Erwin V. Zaretsky

Glenn Research Center, Cleveland, Ohio

Rolling Bearing Steels—A Technical 

and Historical Perspective

NASA/TM—2012-217445

April 2012
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Erwin V. Zaretsky

Glenn Research Center, Cleveland, Ohio

Rolling Bearing Steels—A Technical 

and Historical Perspective

NASA/TM—2012-217445

April 2012

National Aeronautics and

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Glenn Research Center

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NASA/TM—2012-217445 1

Rolling Bearing Steels—A Technical and Historical Perspective

Erwin V. Zaretsky*

National Aeronautics and Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Abstract

Starting about 1920 it becomes easier to track the growth of bearing materials technology. Until 

1955, with few exceptions, comparatively little progress was made in this area. AISI 52100 and some 

carburizing grades (AISI 4320, AISI 9310) were adequate for most applications. The catalyst to quantum 

advances in high-performance rolling-element bearing steels was the advent of the aircraft gas turbine 

engine. With improved bearing manufacturing and steel processing together with advanced lubrication 

technology, the potential improvements in bearing life can be as much as 80 times that attainable in the 

late 1950s or as much as 400 times that attainable in 1940. This paper summarizes the chemical, 

metallurgical and physical aspects of bearing steels and their effect on rolling bearing life and reliability.

The single most important variable that has significantly increased bearing life and reliability is vacuum 

processing of bearing steel. Differences between through hardened, case carburized and corrosion 

resistant steels are discussed. The interrelation of alloy elements and carbides and their effect on bearing 

life are presented. An equation relating bearing life, steel hardness and temperature is given. Life factors 

for various steels are suggested and discussed. A relation between compressive residual stress and bearing 

life is presented. The effects of retained austenite and grain size are discussed. 

Introduction

Jacob Rowe applied for a British patent in 1734 for a rolling-element bearing. In his patent he defined 

the advantages of such a bearing. Rowe stated that with the adoption of such bearings to ...“wheel 

carriages... one Horse now will do the labor of two ... and ... I will suppose that there will be occasion to 

employ only twenty thousand Horses ... instead of the 40,000 existing in the United Kingdom at an 

annual savings of 1,095,000 pounds per year.” B.W. Kelly (Ref. 1) in writing about Rowe said “that he 

was not able to find historical records that reported such sudden prosperity had occurred. As a result, he

concluded that that the cost for keeping a horse that was not working was as much as it was for one that 

was in fact working.” Things really have not changed in over 275 years.

The rolling bearing materials used in Jacob Rowe’s era would have been wood, bronze and iron. 

Modern rolling-element bearing steel and metallurgy do not begin until about 1856 with the disclosure of 

the Bessemer process. In this process air is blown through molten pig iron to produce a relatively high 

grade of steel. This was followed 10 years later by the invention of open-hearth melting, which further 

improved quality and made steel far more accessible to industry. Bicycle bearing manufacturers were 

quick to take advantage of this new material. However, because heat treatment of steel was still an art 

known only to a few, most rolling-element bearings were probably made of unhardened steel. In 1879, 

British patent 869 was issued to J. Harrington and H. Brent for a hardened steel bushing, or inner shaft, 

fitted with a groove for balls. About the same time, W. Hillman of Coventry, England, constructed a 

machine for cutting balls from steel wire (Refs. 2 and 3).

In 1900, according to Stribeck (Ref. 4), the use of carbon and chromium steels for bearings gradually 

increased during the last quarter of the 19th Century as the need for bearings capable of reliably 

supporting heavy loads increased. He reported that water-hardened steel gave higher elastic limits and 

greater capacity than oil-hardened steel. In a discussion of the Stribeck paper, Hess (Ref. 4) presented 

chemical analyses of four bearing steels then in use. Hess (Ref. 4) stated that these bearing steels would 

 *

Distinguished Research Associate
NASA/TM—2012-217445 2

“harden throughout and be uniformly hard and tough where durability and long life are wanted.” The 

chemistry of one of the French steels, no. 88, listed in the table closely matches that of American Iron and 

Steel Institute specification (AISI), AISI 52100. This steel was first specified about 1920 and is the most 

used bearing steel today (Refs. 2 and 3).

Starting about 1920 it becomes easier to track the growth of bearing materials technology. Until 1955, 

with few exceptions, comparatively little progress was made in this area. AISI 52100 and some 

carburizing grades (AISI 4320, AISI 9310) were adequate for most applications. Materials such as AISI 

440 were available in those cases where improved corrosion resistance was required (Ref. 3). In one of 

the classic textbooks on bearing analysis written by Shaw and Macks (Ref. 5) in 1949, the only rollingelement bearing steel discussed was AISI 52100. Even as recently as 1957, in another authoritative text

written by Wilcock and Booser (Ref. 6), the authors made only incidental note of the fact that AISI 52100 

is not useful over 177 °C (350 °F). And, according to Wilcock and Booser (Ref. 6), “For temperatures 

above 177 °C (350 °F) bearing manufacturers have made small lots of bearings of AISI M–l and AISI M–

10 tool steels. These steels retain their hardness to temperatures approaching 538 °C (1000 °F). Evidence 

available to date indicates that they operate satisfactorily, provided lubrication can be maintained.” 

As discussed by Bamberger (Ref. 2) of General Electric Company Engine Division, Cincinnati, Ohio,

the catalyst to quantum advances in all high-performance materials, including those steels used for 

bearings, was the advent of the aircraft gas turbine engine. The impact of the gas turbine engine on the 

growth of the aircraft industry after the Second World War created unprecedented needs for better 

materials and designs for rolling-element bearings. These needs included bearings for higher 

temperatures, higher speeds, and greater loads. The continuously increasing thrust-to-weight ratio for the 

aircraft jet engines required the use of smaller and lighter bearings. The reliability of these bearings 

became a major consideration because of system and mission complexities and because of the high costs 

involved (Refs. 2 and 3).

In order to assure long rolling bearing life and reliability for commercial, industrial and aerospace 

applications, materials, lubricants and design variables must be carefully considered and specified. The 

treatment of an alloy from the ore to the finished bearing can have a very significant effect on bearing 

performance, life and reliability. Experience has shown that different heats of the same material and 

process can produce life differences in the range of 2 to 1. It is therefore the objective of this paper to 

bring together and discuss from both a technical and historical perspective the chemical, metallurgical and

physical aspects of bearing steels and their effect on rolling bearing life and reliability.

Bearing Life

Figure 1 is a schematic of the contact profile of a ball on a bearing race. Figure 2(a) shows the surface 

(Hertz) stress distribution under the ball and the principle stresses at z, a critical location below the 

surface. Figure 2(b) shows the stress distribution below the surface. From these principal stresses, the 

shearing stresses can be calculated.

Three shearing stresses can be applied to bearing life analysis: the orthogonal shearing stress, τo, the 

octahedral shearing stress, τoct, and the maximum shearing stress, τmax. The von Mises stress which is not 

a shearing stress has been inappropriately used by some investigators as a substitute for the octahedral 

shearing stress, τoct. All of these shearing stresses are a function of the maximum Hertz stress where

τ = k1 Smax (1)

The proportionality constant k1 is a variable related to the specific shearing stress and the maximum Hertz 

stress is the maximum value the Hertz stress distribution shown in Figure 2(a). For ball bearings (point 

contact), k1 equal 0.25, 0.28, and 0.32 for the orthogonal shearing stress, τo, the octahedral shearing stress, 

τoct, and the maximum shearing stress, τmax, respectively. For roller bearings (line contact), k1 equal 0.25, 

0.29, and 0.30 for the orthogonal shearing stress, τo, the octahedral shearing stress, τoct, and the maximum 

shearing stress, τmax, respectively.
NASA/TM—2012-217445 3

Figure 1.—Schematic of contact profile of ball on raceway; a and 

b, semiwidths of major and minor axes of Hertzian contact 

area, respectively.

For the analysis reported herein, only the maximum shearing stress is considered. The maximum 

shearing stress is one-half the maximum difference between the principal stresses:

2 max

σz − σx τ = (2) 

Moyer and Zaretsky (Ref. 7) discuss in detail “failure modes related to bearing life. The ultimate 

failure mode limiting bearing life is classical rolling-element fatigue of either a bearing race or a rolling 

element. The failure manifests itself as a spall limited to the width of the running track and the depth of 

the maximum shearing stresses a distance z below the contact surface where

z = k2 b (3)

The proportionality constant k2 is a variable related to the specific shearing stress and b is the semi minor 

axis of the contact ellipse (see Fig. 2). For ball bearings (point contact), k2 equal 0.50, 0.76, and 0.76 for 

the orthogonal shearing stress, τo, the octahedral shearing stress, τoct, and the maximum shearing stress, 

τmax, respectively. For roller bearings, (line contact) k2 equal 0.50, 0.79, and 079 for the orthogonal 

shearing stress, τo, the octahedral shearing stress, τoct, and the maximum shearing stress, τmax, respectively. 

The region or zone of maximum shearing stresses can be defined as the stressed volume beneath the 

Hertzian contact ranging from 0.50b to 0.79b depth below the stressed surface.

Generally, the spall begins in the region of maximum shearing stresses and propagates into a crack 

network. Most bearings, however, fail for other reasons. Failures other than those caused by classical 

rolling-element fatigue are considered avoidable if the bearing is properly designed, handled, installed, 

and lubricated and is not overloaded (Ref. 7).

Rolling-element fatigue is extremely variable but is statistically predictable depending on the steel 

type, steel processing, heat treatment, bearing manufacturing and type, and operating conditions. Sadeghi, 

et al. (Ref. 8) provide an excellent review of this failure mode. 

Alley and Neu (Ref. 9) provide a recent attempt at modeling rolling-element fatigue. With improved 

bearing manufacturing and steel processing together with lubrication technology, the potential 

improvements in bearing life can be as much as 80 times that attainable in the late 1950s or as much as 

400 times that attainable in 1940 (Ref. 3).
NASA/TM—2012-217445 4

Figure 2.—Subsurface stress field under point contact. 

(a) Hertz stress distribution for ball on raceway 

showing principal stresses (σ, stress; τ, shear stress; 

and Smax, maximum Hertz stress) at depth z below 

surface. (b) Distribution of principal and shearing 

stress as function of depth z/b below surface. 

Based on the 1947 work by Lundberg and Palmgren (Refs. 10 and 11) who use the orthogonal 

shearing stress, τo, for their analysis, the life of a ball or roller bearing based on rolling-element fatigue 

can be expressed in its most simplistic form as follows:
NASA/TM—2012-217445 5

p

P

C L LF 



 





10 = (4)

Equation (4) is benchmarked to pre-1940 air melt AISI 52100 steel where LF = 1.

The L10 life, in millions of inner race revolutions, is the theoretical life that 90 percent of a bearing 

population should equal or exceed without failure at their operating load P. C is defined as the theoretical 

load that a bearing can carry for a life of 1 million inner-race revolutions with a 90-percent probability of 

survival. The load-life exponent is p. And, LF is a life factor dependent on the bearing steel and its 

processing (Ref. 11).

Lundberg and Palmgren (Ref. 10) derive the load-life exponent p to be 3 for ball bearings and 4 for 

roller bearings. However, in their 1952 paper (Ref. 12), Lundberg and Palmgren modified their value of 

the load-life exponent p for roller bearings from 4 to 10/3. Their rationale for doing so was that various 

roller bearing types had one contact that is line contact and other that is point contact. They state “... as a 

rule the contacts between the roller and the raceways transforms from a point to a line for some certain 

load so that the life exponent varies from 3 to 4 for differing loading intervals within the same bearing.”

The ANSI/ABMA (Ref. 13) and ISO (Ref. 14) Standards incorporate p = 10/3 for roller bearings. 

Computer codes for rolling-element bearings incorporate p = 4 for roller bearings and p = 3 for ball 

bearings (Ref. 11).

Bearing lives determined by using Equation (4) with the values of C given in bearing manufacturers’

catalogues are based on the “first evidence of fatigue.” This can be a tiny spall that may not significantly 

impair the function of the bearing. Thus, the actual useful life can be much longer. Society of Tribologists 

and Lubrication Engineers (STLE) life factors LF for various bearing steels are given in Table 1 (Ref. 3).

It can be reasonably assumed that these life factors are benchmarked to air melt AISI 52100 steel at a 

maximum Hertz (contact) stress of 1.723 GPa (250 ksi). Table 2 provides the designation and chemistry 

of these and other representative bearing steels (Ref. 15).

TABLE 1.—LIFE FACTORS FOR 

BEARING STEELS (REF. 3)

Material Life factor,

LF

Through-hardened steels

AISI 52100 3

AISI M–10 2

AISI M–50 2

AISI T–1 (18–4–1) 2

Halmo 2

AISI M–1 .6

AISI M–2 .6

Corrosion-resistant steels

AMS 5749 (BG–42) 2

AMS 5900 (CRB7) 2

AISI 440C .6

Case-carburized steels

AMS 6278 (VIM–VAR M50 NiL) 4

AISI 4620 3

AISI 8620 2

AISI 9310 2

CBS 600 2

Vasco X–2 2

CBS 1000 2

AISI 8720 1.5
NASA/TM—2012-217445 6

TABLE 2.—REPRESENTATIVE BEARING AND GEAR STEELS (REF. 15)

Material Alloying element, percent by weight (balance Fe)

Common

designation

Description Reference 

specifications

C P 

(max)

S 

(max)

Mn Si Cr V W Mo Co Nb Ni Other

50100 Cr alloy steel UNS G 50986; AISI

E 50100; AMS 6442 1.00 0.025 0.025 0.35 0.25 0.50 ----- ----- ----- ----- ----- ----- --------

51100 Cr alloy steel

UNS G 51986; AISI

E 51100; AMS 6440,

6444, 6447

1.00 .025 .025 .35 .25 1.00 ----- ----- ----- ----- ----- ----- --------

52100 Cr alloy steel

UNS G 52986; AISI E 

52100; AMS 6440,

6444, 6447

1.00 .025 .025 .35 .39 1.45 ----- ----- ----- ----- ----- ----- --------

MHT Al-modified bearing steel --------------------------- 1.03 .025 .025 .35 .35 1.50 ----- ----- ----- ----- ----- ----- 1.36 Al

Halmo Bearing steel --------------------------- .56 .003 .008 .36 1.12 4.84 0.53 ----- 5.18 ----- ----- ----- ----------

M–1 High-speed tool steel UNS T 11301; AISI 

M–1 .80 .030 .030 .30 .30 4.00 1.00 1.50 8.00 ----- ----- ----- ----------

M–2 High-speed tool steel UNS T 11302; AISI 

M–2 .83 .030 .030 .30 .30 3.85 1.90 6.15 5.00 ----- ----- ----- ----------

M–10 High-speed tool steel UNS T 11310; AISI 

M–10 .85 .030 .030 .25 .30 4.00 2.00 ----- 8.00 ----- ----- ----- ----------

M–42 High-speed tool steel UNS T 11342; AISI 

M–42 1.10 .012 .007 .15 .17 3.77 1.15 1.66 9.51 7.99 ----- ----- ----------

M–50 High-speed tool steel

UNS T 11350; AISI 

M–50; AMS 6490, 

6491

.80 .030 .030 .30 .25 4.00 1.00 ----- 4.25 ----- ----- ----- ----------

M50 NiLa Carburized steel AMS 6278 .13 .030 .030 .30 .25 4.00 1.20 ----- 4.25 ----- ----- 3.50 ----------

T–1(18–4–1) High-speed tool steel UNS T 12001; AISI 

T–1; AMS 5626 .70 .030 .030 .30 .25 4.00 1.00 18.0 ----- ----- ----- ----- ----------

T–15 High-speed tool steel UNS T 12015; AISI 

T–15 1.52 .030 .030 .26 .25 4.70 4.90 12.5 b

1.0 5.10 ----- ----- ----------

440C Hardenable Cr stainless steel

UNS S 44004; AISI

440C; AMS 5618,

5630, 5880, 7445

1.03 .018 .014 b1.0 .41 17.3 .14 ----- .75 ----- ----- ----- ----------

AMS 5749 Martensitic stainless steel UNS S 42700; AMS 

5749 1.15 .015 .010 .50 .30 14.5 1.20 ----- 4.00 ----- ----- ----- b

.35 Cu

Vasco matrix II Gear steel --------------------------- .53 .014 .013 .12 .21 4.13 1.08 1.40 4.80 7.81 ----- .10 ----------

CRB–7 Bearing steel --------------------------- 1.10 .016 .003 .43 .31 14.0 1.03 ----- 2.02 ----- .32 ----- ----------

AMS 5900 Martensitic stainless steel UNS S 42800; AMS 

5900 1.10 .015 .010 .40 .30 14.0 1.00 ----- 2.00 ----- ----- b

.35 .25 Nb

9310a Ni-Cr-Mo alloy steel

UNS G 93106; AISI 

9310; AMS 6260, 

6265, 6267

.10 .025 .025 .54 .28 1.18 ----- ----- .11 ----- ----- 3.15 ----------

CBS 600a Alloy steel UNS K 21940; AMS 

6255 .19 .010 .010 .61 1.05 1.50 ----- ----- .94 ----- ----- .18 .07 Al

CBS 1000a Alloy steel --------------------------- .14 .018 .019 .48 .43 1.12 ----- ----- 4.77 ----- ----- 2.94 ----------

Vasco X–2a Gear steel --------------------------- .14 .011 .011 .24 .94 4.76 .45 1.40 1.40 .03 ----- .10 ----------

8620a Ni-Cr-Mo alloy steel

UHS G 86200; AISI 

8620; AMS 6274, 

6276, 6277

.21 .035 .040 .80 .25 .50 ----- ----- .20 ----- ----- .55 ----------

EX–53a Gear steel --------------------------- .10 .009 .006 .37 .98 1.05 .12 2.13 3.30 ----- ----- ----- 2.07 Cu

3310a Alloy steel UNS G 33106; AISI 

3310 .11 .025 .040 .52 .22 1.58 ----- ----- ----- ----- ----- 3.50 ----------

4320a Ni-Cr-Mo alloy steel UNS G 43200; AISI 

4320 .20 .035 .040 .55 .25 .50 ----- ----- .25 ----- ----- 1.82 ----------

4620a Ni-Mo alloy steel UNS G 46200; AISI 

4620; AMS 6294 .20 .035 .040 .55 .25 ----- ----- ----- .25 ----- ----- 1.82 ----------

4720a Ni-Cr-Ni alloy steel UNS G 47200; AISI 

4720 .20 .035 .040 .55 .25 .45 ----- ----- .20 ----- ----- 1.05 ----------

Pyrowear 675 Carburized stainless steel --------------------------- .07 .005 .003 .65 .40 13.0 .60 ----- 1.80 5.40 ----- 2.60 ----------

a

Carburized grades. 

b

Maximum.
NASA/TM—2012-217445 7

Steel Chemistry

Through-Hardened Steels

In the 1950s through the 1960s the bearing industry assumed that materials with higher alloy content 

would have better hardness retention at elevated temperatures. It was reasoned that this would also result 

in higher ambient-temperature hardness as well as longer bearing life. Based on this assumption steel 

companies and research laboratories within the United States began to develop bearing steels with higher 

alloy content.

It is necessary to compare these steel and processing variables in rolling-element fatigue tests and/or 

actual bearing tests. Standard mechanical tests, such as tension and compression tests or rotating-beam 

tests, could not be correlated with rolling-element fatigue results (Ref. 16). Accordingly, a series of 

studies to verify the effect of increased alloying elements on rolling-element fatigue life was undertaken 

by the author and his colleagues at the NASA Lewis Research Center (now NASA Glenn Research 

Center), Cleveland, Ohio (Refs. 17 and 18).

Figure 3(a) summarizes the results of rolling-element fatigue tests conducted in the NASA Five-Ball 

Fatigue Tester (Ref. 17). Previous studies by others did not maintain the close control of operating and 

processing variables, such as material hardness, melting technique, and lubricant type and batch, required 

for a completely unbiased material comparison (Ref. 18). These tests comprised three groups each of 

eight through-hardened bearing steels. There were a total of 720 tests. All of the specimens for the 

specified steel came from the same heat of material and were manufactured and heat treated to the same 

hardness at the same time. All other variable were as carefully controlled. Contrary to expectation, 

rolling-element fatigue life decreases with increasing total content of alloying element in the steel. When 

present in high percentages these alloying elements appear to significantly decrease rolling element 

fatigue life.

Additional work, shown in Figure 3(b), was performed with 120-mm-bore, angular-contact ball 

bearings made from VAR AISI M–I, AISI M–42, AISI M–50, and WB–49 steels to verify the results in 

the NASA Five-Ball Fatigue Tester (Refs. 17 and 19). Bearings were tested at an outer-race temperature 

of 316 °C (600 °F). These four test series comprised a total of 120 bearings, 30 for each steel. The 

magnitudes of the differences seen in these bearing tests at 316 °C (600 °F) correlate well with the results 

of five-ball fatigue tests that are also shown in Figure 3(a). Using the AISI M–50 L10 life as a comparison, 

the AISI M–1 data from the five-ball fatigue tests and from the bearing tests agree remarkably well. WB–

49 in the bearing tests and AISI M–42 in the five-ball fatigue tests, both of which alloys contain relatively 

high percentages of cobalt and have similar microstructures, show reasonably good agreement (Ref. 20).

These results completely changed previously held assumptions regarding the effect of bearing steel 

alloying elements on rolling-element fatigue life. As a result, by the mid-1980s AISI M-50 steel became 

the steel of choice for most high-temperature bearing applications over 149 °C (300 °F). For bearing 

temperatures less than 149 °C (300 °F), AISI 52100 steel with the lowest alloying content, has a longer 

fatigue life and is probably the most widely used bearing steel throughout the world. These steels are 

usually heat treated to Rockwell C hardness at room temperature of not less than 60. At operating 

temperature, it is a general requirement that the operating hot hardness be greater than Rockwell C 58.

Carburizing-Grade Steels 

Bearings are required to tolerate substantial damage progression without catastrophic fracture during 

the interval between the onset of a problem and when routine maintenance identifies the need for repair 

(Refs. 21 and 22). Material toughness provides this capability. Fracture toughness is the material property 

that defines the stress required to initiate rapid fracture in the presence of a local defect (e.g., a fatigue 

spall). Initial defect size substantially affects fracture characteristics but is beyond the control of the

designer. Tensile stresses, either application induced or residual, are necessary for rapid fracture to occur. 

These are somewhat controllable by the designer, but advanced applications will require tolerance to 

increased stress (Ref. 3).
NASA/TM—2012-217445 8

Figure 3.—Rolling-element fatigue life as a function of 

total content of alloying elements tungsten, chromium, 

vanadium, molybdenum, and cobalt (Ref. 17).

Through-hardened materials, heat treated to Rockwell C 60 hardness, as is typical with bearing 

components, have limited fracture toughness. The KIC is usually less than 24 MPa m1/2 (22 ksi in.1/2), 

depending upon heat treatment (Ref. 23) Materials with fracture toughness this low have limited bulk 

tensile stress capability if rapid fracture is to be avoided. A conservatively safe limit is 172.4 MPa 

(25 ksi). Applications requiring higher toughness will have to be made from a carburizing-grade steel

(Ref. 3).

Carburizing-grade steels have reduced carbon content so that heat treatment normally results in 

moderate hardness and high toughness. High surface hardness, which is required for rolling-element 

bearing performance, is achieved by diffusing carbon into the surface, a process called carburizing, prior 

to heat treatment. Locally the steel is then a high-carbon alloy and is heat treatable to full hardness. The 

resulting structure has a surface layer with mechanical properties that are equivalent to those of traditional 

through-hardened bearing steels and a core that remains at low hardness, with corresponding high 

ductility and high fracture toughness. Surface-initiated defects (e.g., a spall) propagate cracks into the 

tough core before they reach critical size. The tough core prevents rapid and catastrophic fracture (Ref. 3).
NASA/TM—2012-217445 9

Fracture toughness of a material is inversely proportional to its carbon content and hardness. The 

carbon content also determines hardness. Fracture toughness can be improved without affecting hardness 

by adding nickel. When present in high-chromium, low-carbon steels, nickel causes the steel to become 

fully austenitic above 875 °C (1605 °F) where the steel is heat treated or carburized. Adding nickel also 

influences carbide size and distribution within the steel, which affects fatigue life. Recognizing this, 

Bamberger (Ref. 24) modified the chemistry of AISI M–50 steel by decreasing the amount of carbon and 

increasing the amount of nickel. He called this modified AISI M-50 steel M50 NiL (the “Ni” referring to 

increased nickel and the “L” to low carbon). The steel is also designated as Aerospace Material 

Specification (AMS) 6268.

M50 NiL, which is case carburized, has a core with a high fracture toughness KIC (over 60 MPa -m1/2; 

50 ksi-in.1/2) than through hardened AISI M-50 (29 MPa -m1/2; 20 ksi-in.1/2). The M50 NiL core hardness 

is Rockwell C 43 to 45. M50 NiL has finer carbides (compounds of carbon and various alloying 

elements) dispersed more evenly within its microstructure than standard AISI M–50. Compressive 

residual stresses in excess of 210 MPa (30 ksi) are induced in the zone of maximum resolved shear 

stresses during carburization of M50 NiL. These residual stresses combined with the fine carbide 

structure will increase its rolling-element fatigue life over that of conventional AISI M–50 (Ref. 25).

Many carburized gear steels are also used as bearing steels. These carburized steels are primarily used 

for tapered roller bearings or other bearings such as cylindrical roller bearings where tight interference fits 

are required between the bearing bore and shaft. A tight interference fit will induce large tensile (hoop) 

stresses in the bearing inner ring that can cause catastrophic fracture failure of the ring and the bearing.

As with AISI 52100 steel, for temperatures less than 149 °C (300 °F), AISI 9310 and AISI 8620 are 

usually the materials of choice. However, for bearing operating temperatures greater than 149 °C 

(300 °F), M50 NiL is the steel of choice.

Corrosion-Resistant Steels

Although not normally a functional requirement, corrosion resistance is highly desirable because of 

its potentially large effect on life-cycle cost. Alloy steels with high chromium content, greater than 

12 percent, are considered corrosion resistant. However, although the chromium forms a passive 

chromium oxide layer at the surface that provides substantial protection, it is not inert and these alloys 

will corrode in hostile environments (Ref. 3).

Available corrosion-resistant bearing alloys include AISI 440C and the high-temperature variations 

such as AISI 440C Mod, Aerospace Material Specification (AMS) 5749 (VIM–VAR BG-42), AMS 5900 

(VIM–VAR CRB7) and Pyrowear 675. AISI 440C is widely used in instrument bearings and in bearings 

for food-processing equipment. In addition, AISI 440C is the traditional alloy chosen for use in cryogenic 

rocket engine turbopumps such as those in the NASA Space Shuttle. AMS 5749 (VIM–VAR BG-42), 

AMS 5900 (VIM–VAR CRB7) and Pyrowear 675 are more recent developments (Ref. 3). For 

temperatures less than 149 °C (300 °F), AISI 440C is the corrosion resistant steel of choice. For 

temperatures greater than 149 °C (300 °F), AMS 5749(BG-42) is the steel of choice.

Steel Processing

In the early years of the bearing industry acid- and base-refractory, air-melting methods were used to 

process steel. Major advances in steel producing have occurred beginning in the 1950s by the introduction 

of vacuum-melting procedures. Vacuum processing reduces or eliminates the amount of nonmetallic 

inclusions, entrapped gases, and trace elements in structural alloys, resulting in substantially cleaner 

material. The two primary methods of vacuum processing are vacuum induction melting (VIM) and 

consumable-electrode vacuum melting (CEVM) also called vacuum arc remelting (VAR). In the early 

1970s these two methods were combined, whereby the vacuum induction primary melt is vacuum arc 

remelted. This method, called VIM–VAR, produces much cleaner steel than either VIM or CEVM 

individually (Ref. 26). These results are shown in Figure 4.
NASA/TM—2012-217445 10

Figure 4.—Rolling-element fatigue life of AISI M–50 steel in rolling-contact fatigue 

tester as a function of steel processing. Specimen diameter, 9.525 mm (3/8 in.); 

maximum Hertz stress, 4.8 GPa (700 ksi); speed, 12 500 rpm (Ref. 26).

TABLE 3.—LIFE FACTORS FOR MELTING PRACTICE (REF. 3)

Processing Life factor,

LF

Air melting (AM) 1

Vacuum processing (VP) or carbon vacuum degrassing (CVD) 1.5

Vacuum are remelting (VAR)a 3

Electroflux remelting (EFR)b 3

Vacuum are remelting-vacuum are remelting (VAR–VAR) 4.5

Vacuum induction melting-vacuum are remelting (VIM–VAR) 6 a

Also called consumable-electrode vacuum melting (CEVM). b

Also called electroslag remelting (ESR).

Although CEVM and VIM–VAR are the primary methods used today to produce materials such as 

AISI M–50, other vacuum processing methods have been developed, primarily aimed at improving AISI 

52100. The effect of melting practice on rolling-element fatigue life is shown by the Society of 

Tribologists and Lubrication Engineers (STLE) life factors (LF) in Table 3 (Ref. 3). The product of the 

life factors for bearing steel from Table 1 and melting practice from Table 3 is used as a single life factor 

LF in Equation (1) to determine the bearing L10 life 

Vacuum processing of bearing steel increases bearing life by eliminating hard oxide inclusions that 

act as stress raisers to initiate incipient failure. This results in an unforeseen secondary benefit. The 

fatigue life of the bearing steel becomes more sensitive to a reduction in stress. That is, as contact load or 

(Hertz) stress is decreased, bearing fatigue life is increased at a faster rate than with the air melted bearing 

steels.

As previously discussed, the load-life exponent p in Equation (1) is 3 for ball bearings and 4 for roller 

bearings based on pre 1940 air melt AISI 52100 steel. However, a reevaluation of the load-life relation

(Refs. 27 to 29) based on a summary of published data by R.J. Parker and E.V. Zaretsky (Ref. 30)
NASA/TM—2012-217445 11

suggests that for post-1960 vacuum processed bearing steels, the load life exponent p equals 4 and 5 for 

ball and roller bearings, respectively. This accounts for another significant improvement in rolling bearing 

life and reliability.

Heat Treatment

Steel Hardness

Hardness is an influential heat-treatment-induced variable. For most rolling bearing applications it is 

required that the Rockwell C hardness at operating temperature be 58 or higher. In general, the higher the 

hardness of the bearing steel at operating temperature, the longer the life. A relationship that 

approximates the effect of bearing material hardness on fatigue life has been developed (Refs. 20 and 31).

= exp{ [(RC) − 60]} LF m T (5)

where m is an exponent relating material hardness and life (typically m = 0.1) and (RC)T is the Rockwell 

C hardness at operating temperature. It was assumed for the purpose of this relationship, which was 

obtained for AISI 52100, that all components in the rolling-element bearing (i.e., the rolling elements and 

the races) are of the same hardness. It was further assumed that this equation can be extended to other 

bearing steels. A 3 point increase in hardness can result in a 35 percent increase in bearing life. With the 

exception of AISI 52100 and other low-tempering-temperature bearing steels, most bearing steels can be 

expected to maintain their room-temperature hardness after soaking at elevated temperatures. 

As was discussed for through-hardened steels, the bearing industry also assumed that materials with 

higher alloy content would have better hardness retention at elevated temperatures. A study to verify this 

assumption was undertaken at NASA Lewis Research Center (Refs. 32 to 34). Short-term, hot-hardness 

measurements were made for groups of through-hardened specimens of AISI 52100, M–1, M–50, 440C,

Halmo, WB–49, WD65, and Matrix II. Measurements were also made of specimens of Super Nitralloy 

(5Ni-2A1) and case-carburized AISI 8620, CBS 600, CBS 1000, and Vasco X–2. The results for the AISI 

52100 and the other through-hardened steels were normalized and are shown in Figure 5. These 

normalized data show that regardless of the initial hardness, the hot hardness of individual materials 

shows the same functional dependence. These results completely changed previously held assumptions

(Ref. 20). These data can be represented by a straight line having the form:

( ) ( ) α ∆ β RC T = RC RT − T (6)

where (RC)T is the Rockwell C hardness at operating temperature; (RC)RT is the Rockwell C hardness at 

room temperature; ΔT is the difference between operating temperature and room temperature, α is a 

material constant, and β is a material exponent. Values of α and β for various bearing steels are given in 

Table 4. 

To determine hardness effects at the bearing operating temperature, Equations (2) and (3) can be 

combined to obtain a life factor as follows: 

( {[( ) ] ( ) }) β LF = 1 RC RT − 60 − α TT −TRT exp 0. (7)

where TRT is 22 °C (70 °F). Equation (7) is benchmarked to a Rockwell C hardness equal 60 where 

LF = 1.
NASA/TM—2012-217445 12

Figure 5.—Summary of short-term, hot-hardness, 

through-hardened bearing steel data (Ref. 20).
NASA/TM—2012-217445 13

TABLE 4.—TEMPERATURE PROPORTIONALITY FACTORS α

AND EXPONENTS β FOR BEARINGS STEELS

[(RC)T = (RCRT – α ∆Tβ

); from Zaretsky (Ref. 20).]

Material Temperature range,

°C (°F)

α β

°C °F °C °F

AISI 8620 21 to 316 (70 to 600) 73×10–5 26×10–5 1.7 1.7

CBS 600 21 to 316 (70 to 600) .75×10–5 .18×10–5 2.4 2.4

Vasco X–2 21 to 538 (70 to 1000) 1.4×10–5 .38×10–5 2.2 2.2

CBS 1000 21 to 538 (70 to 1000) 93×10–5 38×10–5 1.5 1.5

CBS 1000M 21 to 538 (70 to 1000) 340×10–5 160×10–5 1.3 1.3

Super Nitralloy 21 to 327 (70 to 620) 1.3×10–5 .33×10–5 2.3 2.3

AISI 52100 21 to 260 (70 to 500) 92×10–5 34×10–5 1.6 1.6

AISI M–50 21 to 538 (70 to 1000) 133×10–5 54×10–5 1.4 1.4

AISI M–1

AISI M–2

AISI M–10

AISI M–42

AISI T–1 (18–4–1)

Halmo

WB–49

WD–65

Matrix II

AISI 440C

AMS 5749 (BG42)

AMS 6278 (M50 NiL)

Compressive Residual Stresses

From the late 1920s through to the 1960s, Almen (Ref. 35) and his colleagues (Refs. 36 to 38) at the 

General Motors Research Laboratories pioneered the study of residual stresses in rotating steel 

components that included rolling-element bearings. These residual stresses can either be tensile or 

compressive. They can be induced by producing microscopic and macroscopic deformations and by 

transformations in the microstructure of the steel. Residual stresses can also be induced by heat treating, 

rolling, shot peening, diamond burnishing, and severe grinding. Each of these methods (except heat 

treating) is a separate mechanical process that is performed after heat treating (Ref. 15).

They found that compressive residual stresses induced beneath the surface of ball bearing race 

grooves increase rolling-element fatigue life. According to Gentile and Martin (Ref. 37) ball bearing lives 

were doubled when metallurgically induced (“prenitrided”) compressive residual stresses were present in 

the inner races. Scott et al. (Ref. 38) found that compressive residual stresses induced by unidentified 

“mechanical processing” extend the fatigue life of ball bearings (Ref. 15).

Figure 6 shows representative residual stresses as a function of depth below the surface for three heattreated bearing steels. In general, most, if not all, carburized bearing steels have induced compressive 

residual stresses represented by those shown for AISI 9310 steel in Figure 6. These stresses are induced 

by the carburization process. 
NASA/TM—2012-217445 14

Figure 6.—Representative principal residual stress as a function of depth 

below surface for heat-treated AISI M–50, AISI 9310, and M50 NiL 

(AMS 6278) (Ref. 3).

In 1965 E.V. Zaretsky (Refs. 15 and 39) and his colleagues at the NASA Lewis (now Glenn) 

Research Center published an equation relating rolling-element fatigue life to these compressive residual 

stresses. The maximum shearing stress τmax for a given contact stress is decreased by the presence of a 

compressive residual stress σr. This results in the following life factor due to residual stresses alone:

c

r

LF 



 





τ − σ

τ = max 1/ 2

max (8)

where exponent c is typically 9.

For light to moderately loaded bearings a typical value of τmax is 414 MPa (60 ksi). For heavily loaded 

bearings a typical value of τmax is 724 MPa (105 ksi). From Figure 6, assume AISI 9310 as the bearing 

steel. The compressive residual stress, σr, is 200 MPa (29 ksi). From Equation (8), for a lightly loaded 

bearing, LF ≈ 12. For a heavily loaded bearing, LF ≈ 3.8. These life factors can be applied in Equation (1) 

together with the other life factors discussed. However, when bearing life results are analyzed 

independent of these residual stresses, the load-life exponent p appears to increase from their accepted 

values (Ref. 40).

Investigators have misinterpreted these results caused by the presence of residual stresses as a 

“fatigue limit” (Ref. 41). They have incorporated them into bearing life predictions and in some cases 

bearing manufacturer catalogues (Ref. 42). The concept of a fatigue limit has also been incorporated into 

an ISO standard (Ref. 43) for bearing life prediction for AISI 52100 steel where there are no residual 

stresses in the as heat treated steel (Ref. 42). This can result in bearing life over prediction and/or 

undersizing a bearing for a particular application (Refs. 41 and 42).

There are two problems associated with the use of a fatigue limit for bearing steels. The first problem 

is as discussed above, the form of the equation as expressed in the ISO standard (Ref. 43) may not reflect 
NASA/TM—2012-217445 15

a fatigue limit but the presence of a compressive residual stress. The second problem is that there are no 

data in the open literature that would justify the use of a fatigue limit for through-hardened bearing steels 

such as AISI 52100 and AISI M-50 (Refs. 41 and 42).

In 2007 Sakai (Refs. 44 and 45) presented stress life rotating bending fatigue data from six different 

laboratories in Japan for AISI 52100 steel. He also presented stress-life fatigue data for axial loading. The 

resultant lives were in excess of one billion (>109

) stress cycles at a maximum shearing stress τmax as low 

as 350 MPa (51 ksi) without an apparent fatigue limit.

In 2008 Tosha, et al., (Ref. 46) of Meiji University in Japan reported rotating bending stress-life 

fatigue tests for through hardened bearing steels having Rockwell C hardness above 58. The results of 

these tests at maximum shearing stresses τmax as low as 480 MPa (70 ksi) produced fatigue lives in excess 

of 100 million (>108

) stress cycles without the manifestation of a fatigue limit.

In 2009, in order to verify the results of Sakai (Refs. 44 and 45) and Tosha, et al. (Ref. 46), Shimizu, 

et al. (Ref. 47), also of Meiji University, published the results from six groups of AISI 52100 bearing 

steel specimens using four-alternating torsion fatigue tests rigs to determine whether a fatigue limit exists 

or not and to compare the resultant shear stress-life relation with that used for rolling-element bearing life 

prediction (Ref. 42). The results of these tests at maximum shearing stresses τmax as low as 500 MPa 

(76 ksi) produced fatigue lives in excess of 10 million (>107

) stress cycles without a fatigue limit.

Shimizu, et al. (Ref. 47). reported that the resultant fatigue life was inversely related to the shearing stress 

to the 10.34 power (Ref. 42).

Retained Austenite

In the early 1960s a major U.S. aircraft engine company had to discard unused rolling-element 

bearings made from AISI M-50 because their bore diameter had increased from that specified for the 

engine shaft diameter (in a personal communication with E.N. Bamberger, General Electric Company, 

February 1963). This expansion in bore size was attributed to the presence of large amounts of retained 

austenite in the microstructure of the steel. The retained austenite transformed to martensite and bainite on 

the shelf at room temperature. As a result, for most critical aerospace applications, the retained austenite 

is limited to 2 to 5 percent. However, for noncritical applications, higher amounts of retained austenite are

allowed or may, in some instances, be uncontrolled. Experience has suggested that lower values of 

retained austenite are preferable for reliable bearing operation (Ref. 15).

L.R. Waldmiller of Frost, Inc. (in a personal communication, December 1994) described unrun, 

carburized, 12.7-mm (0.5-in.) diameter AISI 1022 balls having 40 to 50 percent retained austenite. The 

balls lost a portion of the case material while at room temperature due to transformation of the retained 

austenite. The lost material had the appearance of a “skullcap” and the phenomenon was referred to as 

“capping.” He reported that the phenomenon also occurs during bearing operation. The same material 

with a lower amount of retained austenite did not experience capping (Ref. 15).

In general, for a given through-hardened material the amount of retained austenite increases with 

increasing material hardness. Experience has also shown that test rollers made from AISI 52100 of 

Rockwell C hardnesses greater than 63 will have sufficient austenite-to-martensite transformation during 

rolling contact to alter the surface waviness and cause early surface spalling (Ref. 15).

Johnston et al. (Ref. 48) studied the effect of the decomposition of retained austenite and the 

inducement of compressive residual stress as a result of bearing operation. What is unique for their data is 

that the magnitudes of the compressive residual stresses are directly proportional to the decomposition of 

retained austenite (Ref. 15).

Changes in microstructure (phase transformations) have been reported to occur in the same areas as 

the maximum induced residual stress (Refs. 36 and 49). Under some conditions of extremely high contact 

stresses, nonmicrostructural alteration was apparent after significant residual stresses had been induced in 

a few cycles (Ref. 49). Muro and Tsushima (Ref. 50) proposed that the induced residual stresses and the 

microstructural alterations are independent phenomena (Ref. 15).
NASA/TM—2012-217445 16

Research performed by Zhu et al. (Ref. 51) in 1985 on carburized rollers suggested that the structural 

change in the zone of maximum resolved shearing stresses observed by Jones (Ref. 52) in 1947 and later 

by Carter (Ref. 53) in 1960 as well as others is a manifestation of retained austenite transforming to 

martensite under cyclic Hertzian stress conditions. A combination of thermal and strain energy and time is 

believed to cause this change (Ref. 15).

Grain Size

It is generally accepted in the bearing industry that prior austenite grain size should be ASTM number 

8 or finer and that individual grains should not exceed ASTM number 5 (Ref. 54). The higher the ASTM 

number, the finer the grain size. The 1960 work of R.A. Baughman (Ref. 55) suggested that rollingelement fatigue life increases with finer grain sizes (Ref. 20). A recent analysis of grain size and 

orientation on rolling-element fatigue life was performed by N. Weinzapfel, et al. (Ref. 56).

Carbides

Residual carbides are those carbides that do not go completely into solution during austenitizing and 

are a function of the alloying elements and raw material processing. In contrast, hardening carbides 

precipitate upon aging at the tempering temperature. The carbides referred to in the following paragraphs 

are the residual carbides.

Carbide composition has been found to vary among steel producers. Heat treating steel ingots creates 

large, extremely hard metal carbides (MC), considered to be essentially a vanadium carbide, which can 

act as asperities in the bearing surface (Ref. 57). J.E. Bridge, et al. (Ref. 58) identified the primary 

carbides in AISI M–50 as MC and M2C. Pearson and Dickinson (Ref. 57) found that the M2C carbides 

contain a high percentage of molybdenum and that in a bearing ball under thin film elastohydrodynamic 

(EHD) lubrication conditions they can cause distress or peeling of the bearing race surfaces. The carbide 

“stick out” has been attributed in whole or in part to an excessive rate of grinding in the manufacture of

bearing balls made from AISI M-50 steel.

Parker et al. (Refs. 17, 59, and 60) have shown an interrelation among steel alloy content, median 

residual carbide size, number of residual carbide particles per unit area, percentage of residual carbide 

area in through hardened bearing steels, and rolling-element fatigue life. As the percent of alloying 

elements increase in a steel the number and size of the carbides increase (Refs. 61 and 62). Subsequent 

research by Parker and Bamberger (Ref. 63) for AMS 5749 steel further substantiated the negative effect 

of large carbide size and banded carbide distribution on rolling-element fatigue life.

Pearson and Dickinson (Ref. 57) verified the observations of Butterfield and T.R. McNelley

(Ref. 64), who reported voids of the order of 1 μm (40 μin.) adjacent to carbides of AISI M–50 steel. This 

work (Ref. 64) suggested that these voids form during bearing operation at the site of the carbide tip and 

can act as a nucleus for crack initiation in the subsurface zone of maximum shear stresses. The large 

carbides act as stress raisers to initiate an incipient crack that results in a rolling-element fatigue spall. The 

effect of carbides on rolling-element fatigue life is reflected in the life factors of Table 1 (Ref. 15).

In general, case carburized bearing steels, with the exception of M50 NiL, have a more course and 

larger carbide structure when compared to through hardened bearing steels such as AISI 52100 or AISI 

M-50. However, this disadvantage is more than offset by the compressive residual stresses induced into 

the case by the carburization process.

Summary

In order to assure long rolling-element bearing life and reliability for commercial, industrial and 

aerospace applications, materials, lubricants and design variables must be carefully considered and 

specified. The catalyst to quantum advances in high-performance rolling-element bearing steels was the 

advent of the aircraft gas turbine engine. The reliability of these bearings became a major consideration 
NASA/TM—2012-217445 17

because of system and mission complexities and because of the high costs involved. With improved 

bearing manufacturing and steel processing together with advanced lubrication technology, the potential 

improvements in bearing life can be as much as 80 times that attainable in the late 1950s or as much as 

400 times that attainable in 1940. The following summarizes the chemical, metallurgical and physical 

aspects of bearing steels and their effect on rolling bearing life and reliability:

(1) For temperatures less than 149 °C (300 °F) the bearing steels of choice are: through-hardened,

AISI 52100; case-carburized, AISI 8620 and AISI 9310; and corrosion-resistant, AISI 440C. For

temperatures greater than 149 °C (300 °F) the bearing steels of choice are: through-hardened, AISI M-50; 

case-carburized, M50 Nil; and corrosion-resistant, BG-42.

(2) Vacuum processing of bearing steel reduces or eliminates the amount of nonmetallic inclusions, 

entrapped gases, and trace elements in structural alloys, resulting in substantially cleaner material and

significantly longer bearing life.

(3) For a post-1960 vacuum processed bearing steels such as AISI 52100 and AISI M-50, the values 

for the load life-exponent p, where life is inversely proportional to load to the exponent p, increased from 

3 and 4 for ball and roller bearings, respectively, to 4 and 5.

(4) Minimum hardness for bearing steel at operating temperature should not be less than Rockwell 

C 58. Bearing life increases with increasing steel hardness at operating temperature. A 3 point increase in 

hardness can result in a 35 percent increase in bearing life. For M-Series bearing steels, the change in 

hardness with temperature is independent of alloy content.

(5) Bearing steels with high chromium content, greater than 12 percent, such as AISI 440 °C are 

considered corrosion resistant. Although the chromium forms a passive chromium oxide layer at the 

surface that provides substantial protection, it is not inert and these alloys will corrode in hostile 

environments.

(6) Compressive residual stresses induced or present from heat treatment beneath the surface of 

bearing steel components increase rolling-element fatigue life and can alter the Hertz stress-life relation. 

A compressive residual stress of 200 MPa (29 ksi) can increase bearing life for a lightly loaded bearing by 

a life factor, LF ≈ 12. For a heavily loaded bearing, LF ≈ 3.8. 

(7) For most critical aerospace applications, retained austenite is limited to 2 to 5 percent. However, 

for noncritical applications, higher amounts of retained austenite are allowed or may, in some instances, 

be uncontrolled. Experience has suggested that lower values of retained austenite are preferable for 

reliable bearing operation.

(8) Rolling-element fatigue life increases with finer grain sizes. Prior austenite grain size should be 

ASTM number 8 or finer and that individual grains should not exceed ASTM number 5. The higher the 

ASTM number, the smaller the grain size.

(9) There is an interrelation among steel alloy content, median residual carbide size, number of 

residual carbide particles per unit area, percentage of residual carbide area, and rolling-element fatigue 

life. The large carbides act as stress raisers to initiate an incipient crack that results in a rolling-element 

fatigue spall. As the percent of alloying elements increase in through-hardened bearing steel, the number 

and size of the carbides increase.

(10) Case-carburized bearing steels, with the exception of M50 NiL, have a more course and larger 

carbide structure when compared to through hardened bearing steels such as AISI 52100 or AISI M-50. 

However, this disadvantage is more than offset by the compressive residual stresses induced into the case 

by the carburization process.

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Reduction Gearbox System Fatigue Life and Reliability,” Tribology Trans., 50, 4, pp. 225–262.

30. Parker, R.J., and Zaretsky, E.V., (1972), “Reevaluation of the Stress-Life Relation in RollingElement Bearings,” NASA TN D–6745, National Aeronautics and Space Administration, 

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31. Zaretsky, E.V., Parker, R.J., and Anderson, W.J., (1965), “Effect of Component Differential 

Hardness on Rolling-Contact Fatigue and Load Capacity,” NASA TN D–2640, National Aeronautics 

and Space Administration, Washington, D.C.

32. Chevalier, J.L., Dietrich, M.W., and Zaretsky, E.V., (1972), “Short-Term Hot Hardness 

Characteristics of Rolling-Element Steels,” NASA TN D–6632, National Aeronautics and Space 

Administration, Washington, D.C.

33. Chevalier, J.L., Dietrich, M.W., and Zaretsky, E.V., (1973), “Hot Hardness Characteristics of 

Ausformed AISI M–50, Matrix II, WD–65, Modified AISI 440–C, and Super Nitralloy,” NASA TN 

D–7244, National Aeronautics and Space Administration, Washington, D.C.

34. Anderson, N.E. and Zaretsky, E.V., (1975), “Short-term Hot-Hardness Characteristics of Five CaseHardened Steels,” NASA TN D–8031, National Aeronautics and Space Administration, Washington, 

D.C.

35. Almen, J. O., “Effects of Residual Stress on Rolling Bodies,” (1962), Rolling Contact Phenomena, J. 

B. Bidwell, ed., Elsevier, New York, pp. 400–431. 

36. Gentile, A.J., Jordan, E.F., and Martin, A.D., (1965), “Phase Transformations in High-Carbon, HighHardness Steels Under Contact Loads,” AIME Transactions, Vol. 233, No. 6, pp. 1085–1093.

37. Gentile, A.J., and Martin, A.D., 1965, “The Effect of Prior Metallurgically Induced Compressive 

Residual Stress on the Metallurgical and Endurance Properties of Overload-Tested Ball Bearings,” 

ASME Paper 65–WA/CF–7.

38. Scott, R.L., Kepple, R.K., and Miller, M.H., 1962, “The Effect of Processing-Induced Near-Surface 

Residual Stress on Ball Bearing Fatigue,” Rolling Contact Phenomena, J.B. Bidwell, ed., Elsevier, 

pp. 301–316.

39. Zaretsky, E.V., Parker, R.J., Anderson, W.J., and Miller, S.T., (1965), “Effect on Component 

Differential Hardness on Residual Stress and Rolling–Contact Fatigue,” NASA TN D–3101, 

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

40. Oswald, F.B., Zaretsky, E.V. and Poplawski, J.V. (2011), “Relation Between Residual and Hoop 

Stresses and Rolling Bearing Fatigue Life,” Presented at the 2011 STLE Annual Meeting, May 15–

19, 2011, Atlanta, Georgia.
NASA/TM—2012-217445 20

41. Zaretsky, E.V., Poplawski, J.V., and Peters, S.M., (1996), “Comparison of Life Theories for RollingElement Bearings,” STLE Tribology Trans., 39, 2, pp. 237–247, Discussion by E. Ioannides and T.A. 

Harris, pp.247–248; Authors’ reply, pp. 501–503.

42. Zaretsky, E.V., (2010), “In Search of a Fatigue Limit: A Critique of ISO Standard 281:2007,” 

Tribology & Lubrication Technology, STLE, 66, 8, Aug. 2010, pp. 30–40.

43. ISO 281:2007, (2007), “Rolling Bearings—Dynamic Load Ratings and Rating Life,” International 

Organization for Standardization, Geneva.

44. Sakai, T., (2007), Review and Prospects for Current Studies on Very High Cycle Fatigue of Metallic 

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45. Sakai, T., Takeda, M., Shiozawa, K., Ochi, Y., Nakajima, M., Nakamura, T., and Oguma, Y., 

(2000), “Experimental Reconfirmation of Characteristic S-N Property for High Carbon Bearing Steel 

in Wide Life Region in Rotating Bending,” J. Soc. Mat. Sci. Japan, 49, 7, pp. 779–785 (in Japanese).

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49. Bush, J.J., Grube, W.L., and Robinson, G.H., (1962), “Microstructural and Residual Stress Changes 

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50. Muro, H., and Tsushima, N., (1970), “Microstructural Microhardness and Residual Stress Changes 

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NASA/TM—2012-217445 21

60. Parker, R.J., Zaretsky, E.V., and Dietrich, M.W., (1971), “Rolling-Element Fatigue Lives of AISI T–

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3, pp. 352–358.
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4. TITLE AND SUBTITLE

Rolling Bearing Steels--A Technical and Historical Perspective 

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Zaretsky, Erwin, V. 

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Published by the Institute of Materials, Minerals and Mining, in Materials Science and Technology, Vol. 28, No. 1, pp. 58-69, Great Britain, 

January 2012. 

14. ABSTRACT

Starting about 1920 it becomes easier to track the growth of bearing materials technology. Until 1955, with few exceptions, comparatively 

little progress was made in this area. AISI 52100 and some carburizing grades (AISI 4320, AISI 9310) were adequate for most applications. 

The catalyst to quantum advances in high-performance rolling-element bearing steels was the advent of the aircraft gas turbine engine. With 

improved bearing manufacturing and steel processing together with advanced lubrication technology, the potential improvements in bearing 

life can be as much as 80 times that attainable in the late 1950s or as much as 400 times that attainable in 1940. This paper summarizes the 

chemical, metallurgical and physical aspects of bearing steels and their effect on rolling bearing life and reliability. The single most 

important variable that has significantly increased bearing life and reliability is vacuum processing of bearing steel. Differences between 

through hardened, case carburized and corrosion resistant steels are discussed. The interrelation of alloy elements and carbides and their 

effect on bearing life are presented. An equation relating bearing life, steel hardness and temperature is given. Life factors for various steels 

are suggested and discussed. A relation between compressive residual stress and bearing life is presented. The effects of retained austenite 

and grain size are discussed. 

15. SUBJECT TERMS

Rolling bearing materials; Rolling bearing life and reliability; Bearing steel metallurgy; Residual stresses; Heat treatment 

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