Malcolm K. Stanford

Glenn Research Center, Cleveland, Ohio

Charpy Impact Energy and Microindentation 

Hardness of 60-NITINOL

NASA/TM—2012-216029

September 2012
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Malcolm K. Stanford

Glenn Research Center, Cleveland, Ohio

Charpy Impact Energy and Microindentation 

Hardness of 60-NITINOL

NASA/TM—2012-216029

September 2012

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135
Acknowledgments

Thank you to Joy A. Buehler, Dereck F. Johnson, Terry R. McCue, Richard B. Rogers, Fransua Thomas and Walter A. Wozniak 

for their technical assistance during this study and to Dr. Christopher DellaCorte for helpful technological discussions.

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Hanover, MD 21076–1320

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

Level of Review: This material has been technically reviewed by technical management. 

This report contains preliminary findings, 

subject to revision as analysis proceeds.
NASA/TM—2012-216029 1 

Charpy Impact Energy and Microindentation 

Hardness of 60-NITINOL

Malcolm K. Stanford

National Aeronautics and Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Abstract

60-NITINOL (60 wt.% Ni – 40 wt.% Ti) is being studied as a 

material for advanced aerospace components. The Charpy 

impact energy and microindentation hardness has been studied 

for this material, fabricated by vacuum induction skull melting 

(casting) and by hot isostatic pressing. Test specimens were 

prepared in various hardened and annealed heat treatment 

conditions. The average impact energy ranged from 0.33 to 

0.49 J for the hardened specimens while the annealed specimens 

had impact energies ranging from 0.89 to 1.18 J. The average 

hardness values of the hardened specimens ranged from 590 to 

676 HV while that of the annealed specimens ranged from 298 

to 366 HV, suggesting an inverse relationship between impact 

energy and hardness. These results are expected to provide 

guidance in the selection of heat treatment processes for the 

design of mechanical components.

Introduction

The topic of impact toughness rose to prominence after the 

sinking of the Royal Mail Steamship Titanic. The forward 

starboard hull of the ship struck an iceberg at about 11:40 p.m. 

on the night of April 15, 1912, resulting in the sinking of the 

ship in approximately two and a half hours (Ref. 1). Though 

the ship was thought by many to be unsinkable, its designers 

were apparently unaware of a basic, intrinsic materials 

property, namely, the ductile-to-brittle transition temperature. 

The icy waters of the Atlantic embrittled the 

crystallographically body-centered cubic (bcc) iron of the 

ship’s hull, reducing its impact energy (Ref. 2). Unfortunately, 

the human drive for exploration and progress has had a 

number of monumental disasters that, while paralleled by 

some as the inevitable result of Homerian hubris (Ref. 3), have

alternatively served as educational opportunities to enable 

future progress.

In the century that has passed since the fateful first voyage 

of the Titanic, a great deal of effort has been focused on 

studying the response of materials to impact loading. It is now 

well-known that the body-centered cubic crystal structure has 

fewer slip systems than the face-centered cubic system, 

reducing the ability of the bcc system to deform prior to 

fracture and making it particularly susceptible to a transition to 

brittle fracture at lower temperatures. Lots of ongoing work is 

focused on understanding the behavior of ship hull materials 

under dynamic loads (Refs. 4 to 6). Similarly, there are many 

investigations on the response of jet engine and airframe 

structural materials to foreign object damage (Refs. 7 to 9). 

Impact damage is certainly a concern in crewed space 

vehicles (Refs. 10 and 11). However, internal components 

must also be able to withstand shock loads due to high gforce-induced vibration during launch, especially in un-crewed 

vehicles where g-forces can become very high. The National 

Aeronautics and Space Administration is currently studying 

60-NITINOL for incorporation in aerospace gears and 

bearings. 60-NITINOL is an ordered intermetallic material

consisting of 60wt%Ni and 40wt%Ti (Ref. 12). This material 

is of interest to the aerospace community because it is 

corrosion-resistant, has moderate density, can be easily 

hardened for wear resistance and because it is superelastic, 

capable of recoverable elastic strains nearly ten times those of 

conventional materials (Ref. 13). However, as with most 

intermetallics, this material is inherently brittle. 60-NITINOL 

has the B2 crystal structure, very similar to the bcc crystal 

structure, with Ni atoms at the eight corners of the unit cell 

and Ti at the body-centered position. The lack of available slip 

systems and the rigid arrangement of atoms in this material are 

the underlying (atomic-scale) reasons for its brittle behavior. 

However, the mechanisms that allow brittle failure to initiate 

on the micro-scale should be understood so that methods to 

reduce or prevent these types of failures may be developed.

The focus of this investigation was to identify the failure 

mechanisms of this material in various heat treatment 

conditions when subjected to impact loading. While 

mechanical tests are helpful for the specification of an 

operating environment, these tests typically only apply static 

loads to the material, where actual components are often 

subjected to dynamic loads. These results will provide a 

conservative estimation of the behavior of this material in lieu 

of forthcoming static compression strength and fracture 

toughness evaluations. 

Materials and Procedures

60-NITINOL (nominal composition: 60wt% Ni - 40wt% Ti) 

was obtained from commercial sources, fabricated by the 

following methods:

1. Vacuum induction skull melting of bulk 60-NITINOL 

with a water-cooled copper crucible. This method is 
NASA/TM—2012-216029 2 

hereafter referred to as casting (resulting in cast 

material). The cast material was subsequently hot 

isostatically pressed for 2 hr at 900 °C and 100 MPa to 

diminish porosity. The materials from these two heats 

are hereafter designated “Cast-A” and “Cast-B” 

material.

2. Hot isostatic pressing (HIPping) of gas atomized 60-

NITINOL powder. This method is henceforth referred to 

as the powder metallurgy (or PM) method (resulting in 

PM material). The two ingots of material fabricated by 

this technique that were used in this study were obtained 

from two different suppliers using proprietary 

techniques. The materials from these two ingots are 

subsequently identified as “PM-X” and “PM-Y” 

material.

The chemical compositions of the materials used in this 

study were analyzed by inductively-coupled plasma optical 

emission spectroscopy (ICP-OES).

Specimens for Charpy impact testing, depicted 

schematically in Figure 1, were cut from the ingots by wire 

electrical-discharge machining (EDM). The specimens were 

heat treated by water-quenched, aged or furnace-cooled heat 

treatments as listed in Table I, and then tested as specified in 

ASTM E23-07aε1 (Ref. 14). The Charpy impact test measures 

the energy required to fracture a specimen by determining the 

difference between the initial height h and final height h’ of a 

hammer (or striker) on a pendulum when it impacts a 

specimen of the studied material. The setup used in this study 

is depicted schematically in Figure 2. Three to four replicate 

specimens with each of the specified heat treatment conditions 

were tested at room temperature on an apparatus designed to 

measure impact energies from 0 to 2.82 J (0 to 25 in.-lb) with 

0.028 J (0.25 in.-lb) divisions. The speed of the striker at 

impact was approximately 3.46 m/s. The striker profile used 

for these tests had a radius of 8 mm and the chordal distance 

of this radius was approximately 4 mm. After testing, the 

fractured surfaces of the specimens were examined by 

scanning electron microscopy (SEM) and by energydispersive x-ray spectroscopy (EDS).

Cross sections of specimens were sectioned perpendicular to 

the direction of crack propagation with an abrasive cut-off saw 

using a resin bonded aluminum oxide blade. The wheel rotated 

at 3820 rpm and was set to have a traverse speed of 0.1 mm/s.

Figure 1.—Schematic of Charpy U-notch specimen used in this 

investigation.

Figure 2.—Schematic representation of Charpy test of 

impact energy (Ref. 17).

TABLE I.—HEAT TREATMENT CONDITIONS USED ON CHARPY SPECIMENS IN THIS STUDY

Designation Heat treatment for cast material Heat treatment for PM material

Water-quenched 

(partially solution treated)

a

90 min at 980 °C/water quench (WQ) a60 min at 1000 °C/WQ

Aged 

(water-quenched and aged)

90 min at 980 °C/WQ/60 min at 400°C/WQ 60 min at 1000 °C/WQ/60 min at 400 °C/WQ

Furnace-cooled b70 min at 1040 °C/furnace cool (FC) b120 min at 1000 °C/FC a

Water quenching was done in still room temperature water. The specimen was swirled about in the water.

b

Furnace cooling was done by shutting off the power to the furnace after the specified heat soak and allowing it to cool to room 

temperature, which took approximately 24 hr. 
NASA/TM—2012-216029 3 

The specimens were prepared for microstructural examination 

using standard metallographic procedures. A total of 

approximately 1 mm of each specimen was removed from the 

sectioned surface during the grinding and polishing procedure. 

The final polish used colloidal silica on a vibratory polisher. 

The microindentation hardness was measured with a standard 

technique using a Vickers (pyramidal) diamond indenter with 

a 500 gf load and a load apply of approximately 12 s (Ref. 15). 

The polished cross sections were then swab-etched with a 

1vol%HF-10vol%HNO3 solution in deionized water and then 

examined with an optical microscope.

Results and Discussion

The results of the chemical analysis are listed in Table II. 

The composition of each of the studied specimens was

nominally 60wt%Ni and 40wt%Ti with oxygen, iron, copper

and aluminum making up the primary impurities. Due to the 

high reactivity of titanium at high temperatures, some of the 

oxygen was likely adsorbed during primary metallurgical 

processing. In addition, some of the aluminum, copper and 

oxygen are probably the result of low-level contamination 

from alumina- and copper-based melting equipment.

TABLE II.—NOMINAL COMPOSITIONS 

OF THE STUDIED SPECIMENS

Identification Composition Impurities, 

ppm

Cast-A 59.6wt%Ni40.3wt%TiO (680), Fe (130), Al (200), N 

(40), Co (30), Cr (20)

Cast-B 60.1wt%Ni39.8wt%TiO (640), Fe (430), Al (200), N 

(70), Cu (250), Co (25), Zr (30), 

V (25), Mn (4)

PM-X 59.4wt%Ni40.5wt%TiO (580), Fe (100), Al (80), Cu 

(90), Cr (20), Co (5)

PM-Y 60.5wt%Ni39.4wt%TiO (395), Fe (130), Al (90), Cr 

(35), N (30)

Fractography

The fracture surfaces of the furnace-cooled specimens had a 

dull appearance when observed with the unaided eye. As 

shown in Figure 3, under SEM observation the fracture 

surfaces had a grainy texture, which is characteristic of 

intergranular fracture. There was no discernable fracture 

initiation site for these specimens. At high-magnification 

(Figure 4, Figure 5 and Figure 6), the fracture surface is 

faceted and clearly indicative of brittle intergranular fracture. 

Observation of the polished cross sections, as shown in 

Figure 7 and Figure 8, showed that the fracture tended to

follow the second phase Ni3Ti that typically decorates the 

grain boundaries of this material, further indicating that 

fracture was intergranular. 

 

Figure 3.—SEM photomicrographs of representative fracture 

surfaces for (a) Cast-A, (b) PM-X and (c) PM-Y 60-NITINOL 

materials in the furnace-cooled heat treatment condition showing 

intergranular fracture.
NASA/TM—2012-216029 4 

The fracture surface of a PM-X specimen (see Figure 5) 

shows unconsolidated spherical particles and depressions at 

the prior locations of unconsolidated particles. Examination of 

cross sections by optical microscopy (shown in Figure 7) 

corroborated this finding. Unconsolidated particles are thought 

to be a result of particles that were slightly oxidized during the 

atomization process. Work is underway to determine if this 

problem can be mitigated with this material through some 

combination of higher HIP temperatures and cycle times to 

Figure 4.—(a) The fracture surface of a Cast-A Charpy specimen 

in the furnace-cooled heat treatment condition (the highlighted 

area in (a) is shown at higher magnification in (b)). While there is 

no clear fracture initiation site, the brittle nature of the fracture is 

apparent from the sharp, faceted surface morphology. 

force the oxide layer to diffuse away from the particle 

boundaries. It is possible that this issue can be eliminated with 

tighter control of the atomization chamber atmosphere to 

prevent the oxidation in the first place. However, when even 

small amounts of oxygen come into contact with a high 

surface area powder with an oxygen-solvent like Ti, the 

kinetics driving the formation of oxides would be very strong. 

Notwithstanding, PM-Y specimens showed no indications of 

unconsolidated particles.

Figure 5.—The fracture surface of a PM-X Charpy specimen in the 

furnace-cooled heat treatment condition (the highlighted area in 

(a) is shown at higher magnification in (b)). While there is no 

clear fracture initiation site, the brittle nature of the fracture is 

apparent from the sharp, faceted surface morphology.
NASA/TM—2012-216029 5 

Figure 6.—The fracture surface of a PM-Y Charpy specimen in the 

furnace-cooled heat treatment condition (the highlighted area in 

(a) is shown at higher magnification in (b)). While there is no 

clear fracture initiation site, the brittle nature of the fracture is 

apparent from the sharp, faceted surface morphology.

Figure 7.—Cross-sectional image through the U-notch in furnacecooled PM-X material showing intergranular fracture (IGF) and 

unconsolidated particles (U). The rectangle in (a) indicates the 

proximal area shown at higher magnification in (b). 
NASA/TM—2012-216029 6 

Figure 8.—Optical photomicrograph of a cross-section (a) through 

the U-notch of a furnace-cooled PM-Y specimen and (b) at 

higher magnification showing intergranular fracture (second 

phase Ni3Ti decorates the fracture path along the top edge of 

the specimen).

The fracture surfaces of the water-quenched specimens were 

shiny to the naked eye. Observation by SEM revealed classical 

chevron patterns resulting from brittle fracture, which can be 

seen in Figure 9(a) to (c). The fracture origins (labeled “FO”) 

were determined by following the direction of crack 

propagation in reverse. 

The fracture origins in Cast-A specimens had particulate 

contamination, as shown in Figure 10. Upon examination by 

EDS (see Figure 10(d)), the particles were found to consist of 

Zr and O. Further investigation determined that the 

contaminant was zirconia (ZrO2), which had been used as part 

of a crucible lining during the casting process (Ref. 16). CastB material was melted in a crucible without this zirconia liner. 

Figure 9.—SEM photomicrographs of representative fracture 

surfaces for (a) Cast-A, (b) PM-X and (c) PM-Y 60-NITINOL 

materials in the water-quenched heat treatment condition 

showing the fracture origins (FO) and features characteristic of 

intragranular brittle fracture.
NASA/TM—2012-216029 7 

Figure 10.—Contamination in water-quenched Cast-A material in the water-quenched heat treatment condition. The highlighted area in a is 

shown at higher magnification in b. EDS-B represents the parent material (60-NITINOL), while EDS-C indicates that the particulate 

contaminant is composed of Zr and O.
NASA/TM—2012-216029 8 

Figure 11.—SEM photomicrograph of the fracture surface (a), the 

fracture initiation site (b) and a higher magnification image of the 

fracture initiation site (c) of a PM-Y Charpy specimen in the 

water-quenched heat treatment condition. Note the fibrous 

texture at the fracture initiation site, an indication of (micro-scale) 

ductility.

Higher-magnification images of the fracture surfaces of the 

water-quenched PM-Y material, shown in Figure 11 show 

slight indications of ductile fracture where the material 

displays a fibrous texture. This finding was quite surprising, 

given the brittle nature of intermetallics, and it indicates that 

the material had some ductility and yielded prior to failure. 

The fracture surfaces of the aged specimens, shown in 

Figure 12, were basically identical to those of the waterquenched specimens by the examination methods used in this 

study. Examination of the grainy contaminant in the Cast-A 

material seen in Figure 13 again showed that the contaminant 

consisted of Zr and O, as described in Figure 14 and 

Figure 15. Chemical analysis by ICP-OES showed that the 

chemical composition at the fracture origin consisted of 20

ppm Zr. No Zr was detected in the bulk material analysis, as 

listed in Table II, primarily because the contamination is 

sparsely distributed throughout the parent material but 

concentrated at fracture origins. This contamination increases 

the likelihood of fracture initiation since the particles are 

sharply faceted. As seen in Figure 13, the ZrO2 contaminant in 

the aged Cast-A material appears to be less integrated with the 

parent material than in the Cast-A water-quenched material

(see, for example, Figure 10(b)). It is possible that the 

contaminant particles remain trapped in the parent material 

after the water-quenching treatment but the aging treatment 

allows some subsequent residual stress relaxation, resulting in 

free-standing particles. Examination of a cross-section of the 

aged Cast-A specimen (shown in Figure 16) revealed further 

indication of contamination, which was found to be TiN (see 

Figure 17). This could be the result of some exposure to air 

during the casting process.

The fracture surface of the aged Cast-B specimen, shown in 

Figure 18, displays indications of brittle fracture as with the 

Cast-A specimens (e.g., Figure 13). An etched cross-section of 

this specimen (Figure 19) shows second phase Ni3Ti generally 

clustered near the grain boundaries.

Examination of the fracture origin of a PM-X specimen 

revealed faceted unconsolidated particles at the fracture origin, 

as shown in Figure 20. This is further evidence that 

unconsolidated particles are a factor in the fracture of this 

material, whether by reducing the energy required for fracture 

initiation or that for crack propagation. A cross-section of this 

specimen (Figure 21) also showed several unconsolidated 

particles.

Figure 22 shows photomicrographs of the fracture surface 

and a metallographic cross-section of an aged PM-Y 

specimen. The relatively flat fracture initiation site is at the 

center of the divergence of the river patterns and is indicative 

of brittle transgranular fracture. The microstructure 

(Figure 22(b)) also indicates that fracture was transgranular 

and that the material was well-consolidated with some 

indications of prior particle boundaries.
NASA/TM—2012-216029 9 

Figure 12.—SEM photomicrographs of representative fracture surfaces for (a) Cast-A, (b) Cast-B, (c) PM-X and (d) PM-Y 60-

NITINOL materials in the aged heat treatment condition showing the fracture origins (FO).
NASA/TM—2012-216029 10

Figure 13.—Fracture origin and adjacent fracture surface for aged Cast-A 

specimen showing river patterns (R) indicating brittle fracture and 

particulate contamination (C) at fracture origin. 
NASA/TM—2012-216029 11

Figure 14.—Backscattered electron images at (a) low- and 

(b) high-magnification at edge of fracture origin in a Cast-A 

specimen. EDS spectra are shown in the following figure, 

indicate that the contamination is zirconia, while the parent 

material is Ni-Ti.
NASA/TM—2012-216029 12

Figure 15.—EDS spectra from Figure 14 showing that the contamination (a) is primarily Zr and O, while the parent material (b) as well as 

the transition region (c) and (d) is Ni and Ti.
NASA/TM—2012-216029 13

Figure 16.—Cross-section of aged Cast-A specimen (a) perpendicular to 

fracture surface and (b) at higher magnification showing intragranular 

fracture and an unexpected precipitate phase within the grains.
NASA/TM—2012-216029 14

Figure 17.—(a) Secondary electron and (b) backscattered electron SEM images and accompanying EDS spectra ((c) to (e)) of one of the 

second phase features within the grains of an aged Cast-A Charpy specimen indicating that this feature is a contaminant composed of 

TiN (EDS-C) and TiC (EDS-D).
NASA/TM—2012-216029 15

Figure 18.—Fracture surface of aged Cast-B with microstructural 

features indicating brittle fracture. The highlighted area in (a) is 

shown at higher magnification in (b) and likewise for the 

highlighted area in (b). 

Figure 19.—Cross-section of aged Cast-B specimen through 

fracture surface (U-notch on right).

Figure 20.—Unconsolidated particles at the fracture origin of a PMX Charpy specimen in the aged heat treatment condition
NASA/TM—2012-216029 16

Figure 21.—Cross section of aged PM-X through U-notch, near 

fracture surface. Several unconsolidated particles are called out 

with arrows. The pore at P is likely the product of a particle with 

a hollow core.

Figure 22.—(a) SEM of the fracture surface and (b) optical 

photomicrograph of a cross-section through the U-notch of a 

Charpy specimen fabricated from PM-Y in the aged heat treated 

condition.

Hardness and Impact Energy

The impact energy results are listed in Table III. The 

tabulated values are the average and standard deviation of 

three to four replicate specimens. The aged and waterquenched specimens had average impact energies of 0.43 ± 

0.04 and 0.39 ± 0.08 J, respectively. Based on the stated 

statistical variation of the averages for the studies specimens, 

there was essentially no difference between the impact 

energies of materials heat treated by these two methods. The 

furnace-cooled specimens, however, had an average impact 

energy of 1.02 ± 0.14 J, more than twice that of either aged or 

water-quenched specimens. While this difference in impact 

energy is significant with respect to the effect of the studied 

heat treatments on this 60-NITINOL, it should be noted that 

the impact energy of this material is between that of metallic 

and ceramic materials used in bearings. Examples of the 

impact energies of such materials are listed in Table IV. 

The microindentation hardness results are shown in 

Table III. These values represent the average and standard 

deviation of five repeat measurements, approximately 500 μm 

apart. The data shows that the hardness is highest for the 

water-quenched and aged specimens. The aged and waterquenched specimens had average hardness values of 641 ± 42 

and 664 ± 19 HV, respectively. The furnace-cooled specimens 

had an average hardness value of 324 ± 39 HV, approximately 

half that of either the aged or water-quenched specimens 

Based on the given statistical variation, there is no difference 

in the hardness values for specimens with either the waterquenched or aged heat treatment conditions. There was also no 

difference in hardness for water-quenched specimens whether 

they were made by the casting method or by PM. The 

difference between hardness values of aged specimens 

fabricated by different processing methods was approximately 

6 percent.

The results indicate an inverse relationship between impact 

energy and hardness for this material, regardless of the 

processing method. This finding may enable the selection of a 

heat treatment process that optimizes hardness and impact 

resistance. The salient issue, however, is that the impact 

resistance of this material is relatively low due to its inherent 

brittleness. Table IV lists the impact energies of several 

engineered materials, including 60-NITINOL. Clearly, the 

impact energy of 60-NITINOL is low compared to metallic 

bearing materials (e.g., 440C and M50), but in the notched 

condition the impact energy of this material is significantly 

higher than that of many of the engineered ceramics that are 

now used in full ceramic bearings. In fact, the impact energy 

of 60-NITINOL is close to that of partially stabilized zirconia 

(ZrO2 – 3mol% Y2O3), though the impact energy of zirconia 

listed in Table IV was obtained from an unnotched specimen. 

Careful consideration of the property trade-offs for this 

material will be necessary for its successful implementation in 

aerospace components.
NASA/TM—2012-216029 17

TABLE III.—IMPACT ENERGY AND MICROINDENTATION HARDNESS RESULTS. 

[Data for conditions that were not available are listed N/A.]

Impact energy, 

J

Microindentation hardness, 

HV500

Heat treatment↓\

Designation→ Cast-A Cast-B PM-X PM-Y Cast-A Cast-B PM-X PM-Y 

Water-quenched 0.37±0.004 N/A 0.33±0.04 0.49±0.06 673.5±17.2 N/A 675.5±11.7 641.6±18.3

Aged 0.48±0.02 0.45±0.05 0.39±0.02 0.41±0.02 630.7±16.6 589.8±5.7 692.3±14.1 650.2±12.9

Annealed 1.00±0.03 N/A 0.89±0.03 1.18±0.04 288.6±13.8 N/A 365.7±17.8 318.5±7.4

TABLE IV.—IMPACT TEST RESULTS COMPARED TO 

OTHER ENGINEERED MATERIALS AND TO 

EXTRUDED AND HARDENED 60-NITINOL

Material Impact Energy, 

J

L2 tool steel (Ref. 17) 26

410 stainless steel (Ref. 17) 34

Extruded 60-NITINOL (Ref. 18)

Hardened 60-NITINOL (Ref. 18)

6.8 (notched)

2.7 (notched)

Cast 60-NITINOL:

Furnace-cooled

Water-quenched

Aged

1.0 (notched)

0.4 (notched)

0.5 (notched)

Partially stabilized ZrO2 (Ref. 19) 0.9 (unnotched)

Cold isostatically pressed and sintered 

Al2O3 (Ref. 20)

0.1 (unnotched)

Sintered SiC (Ref. 21) 0.1 (unnotched)

Sintered and HIPped Si3N4 (Ref. 22) 0.08 (unnotched)

0.01 (notched)

Conclusions

The impact energy and hardness of 60-NITINOL fabricated 

by casting and by powder metallurgy techniques and subjected 

to various heat treatments has been studied. Based on the 

results of this study, the following conclusions can be drawn:

1. The impact energy of 60-NITINOL after either the 

water-quenching or aging heat treatments is 

approximately half that of the material after furnace 

annealing.

2. The hardness of 60-NITINOL after either waterquenching or aging is approximately twice that of the 

material after furnace annealing.

3. There is an inverse relationship between impact 

resistance and hardness of this material, which suggests 

possible selection of heat treatments for optimum 

properties.

4. The impact energy of 60-NITINOL is intermediate to 

that of conventional metallic bearing materials and 

typical advanced ceramic bearing materials, being most 

comparable to partially stabilized zirconia. 

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