An experimental procedure for surface damage assessment in railway wheel and rail steels

Wear 342-343 (2015) 22–32

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An experimental procedure for surface damage assessment in railway wheel and rail steels A. Mazzù a,n, L. Solazzi a, M. Lancini a, C. Petrogalli a, A. Ghidini b, M. Faccoli a a b

Università degli Studi di Brescia-Dipartimento di Ingegneria Meccanica e Industriale, Via Branze 38, 25123 Brescia, Italy Lucchini RS, Via G. Paglia, 45-24065 Lovere, BG, Italy

art ic l e i nf o

a b s t r a c t

Article history: Received 8 April 2015 Received in revised form 31 July 2015 Accepted 10 August 2015 Available online 18 August 2015

The assessment of damage in rail–wheel cyclic contact requires considering the combined action of different damage mechanisms, such as wear, ratcheting, surface or subsurface crack nucleation and propagation. Models, usually requiring experimental calibration, are available for assessing these phenomena. The best way to calibrate them is based on cyclic contact tests, as these represent the real working conditions more closely. However, some experimental information, such as microstructural changes or crack paths, can be obtained only by destructive methods at the end of the tests, and their evolution cannot be monitored; other parameters, such as the wear rate, can only be determined during time consuming breaks to the tests. In this work, non destructive measurements of vibrations, torque and Barkhausen noise were introduced as indicators of damage evolution in cyclic contact tests on a high performance steel for railway wheels, coupled with a rail steel. In particular, their correlation with surface state, wear rate, subsurface microstructure and presence of cracks was shown. & 2015 Elsevier B.V. All rights reserved.

Keywords: Wheel and rail steel Rolling contact fatigue Wear Non destructive testing Barkhausen noise Vibrations

1. Introduction The damage assessment in wheel–rail contact is a very complex problem, as is affected by many parameters (materials, load history, climatic conditions and other) and different damage mechanisms can occur. Wear, cyclic plasticity, surface origin rolling contact fatigue (RCF) and subsurface origin RCF are the key damage mechanisms occurring on wheels and rails [1,2] that can interact and influence each other, even in competition. Wear, for instance, influences the stress–strain history of the material by removing material layers from the surface; strain hardening can occur due to cyclic plasticity, therefore affecting the material behaviour with respect to wear, fatigue and defect sensitivity. Furthermore, surface cracks nucleated by ratcheting (i.e. cyclic accumulation of unidirectional plastic strain) can be initiation sites for RCF damage [3–5]. Several experimental techniques have been used over the years for studying surface damage of railway wheel and rail steels. Field measurements were collected in the 1950s [6,7], followed by simulated field experiments carried out on specially built test tracks [8]. The laboratory methods have always been used: they included full-scale laboratory experiments [9], scale-down tests [10,11] and bench tests by twin-disc machines [12–14]. Small-scale n

Corresponding author. Tel.: þ 39 030 371 5525; fax: þ39 030 3702448. E-mail address: [email protected] (A. Mazzù).

http://dx.doi.org/10.1016/j.wear.2015.08.006 0043-1648/& 2015 Elsevier B.V. All rights reserved.

laboratory tests allow careful control of the tribo-system whereby the effects of many variables on damage phenomenon can be isolated and determined. They had proved to be good screening techniques for wheels and rails performance even if it is not possible to predict the overall wheel and rail performance at the laboratory level, due to the complexity of the full-scale train dynamics. However, the data obtained in laboratory tests under controlled conditions may help in interpreting the results of field trials, as recently shown in Refs. [15,16] for a rail steel. Besides the experimental techniques, many models were elaborated based on interpretation of experimental results and damage predictions. Some models focus on a single damage mechanism, such as wear [17,18], cyclic plasticity [19,20] or surface crack propagation [21]. Other models integrate two or more damage mechanisms, that can occur simultaneously and in competition [22,23]. Recently, in [24] an integrated model for surface and subsurface damage assessment of railway wheel steels was proposed, considering wear, cyclic plasticity, surface and subsurface crack propagation as concurring and interacting damage phenomena. In particular, the model is able to simulate the evolution of the wear rate, the plastic strain field and the surface crack shape taking into account influencing phenomena such as strain hardening and stress–strain redistribution due to wear. This model can be used either for failure prediction or for material characterization, by the simulation of laboratory rolling and sliding contact tests in controlled conditions. In [24] an example of application for material

A. Mazzù et al. / Wear 342-343 (2015) 22–32

characterization was shown: the wear model was calibrated by experimental wear curves obtained by means of periodical specimen weighing, whereas the cyclic plasticity and the fatigue models were calibrated by subsurface microstructure and crack observations on selected specimens following the experimental testing. However, following this procedure microstructural changes and cracks can only be detected at the test end, whereas they evolve during the tests. An experimental direct verification of damage evolution can only be obtained by a large number of specimens, each one to be tested at different cycle number and destroyed at the end, as shown in Ref. [25]. On the other hand, even the wear rate is not constant, therefore an accurate characterization would require frequent test stops for wear measurements. In this scenario, non destructive techniques (NDT) for getting information about damage evolution can help to speed up the test procedure and obtain additional information along the test duration. In this paper, testing machine vibrations and Barkhausen noise measurements were introduced as NDT for monitoring rolling contact tests executed by a bi-disc machine. The Barkhausen noise use has progressively increased in several industrial fields [26], mainly for the quality control and the evaluation of residual stresses in mechanical components. The Barkhausen effect is a series of sudden changes in the size and orientation of ferromagnetic domains, or microscopic clusters of aligned atomic magnets (spins), that occurs during a continuous process of magnetization or demagnetization within a ferromagnetic material. The magnetization process is not continuous, but is made up of small, abrupt steps caused when the magnetic domains move under an applied magnetic field. When the electrical pulses produced by all domain movements are added together, the noise-like signal called Barkhausen noise is generated. The Barkhausen noise is sensitive to the stress state, which influences the domain arrangement through the magneto-strictive effect, and to micro-structural features such as dislocations network, grain boundaries density, second phase precipitates, texture and cavities [27,28], which act as pinning sites for domain walls. Therefore, the Barkhausen noise measurement is used for investigating material properties correlated to the microstructure and stress state, such as hardness, plastic strain, anisotropy and residual stress [29–31]. In particular, high hardness, being related to finer grain microstructure, tends to lower the Barkhausen noise because of higher boundary density, as well as compressive residual stress field. Considering these properties, the Barkhausen noise method is of particular interest when applied in differential form, i.e. measuring changes with respect to a reference, in order to detect material properties variation from point to point and with time. The main use of this technique is to assess a localized damage, as in the case of grinding burns or decarburized zones, or to assess a progressive material degradation, like that induced by fatigue [32,33]. In this paper it was investigated as a tool for detecting microstructure changes due to cyclic plastic strain in rolling–sliding contact, in particular for predicting material damage due to ratcheting. Vibration measurements are already used as a diagnostic tool for detecting machine components failure [34–36]; in this paper, the vibrational data were handled in a more advanced way, aimed at finding a proportional correlation with damage parameters such as the surface state or the wear rate. Furthermore, their usefulness in revealing hidden failure in specimens was studied. Both measurements can be influenced by various parameters, therefore a procedure for filtration and crosschecking with other experimental data, such as the surface appearance and the torque on the specimen shaft, was elaborated to separate the machine effects from the specimen ones.

23

2. Materials and experimental procedure Rolling contact tests were carried out on cylindrical discs of a railway wheel steel (SUPERLOSs) against discs of a rail steel (UIC 900A). The discs diameter was 60 mm, their thickness was 15 mm. The wheel steel discs were directly obtained from a wheel rim with their axis orthogonal to the wheel tread, the rail steel discs were machined out of a rail head with their axis orthogonal to the rail long axis. All discs were machined by grinding in the same conditions, therefore it was assumed that they had the same contact surface finish. All samples were supplied by the Italian railway components manufacturer Lucchini RS. The chemical composition of the steels is shown in Table 1; their mechanical properties, referred to wheel rims and rail heads [37], are reported in Table 2. The specimens were tested on the two-discs testing machine described in detail in Ref. [25], mounted onto independent shafts, whose one can displace orthogonally acted by a hydraulic cylinder. The rolling speed of each specimen shaft was measured by encoders, the contact load by a load cell located at the piston head. Piezo-accelerometers, whose sensitivity was 0.98 V/(m/s2) and linear bandwidth in the 5 Hz–20 kHz range, were positioned on both mandrel supports, normal to the rotation axis, in the vertical and horizontal planes respectively, as shown in Fig. 1. The two accelerometers, as well as a torque sensor positioned on the displaceable shaft, were acquired by means of a configurable data acquisition system at 5 kHz synchronous sampling frequency. The measured torque signals were also elaborated in order to obtain the friction coefficient between the specimens, by means of a procedure aimed at excluding the contribution of the shaft supporting bearings. During the first 300 cycles, at the beginning of each test, the two discs were pressed one against the other and rotated with no relative tangential velocity, by means of a rolling speed control. The average toque value Ti during this phase was recorded and assumed to be only due to the bearings and the discs’ inertia. After that the proper test phase begins and relative rotation between discs is allowed according to the test specifics, therefore the friction coefficient f is computed by subtracting the Ti value to the measured torque T and dividing the result by the disc radius R and the applied radial force Fr, as shown in Eq. (1).

f=

T − Ti RFr

(1)

Using the acquired data, the transfer function HT/x(ω) between torque and acceleration was computed averaging 5 windows of 1 s, thus producing a spectrum of 1000 lines every 5 s. To detect out changes in the system, the difference DT/x(ω) between the values of the transfer function associated with the current status and a reference transfer function, associated with the initial status, was computed for each frequency resolution ωi. A synthetic index DT/x was proposed, computed as the root mean square summation of the difference DT/x(ω) for all frequencies bins, as shown in Eq. (2):

DT / x =

1 n

n

∑ ( HT / x ( ωi ) − HTref/ x ( ωi ) )( HT / x ( ωi ) − HTref/ x ( ωi ) )*

(2)

i=1

Table 1 Wheel and rail steels chemical composition (wt%). Steel SUPERLOS 900A

s

C

Mn

Cr

Ni

Mo

Si

0.5 0.71

0.93 1.07

0.13 0.03

0.09 0.02

0.02 0.01

0.89 0.27

24

A. Mazzù et al. / Wear 342-343 (2015) 22–32

The reference values HTref/ x (ω), related to the initial dynamic behaviour of the system, were computed for each test by averaging the HT/x(ω) values of the first 2500 cycles. Though all these quantities are related to mechanical properties of the system, these could also be used as damage indicators only, as explained in previous work [38], since connexion to any proper mechanical parameter would be overly complex and beyond the scope of this work. The rolling contact tests were carried out under the working conditions summarized in Table 3. All tests were carried out at 16 °C controlled room temperature. The Hertz contact pressure P0 and sliding ratio s were chosen as mean typical values in real wheels. The tests in dry condition were carried out with a 30 m/s cooling air flow at room temperature directed towards the contact region, with the wheel specimen mounted as driver in order to approach the condition of a power wheelset during traction. The air flow also had the function of cleaning the contact region from any wear debris avoiding third body effects. The Dry 3 test was limited at 3  104 cycles for investigating any subsurface phenomena occurring in the early cycles. The tests in wet condition were carried out with a 6  10  6 m3/s flow of water (added with 10% glycol for protection against corrosion) at room temperature directed towards the mesh inlet, with the wheel specimen as follower. The test in mixed condition was carried out with an initial period of 105 cycles in dry condition (with the wheel specimen as driver) and a subsequent period in wet condition (with the wheel specimen as follower) until severe damage occurrence. Given the water viscosity properties, no complete separation between the contact surfaces of the specimens was expected. The tests in wet and mixed contact were carried out to simulate the effect of rain at the wheel–rail interface: in this condition, as shown in Refs. [39,40], severe shelling can occur by pressurization of the fluid entrapped inside surface cracks. This effect occurs if the crack mouth first enters the contact zone being closed by the load and subsequently the entrapped fluid is pressurized enhancing crack propagation. On the contrary, if the crack tip region is compressed by the load before the crack mouth is closed, the fluid is squeezed out. In mixed tests the fluid pressurization effect is amplified, as the dry phase is expected to accelerate the nucleation Table 2 Wheel and rail steels mechanical properties. Steel

Ultimate strength Rm (MPa)

SUPERLOSs 980 900A 930

Monotonic yield stress Ry (MPa)

Cyclic yield stress Yc (MPa)

Brinell Percentage elongation A hardness (HB) (%)

640 470

525 390

18 14

280 296

of surface cracks able to propagate in the subsequent wet phase. As it is expected that surface micro-cracks are generated by sliding friction and are oriented along the friction direction, the specimen where surface cracks can propagate by fluid pressurization is expected to be the follower. As the steel under investigation was mainly the wheel one, this was the reason for mounting the wheel specimen as follower in wet tests and in the wet phase of mixed tests. However, in purely wet tests the orientation of initial surface cracks can be casually distributed and fatigue can occur even at the driver. Fig. 2 shows the shakedown map for cylindrical contact calculated according to Ref. [20], with a line representing the applied maximum pressure P0 in these experimental tests for rail and wheel specimens, normalized by the cyclic yield shear stress Y k = c . Given the level of P0/k , for the wheel specimen elastic 3

shakedown or ratcheting is expected depending on the friction coefficient. In particular, the threshold friction coefficient fthW ¼0.28 can be determined: if it is exceeded, ratcheting is predicted; otherwise, the material is expected to turn back to elastic behaviour after the initial transient phase where plastic deformation occurs. For the rail specimen, the threshold friction coefficient fthR ¼0.11 can be determined: in this case, if it's exceeded ratcheting is predicted again; otherwise, the material is expected to undergo plastic shakedown, i.e. closed loop plastic strain without cyclic deformation accumulation. The testing procedure included periodical stops for contact surface image acquisition, ultrasonic cleaning, weighing and Barkhausen noise measurement. The contact surface images were realized by means of a high-resolution (1 mm) one-pixel-line camera, with the exception of the longer wet test where the images were obtained by a digital high resolution camera. The specimens were weighed by a precision balance with a resolution of 0.01 g, in order to evaluate the weight loss. The Barkhausen noise was measured by the Rollscan 100-2 instrumentation with an AST-S1-138 miniature sensor. The magnetic field excitation frequency was 120 Hz, the exciting field magnitude was 7.2 Vpp, Table 3 Testing conditions. Test ID

Lubrication condition

Duration

Contact pressure P0 (MPa)

Mean rolling speed (r.p. m.)

Sliding/ rolling ratio (%)

Dry 1 Dry 2 Dry 3 Wet 1 Wet 2 Dry–wet

Dry Dry Dry Wet Wet Mixed

8  105 8  105 3  104 1  106 2  106 1.53  105

1100

500

1

Fig. 1. Coupled rail–wheel specimens and simplified FRF based model.

A. Mazzù et al. / Wear 342-343 (2015) 22–32

the output signal amplification rate was 4, and it was analyzed in the frequency range 70–200 kHz. The sensor detected the signal over an 11  16 mm2, at 0.02 mm below the contact surface. The results of the measurements were taken in terms of root mean square of the Barkhausen noise, also known as magneto-elastic parameter and indicated as MP. The measurements were made on four points of the contact surface and averaged.

6 5

P0 /k

4

Dry 1 (wheel)

Dry 1 (rail)

Dry 2 (wheel)

Dry 2 (rail)

Rail Plastic shakedown

Ratcheting Wheel

Elastic shakedown

3

25

At the end of each test the discs were cut along the mid plane orthogonal to the contact surface. Vickers hardness tests were carried out with a 1000 g load on the polished sections of each sample at varying depth, complying with ASTM E384 (2011) Standard. Afterwards, the polished sections were etched by a specific etchant (2% Nital) for making the microstructure visible. The samples were examined by a Leica DMI 5000M light optical microscope and by a LEO EVO-40XVP scanning electron microscope (SEM) with a Link Analytical eXL microprobe, in order to observe the deformation patterns and the cracks morphology. A few surface residual stress measurements were carried out on some specimens in order to find correlations with the Barkhausen noise; the X-Ray Diffraction (X-RD) technique was used, by means of a TNX Enixe diffractometer, complying with ASTM E2860 (2012) standard.

2 3. Experimental results

1 3.1. Damage analysis

0 0

0.11

0.2

0.28

0.4

0.6

0.8

1

f Fig. 2. Shakedown map for cylindrical contact [20], with the applied load condition in the present tests.

Fig. 3 shows the evolution of the contact surface of the wheel steel disc during one of the RCF tests in dry condition (Dry 1). In the early 2  104 cycles the surface appearance did not change significantly (Fig. 3b); after 2  105 cycles significant flaking appeared on the surface (Fig. 3c) that in the following cycles was

Fig. 3. Contact surface appearance at increasing cycle number of the wheel steel disc in the Dry 1 test: (a) 0 cycles; (b) 2  104 cycles; (c) 2  105 cycles; (d) 4  105 cycles; (e) 6  105 cycles; (f) 8  105 cycles (end of test).

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A. Mazzù et al. / Wear 342-343 (2015) 22–32

progressively smoothed until the end of the test (Fig. 3f). The same results were obtained in the Dry 2 test. Fig. 4 shows the contact surface evolution of the wheel specimen in the longest wet test (Wet 2). The contact surface appeared less damaged at the test end with respect to the discs tested in dry condition: only some shallow pits were visible during the test. Some oxidation bands appeared periodically (Fig. 4a and c), that changed the surface colour but not its smoothness. The origin of these bands is likely related to contact dynamics, i.e. small friction and pressure variations due to alternation of stick and slip phases and to specimen elasticity; this phenomenon occasionally occurs during the tests, without substantially altering the overall damage evolution. The lower surface damage is due to the lower friction coefficient than in dry condition, as discussed more in detail in the next section. Fig. 5 shows the appearance of the rail specimen after 1.6  106 cycles and 1.8  106 cycles in the Wet 2 test: as evident, it was severely damaged; consequently, it was changed after 1.8  106 cycles for continuing the test. This shows that in absence of a dry phase leading to the formation of surface cracks with controlled orientation, surface fatigue can occur even at the driver, due to maybe lower strength or casual distribution of surface defects.

Fig. 6 shows the wheel disc tested in mixed condition. After the dry phase, surface damage similar as in the Dry 1 and Dry 2 tests occurred, including roughening and pitting. In the subsequent wet phase, severe shelling early occurred and the test was stopped after 1.53  105 cycles. As the rail specimen is concerned, similar surface damage as in the wheel one occurred in the initial dry phase, but in the subsequent wet one it was not significantly further damaged. This confirms that the previously described fluid pressurization effect occurred, rapidly leading to heavy surface deterioration of the follower disc. Fig. 7 shows the etched sections of some wheel steel discs at the end of the tests. The ferritic–pearlitic microstructure under the contact surface of the discs tested in dry condition for 8  105 cycles appeared plastically deformed, as shown in Fig. 7a. A layer with unidirectional plastic flow which tends to be aligned towards the direction of the surface friction can be observed. The disc tested in dry condition for 3  104 cycles (Dry 3) did not show any significant microstructure change. The discs tested in wet condition did not show a visibly deformed subsurface layer as well (Fig. 7b). The disc tested in mixed condition showed a thin deformed surface layer, which was produced during the initial dry phase of the test. This layer appeared discontinuous, because it

Fig. 4. Contact surface appearance at increasing cycle number of the wheel steel disc in the Wet 2 test: (a) 1  104 cycles; (b) 8  105 cycles; (c) 2  106 cycles (test end).

Fig. 5. Contact surface appearance at increasing cycle number of the rail steel disc in the Wet 2 test: (a) 1.6  106 cycles; (b) 1.8  106 cycles.

A. Mazzù et al. / Wear 342-343 (2015) 22–32

Fig. 6. Contact surface appearance at increasing cycle number of the wheel steel disc tested in the Dry-Wet test: (a) after 2  104dry cycles; (b) after 105dry cycles; (c) after 105 dry cycles and 5.3  104 wet cycles.

was partially removed by the material detachment resulting by the propagation of surface cracks, as visible in Fig. 7c. Fig. 8 shows the crack paths in the wheel steel discs in the three test conditions. In the discs tested in dry condition (Fig. 8a) only surface cracks were observed; they started with a shallow angle to the surface and followed the plastically deformed material during their growth. This indicates that ratcheting was the main damage mechanism. The very low sulphur content in the wheel steel (0.001%) limited the presence of manganese sulphide inclusions, thus reducing the risk of subsurface cracks nucleation. No surface or subsurface cracks were observed on the sample section of the Wet 1 test after 1 million cycles and of the Dry 3 test after 3  104 cycles. A few subsurface cracks (Fig. 8b) and surface cracks were found in the disc of the Wet 2 test after 2 million cycles. In the disc tested in mixed condition (Fig. 8c), surface cracks with the same morphology as in the dry tests started; subsequently, in the wet phase they propagated to a much higher depth than in the other cases, involving a much thicker layer than the plastic deformation depth (see also Fig. 7c). As rail specimens are concerned, in dry tests the deformation occurred at the subsurface was similar to that observed in the wheel, but it involved a thicker layer, due to lower cyclic yield stress. The specimen tested in mixed condition had similar appearance as those tested in dry condition. In the Wet 2 test (Fig. 9), in addition to the deep shelling already seen on the contact surface, some amount of visible plastic strain involving a

27

Fig. 7. Microstructure of the wheel specimen under the contact surface of the wheel steel disc at the test end in: (a) Dry 1 test (8  105 cycles); (b) Wet 2 test (2  106 cycles); (c) Dry–wet test (1.53  105 cycles).

depth of about 70 μm was observed. This is in agreement with the shakedown map shown in Fig. 2, which predicts plastic strain for the rail specimen even with low friction coefficient. 3.2. Hardness and residual stress The hardness variation at increasing distance from the surface measured at the test end on the wheel and rail discs sections is shown in Fig. 10. Significant hardening can be seen in the wheel discs tested in dry condition for 8  105 cycles (Dry 1 and Dry 2): this result is consistent with the deformation pattern metallographically observed (Fig. 7a). A little increase of hardness was observed in the discs tested in wet condition, in mixed condition and in the “short” dry test (Dry 3) due to the very low plastic deformation of the steel. As the rail steel is concerned, all specimens showed surface hardening: this is in agreement with the fact that the elastic shakedown limit is exceeded in all tests, as shown in Fig. 2, therefore cyclic plastic strain is attended. Many authors [41,42] observed similar hardness profiles studying rail steels and found that the steel hardening rate was higher during the initial stages of the test and decreased with number of rolling cycles until it became zero. The available measurements of residual stress in the tangential direction are reported in Table 4. These show that initial tension due to machining is present in the specimens, rapidly changing to

28

A. Mazzù et al. / Wear 342-343 (2015) 22–32

500

Dry 1 Dry 2 Dry 3 Wet 1 Wet 2 Dry + Wet

450

HV1

400 350 300 250 200

0

1

2

3

Depth (mm) 500

Dry 1 Dry 2 Dry 3 Wet 1 Wet 2 Dry + Wet

450

HV1

400 350 300 250 200

0

1

2

3

Depth (mm) Fig. 10. Vickers hardness profiles on the cross-section of the samples; (a) wheel steel; (b) rail steel.

Fig. 8. Cracks in wheel discs at the test end in: (a) Dry 1 condition (8  105 cycles); (b) Wet 2 condition (2  106 cycles); (c) Dry–wet condition (1.53  105 cycles).

Table 4 Measured residual stress. Test ID

Number of cycles

Residual stress (wheel) [MPa]

Residual stress (rail) [MPa]

New specimen Dry 1 Dry 2 Dry 3 Wet 1

0 8  105 8  105 3  104 106

þ 462  342  345  269  242

þ165  213  377  186  159

3.3. Wear and friction

Fig. 9. Microstructure of the rail discs in the Wet 2 test (1.8  106 cycles).

compression due to the applied load, as the datum of the Dry 3 test shows. Given the low cycle number of the Dry 3 test and the low friction coefficient at this stage (see next section), this condition can be supposed representative of what happens in the initial cycles of the wet test as well. In wet tests, the final residual stress is close to the Dry 3 value, meaning that after the initial change from tension to compression it does not further change significantly. On the contrary, in dry tests the residual compression at the test end is much higher.

Fig. 11 shows the weight loss of the wheel and rail steel discs as a function of cycle number; no data are shown for the Dry 3 tests due to its short duration with respect to the other ones. In dry condition an initial run-in phase of 5  104  1  105 cycles, characterized by very low wear rate, can be recognized in both tests. After this phase, the weight loss profile follows similar curves in both tests: an approximately linear behaviour up to 4  105 cycles, followed by a phase of decreasing weight loss rate until the end of the test. The mixed test in the dry phase had the same behaviour as the dry tests (as expected), whereas in the wet phase dramatic weight loss occurred at the wheel disc due to surface shelling: the dramatic increase of the weight loss in the wet phase allows evaluating the rapidity and severity of the fluid pressurization effect on pre-cracked discs. On the other hand, in both wet tests the weight loss was much lower than in dry tests for both wheel and rail steels, and kept approximately linear for the whole test duration.

A. Mazzù et al. / Wear 342-343 (2015) 22–32

14

0.5

Dry 1

0.45

Dry 2

0.4

10

0.35

8

0.25

ratcheting threshold (wheel)

0.2

6

0.15

4 2 0

Dry-Wet

0.3 f

Weight loss (g)

12

29

0

0.2

Dry 1-Wheel

Dry 1-Rail

Dry 2-Wheel

Dry 2-Rail

Dry-Wet-Wheel

Dry-Wet-Rail

0.4 0.6 Millions of cycles

ratcheting threshold (rail)

0.1 0.05 0

0

0.2

0.8

0.4 Millions of cycles

0.6

0.8

0.5

14

0.45

0.3

8 6

Wet 1-Wheel

Wet 1-Rail

Wet 2-Wheel

Wet 2-Rail

f

Weight loss (g)

Wet 2

0.35

10

ratcheting threshold (wheel)

0.25 0.2 0.15 0.1

4

ratcheting threshold (rail)

0.05 0

2 0

Wet 1

0.4

12

0

0.5

1 1.5 Millions of cycles

2

0

0.4

0.8 1.2 Millions of cycles

1.6

2

Fig. 12. Friction coefficient between the discs tested in: (a) dry and mixed condition; (b) wet condition.

Fig. 11. Weight loss of the wheel and rail steel discs tested in: (a) dry and mixed condition; (b) wet condition.

Fig. 12 shows the friction coefficient behaviour in the tests, obtained considering the measured torque time history, the nominal applied load and the specimens geometry. No data are shown for the Dry 3 test due to its short duration. For the wheel specimens, in dry tests and in the dry phase of the mixed tests the friction coefficient value kept below the elastic shakedown limit (i.e. the ratcheting threshold) in the early 2  104–7  104 cycles, subsequently raising up to about 0.45 within the early 1  105 cycles. In the following cycles, in dry tests, it gradually decayed down to 0.2–0.25, falling below the elastic shakedown limit after 5.5  105–6  105 cycles. In wet tests the friction coefficient fluctuated between 0.1 and 0.25, showing a gradual evident increment in the last phase of the longest wet test (Wet 2). Considering the overall results, in dry tests the material worked mostly in the ratcheting region, whereas in wet tests the elastic shakedown limit never was exceeded. This is coherent with the damage experimentally observed. For the rail specimens, after a very early phase, the ratcheting threshold was exceeded in most part of the tests, by a much higher ratio in dry than in wet condition. This is coherent with the strain hardening and ratcheting experimentally observed in rail specimens even in wet tests. 3.4. Vibrations Fig. 13 shows the DT/x index (as usual, the Dry 3 test was excluded). The test Wet 1 recording was suspended since 6  105–8  105 cycles approximately, due to a system malfunction, hence the data relative to this cycle interval are missing. In

general, in dry tests this parameter, which quantifies the dynamic behaviour changes with respect to the initial condition, was much higher than in wet tests. In dry tests, after an initial transitory of about 3  104–7  104 cycles, DT/x rose up with the same behaviour as the friction coefficient. After about 5.5  105–6  105 cycles, i.e. when the friction coefficient fell below the ratcheting threshold, DT/x gradually decreased. In the mixed test DT/x had a peak at about 1.15  105 cycles, related to the shelling occurrence. In wet tests DT/x fluctuated at low levels; however, in the longer test (Wet 2) DT/x increased and became instable after 1.5  106 cycles. 3.5. Barkhausen noise Fig. 14 shows the average values of the Barkhausen noise during the contact tests in all conditions; the results for the wheel specimen in mixed test condition are shown only up to 105 cycles, because after the subsequent wet phase the contact surface was too severely damaged for a consistent Barkhausen noise measurement. For the same reason the rail data in the Wet 2 condition are limited at 1.5  106cycles. The data referring to the Dry 3 tests are not shown due to its short duration, but they are consistent with the other dry tests. The measurements standard deviation was lower than 20% of the mean value, decreasing as far as the number of cycles increased due to increased material state uniformity. In dry and mixed tests, in the early 105 cycles the value of Barkhausen noise showed wide scatter. In dry tests, in the following cycles the Barkhausen noise gradually decreased, stabilising after approximately 4  105cycles at about 25–30 MP for the wheel specimens and about 50–60 MP for the rail one, with lower

A. Mazzù et al. / Wear 342-343 (2015) 22–32

8

Dry 1 Dry 2 Dry-Wet

7

DT/x (kg m)

6 5 4 3 2

180 Barkhausen noise (MP)

30

1 0

0.2

0.4 Millions of cycles

0.6

140

Dry 2 (wheel)

Dry 2 (rail)

120

Dry-Wet (wheel)

Dry-Wet (rail)

100 80 60 40 0

0.8

Wet 1 Wet 2

7

0

0.2

0.4 Millions of cycles

0.6

0.8

160

6 DT/x (kg m)

Dry 1 (rail)

180

8

5 4 3 2 1 0

Dry 1 (wheel)

20

Barkhausen noise (MP)

0

160

140 120 100 80 60 40 20 0

0

0.4

0.8 1.2 Millions of cycles

1.6

2

Fig. 13. DT/x parameter in tests in: (a) dry and mixed condition; (b) wet condition.

scatter. In wet tests the behaviour was much less stable than in dry ones, and a significant difference between the mean values of the two tests is evident. However, the Barkhausen noise values in wet tests varied within a band (50–100 MP for the wheel and 110– 150 MP for the rail) clearly higher than the stabilised value in dry tests, therefore allowing univocal association of the Barkhausen noise values measured after 4  105 cycles to wet or dry working condition.

4. Discussion The sections of the specimens of Figs. 7a and 8a show that the main damage phenomenon occurred during the dry tests was ratcheting, leading to intense unidirectional plastic strain, surface crack formation and strain hardening, as confirmed by the Vickers hardness profiles of Fig. 10. In mixed tests, the cracks nucleated in the initial dry phase were initiation sites for the shelling occurred in the subsequent wet phase, which rapidly led to severe damage by means of pressurization of the fluid entrapped inside the cracks (Figs. 7a and 8a). In wet condition the main damage was RCF, as surface and subsurface cracks in absence of overall plastic strain appeared (Fig. 7b); however this phenomenon was observed only in the longer test, meaning that it occurred later than the duration of the shorter test (1 million cycles). These damages are produced by mechanisms evolving during the test, occurring mainly at the subsurface region. The destructive methods such as the microscopy on the specimen sections, even though they give important information about the damage mechanism typology and the extension, cannot give complete

0

Wet 1 (wheel)

Wet 1 (rail)

Wet 2 (wheel)

Wet 2 (rail)

0.4

0.8 1.2 Millions of cycles

1.6

2

Fig. 14. Barkhausen noise values vs. number of cycles in: (a) dry and mixed tests; (b) wet tests.

information about the transient phase leading to the damage appearance as observed at the end of the test. The analysis of the monitored parameters (friction coefficient, vibrations, Barkausen noise, wear) allows interesting correlations between themselves and with the observed damage, so that the damage history can be reconstructed a posteriori. In dry tests three main phases can be identified. The first phase took about 3  104–1  105 cycles, and was characterized by low wear, low friction and low vibration index, whereas the Barkhausen noise increased and the surface residual stress changed from tension to compression. It was hypothesized that in this phase the material state was not significantly altered (unless the residual stress switch), as confirmed by the Dry 3 test. The increment of the Barkhausen noise can be explained by the reduction of the surface roughness due to peak smoothing, which improved the effective contact area between the probe and the specimen; however, this hypothesis could not be proven. The second phase followed, starting with significant changes occurred in the contact working conditions: the friction coefficient rapidly rose up to its maximum value and wear as well increased reaching a constant rate. Correspondingly, the DT/x parameter dramatically increased, whereas the Barkhausen noise decreased; simultaneously, the contact surface became rougher. After this rapid change, the friction coefficient started to gradually decrease until it fell beneath the elastic shakedown limit of the wheel steel, after about 5.5  105–6  105 total cycles; coherently, also the contact surface of both steels appeared smoother as far as the test went on, as documented in Fig. 3. On the other hand, in this phase the DT/x parameter kept oscillating within a relatively high constant band, whereas the wear rate kept constant until 4  105

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cycles. However, the wear measurement at 5  105 cycles is missing because a tests stop was not scheduled at that time, but it is reasonable to suppose that wear rate could keep constant as far as the friction coefficient was higher than the wheel ratcheting threshold. Furthermore, this suggests that the vibration profile can be used to reconstruct the wear curve in the intervals between two consecutive measurements. These evidences can be correlated with ratcheting: after the initial phase, the contact surfaces may have lost some smoothness due to wear, increasing both friction coefficient and wear rate; this increased the shear stresses on and beneath the contact surface, starting ratcheting and consequent surface cracking. As far as the grains were strained along the friction direction, the ratio of their external border to their volume decreased: this is in agreement with the Barkhausen noise decrease, because the grain border is a factor hindering the magnetic dominions rotation. Also other phenomena occurring with ratcheting, such as the strain hardening and compressive residual stress increment observed at the test end, are phenomena concurring to lower the Barkhausen noise, therefore this parameter can be considered an indicator of the ratcheting progress. On the contrary, the surface smoothing occurring after 2  105 cycles is expected to improve the contact between the specimen surface and the Barkhausen noise probe, therefore it can be excluded a correlation between the Barkhausen noise decrease and the changes in the specimen-probe effective contact area. The correlation of DT/x with ratcheting can be explained if ratcheting is thought as a process linked to contact stick–slip alternation rather than a continuous phenomenon. In fact the DT/x parameter takes into account the relationship between torque and displacement vibrations of the mandrel, therefore is expected to be particularly sensitive to torsional vibrations. The Barkhausen noise stabilized at a constant value after 4  105 cycles, meaning that the microstructure near the contact surface became stable as well. As shown in Ref. [20,25] by means of experiments and simulations of the cyclic strain process, this condition often occurs in rolling/sliding contact, under constant load and sliding ratio, when equilibrium between the wear rate and the ratcheting rate is reached. The third phase started when the friction coefficient fell below the wheel ratcheting threshold, i.e. after about 5.5  105–6  105 total cycles. This phase was characterized by lower wear rate, further decrease of the friction coefficient and further surface smoothing; correspondingly, the DT/x parameter significantly decreased. The rail specimen was expected to be still subjected to ratcheting, even by a lower rate with respect to the previous phase due to lower friction coefficient. This is in agreement with the DT/x profile, which in the third phase is lower than in the second one, when both steels were subjected to ratcheting, but higher than in the initial one, when none was subjected. In the third phase the Barkhausen noise still kept approximately constant, meaning that the mild wear still present in this phase was not sufficient for removing material enough for allowing the emersion of significantly different microstructure. In mixed tests, in the dry phase the measured parameters showed the same behaviour as in dry tests. In the wet phase, the friction coefficient fell down to about 0.12, due to the effect of lubricant, whereas DT/x, after a few cycles, had a sudden increment: this can be related to the beginning of RCF, that in less than 4  104 cycles led to the very severe damage shown in Fig. 6. The lubricated tests were characterized by low wear with approximately constant rate and slight surface deterioration. The microstructure, as well, did not show signs of significant plastic strain, in agreement with the slight hardness profile alteration. The friction coefficient was quite unstable. The band of oscillation was ever below the ratcheting threshold for the wheel steel, therefore justifying the absence of visible plastic strain at the test end; on the contrary, it was slightly over the ratcheting threshold

31

for the rail specimens, in agreement with the surface hardness increment measured in wet tests. The Barkhausen noise showed a wide band of oscillation in both lubricated tests; furthermore, large difference between the results of the two tests was found. This cannot be correlated to any significant difference between the two wet tests. However, the difference with respect to the dry tests is evident, as the Barkhausen noise, although varying with high scatter, is included in a band clearly separated from the one of the dry tests; therefore an association between the material strain state and the measured Barkhausen level is possible. The onset of surface damage in the rail specimen in the Wet 2 test was revealed by the DT/x profile, that started to raise up and become instable correspondingly to the deterioration of the rail specimen surface; furthermore, comparing Fig. 5 with Fig. 13, a correlation can be found between the DT/x values and the grade of the contact surface deterioration. These evidences show the usefulness of the investigated NDT for monitoring the material damage. A direct correlation between the wear rate, the friction coefficient and the DT/x parameter is evident. The wear measurements require specimen dismounting and weighing: frequent test stops are necessary for accurately tracking the wear evolution, with much time consuming. By means of the friction coefficient and DT/x acquisition, the number of test stops for weight measurements can be reduced to the minimum necessary for calibrating the wear curve; the wear rate evolution can be reconstructed as analogous to the friction coefficient variation. Furthermore, when the friction coefficient falls under the elastic shakedown limit at least for one specimen and DT/x rapidly decreases,the wear rate is expected to change, therefore that is the right time for taking wear measurements. The DT/x parameter in dry contact, being correlated with ratcheting occurrence at one specimen, is useful for reconstructing the strain history during the test. In wet contact, it is correlated to the occurrence of RCF, even when the damage is still light. This is very useful for determining the right time for stopping the tests, both in mixed and wet tests: in fact, usually, observing fatigue cracks in the initial phase of the damage process is much more interesting than waiting for the complete destruction of the surface layer, as crack nucleation sites, propagation paths and modes can be more easily determined. On the other hand, the Barkhausen noise is directly correlated with the deformation state of the layer beneath the contact surface: low Barkhausen noise levels imply severely strained microstructure, with the probable presence of surface cracks. Furthermore, when the Barkhausen noise stabilizes, also the material microstructure and properties under the contact surface are expected to keep constant as far as the same working condition occur, therefore the test duration can be set on the basis of this parameter. All these parameters are useful for improving the simulations of material damage by means of numerical methods, such as in Ref. [24], especially when these are carried out for material cyclic plasticity characterisation. In fact, the continuous torque measurement, coupled with periodical weight loss measurement, allows determining wear rate and friction coefficient histories that can be introduced in the simulations instead of constant mean parameters, as is usually done. The DT/x parameter allows verifying the simulated deformation process in relation to the duration of ratcheting; on the other hand, the Barkhausen noise measurements give information about the number of cycles necessary for the strain stabilization. 5. Conclusions Non destructive measurements techiniques were introduced in a two-disc testing procedure applied to couples of specimens in

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rail and railway wheel steels. In particular, Barkhausen noise, torque and vibrations measurements were added to periodical weighings and final destructive techiques that are usually used for wear characterisation and subsurface region investigation. Tests in dry, wet and mixed dry–wet condition were carried out. In dry tests correlations were found between the wear rate, the friction coefficient and the vibration-based parameter DT/x. In particular, DT/x was found sensitive to wear rate and ratcheting. In mixed and wet tests it was found that the vibration parameter is sensitive to surface damage, showing a qualitative correlation with the surface deterioration grade. The Barkhausen noise was found to be correlated with the deformation state of the microstructure just below the contact surface: in particular, phenomena related to ratcheting (high plastic strain, high compressive residual stresses, strain hardening) were found to cause a significant Barkhausen noise reduction. These experiments showed the usefulness of the investigated non destructive measurements techniques. Vibrations and torque measurements are useful for reconstructing the strain history in high friction processes, as typical in dry contact. Furthermore, given the correlation with the wear rate, they are useful for determining the wear hystory during a test with a reduced number of stops for specimen weighing. The strain and wear hystory can be used, for instance, in material chracterisation by means of numerical simulations. Vibration measurements allow determining the time of rolling contact fatigue phenomena initiation: this way, the tests can be stopped before the occurrence of very severe damage and the subsurface cracks can be better studied. The Barkhausen noise measurement allows determining when the material under the contact surface is heavily strained and likely affected by surface cracks, which can enhance rolling contact fatigue in presence of water. This parameter, in particular, can have a development as diagnostic tool for best scheduling of maintenance stops of real train wheels, as it can help to determine the right time for reprofiling the wheel thread in order to remove surface cracks and prevent fluid driven rolling contact fatigue.

Acknowledgements The authors are grateful to Lucchini RS – Italy for materials supplying and technical support, and to 2 Effe Engineering – Italy for the X-RD measurements of the residual stresses. They also wish to thank Silvio Bonometti and Valentina Ferrari for their support in the experimental activities.

References [1] A. Ekberg, E. Kabo, Fatigue of railway wheels and rails under rolling contact, Wear 258 (2005) 1288–1300. [2] Bharat Bhushan (Ed.), Modern Tribology Handbook Vol. II, CRC Press, Boca Raton, Florida, USA, 2001, pp. 1271–1326. [3] G. Donzella, M. Faccoli, A. Ghidini, A. Mazzù, R. Roberti, The competitive role of wear and RCF in a rail steel, Eng. Fract. Mech. 72 (2005) 287–308. [4] G. Donzella, A. Mazzù, C. Petrogalli, Competition between wear and rolling contact fatigue at the wheel-rail interface: some experimental evidences on rail steel, Proc. Inst. Mech. Eng. F: J. Rail Rapid Transit. 223 (F1) (2009) 31–44. [5] F.J. Franklin, A. Kapoor, Modeling wear and crack initiation in rails, Proc. Inst. Mech. Eng. Part F: J. Rail Rapid Transit. 221 (2007) 23–33. [6] J. Dearden, The wear of steel rails: a review of the factors involved, Proc. Inst. Civil Eng.: Eng. Div. 3 (5) (1954) 456–481. [7] J. Dearden, The wear of steel rails and tyres in railway service, Wear 3 (1960) 43–49. [8] R.K. Steele, Observations of in-service wear of railroad wheels and rails under conditions of widely varying lubrication, ASLE Trans. 25 (3) (1982) 400–409. [9] I.J. McEwen, R.F. Harvey, Full-scale wheel-on-rail testing: comparison with service wear and a developing theoretical predictive model, Lubr. Eng. 41 (2) (1985) 80–88.

[10] S. Kumar, D.L.P. Rao, Wheel–rail contact wear, work and lateral force for zero angle of attack – a laboratory study, ASME Trans. J. Dyn. Syst. Meas. Control 106 (1984) 319–326. [11] J. Kalousek, D.M. Fegredo, E.E. Laufer, Wear Resistance and Worn Metallography of Pearlite, Bainite and Tempered Microstructures of high Hardness, in: K.C. Ludema (Ed.), ASME Trans., 1985, pp. 212–231. [12] P.J. Bolton, P. Clayton, Rolling–sliding wear damage in rail and tyre steels, Wear 93 (1984) 145–165. [13] H. Krause, G. Poll, Wear of wheel–rail surfaces, Wear 113 (1986) 103–122. [14] J.E. Garnham, J.H. Beynon, Dry rolling–sliding wear of bainitic and pearlitic steels, Wear 57 (1992) 81–109. [15] F.C. Robles Hernandez, N.G. Demas, D.D. Davis, A.A. Polycarpou, L. Maal, Mechanical properties and wear performance of premium rail steels, Wear 263 (2007) 766–772. [16] F.C. Robles Hernandez, N.G. Demas, K. Gonzales, A.A. Polycarpou, Correlation between laboratory ball-on-disc and full-scale rail performance tests, Wear 270 (2011) 479–491. [17] F. Braghin, S. Bruni, F. Resta, Wear of railway wheel profiles: a comparison between experimental results and a mathematical model, Veh. Syst. Dyn. 37 (2003) 478–489. [18] R. Enblom, M. Berg, Simulation of railway wheel profile development due to wear-influence of disc braking and contact environment, Wear 258 (2005) 1055–1063. [19] F.J. Franklin, I. Widiyarta, A. Kapoor, Computer simulation of wear and rolling contact fatigue, Wear 251 (2001) 949–955. [20] A. Mazzù, Surface plastic strain in contact problems: prediction by a simplified non-linear kinematic hardening model, J. Strain Anal. Eng. Des. 44 (2009) 187–199. [21] G. Fajdiga, J. Flašker, S. Glodež, The influence of different parameters on surface pitting of contacting mechanical elements, Eng. Fract. Mech. 71 (2004) 747–758. [22] J. Brouzulis, Wear impact on rolling contact fatigue crack growth in rails, Wear 314 (2014) 13–19. [23] A. Ekberg, E. Kabo, H. Andersson, An engineering model for prediction of rolling contact fatigue of railway wheels, Fatigue Fract. Eng. Mater. Struct. 25 (10) (2002) 899–909. [24] A. Mazzù, C. Petrogalli, M. Faccoli, An integrated model for competitive damage mechanisms assessment in railway wheel steels, Wear 322–323 (2015) 181–191. [25] G. Donzella, M. Faccoli, A. Mazzù, C. Petrogalli, R. Roberti, Progressive damage assessment in the near-surface layer of railway wheel–rail couple under cyclic contact, Wear 271 (1–2) (2011) 408–416. [26] C.C.H. Lo, A review of the Barkhausen effect and its applications to nondestructive testing, Mater. Eval. 62 (2004) 741–748. [27] J. Anglada-Rivera, L.R. Padovese, J. Capo-Sanchez, Magnetic Barkhausen noise and hysteresis loop in commercial carbon steel: influence of applied tensile stress and grain size, J. Magn. Magn. Mater. 231 (2001) 299–306. [28] S. Yamaura, Y. Furuya, T. Watanabe, The effect of grain boundary microstructure on Barkhausen noise in ferromagnetic materials, Acta Mater. 49 (2001) 3019–3027. [29] J. Gauthier, T.W. Krause, D.L. Atherton, Measurement of residual stress in steel using the magnetic Barkahusen noise technique, NDT E Int. 31 (1998) 23–31. [30] G. Bertotti, A. Montorsi, Dependence of Barkhausen noise on grain size in ferromagnetic materials, J. Magn. Magn. Mater. 83 (1990) 214–216. [31] A.J. Birkett, W.D. Corner, B.K. Tanner, S.M. Thompson, Influence of plastic deformation on Barkhausen power spectra in steels, J. Phys. D: Appl. Phys., 22, (1989) 1240–1242. [32] M.M. Blaow, B.A. Shaw, Magnetic Barkhausen noise profile analysis: effect of excitation field strength and detection coil sensitivity in case carburized steel, Mater. Sci. Appl. 5 (2014) 258–266. [33] H. Kikuchi, K. Ara, Y. Kamada, S. Kobayashi, Effect of microstructure Changes on Barkhausen noise properties and hysteresis loop in cold rolled low carbon steel, IEEE Trans. Magn. 45 (6) (2009) 2744–2747. [34] L. Nohál, F. Hort, J. Dvořáček, P. Mazal, An experimental investigation of rolling contact fatigue of steels using acoustic emission method, Insight: Nondestr. Test. Cond. Monit. 55 (12) (2013) 665–669. [35] C.S. Byington, M.J. Watson, J.S. Sheldon, G.M. Swerdon, Shaft coupling modelbased prognostics enhanced by vibration diagnostics, Insight: Nondestr. Test. Cond. Monit. 51 (2009) 420–425. [36] J. Lee, F. Wu, W. Zhao, M. Ghaffari, L. Liao, D. Siegel, Prognostics and health management design for rotary machinery systems—reviews, methodology and applications, Mech. Syst. Sign. Process. 42 (2014) 314–334. [37] A. Ghidini, M. Diener, A. Gianni, J. Schneider (Eds.), Superloss: Innovative Steel by Lucchini RS for High-Speed Wheel Application, Lucchini RS, Lovere, Italy, 2012. [38] L. Solazzi, C. Petrogalli, M. Lancini, Vibration based diagnostics on rolling contact fatigue test bench, Procedia Eng. 10 (2011) 3465–3470. [39] K. Farhangdoost, M. Kavoosi, Effect of lubricant on surface rolling contact fatigue cracks, Adv. Mater. Res. 97–101 (2010) 793–796. [40] S. Bogdański, P. Lewicki, 3D model of liquid entrapment mechanism for rolling contact fatigue cracks in rails, Wear 265 (9) (2008) 1356–1362. [41] W.R. Tyfour, J.H. Beynon, A. Kapoor, Deterioration of rolling contact fatigue life of pearlitic rail steel due to dry–wet rolling–sliding line contact, Wear 197 (1996) 255–265. [42] J.H. Beynon, J.E. Garnham, K.J. Sawley, Rolling contact fatigue of three pearlitic rail steels, Wear 192 (1996) 94–111.