Adhesive tool wear in the cold roll forming process

Wear 271 (2011) 2728–2745

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Adhesive tool wear in the cold roll forming process Alexander S. Galakhar a,∗ , Jeffrey D. Gates b , William J.T. Daniel a , Paul A. Meehan a a b

School of Mechanical and Mining Engineering, the University of Queensland, Brisbane, QLD, 4072, Australia UQ Materials Performance, the University of Queensland, Brisbane, QLD, 4072, Australia

a r t i c l e

i n f o

Article history: Received 14 September 2010 Received in revised form 16 April 2011 Accepted 10 May 2011 Available online 12 June 2011 Keywords: Other manufacturing processes Sliding friction Traction Sliding wear Steel Wear modelling

a b s t r a c t Tool wear manifests in change of roughness of roll surfaces and governs the surface quality of both coated and uncoated products. Decrease of friction and wear in the roll–strip interface will stabilize the performance of the cold roll forming (CRF) mill and ensure uniform process output. The scope of this paper includes an experimental study of adhesive wear of hardened tool steel blocks sliding against a mild steel wheel and field observation of tool wear in the CRF of zinc coated mild steel. Wear mechanisms are identified for the materials interacting during the CRF process. The results include an empirical value of tool wear coefficient for D2 steel sliding against unlubricated uncoated hot rolled AISI C1020 mild steel. © 2011 Elsevier B.V. All rights reserved.

1. Introduction Cold roll forming (CRF) is a process whereby a flat strip of metal is progressively formed into a desired cross-section by passing through a series of rolls arranged in tandem [1]. An inherent limitation of the CRF process is that sections of the forming roll pairs are in full sliding contact due to differences in sectional roll radii, wherever the strip cross-section is non-flat. This causes significant wear of tool surface and affects surface quality of the product. Hence tool wear prediction and friction modelling are important for optimising the future design of the CRF processes. The first research on wear in the CRF process [2] was devoted to study of roll–strip friction and strip surface damage on a slider-onstrip tribometer, and did not address the phenomenon of tool wear. Although the previous research on tool wear in sheet metal forming (SMF) [3–16] has in general not been related to CRF, some studies [4,5,10,16] have focused on reducing friction and have considered lubricant performance or wear-resistant tool coatings applicable to conditions of the CRF process. Early empirical studies of galling mechanisms in SMF [17,18] contained observations of sheet surface damage caused by galling, but did not give a cogent explanation of tool material loss in interaction with the soft strip coating. Later Christiansen and De Chiffre [6] identified three wear mechanisms: adhesive, abrasive and fatigue observed in SMF for deep draw-

∗ Corresponding author. Tel.: +61 04 1580 0431; fax: +61 07 3365 4799. E-mail addresses: [email protected] (A.S. Galakhar), [email protected] (J.D. Gates), [email protected] (W.J.T. Daniel), [email protected] (P.A. Meehan). 0043-1648/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2011.05.047

ing. An explanation of tool wear in SMF on the basis of asperity interactions clarifies the much stronger dependence of galling on roughness of a tool surface than on roughness of a sheet surface [19], but it leaves obscure the effects of humidity and atmospheric gas composition on friction and wear [20]. Hence Van der Heide and Schipper [21] add tribochemical wear to the above three wear mechanisms in SMF [6]. Boher et al. [3] conducted an empirical study of wear mechanisms for AISI D2 tool steel sliding conducted an empirical study of wear mechanisms for AISI D2 tool steel sliding against mild steel and stainless steel on a deep-drawing process simulator and distinguished adhesion and ploughing wear mechanisms on the surface of the D2 steel die. In order to quantify and/or model the development of tool wear in CRF, material and wear mechanism dependent values for wear coefficients are required. Vuong and Meehan [22] provided a means of predicting transitions in the wear mechanisms with frictional power density. The wear coefficient of a wide range of materials has been obtained in the past (e.g. [23,24]), but not for CRF tool conditions. Ersoy-Nürnberg et al. [7] studied tool wear in dies during deep drawing processes and found the wear coefficient to depend on accumulated wear work. They found that the wear coefficient of tool steel changes with sliding distance in three stages: decreases in the first stage, remains nearly constant during the second stage and increases quickly in the third stage. Such change of wear coefficient was explained by Ersoy-Nürnberg et al. in terms of accumulated mechanical fatigue leading to growth of fatigue fractures on the tool surface. Unfortunately, their article contains no data about change of surface roughness of the die. Gåård et al. [8,25] empirically studied wear of D2 tool steel and nodular iron against sheets of ferritic-martensitic steel DP600 on a slider-on-flat surface (SOFS)

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tribometer designed to model typical SMF conditions. Their work was devoted to study of tool surface degradation caused by wear processes in different contact conditions and did not include any estimates of wear coefficient. Formation of pits on the surface was observed for the nodular iron and explained by removal of graphite nodules [8], but they did not observe pit formation on the surface of the D2 tool steel. The CRF process typically utilises cold work tool steels, hence research on wear modes in SMF for these steels are potentially applicable to wear analysis in CRF. Adhesive wear of AISI D2 and O1 cold work tool steels of equal hardness was studied in a pin-on-disc test [26] with a stationary spherical pin made of Al2 O3 [27] loaded against specimens attached to a rotating disk. The wear coefficient was found to be different for these two tool steels under the same sliding conditions. This finding differs from earlier data [28] suggesting that wear coefficient is constant and the same for any tool steel sliding against any mild steel. Friction in CRF is assumed to obey Amontons’s and Coulomb’s laws of friction [2] used in most simulations of SMF processes [29]. Empirical models of friction developed for particular SMF processes like deep drawing [30–33] are not based on analysis of physical mechanisms of friction and cannot be extrapolated to friction modelling for all SMF processes. The mechanism of friction and adhesive wear [34,35] and explanations of friction phenomena [29,36,37] proposed earlier for SMF are based on the theory of interlocking asperities suggested initially by Euler [38] to explain friction and developed further by Hirn [39] to explain both friction and wear phenomena at once. Although the theory of interlocking asperities is well developed it cannot explain why the friction coefficient sometimes depends on sliding distance but sometimes does not. The friction coefficient determined by Boher et al. [3] steadily decreased for a cold rolled mild steel strip sliding up to 100 m against the fillet of a die made from D2 tool steel at different strip exit angles with a contact pressure smaller than 120 MPa. Boher et al. did not study the dependence of friction coefficient on roughness or contact pressure and did not explain the decrease of friction coefficient with sliding distance. Hashimoto et al. [9] analysed the influence of sliding distance and surface cracks on friction coefficient in SMF, concluding that friction coefficient depends mainly on sliding distance and proposed a non-linear sliding friction law which takes into account the frictional work quantity and normal pressure. Gåård et al. [8,25] studied the dependence of friction coefficient on contact force and sliding distance for D2 tool steel sliding against mild steel on their SOFS tribometer. Gåård et al. [8] found the friction coefficient dependent on sliding distance at contact load of 300 N within 40 m of sliding, and virtually independent of both distance and contact force within the first 60 mm of sliding under contact load of 50 N and 500 N [25]. Gåård et al. did not analyse the effect of surface roughness on the kinetic friction coefficient. Experiments on a slider-on-strip tribometer [2] showed that friction coefficient can depend on roughness of both sliding surfaces, contact force and sliding distance. Määttä et al. [2] concluded that wear debris particles can either increase the friction coefficient (due to abrasion) or reduce friction coefficient under some conditions. This conclusion correlates with earlier data obtained by Suh [40], who showed that increase of friction coefficient with sliding distance may be caused by debris agglomeration. The process of debris formation, evolution of strip roughness, and transfer of roll roughness to a strip were analysed both empirically and theoretically for the case of cold rolling (CR) [41–43]. In spite of significant differences between CR and CRF, the mechanisms of debris formation and roughness change distinguished in CR [41–43] may occur in CRF. According to a theory of dry friction presented recently by Deulin et al. [44], absorbed gases may in some cases have a signifi-

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cant contribution to friction coefficient. Deulin et al. explained the influence of humidity and atmospheric pressure on friction coefficient for smooth dry surfaces via cumulative coverage of sliding surfaces with sorbed gases on the basis of empirical data obtained for silica and stainless steel. Suzuki et al. [45] studied empirically typical contact conditions occurring in the CRF process and ascertained the following: 1) apparent roll–strip contact area is governed by bending of strip on a roll surface; 2) areas of plastic forming of the strip may not coincide with apparent roll–strip contact areas in general case; 3) transverse distribution of force in a roll–strip contact is highly non-linear. Geometries of typical apparent roll–strip contact areas obtained empirically by Suzuki et al. [45] allow estimating distributions of roll–strip sliding velocities and making a conclusion that the main type of roll–strip interaction in CRF is sliding [46]. Rolling occurs in the CRF process rarer and takes place in contact areas with uniform diameter of the roll. The aims of the study reported in the current paper are to investigate parameters that can be used for predictive modelling of tool surface degradation, obtain a value of tool wear coefficient, detect the phenomena governing the value of friction coefficient for a CRF application, and analyse tool wear mechanisms for CRF of Zincalume coated mild steel. The contributions of this paper are: • experimental study of the effect of surface roughness and sliding distance on friction coefficient for the materials interacting during the CRF process; • empirical study of wear coefficient of AISI D2 tool steel sliding without lubrication against uncoated AISI C1020 hot rolled mild steel; • analysis of tool wear mechanisms for CRF of Zincalume coated mild steel; • study of debris size distribution obtained in laboratory conditions. 2. Procedures – experimental techniques 2.1. Wear test This research is an early exploration of the consequences of performing CRF operation without lubricant. The simulated CRF process is used to manufacture hollow sections from uncoated hot rolled 2.5 mm thick strip of AISI C1020 mild steel for construction industry. The studied CRF velocity was 1.5 m/s. Forming rolls simulated are made from AISI D2 tool steel and have a diameter increasing by 34% along a roll axis. Wear is defined here as progressive loss of material due to relative motion between a block tested and a wheel. Presently, both tool and mild steels were assumed to be obeying Archard’s wear law [47] in the form of Eq. (1). Pc · Vs ∂w∗ =K· , Hr ∂t

(1)

where w* is a specific wear [m3 /m2 ], t is time of sliding [s], K is a dimensionless empirical wear coefficient, Pc is average contact pressure in a block–wheel contact [MPa]; Vs is sliding velocity [m/s]; Hr is surface hardness of the material worn obtained from Eq. (2) [MPa]. The Rockwell hardness value HRC is converted into MPa (Hr ) using a least squares interpolation of hardness conversion tables [48] expressed with Eq. (2).

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Lever

Electric motor

Control console

Gearbox ON

Ra

26

OFF

AISI D2

1.6

AISI C1020

15 Block holder

y 0

56

Block Rotating sense 0

.2 Ra 3

x

z

z Wheel

Weight

x

Debris collector

13

Fig. 1. Global schema of the block-on-wheel test rig.

φ7

Hr = 7.537 × 10−5 · HRC4 − 3.018 × 10−3 · HRC3 + 0.205 · HRC2 +10.487 · HRC + 513.935.

0

(2)

The empirical longitudinal distribution of the roll–strip contact force during the CRF process [49] is most similar to the longitudinal contact force distribution achieved in the block-on wheel test scheme (see Appendix A). This similarity allows choosing the blockon-wheel test scheme [50] for measurement of wear coefficient for D2 steel (the blocks) sliding without abrasive and lubrication against hot rolled C1020 steel (the wheel). It is noted that in this case the material in contact with the tool was not kept fresh (according to recommendations from [21] and remark from [2]) to be consistent with the CRF process itself. The sliding velocity and contact pressure were chosen to be 0.78 m/s and between 43 and 72 MPa in accordance with typical CRF conditions for localised high wear sections of the rolls (see [46,45] and details in Appendix A). The global schema of the block-on-wheel test rig is presented in Fig. 1. An electric motor actuates a gearbox that rotates the wheel which slides on a surface of the block. The block is pressed against the wheel with the force of 42 N or 121 N. The force is set up using a weight applied to a lever with a block holder attached. A control console serves for setting up sliding distances required for wear tests. A changeable debris collector is meant to collect debris at a certain segment of sliding distance. The friction force was measured via power consumed by the electric motor of the test machine during the tests. The systematic measurement error for the friction coefficient was estimated to be less than 10% of its average value for such measurement (see Appendix B). Such error allowed measuring friction coefficient with an accuracy which is not worse than the accuracy of the earlier results for friction of stainless steel on different tool steels in CRF conditions [2]. Three sliding distances were selected for wear tests at each contact load. Six wear and six scar roughness measurements were taken for each combination of contact load and sliding distance. A total of 18 blocks were tested on both faces and 3 blocks were used as controls for wear detection. The result of the tests of 18 blocks was checked via testing additional 4 blocks for 4 additional combinations of load and sliding distance. The same 3 control blocks were used for wear detection as in the tests of 18 blocks. A preliminary 90 km (peripheral accumulated circumference) run-in of the wheel was implemented under the contact load of 121 N before the start of the main series of the block-on-wheel wear tests to ensure more uniform transverse distribution of a

Fig. 2. Blocks and the wheel used.

block–wheel contact pressure. The wheel was not removed from test machine or redressed until the tests of 18 blocks had been finished. Loose debris discharged from the contact zone in the wear tests was collected in two 1.8 km sliding intervals under four conditions, namely early stage (0.5–2.3 km) and established (13.7–15.5 km) wear at each of the two contact loads. The debris was collected in plastic boxes and hermetically sealed immediately upon completion of the collection interval. 23 measurements of contact temperature were taken on the upper edge of the block–wheel contact zone during each wear test at regular sliding intervals. All tests were performed in uncontrolled laboratory atmosphere. 2.2. Materials studied Typical published [24] chemical compositions of AISI steels are given in Table 1. The samples of AISI D2 steel were cut into blocks as illustrated in Fig. 2 then quenched and tempered to a hardness of HRC 60–62 and polished to a roughness value of Ra 1.6 ␮m. The wheel (see Fig. 2) was made of hot rolled mild steel AISI C1020 with the hardness of HRC 5–7.5. Ethanol was used to clean the surfaces of the block and the wheel before the start of each test in compliance with requirements of the test procedure [50]. 2.3. Rolls analysed The laboratory equipment able to simulate all aspects of in-plant tool wear is not available to us at the present time. Therefore observation of tool wear during the existing CRF process was performed to identify types of wear occurred in existing process conditions and to understand the wear mechanisms. When a strip with a soft coating is formed, the tool wear does not appreciably alter the dimensions, but is manifested as surface degradation of the forming rolls. The analysis of tool surface degradation during the CRF of Zincalume coated strip was implemented

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Table 1 Typical chemical compositions of AISI steels, wt%. C C1020 D2 H13

0.17–0.23 1.40–1.60 0.32–0.45

Cr 11.0–13.0 4.75–5.50

Mo 0.70–1.20 1.10–1.75

via visual inspection of main forming rolls and scanning electron microscope (SEM) examination of the surface of side forming rolls used for the process. The main forming rolls were manufactured from AISI D2 steel quenched and tempered to 60–62 HRC according to the specification of Smorgon Steel. The side forming roll studied by SEM was manufactured from AISI H13 steel (see Table 1), quenched and tempered to 52–55 HRC according to Smorgon’s workshop drawing. The worn surface of the industrial side forming roll shown in Fig. 3(a) was examined after forming of 74,500 m of Zincalume coated strip at the temperature of 150–200 ◦ C on the roll surface [51,52]. A profile section was cut from the roll as shown in Fig. 3(b) then the section was cut into segments as shown in Fig. 3(c) to facilitate microscopic examination. The profile section and its segments were cut from the forming roll using the wire cutting process under kerosene flow. 2.4. Characterisation of wear Hardness of the blocks and the wheel used in wear tests and hardness of the specimens cut from the side forming roll were measured with Buehler Macromet® II Rockwell hardness tester. The blocks tested were weighed before and after the test to within 10−4 g to calculate the wear. The wear coefficient was determined with the same procedure as used for measuring abrasion of blocks [50]. The volumetric loss of block material was determined via change of the block masses

V ≤1.10 0.80–1.20

Mn

Ni

Si

0.30–0.60 ≤0.60 0.20–0.50

≤0.30 ≤0.30

0.10–0.25 ≤0.60 0.80–1.20

during the test. In the tests of 18 blocks the volumetric wear of the wheel was determined from change of its outer diameter measured using a calliper providing measurements of 0.01 mm. In four additional tests the volumetric wear of the wheel was determined from change of its mass. The temperature of the block–wheel contact was measured using an Optris MiniSight infrared thermometer with a precision of ±1 ◦ C. Surface roughness of the forming rolls and the blocks tested was measured with a Taylor Hobson Surtronic 3+ roughness meter with an error of 2% of measured value. The forming rolls were inspected with a Celestron 44300 handheld digital optical microscope. The blocks tested, debris generated during the tests in laboratory and specimens prepared from surface fragments of the side roll worn (see Fig. 3) were examined using a JEOL 6460LA SEM, and chemically analysed by energy dispersive X-ray spectroscopy (EDS). Fragments of the wheel surface were inspected and photographed using a Wild M8 stereomicroscope and SEM. The debris particle size distribution was obtained using a Malvern Mastersizer 2000. 3. Wear test results For D2 tool steel sliding against C1020 mild steel, the results presented below include measurements and observations of wear coefficients, friction coefficients, debris generated and worn surfaces. 3.1. Wear coefficient The volumetric wear rate for the blocks of D2 tool steel in the block-on-wheel wear test is shown in Fig. 4. In Fig. 5 these data are converted to wear coefficient using Eq. (3): K=

Fig. 3. Layout of a side roll sectioned for SEM examination.

w · Hr , Fc · Ls

(3)

Fig. 4. Cumulative volumetric wear of D2 block.

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Fig. 5. Wear coefficient of D2.

where w is the volumetric wear measured [mm3 ], Fc is contact force [N] and Ls is total sliding distance [mm]. Referring to Fig. 5, the wear coefficient for the block was considerably higher for 42 N than for 121 N, and at the lower contact force continued to increase with sliding distance whereas at the higher force it quickly reached a plateau value. This increase in wear coefficient with decreasing load might be taken to suggest a transition from mild wear to severe wear. Possible mechanisms for such a wear transition are considered in Section 5. For the mild steel wheel, the wear data measured during the main body of experiments displayed large statistical scatter because they were measured via change of the outer diameter of the wheel, so that the data were not of sufficient quality to warrant graphical presentation. A subsequent experiment measured wheel wear via weight loss, resulting in more reliable values, but without statistical validation. The latter data points combined with data from the main experiment allowed estimates of wheel wear coefficient as presented in Table 2. Conversely to that for the block, the wear coefficient of the wheel is apparently higher for the larger contact load and decreases with sliding distance.

Fig. 6. Friction coefficient diagram for AISI D2 tool steel sliding against. AISI C1020 hot rolled mild steel in the block-on-wheel wear test.

3.2. Friction coefficient The coefficient of sliding friction for D2 tool steel sliding against C1020 mild steel was measured as described in Section 2.1 and the results are displayed in Fig. 6. For the contact load of 42 N the friction coefficient rose from 0.3 at the start of the test, to about Table 2 Estimates of wear coefficient for mild steel wheel sliding against block of D2 tool steel. Force, N/Sliding distance, km

7.7

16

50

93

42 121

1.7 × 10−5 4.2 × 10−5

1.6 × 10−5 3.9 × 10−5

1.1 × 10−5 3.1 × 10−5

0.57 × 10−5

Fig. 7. Contact temperature for blocks of AISI D2 tool steel when sliding against. AISI C1020 hot rolled mild steel wheel in the block-on-wheel wear test.

0.4 after 18 km of sliding and eventually up to 0.6–0.7 (although with short excursions down to 0.51 ) for sliding distances greater than about 40 km. For the contact load of 121 N, however, no such increase was evident, the friction coefficient remaining below 0.4 for sliding distances up to 32 km. Although the friction coefficient

1 These excursions correlated with nightly interruptions to testing, necessary to satisfy laboratory safety regulations.

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Fig. 9. Friction coefficient corresponding to measured roughness obtained during the block-on-wheel wear test.

Fig. 8. Block roughness in x and y directions as a function of sliding distance.

was not measured for longer sliding distances at the higher load, the 121 N results show no evidence of strong dependence on sliding distance as occurs for the lower load. The measurable temperature at the upper edge of the block–wheel contact zone rose to an equilibrium value within the first 9 km of sliding and then varied slightly, as shown in Fig. 7. As would be expected the temperature was somewhat higher for

the larger contact load, but never exceeded +66 ◦ C throughout the experiment. Such temperatures could not and did not cause measureable block softening. For all blocks hardness was measured before and after the wear test, finding that scar hardness was the same as that of unworn block surfaces within the measurement error for all tests, namely 60–62 HRC. The surface roughness values of wear scars on the D2 blocks were measured as described in Section 2.1, with results shown in Fig. 8. The roughness values Rax (in the sliding direction x as per Fig. 2) and Ray (in the transverse direction y) increase from their as-received values to higher values depending on cumulative sliding distance. The transverse roughness Ray appears to rise steadily with sliding distance, from the initial 0.7 ± 0.1 ␮m to 1.8 ± 0.4 ␮m after 8 km and continuing to rise to a maximum recorded value of

Fig. 10. Wear scar surfaces of AISI D2 blocks after sliding under contact force of 42 N (a), (b) and 121 N (c), (d).

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Fig. 11. Trailing (a), (b) and leading (c), (d) edges of a scar after sliding under contact force of 42 N.

3.8 ± 0.8 ␮m after 93 km. The longitudinal roughness Rax seems to rise quickly from the initial 0.35 ± 0.11 ␮m to a plateau at about 1.8 ␮m by the time of the first subsequent measurement. These roughness results show little if any statistically significant dependence on contact load. If the friction coefficient data from Fig. 6 are juxtaposed with the surface roughness data from Fig. 8 then the correlation of friction coefficient with roughness can be plotted. Fig. 9 suggests that friction coefficient rises with increasing surface roughness at the contact load of 42 N, but is independent of roughness at 121 N. 3.3. Wheel surface, scars and debris particles SEM micrographs were taken from a wear scar on the block after sliding of 23 km on the wheel at 42 N contact force, and after 31 km at 121 N. The transverse (y) and sliding (x) directions (see Fig. 2) are marked on all micrographs below (Figs. 10–13). At the contact load of 42 N the tool steel surface appears relatively smooth, as shown in Fig. 10(a), (b). There are shallow linear striations oriented parallel to the sliding direction, resulting in broad undulations in the transverse direction. This appears consistent with the measured roughness data (Fig. 8), in which Ray is typically approximately double the corresponding Rax . The measured Ray values indicate that the undulations of the worn surface are deeper than the sharper grooves in the initial ground surface (as visible outside the wear scar in Figs. 11 and 12). When imaged in backscattered electron (BSE) mode the surface clearly reveals the D2 tool steel’s chromium-rich carbides, the identity of which was easily confirmed by EDS analysis. The clean metal surface and

replacement of the original ground profile with broader striations suggest removal of some metal from the D2 block surface. With reference to the original ground profile the wear mechanism could be described as “asperity truncation”. At 121 N (Fig. 10(c), (d)) the wear scar surface no longer appears clean, and the microstructure of the D2 steel (with its characteristic chromium-rich carbides) is no longer visible. EDS analysis of the rough material indicates that there has been transfer of material from the C1020 mild steel wheel to the D2 tool steel block at this higher load. The 121 N scar surface with the adhering transfer material visually appears rougher than the 42 N scar, but this visual impression is not reflected in the quantitative measurements in Fig. 8. Low-magnification images of the leading and trailing edges of the wear scar (Figs. 11 and 12) indicate that debris generated during the sliding process adhered partly to the wheel surface, permitting it to be transported from the trailing edge of the wear scar (Fig. 11(a), (b) and Fig. 12(a), (b)) around the wheel and deposited at the leading edge of the scar, forming mounds as shown in Fig. 11(c), (d) and Fig. 12(c), (d). The debris mounds display large concentrations of oxygen and carbon. Similar large concentrations of oxygen and carbon were detected in the debris particles picked up by central regions of the scar surface shown in Fig. 10(c), (d). The trailing edge of the scar retained more debris under the contact load of 121 N (see Fig. 11(a), (b)), than at 42 N (see Fig. 12(a), (b)). Visually it was observed that the wheel surface had distinctive red deposits after the test under the contact load of 42 N, but was of dark colour with metallic gloss after the test under the contact load of 121 N.

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Fig. 12. Trailing (a), (b) and leading (c), (d) edges of a scar after sliding under contact force of 121 N.

After completion of testing, two segments of wheel surface were examined by SEM. The wheel showed two distinct types of surface finish: (i) Relatively flat regions with fine striations; (ii) Broad grooves, mostly filled with agglomerated debris. The debris appeared partly metallic and partly non-metallic, apparently mostly oxide. Various points were analysed by EDS as illustrated in Fig. 13. The metallic material (with high average atomic number hence appearing bright in the BSE image), as analysed at point 003 in Fig. 13(a) and at point 003 in Fig. 13(b), contained zero or negligible Cr, confirming the origin of this debris at the mild steel wheel. However, the oxide (or heavily oxidised regions of) debris contained significant (0.2–1.6%) Cr, suggesting some contribution from the block of D2 tool steel. Turning to the loose debris particles collected, at both 42 N and 121 N the debris generated during early-stage sliding (0.5–2.3 km) had a black colour with metallic gloss. The debris collected during established wear (13.7–15.5 km) at 121 N was again black with metallic gloss. In contrast to these three conditions, the debris collected during established wear at 42 N had a predominantly red colouration. This colouration of the loose debris correlated with the observed colour of the wheel surface as described above. These differences in colour seem to suggest that at the lower load there is a transition in tribochemical conditions from early stage to established conditions; a transition which does not occur at the higher load. SEM examination of the debris particles revealed two apparently distinct morphologies, visible in Fig. 14 and Fig. 15. Prominent

were a relatively small number of large, flat metallic flakes, suggestive of some kind of delamination wear process. Their size and appearance could be consistent with an origin in the broad grooves seen on the surface of the wheel, as seen most clearly in the secondary electron image (SEI) at left in Fig. 13(a). Then there was a much larger number of fine debris particles of irregular shape. The large particles were clearly metallic, and the great majority of fine particles also appeared to be substantially metallic, judging from contrast in BSE imaging mode. However, most particles had a surface oxide film, which varied in thickness. It was possible to find occasional regions where the oxide film was absent or thin, and in these regions the EDS analysis was consistent with mild steel, characteristically lacking in Cr – see the example analysis point 001 in Fig. 15(a) and points 001 and 003 in Fig. 15(b). By contrast, where the oxide film was thicker, significant Cr (0.2–1.5%) was usually detected. These differences in Cr levels for the loose debris were consistent with those measured for the debris adhering to the wheel surface. Diagrams of particle size distributions obtained for the contact loads of 42 N and 121 N are shown in Fig. 16 and Fig. 17 respectively. Although visually in the SEI the particle types appeared much the same in the four samples (low and high loads and early and established stage wear), the quantitative particle size distributions show some significant differences. For the lower contact load, Fig. 16 shows a strongly bimodal size distribution. Both early and established stage wear produced a group of large particles in the range of about 200–350 ␮m, a much

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Fig. 13. SEM micrographs and EDS analysis results from the wheel surface after sliding under contact force of (a) 42 N and (b) 121 N.

larger volume of finer particles in the range 0.2–80 ␮m, and relatively few particles in the intermediate range 80–200 ␮m. The discernible change from the early stage to advanced stage wear is that the 80–200 ␮m range is even more strongly depopulated, and the previously large volume of particles in the 20–80 ␮m range is greatly reduced in favour of finer particles of 0.5–20 ␮m. For the higher load, neither of the graphs (see Fig. 17) shows an absence of particles in the range 80–200 ␮m as seen at the lower load; in fact this size range is well represented. The notable change from early stage to established stage wear is a marked shift from predominantly finer (1.5–20 ␮m) particles to predominantly coarser (20–400 ␮m) particles. In broad summary, at 42 N the intermediate size particles seem to be broken up in favour of finer particles as wear proceeds, whereas at 121 N there is a strong increase in the volume fraction of coarser particles. In addition, the standard deviation of volumetric distribution of debris sizes measured for each debris specimen was

generally smaller for the larger contact load and decreased with the distance slid under both contact loads of 42 N and 121 N. 4. Field observations The CRF process taken for field observations is used to manufacture hollow girders from hot rolled 2.5 mm thick strip of Zincalume coated mild steel for construction industry. The velocity of CRF is 0.5 m/s. The sketches of work forming rolls presented in Fig. 18 demonstrate variation of roll diameters along axes. 4.1. Forming roll wear mechanism The adhesive wear of tools sliding against zinc-coated steel sheet has been modelled using tools made of wood and plastic [15]. However, in industry the Zincalume coated strip is usually roll formed by rolls made of uncoated tool steel. The wear of tool steel against

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Fig. 14. Typical images and EDS spectra of debris generated at 42 N during sliding from 0.5 to 2.3 km (a) and from 13.7 to 15.5 km (b).

Zincalume coated steel cannot be modelled correctly in laboratory environments at the present time, because of strong adhesion of zinc or Zincalume soft coatings to the hard surface of the tool steel. The soft coating quickly forms large mounds during any of the standard wear tests. The mass of adhered soft material exceeds the wear of tool steel by orders of magnitude rendering the wear of the tool

steel undetectable. However, degradation of the tool surface can be assessed via field observation. Relative motion of a forming roll over the formed strip includes both sliding and rolling. Rolling prevails in regions with larger contact pressures and constant radius (flat horizontal parts of the rolls shown in Fig. 18). Sliding prevails in regions with changing radius

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Fig. 15. Typical images and EDS spectra of debris generated at contact load of 121 N during from 0.5 to 2.3 km (a) and from 13.7 to 15.5 km (b).

of the roll and smaller contact pressures – examples of such regions are shaded with grey in Fig. 18. As has been shown earlier [46] the transverse component of the roll–strip sliding velocity is significant. The roll–strip sliding leads to the largest power dissipation in contact during the CRF process and is reflected in localisation of tool wear.

Field observation of surfaces of industrial forming rolls used for CRF of Zincalume coated strip revealed surface scratches and Zincalume coating adhered in the regions where sliding is dominant. In these regions the scratches on the roll surface are oriented at an angle to the forming direction. Although photography was difficult, live observations of the surface features were interpreted to indi-

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Fig. 16. Size distribution of debris generated during the wear test under the contact load of 42 N.

cate a combination of asperity truncation and groove formation by microploughing. Apparent transfer of the strip coating material was observed on the surface of some forming rolls, as shown to the right of the scratched region in Fig. 18(a). No surface scratches or material transfer were observed in rolling-dominant regions (where the radius of the forming rolls was constant, Fig. 18(b)). Both ploughing and strip coating pickup lead to increase of roughness of the forming roll surface and eventually to damage of the formed strip, rendering the forming roll unusable. 4.2. Side roll surface degradation The side roll shown in Fig. 3 was removed from service after forming of Zincalume coated strip under working conditions shown in Fig. 19. Degradation of the side roll surface is illustrated by comparison of Fig. 20 (as-machined) with Fig. 21 (worn). Referring to the sampling plan shown in Fig. 3(c), SEM examination was carried out on specimens 5, 6, 10 and 13. Specimens 6 and 10 are from parts of the roll which contact the strip with varying sliding velocities and contact pressure, while specimens 5 and 13 shown in Fig. 20 are from a non-contacting regions, enabling comparison with either the original as-machined surface or the surface affected by any degradation mechanisms unrelated to sliding. The electron micrographs in Fig. 20(a), (b) show the surface of the roll from specimen 13, which was not in contact with the strip – essentially the original machined surface. Specimen 5 shown in Fig. 20(c), (d) is from the bottom part of the side roll surface. Surface degradation in this area includes deposits of debris elongated in the

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Fig. 17. Size distribution of debris generated during the wear test under the contact load of 121 N.

direction of a lubricoolant. The deposits contain mainly aluminium, zinc and oxygen, with admixture of iron, chromium, molybdenum, and vanadium. These features are consistent with corrosion products. In specimens 6 and 10, which were in contact with the formed strip, significant coating pickup could be seen as illustrated in Fig. 21. The globular particles shown in Fig. 21(a), (b) were analysed by EDS and found to consist mainly of iron and aluminium, with some containing copper, chromium, zinc, molybdenum, nickel and vanadium. The most obvious origin of the globular particles in Fig. 21(a), (b) is metal spray from the welding point, visible in Fig. 19. Interaction with the soft strip coating changes the original side roll surface. In zones where the roll surface slips over the coated strip, as shown in Fig. 21(c), (d), the pre-existing asperities become truncated while new grooves are formed. The flattened metallic particles shown in Fig. 21(c), (d) consist mainly of iron with copper, chromium, zinc and aluminium. Inside the grooves there are increased concentrations of aluminium, zinc and oxygen. 4.3. Side roll hardness change The side roll hardness measured on the section shown in Fig. 3(b) ranged from 42.5 to 54.5 HRC. Most of this range is lower than the specified range of 52–55 HRC in Smorgon’s workshop drawing. The higher values of hardness were measured near the outer surface of the side roll, which was cooled during the CRF process. The lower hardness values were measured near the inner surface of the side roll, with the lowest hardness value (42.5 HRC) being measured near the housing of the upper bearing. Although the side roll was

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5. Discussion 5.1. Wear mechanisms observed in laboratory tests

Fig. 18. Wear of a forming roll during CRF of Zincalume coated strip.

cooled intensively with lubricoolant containing about 96% water and 4% mineral based additive [51], the most obvious explanation for the observed roll softening was the elevated working temperature due to heating by high-frequency induction from the welding point in this case.

Fig. 19. Side roll working conditions. Sparks are from strip welding.

Examination of wear scars and debris particles from the blockon-wheel wear test revealed two wear mechanisms [53] in the block of D2 tool steel, namely asperity truncation and material transfer. At the lower contact load of 42 N, the contact zone on the block of D2 tool steel developed a clean and apparently relatively smooth surface as shown in Fig. 10(a), (b). The apparent smoothing evidently results from a process of asperity truncation, as illustrated at the well-defined boundary between the original machined surface and the wear scar surface in Fig. 11(a). The visual appearance of smoothness is illusory, unsupported by the quantitative surface roughness measurements, indicating that the broader undulations have greater depth than do the sharp grooves in the as-ground surface. By contrast, at the higher load of 121 N the scar surface becomes coated with mild steel debris, evidently transferred from the wheel, as shown in Fig. 10(c), (d). In adhesive wear of dissimilar metal couples it is commonly observed that material is transferred primarily from the softer to the harder surface. In particular, similar material transfer has been observed while studying tool wear during deep drawing [3,17,35]. The adhesive mechanism of wear of the wheel was confirmed via debris examination. The larger flakes in Fig. 14 and Fig. 15 were clearly consistent with a severe adhesive wear mechanism [20,54]. The BSE images indicate that the fine debris particles have as high average atomic number as the large flakes do. From this it can be concluded that if the large flakes are metallic and due to severe adhesive wear, then the smaller particles are also metallic and presumably also due to severe adhesive wear. The very flat flake morphology of the larger debris particles clearly shows evidence of a delamination phenomenon. On the wheel surface, the broad grooves (Fig. 13(b)) suggest galling. The copious material transfer can be expected to effectively protect the D2 block from wear. Since the transfer only occurred at the higher load, this observation can explain the much lower wear coefficients measured for the block at the higher load as shown in Fig. 5. Due to this transfer the wear coefficient of the mild steel wheel was higher at the higher load as shown in Table 2. The greater material transfer at the higher load might be explained in terms of local heating caused by larger frictional power dissipated under larger contact loads for a given sliding velocity. One effect of this heating could be to promote outgassing of adsorbed moisture, facilitating a larger amount of true metal-to-metal contact and adhesion. Table 2 shows that the wear coefficient of the wheel decreases with increasing sliding distance, while Fig. 5 shows conversely that the wear coefficient of the block increases with increasing sliding distance. As wear proceeds, the scar length and hence the contact area increases, decreasing the contact pressure for a given load. The Archard wear coefficient obtained for the block–wheel contact according to Eq. (3) depends on contact force rather than on contact pressure. For a given force the progressively increasing contact area and decreasing pressure neither increase nor decrease the wear rate, because friction and adhesive wear are proportional to the true contact area (asperity contacts), which is much less than the apparent contact area. The true contact area is determined by the load and by the yield strength of the metal – hence is expected to be independent of the apparent contact area. The above describes the first-order effects in a loaded sliding like-metal couple. In a dissimilar metal couple there could conceivably be second-order effects, such that the decrease of average contact pressure usually shifts the wear from the hard block to the wheel of mild steel, while size reduction of the debris obviously promotes pickup of hard fine particles by the surface of mild steel

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Fig. 20. Original machined surface and debris deposits.

(apparently points 2 and 4 in Fig. 13(a)) which impedes the wear of the wheel and causes wear coefficient of the hard block approaching the values observed in tests of wheel and rail steels of similar hardness (Fig. 23). In addition to the effect of increased length of contact zone, other factors which might promote an increase of wear coefficient of a block at longer sliding distances include interaction of the metal with gases and water absorbed on or reacting with the sliding surfaces [55]. The lower the contact surface temperature, the more gases are absorbed by the surfaces. The smaller the size of debris generated, the larger is their gas absorbing surface and hence the potential effect on wear if these particles are not removed from contact. The red colour of the debris and the red deposits on the wheel after long tests under the contact load of 42 N presumably represent ferric oxides (Fe2 O3 ·xH2 O). The dark colour with metallic gloss at the contact load of 121 N could in principle be either ferrousferric oxide (Fe3 O4 ) formed at elevated temperature or metallic debris indicating severe adhesive wear; but the BSE imaging and EDS analyses of the debris particles showed that the debris is primarily metallic at both low and high load. The red oxide formation most probably involves atmospheric moisture. EDS analysis of oxide films on mild steel debris particles (both in collected loose debris and adhering to the wheel) consistently show significant Cr content, even though the underlying metal particles contain negligible Cr. This seems to indicate a tribochemical reaction involving diffusion of Cr sourced from the D2 block. Although

the effect of water and gases absorbed from the atmosphere was not specifically studied in this research, some earlier works suggest that sorbed water can accelerate the wear process [56–58]. 5.2. Values of friction coefficient, roughness, and debris Friction coefficient is determined by several components of solid friction, which can be loosely grouped under the categories of “adhesion” and “mechanical deformation”. The adhesion component depends on inter-atomic bonding between the contacting bodies, bearing in mind the usual presence of contaminants (such as oxides and sorbed gases) which greatly reduce the amount of direct metal-to-metal contact between the nominally contacting surfaces. The mechanical component depends on interaction and deformation of asperities, oxide films and third bodies (such as agglomerated wear debris). Usually, one of these two friction components is dominating. One of the factors influencing the mechanical component of friction is surface roughness. Along with the contact pressure surface roughness governs a difference between apparent and true contact areas. When the contact pressure exceeds a critical value related with the yield point of the weakest of contacting materials, then apparent and true contact areas coincide and the Tresca model becomes more suitable for characterizing the friction [29]. However the block–wheel contact pressure did not reach the yield point of mild steel in the tests as in most part of contact areas

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Fig. 21. Strip coating pickup, metal spray, scratch formation and asperity truncation.

where sliding dominates in the CRF process. Hence, Amontons’s and Coulomb’s laws of friction are considered as more appropriate for characterizing the friction in the tests and in application to CRF [2,29]. The laboratory experiments show that surface roughness of a block of tool steel sliding against a mild steel wheel is not a monotonic function of sliding distance in the general case. The friction force is not proportional to surface roughness of D2 steel for Ra values of the order of 1.6 ␮m as usually specified for the working surfaces of forming rolls in CRF (see Fig. 9). As described in Section 1, Suh [40,59] showed that debris particles which remain in the contact zone may agglomerate and contribute significantly to friction coefficient. In Suh’s experiments the friction coefficient increased with distance when the debris particles were trapped in the contact zone, but remained constant when the debris particles were able to discharge into grooves. Our observation (Fig. 6) that friction coefficient increases with sliding distance (apparently true only for the contact load of 42 N) might be explained by such debris agglomeration. The independence of friction coefficient from sliding distance at the contact load of 121 N might be explained in terms of the greater volume of larger particles in the distribution observed at this contact load (Fig. 17). Rather than agglomeration of finer particles, the observed coarsening for 121 N may be due to an increase in primary generation of coarser particles due to galling and delamination on the wheel surface as observed in the microscopy of wear scars. The observed larger standard deviation of both friction coefficient and volumetric distribution of debris size for the smaller load (Figs. 6 and 16)

might be explained in terms of instability of the friction process, perhaps caused by debris agglomeration. According to empirical data [20,53] and a recent theory of dry friction [44], the amount of gases including water vapour sorbed by sliding surfaces can have a strong effect on friction coefficient. From this consideration, thermal desorption would be expected lead to increase of friction coefficient as shown in Fig. 6. The results presented in Fig. 7 do not reflect such an expectation, but local temperature transients are notoriously difficult to measure. At present it is uncertain whether desorption is instrumental in the apparent transition of friction coefficient with sliding distance at 42 N. From the balance of evidence it might be postulated that the mechanical component of friction coefficient was dominating in these experiments, such that interaction of contact surfaces with debris particles had the strongest effect on friction coefficient in the block-on-wheel wear test. 5.3. Wear of industrial forming rolls Field observations show that when a strip with a soft coating is roll formed, the wear of forming rolls (made from AISI D2 or AISI H13 tool steels) is influenced by material transfer, asperity truncation and surface ploughing, as shown in Figs. 18, 20 and 21. The groove formation, and perhaps also the asperity truncation, can be postulated to have been caused by hard debris adhering to the soft strip coating at previous roll stands. The position of surface scratches (Fig. 18(a), (b)) correlates with the smaller linear load in contact while the position of soft coating adhered to the hard tool-

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(b) suggests that they are metal spray droplets from the welding point shown in Fig. 19, adhering to the soft coating of the strip subsequently flattened between the roll and the strip. Such trapped particles might contribute to the creation of grooves and scratches such as seen in Fig. 21(c), (d). 6. Conclusions The most important results of this work are the following. 1. A block-on-wheel wear test offers low-cost determination of tool wear coefficient in geometries relevant to CRF. 2. The wear coefficient has been determined for AISI D2 tool steel sliding against AISI C1020 hot rolled mild steel without lubrication or surface coating. 3. An apparent transition of wear mechanisms in the blocks of D2 tool steel, mediated by contact load and sliding distance, has been explained in terms of material transfer from the wheel. 4. It was demonstrated that for AISI D2 tool steel sliding against AISI C1020 in the block-on-wheel wear test, the friction coefficient is independent of roughness and of sliding distance, but is influenced by the interaction of contact surfaces with debris. 5. A variety of possible mechanisms have been identified for degradation of tool surfaces during the CRF process of Zincalume coated mild steel. Acknowledgements Fig. 22. Longitudinal normalized force distribution in a roll–strip and a cylinder–plane contact from empirical data [49,60,61].

ing (Fig. 18(a)) conforms with the larger linear load in contact (the strip is bent by the right conical part of the rolls presented in Fig. 18) as compared with empirical transverse distribution of contact force obtained for similar roll profiles [45]. This observation agrees with the results of the block-on-wheel wear tests discussed in Section 5.1. The removal of hard asperities by the soft strip coating is likely to have been initiated at previous roll stands by a form of fatigue wear [6,21] under periodic hydrostatic load. Under high contact pressure the soft coating might behave as a hydrostatic fluid and act as a hydrowedge to promote opening of incipient cracks. This could potentially promote propagation of fatigue cracks permitting separation of debris particles from the forming roll. The hard asperities thus removed from the roll surface and embedded within the soft strip coating could then act as abrasive particles. Material transfer and surface ploughing dominated on the surface of the side roll shown in Fig. 3. The material transfer took the form of Zincalume debris sediments on parts of the roll surface that were not in contact with the strip, as a result of lubricoolant flow, as shown in Fig. 20(c), (d). Direct transfer of soft Zincalume coating from the strip to the side roll surface took place in the roll–strip contact area (see Fig. 21(a), (b)). The chemical composition and morphology of the spheroidal particles shown in Fig. 21(a), p, Pa

K·10-4 300 - 400 severe

0.8 · Hσ 1 - 10 mild

30 - 40 borderline 0.2

1 - 10 mild 0.7

vslip, m/s

Fig. 23. Empirical wear coefficient for wheel and rail steels [63].

The authors want to thank Australian Research Council and Smorgon Steel Tube Mills for financial support of this research. The authors express their gratitude to Mr. Ross Bartlett for his kind support of tool wear observation at Smorgon’s CRF mill. The authors appreciate Dr. Shichao Ding and Mr. Tien Vuong for their valuable help in search for publications referenced in this paper, Mr. Leonard J. McInnes for his assistance in the start of wear tests, Mrs. Glenda Zemanek for her friendly assistance in hardness tests and optical microscopy, Mr. David Page for his kind help with measurement of debris size distribution and the staff of the Centre for Microscopy and Microanalysis of the University of Queensland for their amicable support of this research. Appendix A. Wear test selection A wear test method is required to simulate the real interaction of contacting bodies as accurately as possible. Therefore the contact force distribution in the test must be similar to that observed in the operation being simulated. The transverse roll–strip contact force distribution varies greatly depending on the section formed and the profile of the forming roll [45]. Therefore it is reasonable to consider short segments of contact with force distributions uniform in transverse direction for approximation of transverse distribution of force in any roll–strip contact. The normalized longitudinal distribution of roll–strip contact force would not depend on the section formed unless it possesses bicurvature. Such a longitudinal contact force distribution obtained from empirical data [45] is most similar to the normalized longitudinal contact force distribution obtained for a flat surface contacting with a rolling cylinder from experimental data [60,61] (see Fig. 22). The normalized forces F1 and F2 in Fig. 22 were obtained for the roll–strip contact [49]. These forces are the same within the accuracy of the experiment. The normalized force FT in Fig. 22 was obtained for a flat surface contacting with a rolling cylinder [60,61]. Centroids of F1 and F2 coincide within the accuracy of the experiment and are close to centroid FT as illustrated in Fig. 22. Similarity of interactions in a roll–strip contact and in a plane–rolling cylinder contact shown in Fig. 22 makes

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the block-on-wheel wear test ideally suited for simulation of tool wear conditions in the CRF process. The force in a real roll–strip contact varies over a wide range of values. Published empirical data obtained on roll forming test rigs give values of contact force up to 1980 N/mm across a 3 mm thick strip formed [45]. However, Archard’s wear law used widely for wear prediction is linear relative to contact pressure. Hence only two contact force values of 121 N and 42 N were chosen to measure the empirical tool wear coefficient in this research. As shown earlier [46] the roll–strip slip velocity has a significant transverse component during the CRF process and thus depends on the section formed, and can vary over a wide range even within one contact spot. The empirical diagram in Fig. 23 shows that the wear coefficient for friction pairs of wheel and rail steels changes depending on contact pressure and slip velocity. If the data given in Fig. 23 are extrapolated to tool steel sliding against mild steel then the values of wear coefficient for nonzero slip velocities less than 0.2 m/s, and for those greater than 0.7 m/s but less than 0.9 m/s, would be symmetrical about 0.45 m/s. Such extrapolation allowed the selection of 0.78 m/s sliding velocity for the wear test presented in this paper. The sliding velocity was not changed in the test, as the linear Archard wear law was adopted for tool wear modelling. Appendix B. Wear test accuracy An appropriate sample size was calculated in accordance with ASTM E122 [62]. The ratio of standard deviation to maximum allowable deviation was taken to be 0.8. The probability of deviation beyond maximum allowable was taken to be 0.003. The calculated sample size for each point of measurement was rounded to 6. The accuracy of friction coefficient values measured was determined by the performance index of the torque transmission from the electric motor to the test wheel and by the electric motor efficiency. The performance index of the torque transmission used varies slightly with the value of radial load applied to the wheel. The total performance index ktp of the electric motor and torque transmission of the testing machine was calculated from measured idle power of the test machine Pi by Eq. (B.1): ktp = 1 −

Pi , Pn

(B.1)

where Pn = 1.5 × 103 W is the nominal power consumption of the electric motor used in the test machine. The friction coefficient in the block–wheel contact is expressed by Eq. (B.2): kf =

ktp · (Pt.m. − Pi ) ,  · dw · nw · Fl

(B.2)

where Pt.m. is the power consumption of the test machine during the wear test (W); dw is the outer diameter of the wheel (m); nw is the number of wheel revolutions per second; Fl is the load applied to the block tested (N). The power consumption of the testing machine rose from 5% to 15% when the loaded block was applied to the wheel, in dependence on the voltage of power supply and the applied load. This increase in power consumption was divided between electric motor, transmission and the block–wheel contact. A typical performance index of an electric motor of the same power changes in the range 0.5–1.5% after such increase of load and it was giving a systematic measurement error εe.m. = 1.5%. The remaining part of power consumption change fell to the proportional share of the transmission and the block–wheel contact according to Eq. (B.2). The systematic measurement error introduced with change of the performance index of the transmission under load was neglected. The electric motor power supply was not stabilized and varied in the range of 235–245 V within a day. This added 5% system-

atic friction measurement error to the value of friction coefficient εp.s. = 4%. Total systematic friction measurement error in the test estimated by Eq. (B.3) was about 6%. εst = εe.m. + εp.s.

(B.3)

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