Timothy P. Gabb and John Gayda

Glenn Research Center, Cleveland, Ohio

Tensile and Creep Property Characterization

of Potential Brayton Cycle Impeller

and Duct Materials

NASA/TM—2006-214110

February 2006
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Timothy P. Gabb and John Gayda

Glenn Research Center, Cleveland, Ohio

Tensile and Creep Property Characterization

of Potential Brayton Cycle Impeller

and Duct Materials

NASA/TM—2006-214110

February 2006

National Aeronautics and

Space Administration

Glenn Research Center
Available from

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Available electronically at http://gltrs.grc.nasa.gov
Tensile and Creep Property Characterization of Potential Brayton 

Cycle Impeller and Duct Materials 

Timothy P. Gabb and John Gayda 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Abstract 

This paper represents a status report documenting the work on creep of superalloys performed under 

Project Prometheus. Cast superalloys have potential applications in space as impellers within closed-loop 

Brayton cycle nuclear power generation systems. Likewise wrought superalloys are good candidates for 

ducts and heat exchangers transporting the inert working gas in a Brayton-based power plant. Two cast 

superalloys, Mar-M247LC and IN792, and a NASA Glenn Research Center powder metallurgy 

superalloy, LSHR, are being screened to compare their respective capabilities for impeller applications. 

Several wrought superalloys including Hastelloy X (Haynes International, Inc., Kokomo, IN), Inconel 

617, Inconel 740, Nimonic 263, and Incoloy MA956 (Special Metals Corporation, Huntington, WV), are 

also being screened to compare their capabilities for duct applications. These proposed applications would 

require sufficient strength and creep resistance for long term service at temperatures up to 1200 K, with 

service times to 100,000 h or more. Conventional tensile and creep tests were performed at temperatures 

up to 1200 K on specimens extracted from the materials. Initial microstructure evaluations were also 

undertaken. 

Introduction 

Certain space operations require an electric power source that is not dependent on the sun and can 

perform for long time periods. A promising concept fulfilling these requirements uses a nuclear reactor to 

heat an inert He-Xe working gas, which is expanded in a closed-loop Brayton cycle to power a turbogenerator for electricity production (ref. 1). There are multiple combinations of reactor and heat 

exchanger design that can be used, but in each scenario power conversion efficiency increases with higher 

turbine temperature. Also space power systems may be required to operate for very long service times. In 

the Jupiter Icy Moons Orbiter (JIMO) vehicle concept, the turbo-generator would require an impeller 

(radial turbine wheel) which can operate with sustained centrifugal loads for times extending to 

100,000 h and maximum temperatures approaching 1200 K. Cast superalloys such as Mar-M247LC and 

IN792 could potentially be used for these impellers, provided creep and alloy phase stability are 

sufficient. LHSR is an advanced powder metallurgy superalloy that could also have potential for these 

impellers, offering higher strength and fatigue resistance but somewhat lower temperature capability. 

Ducts subjected to pressure and thermal loads at similar temperatures will be needed to route and cool 

the working gas (ref. 2). The associated materials will need to be shaped, brazed, and welded into duct 

and turbine housing components. Solid solution strengthened wrought superalloys such as Hastelloy X, 

and Inconel 617 could be readily fabricated in these ways. However, their high temperature creep 

resistance may be insufficient. Precipitation hardened wrought superalloys such as Nimonic 263 and 

Inconel 740 have improved creep resistance, but could be harder to shape, join, and subsequently 

precipitation heat treat. Mechanically alloyed, precipitation hardened wrought superalloys such as Incoloy 

MA956 can have even higher creep resistance at high temperatures near 1200 K, but are very difficult to 

shape and join. Long term creep tests are therefore needed at realistic stresses and temperatures to 

compare these alloys at simulated the service conditions. 

The objective of this study was to begin screening the mechanical properties of these materials for 

potential space applications. Potential impeller alloys Mar-M247LC, IN792, and LSHR, and potential 

NASA/TM—2006-214110 1
duct alloys Hastelloy X, Inconel 617, Inconel 740, Nimonic 263, and Incoloy MA956 were screened with 

selective tensile and creep testing. Typical material data bases do not include creep tests beyond about 

10,000 h or consider such long term exposure in inert gases (ref. 3). Based on the long durations 

envisioned in some of these space applications, short term creep tests are being run in order to enable 

selection of allowable creep stresses and temperatures conducive to such long durations. Some of these 

tests are still ongoing, but results generated to date will be reviewed. 

Materials and Procedure 

Impeller Materials 

Most Mar-M247LC (ref. 4) bars were cast at PCC, Inc. in a 1 by 18 bar mold using standard foundry 

practices. Additional bars were cast using the Microcast process to produce finer grain size. The bars had 

diameters about 1.9 cm and lengths of about 15 cm. These bars were subsequently hot isostatically 

pressed at 1458 K using a pressure of 172 MPa for 4 h. They were then solution heat treated in a vacuum 

furnace at 1494 K/2 h/Argon quench, and aged at 1352 K/4 h/Argon quench + 1144 K/20 h/air quench. 

An impeller from an auxiliary power unit of the superalloy IN792 (ref. 5) was provided by Sundstrand 

Aerospace. It had been cast using the Grainex process, another fine grain casting process, at Howmet. It 

was then hot isostatically pressed and heat treated using proprietary processes. Powder of LSHR 

superalloy (ref. 6) was atomized by Special Metals Corp. in argon and passed through screens of 

–270 mesh to give powder particle diameters of no more than about 55 μm. The powder was then sealed 

in a stainless steel container, hot compacted, and extruded at a reduction ratio of 6:1. Segments of the 

extrusion billet were machined into sections 15 cm in diameter and 20 cm long, then isothermally forged 

into a flat disk about 30 cm in diameter and 4 cm thick by Wyman-Gordon Forgings. The disk was 

solution heat treated at about 1444 K for 2.5 h, then transferred in 2 min for a quench in air by fans on 

each side of the disk. The disk was then given an aging heat treatment of 1128 K/8 h. 

Duct Materials 

Most of the wrought alloys were supplied by Special Metals Corp. Hot rolled plates of Hastelloy X 

(refs. 7 and 8), Inconel 617 (ref. 9), and Incoloy MA956 (ref. 10) of about 0.64 cm in thickness were 

supplied. They had been given a solution annealing heat treatment, and were tested in this state. A hot 

rolled plate of Inconel 740 (refs. 11 and 12) plate of 1.6 cm thickness and a Nimonic 263 (ref. 13) section 

of 4 cm thickness were also supplied in solution annealed and air quenched form. Unlike the other 

wrought duct materials, Inconel 740 and Nimonic 263 require an aging heat treatment in order to 

precipitate strengthening γ ′ precipitates. Therefore, they were subsequently given an aging heat treatment 

of 1073 K/8 h before testing. 

Testing 

Microstructure.—Actual compositions of the test materials were measured using inductively coupled 

plasma spectroscopy. Grain sizes were determined from metallographically prepared grip sections of 

tensile specimens according to ASTM E112 linear intercept procedures using circular grid overlays, and 

As-Large-As (ALA) grain sizes were determined using ASTM E930. 

Creep Tests.—Machining of creep specimens was performed by Metcut Research Associates. Creep 

specimens with a gage diameter of 0.63 cm and gage length of 3.8 cm were machined from the candidate 

impeller materials. Smaller creep specimens with a gage diameter of 0.32 cm and gage length of 0.76 cm 

were machined from the plate materials. Most tests are being performed in conventional uniaxial lever 

arm constant load creep frames using resistance heating furnaces and shoulder-mounted extensometers at 

Metcut and GRC. However, two tests were run in specialized creep testing machines at GRC, which had 

NASA/TM—2006-214110 2
the specimens sealed within environmental chambers containing inert helium gas of 99.999 percent purity 

held slightly above atmospheric pressure. All creep tests were performed according to ASTM E139. 

Tensile Tests.—Tensile test specimens were machined by Metcut having a gage diameter of 

about 0.40 cm and gage length of about 2.3 cm. Tensile tests were performed at Metcut and GRC in 

uniaxial test machines employing a resistance heating furnace and axial extensometer. The tests were 

performed according to ASTM E21, using an initial test segment with strain increased at a uniform rate of 

0.2 percent per min, followed by a segment with displacement increased at a uniform rate of 5 mm/min. 

Results and Discussion 

Material and Microstructures 

The actual chemistries in weight percent of all tested alloys are listed in table 1. Typical grain 

microstructures in optical images of etched metallographic sections of impeller materials are shown in 

figures 1 and 2. Grain width was approximated using grain linear intercept distance according to ASTM 

E112. LSHR had a fine, uniform grain microstructure, with an average linear intercept distance of about 

29 μm, equivalent to a grain size of 7.0. Conventionally cast Mar-M247LC had more irregular, very 

coarse grains of near 700 μm width and near 800 to 12000 μm length, with grains often longer in the 

direction of primary dendrite growth. However, Microcast Mar-M247LC had a more uniform grain 

microstructure, with smaller grains of about 96 μm width (grain size = 3.5). IN792 had intermediate grain 

uniformity and size compared to the two versions of Mar-M247LC, with grain about 320 μm in width 

(grain size = 0). All the impeller alloys had microstructures predominated by 55 to 70 percent γ′

precipitates in a γ matrix. MC and M23C6 carbides were also present. LSHR microstructure was 

predominated by about 55 vol% γ′ precipitates in a γ matrix, but also contained minor MC and M23C6

carbides, M3B2 borides, and very fine oxides. Cooling γ′ precipitate size was uniformly about 0.2 μm. 

Mar-M247LC and IN792 microstructure was predominated by about 65 to 70 vol% γ′ precipitates in a γ

matrix, with minor MC carbides. γ′ precipitate sizes varied widely in these cast alloys, between about 0.4 

and 3.0 μm, due to dendritic growth within grains. 

Typical microstructures for plate materials are shown in figures 3 to 4. Hastelloy X, Inconel 617, 

Inconel 740, and Nimonic 263 had comparable grain sizes, but MA956 had much coarser grains. 

Hastelloy X and Inconel 617 had fairly uniform grain structures with mean grain sizes of about 69 and 

75 μm (grain sizes of 4.4 and 4.2), respectively. They both were predominantly made up of γ phase, with 

minor contents of carbides. Inconel 617 also had about 2 vol% of γ′ precipitates. However, these 

precipitates were very large at 1 to 10 μm diameter, and would not be expected to provide significant 

strengthening. Inconel 740 and Nimonic 263 had similar grain sizes of about 69 and 62 μm (grain sizes of 

4.4 and 4.7), respectively. They both were predominantly made up of γ phase strengthened by about 20 to 

25 vol% γ′ precipitates, along with minor contents of carbides. The γ′ precipitates were smaller in Inconel 

740 than for Nimonic 263. Incoloy MA956 had a much coarser grain structure than the other duct alloys, 

with grains of flattened aspect ratio in the plane of the plate. Grains were typically 100 to 

200 μm thick, but 300 to 1500 μm wide in the plane of the plate. This microstructure was predominantly 

made up of BCC ferrite phase, with minor contents of fine carbides and Y2O3 – Al2O3 dispersion 

strengthening particles. These chemistries and microstructures were typical of those reported in supplier 

references (refs. 3 to 12). 

Mechanical Properties of Impeller Materials 

Tensile Response.—Yield strengths at 0.2 percent offset, ultimate strengths, percent elongation after 

failure, and percent reduction in area after failure are compared as functions of temperature in table 3 and 

figure 5. LSHR had higher strength than Mar-M247LC at 978 K, but comparable strengths at 1090 and 

1200 K. This could be related in part to the finer grain size of LSHR compared to both conventionally 

NASA/TM—2006-214110 3
cast and Microcast Mar-M247LC. Decreasing grain size has often been demonstrated to strongly improve 

strength in many superalloys (ref. 14) as well as other materials. However, the chemistries and precipitate 

contents of the two alloys would also influence their strengths. Microcast Mar-M247 had nearly identical 

strengths to conventionally cast Mar-M24LC at 1090 and 1200 K, even though the finer grain size of 

Microcast material would be expected to encourage higher strengths. IN792 had consistently lower 

strengths than both forms of Mar-M247LC and LSHR. 

Ductility as indicated by elongation and reduction in area after failure is also compared as functions 

of temperature in table 3 and figure 5. LSHR had higher ductility than Mar-M247LC at 978 °C, but lower 

ductility at 1090 and 1200 K. Microcast Mar-M247 had lower ductility than conventionally cast MarM247LC at 1090 and 1200 K even though the strengths were comparable. IN792 had similar ductility to 

conventionally cast Mar-M247LC. 

Creep Response.—Creep tests of impeller materials were designed to first screen the materials at 

temperatures of 978, 1090, and 1200 K. Subsequent tests were designed to ultimately determine allowable 

creep stresses for each material and temperature that would give 1 percent creep in 10 years of service, a 

goal then projected for JIMO applications. This service goal represented a target strain rate of 0.1 percent 

per year. Therefore, tests were performed to determine the allowable stress that would produce this strain 

rate at each of the test temperatures. Test conditions and results-to-date are listed for Mar-M247LC and 

LSHR in table 4. Initial tests of the third impeller material, IN792, are also included. 

Tests to estimate the effects of air versus inert environments on creep resistance were initiated early 

on, using conventionally cast Mar-M247LC. The results of single tests in air at one atmosphere pressure 

and in helium slightly above one atmosphere at 1089 and 1200 K are compared in figure 6. Creep life in 

helium was lower than that for air in tests at 1089 K, but higher in tests at 1200 K. The differences were 

small at only 2 to 4×, which is near the normal scatter expected in creep response. More tests are needed 

for confirmation, but this appears to show relatively modest differences in life to 0.2 percent creep due to 

changing environment. A previous study that compared creep for superalloys in helium and air 

environments (ref. 3) also found only modest differences. However, other creep testing on LSHR has 

conclusively shown that creep loading in air encourages cracks to initiate at oxidized surfaces, causing 

failure (ref. 5). Therefore, LSHR and Mar-M247LC specimens tested to higher creep strains in air could 

initiate similar surface cracks at surface oxidation, giving accelerated creep response compared to 

specimens tested in helium. This could produce shorter creep lives in air than in helium. Eventually, creep 

tests in helium and air need to be continued to rupture at higher strains, in order to assess this issue. 

Typical comparisons of time to 0.2 percent creep of LSHR and Mar-M247LC for temperatures of 

977, 1089, and 1200 K are shown in figure 7. LSHR has slightly higher creep life at 977 K, but MarM247LC is superior at 1089 and 1200 K. This could be related in part to the much coarser grain size in 

Mar-M247LC compared to LSHR. The conventionally cast Mar-M247LC had a larger typical grain size 

than LSHR, as shown in figure 1. Increasing grain size has often been demonstrated to strongly improve 

superalloy creep resistance in these alloys (ref. 6). However, Mar-M247LC also has a different chemistry 

containing more aluminum and titanium to produce a content of γ′ precipitates near 70 percent, compared 

to that of LSHR containing about 55 to 60 percent γ′. Mar-M247LC also has higher levels of the 

refractory elements tungsten, tantalum, niobium, and hafnium than does LSHR, which can all improve 

creep resistance of superalloys (ref. 4). 

In order to examine the effects of only grain size on creep resistance, specimens from a bar of MarM247LC cast using the Microcast process were obtained, containing finer grains compared to 

conventionally cast material, figure 1. Limited comparison tests indicate the Microcast specimens had 2 to 

10× lower creep life than conventionally cast Mar-M247LC, figure 8. The initial tests of IN792 are also 

included for comparison, as its content of γ ′ and microstructure were similar to Mar-M247LC. IN792 

material was cast using a proprietary Grainex process that also produced a finer grain size than 

conventionally cast Mar-M247LC. It also had lower creep resistance, comparable to Microcast MarM247LC. This illustrated the strong effect of increasing grain size on creep resistance at high 

NASA/TM—2006-214110 4
temperatures, and should be kept in mind when selecting processing paths for both impeller and duct 

alloys. This grain size difference more strongly affected creep resistance than strength for this alloy. 

Creep results of all tests were used to generate conventional Larson-Miller curves of stress versus 

Larson-Miller parameter (LMP) using temperature (T) and time (t) in the equation (ref. 15): 

LMP = (T)(20 + log t)/1000 

This parameter has a direct dependence with temperature, but a weaker logarithmic dependence of time. 

The resulting plot is shown in figure 9. The LMP curves confirm consistently higher comparative 

response for LSHR at combinations of high stress and low temperatures, but higher response for MarM247LC at combinations of low stress and high temperatures. Simple regression equations are included 

for estimating creep response at intermediate conditions. However, it is not advisable to use these 

equations to estimate creep response at times greater than 10 times over that of the input data, which 

extended to 7,295 h. 

Creep strain rate to 0.2 percent is shown versus stress in figures 10 to 11. These plots allowed initial 

estimation of allowable stresses for achieving the target strain rate of 0.1 percent per yr. for each material. 

At 978 K, a stress of about 350 MPa was estimated for both LSHR and Mar-M247LC. However, at higher 

temperature the superior creep resistance of Mar-M247LC over LSHR was clearly reflected in the 

estimated stresses. A stress of 100 MPa was estimated for Mar-M247LC at 1090 K, while no stress could 

be estimated for LSHR. At 1200 K, a stress could no longer be clearly estimated for either material, 

although Mar-M247LC again had much better creep resistance. Additional creep tests and analyses are 

necessary, but a preliminary creep analysis using current test results indicates quite good potential for an 

impeller fabricated of Mar-M247LC to meet Brayton cycle requirements, for maximum temperatures 

below 1200 K (ref. 16). 

Mechanical Properties of Duct Materials 

Tensile Response.—Comparative duplicate tensile tests were performed at 1090 and 1200 K for most 

of the duct materials. Incoloy MA956 was not tensile tested due to an insufficient material supply. 

Average yield strengths at 0.2 percent offset and ultimate strengths are compared as functions of 

temperature in figure 12. Inconel 740 had highest strength among the duct alloys tested. Hastelloy X had 

lowest strength at 1090 K, while Nimonic 263 had lowest strength at 1200 K. As a group, the 

precipitation strengthened alloys Inconel 740 and Nimonic 263 had significantly higher strength than the 

solid solution alloys Hastelloy X and Inconel 617 at 1090 K. At 1200 K the precipitation strengthening 

improvement over the predominantly solid solution alloys was still evident for Inconel 740, but not for 

Nimonic 263. Now both these alloys would be expected to have comparable amounts of γ′ precipitates at 

1200 K, based on their similar compositions including the amounts of Al + Ti which stabilize γ′. 

However, Nimonic 263 appeared to have significantly larger precipitates even before testing at 1200 K, 

and this larger precipitate size may be less effective in strengthening at high temperatures. 

Ductility as indicated by elongation and reduction in area after failure is also compared as functions 

of temperature in figure 12. Hastelloy X had the highest ductility among the duct alloys at 1090 K, 

while Inconel 740 had the lowest ductility. As a group, the predominantly solid solution strengthened 

alloys Hastelloy X and Inconel 617 had superior ductilities to the precipitation strengthened alloys 

Inconel 740 and Nimonic 263 at 1090 K. At 1200 K, all duct alloys had very good ductilities 

exceeding 50 percent reduction in area, with no consistent trends among the alloys. 

Additional tensile tests were performed on Inconel 617 to compare tensile properties in the plate 

longitudinal and width directions. Averaged yield and ultimate strengths are compared with associated 

elongations and reductions in area in figure 13. Strengths and ductilities were quite comparable in the two 

plate directions, for tests performed at 300, 1090, and 1200 K.

Creep Response.—Vendor literature contained appreciable information on the creep resistance of 

these materials for temperatures near 977 K, but had little data at higher temperatures. Therefore, creep 

NASA/TM—2006-214110 5
tests of duct/heat exchanger materials were designed to screen the materials at temperatures of 1090 and 

1200 K. Subsequent tests designed to determine allowable stresses for each alloy based on service goals 

have not yet been started, due to the variety of duct and heat exchanger applications possible here, as well 

as recent changes in program goals. Current test conditions and results are listed for Hastelloy X, Inconel 

617, Inconel 740, Nimonic 263, and Incoloy MA956 in table 3. 

Typical comparisons of time to 0.1, 0.2, and 0.3 percent creep of these materials at 1089 K are shown 

in figure 14, for duplicate tests at a stress of 103 MPa. Of the solid solution strengthened superalloys, 

Inconel 617 had about 10x better creep resistance than Hastelloy X. However, the γ′ precipitate 

strengthened superalloys Inconel 740 and Nimonic 263 had about 20x improved creep resistance over 

Inconel 617. These results indicate Inconel 617, Inconel 740, and Nimonic 263 had good creep resistance 

for potential applications at 1089 K, while Hastelloy X would be at a definite disadvantage. Additional 

testing would be necessary to determine allowable stresses for each alloy. 

Typical comparisons of time to 0.1 and 0.2 percent creep of these materials at 1200 K are shown in 

figure 15, for single tests at a very low stress of 14 MPa. The materials had much lower creep resistance 

at this temperature. Of the solid solution strengthened superalloys, Inconel 617 had 2 to 3 times better 

creep resistance than Hastelloy X. However, the γ′ precipitate strengthened superalloys Inconel 740 and 

Nimonic 263 had mixed response. Nimonic 263 had comparable creep resistance to Inconel 617, while 

Inconel 740 had lower creep resistance than Hastelloy X as well as Inconel 617. This could be related to 

the potential for precipitation of harmful phases in Inconel 740 and Nimonic 263 at this temperature 

(ref. 12). Incoloy MA956 had much better creep resistance than the other materials at 1200 K, but does 

have many component processing issues. It would therefore be expected that Inconel 617 and Nimonic 

263 have potential for 1200 K applications, though at very low stresses. Incoloy MA956 could have 

potential for application for higher stresses, but would require an extensive process development program 

in order to form and join duct components from this alloy. 

Several creep tests were performed at 1200 K to compare the relative creep resistance of Hastelloy X 

and Inconel 617 in the plate longitudinal and width directions. Specimens in the plate longitudinal 

direction had superior response for both alloys, as shown in figure 16. While the differences were only 2 

to 3×, this direction dependence could be an important consideration in duct design, where internal 

pressure-induced stresses for tubes rolled from flat plates are often highest in the hoop direction. 

Impeller and Duct Material Selection 

Impeller material.—For applications having maximum temperatures near 978 K, current results 

indicate LSHR would appear to be the preferable impeller material. It has higher strength than MarM247LC and IN792, and creep resistance as good as Mar-M247LC. Other work currently being prepared 

for publication indicates LSHR also has far superior high cycle fatigue resistance to Mar-M247LC in tests 

at 978 and 1090 K. For applications having maximum temperatures of 1090 K and higher, conventionally 

cast Mar-M247LC would appear to be the material of choice. It had far superior creep response to LSHR 

and IN792, with roughly comparable tensile strength at these temperatures. This was related in part to the 

coarse grain size of the conventionally cast Mar-M247LC bars. Indeed, the differences in response 

between Mar-M247LC and IN792 may be largely due to their different casting processes. Inspection of 

table 1 indicates that Mar-M247LC and IN792 have very similar compositions with respect to total 

content of γ′ stabilizing elements (Al+Ti+Ta+Nb+Hf), although hafnium is absent in IN792. Yet IN792 

was cast as an actual impeller for an auxiliary power unit using a common fine grain casting process, 

while most Mar-M247LC were conventionally cast as simple bars. The IN792 properties may therefore 

reflect those to be expected in Mar-M247LC, when the latter alloy is cast into impellers using typical 

production casting practices (ref. 17). 

Duct Materials.—For applications having maximum temperatures of 978 to 1089 K, current results 

indicate Inconel 617 would be the solid solution strengthened material of choice over Hastelloy X, due to 

superior creep resistance. This alloy would be relatively easy to fabricate into duct components. The 

precipitation strengthened alloys Inconel 740 and Nimonic 263 offer further improvements in creep 

NASA/TM—2006-214110 6
resistance. However, significant additional process development efforts would be necessary to shape, join, 

and heat treat duct components for these alloys. For applications having maximum temperatures of 

1200 K, none of the tested duct alloys looked overly promising. Inconel 617 could be the preferable 

selection, albeit at low creep stresses. Inconel 740 and Nimonic 263 offered no evident improvements in 

creep resistance, and would again have the processing development issues and also phase stability 

questions. Incoloy MA956 might offer improvements in creep resistance, but would have extensive 

processing issues to be addressed. 

Summary and Conclusions 

Two cast superalloys, Mar-M247LC and IN792, and a NASA powder metallurgy superalloy, LSHR, 

were screened to compare their respective capabilities for impeller applications. Several wrought 

superalloys including Hastelloy X, Inconel 617, Inconel 740, Nimonic 263, and Incoloy MA956 were 

also screened to compare their capabilities for duct applications. Conventional tensile and creep tests were 

performed at temperatures up to 1200 K on specimens extracted from the materials. Initial microstructure 

evaluations were also undertaken. The results can be summarized as follows: 

(1) Impeller materials: LSHR had higher strength than Mar-M247LC and IN792 at 978 K, but 

comparable strengths at 1090 and 1200 K. LSHR had comparable creep resistance to Mar-M247LC at 

978 K. At higher temperatures, conventionally cast Mar-M247LC had superior creep resistance. Initial 

estimates suggest a target strain rate of 0.1 percent per year for Mar-M247 could be achieved at 100 MPa 

and 1090 K or 350 MPa and 978 K. 

(2) Duct/heat exchanger materials: The precipitate strengthened superalloys Inconel 740 and Nimonic 

263 had improved creep response over solid solution strengthened superalloys Inconel 617 and Hastelloy 

X at temperatures to 1090 K. At temperatures near 1200 K, all these alloys had low creep stress bearing 

capabilities. Mechanically alloyed Incoloy MA956 had improved creep resistance at this temperature. 

It can be concluded from this evaluation that LSHR should be further evaluated for impeller 

applications with maximum temperatures near 978 K. Mar-M247LC and IN792 should be further 

evaluated for impeller applications requiring higher temperatures. The solid solution strengthened 

superalloy Inconel 617 should be further evaluated for duct applications at maximum temperatures up 

to 1089 K. A precipitation hardened alloy such as Nimonic 263 should also be further studied for duct 

applications. For both impeller and duct materials, the material processing and fabrication procedures 

necessary to produce these components should also be considered and screened, in order to assess 

material potentials and properties as real components. 

References 

1. L.S. Mason, “A Power Conversion Concept for the Jupiter Icy Moons Orbiter,” Journal of Propulsion 

and Power, vol. 20, no. 5, Sept. 2005, pp. 902–910. 

2. M.J. Barrett, “Expectations of Closed-Brayton-Cycle Heat Exchangers in Nuclear Space Power 

Systems,” Journal of Propulsion and Power, vol. 21, no. 1, January 2005, pp. 152–157. 

3. R.L. Ammon, L.R. Eisenstatt, and G.O. Yatsko, “Creep Rupture Behavior of Selected Turbine 

Materials in Air, Ultra-High Purity Helium, and Simulated Closed Cycle Brayton Helium Working 

Fluids,” Journal of Engineering for Power, vol. 103, April 1981, pp. 331–337. 

4. K. Harris, G.L. Erickson, and R.E. Schwer, “Development of a High Creep Strength, High Ductility, 

Cast Superalloy for Integral Turbine Wheels,” 1982 AIME Annual Meeting, Feb. 1982. 

5. C.T. Sims, N.S. Stoloff, and W.C. Hagel, Superalloys II, John Wiley & Sons, New York, NY, 1987, 

pp. 583–585. 

NASA/TM—2006-214110 7
6. T.P. Gabb, J. Gayda, J. Telesman, and P.T. Kantzos, “Thermal and Mechanical Property 

Characterization of the Advanced Disk Alloy LSHR,” NASA/TM—2005-213645, Washington, D.C., 

June 2005. 

7. “Hastelloy X Alloy,” technical publication by Haynes International, Inc., 1997. 

8. “Inconel alloy HX,” technical publication by Special Metals Corp., 2003. 

9. “Inconel alloy 617,” technical publication by Special Metals Corp., 2005. 

10. “Incoloy alloy MA956,” technical publication by Special Metals Corp., 2003. 

11. “Inconel alloy 740,” technical publication by Special Metals Corp., 2004. 

12. G. Smith, L. Shoemaker, “Advanced Nickel Alloys For Coal-Fired Boiler Tubing,” Advanced 

Materials and Processes, July 2004, pp. 23–26. 

13. “Nimonic alloy 263,” technical publication by Special Metals Corp., 2003. 

14. C.T. Sims, N.S. Stoloff, and W.C. Hagel, Superalloys II, John Wiley & Sons, New York, NY, 1987, 

p. 477. 

15. F.R. Larson and J. Miller, Trans. ASME, vol. 74, 1952, pp. 765–766. 

16. J. Gayda, and T.P. Gabb, “Two-Dimensional Viscoelastic Stress Analysis of a Prototypical JIMO 

Turbine Wheel,” NASA/TM—2005-213650, Washington, D.C., June, 2005. 

17. I.R. Delgado, G.R. Halford, B.M. Steinetz, and C.M. Rimnac, “Strain-Life Assessment of Grainex 

Mar-M247 for NASA’s Turbine Seal Test Facility,” NASA/TM—2004-212985 (ARL–TR–3178), 

National Aeronautics and Space Administration, Washington, D.C., April 2004. 

NASA/TM—2006-214110 8
TABLE 1.—MEASURED CHEMISTRIES OF IMPELLER AND DUCT SUPERALLOYS 

Alloy, 

wt% 

Al B C Co Cr Cu Fe Hf Mh Mo Ni Nb O Si Ta Ti V W Y Zr 

NASA LSHR 3.46 0.028 0.029 20.7 12.52 ----- 0.07 ----- ----- 2.73 Bal. 1.45 ------- 0.03 1.6 3.5 ------- 4.33 ----- 0.049 Conv. Cast Mar-M247LC 5.6 0.015 0.09 9.4 8.3 ----- ------- 1.4 0.02 0.5 Bal. ------- ------- 0.01 3.2 0.7 ------- 9.5 ----- 0.011 Microcast Mar-M247LC 5.52 0.002 0.0855 9.31 8.52 ----- 0.04 1.46 ----- 0.51 Bal. ------- ------- 0.05 3.15 0.75 ------- 9.47 ----- 0.011 IN792 3.45 0.0015 0.08385 8.95 12.635 ----- 0.01 0.02 ----- 1.83 Bal. ------- ------- 0.03 3.975 3.835 ------- 4.08 ----- 0.015 Hastelloy X 0.19 -------- 0.0855 1.01 21.435 0.24 19.485 ----- 0.525 8.5 Bal. 0.115 ------- 0.31 0.008 0.02 0.025 0.455 ----- ------- Incone1 617 0.985 -------- 0.0945 11.64 22.65 0.05 0.905 ----- 0.04 9.45 Bal. 0.12 ------- 0.155 0.008 0.36 0.0055 0.035 ----- ------- Incone1 740 0.93 -------- 0.031 19.945 24.24 0.04 0.445 ----- 0.25 0.51 Bal. 1.96 ------- 0.525 ------- 1.71 ------- 0.01 ----- 0.011 Nimonic 263 0.45 -------- 0.061 19.3 20.45 0.02 0.38 ----- 0.33 5.8 Bal. 0.05 ------- 0.125 0.0065 2.19 ------- 0.01 ----- ------- Incoloy MA956 4.555 -------- 0.01675 -------- 19.7 0.01 Bal. ----- 0.09 0.01 0.13 0.005 0.2035 0.08 ------- 0.365 0.025 0.01 0.29 ------- 

TABLE 2.—TENSILE TEST RESULTS FOR IMPELLER SUPERALLOYS

 Temperature, 

°F 

Temperature, 

K 

0.2% yield, 

ksi 

Yield, 

MPa 

Ultimate tensile 

strength, 

ksi 

Ultimate tensile 

strength, 

MPa 

Elongation, 

% 

Reduction in 

area, 

% 

LSHR Tensile 1300 978 136.9 944 184.9 1275 24.0 32.6 LSHR Tensile 1300 978 155.6 1073 200.1 1380 10.4 11.7 Mar-M247LC Tensile 1300 978 111.3 767 151.6 1045 12.9 23.0 Mar-M247LC Tensile 1300 978 111.1 766 145.7 1005 12.1 19.1 LSHR Tensile 1500 1089 119.3 823 143.6 990 13.8 20.6 LSHR Tensile 1500 1089 133.9 923 157.0 1083 5.6 10.6 Mar-M247LC (Conv.) Tensile 1500 1089 121.4 837 147.2 1015 9.3 23.0 Mar-M247LC (Conv.) Tensile 1500 1089 120.2 829 147.9 1020 11.9 30.7 Mar-M247LC (Micro.) Tensile 1500 1089 117.1 807 147.8 1019 5.4 10.2 IN792 Tensile 1500 1089 104.5 721 135.4 934 9.0 21.5 LSHR Tensile 1700 1200 71.8 495 94.8 654 10.4 15.6 LSHR Tensile 1700 1200 76.9 530 95.4 658 5.6 8.7 Mar-M247LC (Conv.) Tensile 1700 1200 73.3 505 106.2 732 9.8 15.6 Mar-M247LC (Conv.) Tensile 1700 1200 73.2 505 106.7 736 12.1 20.8 Mar-M247LC (Micro.) Tensile 1700 1200 75.5 521 107.9 744 4.9 7.2 IN792 Tensile 1700 1200 61.6 425 86.7 598 18.5 25.1 

NASA/TM—2006-214110 9
TABLE 3.—CREEP TEST RESULTS FOR IMPELLER SUPERALLOYS 

Alloy Process Temperature, 

°F 

Temperature,

K 

Stress, 

ksi 

Stress, 

MPa 

0.1% creep 

life, 

h 

0.2% creep 

life, 

h 

0.3% creep 

life, 

h 

Rupture 

life, 

h 

Elongation, 

% 

Reduction

in area, 

% 

Notes 

LSHR PM Disk 1200 922 90 621 90.9 3769.4 --- -------- --- --- LSHR PM Disk 1300 978 100 690 115.0 174.0 --- -------- --- --- LSHR PM Disk 1300 978 100 690 149.0 254.0 --- -------- --- --- LSHR PM Disk 1300 978 80 552 1877.0 3082.0 --- -------- --- --- LSHR PM Disk 1500 1089 45 310 61.0 281.0 603 1068.7 5.8 6.5 LSHR PM Disk 1500 1089 30 207 414.0 1098.0 --- -------- --- --- LSHR PM Disk 1500 1089 15 103 1805.0 3356.0 --- -------- --- --- LSHR PM Disk 1500 1089 3 21 478.0 3888.0 --- -------- --- --- LSHR PM Disk 1500 1089 25 172 6.8 80.0 --- -------- --- --- LSHR PM DiSk 1700 1200 20 138 5.7 13.6 --- 121.7 8.8 10 LSHR PM Disk 1700 1200 10 69 9.9 24.7 40 664.9 26.9 28.4 LSHR PM Disk 1700 1200 5 34 30.5 86.0 --- 3109.6 83 56.9 

MM247LC ConvCastBar 1300 978 100 690 30.0 104.0 --- -------- --- --- MM247LC ConvCastBar 1300 978 100 690 18.6 55.0 --- -------- --- --- MM247LC ConvCastBar 1300 978 80 552 1424.0 3416.0 --- -------- --- --- MM247LC ConvCastBar 1500 1089 45 310 515.0 1014.0 --- -------- --- --- MM247LC ConvCastBar 1500 1089 30 207 1296.0 7295.0 --- -------- --- --- MM247LC ConvCastBar 1500 1089 30 207 502.5 1952.0 --- -------- --- --- Helium env. MM247LC ConvCastBar 1700 1200 30 207 29.0 74.0 187 546.4 7.6 12.4 MM247LC ConvCastBar 1700 1200 20 138 837.0 1422.0 --- -------- --- --- MM247LC ConvCastBar 1700 1200 10 69 334.1 3454.1 --- -------- --- --- MM247LC ConvCastBar 1700 1200 10 69 2274.0 5891.0 --- -------- --- --- Helium env. MM247LC MicroCastBar 1300 978 80 552 634.7 --- -------- --- --- MM247LC MicroCastBar 1500 1089 45 310 48.7 188.0 --- -------- --- --- 

IN792GX Grainex Cast 

Impeller 

1300 978 80 552 1153.8 --- -------- --- --- 

IN792GX Grainex Cast 

Impeller 

1300 978 80 552 146.0 940.0 --- -------- --- --- 

IN792GX Grainex Cast 

Impeller 

1500 1089 45 310 126.0 340.0 --- -------- --- --- 

IN792GX Grainex Cast 

Impeller 

1500 1089 45 310 110.0 272.0 764 -------- --- --- 

NASA/TM—2006-214110 10
TABLE 4.—TENSILE TEST RESULTS FOR DUCT SUPERALLOYS

Alloy Test Orientation Temperature, 

°F 

Temperature, 

K 

0.2% yield, 

ksi 

Yield, 

MPa 

Ultimate tensile 

strength, 

ksi 

Ultimate tensile 

strength 

MPa 

Elongation,

% 

Reduction 

in area, 

% 

Inconel 617 Tensile Longitudinal 75 297 70 483 123 848 47 58 Inconel 617 Tensile Longitudinal 75 297 70 483 124 855 46 58 Inconel 617 Tensile Width 75 297 66.5 459 122 841 49 59 Inconel 617 Tensile Width 75 297 66.5 459 122 841 49 59 Hastelloy X Tensile Longitudinal 1500 1089 24.6 170 48 331 67 78 Hastelloy X Tensile Longitudinal 1500 1089 24.4 168 48.6 335 66 75 Inconel 617 Tensile Longitudinal 1500 1089 40.6 280 57 393 45 55 Inconel 617 Tensile Longitudinal 1500 1089 40.6 280 57 393 51 54 Inconel 617 Tensile Width 1500 1089 41.7 288 56 386 50 56 Inconel 617 Tensile Width 1500 1089 42.5 293 56.5 390 49 55 Inconel 740 Tensile Longitudinal 1500 1089 73 503 90.5 624 8.5 18 Inconel 740 Tensile Longitudinal 1500 1089 70 483 85.5 590 8 16 Nimonic 263 Tensile Longitudinal 1500 1089 62 427 80 552 27 47 Nimonic 263 Tensile Longitudinal 1500 1089 60 414 79 545 28 43 Hastelloy X Tensile Longitudinal 1700 1200 18.6 128 26.9 185 77 81 Hastelloy X Tensile Longitudinal 1700 1200 19.3 133 27.5 190 80 84 Inconel 617 Tensile Longitudinal 1700 1200 21.4 148 31.1 214 54 59 Inconel 617 Tensile Longitudinal 1700 1200 21.8 150 31.6 218 52 55 Inconel 617 Tensile Width 1700 1200 22.9 158 31.2 215 49 53 Inconel 617 Tensile Width 1700 1200 24.9 172 31.9 220 49 60 Inconel 740 Tensile Longitudinal 1700 1200 29.1 201 35.8 247 47 86 Inconel 740 Tensile Longitudinal 1700 1200 30.1 208 35.7 246 46 79 Nimonic 263 Tensile Longitudinal 1700 1200 16.1 111 24.6 170 78 96 Nimonic 263 Tensile Longitudinal 1700 1200 16.4 113 24.7 170 73 95 

NASA/TM—2006-214110 11
TABLE 5.—CREEP TEST RESULTS FOR DUCT SUPERALLOYS

Alloy Process Temperature, 

°F 

Temperature, 

K 

Stress, 

ksi 

Stress, 

MPa 

0.1% creep 

life, 

h 

0.2% creep 

life, 

h 

0.3% creep 

life, 

h 

Rupture 

life, 

h 

Elongation, 

% 

Reduction 

in area, 

% 

Hastelloy X Plate 1500 1089 15 103 0.5 0.8 1.8 97.3 64.6 76.1 Hastelloy X Plate 1500 1089 15 103 0.4 0.7 1.7 ------- ----- ----- Hastelloy X Plate 1700 1200 5 34 0.6 3.1 9.8 459.6 77.8 ----- Hastelloy X Plate 1700 1200 2 14 107.0 314.0 ------- ------- ----- ----- Hastelloy X Plate 1700 1200 2 14 48.0 254.0 941 ------- ----- ----- Inconel 617 Plate 1500 1089 15 103 1.5 8.3 19 ------- ----- ----- Inconel 617 Plate 1500 1089 15 103 1.8 10.9 ------- 769.3 29.4 34.6 Inconel 617 Plate 1700 1200 5 34 157.0 480.0 785 2028.1 13.2 19.9 Inconel 617 Plate 1700 1200 2 14 216.0 1092.0 ------- ------- ----- ----- Inconel 617 Plate 1700 1200 2 14 58.0 361.0 1737 ------- ----- ----- Inconel 740 Plate 1500 1089 15 103 115.0 277.0 741 ------- ----- ----- Inconel 740 Plate 1500 1089 15 103 245.0 470.0 667 ------- ----- ----- Inconel 740 Plate 1700 1200 5 34 8.7 26.0 71 1563.6 40 51.3 Inconel 740 Plate 1700 1200 2 14 35.8 96.0 145 ------- ----- ----- Nimonic 263 Plate 1500 1089 15 103 193.0 355.0 474 ------- ----- ----- Nimonic 263 Plate 1500 1089 15 103 78.0 257.0 358 1978.8 20.3 19.4 Nimonic 263 Plate 1700 1200 5 34 0.9 2.4 ------- 239.2 35.5 33.2 Nimonic 263 Plate 1700 1200 2 14 400.0 754.0 1042 ------- ----- ----- MA956 Plate 1700 1200 2 14 5995.0 -------- ------- ------- ----- ----- 

NASA/TM—2006-214110 12
Figure 1.—Typical microstructures of impeller alloys: a) LSHR, b) conventionally cast 

Mar-M247LC, c) Microcast Mar-M247LC, d) IN792.

a) LSHR b) Conventionally Cast Mar-M247LC

c) Microcast Mar-M247LC c) IN792

NASA/TM—2006-214110 13
Figure 2.—Typical microstructures of impeller alloys: a) LSHR, b) conventionally cast

Mar-M247LC, c) Microcast Mar-M247LC, d) IN792.

a) LSHR b) Conventionally Cast Mar-M247LC

c) Microcast Mar-M247LC c) IN792

NASA/TM—2006-214110 14
Figure 3.—Typical microstructures of duct superalloys: a) Hastelloy X, b) Inconel 617,

c) Inconel 740, d) Nimonic 263, e) Incoloy MA 956.

a) Hastelloy X b) Inconel 617

c) Inconel 740 d) Nimonic 263

e) Incoloy MA956

NASA/TM—2006-214110 15
Figure 4.—Typical microstructures of duct superalloys: a) Hastelloy X, b) Inconel 617,

c) Inconel 740, d) Nimonic 263, e) Incoloy MA 956.

a) Hastelloy X b) Inconel 617

c) Inconel 740 d) Nimonic 263

e) Incoloy MA956

NASA/TM—2006-214110 16
Figure 5.—Tensile response of impeller alloys: a) yield strength, b) ultimate strength, c) elongation,

d) reduction in area.

a) b)

c) d)

NASA/TM—2006-214110 17
Figure 7.—Comparison of creep response of LSHR and conventionally

cast Mar-M247LC.

Figure 6.—Comparison of creep response of conventionally 

cast Mar-M247LC in helium and air.

NASA/TM—2006-214110 18
Figure 9.—Larson-Miller parameter comparison of creep response

of LSHR and Mar-M247LC.

Figure 8.—Comparison of creep response of conventionally 

cast and Microcast® Mar-M247LC, along with IN792.

NASA/TM—2006-214110 19
Figure 10.—Creep rate versus stress for LSHR.

Figure 11.—Creep rate versus stress for conventionally cast Mar-M247LC.

NASA/TM—2006-214110 20
Figure 12.—Tensile response of selected duct alloys: a) yield strength, b) ultimate strength, 

c) elongation, d) reduction in area.

a) b)

c) d)

NASA/TM—2006-214110 21
Figure 13.—Tensile response of Inconel 617 in plate longitudinal versus width directions: 

a) yield strength, b) ultimate strength, c) elongation, d) reduction in area.

a) b)

c) d)

NASA/TM—2006-214110 22
Figure 15.—Comparison of creep resistance for duct alloys at 1200K/14 MPa.

Figure 14.—Comparison of creep resistance for duct alloys at 1089K/103 MPa.

NASA/TM—2006-214110 23
L

L

W W

Figure 16.—Comparison of creep resistance of duct alloys Hastelloy X and 

Inconel 617 in the plate longitudinal and width directions at 1200K/14 MPa.

NASA/TM—2006-214110 24
This publication is available from the NASA Center for AeroSpace Information, 301–621–0390.

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E–15444

WBS–22–973–80–50

30

Tensile and Creep Property Characterization of Potential Brayton Cycle

Impeller and Duct Materials

Timothy P. Gabb and John Gayda

Superalloys; Rotating impellers; Creep

Unclassified -Unlimited

Subject Category: 26

Responsible person, Timothy P. Gabb, organization code RXA, 216–433–3272.

This paper represents a status report documenting the work on creep of superalloys performed under Project

Prometheus. Cast superalloys have potential applications in space as impellers within closed-loop Brayton cycle

nuclear power generation systems. Likewise wrought superalloys are good candidates for ducts and heat exchangers

transporting the inert working gas in a Brayton-based power plant. Two cast superalloys, Mar-M247LC and IN792,

and a NASA GRC powder metallurgy superalloy, LSHR, are being screened to compare their respective capabilities

for impeller applications. Several wrought superalloys including Hastelloy X, (Haynes International, Inc., Kokomo, IN),

Inconel 617, Inconel 740, Nimonic 263, and Incoloy MA956 (Special Metals Corporation, Huntington, WV) are also

being screened to compare their capabilities for duct applications. These proposed applications would require sufficient

strength and creep resistance for long term service at temperatures up to 1200 K, with service times to 100,000 h or

more. Conventional tensile and creep tests were performed at temperatures up to 1200 K on specimens extracted from

the materials. Initial microstructure evaluations were also undertaken.