of testing can detect
equivalent behavior.
Footnote
*
The section “Creep Properties” was adapted from Ref 1
with additional content by R.W. Hayes, Metals
Technology, Inc. The remaining sections of this article were adapted from “Creep, Stress-Rupture, and Stress-
Relaxation Testing” in Volume 8 of the 9th Edition Metals Handbook. The section “Constant-Stress Testing”
was updated by N.L. Carroll, Applied Test Systems Inc.
Creep and Creep-Rupture Testing*
References
1. R. Viswanathan, Damage Mechanisms and Life Assessment of High-Temperature Components, ASM
International, 1989, p 59–69
2. F. Garofalo, Fundamentals of Creep and Creep Rupture in Metals, M. Fine, J. Weertman, and J.R.
Weertman, Ed., MacMillan Series in Materials Science, The MacMillan Co., New York, 1965
3. S.K. Mitra and D. McLean, Work Hardening and Recovery in Creep, Proc. R. Soc., Vol 295, 1966, p
288–299
4. G. Bernasconi and G. Piatti, Ed., Creep of Engineering Materials and Structures, Applied Science
Publishers, London, 1979
5. O.D. Sherby, Acta Metallurgica, Vol 10, 1962, p 135–147
6. O.D. Sherby and A.K. Miller, ASME Journal of Materials and Technology, Vol 101, 1979, p 387
7. F.H. Norton, The Creep of Steel at High Temperatures, McGraw-Hill, 1929
8. R.W. Bailey, Creep of Steel under Simple and Compound Stress, Engineering, Vol 121, 1930, p 265
9. O.D. Sherby and P.M. Burke, Mechanical Behavior of Crystalline Solids at Elevated Temperatures,
Prog. Mater. Sci., Vol 13, 1967, p 325
10. R. Viswanathan, The Effect of Stress and Temperature on the Creep and Rupture Behavior of a 1.25 pct.
Chromium -0.5 pct. Molybdenum Steel, Metall. Trans. A, Vol 8A, 1977, p 877–883
11. B.J. Cane, “The Process Controlling Creep and Creep Fracture of 2 Cr-1Mo Steel,” CEGB Report
RD/LR 1979, Central Electricity Generating Board Research Laboratories, Leatherhead, England, 1979
12. M.S. Shammas, “Estimating the Remaining Life of Boiler Pressure Parts,” EPRI Final Report on RP
2253-1, Vol 4, Electric Power Research Institute, Palo Alto CA, 1987
13. M.C. Askins et al., EPRI Final Report on RP 2253-1, Vol 5, Electric Power Research Institute, Palo
Alto, CA, 1987
14. I.R. McLauchlin, in Creep Strength in Steel and High Temperature Alloys, The Metals Society, London,
1974, p 8, 71, 86, 106, 122, 217, 230
15. K.E. Amin and J.E. Dorn, Acta Metall., Vol 17, 1969, p 1429
16. V. Foldyna, A. Jakobova, T. Prnka, and J. Sabotka, Creep Strength in Steels and High Temperature
Alloys, The Metals Society, London, 1974, p 130
17. K.R. Williams and B. Wilshire, Metall. Sci. J., Vol 7, 1973, p 176
18. C.N. Aneceuist, R. Gasca-Neri, and W.D. Nix, A Phenomenological Theory of Steady State Creep
Based on Average Internal and Effective Stress, Acta Metall., Vol 18, June 1970, p 663–671
19. A.K. Mukherjee, J.E. Bird, and J.E. Dorn, ASM Trans. Quart., Vol 62, 1969, p 155
20. S.L. Robinson and O.D. Sherby, Acta Metall., Vol 17, 1969, p 109
21. H. Luthy, A.K. Miller, and O.D. Sherby, Acta Metall., Vol 28, 1980, p 169
22. J.R. Spingarn, D.M. Barnett, and W.D. Nix, Acta Metall., Vol 27, 1979, p 1549
23. W.D. Nix, J.C. Gibeling, and D.A. Hughes, Metall. Trans. A., Vol 16, 1985, p 2215
24. M. Biberger and J.C. Gibeling, Acta Metall. Mater., Vol 43, 1995, p 3247
25. U.F. Kocks, A.S. Argon, and M.F. Ashby, Thermodynamics and Kinetics of Slip, Pergamon Press,
Oxford, 1975, p 141
26. G.S. Nakayama and J.C. Gibeling, Acta Metall., Vol 38, 1990, p 2023
27. G.S. Nakayama and J.C. Gibeling, Scr. Metall. Mater., Vol 24, 1990, p 2031
28. G.S. Nakayama and J.C. Gibeling, Modeling the Deformation of Crystalline Solids, T.C. Lowe et al.,
Ed., TMS, Warrendale, PA, 1991, p 589
29. J.B. Conway, Stress Rupture Parameters: Origin, Calculations and Use, Gordon and Breach, 1969
30. S.S. Manson, Time-Temperature Parameters—A Re-Evaluation of Some New Approaches, ASM
Publication D-8-100, American Society for Metals, 1968, p 1–115
31. S.S. Manson and C.R. Ensign, A Quarter Century of Progress in the Development of Correlation and
Extrapolation Methods for Creep Rupture Data, ASME J. Eng. Mater. Technol., Vol 101, 1979, p 317–
325
32. G.P. Tilly and G.F. Harrison, Interpretation of Tensile and Compressive Creep Behavior of Two Nickel
Alloys, J. Strain Anal., Vol 8, 1973, p 124–131
33. H.H. Gray, “Transverse Tensile and Stress Rupture Properties of γ/γ′ - δ Directionally Solidified
Eutectic,” NASA TMX-73451, 1979
34. J.M. Thomas and J.F. Carlson, Errors in Deformation Measurements for Elevated Temperature Tension
Tests, ASTM Bullein, May 1955, p 47–55
35. A. Fox, Lecture Notes, ASM Continuing Education Seminar, Creep and Stress Relaxation Testing,
March 1980
36. E.N. Da C. Andrade, On the Viscous Flow in Metals and Allied Phenomena, Proc. R. Soc. (London)
Vol 84, Series A, 1910–11, p 1
37. E.N. Da C. Andrade and G. Chalmers, The Resistivity of Polycrystalline Wires in Relation to Plastic
Deformation and the Mechanism of Plastic Flow, Proc. R. Soc. (London) Vol 138, Series A, 1932, p
348
38. F. Garofalo, O. Richmond, and W.F. Domis, Design of Apparatus for Constant-Stress or Constant-Load,
J. Basic Eng., Vol 84, 1962, p 287–293
39. M.F. Holmes and P.J. Wray, A Constant Stress Apparatus Suitable for Use with Small Forces, J. Phys.
E, Sci. Instrum., Vol 3, 1970
40. L.M.T. Hopkin, A Simple Constant Stress Apparatus for Creep Testing, Proc. Phys. Soc. B, Vol 63,
1950, p 346–349
41. C.H Carter, Jr., C.A. Stone, and R.F. Davis, High-Temperature Multi-atmosphere Constant Stress
Compression Creep Apparatus, Rev. Sci. Instrum., Vol 51 (No. 10), Oct 1980, p 1352–1357
42. R.L. Fullman, R.P. Carreker, and J.C. Fisher, Simple Devices for Approximating Constant Stress
During Tensile Creep Tests, J. Met., Vol 5, 1953, p 657
43. R.B. Grishaber, R.H. Lu, D.M. Farkas, and A.K. Mukherjee, A Novel Computer Controlled Constant
Stress Lever Arm Creep Testing Machine, Rev. Sci. Instrum., Vol 68 (No. 7), July 1997, p 2812–2817
Footnote
*
The section “Creep Properties” was adapted from Ref 1
with additional content by R.W. Hayes, Metals
Technology, Inc. The remaining sections of this article were adapted from “Creep, Stress-Rupture, and Stress-
Relaxation Testing” in Volume 8 of the 9th Edition Metals Handbook. The section “Constant-Stress Testing”
was updated by N.L. Carroll, Applied Test Systems Inc.
Assessment and Use of Creep-Rupture Properties
Howard R. Voorhees, Materials Technology Corporation Martin Prager, Materials Properties Council
Introduction
USE OF CREEP-RUPTURE PROPERTIES to determine allowable stresses for service parts has evolved with
experience, although guidelines for use differ among specifications.
Pioneering efforts by ASME (Ref 1 and earlier versions) to improve safety in use of metals for long-time
operation at temperatures in the creep range set allowable stresses by considering the following criteria among
others:
• 100% of the stress to produce a creep rate of 0.01%/1000 h, based on a conservative average of reported
tests as evaluated by an ASME committee. In assessing data, greater weight is given to those tests run
for longer times.
• 63% of the average stress or 80% of the minimum stress reported to produce rupture at the end of
100,000 h, as determined from available extrapolated data
These criteria continue to find wide acceptance.
For most commercial steels and alloys, available raw data come from tests with durations ranging from a few
hundred to several thousand hours. Tests seldom run longer than 10,000 h, and durations of 100,000 h are rare.
Measured secondary creep rates as low as 0.01%/1000 h are also unusual. Allowable stresses recommended in
existing specifications usually are derived from extrapolations. Considerable scatter in test results may be
observed even for a given heat of an alloy, so interpolated creep and rupture strengths are not precisely
established.
Although measurement and application of creep-rupture properties is often imprecise, general trends are
evident. Methods for assessing creep-rupture properties (including nonclassical creep behavior), common
interpolation and extrapolation procedures, and properties estimation based on insufficient data are discussed in
this article.
Reference cited in this section
1. Basis for Estimating Stress Values, Paragraph A-150, Rules for Construction of Power Boilers, ASME
Boiler and Pressure Vessel Code, Section I, American Society of Mechanical Engineers, New York,
1983
Assessment and Use of Creep-Rupture Properties
Howard R. Voorhees, Materials Technology Corporation Martin Prager, Materials Properties Council
Scatter of Creep-Rupture Test Data
Reliable creep-rupture property measurements require that the test specimens be representative of the material
to be used in service, preferably in the product form and condition of intended service. Testing results may vary
significantly with sampling procedures. For example, heavy sections may exhibit variations in strength level
with depth after normalizing or quenching and tempering treatments. A common practice is to take samples
midway between the center and the surface of the specimen.
Similar variations in strength can occur for fully annealed plate or bar that has been cold straightened.
Reannealing may be required before valid testing results can be obtained. For some materials (e.g., stainless
steel castings several inches thick, or alloys subjected to elevated temperature after critical prior plastic
deformation), the grains may become so coarse that a specimen cross section contains only one or several
grains. Local creep rates may then vary considerably along the specimen gage length according to the
orientation of the individual large grains with respect to the loading direction.
Sampling direction seldom affects creep and rupture properties for materials with uniform single-phase
structures of small equiaxed grains. For such materials, specimen size also exhibits minimal influence, except
for the greater relative importance of surface oxidation and similar effects in small specimens. Due to
solidification and processing conditions, preferred orientation of secondary phases or grains can alter test
results. In critical situations, the loading direction in the test specimen should parallel the major stress direction
expected under service loading. For rolled materials, ASTM E 139 (Ref 2) recommends that specimens be
taken in the direction of rolling, unless otherwise specified.
The presence of large discrete particles of lower ductility and strength (e.g., graphite flakes in gray cast iron, or
large glassy inclusions in steel) may significantly lower the sound cross section of a small specimen, but have a
lesser influence in a larger specimen. Specimen size may have the opposite effect if only a few scattered low-
strength particles are present; consequently, a specimen small enough could be free of these large-scale areas of
weakness. More subtle variations in local creep and rupture strengths arise at grain boundaries, precipitates,
voids, composition gradients, and other regions of microscale nonuniformity.
Specimen Size. For many steels, the influence of specimen size is minimal compared to other variables. During
testing of a low-alloy steel (ASTM A 193, grade B16) and a high-alloy steel (ASTM A 453, grade 660) using
five different specimen diameters ranging from 5 to 84 mm (0.2–3.3 in.), rupture strength appeared to be
independent of size for unnotched specimens (Ref 3). Variation in rupture time with size was erratic for notched
specimens, but the largest size had about three quarters of the strength of the smallest geometrically similar
specimen.
Scatter from Heterogeneities. Heterogeneities in a low-alloy steel of commercial quality have been found to
cause much of the typical 30% scatter encountered in determining creep rates under strict testing conditions
(Ref 4). Scatter for rupture times is approximately half that magnitude with materials that are specially prepared
to be uniform.
Temperature and Other Testing Variables. Even with the use of precise temperature controllers and high-quality
pyrometric practice, care is required to hold the average indicated specimen temperature within usual
specifications of ±2 °C (±3.6 °F) (Ref 2) with time. Variation from one location to another along the specimen
gage length typically approaches this magnitude, even in furnaces with independent adjustment of power input
within zones.
These variations in indicated temperature do not include errors in initial thermocouple and pyrometer
calibrations, leadwire mismatch, or drift in thermocouple output with time. Actual temperature can differ from
the reported value by at least 5 °C (9 °F) during some portion of a representative creep or rupture test.
For steel at 450 °C (842 °F), a temperature disparity of 5 °C (9 °F) corresponds approximately to 20% variation
in rupture time (Ref 4). A 10% change in creep rate results from a load change of 1.5%, or a temperature
change of 2.5 °C (4.5 °F) (Ref 3). An error of more than 1% in the applied load is unlikely under typical
conditions.
Equipment for creep-rupture testing commonly includes features to promote uniform axial loading. However,
bending loads may still occur. The increase in secondary creep rates and reduction in rupture life with
eccentricity of loading is analyzed in Ref 5, which reports that extreme cases of bending introduced by threaded
ends on cylindrical specimens reduce rupture life by as much as 60%. The largest effects occur for short or
notched specimens and for materials with low ductility to rupture.
Eccentric loading exerts maximum influence on creep measurements in the early stages of a test. With initially
bowed specimens, which are common when specimens are as-cast or are strip-type machined from thin
material, early creep indications may be erroneous if the extensometer fails to accurately average the strains at
opposite sides of the specimen.
Typical Data Scatter. Rolled bars from a single heat of 2.25Cr-1Mo steel were tested by 21 laboratories in eight
countries (Ref 6). The largest deviations in average results for one laboratory from the arithmetic mean of
values from all laboratories for the stress to cause rupture in 1000, 3000, and 10,000 h were:
Temperature
°C °F
Deviation
550 1022
+14.3 and -16.6%
600 1112 +21.2 and -12.1%
Another group of 18 laboratories in seven countries cooperated in tests of Nimonic 105 alloy bars at 900 °C
(1652 °F) (Ref 7). Mean rupture times adjusted to a common stress for individual laboratories (four tests each)
ranged from a high of 1491.4 h to a low of 1090.2 h (15.5% deviation). Time to 0.5% total deformation varied
from 75.4 to 182.3 h for 16 of these laboratories; deviation from the mean value thus increased to 41.5%.
Temperature control was found to be the most serious source of variation in rupture measurement. In particular,
calibration drift of Chromel-Alumel thermocouples at this high test temperature was responsible for the long
mean rupture times found by five of the laboratories. Not more than about 15% scatter generally can be
attributed to testing variables, if a laboratory has followed standard procedures.
Multiple Heats and Product Forms. Scatter bands become much broader when data originate from tests on
numerous heats, particularly when data include a broad range of product forms and sizes. Elevated-temperature
properties for steel and superalloys are available (see, for example, Ref 8, 9, 10, 11, 12, 13, 14, 15, 16, 17, and
18). Typically, data for each grade of steel include plots of the stress for rupture in 1000, 10,000, or 100,000 h
versus the test temperature. Where appropriate, distinctive shapes of data points identify different product forms
or heat treatments.
When data at a given temperature are extensive, the reported range of derived stresses to produce rupture in a
given time typically consists of a two-fold ratio of the highest to the lowest stress level. The corresponding
spread of rupture times for a given test stress is on the order of 100-fold.
If a new set of test results does not fall within the broad scatter bands, careful review and confirmation is
required before the new data are accepted as valid. Bias of new data tends toward the upper half of the scatter
bands compiled for older data because the increase of residual alloying elements from scrap recycling, higher
nitrogen content, and improved alignment and temperature control in modern testers tend to raise indicated
creep and rupture properties.
At a test temperature of 593 °C (1100 °F) and at a stress of 207 MPa (30 ksi), rupture life ranged from 84 to
2580 h and secondary creep rate ranged from 0.16 to 0.00077%/h for 20 heats of type 304 stainless steel in the
reannealed condition (Ref 19). Corresponding large variations were observed at all test temperatures and for