Effects of full-stops on shoe-braked railway wheel wear damage

Author’s Accepted Manuscript Effects of full-stops on shoe-braked railway wheel wear damage Michela Faccoli, Luca Provezza, Candida Petrogalli, Andrea Ghidini, Angelo Mazzù www.elsevier.com/locate/wear

PII: DOI: Reference:

S0043-1648(18)31496-0 https://doi.org/10.1016/j.wear.2019.03.006 WEA102834

To appear in: Wear Received date: 6 December 2018 Revised date: 4 March 2019 Accepted date: 4 March 2019 Cite this article as: Michela Faccoli, Luca Provezza, Candida Petrogalli, Andrea Ghidini and Angelo Mazzù, Effects of full-stops on shoe-braked railway wheel wear damage, Wear, https://doi.org/10.1016/j.wear.2019.03.006 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. 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.

Effects of full-stops on shoe-braked railway wheel wear damage Michela Faccolia*, Luca Provezzaa, Candida Petrogallia, Andrea Ghidinib, Angelo Mazzùa 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], voice: (+39) 030 371 5572

Abstract The purpose of this work is to gain a better understanding of the complex damage phenomena taking place at the railway wheel/brake block interface due to thermo-mechanical loading. Initially, full-stop braking was studied using Finite Element (FE) simulations to estimate the temperature reached in the wheel rim. Experiments to reproduce wheel damage were conducted with a two-disc machine using test conditions that were based on the results of the FE simulations. Three different wheel steels were tested against the same cast iron shoe material. The evolution of the wheel disc damage was studied at various numbers of cycles under fixed contact pressure and sliding speed. The friction coefficient and the temperature on the wheel disc surface were measured during the tests. At the end of the experiments, the wheel disc was examined and characterized. Cross-sections were observed with an optical microscope and the hardness was measured as a function of the depth to investigate the damage mechanisms that occurred at surface and subsurface. Material transfer from the shoe specimen to the wheel specimen results in the formation of a discontinuous “third body” layer, and that layer plays a key role in the evolution of the wheel disc damage. When the transferred layer of brake material is worn away, detachment of steel from the wheel disc surface occurs, probably promoting the crack nucleation. In addition, wear debris from both disc materials promotes three-body abrasive wear of the wheel disc surface. 1

Keywords: Railway wheel steel, tread braking, stop braking, wheel damage, rolling contact fatigue.

1. Introduction Shoe braking (also known as “block braking” or “tread braking”) is achieved by means of brake blocks applying pressure and friction directly to the wheel tread, thereby dissipating mechanical energy. Shoe braking is commonly used on freight trains and also finds wide application in metro and suburban trains, whose mode of operation involves frequent stops. The shoe brake may be either the sole braking system, or it may be combined with disc brakes and/or electrodynamic brakes. In high-speed trains, non-friction braking systems, specifically electro-dynamic brakes, are recommended as main braking systems by the EN 15734-1 European standard [1]. Alternative braking systems, including friction braking, are admitted in coupling with these ones, especially for emergency braking purposes. As a consequence, friction braking systems are usually added to electrodynamic brakes and come into operation only when the speed is below 120 km/h. Among the various friction braking systems, disc brakes are preferred to shoe brakes, as they do not involve the direct action of the brakes on the wheel; block brakes are admitted provided that the energy input to the wheel can be limited appropriately. However, block braking systems do have several advantages that need to be taken into consideration. First, they have a cleaning action on the wheel tread, as they remove third bodies such as dust, leaves, ice or other contaminants which are known to be damaging factors [2]-[4]. Secondly, their use eliminates the need for braking discs, which are rotating dead masses that increase the axle weight and, if they have a small eccentricity, at high speed are a potential source of vibration. Such reasons explain the recent interest in the application of shoe braking to high speed trains.

2

In Europe, in accordance with the European standard EN13262 [5], the most widely diffused steel for tread braked wheels is ER7. Outside Europe, in conformity with the American standard AAR M-107/M-208 [6], Class B, Class C and Class D steels are recommended for tread braking applications. However, given the harmful thermomechanical loads acting on the wheel tread, specific steels are now being designed for tread-braked wheels. For instance, an improved railway wheel steel with higher contents of C, Mn, Si and V than the traditional ER7 steel, and fulfilling the requirements of EN13262, was recently developed and put on the market for this purpose [7]. The energy dissipation due to shoe braking generates heat at the shoe-tread interface; consequently, there is heat entering both the wheel and the brake block. The heat transferred to the wheel through the tread increases the temperature and modifies the local and global stress and strain fields. The thermomechanical problem in shoe-braked wheels has been studied by both numerical and experimental approaches. Moyar and Stone [8], using a simplified elastic model, showed the role of thermal loading in surface cracking, especially in the case of non-uniformly distributed braking. Other experimental works [9]-[16] highlighted the microstructural modifications of the steel near the contact surface, such as pearlite spheroidization or occasionally White Etching Layer (WEL) formation, depending on the heating and cooling rates, the temperature reached, the effect of alloying elements on the austenitization temperature and the kinetics of the pearlite reaction during cooling. Vernersson et al. [17]-[19] used experiments, FE simulations and in-field reporting to study the role of thermal loads due to block braking on tread roughness (corrugation or waviness) by noneven wear mechanisms on the tread, due to the formation of hot spots on the tread surface. The nucleation of surface cracks on the wheel tread is promoted by the above phenomena [8], [15], [16], [20] also due to the change of the hoop residual stress from compression to tension as a result of the thermal loading. The thermal problem associated with block braking was studied by means of various modelling techniques. Vernersson et al. [21]-[23] calculated the temperature distribution 3

and the heat partition by means of increasingly complex FE models, considering the thermal interaction between the wheel, the brake blocks, the rail and the surrounding environment under various working conditions and with various brake block materials. They also validated their model by comparison with laboratory and in-field tests. Their analyses showed that cast iron blocks minimize the heat portion flowing into the wheel, whereas composite blocks maximize it: the percentage of the generated heat flowing into the wheel varied from 77% to 87% for cast iron blocks, rising to 96% with composite blocks. Tudor and Khonsari [24] also developed a thermal model and, among their results, showed that, during a rotation of the wheel, temperature peaks occur at the passage under the brake shoes. However, they also showed that the depth of thermal penetration is very small, so the temperature peaks affect a very thin layer. Kolonits [25] simplified the 3D transient problem by means of a half space model with a moving heat source on the edge, obtaining realistic results for heat sources with a limited width. A review of thermal models for approaching the problem of tread braking can be found in [26]. Peng et al. [27], [28] introduced a FE thermomechanical model to study the effect of thermal loads on crack propagation. The authors studied two types of braking: “stop braking”, e.g. the arrest of a train from an initial velocity, and “drag braking”, e.g. a continuous braking to keep the train speed constant along a slope. They found that drag braking is much more severe for the wheel tread, as far higher temperatures are reached: about 680° C in drag braking, compared with about 200 °C in stop braking. Faccoli et al. [29], investigated the effect of the various temperatures (up to 970 °C) that can be reached in block braking, on the microstructure and mechanical properties of various wheel steels: they found that the temperatures reached in drag braking are particularly dangerous, as they can induce phase transformation in the material microstructure. These phenomena have been studied especially for freight and urban use, because these are the main applications of block braking. This paper investigates its application specifically to high speed trains. High speed trains have a lower axle 4

load than freight trains but, on the other hand, are subject to a higher number of cycles. For this reason, the initiation of defects which cause high cycle Rolling Contact Fatigue (RCF) phenomena, in particular under the subsequent effect of environmental factors like rain or ice, are of particular interest for this application. On the other hand, the wheel treads of shoe-braked high speed trains are presumed to be subject to lower temperatures than freight trains, because the load-per-axle is significantly lower and drag braking operated by shoe braking systems is not foreseen. The approach of this study is substantially experimental, aimed at investigating the effect of pairing cast iron brake blocks with three different steels used for tread braked wheels. 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 [30]. Initially, a thermal Finite Element (FE) simulation was carried out by a simplified model in order to estimate the temperature field in the wheel rim during an emergency full-stop braking. Once this temperature field was determined, the experimental study of the tread damage was carried out by means of a two-disc testing machine, under working conditions that reproduced temperatures similar to those estimated by FE simulations on the contact surface of the wheel steel specimen. The behaviour of the three tested steels and the damage occurring at their respective specimens was compared.

2. Thermal finite element simulation of shoe braking Preliminary FE simulations of the heat transfer process during the full-stop of a high speed train by shoe braking were carried out, in order to estimate the temperature map in the wheel rim. The aim of these simulations was to determine a temperature target range to be reached on the contact surface of the wheel steel specimens in experimental bi-disc tests that are described in the following sections. A simplified model, based on the results obtained in the previously cited thermal and thermomechanical studies, was introduced. An axisymmetric model with a heat flux uniformly 5

distributed along the circumferential direction was considered, in accordance with the limited effect of the temperature peaks along the circumference found by Tudor and Khonsari [24] and with the brief persistence of such peaks over each point of the surface tread. The heat generated by the block-tread friction was estimated by assuming the braking force to be constant, thereby generating a constant deceleration. A train with a load-per-axle of 18 metric tons (a typical value for high speed trains) was supposed to be stopped from an initial velocity of hypothesis, the kinetic energy

120 km/h. Under this

to be dissipated at each wheel was estimated according to

Equation (1):

(1)

where

is the load-per-wheel (equal to a half of the load-per-axle). The potential energy to be

dissipated was considered null, as the train was presumed to be running across flat land; the kinetic energy related to the wheelsets rotational inertia was considered negligible.

is assumed to be

transformed into heat by the friction between the shoe and the wheel. If

is the constant braking deceleration, the braking time

can be calculated as follows:

(2)

The average dissipated power per wheel W is therefore:

(3)

6

of which a fraction

is dissipated by the wheel, the rest by the brake blocks.

The generated heat, equal to the dissipated power, is presumed to flow across the total contact area between the blocks and the wheel. The heat flux

, presumed uniform over this area, is therefore

calculated by the following equation:

(4)

where

is the wheel diameter and

is the width of the area of contact between the wheel tread and

the brake block. The heat generated at the tread-block interface was assumed to be exchanged with the surrounding elements and environment by convection and conduction. Radiation and transmission to the rail (rail chill) were considered negligible: this approximation is enforced by various numerical results (see for instance [23] and [28]) showing that convection and conduction are far and away the most common means of heat exchange in train wheels. Both the instantaneous heat generated by friction and the instantaneous heat exchanged by convection depend on the train velocity, so they decrease during stop braking. Whereas estimating the former is relatively simple, estimating the latter is not, as convection depends on a number of working and environmental parameters that are difficult to obtain without a specific experimental test. The same goes for the heat partition generated between the brake blocks and the wheel. In order to get a rough estimate of these quantities, the recommendations provided by the UIC B 169/RP 17 [31] standard (which advises considering only conduction and convection heat exchange) were considered. This standard, which is provided by the International Railway Union, recommends certain settings and parameter values to be used in thermal FE simulations of block braked wheels. These parameters were obtained by comparing FE simulation results with laboratory test results under a specific working condition, in particular on 920 mm diameter wheels rotating at a tangential speed of 60 km/h with 7

cast iron brake blocks. Although these conditions are not the same as in the present stop braking simulations, in particular with regard to the train speed time variation, the recommended parameters were obtained on real train wheels rotating at a constant speed corresponding exactly to the mean speed value during the stop braking to be simulated. For this reason, the approximation that considered heat partition, heating power and convection as constant, even in stop braking simulation was judged acceptable. According to the standard, the heat fraction dissipated by the shoes is estimated as 17% of the total heat power in the case of a new wheel and 18.6% of the total heating power in the case of a wheel at the end of its life, e.g. whose diameter has been reduced by the periodical regenerating reprofilings. Here, a wheel with tread profile conforming to the EN 13715 European standard [32] was considered; a diameter Dw = 920 mm was considered for a new wheel and a diameter Dw = 840 mm was considered for a wheel at the end of life. All of these working parameters are summarized in Table 1.

Table 1. Input parameters for stop braking in high speed trains. New End of life Wheel state 920 840 Wheel diameter Dw (mm) 80 80 Block width b (mm) 9 9 Load-per-wheel Mak (t) 2 0.5 0.5 Deceleration a (m/s ) 120 120 Initial velocity v (km/h) 67 67 Braking time T (s) 5000 5000 Dissipated energy per wheel E (kJ) 75 75 Average dissipated power per wheel W (kW) 0.83 0.814 Wheel heat fraction fW 2 0.269 0.289 Wheel heat flux HW (W/mm )

The wheel was modeled by an axisymmetric model, with the heat flux

across the surface 3 of

Figure 1. The convection coefficient h changes from the center to the periphery of the wheel depending on the interaction between the wheel body and the surrounding air. According to the UIC 8

B 169/RP 17 [31] standard, the gradient of h was simplified by a rough discretization into three regions with constant convection coefficient. The discretization regions are shown in Figure 1; the local values of h are listed in Table 2.

3 2

2

1

1

Figure 1. Regions of different convection coefficient h.

Table 2. Convection coefficient values in surface zones of Figure 1. Zone 1 2 3

h (W m-2 K-1) 65.3 55.9 32.6 (new wheel) 31.8 (wheel at end of life) 9

The value of h in region 3, which corresponds to the wheel-brake contact surface, was estimated according to Equation (5) [31]:

(

where

)

⁄

(

)

(5)

, with n = 2, is the number of brake shoes and L = 320 mm is a typical shoe-

wheel contact length. As

depends on the wheel diameter D, h in region 3 varies depending on the

wheel state (new or at end of life). According, once again, to the UIC B 169/RP 17 [31] standard, the temperature dependent thermal conductivity k and specific heat c reported in Table 3 were introduced. The initial room temperature was presumed to be 20°C.

Table 3. Temperature depending values of conductivity k and specific heat c. Temperature k (°C) (W m-1 K-1) 0 47.3 200 44.1 400 39.3 600 32.9 800 25.0

c (J kg-1 K-1) 440 510 570 630 700

In Figure 2 the temperature field in the wheel rim section is shown for the two analyzed cases, at the instant when the maximum surface temperature is reached: the peak temperature ranges from about 200°C in the case of a new wheel to about 270°C in the case of a wheel at the end of life.

10

Wheel at end of life

New wheel

Figure 2. Temperature maps during stop braking in the wheel rim sections of the analyzed cases.

3. Small scale experimental tests 3.1 Materials and methods Small scale experimental tests were performed with a two-disc test rig, to simulate experimentally the effect of shoe braking on the wheel tread. The test rig is described in detail in [33] and its layout is shown in Figure 3.

Specimen 2

Mobile mandrel

Torque sensor

Fixed mandrel

Specimen 1

Encoder Hydraulic cylinder

Figure 3. Two-disc testing machine.

11

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. The rolling contact tests were carried out pairing a wheel steel disc with a brake disc. The wheel discs were cylindrical with a diameter of 80 mm and a contact track width of 20 mm. They were machined out of new wheel rims, as close as possible to the running surface, with their axis perpendicular to the wheel tread. Three different wheel steels were investigated: HYPERLOS®, CLASS B and SANDLOS® steels. HYPERLOS® is improved ER7 steel, containing higher contents of C, Mn, Si and V, recently developed for shoe braked high speed train applications [7]. ER7 steel is one of the two steels permitted by the technical specification UIC 812-3 [34] for shoe braked wheels for freight cars or passenger transportation in Europe. CLASS B steel is widely used in North America. As specified by the American standard AAR M107/M208 [6], it must be used for freight cars in interchange service and is recommended 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 type CLASS B steel modified [2], [35], with a higher content of Mn and Si. All steels are supplied by Lucchini RS. Their chemical composition and mechanical properties are reported in Table 4. The tensile properties were obtained by the supplier using standard specimens extracted from wheel rims, according to European standard EN ISO 6892-1 [36]. The Brinell hardness was measured on the radial section of the rims in accordance with European standard EN ISO 6506-1 [37], in the same position from which the discs were extracted. The brake discs were extracted from cast iron brake blocks with the chemical composition shown in Table 5. They had a diameter of 60 mm, a contact track width of 15 mm and a Brinell hardness of 230 HB.

Table 4. Main chemical elements (wt %) and mechanical properties of the wheel steels. 12

[wt%]

Chemical composition

C Mn Si S P Ultimate tensile strength [MPa] Yield strength [MPa] Elongation [%] Hardness HB

HYPERLOS® CLASS B SANDLOS® 0.51 0.65 0.63 0.78 0.63 0.84 0.38 0.26 0.88 0.002 0.001 0.001 0.015 0.012 0.009 885 568 19 280

990 642 14 315

1142 690 15 322

Table 5. Chemical composition (wt %) of the brake block material. Chemical composition

C

S

P

Mn

Cr

Ni

Mo

Cu

Si

V

Al

Ti

[wt%]

Cast iron

3.03 0.18 1.70 0.61 0.10 0.05 0.01 0.15 1.66 0.006 0.04 0.05

The applied test conditions were thus set as summarized in Table 6; in particular, as shown in Figure 4, the specimens were put in rotation in the same sense, in order to increase the sliding speed. As detailed in next section, these conditions are such that the obtained surface temperature of the wheel specimens is in the range of the tread temperature of train wheels during a stop braking simulated in the previous section.

Figure 4. Rolling direction of the specimens.

13

Table 6. Test working conditions. Specimen

Wheel

Rolling speed (rpm) Diameter Width (mm) Tangential speed (m/s) Sliding speed (m/s2) Contact load (N)

175 80 20 0.73

Brake block -175 60 15 -0.55 1.28 2000

The test duration was 2000, 4000 and 8000 cycles. 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 extracted by the signal acquisition system and the procedure detailed in [38]. A thermographic VARIOSCAN 3022 camera was used to measure and record the temperature of the wheel disc contact surface during the brake-wheel tests. As the emissivity coefficient of the disc varies during the test with the surface state and temperature [39], [40], no fixed value can be set to get the correct temperature value from beginning to end of the test. The goal was to know the steady state contact temperature that matched the simulation temperature range. Therefore, an emissivity coefficient of 0.8 was set, compatible with an oxidized steel surface. Thermal images were recorded during the test and analyzed to find the steady state contact temperature. A post-test check was then performed to convert the temperature value of the images into the real temperature. The contact surface of a tested worn disc was partially covered with a known emissivity high temperature paint (ε = 0.98±0.01). The disc was heated up on a hot plate to get the same temperature value on the bare surface as observed during the test, still using ε = 0.8. At this point the temperature measured on the paint with ε = 0.98 leads to the real temperature value of 230 °C. The above process also allowed for the estimation of ε = 0.65 on the bare steel surface in this condition. Before the test the wheel and brake block discs were cleaned in a bath of ethanol with ultrasonic vibrations and weighed by a precision balance with a resolution of 0.001 g. The weight 14

variation (i.e. the difference between the final weight and the starting weight) of both wheel and brake discs was measured at the end of the tests. Then, the wheel disc was cut along the mid plane orthogonally to the contact surface. The section obtained was ground, mechanically polished to a 1 μm finish, etched with 2% Nital and examined using 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 damage evolution of the wheel disc at an increasing number of cycles was documented. Vickers hardness tests were carried out on the wheel disc section at varying distances from the contact surface in order to evaluate the steel work-hardening phenomenon and correlate it with the deformation beneath the contact surface. The hardness tests were performed using a 1000 g load and a dwell time of 15 s, complying with the American standard ASTM E384 [41]. 3.2 Experimental results 3.2.1 Weight variation The plots of weight variation (i.e. the difference between the final weight and the starting weight) versus number of cycles for the wheel and brake discs are shown in Figure 5. At first sight, the weight of all wheel discs slightly increases after the tests with a duration of 2000 cycles. The weight of SANDLOS® discs also increases after the tests that last 4000 cycles. All wheel discs show a weight loss at the end of the tests that last 8000 cycles. The brake discs always lose weight and the weight loss generally increases with the number of cycles due to progressive wear. These results can be explained as follows (the phenomena will be documented in section 3.2.3). During the shortest tests (2000 cycles), as a result of the ploughing action of the wheel steel disc asperities and the heating of the contact surfaces, fragments from the cast iron brake disc were first removed and then attached by adhesion to various preferential sites on the wheel disc contact surface. These preferential sites could be the highest asperities that made earliest contact with the counter-surface. 15

Repeated contact during the test resulted in the accumulation of fragments, which then united and formed a discontinuous “third body” layer on the wheel disc. These transferred patches are visible with the naked eye on the wheel disc contact surface at the end of the tests thanks to the darker colour of the cast iron compared with the steel. The material transfer results in the initial weight increase of the wheel disc. It continues after 2000 cycles, at the same time cracking and removal of the “third body” layer take place on the wheel disc contact surface also causing the detachment of steel particles and probably the initiation of surface cracks. In addition, the wear debris of both materials leads to the abrasive wear of the wheel disc surface. When these phenomena prevail over the brake material transfer, the overall result is the weight loss of the wheel discs. The weight changes measured at the end of the three repeated longest tests (8000 cycles) can almost be superimposed for all materials (except for one SANDLOS® disc), qualitatively indicating a good reproducibility of the tests. At the end of the longest tests, all wheel discs lose weight but the behaviour of the three steels is different: the HYPERLOS® discs show lower weight loss than the other two steels. The brake discs paired with the HYPERLOS® discs have the lowest weight loss compared with those paired with the other steels. These results were commented on, taking into account that the wear resistance of two surfaces in sliding contact depends on material hardnesses [42] and also considering the presence of the “third body” layer. HYPERLOS® disc/brake disc pairs underwent to the lowest wear during the tests because HYPERLOS® steel hardness (280 HB) is more similar to the brake material hardness (230 HB) than that of CLASS B and SANDLOS® steels (315 HB and 322 HB, respectively). The higher weight variation of SANDLOS discs at the end of the longest tests compared with CLASS B discs despite their similar hardness could be explained by the almost complete detachment of the third body layer, which will be documented in section 3.2.2.

16

Weight variation (g)

test 2

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0 -0.1 0 -0.2 test 1

Weight variation (g)

2000

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0 -0.1 0 -0.2 test 1

test 3

test 4

test 5

test 6

test 7

test 2

15 10 5 0

test 1

test 3

test 4

test 5

4000

6000

test 2

4000 6000 8000 10000 Number of cycles test 3 test 4 test 5 test 6 test 7

CAST IRON

20 15 10 5 0 0

test 6

2000

25

4000 6000 8000 10000 Number of cycles

test 1

SANDLOS®

2000

20

0

CLASS B

2000

CAST IRON

25

4000 6000 8000 10000 Number of cycles

Weight variation (g)

Weight variation (g)

test 1

HYPERLOS®

Weight variation (g)

Weight variation (g)

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0 -0.1 0 -0.2

2000 test 2

test 3

test 4

test 4

test 5

test 6

20 15 10

8000 10000 test 5

test 3

CAST IRON

25

5 0 0

2000

test 1

test 2

Number of cycles test 2

4000 6000 8000 10000 Number of cycles

test 6

4000 6000 8000 10000 Number of cycles

test 3

test 4

test 5

test 6

Figure 5. Weight variation of the wheel and brake discs.

These results highlight the positive effects of the correct choice of the wheel material/brake material pair on the service life, costs and also the environment. A decrease in wear means lower of wheel 17

maintenance and replacement costs, so it is extremely important for wheel life. Furthermore, it limits the wear of the brake block and consequently reducing the fine particle pollution caused by the friction. 3.2.2 Friction and disc surface temperature during the tests Figure 6 shows the coefficient of friction during the 8000 cycle tests for each steel (for the HYPERLOS® steel the data of test 4 are missing as the acquisition system failed during the test), obtained from the RMS of the torque signal acquired at 5 kHz over time windows of 1 s. In the initial hundreds of cycles the coefficient of friction shows very high fluctuations, likely due to an initial small error in the shape of the specimens and in the coupling during mounting. Subsequently, following the reciprocal shape adaptation of the specimens due to wear, the amplitude of the fluctuations of the coefficient of friction was reduced, although some cyclic variation still remained even after the initial run-in phase. This phenomenon can be correlated to the material transfer from the cast iron disc to the steel one, described in the previous section and documented in the following one. The material transfer from the brake disc to the wheel disc is hypothesized to be a discontinuous phenomenon, cyclically altering the coefficient of friction by means of a “third body” layer which is deposed and removed from the steel disc surface. At the end of the tests, the coefficient of friction oscillates around a mean value of 0.3 for HYPERLOS® and CLASS B, and 0.25 for SANDLOS®. The band of oscillation is wider for HYPERLOS®, whereas it is much narrower for CLASS B and SANDLOS®, sign of a more stable contact interaction. These results are slightly lower than the value reported by Bowden et al. [43]. According to the authors, cast iron shows a friction coefficient of 0.4 in sliding against steel due to the graphite in the cast iron matrix, which forms a layer at the sliding interface [42].

18

HYPERLOS® Coefficient of friction

0.8 Test 3

0.7

Test 5

0.6

0.5 0.4 0.3 0.2 0.1 0 0

2000

4000 Number of cycles

6000

8000

6000

8000

6000

8000

CLASS B Coefficient of friction

0.8 Test 1

0.7

Test 2

Test 3

0.6 0.5 0.4 0.3 0.2 0.1 0 0

2000

SANDLOS®

0.8 Coefficient of friction

4000 Number of cycles

Test 1

0.7

Test 2

Test 3

0.6 0.5 0.4 0.3 0.2 0.1 0 0

2000

4000 Numebr of cycles

Figure 6. Coefficient of friction of the tests. 19

The variation of the wheel disc surface temperature was monitored by probing the temperature on thermographic images near the contact region avoiding the brake disc reflection. The effective contact area is hidden in the images, therefore these values have to be considered as a lower boundary for the maximum surface temperature. The wheel disc surface reaches and then maintains a temperature of about 230 °C after 1750 cycles in all of the tests (Figure 7), reproducing the temperature of the wheel rim during stop braking obtained by the simulations. This temperature is not high enough to induce microstructural changes in the wheel steels, but it probably promotes the adhesion of the brake material on the wheel disc.

CLASS B

HYPERLOS®

300 Temperature ( C)

Temperature ( C)

300 250 200 150 100 50

250 200 150 100 50 0

0

test 1

2000 4000 6000 8000 10000 Number of cycles test 2 test 3 test 4 test 5 test 6 test 7

0

0

10000 test 6

SANDLOS®

300 Temperature ( C)

test 1

2000 4000 6000 8000 Number of cycles test 2 test 3 test 4 test 5

250 200 150 100

50 0 0

2000 4000 6000 8000 10000 Number of cycles test 1 test 3 test 4 test 5 test 6

Figure 7. Steady state temperature at the contact surface of the wheel disc. 20

3.2.3 Wheel disc cross-section observations Several representative cross-sections of the wheel discs after the tests lasting for 2000, 4000 and 8000 cycles are shown in Figures 8 - 10. Many of them were obtained by connecting several micrographs taken with the optical microscope. Nital etch highlights the microstructure of the three steels, which is ferritic-pearlitic with bainitic traces for HYPERLOS® and SANDLOS® steels. A thin layer with unidirectional plastic flow that tends to be aligned in the direction of the surface friction can be seen below the contact surface of all specimens. This layer is due to the unidirectional plastic strain accumulation (ratcheting) during the tests caused by the frictional sliding. Transferred brake material, which forms a discontinuous “third body” layer, can be observed on the contact surface of the specimens. Some micrographs of the HYPERLOS® disc cross-sections are shown in Figure 8. The “third body” layer is more extended after the 2000 cycle test than after the tests lasting for 4000 and 8000 cycles. This observation is consistent with the weight variation of the HYPERLOS® discs shown in Figure 5 and suggests that the “third body” layer is formed continuously and then removed during the test. Some small surface cracks in the wheel disc near the “third body” layer are observed in the samples at the end of the 8000 cycle test. The crack nucleation can probably be promoted by the “third body” layer detachment in addition to the plastic deformation (ratcheting) of the surface layer. It is well known that surface cracks are extremely dangerous in railway wheels. In particular, when they undergo wet contact due to rain or snow falling at the wheel-rail interface, they can grow rapidly, driven by the pressurization of the fluid trapped inside the crack, causing severe damage such as shelling [44], [45]. Some CLASS B disc cross-sections are shown in Figure 9: cracking and detachment of the “third body” layer can already be observed after the shortest tests. Some steel detachments were also seen after 4000 cycles. The formation of surface cracks in the wheel disc 21

near the “third body” layer is evident in the samples at the end of the longest tests. Figure 10 shows some SANDLOS® disc cross-sections. An extended “third body” layer is observed after 2000 and 4000 cycle tests, which is consistent with the increased weight shown in Figure 5. The “third body” layer is almost absent after the tests lasting for 8000 cycles and many steel detachments can be observed, which are coherent with the considerable weight loss shown in Figure 5. It is worth noting that the damage observed in the wheel discs is similar to that found in a real wheel.

22

2000 cycles

4000 cycles

8000 cycles Figure 8. Micrographs of HYPERLOS® disc sections of the tests lasting for 2000, 4000 and 8000 cycles.

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2000 cycles

4000 cycles

8000 cycles Figure 9. Micrographs of CLASS B disc sections of the tests lasting for 2000, 4000 and 8000 cycles.

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2000 cycles

4000 cycles

8000 cycles Figure 10. Micrographs of SANDLOS® disc sections of the tests lasting for 2000, 4000 and 8000 cycles.

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3.2.4 Subsurface hardness The hardness variation at increasing distances from the contact surface measured on the section of the wheel discs at the end of the tests is shown in Figure 11. An increase in hardness can be observed under the contact surface in all tested discs. This result correlates well with the pattern of deformation observed metallographically in Figures 8 - 10, because the progressive accumulation of plastic strain under the contact surface during the test resulted in the steel hardening. The maximum hardening is close to the contact surface where the plastic deformation is more severe, then the hardening gradually decreases at increasing distances from the surface as a consequence of the smaller deformation. The maximum hardness measured was generally slightly higher in the discs tested at 8000 cycles than those at 2000 and 4000 cycles because of the highest accumulation of plastic strain. The SANDLOS® discs reached the maximum hardness compared with the other steels, whereas the CLASS B discs show the maximum hardening after 8000 cycles (HV1CLASS B = 61, HV1SANDLOS® = HV1HYPERLOS® ~ 35). Moreover, the depth of the hardened layer is almost the same in the discs of all steels (~ 0.25 mm).

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test 7 - 2000 cycles test 6 - 4000 cycles test 3 - 8000 cycles HV1

HV1

HYPERLOS®

450 425 400 375 350 325 300 275 250

1

2 Depth (mm)

HV1

0

CLASS B

450 425 400 375 350 325 300 275 250

test 5 - 2000 cycles test 4 - 4000 cycles test 1 - 8000 cycles

0

3

1

2 Depth (mm)

3

SANDLOS®

450 425 400 375 350 325 300 275 250

test 5 - 2000 cycles test 4 - 4000 cycles test 2 - 8000 cycles

0

1

2 Depth (mm)

3

Figure 11. Vickers hardness profiles on the cross-section of the wheel discs

4. Discussion The investigations carried out on the wheel disc cross-sections make it easier to understand the effects on the tread damage of the thermomechanical interaction between wheel and brake block during full-stop braking. The same surface damage types, defined by Hogmark et al. [46], were observed in the discs of the three different wheel steels: - plastic deformation caused by mechanical stresses at the surface and accumulated during the tests (ratcheting) is of the utmost importance in the creation of surface damage that may lead to failure;

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- gain of material, such as material transfer from the counter-surface, resulting from the mechanical stresses at the surface, the heating of the contact surfaces and the agglomeration of debris. The resulting “third body” layer is typical of conformal sliding contacts [47]; - wear due to mechanical stresses at the surface involving continuous material loss; - surface cracking caused by ratcheting can lead to large-scale damage as RCF. The crack nucleation is probably promoted by the detachment of the “third body” layer. The overall damage of the wheel discs is the result of these coexisting and interacting damage types. The transfer of material from the brake to the wheel disc with the formation of a discontinuous “third body” layer plays a key role in the damage evolution of the wheel disc. When the transferred layer of brake material is worn away, it contributes to the detachment of steel from the wheel disc surface probably promoting the crack nucleation. In addition, the resulting debris of both materials leads to the abrasive wear of the wheel and brake disc surfaces. The phenomenon of material transfer from the brake to the wheel specimen was also mentioned by Vernersson et al. [17]-[19], [48]. The authors studied the evolution of the wheel roughness due to tread braking, which has deleterious effects on passenger safety and comfort due to the consequent vibration, noise and fatigue. Full-scale block braking experiments were performed on an inertia dynamometer both in drag braking and stop braking cycles. Higher levels of roughness were found using cast iron blocks than composition and sinter blocks, linked to the material transfer from the block to the ER7 steel wheel. The authors did not examine the material transfer phenomenon in depth, because it was beyond the range of their research. However, they reported that the transferred material was not very well bonded to the wheel surface and it was easily detached from the surface during cycling. It is worth noting that in the present work, despite the reduced scale of the experiments, a mechanism of material transfer and subsequent detachment similar to the one observed on full-scale experiment was observed. No evidence of the occurrence of this phenomenon was found from data reports 28

coming from in-field observations. However, it is reasonable to infer that similar phenomena occur even in a real wheel during stop braking, even though they cannot be witnessed by the final wheel status. The repeated contact of the wheel with the brake and the rail, also bearing in mind that the wheel-rail contact patch displaces laterally in curve covering a wide area on the wheel tread, is expected to completely remove the “third body” layer during service, thereby making it difficult to observe on serviced wheels. This mechanism is expected to promote the wheel damage in terms of wear and surface crack nucleation. The results obtained in this research suggest that during stop braking the interaction between the wheel and cast iron brake can promote the damage phenomena even at temperatures lower than those which can cause the hot spot formation on the wheel tread [48]. Finally, the comparison between the three investigated steels showed that HYPERLOS® steel has the best behaviour when coupled with cast iron, probably thanks to their similar hardness.

5. Conclusions The effects of full-stop braking on shoe-braked wheel damage were investigated by means of experimental tests, carried out by pairing discs in steels for shoe-braked wheels with discs in cast iron for brake blocks. The discs were tested in a rolling/sliding condition with very high sliding speed, so as to obtain the temperature typically present at the wheel tread during stop braking from 120 km/h. Three steels were investigated: HYPERLOS® steel, which has recently been developed for shoe braked high speed train applications complying with the European EN13262 standard, and CLASS B and SANDLOS® steels, complying with the American AAR M-107/M-208 standard. The main findings follow. 

The damage mechanisms occurring at the surface of the wheel specimens were wear, ratcheting and surface crack nucleation. Material transfer from the brake specimens to the wheel specimens was observed as well. 29



The mechanism of material transfer is thought to play a key role in the tread damage: in fact, a “third body” layer is generated. When the transferred layer of brake material is detached, it also involves the steel substrate, probably promoting surface crack nucleation; furthermore, the wear debris produced this way promotes abrasive wear.



The surface cracks generated this way can be preferential sites for rolling contact fatigue occurrence: in fact, especially in the presence of deposited fluids that penetrate the cracks and are pressurized at the passage of the load, these cracks can propagate upwards to cause spalling.



Considering the three tested steels, HYPERLOS® shows the best performance in terms of wear. This is probably due to the lower hardness, which is comparable to that of the brake specimens. In fact, the total wear (i.e. considering both the cast iron and the steel specimens), is lower with this pairing, reducing the presence of wear debris at the contact interface.

ACKNOWLEDGEMENTS The authors wish to thank Silvio Bonometti for his support in the experimental activities.

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Highlights

   

The damage phenomena taking place at the wheel/brake block interface were investigated The temperature map in the wheel rim was estimated by thermal FE simulations Experimental tests were performed with a two-disc machine to reproduce the wheel damage The brake material transfer plays a key role in the wheel disc damage evolution

36