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Effect of shoe braking on wear and fatigue damage of various railway wheel steels for high speed applications
Journal Pre-proof Effect of shoe braking on wear and fatigue damage of various railway wheel steels for high speed applications Angelo Mazzù, Luca Provezza, Nicola Zani, Candida Petrogalli, Andrea Ghidini, Michela Faccoli PII:
S0043-1648(19)30663-5
DOI:
https://doi.org/10.1016/j.wear.2019.203005
Reference:
WEA 203005
To appear in:
Wear
Received Date: 12 April 2019 Revised Date:
2 July 2019
Accepted Date: 7 August 2019
Please cite this article as: A. Mazzù, L. Provezza, N. Zani, C. Petrogalli, A. Ghidini, M. Faccoli, Effect of shoe braking on wear and fatigue damage of various railway wheel steels for high speed applications, Wear (2019), doi: https://doi.org/10.1016/j.wear.2019.203005. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2019 Published by Elsevier B.V.
Effect of shoe braking on wear and fatigue damage of various railway wheel steels for high speed applications Angelo Mazzùa*, Luca Provezzaa, Nicola Zania, Candida Petrogallia, Andrea Ghidini b, Michela Faccolia a
Department of Mechanical and Industrial Engineering, University of Brescia, via Branze 38, 25123 Brescia, Italy b
Lucchini RS, Via G. Paglia 45, 24065 Lovere (BG), Italy * [email protected], tel: (+39) 030 371 5525
Abstract The effect of shoe braking on three railway wheel steels was studied by means of bi-disc tests, in order to evaluate the applicability of shoe braking as an emergency stop braking system in highspeed passenger trains. Disc-shaped specimens, extracted from real wheels, were initially paired with discs extracted from cast iron brake blocks, in condition of rolling and sliding such to reach the typical temperature of the surface of a wheel rim during stop braking. Subsequently, the wheel discs were coupled with rail steel discs in rolling and sliding dry contact. Some of them were cut and analysed; other ones were subsequently subjected to rolling and sliding contact with rail specimens in wet condition, with various test duration. Measurements of the coefficient of friction and of the weight loss were carried out during the tests; at the end of the tests, hardness measurements and microstructure observation were made on the disc cross-sections. The main damage phenomena observed in the specimens subjected to wheel-brake contact and to dry wheel-rail contact were ratcheting, wear and nucleation of surface cracks. Traces of cast iron, previously transferred from the brake specimens to the wheel ones, were detected. The main damage phenomenon observed in the specimens subjected to wheel-brake, dry wheel-rail and wet wheel-rail contact was rolling contact fatigue, developing by the propagation of surface cracks due to the pressurization of the water entrapped inside the cracks. All of these phenomena were found to be similar to those 1
observed in wheel specimens subjected to wheel-rail contact only. The effect of the cast iron traces on surface crack propagation was evaluated by means of finite element simulations and it was found to be limited. Therefore, the application of shoe braking as emergency stop braking system to high speed trains did not evidence any relevant contraindications.
Keywords: High speed trains, railway wheels, shoe braking, rolling contact fatigue, wear.
1. Introduction Shoe braking is obtained by direct pressure of brake blocks on the wheel tread. It is typical for freight trains and frequent for metro and suburban trains, either as unique braking system or coupled with other ones. For high-speed passenger transportation, the EN 15734-1 standard recommends electro-dynamic braking as principal system; however, shoe braking is allowed as emergency braking system in coupling with the main one, provided that the braking power is limited and it is used only when the train speed is below 120 km/h. Notwithstanding the problems that can be related to the direct action of the brake blocks on the wheel tread, an interest on the application of shoe braking to high speed trains has recently raised, due to its cleaning effect from environmental contaminants that can have a role on wheel damage [1]-[3]. Furthermore, in comparison with other friction braking systems such as disc braking, shoe braking is advantageous in terms of axle weight and cost. Many authors studied the damage induced by the contact loads coupled with the heat generated by friction at the block-tread interface. Some authors [4]-[11] focused especially on the effects of thermomechanical loads on the material microstructure, including pearlite spheroidization or occasionally White Etching Layer (WEL) formation. Vernersson et al. [12]-[14] studied both numerically and experimentally the tread corrugation due to thermal loads, identifying a non-even wear mechanisms on the tread. Some studies [10], [11], [15] highlighted the role of these 2
phenomena in favoring the nucleation of surface cracks on the wheel, also due to the change of the residual stresses from tension to compression as a consequence of thermal cycles. Peng et al. [16], [17], studying the effect of thermal loads on crack propagation, distinguished stop braking, which is the stop of a train from a given speed, from drag braking, which is acted when the train speed has to be kept constant along a slope. They found that in drag braking the tread temperature can rise up to about 680° C, whereas in stop braking it is limited to about 200 °C. Faccoli et al. [18], [19] studied the changes of microstructure and mechanical properties of some wheel steels under different thermal histories reproducing various braking operations. They found that, whereas the temperatures that can be reached in drag braking can induce phase transformation, under temperatures typical of stop braking only slight changes can be observed. Teimourimanesh et al. [20] elaborated a temperature-dependent elastic-plastic material model, coupled with a fatigue model, for predicting the wheel life in presence of thermomechanical loads due to tread braking and to wheel-rail contact. In [21], they applied this model to the case of a metro train, considering both stop braking and drag braking, finding also that in the case of repeated stop brakings the fatigue life is controlled by mechanical loads rather than by thermal loads. In [22] they measured the temperature reached in a metro train, finding that it does not exceed 250°C. Vernersson et al. [23] studied the effect of thermomechanical load cycles on Rolling Contact Fatigue (RCF), especially considering the effect of tread temperatures up to 300°C on ratcheting. Caprioli et al. [24] studied the propagation of cracks due to thermal loading induced by tread braking in heavy haul applications. They found that fully functional brake systems on heavy haul trains are not likely to induce thermal crack propagation in stop braking, unless in the case of severe drag braking due to malfunctioning brakes. Caprioli et al. [25] studied RCF induced by ratcheting in stop braking with both a numerical and experimental approach, highlighting the role of repeated braking cycles in the formation of surface cracks.
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Overall, these studies identified the drag braking on heavy haul trains as the most severe application for tread-braked wheels: indeed, the high temperatures reached in this condition are able to induce microstructural changes, tensile residual hoop stress and thermal fatigue. Stop braking, especially in passenger trains, at most is expected to induce surface cracks, in addition to those due to ordinary mechanical loads. However, in high speed applications the damage related to stop braking cannot be neglected. High speed trains are characterized by a lower axle load than freight trains and by rarer stop braking operations with respect to metro trains. On the other hand, many operators are interested on extending the mileage between maintenance operations on high speed train wheels, in order to reduce the costs of train stops. This way, the wheels tend to be subjected to more loading cycles between the wheel maintenance operations, due to their higher speed. Even though microstructure changes or significant thermal cracks are not expected, small surface cracks generated during tread braking can be preferential sites for RCF initiation, especially under the action of fluid contaminants, which promote crack propagation by means of the pressurization of the fluid entrapped inside the cracks [2], [26], [27]. In a recent study [28], the effect of shoe braking by cast iron blocks on various wheel steels has been studied by means of bi-disc contact tests. Cylindrical specimens extracted from real wheels were put in rolling and sliding contact against cylindrical specimens extracted from cast iron brake blocks, in such a working condition to reproduce on the wheel specimens the typical temperature of a real wheel tread during the stop braking of a high speed train (around 240°C). The authors found that the damage mechanisms occurring at the surface of the wheel specimens were wear, ratcheting and surface crack nucleation. In addition, they documented a mechanism of material transfer from the brake specimens to the wheel ones, which is thought to play a key role in the tread damage. A similar phenomenon was observed even by Vernersson et al. [29] on real wheels, although they did not study it in depth. Due to the material transfer from the brake specimen to the wheel one, a “third 4
body” layer is generated on the surface of the latter. When it is detached, it also involves the steel substrate, probably promoting the nucleation of surface cracks. Indeed, in [28] there was evidence of surface cracks in the wheel steel layer starting from zones where the third body layer had been detached. This paper illustrates a forward step from the achievements described in [28]. The aim is to study the evolution of the surface damage when shoe braked wheels are put in contact with rails, both in dry and wet contact. This investigation was carried out again by means of bi-disc tests, using specimens of various wheel steels. These specimens were first put in dry contact with cast iron brake block specimens, then again in dry contact with rail steel specimens. For some of them, this procedure was followed by a session of wet contact with rail steel specimens. The damage in the tested wheel materials was evaluated in terms of wear, strain hardening, crack nucleation and propagation.
2. Experimental tests: materials and methods A test procedure was designed to experimentally simulate the effect of shoe braking on the wheel tread, considering its consequences on dry and wet wheel-rail contact subsequently occurring. The tests were carried out on a bi-disc bench whose schematic is shown in Figure 1. The discs were mounted on two shafts driven by independent engines, one of which can be displaced orthogonally to the shaft axis by means of a hydraulic piston, which also applies the contact load.
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Specimen 2 (wheel)
Mobile mandrel
Torque sensor
Fixed mandrel
Specimen 1 (brake-rail)
Encoder Hydraulic cylinder
Figure 1. Schematic of the bi-disc testing machine. Three different wheel steels were the object of the investigation, identified in the following as HYPERLOS®, CLASS B and SANDLOS®. HYPERLOS® is an upgrade of the EN13262 ER7 steel [18], [19]. The HYPERLOS® steel fulfils the requirements of the EN13262 standard, it has a better combination of strength and toughness with respect to the traditional ER7 steel thanks to the higher content of C, Mn, Si and V and a well-balanced chemical composition. The AAR M107/M208 CLASS B steel is widely used in North America, recommended for freight cars in interchange service and for use on locomotives. For passenger car service, it is used for high-speed service with severe braking conditions and heavier wheel loads. SANDLOS® steel is an upgrade of the CLASS B steel [1], [30], with a higher content of Mn and Si. Their chemical composition and mechanical properties are reported in Table 1. The tensile properties were obtained by the supplier using standard specimens extracted from the components, according to EN 6892-1 Standard. The Brinell hardness was measured on the radial section of the wheel rims in accordance with EN ISO 6506-1 Standard, in the same position as that of the disc extraction. All of the wheel steels were supplied by
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Lucchini RS. The wheel steels were tested in coupling with a brake block cast iron and subsequently with the rail steel R350HT EN 13674-1. Cast iron was chosen for the block material as its properties are well-known and less subject to the production process than other brake block materials [31]. The chemical composition and the nominal mechanical properties of the rail steel are shown again in Table 1; the chemical composition and the Brinell hardness of the brake block cast iron are shown in Table 2. The wheel discs had 80 mm diameter and 20 mm thickness; the brake and the rail discs had 60 mm diameter and 15 mm thickness. All of the specimens were extracted from real components according to the schematic shown in Figure 2
Chemical composition [wt%]
HYPERLOS® CLASS B SANDLOS® R350HT 0.51 0.65 0.63 0.63 C 0.78 0.63 0.84 1.095 Mn 0.38 0.26 0.88 0.296 Si 0.002 0.001 0.001 0.018 S 0.015 0.012 0.009 0.01 P Ultimate tensile strength [MPa] 885 990 1142 ≥1175 Yield strength [MPa] 568 642 690 Elongation [%] 19 14 15 ≥9 Hardness HB 280 315 322 355 Table 1. Main chemical elements and mechanical properties of the wheel steels. Cast iron chemical composition [wt%] C S P Mn Cr Ni Mo Cu Si V Al Ti 3.03 0.18 1.70 0.61 0.10 0.05 0.01 0.15 1.66 0.006 0.04 0.05 Table 2. Chemical composition and Brinell hardness of the brake block cast iron.
HB 230
7
brake block
wheel rail
Figure 2. Schematic of specimen extraction from a wheel, a brake block and a rail Two series of tests were carried out. In the first one a wheel disc was initially paired with a brake disc and cycled for 2000 cycles (“braking step” in the following); subsequently, the brake disc was replaced with a rail disc and the wheel disc was cycled with follower role for 10000 cycles in dry contact (“dry step” in the following). This test was repeated three times for each wheel material (test 1, test 2 and test 3). The second series of tests consisted of the same two steps of the first one, followed by another step with the addition of liquid to the wheel-rail interface (“wet step” in the following). The liquid was water (with the addition of 10% glycol to protect the test bench from corrosion), ejected towards the contact interface with a flow of 6×10-6 m3/s. For each material, four tests were carried out, characterized by different durations: 30000 cycles (test 4), 20000 cycles (test 5), 10000 cycles (test 6) and 5000 cycles (test 7). In all the tests, the cycles were counted as the rotations of the wheel steel specimen. The working conditions of the braking step, in particular the sliding speed and the contact load, were chosen to have a friction dissipated power able to obtain a steady temperature on the specimen contact surface as that resulted from the thermal simulation presented in [28], which was aimed at estimating the typical temperatures reached on the tread of a full-scale wheel during stop braking. The number of cycles was chosen to let the surface temperature of the specimen stabilize. These 8
working conditions were first estimated and subsequently tuned per trials in preliminary tests. The contact load in the dry and wet steps was chosen in order to reproduce a typical value of the wheelrail contact pressure in curve [32]; the rolling and sliding speeds in these steps were chosen to have a sliding-to-rolling ratio of 1% and, with the given load, to let the discs work in the field of “mild wear” according to the wear maps given for instance by Lewis et al. [33]. Overall, these working conditions are expected to be more severe than in real wheels, which are subjected to such contact pressure and sliding only in curve, e.g. intermittently and not continuously: these were chosen just to compare the behavior of the three steels. The working conditions in the different steps of the tests are given in Table 3.
Specimen Rolling speed (rpm) Diameter Tangential speed (m/s) Sliding speed ݒ௦ (m/s) Contact width (mm) Contact load ( ܨN) Contact Hertz pressure (MPa)
Braking step Wheel Brake 175 -175 80 60 0.73 -0.55 1.28 15 2000 529
Dry and wet step Wheel Brake 373 502.5 80 60 1.562 1.579 0.017 15 8636 1100
Table 3. Working conditions in the different step of the tests The coefficient of friction was obtained from the torque signal coming from a sensor mounted between the displaceable specimen shaft and the transmission, which was elaborated through the signal acquisition system and the procedure detailed in [34], [35]. The temperature on the contact surface of the specimens during the braking step was checked by means of a thermographic camera following the procedure detailed in [28]. The weight variation of the specimens was measured by means of a precision balance with a resolution of 0.001 g after cleaning in a bath of ethanol with ultrasonic vibrations; the weight measurements were taken before the tests and at the end of each step.
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At the end of the tests, the wheel discs were cut along the mid plane orthogonally to the contact surface. The disc sections were ground, mechanically polished to a 1 µm finish, etched with 2% Nital and examined with a Leica DMI 5000 M light optical microscope. The deformation under the contact surface and the crack morphology were investigated and the damage mechanisms were identified. The length and the inclination of all cracks longer than 20 µm detected over a 120° sector of each section were measured. The Vickers hardness was measured on the wheel disc sections at varying distances from the contact surface to evaluate the steel work-hardening phenomenon and correlate it with the deformation beneath the contact surface. The tests were carried out using a 1000 g load and a dwell time of 15 s, in compliance with ASTM E384.
3. Experimental results 3.1. First test series (braking step + dry step) 3.1.1
Weight loss
Figure 3 shows the plots of weight loss versus number of cycles for the discs of the three wheel steels at the end of each step of the test (1= braking step, 2 = dry step). Table 4 shows the weight loss of the brake discs at the end of the braking step and the weight loss of the rail discs at the end of the dry step. At a first sight, the weight of all wheel discs decreases after both steps, as a consequence of wear. During the braking step, material transfer from the brake disc to the wheel disc resulting in the formation of a discontinuous “third body” layer was observed by the naked eye thanks to the darker colour of cast iron compared with steel. The brake material transfer resulted from the mechanical stresses at the contact surface, the heating of the contact surfaces and the agglomeration of wear debris [36]. The “third body” layer, as previously observed by the authors in similar tests [28], is worn away due to cycling, and detachment of steel from the wheel disc surface 10
also occurs, probably promoting crack nucleation. In addition, wear debris from both disc materials leads to three-body abrasive wear of the wheel disc surface. In the dry step of the test, wear and crack growth due to the contact with the rail disc lead to further material detachment (the mentioned phenomena are documented in the following). The HYPERLOS® discs are the most affected by wear, whereas the SANDLOS® discs are the least affected. This result is due to their different hardness [37], being the former the softest steel and the latter the hardest one. The weight loss of the brake discs is much higher than that of the wheel discs at the end of the braking step because of the lower hardness of the cast iron compared with the steel. The brake discs tested against the HYPERLOS® discs show slightly higher weight loss than those tested against SANDLOS® and CLASS B discs. The rail discs show similar weight loss regardless the counterpart. For all materials, the weight loss measured in the three repeated tests is similar, indicating a good reproducibility of the tests.
11
1
1
2
1
2
2
Figure 3. Weight loss of the wheel discs in the first test series (1= braking step, 2 = dry step). Weight loss [g] Brake discs
vs HYPERLOS®
vs CLASS B
vs SANDLOS®
17.83 ± 0.70
16.61 ± 0.88
16.50 ± 1.77
Rail discs
0.05 ± 0.02
0.04 ± 0.02
0.06 ± 0.03
Table 4. Weight loss of the brake discs at the end of the braking step and of the rail discs at the end of the dry step.
3.1.2
Coefficient of friction and temperature
Figure 4 shows the coefficient of friction measured during the tests. For all of the steels, in the braking step the coefficient of friction starts from a value ranging from 0.4 to 0.5 and subsequently tends to a value around 0.2-0.3, similarly to the tests performed in [28]. In the subsequent dry step, 12
the coefficient of friction increases, with some differences between the steels: it stabilizes at 0.350.45 for the HYPERLOS®, 0.3-0.4 for the SANDLOS® and 0.30-0.55 for the CLASS B, with more dispersion for the latter. The increment of the coefficient of friction can be reasonably correlated to the removal of the deposited layer of cast iron, which leads the two steels to direct contact. The variation of the wheel disc surface temperature was monitored during the braking step of the test by probing the point temperature on thermographic images near the contact region. The wheel disc surface reaches and then maintains about 230 °C after 1750 cycles in all tests, reproducing the temperature of the wheel rim during stop braking estimated by finite elements in a previous work [28]. This temperature is not high enough to induce microstructural changes in the wheel steels, but it leads to an increasing of the wear rate of the brake cast iron, as shown by Abbasi et al. [38], and it probably also promotes the adhesion of the “third body” layer on the wheel disc.
13
Figure 4. Coefficient of friction in the first test series (1= braking step, 2 = dry step).
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3.1.3
Subsurface morphology and crack mapping
Figure 5 shows some representative cross-sections of the wheel and the rail discs cut at the end of the tests. The ferritic-pearlitic microstructure (with traces of bainite for the HYPERLOS® and SANDLOS® steels) is highlighted by the Nital etch. A layer with unidirectional plastic flow that tends to be aligned in the direction of the surface friction can be seen below the contact surface. This layer, as shown for instance by Mazzù [39] and Ponter et al. [40], is due to ratcheting, intended as the plastic shear strain accumulation at each load cycle during the tests, due to the high friction at the contact surface. Several little surface cracks can be observed in the samples of all three steels. These cracks start with a shallow angle to the surface and follow the plastically deformed material during their growth. They are caused by ratcheting and probably also promoted by the detachment of the material transferred from the brake disc to the wheel disc during the braking step, as also observed in [28]. The brake material is almost completely removed during the repeated wheel disc/rail disc contact (dry step of the test), only few traces were observed on the wheel discs at the end of the tests. A detail of brake material inside a surface crack in a HYPERLOS® specimen is shown in Figure 5. These results agree with the works of Vernersson et al. [12]-[14], who studied the evolution of the wheel roughness due to tread braking by means of full-scale block braking experiments. They found that cast iron blocks caused in stop braking condition higher roughness on the wheel surface than composition and sinter blocks did, and they explained this effect with the material transfer from the cast iron blocks to the wheel, although they did not study in depth this phenomenon. Given these evidences, it is reasonable to infer that similar phenomena also occur in a real wheel during stop braking, even though they cannot be easily observed on the wheel tread because the material transfer and removal occur subsequently at each wheel revolution.
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Figure 5. Micrographs of the wheel discs (on the left) and paired rail discs (on the right). The wheel steels are: a) HYPERLOS®, b) CLASS B and c) SANDLOS®; the rail steel is 350HT (d).
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As surface cracks can propagate when a liquid contaminant is added to the contact surface, the three wheel materials were characterized also on the basis of the surface cracks generated. All cracks exceeding 20 µm in length, detected over an arc of 120° on the cross-sections of all of the specimens, were mapped in terms of geometry, e.g. in terms of their extension x and z along the directions tangential and radial respectively of the disc. The HYPERLOS® developed the lowest number of cracks, CLASS B the highest. At a first sight, CLASS B was also the material that developed the longest cracks, especially in test 1, which was characterized by the highest coefficient of friction. This result is consistent with the weight loss shown in Figure 3: the higher weight loss of the HYPERLOS® discs implied the removal of short surface cracks, leaving fewer and shorter cracks than in the CLASS B discs. As a consequence, the HYPERLOS® steel offers fewer initiation points to major fatigue damage, such as shelling, when wet contact occurs.
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Figure 6: Map of the measured extension of the cracks along the radial (z) and tangential (x) directions of the discs at the end of all of the first series tests. 3.1.4
Subsurface hardness
Figure 7 shows the hardness variation at increasing distance from the contact surface measured on the cross-section of the wheel steel discs. An increase of hardness can be observed under the contact surface in all tested discs. This effect is consistent with the pattern of deformation shown in Figure 5, as the material hardening is caused by the progressive accumulation of plastic strain under the contact surface. The maximum hardness is close to the contact surface where the plastic deformation is more severe; then the hardness gradually decreases at increasing distance from the contact surface as a consequence of the decreasing deformation. The HYPERLOS® and CLASS B discs show higher hardening compared with the SANDLOS® discs (∆HV1HYPERLOS® ~ 95, 18
∆HV1CLASS B ~ 108 and ∆HV1SANDLOS® ~ 35) an also larger hardened layer (~ 0.45 mm for HYPERLOS®, ~ 0.55 mm for CLASS B and ~ 0.35 mm for SANDLOS®). These results suggest better work-hardening behaviour of the former two steels compared with the latter. Furthermore, the CLASS B discs show a maximum hardness slightly higher than the other steels, which reach similar values. For all of the three steels, the hardness profile measured in the repeated tests is similar, confirming a good reproducibility of the tests.
HYPERLOS®
500
test 1 test 2 test 3
450
test 1 test 2 test 3
450 400
HV1
HV1
400
CLASS B
500
350
350
300
300
250
250
200
200 0
1
2 Depth (mm)
3
0
2 Depth (mm)
3
SANDLOS®
500
test 1 test 2 test 3
450 400
HV1
1
350 300 250 200 0
1
2 Depth (mm)
3
Figure 7: Vickers hardness profiles on the cross-section of wheel steels discs at the end of the first series of the tests.
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3.2 Second series of tests (braking step + dry step + wet step) 3.2.1
Weight loss
The weight loss of the wheel discs was measured at the end of each step of the test (1= braking step, 2 = dry step, 3 = wet step). The weight loss of the rail discs was measured at the end of the dry and the wet step. Figure 8 shows the plots of weight loss versus number of cycles for the wheel and the rail discs. Generally, the weight loss of all the wheel discs increases slowly up to 22000 cycles (corresponding to the total duration of the braking + dry + wet steps, with last 10000 cycles in wet contact), then it increase more quickly. The increase of the weight loss in the wheel discs tested longer than 22 000 cycles is related to the onset of RCF due to the pressurization of the fluid entrapped inside the surface cracks. The weight loss of all of the rail discs slightly increased with the number of cycles (except for the rail disc paired with the CLASS B disc in test 4), but it was a magnitude order lower than for the wheel discs, as the entrapped fluid pressurization takes place only in the follower disc. The rail discs paired with the SANDLOS® discs showed the highest weight loss, being this wheel steel the hardest one.
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HYPERLOS® test 4
test 4
test 5 2
test 6 test 7
1 1
2
3
Weight loss (g)
Weight loss (g)
RAIL 0.3
3
test 5
0.2
test 6 test 7
0.1
2
0
0 0
0
10000 20000 30000 40000 50000 Number of cycles
CLASS B
3
RAIL
test 6 test 7
1
2
3
Weight loss (g)
Weight loss (g)
test 4
test 5
2
0
test 5
0.2
test 6 test 7 0.1
2
3
0 0
10000 20000 30000 40000 50000 Number of cycles
0
SANDLOS®
3
RAIL test 4
test 6 1 1
test 7 2
3
0
Weight loss (g)
test 5
2
10000 20000 30000 40000 50000 Number of cycles
0.3
test 4 Weight loss (g)
10000 20000 30000 40000 50000 Number of cycles
0.3
test 4
1
3
test 5 0.2
test 6 test 7 2
0.1
3
0 0
10000 20000 30000 40000 50000 Number of cycles
0
10000 20000 30000 40000 50000 Number of cycles
Figure 8: Weight loss of the wheel and paired rail discs in the second series of tests (1 = braking step, 2 = dry step, 3 = wet step).
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3.2.2
Coefficient of friction and temperature of the wheel disc surface
Figure 9 shows the coefficient of friction during the longest tests (42000 cycles) of the second test series for the three steels. In the braking and dry steps the behaviour of the steels was similar to that of the first test series. In the wet step, initially it fell down to about 0.15-0.18 due to the lubrication effect of the added fluid; however, it progressively increased stabilizing at about 0.2 due to the occurrence of surface rolling contact fatigue which increased the surface roughness. A similar behaviour was observed in the shorter tests of the second series. The surface temperature was measured again in the braking step and it was in the same range of the first test series. 0.5 Coefficient of friction
HYPERLOS® CLASS B
0.4
SANDLOS® 0.3 0.2 0.1
1
2
3
0 0
5000
10000
15000 20000 25000 30000 Number of cycles
35000
40000
45000
Figure 9: Coefficient of friction in the longest tests (42000 cycles) of the second test series for the three tested wheel steels (1 = braking step, 2 = dry step, 3 = wet step). 3.2.3
Subsurface morphology
Figure 10 shows some representative sections of the wheel discs cut at the end of test 6 (22000 total cycle, with last 10000 cycles in wet contact). In addition to the plastic deformation accumulated during the tests, incipient shelling is observed below the contact surface in all wheel discs. The surface cracks nucleated during the braking + dry steps of the test initially propagated obliquely from the surface to a certain depth following the plastic deformed material. Then, in some cases, 22
they branched towards the surface, even creating connections between adjacent cracks. This behaviour, observed only in the follower (in this case the wheel disc) is typical of crack propagation driven by the entrapped fluid pressurization, as shown in previous works [2], [41] and studied in terms of Fracture Mechanics by Makino et al. [27]. The damage observed on the wheel disc sections is consistent with the weight loss shown in Figure 8. The surface cracks originated during the dry steps of the test grew in the wet step involving a much thicker layer than the plasticised depth; however shelling was not detected up to 22000 cycles (10000 cycles in wet contact), as shown by the micrographs taken after 5000 wet cycles (see for example the HYPERLOS® disc in Figure 11). Therefore, wear is the main damage mechanism of all of the wheel discs up to 22000 cycles. RCF prevails over wear in wheel discs tested for longer durations leading to shelling and consequently to relevant damage and weight loss (see also the HYPERLOS® disc at the end of the test 4 in Figure 12a). It is worth noting that the damage observed on the wheel discs, at least as the crack morphology is concerned, is similar to the one that affects the real shoe-braked wheels, as proved by the comparison between the HYPERLOS® disc section at the end of test 4 (Figure 12a) and the section of an ER7 shoe-braked wheel (Figure 12b), despite the working conditions in the laboratory tests, due to the continuous sliding imposed, are probably more severe than in real wheels. According to Figure 8, the SANDLOS® discs perform better than the other wheel discs up to 22000 cycles. This result is due to the higher hardness of SANDLOS® steel compared with the other two steels, which guarantees better wear behaviour. When RCF becomes the main damage mechanism the behaviour of the three steels is similar.
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Figure 10: Incipient shelling in the HYPERLOS® (a), CLASS B (b) and SANDLOS® (c) discs at the end of test 6 (22000 cycle).
Figure 11: Surface cracks in the HYPERLOS® disc at the end of test 7 (17000 cycles). 24
Figure 12: Shelling in the HYPERLOS® disc at the end of test 4 (42000 cycles) (a) and in a real shoe-braked ER7 wheel [courtesy of Lucchini RS] (b).
3.2.4
Subsurface hardness
Figure 13 shows the hardness variation at increasing distance from the contact surface measured on the section of the wheel steel discs at the end of test 6 (22000 cycles) and test 4 (42000 cycles). An increase of hardness is observed in all tested discs, which is consistent with the pattern of deformation shown in Figure 10 and Figure 11. The maximum hardness reached is sometimes lower than that measured at the end of the first series of tests due to the considerable detachment of material caused by RCF.
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HYPERLOS®
500
test 6 - 22 000 cycles test 4 - 42 000 cycles
450
test 6 - 22 000 cycles test 4 - 42 000 cycles
450 400 HV1
400 HV1
CLASS B
500
350
350
300
300
250
250
200
200 0
1
2 Depth (mm)
3
0
1
2 Depth (mm)
3
SANDLOS®
500
test 6 - 22 000 cycles test 4 - 42 000 cycles
450 HV1
400 350 300 250 200 0
1
2 Depth (mm)
3
Figure 13: Vickers hardness profiles on the cross-section of wheel steels discs at the end of the 22 000 and 42 000 cycle tests.
4. Evaluation of crack propagation and discussion The experimental evidences of the tests show damage mechanisms similar to those observed on wheel steels subjected to dry and to alternated dry-wet contact, such as those published in [2] and [41], for instance. As shown in [28], shoe braking causes material transfer from the brake blocks to the wheel rim. However, the evidences of the present tests show that such layer is removed in a relatively low number of cycles in rolling-sliding contact with the rail, only leaving few traces of the block material, as the one shown in Figure 5a. The removal of the transferred cast iron presumably contributes to the formation of surface cracks, in addition to ratcheting, as suggested by 26
the micrograph shown again in Figure 5a. However, the crack size distribution shown in Figure 6 does not highlight a clear distinction between the cracks generated by the two different mechanisms. Surface cracks, whatever the mechanism that generated them, propagate when fluid is added, causing severe RCF damage. In order to evaluate whether the presence of cast iron traces into a surface crack have a specific role in its propagation, the case depicted in Figure 5a was simulated by Finite Elements (FE). In particular, two models of the cracked regions were built: one considering uniform steel properties for the whole mesh, the other one considering cast iron and steel properties for the respective regions of Figure 5a. The model, shown in Figure 14, is built as a fixed 2D rectangular body, representing the wheel specimen, surmounted by a 2D traveling roller, representing the rail specimen. The radius of the rail roller was 17.14 mm, correspondent to the equivalent radius of the contact in the experimental tests, according to the Hertz theory. The wheel body was modelled by quadratic plane strain elements, the rail roller by linear ones. The bottom body had a 100 µm long crack, inclined by 15° with respect to the contact surface with a 1.5 µm wide opening. The crack tip was modelled by the quarter point technique: this technique is able to create a singularity at the crack tip generating an accurate stress field for a linear elastic model. A fluid cavity interaction was imposed inside the crack, meaning that it was considered full of incompressible fluid and consequently its volume was kept constant during the simulation. In addition, unilateral contact between the crack faces was imposed; no friction interaction was imposed, due to the lubrication effect of the entrapped fluid. As the material properties are concerned, the darker region of Figure 14 was modelled with the cast iron elastic properties in the first simulation (elastic modulus 120 GPa and Poisson ratio 0.23), and with the steel elastic properties in the second simulation (elastic modulus 206 GPa and Poisson ratio 0.29). Steel properties were attributed to the rest of the model. A unit contact load of 575.7 N/mm was imposed between the rail and the wheel bodies, able to generate the same pressure distribution as in the experimental tests. 27
cast iron or steel
steel
Figure 14: Finite Element model of a cracked wheel specimen portion with a traveling rail specimen. The gray part represents the portion of material that was modelled as steel in the first simulation and cast iron in the second one. The applied Stress Intensity Factors (SIF) in mode I (ܭூ ) and mode II (ܭூூ ) are obtained from the displacements of the quarter nodes on the crack face of the collapsed elements, according to the Williams equations:
ாට ۓ ೝ ۖ ܭூ = ሺݒଵ − ݒଶ ሻ ଼ሺଵିఔమ ሻ మഏ
మഏ ۔ ாට ೝ ۖܭூூ = ሺݑଵ − ݑଶ ሻ ە ଼ሺଵିఔ మ ሻ
(3)
Where v1 and v2 are the Crack Opening Displacements (COD) of the quarter nodes of the elements at the crack tip, u1 and u2 are the Crack Shearing Displacements (CSD) of the same nodes, r is the distance of the quarter nodes from the crack tip, E and ν are the material elastic modulus and coefficient of Poisson respectively. In this equation the steel properties were used for both of the models, as the region surrounding the crack tip was in steel for both.
Figure 15 shows the ܭூ and ܭூூ variation during a load passage for the two models, where the x abscissa represents the distance of the contact point from the crack mouth on the contact surface. 28
The difference between the results obtained with the two models is limited: the mode I and mode II SIF ranges ∆ܭூ and ∆ܭூூ are about 11% and 6% higher for the model with the cast iron layer with
respect to the model with uniform material. This means that the cast iron traces left after the rollingsliding contact between the wheel and the rail do not substantially affect crack propagation. The zone of adhesion between the steel matrix and the cast iron layer could be a preferential site for crack propagation due to debonding; however, the experimental evidences after the wet sessions did not highlight a significance of such phenomenon, being rather the in-depth propagation and subsequent branching and crack coalescence the main damage mechanism. 2.5 2 1.5 K [MPa m0.5]
1 0.5 0 -0.5 KI (with cast iron layer) KI (uniform steel) KII (with cast iron layer) KII (uniform steel)
-1 -1.5 -2 -2.5 -0.75
-0.5
-0.25
0 x [mm]
0.25
0.5
0.75
Figure 15: Mode I and mode II SIF variation during a load passage with uniform steel material and with a layer in cast iron. 5. Conclusions The damage occurring in three wheel steels due to shoe braking and to subsequent rolling-sliding contact with the rail was evaluated by means of bi-disc tests. Both dry contact and consecutive drywet contact were applied to the wheel-rail pairings. Ratcheting and wear were the main damage 29
mechanisms observed in the specimens subjected to brake-wheel contact and subsequently to railwheel dry contact, evidenced by the weight loss, the surface hardening and the microstructural plastic flow. Several shallow surface cracks were generated as well. Only traces of the cast iron transferred onto the wheel specimens during the previous brake-wheel contact step were found: most of the transferred cast iron, indeed, was removed in the initial cycles of the dry wheel-rail contact, as witnessed by the increment of the coefficient of friction. In tests with consecutive drywet wheel-rail contact after the wheel-brake step, the final failure occurred by fatigue due to propagation of surface cracks enhanced by the pressurization of the entrapped fluid. No traces of residual brake material were detected after this phase. The effect of the cast iron traces on crack propagation was evaluated by finite elements, and only a slightly increment of the applied stress intensity factor due to the stuck cast iron was found. In conclusion, the material state at the end of the dry wheel-rail contact phase was found not dissimilar to what was observed in previous rollingsliding wheel-rail tests, and the final failure mechanism in the wet phase was the same as well. No specific damage mechanism related to the braking phase was found; given also the results of the FE simulation, tread-braked wheels are expected to fail with the same mechanisms of non tread-braked ones, and the concerns about the detrimental effect of tread-braking on high-speed train wheels appear unconfirmed, at least in the conditions admitted by the standard for friction brakes in high speed trains.
ACKNOWLEDGEMENTS The authors wish to thank Silvio Bonometti and Andrea Danesi for their support in the experimental activities.
30
REFERENCES [1] M. Faccoli, C. Petrogalli, M. Lancini, A. Ghidini, A. Mazzù, Effect of desert sand on wear and rolling contact fatigue behaviour of various railway wheel steels, Wear 396-397, 2018, 146 161. [2] A. Mazzù, C. Petrogalli, M. Lancini, A. Ghidini, M. Faccoli, Effect of wear on surface crack propagation in rail-wheel wet contact, Journal of Material Engineering and Performance 27(2), 2018, 630-639. [3] A. Ekberg, E. Kabo, K. Karttunen, B. Lindqvist, R. Lunden, T. Nordmark, J. Olovsson, O. Salomonsson, T. Vernersson, Identifying the root causes of damage on the wheels of heavy haul locomotives and its mitigation, Proceedings of the Institution of Mechanical Engineers Part F: Journal of Rail and Rapid Transit 228(6), 2014, 663-672. [4] D. Nikas, J. Ahlström, A. Malakizadi, Mechanical properties and fatigue behaviour of railway wheel steels as influenced by mechanical and thermal loadings, Wear 366-367, 2016, 407 - 415. [5] K. Cvetkovski, J. Ahlström, B. Karlsson, Thermal degradation of pearlitic steels: influence on mechanical properties including fatigue behaviour, Material Science and Technology (27), 2011, 648 – 654. [6] J. Ahlström, B. Karlsson, Microstructural evaluation and interpretation of the mechanically and thermally affected zone under railway wheel flats, Wear 232 (1), 1999, 1-14. [7] J. Ahlström, B. Karlsson, Analytical 1D model for analysis of the thermally affected zone formed during railway wheel skid, Wear 232 (1), 1999, 15-24. [8] S. H. Avner: Introduction to Physical Metallurgy, 2nd ed. India: Tata Mcgraw Hill Pub., 1997. [9] J. Jergéus, Martensite formation and residual stress around railway wheel flats, Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science 212(1), 1998, 69 - 79
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[10] K. Handa, Y. Kimura, and Y. Mishima, Surface cracks initiation on carbon steel railway wheels under concurrent load of continuous rolling contact and cyclic frictional heat, Wear 268 (1), 2010, 50 – 58. [11] O. Orringer, D. E. Gray, Thermal cracking in railroad vehicle wheels subjected to high performance stop braking, Theoretical and Applied Fracture Mechanics 23(1), 1995, 55 - 65. [12] T. Vernersson, Thermally induced roughness of tread-braked railway wheels Part 1: brake rig experiments, Wear 236, 1999, 96–105. [13] T. Vernersson, Thermally induced roughness of tread-braked railway wheels Part 2: modelling and field measurements, Wear 236, 1999, 106-116. [14] M. Petersson, T. Vernersson, Noise-related roughness on tread braked railway wheelsexperimental measurements and numerical simulations, Wear 253, 2002, 301- 307. [15] G. Donzella, M. Scepi, L. Solazzi, F. Trombini, The effect of block braking on the resiudal stress state of a solid railway wheel, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 212, 1998, 145-158. [16] D. Peng, R. Jones, T. Constable, A study into crack growth in a railway wheel under thermal stop brake loading spectrum, Engineering Failure Analysis 25, 2012, 280–290. [17] D. Peng, R. Jones, T. Constable, An investigation of the influence of rail chill on crack growth in a railway wheel due to braking loads, Engineering Fracture Mechanics 98, 2013, 114. [18] M. Faccoli, A. Ghidini, A. Mazzù, Experimental and numerical investigation of the thermal effects on railway wheels for shoe-braked high-speed train applications, Metallurgical and Materials Transactions A 49(10), 2018, 4544 - 4554. [19] M. Faccoli, A. Ghidini, A. Mazzù, Changes in the microstructure and mechanical properties of railway wheel steels as a result of the thermal load caused by shoe braking, Metallurgical and Material Transaction A (2019), 50(4), 1701-1714. 32
[20] S. Teimourimanesh, T. Vernersson and R. Lundén, Thermal capacity of tread-braked railway wheels. Part 1: Modelling, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 230(3), 2016, 784-797. [21] S. Teimourimanesh, T. Vernersson and R. Lundén, Thermal capacity of tread-braked railway wheels. Part 2: Applications, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 230(3), 2016, 798-812. [22] S. Teimourimanesh, T. Vernersson, R. Lundén, F. Blennow and M. Meinel, Tread braking of railway wheels – temperatures generated by a metro train, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 228(2), 2014, 210-221. [23] T. Vernersson, S. Caprioli, E. Kabo, H. Hansson and A. Ekberg, Wheel tread damage: a numerical study of railway wheel tread plasticity under thermomechanical loading, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 224, 2010, 435-443. [24] S. Caprioli, T. Vernersson and A. Ekberg, Thermal cracking of a railway wheel tread due to tread braking – critical crack sizes and influence of repeated thermal cycles, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 227(1), 2013, 1018. [25] S. Caprioli, T. Vernersson, K. Handa, K. Ikeuchi, Thermal cracking of railway wheels: Towards experimental validation, Tribology International 94, 2016, 409–420. [26] F. Khalil, M. Kavoosi, Effect of lubricant on surface rolling contact fatigue cracks, Advanced Materials Research 97-101, 2010, 793 - 796. [27] T. Makino, T. Kato, K. Hirakawa, The effect of slip ratio on the rolling contact fatigue property of railway wheel steel, International Journal of Fatigue 36(1), 2012, 68 - 79. [28] M. Faccoli, L. Provezza, C. Petrogalli, A. Ghidini and A. Mazzù, Effects of full-stops on shoe-braked railway wheel wear damage, Wear 428-429, 2019, 64-75. 33
[29] T. Vernersson, M. Petersson, M. Hiensch, Thermally induced roughness of tread braked railway wheels, 12th International Wheelset Congress, Qingdao, China, 21–25 September 1998, 68-75. [30] A. Ghidini, A. Mazzù, C. Petrogalli, M. Faccoli, Una famiglia di acciai innovativi per ruote monoblocco forgiate-laminate progettati per ambienti sabbiosi (An innovative steel grade family for forged-rolled solid wheels designed for sandy environments), Ingegneria Ferroviaria 9 (2018), 729-742. [31] M. R. K. Vakkalagadda, D. K. Srivastava, A. Mishra, V. Racherla, Performance analyses of brake blocks used by Indian Railways, Wear 328-329, 2015, 64–76. [32] A. Mazzù, A. Ghidini, N. Zani, M. Faccoli, Study of wheel/rail material coupling in presence of solid contaminants, Proceedings of the 11th International Conference on Contact Mechanics and Wear of Rail/wheel Systems, CM 2018, 24-27 September 2018, Delft, (Netherlands), 701710. [33] R. Lewis, R. S. Dwyer-Joyce, U. Olofsson, J. Pombo, J. Ambrosio, M. Pereira, C. Ariaudo, N. Kuka, Mapping railway wheel material wear mechanisms and transitions, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 224(3), 2010, 125-137. [34] A. Mazzù, L. Solazzi, M. Lancini, C. Petrogalli, A. Ghidini, M. Faccoli, An experimental procedure for surface damage assessment in railway wheel and rail steels, Wear 342-343 (2015) 22-32. [35] I. Bodini, C. Petrogalli, M. Faccoli, M. Lancini, S. Pasinetti, G. Sansoni, F. Docchio, A. Mazzù, Evaluation of wear in rolling contact tests by means of 2D image analysis, Wear 400401 (2018), 156-168. [36] S. Hogmark, S. Jacobson, O. Vingsbo, ASM Handbook on Friction, Lubrication and Wear Technology, P. Blau (Ed.), 1992, 18, 176-183. 34
[37] Bharat Bhushan Ed., Modern tribology handbook vol. II, 2001, CRC Press, Boca Raton, Florida, USA, 774 - 776. [38] S. Abbasi, S. Teimourimanesh, T. Vernersson, U. Sellgren, U. Olofsson, R. Lundén, Temperature and thermoelastic instability at tread braking using cast iron friction material, Wear 314, 2014, 171–180. [39] A. Mazzù, Surface plastic strain in contact problems: prediction by a simplified non-linear kinematic hardening model, Journal of Strain Analysis for Engineering Design 44 (2009), 187199. [40] A. R. S. Ponter, L. Afferrante, M. Ciavarella, A note on Merwin’s measurements of forward flowin rolling contact, Wear 256 (2004), 321-328. [41] M. Faccoli, C. Petrogalli, M. Lancini, A. Ghidini, A. Mazzù, Rolling contact fatigue and wear behavior of high-performance railway wheel steels under various rolling-sliding contact conditions, Journal of Materials Engineering and Performance 26(7) (2017) 3271-3284.
35
• • • • •
Cast iron is transferred from brake blocks to wheel rim during arrest shoe braking. The transferred cast iron layer is removed during wheel-rail dry contact. Ratcheting, wear and formation of surface cracks occur in dry wheel-rail contact. Surface cracks propagation occurs in subsequent wet contact. The role of arrest shoe braking on final damage appears not relevant.
S0043-1648(19)30663-5
DOI:
https://doi.org/10.1016/j.wear.2019.203005
Reference:
WEA 203005
To appear in:
Wear
Received Date: 12 April 2019 Revised Date:
2 July 2019
Accepted Date: 7 August 2019
Please cite this article as: A. Mazzù, L. Provezza, N. Zani, C. Petrogalli, A. Ghidini, M. Faccoli, Effect of shoe braking on wear and fatigue damage of various railway wheel steels for high speed applications, Wear (2019), doi: https://doi.org/10.1016/j.wear.2019.203005. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2019 Published by Elsevier B.V.
Effect of shoe braking on wear and fatigue damage of various railway wheel steels for high speed applications Angelo Mazzùa*, Luca Provezzaa, Nicola Zania, Candida Petrogallia, Andrea Ghidini b, Michela Faccolia a
Department of Mechanical and Industrial Engineering, University of Brescia, via Branze 38, 25123 Brescia, Italy b
Lucchini RS, Via G. Paglia 45, 24065 Lovere (BG), Italy * [email protected], tel: (+39) 030 371 5525
Abstract The effect of shoe braking on three railway wheel steels was studied by means of bi-disc tests, in order to evaluate the applicability of shoe braking as an emergency stop braking system in highspeed passenger trains. Disc-shaped specimens, extracted from real wheels, were initially paired with discs extracted from cast iron brake blocks, in condition of rolling and sliding such to reach the typical temperature of the surface of a wheel rim during stop braking. Subsequently, the wheel discs were coupled with rail steel discs in rolling and sliding dry contact. Some of them were cut and analysed; other ones were subsequently subjected to rolling and sliding contact with rail specimens in wet condition, with various test duration. Measurements of the coefficient of friction and of the weight loss were carried out during the tests; at the end of the tests, hardness measurements and microstructure observation were made on the disc cross-sections. The main damage phenomena observed in the specimens subjected to wheel-brake contact and to dry wheel-rail contact were ratcheting, wear and nucleation of surface cracks. Traces of cast iron, previously transferred from the brake specimens to the wheel ones, were detected. The main damage phenomenon observed in the specimens subjected to wheel-brake, dry wheel-rail and wet wheel-rail contact was rolling contact fatigue, developing by the propagation of surface cracks due to the pressurization of the water entrapped inside the cracks. All of these phenomena were found to be similar to those 1
observed in wheel specimens subjected to wheel-rail contact only. The effect of the cast iron traces on surface crack propagation was evaluated by means of finite element simulations and it was found to be limited. Therefore, the application of shoe braking as emergency stop braking system to high speed trains did not evidence any relevant contraindications.
Keywords: High speed trains, railway wheels, shoe braking, rolling contact fatigue, wear.
1. Introduction Shoe braking is obtained by direct pressure of brake blocks on the wheel tread. It is typical for freight trains and frequent for metro and suburban trains, either as unique braking system or coupled with other ones. For high-speed passenger transportation, the EN 15734-1 standard recommends electro-dynamic braking as principal system; however, shoe braking is allowed as emergency braking system in coupling with the main one, provided that the braking power is limited and it is used only when the train speed is below 120 km/h. Notwithstanding the problems that can be related to the direct action of the brake blocks on the wheel tread, an interest on the application of shoe braking to high speed trains has recently raised, due to its cleaning effect from environmental contaminants that can have a role on wheel damage [1]-[3]. Furthermore, in comparison with other friction braking systems such as disc braking, shoe braking is advantageous in terms of axle weight and cost. Many authors studied the damage induced by the contact loads coupled with the heat generated by friction at the block-tread interface. Some authors [4]-[11] focused especially on the effects of thermomechanical loads on the material microstructure, including pearlite spheroidization or occasionally White Etching Layer (WEL) formation. Vernersson et al. [12]-[14] studied both numerically and experimentally the tread corrugation due to thermal loads, identifying a non-even wear mechanisms on the tread. Some studies [10], [11], [15] highlighted the role of these 2
phenomena in favoring the nucleation of surface cracks on the wheel, also due to the change of the residual stresses from tension to compression as a consequence of thermal cycles. Peng et al. [16], [17], studying the effect of thermal loads on crack propagation, distinguished stop braking, which is the stop of a train from a given speed, from drag braking, which is acted when the train speed has to be kept constant along a slope. They found that in drag braking the tread temperature can rise up to about 680° C, whereas in stop braking it is limited to about 200 °C. Faccoli et al. [18], [19] studied the changes of microstructure and mechanical properties of some wheel steels under different thermal histories reproducing various braking operations. They found that, whereas the temperatures that can be reached in drag braking can induce phase transformation, under temperatures typical of stop braking only slight changes can be observed. Teimourimanesh et al. [20] elaborated a temperature-dependent elastic-plastic material model, coupled with a fatigue model, for predicting the wheel life in presence of thermomechanical loads due to tread braking and to wheel-rail contact. In [21], they applied this model to the case of a metro train, considering both stop braking and drag braking, finding also that in the case of repeated stop brakings the fatigue life is controlled by mechanical loads rather than by thermal loads. In [22] they measured the temperature reached in a metro train, finding that it does not exceed 250°C. Vernersson et al. [23] studied the effect of thermomechanical load cycles on Rolling Contact Fatigue (RCF), especially considering the effect of tread temperatures up to 300°C on ratcheting. Caprioli et al. [24] studied the propagation of cracks due to thermal loading induced by tread braking in heavy haul applications. They found that fully functional brake systems on heavy haul trains are not likely to induce thermal crack propagation in stop braking, unless in the case of severe drag braking due to malfunctioning brakes. Caprioli et al. [25] studied RCF induced by ratcheting in stop braking with both a numerical and experimental approach, highlighting the role of repeated braking cycles in the formation of surface cracks.
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Overall, these studies identified the drag braking on heavy haul trains as the most severe application for tread-braked wheels: indeed, the high temperatures reached in this condition are able to induce microstructural changes, tensile residual hoop stress and thermal fatigue. Stop braking, especially in passenger trains, at most is expected to induce surface cracks, in addition to those due to ordinary mechanical loads. However, in high speed applications the damage related to stop braking cannot be neglected. High speed trains are characterized by a lower axle load than freight trains and by rarer stop braking operations with respect to metro trains. On the other hand, many operators are interested on extending the mileage between maintenance operations on high speed train wheels, in order to reduce the costs of train stops. This way, the wheels tend to be subjected to more loading cycles between the wheel maintenance operations, due to their higher speed. Even though microstructure changes or significant thermal cracks are not expected, small surface cracks generated during tread braking can be preferential sites for RCF initiation, especially under the action of fluid contaminants, which promote crack propagation by means of the pressurization of the fluid entrapped inside the cracks [2], [26], [27]. In a recent study [28], the effect of shoe braking by cast iron blocks on various wheel steels has been studied by means of bi-disc contact tests. Cylindrical specimens extracted from real wheels were put in rolling and sliding contact against cylindrical specimens extracted from cast iron brake blocks, in such a working condition to reproduce on the wheel specimens the typical temperature of a real wheel tread during the stop braking of a high speed train (around 240°C). The authors found that the damage mechanisms occurring at the surface of the wheel specimens were wear, ratcheting and surface crack nucleation. In addition, they documented a mechanism of material transfer from the brake specimens to the wheel ones, which is thought to play a key role in the tread damage. A similar phenomenon was observed even by Vernersson et al. [29] on real wheels, although they did not study it in depth. Due to the material transfer from the brake specimen to the wheel one, a “third 4
body” layer is generated on the surface of the latter. When it is detached, it also involves the steel substrate, probably promoting the nucleation of surface cracks. Indeed, in [28] there was evidence of surface cracks in the wheel steel layer starting from zones where the third body layer had been detached. This paper illustrates a forward step from the achievements described in [28]. The aim is to study the evolution of the surface damage when shoe braked wheels are put in contact with rails, both in dry and wet contact. This investigation was carried out again by means of bi-disc tests, using specimens of various wheel steels. These specimens were first put in dry contact with cast iron brake block specimens, then again in dry contact with rail steel specimens. For some of them, this procedure was followed by a session of wet contact with rail steel specimens. The damage in the tested wheel materials was evaluated in terms of wear, strain hardening, crack nucleation and propagation.
2. Experimental tests: materials and methods A test procedure was designed to experimentally simulate the effect of shoe braking on the wheel tread, considering its consequences on dry and wet wheel-rail contact subsequently occurring. The tests were carried out on a bi-disc bench whose schematic is shown in Figure 1. The discs were mounted on two shafts driven by independent engines, one of which can be displaced orthogonally to the shaft axis by means of a hydraulic piston, which also applies the contact load.
5
Specimen 2 (wheel)
Mobile mandrel
Torque sensor
Fixed mandrel
Specimen 1 (brake-rail)
Encoder Hydraulic cylinder
Figure 1. Schematic of the bi-disc testing machine. Three different wheel steels were the object of the investigation, identified in the following as HYPERLOS®, CLASS B and SANDLOS®. HYPERLOS® is an upgrade of the EN13262 ER7 steel [18], [19]. The HYPERLOS® steel fulfils the requirements of the EN13262 standard, it has a better combination of strength and toughness with respect to the traditional ER7 steel thanks to the higher content of C, Mn, Si and V and a well-balanced chemical composition. The AAR M107/M208 CLASS B steel is widely used in North America, recommended for freight cars in interchange service and for use on locomotives. For passenger car service, it is used for high-speed service with severe braking conditions and heavier wheel loads. SANDLOS® steel is an upgrade of the CLASS B steel [1], [30], with a higher content of Mn and Si. Their chemical composition and mechanical properties are reported in Table 1. The tensile properties were obtained by the supplier using standard specimens extracted from the components, according to EN 6892-1 Standard. The Brinell hardness was measured on the radial section of the wheel rims in accordance with EN ISO 6506-1 Standard, in the same position as that of the disc extraction. All of the wheel steels were supplied by
6
Lucchini RS. The wheel steels were tested in coupling with a brake block cast iron and subsequently with the rail steel R350HT EN 13674-1. Cast iron was chosen for the block material as its properties are well-known and less subject to the production process than other brake block materials [31]. The chemical composition and the nominal mechanical properties of the rail steel are shown again in Table 1; the chemical composition and the Brinell hardness of the brake block cast iron are shown in Table 2. The wheel discs had 80 mm diameter and 20 mm thickness; the brake and the rail discs had 60 mm diameter and 15 mm thickness. All of the specimens were extracted from real components according to the schematic shown in Figure 2
Chemical composition [wt%]
HYPERLOS® CLASS B SANDLOS® R350HT 0.51 0.65 0.63 0.63 C 0.78 0.63 0.84 1.095 Mn 0.38 0.26 0.88 0.296 Si 0.002 0.001 0.001 0.018 S 0.015 0.012 0.009 0.01 P Ultimate tensile strength [MPa] 885 990 1142 ≥1175 Yield strength [MPa] 568 642 690 Elongation [%] 19 14 15 ≥9 Hardness HB 280 315 322 355 Table 1. Main chemical elements and mechanical properties of the wheel steels. Cast iron chemical composition [wt%] C S P Mn Cr Ni Mo Cu Si V Al Ti 3.03 0.18 1.70 0.61 0.10 0.05 0.01 0.15 1.66 0.006 0.04 0.05 Table 2. Chemical composition and Brinell hardness of the brake block cast iron.
HB 230
7
brake block
wheel rail
Figure 2. Schematic of specimen extraction from a wheel, a brake block and a rail Two series of tests were carried out. In the first one a wheel disc was initially paired with a brake disc and cycled for 2000 cycles (“braking step” in the following); subsequently, the brake disc was replaced with a rail disc and the wheel disc was cycled with follower role for 10000 cycles in dry contact (“dry step” in the following). This test was repeated three times for each wheel material (test 1, test 2 and test 3). The second series of tests consisted of the same two steps of the first one, followed by another step with the addition of liquid to the wheel-rail interface (“wet step” in the following). The liquid was water (with the addition of 10% glycol to protect the test bench from corrosion), ejected towards the contact interface with a flow of 6×10-6 m3/s. For each material, four tests were carried out, characterized by different durations: 30000 cycles (test 4), 20000 cycles (test 5), 10000 cycles (test 6) and 5000 cycles (test 7). In all the tests, the cycles were counted as the rotations of the wheel steel specimen. The working conditions of the braking step, in particular the sliding speed and the contact load, were chosen to have a friction dissipated power able to obtain a steady temperature on the specimen contact surface as that resulted from the thermal simulation presented in [28], which was aimed at estimating the typical temperatures reached on the tread of a full-scale wheel during stop braking. The number of cycles was chosen to let the surface temperature of the specimen stabilize. These 8
working conditions were first estimated and subsequently tuned per trials in preliminary tests. The contact load in the dry and wet steps was chosen in order to reproduce a typical value of the wheelrail contact pressure in curve [32]; the rolling and sliding speeds in these steps were chosen to have a sliding-to-rolling ratio of 1% and, with the given load, to let the discs work in the field of “mild wear” according to the wear maps given for instance by Lewis et al. [33]. Overall, these working conditions are expected to be more severe than in real wheels, which are subjected to such contact pressure and sliding only in curve, e.g. intermittently and not continuously: these were chosen just to compare the behavior of the three steels. The working conditions in the different steps of the tests are given in Table 3.
Specimen Rolling speed (rpm) Diameter Tangential speed (m/s) Sliding speed ݒ௦ (m/s) Contact width (mm) Contact load ( ܨN) Contact Hertz pressure (MPa)
Braking step Wheel Brake 175 -175 80 60 0.73 -0.55 1.28 15 2000 529
Dry and wet step Wheel Brake 373 502.5 80 60 1.562 1.579 0.017 15 8636 1100
Table 3. Working conditions in the different step of the tests The coefficient of friction was obtained from the torque signal coming from a sensor mounted between the displaceable specimen shaft and the transmission, which was elaborated through the signal acquisition system and the procedure detailed in [34], [35]. The temperature on the contact surface of the specimens during the braking step was checked by means of a thermographic camera following the procedure detailed in [28]. The weight variation of the specimens was measured by means of a precision balance with a resolution of 0.001 g after cleaning in a bath of ethanol with ultrasonic vibrations; the weight measurements were taken before the tests and at the end of each step.
9
At the end of the tests, the wheel discs were cut along the mid plane orthogonally to the contact surface. The disc sections were ground, mechanically polished to a 1 µm finish, etched with 2% Nital and examined with a Leica DMI 5000 M light optical microscope. The deformation under the contact surface and the crack morphology were investigated and the damage mechanisms were identified. The length and the inclination of all cracks longer than 20 µm detected over a 120° sector of each section were measured. The Vickers hardness was measured on the wheel disc sections at varying distances from the contact surface to evaluate the steel work-hardening phenomenon and correlate it with the deformation beneath the contact surface. The tests were carried out using a 1000 g load and a dwell time of 15 s, in compliance with ASTM E384.
3. Experimental results 3.1. First test series (braking step + dry step) 3.1.1
Weight loss
Figure 3 shows the plots of weight loss versus number of cycles for the discs of the three wheel steels at the end of each step of the test (1= braking step, 2 = dry step). Table 4 shows the weight loss of the brake discs at the end of the braking step and the weight loss of the rail discs at the end of the dry step. At a first sight, the weight of all wheel discs decreases after both steps, as a consequence of wear. During the braking step, material transfer from the brake disc to the wheel disc resulting in the formation of a discontinuous “third body” layer was observed by the naked eye thanks to the darker colour of cast iron compared with steel. The brake material transfer resulted from the mechanical stresses at the contact surface, the heating of the contact surfaces and the agglomeration of wear debris [36]. The “third body” layer, as previously observed by the authors in similar tests [28], is worn away due to cycling, and detachment of steel from the wheel disc surface 10
also occurs, probably promoting crack nucleation. In addition, wear debris from both disc materials leads to three-body abrasive wear of the wheel disc surface. In the dry step of the test, wear and crack growth due to the contact with the rail disc lead to further material detachment (the mentioned phenomena are documented in the following). The HYPERLOS® discs are the most affected by wear, whereas the SANDLOS® discs are the least affected. This result is due to their different hardness [37], being the former the softest steel and the latter the hardest one. The weight loss of the brake discs is much higher than that of the wheel discs at the end of the braking step because of the lower hardness of the cast iron compared with the steel. The brake discs tested against the HYPERLOS® discs show slightly higher weight loss than those tested against SANDLOS® and CLASS B discs. The rail discs show similar weight loss regardless the counterpart. For all materials, the weight loss measured in the three repeated tests is similar, indicating a good reproducibility of the tests.
11
1
1
2
1
2
2
Figure 3. Weight loss of the wheel discs in the first test series (1= braking step, 2 = dry step). Weight loss [g] Brake discs
vs HYPERLOS®
vs CLASS B
vs SANDLOS®
17.83 ± 0.70
16.61 ± 0.88
16.50 ± 1.77
Rail discs
0.05 ± 0.02
0.04 ± 0.02
0.06 ± 0.03
Table 4. Weight loss of the brake discs at the end of the braking step and of the rail discs at the end of the dry step.
3.1.2
Coefficient of friction and temperature
Figure 4 shows the coefficient of friction measured during the tests. For all of the steels, in the braking step the coefficient of friction starts from a value ranging from 0.4 to 0.5 and subsequently tends to a value around 0.2-0.3, similarly to the tests performed in [28]. In the subsequent dry step, 12
the coefficient of friction increases, with some differences between the steels: it stabilizes at 0.350.45 for the HYPERLOS®, 0.3-0.4 for the SANDLOS® and 0.30-0.55 for the CLASS B, with more dispersion for the latter. The increment of the coefficient of friction can be reasonably correlated to the removal of the deposited layer of cast iron, which leads the two steels to direct contact. The variation of the wheel disc surface temperature was monitored during the braking step of the test by probing the point temperature on thermographic images near the contact region. The wheel disc surface reaches and then maintains about 230 °C after 1750 cycles in all tests, reproducing the temperature of the wheel rim during stop braking estimated by finite elements in a previous work [28]. This temperature is not high enough to induce microstructural changes in the wheel steels, but it leads to an increasing of the wear rate of the brake cast iron, as shown by Abbasi et al. [38], and it probably also promotes the adhesion of the “third body” layer on the wheel disc.
13
Figure 4. Coefficient of friction in the first test series (1= braking step, 2 = dry step).
14
3.1.3
Subsurface morphology and crack mapping
Figure 5 shows some representative cross-sections of the wheel and the rail discs cut at the end of the tests. The ferritic-pearlitic microstructure (with traces of bainite for the HYPERLOS® and SANDLOS® steels) is highlighted by the Nital etch. A layer with unidirectional plastic flow that tends to be aligned in the direction of the surface friction can be seen below the contact surface. This layer, as shown for instance by Mazzù [39] and Ponter et al. [40], is due to ratcheting, intended as the plastic shear strain accumulation at each load cycle during the tests, due to the high friction at the contact surface. Several little surface cracks can be observed in the samples of all three steels. These cracks start with a shallow angle to the surface and follow the plastically deformed material during their growth. They are caused by ratcheting and probably also promoted by the detachment of the material transferred from the brake disc to the wheel disc during the braking step, as also observed in [28]. The brake material is almost completely removed during the repeated wheel disc/rail disc contact (dry step of the test), only few traces were observed on the wheel discs at the end of the tests. A detail of brake material inside a surface crack in a HYPERLOS® specimen is shown in Figure 5. These results agree with the works of Vernersson et al. [12]-[14], who studied the evolution of the wheel roughness due to tread braking by means of full-scale block braking experiments. They found that cast iron blocks caused in stop braking condition higher roughness on the wheel surface than composition and sinter blocks did, and they explained this effect with the material transfer from the cast iron blocks to the wheel, although they did not study in depth this phenomenon. Given these evidences, it is reasonable to infer that similar phenomena also occur in a real wheel during stop braking, even though they cannot be easily observed on the wheel tread because the material transfer and removal occur subsequently at each wheel revolution.
15
Figure 5. Micrographs of the wheel discs (on the left) and paired rail discs (on the right). The wheel steels are: a) HYPERLOS®, b) CLASS B and c) SANDLOS®; the rail steel is 350HT (d).
16
As surface cracks can propagate when a liquid contaminant is added to the contact surface, the three wheel materials were characterized also on the basis of the surface cracks generated. All cracks exceeding 20 µm in length, detected over an arc of 120° on the cross-sections of all of the specimens, were mapped in terms of geometry, e.g. in terms of their extension x and z along the directions tangential and radial respectively of the disc. The HYPERLOS® developed the lowest number of cracks, CLASS B the highest. At a first sight, CLASS B was also the material that developed the longest cracks, especially in test 1, which was characterized by the highest coefficient of friction. This result is consistent with the weight loss shown in Figure 3: the higher weight loss of the HYPERLOS® discs implied the removal of short surface cracks, leaving fewer and shorter cracks than in the CLASS B discs. As a consequence, the HYPERLOS® steel offers fewer initiation points to major fatigue damage, such as shelling, when wet contact occurs.
17
Figure 6: Map of the measured extension of the cracks along the radial (z) and tangential (x) directions of the discs at the end of all of the first series tests. 3.1.4
Subsurface hardness
Figure 7 shows the hardness variation at increasing distance from the contact surface measured on the cross-section of the wheel steel discs. An increase of hardness can be observed under the contact surface in all tested discs. This effect is consistent with the pattern of deformation shown in Figure 5, as the material hardening is caused by the progressive accumulation of plastic strain under the contact surface. The maximum hardness is close to the contact surface where the plastic deformation is more severe; then the hardness gradually decreases at increasing distance from the contact surface as a consequence of the decreasing deformation. The HYPERLOS® and CLASS B discs show higher hardening compared with the SANDLOS® discs (∆HV1HYPERLOS® ~ 95, 18
∆HV1CLASS B ~ 108 and ∆HV1SANDLOS® ~ 35) an also larger hardened layer (~ 0.45 mm for HYPERLOS®, ~ 0.55 mm for CLASS B and ~ 0.35 mm for SANDLOS®). These results suggest better work-hardening behaviour of the former two steels compared with the latter. Furthermore, the CLASS B discs show a maximum hardness slightly higher than the other steels, which reach similar values. For all of the three steels, the hardness profile measured in the repeated tests is similar, confirming a good reproducibility of the tests.
HYPERLOS®
500
test 1 test 2 test 3
450
test 1 test 2 test 3
450 400
HV1
HV1
400
CLASS B
500
350
350
300
300
250
250
200
200 0
1
2 Depth (mm)
3
0
2 Depth (mm)
3
SANDLOS®
500
test 1 test 2 test 3
450 400
HV1
1
350 300 250 200 0
1
2 Depth (mm)
3
Figure 7: Vickers hardness profiles on the cross-section of wheel steels discs at the end of the first series of the tests.
19
3.2 Second series of tests (braking step + dry step + wet step) 3.2.1
Weight loss
The weight loss of the wheel discs was measured at the end of each step of the test (1= braking step, 2 = dry step, 3 = wet step). The weight loss of the rail discs was measured at the end of the dry and the wet step. Figure 8 shows the plots of weight loss versus number of cycles for the wheel and the rail discs. Generally, the weight loss of all the wheel discs increases slowly up to 22000 cycles (corresponding to the total duration of the braking + dry + wet steps, with last 10000 cycles in wet contact), then it increase more quickly. The increase of the weight loss in the wheel discs tested longer than 22 000 cycles is related to the onset of RCF due to the pressurization of the fluid entrapped inside the surface cracks. The weight loss of all of the rail discs slightly increased with the number of cycles (except for the rail disc paired with the CLASS B disc in test 4), but it was a magnitude order lower than for the wheel discs, as the entrapped fluid pressurization takes place only in the follower disc. The rail discs paired with the SANDLOS® discs showed the highest weight loss, being this wheel steel the hardest one.
20
HYPERLOS® test 4
test 4
test 5 2
test 6 test 7
1 1
2
3
Weight loss (g)
Weight loss (g)
RAIL 0.3
3
test 5
0.2
test 6 test 7
0.1
2
0
0 0
0
10000 20000 30000 40000 50000 Number of cycles
CLASS B
3
RAIL
test 6 test 7
1
2
3
Weight loss (g)
Weight loss (g)
test 4
test 5
2
0
test 5
0.2
test 6 test 7 0.1
2
3
0 0
10000 20000 30000 40000 50000 Number of cycles
0
SANDLOS®
3
RAIL test 4
test 6 1 1
test 7 2
3
0
Weight loss (g)
test 5
2
10000 20000 30000 40000 50000 Number of cycles
0.3
test 4 Weight loss (g)
10000 20000 30000 40000 50000 Number of cycles
0.3
test 4
1
3
test 5 0.2
test 6 test 7 2
0.1
3
0 0
10000 20000 30000 40000 50000 Number of cycles
0
10000 20000 30000 40000 50000 Number of cycles
Figure 8: Weight loss of the wheel and paired rail discs in the second series of tests (1 = braking step, 2 = dry step, 3 = wet step).
21
3.2.2
Coefficient of friction and temperature of the wheel disc surface
Figure 9 shows the coefficient of friction during the longest tests (42000 cycles) of the second test series for the three steels. In the braking and dry steps the behaviour of the steels was similar to that of the first test series. In the wet step, initially it fell down to about 0.15-0.18 due to the lubrication effect of the added fluid; however, it progressively increased stabilizing at about 0.2 due to the occurrence of surface rolling contact fatigue which increased the surface roughness. A similar behaviour was observed in the shorter tests of the second series. The surface temperature was measured again in the braking step and it was in the same range of the first test series. 0.5 Coefficient of friction
HYPERLOS® CLASS B
0.4
SANDLOS® 0.3 0.2 0.1
1
2
3
0 0
5000
10000
15000 20000 25000 30000 Number of cycles
35000
40000
45000
Figure 9: Coefficient of friction in the longest tests (42000 cycles) of the second test series for the three tested wheel steels (1 = braking step, 2 = dry step, 3 = wet step). 3.2.3
Subsurface morphology
Figure 10 shows some representative sections of the wheel discs cut at the end of test 6 (22000 total cycle, with last 10000 cycles in wet contact). In addition to the plastic deformation accumulated during the tests, incipient shelling is observed below the contact surface in all wheel discs. The surface cracks nucleated during the braking + dry steps of the test initially propagated obliquely from the surface to a certain depth following the plastic deformed material. Then, in some cases, 22
they branched towards the surface, even creating connections between adjacent cracks. This behaviour, observed only in the follower (in this case the wheel disc) is typical of crack propagation driven by the entrapped fluid pressurization, as shown in previous works [2], [41] and studied in terms of Fracture Mechanics by Makino et al. [27]. The damage observed on the wheel disc sections is consistent with the weight loss shown in Figure 8. The surface cracks originated during the dry steps of the test grew in the wet step involving a much thicker layer than the plasticised depth; however shelling was not detected up to 22000 cycles (10000 cycles in wet contact), as shown by the micrographs taken after 5000 wet cycles (see for example the HYPERLOS® disc in Figure 11). Therefore, wear is the main damage mechanism of all of the wheel discs up to 22000 cycles. RCF prevails over wear in wheel discs tested for longer durations leading to shelling and consequently to relevant damage and weight loss (see also the HYPERLOS® disc at the end of the test 4 in Figure 12a). It is worth noting that the damage observed on the wheel discs, at least as the crack morphology is concerned, is similar to the one that affects the real shoe-braked wheels, as proved by the comparison between the HYPERLOS® disc section at the end of test 4 (Figure 12a) and the section of an ER7 shoe-braked wheel (Figure 12b), despite the working conditions in the laboratory tests, due to the continuous sliding imposed, are probably more severe than in real wheels. According to Figure 8, the SANDLOS® discs perform better than the other wheel discs up to 22000 cycles. This result is due to the higher hardness of SANDLOS® steel compared with the other two steels, which guarantees better wear behaviour. When RCF becomes the main damage mechanism the behaviour of the three steels is similar.
23
Figure 10: Incipient shelling in the HYPERLOS® (a), CLASS B (b) and SANDLOS® (c) discs at the end of test 6 (22000 cycle).
Figure 11: Surface cracks in the HYPERLOS® disc at the end of test 7 (17000 cycles). 24
Figure 12: Shelling in the HYPERLOS® disc at the end of test 4 (42000 cycles) (a) and in a real shoe-braked ER7 wheel [courtesy of Lucchini RS] (b).
3.2.4
Subsurface hardness
Figure 13 shows the hardness variation at increasing distance from the contact surface measured on the section of the wheel steel discs at the end of test 6 (22000 cycles) and test 4 (42000 cycles). An increase of hardness is observed in all tested discs, which is consistent with the pattern of deformation shown in Figure 10 and Figure 11. The maximum hardness reached is sometimes lower than that measured at the end of the first series of tests due to the considerable detachment of material caused by RCF.
25
HYPERLOS®
500
test 6 - 22 000 cycles test 4 - 42 000 cycles
450
test 6 - 22 000 cycles test 4 - 42 000 cycles
450 400 HV1
400 HV1
CLASS B
500
350
350
300
300
250
250
200
200 0
1
2 Depth (mm)
3
0
1
2 Depth (mm)
3
SANDLOS®
500
test 6 - 22 000 cycles test 4 - 42 000 cycles
450 HV1
400 350 300 250 200 0
1
2 Depth (mm)
3
Figure 13: Vickers hardness profiles on the cross-section of wheel steels discs at the end of the 22 000 and 42 000 cycle tests.
4. Evaluation of crack propagation and discussion The experimental evidences of the tests show damage mechanisms similar to those observed on wheel steels subjected to dry and to alternated dry-wet contact, such as those published in [2] and [41], for instance. As shown in [28], shoe braking causes material transfer from the brake blocks to the wheel rim. However, the evidences of the present tests show that such layer is removed in a relatively low number of cycles in rolling-sliding contact with the rail, only leaving few traces of the block material, as the one shown in Figure 5a. The removal of the transferred cast iron presumably contributes to the formation of surface cracks, in addition to ratcheting, as suggested by 26
the micrograph shown again in Figure 5a. However, the crack size distribution shown in Figure 6 does not highlight a clear distinction between the cracks generated by the two different mechanisms. Surface cracks, whatever the mechanism that generated them, propagate when fluid is added, causing severe RCF damage. In order to evaluate whether the presence of cast iron traces into a surface crack have a specific role in its propagation, the case depicted in Figure 5a was simulated by Finite Elements (FE). In particular, two models of the cracked regions were built: one considering uniform steel properties for the whole mesh, the other one considering cast iron and steel properties for the respective regions of Figure 5a. The model, shown in Figure 14, is built as a fixed 2D rectangular body, representing the wheel specimen, surmounted by a 2D traveling roller, representing the rail specimen. The radius of the rail roller was 17.14 mm, correspondent to the equivalent radius of the contact in the experimental tests, according to the Hertz theory. The wheel body was modelled by quadratic plane strain elements, the rail roller by linear ones. The bottom body had a 100 µm long crack, inclined by 15° with respect to the contact surface with a 1.5 µm wide opening. The crack tip was modelled by the quarter point technique: this technique is able to create a singularity at the crack tip generating an accurate stress field for a linear elastic model. A fluid cavity interaction was imposed inside the crack, meaning that it was considered full of incompressible fluid and consequently its volume was kept constant during the simulation. In addition, unilateral contact between the crack faces was imposed; no friction interaction was imposed, due to the lubrication effect of the entrapped fluid. As the material properties are concerned, the darker region of Figure 14 was modelled with the cast iron elastic properties in the first simulation (elastic modulus 120 GPa and Poisson ratio 0.23), and with the steel elastic properties in the second simulation (elastic modulus 206 GPa and Poisson ratio 0.29). Steel properties were attributed to the rest of the model. A unit contact load of 575.7 N/mm was imposed between the rail and the wheel bodies, able to generate the same pressure distribution as in the experimental tests. 27
cast iron or steel
steel
Figure 14: Finite Element model of a cracked wheel specimen portion with a traveling rail specimen. The gray part represents the portion of material that was modelled as steel in the first simulation and cast iron in the second one. The applied Stress Intensity Factors (SIF) in mode I (ܭூ ) and mode II (ܭூூ ) are obtained from the displacements of the quarter nodes on the crack face of the collapsed elements, according to the Williams equations:
ாට ۓ ೝ ۖ ܭூ = ሺݒଵ − ݒଶ ሻ ଼ሺଵିఔమ ሻ మഏ
మഏ ۔ ாට ೝ ۖܭூூ = ሺݑଵ − ݑଶ ሻ ە ଼ሺଵିఔ మ ሻ
(3)
Where v1 and v2 are the Crack Opening Displacements (COD) of the quarter nodes of the elements at the crack tip, u1 and u2 are the Crack Shearing Displacements (CSD) of the same nodes, r is the distance of the quarter nodes from the crack tip, E and ν are the material elastic modulus and coefficient of Poisson respectively. In this equation the steel properties were used for both of the models, as the region surrounding the crack tip was in steel for both.
Figure 15 shows the ܭூ and ܭூூ variation during a load passage for the two models, where the x abscissa represents the distance of the contact point from the crack mouth on the contact surface. 28
The difference between the results obtained with the two models is limited: the mode I and mode II SIF ranges ∆ܭூ and ∆ܭூூ are about 11% and 6% higher for the model with the cast iron layer with
respect to the model with uniform material. This means that the cast iron traces left after the rollingsliding contact between the wheel and the rail do not substantially affect crack propagation. The zone of adhesion between the steel matrix and the cast iron layer could be a preferential site for crack propagation due to debonding; however, the experimental evidences after the wet sessions did not highlight a significance of such phenomenon, being rather the in-depth propagation and subsequent branching and crack coalescence the main damage mechanism. 2.5 2 1.5 K [MPa m0.5]
1 0.5 0 -0.5 KI (with cast iron layer) KI (uniform steel) KII (with cast iron layer) KII (uniform steel)
-1 -1.5 -2 -2.5 -0.75
-0.5
-0.25
0 x [mm]
0.25
0.5
0.75
Figure 15: Mode I and mode II SIF variation during a load passage with uniform steel material and with a layer in cast iron. 5. Conclusions The damage occurring in three wheel steels due to shoe braking and to subsequent rolling-sliding contact with the rail was evaluated by means of bi-disc tests. Both dry contact and consecutive drywet contact were applied to the wheel-rail pairings. Ratcheting and wear were the main damage 29
mechanisms observed in the specimens subjected to brake-wheel contact and subsequently to railwheel dry contact, evidenced by the weight loss, the surface hardening and the microstructural plastic flow. Several shallow surface cracks were generated as well. Only traces of the cast iron transferred onto the wheel specimens during the previous brake-wheel contact step were found: most of the transferred cast iron, indeed, was removed in the initial cycles of the dry wheel-rail contact, as witnessed by the increment of the coefficient of friction. In tests with consecutive drywet wheel-rail contact after the wheel-brake step, the final failure occurred by fatigue due to propagation of surface cracks enhanced by the pressurization of the entrapped fluid. No traces of residual brake material were detected after this phase. The effect of the cast iron traces on crack propagation was evaluated by finite elements, and only a slightly increment of the applied stress intensity factor due to the stuck cast iron was found. In conclusion, the material state at the end of the dry wheel-rail contact phase was found not dissimilar to what was observed in previous rollingsliding wheel-rail tests, and the final failure mechanism in the wet phase was the same as well. No specific damage mechanism related to the braking phase was found; given also the results of the FE simulation, tread-braked wheels are expected to fail with the same mechanisms of non tread-braked ones, and the concerns about the detrimental effect of tread-braking on high-speed train wheels appear unconfirmed, at least in the conditions admitted by the standard for friction brakes in high speed trains.
ACKNOWLEDGEMENTS The authors wish to thank Silvio Bonometti and Andrea Danesi for their support in the experimental activities.
30
REFERENCES [1] M. Faccoli, C. Petrogalli, M. Lancini, A. Ghidini, A. Mazzù, Effect of desert sand on wear and rolling contact fatigue behaviour of various railway wheel steels, Wear 396-397, 2018, 146 161. [2] A. Mazzù, C. Petrogalli, M. Lancini, A. Ghidini, M. Faccoli, Effect of wear on surface crack propagation in rail-wheel wet contact, Journal of Material Engineering and Performance 27(2), 2018, 630-639. [3] A. Ekberg, E. Kabo, K. Karttunen, B. Lindqvist, R. Lunden, T. Nordmark, J. Olovsson, O. Salomonsson, T. Vernersson, Identifying the root causes of damage on the wheels of heavy haul locomotives and its mitigation, Proceedings of the Institution of Mechanical Engineers Part F: Journal of Rail and Rapid Transit 228(6), 2014, 663-672. [4] D. Nikas, J. Ahlström, A. Malakizadi, Mechanical properties and fatigue behaviour of railway wheel steels as influenced by mechanical and thermal loadings, Wear 366-367, 2016, 407 - 415. [5] K. Cvetkovski, J. Ahlström, B. Karlsson, Thermal degradation of pearlitic steels: influence on mechanical properties including fatigue behaviour, Material Science and Technology (27), 2011, 648 – 654. [6] J. Ahlström, B. Karlsson, Microstructural evaluation and interpretation of the mechanically and thermally affected zone under railway wheel flats, Wear 232 (1), 1999, 1-14. [7] J. Ahlström, B. Karlsson, Analytical 1D model for analysis of the thermally affected zone formed during railway wheel skid, Wear 232 (1), 1999, 15-24. [8] S. H. Avner: Introduction to Physical Metallurgy, 2nd ed. India: Tata Mcgraw Hill Pub., 1997. [9] J. Jergéus, Martensite formation and residual stress around railway wheel flats, Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science 212(1), 1998, 69 - 79
31
[10] K. Handa, Y. Kimura, and Y. Mishima, Surface cracks initiation on carbon steel railway wheels under concurrent load of continuous rolling contact and cyclic frictional heat, Wear 268 (1), 2010, 50 – 58. [11] O. Orringer, D. E. Gray, Thermal cracking in railroad vehicle wheels subjected to high performance stop braking, Theoretical and Applied Fracture Mechanics 23(1), 1995, 55 - 65. [12] T. Vernersson, Thermally induced roughness of tread-braked railway wheels Part 1: brake rig experiments, Wear 236, 1999, 96–105. [13] T. Vernersson, Thermally induced roughness of tread-braked railway wheels Part 2: modelling and field measurements, Wear 236, 1999, 106-116. [14] M. Petersson, T. Vernersson, Noise-related roughness on tread braked railway wheelsexperimental measurements and numerical simulations, Wear 253, 2002, 301- 307. [15] G. Donzella, M. Scepi, L. Solazzi, F. Trombini, The effect of block braking on the resiudal stress state of a solid railway wheel, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 212, 1998, 145-158. [16] D. Peng, R. Jones, T. Constable, A study into crack growth in a railway wheel under thermal stop brake loading spectrum, Engineering Failure Analysis 25, 2012, 280–290. [17] D. Peng, R. Jones, T. Constable, An investigation of the influence of rail chill on crack growth in a railway wheel due to braking loads, Engineering Fracture Mechanics 98, 2013, 114. [18] M. Faccoli, A. Ghidini, A. Mazzù, Experimental and numerical investigation of the thermal effects on railway wheels for shoe-braked high-speed train applications, Metallurgical and Materials Transactions A 49(10), 2018, 4544 - 4554. [19] M. Faccoli, A. Ghidini, A. Mazzù, Changes in the microstructure and mechanical properties of railway wheel steels as a result of the thermal load caused by shoe braking, Metallurgical and Material Transaction A (2019), 50(4), 1701-1714. 32
[20] S. Teimourimanesh, T. Vernersson and R. Lundén, Thermal capacity of tread-braked railway wheels. Part 1: Modelling, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 230(3), 2016, 784-797. [21] S. Teimourimanesh, T. Vernersson and R. Lundén, Thermal capacity of tread-braked railway wheels. Part 2: Applications, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 230(3), 2016, 798-812. [22] S. Teimourimanesh, T. Vernersson, R. Lundén, F. Blennow and M. Meinel, Tread braking of railway wheels – temperatures generated by a metro train, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 228(2), 2014, 210-221. [23] T. Vernersson, S. Caprioli, E. Kabo, H. Hansson and A. Ekberg, Wheel tread damage: a numerical study of railway wheel tread plasticity under thermomechanical loading, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 224, 2010, 435-443. [24] S. Caprioli, T. Vernersson and A. Ekberg, Thermal cracking of a railway wheel tread due to tread braking – critical crack sizes and influence of repeated thermal cycles, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 227(1), 2013, 1018. [25] S. Caprioli, T. Vernersson, K. Handa, K. Ikeuchi, Thermal cracking of railway wheels: Towards experimental validation, Tribology International 94, 2016, 409–420. [26] F. Khalil, M. Kavoosi, Effect of lubricant on surface rolling contact fatigue cracks, Advanced Materials Research 97-101, 2010, 793 - 796. [27] T. Makino, T. Kato, K. Hirakawa, The effect of slip ratio on the rolling contact fatigue property of railway wheel steel, International Journal of Fatigue 36(1), 2012, 68 - 79. [28] M. Faccoli, L. Provezza, C. Petrogalli, A. Ghidini and A. Mazzù, Effects of full-stops on shoe-braked railway wheel wear damage, Wear 428-429, 2019, 64-75. 33
[29] T. Vernersson, M. Petersson, M. Hiensch, Thermally induced roughness of tread braked railway wheels, 12th International Wheelset Congress, Qingdao, China, 21–25 September 1998, 68-75. [30] A. Ghidini, A. Mazzù, C. Petrogalli, M. Faccoli, Una famiglia di acciai innovativi per ruote monoblocco forgiate-laminate progettati per ambienti sabbiosi (An innovative steel grade family for forged-rolled solid wheels designed for sandy environments), Ingegneria Ferroviaria 9 (2018), 729-742. [31] M. R. K. Vakkalagadda, D. K. Srivastava, A. Mishra, V. Racherla, Performance analyses of brake blocks used by Indian Railways, Wear 328-329, 2015, 64–76. [32] A. Mazzù, A. Ghidini, N. Zani, M. Faccoli, Study of wheel/rail material coupling in presence of solid contaminants, Proceedings of the 11th International Conference on Contact Mechanics and Wear of Rail/wheel Systems, CM 2018, 24-27 September 2018, Delft, (Netherlands), 701710. [33] R. Lewis, R. S. Dwyer-Joyce, U. Olofsson, J. Pombo, J. Ambrosio, M. Pereira, C. Ariaudo, N. Kuka, Mapping railway wheel material wear mechanisms and transitions, Proceedings of the Institution of Mechanical Engineers, Part F: Journal of Rail and Rapid Transit 224(3), 2010, 125-137. [34] A. Mazzù, L. Solazzi, M. Lancini, C. Petrogalli, A. Ghidini, M. Faccoli, An experimental procedure for surface damage assessment in railway wheel and rail steels, Wear 342-343 (2015) 22-32. [35] I. Bodini, C. Petrogalli, M. Faccoli, M. Lancini, S. Pasinetti, G. Sansoni, F. Docchio, A. Mazzù, Evaluation of wear in rolling contact tests by means of 2D image analysis, Wear 400401 (2018), 156-168. [36] S. Hogmark, S. Jacobson, O. Vingsbo, ASM Handbook on Friction, Lubrication and Wear Technology, P. Blau (Ed.), 1992, 18, 176-183. 34
[37] Bharat Bhushan Ed., Modern tribology handbook vol. II, 2001, CRC Press, Boca Raton, Florida, USA, 774 - 776. [38] S. Abbasi, S. Teimourimanesh, T. Vernersson, U. Sellgren, U. Olofsson, R. Lundén, Temperature and thermoelastic instability at tread braking using cast iron friction material, Wear 314, 2014, 171–180. [39] A. Mazzù, Surface plastic strain in contact problems: prediction by a simplified non-linear kinematic hardening model, Journal of Strain Analysis for Engineering Design 44 (2009), 187199. [40] A. R. S. Ponter, L. Afferrante, M. Ciavarella, A note on Merwin’s measurements of forward flowin rolling contact, Wear 256 (2004), 321-328. [41] M. Faccoli, C. Petrogalli, M. Lancini, A. Ghidini, A. Mazzù, Rolling contact fatigue and wear behavior of high-performance railway wheel steels under various rolling-sliding contact conditions, Journal of Materials Engineering and Performance 26(7) (2017) 3271-3284.
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Cast iron is transferred from brake blocks to wheel rim during arrest shoe braking. The transferred cast iron layer is removed during wheel-rail dry contact. Ratcheting, wear and formation of surface cracks occur in dry wheel-rail contact. Surface cracks propagation occurs in subsequent wet contact. The role of arrest shoe braking on final damage appears not relevant.