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Rolling-sliding behavior of rail steels
159
Wear, 162-164 (1993) 159-172
Rolling-sliding M. Sate”,
behavior of rail steels
P. M. Anderson
and D. A. Rigney
Materials Science and Engineering, The Ohio State Universi& 116 West 19th Avenue, Columbus, OH 43210-1179 USA
Abstract A twin disk-rolling-sliding contact machine has been used to investigate the deformation, wear and cracking that occur during rolling-sliding of high-carbon rail and wheel steels. A martensitic wheel steel with initial hardness of 360 HV contacted pearlitic or martensitic rail steels of different initial hardness, 280-360 HV. Other variables included test time, test environment and slip ratio (0%, 5%, 10% and 26%). Tests were run in air, but with different lubrication conditions, e.g. dry, with water, water with inhibitor, silicone oil, and various combinations. Wear was measured by weight loss. Wear surfaces, sample cross-sections and wear debris were
observed by optical and scanning electron microscopy. Principal conclusions are as follows. The thickness of the highly deformed layer was very small for pure rolling compared with cases involving slip. Wear debris was generated from near-surface material in which the cementite phase had broken into small particles. Cracks initiated at the surface and propagated along lines of earlier plastic flow, rather than normal to the local direction of maximum tensile stress. The results of rolling-sliding in water or other liquids after dry rolling-sliding indicate that extensive deformation from sliding during the dry stage prepares the material for extensive cracking during subsequent operation with a lubricant.
1. Introduction Railroads provide energy-efficient systems for transport of both freight and passengers. Trends to increase axle load for the former and train speed for the latter require renewed attention to wear and other kinds of damage produced during operation of rail systems [l]. Steele and Stone [2] and Steele [3] have reviewed the many kinds of defects observed in rail systems. Some of these involve obvious plastic deformation. For others, the effects of the environment are clearly important, but the detailed mechanisms are not well understood. Part of the difficulty is related to the many variables which are involved. These include load, speed, slip ratio, materials and environmental conditions. It is convenient to think of wear processes as those which occur fairly steadily over extended periods of time. Material is typically lost by removal of small debris particles. In contrast, some damage to rails occurs more dramatically, with rather sudden removal of larger pieces of material. Such damage is commonly called rolling contact fatigue. This term may not be adequate to cover the full range of phenomena reported, but it is well established among those who work on tribological problems of railroads. It should be noted that results reported *Present address: Yawata R & D Laboratory, Nippon Corp., l-l Tobihata, Tobata, Kitakyushu 804, Japan.
Steel
in the literature show that wear and rolling contact fatigue processes can interact [2], so their separation is somewhat artificial. Previous work suggests that pitting or rolling contact fatigue occurs when unlubricated testing conditions are followed by continued testing with lubrication. Way [4] has reported that pitting cracks appeared soon after oil was added to a bearing system which had first been run without lubricant. He concluded that oil was necessary for pitting, but “some form of destruction leading to the formation of pitting cracks takes place independently of the oil.” Sugino and co-workers [5, 61 found no rolling contact fatigue for either dry or lubricated tests of rail steels, but rolling contact fatigue occurred readily when dry conditions were followed by operation with water or grease. These observations are important, because rails in use are exposed to alternating periods of dry and wet conditions. This research program was designed to provide new information on the relationships among wear, plastic deformation and fracture and the effects of hardness, microstructure, slip ratio and lubrication. Initial tests were run under dry or lubricated conditions. Tests with various combinations of unlubricated and lubricated conditions were also run. These included unlubricated tests with slip, followed by continued testing with lubrication (water, water with inhibitor, silicone oil).
Elsevier
Sequoia
M. &to
160
et al. i Rolling-siid~ng of rail steels
2. Experimental procedure A twin-disk rolling contact machine was used for all tests. The wheel specimen was mounted on the lower shaft, which was driven directly by the motor, as shown in Fig. 1 [7]. The rail specimen was mounted on the upper shaft, which was driven by a simple gear system. Both gears had the same number of teeth, so the rotation speeds of the two shafts were the same. To obtain different slip ratios, the circumferential speeds of the rail and wheel specimen disks were made different 1 by changing their relative diameters. A simple definition of the slip ratio is provided by the expression 2(L), - D,)/(D, +D,), where I),., and D, are the diameters of the wheel and rail respectively. The following values were used for this nominal slip ratio: SR(nom) =O%, 5%, 10% and 26%. To focus attention on the loading conditions for the rail specimen, another expression for slip ratio is also used in this paper. It is equal to the ratio of distance slid to distance rolled at the rail specimen surface, defined as SR(rai1) = (D, -D,)/D, = 0%, 5.1%, 10.5% and 29.9%. In the discussion to follow, nominal and rail slip ratios are clearly distinguished from each other by the numbers cited, e.g. 5% compared with 5.1%. The normal load was supplied by the dead-weight of the upper shaft and holder (314 N) and compressed air which pressed a loading bar against the upper shaft. Two loads were used for these tests, 750 N (maximum elastic contact stress, 525 MPa) and 500 N (430 MPa). Rail and wheel specimens were cut from the base portion of a Japanese Industrial Standard 60 kg (60 kg m-l) rail and a 50 kg “N” (50 kg m-r, new type) rail respectively. Rail specimens were 66.3-76.2 mm in diameter, with a loading surface 5 mm wide. The corresponding dimensions for the wheel specimens were 76.2-86.1 mm and 12.7 mm. Except for the 0% slip case, the rail specimen diameter was smaller than that of the wheel specimen. Therefore, the twin-disk machine produced a situation similar to acceleration in a wheel-rail system, in that the motion of the contact point and shear traction were in opposite directions
for the rail specimen and in the same direction for the wheel specimen. For a given specimen, the relative direction of motion of the counterface defined the direction of the traction force. The chemical composition for the steel used is presented in Table 1. Rail specimens were heat treated to provide two different microstructures, each having three different hardnesses, 280, 340 or 300, and 360 HV. The pearlitic specimens (PL280, PL340, PL360) were produced by induction heating followed by cooling with high pressure air. The tempered martensite specimens (TM280, TM300, TM360) were produced by induction heating followed by oil quenching and tempering at three different temperatures (600, 580, 525 “C). Typical microstructures are shown in Fig. 2. All wheel specimens were processed to give 360 HV (TM360). Double-distilled water was used for tests involving lubrication with water. In some cases, sodium nitrite TABLE 1. Chemical composition (weight per cent) of the steel used for rail and wheel test specimens C
Si
Mn
s
P
0.77
0.24
0.88
0.013
0.008
(4
Gears
0’) Wheel specimen
Fig. 1. Schematic diagram of twin-disk rolling contact test machine.
Fig. 2. Secondary electron images showing bulk microstructures for (a) pearlite, 280 IW, and (b) tempered martensite, 280 HV.
M. Sato et al. I Rolling-sliding
(0.01 M) was added as a corrosion inhibitor. A low viscosity silicone oil (Dow Coming 200, polydimethylsiloxane, 9 cS, 9 x lO-‘j m2 s-’ at 38 “C) was also used for comparison with the water-lubricated tests. For convenience, the tests run entirely without lubricant will be referred to as unlubricated or dry wear tests. None of these resulted in damage of the type broadly defined earlier as rolling contact fatigue. All wear tests were conducted at 750 N in air at room temperature. The number of cycles was 300 000 at 1725 rev min-‘, except for the test with a slip ratio of 26%. In that case, there was much heat generation, and the test was stopped after 126 000 cycles. Conditions for the unlubricated tests are summarized within three groups in Table 2. In the first group, the principal variable is the slip ratio. In the other two groups, the principal variable is the hardness of pearlite and of tempered martensite. The tests with lubricant are listed in four groups in Table 3. These emphasize the effects of environment, slip ratio, hardness of pearlite and hardness of tempered martensite. In most cases, the tests were run dry for 300 000 cycles and then with lubricant for 300 000 additional cycles. For comparison, all-dry and all-lubricated tests were also run for 600 000 cycles. The first six tests of environment effects were carried out at 750 N and the last four tests at 500 N. Three of the six tests at 750 N involved testing without changing the environment during testing; the other three involved combinations of dry testing followed by continued testing with lubricant. Most of the tests of environment used TABLE
2. List of conditions
used
Microstructure and sample designation
for unlubricated Hardness 0-W
of railsteeh
161
a slip ratio of 5%; however, in some tests 0% slip was used in the second stage. Weight changes were determined by measuring mass to within f5 mg before and after testing. A Buehler microhardness tester with a load of 0.2 kgf (1.96 N) was used to measure Vickers microhardness. These measurements provided information on hardness gradients and the depth of work hardening below the specimen surface. Indentations were spaced approxi: mately 50 pm apart. Polished but unetched sections were first observed by optical microscopy to reveal cracks and inclusions. The samples were then etched with nital to reveal changes in microstructure, the thickness of the highly deformed material and correlations of microstructure with fracture. For scanning electron microscopy (SEM) observations, etching with 4% picral was also used to obtain better detail.
3. Experimental results 3.1. Tests without lubricant 3.1.1. Effect of slip ratio Figure 3(a) shows the volume wear for the PI280 rail and TM360 wheel specimens at different values of nominal slip ratio, as given in group 1 of Table 2. The data for 26% slip have been extrapolated from 126 000 cycles to 300 000 cycles, assuming a constant wear rate. For each slip ratio the wear is larger for the rail specimens than for the wheel specimens. The
(dry) wear tests Load (N)
Slip ratio W)
Environment
Number Cycles (Xl@)
Slip ratio Pearlite (PL280)
280
750
0
300
Pearlite Pearlite
(PL280) (PL280)
280 280
750 750
5 5
300 300
Pearlite Pearlite
(PL280) (PL280)
280 280
750 750
10 10
300 300
Pearlite
(PL280)
280
750
26
126
Hardness of pearlite Pearlite (PL340) Pearlite (PL340)
340 340
750 750
300 300
Pearlite Pearlite
360 360
750 750
300 300
280
750 750 750
(PL360) (PL_360)
Hardness of tempenzd Tempered martensite Tempered martensite Tempered martensite
marten&e (TM280) (TM300) (TM360)
300 360
5 5 5
300 300 300
of
M. Sate et al. i Rolling-sliding of rail steels
162 TABLE 3 List of conditions
for two-stage
(“rolling contact fatigue”)
tests, most involving lubricant
Microstructure and sample designation
Hardness (HV)
Load 0)
Environment Pearlite (PL280) Pearlite (PL280) Pearlite (PL280)
280 280 280
7.50 750 750
5+5 5+0 5+5
Pearlite Pearlite Pearlite
(PL280) (PL280) (PL280)
280 280 280
750 750 750
Pearlite Pearlite Pearlite Pearlite
(PL280) (PL280) (PL280) (PL280)
280 280 280 280
so0 500 500 500
Slip ratio Pearlite (PL280) Pearlite (PL280) Pearlite (PL280)
280 280 280
750 750 500
Hardness of pearlite Pearlite (PL.340) Pearlite (PL360)
340 360
Hardness of tempered martensite Tempered martensite (TM280) Tempered martensite (TM280) Tempered Tempered
iI
(4
martensite martensite
5
10
15
20
si,p ram inem)
(TM300) (TM360)
25
%
30
Slip ratio (%)
in at least one
stage
Environment
Number of Cycles (X 10’)
Dry+dry Dry+dry Water + water
300+300 3!30+300 300 + 300
5+0 51-5 5+5
Dry+water Dry + water Dry + water
300+300 300+300 300+700
5+5 5+5 5+0 5+5
Dry + water Dry + water (Inhib.) Dry+oil Dry-toil
300+300 3OOt300 300 + 300 300+300
oco 10+ 10 26+0
Dry+water Dry + water Dry f water
300+300 300-t-300 126+300
750 750
5+5 5+5
Dry f water Dry+water
300 + 300 300+300
280 280
750 750
5+5
Dry+ water Dry + water
300+300 300+300
300 360
750 750
Dry + water Dry f water
300 + 300 300 + 300
35
II
tb)
5
10
15
20
sip ram
(rail!
25 %
5+5 5+5
5+s
30
0
35
fc)
5
10
15
20
Sbp ratio (mill
25
30
75
9b
Fig. 3. Dependence of volume wear of rail specimens (PL280) and wheel specimens (TM360) on slip ratio. Values of slip ratio (defined in text) were 0%, 5%, 10% and 26% (nominal) or O%, 5.1%, 10.5% and 29.9% (rail). The normal load was 750 N. The number of cycles was 300 000, except for the case of 26% slip, which was normalized from 126 OOflto 300 000 cycles. (a) Wear z~s. slip ratio (nominal) for rail and wheel specimens; (b) wear US. slip ratio (rail) for rail specimens; (c) wear normalized by slip ratio (rail) for rail specimens.
data for the rail wear are replotted in Fig. 3(b) using the rail rather than the nominal slip ratio. In Fig. 3(c), the same data are plotted with the wear normalized by the rail slip ratio; the ordinate in this case should be proportional to wear rate. The normalized wear (Fig. 3(c)) is appro~mately the same for the 5% and 10% slip ratios, but somewhat higher for the 26% ratio. Figure 4 shows typical optical micrographs of unetched and nital-etched lo~~tudinal sections taken through the center of rail specimens tested at different slip
ratios. The depth of the highly deformed material can be estimated from the etched sections. It is less than 10 pm for the case of 0% slip, but clearly larger when slip is finite. Only very small cracks were observed for 0% and 26% slip. For 5% and 10% slip, a few larger cracks were found. They followed microst~~tural features which indicated local directions of plastic flow. The largest of these cracks is shown in Figure 4(b) (5% slip). The average value of the deepest cracks is shown in Fig. 5 for 5% and 10% slip. Also shown there
163
iU. Sato et al. 1 Rolling-sliding of rail steels
L
a, 1.00
A
0 ‘t 2 g
8
0.10
0
Crack depth
0
O
IO
5
15
20
25
30
35
Slip ratio (rail), % Fig. 5. Comparison of depth of work hardening (estimated from microhardness) and depths of plastic flow and cracks (both estimated from longitudinal sections) for PL280 rail specimens. The logarithmic scale for the ordinate should be noted. Conditions: 750 N, 300000 cycles, unlubricated.
A Wheel (TM360) wHh PL rail A Wheel (TM360) wHhTM rail
260
200
300
320
340
360
Initial rail hardness ,
(4
(4
Fig. 4. Optical micrographs of unetched and etched longitudinal sections of worn rail specimens (750 N, unlubricated) for different slip ratios: (a) 0%; (b) 5%; (c) 10%; (d) 26%.
are data for the depth of plastic flow estimated visually from Fig. 4 and the depth of work hardening defined as the depth at which the microhardness has fallen to within 50 HV of the bulk value. The depth of work hardening is roughly proportional to the rail slip ratio. 3.1.2. Effects of hardness and microstructure A constant slip ratio of 5% was selected for the tests listed in groups 2 and 3 of Table 2 to determine the influence of hardness and microstructure on wear. Data for pearlite and tempered martensite are shown in Fig. 6 for different values of initial hardness. For both pearlite and tempered martensite rail specimens, wear decreased approximately linearly with increases in hard-
380
Hv
Fig. 6. Comparison of wear of rail and wheel specimens for different microstructures and different values of rail hardness. Conditions: 750 N, 5% slip, unlubricated. Error bars estimated from spread of two or three data points.
ness. The wear of pearlite was slightly lower than for tempered martensite specimens, but the difference was less than typical error bars estimated from the spread of two or three data points. The wear of the wheel specimens was smaller and approximately independent of the microstructure and of the initial hardness of the rail specimens. A stronger trend is evident from Fig. 7, in which the wear of rail specimens is plotted 2)s.surface hardness (Vickers) after testing. The data for pearlite and tempered martensite lie close to a single smooth curve which falls more steeply at higher surface hardness levels. The increase in surface hardness ranges from 267 to 327 HV for pearhte and from 240 to 310 HV for tempered martensite.
M. Sate et al. / Rolling-sliding of rail steels
164
500
540
580
620
660
700
Surface hardness after test , Hv Fig. 7. Wear of rail specimens for different microstructures and different values of surface hardness after testing. Conditions: 750 N, 5% slip, unlubricated, The trends are similar for the two microstructures.
(h)
(cl (4
@I
Fig. 8. Optical micrographs of unetched and etched longitudinal sections of worn rail specimens (750 N, 5% slip, unlubricated): (a) PL340; (b) PL360.
Figures 4(b) and 8 show optical micrographs of etched and unetched longitudinal sections of pearlite rail specimens with different initial hardness. The depth of plastic flow estimated from the micrographs is much less for the PL340 and PL360 than for the PL280 specimens. The PL280 specimen (Fig. 4(b)) has the deepest crack of the type clearly aligned with plastic flow lines. The crack in the PL360 (Fig. 8) seems to be related to inclusions. Except for that special crack, little fracture was evident for PL340 and PL360. Longitudinal sections for the tempered martensite rail specimens and for one TM360 wheel specimen are shown in Fig. 9. The etched samples show that TM280
Fig. 9. Optical micrographs sections of worn rail and unlubricated): (a) TM280 (d) TM360 wheel (against
(4 of unetched and etched longitudinal wheel specimens (7.50 N, 5% slip, rail; (b) TM300 rail; (c) TM360 rail; TM360 rail).
has the largest depth of visible plastic flow. For the sections illustrated, the depth of flow is greater for the TM360 wheel specimen than for the TM360 rail specimen. Comparison of the unetched samples indicates that the only crack is a small crack in the TM360 rail specimen. 3.1.3. Scanning electron microscopy observations Figure 10 shows secondary electron images of longitudinal sections of a PL280 rail specimen after etching. In Fig. 10(a) there is a layer of material approximately 5-8 pm thick in which the cementite lamellae are lined up parallel to the wear surface. Most of the lamellae in this region are broken into small particles. Below
M. Sato et al. / Rolling+liding of rail steels
16.5
and sott regions which would influence crack propagation. Both SEM and optical microscopy revealed that cracks stopped within the region in which evidence of plastic flow was visible. This was characteristic of specimens tested without lubrication. Figure 11(a) shows a secondary electron image of a typical wear surface (PL280, 5% slip). The craters visible in Fig. 11(a) are similar in size to the wear debris shown in Fig. 11(b). The debris are typically 2-3 pm thick and l-30 pm long. Some of the particles have fine grooves which are similar in width to those on the wear surface. Also, the thicknesses of debris particles are similar to the thicknesses of layers of broken cementite adjacent to cracks at the surface. (4 3.2. Tests with lubricant 3.2.1. Eficts of environment Tests on a series of PL280 specimens, as summarized in group 1 of Table 3, were run to determine the effects of different environments, including air, water, water with inhibitor or silicone oil. Volume loss data, calculated from measured changes in mass, are shown in Fig. 12 for different conditions. The generic term “volume loss”
Fig. 10. Secondary electron images of etched longitudinal section of PL280 rail specimen (750 N, 300 000 cycles, 5% slip, unlubricated): (a) near surface, showing layer in which cementite lamellae are aligned with the surface and broken into small particles; (b) near a crack, showing alignment of crack with microstructure.
this layer (Fig. 10(b)), the lamellae are aligned with each other, but they are inclined more steeply to the surface. Figure 10(b) also shows cracks which seem to start at or near the surface. A typical crack follows flow lines marked by aligned pearlite for most of its path. However, the deepest part of the crack deviates from the local flow lines toward a direction more steeply inclined to the surface. Examination of optical micrographs of the highly deformed material also reveals variations in the spacing of cementite lamellae. In the bulk material, some of these variations could be associated with sets of lamellae inclined at different angles to the plane of the longitudinal section. However, closer to the surface, the principal rotation is known to be about the transverse axis [8], so some of the variations there are likely to be associated with real differences in lamellar spacing. Such an inhomogeneous structure would provide hard
(a)
@I Fig. 11. Secondary electron images of (a) wear surface and (b) wear debris for PL280 rail specimen (750 N, 300 000 cycles, 5% slip, unlubricated).
166
M. Sate et al. I Rolling-sliding
I I I ( I I I , I I 1, I I I , 1 I I , I I
250
Dry 5%+dty 5% Dry 5%+dry 0% Water 5%+water 5%
. 200
-
A n
mE E_
150
-
-0 9) 100
-
0
I%
3E p
A 50
A
n
0 il,,‘,‘t”‘l’l’l’,,l’,,, 0
200
(a)
2 a E 2 3 9
400
600
600
1000
1200
Rotation cycles (x1000)
100
50
!
A
:
l
A n
Dry 5%+Water Drv 5%+Water Dry S%+Water
0% 5% 5%
0~““‘“““‘“““““’ 0
200
600
1000
1200
1000
1200
600
Rotation cycles (xl 000)
@I
0
Cc)
400
200
400
600
600
Rotation cycles (x1000)
Fig. 12. Volume loss of rail PL2M specimens (5% and/or 0% slip) vs. number of rotation cycles for different combinations of unlubricated and lubricated conditions: (a) all dry or all wet (water), 750 N, (b) dry+water, 750 N, (c) dry+water (with or without inhibitor) or dty+silicone oil, 500 N.
has been used here because the damage mechanism depends on the conditions. Only in the case of the unlubricated test with 5% slip ratio (labelled dry
of rail steels
5%+dry 5% in Fig. 12(a)) does the volume loss have the same meaning as the volume wear in Fig. 3(a). In Fig. 12(a), the data are for tests in which the environment remained constant during each test, i.e. all dry or all water. The volume loss is about five times less with water than without. Also, there is negligible continuing loss of material when the slip ratio is changed from 5% to 0% (unlubricated). The data shown in Figs. 12(b) and 12(c) resulted from combination tests: 300 000 cycles, unlubricated, 5% slip, followed by additional testing with a lubricant. Material continues to be lost at a high rate during the second stage of testing. Comparison of these results with those in Fig. 12(a) shows that volume loss during lubricated stages can be comparable with wear rates under dry conditions, provided that the lubricated stage was preceded by a dry sliding stage. Similar results were obtained for each choice of lubricant used in the second stage (water, water + inhibitor, silicone oil) (Fig. 12(c)). The micrographs of longitudinal sections for rail specimens for the all dry or all wet tests are shown in Fig. 13. They are as expected from the volume loss data shown in Fig. 12(a). For the continuously waterlubricated test, neither a deformed layer nor cracks are visible in the micrograph (Fig. 13(c)). Figures 14(a)-14(c) show longitudinal sections of rail specimens tested first under dry conditions (5% slip) followed by testing in water. The damage is dramatically different from the cases presented earlier. Although some cracks appeared during the dry stage, many more developed and propagated during the second (wet) stage. Even when no cracks appeared during the dry stage, there was much cracking during the second stage. Similar damage occurred even when the slip ratio in the second stage was 0%. Even when the rail specimen had much damage, the wheel specimen (TM360) had very few cracks. A typical example is shown in Fig. 14(d). Figure 15(a) shows that damage similar to that shown in Figs. 14(a)-14(c) was evident in dry+water (5% slip) tests even when the load was reduced from 750 N to 500 N. The other parts of Fig. 15 show that similar damage was produced when the lubricant in the second stage was changed from water to water with inhibitor or to silicone oil. Data for the depths of work hardening (defined in Section 3.1.1), visible plastic flow and fracture are presented in Figs. 16(a)-16(c) which correspond to the tests of Figs. 12(a)-12(c). Again, with no lubricant present, the depth of work hardening increased during testing with 5% slip but did not increase further during a subsequent stage of 0% slip. Also, for tests with water lubrication at all times, the depth of work hardening was small. Figure 16(a) shows that cracks remained
167
M. Sate et al. / Rolling-sliding of rail steek
(cl Fig. 13. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens (750 N). These correspond to the three data points shown in Fig. 12(a) for 600 000 total cycles: (a) dry, 5% slip+dry, 5% slip; (b) dry, 5% slip+ dry, 0% slip; (c) water, 5% slip+water, 5% slip.
within the visibly deformed material when the environment for a given test remained constant. However, the behavior is strikingly different when a second stage with a lubricant followed a dry stage with slip (Figs. 16(b) and 16(c)). In those cases cracks propagated well into the region of little or no work hardening. 3.2.2. Scanning electron microscopy observations Figure 17 shows secondary electron images for a longitudinal section of a PL280 rail specimen after testing with 5% slip, first without lubricant and then with water. The cracks penetrate very deeply, as noted in the previous section. Also, their paths are strongly influenced by microstructural features such as cementite
(cl
(4
Fig. 14. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens (750 N). These correspond to the data shown in Fig. 12(b). Also shown is one wheel specimen (TM360) for comparison. (a) Dry, 5% slip+water, 0% slip, rail specimen; (b) dry, 5% slip+water, 5% slip, rail specimen; (c) dry, 5% slip +water, 5% slip, rail specimen, extended test; (d) dry, 5% slip+water, 5% slip, wheel specimen.
lamellae, pearlite colonies and grain boundaries. This results in much branching of the cracks. There are even places such as that marked by an arrow where the cracks propagate toward the surface after first growing away from the surface. The microhardness measured near such cracks was not significantly different from microhardness of adjacent material at the same distance from the surface. 3.2.3. Effect of slip ratio The importance of slip ratio in the first (dry) stage is emphasized in Fig. 18, which shows volume loss VS.
o
1.0
-
Work hardening
0
Plastic flow
A
Crack depth
0.8 0.8
-
0.4
-
0.2
-
0
D
A
0.0
I
t
05%
I
D5%+05%
0” +
Q
05%+00%
WS%+WS%
1
Testing conditions 1.2
(b) 1 .O
[
I
I
:
0
Work hardening
.
I
Crack depth
I
0.0 t
05%
I
I
A
I D5%+WO%
I
OF&W596
f
I
1
DEi%+W5%(700)
Testing conditions 1.2
L
I
I
1.0
(4
td)
Fig. 15. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens (500 N). These correspond to the data shown in Fig. 12(c). (a) Dry, 5% siipiwater, 5% slip; (b) dry, 5% shp+water with inhibitor, 5% slip; (c) dry, 5% slip + silicone oil, 0% slip; (d) dry, 5% slip + silicone oil, 5% slip.
number of rotation cycles for combination tests. When the slip ratio was 0% in the first stage, the volume loss in both stages was very small. It was much larger when 5%, 10% or 26% slip was used in the first stage, and, as shown in Fig. 19, there was much fracture damage in the second stage (wet), even when the slip ratio in that stage was 0%. 32.4. Efect of hardness and microstructure The volume losses for pearlite and tempered martensite rail specimens after combination tests (dry, 5% slip+ water, 5% slip) are shown in Fig. 20 for different values of initial hardness. The volume losses for all of
I
I
J
I
1
A
0.8
-
0.6
-
0
-
A crackdepth
0.4
-
0.2
-
0.0
t
Work hardening
P
?
DS%+WS% D5%+W(t)5%
I
D5*LMX%
D5%+05%
Testing conditions Fig. 16. Comparisons of depth of work hardening (estimated from microhardness) and depths of visible plastic flow and cracks (both estimated fram longitudinal sections) for PL280 rail specimens. These correspond to the data shown in Figs. 12(a)-12(c). In each case the slip was 5% in the first stage. D represents dry or unlubricated and W and 0 represent lubrication with water and oil respectively. (a) All dry or all lubricated (water), 750 N; (b) d~+lub~cated (water), 0% or 5% slip, 750 N; (c) dry +Iubricated (water, water with inhibitor, or silicone oil), 0% or 5% slip, 500 N.
M. Sat0 ef al. I Rollingdiding
of rail sreeh
169
(4
Fig. 17. Secondary electron images of etched longitudinal section of PL280 rail specimen (unlubricated, 5% slip, 300 000 cycles+ lubricated, 5% slip, 300 CKJOcycles; 750 N): (a) lower magnification, showing branching of cracks, (b) higher magnification of region A, showing crack propagating toward surface.
0 0 n
IJ
Dry Dry Dty Dry
O%+Water O%, 750N S%+Water 5%, 750N lO%+Waler lo%, 750N 26%+Water O%, 500N
n
Cc)
_
0
Fig. 19. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens after testing first without lubricant followed by continued testing with water. These correspond to the data shown in Fig. 18. (a) Dry, 0% slip+water, 0% slip, 750 N (negligible damage); (b) dry, 10% slip+water, 10% slip, 750 N (much damage); (c) dry, 26% slip+water, 0% slip, 500 N (much damage).
these cases were similar, except that the PL280 specimen showed a larger volume loss in the second stage.
4. 0
“,‘,““,,,~‘,,‘,,~,‘,,, 100 200
300
400
500
600
700
Rotation cvcles (xl 000) Fig. 18. Comparison of volume losses of PL280 rail specimens after testing first without lubricant followed by continued testing with water. Note especially the results when 0% slip was used in the first (unlubricated) stage.
Discussion
4.1. Unlubricated tests Beagley [9] has described mild (type I) and severe (type II) wear modes for rail steels. Clayton and coworkers [l&12] reported an additional wear mode (type III) in which the surface was roughened more and the debris particles were larger. Their tests involved 500
170
M. Sato et al. f Rolling-sliding
700
in a dry environment. The present work is in agreement with this observation. In contrast, Way [4] has concluded from tests on bearings that cracking only occurs when lubricant is present. The difference in behavior is not surprising in view of the marked differences in contact stress, geometry, materials, microstructure and hardness. In the present work, some cracks did appear during unlubricated testing. However, they propagated along flow lines until they were blocked or diverted by material which was not suitably aligned. That is, these cracks remained within the region of visible flow. The mechanism for initiation of cracks during unlubricated testing cannot be determined with certainty from the present work. However, local variations in the extent of surface flow have been invoked by Kjer [15] and by Oxley and co-workers [16]. In both cases, material from one part of the surface advances over adjacent surface. According to Oxley and co-workers, this creates a suitable site for initiation of fatigue cracks. A simple elastic analysis of a combined normal and shear line contact load moving past an inclined surface crack as shown in Fig. 21 does not predict either the observed crack growth direction or significant values of crack tip driving force. In the absence of a crack, the radial stress produced by a compressive normal force P and corresponding shear force Q per unit depth is given by [17]
700
a,=
o
0
100
300
400
500
600
Rotation cycles (x1000)
(a)
01,,,, 0
(b)
200
100
200
300
400
0
TM280
A
TM300
l
TM360
500
600
of rail steels
Rotation cycles (xl 000)
Fig. 20. Volume loss of rail specimens with different microstructures after testing first without lubricant (5% slip) followed by continued testing with water (5% slip): (a) pearlite: (b) tempered martensite.
MPa and 25% slip, which is close to the 525 MPa used in the present study. Therefore the present conditions with 26% slip would correspond to the conditions used by Clayton and co-workers for the tests in which the wear was described as type III. This could be the reason that the result shown in Fig. 3(c) for 26% slip is higher than the normalized results for 5% and 10% slip. Sugino et al. [13] reported better wear resistance of pearlite rail steels compared with tempered martensite in rolling contact testing. They suggested that the difference in performance was related to the greater increase in hardness of pearlite during testing compared with tempered martensite. The results of the present work are similar, except that the differences reported here are smaller. Sugino et al. also reported that wear correlated with the hardness measured after testing, in agreement with the results shown in Fig. 7. Akama and Matsuyama [14] have noted that some cracks are generated during testing of rail steels even
21 -P cos tl+Q sin 8 , u~~=O, u,=vu, -7rr (
(1)
where (r, 8, z) are polar coordinates centered on the contact line, such that fI= (0, -r/2) denotes the directions of (P, Q) respectively. Plane strain conditions are assumed along the direction of line contact (z axis), and v is Poisson’s ratio. For 8> &= tan-‘(P/Q), a,, is positive, with a maximum value of 2Q/m at 8= n/2. In the homogeneous, isotropic, elastic material modeled, cracks would tend to nucleate at the surface where the stress is largest, at the inclination 8=0. Typical crack angles 6, vary from near 90” at the surface to about 60” further into the highly deformed material. The difference between
Fig. 21. Surface crack and line contact geometry used for fracture analysis.
M, Sato et al. / Rolling-sliding of rail steels
predicted and observed crack angles may be caused by the substantial material rotation and large orientation dependence of fracture toughness in the textured material near the surface. Estimates of mode I and mode II stress intensity factors for an edge crack of length L are obtained according to [18] F(x’/L)a,,&‘/L)
d(x’/L)
(2) P(x’/L)a,,&‘/L)
d(x’/L)
(3) where @‘IL)
= 1.3 - 0.3x’lL
(4)
and a,,,(x’/L) and q,,,,(x’/L) denote normal and shear components of traction along the crack surface, from x’lL=O to xl/L= 1. This solution is applied to a surface crack at inclination 0, and position x=u relative to the contact load by using eqn. (1) to estimate the normal and tangential components of traction in eqns. (2) and (3). Accordingly, portions of the crack near enough to the surface, so that x’/L < (u/L) tan $/(tan 0, -tan Q, are predicted to be loaded in tension, and the remaining portions of the crack nearer to the tip are normally in compression. The results for a crack at &= 68” and Q/P= 0.6 are that KI increases to a maximum of about 0.02P/L”2 as the contact load approaches the crack, corresponding to a/L +O. In this same interval, K,, reaches a substantially larger value of 0.12P/L’n. As the load passes the crack, both KI and K,, reach large negative values of - 2.2PIL’R and about 0.7P/Lln respectively at alL = - 1, and rapidly decay to zero as a/L + - m. Assuming P= 0.15 MN m-’ and L =O.Ol mm, KI ,.,,== 0.95 MPa min. This is significantly lower than nominal fracture toughness values for high strength steels, but seems to be comparable with threshold fatigue values in textured steels [13]. However, this analysis does not provide a suitable explanation for continued crack growth, since K decreases with crack length. Also, it does not provide a suitable explanation for observed crack growth directions. The residual stresses produced by near-surface plasticity and large anisotropy in fracture toughness due to texturing, although not incorporated into the model, may be important features that control crack extension force and crack growth direction.
171
Sugino et al. [5] have discussed a possible interaction between wear and the extent of cracking. If wear is high enough, damaged material is removed before cracks can propagate deeply. This could explain the observation that very little cracking was observed when the slip ratio was high (see Figs. 3-5). 4.2. Tests with lubricant The worst damage was observed when two conditions were satisfied: (1) the slip ratio during the unlubricated stage was greater than zero, and (2) further testing, with or without slip, was in the presence of a liquid lubricant. This was consistently true, whether or not any cracks were observed after unlubricated testing. The density of cracks was much greater when both conditions (1) and (2) were satisfied than for condition (1) alone. Deformation during the dry stage seemed to prepare the material for extensive cracking during the lubricated stage. Whereas cracks produced during unlubricated tests remained within the region of visible flow, cracks produced with lubricant, preceded by sliding without lubricant, extended well beyond that region. This suggests that a fatigue process is involved in the latter case. It would be tempting to turn to stress-corrosion cracking to explain these results. However, similar behavior was found when a corrosion inhibitor was used with water as a lubricant and when a low viscosity silicone oil was used. In the case of the inhibitor, the results are not conclusive, because new surfaces generated by cracking and debris formation may have depleted the local concentration of inhibitor in the lubricant. However, the behavior with silicone oil was similar to that with water, despite the expected differences in chemical reactivity. This result suggests that some mechanism other than hydrogen embrittlement and stress-corrosion may be involved. Among the candidates affecting the stress intensity are Way’s pressure mechanism [4, 19, 201, a reduction in friction between crack faces [20] and surface energy effects [21].
5. Conclusions Effects of slip ratio, hardness, microstructure (pearlite and tempered martensite) and environment (dry, water and oil) on wear and rolling contact fatigue were examined. The main conclusions are summarized here. 5.1. Tests without lubrication (1) The thickness of the highly deformed layer is very small for pure rolling compared with cases involving slip. (2) Wear decreases as hardness of the rail steel increases.
172
M. Sato et al. I Rolling-sliding of rail steels
(3) For a given initial hardness, the wear rate of pearlite is slightly smaller than that of tempered martensite. (4) Wear debris are generated from near-surface material in which the cementite phase has broken into small particles. (5) Cracks initiate at the surface and propagate along lines of plastic flow, rather than normal to the local direction of maximum tensile stress. These cracks stop within the plastically deformed material. (6) Cracking is more extensive in material with low initial hardness. In such cases, tempered martensite is more resistant to cracking than pearlite. 5.2. Tests with lubricant (1) The results of rolling with or without sliding in water after dry rolling-sliding indicate that (i) many cracks appear during the second (wet) stage, (ii) cracks extend well beyond the region of visible plastic flow, suggesting that fatigue is involved, (iii) crack propagation and branching are influenced by local microstructural features, e.g, pearlite colony size and pearlite lamellar spacing and orientation, (iv) sliding is not needed in the second (lubricated) stage if sufficient deformation has occurred in the first (unlubricated) stage, and (v) extensive deformation during the dry stage seems to prepare the material for cracking during subsequent operation in water. (2) Behavior is similar for all fluids used as lubricants (water, water plus inhibitor, silicone oil).
Acknowledgments
The authors are pleased to acknowledge the support and encouragement of the Bethlehem Steel Corporation, which supplied the test machine, and the Nippon Steel Corporation, which supplied the test specimens. The following were also helpful in discussions or with experimental aspects of the research: W. A. Glaeser, A. R. Rosenfield, R. Farrar, J. Kocka, J. Ellis and C. MacDonald.
References 1 L&ton heavy axle load test: preliminary results come in, Railw. Track Struct., 85 (July 1989) 29-30.
2 R. K. Steele and D. H. Stone, Developments in railroad rail, Am. Railw. Eng. Assoc., Proc., 87 (1986) 311-358. 3 R. K. Steele, Recent North American experience with shelling in railroad rails (part I: observations), Am. Railw. Eng. ASSOC., Proc., 90 (1989) 311-345. 4 S. Way, Pitting due to rolling contact, J. Appl. Mech., 2 (1935) A49-A58. 5 K. Sugino, H. Kageyama and M. Sato, Effect of wheel and rail profiles on wear and fatigue behavior of rail, Proc. 2nd Int. Symp. on Contact Mechanics and Wear of Rail-Wheel Systems, 1986, University of Waterloo Press, Waterloo, Ont., 1987, pp. 435450. 6 M. Sato, K. Sugino, K. Tanikawa and H. Iida, The nature of lubricants and their influence on wear and fatigue behavior of rail, Proc. 2nd Int. Symp. on Wheel/Rail Lubrication, Memphis, TN, 1987, Association of American Railroads. 7 J. P. Sheehan and M. A. H. Howes, The effect of case carbon content and heat treatment on the pitting fatigue of 8620 steel, SAE Trans., 81 (1972) 1031-1045. 8 P. Heilmann, W. A. T. Clark and D. A. Rigney, Orientation determination of subsurface cells generated by sliding, Acta Metall., 31 (1983) 1293-1305. 9 T. M. Beagley, Severe wear of rolling-sliding contacts, Wear, 36 (1976) 317-335. 10 P. J. Bolton and P. Clayton, Rolling-sliding wear damage in rail and tyre steels, Wear, 93 (1984) 145-165. 11 D. Danks and P. Clayton, Comparison of wear processes for eutectoid rail steels: field and laboratory tests, Wear, 120 (1987) 233-250. 12 P. Clayton and D. Danks, Effect of interlamellar spacing on the wear resistance of eutectoid steels under rolling-sliding conditions, Wear, I35 (1990) 369-389. 13 K. Sugino, H. Kageyama, T. Urashima, S. Nishida, H. Masumoto and M. Hattori, Basic properties of high strength rail steels and new rail development, Nippon Steel Tech. Rep., I6 (1980) 103-119. 14 M. Akama and S. Matsuyama, Wear characteristics of wheel and rail steels, J. Jpn. Sot. Lubr. Eng., Znt. Ed., 8 (1987) 75-80. 15 T. Kjer, A lamination wear mechanism based on plastic waves, Proc. Int. Con& on Wear of Materials, American Society of Mechanical Engineers, New York, 1987, pp, 191-198. 16 B. S. Hockenhull, E. M. Kopalinsky, and P. L. B. Oxley, An investigation of the role of low cycle fatigue in producing surface damage in sliding metallic friction, Wear, I48 (1991) 135-146. 17 K. L. Johnson, ContactMechanics,Cambridge University Press, Cambridge, 1985, p. 16. 18 H. Tada, P. C. Paris and G. R. Irwin, The Stress Ann&s of Cracks Handbook, Del Research, Hellertown, PA, 1973, Section 8.3. 19 M. Kaneta and Y. Murakami, Effects of oil hydraulic pressure on surface crack growth in rolling/sliding contact. Tribal. Int., 20 (1987) 210-217. 20 A. F. Bower, The influence of crack face friction and trapped fluid on surface initiated rolling contact fatigue cracks, J. Lubr. Technol., II0 (1988) 704-711. due to surface 21 N. J. Petch, The lowering of fracture-stress adsorption, Philos. Mag., I (1956) 331-337.
Wear, 162-164 (1993) 159-172
Rolling-sliding M. Sate”,
behavior of rail steels
P. M. Anderson
and D. A. Rigney
Materials Science and Engineering, The Ohio State Universi& 116 West 19th Avenue, Columbus, OH 43210-1179 USA
Abstract A twin disk-rolling-sliding contact machine has been used to investigate the deformation, wear and cracking that occur during rolling-sliding of high-carbon rail and wheel steels. A martensitic wheel steel with initial hardness of 360 HV contacted pearlitic or martensitic rail steels of different initial hardness, 280-360 HV. Other variables included test time, test environment and slip ratio (0%, 5%, 10% and 26%). Tests were run in air, but with different lubrication conditions, e.g. dry, with water, water with inhibitor, silicone oil, and various combinations. Wear was measured by weight loss. Wear surfaces, sample cross-sections and wear debris were
observed by optical and scanning electron microscopy. Principal conclusions are as follows. The thickness of the highly deformed layer was very small for pure rolling compared with cases involving slip. Wear debris was generated from near-surface material in which the cementite phase had broken into small particles. Cracks initiated at the surface and propagated along lines of earlier plastic flow, rather than normal to the local direction of maximum tensile stress. The results of rolling-sliding in water or other liquids after dry rolling-sliding indicate that extensive deformation from sliding during the dry stage prepares the material for extensive cracking during subsequent operation with a lubricant.
1. Introduction Railroads provide energy-efficient systems for transport of both freight and passengers. Trends to increase axle load for the former and train speed for the latter require renewed attention to wear and other kinds of damage produced during operation of rail systems [l]. Steele and Stone [2] and Steele [3] have reviewed the many kinds of defects observed in rail systems. Some of these involve obvious plastic deformation. For others, the effects of the environment are clearly important, but the detailed mechanisms are not well understood. Part of the difficulty is related to the many variables which are involved. These include load, speed, slip ratio, materials and environmental conditions. It is convenient to think of wear processes as those which occur fairly steadily over extended periods of time. Material is typically lost by removal of small debris particles. In contrast, some damage to rails occurs more dramatically, with rather sudden removal of larger pieces of material. Such damage is commonly called rolling contact fatigue. This term may not be adequate to cover the full range of phenomena reported, but it is well established among those who work on tribological problems of railroads. It should be noted that results reported *Present address: Yawata R & D Laboratory, Nippon Corp., l-l Tobihata, Tobata, Kitakyushu 804, Japan.
Steel
in the literature show that wear and rolling contact fatigue processes can interact [2], so their separation is somewhat artificial. Previous work suggests that pitting or rolling contact fatigue occurs when unlubricated testing conditions are followed by continued testing with lubrication. Way [4] has reported that pitting cracks appeared soon after oil was added to a bearing system which had first been run without lubricant. He concluded that oil was necessary for pitting, but “some form of destruction leading to the formation of pitting cracks takes place independently of the oil.” Sugino and co-workers [5, 61 found no rolling contact fatigue for either dry or lubricated tests of rail steels, but rolling contact fatigue occurred readily when dry conditions were followed by operation with water or grease. These observations are important, because rails in use are exposed to alternating periods of dry and wet conditions. This research program was designed to provide new information on the relationships among wear, plastic deformation and fracture and the effects of hardness, microstructure, slip ratio and lubrication. Initial tests were run under dry or lubricated conditions. Tests with various combinations of unlubricated and lubricated conditions were also run. These included unlubricated tests with slip, followed by continued testing with lubrication (water, water with inhibitor, silicone oil).
Elsevier
Sequoia
M. &to
160
et al. i Rolling-siid~ng of rail steels
2. Experimental procedure A twin-disk rolling contact machine was used for all tests. The wheel specimen was mounted on the lower shaft, which was driven directly by the motor, as shown in Fig. 1 [7]. The rail specimen was mounted on the upper shaft, which was driven by a simple gear system. Both gears had the same number of teeth, so the rotation speeds of the two shafts were the same. To obtain different slip ratios, the circumferential speeds of the rail and wheel specimen disks were made different 1 by changing their relative diameters. A simple definition of the slip ratio is provided by the expression 2(L), - D,)/(D, +D,), where I),., and D, are the diameters of the wheel and rail respectively. The following values were used for this nominal slip ratio: SR(nom) =O%, 5%, 10% and 26%. To focus attention on the loading conditions for the rail specimen, another expression for slip ratio is also used in this paper. It is equal to the ratio of distance slid to distance rolled at the rail specimen surface, defined as SR(rai1) = (D, -D,)/D, = 0%, 5.1%, 10.5% and 29.9%. In the discussion to follow, nominal and rail slip ratios are clearly distinguished from each other by the numbers cited, e.g. 5% compared with 5.1%. The normal load was supplied by the dead-weight of the upper shaft and holder (314 N) and compressed air which pressed a loading bar against the upper shaft. Two loads were used for these tests, 750 N (maximum elastic contact stress, 525 MPa) and 500 N (430 MPa). Rail and wheel specimens were cut from the base portion of a Japanese Industrial Standard 60 kg (60 kg m-l) rail and a 50 kg “N” (50 kg m-r, new type) rail respectively. Rail specimens were 66.3-76.2 mm in diameter, with a loading surface 5 mm wide. The corresponding dimensions for the wheel specimens were 76.2-86.1 mm and 12.7 mm. Except for the 0% slip case, the rail specimen diameter was smaller than that of the wheel specimen. Therefore, the twin-disk machine produced a situation similar to acceleration in a wheel-rail system, in that the motion of the contact point and shear traction were in opposite directions
for the rail specimen and in the same direction for the wheel specimen. For a given specimen, the relative direction of motion of the counterface defined the direction of the traction force. The chemical composition for the steel used is presented in Table 1. Rail specimens were heat treated to provide two different microstructures, each having three different hardnesses, 280, 340 or 300, and 360 HV. The pearlitic specimens (PL280, PL340, PL360) were produced by induction heating followed by cooling with high pressure air. The tempered martensite specimens (TM280, TM300, TM360) were produced by induction heating followed by oil quenching and tempering at three different temperatures (600, 580, 525 “C). Typical microstructures are shown in Fig. 2. All wheel specimens were processed to give 360 HV (TM360). Double-distilled water was used for tests involving lubrication with water. In some cases, sodium nitrite TABLE 1. Chemical composition (weight per cent) of the steel used for rail and wheel test specimens C
Si
Mn
s
P
0.77
0.24
0.88
0.013
0.008
(4
Gears
0’) Wheel specimen
Fig. 1. Schematic diagram of twin-disk rolling contact test machine.
Fig. 2. Secondary electron images showing bulk microstructures for (a) pearlite, 280 IW, and (b) tempered martensite, 280 HV.
M. Sato et al. I Rolling-sliding
(0.01 M) was added as a corrosion inhibitor. A low viscosity silicone oil (Dow Coming 200, polydimethylsiloxane, 9 cS, 9 x lO-‘j m2 s-’ at 38 “C) was also used for comparison with the water-lubricated tests. For convenience, the tests run entirely without lubricant will be referred to as unlubricated or dry wear tests. None of these resulted in damage of the type broadly defined earlier as rolling contact fatigue. All wear tests were conducted at 750 N in air at room temperature. The number of cycles was 300 000 at 1725 rev min-‘, except for the test with a slip ratio of 26%. In that case, there was much heat generation, and the test was stopped after 126 000 cycles. Conditions for the unlubricated tests are summarized within three groups in Table 2. In the first group, the principal variable is the slip ratio. In the other two groups, the principal variable is the hardness of pearlite and of tempered martensite. The tests with lubricant are listed in four groups in Table 3. These emphasize the effects of environment, slip ratio, hardness of pearlite and hardness of tempered martensite. In most cases, the tests were run dry for 300 000 cycles and then with lubricant for 300 000 additional cycles. For comparison, all-dry and all-lubricated tests were also run for 600 000 cycles. The first six tests of environment effects were carried out at 750 N and the last four tests at 500 N. Three of the six tests at 750 N involved testing without changing the environment during testing; the other three involved combinations of dry testing followed by continued testing with lubricant. Most of the tests of environment used TABLE
2. List of conditions
used
Microstructure and sample designation
for unlubricated Hardness 0-W
of railsteeh
161
a slip ratio of 5%; however, in some tests 0% slip was used in the second stage. Weight changes were determined by measuring mass to within f5 mg before and after testing. A Buehler microhardness tester with a load of 0.2 kgf (1.96 N) was used to measure Vickers microhardness. These measurements provided information on hardness gradients and the depth of work hardening below the specimen surface. Indentations were spaced approxi: mately 50 pm apart. Polished but unetched sections were first observed by optical microscopy to reveal cracks and inclusions. The samples were then etched with nital to reveal changes in microstructure, the thickness of the highly deformed material and correlations of microstructure with fracture. For scanning electron microscopy (SEM) observations, etching with 4% picral was also used to obtain better detail.
3. Experimental results 3.1. Tests without lubricant 3.1.1. Effect of slip ratio Figure 3(a) shows the volume wear for the PI280 rail and TM360 wheel specimens at different values of nominal slip ratio, as given in group 1 of Table 2. The data for 26% slip have been extrapolated from 126 000 cycles to 300 000 cycles, assuming a constant wear rate. For each slip ratio the wear is larger for the rail specimens than for the wheel specimens. The
(dry) wear tests Load (N)
Slip ratio W)
Environment
Number Cycles (Xl@)
Slip ratio Pearlite (PL280)
280
750
0
300
Pearlite Pearlite
(PL280) (PL280)
280 280
750 750
5 5
300 300
Pearlite Pearlite
(PL280) (PL280)
280 280
750 750
10 10
300 300
Pearlite
(PL280)
280
750
26
126
Hardness of pearlite Pearlite (PL340) Pearlite (PL340)
340 340
750 750
300 300
Pearlite Pearlite
360 360
750 750
300 300
280
750 750 750
(PL360) (PL_360)
Hardness of tempenzd Tempered martensite Tempered martensite Tempered martensite
marten&e (TM280) (TM300) (TM360)
300 360
5 5 5
300 300 300
of
M. Sate et al. i Rolling-sliding of rail steels
162 TABLE 3 List of conditions
for two-stage
(“rolling contact fatigue”)
tests, most involving lubricant
Microstructure and sample designation
Hardness (HV)
Load 0)
Environment Pearlite (PL280) Pearlite (PL280) Pearlite (PL280)
280 280 280
7.50 750 750
5+5 5+0 5+5
Pearlite Pearlite Pearlite
(PL280) (PL280) (PL280)
280 280 280
750 750 750
Pearlite Pearlite Pearlite Pearlite
(PL280) (PL280) (PL280) (PL280)
280 280 280 280
so0 500 500 500
Slip ratio Pearlite (PL280) Pearlite (PL280) Pearlite (PL280)
280 280 280
750 750 500
Hardness of pearlite Pearlite (PL.340) Pearlite (PL360)
340 360
Hardness of tempered martensite Tempered martensite (TM280) Tempered martensite (TM280) Tempered Tempered
iI
(4
martensite martensite
5
10
15
20
si,p ram inem)
(TM300) (TM360)
25
%
30
Slip ratio (%)
in at least one
stage
Environment
Number of Cycles (X 10’)
Dry+dry Dry+dry Water + water
300+300 3!30+300 300 + 300
5+0 51-5 5+5
Dry+water Dry + water Dry + water
300+300 300+300 300+700
5+5 5+5 5+0 5+5
Dry + water Dry + water (Inhib.) Dry+oil Dry-toil
300+300 3OOt300 300 + 300 300+300
oco 10+ 10 26+0
Dry+water Dry + water Dry f water
300+300 300-t-300 126+300
750 750
5+5 5+5
Dry f water Dry+water
300 + 300 300+300
280 280
750 750
5+5
Dry+ water Dry + water
300+300 300+300
300 360
750 750
Dry + water Dry f water
300 + 300 300 + 300
35
II
tb)
5
10
15
20
sip ram
(rail!
25 %
5+5 5+5
5+s
30
0
35
fc)
5
10
15
20
Sbp ratio (mill
25
30
75
9b
Fig. 3. Dependence of volume wear of rail specimens (PL280) and wheel specimens (TM360) on slip ratio. Values of slip ratio (defined in text) were 0%, 5%, 10% and 26% (nominal) or O%, 5.1%, 10.5% and 29.9% (rail). The normal load was 750 N. The number of cycles was 300 000, except for the case of 26% slip, which was normalized from 126 OOflto 300 000 cycles. (a) Wear z~s. slip ratio (nominal) for rail and wheel specimens; (b) wear US. slip ratio (rail) for rail specimens; (c) wear normalized by slip ratio (rail) for rail specimens.
data for the rail wear are replotted in Fig. 3(b) using the rail rather than the nominal slip ratio. In Fig. 3(c), the same data are plotted with the wear normalized by the rail slip ratio; the ordinate in this case should be proportional to wear rate. The normalized wear (Fig. 3(c)) is appro~mately the same for the 5% and 10% slip ratios, but somewhat higher for the 26% ratio. Figure 4 shows typical optical micrographs of unetched and nital-etched lo~~tudinal sections taken through the center of rail specimens tested at different slip
ratios. The depth of the highly deformed material can be estimated from the etched sections. It is less than 10 pm for the case of 0% slip, but clearly larger when slip is finite. Only very small cracks were observed for 0% and 26% slip. For 5% and 10% slip, a few larger cracks were found. They followed microst~~tural features which indicated local directions of plastic flow. The largest of these cracks is shown in Figure 4(b) (5% slip). The average value of the deepest cracks is shown in Fig. 5 for 5% and 10% slip. Also shown there
163
iU. Sato et al. 1 Rolling-sliding of rail steels
L
a, 1.00
A
0 ‘t 2 g
8
0.10
0
Crack depth
0
O
IO
5
15
20
25
30
35
Slip ratio (rail), % Fig. 5. Comparison of depth of work hardening (estimated from microhardness) and depths of plastic flow and cracks (both estimated from longitudinal sections) for PL280 rail specimens. The logarithmic scale for the ordinate should be noted. Conditions: 750 N, 300000 cycles, unlubricated.
A Wheel (TM360) wHh PL rail A Wheel (TM360) wHhTM rail
260
200
300
320
340
360
Initial rail hardness ,
(4
(4
Fig. 4. Optical micrographs of unetched and etched longitudinal sections of worn rail specimens (750 N, unlubricated) for different slip ratios: (a) 0%; (b) 5%; (c) 10%; (d) 26%.
are data for the depth of plastic flow estimated visually from Fig. 4 and the depth of work hardening defined as the depth at which the microhardness has fallen to within 50 HV of the bulk value. The depth of work hardening is roughly proportional to the rail slip ratio. 3.1.2. Effects of hardness and microstructure A constant slip ratio of 5% was selected for the tests listed in groups 2 and 3 of Table 2 to determine the influence of hardness and microstructure on wear. Data for pearlite and tempered martensite are shown in Fig. 6 for different values of initial hardness. For both pearlite and tempered martensite rail specimens, wear decreased approximately linearly with increases in hard-
380
Hv
Fig. 6. Comparison of wear of rail and wheel specimens for different microstructures and different values of rail hardness. Conditions: 750 N, 5% slip, unlubricated. Error bars estimated from spread of two or three data points.
ness. The wear of pearlite was slightly lower than for tempered martensite specimens, but the difference was less than typical error bars estimated from the spread of two or three data points. The wear of the wheel specimens was smaller and approximately independent of the microstructure and of the initial hardness of the rail specimens. A stronger trend is evident from Fig. 7, in which the wear of rail specimens is plotted 2)s.surface hardness (Vickers) after testing. The data for pearlite and tempered martensite lie close to a single smooth curve which falls more steeply at higher surface hardness levels. The increase in surface hardness ranges from 267 to 327 HV for pearhte and from 240 to 310 HV for tempered martensite.
M. Sate et al. / Rolling-sliding of rail steels
164
500
540
580
620
660
700
Surface hardness after test , Hv Fig. 7. Wear of rail specimens for different microstructures and different values of surface hardness after testing. Conditions: 750 N, 5% slip, unlubricated, The trends are similar for the two microstructures.
(h)
(cl (4
@I
Fig. 8. Optical micrographs of unetched and etched longitudinal sections of worn rail specimens (750 N, 5% slip, unlubricated): (a) PL340; (b) PL360.
Figures 4(b) and 8 show optical micrographs of etched and unetched longitudinal sections of pearlite rail specimens with different initial hardness. The depth of plastic flow estimated from the micrographs is much less for the PL340 and PL360 than for the PL280 specimens. The PL280 specimen (Fig. 4(b)) has the deepest crack of the type clearly aligned with plastic flow lines. The crack in the PL360 (Fig. 8) seems to be related to inclusions. Except for that special crack, little fracture was evident for PL340 and PL360. Longitudinal sections for the tempered martensite rail specimens and for one TM360 wheel specimen are shown in Fig. 9. The etched samples show that TM280
Fig. 9. Optical micrographs sections of worn rail and unlubricated): (a) TM280 (d) TM360 wheel (against
(4 of unetched and etched longitudinal wheel specimens (7.50 N, 5% slip, rail; (b) TM300 rail; (c) TM360 rail; TM360 rail).
has the largest depth of visible plastic flow. For the sections illustrated, the depth of flow is greater for the TM360 wheel specimen than for the TM360 rail specimen. Comparison of the unetched samples indicates that the only crack is a small crack in the TM360 rail specimen. 3.1.3. Scanning electron microscopy observations Figure 10 shows secondary electron images of longitudinal sections of a PL280 rail specimen after etching. In Fig. 10(a) there is a layer of material approximately 5-8 pm thick in which the cementite lamellae are lined up parallel to the wear surface. Most of the lamellae in this region are broken into small particles. Below
M. Sato et al. / Rolling+liding of rail steels
16.5
and sott regions which would influence crack propagation. Both SEM and optical microscopy revealed that cracks stopped within the region in which evidence of plastic flow was visible. This was characteristic of specimens tested without lubrication. Figure 11(a) shows a secondary electron image of a typical wear surface (PL280, 5% slip). The craters visible in Fig. 11(a) are similar in size to the wear debris shown in Fig. 11(b). The debris are typically 2-3 pm thick and l-30 pm long. Some of the particles have fine grooves which are similar in width to those on the wear surface. Also, the thicknesses of debris particles are similar to the thicknesses of layers of broken cementite adjacent to cracks at the surface. (4 3.2. Tests with lubricant 3.2.1. Eficts of environment Tests on a series of PL280 specimens, as summarized in group 1 of Table 3, were run to determine the effects of different environments, including air, water, water with inhibitor or silicone oil. Volume loss data, calculated from measured changes in mass, are shown in Fig. 12 for different conditions. The generic term “volume loss”
Fig. 10. Secondary electron images of etched longitudinal section of PL280 rail specimen (750 N, 300 000 cycles, 5% slip, unlubricated): (a) near surface, showing layer in which cementite lamellae are aligned with the surface and broken into small particles; (b) near a crack, showing alignment of crack with microstructure.
this layer (Fig. 10(b)), the lamellae are aligned with each other, but they are inclined more steeply to the surface. Figure 10(b) also shows cracks which seem to start at or near the surface. A typical crack follows flow lines marked by aligned pearlite for most of its path. However, the deepest part of the crack deviates from the local flow lines toward a direction more steeply inclined to the surface. Examination of optical micrographs of the highly deformed material also reveals variations in the spacing of cementite lamellae. In the bulk material, some of these variations could be associated with sets of lamellae inclined at different angles to the plane of the longitudinal section. However, closer to the surface, the principal rotation is known to be about the transverse axis [8], so some of the variations there are likely to be associated with real differences in lamellar spacing. Such an inhomogeneous structure would provide hard
(a)
@I Fig. 11. Secondary electron images of (a) wear surface and (b) wear debris for PL280 rail specimen (750 N, 300 000 cycles, 5% slip, unlubricated).
166
M. Sate et al. I Rolling-sliding
I I I ( I I I , I I 1, I I I , 1 I I , I I
250
Dry 5%+dty 5% Dry 5%+dry 0% Water 5%+water 5%
. 200
-
A n
mE E_
150
-
-0 9) 100
-
0
I%
3E p
A 50
A
n
0 il,,‘,‘t”‘l’l’l’,,l’,,, 0
200
(a)
2 a E 2 3 9
400
600
600
1000
1200
Rotation cycles (x1000)
100
50
!
A
:
l
A n
Dry 5%+Water Drv 5%+Water Dry S%+Water
0% 5% 5%
0~““‘“““‘“““““’ 0
200
600
1000
1200
1000
1200
600
Rotation cycles (xl 000)
@I
0
Cc)
400
200
400
600
600
Rotation cycles (x1000)
Fig. 12. Volume loss of rail PL2M specimens (5% and/or 0% slip) vs. number of rotation cycles for different combinations of unlubricated and lubricated conditions: (a) all dry or all wet (water), 750 N, (b) dry+water, 750 N, (c) dry+water (with or without inhibitor) or dty+silicone oil, 500 N.
has been used here because the damage mechanism depends on the conditions. Only in the case of the unlubricated test with 5% slip ratio (labelled dry
of rail steels
5%+dry 5% in Fig. 12(a)) does the volume loss have the same meaning as the volume wear in Fig. 3(a). In Fig. 12(a), the data are for tests in which the environment remained constant during each test, i.e. all dry or all water. The volume loss is about five times less with water than without. Also, there is negligible continuing loss of material when the slip ratio is changed from 5% to 0% (unlubricated). The data shown in Figs. 12(b) and 12(c) resulted from combination tests: 300 000 cycles, unlubricated, 5% slip, followed by additional testing with a lubricant. Material continues to be lost at a high rate during the second stage of testing. Comparison of these results with those in Fig. 12(a) shows that volume loss during lubricated stages can be comparable with wear rates under dry conditions, provided that the lubricated stage was preceded by a dry sliding stage. Similar results were obtained for each choice of lubricant used in the second stage (water, water + inhibitor, silicone oil) (Fig. 12(c)). The micrographs of longitudinal sections for rail specimens for the all dry or all wet tests are shown in Fig. 13. They are as expected from the volume loss data shown in Fig. 12(a). For the continuously waterlubricated test, neither a deformed layer nor cracks are visible in the micrograph (Fig. 13(c)). Figures 14(a)-14(c) show longitudinal sections of rail specimens tested first under dry conditions (5% slip) followed by testing in water. The damage is dramatically different from the cases presented earlier. Although some cracks appeared during the dry stage, many more developed and propagated during the second (wet) stage. Even when no cracks appeared during the dry stage, there was much cracking during the second stage. Similar damage occurred even when the slip ratio in the second stage was 0%. Even when the rail specimen had much damage, the wheel specimen (TM360) had very few cracks. A typical example is shown in Fig. 14(d). Figure 15(a) shows that damage similar to that shown in Figs. 14(a)-14(c) was evident in dry+water (5% slip) tests even when the load was reduced from 750 N to 500 N. The other parts of Fig. 15 show that similar damage was produced when the lubricant in the second stage was changed from water to water with inhibitor or to silicone oil. Data for the depths of work hardening (defined in Section 3.1.1), visible plastic flow and fracture are presented in Figs. 16(a)-16(c) which correspond to the tests of Figs. 12(a)-12(c). Again, with no lubricant present, the depth of work hardening increased during testing with 5% slip but did not increase further during a subsequent stage of 0% slip. Also, for tests with water lubrication at all times, the depth of work hardening was small. Figure 16(a) shows that cracks remained
167
M. Sate et al. / Rolling-sliding of rail steek
(cl Fig. 13. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens (750 N). These correspond to the three data points shown in Fig. 12(a) for 600 000 total cycles: (a) dry, 5% slip+dry, 5% slip; (b) dry, 5% slip+ dry, 0% slip; (c) water, 5% slip+water, 5% slip.
within the visibly deformed material when the environment for a given test remained constant. However, the behavior is strikingly different when a second stage with a lubricant followed a dry stage with slip (Figs. 16(b) and 16(c)). In those cases cracks propagated well into the region of little or no work hardening. 3.2.2. Scanning electron microscopy observations Figure 17 shows secondary electron images for a longitudinal section of a PL280 rail specimen after testing with 5% slip, first without lubricant and then with water. The cracks penetrate very deeply, as noted in the previous section. Also, their paths are strongly influenced by microstructural features such as cementite
(cl
(4
Fig. 14. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens (750 N). These correspond to the data shown in Fig. 12(b). Also shown is one wheel specimen (TM360) for comparison. (a) Dry, 5% slip+water, 0% slip, rail specimen; (b) dry, 5% slip+water, 5% slip, rail specimen; (c) dry, 5% slip +water, 5% slip, rail specimen, extended test; (d) dry, 5% slip+water, 5% slip, wheel specimen.
lamellae, pearlite colonies and grain boundaries. This results in much branching of the cracks. There are even places such as that marked by an arrow where the cracks propagate toward the surface after first growing away from the surface. The microhardness measured near such cracks was not significantly different from microhardness of adjacent material at the same distance from the surface. 3.2.3. Effect of slip ratio The importance of slip ratio in the first (dry) stage is emphasized in Fig. 18, which shows volume loss VS.
o
1.0
-
Work hardening
0
Plastic flow
A
Crack depth
0.8 0.8
-
0.4
-
0.2
-
0
D
A
0.0
I
t
05%
I
D5%+05%
0” +
Q
05%+00%
WS%+WS%
1
Testing conditions 1.2
(b) 1 .O
[
I
I
:
0
Work hardening
.
I
Crack depth
I
0.0 t
05%
I
I
A
I D5%+WO%
I
OF&W596
f
I
1
DEi%+W5%(700)
Testing conditions 1.2
L
I
I
1.0
(4
td)
Fig. 15. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens (500 N). These correspond to the data shown in Fig. 12(c). (a) Dry, 5% siipiwater, 5% slip; (b) dry, 5% shp+water with inhibitor, 5% slip; (c) dry, 5% slip + silicone oil, 0% slip; (d) dry, 5% slip + silicone oil, 5% slip.
number of rotation cycles for combination tests. When the slip ratio was 0% in the first stage, the volume loss in both stages was very small. It was much larger when 5%, 10% or 26% slip was used in the first stage, and, as shown in Fig. 19, there was much fracture damage in the second stage (wet), even when the slip ratio in that stage was 0%. 32.4. Efect of hardness and microstructure The volume losses for pearlite and tempered martensite rail specimens after combination tests (dry, 5% slip+ water, 5% slip) are shown in Fig. 20 for different values of initial hardness. The volume losses for all of
I
I
J
I
1
A
0.8
-
0.6
-
0
-
A crackdepth
0.4
-
0.2
-
0.0
t
Work hardening
P
?
DS%+WS% D5%+W(t)5%
I
D5*LMX%
D5%+05%
Testing conditions Fig. 16. Comparisons of depth of work hardening (estimated from microhardness) and depths of visible plastic flow and cracks (both estimated fram longitudinal sections) for PL280 rail specimens. These correspond to the data shown in Figs. 12(a)-12(c). In each case the slip was 5% in the first stage. D represents dry or unlubricated and W and 0 represent lubrication with water and oil respectively. (a) All dry or all lubricated (water), 750 N; (b) d~+lub~cated (water), 0% or 5% slip, 750 N; (c) dry +Iubricated (water, water with inhibitor, or silicone oil), 0% or 5% slip, 500 N.
M. Sat0 ef al. I Rollingdiding
of rail sreeh
169
(4
Fig. 17. Secondary electron images of etched longitudinal section of PL280 rail specimen (unlubricated, 5% slip, 300 000 cycles+ lubricated, 5% slip, 300 CKJOcycles; 750 N): (a) lower magnification, showing branching of cracks, (b) higher magnification of region A, showing crack propagating toward surface.
0 0 n
IJ
Dry Dry Dty Dry
O%+Water O%, 750N S%+Water 5%, 750N lO%+Waler lo%, 750N 26%+Water O%, 500N
n
Cc)
_
0
Fig. 19. Optical micrographs of unetched and etched longitudinal sections of PL280 rail specimens after testing first without lubricant followed by continued testing with water. These correspond to the data shown in Fig. 18. (a) Dry, 0% slip+water, 0% slip, 750 N (negligible damage); (b) dry, 10% slip+water, 10% slip, 750 N (much damage); (c) dry, 26% slip+water, 0% slip, 500 N (much damage).
these cases were similar, except that the PL280 specimen showed a larger volume loss in the second stage.
4. 0
“,‘,““,,,~‘,,‘,,~,‘,,, 100 200
300
400
500
600
700
Rotation cvcles (xl 000) Fig. 18. Comparison of volume losses of PL280 rail specimens after testing first without lubricant followed by continued testing with water. Note especially the results when 0% slip was used in the first (unlubricated) stage.
Discussion
4.1. Unlubricated tests Beagley [9] has described mild (type I) and severe (type II) wear modes for rail steels. Clayton and coworkers [l&12] reported an additional wear mode (type III) in which the surface was roughened more and the debris particles were larger. Their tests involved 500
170
M. Sato et al. f Rolling-sliding
700
in a dry environment. The present work is in agreement with this observation. In contrast, Way [4] has concluded from tests on bearings that cracking only occurs when lubricant is present. The difference in behavior is not surprising in view of the marked differences in contact stress, geometry, materials, microstructure and hardness. In the present work, some cracks did appear during unlubricated testing. However, they propagated along flow lines until they were blocked or diverted by material which was not suitably aligned. That is, these cracks remained within the region of visible flow. The mechanism for initiation of cracks during unlubricated testing cannot be determined with certainty from the present work. However, local variations in the extent of surface flow have been invoked by Kjer [15] and by Oxley and co-workers [16]. In both cases, material from one part of the surface advances over adjacent surface. According to Oxley and co-workers, this creates a suitable site for initiation of fatigue cracks. A simple elastic analysis of a combined normal and shear line contact load moving past an inclined surface crack as shown in Fig. 21 does not predict either the observed crack growth direction or significant values of crack tip driving force. In the absence of a crack, the radial stress produced by a compressive normal force P and corresponding shear force Q per unit depth is given by [17]
700
a,=
o
0
100
300
400
500
600
Rotation cycles (x1000)
(a)
01,,,, 0
(b)
200
100
200
300
400
0
TM280
A
TM300
l
TM360
500
600
of rail steels
Rotation cycles (xl 000)
Fig. 20. Volume loss of rail specimens with different microstructures after testing first without lubricant (5% slip) followed by continued testing with water (5% slip): (a) pearlite: (b) tempered martensite.
MPa and 25% slip, which is close to the 525 MPa used in the present study. Therefore the present conditions with 26% slip would correspond to the conditions used by Clayton and co-workers for the tests in which the wear was described as type III. This could be the reason that the result shown in Fig. 3(c) for 26% slip is higher than the normalized results for 5% and 10% slip. Sugino et al. [13] reported better wear resistance of pearlite rail steels compared with tempered martensite in rolling contact testing. They suggested that the difference in performance was related to the greater increase in hardness of pearlite during testing compared with tempered martensite. The results of the present work are similar, except that the differences reported here are smaller. Sugino et al. also reported that wear correlated with the hardness measured after testing, in agreement with the results shown in Fig. 7. Akama and Matsuyama [14] have noted that some cracks are generated during testing of rail steels even
21 -P cos tl+Q sin 8 , u~~=O, u,=vu, -7rr (
(1)
where (r, 8, z) are polar coordinates centered on the contact line, such that fI= (0, -r/2) denotes the directions of (P, Q) respectively. Plane strain conditions are assumed along the direction of line contact (z axis), and v is Poisson’s ratio. For 8> &= tan-‘(P/Q), a,, is positive, with a maximum value of 2Q/m at 8= n/2. In the homogeneous, isotropic, elastic material modeled, cracks would tend to nucleate at the surface where the stress is largest, at the inclination 8=0. Typical crack angles 6, vary from near 90” at the surface to about 60” further into the highly deformed material. The difference between
Fig. 21. Surface crack and line contact geometry used for fracture analysis.
M, Sato et al. / Rolling-sliding of rail steels
predicted and observed crack angles may be caused by the substantial material rotation and large orientation dependence of fracture toughness in the textured material near the surface. Estimates of mode I and mode II stress intensity factors for an edge crack of length L are obtained according to [18] F(x’/L)a,,&‘/L)
d(x’/L)
(2) P(x’/L)a,,&‘/L)
d(x’/L)
(3) where @‘IL)
= 1.3 - 0.3x’lL
(4)
and a,,,(x’/L) and q,,,,(x’/L) denote normal and shear components of traction along the crack surface, from x’lL=O to xl/L= 1. This solution is applied to a surface crack at inclination 0, and position x=u relative to the contact load by using eqn. (1) to estimate the normal and tangential components of traction in eqns. (2) and (3). Accordingly, portions of the crack near enough to the surface, so that x’/L < (u/L) tan $/(tan 0, -tan Q, are predicted to be loaded in tension, and the remaining portions of the crack nearer to the tip are normally in compression. The results for a crack at &= 68” and Q/P= 0.6 are that KI increases to a maximum of about 0.02P/L”2 as the contact load approaches the crack, corresponding to a/L +O. In this same interval, K,, reaches a substantially larger value of 0.12P/L’n. As the load passes the crack, both KI and K,, reach large negative values of - 2.2PIL’R and about 0.7P/Lln respectively at alL = - 1, and rapidly decay to zero as a/L + - m. Assuming P= 0.15 MN m-’ and L =O.Ol mm, KI ,.,,== 0.95 MPa min. This is significantly lower than nominal fracture toughness values for high strength steels, but seems to be comparable with threshold fatigue values in textured steels [13]. However, this analysis does not provide a suitable explanation for continued crack growth, since K decreases with crack length. Also, it does not provide a suitable explanation for observed crack growth directions. The residual stresses produced by near-surface plasticity and large anisotropy in fracture toughness due to texturing, although not incorporated into the model, may be important features that control crack extension force and crack growth direction.
171
Sugino et al. [5] have discussed a possible interaction between wear and the extent of cracking. If wear is high enough, damaged material is removed before cracks can propagate deeply. This could explain the observation that very little cracking was observed when the slip ratio was high (see Figs. 3-5). 4.2. Tests with lubricant The worst damage was observed when two conditions were satisfied: (1) the slip ratio during the unlubricated stage was greater than zero, and (2) further testing, with or without slip, was in the presence of a liquid lubricant. This was consistently true, whether or not any cracks were observed after unlubricated testing. The density of cracks was much greater when both conditions (1) and (2) were satisfied than for condition (1) alone. Deformation during the dry stage seemed to prepare the material for extensive cracking during the lubricated stage. Whereas cracks produced during unlubricated tests remained within the region of visible flow, cracks produced with lubricant, preceded by sliding without lubricant, extended well beyond that region. This suggests that a fatigue process is involved in the latter case. It would be tempting to turn to stress-corrosion cracking to explain these results. However, similar behavior was found when a corrosion inhibitor was used with water as a lubricant and when a low viscosity silicone oil was used. In the case of the inhibitor, the results are not conclusive, because new surfaces generated by cracking and debris formation may have depleted the local concentration of inhibitor in the lubricant. However, the behavior with silicone oil was similar to that with water, despite the expected differences in chemical reactivity. This result suggests that some mechanism other than hydrogen embrittlement and stress-corrosion may be involved. Among the candidates affecting the stress intensity are Way’s pressure mechanism [4, 19, 201, a reduction in friction between crack faces [20] and surface energy effects [21].
5. Conclusions Effects of slip ratio, hardness, microstructure (pearlite and tempered martensite) and environment (dry, water and oil) on wear and rolling contact fatigue were examined. The main conclusions are summarized here. 5.1. Tests without lubrication (1) The thickness of the highly deformed layer is very small for pure rolling compared with cases involving slip. (2) Wear decreases as hardness of the rail steel increases.
172
M. Sato et al. I Rolling-sliding of rail steels
(3) For a given initial hardness, the wear rate of pearlite is slightly smaller than that of tempered martensite. (4) Wear debris are generated from near-surface material in which the cementite phase has broken into small particles. (5) Cracks initiate at the surface and propagate along lines of plastic flow, rather than normal to the local direction of maximum tensile stress. These cracks stop within the plastically deformed material. (6) Cracking is more extensive in material with low initial hardness. In such cases, tempered martensite is more resistant to cracking than pearlite. 5.2. Tests with lubricant (1) The results of rolling with or without sliding in water after dry rolling-sliding indicate that (i) many cracks appear during the second (wet) stage, (ii) cracks extend well beyond the region of visible plastic flow, suggesting that fatigue is involved, (iii) crack propagation and branching are influenced by local microstructural features, e.g, pearlite colony size and pearlite lamellar spacing and orientation, (iv) sliding is not needed in the second (lubricated) stage if sufficient deformation has occurred in the first (unlubricated) stage, and (v) extensive deformation during the dry stage seems to prepare the material for cracking during subsequent operation in water. (2) Behavior is similar for all fluids used as lubricants (water, water plus inhibitor, silicone oil).
Acknowledgments
The authors are pleased to acknowledge the support and encouragement of the Bethlehem Steel Corporation, which supplied the test machine, and the Nippon Steel Corporation, which supplied the test specimens. The following were also helpful in discussions or with experimental aspects of the research: W. A. Glaeser, A. R. Rosenfield, R. Farrar, J. Kocka, J. Ellis and C. MacDonald.
References 1 L&ton heavy axle load test: preliminary results come in, Railw. Track Struct., 85 (July 1989) 29-30.
2 R. K. Steele and D. H. Stone, Developments in railroad rail, Am. Railw. Eng. Assoc., Proc., 87 (1986) 311-358. 3 R. K. Steele, Recent North American experience with shelling in railroad rails (part I: observations), Am. Railw. Eng. ASSOC., Proc., 90 (1989) 311-345. 4 S. Way, Pitting due to rolling contact, J. Appl. Mech., 2 (1935) A49-A58. 5 K. Sugino, H. Kageyama and M. Sato, Effect of wheel and rail profiles on wear and fatigue behavior of rail, Proc. 2nd Int. Symp. on Contact Mechanics and Wear of Rail-Wheel Systems, 1986, University of Waterloo Press, Waterloo, Ont., 1987, pp. 435450. 6 M. Sato, K. Sugino, K. Tanikawa and H. Iida, The nature of lubricants and their influence on wear and fatigue behavior of rail, Proc. 2nd Int. Symp. on Wheel/Rail Lubrication, Memphis, TN, 1987, Association of American Railroads. 7 J. P. Sheehan and M. A. H. Howes, The effect of case carbon content and heat treatment on the pitting fatigue of 8620 steel, SAE Trans., 81 (1972) 1031-1045. 8 P. Heilmann, W. A. T. Clark and D. A. Rigney, Orientation determination of subsurface cells generated by sliding, Acta Metall., 31 (1983) 1293-1305. 9 T. M. Beagley, Severe wear of rolling-sliding contacts, Wear, 36 (1976) 317-335. 10 P. J. Bolton and P. Clayton, Rolling-sliding wear damage in rail and tyre steels, Wear, 93 (1984) 145-165. 11 D. Danks and P. Clayton, Comparison of wear processes for eutectoid rail steels: field and laboratory tests, Wear, 120 (1987) 233-250. 12 P. Clayton and D. Danks, Effect of interlamellar spacing on the wear resistance of eutectoid steels under rolling-sliding conditions, Wear, I35 (1990) 369-389. 13 K. Sugino, H. Kageyama, T. Urashima, S. Nishida, H. Masumoto and M. Hattori, Basic properties of high strength rail steels and new rail development, Nippon Steel Tech. Rep., I6 (1980) 103-119. 14 M. Akama and S. Matsuyama, Wear characteristics of wheel and rail steels, J. Jpn. Sot. Lubr. Eng., Znt. Ed., 8 (1987) 75-80. 15 T. Kjer, A lamination wear mechanism based on plastic waves, Proc. Int. Con& on Wear of Materials, American Society of Mechanical Engineers, New York, 1987, pp, 191-198. 16 B. S. Hockenhull, E. M. Kopalinsky, and P. L. B. Oxley, An investigation of the role of low cycle fatigue in producing surface damage in sliding metallic friction, Wear, I48 (1991) 135-146. 17 K. L. Johnson, ContactMechanics,Cambridge University Press, Cambridge, 1985, p. 16. 18 H. Tada, P. C. Paris and G. R. Irwin, The Stress Ann&s of Cracks Handbook, Del Research, Hellertown, PA, 1973, Section 8.3. 19 M. Kaneta and Y. Murakami, Effects of oil hydraulic pressure on surface crack growth in rolling/sliding contact. Tribal. Int., 20 (1987) 210-217. 20 A. F. Bower, The influence of crack face friction and trapped fluid on surface initiated rolling contact fatigue cracks, J. Lubr. Technol., II0 (1988) 704-711. due to surface 21 N. J. Petch, The lowering of fracture-stress adsorption, Philos. Mag., I (1956) 331-337.