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Sliding wear and fatigue cracking damage mechanisms in reciprocal and unidirectional sliding of high-strength steels in dry contact
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Sliding wear and fatigue cracking damage mechanisms in reciprocal and unidirectional sliding of high-strength steels in dry contact €m Abdulbaset Mussa *, Pavel Krakhmalev, Jens Bergstro Karlstad University, Department of Engineering and Physics, SE-658 88, Karlstad, Sweden
A R T I C L E I N F O
A B S T R A C T
Keywords: Rock drill rods Reciprocal sliding contact High strength steel Fatigue cracks Sliding wear and plastic deformation
Rock drill components operate under tough contact conditions during rock drilling. Reciprocal and unidirec tional motion under high contact stresses are the common contact conditions between interconnected compo nents. It will result in component damage and often the observed surface damage of rock drill tools is due to wear and fatigue cracks. Nevertheless, the effects of the properties and structure of the mating materials on tribo logical performance, is not fully understood. The present study is dedicated to simulation and investigation of the wear mechanisms observed in reciprocal and unidirectional sliding of high strength steels for rock drill com ponents. A high strength martensitic steel, 22NiCrMo12–F, commonly used in rock drills was tested in selfmating contact. Wear mechanisms were investigated by means of electron microscopy and wear damage was quantified by a 3D optical interferometer. Total damage, as a result of adhesive wear, severe plastic deformation and nucleation and propagation of fatigue cracks, was discussed in relation to test conditions and material properties. It was observed that the coefficient of friction decreased with increasing normal load. Moreover, the results showed that the type of motion had a significant influence on the worn volume and crack nucleation of the specimens in sliding contact. In addition, the reciprocal motion resulted in higher wear than unidirectional motion under the same test conditions.
1. Introduction In rotary and percussive rock drilling high energy waves are used to break a rock material into smaller pieces by impacting the rock. Rock drills are usually driven by a hydraulic system that generates energy in terms of shock waves and transmits it to a drill bit through a rotating drill rod components. In order to drill a hole with a diameter of 35–89 mm, a hydraulic pressure of 20 MPa is used. It provides a fre quency up to 50 Hz, rotating rate up to 200 rpm and a feeding thrust, in order to maintain contact between rock and drill bit, about 10 MPa [1]. These application parameters expose rock drill components to extreme contact situations due to high pressure sliding contact and repeated cyclic load. One example of such components is thread joints, Fig. 1, used to connect extended drill rods or to connect a drill rod to a drill bit. During rock drilling, high pressure sliding contact between threaded surfaces takes place. It leads to successive surface damages that worn out surface material. Wear of the thread joints is of an important techno logical and economic significance. It deteriorates the surface material and reduces the tightness between connecting components that results
in an inefficient drilling process with increased drilling time and costs. Commonly, materials with high hardness and high wear resistance are used to minimize wear rate between contacting components. How ever, the high hardness of a component should not compromise its fa tigue life. Therefore, it is important to select materials with a proper balance between the hardness and the toughness. In rock drill components, it is of high importance to use materials with high surface hardness in order to achieve an adequate wear resis tance and high core toughness to have longer fatigue life of a compo nent. To this matter, martensitic high strength steels with low carbon content are used in rock drills [2,3]. Surface hardening of the involved components is frequently used to achieve high wear resistance [4]. Carburizing is a quite common surface treatment to generate a hard surface layer without altering mechanical properties of the core material. In rock drill rods, connectors and other components, the relative motion depends on a location and a component. Often a unidirectional motion is dominating, but in some components, reciprocal motion is observed. A case study previously performed by the authors [5], on the
* Corresponding author. E-mail addresses: [email protected] (A. Mussa), [email protected] (P. Krakhmalev), [email protected] (J. Bergstr€ om). https://doi.org/10.1016/j.wear.2019.203119 Received 25 January 2019; Received in revised form 11 September 2019; Accepted 8 November 2019 Available online 9 November 2019 0043-1648/© 2019 Elsevier B.V. All rights reserved.
Please cite this article as: Abdulbaset Mussa, Wear, https://doi.org/10.1016/j.wear.2019.203119
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2.2. Specimen preparation Two different specimen geometries were used in the present study, discs and plates. The disc specimens had a major diameter of 50 mm and minor diameter of 10 mm. The thickness of the discs was 10 mm. The plate specimens had a length of 600 mm, a width of 100 mm and 10 mm in thickness. All tested specimens were carburized in a well-controlled and carbon rich atmosphere (N2 þ CH3OH) followed by water quench ing and tempering. The carbon content in the case hardened layer was up to 0.6 wt %, the tempered martensitic microstructure of the case hardened layer is presented in Fig. 2 a. The carburized case depth was 1.2 mm, the surface hardness was up to 720 HV and the core hardness was around 450 HV, Fig. 2b. As a result of the case hardening by carburizing, an oxide layer was formed on the surface of the heat treated specimens. In order to inves tigate the influence of the oxide layer on the material performance in sliding contact the specimens were tested under two different surface conditions, as-delivered and ground surface conditions. On the ground surface, the most intensive oxide was removed from the surface by manual grinding using SiC paper, started with rough to smoother using 320, 500 and 2400 grit papers. Prior to wear testing, the surface roughness of the tested specimens were measured and evaluated by the means of a 3D optical profiler, Bruker 3D.
Fig. 1. Schematic of the thread joint between the drill rod and drill bit with rotational drilling and axial percussion movements during the rock dril ling process.
failure mechanisms of the drill rod and thread joints failed under in dustrial conditions has revealed that the surface material in the threaded parts is often plastically deformed when it is used in the application. Severe adhesive wear due to a high-pressure sliding contact between connected threads took place. It has also been revealed that grain refinement of the surface material occurred in locations where plastic deformation took place. Additionally, fatigue cracks as a result of sliding contact between threaded parts have been observed [5]. The complexity of the contact situation in rock drilling, the type of relative motion, high mechanical stresses, temperature and corrosion, make it difficult to predict the performance and wear resistance of a particular material in sliding contact. Therefore, experimental tests at a laboratory level under well-controlled test parameters are of high importance to enable pre diction of the wear resistance of a tested material in a particular appli cation. Experimental studies have declared several key parameters that have an essential effect on the surface performance in sliding contacts, e. g. mechanical properties of the surface material, microstructure, surface roughness, contact pressure and sliding distance [6]. Additionally, the contact conditions change over time and, therefore, the wear mecha nisms change continuously. The aim of the present study is to under controlled laboratory tests simulate the wear mechanisms of the rock drill steels in sliding contact. Commonly, the tested steel is used in the thread joints to connect drill rods with drill bits. Furthermore, the aim was also to investigate the influence of surface condition and type of motion on friction and wear resistance in self-mating contact of a case hardened martensitic steel. In order to achieve the aim of the present study, wear tests under dry reciprocal and unidirectional sliding contact were conducted. A range of loads and sliding distances were used in order to study the correlation between contact pressures and different wear regimes on the tested specimens.
2.3. Wear testing and characterization Wear tests were performed using the Slider On Flat Surface (SOFS) tribometer [7]. The SOFS has been proven to simulate adhesive wear processes, and provide sufficient characterization of adhesive wear resistance of steels in sliding contact at high pressures [7–9]. Since the main wear mechanisms revealed by the case study [5] was severe ad hesive wear, the SOFS was decided to be suitable for investigation and evaluation of wear resistance of the steel in the present research. In the SOFS, a disc-shaped specimen is pressed against a flat plate, Fig. 3a, under controlled normal load and slid until the desired sliding distance (S) is achieved. The disc is fixed in a way that rotation movement during sliding is avoided. The tests were conducted under dry reciprocal and unidirectional repeating sliding. In the present investigation, the reciprocal sliding tests involved pressing the disc with a predetermined normal load (FN) against the plate and sliding with a constant speed of 0.3 m/s and a stroke length of 150 mm. The reciprocal movement of the disc was repeated in the same track until the predetermined total sliding distance was achieved. In the unidirectional sliding tests, the disc was pressed and slid against the plate under a predetermined normal load FN with a constant speed of 0.3 m/s. When the disc reached the end of the stroke, 150 mm, it was lifted up from the plate and moved back to the starting position. This operation was repeated in the same track until the predefined total sliding distance was reached. Wear tests were conducted at three different normal loads, of 100, 300 and 500 N, and the sliding distances of 100, 200 and 300 m. In the real application, the contact pressures between the threaded parts in some extreme locations exceed the yield strength of the surface material. This was revealed when a detailed examination of wear characteristics of the failed thread joints used in the field application has been per formed [5]. Plastic deformation of the surface material and severe sliding wear were the dominant wear mechanisms. The selection of the normal load range between 100 – 500 N ensures elastic initial contact at 100 N and elastic-plastic initial contact at 500 N between the discs and
2. Experimental details 2.1. Material All specimens used in the present study, the discs and plates, were made of an advanced high strength steel with low carbon content and high hardenability, 22NiCrMo12–F. The chemical composition and the typical mechanical properties of the steel grade are presented in Table 1.
Table 1 Chemical composition, wt. %, and mechanical properties of the core material. C
Si
Mn
P
S
Cr
Ni
Mo
Al
Rp0.2a [MPa]
Rm [MPa]
0.20
0.30
0.70
0.02
0.02
1.30
2.95
0.25
0.03
1050
1500
a
Rp0.2. and Rm are the proof strength and the ultimate strength of the tested material, respectively. 2
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Fig. 2. a) Case hardened tempered martensitic microstructure at 0.5 mm depth and b) hardness profile from the surface into the bulk material.
Fig. 3. a) The SOFS tribometer configuration and b) schematic of a wear test. Specimens, sliding direction, normal load FN and friction force Ff are marked with black arrows.
plates. The sliding distances selected for the present work were sufficient to produce a wear damage that can be characterized. Moreover, these sliding distances were selected based on the practical experimental reasons and also based on the pervious experimental experiences. Dur ing the wear tests the instantaneous coefficient of friction (CoF) between the movable disc and fixed plate was recorded and monitored. In order to visualize and quantify wear of the worn surfaces, the 3D profiler was used. Samples from locations near the starting, middle and ending point of the wear track, of each test, were cut from the plates with dimensions 6 � 6 mm2 and used for wear analysis. The worn depth and the worn volume, V, from the wear track of each test were measured and analyzed. Five measurements from five different locations for each tested specimen were conducted. The average of the measured values combined with its standard deviations were plotted against the wear load, according to the Archard’s wear equation [6]. V¼
K FN S H
mounted in thermoplastic, ground and polished according to the stan dard metallographic sample preparation method. The prepared crosssection samples were etched in 5% Nital and examined using RiechertJung Polyvar metallography microscope. The hardness of the worn surfaces were measured and analyzed by a microhardness tester, FutureTech FM, at 25 gf. 2.4. Contact pressure calculations When running laboratory wear tests, it is important that the tests simulate key parameters involved in a real application such as contact pressure, contact mode, etc. The contact pressures obtained between the contacting bodies in the present study were estimated using analytical and numerical methods. The initial static contact pressure between the disc and plate was calculated using the Hertzian elastic contact theory and the Finite Element Method (FEM). A simplified 3D FEM-model of the contact was created in ABAQUS. Both disc and plate were modeled as elastic-plastic von Mises materials with strain hardening according to the stress-strain behavior of the core material, Fig. 4a. Note that the carburized layer has other mechanical properties than the core material. Due to the symmetry only a one-quarter part of the contact area was used in the calculations. The FEM-model was created based on the fol lowings. Both disc and plate, face 1 and 2 in Fig. 4b, had their translation locked in the X- and Y-directions and the bottom face of the plate was constrained in all X-, Y- and Z-directions. A CoF equal to 0.2 was assumed between contacting surfaces. The normal load was adjusted to a uniform distributed pressure and applied to the top surface of the disc segment, red arrows in Fig. 4b. A C3D10 Element type with a 10-node
(1)
K is the coefficient of wear and is dimensionless and always less than unity, H is the hardness of the softer material, note that the contacting specimens have the same hardness. The ratio K/H is assumed to be constant because of the self-mating contact. The product of the normal load and the sliding distance (FN � S) is designated as the wear load in the present study. The operating wear mechanisms of the worn specimens were investigated. Microscopic analysis of the wear track on the plate and the worn contact of the disc were performed with Leo 1530 FEG scanning electron microscope (SEM) equipped with Oxford EDX INCA-sight sys tem. For the metallographic study samples were cross-sectioned, 3
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Fig. 4. a) the tensile stress-strain diagram of the tested material, b) 3D FE-model represents one-quarter of the contact between the disc and the plate, c) the refined mesh in the contact area.
quadratic tetrahedron with controlled hourglass shape was used. The mesh was refined in the contact area, to approximately 0.2 mm, in order to obtain higher precision of the contact pressure, Fig. 4c.
3.2. Surface conditions Prior to testing the surface roughness for the selected specimens were measured and evaluated by the 3D optical profiler. It was found that the discs in as-delivered condition had a surface roughness Ra around 0.65 μm while the ground discs had the roughness of 0.51 μm. For plates the difference between the roughness values was larger, Fig. 6a and b. In the case of as-delivered surface the roughness values Ra were around 7.7 μm whereas for the ground surface the roughness values Ra were around 0.8 μm. SEM analysis revealed that the initial oxide layer was porous at the very top and it became denser towards the interface between the oxide layer and the steel surface, Fig. 6c. The layer thickness varied along the surface with an average thickness of 15 μm, Fig. 6c. For the ground surface, the oxide layer depth was reduced to a thickness of approxi mately 5 μm, Fig. 6d. The hardness of the oxide layer on the as-delivered was 290 HV. In both surface conditions, as-delivered and ground sur face, grain boundary oxides were observed, marked by black arrows in Fig. 6c and d. An EDS analysis at higher magnification was performed on the grain boundaries and the surrounding material. It revealed enhanced content of silicon and oxygen at the grain boundaries, Fig. 7.
3. Results 3.1. Contact pressure The maximum static contact pressure for normal loads between 20 700 N were calculated by two different methods, a numerical method by FEM calculations and an analytical method by Hertzian elastic contact theory for elliptical contacts. The FEM calculations showed that width of the initial contact between the disc and the plate was around 0.8 mm when a normal load of 500 N was used. The calculated width of the contact was in a good agreement with the measured width during experimental tests that was close to 0.72 mm. Moreover, the FEM results revealed that the maximum contact pressure was concentrated to the center of the ellipse, contact area between the disc and plate, and the contact pressure decreased as moving towards the circumferences, Fig. 5a. The results, calculated by FEM and Hertzian analyses, are plotted against the corresponding normal loads in Fig. 5b. At low normal loads, only a small deviation was observed between the two methods. The deviation became larger with increased normal load up to 500 N and higher, Fig. 5b. Equivalent plastic strain (PEEQ) values were also plotted against the normal loads. When PEEQ value for a certain element ex ceeds zero, plastic strain occurred. As observed in the obtained results plastic deformation occurred immediately when the disc was pressed against the plate with a normal load of 400 N.
3.3. Friction The instantaneous CoF between the discs and plates was recorded during the tests. The CoF values of all tests showed three distinct regions, the running-in and steady-state regions. Between these two regions a third region was presented where the CoF increased. The running-in region was relatively short where the CoF started with lower values
Fig. 5. Disc and plate contact pressure analysis, a) FEM contact pressure distribution for a selected test at 500 N as normal load and b) Equivalent plastic strain and maximum contact pressure as the function of the normal load, comparison between Hezrtian contact calculation and FEM calculation. 4
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Fig. 6. SEM images of the plates, a-b) shows the topography of the oxide layer a) as-delivered and b) ground surface, and c-d) shows the thickness of the oxide layer, c) as-delivered and d) ground surface. Arrows indicating grain boundary oxides.
Fig. 7. EDS analysis of the grain boundary oxide, a) SEM image of the grain boundaries and b) EDS spectrum of the area of interest.
and increased, in the first meters of the sliding contact, to a higher value corresponding to the steady-state region, Fig. 8. The development of the CoF over the first millimeters of sliding for each test was studied in order to investigate the influence of surface roughness on friction between contacting surfaces. For all the tests conducted it was observed that the starting friction values were around 0.3–0.4. Then it increased to slightly higher values of 0.6–0.7. After sliding for 30–50 mm and longer, the CoF decreased to 0.2–0.3. This behavior was the same for as-delivered and ground surfaces. After that the CoF values were about the same during running-in region. At low normal loads, 100 N, and particularly under unidirectional sliding mode, the running-in region was longer than in the tests performed at higher normal loads. Moreover, the increase in the friction values to reach the steady state, for the tests performed at low loads, required
longer sliding distance. The steady state of CoF at low load showed a stable mean value until the test was stopped. At higher normal loads, the CoF increased immediately after short distance to higher values where the steady state region of CoF started. Generally, the mean CoF value in the running-in region for all tests was between 0.2 and 0.3 and it increased with increasing sliding distance. Thereafter the CoF during the steady-state did not change remarkably and remained at the same level until the test was finished. The mean value of the CoF at the steady-state obtained from the wear tests was plotted against the normal load and sliding distance, Fig. 9. It was observed that the CoF was strongly dependent on the normal load. In the case of the as-delivered surface tested under the reciprocal sliding, the CoF value decreased from 0.72 to 0.45 with increased normal load from 100 to 500 N, Fig. 9a. Similar trend was observed for other tests, 5
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Fig. 8. CoF between the disc and plate tested at a normal load of 100 N and a sliding distance of 100 m during unidirectional sliding contact.
Fig. 9. Influence of normal load and sliding distance on the CoF, a) as-delivered surface, reciprocal sliding, b) as-delivered surface, unidirectional sliding, c) ground surface, reciprocal sliding and d) ground surface, unidirectional sliding.
Fig. 9b, c and 9d. The type of motion, unidirectional or reciprocal sliding, did have a certain influence on the development of CoF values but not as the normal load. The reciprocal sliding contact gave higher friction values than unidirectional. Note that the surface condition did not show any remarkable influence on the steady-stage friction values.
3.4. Wear characterization 3.4.1. Discs The worn surface of the discs for each test was studied and analyzed by means of SEM. At low load of 100 N, the dominant damage mecha nism for both cases, reciprocal and unidirectional sliding, was mild adhesive wear with removal of the surface material. For specimens tested in as-delivered surface conditions, the oxide layer was completely 6
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removed from the surface and the metallic surface was exposed to the sliding contact, Fig. 10a and b. The oxide layer did not reveal any pro tective properties on the as-delivered disc surfaces during the sliding contact. However, in the reciprocal sliding contact an accumulation of oxide particles was observed close to the center part of the worn area, Fig. 10a. With the increased normal load, a transition from mild to se vere adhesive wear was observed and the wear mechanisms were sur face cracks, delamination and plastic deformation, Fig. 10c and d.
contact, the degree of surface delamination was less and the material removal from the surface occurred as removal of thinner layers. 3.5. 3D profilometer examination 3D surface examination of the wear tracks showed that the worn depth increased with increasing the wear load, (FN � S). The average of the worn depth of each test is plotted against the wear load, Fig. 13. For the as-delivered surface under the unidirectional sliding, the worn depth at low loads was less than the depth of the oxide layer, B in Fig. 13. When the wear load increased, the wear track penetrated the oxide layer and the metal started to wear out. For the reciprocating sliding contact only at the lowest wear load the observed worn depth was less than the depth of the oxide layer, A in Fig. 13. In general, the worn depth reached higher values under the reciprocating sliding than the unidirectional sliding and higher initial wear observed for the ground surface. It was also observed that the width of the wear track was depended on the wear load and the type of motion. The width of the wear track extended from approximately 2 mm–3 mm when the total sliding dis tance increased from 100 m to 300 m under the same normal load, 500 N, Fig. 14. The influence of the type of motion on the wear track became more remarkable at higher loads. To illustrate that the worn volume for each test was measured and plotted against the wear load. It showed higher values in the case of reciprocal sliding compared to the unidirectional sliding, Fig. 15. It is seen from the worn volume diagram that the lines corresponding to the reciprocal tests, A and C, are above the lines corresponding to the unidirectional tests, B and D, Fig. 15. For the as-delivered surface it was found that the difference of worn volume between reciprocal and uni directional contact increased with increasing wear load. Whereas, for the ground surface, the difference in worn volume seemed to be constant and did not increase with an increase of the wear load. Nevertheless, despite the surface condition, it was found that the worn volume in
3.4.2. Plates SEM analysis revealed that the type of motion, normal load and sliding distance were the key parameters influencing the morphology of the wear tracks. At low loads, two distinguished morphologies were observed and they were strongly related to the type of motion. In the case of the reciprocal sliding, the surface was rough and the oxide layer was sheared in both directions, Fig. 11a. Whereas, for the unidirectional sliding contact, the worn surface was smoother and covered by the oxide layer. During the unidirectional sliding, the soft and porous initial oxide layer transformed into a thinner and denser layer containing surface cracks, Fig. 11b. SEM images taken by back-scattered electron detector revealed also brighter region at the surface that corresponded to the metallic material that had a higher atomic number than the oxide-rich material. It was also confirmed by the EDS analysis performed on the cross-section near the worn surface that the thickness of the oxide layer was reduced, Fig. 12. At high normal loads the worn surface was plastically deformed with clear traces of removed surface material. During sliding and at higher normal loads the oxide layer was removed completely from the surface, through crack propagation and delamination. The dominant wear mechanisms were plastic deformation and surface delamination when the contacting surfaces were under reciprocal sliding contact, Fig. 11c and d. Surface delamination was mainly caused due to the nucleation and propagation of subsurface cracks. For the unidirectional sliding
Fig. 10. Typical wear damage on the worn surface of the discs tested a) at 100 N and 200 m, b) 100 N and 200 m, c) 500 N and 300 m and d) 500 N and 100 m. Thick arrows indicate sliding direction. Small arrows and dashed shapes indicate surface cracks and surface delamination, respectively. 7
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Fig. 11. Typical wear damage on the worn surface of the as-delivered plates, a) and b) tested at 100 N and 100 m under, a) reciprocal sliding, b) unidirectional sliding, and c) and d) tested at 500 N and 100 m under c) reciprocal sliding and d) unidirectional sliding. White arrows indicate cracks at the surface.
Fig. 12. EDS analysis of the worn surface showing a thin and dense oxide layer formed at the worn surface.
reciprocal sliding contact was higher than the worn volume in unidi rectional contact. It is difficult to quantify the difference in terms of worn volume, it is seen that for the wear load of 150 � 103 N � m, the worn volume after reciprocal sliding was about 25–35% more than the worn volume of the unidirectional sliding at the same wear load. Further examination of the cross-sections near the worn surface showed that the subsurface material response was dependent on the type of motion. Under the unidirectional sliding contact the plastically deformed depth, for the surface tested at 500 N and 300 m, was about 40 μm. The microstructure of this region was plastically deformed and grains were elongated along the sliding direction, Fig. 16a. At the depth of 10 μm from the worn surface, the microstructure was transformed to an ultra-fined elongated grain structure. Under the reciprocated sliding contact the plastically deformed depth at high wear loads of 500 N and 300 m was about 18 μm. Grain elongation along the sliding direction particularly in reciprocal sliding
was observed towards only one direction even though the motion was reciprocating. Subsurface cracks were observed at the plastically deformed surface layer, Fig. 16b. Hardness profiles for the examined cross-sections are shown in, Fig. 16c. In the case of the unidirectional sliding a significant increase in hardness was observed. The depth of higher hardness region was about 100 μm that was bigger than the plastically deformed depth, 40 μm and 18 μm for unidirectional and reciprocal contact, respectively, Fig. 16. The increase in hardness was related to the work hardening of the sur face material. Unexpectedly, near the worn surface a reduction in the surface hardness was observed for the reciprocating sliding contact.
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pressure was at the maximum in the center of the elliptical contact area between the disc and plate and it decreases as moving towards the contact circumferences, Fig. 5a. The numerical calculations confirmed that the selection of minimum and maximum normal loads of 100 and 500 N ensures elastic contact at lower loads, 100 N, and plastic defor mation at high normal loads, 500 N. It is important that sliding tests can be performed at the contact pressures similar to the contact pressure presented in the thread joints. In an earlier study [5] it has been observed that the contact pressures between the threaded parts operated in real application, in certain locations exceeded the yield point of the surface material. Therefore, sliding tests at elastic and plastic contact conditions were performed. Note that the contact pressure calculations were performed in accordance to the stress-strain curve of the core material. The carburized layer that at actual surface in contact has other mechanical properties than the core material. Thus, if the calculations were performed based on the mechanical properties of the carburized layer the results will be slightly different due to the higher hardness of this layer. The onset of plastic deformation could be delayed for the carburized layer compared to the core material. However, as it is seen in Fig. 5b there is no large deviation between Hertzian and FEM results even at the highest normal loads used here. Moreover, the initial contact, between the disc and the plate, calculated by FEM was around 0.8 mm when a normal load of 500 N was used. The calculated width of the contact was in a good agreement with the measured width during experimental tests that was close to 0.72 mm. The experimental contact width was measured when
Fig. 13. Worn depth as a function of the sliding distance and normal load, A is as-delivered surface, reciprocal sliding, B is as-delivered surface, unidirectional sliding, C is ground surface, reciprocal sliding, D is ground surface, unidirec tional sliding.
4. Discussion 4.1. Contact pressure and CoF It is well known that wear of contacting surfaces is related to contact pressure and other factors as sliding speed, sliding distance, surface texture and material, humidity and temperature. Thus it was found of importance to estimate the initial static contact pressure between the contacting bodies. To this purpose an analytical and a numerical methods were employed, Hertzian theory and FEM respectively. The results obtained from both methods were to some extent overlapping, especially, at low normal loads 20–300 N. The Hertzian theory assumes a frictionless contact between the contacting bodies and it does not consider plastic deformation [10]. While for the FEM calculations, a coefficient of friction equal to 0.2 was introduced and the contact was modeled as elastic-plastic with strain hardening according to the stress-strain diagram of the tested material. Therefore, in the present results when the plastic deformation took place at higher loads, devia tion between the analytical and numerical results was noticed, Fig. 5b. According to an earlier study, the onset of plasticity under compression load starts when the maximum contact pressure reaches 1.67 σY , where σY is the yield strength of the loaded material [10]. The obtained results by numerical method, FEM, revealed that the plastic deformation occurred at the loads equal to and above 400 N, as it is seen from a rise of PEEQ values as plotted against the corresponding normal loads in Fig. 5b. The numerical results moreover showed that the contact
Fig. 15. Worn volume as a function of the sliding distance and normal load, A is as-delivered surface, reciprocal sliding, B is as-delivered surface, unidirec tional sliding, C is ground surface, reciprocal sliding, D is ground surface, unidirectional sliding.
Fig. 14. 3D optical profilometer images of 6 � 6 mm2 specimens of the wear track of the as-delivered surface tested under reciprocal sliding at 500 N after sliding, a) 100 m, b) 300 m. 9
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Fig. 16. a)-b) deformed texture of the ground surface tested at 500 � 300 Nm, a) under unidirectional sliding b) under reciprocal sliding, c) the hardness profiles of the worn surface shown in a and b.
the disc at 500 N load was pressed against the plate and then removed from plate. As explained before, the estimation of initial contact pres sures is of importance to define whether the initial contact will be in elastic or in plastic region. However, during the sliding contact and due to the tribological effects that take place between the contacting surfaces the contact area is continuously changing and so the contact pressure. Therefore, it would be difficult to model the precise contact pressures acting during sliding contact. Even though the obtained results are derived based on the core material still it is useful for the present work in order to define the level of the contact pressure when the disc is pressed against the plate. Carburizing is a quite common surface treatment to generate a harder surface with high wear resistance [4]. As a result of the case hardening by carburizing, an oxide layer was produced on the surface of all heat treated specimens. SEM analysis revealed that the oxide was not only produced on the surface but also penetrated along the grain boundaries. The oxide embrittles the grain boundaries, and it can be detrimental in applications where a cyclic tensile or bending load is applied. Nevertheless, in sliding contact, the grain boundary oxides may have less significant impact on the wear performance as the sliding contact involves a progressive material removal from one or both con tacting surfaces. However, in general the overall influence of the oxide layer on the performance of a surface in sliding contact is not fully un derstood and it was found of interest to investigate its impact on surfaces exposed to sliding contact. Sliding tests were performed on specimens with as-delivered and ground surfaces. The as-delivered specimens had an oxide layer at the surface with a mean thickness of 15 μm, while the specimens with the ground surface had an oxide layer of about 5 μm thick. It was observed
that the friction values, over the first 150 mm of sliding contact, for the as-delivered surface were the same as for the ground surface values. This behavior can be associated with surface flattening and wear of asperities and allows us to conclude that the initial surface roughness had an insignificant effect at sliding distances longer than 150 mm. It was also observed that the running-in region was relatively short compared to the steady state region and the transition region where friction increased. Therefore, no remarkable distinction between friction values for asdelivered and ground surfaces was observed. As sliding distance increased, the surface roughness was not influ encing the friction level as the high contact pressure did. Further fric tional behavior was rather controlled by adhesive wear and formation of mixed oxide-metal debris. Nevertheless, the level of the friction for asdelivered and ground surfaces was close to each other and it is believed that the thickness of the oxide layer did not have significant impact on the friction characteristics. At the same time the observed CoF values in the present study are lower than what have been reported in other studies for steel in dry contact. In Ref. [11] the reported CoF of dry steel to steel contact was around 0.8–1.5. The lower values observed in the present study may be related to the presence of the oxide layer that formed a tribolayer and thus prevented metal-to-metal contact. The CoF values of all tests showed two distinct regions, running-in and steady-state regions typical in dry sliding contact [12]. In the pre sent study, in the running-in region, it is believed that the first contact between specimens was oxide-to-oxide contact. The running-in region was short for all the tests. Though, it was to some extent longer for the tests performed at low load, 100 N and particularly in the case of uni directional sliding. Under further sliding the oxide layer broke into wear debris and it resulted in an increase in CoF values. In the steady-state 10
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region, the contact was a mixture of oxide-to-metal and metal-to-metal contact and the CoF found to be high and stable, 0.75. For the tests performed at higher loads, 500 N, the running-in region was much shorter compared to the tests performed at low loads. The running-in region was followed by an immidate increase and a steadystate region. The CoF values in this region were lower, around 0.45. Sliding contact under higher loads resulted in higher contact pressures and higher shear stresses that deformed surface material to a higher degree compared to the tests performed at lower loads. Plastic defor mation also resulted in grain refinement and thereby work hardening of the surface material. These factors may contribute to the lower friction values observed for the tests performed at high loads. Similar behavior of friction, for steels tested under dry sliding contact, where the CoF values in the steady state decreased with an increase in normal loads has been observed in other studies [12–15]. Generally the CoF is the sum of two components, adhesion friction (μa) and deformation friction (μd). Adhesion friction is the force required to break junctions when the opposite asperities are cold weld together during surface contact [16]. Bowden and Tabor developed a theory of adhesion friction and stated that the μa is dependent on the material properties of the contacting surface, μa ¼ Hτ [17]. Where τ is the shear strength of the junction and is close to the maximum shear strength of the softer material. H is the hardness of the softer surface material. The qffiffi second friction component is equal to μd ¼ 0:6 Rh. Where h is the
layer, Fig. 11b. It was also observed that the oxide layer became denser and thinner when the plate was exposed to the unidirectional sliding contact. As the normal load increased, larger shear strains were accu mulated at and beneath the surface resulting in crack formation and delamination of the surface material, Fig. 11d. In the case of reciprocal sliding contact, a rougher morphology was found at the worn track. Due to the type of motion the oxide layer was sheared in both direction. At high normal loads, the oxide layer mechanically mixed with the surface material forming wear debris. Under further sliding the wear debris were compacted and flattened. As well, further sliding resulted in the surface delamination, Fig. 11c. The significant surface delamination observed on the worn surface of the discs was not found on the plates. For the plates, delamination was observed only in the case of reciprocal sliding motion. The reason for the remarkable surface delamination of the discs was that the same area was in contact with the plate until the test was stopped. Whereas for the plates each point at the worn surface only experience the sliding contact loads once when the disc slid along the plate surface. It means that the material at each point was resting for a specific time before the next sequence of sliding contact taking place. Therefore, surface cracking and delamination were the dominant damage features for the worn surface of the discs at higher loads. Essentially, the dominant wear mechanisms at high normal loads for plates were plastic deformation and surface smoothing accompanied by surface shearing and delamination of wear debris, Fig. 11c and d.. These wear mechanisms were observed in all wear tests regardless of the initial thickness of the oxide layer. Earlier studies have also confirmed the smoothening of wear track at high normal loads as the effect of plastic deformation [13,19]. Concerning the subsurface shear stress distribution and its influence on CoF, a study has proposed that when the CoF values are smaller than 0.25, the maximum shear stress will be in the subsurface. In the other hand, when the CoF values are higher than 0.25 the maximum shear stresses are at the surface [20]. Nevertheless, in the present study, as revealed by cross-section examinations, plastic deformation occurred at the surface and in the subsurface. What was notable is that the plastic deformation depth in unidirectional sliding was twice deeper than in the reciprocal sliding contact, Fig. 16a and b.. In the unidirectional sliding, the plastically deformed surface material consisted of two regions. The first regions corresponded to ultrafine grains and had a depth of about 10 μm. The second region, 30 μm in depth, corresponded to the plasti cally elongated texture along the sliding direction. Similar regions were observed for the reciprocal sliding contact. Contrary, the plastic flow depth along the sliding direction was shal lower than the corresponding depth in the unidirectional sliding contact. The depth and number of subsurface cracks were also higher in the reciprocal contact. The reason for more extensive subsurface cracking in the reciprocal contact could be the reversed plastic deformation of the surface material. In the unidirectional contact the surface was sheared in one direction and the material was responding to the loading by strengthening its texture through work hardening. In fact, finer grains and plastic deformation observed beneath the worn surface are two fundamental strengthening mechanisms for metals. In the present work, the measured hardness profiles of the worn surface tested under unidi rectional contact had higher hardness values at a depth up to 100 μm, Fig. 16c It confirmed that work hardening phenomenon occurred. However, in the reciprocal sliding a similar work hardening and increase in hardness was not observed. Instead, a lower surface hardness was obtained. The explanation for this unexpected behavior could be that the high density of the subsurface cracks affected the hardness values and these cracks, served as voids, decreased the hardness. According to an earlier study, more subsurface crack will result in high degree of surface delamination during further sliding [21]. Higher surface delamination then leads to higher worn volume because of the proportional rela tionship between surface delamination and wear rate. 3D surface examinations were performed to quantify the worn depth and the worn volume of the wear track as a function of the wear load.
depth that a circular body, an asperity, makes when it is pressed against another surface and R is the radius of the asperity. However, Bowden and Tabor stated that in the case of low surface roughness the defor mation component is relatively small compared to the adhesion component. Anyway, during the sliding contact and due to the tribo logical effects temperature may rise, surface hardness and shear strength may change, cracks may form. All these factors will for sure influence the friction level between the contacting surfaces. In the present study, it is clearly seen that CoF values decreased with increased normal load, Fig. 8. The decrease in friction at high normal loads could be related to smoothing of the contact area and work hardening. Moreover, the high plastic deformation invoked by the high contact pressure, have removed and/or flattened the contact asperities as was experimentally observed. 4.2. Wear characterization and quantification
The wear characterization of the worn track revealed different type of wear damages: mild to severe adhesive wear, abrasive wear and crack formation and surface delamination. Depending on the type of motion, initial surface condition and the wear load the different wear mecha nisms will take place. The type of motion influences strongly the surface texture and de termines how the surface material deformed, e.g. the direction of the deformed texture [18]. In the unidirectional sliding in the present work, the surface material was sheared in one direction. For each passage of the disc over the plate an increment of shear strain was accumulated at the surface, and also in the subsurface depending on where the maximum shear stresses are located. Whereas in the reciprocal sliding, the material flow was reversed due to the reciprocating motion of the disc. However, for the discs the worn morphology was quite similar despite the type of motion. At low loads, the oxide layer was removed from the surface and the metallic surface appeared, Fig. 10a and b. At high load, 500 N, where the maximum contact pressures exceed the yield strength of the surface material, the strain accumulation was high enough to cause surface, and subsurface cracks and surface delamina tion, Fig. 10c and d. In contrary, the worn morphology of the plates was dependent on the type of motion. At low normal loads where the initial contact pressure was around the elastic limit, surface cracks and surface delamination were observed at low loads. However, these damage mechanisms occurred in the oxide layer and resulted in the surface cracks of the oxide 11
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Both the worn depth and worn volume increased with increasing wear load, Figs. 13 and 15. Moreover, the width of the wear track extended from approximately 2 mm–3 mm when the total sliding distance increased from 100 m to 300 m under the same normal load, 500 N, Fig. 14. The width increase of the wear track was the result of a com bination of material removal and plastic deformation. Material removal was mostly due to adhesive wear and surface delamination. Accumu lation of plastic deformation during each load passage made the surface material to plough to the sides of the wear track. The worn depth and worn volume of the wear track became pro portional to the wear load regardless the type of sliding mode and sur face condition. The general trend of the worn volume observed in the present study does follow the Archard’s wear equation, to some extent, Fig. 15. However, this wear equation is limited to only one wear mechanism where the harder surface ploughs the softer. Commonly, the worn volume is linearly proportional to the normal load and sliding distance and inversely proportional to the surface hardness of the softer material [6]. Whilst in the present case several wear mechanisms such as adhesive wear, abrasive wear and plastic deformation were acting simultaneously and contributed to the total wear of the surface material. The measured worn volume values in the reciprocal sliding tests were higher than in the unidirectional sliding in every wear load, Fig. 15. During the forward sliding contact, material transfer preferred to accumulate at the back side of the contact area of the disk. As the disk slid back, the transferred material resulted in a certain increase in the friction values and also ploughed the plate material. This behavior was not presented in the case of unidirectional sliding since the disk slid only one direction. In addition, higher worn volume in the case of reciprocal sliding could be related to the loading condition and the strain accu mulation in the material during sliding contact. In the unidirectional sliding, the surface material was deformed in one direction and loading condition was repeated uniaxial loading that resulted in hardening of the surface material. Whereas in the reciprocal sliding contact, the loading condition was repeated reversed cyclic loading. It resulted in higher surface cracking and delamination of the surface material. Worn volume for the as-delivered and ground surface were compared to each other based on the type of motion. It was found that the dif ference in the worn volume between the reciprocal and unidirectional contact for the as-delivered surface increased with increasing wear load. While for the ground surface, the difference in worn volume between these two sliding modes seemed to be constant. The reason for the observed behavior could be due to the initial oxide layer presented at the as-delivered surface. This layer served as a protection of the surface material and, thereby, the wear rate was low for both the reciprocal and unidirectional contact. As the wear load increased a higher wear rate was obtained for the reciprocal contact. Indeed, increasing normal load increases worn volume that in turn will increase the wear rate, as it has been observed by other researchers [15,19]. In the present study, tests using the same contact parameters, normal load 500 N and sliding distance 300 m, in different sliding modes, reciprocal and unidirectional sliding, resulted in the different total worn volumes. The maximum worn depth and worn width were 50 μm and 2.77 mm and 37 μm and 2.95 in the case of reciprocal and unidirectional sliding, respectively. During the reciprocal sliding, it is believed that wear debris were formed and stayed at the interface resulting in three-body abrasive wear. These loose abrasive particles presumably resulted in higher wear rates. Thus, larger worn depth was observed in the reciprocal sliding than in the unidirectional sliding. However, at low wear loads a smaller worn depth was observed in the as-delivered conditions, and it points towards an initial protective effect of the thin oxide layer.
The influence of the type of motion, the reciprocal and unidirectional, and the oxide layer on the steel performance in sliding contact were studied. It was found that the type of motion had a much higher effect on the tribological performance of the steel in sliding contact than the other test parameters. Obtained results can be summarized as follow: � The thickness of the oxide layer did not influence much the tribo logical performance of the investigated steel. Instead, the type of motion had a much higher effect on the tribological performance. � The reciprocal sliding tests resulted in 25–35% higher worn volume than the unidirectional sliding tests under the same contact condi tions, normal load and sliding distance. � CoF decreased with increasing normal load and its values for the reciprocal sliding were slightly higher than for the unidirectional sliding. � At low load, 100 N, and short sliding distances, the oxide layer of the as-delivered specimens tended to cover the contact. Ground speci mens tested at the same conditions a worn surface typical for adhe sive wear, with some islands of oxides. � At higher normal loads 300–500, the deformed texture consisted of two regions. One region characterized with ultra-fine grains with a depth up to 10 μm and the other with elongated grains along sliding direction. The depth of the deformed texture was depended on the type of motion. Higher depth observed at the unidirectional sliding contact. � Crack formation and propagation was observed in both the unidi rectional and reciprocal sliding contact. Though more subsurface cracks observed at the reciprocal sliding contact. Acknowledgments The funding of the present work by the Swedish Knowledge Foun dation, project no 20150090, Epiroc Rock Drills AB, Ovako AB, Sandvik Mining and Rock technology AB is gratefully acknowledged. References [1] D. Zou, Theory and Technology of Rock Excavation for Civil Engineering, Springer, 2017. [2] G. Krauss, D. Matlock, Effects of strain hardening and fine structure on strength and toughness of tempered martensite in carbon steels, J. Phys. IV 5 (C8) (1995). C8C51-C8-60. [3] G. Krauss, Heat treated martensitic steels: microstructural systems for advanced manufacture, ISIJ Int. 35 (4) (1995) 349–359. [4] S. Lampman, Introduction to surface hardening of steel ASM Handbook, Heat Treating, ASM International, Metals Park, Ohio 4 (1991), 1991. [5] A. Mussa, P. Krakhmalev, J. Bergstr€ om, Failure analyses and wear mechanisms of rock drill rods, a case study, Eng. Fail. Anal. 102 (2019) 69–78. [6] I. Hutchings, P. Shipway, Tribology: Friction and Wear of Engineering Materials, Butterworth-Heinemann, 2017. [7] A. Gaard, P.V. Krakhmalev, J. Bergstrom, N. Hallback, Galling resistance and wear mechanisms - cold work tool materials sliding against carbon steel sheets, Tribol. Lett. 26 (1) (2007) 67–72. [8] A. Gaard, P. Krakhmalev, J. Bergstrom, Wear mechanisms in galling: cold work tool materials sliding against high-strength carbon steel sheets, Tribol. Lett. 33 (1) (2009) 45–53. [9] A. Gåård, P. Krakhmalev, J. Bergstr€ om, Wear mechanisms in deep drawing of carbon steel - correlation to laboratory testing, Tribotest 14 (1) (2008) 1–9. [10] K. Johnson, Contact Mechanics, Cambridge University Press, Cambridge, UK, 1985. [11] V. Angelini, I. Boromei, C. Martini, C. Scheuer, R. Cardoso, S. Brunatto, L. Ceschini, Dry sliding behavior (block-on-ring tests) of AISI 420 martensitic stainless steel, surface hardened by low temperature plasma-assisted carburizing, Tribol. Int. 103 (2016) 555–565. [12] P.J. Blau, Interpretations of the friction and wear break-in behavior of metals in sliding contact, Wear 71 (1) (1981) 29–43. [13] M. Ulutan, O.N. Celik, H. Gasan, U. Er, Effect of different surface treatment methods on the friction and wear behavior of AISI 4140 steel, J. Mater Sci Tech 26 (2010) 251–257. [14] M. Ruiz-Andres, A. Conde, J. De Damborenea, I. Garcia, Friction and wear behaviour of dual phase steels in discontinuous sliding contact conditions as a function of sliding speed and contact frequency, Tribol Int 90 (2015) 32–42. [15] R. Tyagi, S. Nath, S. Ray, Effect of martensite content on friction and oxidative wear behavior of 0.42 Pct carbon dual-phase steel, Metall. Mater. Trans. A 33 (2002) 3479–3488, 2002.
5. Conclusions The friction and wear characteristics of a high strength steel, 22NiCrMo12–F, commonly used as rock drill steel were investigated. 12
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[16] R. Gohar, H. Rahnejat, Fundamentals of Tribology, World Scientific Publishing Company, 2012. [17] F. Bowden, A. Moore, D. Tabor, The ploughing and adhesion of sliding metals, J. Appl. Phys. 14 (2) (1943) 80–91. [18] D.H. Persson, S. Jacobson, S. Hogmark, Antigalling and low friction properties of a laser processed Co-based material, J. Laser Appl. 15 (2003) 115–119, 2003.
[19] H. So, D. Yu, C. Chuang, Formation and wear mechanism of tribo-oxides and the regime of oxidational wear of steel, Wear 253 (9) (2002) 1004–1015. [20] A. Bower, K. Johnson, The influence of strain hardening on cumulative plastic deformation in rolling and sliding line contact, J. Mech. Phys. Solids 37 (4) (1989) 471–493. [21] N.P. Suh, The delamination theory of wear, Wear 25 (1) (1973) 111–124.
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Wear journal homepage: http://www.elsevier.com/locate/wear
Sliding wear and fatigue cracking damage mechanisms in reciprocal and unidirectional sliding of high-strength steels in dry contact €m Abdulbaset Mussa *, Pavel Krakhmalev, Jens Bergstro Karlstad University, Department of Engineering and Physics, SE-658 88, Karlstad, Sweden
A R T I C L E I N F O
A B S T R A C T
Keywords: Rock drill rods Reciprocal sliding contact High strength steel Fatigue cracks Sliding wear and plastic deformation
Rock drill components operate under tough contact conditions during rock drilling. Reciprocal and unidirec tional motion under high contact stresses are the common contact conditions between interconnected compo nents. It will result in component damage and often the observed surface damage of rock drill tools is due to wear and fatigue cracks. Nevertheless, the effects of the properties and structure of the mating materials on tribo logical performance, is not fully understood. The present study is dedicated to simulation and investigation of the wear mechanisms observed in reciprocal and unidirectional sliding of high strength steels for rock drill com ponents. A high strength martensitic steel, 22NiCrMo12–F, commonly used in rock drills was tested in selfmating contact. Wear mechanisms were investigated by means of electron microscopy and wear damage was quantified by a 3D optical interferometer. Total damage, as a result of adhesive wear, severe plastic deformation and nucleation and propagation of fatigue cracks, was discussed in relation to test conditions and material properties. It was observed that the coefficient of friction decreased with increasing normal load. Moreover, the results showed that the type of motion had a significant influence on the worn volume and crack nucleation of the specimens in sliding contact. In addition, the reciprocal motion resulted in higher wear than unidirectional motion under the same test conditions.
1. Introduction In rotary and percussive rock drilling high energy waves are used to break a rock material into smaller pieces by impacting the rock. Rock drills are usually driven by a hydraulic system that generates energy in terms of shock waves and transmits it to a drill bit through a rotating drill rod components. In order to drill a hole with a diameter of 35–89 mm, a hydraulic pressure of 20 MPa is used. It provides a fre quency up to 50 Hz, rotating rate up to 200 rpm and a feeding thrust, in order to maintain contact between rock and drill bit, about 10 MPa [1]. These application parameters expose rock drill components to extreme contact situations due to high pressure sliding contact and repeated cyclic load. One example of such components is thread joints, Fig. 1, used to connect extended drill rods or to connect a drill rod to a drill bit. During rock drilling, high pressure sliding contact between threaded surfaces takes place. It leads to successive surface damages that worn out surface material. Wear of the thread joints is of an important techno logical and economic significance. It deteriorates the surface material and reduces the tightness between connecting components that results
in an inefficient drilling process with increased drilling time and costs. Commonly, materials with high hardness and high wear resistance are used to minimize wear rate between contacting components. How ever, the high hardness of a component should not compromise its fa tigue life. Therefore, it is important to select materials with a proper balance between the hardness and the toughness. In rock drill components, it is of high importance to use materials with high surface hardness in order to achieve an adequate wear resis tance and high core toughness to have longer fatigue life of a compo nent. To this matter, martensitic high strength steels with low carbon content are used in rock drills [2,3]. Surface hardening of the involved components is frequently used to achieve high wear resistance [4]. Carburizing is a quite common surface treatment to generate a hard surface layer without altering mechanical properties of the core material. In rock drill rods, connectors and other components, the relative motion depends on a location and a component. Often a unidirectional motion is dominating, but in some components, reciprocal motion is observed. A case study previously performed by the authors [5], on the
* Corresponding author. E-mail addresses: [email protected] (A. Mussa), [email protected] (P. Krakhmalev), [email protected] (J. Bergstr€ om). https://doi.org/10.1016/j.wear.2019.203119 Received 25 January 2019; Received in revised form 11 September 2019; Accepted 8 November 2019 Available online 9 November 2019 0043-1648/© 2019 Elsevier B.V. All rights reserved.
Please cite this article as: Abdulbaset Mussa, Wear, https://doi.org/10.1016/j.wear.2019.203119
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2.2. Specimen preparation Two different specimen geometries were used in the present study, discs and plates. The disc specimens had a major diameter of 50 mm and minor diameter of 10 mm. The thickness of the discs was 10 mm. The plate specimens had a length of 600 mm, a width of 100 mm and 10 mm in thickness. All tested specimens were carburized in a well-controlled and carbon rich atmosphere (N2 þ CH3OH) followed by water quench ing and tempering. The carbon content in the case hardened layer was up to 0.6 wt %, the tempered martensitic microstructure of the case hardened layer is presented in Fig. 2 a. The carburized case depth was 1.2 mm, the surface hardness was up to 720 HV and the core hardness was around 450 HV, Fig. 2b. As a result of the case hardening by carburizing, an oxide layer was formed on the surface of the heat treated specimens. In order to inves tigate the influence of the oxide layer on the material performance in sliding contact the specimens were tested under two different surface conditions, as-delivered and ground surface conditions. On the ground surface, the most intensive oxide was removed from the surface by manual grinding using SiC paper, started with rough to smoother using 320, 500 and 2400 grit papers. Prior to wear testing, the surface roughness of the tested specimens were measured and evaluated by the means of a 3D optical profiler, Bruker 3D.
Fig. 1. Schematic of the thread joint between the drill rod and drill bit with rotational drilling and axial percussion movements during the rock dril ling process.
failure mechanisms of the drill rod and thread joints failed under in dustrial conditions has revealed that the surface material in the threaded parts is often plastically deformed when it is used in the application. Severe adhesive wear due to a high-pressure sliding contact between connected threads took place. It has also been revealed that grain refinement of the surface material occurred in locations where plastic deformation took place. Additionally, fatigue cracks as a result of sliding contact between threaded parts have been observed [5]. The complexity of the contact situation in rock drilling, the type of relative motion, high mechanical stresses, temperature and corrosion, make it difficult to predict the performance and wear resistance of a particular material in sliding contact. Therefore, experimental tests at a laboratory level under well-controlled test parameters are of high importance to enable pre diction of the wear resistance of a tested material in a particular appli cation. Experimental studies have declared several key parameters that have an essential effect on the surface performance in sliding contacts, e. g. mechanical properties of the surface material, microstructure, surface roughness, contact pressure and sliding distance [6]. Additionally, the contact conditions change over time and, therefore, the wear mecha nisms change continuously. The aim of the present study is to under controlled laboratory tests simulate the wear mechanisms of the rock drill steels in sliding contact. Commonly, the tested steel is used in the thread joints to connect drill rods with drill bits. Furthermore, the aim was also to investigate the influence of surface condition and type of motion on friction and wear resistance in self-mating contact of a case hardened martensitic steel. In order to achieve the aim of the present study, wear tests under dry reciprocal and unidirectional sliding contact were conducted. A range of loads and sliding distances were used in order to study the correlation between contact pressures and different wear regimes on the tested specimens.
2.3. Wear testing and characterization Wear tests were performed using the Slider On Flat Surface (SOFS) tribometer [7]. The SOFS has been proven to simulate adhesive wear processes, and provide sufficient characterization of adhesive wear resistance of steels in sliding contact at high pressures [7–9]. Since the main wear mechanisms revealed by the case study [5] was severe ad hesive wear, the SOFS was decided to be suitable for investigation and evaluation of wear resistance of the steel in the present research. In the SOFS, a disc-shaped specimen is pressed against a flat plate, Fig. 3a, under controlled normal load and slid until the desired sliding distance (S) is achieved. The disc is fixed in a way that rotation movement during sliding is avoided. The tests were conducted under dry reciprocal and unidirectional repeating sliding. In the present investigation, the reciprocal sliding tests involved pressing the disc with a predetermined normal load (FN) against the plate and sliding with a constant speed of 0.3 m/s and a stroke length of 150 mm. The reciprocal movement of the disc was repeated in the same track until the predetermined total sliding distance was achieved. In the unidirectional sliding tests, the disc was pressed and slid against the plate under a predetermined normal load FN with a constant speed of 0.3 m/s. When the disc reached the end of the stroke, 150 mm, it was lifted up from the plate and moved back to the starting position. This operation was repeated in the same track until the predefined total sliding distance was reached. Wear tests were conducted at three different normal loads, of 100, 300 and 500 N, and the sliding distances of 100, 200 and 300 m. In the real application, the contact pressures between the threaded parts in some extreme locations exceed the yield strength of the surface material. This was revealed when a detailed examination of wear characteristics of the failed thread joints used in the field application has been per formed [5]. Plastic deformation of the surface material and severe sliding wear were the dominant wear mechanisms. The selection of the normal load range between 100 – 500 N ensures elastic initial contact at 100 N and elastic-plastic initial contact at 500 N between the discs and
2. Experimental details 2.1. Material All specimens used in the present study, the discs and plates, were made of an advanced high strength steel with low carbon content and high hardenability, 22NiCrMo12–F. The chemical composition and the typical mechanical properties of the steel grade are presented in Table 1.
Table 1 Chemical composition, wt. %, and mechanical properties of the core material. C
Si
Mn
P
S
Cr
Ni
Mo
Al
Rp0.2a [MPa]
Rm [MPa]
0.20
0.30
0.70
0.02
0.02
1.30
2.95
0.25
0.03
1050
1500
a
Rp0.2. and Rm are the proof strength and the ultimate strength of the tested material, respectively. 2
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Fig. 2. a) Case hardened tempered martensitic microstructure at 0.5 mm depth and b) hardness profile from the surface into the bulk material.
Fig. 3. a) The SOFS tribometer configuration and b) schematic of a wear test. Specimens, sliding direction, normal load FN and friction force Ff are marked with black arrows.
plates. The sliding distances selected for the present work were sufficient to produce a wear damage that can be characterized. Moreover, these sliding distances were selected based on the practical experimental reasons and also based on the pervious experimental experiences. Dur ing the wear tests the instantaneous coefficient of friction (CoF) between the movable disc and fixed plate was recorded and monitored. In order to visualize and quantify wear of the worn surfaces, the 3D profiler was used. Samples from locations near the starting, middle and ending point of the wear track, of each test, were cut from the plates with dimensions 6 � 6 mm2 and used for wear analysis. The worn depth and the worn volume, V, from the wear track of each test were measured and analyzed. Five measurements from five different locations for each tested specimen were conducted. The average of the measured values combined with its standard deviations were plotted against the wear load, according to the Archard’s wear equation [6]. V¼
K FN S H
mounted in thermoplastic, ground and polished according to the stan dard metallographic sample preparation method. The prepared crosssection samples were etched in 5% Nital and examined using RiechertJung Polyvar metallography microscope. The hardness of the worn surfaces were measured and analyzed by a microhardness tester, FutureTech FM, at 25 gf. 2.4. Contact pressure calculations When running laboratory wear tests, it is important that the tests simulate key parameters involved in a real application such as contact pressure, contact mode, etc. The contact pressures obtained between the contacting bodies in the present study were estimated using analytical and numerical methods. The initial static contact pressure between the disc and plate was calculated using the Hertzian elastic contact theory and the Finite Element Method (FEM). A simplified 3D FEM-model of the contact was created in ABAQUS. Both disc and plate were modeled as elastic-plastic von Mises materials with strain hardening according to the stress-strain behavior of the core material, Fig. 4a. Note that the carburized layer has other mechanical properties than the core material. Due to the symmetry only a one-quarter part of the contact area was used in the calculations. The FEM-model was created based on the fol lowings. Both disc and plate, face 1 and 2 in Fig. 4b, had their translation locked in the X- and Y-directions and the bottom face of the plate was constrained in all X-, Y- and Z-directions. A CoF equal to 0.2 was assumed between contacting surfaces. The normal load was adjusted to a uniform distributed pressure and applied to the top surface of the disc segment, red arrows in Fig. 4b. A C3D10 Element type with a 10-node
(1)
K is the coefficient of wear and is dimensionless and always less than unity, H is the hardness of the softer material, note that the contacting specimens have the same hardness. The ratio K/H is assumed to be constant because of the self-mating contact. The product of the normal load and the sliding distance (FN � S) is designated as the wear load in the present study. The operating wear mechanisms of the worn specimens were investigated. Microscopic analysis of the wear track on the plate and the worn contact of the disc were performed with Leo 1530 FEG scanning electron microscope (SEM) equipped with Oxford EDX INCA-sight sys tem. For the metallographic study samples were cross-sectioned, 3
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Fig. 4. a) the tensile stress-strain diagram of the tested material, b) 3D FE-model represents one-quarter of the contact between the disc and the plate, c) the refined mesh in the contact area.
quadratic tetrahedron with controlled hourglass shape was used. The mesh was refined in the contact area, to approximately 0.2 mm, in order to obtain higher precision of the contact pressure, Fig. 4c.
3.2. Surface conditions Prior to testing the surface roughness for the selected specimens were measured and evaluated by the 3D optical profiler. It was found that the discs in as-delivered condition had a surface roughness Ra around 0.65 μm while the ground discs had the roughness of 0.51 μm. For plates the difference between the roughness values was larger, Fig. 6a and b. In the case of as-delivered surface the roughness values Ra were around 7.7 μm whereas for the ground surface the roughness values Ra were around 0.8 μm. SEM analysis revealed that the initial oxide layer was porous at the very top and it became denser towards the interface between the oxide layer and the steel surface, Fig. 6c. The layer thickness varied along the surface with an average thickness of 15 μm, Fig. 6c. For the ground surface, the oxide layer depth was reduced to a thickness of approxi mately 5 μm, Fig. 6d. The hardness of the oxide layer on the as-delivered was 290 HV. In both surface conditions, as-delivered and ground sur face, grain boundary oxides were observed, marked by black arrows in Fig. 6c and d. An EDS analysis at higher magnification was performed on the grain boundaries and the surrounding material. It revealed enhanced content of silicon and oxygen at the grain boundaries, Fig. 7.
3. Results 3.1. Contact pressure The maximum static contact pressure for normal loads between 20 700 N were calculated by two different methods, a numerical method by FEM calculations and an analytical method by Hertzian elastic contact theory for elliptical contacts. The FEM calculations showed that width of the initial contact between the disc and the plate was around 0.8 mm when a normal load of 500 N was used. The calculated width of the contact was in a good agreement with the measured width during experimental tests that was close to 0.72 mm. Moreover, the FEM results revealed that the maximum contact pressure was concentrated to the center of the ellipse, contact area between the disc and plate, and the contact pressure decreased as moving towards the circumferences, Fig. 5a. The results, calculated by FEM and Hertzian analyses, are plotted against the corresponding normal loads in Fig. 5b. At low normal loads, only a small deviation was observed between the two methods. The deviation became larger with increased normal load up to 500 N and higher, Fig. 5b. Equivalent plastic strain (PEEQ) values were also plotted against the normal loads. When PEEQ value for a certain element ex ceeds zero, plastic strain occurred. As observed in the obtained results plastic deformation occurred immediately when the disc was pressed against the plate with a normal load of 400 N.
3.3. Friction The instantaneous CoF between the discs and plates was recorded during the tests. The CoF values of all tests showed three distinct regions, the running-in and steady-state regions. Between these two regions a third region was presented where the CoF increased. The running-in region was relatively short where the CoF started with lower values
Fig. 5. Disc and plate contact pressure analysis, a) FEM contact pressure distribution for a selected test at 500 N as normal load and b) Equivalent plastic strain and maximum contact pressure as the function of the normal load, comparison between Hezrtian contact calculation and FEM calculation. 4
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Fig. 6. SEM images of the plates, a-b) shows the topography of the oxide layer a) as-delivered and b) ground surface, and c-d) shows the thickness of the oxide layer, c) as-delivered and d) ground surface. Arrows indicating grain boundary oxides.
Fig. 7. EDS analysis of the grain boundary oxide, a) SEM image of the grain boundaries and b) EDS spectrum of the area of interest.
and increased, in the first meters of the sliding contact, to a higher value corresponding to the steady-state region, Fig. 8. The development of the CoF over the first millimeters of sliding for each test was studied in order to investigate the influence of surface roughness on friction between contacting surfaces. For all the tests conducted it was observed that the starting friction values were around 0.3–0.4. Then it increased to slightly higher values of 0.6–0.7. After sliding for 30–50 mm and longer, the CoF decreased to 0.2–0.3. This behavior was the same for as-delivered and ground surfaces. After that the CoF values were about the same during running-in region. At low normal loads, 100 N, and particularly under unidirectional sliding mode, the running-in region was longer than in the tests performed at higher normal loads. Moreover, the increase in the friction values to reach the steady state, for the tests performed at low loads, required
longer sliding distance. The steady state of CoF at low load showed a stable mean value until the test was stopped. At higher normal loads, the CoF increased immediately after short distance to higher values where the steady state region of CoF started. Generally, the mean CoF value in the running-in region for all tests was between 0.2 and 0.3 and it increased with increasing sliding distance. Thereafter the CoF during the steady-state did not change remarkably and remained at the same level until the test was finished. The mean value of the CoF at the steady-state obtained from the wear tests was plotted against the normal load and sliding distance, Fig. 9. It was observed that the CoF was strongly dependent on the normal load. In the case of the as-delivered surface tested under the reciprocal sliding, the CoF value decreased from 0.72 to 0.45 with increased normal load from 100 to 500 N, Fig. 9a. Similar trend was observed for other tests, 5
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Fig. 8. CoF between the disc and plate tested at a normal load of 100 N and a sliding distance of 100 m during unidirectional sliding contact.
Fig. 9. Influence of normal load and sliding distance on the CoF, a) as-delivered surface, reciprocal sliding, b) as-delivered surface, unidirectional sliding, c) ground surface, reciprocal sliding and d) ground surface, unidirectional sliding.
Fig. 9b, c and 9d. The type of motion, unidirectional or reciprocal sliding, did have a certain influence on the development of CoF values but not as the normal load. The reciprocal sliding contact gave higher friction values than unidirectional. Note that the surface condition did not show any remarkable influence on the steady-stage friction values.
3.4. Wear characterization 3.4.1. Discs The worn surface of the discs for each test was studied and analyzed by means of SEM. At low load of 100 N, the dominant damage mecha nism for both cases, reciprocal and unidirectional sliding, was mild adhesive wear with removal of the surface material. For specimens tested in as-delivered surface conditions, the oxide layer was completely 6
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removed from the surface and the metallic surface was exposed to the sliding contact, Fig. 10a and b. The oxide layer did not reveal any pro tective properties on the as-delivered disc surfaces during the sliding contact. However, in the reciprocal sliding contact an accumulation of oxide particles was observed close to the center part of the worn area, Fig. 10a. With the increased normal load, a transition from mild to se vere adhesive wear was observed and the wear mechanisms were sur face cracks, delamination and plastic deformation, Fig. 10c and d.
contact, the degree of surface delamination was less and the material removal from the surface occurred as removal of thinner layers. 3.5. 3D profilometer examination 3D surface examination of the wear tracks showed that the worn depth increased with increasing the wear load, (FN � S). The average of the worn depth of each test is plotted against the wear load, Fig. 13. For the as-delivered surface under the unidirectional sliding, the worn depth at low loads was less than the depth of the oxide layer, B in Fig. 13. When the wear load increased, the wear track penetrated the oxide layer and the metal started to wear out. For the reciprocating sliding contact only at the lowest wear load the observed worn depth was less than the depth of the oxide layer, A in Fig. 13. In general, the worn depth reached higher values under the reciprocating sliding than the unidirectional sliding and higher initial wear observed for the ground surface. It was also observed that the width of the wear track was depended on the wear load and the type of motion. The width of the wear track extended from approximately 2 mm–3 mm when the total sliding dis tance increased from 100 m to 300 m under the same normal load, 500 N, Fig. 14. The influence of the type of motion on the wear track became more remarkable at higher loads. To illustrate that the worn volume for each test was measured and plotted against the wear load. It showed higher values in the case of reciprocal sliding compared to the unidirectional sliding, Fig. 15. It is seen from the worn volume diagram that the lines corresponding to the reciprocal tests, A and C, are above the lines corresponding to the unidirectional tests, B and D, Fig. 15. For the as-delivered surface it was found that the difference of worn volume between reciprocal and uni directional contact increased with increasing wear load. Whereas, for the ground surface, the difference in worn volume seemed to be constant and did not increase with an increase of the wear load. Nevertheless, despite the surface condition, it was found that the worn volume in
3.4.2. Plates SEM analysis revealed that the type of motion, normal load and sliding distance were the key parameters influencing the morphology of the wear tracks. At low loads, two distinguished morphologies were observed and they were strongly related to the type of motion. In the case of the reciprocal sliding, the surface was rough and the oxide layer was sheared in both directions, Fig. 11a. Whereas, for the unidirectional sliding contact, the worn surface was smoother and covered by the oxide layer. During the unidirectional sliding, the soft and porous initial oxide layer transformed into a thinner and denser layer containing surface cracks, Fig. 11b. SEM images taken by back-scattered electron detector revealed also brighter region at the surface that corresponded to the metallic material that had a higher atomic number than the oxide-rich material. It was also confirmed by the EDS analysis performed on the cross-section near the worn surface that the thickness of the oxide layer was reduced, Fig. 12. At high normal loads the worn surface was plastically deformed with clear traces of removed surface material. During sliding and at higher normal loads the oxide layer was removed completely from the surface, through crack propagation and delamination. The dominant wear mechanisms were plastic deformation and surface delamination when the contacting surfaces were under reciprocal sliding contact, Fig. 11c and d. Surface delamination was mainly caused due to the nucleation and propagation of subsurface cracks. For the unidirectional sliding
Fig. 10. Typical wear damage on the worn surface of the discs tested a) at 100 N and 200 m, b) 100 N and 200 m, c) 500 N and 300 m and d) 500 N and 100 m. Thick arrows indicate sliding direction. Small arrows and dashed shapes indicate surface cracks and surface delamination, respectively. 7
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Fig. 11. Typical wear damage on the worn surface of the as-delivered plates, a) and b) tested at 100 N and 100 m under, a) reciprocal sliding, b) unidirectional sliding, and c) and d) tested at 500 N and 100 m under c) reciprocal sliding and d) unidirectional sliding. White arrows indicate cracks at the surface.
Fig. 12. EDS analysis of the worn surface showing a thin and dense oxide layer formed at the worn surface.
reciprocal sliding contact was higher than the worn volume in unidi rectional contact. It is difficult to quantify the difference in terms of worn volume, it is seen that for the wear load of 150 � 103 N � m, the worn volume after reciprocal sliding was about 25–35% more than the worn volume of the unidirectional sliding at the same wear load. Further examination of the cross-sections near the worn surface showed that the subsurface material response was dependent on the type of motion. Under the unidirectional sliding contact the plastically deformed depth, for the surface tested at 500 N and 300 m, was about 40 μm. The microstructure of this region was plastically deformed and grains were elongated along the sliding direction, Fig. 16a. At the depth of 10 μm from the worn surface, the microstructure was transformed to an ultra-fined elongated grain structure. Under the reciprocated sliding contact the plastically deformed depth at high wear loads of 500 N and 300 m was about 18 μm. Grain elongation along the sliding direction particularly in reciprocal sliding
was observed towards only one direction even though the motion was reciprocating. Subsurface cracks were observed at the plastically deformed surface layer, Fig. 16b. Hardness profiles for the examined cross-sections are shown in, Fig. 16c. In the case of the unidirectional sliding a significant increase in hardness was observed. The depth of higher hardness region was about 100 μm that was bigger than the plastically deformed depth, 40 μm and 18 μm for unidirectional and reciprocal contact, respectively, Fig. 16. The increase in hardness was related to the work hardening of the sur face material. Unexpectedly, near the worn surface a reduction in the surface hardness was observed for the reciprocating sliding contact.
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pressure was at the maximum in the center of the elliptical contact area between the disc and plate and it decreases as moving towards the contact circumferences, Fig. 5a. The numerical calculations confirmed that the selection of minimum and maximum normal loads of 100 and 500 N ensures elastic contact at lower loads, 100 N, and plastic defor mation at high normal loads, 500 N. It is important that sliding tests can be performed at the contact pressures similar to the contact pressure presented in the thread joints. In an earlier study [5] it has been observed that the contact pressures between the threaded parts operated in real application, in certain locations exceeded the yield point of the surface material. Therefore, sliding tests at elastic and plastic contact conditions were performed. Note that the contact pressure calculations were performed in accordance to the stress-strain curve of the core material. The carburized layer that at actual surface in contact has other mechanical properties than the core material. Thus, if the calculations were performed based on the mechanical properties of the carburized layer the results will be slightly different due to the higher hardness of this layer. The onset of plastic deformation could be delayed for the carburized layer compared to the core material. However, as it is seen in Fig. 5b there is no large deviation between Hertzian and FEM results even at the highest normal loads used here. Moreover, the initial contact, between the disc and the plate, calculated by FEM was around 0.8 mm when a normal load of 500 N was used. The calculated width of the contact was in a good agreement with the measured width during experimental tests that was close to 0.72 mm. The experimental contact width was measured when
Fig. 13. Worn depth as a function of the sliding distance and normal load, A is as-delivered surface, reciprocal sliding, B is as-delivered surface, unidirectional sliding, C is ground surface, reciprocal sliding, D is ground surface, unidirec tional sliding.
4. Discussion 4.1. Contact pressure and CoF It is well known that wear of contacting surfaces is related to contact pressure and other factors as sliding speed, sliding distance, surface texture and material, humidity and temperature. Thus it was found of importance to estimate the initial static contact pressure between the contacting bodies. To this purpose an analytical and a numerical methods were employed, Hertzian theory and FEM respectively. The results obtained from both methods were to some extent overlapping, especially, at low normal loads 20–300 N. The Hertzian theory assumes a frictionless contact between the contacting bodies and it does not consider plastic deformation [10]. While for the FEM calculations, a coefficient of friction equal to 0.2 was introduced and the contact was modeled as elastic-plastic with strain hardening according to the stress-strain diagram of the tested material. Therefore, in the present results when the plastic deformation took place at higher loads, devia tion between the analytical and numerical results was noticed, Fig. 5b. According to an earlier study, the onset of plasticity under compression load starts when the maximum contact pressure reaches 1.67 σY , where σY is the yield strength of the loaded material [10]. The obtained results by numerical method, FEM, revealed that the plastic deformation occurred at the loads equal to and above 400 N, as it is seen from a rise of PEEQ values as plotted against the corresponding normal loads in Fig. 5b. The numerical results moreover showed that the contact
Fig. 15. Worn volume as a function of the sliding distance and normal load, A is as-delivered surface, reciprocal sliding, B is as-delivered surface, unidirec tional sliding, C is ground surface, reciprocal sliding, D is ground surface, unidirectional sliding.
Fig. 14. 3D optical profilometer images of 6 � 6 mm2 specimens of the wear track of the as-delivered surface tested under reciprocal sliding at 500 N after sliding, a) 100 m, b) 300 m. 9
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Fig. 16. a)-b) deformed texture of the ground surface tested at 500 � 300 Nm, a) under unidirectional sliding b) under reciprocal sliding, c) the hardness profiles of the worn surface shown in a and b.
the disc at 500 N load was pressed against the plate and then removed from plate. As explained before, the estimation of initial contact pres sures is of importance to define whether the initial contact will be in elastic or in plastic region. However, during the sliding contact and due to the tribological effects that take place between the contacting surfaces the contact area is continuously changing and so the contact pressure. Therefore, it would be difficult to model the precise contact pressures acting during sliding contact. Even though the obtained results are derived based on the core material still it is useful for the present work in order to define the level of the contact pressure when the disc is pressed against the plate. Carburizing is a quite common surface treatment to generate a harder surface with high wear resistance [4]. As a result of the case hardening by carburizing, an oxide layer was produced on the surface of all heat treated specimens. SEM analysis revealed that the oxide was not only produced on the surface but also penetrated along the grain boundaries. The oxide embrittles the grain boundaries, and it can be detrimental in applications where a cyclic tensile or bending load is applied. Nevertheless, in sliding contact, the grain boundary oxides may have less significant impact on the wear performance as the sliding contact involves a progressive material removal from one or both con tacting surfaces. However, in general the overall influence of the oxide layer on the performance of a surface in sliding contact is not fully un derstood and it was found of interest to investigate its impact on surfaces exposed to sliding contact. Sliding tests were performed on specimens with as-delivered and ground surfaces. The as-delivered specimens had an oxide layer at the surface with a mean thickness of 15 μm, while the specimens with the ground surface had an oxide layer of about 5 μm thick. It was observed
that the friction values, over the first 150 mm of sliding contact, for the as-delivered surface were the same as for the ground surface values. This behavior can be associated with surface flattening and wear of asperities and allows us to conclude that the initial surface roughness had an insignificant effect at sliding distances longer than 150 mm. It was also observed that the running-in region was relatively short compared to the steady state region and the transition region where friction increased. Therefore, no remarkable distinction between friction values for asdelivered and ground surfaces was observed. As sliding distance increased, the surface roughness was not influ encing the friction level as the high contact pressure did. Further fric tional behavior was rather controlled by adhesive wear and formation of mixed oxide-metal debris. Nevertheless, the level of the friction for asdelivered and ground surfaces was close to each other and it is believed that the thickness of the oxide layer did not have significant impact on the friction characteristics. At the same time the observed CoF values in the present study are lower than what have been reported in other studies for steel in dry contact. In Ref. [11] the reported CoF of dry steel to steel contact was around 0.8–1.5. The lower values observed in the present study may be related to the presence of the oxide layer that formed a tribolayer and thus prevented metal-to-metal contact. The CoF values of all tests showed two distinct regions, running-in and steady-state regions typical in dry sliding contact [12]. In the pre sent study, in the running-in region, it is believed that the first contact between specimens was oxide-to-oxide contact. The running-in region was short for all the tests. Though, it was to some extent longer for the tests performed at low load, 100 N and particularly in the case of uni directional sliding. Under further sliding the oxide layer broke into wear debris and it resulted in an increase in CoF values. In the steady-state 10
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region, the contact was a mixture of oxide-to-metal and metal-to-metal contact and the CoF found to be high and stable, 0.75. For the tests performed at higher loads, 500 N, the running-in region was much shorter compared to the tests performed at low loads. The running-in region was followed by an immidate increase and a steadystate region. The CoF values in this region were lower, around 0.45. Sliding contact under higher loads resulted in higher contact pressures and higher shear stresses that deformed surface material to a higher degree compared to the tests performed at lower loads. Plastic defor mation also resulted in grain refinement and thereby work hardening of the surface material. These factors may contribute to the lower friction values observed for the tests performed at high loads. Similar behavior of friction, for steels tested under dry sliding contact, where the CoF values in the steady state decreased with an increase in normal loads has been observed in other studies [12–15]. Generally the CoF is the sum of two components, adhesion friction (μa) and deformation friction (μd). Adhesion friction is the force required to break junctions when the opposite asperities are cold weld together during surface contact [16]. Bowden and Tabor developed a theory of adhesion friction and stated that the μa is dependent on the material properties of the contacting surface, μa ¼ Hτ [17]. Where τ is the shear strength of the junction and is close to the maximum shear strength of the softer material. H is the hardness of the softer surface material. The qffiffi second friction component is equal to μd ¼ 0:6 Rh. Where h is the
layer, Fig. 11b. It was also observed that the oxide layer became denser and thinner when the plate was exposed to the unidirectional sliding contact. As the normal load increased, larger shear strains were accu mulated at and beneath the surface resulting in crack formation and delamination of the surface material, Fig. 11d. In the case of reciprocal sliding contact, a rougher morphology was found at the worn track. Due to the type of motion the oxide layer was sheared in both direction. At high normal loads, the oxide layer mechanically mixed with the surface material forming wear debris. Under further sliding the wear debris were compacted and flattened. As well, further sliding resulted in the surface delamination, Fig. 11c. The significant surface delamination observed on the worn surface of the discs was not found on the plates. For the plates, delamination was observed only in the case of reciprocal sliding motion. The reason for the remarkable surface delamination of the discs was that the same area was in contact with the plate until the test was stopped. Whereas for the plates each point at the worn surface only experience the sliding contact loads once when the disc slid along the plate surface. It means that the material at each point was resting for a specific time before the next sequence of sliding contact taking place. Therefore, surface cracking and delamination were the dominant damage features for the worn surface of the discs at higher loads. Essentially, the dominant wear mechanisms at high normal loads for plates were plastic deformation and surface smoothing accompanied by surface shearing and delamination of wear debris, Fig. 11c and d.. These wear mechanisms were observed in all wear tests regardless of the initial thickness of the oxide layer. Earlier studies have also confirmed the smoothening of wear track at high normal loads as the effect of plastic deformation [13,19]. Concerning the subsurface shear stress distribution and its influence on CoF, a study has proposed that when the CoF values are smaller than 0.25, the maximum shear stress will be in the subsurface. In the other hand, when the CoF values are higher than 0.25 the maximum shear stresses are at the surface [20]. Nevertheless, in the present study, as revealed by cross-section examinations, plastic deformation occurred at the surface and in the subsurface. What was notable is that the plastic deformation depth in unidirectional sliding was twice deeper than in the reciprocal sliding contact, Fig. 16a and b.. In the unidirectional sliding, the plastically deformed surface material consisted of two regions. The first regions corresponded to ultrafine grains and had a depth of about 10 μm. The second region, 30 μm in depth, corresponded to the plasti cally elongated texture along the sliding direction. Similar regions were observed for the reciprocal sliding contact. Contrary, the plastic flow depth along the sliding direction was shal lower than the corresponding depth in the unidirectional sliding contact. The depth and number of subsurface cracks were also higher in the reciprocal contact. The reason for more extensive subsurface cracking in the reciprocal contact could be the reversed plastic deformation of the surface material. In the unidirectional contact the surface was sheared in one direction and the material was responding to the loading by strengthening its texture through work hardening. In fact, finer grains and plastic deformation observed beneath the worn surface are two fundamental strengthening mechanisms for metals. In the present work, the measured hardness profiles of the worn surface tested under unidi rectional contact had higher hardness values at a depth up to 100 μm, Fig. 16c It confirmed that work hardening phenomenon occurred. However, in the reciprocal sliding a similar work hardening and increase in hardness was not observed. Instead, a lower surface hardness was obtained. The explanation for this unexpected behavior could be that the high density of the subsurface cracks affected the hardness values and these cracks, served as voids, decreased the hardness. According to an earlier study, more subsurface crack will result in high degree of surface delamination during further sliding [21]. Higher surface delamination then leads to higher worn volume because of the proportional rela tionship between surface delamination and wear rate. 3D surface examinations were performed to quantify the worn depth and the worn volume of the wear track as a function of the wear load.
depth that a circular body, an asperity, makes when it is pressed against another surface and R is the radius of the asperity. However, Bowden and Tabor stated that in the case of low surface roughness the defor mation component is relatively small compared to the adhesion component. Anyway, during the sliding contact and due to the tribo logical effects temperature may rise, surface hardness and shear strength may change, cracks may form. All these factors will for sure influence the friction level between the contacting surfaces. In the present study, it is clearly seen that CoF values decreased with increased normal load, Fig. 8. The decrease in friction at high normal loads could be related to smoothing of the contact area and work hardening. Moreover, the high plastic deformation invoked by the high contact pressure, have removed and/or flattened the contact asperities as was experimentally observed. 4.2. Wear characterization and quantification
The wear characterization of the worn track revealed different type of wear damages: mild to severe adhesive wear, abrasive wear and crack formation and surface delamination. Depending on the type of motion, initial surface condition and the wear load the different wear mecha nisms will take place. The type of motion influences strongly the surface texture and de termines how the surface material deformed, e.g. the direction of the deformed texture [18]. In the unidirectional sliding in the present work, the surface material was sheared in one direction. For each passage of the disc over the plate an increment of shear strain was accumulated at the surface, and also in the subsurface depending on where the maximum shear stresses are located. Whereas in the reciprocal sliding, the material flow was reversed due to the reciprocating motion of the disc. However, for the discs the worn morphology was quite similar despite the type of motion. At low loads, the oxide layer was removed from the surface and the metallic surface appeared, Fig. 10a and b. At high load, 500 N, where the maximum contact pressures exceed the yield strength of the surface material, the strain accumulation was high enough to cause surface, and subsurface cracks and surface delamina tion, Fig. 10c and d. In contrary, the worn morphology of the plates was dependent on the type of motion. At low normal loads where the initial contact pressure was around the elastic limit, surface cracks and surface delamination were observed at low loads. However, these damage mechanisms occurred in the oxide layer and resulted in the surface cracks of the oxide 11
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Both the worn depth and worn volume increased with increasing wear load, Figs. 13 and 15. Moreover, the width of the wear track extended from approximately 2 mm–3 mm when the total sliding distance increased from 100 m to 300 m under the same normal load, 500 N, Fig. 14. The width increase of the wear track was the result of a com bination of material removal and plastic deformation. Material removal was mostly due to adhesive wear and surface delamination. Accumu lation of plastic deformation during each load passage made the surface material to plough to the sides of the wear track. The worn depth and worn volume of the wear track became pro portional to the wear load regardless the type of sliding mode and sur face condition. The general trend of the worn volume observed in the present study does follow the Archard’s wear equation, to some extent, Fig. 15. However, this wear equation is limited to only one wear mechanism where the harder surface ploughs the softer. Commonly, the worn volume is linearly proportional to the normal load and sliding distance and inversely proportional to the surface hardness of the softer material [6]. Whilst in the present case several wear mechanisms such as adhesive wear, abrasive wear and plastic deformation were acting simultaneously and contributed to the total wear of the surface material. The measured worn volume values in the reciprocal sliding tests were higher than in the unidirectional sliding in every wear load, Fig. 15. During the forward sliding contact, material transfer preferred to accumulate at the back side of the contact area of the disk. As the disk slid back, the transferred material resulted in a certain increase in the friction values and also ploughed the plate material. This behavior was not presented in the case of unidirectional sliding since the disk slid only one direction. In addition, higher worn volume in the case of reciprocal sliding could be related to the loading condition and the strain accu mulation in the material during sliding contact. In the unidirectional sliding, the surface material was deformed in one direction and loading condition was repeated uniaxial loading that resulted in hardening of the surface material. Whereas in the reciprocal sliding contact, the loading condition was repeated reversed cyclic loading. It resulted in higher surface cracking and delamination of the surface material. Worn volume for the as-delivered and ground surface were compared to each other based on the type of motion. It was found that the dif ference in the worn volume between the reciprocal and unidirectional contact for the as-delivered surface increased with increasing wear load. While for the ground surface, the difference in worn volume between these two sliding modes seemed to be constant. The reason for the observed behavior could be due to the initial oxide layer presented at the as-delivered surface. This layer served as a protection of the surface material and, thereby, the wear rate was low for both the reciprocal and unidirectional contact. As the wear load increased a higher wear rate was obtained for the reciprocal contact. Indeed, increasing normal load increases worn volume that in turn will increase the wear rate, as it has been observed by other researchers [15,19]. In the present study, tests using the same contact parameters, normal load 500 N and sliding distance 300 m, in different sliding modes, reciprocal and unidirectional sliding, resulted in the different total worn volumes. The maximum worn depth and worn width were 50 μm and 2.77 mm and 37 μm and 2.95 in the case of reciprocal and unidirectional sliding, respectively. During the reciprocal sliding, it is believed that wear debris were formed and stayed at the interface resulting in three-body abrasive wear. These loose abrasive particles presumably resulted in higher wear rates. Thus, larger worn depth was observed in the reciprocal sliding than in the unidirectional sliding. However, at low wear loads a smaller worn depth was observed in the as-delivered conditions, and it points towards an initial protective effect of the thin oxide layer.
The influence of the type of motion, the reciprocal and unidirectional, and the oxide layer on the steel performance in sliding contact were studied. It was found that the type of motion had a much higher effect on the tribological performance of the steel in sliding contact than the other test parameters. Obtained results can be summarized as follow: � The thickness of the oxide layer did not influence much the tribo logical performance of the investigated steel. Instead, the type of motion had a much higher effect on the tribological performance. � The reciprocal sliding tests resulted in 25–35% higher worn volume than the unidirectional sliding tests under the same contact condi tions, normal load and sliding distance. � CoF decreased with increasing normal load and its values for the reciprocal sliding were slightly higher than for the unidirectional sliding. � At low load, 100 N, and short sliding distances, the oxide layer of the as-delivered specimens tended to cover the contact. Ground speci mens tested at the same conditions a worn surface typical for adhe sive wear, with some islands of oxides. � At higher normal loads 300–500, the deformed texture consisted of two regions. One region characterized with ultra-fine grains with a depth up to 10 μm and the other with elongated grains along sliding direction. The depth of the deformed texture was depended on the type of motion. Higher depth observed at the unidirectional sliding contact. � Crack formation and propagation was observed in both the unidi rectional and reciprocal sliding contact. Though more subsurface cracks observed at the reciprocal sliding contact. Acknowledgments The funding of the present work by the Swedish Knowledge Foun dation, project no 20150090, Epiroc Rock Drills AB, Ovako AB, Sandvik Mining and Rock technology AB is gratefully acknowledged. References [1] D. Zou, Theory and Technology of Rock Excavation for Civil Engineering, Springer, 2017. [2] G. Krauss, D. Matlock, Effects of strain hardening and fine structure on strength and toughness of tempered martensite in carbon steels, J. Phys. IV 5 (C8) (1995). C8C51-C8-60. [3] G. Krauss, Heat treated martensitic steels: microstructural systems for advanced manufacture, ISIJ Int. 35 (4) (1995) 349–359. [4] S. Lampman, Introduction to surface hardening of steel ASM Handbook, Heat Treating, ASM International, Metals Park, Ohio 4 (1991), 1991. [5] A. Mussa, P. Krakhmalev, J. Bergstr€ om, Failure analyses and wear mechanisms of rock drill rods, a case study, Eng. Fail. Anal. 102 (2019) 69–78. [6] I. Hutchings, P. Shipway, Tribology: Friction and Wear of Engineering Materials, Butterworth-Heinemann, 2017. [7] A. Gaard, P.V. Krakhmalev, J. Bergstrom, N. Hallback, Galling resistance and wear mechanisms - cold work tool materials sliding against carbon steel sheets, Tribol. Lett. 26 (1) (2007) 67–72. [8] A. Gaard, P. Krakhmalev, J. Bergstrom, Wear mechanisms in galling: cold work tool materials sliding against high-strength carbon steel sheets, Tribol. Lett. 33 (1) (2009) 45–53. [9] A. Gåård, P. Krakhmalev, J. Bergstr€ om, Wear mechanisms in deep drawing of carbon steel - correlation to laboratory testing, Tribotest 14 (1) (2008) 1–9. [10] K. Johnson, Contact Mechanics, Cambridge University Press, Cambridge, UK, 1985. [11] V. Angelini, I. Boromei, C. Martini, C. Scheuer, R. Cardoso, S. Brunatto, L. Ceschini, Dry sliding behavior (block-on-ring tests) of AISI 420 martensitic stainless steel, surface hardened by low temperature plasma-assisted carburizing, Tribol. Int. 103 (2016) 555–565. [12] P.J. Blau, Interpretations of the friction and wear break-in behavior of metals in sliding contact, Wear 71 (1) (1981) 29–43. [13] M. Ulutan, O.N. Celik, H. Gasan, U. Er, Effect of different surface treatment methods on the friction and wear behavior of AISI 4140 steel, J. Mater Sci Tech 26 (2010) 251–257. [14] M. Ruiz-Andres, A. Conde, J. De Damborenea, I. Garcia, Friction and wear behaviour of dual phase steels in discontinuous sliding contact conditions as a function of sliding speed and contact frequency, Tribol Int 90 (2015) 32–42. [15] R. Tyagi, S. Nath, S. Ray, Effect of martensite content on friction and oxidative wear behavior of 0.42 Pct carbon dual-phase steel, Metall. Mater. Trans. A 33 (2002) 3479–3488, 2002.
5. Conclusions The friction and wear characteristics of a high strength steel, 22NiCrMo12–F, commonly used as rock drill steel were investigated. 12
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[16] R. Gohar, H. Rahnejat, Fundamentals of Tribology, World Scientific Publishing Company, 2012. [17] F. Bowden, A. Moore, D. Tabor, The ploughing and adhesion of sliding metals, J. Appl. Phys. 14 (2) (1943) 80–91. [18] D.H. Persson, S. Jacobson, S. Hogmark, Antigalling and low friction properties of a laser processed Co-based material, J. Laser Appl. 15 (2003) 115–119, 2003.
[19] H. So, D. Yu, C. Chuang, Formation and wear mechanism of tribo-oxides and the regime of oxidational wear of steel, Wear 253 (9) (2002) 1004–1015. [20] A. Bower, K. Johnson, The influence of strain hardening on cumulative plastic deformation in rolling and sliding line contact, J. Mech. Phys. Solids 37 (4) (1989) 471–493. [21] N.P. Suh, The delamination theory of wear, Wear 25 (1) (1973) 111–124.
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