White Etching Areas and Cracking

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Bearing crack from White Etching Crack ("WEC"). Photo by BRL.

A well-known problem for wind turbine gearbox bearings is axial cracking. Research on this significant problem is presented below.


Investigations of Bearing Failures Associated with White Etching Areas (WEAs) in Wind Turbine Gearboxes

Published here with permission from the authors, ROBERT ERRICHELLO, ROBERT BUDNY and RAINER ECKERT

A critical problem for wind turbine gearboxes is failure of rolling element bearings where axial cracks form on the inner rings. Metallurgical analyses show that the failure mode is associated with microstructural alterations manifested by white etching areas (WEAs) and white etching cracks (WECs). This article presents field experience from operating wind turbines that compares performance of through-hardened and carburized materials. It shows that through-hardened bearings develop WEA/WECs and fail with axial cracks, whereas carburized bearings do not. In another comparison of two rotor bearings with different carburized metallurgies, one bearing developed WEA/WECs and failed by macropitting, whereas the other bearing did not develop WEAs or WECs and did not fail. The field experience shows that a carburized bearing that has a core with low carbon content, high nickel content, greater compressive residual stresses, and a higher amount of retained austenite provides higher fracture resistance and makes carburized bearings more durable than through-hardened bearings in the wind turbine environment.


KEY WORDS
Fatigue Crack Propagation; Failure Analysis; Rolling Bearings; Cylindrical Roller Bearings; Tapered Roller Bearings; Carburizing; Through-Hardening; Rolling Contact Fatigue; Residual stress; Retained Austenite


INTRODUCTION
Microstructural alterations have been studied since 1947 (Warhapande, et al. (1)). From the late 1960s through the 1980s classic subsurface-initiated fatigue was investigated (Gentile, et al. (2); Martin, et al. (3); O’Brien and King (4); Swahn, et al. (5); Voskamp, et al. (6); Zwirlein and Schlicht (7); Voskamp (8); Schlicht, et al. (9)), including the well-documented slow structural breakdown (martensite decay), a progressive change in the steel matrix that occurs under moderately high Hertzian stresses. The decay creates dark etching areas followed by white etching bands (WBs; Warhapande, et al. (1)). Flat WBs first form at an angle of 30–40◦ to the surface (Warhapande, et al. (1); Zwirlein and Schlicht (7)) and steep WBs form later at an angle of 70–80◦ to the surface (Warhapande, et al. (1); Zwirlein and Schlicht (7)). Steep bands are located closer to the surface in the area of the greatest density of the flat WBs. Hertzian stress and the number of cycles are the controlling parameters for dark etching areas and WBs, and as the number of load cycles increase, the hardness drops in areas of structural changes, the hardness minima displace toward the surface, and the X-ray diffraction half-value breadth decreases (Warhapande, et al. (1); Martin, et al. (3); Zwirlein and Schlicht (7)). The Hertzian stress limit for the development of WBs is po ∼ 2,500 MPa (Warhapande, et al. (1); Zwirlein and Schlicht (7)).
WBs consist of nanosized ferrite cellular structures that, due to their fine-grained cellular structure, have a hardness that is 30–50% higher than the hardness of the surrounding matrix (Grabulov, et al. (10); Lund (11); M. H. Evans (12); R. D. Evans (13)).
When residual stresses are superimposed on Hertzian stresses, it is found that flatWBs are perpendicular to the effective tensile stress, and steep WBs coincide with the direction of the maximum shear stress (Zwirlein and Schlicht (7)). Figure 1 shows that it is possible to draw conclusions about the Hertzian stress that was effective at the time of formation ofWBs from the position, density, and direction of the WBs (Zwirlein and Schlicht (7)).


WEA MORPHOLOGY AND CHARACTERISTICS OF BUTTERFLIES
Stress concentrations occur around inhomogeneities (nonmetallic inclusions and large carbides) due to elastoplastic strain incompatibility between the inhomogeneities and the martensitic matrix (Bohmer (14)). Once the yield strength of the matrix is exceeded, a plastic strain is induced in a small domain surrounding the inhomogeneities. Under repeated Hertzian stress, dislocations shuttle back and forth and accumulate in this domain. This process causes localized changes in microstructure, such as white etching areas (WEAs), with the appearance of “butterfly wings.” Cracks nucleate in this domain once a critical density of dislocations is reached. The origin of a subsurface-initiated macropit most often occurs at the depth of the maximum orthogonal alternating shear stress, which is the basis of the Lundberg-Palmgren theory for rating bearing life.

Fig. 1 pressure vs load cycles.jpeg
Fig. 1—Correlation between Hertzian stress, load cycles, andmicrostructural alterations (from Voskamp, et al. (6)).



WEAs with the appearance of butterflies form adjacent to nonmetallic inclusions in planes 40–50◦ from the surface corresponding to planes of maximum unidirectional shear stresses (Warhapande, et al. (1); Martin, et al. (3); Hiraoka, et al. (15)). According to Lund (11) and Gegner and Nierlich (16), with AISI 52100 steel, butterflies form when the maximum shear stress τmax > 400 MPa.
No simple relationship between Hertzian stress and microcracks exists because the relationship is related to applied stress, local matrix conditions, and the composition, shape, size, and alignment of the inclusion with the stress field (Lund (11)). Heat treatment conditions do not significantly affect butterfly development. However, inclusion type, inclusion size, inclusion distribution, and forging reduction ratio do (Lund (11)). Welldispersed, small spherical inclusions (<13 μm) and high forging reduction ratios (≥7:1) maximize resistance to cracks associated with butterflies (Lund (11)).
Platelike (lenticular) carbides are thin carbide discs sandwiched between WEAs (Warhapande, et al. (1); Martin, et al. (3)). Lenticular carbides appear dark black on etching (Warhapande, et al. (1); Martin, et al. (3)). WEAs increase in number with longer cycling time. Their density varies in a manner consistent with the variation in shear stress, and WEAs form at a rapid rate around nonmetallic inclusions (Martin, et al. (3)). The maximum density of WEAs occurs at the depth of the maximum unidirectional shear stress (Martin, et al. (3)).
Butterflies are the result of an accumulation of localized plastic deformation at inhomogeneities such as nonmetallic inclusions and large carbides (Bohmer (14); Harada, et al. (17)).
Cracks tend to propagate along the boundary of WEAs. It is not resolved whether WEAs precede cracking (Harada, et al. (17)) or cracking precedes WEAs (M. H. Evans (12); Hiraoka, et al. (15); Vegter and Slycke (18)).
Many microcracks are nucleated at inclusion butterflies early in the fatigue life, but most do not grow beyond the WEAs unless the inclusions are located at Hertzian depths where the cyclic shear stress is high (Lund (11); M. H. Evans (12)).
Transmission electron microscope investigations (Grabulov, et al. (10); Harada, et al. (17)) show that WEAs consist of ultrafine nanocrystalline ferrite grains. Researchers (Grabulov, et al. (10); Lund (11)) concluded that WEAs result from recrystallization where new grains grow from a highly deformed steel matrix.
Though WEAs have been the subject of considerable study, there has been no clear link between WEAs and Hertzian fatigue. Grabulov, et al. (10) have found that the microstructure of classic subsurface-initiated macropitting is very different from the WEAs observed at inclusion butterflies. Therefore, WEAs are not an essential step in classic subsurface-initiated macropitting.
WEAs are not limited to wind turbine gearbox bearings and they occur in many other industries (M. H. Evans (12); Gegner and Nierlich (16); Holweger (19); Luyckx (20)). Furthermore, WEAs are not limited to any one gear manufacturer, bearing manufacturer, or wind turbine manufacturer (Gegner and Nierlich (16); Holweger (19); Luyckx (20); Uyama (21)). Currently, there is no calculation method that is recognized for predicting WEAs, and the root cause and significance of WEAs are not clearly understood (R. D. Evans (13); Gegner and Nierlich (16); Holweger (19); Luyckx (20); Uyama (21); Doll (22); Errichello (24), Oliver (25)).


MORPHOLOGY AND CHARACTERISTICS OF IRREGULAR WHITE ETCHING AREAS AND WHITE ETCHING CRACKS
A current critical problem for wind turbine gearboxes is the failure of rolling element bearings where axial cracks form on the inner rings.Metallurgical analyses show that the failure mode is associated with microstructural alterations manifested by irregular white etching areas (irWEAs) and white etching cracks (WECs). IrWEAs are branching crack networks that follow preaustenite grain boundaries and form crack networks with white etching borders. WECs can be straight-growing cracks that are parallel to the surface or branching crack networks (M. H. Evans (12)); that is, irWEAs. Both are associated with axial cracks in wind turbine bearings. There are several hypotheses for the root cause of irWEAs, WECs, and axial cracks, including impact loads (R. D. Evans (13); Holweger (19); Luyckx (20); Doll (22)), sliding (Gegner and Nierlich (16); Holweger (19); Doll (22)), hydrogen embrittlement (Gegner and Nierlich (16); Vegter and Slycke (18); Uyama (21)), electrostatic discharge (Holweger (19); Uyama (21)), corrosion fatigue (Gegner and Nierlich (16)), and adiabatic shear (Luyckx (20)). However, none of the hypotheses have been proven, and it is currently an active field of research (R. D. Evans (13);Gegner and Nierlich (16);Holweger (19); Luyckx (20); Uyama (21); Doll (22)).
It is known that WBs, WECs, and irWEAs share similar microstructural morphologies. They are nanosized ferrite cellular structures that result from recrystallization of new grains that grow from the highly deformed steel matrix. Therefore, WBs, WECs, and irWEAs are different development stages of the same phenomenon (Lund (11)).
How irWEAs and WECs develop and progress is not yet understood. Both have been characterized as brittle fracture modes that generate cleavage fractures. It might be a single-step or a multiple-step process that generates cleavage cracks andWBs (R. D. Evans (13); Gegner and Nierlich (16); Holweger (19); Luyckx (20); Uyama (21); Doll (22)).
It is not understood how a chemical conversion coating such as black oxide helps to prevent irWEAs and axial cracks. Itmight reduce tractional stresses, damp vibrations, or prevent diffusion of hydrogen. On the other hand, the temperatures used to treat the componentsmight beneficially alter the bearingmetallurgy (Gegner and Nierlich (16); Holweger (19); Luyckx (20); Doll (22)).
We usually do not find moisture corrosion, severe wear, or electric discharge damage in wind turbine gearboxes. Therefore, hydrogen absorption due to water in oil can be excluded and hydrogen generation due to sliding or electric discharge and diffusion into the bearing seems unlikely. Furthermore, although the authors have seen butterflies in wind turbine bearings, there is little evidence to support butterfly cracks propagating into irWEAs or WEC networks.


DESCRIPTION OF WIND TURBINE BEARING FAILURES
In the following sections, we present actual field experience from operating wind turbines that compares the performance of rolling element bearings manufactured from through-hardened and carburized materials. The wind turbines are utility scale and have been operating for up to 6 years. There are over 500 turbines of this type currently in operation.


Through-Hardened versus Carburized Intermediate Bearings
The wind turbine utilizes an NJ 2334 cylindrical roller bearing at each end of the four intermediate (INT) shafts in the gearbox. The bearings are manufactured by two different manufacturers, designated here as bearings INT-A and INT-B. The INT-B bearing is through-hardened and the INT-A bearing is carburized. The failure rate of the INT-B through-hardened bearings is 16% with a mean time to failure of 27,200 h (1.4 × 108 cycles). So far, there has been only one failure of an INT-A carburized bearing, and the failure is believed to be a secondary failure that occurred due to the presence of a surface defect of unknown origin. Table 1 compares data for the INT bearings and Table 2 compares chemistry for the INT bearings.


Figures 2a and 2b show a failed INT-B bearing inner ring (IR) that had operated for 18,000 h (9.3 × 107 cycles) when the failure was discovered. There are numerous axial cracks concentrated toward the flange end of the IR. Figure 3 is a micrograph of a circumferential metallurgical section through an axial crack. The section has been nital etched to display the irWEAs associated with the crack. Figure 4 is a scanning electron microscope image of an axial crack that was opened in the laboratory to expose the fracture surface. The crack morphology is typical of axial cracks found on through-hardened bearings from wind turbines (R. D. Evans (13); Gegner and Nierlich (16); Holweger (19); Luyckx (20); Uyama (21); Doll (22)). The origin of the fracture is believed to be at the center of the smooth circular “lens” (Gegner and Nierlich (16)).

TABLE 1—DATA FOR INTERMEDIATE BEARINGS
Parameter INT-A INT-B
Bore, d (mm) 170 170
Outside diameter, D (mm) 360 360
Width, T (mm) 120 120
Dynamic capacity, C (kN) 1,840 1,660
Static capacity, Co (kN) 2,110 2,040
Fatigue limit load, Pu (kN) 332 204
Number of rollers, z 14 14
Roller pitch diameter, Dpw (mm) 268 266
Roller mean diameter, Dw (mm) 52 50
Roller total length, lw (mm) 85 85
Roller effective length, lweff (mm) 81 81
Roller crown type Circular End reliefs
Roller crown magnitude (mm) 0.030 0.036
Nominal contact angle, αo (◦) 0 0
Heat treatment Carburized Through-hardened
IR case depth, Eht (mm) 3.5 N/A
Cage type Brass Brass
Hertzian stress (MPa) 1,583 1,756
Hertzian stress max (MPa) 1,869 2,056
Shaft speed (rpm) 86.12 86.12
DIN ISO 281-4 life (h) 212,200 178,500


Residual stresses and retained austenite were measured on new, unused intermediate bearings. Figure 5 shows residual stresses on INT bearings determined by X-ray diffraction (XRD) per ASTM E915. The residual stress for the INT-B throughhardened bearing is compressive up to −700 MPa at the surface, decreases to zero at a depth of 12 μm, and is tensile (ranging from 35 to 100 MPa) at depths greater than 13 μm. The residual stress for the INT-A carburized bearing is entirely compressive up to −1,000 MPa at the surface and greater than −400 MPa to a depth of 500 μm.

TABLE 2—CHEMICAL COMPOSITION FOR INTERMEDIATE BEARINGS
Element INT-A IR (% wt.) INT-B IR (% wt.)
Carbon, C 0.227 1.02
Silicon, Si 0.288 0.270
Manganese, Mn 0.774 0.280
Chromium, Cr 0.660 1.69
Nickel, Ni 1.67 0.130
Molybdenum, Mo 0.226 0.210
Sulfur, S 0.017 0.013
Phosphorus, P 0.010 0.015
Copper, Cu 0.190 0.260
Aluminum, Al 0.045 Not reported
Cobalt, Co 0.019 Not reported
Columbium, Cb 0.011 Not reported
Titanium, Ti 0.005 Not reported
Boron, B <0.001 Not reported
Vanadium, V 0.006 Not reported
Tin, Sn 0.005 Not reported


Figure 6 shows retained austenite on INT bearings determined by XRD per ASTM E975. The retained austenite for the INT-B through-hardened bearing is less than 1%. The retained austenite for the INT-A carburized bearing ranges from 23 to 31%.


Fig. 2a - roller path.jpg


Fig. 2b - axial cracks.jpg
Fig. 2—(a) Axial cracks on intermediate bearing INT-B IR. (b) Axial cracks on intermediate bearing INT-B IR (color figure available online).


Fig. 3 - propagation.jpg
Fig. 3—Irregular white etching areas on axial crack (color figure available online).


Fig. 4 - fatigue progression.jpg
Fig. 4—Opened axial crack on intermediate bearing INT-B IR (color figure available online).


Fig. 5 - residual stress.jpg
Fig. 5—Residual stresses for intermediate bearings.


Carburized Rotor Bearings
The wind turbine utilizes a tapered roller bearing at each end of the rotor shaft. The bearings aremanufactured by two different manufacturers, designated here as bearings ROT-C and ROT-D, and both are carburized. The failure rate of the ROT-D bearings is 17% with a mean time to failure of 26,690 h (2.2 × 107 cycles). So far, there have been no failures of a ROT-C bearing.
Table 3 compares data for the rotor bearings and Table 4 compares chemistry for the rotor bearings.
Figure 7 shows a failed ROT-D rotor bearing IR that had operated for 22,000 h (1.8 × 107 cycles) when the failure was discovered. The entire circumference of the IR is covered with severe macropitting.

Fig. 6 - retained austenite.jpg
Fig. 6—Retained austenite for intermediate bearings.



Figure 8a is a micrograph of a circumferential metallurgical section through a macropit. The section has been nital etched to display the irWEAs associated with the macropits. Figure 8b is at higher magnification and shows cracks within the irWEAs. The crack morphology is typical of macropitting found on carburized bearings from wind turbines (M. H. Evans (12); R. D. Evans (13); Holweger (19)).
Residual stresses and retained austenite were measured on new, unused rotor bearings. Figure 9 shows residual stresses on rotor bearings determined by XRD per ASTM E915. The residual stresses for the ROT-C and ROT-D bearings are entirely compressive to a depth of 500 μm. For depths below 20 μm, the residual stress for the ROT-D bearing is −250 MPa near the surface and it fades to 0 MPa at a depth of 500 μm, whereas the residual stress for the ROT-C bearing increases from −250 MPa near the surface to a maximum of −350 MPa at a depth of 300 μm and fades to −100 MPa at a depth of 500 μm.

TABLE 3—DATA FOR ROTOR BEARINGS
Parameter ROT-C ROT-D
Bore, d (mm) 749.3 749.3
Outside diameter, D (mm) 990.6 990.6
Width, T (mm) 159.5 159.5
Dynamic capacity, C (kN) 4,460 5,330
Static capacity, Co (kN) 13,200 13,110
Fatigue limit load, Pu (kN) 1,477 1,467
Number of rollers, z 43 43
Roller pitch diameter, Dpw (mm) 868.2 869.95
Roller mean diameter, Dw (mm) 58.5 60.15
Roller total length, lw (mm) 118.1 118.1
Roller effective length, lweff (mm) 114.1 107.1
Roller crown type Circular End reliefs
Roller crown magnitude (mm) 0.030 0.053
Nominal contact angle, αo (◦) 12.5 12.5
Axial preload, Gao (mm) 0.2 0.3
Heat treatment Carburized Carburized
IR case depth, Eht (mm) 3.5 3.5
Cage type Stamped Pinned
Hertzian stress (MPa) 1,634 1,798
Hertzian stress max (MPa) 1,886 2,075
Shaft speed (rpm) 13.66 13.66
DIN ISO 281-4 life (h) 146,000 121,300


TABLE 4—CHEMICAL COMPOSITION FOR ROTOR BEARINGS
Element ROT-C IR (% wt.) ROT-D IR (% wt.)
Carbon, C 0.126 0.199
Silicon, Si 0.217 0.265
Manganese, Mn 0.405 0.453
Chromium, Cr 1.29 1.32
Nickel, Ni 3.30 3.44
Molybdenum, Mo 0.099 0.190
Sulfur, S 0.016 0.003
Phosphorus, P 0.013 0.009
Copper, Cu 0.218 0.130
Aluminum, Al 0.030 0.029
Cobalt, Co 0.017 0.018
Columbium, Cb 0.013 <0.001
Titanium, Ti 0.005 <0.001
Boron, B <0.001 <0.001
Vanadium, V 0.008 <0.001
Tin, Sn 0.023 0.018



Figure 10 shows retained austenite on rotor bearings determined by XRD per ASTM E975. The retained austenite for the ROT-D bearing ranges from 12 to 17%. The retained austenite for the ROT-C bearing ranges from 20 to 26%.
Two unfailed rotor bearings (a ROT-C and a ROT-D) were removed from service after 25,000 h and metallurgically examined. The ROT-C bearing had butterflies at depths ranging from 100 to 400 μm but no irWEAs. The ROT-D bearing had numerous irWEAs at depths ranging from 400 to 600 μm as shown in Figure 11.


Fig. 7 - macropitting.jpg
Fig. 7—Macropitting on rotor bearing ROT-D IR (color figure available online).


Fig. 8a.jpg


Fig. 8b.jpg
Fig. 8—(a) Macropits and irWEAs on rotor bearing ROT-D IR. (b)Macropits and irWEAs on rotor bearing ROT-D IR (color figure available online).


DISCUSSION
The wind turbine gearbox failures have shown that throughhardened bearings fail by axial cracks, whereas carburized bearings fail by macropitting. This is consistent with the findings of other investigators (Gegner and Nierlich (16); Holweger (19); Luyckx (20); Doll (22)).

Fig. 9 - residual strss depth.jpg
Fig. 9—Residual stresses for rotor bearings.


Fig. 10 - retained austenite depth.jpg
Fig. 10—Retained austenite for rotor bearings.


Through-Hardened versus CarburizedIntermediate Bearings
Table 1 shows that the INT-A and INT-B bearings have similar geometries. However, Table 2 shows that the INT-B throughhardened bearing and INT-A carburized bearing have very different chemistries. Furthermore, Figs. 5 and 6 show that the INTB through-hardened bearings have tensile residual stresses and very little retained austenite, whereas the INT-A carburized bearings have significantly higher compressive residual stresses and greater amounts of retained austenite.
To compare the properties of the bearings to other industries, the evolution of bearing materials for gas turbines has shown that a carburized bearing with a core having a low carbon, high nickel content such as M50NiL (Forster, et al. (23)) has relatively high fracture resistance compared to an M50 through-hardened bearing. Furthermore, Forster, et al. (23) found WEAs in AISI 52100 and AISI M50 bearings but no WEAs in M50NiL bearings.


Fig. 11 - white area crack.jpg
Fig. 11—irWEAs on rotor bearing ROT-D IR (color figure available online).



Figure 4 shows that the origin of the brittle fracture lens is about 300 μm below the surface. Figure 5 shows that the residual stress at this depth is about +85 MPa tensile for the INT-B through-hardened bearing and about −450 MPa compressive for the INT-A carburized bearing. We believe that the compressive residual stress increases fracture resistance and is the principal reason the INT-A carburized bearing is immune to axial cracks.
In summary, the intermediate bearings have similar geometry but very different chemistry, residual stresses, and retained austenite. Therefore, it is believed that the better performance of the INT-A carburized bearing is due to different chemistry with a core having a low carbon, high nickel content; greater compressive residual stress; and a higher amount of retained austenite. All of these properties provide higher fracture resistance and make carburized bearings more durable in the wind turbine environment than through-hardened bearings.

Carburized Rotor Bearings
Tables 3 and 4 show that the ROT-C and ROT-D rotor bearings have similar geometries and chemistries. However, Figure 10 shows that the ROT-C bearings have a significantly greater amount of retained austenite at all depths. Furthermore, metallurgical analyses of unfailed bearings have shown that no irWEAs form in the ROT-C bearing, whereas irWEAs form in the ROTD bearing at depths ranging from 400 to 600 μm, as shown in Fig. 11. Figure 9 shows that the residual stresses for the ROTC and ROT-D bearings are very similar at the depths of the ir- WEAs. In summary, the carburized rotor bearings ROT-C and ROT-D have similar macrogeometries, chemistries, and residual compressive stresses but different levels of retained austenite. There are also differences in the design roller profiles between the two bearings, and analysis of the profiles predicts higher peak stress for the ROT-D profile under many operating conditions. Furthermore, coordinate-measuring machine (CMM) measurements made on actual components suggest that the as-built roller geometry of ROT-D bearing results in even higher peak stress under many operating conditions than do the nominal profiles. Therefore, it is believed that the better performance of the ROTC bearing is due to its higher amount of retained austenite, superior roller profile design, and smaller variation in manufacturing.

Summary of Through-Hardened Versus Carburized Bearing Performance
The authors’ experience shows that through-hardened bearings display irWEAs on axial cracks that propagate radially through the bearing IRsection as shown in Fig. 3, whereas carburized bearings display irWEAs on crack networks that occur over large subsurface areas at depths ranging from near the surface to the depth of the maximum shear stress as shown in Fig. 11. Eventually, the cracks reach the surface where they form macropits. This is consistent with the findings of other investigators (Gegner and Nierlich (16); Luyckx (20); Doll (22)). The failure mode has been called white structure flaking by Uyama (21).
The authors’ experience, and that of others (Kiraoka, et al. (15); Luyckx (20); Doll (22)), has shown that carburized bearings with a proper microstructure can be immune to the axial crack failure mode and are more durable in a wind turbine environment than through-hardened bearings. Furthermore, the authors’ results show that if the carburized microstructure has at least 20% retained austenite, irWEAs do not form and premature macropitting is avoided.

CONCLUSIONS
The following conclusions are drawn from field experience from operating wind turbines that compares the performance of through-hardened and carburized intermediate bearings and two rotor bearings with different carburized metallurgies. The conclusions are intended to apply to wind turbine gearbox bearings and they may or may not apply to other applications.
1. Through-hardened bearings fail by axial cracks, whereas carburized bearings fail by macropitting.
2. Through-hardened bearings display irWEAs on axial cracks that propagate radially through the bearing IR section.
3. Carburized bearings are more durable in the wind turbine environment than through-hardened bearings and might be immune to irWEAs and the axial crack failure mode if they have at least 20% retained austenite.
4. Carburized bearings with less than 20% retained austenite display irWEAs on crack networks that occur over large subsurface areas at depths ranging from the near surface to the depth of the maximum shear stress. When the cracks reach the surface, they form macropits.
5. Carburized bearings with at least 20% retained austenite might be immune to irWEAs and avoid premature macropitting.

ACKNOWLEDGEMENT The authors express their deepest appreciation to Tibor Tallian for his many helpful suggestions.

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(21) Uyama, H. (2011), “The Mechanism of White Structure Flaking in Rolling Bearings,” Available at: http://www.nrel.gov/wind/pdfs/2011 wind turbine tribology seminar.pdf
(22) Doll, G. L. (2011), “Tribological Challenges in Wind Turbine Technology,” Available at: http://www.nrel.gov/wind/pdfs/2011 wind turbine tribology seminar.pdf
(23) Forster, N., Rosado, L., Ogden, W. P., and Trivedi, H. (2009), “Rolling Contact Fatigue Life and Spall Propagation Characteristics of AISI M50, M50NiL, and AISI 52100, Part III, Metallurgical Examination,” Tribology Transactions, 53(1), pp 52–59.
(24) Errichello, R. (2011), “Microstructural Alterations in Hertzian Fatigue,” Available at: http://www.nrel.gov/wind/pdfs/2011 wind turbine tribology seminar.pdf
(25) Olver, A. (2011), “Microstructural Alteration in Rolling Contact,” Available at: http://www.nrel.gov/wind/pdfs/2011 wind turbine tribology seminar.pdf


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