William S. Loewenthal

Ohio Aerospace Institute, Brook Park, Ohio

David L. Ellis

Glenn Research Center, Cleveland, Ohio

Sources of Variation in Creep Testing

NASA/TM—2011-215493

December 2011
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William S. Loewenthal

Ohio Aerospace Institute, Brook Park, Ohio

David L. Ellis

Glenn Research Center, Cleveland, Ohio

Sources of Variation in Creep Testing

NASA/TM—2011-215493

December 2011

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135
Acknowledgments

The authors acknowledge and express their appreciation to the efforts of Sharon Y. Thomas and John W. Zoha for conducting 

the creep testing and the important contributions by Ronald E. Phillips, John J. Juhas, and Adrienne H. Veverka for their help

in identifying sources of equipment variation and quickly implementing the needed modifi cations.

Available from

NASA Center for Aerospace Information

7115 Standard Drive

Hanover, MD 21076–1320

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Available electronically at http://www.sti.nasa.gov

Trade names and trademarks are used in this report for identifi cation 

only. Their usage does not constitute an offi cial endorsement, 

either expressed or implied, by the National Aeronautics and 

Space Administration.

Level of Review: This material has been technically reviewed by technical management. 
NASA/TM—2011-215493 1 

Sources of Variation in Creep Testing 

William S. Loewenthal1

Ohio Aerospace Institute 

Brook Park, Ohio 44142 

David L. Ellis 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Summary 

This report is a section of the final report on the GRCop-84 task of the Constellation Program and 

incorporates the results obtained between October 2000 and September 2005 when the program ended. 

Creep rupture is an important material characteristic for the design of a variety of rocket engines, 

including those for reusable vehicles, the Earth Departure Stage (EDS), and the Lunar Surface Access 

Module (LSAM). It was observed during the earlier characterization of GRCop-84 that the complete data set 

had nearly 4 orders of magnitude of scatter. This large scatter may have confounded attempts to accurately 

determine how creep performance was influenced by manufacturing. It was unclear if this variation was 

from the testing, the material, or a combination of both. Sources of variation were examined for constantload vacuum creep tests. Special attention was paid to the repeatability between creep tests conducted on 

different creep frames. Tests were conducted on identically processed specimens at the same specified 

stresses and temperatures. 

It was found that significant differences existed between the five constant-load Brew creep test frames 

examined and that the specimen temperature was higher than the desired temperature by as much as 43 °C 

(77 °F). It was also observed that the top-to-bottom temperature profile varied as much as 44 °C (79 °F). 

Results showed that GRCop-84 creep behavior is more sensitive to small furnace temperature changes than 

was previously realized. The higher actual temperatures suggest that the previously collected data are a 

conservative prediction of creep rate and life. 

Improved specimen temperature measurement and control were incorporated as part of an effort to 

minimize the temperature variations. To verify that these modifications decreased variation, additional tests 

were conducted, and the data between Brew vacuum creep test frames were shown to be comparable. The 

variation decreased to 1/2 order of magnitude from the previously observed 2 orders of magnitude for the 

baseline extruded GRCop-84 data set. Independent determination of creep rates by a step-loading method in 

a reference load frame closely matches the creep rates determined after the Brew modifications. Substituting 

a helium atmosphere for the vacuum tended to decrease the sample temperature gradient by improving heat 

transfer, but helium was not a significant improvement over vacuum. 

Introduction 

Creep rupture behavior is an important material characteristic needed for the design of rocket engines. 

GRCop-84 (Cu-8 at.% Cr-4 at.% Nb), a Cu alloy developed by the NASA Glenn Research Center, has 

excellent elevated temperature strength, low cycle fatigue (LCF), and creep properties. This makes it an 

excellent replacement candidate for NARloy-Z (Cu-3 wt% Ag-0.5 wt% Zr), the current Space Shuttle Main 

Engine main combustion chamber liner material, and makes it competitive with many other highconductivity Cu alloys (Ref. 1). Work to understand and optimize GRCop-84 has produced a large database 

of creep data, which are reported elsewhere in the final report (Ref. 2, to be published). 

 

1

William Loewenthal worked for the Ohio Aerospace Institute at the time that this research was done. He now works 

for H.C. Starck in Euclid, Ohio. 
NASA/TM—2011-215493 2 

It was observed that GRCop-84 creep data, regardless of process variants, consistently had nearly 

2 orders of magnitude of scatter—even when tested on the same creep frame—and up to 4 orders of 

scatter overall (Ref. 2 and Refs. 3 to 5). The large scatter has been an ongoing concern. One example 

showing large creep variation was observed during a study of GRCop-84 rolling parameters (Ref. 5). In 

this study, GRCop-84 sheet was rolled at 215, 300, and 415 °C (419, 572, and 779 °F), and creep life was 

determined at 500, 650, and 800 °C (932, 1202, and 1472 °F). The results shown in Figure 1 illustrate the 

large scatter, with considerable overlap of the results for the three rolling temperatures. Similarly, it is 

difficult to estimate the actual performance difference between the rolled sheet and the baseline extruded 

material (lines in Fig. 1). Although it appears that the rolled material has a lower creep life, the scatter 

does not allow such a conclusion to be reached statistically. 

The problem is compounded when examining the complete historical data set. When all data are 

plotted as shown in Figure 2, the range of results approaches 4 orders of magnitude. However, the 

Monkman-Grant plot of the historical data had an excellent fit with only a few data points identified as 

probable outliers (Ref. 2). Given this variability within individual data sets and in the overall data, the 

statistical analysis was unable to detect significant differences between processing methods, heat treatments, or any other effects besides the primary ones of stress and temperature. Creep rate data are not 

presented, but the variability and range were almost identical to the creep lives presented in Figure 2. 

It was unclear if this large variation was due to the test methods, the material, or a combination of 

both. Possible sources of variation besides material effects could include temperature and stress effects 

that were not accounted for in the original test method as well as instrumentation and calibration errors. 
NASA/TM—2011-215493 3 
NASA/TM—2011-215493 4 

Creep rupture lives and rates were measured under a constant load using a simple lever arm machine 

with a vacuum chamber manufactured by the Brew Corporation, referred to as a Brew test frame. These 

frames were manufactured circa 1957, have sequential serial numbers indicating manufacture at the same 

time, and had been upgraded once prior to the baseline GRCop-84 testing and twice prior to the current 

creep testing. The upgrades were designed to improve the data collection, temperature control, and 

vacuum systems by bringing them up to then current standards for electronics and data acquisition. All 

units were upgraded at the same time using the same components by the same technicians, so variation 

from frame to frame was anticipated to be minimal. However, the load, the load train alignment, the 

temperature, and other frame-specific variables were considered in this study. 

From the baseline data shown in Figure 2, it was known that stress and temperature are the major 

factors affecting creep. Efforts were undertaken to experimentally characterize these and other potential 

machine-related factors affecting creep property variability through a design of experiments (DOE) and 

statistical analysis of the results. 

Experimental Procedure 

GRCop-84 and other competitive materials were characterized in five of the six Brew frames available at Glenn. The Brew frames are numbered 1 to 6. Brew frame 6 (serial no. 31–309) was not used in 

these studies because of a nonfunctioning vacuum system. In addition, a calibrated data set for creep rate 

was created using a converted Instron tensile test frame (TM–4) whose load and temperature had been 

carefully characterized. The load measurement and temperature gradients were known to be superior for 

this unit. This data set was used as the baseline for determining the true creep rate of GRCop-84 at these 

temperatures and stresses. 
NASA/TM—2011-215493 5 

Special attention was given to examining the repeatability between supposedly identical creep test 

conditions conducted on different Brew frames. So that the sources of variation between each Brew frame 

could be determined, a DOE was conducted on specimens from the same sheet of GRCop-84 at the same 

stresses and temperatures. 

This work was conducted in three sequential stages, where the results of the prior stage were used 

to guide the work in the next stage. Details of the components of each stage are given following an 

overview. 

In stage I, an initial survey was conducted to determine if the measured GRCop-84 creep properties 

for a single stress and temperature combination differed or were statistically the same between Brew 

frames. The findings from the initial survey confirmed that the mean value varied between Brew frames 

more than could be accounted for by the variation of results within a frame. This result led to a second 

stage to examine the potential sources of variation. 

From the stage II (root cause) study, it was learned that variation in specimen temperature was larger 

and more important than previously recognized. Modifications were made to each Brew frame to monitor 

actual specimen temperature and thermal gradients by installing additional thermocouples. The operating 

procedure was also modified to control the furnace from the middle specimen thermocouple to achieve 

equivalent specimen temperatures for all tests in all Brew frames. 

In stage III, the final stage, tests were conducted to validate that each Brew frame provided results 

similar to those of the other Brew frames and to the calibrated data set from TM–4. The updated procedures and test frames also were used to determine if the expected changes in creep properties with 

processing—which previously were not statistically significant, though probably observed—became 

statistically significant. The statistical variations of the results also were examined to determine if they 

had decreased. 

Initial Survey Design of Experiments 

A DOE was conducted on identically processed specimens at the same applied stress and temperature 

combination to determine differences between creep test frames. It was assumed that conducting all of the 

creep tests from the same material lot would minimize any specimen-to-specimen differences; and that, if 

differences were observed, they could be attributed to the creep frame rather than to the material. Material 

effects were minimized further by taking the samples from random locations within a single sheet of 

material. 

The test variables selected were Brew frame, temperature, and stress, with only the Brew frame 

purposely varied. From prior experience, the temperature and stress were well known to affect the creep 

rupture life of GRCop-84, as shown in Figures 1 and 2. It was suspected, but not known, that frame-toframe variations existed despite the identical manufacture and operation of the units. 

The response variables were creep life and creep rate. For the stage I survey to determine if there was 

variation between Brew frames, triplicate tests were conducted at 500 °C/105 MPa (932 °F/15.2 ksi) on 

each test frame for a total of 15 creep tests. 

Investigation of Root Causes of Variation 

Various root causes that could produce different creep responses were investigated. These investigations centered on temperature and load, but other test aspects also were considered. A detailed evaluation 

was conducted of the creep test procedure, thermocouple and instrumentation calibration, frame alignment, friction in the mechanical feedthrough ball joints, and resistance from the weight of the rubber 

cooling hose hanging on the upper mechanical feedthrough ball joint. The procedure appeared to be adequate, the calibrations and alignments were good, and no extraneous loads or bending moments could be 

found. The results indicated that there were no major contributions from these factors, so the investigation 

of root causes which is reported here focuses on temperature and load. 
NASA/TM—2011-215493 6 

During the investigation of root causes, thermocouples were installed on Brew frames 3 and 4 to 

monitor the specimen temperature and thermal gradients. These values were compared with the control 

thermocouple to determine variations between the control thermocouple and the sample temperature. 

Three thermocouples attached to the sample allowed measurement of the thermal gradients in the samples 

as well. Also on Brew frames 3 and 4, an 8.9-kN- (2000-lbf-) capacity Sensotec model 31 series load cell 

and Model GM controller from Honeywell International Inc. were installed in the lower pull rod to 

examine loads and determine gross variations in the loads applied to the samples because of mechanical 

differences in the load frames. 

The Brew frame test chambers were designed to allow pressurization of up to 1 atm as well as 

vacuum operation. An Omega Engineering DPF60 series controller and automated supply and bleed 

solenoid valves were added to Brew frame 3 to pressurize the chamber to 89.7 kPa (13 psi) with highpurity helium. It was thought that a helium atmosphere would improve specimen temperature uniformity 

by allowing heat transfer by convection as well as radiation and conduction within the sample. Helium 

was selected over argon and other inert gases because of its higher thermal conductivity. A series of tests 

using the GRCop-84 sheet samples shown in Table I were conducted following the same setup and test 

procedures used previously. The testing was conducted sequentially, except for tests 10 to 15, with the 

results of prior tests being used to determine the conditions to be used on the subsequent tests. 

Results were analyzed using analysis of variance (ANOVA) tests to compare the average creep rate 

and creep life for the five test frames simultaneously. When differences in the means were detected, a 

Student-Newman-Keuls (SNK) test was used to compare and rank the means. 

TABLE I.—SPECIMEN TEST CONDITIONS FOR INVESTIGATING 

ROOT CAUSES OF VARIATION 

Test Creep 

temperature 

Creep stress Brew 

frame 

Atmosphere 

°C °F MPa ksi 

 1 500 932 105 15.2 3 Vacuum 

 2 500 932 105 15.2 3 Vacuum 

 3 500 932 105 15.2 4 Vacuum 

 4 800 1472 17 2.5 3 Vacuum 

 5 800 1472 17 2.5 4 Vacuum 

 

 6 650 1202 44.3 6.9 3 Vacuum 

 7 650 1202 44.3 6.9 4 Vacuum 

 8 500 932 105 15.2 3 Vacuum 

 9 500 932 105 15.2 4 Vacuum 

 10 500 932 105 15.2 3 Helium 

 

 11 500 932 105 15.2 3 Helium 

 12 650 1202 44.3 6.9 3 Helium 

 13 650 1202 44.3 6.9 3 Helium 

 14 800 1472 17 2.5 3 Helium 

 15 800 1472 17 2.5 3 Helium 

Designs of Experiments for Verifying Improvements 

The stage II root cause determination found that temperature was the most important variable and that 

there were large differences between the control thermocouple temperature and the sample temperature. 

Furthermore, the differences did not appear to be consistent even within a single load frame. Thermocouples to measure the temperature of the top, middle, and bottom of the reduced section of the specimen 

were permanently installed in all Brew frames. The middle thermocouple temperature was used to control 

the temperature of the specimen after installation. The operating procedure was updated to include the 

newly determined best practices, including the use of three thermocouples for temperature measurement. 
NASA/TM—2011-215493 7 

After these changes, another series of tests with GRCop-84 sheet specimens was conducted to confirm that the variation in creep properties was reduced and that the experimental error had been minimized. Duplicate tests of material from sheet 3A1B were measured at test conditions similar to those for 

historical tests. Creep conditions of 500 °C/105 MPa (932 °F/15.2 ksi), 650 °C/44.3 MPa (1202 °F/ 

6.9 ksi), and 800 °C/17 MPa (1472 °F/2.5 ksi) were tested on Brew frames 2, 3, and 4. Brew frames 1, 5, 

and 6 were unavailable because of other testing. An additional series of tests was conducted at the same 

conditions except that a helium atmosphere was used instead of a vacuum and that the material came from 

sheet 4A2. Table II presents the verification test matrix. 

TABLE II.—CONDITIONS AND SAMPLES FOR GRCop-84 SHEET VERIFICATION 

SURVEY DESIGN OF EXPERIMENTS 

Run Sample Creep stress, 

MPa

Temperature,

°C

Run Sample Creep stress, 

MPa

Temperature,

°C

Vacuum tests Helium atmosphere tests

1 3BA1–1 105 500 1 4A2–1 105 500 

2 3BA1–2 105 500 2 4A2–2 105 500 

3 3BA1–3 44.3 650 3 4A2–3 44.3 650 

4 3BA1–4 44.3 650 4 4A2–4 44.3 650 

5 3BA1–5 17 800 5 4A2–5 17 800 

6 3BA1–6 17 800 6 4A2–6 17 800 

7 3BA1–7 105 500 7 4A2–7 105 500 

8 3BA1–8 105 500 8 4A2–8 105 500 

9 3BA1–9 44.3 650 9 4A2–9 44.3 650 

10 3BA1–10 44.3 650 10 4A2–10 44.3 650 

11 3BA1–11 17 800 11 4A2–11 17 800 

12 3BA1–12 17 800 12 4A2–12 17 800 

13 3BA1–13 105 500 13 4A2–13 105 500 

14 3BA1–14 105 500 14 4A2–14 105 500 

15 3BA1–15 44.3 650 15 4A2–15 44.3 650 

16 3BA1–16 44.3 650 16 4A2–16 44.3 650 

17 3BA1–17 17 800 17 4A2–17 17 800 

18 3BA1–18 17 800 18 4A2–18 17 800 

Two additional sets of creep tests were completed using the updated practices. The first was a survey 

of the creep properties of GRCop-84 plate made from standard –140 mesh powder. The second was an 

analysis of the effects of a braze thermal cycle on the creep properties of GRCop-42 (Cu-4 at.% 

Cr-2 at.% Nb). 

The –140 mesh powder GRCop-84 plate is from material used in the production of several 177.9-kN 

(40 000-lbf) class rocket engine main combustion chamber liners and was used to verify and certify the 

liner material’s creep properties. It provides a set of data for comparison of the observed variation in 

creep test results with a reasonably large data set. As such it provides an estimate of the material and 

remaining test method variability. Table III shows the test conditions and DOE. 

Following testing, a model with the form 

  10   2

10 0 1 10 3

log log log 273 273

Y

T T

 

    

 

 (1) 

was entertained to determine the dependency of the dependent variable Y (creep rate in reciprocal seconds 

(s–1) or creep life in hours) on the stress  and test temperature T. Here 0 to 3 are the coefficients to be 

fit by the regression analysis. Equation (1) allows prediction of the life and rate for intermediate temperatures and stresses, so it can be used to explore the relative dependency of the rate and life on these 

independent variables. 

TABLE III.—CONDITIONS AND SAMPLES FOR GRCop-84 SHEET VERIFICATION 
NASA/TM—2011-215493 8 

SURVEY DESIGN OF EXPERIMENTS 

Run Sample Creep stress, 

MPa

Temperature,

°C

Run Sample Creep stress, 

MPa

Temperature,

°C

Vacuum tests Helium atmosphere tests

1 3BA1–1 105 500 1 4A2–1 105 500 

2 3BA1–2 105 500 2 4A2–2 105 500 

3 3BA1–3 44.3 650 3 4A2–3 44.3 650 

4 3BA1–4 44.3 650 4 4A2–4 44.3 650 

5 3BA1–5 17 800 5 4A2–5 17 800 

6 3BA1–6 17 800 6 4A2–6 17 800 

7 3BA1–7 105 500 7 4A2–7 105 500 

8 3BA1–8 105 500 8 4A2–8 105 500 

9 3BA1–9 44.3 650 9 4A2–9 44.3 650 

10 3BA1–10 44.3 650 10 4A2–10 44.3 650 

11 3BA1–11 17 800 11 4A2–11 17 800 

12 3BA1–12 17 800 12 4A2–12 17 800 

13 3BA1–13 105 500 13 4A2–13 105 500 

14 3BA1–14 105 500 14 4A2–14 105 500 

15 3BA1–15 44.3 650 15 4A2–15 44.3 650 

16 3BA1–16 44.3 650 16 4A2–16 44.3 650 

17 3BA1–17 17 800 17 4A2–17 17 800 

18 3BA1–18 17 800 18 4A2–18 17 800 

Three sets of GRCop-42 samples from the baseline extruded material (extrusion lots 192 and 211) 

were tested in the as-extruded condition and following simulated braze cycles at 935 and 1000 °C 

(1715 and 1832 °F). These simulated braze cycles are detailed in Tables IV and V. The objective of the 

test program was to determine if the creep rate increased and the creep life decreased as was expected 

from a decrease in tensile strength. Prior statistical analysis had not been able to detect such a difference 

even though the simulated braze cycle material tended to have the expected differences. The likely cause 

of the failure to detect the difference was the scatter in the data. With reduced scatter, it was expected that 

the differences would now be detectable if the changes had actually improved the test results. Such a 

result would validate the new test procedures. 

Tests were conducted at 500, 650, and 800 °C (932, 1202, and 1472 °F) using multiple stresses. 

Table VI lists the test details. Following testing, a model was entertained to determine which independent 

variables were statistically significant. The model had the form 

   

 

2 10

10 0 1 10 3

61 72

4 1 10 5 2 10

log log log 273 273

log log 273 273

Y

T T

B B

B B

T T

 

    

 

       

 

 (2) 

where Y is the dependent response variable (creep life in hours or creep rate in reciprocal seconds (s–1)); 

 is the applied stress in megapascals (MPa); T is the temperature in degrees Celsius; B1 and B2 are 

blocking variables corresponding to simulated brazes of 935 and 1000 °C, respectively; and 0 to 7 are 

the coefficients determined by the regression analysis.

A blocking variable took a value of 1 if the sample had a particular heat treatment and 0 if it did not. 

If both B1 and B2 were 0, the sample was in the as-extruded condition. The blocking variables allowed the 

use of forward stepwise regression to determine if a heat treatment had a significant effect. If no term 

with the blocking variable corresponding to a heat treatment entered the regression, the conclusion was 

that the heat treatment did not have a significant effect on creep properties relative to the as-extruded 

condition.
NASA/TM—2011-215493 9 

TABLE IV.—SIMULATED 935 °C (1715 °F) BRAZE CYCLE 

Step Temperatures and rates 

°C at °C/min °F at °F/min 

1 Room temperature to 538 °C at 5.6 °C/min Room temperature to 1000 °F at 10 °F/min 

2 538 to 871 °C at 2.8 °C/min 1000 to 1600 °F at 5 °F/min 

3 871 to 935 °C at 1.7 °C/min 1600 to 1715 °F at 3 °F/min 

4 935±8.3 °C/hold for 22.5±2.5 min 1715±15 °F/hold for 22.5±2.5 min 

5 935 to 871 °C at 1.7 °C/min 1715 to 1600 °F at 3 °F/min 

6 871 to 538 °C at 2.8 °C/min 1600 to 1000 °F at 5 °F/min 

7 538 °C to room temperature free cool 1000 °F to room temperature free cool 

TABLE V.—SIMULATED 1000 °C (1832 °F) BRAZE CYCLE 

Step Temperatures and rates 

°C at °C/min °F at °F/min 

1 Room temperature to 538 °C at 5.6 °C/min Room temperature to 1000 °F at 10 °F/min 

2 538 to 816 °C at 2.8 °C/min 1000 to 1500 °F at 5 °F/min 

3 816 to 1000 °C at 1.7 °C/min 1500 to 1832 °F at 3 °F/min 

4 1000±8.3 °C/hold for 22.5±2.5 min 1832±15 °F/hold for 22.5±2.5 min 

5 1000 to 260 °C at 2.8 °C/min 1832 °F to 500 °F at 5 °F/min 

6 260 to 121 °C at 2.8 °C/min 

 or best possible cooling rate 

500 to 250 °F at 5 °F/min 

 or best possible cooling rate 

7 121 °C to room temperature free cool 250 °F to room temperature free cool 

TABLE VI.—CONDITIONS AND SAMPLES FOR GRCop-42 BRAZE EFFECT DESIGN OF EXPERIMENTS

Run Sample Condition Temperature,

°C 

Applied 

stress, 

MPa 

Run Sample Condition Temperature, 

°C 

Applied 

stress, 

MPa 

1 5–23 1000 °C braze 500 77.3 27 5–15 1000 °C braze 800 17.4 

2 2–5 1000 °C braze 500 77.3 28 2–7 1000 °C braze 800 18.6 

3 4–13 935 °C braze 500 70.7 29 4–12 935 °C braze 800 18.7 

4 6–5 As extruded 500 97.8 30 1–22 As extruded 800 20.4 

5 1–24 As extruded 500 92.0 31 4–11 935 °C braze 800 19.9 

6 2–8 1000 °C braze 500 87.6 32 6–6 As extruded 800 21.8 

7 3–5 935 °C braze 500 87.0 33 6–1 As extruded 800 20.4 

8 5–19 1000 °C braze 500 82.4 34 4–25 935 °C braze 800 18.7 

9 1–28 As extruded 500 97.8 35 1–17 As extruded 800 21.8 

10 3–3 935 °C braze 500 92.5 36 3–1 935 °C braze 800 19.9 

11 1–26 As extruded 500 92.0 37 5–24 1000 °C braze 650 35.0 

12 3–4 935 °C braze 500 87.0 38 5–13 1000 °C braze 650 35.0 

13 2–2 1000 °C braze 650 51.5 39 3–6 935 °C braze 650 35.0 

14 1–23 As extruded 650 60.4 40 1–32 As extruded 650 35.0 

15 1–16 As extruded 650 56.8 41 1–13 1000 °C braze 650 35.0 

16 5–27 1000 °C braze 650 48.5 42 5–21 1000 °C braze 800 10.0 

17 4–29 935 °C braze 650 49.7 43 6–7 As extruded 800 10.0 

18 4–14 935 °C braze 650 53.0 44 4–19 935 °C braze 650 35.0 

19 3–7 935 °C braze 650 49.7 45 5–28 1000 °C braze 500 130.0 

20 1–19 As extruded 650 53.3 46 4–18 935 °C braze 800 10.0 

21 2–4 1000 °C braze 650 45.5 47 1–31 As extruded 500 130.0 

22 2–3 1000 °C braze 650 48.5 48 5–25 1000 °C braze 800 13.0 

23 1–25 As extruded 650 56.8 49 5–26 1000 °C braze 800 10.0 

24 4–28 935 °C braze 650 53.0 50 4–15 935 °C braze 500 130.0 

25 5–16 1000 °C braze 800 17.4 51 1–18 As extruded 800 13.0 

26 5–18 1000 °C braze 800 18.6 52 4–27 935 °C braze 800 13.0 
NASA/TM—2011-215493 10 

Stepwise forward analysis was conducted using Systat Software’s2 SigmaStat Version 3.1 statistical 

software. The raw data were converted to the form needed for Equation (2), and blocking variables were 

added. Analyses were conducted for both the creep rate and creep life. The stepwise regression analysis 

used F statistic values (the value of the F distribution for X degrees of freedom associated with the 

numerator and Y degrees of freedom associated with the denominator) of 4 to enter and 3.9 to leave. 

These values were selected because they correspond to approximately a 95-percent probability that the 

variable should or should not be in the model, respectively. 

Specimen Designs 

Quantities of standard sheet tensile specimens conforming to ASTM Standard E8 (Ref. 6) were wire 

electrical discharge machined. The sheet specimen design used is shown in Figure 3. The sheet samples 

for stages I and II of the study were commercially rolled by H.C. Starck GmbH in Coldwater, Michigan, 

to a nominal thickness of 0.762 mm (0.030 in.) and were annealed for 30 min at 600 °C (1112 °F). Details 

of the manufacturing process are reported in Reference 7. The stage II testing used materials from the 

rolling parameter optimizations study (Ref. 5). Neither set of samples had the recast layer removed 

because the volume of affected material was very small relative to the total volume and because the recast 

layer and base material should have similar microstructures since they were both rapidly solidified. In this 

context, the small amount of recast material was deemed unlikely to affect the test results. The procedure 

is also consistent with the preparation of prior test specimens. 

The GRCop-42 bar stock used for creep testing was made by extruding two 22.7-kg (50-lb) lots of 

–140 mesh loose powder at H.C. Starck. The lots were designated 192 and 211. A nominal 30:1 reduction 

in area was used to produce 2.5-cm- (1-in.-) diameter rods. The GRCop-84 plate was made by extruding 

800 lb of –140 mesh powder and rolling the material to 12.7-mm- (0.5-in.-) thick plate. Round samples 

conforming to ASTM Standard E8 were machined from the rods and plate for creep testing. The design 

used for these samples is shown in Figure 4. 

 

2

Systat Software, Inc., 1735, Technology Drive, Ste. 430, San Jose, CA 95110. 
NASA/TM—2011-215493 11 

Creep Testing 

Two types of creep tests were conducted: constant-load creep rupture tests and step-loading creep rate 

tests. Constant-load creep tests give a steady-state creep rate for a given stress and temperature as well as 

a rupture life. These were the tests conducted in the Brew frames. The step-loading tests were conducted 

using a separate modified tensile test load frame designated TM–4 with vacuum, heating, and data collection capabilities similar to those of the Brew frames. Step-loading tests produce many creep rates at different stresses but do not produce any creep lives. The procedures for conducting both types of creep tests 

are discussed in the following sections. 

Constant-load vacuum creep testing.—Constant-load creep tests were conducted using a simple 

lever arm machine with a vacuum chamber. As noted before, the units were manufactured by the Brew 

Corporation circa 1957 and have undergone multiple upgrades to the electronics and controls since then. 

The current configuration is shown in Figure 5. 

The test specimens were placed into a water-cooled stainless steel chamber with vacuum-tight 

mechanical feedthrough ball joints at the top and bottom. The ball joints ensure transmission of the load 

to the specimen with a minimal bending moment or other misalignments. The specimen was heated by a 

tungsten mesh heater surrounded by multiple tungsten heat shields. For GRCop-84 and GRCop-42 testing, Type 310 stainless steel fixtures were normally used, although MAR M–246 fixtures were used when 

310 stainless steel fixtures were not available. The fixturing was adjusted in length so that the sample was 

centered in the hot zone. 

Temperature control was initially done using an Inconel sheathed Type R (Pt/Pt-13% Rh) thermocouple with an exposed bead introduced through the back of the chamber. The control thermocouple was 

placed 1.6 to 3.2 mm (0.063 to 0.125 in.) from the sample. The thermocouple was not allowed to touch 

the specimen to prevent the development of a bending moment upon heating the specimen. An overtemperature thermocouple was coaxially located with the control thermocouple. 

During this program, three additional flexible Type R thermocouples were attached to the top, middle, 

and bottom of the gauge section of the sample as shown in Figure 5(c). The thermocouples used highpurity, high-density alumina beads for insulation and sheathing. The beads also allowed the thermocouples to flex easily. Initially, these attached thermocouples were used only for sample temperature 

measurements. Later, the middle thermocouple also was used as the control thermocouple for stage III, 

the final stage of the study, when the improved testing procedures were verified. A vacuum-tight 

thermocouple feedthrough was used to connect the thermocouples to exterior instrumentation and data 

acquisition. 
NASA/TM—2011-215493 12 
NASA/TM—2011-215493 13 

The load was applied to the specimen using a 10:1 lever arm and a dead weight consisting of various 

size weights. A load pan was attached to the opposite end of the lever arm from the sample to allow the 

weights to be stacked. Prior to loading the sample, the arm was leveled and balanced so that no moments or 

loads were induced in the samples from the lever arm. The samples were heated in vacuum to the desired 

test temperature with a minimal load (typically 4.5 N or 1 lbf) to stabilize the sample and fixtures during 

heating. When the sample’s temperature stabilized, a scissor jack was raised underneath the pan to remove 

all weight on the lever arm, the desired pan load equal to 1/10 the desired sample load was placed on the 

pan, and the scissor jack was lowered. The load on the sample was measured using an 8.9-kN- (2000-lbf-) 

capacity Honeywell Sensotec Type 31 miniature load cell placed in the load train beneath the chamber 

during tests where the load cell had been added to the test frames. The load cell allowed a quantitative 

measure of the load and acted as a way for the data acquisition system to detect failure of the sample. The 

maximum error for the load cells based on the manufacturer-supplied specifications was 0.25 percent of full 

scale, or 22 N (5 lbf). 

Displacement was measured by a linear variable resistor (LVR) attached to the load train rather than the 

sample. The LVR used has a resolution of 2.54 m (0.0001 in.) and, unlike prior optical measurements, 

allowed the displacement to be recorded at a frequent, set time interval. So that the strain could be calculated, the gauge length of the reduced section was measured using a Focus Contour Projector optical comparator from Optical Gaging Products, Inc. The displacement measured by the LVR was divided by the 

initial room temperature gauge length to compute the strain. 

At least the first data point included the displacement caused by the slack of the load frame being 

removed through the application of the load. No adjustments to the LVR displacement values and measured 

strain were made for the loading displacement because there was no direct measurement of the sample 

displacement from which to assign definitive values to the duration and magnitude of the loading displacement. Prior in-house efforts to measure the strain in the reduced section using an optical cathetometer and to 

correlate the LVR strains to sample strains showed that there was a strong correlation between the two 

measurements, but the scatter in the data was too great to provide the degree of confidence desired for 

conversion of the LVR strain into sample strain and ultimately creep strain (Ref. 2, to be published). 

The correlation and the individual creep curves generated for the samples used to develop the 

correlation did show that the relative displacements measured by the LVRs between consecutive data points 

accurately represented the actual change in the sample’s strain as measured by the cathetometers. Because 

of this, the absolute magnitude of the LVR strain without any adjustments was uncertain and inaccurate, but 

the change in strain after the first few data points was accurate. Since the steady-state creep rate was calculated from the relative change of the strains well after the initial loading, the creep rate was accurate. 

All signals were collected by data acquisition using internally developed NASA data acquisition software running on a Dell 2400 desktop computer. The signals were converted from analog to digital using 

I–7019R eight-channel universal analog input modules from MicroDAQ.com, Ltd., connected to an 

I–7561 USB (universal serial bus) to an RS–232/422/485 converter. Data were collected every second for 

the first 5 min of the test, every 5 min for the next 30 min of the test, and once every hour until the end of 

the test. In addition, the software queried the modules once every 10 s and stored the data in a ring buffer. 

Upon failure, as defined by a 90-percent or more drop in the load from the peak load, the contents of the 

ring buffer were written to the file, and the 10-s-interval data for the last hour of the test were recorded. 

During testing, using a helium atmosphere instead of vacuum was investigated to determine if it would 

improve specimen temperature uniformity. An Omega Engineering DPF60 series process controller was 

used to control the helium pressure into the chamber. An electrically operated inlet and outlet solenoid valve 

and a pressure transducer were connected to the process controller. The process controller was set to maintain 93- to 97-Pa (13.5- to 14-psia) pressure in the chamber as measured by the pressure transducer. The 

outlet line had to be connected to an existing mechanical vacuum pump to remove extra helium since the 

chamber pressure was slightly below 1 atm. This design resulted in a very low flow of helium through the 

chamber and hence very little temperature fluctuation due to the sudden inrush of gas. The selected system 

also had the advantage of minimizing oxygen contamination and helium usage. 
NASA/TM—2011-215493 14 

Step-loading vacuum creep testing.—An Instron TT series tensile test frame designated TM–4 

(Fig. 6) was upgraded to computer data acquisition and control by the Instru-Met Corporation3

 using MTS 

Systems Corporation4

 TestWorks 4 testing software and associated MTS hardware. The software allows a 

load and recording displacement to be held over extended periods. The programming portion of the 

software allowed for developing a stepped sequence of loads and times. The sample was monitored 

continually, and data points were logged when the sample underwent a 0.01-percent change in strain. 

Each data point consisted of the elapsed time, load, and strain. The desired load was placed on the sample 

by moving the crosshead beneath the chamber downwards and pulling on the sample until the load cell 

measured the desired load. The movement of the crosshead was measured using an optical encoder to 

determine the displacement of the sample during creep. 

One notable difference in the two types of load frames is that TM–4 uses a solid heater element rather 

than a mesh heater element. Temperature was controlled using a Type R thermocouple tied to the center 

of the specimen’s gauge section. The over-temperature thermocouple was attached to the top of the specimen’s gauge section, and a recording thermocouple was attached to the bottom of the specimen’s gauge 

section. The gradient was not recorded, but general observations of the three temperature readouts 

indicated a temperature gradient of less than 5 °C (9 °F). 

As with the LVR data for the constant-load tests, all displacement was assumed to occur in the 

reduced section of the specimens. Some creep occurs in the transition regions or sample shoulders, and 

that creep depends on the specimen geometry. If sufficiently large, the creep strain in the transition 

regions leads to an overestimate of the total creep since the total displacement divided by just the reduced 

gauge section length is used to calculate the creep strain. 

 

3

Instru-Met Corporation, 999 Rahway Avenue, Union, NJ 07083. 

4

MTS Systems Corporation, 14000 Technology Drive, Eden Prairie, MN 55344.
NASA/TM—2011-215493 15 

So that the error introduced by this assumption and the difference in specimen geometries could be 

estimated, the area of the sample for both specimen designs was calculated as a function of the distance 

along the long axis of the specimen. For this calculation, the zero point was assumed to be the transition 

from the straight uniform gauge section to the curved shoulder region. Only one-half of the specimen was 

considered since the samples were symmetric about the center point of the samples. The shoulders were a 

segment of a circle with the center at zero along the length of the sample and offset by the radius of the 

circle from the edge or surface of the uniform gauge section. This allowed the specimen shoulder width 

(flat-sheet specimens) or radius (round specimens) to be expressed as a function of the distance from the 

zero point. Knowing the thickness and width of the sheet specimens or the radius for the round specimens 

allowed the area of the specimen in the shoulder to be calculated as a function of distance along the long 

axis of the specimen. 

Since the load was constant at all points in the specimen, the stress was calculated at each point in the 

transition region by dividing the applied load by the area of the cross section in the shoulders at that point. 

The creep rate for each point was estimated using the regression equation for the creep rate and the calculated stress at each point. The total displacement at each point was estimated by multiplying the creep rate 

by the specimen life. Summing the displacements at each point over the life of the test allowed the total 

strain in the transition region to be estimated. 

Figure 7 shows the calculated total strains for tests conducted at 500 °C/100 MPa, 650 °C/50 MPa, 

and 800 °C/25 MPa (932 °F/14.5 ksi, 1202 °F/7.2 ksi, and 1472 °F/3.6 ksi). These stresses were selected 

because they were near the high end of the experimental values tested. A high value was desired since the 

stress decreases as the area increases and could easily drop below previously tested stress levels. The 

selected values allowed the rates to be calculated over the entire stress range experienced by the transition 

regions based on experimental data without the need to extrapolate except for the lowest stresses. The 

worst case as defined by the most creep strain, 650 °C/50 MPa, had a strain of 1.0210–5 for the sheet 

specimen and 3.5610–6 for the round specimen. The strain in both specimen geometries was miniscule in 

comparison to the typical total strains of 0.05 to 0.15 at failure observed by the LVR or optical encoder. 

Because the frames measure displacement remote from the specimen, displacements also can be 

caused by creep of the fixtures, compliance of the machine, and other factors. Large-diameter, highly 

creep resistant materials were used for the fixtures. For the portion in the hot zone, no attempt was made 

to calculate the displacement from creep, since the area of the fixtures was at least 9 times greater than the 

maximum area of the test specimens and their creep rates were much lower at the test temperatures. On 

the basis of the low strain for the GRCop-84 in the transition regions and much lower creep rates for the 

fixture materials, it was assumed that the creep in the fixtures was nil. 
NASA/TM—2011-215493 16 

It is difficult to measure compliance of the machine without an extensometer or other direct measurement of the specimen strain, but the low loads («4.4 kN) on a 44.5-kN- (10 000-lbf-) capacity frame 

probably mean that the contribution from the compliance of the machine is minimal as well. In addition, 

the 0.2-percent-offset yield strength of GRCop-84 is low at the test temperatures (Ref. 4), and the onset of 

plastic deformation occurs at even lower stresses (as shown in the stress-strain curves in Ref. 4), so specimens tend to yield plastically rather than stretch elastically. The low loads and plastic deformation make 

the error from machine compliance small and probably unobservable. 

The only other known source of error in the displacement measurements is the slack in the load train, 

which leads to an initial large displacement upon loading. All subsequent strains have this initial strain 

incorporated into the measurement. Since that unknown initial strain value is constant for all subsequent 

displacement measurements, subtracting one measurement from another gives an accurate displacement 

between the two data points. If this difference is divided by the time between the two measurements, the 

creep rate can be accurately calculated in a manner similar to the calculation of creep rates for the Brew 

frames using the LVR data points. 

So although two different load frames were used to test two different specimen geometries, the 

differences in the remotely measured displacements, and hence the strains used to calculate the creep 

rates, should not affect the results. 

Step-loading vacuum creep testing was conducted to independently determine creep rate for comparison with the Brew frame data sets and to provide a calibrated standard for the creep test results. Varying 

the temperatures from test to test over a narrow range allowed for estimation of the deviation of the Brew 

frame sample temperature from the setpoint since the observed creep rate at a given load could be compared with the measured creep rate at various temperatures for that load. It also allowed for a better 

examination of the load and temperature sensitivity of GRCop-84 creep. 

A series of tests were conducted at temperatures of 500, 525, 550, 650, 675, and 700 °C (932, 977, 

1022, 1202, 1247, and 1292 °F) for specimens taken from sheet 3A1B. Table VII presents the load and 

hold times used. 

TABLE VII.—STEP-LOADING CREEP RATE TEST PARAMETERS 

Step Applied 

stress, 

MPa 

Time, 

h 

 Step Applied 

stress, 

MPa 

Time,

h 

 Step Applied 

stress, 

MPa 

Time, 

h 

500 °C (932 °F) 525 °C (977 °F) 550 °C (1022 °F) 

1 70.0 5 1 70.0 5 1 70.0 5 

2 73.0 5 2 73.0 5 2 73.0 5 

3 76.0 5 3 76.0 5 3 76.0 5 

4 79.0 5 4 79.0 5 4 79.0 5 

5 82.0 5 5 82.0 5 5 82.0 5 

 

6 85.0 3 6 85.0 3 6 85.0 3 

7 88.0 3 7 88.0 3 7 88.0 3 

8 91.0 3 8 91.0 3 8 91.0 3 

9 94.0 3 9 94.0 3 9 94.0 3 

10 97.0 3 10 97.0 3 10 97.0 3 

 

11 100.0 2 11 100.0 2 11 100.0 2 

12 103.0 2 12 103.0 2 12 103.0 2 

13 106.0 2 13 106.0 2 13 106.0 2 

14 109.0 2 14 109.0 2 14 109.0 2 

15 112.0 2 15 112.0 2 15 112.0 2 

 

16 115.0 1 16 115.0 1 16 115.0 1 

17 118.0 1 17 118.0 1 17 118.0 1 

18 121.0 1 18 121.0 1 18 121.0 1 

19 124.0 1 19 124.0 1 19 124.0 1 

20 127.0 1 20 127.0 1 20 127.0 1 
NASA/TM—2011-215493 17 

TABLE VII.—Concluded. 

Step Applied 

stress, 

MPa 

Time, 

h 

 Step Applied 

stress, 

MPa 

Time,

h 

 Step Applied 

stress, 

MPa 

Time, 

h 

650 °C (1202 °F) 675 °C (1247 °F) 700 °C (1292 °F) 

1 27.5 5 1 27.5 5 1 27.5 5 

2 30.0 5 2 30.0 3 2 30.0 3 

3 32.5 5 3 32.5 2 3 32.5 2 

4 35.0 5 4 35.0 1 4 35.0 1 

5 37.5 5 5 37.5 1 5 37.5 1 

 

6 40.0 3 6 40.0 1 6 40.0 1 

7 42.5 3 7 42.5 .5 7 42.5 .5 

8 45.0 3 8 45.0 .5 8 45.0 .5 

9 47.5 3 9 47.5 .5 9 47.5 .5 

10 50.0 3 10 50.0 .5 10 50.0 .5 

 

11 52.5 2 11 52.5 .2 11 52.5 .2 

12 55.0 2 12 55.0 .2 12 55.0 .2 

13 57.5 2 13 57.5 .2 13 57.5 .2 

14 60.0 2 14 60.0 .2 14 60.0 .2 

15 62.5 2 15 62.5 .2 15 62.5 .2 

 

16 65.0 1 16 65.0 .2 16 65.0 .2 

17 67.5 1 17 67.5 .2 17 67.5 .2 

18 70.0 1 18 70.0 .2 18 70.0 .2 

19 72.5 14 19 72.5 .2 19 72.5 .2 

20 75.0 1 20 75.0 .2 20 75.0 .2 

Results and Discussion 

This work was conducted in three sequential stages. During stage I, an initial survey was conducted to 

determine if data were identical among Brew frames. Stage II postulated various root causes, examined 

and identified the main sources of variation, and independently determined creep rate through step loading. In stage III, the final stage, tests were conducted to validate that each Brew frame provided similar 

results and to determine if the anticipated effects of processing could be discerned. The results of each 

stage are presented and discussed separately since each stage builds on the preceding stage. 

Stage I—Initial Survey 

The initial survey compared creep response as a function of Brew frame at a constant stress 

(105 MPa) and temperature (500 °C) for material from sheet 3HCS. Table VIII shows the results. The 

steady-state creep rate was determined using the slope of the linear portion of the creep curve that had the 

minimum slope. 

On the basis of the Monkman-Grant relationship (Ref. 8), the data plotted as log10(creep life), with the 

life in seconds, versus log10(creep rate), with the rate in reciprocal seconds, should produce a straight line. 

Figure 8 shows that the data set follows the Monkman-Grant relationship well. The Monkman-Grant 

model was used to check the data for outliers since it is independent of stress and temperature—the two 

independent variables typically most influential in creep testing. A least-squares linear regression with a 

two-way 95-percent confidence interval was conducted to determine the power law equation for the 

observed relationship. The fit of the data indicates that there are no apparent outliers. Therefore, all the 

tests were considered to be valid. 
NASA/TM—2011-215493 18 

TABLE VIII.—CREEP BEHAVIOR AS A FUNCTION OF BREW FRAME 

[Stress, 105.0 MPa; setpoint temperature, 500 °C.] 

Test Brew 

frame 

Specimen number 

of sample 3HCS 

Steady-state 

creep rate, 

s

–1

Creep 

life, 

h 

Total strain, 

percent 

2A 1 22T 1.3210–5 2.7 18.6 

3A 1 4T 8.1110–6 4.6 21.5 

1A 1 20L 5.7110–6 5.6 22.8 

6A 2 1T 1.5210–4 .4 37.4 

4A 2 25L 8.0810–5 .6 28.7 

5A 2 13T 7.9910–5 .6 33.9 

 

7A 3 15L 4.3610–5 1.0 22.1 

9A 3 26L 3.5610–5 1.1 18.3 

8A 3 1L 1.8010–5 2.2 25.8 

11A 4 14L 3.1410–5 1.2 27.4 

10A 4 15T 2.9810–5 1.4 32.0 

1 4 24T 2.3910–5 1.5 23.6 

 

12A 4 20T 2.6710–5 1.5 31.8 

2 4 22L 1.7610–5 1.9 25.3 

14A 5 18L 6.8110–6 4.7 33.1 

13A 5 17L 2.2710–7 69.5 33.6 a

15A 5 7L 1.8210–8 330.6 20.7 

a

Sample 15A–7L did not break, but the test was stopped after 330.6 h. 
NASA/TM—2011-215493 19 

However, when creep life at 500 °C/105 MPa was plotted by Brew frame, as shown in Figure 9, each 

frame had a different mean creep life. If test conditions had been identical, the lives should have been 

similar with a small standard deviation for the log10(creep life) values. The absolute range for each Brew 

frame on a logarithmic scale was similar among the frames. The single exception was for Brew frame 5, 

which had one test that did not fail after 330.6 h, the time when the test was terminated. Figure 10 shows 

creep curves for all the tests conducted at 500 °C/105 MPa. The variability in the shapes and magnitudes 

of the curves confirms that each Brew frame had a very different response. This indicates that the test 

conditions were not identical as assumed. 
NASA/TM—2011-215493 20 

A one-way ANOVA was conducted on the life data using log10(creep life) as the dependent variable 

to confirm that there was a statistical difference in the means of the five Brew frames as suspected. 

Table IX shows the results. An SNK test was used to compare the means. As expected, Brew frame 5 had 

a much different mean, but there were also statistical differences detected between Brew frame 1 and 

Brew frames 2, 3, and 4. Brew frames 2, 3, and 4 gave statistically equal results. 

This result for identical test conditions of specimens with the same process history was not expected. 

It was reasoned that whatever was causing this difference was responsible for much of the variation seen 

in the data shown in Figures 1 and 2. Since the specimen material is assumed to have been identical, each 

Brew frame should have had a similar creep curve, creep rate, and creep life for a given temperature and 

creep stress. For the secondary, or steady-state, slope to change significantly, either temperature or stress 

would need to be different among these tests. Figure 11 illustrates how the creep response would change 

with increasing temperature or stress (Ref. 9). 

These results clearly demonstrate the need for additional testing to determine the source of variation 

and to develop corrective actions to improve the creep tests. 

TABLE IX.—RESULTS OF ONE-WAY ANALYSIS OF VARIANCE (ANOVA) FOR CREEP LIVES

Source of variation Degrees of 

freedom 

Sum of squares Mean 

squares 

F valuea

 Probability 

Between subjects 2 0.122 0.0610 ------- ------- 

Between treatments 4 3.062 .765 7.463 0.011 

Residual 7 .718 .103 ------ ------- 

Total 13 4.119 .317 ------ ------- 

a

The value of the F distribution for X degrees of freedom associated with the numerator and Y degrees of 

freedom associated with the denominator. 
NASA/TM—2011-215493 21 

Stage II—Identifying Sources of Variation 

Constant-load vacuum creep testing.—Investigations of various root causes that could produce 

different creep responses were conducted. Although these were centered on temperature and load, other 

test aspects were also considered. A detailed review of the creep test procedure, thermocouple and 

instrumentation calibration, friction in the mechanical feedthrough ball joint, and resistance from the 

weight of the rubber cooling hose hanging on the upper mechanical feedthrough ball joint were examined 

qualitatively and, when possible, quantitatively. They were found not to be detectably different among 

test frames on the basis of a review of the calibration data sheets, calibrated simulated input data, measured loads, and when quantitative measurements were not possible, optical and physical examination of 

the frames. Minor changes, such as supporting the hoses closer to the upper pull rod, produced no 

detectable changes in loads or temperatures. 

Operator error was assumed to be random, but some systematic factors could have been introduced 

through an operator bias. A systematic error would result in a shift of the creep life to a higher or lower 

value, but it should have been relatively uniform and not have affected the scatter in the data set. Differences between various operators were recognized as a potential source of error and scatter. Only one 

operator was used in this study to remove operator-to-operator variability from the possible sources of 

errors. 

The measurement technique for the displacements was also examined. Measuring the displacement 

by measuring the travel of the load train introduces a random variability from equipment measurement 

errors. There is also a systematic error because the total displacement includes deformation outside of the 

gauge section. However, as was pointed out in the Experimental Procedure section, these displacements 

are miniscule in comparison to the displacements from creep. The errors in measurement would tend to be 

small, a few percent or less of the value, and would only affect the creep rates, not the creep lives. Since 

the creep rates and the creep lives had similar variability, the contribution of this assumption was shown 

to be small relative to the observed variability. As such, the measurement of the displacements remotely 

was not considered to be a primary source of variability. 

It was determined that, to examine temperature and load differences, additional instrumentation 

would be needed. On Brew frames 3 and 4, provisions were made to install thermocouples to monitor the 

specimen temperature, and a load cell was placed in the lower pull rod. In addition, provisions were made 

to pressurize the chamber of Brew frame 3 with helium at slightly less than 1 atm in an attempt to try to 

improve the heat transfer rate and uniformity from the heater to the sample. These frames then served as 

prototypes for potential equipment upgrades to the load frames. A series of tests to examine the variability 

in loads and temperatures were conducted at various temperatures and loads following the same setup and 

test procedures for each test. Table X shows the test conditions and results. 

Test 1 was conducted on Brew frame 3 with a temperature setpoint of 500 °C (932 °F) and an initial 

stress of 105 MPa (15.2 ksi). The load was monitored periodically and was a constant value throughout 

the test. The load was compared with the balance pan weight and differed slightly, which was attributed 

to errors arising from zeroing the electronics, friction in the load train, and the inherent error of the load 

cell based on the load cell specifications. Because the load did not vary during the test and the small load 

differences would not explain the different shaped creep curves, the main emphasis shifted to temperature 

differences. The subsequent creep tests confirmed the accuracy, reproducibility, and stability of the loads. 

During test 1, the specimen temperature was found to be 43 °C (77 °F) higher than the controller 

setpoint, and from top to bottom along the sample, a temperature gradient of 44 °C (79 °F) was observed. 

Such a large temperature difference and gradient would have a profound effect on the creep properties, 

and they could produce great variability if either varied from test to test or from frame to frame while the 

setpoint remained the same. However, if the temperature difference and the thermal gradient were both 

consistent for all tests, they would not produce variability in the creep responses. Instead, they would 

introduce a systematic error into the creep testing. The creep rate would be higher and the creep life 

shorter than for a sample tested at the proper, lower sample temperature. 
NASA/TM—2011-215493 22 

TABLE X.—GRCop-84 SHEET SPECIMEN TEST CONDITIONS AND RESULTS 

USED FOR IDENTIFYING SOURCES OF VARIATION 

(a) Test comments. 

Test 

1 First run with top, middle, and bottom thermocouple; control thermocouple might be as much as 1/8 to 1/4 in. away 

from sample 

2 Control thermocouple located 1/16 in. away from specimen center; grips not symmetrical 

3 Control thermocouple located 1/16 in. away from specimen center; grips not symmetrical and different from Brew 

frame 3 combination 

4 Control thermocouple located 1/16 in. away from specimen center; grips not symmetrical 

5 Control thermocouple located 1/16 in. away from specimen center; grips not symmetrical and different from Brew 

frame 3 combination 

 

6 Control thermocouple located 1/16 in. away from specimen center; grips not symmetrical 

7 Control thermocouple located 1/16 in. away from specimen center; grips not symmetrical and different from Brew 

frame 3 combination 

8 Symmetrical grips and control thermocouple 1/16 in. away from specimen center 

9 Symmetrical grips and control thermocouple 1/16 in. away from specimen center 

 

10 Helium atmosphere 

11 Helium atmosphere 

12 Helium atmosphere 

13 Helium atmosphere 

14 Helium atmosphere 

15 Helium atmosphere 

(b) Conditions and results.

Test Brew 

frame 

Specimen Applied 

stress, 

MPa 

Setpoint 

temp., 

°C 

Average temperature of 

various thermocouples, 

°C 

Temperature difference 

between thermocouples, 

°C 

Life, 

h 

Creep rate,

s

–1

Top Middle Bottom Bottom to 

top 

Bottom to 

control 

1 3 12 1a5c-350 105 500 499.3 537.0 542.9 43.6 42.9 2.30 1.2210–5

2 3 15 1a5c-450 105 500 499.2 523.7 532.8 33.6 32.8 1.62 1.9910–6

3 4 18 1a3b-450 105 500 499.1 504.1 511.0 11.9 11.0 1.85 1.5510–6

4 3 16 1a3c-450 17 800 800.0 805.2 812.0 12.0 12.0 24.62 2.3110–7

5 4 17 1a5c-450 17 800 799.4 826.1 832.1 32.8 32.1 5.04 7.1810–6

 

6 3 24 1a3b-450 44.3 650 649.9 658.8 668.1 18.2 18.1 4.98 3.8210–6

7 4 22 1a5c-350 44.3 650 649.5 674.9 682.1 32.6 32.1 8.78 1.5310–6

8 3 14 1a3c-450 105 500 499.6 529.2 534.1 34.5 34.1 4.82 6.5410–6

9 4 21 1a5c-350 105 500 499.2 524.3 537.0 37.8 37.0 6.18 4.0810–6

 

10 3 20 DLE4a2 105 500 499.2 500.6 506.6 7.3 6.6 4.57 6.3410–6

11 3 21 DLE4a2 105 500 499.0 496.9 504.7 5.7 4.7 6.74 5.7910–6

12 3 22 DLE4a2 44.3 650 649.4 643.7 653.4 4.1 3.4 3.11 3.9310–6

13 3 23 DLE4a2 44.3 650 649.4 643.1 651.8 2.3 1.8 3.68 3.1910–6

14 3 24 DLE4a2 17 800 Specimen broke on loading 

15 3 25 DLE4a2 17 800 800.1 794.0 801.6 1.5 1.6 5.35 1.9610–6
NASA/TM—2011-215493 23 

Possible reasons for the large temperature differences between the control and specimen were 

investigated first. It was identified that the control thermocouple was not being located precisely relative 

to the specimen. The farther the control thermocouple was from the specimen, the greater the temperature 

difference was between the control thermocouple and the specimen. Ideally, the control thermocouple 

should touch the specimen, but this is not suitable for these tests because a bending moment would be 

introduced. The closest practical control thermocouple position was 1.57 mm (0.062 in.) away from the 

specimen when the specimen was at room temperature. A 1.57-mm feeler gauge was fabricated to aid 

in consistent positioning, and the test procedure was modified to include positioning the chamber thermocouple with the feeler gauge. All subsequent tests were run with the control thermocouple at this standoff 

distance. 

Test 2 was again conducted on Brew frame 3 using a 500 °C (932 °F) setpoint and a 105-MPa 

(15.2-ksi) initial stress. The control thermocouple standoff was set using the new thermocouple positioning technique. The specimen temperature was 33 °C (59 °F) higher than the controller setpoint, and the 

temperature gradient along the specimen was 34 °C (61 °F). This was a slight improvement over the first 

test, but it also highlighted the potential variability of both the temperature difference and the thermal 

gradient with even small changes in control thermocouple position. 

Test 3 was conducted on Brew frame 4 still using a control thermocouple setpoint of 500 °C (932 °F) 

and an initial stress of 105 MPa (15.2 ksi). The same feeler gauge used for the Brew frame 3 test was used 

to position the control thermocouple. The specimen temperature was 11 °C (20 °F) higher than the controller setpoint, and the temperature gradient along the specimen was 12 °C (22 °F). This showed that 

even when the same gap was used there were large differences in the specimen temperature and gradient 

from load frame to load frame. 

Tests 4 and 5 were conducted using a control thermocouple setpoint of 800 °C (1472 °F) and an 

initial stress of 17 MPa (2.5 ksi) employing Brew frames 3 and 4, respectively. The higher temperature 

was expected to improve radiative heat transfer and to minimize temperature differences and gradients. 

For Brew frame 3, the specimen temperature was 12 °C (22 °F) higher than the controller setpoint, 

and the temperature gradient along the specimen was 12 °C (22 °F). For Brew frame 4, the specimen 

temperature was 32 °C (58 °F) higher than the controller setpoint, and the temperature gradient along the 

specimen was 33 °C (59 °F). These differences and gradients had Brew frame 4 now being hotter with a 

larger gradient than Brew frame 3. This is the opposite of the results obtained in tests 2 and 3. This highlighted the frame-to-frame and specimen-to-specimen variability in temperature that can contribute to the 

variability in creep rate and creep life. 

Tests 6 and 7 were conducted using a control thermocouple setpoint of 650 °C (1202 °F) and an initial 

applied stress of 44.3 MPa (6.4 ksi) employing Brew frames 3 and 4, respectively. For Brew frame 3, the 

specimen temperature was 18 °C (32 °F) higher than the controller setpoint, and the temperature gradient 

along the specimen was 18 °C (32 °F). For Brew frame 4, the specimen temperature was 32 °C (58 °F) 

higher than the controller setpoint, and the temperature gradient along the specimen was 33 °C (59 °F). 

From these seven tests, it was observed that the specimen temperature and temperature gradients were 

different between Brew frames 3 and 4 despite the consistent location of the control thermocouple. In 

both Brew frames, the specimen temperature was consistently hotter than the setpoint control temperature, and the temperature gradient along the length of the specimen from top to bottom was consistently 

hottest at the bottom of the gauge length and coldest at the top. The top thermocouple most closely 

matched the control thermocouple. The differences and the gradients were also noted to vary even within 

a single Brew frame. 

It was observed that the specimen grip fixtures were not fully symmetrical and that the tooling 

arrangement was different between Brew frames 3 and 4. It was reasoned that the different mass of the 

grips that are attached to the water-cooled pull rods transferred heat unevenly away from the specimen 

and affected the specimen temperature distribution. Similarly, the different tooling arrangement between 

Brew frames 3 and 4 would produce different temperature distributions between the frames. Sufficient 

310 stainless steel tooling was fabricated using the same designs so that the top and bottom tooling would 

be identical in all of the Brew frames to eliminate this potential source of variation. 
NASA/TM—2011-215493 24 

Tests 8 and 9 were conducted using symmetrical grips, again using a control thermocouple setpoint of 

500 °C (932 °F) and an initial stress of 105 MPa (15.2 ksi) employing Brew frames 3 and 4, respectively. 

The specimen temperature gradients were now similar for both Brew frames. For Brew frame 3, the 

specimen temperature was 34 °C (61 °F) higher than the controller setpoint, and the temperature gradient 

along the specimen was 35 °C (63 °F). For Brew frame 4, the specimen temperature was 37 °C (67 °F) 

higher than the controller setpoint, and the temperature gradient along the specimen was 38 °C (68 °F). 

On the basis of these results, which indicate that grip material and geometry can affect the specimen 

gradients, the testing procedure was updated to require symmetrical fixturing made from the same 

material. This change was implemented for all subsequent testing. Through trial and error it may be 

possible to use asymmetrical grips to reduce the gradient, but the potential improvement was not 

addressed in this study because of limited time and resources. In addition, asymmetrical grips were not 

considered to be suitable for standard testing given the wide range of materials tested in the laboratory, 

their greatly varying test temperatures, and the normally low volume of samples. 

The heaters were also examined. The bottom of each heater had a solid band rather than a mesh (see 

Fig. 6(b)). The top also had a band, but it was split and attached to water-cooled Cu electrical power 

supplies that would tend to cool the top. The authors believe that the solid band on the heater base 

provided a higher heat flux to the sample and pull rods than the open mesh provided. The upper section 

was in contact with the water-cooled electrical feedthroughs, which may have reduced its heat flux as 

well. The solid heater used for the step-loading calibration runs had a thermal gradient typically less than 

5 °C (9 °F) as measured by the three attached thermocouples. This improvement, while using very similar 

geometries for the heaters, indicates that the design of the heater is important. 

The mesh heaters also tended to bow and otherwise deform more than the solid heater. This variability of the gap between the sample and the heater introduced some variability into the temperature and 

specimen temperature gradients as well. Solid heaters with a more consistent geometry from heater to 

heater that retain that geometry from specimen to specimen could help to reduce the observed variability. 

Replicate tests were not conducted because of equipment schedule constraints, but the specimen 

temperature gradients were now similar for both Brew frames and were believed to be reproducible. 

Although the thermal gradients were larger than desired, the consistency of the gradients means that they 

will not contribute to the variability in the creep rates and lives. Additional work is planned to determine 

if the gradients can be reduced through improved heater design and if the applicable ASTM International 

standards for temperature gradients in creep tests can be met. 

Tests 10 to 15 were conducted in Brew frame 3 with a helium atmosphere and symmetrical grips. 

Duplicate tests were performed with the control thermocouple setpoint and pressure at 500 °C/105 MPa, 

650 °C/44.3 MPa, and 800 °C/17 MPa (932 °F/15.2 ksi, 1202 °F/6.4 ksi, and 1472 °F/2.5 ksi). No corrections were made to match the specimen temperatures to the setpoint temperatures. 

When helium was used, the average specimen temperature for all of these conditions was 4 °C (7 °F) 

higher than the controller setpoint. The temperature gradient along the specimen was reduced to 4 °C 

(7 °F). The helium clearly improved the specimen temperature uniformity and more closely matched the 

sample temperature to the control thermocouple temperature. 

Although the life and creep results do not lend themselves well to statistical analysis because there are 

no true repeats, general observations can be made regarding the differences in what should be identical 

tests. The creep rates and lives for the 500 °C/105 MPa and 800 °C/17 MPa tests with fairly large temperature differences and gradients varied by almost 1 order of magnitude, and the creep life varied by 

about the same amount. Even the 650 °C/44.3 MPa data set, with its more consistent temperatures and 

gradients, shows a 1/2 order of magnitude variability. In contrast, the subset of tests conducted in a 

helium atmosphere with small temperature gradients and differences had reproducible creep rates and 

creep lives. This underscores the effect of temperature on the variability of creep properties. 
NASA/TM—2011-215493 25 

Step-loading vacuum creep testing.—Concurrent with the constant-load creep testing, step-loading 

vacuum creep testing was conducted to independently determine creep rate as a way to establish a standard for the constant-load tests. The variations in temperature and stress investigated would also assist in 

quantifying the sensitivity of GRCop-84 to small changes in temperature and stress. 

A limited series of tests were conducted at the temperatures and loads listed in Table VII using sheet 

specimens from the same source as the Brew frame creep test specimens. During these tests, the combination of loads and hold times generated failure of the samples in all cases before the entire series of 20 

steps were completed. This did not affect the validity of the tests or the results, but it did result in a 

smaller data set for each test. 

The results of the step-loading creep tests are presented in Table XI and plotted in Figure 12. A leastsquares-fit power law regression was used to estimate the creep rate versus stress. The general form of the 

equation used is 

n    A (3) 

where  is creep rate in reciprocal seconds, A is a constant, σ is load (in MPa), and n is the stress exponent. Table XII presents values for A and n. Included for comparison are the results for an as-extruded 

GRCop-84 bar and an annealed 6.5-mm (0.255-in.) GRCop-84 plate from other creep testing during this 

program (Refs. 2 and 10). The results for n indicate that power law creep predominated as expected. 

Several important points are illustrated in Figure 12. First, the creep rate is highly sensitive to 

temperature. A 50 °C (90 °F) temperature increase results in a 1 order of magnitude increase in the 

steady-state creep rate. 

TABLE XI.—STEP-LOADING CREEP TEST RESULTS FOR VARIOUS TEMPERATURES 

Step Stress, 

MPa 

Creep rate, 

s

–1

 Step Stress,

MPa 

Creep rate, 

s

–1

 Step Stress, 

MPa 

Creep rate, 

s

–1

500 °C (932 °F) 525 °C (977 °F) 550 °C (1022 °F) 

1 70.0 3.210–7 1 70.0 5.810–7 1 69.9 2.310–6

2 73.0 3.010–7 2 73.0 4.510–7 2 72.9 3.310–6

3 76.0 3.010–7 3 75.9 4.910–7 3 75.9 5.610–6

4 79.0 3.510–7 4 78.9 5.810–7 700 °C (1292 °F) 

5 82.0 4.210–7 5 81.9 9.710–7 1 27.5 1.410–6

6 85.0 4.710–7 6 84.9 1.410–6 2 30.0 2.110–6

7 88.0 5.910–7 7 87.9 2.210–6 3 32.5 4.110–6

8 91.0 7.610–7 675 °C (1247 °F) 4 35.0 9.810–6

9 93.9 9.810–7 1 27.4 9.510–7 

10 96.9 1.310–6 2 29.9 1.010–6 

11 99.9 1.910–6 3 32.3 1.510–6 

12 102.9 3.010–6 4 34.8 2.910–6 

650 °C (1202 °F) 5 37.3 4.510–6 

1 27.5 3.410–7 6 39.8 8.110–6 

2 30.0 4.110–7 7 42.1 1.710–5 

3 32.5 4.310–7 

4 35.0 6.310–7 

5 37.5 9.510–7 

6 37.5 1.610–6 

7 42.5 3.010–6 
NASA/TM—2011-215493 26 

TABLE XII.—REGRESSION COEFFICIENTS AND 

GOODNESS OF FIT

Creep test 

temperature 

Sample type Equation coefficients,a

n    A

R2

°C °F A n 

500 932 Rolled 1.1-mm sheet 3A1B 8.0610–22 7.66 0.89 

525 977 5.3510–27 10.59 .98 

550 1022 2.2110–26 10.85 .99 

650 1202 1.3610–14 5.05 .87 

675 1247 1.0910–16 6.80 .94 

700 1292 2.5710–18 8.11 .96 

Prior testing 

500 932 Extruded 25-mm bar 7.4810–32 12.10 0.98 

500 932 Rolled 6.5-mm plate 1.2510–24 8.78 .96 

650 1202 Rolled 6.5-mm plate 1.9310–21 8.86 .94 

800 1472 Rolled 6.5-mm plate 1.1510–16 7.45 .91 

a , creep rate; A, constant;, load; n, stress exponent; R2

, simple coefficient of 

determination.

The creep rate is less sensitive to stress. A 10-MPa (1.4-ksi) change in stress would result in about 

1/10 order of magnitude change in steady-state creep rate. Generally, the load is controllable to less than 

4.4 N (1 lbf) using the dead weights. Although the variability in stress depends on specimen cross section, 

the variability in stress is normally 0.1 MPa (0.01 ksi) or less. Results of the sensitivity analysis indicate 

that this uncertainty in stress should not produce observable variability. 

These results indicate that GRCop-84 creep behavior is more sensitive to small temperature changes 

than was previously appreciated. It also demonstrates that temperature is more likely than stress to be the 

source of the observed variability. 
NASA/TM—2011-215493 27 

Stage III—Verification 

GRCop-84 sheet.—The preceding findings show that the specimen could be much hotter than the 

control thermocouple and that GRCop-84 creep behavior is more temperature sensitive than expected. We 

concluded that thermocouples had to be tied directly to the specimens in such a way that the beads were 

in intimate contact with the specimen to minimize the difference between the temperature controller 

setpoint temperature and the specimen temperature. The attachment also had to be done without having 

the thermocouples introduce a bending moment or other spurious stresses in the specimens. 

A design using thermocouples with small alumina beads introduced from the bottom of the chamber 

was selected. Stainless steel wire was used to tie the thermocouples to the lower pull rod and the sample. 

As shown in Figure 6(c), the beads were in intimate physical contact with the sample. The flexible nature 

of the thermocouples minimized bending moments. The weight of the thermocouples was almost entirely 

supported by tying them to the lower pull rod to minimize any contribution that they may have had to the 

loads on the sample. Vacuum thermocouple feedthroughs were added to an existing port, and the three 

thermocouples were tied into the data acquisition system for recording. 

Modifications to add thermocouples that could be tied directly to the specimens for each Brew frame 

were expected to permit more accurate and consistent specimen temperature control from test to test and 

from frame to frame. Therefore, a series of tests in vacuum and helium were conducted to demonstrate the 

improvements and to determine if the variations observed between Brew frames had been eliminated. 

Table XIII presents the results. 

The vacuum and helium test creep life data for all test conditions were plotted against creep rate on a 

log-log scale and are shown in Figure 13. The Monkman-Grant relationship was again used to determine 

if there were any obvious outliers. Also included are data from stage I of the creep testing (original procedure) for comparison. Again, the new data fall on a straight line, indicating that there are no outliers. 

The least-squares fit for the power law equation is also nearly identical to the prior results, indicating that 

the changes in test methods and procedures did not change the creep behavior of GRCop-84. 

Figure 14 shows the creep life data from the vacuum and helium tests plotted as a function of Brew 

frame for each test temperature. The term “stage I survey” in the figure refers to testing done in stage I to 

establish if there were detectable differences between creep frames. Because of an error in testing, two 

samples were run at 800 °C (1472 °F) using the stress for the 650 °C (1202 °F) tests. Both samples failed 

upon loading. The limited results obtained are reported in Table XIII but are not used in any statistical 

comparisons. Enough data points for the temperature gradients and sample temperature were gathered 

during the preloading period to provide meaningful data. 

Visual observation from Figure 14 suggests that Brew frames 2, 3, and 4 were equal. To confirm this, 

the creep rate and creep life data for each frame were analyzed statistically to determine if there was a 

difference in the data sets. A one-way ANOVA was attempted, but the data sets all lacked constant 

variance, so a Kruskal-Wallis one-way analysis of variance by ranks was used instead. This method 

generates a test statistic H and determines if it is sufficiently large to reject the null hypothesis that the 

data from all three Brew frames are equivalent. The distribution of the actual data is used for small data 

sets such as these to determine a critical value for H. 
NASA/TM—2011-215493 28 

TABLE XIII.—GRCop-84 CREEP RESULTS AFTER MODIFICATIONS 

Run Brew 

frame 

Sample Set-point 

temperature, 

°C 

Actual 

stress, 

MPa 

Life, 

h 

Creep 

rate, 

s

–1

Elongation, 

percent 

Temperature measured by 

various thermocouples, 

°C 

Top Middle Bottom Old 

control 

Vacuum tests 

1 2 1 500 105.8 11.2 1.810–6 8.4 513.0 498.7 512.3 438.9 

2 2 2 500 105.2 17.5 1.110–6 8.5 497.2 502.2 504.1 432.5 

7 3 7 500 106.4 10.4 1.210–6 5.5 496.8 499.6 502.1 491.9 

8 3 8 500 105.2 26.6 8.510–7 11.8 502.0 501.8 506.9 499.4 

13 4 13 500 105.5 21.0 8.210–7 14.7 495.4 497.6 509.8 447.2 

14 4 14 500 105.6 31.8 6.610–7 15.2 494.5 497.3 500.3 441.6 

 

3 2 3 650 44.2 13.7 1.310–6 9.3 645.3 653.1 651.4 603.5 

4 2 4 650 44.5 10.7 1.610–6 8.6 639.6 649.2 654.0 609.2 

9 3 9 650 44.9 14.6 1.110–6 9.4 640.3 649.0 651.7 648.9 

10 3 10 650 44.2 7.1 1.610–6 7.2 643.2 651.5 653.3 653.2 

15 4 15 650 43.9 7.2 2.510–6 18.0 645.9 651.9 656.9 600.2 

16 4 16 650 44.3 10.3 1.610–6 15.4 638.1 647.9 653.3 600.7 

 a

5 2 5 800 44.5 0.1 1.410–3 49.5 790.9 799.8 804.1 778.8 

6 2 6 800 16.9 111.4 7.110–8 6.8 791.5 800.5 806.4 781.5 

11 3 11 800 16.8 135.9 6.110–8 9.1 793.4 798.9 797.2 797.9 

12 3 12 800 16.9 12.2 1.310–6 12.1 792.1 798.8 798.1 805.0 

17 4 17 800 16.9 356.2 4.410–8 22.1 790.5 797.6 805.3 761.2 

18 4 18 800 16.6 33.6 4.710–7 20.6 792.6 799.4 804.8 754.7 

Helium atmosphere tests 

1 2 1 500 105.5 7.2 3.510–6 11.2 509.0 500.4 493.6 507.0 

2 2 2 500 105.4 6.4 3.8010–6 10.7 505.6 500.3 492.1 505.8 

7 3 3 500 104.8 11.7 3.410–6 27.7 497.1 499.0 500.6 501.0 

8 3 4 500 105.5 7.0 4.210–6 13.9 496.3 500.7 501.2 504.9 

13 4 5 500 105.7 7.9 4.110–6 14.0 491.1 498.2 499.9 456.7 

14 4 6 500 104.8 8.8 4.010–6 16.2 494.5 500.2 501.1 461.9 

 

3 2 7 650 44.5 7.3 1.410–6 5.4 655.9 649.5 636.8 653.0 

4 2 8 650 44.7 5.3 2.110–6 5.7 658.6 650.1 637.3 656.3 

9 3 9 650 44.3 7.8 2.210–6 9.0 648.4 649.5 652.2 654.6 

10 3 10 650 44.1 7.1 2.410–6 9.1 646.5 650.4 649.1 650.7 

15 4 11 650 44.3 10.9 1.210–6 7.6 647.8 652.2 654.7 603.5 

16 4 12 650 44.7 11.9 1.010–6 6.9 644.8 650.0 651.9 601.9 

 a

5 2 13 800 44.7 0.1 (b) 40.6 799.6 800.3 795.9 797.0 

6 2 14 800 17.6 16.5 4.2610–7 5.8 797.8 797.2 794.3 797.0 

11 3 15 800 17.6 6.0 1.410–6 4.6 799.5 801.0 800.4 804.1 

12 3 16 800 17.6 11.1 7.010–7 4.2 798.9 801.1 801.5 805.7 

17 4 17 800 16.5 56.6 1.310–7 4.1 793.1 801.3 801.9 752.4 

18 4 18 800 17.3 22.0 4.310–7 5.2 798.4 802.2 802.3 754.2 a

Sample failed prematurely upon loading because of the wrong test load. 

b

Not determined. 
NASA/TM—2011-215493 29 
NASA/TM—2011-215493 30 
NASA/TM—2011-215493 31 

Table XIV shows the H values generated from the statistical analysis and the probability P of falsely 

rejecting the null hypothesis that all the data sets are equivalent. The lower the value of P, the greater the 

probability that the data sets are different. A lower limit of 0.05 was selected for P to give a 95-percent 

confidence in accepting the null hypothesis. In all cases, the analysis indicated that the data sets were 

statistically equal, but the analysis for the creep rate at 650 °C (1202 °F) in helium has a P value of only 

0.067. This indicates that the data are very close to being statistically significantly different. 

On the basis of this analysis, Brew frames 2, 3, and 4 produced statistically equal results for creep rate 

and creep life in the same environment. This allowed the data sets for each test temperature for both 

environments to be pooled into two data sets for each test temperature. The averages of the data sets were 

compared using a t-test to determine if they were statistically equal. If the tests for normality or equal 

variance failed, the data sets were compared using a Mann-Whitney Rank Sum Test. Table XV shows the 

results. The 500 °C (932 °F) data show a difference between the vacuum and helium creep rate and creep 

life, but the 650 °C (1202 °F) and 800 °C (1472 °F) data sets are statistically equal. 

TABLE XIV.—STATISTICAL RESULTS FOR 

COMPARISON OF BREW FRAMES 

[Two degrees of freedom.] 

Property Test 

temperature,

°C 

Test 

statistic,

H 

Equal? Probability,

P 

Vacuum tests 

Creep life 500 2.000 Yes 0.533 

650 1.143 Yes .667 

800 .400 Yes .933 

Creep rate 500 3.714 Yes .200 

650 2.258 Yes .533 

800 .400 Yes .933 

Helium atmosphere tests 

Creep life 500 2.000 Yes 0.533 

650 3.714 Yes .200 

800 3.600 Yes .200 

Creep rate 500 1.143 Yes .667 

650 4.571 Yes .067 

800 3.000 Yes .330 

 TABLE XV.—STATISTICAL COMPARISON OF AVERAGE 

VACUUM AND HELIUM CREEP PROPERTIES 

(a) Creep life. 

Test 

temperature 

Normality Equal 

variance 

Average/median 

of log10(life) 

Equal? Probability,

P 

°C °F Vacuum Helium 

500 932 Pass Pass 1.275 0.903 No 0.001 

650 1202 Fail ----- 0.935 0.878 Yes 0.937 

800 1472 Pass Pass 1.869 1.227 Yes 0.066 

(b) Creep rate. 

Test 

temperature 

Normality Equal 

variance 

Average/median 

of log10(life) 

Equal? Probability,

P 

°C °F Vacuum Helium 

500 932 Pass Fail –6.015 –5.456 No 0.002 

650 1202 Fail ----- –5.796 –5.766 Yes 0.589 

800 1472 Pass Pass –6.787 –6.326 Yes 0.205 
NASA/TM—2011-215493 32 

The higher creep rate and shorter life of the helium environment samples was unexpected since the 

average temperature of the helium specimens was 2.8 °C (5.0 °F) lower and the average gradient was 

5.2 °C (9.4 °F) less. With six samples in each data set, the sample size should not have played a major 

role. The cause for this difference remains unknown and will need to be investigated further in the future. 

The helium tests shown in Table XIII did not show a clear benefit for helium over the vacuum in 

reducing the specimen temperature gradient as was expected. It is likely that small differences in furnace 

construction and/or geometry between each Brew frame had the biggest remaining effect on the observed 

temperature gradients. 

The creep rates for all tests are plotted by Brew frame at each test temperature in Figure 15. As 

expected from the strong dependency of the life on rate, after the modifications, all Brew frames have 

very similar averages and ranges for both the vacuum and helium tests. Data from the stage I tests also are 

shown in the plots and are designated “stage I survey.” In addition, the creep rates independently determined by step loading at 500, 525, 550, 650, 675, and 700 °C (932, 977, 1022, 1202, 1247, and 1292 °F) 

are shown in the plots. After the modifications, there was good agreement between creep rates determined 

by constant loading and step loading. The comparison of creep rates also gives an indication of the actual 

specimen temperature for the stage I tests. The survey creep rate can be compared with the creep rates 

determined by step loading at the various temperatures. It is likely that the survey sample was near the 

temperature with an equal creep rate. The comparisons indicate that the survey samples were generally 

hotter than the setpoints. These differences in temperature are consistent with the results from the stage II 

measurement of specimen temperatures. 

The creep curves shown in Figure 16 for vacuum tests and in Figure 17 for helium now have similar 

shapes and magnitudes, confirming that the test conditions are as close to identical as is practical to 

achieve in the Brew frames. This is in sharp contrast to Figure 9, where there is no consistency in the 

creep curves. The remaining small differences are explained mostly by the slight variation in specimen 

temperature, the temperature gradient along the specimen, and inherent material and testing variability. 

GRCop-84 plate testing.—The results for the vacuum creep testing of GRCop-84 plate made from 

–140 mesh powder are plotted in Figures 18 and 19. With the larger number of tests and the wider range 

of stresses, a more accurate assessment of the remaining variability for the creep properties can be made. 

Unlike the 2 orders of magnitude scatter for the survey creep test and the 4 orders of magnitude 

scatter for the combined historical data (Fig. 2), the new data set has a scatter of less than 1 order of 

magnitude for both the creep rate and the creep life. Even the two-sided 95-percent confidence interval 

(dashed lines in figures) has a range of less than 1 order of magnitude now. 

The creep rate determined using TM–4 for GRCop-84 plate, which was presented in Figure 12, is also 

plotted in Figure 18. The Brew frame results show excellent agreement with the data for TM–4 and a 

similar scatter. This reinforces the concept of using TM–4 as a standard. It also demonstrates the ability to 

generate reproducible and consistent data among multiple Brew frames using the updated creep-testing 

procedures. 

The improved data set was used to quantitatively estimate the dependency of the creep rate and creep 

life on stress and temperature. The model presented in Equation (1) was examined first with forward 

stepwise regression. Problems were encountered with the statistical analysis program because both the 

log10() and 1/(T + 273) terms had equal F values to enter. Backward stepwise regression was used to 

examine the model. The log10()/(T + 273) term was eliminated in the first step, and the log10() and 

1/(T + 273) terms were both determined to still be statistically significant in the second step. Since the 

model was reduced to two first-order terms and there were concerns about potential multicolinearity, 

multilinear regression analysis was conducted using the two remaining variables. 

Table XV presents the results of the multilinear regression analysis. The regression analysis suffers 

from multicolinearity of the independent variables (temperature and stress) as measured by the variance 

inflation factor (VIF). A VIF value greater than 4 is generally an indication of multicolinearity, and both 

regression analyses have VIFs of 22.8 for both variables. 
NASA/TM—2011-215493 33 
NASA/TM—2011-215493 34 
NASA/TM—2011-215493 35 
NASA/TM—2011-215493 36 
NASA/TM—2011-215493 37 

The multicolinearity was caused by the need to limit the stress range within each test temperature to 

achieve tests with reasonable durations. As a result, there is a lack of overlap in the stresses at different 

temperatures and a strong correlation of stress to temperature in the data set. It was decided to still use the 

regression so that the relative effects of changes in temperature and stress could be assessed. 

Figure 20 presents the calculated creep rate as a function of stress at a series of temperatures ranging 

from 450 to 850 °C. At any given stress, increasing the temperature increases the creep rate as expected. 

The largest changes are experienced at the lowest temperatures. In the plot, this is manifested as the lines 

becoming closer together as the temperature increases. A 50 °C increase from 500 to 550 °C increases the 

creep rate by approximately 1 order of magnitude. The effect decreases as the temperature increases, and 

the creep rate increases by only about 1/2 order of magnitude when the temperature increases from 800 to 

850 °C. The effect of temperature is relatively insensitive to the magnitude of the applied stress. 

Figure 21 presents the calculated creep rate as a function of temperature for a series of stresses. As 

would be expected, the creep rate increases when the stress increases. However, it takes a very large 

change in stress to change the creep rate by 1 order of magnitude even at the lowest stresses used in this 

series of creep tests. At the higher stresses, the creep rate is relatively insensitive to even a 10-MPa 

(1.4-ksi) change in stress. 

Typically the applied stress is within 1 MPa (0.1 ksi) or less of the desired stress and is always calculated using the actual load for each test. The lines in Figure 21 represent 10-MPa steps. On the basis of 

this analysis and the observed consistency of the applied stresses, the effect of small stress differences due 

to frictional losses and other minor mechanical differences on creep rate will be minimal. The magnitude 

of the effect of stress is also relatively insensitive to the creep test temperature. 
NASA/TM—2011-215493 38 

Creep life follows the same trends, with the life being relatively insensitive to small changes in stress 

but very sensitive to small changes in temperature. This is consistent with the observations of the testing 

to determine the sources of variation in creep properties. 

The average difference between the chamber thermocouple, which historically served as the specimen 

temperature control thermocouple, and the middle thermocouple, which now serves as the control thermocouple, was calculated at each temperature for this data set. The average gradients along the sample 

length also were calculated. The summary is listed in Tables XVI and XVII. 

TABLE XVI.—CREEP RATE RESULTS OF MULTILINEAR 

REGRESSION ANALYSIS 

(a) Summary. 

Number of observations, N ................. 29 

Multiple correlation coefficent, R .. 0.899 

Standard error of estimate, SY,X ...... 0.243 

Missing observations ............................ 1 

Coefficient of determination, R2 .... 0.808 

Adjusted R2 .................................... 0.793

(b) Coefficients determined by multilinear regression, where T is temperature and  is stress. 

 Coefficient Standard 

error 

t-Test 

value 

Probability Variance 

inflation factor 

Constant –2.924 0.455 –6.426 <0.001 -------- 

1/(T – 273) –13982.578 1425.634 –9.808 <0.001 22.782 

log10() 7.471 0.722 10.343 <0.001 22.782 

 
NASA/TM—2011-215493 39 

(c) Analysis of variance. 

 Degrees of 

freedom 

Sum of 

squares 

Mean 

square 

F valuea

 Probability, 

P

Regression 2 6.420 3.210 54.558 <0.001 

Residual 26 1.530 0.059 -------- ------- 

Total 28 7.949 0.284 -------- ------- 

a

The value of the F distribution for X degrees of freedom associated with the 

numerator and Y degrees of freedom associated with the denominator. 

(d) Incremental change in sum of squares.

Variable Sum of squares

incremental 

Sum of squares

marginal 

1/(T – 273) 0.125 5.660 

log10() 6.294 6.294 

The dependent variable log10(creep rate) can be predicted from a linear combination of the independent 

variables: 

(e) Probability that independent variable should 

not be included.a

Independent variable Probability 

1/(T – 273) <0.001 

log10() <0.001 a

All independent variables appear to contribute to 

predicting log10(creep rate) for P < 0.05. 

(f) Assumption checks and power of regression.

Test Pass or fail? Probability, 

P 

Normality Passed 0.723 

Constant variance Passed 0.867 

Power of performed test with probability  = 0.050.......1.000 a

, the acceptable probability of incorrectly concluding that 

there is a difference. 

TABLE XVII.—CREEP LIFE RESULTS OF MULTILINEAR 

REGRESSION ANALYSIS

(a) Summary.

Number of observations, N ............................. 29 

Multiple correlation coefficent, R .............. 0.938 

Standard error of estimate, SY,X ......... 0.1550.243 

Missing observations ........................................ 1 

Coefficient of determination, R2 ................ 0.879 

Adjusted R2 ................................................ 0.870

(b) Coefficients determined by multilinear regression, where T is temperature and  is stress. 

 Coefficient Standard 

error 

t Value Probability Variance 

inflation factor 

Constant –1.838 0.291 –6.320 <0.001 -------- 

1/(T + 273) 12292.432 911.383 13.488 <0.001 22.782 

log10() –6.348 0.462 –13.746 <0.001 22.782 

 
NASA/TM—2011-215493 40 

(c) Analysis of variance. 

 Degrees of 

freedom 

Sum of 

squares 

Mean 

square 

F valuea

 Probability, 

P

Regression 2 4.545 2.272 94.506 <0.001 

Residual 26 0.625 0.024 ------- -------- 

Total 28 5.170 0.185 ------- -------- 

a

The value of the F distribution for X degrees of freedom associated with the 

numerator and Y degrees of freedom associated with the denominator. 

(d) Incremental change in sum of squares.

Independent variable Sum of squares

incremental 

Sum of squares

marginal 

1/(T + 273) 0.00118 4.374 

log10() 4.544 4.544 

The dependent variable log10(life) can be predicted from a linear combination of the independent 

variables: 

(e) Probability that independent variable 

should not be included.a

Independent variable Probability 

1/(T + 273) <0.001 

log10() <0.001 a

All independent variables appear to contribute to predicting 

log10(life) for P < 0.05. 

(f) Assumption checks and power of regression.

Test Pass or fail? Probability, 

P 

Normality Passed 0.788 

Constant variance Passed 0.186 

Power of performed test with probability  = 0.050.. .. 1.000 a

Probability , the acceptable probability of incorrectly 

concluding that there is a difference. 

The average difference between the chamber thermocouple and the middle sample thermocouple was 

12 °C (22 °F), but examination of the differences for the individual Brew frames shows that Brew frame 3 

had a very small difference at all temperatures, whereas Brew frame 2 had a much larger difference at all 

temperatures. Only one test was conducted in Brew frame 4, but it also had a large difference between the 

chamber and middle thermocouples. In general, the differences decreased with increasing test temperature. The authors believe that this was caused by radiative heating becoming more effective as the test 

temperature increased and produced more uniform and consistent heating. 

As can be seen in Figure 20, such a difference between the specimen temperature and the control 

thermocouple would have resulted in the sample creeping almost 1 order of magnitude faster in Brew 

frames 2 and 4 than in Brew frame 3 if the chamber thermocouple had been used to control the test temperature. This highlights how large a role the observed differences in specimen temperature play in the 

variability observed in the current and historical GRCop-84 data sets. 

The gradients in the sample gauge lengths were larger than desired, with average gradients for 

individual Brew frames being as high as 13.6 °C (24.5 °F) from the top of the sample to the bottom of the 

sample. In all cases, the bottom was hotter than the top, which is consistent with prior observations. It was 

also noted that the majority of the gradient occurred between the bottom and the middle of the sample. 

Average gradients from the middle to the top of the sample were relatively modest and typically in the 

2 to 4 °C (4 to 7 °F) range. 

Although not desirable, the consistency of the gradients indicates that they were not a major source 

of variation in the tests for a given Brew frame and were probably a minor contributor to the combined 
NASA/TM—2011-215493 41 

variability of tests using several Brew frames. It does highlight the need to improve the heaters to 

eliminate or minimize the thermal gradients. 

On the basis of this data set, it appears that the majority of the specimen-to-specimen and frame-toframe variability was eliminated by the new test procedures. Improvements in the thermal gradients still 

observed in the specimens may further improve these results, but the inherent variability of the creep of 

the material now appears to be the dominant source of variability. As will be demonstrated in the next 

section, the level of variability is now sufficiently low to detect relatively small changes in the creep 

properties of GRCop-84 caused by processing using statistical analysis. 

GRCop-42 heat treatment effects.—A regeneratively cooled rocket engine main combustion chamber 

liner may be installed into its jacket by brazing the liner into the jacket. Such a brazing operation will 

result in a small, but definite, decrease in the 0.2-percent-offset yield strength of GRCop-84 and its companion alloy GRCop-42 (Ref. 4). Such a decrease in strength should result in a decrease in creep properties since the material is now being tested at a higher fraction of the yield strength. Prior creep testing 

has not been able to detect differences in the creep behavior of GRCop-42 or GRCop-84 after a hightemperature simulated braze cycle. A new set of GRCop-42 specimens tested using the improved creep 

test procedures was examined to determine if statistically significant differences could be discerned. If 

so, the utility of the creep testing improvements in allowing observation of small differences could 

be validated. 

Figures 22 and 23 present the creep properties of extruded GRCop-42 in the as-extruded and brazed 

conditions. Visually there are strong indications that the expected trends are occurring. The simulated 

braze cycles tend to lower life and higher rate, and a 1000 °C (1832 °F) braze cycle results in a larger 

change than a 935 °C (1715 °F) braze cycle. 

To confirm that these apparent trends are correct, a forward stepwise regression using the model 

presented in Equation 2 was conducted. Again, F values of 4 to enter and 3.9 to leave were used to determine which variables should be in the model. For both the rate and life models, the log10() term was 

marginal for inclusion (P = 0.046 for the creep life model) or exclusion (P = 0.059 for the creep rate 

model) since the desired P value is 0.05 for 95-percent confidence. It was decided for consistency 

between the models that the log10() term would be excluded from the creep life model. This was not 

considered to be a major problem even though the stress was known to be a prime determiner of creep 

rate and life. The stress dependency was instead incorporated in the creep models through the stresstemperature interaction variable (log10()/(T + 273)) and through the 935 °C (1715 °F) braze-stress 

variable (B1 log10()), so that the stress dependency of the creep of GRCop-42 would still be captured. 

The forward stepwise regression was rerun, and the fit of the model to the data was slightly decreased as 

shown by the R2

 value decreasing from 0.939 to 0.930. This change was considered to be insignificant, 

and the model was kept. 

Table XVIII presents the ANOVA and F statistics from the last step of the regression analysis for 

creep rate and creep life, as well as the coefficients for each variable. The P values for the variables in the 

model are the probability of a false positive (the variable is actually not statistically significant but enters 

the regression model). The P values for the variables not in the model are the probability that the variable 

should not be in the model. 
NASA/TM—2011-215493 42 

 
NASA/TM—2011-215493 43 

TABLE XVIII.—AVERAGE TEMPERATURE DIFFERENCES AND GRADIENTS 

FOR GRCop-84 TESTS 

Temperature 

 

Brew 

frame 

Temperature difference, 

middle to chamber, 

°C 

Temperature gradient, 

°C 

°C °F Top to bottom Top to middle Middle to bottom 

All All All 12.0 –9.3 –5.7 –3.4 

 2 27.6 –12.3 –7.5 –4.2 

 3 0.2 –6.9 –4.2 –2.7 

 4 25.8 –13.6 –8.7 –4.9 

500 932 All 15.4 –8.0 –6.1 –1.9 

 2 48.0 –10.2 –8.2 –2.0 

 3 4.6 –4.8 –3.0 –1.8 

 4 25.8 –13.6 –8.7 –4.9 

650 1202 All 17.6 –8.0 –6.1 –1.9 

 2 31.4 –10.2 –8.2 –2.0 

 3 –3.2 –4.8 –3.0 –1.8 

800 1472 All 3.0 –7.2 –4.2 –3.0 

 2 11.5 –10.4 –5.6 –4.9 

 3 –2.7 –5.0 –3.2 –1.8 

In examining the statistically important variables, the influence of the simulated braze cycle on creep 

rate and creep life is now strongly statistically significant as demonstrated by the inclusion of variables 

with the blocking variables (B1, B2) and their large F values. Furthermore, the effects of the simulated 

braze cycle can be quantified through the coefficients of the models. 

These results are in direct contrast to historical efforts (Ref. 2, to be published, and Ref. 4), which 

could not detect a statistically significant effect for heat treatments including the simulated braze cycles. It 

clearly demonstrates the improvements in both the experimental procedures and the quality and utility of 

the creep results from Brew frame creep tests. It also demonstrates that data from multiple Brew frames 

can be used to detect these small differences. 

As with the GRCop-84 plate data set, the average temperature difference and gradients were calculated. The results appear in Tables XIX and XX. Brew frames 2 and 4 have a sample temperature far 

higher than the chamber thermocouple previously used for control. Brew frame 3 has nearly the same 

temperature reading from both thermocouples. The difference between the chamber and middle sample 

thermocouples can average as much as 42.2 °C. For Brew frame 2, the difference decreased with increasing temperature as was observed for the GRCop-84 plate samples. However, Brew frame 4 did not have 

the same trend. If radiative heat transfer was becoming more effective in Brew frame 2 and was reducing 

the difference, it would be expected that the same trend would be observed in Brew frame 4. The lack of 

such a trend suggests that another, unknown mechanism might have been operative and that there might 

have been a lurking variable not in the model or analyses. For now the analysis of the creep properties 

indicates that the variable does not have a large enough effect to prevent good analysis of the test results. 

 
NASA/TM—2011-215493 44 

TABLE XIX.—STATISTICAL ANALYSIS OF GRCop-42 CREEP RATE 

(a) Summary 

Number of observations, N ............................. 50

Multiple correlation coefficient, R ............. 0.969 

Standard error of estimate, SY,X .................. 0.238 

Coefficient of determination, R2 ................ 0.939 

Adjusted R2 ................................................ 0.934

(b) Analysis of variance. 

Group Degrees of

freedom 

Sum of

squares 

Mean 

square 

F valuea

 Probability, 

P

Regression 4 40.093 10.023 177.208 <0.001 

Residual 46 2.602 0.0566 --------- -------- 

a

The value of the F distribution for X degrees of freedom associated with the 

numerator and Y degrees of freedom associated with the denominator. 

(c) Variables in model, where T is temperature,  is stress, and B1 and B2 are the 

blocking variables corresponding to brazes of 935 and 1000 °C, respectively. 

Variable Coefficient Standard error F valuea

 Probability 

Constant 13.192 0.808 ----------- -------- 

1/T + 273 –30765.830 1224.541 631.234 <0.001 

log10()/(T + 273) 7868.689 302.107 678.397 <0.001 

B2/(T + 273) 692.479 76.833 81.230 <0.001 

B1 log10() 0.285 0.0507 31.504 <0.001 

(d) Variables not in model. 

Variable Coefficient Standard error 

log10() 3.754 0.059 

B1/(T + 273) 1.337 0.254 

B2 log10() 0.106 0.746 

(e) Assumption checks and power of regression. 

Test Pass or fail? Probability, 

P 

Normality Passed 0.358 

Constant variance Passed 0.859 

Power of performed test with probability  = 0.050.. ....... 1.000 a

Probability , the acceptable probability of incorrectly 

concluding that there is a difference. 

TABLE XX.—STATISTICAL ANALYSIS OF GRCop-42 CREEP LIFE 

(a) Summary. 

Number of observations, N ............................. 50

Multiple correlation coefficient, R............. 0.966

Standard error of estimate, SY,X .................. 0.220

Coefficient of determination, R2 ................ 0.930

Adjusted R2 ............................................... 0.924

(b) Analysis of variance. 

Group Degrees of

freedom 

Sum of

squares 

Mean 

square 

F valuea

 Probability, 

P

Regression 4 29.737 7.434 153.936 <0.001 

Residual 46 2.222 0.0483 --------- -------- 

a

The value of the F distribution for X degrees of freedom associated with the 

numerator and Y degrees of freedom associated with the denominator. 
NASA/TM—2011-215493 45 

(c) Variables in model, where T is temperature,  is stress, and B1 and B2 are the 

blocking variables corresponding to brazes of 935 and 1000 °C, respectively. 

Variable Coefficient Standard error F valuea

 Probability 

Constant –15.593 0.746 ---------- -------- 

1/T + 273 26866.373 1131.508 563.771 <0.001 

log10()/(T + 273) –6830.686 279.155 598.741 <0.001 

B2/(T + 273) –543.622 70.996 58.631 <0.001 

B1 log10() –0.231 0.0469 24.221 <0.001 

(d) Variables not in model. 

Variable Coefficient Standard error 

B1/(T + 273) 0.803 0.375 

B2 log10() 0.202 0.655 

log10() Excluded from model 

(e) Assumption checks and power of regression. 

Test Pass or fail? Probability, 

P 

Normality Passed 0.277 

Constant variance Passed 0.840 

Power of performed test with probability  = 0.050.. ...... 1.000 a

Probability , the acceptable probability of incorrectly 

concluding that there is a difference. 

The temperature gradients along the gauge section of the specimens were again higher than desired, 

as shown in Table XXI. The average gradients were as high as 22.1 °C (39.8 °F), and overall they 

averaged 12.5 °C (22.5 °F) from the top to the bottom of the sample. Although the maximum gradient 

exceeds that observed for the GRCop-84 specimens, the overall average is similar to the overall 

GRCop-84 average gradient of 9.3 °C (16.7 °F). 

The average gradients for the individual Brew frames at each test temperature were examined next. 

The results for Brew frames 3 and 4 were generally consistent with the GRCop-84 plate data set. 

However, the average gradients for the GRCop-42 specimens tested in Brew frame 2 were nearly twice 

those of GRCop-84 plate specimens tested in the same frame. The larger gradients in Brew frame 2 were 

the primary source of the difference in the observed differences in the average gradients at all temperatures. The gradient data for Brew frame 2 may have been influenced by having only five GRCop-42 tests 

from Brew frame 2 in this data set. The small number of samples with only the 500 °C (932 °F) and 

650 °C (1202 °F) test temperatures having a repeat may skew the averages. 

Most of the increase in the top-to-bottom gradient occurred between the top and middle thermocouples. The implication of this is that the top section of the specimen was cooler in the GRCop-42 tests 

than in the GRCop-84 tests. As shown in Figure 20, the lower temperature would result in a lower creep 

rate, which would also result in a longer creep life. For a 5 to 10 °C (9 to 18 °F) change, the creep rate 

could decrease by 1/10 to 1/3 order of magnitude. The variation in gradients from data set to data set is 

undesirable, and it appears to be a contributor to the remaining variability in the data sets. A goal of any 

new furnace heater design will need to be to not only reduce the gradients overall but to make them 

consistent from frame to frame so that this source of error is minimized. 

In Tables XVIII and XXI, the magnitude of the gradient from the bottom thermocouple to the middle 

thermocouple normally exceeds that of the gradient from the middle thermocouple to the top thermocouple. These results are consistent with the observations in the prior testing to seek out sources of variability. Because of the consistency of the gradient direction and the difference in magnitude between the 

top half and the bottom half, the total gradient is reported rather than the middle thermocouple value plus 

or minus half of the average measured gradient. It is felt that this is a more appropriate and meaningful 

way to present the data at this point since it captures the asymmetry of the gradient, whereas a single 

value with a simple plus/minus range does not. If the heater is redesigned and the two gradients are 

equalized, then the best method of reporting the gradients will be revisited. 
NASA/TM—2011-215493 46 

 TABLE XXI.—AVERAGE TEMPERATURE DIFFERENCES AND GRADIENTS 

FOR GRCop-42 TESTS 

Temperature Brew 

frame 

Temperature difference, 

middle to chamber, 

°C 

Temperature gradient, 

°C 

°C °F Top to bottom Top to middle Middle to bottom 

All All All 15.9 –12.5 –6.8 –5.8 

 2 33.7 –19.3 –14.5 –4.8 

 3 –1.6 –7.3 –5.3 –2.1 

 4 20.1 –13.7 –6.3 –7.5 

500 932 All 21.1 –14.0 –7.4 –6.6 

 2 42.2 –22.1 –17.3 –4.8 

 3 –0.7 –10.9 –6.6 –4.3 

 4 21.2 –13.0 –5.7 –7.3 

650 1202 All 16.0 –15.0 –8.5 –6.5 

 2 34.5 –21.0 –15.3 –5.7 

 3 –2.8 –7.6 –5.7 –1.9 

 4 20.4 –16.8 –8.8 –8.0 

800 1472 All 11.7 –9.1 –4.6 –4.5 

 2 21.7 –12.7 –10.0 –2.7 

 3 –1.1 –6.2 –4.6 –1.5 

 4 18.7 –10.5 –3.6 –6.9 

Summary and Conclusions 

It was shown that significant differences existed between results obtained for five Brew creep test 

frames when identical specimens with the same process history were tested under presumably identical 

conditions. It was determined that the primary cause for the variability was that the setpoint control temperature measured by a noncontact thermocouple was different from the actual sample temperature by a 

large and variable amount. 

Equipment and procedure modifications that were implemented resulted in the creep rate between 

Brew frames now being equivalent to an independently determined calibration standard. The variation 

decreased to about 1/2 order of magnitude from 2 to 4 orders of magnitude. 

Substituting a helium atmosphere for vacuum generally decreased the temperature gradient along the 

specimen by improving the uniformity of heat transfer from the heater to the specimen. This further 

decreased variability in the results especially at 500 °C. It does not appear that the helium consistently 

helped at higher test temperatures. 

It was also shown that GRCop-84 creep behavior is much more sensitive to small temperature 

changes than to small stress changes. This magnified the effects of differences in specimen temperature 

on test-to-test variability. Higher test-to-test variability increases the total variability observed in the creep 

properties. 

It is probable that most creep tests conducted prior to modifying the creep test methods and procedures were done at sample temperatures higher than reported. The magnitude of the temperature 

difference may have been as high as 50 °C (90 °F) for some tests. The estimated average temperature 

difference was approximately 15 °C (27 °F) on the basis of observed temperature differences between the 

control thermocouple and the middle specimen thermocouple for the GRCop-84 and GRCop-42 tests in 

this report. The differences varied from frame to frame and were fairly strongly dependent on the test 

temperature for Brew frame 2 but not very dependent on the test temperature for Brew frames 3 and 4. 

These combined findings strongly indicate that previous tests prior to implementing these modifications resulted in reporting GRCop-84 creep lives shorter than the actual value and creep rates greater 

than the actual values at the reported setpoint temperature. This means that the prior reported work 

(Ref. 4) is overly conservative. On the basis of the analysis of the dependency of creep rate and creep life 
NASA/TM—2011-215493 47 

on temperature, the temperature difference is sufficient to increase creep rates by 1/3 to 1 order of 

magnitude and to decrease creep lives by 1/3 to 1 order of magnitude. 

The verification testing in stage III using the updated methods shows that the variability in the creep 

results has been greatly decreased. The testing was also demonstrated to be enhanced by being able to 

detect and quantify small differences in creep rate and creep life using statistical analyses. Analysis of 

GRCop-84 plate specimens showed that the variability for a single material condition was reduced to 

1 order of magnitude or less as measured by the 95-percent confidence interval for the regression 

analysis. Testing GRCop-42 specimens in three conditions demonstrated that small differences in creep 

behavior previously masked by the large variability in the data were now distinguishable and that 

quantifiable changes in the creep properties could be calculated. The improvements also imply that 

smaller numbers of samples will be needed for future testing to accurately characterize the response 

of GRCop-42, GRCop-84, and other samples tested in these creep frames. 

Future Work 

It would be desirable to replace the current wire mesh heaters with solid heaters to determine if they 

produce smaller thermal gradients as were observed with TM–4. The gradients need to be reduced to 

±5 °C or less to produce the best possible creep results. Such work is currently not planned but will be 

proposed in the future. 

References 

1. de Groh, Henry C., III; Ellis, David L.; and Loewenthal, William S.: Comparison of GRCop-84 to 

Other High Thermal Conductive Cu Alloys. NASA/TM—2007-214663, 2007. http://gltrs.grc.nasa.gov 

Accessed Feb. 6, 2009. 

2. Ellis, D.L.; Loewenthal, W.S.; and Haller, H.: Creep of Commercially Produced GRCop-84. 

NASA/TM—2012-216998, 2012, to be published. 

3. Ellis, David L.; and Michal, Gary M.: Mechanical and Thermal Properties of Two Cu-Cr-Nb Alloys 

and NARloy-Z. NASA/CR198529, 1996. http://gltrs.grc.nasa.gov Accessed Feb. 6, 2009. 

4. Ellis, D.L.: GRCop-84. Aerospace Structural Materials Handbook Supplement, 40th ed., vol. 6, 

CINDAS LLC, West Lafayette, IN, 2001. 

5. Loewenthal, William S.; and Ellis, David L.: GRCop-84 Rolling Parameter Study. NASA/TM—2008-

215213, 2008. http://gltrs.grc.nasa.gov Accessed Feb. 6, 2009. 

6. ASTM Standard E–8. Standard Test Methods for Tension Testing of Metallic Materials. ASTM 

International, West Conshohocken, PA, 2008. 

7. Ellis, David, et al.: A Practical Guide to the Production of Metal Spun NiCrAlY Coated GRCop-84 

Liners. NASA/TM—2006-214123, 2006. Available from the NASA Center for Aerospace Information. 

8. Monkman, F.C.; and Grant, N.J.: An Empirical Relationship Between Rupture Life and Minimum 

Creep Rate in Creep-Rupture Tests. Proc. ASTM, vol. 56, 1956, pp. 593–620. 

9. Reed-Hill, Robert E.: Physical Metallurgy Principles. 2nd ed., Van Nostrand, New York, NY, 1973, 

pp. 871–877. 

10. Ellis, D.L.; and Garg, A.: Annealing of GRCop-84. To be published as a NASA TM. 
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2. REPORT TYPE

Technical Memorandum 

3. DATES COVERED (From - To)

4. TITLE AND SUBTITLE

Sources of Variation in Creep Testing 

5a. CONTRACT NUMBER

5b. GRANT NUMBER

5c. PROGRAM ELEMENT NUMBER

6. AUTHOR(S)

Loewenthal, William, S.; Ellis, David, L. 

5d. PROJECT NUMBER

5e. TASK NUMBER

5f. WORK UNIT NUMBER

WBS 599489.02.07.03.02.04.01 

7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS(ES)

National Aeronautics and Space Administration 

John H. Glenn Research Center at Lewis Field 

Cleveland, Ohio 44135-3191 

8. PERFORMING ORGANIZATION

 REPORT NUMBER

E-16689 

9. SPONSORING/MONITORING AGENCY NAME(S) AND ADDRESS(ES)

National Aeronautics and Space Administration 

Washington, DC 20546-0001 

10. SPONSORING/MONITOR'S

 ACRONYM(S)

NASA 

11. SPONSORING/MONITORING

 REPORT NUMBER

NASA/TM-2011-215493 

12. DISTRIBUTION/AVAILABILITY STATEMENT

Unclassified-Unlimited 

Subject Category: 26 

Available electronically at http://www.sti.nasa.gov 

This publication is available from the NASA Center for AeroSpace Information, 443-757-5802

13. SUPPLEMENTARY NOTES

14. ABSTRACT

Creep rupture is an important material characteristic for the design of rocket engines. It was observed during the characterization of GRCop84 that the complete data set had nearly 4 orders of magnitude of scatter. This scatter likely confounded attempts to determine how creep 

performance was influenced by manufacturing. It was unclear if this variation was from the testing, the material, or both. Sources of 

variation were examined by conducting tests on identically processed specimens at the same specified stresses and temperatures. Significant 

differences existed between the five constant-load creep frames. The specimen temperature was higher than the desired temperature by as 

much as 43 °C. It was also observed that the temperature gradient was up to 44 °C. Improved specimen temperature control minimized 

temperature variations. The data from additional tests demonstrated that the results from all five frames were comparable. The variation 

decreased to 1/2 order of magnitude from 2 orders of magnitude for the baseline data set. Independent determination of creep rates in a 

reference load frame closely matched the creep rates determined after the modifications. Testing in helium tended to decrease the sample 

temperature gradient, but helium was not a significant improvement over vacuum.

15. SUBJECT TERMS

Creep properties; Temperature; Temperature gradients; Variability 

16. SECURITY CLASSIFICATION OF: 17. LIMITATION OF

 ABSTRACT

UU 

18. NUMBER

 OF 

 PAGES

53 

19a. NAME OF RESPONSIBLE PERSON

STI Help Desk (email:help@sti.nasa.gov) 

a. REPORT

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b. ABSTRACT

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c. THIS 

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19b. TELEPHONE NUMBER (include area code)

443-757-5802 

Standard Form 298 (Rev. 8-98)

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