- Email: [email protected]
ABS blend
Tribology International 143 (2020) 106090
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Tribology International journal homepage: http://www.elsevier.com/locate/triboint
Torsional fretting wear experimental analysis of a R3 offshore steel against a PC/ABS blend T. Pandim a, T. Doca a, *, A.R. Figueiredo a, F.M. Andrade Pires b a b
ENM - Department of Mechanical Engineering, Faculty of Technology, University of Brasilia: Campus Darcy Ribeiro, 70910-900, Brasilia, DF, Brazil DEMEC - Department of Mechanical Engineering, Faculty of Engineering, University of Porto: Rua Dr. Roberto Frias, 4200-465, Porto, Portugal
A R T I C L E I N F O
A B S T R A C T
Keywords: Torsional fretting wear R3 offshore steel Polycarbonate Acrylonitrile butadiene styrene
The mooring lines of offshore vessels are composed of metallic chains coupled with polyester cables. In this setting, failure of a single chain link can result in severe financial losses, environmental disaster and deadly hazard to personnel. It is known that fretting wear due to torsional reciprocating movement is one of the main causes of early failure of a R3 offshore steel chain link. In this study, a possible option to increase the lifespan and reliability of mooring chains is presented. The torsional fretting wear behavior of a Polycarbonate/AcrylonitrileButadiene-Styrene (PC/ABS) 60:40 blend is compared to the reference configuration. Wear scars are investigated using confocal microscopy analysis. The PC/ABS, despite having a much weaker structure and higher wear rate, shows a considerably more predictable wear behavior. It is concluded that, a PC/ABS coating can lead to a significant increase in reliability of operation and lifespan of a mooring system.
1. Introduction Floating Production Storage and Off-loading (FPSO) vessels are often employed for the exploration of oil and gas. The mooring system, used to attach such vessels to the sea floor, is typically composed of 8–12 mooring lines. Each line has a metallic chain on the dry section and several polyester cables below the water line. The failure of even one of those lines can result in structural damage, loss of lives and environ mental issues. Those consequences range from production stoppage for mechanical repairs to severe oil spills into the oceans [1]. Failures of mooring chains may happen for a number of reasons and on different parts of the structure. The most common parts for failure are the chain links located in the vicinity of fairleads, connectors and wire ropes [2,3]. Moreover, besides the usual failure mechanisms of chains on such ap plications, it has been noted that other novel failure mechanisms play an important role, such as out-of-plane bending [4] and chain twisting [5]. The wear phenomena is typically found between chain links and can promote significant lifespan reduction. For instance, the shackle chains employed in fishing nets are submitted to planar and circular motions that often enable abrasive wear [6] while the loading conditions in general mooring systems [7,8] favors the development of fretting wear [9] when the relative motion exceeds 0.5∘ [10]. Fig. 1 shows a typical metallic section of a mooring system (chain links and D-type connectors)
where forces, rotations and displacements can be observed. The traction required to restrain a FPSO is significant. The number of lines, the dimensions of the chain links and the materials employed are based on the concept of a Minimum Breaking Load (MBL). Depending on the steel grade used [11], a standard 120D offshore chain link might be loaded with up to 70% of its MBL. At this loading condition, the normal contact forces produce plastic strains [12] that flattens the contact interface and inhibit some of the movements produced by the currents, waves and tides. Depending on the curvatures of the many components (chain links, bolts, connectors, kenters, squared pins, etc) several con tact configurations can be found on a mooring line. However, as the actual contact zone develops in a rather small region, these contact configurations can be fairly described as crossed-cylinders or as cylinder/sphere-to-flat configurations with an equivalent effective radius. Each individual components is able to display six degrees of freedom (z-axis rotation, y-axis rotation, x-axis rotation, z-direction displacement, y-direction displacement and x-direction displacement). Nevertheless, due to the high levels of z-axis load (traction) employed in an ultra deep mooring system, it can be stated that only the z-axis rotation (torsion) is able to produce a sufficient relative motion to enable wear [13] while the remaining degrees of freedom are considered con strained. Moreover, assuming the traction on the system is up to the standards, both an adhesion zone and a mixed-slip zone (small
* Corresponding author. E-mail address: [email protected] (T. Doca). https://doi.org/10.1016/j.triboint.2019.106090 Received 24 July 2019; Received in revised form 2 November 2019; Accepted 25 November 2019 Available online 30 November 2019 0301-679X/© 2019 Elsevier Ltd. All rights reserved.
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polymer composites, the alloy would show wear marks compatible with abrasive wear. The more general conclusion of his work was that the type of wear depends on the polymer’s (or composite’s) ability to form friction transfer layers (FTL) on the metallic surface. Therefore, an adequate choice for the PC/ABS ratio could provide a suitable coating for the chain links of a mooring system. The assessment of coatings is commonly carried-out using a tradi tional tribometer in unidirectional linear motion. Usually, the contact geometry is not considered relevant for the purposes of testing, since it is argued that the wear phenomenon only depends on the true area of contact, generated by micro asperities on the surfaces of the materials in contact, which is assumed to be proportional to the normal load. By that reasoning, regardless of the initial contact configuration, after the true contact area is established, the wear mechanism would evolve the same way it would with any other initial geometry. However, Warmuth et al. [23] has shown experimentally that fretting wear can, indeed, be geometry-dependent. They observed that for less-conforming contact configurations, the dominant wear mechanism results in damage by bulk material removal, while for more conforming contacts, damage caused by subsurface deformation and adhesive transfer is dominant. They argue that, despite the friction coefficient being independent from contact geometry, at least for fretting displacement amplitudes, the wear coefficient is indeed dependent on the initial configuration. This hap pens not because of the changes in contact pressure, but rather due to the debris flowing out or being retained within the contact zone, a condition commonly found in torsional wear problems. Despite the fact that torsional wear happens in many mechanical and biological systems, such as human hip, knee joints and controllable pitch propellers, studies regarding this topic are still lacking. In particular, the interaction between polymers and metals in such condition is rarely addressed. Therefore, given the clear importance of the topic and considering the lack of research on this particular area, an analysis of the torsional fretting wear behavior of the R3-Offshore steel and a Polycarbonate/Acrylonitrile-Butadiene-Styrene (PC/ABS) blend is pre sented. In order to encompass a possible geometrical effect of the cur vature of the contact interface while still allowing for an microscopy analysis, a sphere-to-flat contact configuration has been chosen. The goal of this study is to identify surface properties (i.e. hardness and wear rate), wear morphology and quantify a possible life extension of a chain link due to a PC/ABS coating.
Fig. 1. Illustration of the general loading system found in a mooring line: (a) Forces due to weights, currents, waves and buoyancy; (b) Displacements and rotations.
amplitude reciprocate angular motion) are to be formed. Therefore, a torsional fretting wear condition is attained. Most alternatives proposed to mitigate wear on chain links are concerned with increasing the strength of the steel employed in the mooring lines. However, elastic chains have been suggested as an alternative to the metallic ones. Such chains would be manufactured by burying steel chain links within a rubber column, so that the elastic properties of the rubber would reduce wear and the impulsive tension acting on the chain. Simulations with this design have shown up to a 50% reduction in the impulsive tension [14]. Mooring line systems with advanced dynamic behaviors were also considered [15]. Nevertheless, a cheaper and simpler solution might be to coat the contact zone of critical chain links with a soft yet resistant material. Coatings have distinct surface properties that enhance the mechan ical performance of its substrate. They are often hard, resilient, thermalprotective and non-conductive. For instance, Polycarbonate/ Acrylonitrile-Butadiene-Styrene (PC/ABS) is a polymer blend widely used in the automotive and consumer electronics industries due to its availability, good impact resistance, high strength modulus and high thermal stability [16–18]. Its popularity can be explained by the nearly additive nature of the blend’s properties with respect to its components [19,20]. A blend with a higher concentration of PC with respect to ABS has the tendency to have higher hardness and overall mechanical strength. On the other hand, a blend with a higher percentage of ABS should be more resilient, impact resistant and easier to machine. A study on the friction and wear behavior of 18 polymers fretted against steel in dry air and humid environments can be found in Ref. [21]. It was observed that the addition of fillers in a polymer matrix has different effects on the wear rate depending on the environmental conditions. Zaitsev [22] investigated the contact of a cobalt and tungsten carbide alloy against a number of polymers and their composites. He concluded that the predominant wear mechanism depended on the chemical composition of the polymer. More specifically, he discovered that for polymers containing active oxygen groups, the most relevant wear mechanism was based on oxidation-fatigue. When in contact with
2. Materials and methods Two material configurations are employed in this study: i) R3 offshore-grade steel versus R3 offshore-grade steel (reference condi tion); and ii) R3 offshore-grade steel versus a PC/ABS 60:40 blend. Most of the properties of the R3 steel and PC/ABS blend were provided by their manufacturers and are listed in Table 1. The only exception is the Poisson’s ratio of the 60:40, which was retrieved from Ref. [24]. Representative samples of the hemispherical pad and flat counter part are depicted in Fig. 2. The pads, made from R3 steel, have a cylindrical body and a hemi spherical tip with a 7.5 mm radius. The flat counterparts, made from R3 steel and PC/ABS, are prismatic shaped (20 mm � 20 mm x 5 mm). The R3 samples were manufactured in a CNC machine (tolerance of �0.1 mm) using material removed from a standard mooring chain link. The PC/ABS samples were produced by SABIC using injection technique. Table 1 Mechanical properties of the R3 offshore grade steel and the PC/ABS 60:40 blend.
2
Material
R3 offshore steel
PC/ABS 60:40
Young modulus, E(GPa) Poisson’s coefficient, ν Yield stress, σy (MPa)
207 0.34 410
2.2 0.35 54
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Fig. 2. Depiction of representative specimens of each test configuration: (a) R3 steel hemispherical pad and R3 steel flat counterpart; (b) R3 steel hemispherical pad and PC/ABS 60:40 flat counterpart.
After production, the flat counterparts are polished in a rotational wet sanding bench at 300 RPM using a procedural method with six grades of sandpaper: 120, 240, 400, 600, 800 and 1200 grid. Afterward, the samples were controlled for roughness (Ramin ¼ 0:4) and dimensional distortions (angle ¼ 1∘). All specimens were carefully cleaned before testing using isopropyl alcohol to avoid unwanted particles (e.g. con taminants, small polishing debris and oxidation) on the contact surfaces. Previous to the wear tests, the hardness of the counterpart samples were analyzed. Two indentors/techniques in two different testers were employed to evaluate the material hardness and check for discrepancies. For the analysis of the micro-hardness, a Vickers micro-indentor (HM2000 Fischer Technology) with a load of 0.1 N was employed. This first method is able to capture hardness variations on the top layer and indicate a possible non-homogeneous surface. The evaluation of the standard hardness was performed in an Universal hardness tester (ZHU 250). The chosen indentation method was the HBW2.5/300 (Brinell test with a 2.5 mm tungsten carbide sphere and a load of 300 N), this indentor has been picked not only for its shape configuration (similar to the one used in the wear test) but also for having a wide scale for evaluation of mean combined hardness of both surface and substrate. Three measurements in three different flat counterparts of each material were carried out. A MTS 809 Axial-Torsional Testing Machine (Fig. 3a) and a sample holder (Fig. 3b) device were used to run the wear tests. The experimental procedure can be summarized as follows. A flat sample is fit inside the holder which is closed with a bolted lid. The holder is then placed and held by the lower grip while the hemispherical pad is held by the upper grip. The pad’s lateral surface (red zone) is fully restrained while each side facet (such as the one in orange) and bottom of the flat sample has a normal displacement restraint, see Fig. 4a. The schematic view of the loading conditions is depicted in Fig. 4b. The crosshead is positioned at a 1 mm gap distance from the flat sample. When the test is stated, the lower grip is programmed to move the flat sample towards the pad until a preselected normal force, Fz , is reached. The normal forces selected are: 1.0 and 2.0 kN. These values were chosen such that two distinct levels of plastic strains, similar to the ones observed in real scenarios, would be generated. As soon as the desired contact force is established, the normal load is maintained and a prescribed rotation (θz ¼ �1∘ at 5 Hz) is applied to the pad during a defined number of cycles (0, 50, 100 and 200 thousand cycles). Each test condition (two material pairings, two loads and four number of cycles) is
Fig. 3. Experimental setup for the sphere-on-flat torsional fretting wear tests. (a) Equipment: 1) Crosshead; 2) Force transducer; 3) Top cell; 4) Upper grip; 5) Lower grip; 6) Bottom cell. Detailed view: 7) Hemispherical pad; 8) Flat sample; 9) Holder.
repeated once for standard deviation assessment. A total of 32 torsional fretting wear tests were performed. After each test, the sample holder device is detached from the grips and disassembled. Then, the specimen is visually inspected for debris and cleaned using a compressed dry air pistol. In the following, a confocal laser microscope (OLYMPUS LEXT OLS4100) is employed to produce profiles of the wear marks (see Fig. 5). The height and lateral resolutions are 10 nm and 12 μm, respectively. A stitching procedure is 3
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Fig. 4. Schematic views of the contact configuration, restraints and loadings: (a) Encastre on the pad (red area) and displacement constraints on the flat sample (orange area); (b) normal force and z-axis rotation. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)
performed in both laser and color spectrum using 1 μm steps and a 20x magnification lens. Due to its darker aspect, a brightness of approxi mately 60% is required for the analysis of PC/ABS samples. For the R3 steel samples, brightness levels ranging from 20%(color) to 40%(laser) were employed. The wear volume is evaluated using two techniques. The first volume estimation method consists of a geometrical approach, employed only for the PC/ABS flat samples (which showed regular, symmetrical and hemispherical shaped scars). Having measured the scar’s depth, h, and diameter, d, (see Fig. 5a), the worn volume V was estimated using the definition of a spherical cap geometry, as follows, �� �2 � 1 d V ¼ ⋅ π ⋅ h⋅ þ h2 : (1) 6 2 Herein, the measurements of depth and diameter were performed in both XZ and YZ planes to evaluate a possible eccentricity of the contact zone. Similar consolidated procedures, regarding micro-scale abrasion and wear tests, can be found in Refs. [25,26]. For the analysis of the R3 steel specimens, which showed irregular scars (see the 2D profile in Fig. 5b), the companion software of the confocal microscope was used. This second method requires the definition of a reference surface and an interest zone, depicted in Fig. 6. Herein, these zones are defined as the unscathed surface (outside the wear region) and a polygon surrounding the outer-border of the wear region, respectively. Then, using a computer-aided profile analysis, the volume (removed or deposited) is measured.
Fig. 5. Representative 3D (top) and 2D (bottom) profiles of flat samples for morphology analysis and quantification of wear: (a) PC/ABS; (b) R3 steel.
pairing solids with such distinct levels of hardness in a wear condition, it is expected that the soft material will endure the majority of the damage. The highest standard deviation observed during the measurements was equal to 11%, meaning that the hardness of both materials remains fairly constant throughout the extension of their surfaces. Moreover, the dif ference between the results provided by the micro-indentation and the standard hardness test were small. For instance, the R3 steel shows a micro-hardness only 7% higher than the micro-hardness while the dif ference for the PC/ABS is 14%. This variation is likely related to a size effect phenomenon and a small density difference between top layer and substrate. Nevertheless, as this variation was rather low, it indicates that the solids are mostly homogeneous over depth. Therefore, the particles produced over the wear cycles should have a fairly constant hardness.
3. Results and discussion In this section, three aspects are discussed: i) hardness; ii) scar morphology; iii) wear analysis; and iv) life estimations. 3.1. Hardness The hardness test results, listed in Table 2, revealed the R3 steel as a �77% harder surface when compared to the PC/ABS 60:40. When 4
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preserved, see Fig’s. 7d-7f. Similar results were obtained by Wang [27], whose research on torsional fretting wear of 1045 steel against a monomer casting nylon filled with glass fiber resulted in analogous profiles: a central region with almost no indentation and an annular area with severe wear grooves. Nevertheless, in all tests in the reference configuration, erosion marks produced by a chipping effect are observed inside the adhesion zone. This pattern is evident in both sides of the interface and becomes more prominent at 2.0 kN (see Fig.s 7j-7l) since under this condition the adhesion zone is 18–22% bigger. In most cases, chipping marks with this aspect are created by the smallest debris pro duced in the mixed-slip zone which, due to the angular movement of the pad, are brought into the adhesion and crushed to even smaller particles. Three trends can be noticed when analyzing these marks: i. as the number of cycles increases the chipping marks become more slender; ii. as one moves away from the first point of contact (center of the adhesion zone) the chipping marks get larger; and iii. the depth of these chipping marks remained fairly constant as the number of cycles increased. These three features corroborate the assumption that the debris decrease in size as the number of cycles increase. Moreover, a dark red powder (typical of oxidised debris) was found inside the adhesion zone. Since the chipping marks were rather mild it is also reasonable to assume that their hardness was not significantly altered by the oxidation process, rendering the debris almost harmless. It is important to emphasize that, if any oxide particles were present in the wear mark after test, they were completely removed during the cleaning process (compressed dry air) as no trace of oxidation (oxide debris/oxidation spots) was found on the samples during the microscopy analysis. On the other, the partial-slip zone shows severe damage. The scars contain protrusions, grooves and lumps unevenly spread indicating an intricate exchange of material. Although the overall shape of the scars is the expected circular pattern, in some cases a rather elliptical mark is found (see Fig’s. 7b and 7e). The exact opposite behavior is found in the PC/ABS marks, as can be seen in Fig. 8. Overall, the wear marks observed in the PC/ABS samples have a smooth bowl-shape (crater). Therefore, a continuous transition from the adhesion zone to the mixed-slip zone is observed. Although a few dispersed dragging scars (marked by black arrows) are found, the damage promoted by the debris is mostly small scratches. Nevertheless, cracks and micro-fractures are found in the corners of all samples (as indicated by the white circles such as the one in Fig. 8a). We believe that these scratch marks were formed at the early stage of the wear process, when the debris produced in the mixed-slip zone would damage the borders of the contact zone. This condition is only evident in the PC/ABS samples, as the material is rather softer and weaker when compared to the R3 steel. As the load increases, the central adhesion area, which used to pre sent several scratches, starts to show a more smooth pattern, see Fig’s 8b-8e. This behavior is notable at the 200k cycles condition depicted in Fig. 8f. Another peculiar feature is the polishing effect created at the borders of the contact zone (denoted by the black circles). These ob servations, once again, indicate a decrease of the debris over load and over number of cycles as observed in the reference condition.
Fig. 6. Computer-aided volume measurement of the material removed (dark yellow: top view) and deposited (yellow: top view and pink: 2D view). (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.) Table 2 Hardness of the flat counterpart batch. HV0.1
HBW 25/300
R3 steel
PC/ABS 60:40
198�5
122�14
199�2
120�8
198�4
115�3
183�7
103�2
188�12
105�3
186�2
104�2
3.2. Scar morphology The images of the scars provide a great level of detail and revealed interesting aspects of the morphology of the worn surfaces. Three important features of the configurations studied are observed right after cleaning the samples: 1. For compression test (i.e. zero cycles) a regular smooth hemispher ical mark is obtained, which is barely visible due to the low inden tation depth produced. Therefore, these results are only detailed in subsection 3.3; 2. In all tests against the PC/ABS flat counterparts and all compression tests the hemispherical pad remained intact; 3. The fretting wear tests in the reference configuration lead to signif icant damage in both pad and flat counterpart.
3.3. Wear analysis As previously mentioned, the analysis of the 3D profiles reveals that material transfer occurred during the tests in the reference configura tion. In some cases, one side of the interface is basically unscathed and only material deposition occurred, as seen in Fig. 9. However, in most cases a complex and nonlinear exchange of material is observed. Therefore, in order to quantify the net wear volume of the fretting wear tests, each side of the interface is evaluated independently and the measurements are taken with respect to a reference surface (e.g. a plane for the flat counterpart or a spherical cap for the hemispherical pad). Results for the volume above the reference surface, Vþ , void below the reference surface, V , and net wear volume of the interface, Vw , are
Representative top views of the profiles at reference condition wear are shown in Fig. 7, we warn the reader that (due to changes in shape and overall dimensional of the wear marks) the scales used were not kept constant. The wear marks on R3 steel samples show, as expected, two distinct zones: an adhesion zone located at the center of the contact region; and a thick outer annular region representing the partial-slip zone. The former shows a shallow indentation and minute signs of damage while the latter presents severe fretting wear scars. On one hand, results with the reference configuration at 1 kN reveal an adhesion zone (inner region of the white dashed circle) well 5
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Fig. 7. Top view of the R3 steel samples.
6
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Fig. 8. Top view of PC/ABS 60:40 flat counterpart samples.
listed in Table 3. As already stated, the pads used in the compression tests did not presented any sign of damage. Moreover, despite the fact that wear marks are found on the pads used in the 50k cycles tests none of them actually lost material. Furthermore, at 50k cycles the material removed from the flat counterpart can be found deposited on itself and at the center of the pad’s tip. At 100k and 200k cycles both surfaces presented wear damage and a considerable amount of material deposited on their surfaces. Although it is not trivial to discern the origin of the material deposited in each side, it is reasonable to assume that they were removed from their respective surfaces. A graphical representation of the reference configuration net wear volume over number of cycles is provided in Fig. 10. An important observation is that the initial mark created by the compression phase of the tests with 2 kN is 6x bigger in volume than ones produced in the tests with 1 kN. Moreover, the running-in phase is observed and it is likely to have lasted less than 50k cycles. Following [28], this non-linear phase of the tests will be disregarded for a coherent analysis of the wear coefficient after its stabilization. The wear rate observed on the 2 kN tests (1.73E-4mm3/kilocycle) is approximately 2x the one observed in the 1 kN tests (8.00E-5mm3/kilocycle). The high R-squared values obtained for the linear regressions (0.83 for 1 kN and 0.90 for 2 kN) indicate that, despite the complex material exchange observed in the R3 steel pairing, a predictive behavior (from 50k to 200k cycles) is achieved. The measurements of the wear craters produced on the PC/ABS 60:40 flat counterparts are listed in Table 4. The wear volumes observed in this second configuration reached values two orders of magnitude larger than the reference configuration. This behavior was expected as a polymer has a rather softer and weaker structure when compared to a metal, which is evident in the graphical representation of the data dis played in Fig. 11.
The wear rate observed (during the stable phase) was equal to 2.22E2mm3/kilocycle for the 2 kN tests and 1.74E-3mm3/kilocycle for the 1 kN tests, a difference of over 10 times. Therefore, since the R3 steel shown a �2x increase on its wear rate at the same loads, one could state that the PC/ABS is 5x more susceptible to load variations. However, results also show two paramount advantages: a smoother transition from the running-in phase to the stable phase and an even better linear fitting. Although in this analysis it is also assumed that the stable wear phase of the PC/ABS samples started after 50k cycles, we believe that the required number of cycles for achieving this condition might be lower. The R-squared values of 0.86 and 0.91, for 1 kN tests and for 2 kN tests respectively, indicate a reliable linear trend. Furthermore, as no damage is induced on the pads during the PC/ABS tests, this same predictable behavior is observed on the evolution of wear depth of the flat coun terparts, see Fig. 12. 3.4. Life estimations In the guidelines for failure analysis of mooring components [29] it is stated that the dimension of surface defects must be inferior to 5% of the component’s nominal diameter. For instance, hemispherical defects on standard offshore chain links (nominal diameter of 120 mm) must have less than 6 mm in diameter or a perfect hemispherical scar with a radius of 3 mm and a 56 mm3 in volume. In this setting, the wear marks observed at 200k cycles (0.04 mm3 for 1 kN and 0.09 mm3 for 2 kN) would represent 0.07% and 0.16% of the tolerance for surface defects in mooring chains. Considering the most optimistic scenario observed in this study, i.e. a one-sided linear evolution of the wear volume at 1 kN load, at approximately 285.7 M cycles a standard chain link would be flag for replacement. However, given the complex material exchange and non-linear behavior observed in this study, such high lives are never attained. In order to address this issue a rather complicated and costly 7
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Fig. 10. Net wear volume versus number of cycles: reference configuration. Table 4 Volume estimations: PC/ABS 60:40 flat counterpart. Condition
dx [mm]
dy [mm]
hx [mm]
hy [mm]
V[mm3]
1 kN, 0 cycles
2.879 3.186 3.869 4.053 4.159 4.058 4.229 4.307 4.480 4.192 5.579 5.773 6.252 6.194 6.785 6.149
3.102 3.207 3.978 3.996 4.107 4.039 4.271 4.299 4.271 4.335 5.625 5.809 6.297 6.244 6.874 6.242
0.013 0.013 0.053 0.055 0.059 0.056 0.073 0.084 0.052 0.031 0.149 0.203 0.261 0.222 0.345 0.309
0.011 0.016 0.048 0.050 0.053 0.051 0.076 0.085 0.054 0.033 0.142 0.202 0.252 0.221 0.361 0.322
0.042 0.059 0.307 0.333 0.379 0.347 0.529 0.615 0.404 0.229 1.799 2.670 3.970 3.370 6.485 4.773
1 kN, 50k cycles 1 kN, 100k cycles
Fig. 9. Height profiles of the wear marks on R3 steel samples: (a) Hemi spherical pad with severe deterioration at the borders of the contact zone; (b) Flat counterpart with unscathed surface and material deposition.
1 kN, 200k cycles 2 kN, 0 cycles 2 kN, 50k cycles
Table 3 Volume measurements: reference configuration. Flat
2 kN, 100k cycles
Pad
Interface
Condition
Vþ [mm3]
V [mm3]
Vþ [mm3]
V [mm3]
1 kN, 0 cycles
– – 0.020 0.025 0.031 0.042 0.035 0.030 – – 0.036 0.023 0.047 0.061 0.077 0.062
– –
– – 0.020 0.026 0.018 0.036 0.029 0.025 – – 0.071 0.049 0.082 0.055 0.032 0.066
– – 0.000 0.000 0.000 0.028 0.014 0.005 – – 0.000 0.000 0.040 0.037 0.002 0.011
1 kN, 50k cycles 1 kN, 100k cycles 1 kN, 200k cycles 2 kN, 0 cycles 2 kN, 50k cycles 2 kN, 100k cycles 2 kN, 200k cycles a b
– –
0.012 0.019 0.012 0.011 0.008 0.007 0.046 0.013 0.013 0.012 0.017 0.034
2 kN, 200k cycles
Vw [mm3] 0.003 0.003 0.028 0.032 0.037 0.039 0.042 0.043 0.018 0.018 0.061 0.059 0.076 0.067 0.090 0.083
(a) (a)
more than 4.6 M cycles would be required. The simplicity of this assessment is a decisive feature as it greatly facilitates maintenance planing and the prediction of safe limits to avoid unexpected failures. 4. Conclusions
(b) (b)
This study presents a sphere-to-flat torsional fretting wear analysis of two material pairings: a R3 steel pad against a R3 steel counterpart; and a R3 steel pad against a PC/ABS 60:40 counterpart. Analysis of hardness, wear morphology and wear rate are provided. The main conclusions drawn from the observations made in this study can be summarized as follows:
Spherical indentation volume for the 1 kN load. Spherical indentation volume for the 2 kN load.
� In both configurations addressed, a direct relation between wear volume, applied load and number of cycles is observed. After a stable condition is reached (50k cycles), this relation is linear in both configurations; � R3 steel shows higher hardness and lower wear rate when compared to the PC/ABS 60:40. However, the former shows a rather unpre dictable behavior as complex material exchange are observed. This condition becomes even more evident as the load is increased.
maintenance technical procedure is employed. A much more reliable scenario is observed in the proposed solution. The linear relation between wear depth and number of cycles (observed in Fig. 12) enables an easy evaluation of the lifespan of a PC/ABS 60:40 coating. For instance, at 1 kN, over 25.4 M cycles would be needed to completely remove a 5 mm thick protective layer, where as for 2 kN, 8
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� The torsional fretting wear test have shown that the PC/ABS 60:40 flat counterparts are able to protect the interface for a substantial number of cycles while displaying a predictable wear behavior, which is essential for a coating material; � Finally, the linear relation between wear depth and number of cycles on the polymer tested enables a straightforward estimation of wear life of a mooring component. The findings of this study are promising. However, more research is needed. Analysis considering different blend ratios and types of poly mers will follow. Acknowledgments This work was financially supported by the Brazilian National Council for Scientific and Technological Development (CNPq) under the project [406724/2016-4], by the Foundation for the Support of Research (FAPDF) under the project [0193.001521/2016] and by the Coordenaç~ ao de Aperfeiçoamento de Pessoal de Nível Superior (CAPES) [Finance Code 001]. The authors gratefully acknowledge this support. Our thanks are also extended to Eng. Marcelo Fonseca dos Santos and Petrobras/CENPES for the offshore components supplied. References
Fig. 11. Net wear volume versus number of cycles: PC/ABS 60:40 flat counterpart.
[1] Gordon R, Brown M, Allen E. Mooring integrity management: a state-of-the-art review. In: OTC 25134-MS. Offshore Technology conference; 2014. p. 1–19. [2] Bjørnsen E. Chains in mooring systems. Master’s thesis. Norwegian University of Science and Technology; 2014. [3] Yaghin A, Melchers R. Long-term inter-link wear of model mooring chains. Mar Struct 2015;44:61–84. [4] Jean P, Goessens K, L’Hostis D. Failure of chains by bending on deepwater mooring systems. In: OTC-17238-MS. Offshore Technology conference; 2005. p. 1–12. [5] Ma K. Historical review on integrity issues of permanent mooring systems. In: OTC24025-MS. Offshore Technology conference; 2013. p. 1–13. [6] De Pauw J, De Baets P, Delgado YP, Sukumaran J, Ost W. A full scale test rig for assessment of abrasive wear of shackle chains. Wear 2013;302:1017–25. [7] Faltinsen OM. Sea loads on ships and offshore structures. 1 ed. Cambridge University Press; 1990. [8] Guide for position mooring systems. American Bureau of Shipping; 2018. [9] Waterhouse RB. Fretting wear. Wear 1984;100:107–18. [10] Duggal A, Fontenot W. Anchor leg system integrity – from design through service life. In: OTC-21012-MS. Offshore Technology conference; 2010. p. 1–15. [11] Offshore standard DNV-OS-E302: offshore mooring chain. Det Norske Veritas; 2008. [12] Zegatti M, Doca T. 2D fretting wear analysis of mooring connectors under plastic strain. RIPE 2016;2. [13] Ridge I, Hobbs R, Fernandez J. Predicting the torsional response of large mooring chains. In: OTC-17789-MS. Offshore Technology conference; 2006. p. 1–13. [14] Takayama T, Tokubuchi K. Economical design of elastic chain for mooring through the numerical simulation. International Society of Offshore and Polar Engineers; 2002. [15] Brown DT, Mavrakos S. Comparative study on mooring line dynamic loading. Mar Struct 1999;12:131–51. [16] Hund J, Naumann J, Seelig T. An experimental and constitutive modeling study on the large strain deformation and fracture behavior of PC/ABS blends. Mech Mater 2018;124:132–42. [17] Ishikawa M. Stability of plastic deformation and toughness of polycarbonate blended with poly(acrylonitriie-butadiene-styrene) copolymer. Polymer 1994;36 (11):2203–10. [18] Win KN, Liu E. Thermal, mechanical and tribological properties of polycarbonate/ acrylonitrile-butadiene-styrene blends. J Polym Eng 2013;33:535–43. [19] Krache R, Debbah I. Some mechanical and thermal properties of PC/ABS blends. Mater Sci Appl 2011;2:404–10. [20] Utracki LA. Commercial polymer blends. 2 ed. Springer; 2013. [21] Mens JWM, de Gee AWJ. Friction and wear behaviour of 18 polymers in contact with steel in environments of air and water. Wear 1991;149:255–68. [22] Zaitscv AL. Mechanisms of hard alloy wear in frictional processes with polymers and composite materials. Wear 1993;162:40–6. [23] Warmuth AR, Pearson SR, Shipway PH, Sun W. The effect of contact geometry on fretting wear rates and mechanisms for a high strength steel. Wear 2013;301: 491–500. [24] Grellmann W, Seidler S. Landolt-b€ ornstein numerical data and functional relationships in science and Technology, 6A. Springer-Verlag; 2014. [25] Gee M, Gant A, Hutchings I, Kusano Y, Schiffman K, Van Acker K, et al. Results from an interlaboratory exercise to validate the micro-scale abrasion test. Wear 2005;259:27–35.
Fig. 12. Wear depth versus number of cycles: PC/ABS 60:40 flat counterpart.
Moreover, the wear volume of the reference condition required a more costly and time consuming evaluation method; � Both configurations have shown scars with two distinct zones: a central circular adhesion zone and an annular mixed-slip zone. The size of these zones were proportional to the applied load; � In the reference configuration, the majority of the damage is found in the mixed-slip zone whereas in the PC/ABS 60:40 flat counterpart the damage is more evenly distributed; � Interesting scratching marks were found in both configurations and they evolved as the number of cycles increased. A dark red powder (oxidised debris) and a black powder (polished particles) were found on the wear scars of the reference configuration and the PC/ABS 60:40 flat counterpart, respectively; 9
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[26] Steier V, Pires M, Doca T. The influence of diamond-like carbon and anodised aluminium oxide coatings on the surface properties of the SAE 305 aluminium alloy. J Braz Soc Mech Sci Eng 2018;40:41. [27] Wang S. Torsional wear behavior of monomer casting nylon composites reinforced with GF: effect of content of glass fiber. Tribol Trans 2013;56(2):178–86.
[28] Cruzado A, Hartelt M, W€ asche R, Urchegui M, G� omez X. Fretting wear of thin steel wires. part 1: influence of contact pressure. Wear 2010;11–12:1409–16. [29] Guide for the certification of offshore mooring chain. American Bureau of Shipping; 2017.
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Contents lists available at ScienceDirect
Tribology International journal homepage: http://www.elsevier.com/locate/triboint
Torsional fretting wear experimental analysis of a R3 offshore steel against a PC/ABS blend T. Pandim a, T. Doca a, *, A.R. Figueiredo a, F.M. Andrade Pires b a b
ENM - Department of Mechanical Engineering, Faculty of Technology, University of Brasilia: Campus Darcy Ribeiro, 70910-900, Brasilia, DF, Brazil DEMEC - Department of Mechanical Engineering, Faculty of Engineering, University of Porto: Rua Dr. Roberto Frias, 4200-465, Porto, Portugal
A R T I C L E I N F O
A B S T R A C T
Keywords: Torsional fretting wear R3 offshore steel Polycarbonate Acrylonitrile butadiene styrene
The mooring lines of offshore vessels are composed of metallic chains coupled with polyester cables. In this setting, failure of a single chain link can result in severe financial losses, environmental disaster and deadly hazard to personnel. It is known that fretting wear due to torsional reciprocating movement is one of the main causes of early failure of a R3 offshore steel chain link. In this study, a possible option to increase the lifespan and reliability of mooring chains is presented. The torsional fretting wear behavior of a Polycarbonate/AcrylonitrileButadiene-Styrene (PC/ABS) 60:40 blend is compared to the reference configuration. Wear scars are investigated using confocal microscopy analysis. The PC/ABS, despite having a much weaker structure and higher wear rate, shows a considerably more predictable wear behavior. It is concluded that, a PC/ABS coating can lead to a significant increase in reliability of operation and lifespan of a mooring system.
1. Introduction Floating Production Storage and Off-loading (FPSO) vessels are often employed for the exploration of oil and gas. The mooring system, used to attach such vessels to the sea floor, is typically composed of 8–12 mooring lines. Each line has a metallic chain on the dry section and several polyester cables below the water line. The failure of even one of those lines can result in structural damage, loss of lives and environ mental issues. Those consequences range from production stoppage for mechanical repairs to severe oil spills into the oceans [1]. Failures of mooring chains may happen for a number of reasons and on different parts of the structure. The most common parts for failure are the chain links located in the vicinity of fairleads, connectors and wire ropes [2,3]. Moreover, besides the usual failure mechanisms of chains on such ap plications, it has been noted that other novel failure mechanisms play an important role, such as out-of-plane bending [4] and chain twisting [5]. The wear phenomena is typically found between chain links and can promote significant lifespan reduction. For instance, the shackle chains employed in fishing nets are submitted to planar and circular motions that often enable abrasive wear [6] while the loading conditions in general mooring systems [7,8] favors the development of fretting wear [9] when the relative motion exceeds 0.5∘ [10]. Fig. 1 shows a typical metallic section of a mooring system (chain links and D-type connectors)
where forces, rotations and displacements can be observed. The traction required to restrain a FPSO is significant. The number of lines, the dimensions of the chain links and the materials employed are based on the concept of a Minimum Breaking Load (MBL). Depending on the steel grade used [11], a standard 120D offshore chain link might be loaded with up to 70% of its MBL. At this loading condition, the normal contact forces produce plastic strains [12] that flattens the contact interface and inhibit some of the movements produced by the currents, waves and tides. Depending on the curvatures of the many components (chain links, bolts, connectors, kenters, squared pins, etc) several con tact configurations can be found on a mooring line. However, as the actual contact zone develops in a rather small region, these contact configurations can be fairly described as crossed-cylinders or as cylinder/sphere-to-flat configurations with an equivalent effective radius. Each individual components is able to display six degrees of freedom (z-axis rotation, y-axis rotation, x-axis rotation, z-direction displacement, y-direction displacement and x-direction displacement). Nevertheless, due to the high levels of z-axis load (traction) employed in an ultra deep mooring system, it can be stated that only the z-axis rotation (torsion) is able to produce a sufficient relative motion to enable wear [13] while the remaining degrees of freedom are considered con strained. Moreover, assuming the traction on the system is up to the standards, both an adhesion zone and a mixed-slip zone (small
* Corresponding author. E-mail address: [email protected] (T. Doca). https://doi.org/10.1016/j.triboint.2019.106090 Received 24 July 2019; Received in revised form 2 November 2019; Accepted 25 November 2019 Available online 30 November 2019 0301-679X/© 2019 Elsevier Ltd. All rights reserved.
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polymer composites, the alloy would show wear marks compatible with abrasive wear. The more general conclusion of his work was that the type of wear depends on the polymer’s (or composite’s) ability to form friction transfer layers (FTL) on the metallic surface. Therefore, an adequate choice for the PC/ABS ratio could provide a suitable coating for the chain links of a mooring system. The assessment of coatings is commonly carried-out using a tradi tional tribometer in unidirectional linear motion. Usually, the contact geometry is not considered relevant for the purposes of testing, since it is argued that the wear phenomenon only depends on the true area of contact, generated by micro asperities on the surfaces of the materials in contact, which is assumed to be proportional to the normal load. By that reasoning, regardless of the initial contact configuration, after the true contact area is established, the wear mechanism would evolve the same way it would with any other initial geometry. However, Warmuth et al. [23] has shown experimentally that fretting wear can, indeed, be geometry-dependent. They observed that for less-conforming contact configurations, the dominant wear mechanism results in damage by bulk material removal, while for more conforming contacts, damage caused by subsurface deformation and adhesive transfer is dominant. They argue that, despite the friction coefficient being independent from contact geometry, at least for fretting displacement amplitudes, the wear coefficient is indeed dependent on the initial configuration. This hap pens not because of the changes in contact pressure, but rather due to the debris flowing out or being retained within the contact zone, a condition commonly found in torsional wear problems. Despite the fact that torsional wear happens in many mechanical and biological systems, such as human hip, knee joints and controllable pitch propellers, studies regarding this topic are still lacking. In particular, the interaction between polymers and metals in such condition is rarely addressed. Therefore, given the clear importance of the topic and considering the lack of research on this particular area, an analysis of the torsional fretting wear behavior of the R3-Offshore steel and a Polycarbonate/Acrylonitrile-Butadiene-Styrene (PC/ABS) blend is pre sented. In order to encompass a possible geometrical effect of the cur vature of the contact interface while still allowing for an microscopy analysis, a sphere-to-flat contact configuration has been chosen. The goal of this study is to identify surface properties (i.e. hardness and wear rate), wear morphology and quantify a possible life extension of a chain link due to a PC/ABS coating.
Fig. 1. Illustration of the general loading system found in a mooring line: (a) Forces due to weights, currents, waves and buoyancy; (b) Displacements and rotations.
amplitude reciprocate angular motion) are to be formed. Therefore, a torsional fretting wear condition is attained. Most alternatives proposed to mitigate wear on chain links are concerned with increasing the strength of the steel employed in the mooring lines. However, elastic chains have been suggested as an alternative to the metallic ones. Such chains would be manufactured by burying steel chain links within a rubber column, so that the elastic properties of the rubber would reduce wear and the impulsive tension acting on the chain. Simulations with this design have shown up to a 50% reduction in the impulsive tension [14]. Mooring line systems with advanced dynamic behaviors were also considered [15]. Nevertheless, a cheaper and simpler solution might be to coat the contact zone of critical chain links with a soft yet resistant material. Coatings have distinct surface properties that enhance the mechan ical performance of its substrate. They are often hard, resilient, thermalprotective and non-conductive. For instance, Polycarbonate/ Acrylonitrile-Butadiene-Styrene (PC/ABS) is a polymer blend widely used in the automotive and consumer electronics industries due to its availability, good impact resistance, high strength modulus and high thermal stability [16–18]. Its popularity can be explained by the nearly additive nature of the blend’s properties with respect to its components [19,20]. A blend with a higher concentration of PC with respect to ABS has the tendency to have higher hardness and overall mechanical strength. On the other hand, a blend with a higher percentage of ABS should be more resilient, impact resistant and easier to machine. A study on the friction and wear behavior of 18 polymers fretted against steel in dry air and humid environments can be found in Ref. [21]. It was observed that the addition of fillers in a polymer matrix has different effects on the wear rate depending on the environmental conditions. Zaitsev [22] investigated the contact of a cobalt and tungsten carbide alloy against a number of polymers and their composites. He concluded that the predominant wear mechanism depended on the chemical composition of the polymer. More specifically, he discovered that for polymers containing active oxygen groups, the most relevant wear mechanism was based on oxidation-fatigue. When in contact with
2. Materials and methods Two material configurations are employed in this study: i) R3 offshore-grade steel versus R3 offshore-grade steel (reference condi tion); and ii) R3 offshore-grade steel versus a PC/ABS 60:40 blend. Most of the properties of the R3 steel and PC/ABS blend were provided by their manufacturers and are listed in Table 1. The only exception is the Poisson’s ratio of the 60:40, which was retrieved from Ref. [24]. Representative samples of the hemispherical pad and flat counter part are depicted in Fig. 2. The pads, made from R3 steel, have a cylindrical body and a hemi spherical tip with a 7.5 mm radius. The flat counterparts, made from R3 steel and PC/ABS, are prismatic shaped (20 mm � 20 mm x 5 mm). The R3 samples were manufactured in a CNC machine (tolerance of �0.1 mm) using material removed from a standard mooring chain link. The PC/ABS samples were produced by SABIC using injection technique. Table 1 Mechanical properties of the R3 offshore grade steel and the PC/ABS 60:40 blend.
2
Material
R3 offshore steel
PC/ABS 60:40
Young modulus, E(GPa) Poisson’s coefficient, ν Yield stress, σy (MPa)
207 0.34 410
2.2 0.35 54
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Fig. 2. Depiction of representative specimens of each test configuration: (a) R3 steel hemispherical pad and R3 steel flat counterpart; (b) R3 steel hemispherical pad and PC/ABS 60:40 flat counterpart.
After production, the flat counterparts are polished in a rotational wet sanding bench at 300 RPM using a procedural method with six grades of sandpaper: 120, 240, 400, 600, 800 and 1200 grid. Afterward, the samples were controlled for roughness (Ramin ¼ 0:4) and dimensional distortions (angle ¼ 1∘). All specimens were carefully cleaned before testing using isopropyl alcohol to avoid unwanted particles (e.g. con taminants, small polishing debris and oxidation) on the contact surfaces. Previous to the wear tests, the hardness of the counterpart samples were analyzed. Two indentors/techniques in two different testers were employed to evaluate the material hardness and check for discrepancies. For the analysis of the micro-hardness, a Vickers micro-indentor (HM2000 Fischer Technology) with a load of 0.1 N was employed. This first method is able to capture hardness variations on the top layer and indicate a possible non-homogeneous surface. The evaluation of the standard hardness was performed in an Universal hardness tester (ZHU 250). The chosen indentation method was the HBW2.5/300 (Brinell test with a 2.5 mm tungsten carbide sphere and a load of 300 N), this indentor has been picked not only for its shape configuration (similar to the one used in the wear test) but also for having a wide scale for evaluation of mean combined hardness of both surface and substrate. Three measurements in three different flat counterparts of each material were carried out. A MTS 809 Axial-Torsional Testing Machine (Fig. 3a) and a sample holder (Fig. 3b) device were used to run the wear tests. The experimental procedure can be summarized as follows. A flat sample is fit inside the holder which is closed with a bolted lid. The holder is then placed and held by the lower grip while the hemispherical pad is held by the upper grip. The pad’s lateral surface (red zone) is fully restrained while each side facet (such as the one in orange) and bottom of the flat sample has a normal displacement restraint, see Fig. 4a. The schematic view of the loading conditions is depicted in Fig. 4b. The crosshead is positioned at a 1 mm gap distance from the flat sample. When the test is stated, the lower grip is programmed to move the flat sample towards the pad until a preselected normal force, Fz , is reached. The normal forces selected are: 1.0 and 2.0 kN. These values were chosen such that two distinct levels of plastic strains, similar to the ones observed in real scenarios, would be generated. As soon as the desired contact force is established, the normal load is maintained and a prescribed rotation (θz ¼ �1∘ at 5 Hz) is applied to the pad during a defined number of cycles (0, 50, 100 and 200 thousand cycles). Each test condition (two material pairings, two loads and four number of cycles) is
Fig. 3. Experimental setup for the sphere-on-flat torsional fretting wear tests. (a) Equipment: 1) Crosshead; 2) Force transducer; 3) Top cell; 4) Upper grip; 5) Lower grip; 6) Bottom cell. Detailed view: 7) Hemispherical pad; 8) Flat sample; 9) Holder.
repeated once for standard deviation assessment. A total of 32 torsional fretting wear tests were performed. After each test, the sample holder device is detached from the grips and disassembled. Then, the specimen is visually inspected for debris and cleaned using a compressed dry air pistol. In the following, a confocal laser microscope (OLYMPUS LEXT OLS4100) is employed to produce profiles of the wear marks (see Fig. 5). The height and lateral resolutions are 10 nm and 12 μm, respectively. A stitching procedure is 3
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Fig. 4. Schematic views of the contact configuration, restraints and loadings: (a) Encastre on the pad (red area) and displacement constraints on the flat sample (orange area); (b) normal force and z-axis rotation. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)
performed in both laser and color spectrum using 1 μm steps and a 20x magnification lens. Due to its darker aspect, a brightness of approxi mately 60% is required for the analysis of PC/ABS samples. For the R3 steel samples, brightness levels ranging from 20%(color) to 40%(laser) were employed. The wear volume is evaluated using two techniques. The first volume estimation method consists of a geometrical approach, employed only for the PC/ABS flat samples (which showed regular, symmetrical and hemispherical shaped scars). Having measured the scar’s depth, h, and diameter, d, (see Fig. 5a), the worn volume V was estimated using the definition of a spherical cap geometry, as follows, �� �2 � 1 d V ¼ ⋅ π ⋅ h⋅ þ h2 : (1) 6 2 Herein, the measurements of depth and diameter were performed in both XZ and YZ planes to evaluate a possible eccentricity of the contact zone. Similar consolidated procedures, regarding micro-scale abrasion and wear tests, can be found in Refs. [25,26]. For the analysis of the R3 steel specimens, which showed irregular scars (see the 2D profile in Fig. 5b), the companion software of the confocal microscope was used. This second method requires the definition of a reference surface and an interest zone, depicted in Fig. 6. Herein, these zones are defined as the unscathed surface (outside the wear region) and a polygon surrounding the outer-border of the wear region, respectively. Then, using a computer-aided profile analysis, the volume (removed or deposited) is measured.
Fig. 5. Representative 3D (top) and 2D (bottom) profiles of flat samples for morphology analysis and quantification of wear: (a) PC/ABS; (b) R3 steel.
pairing solids with such distinct levels of hardness in a wear condition, it is expected that the soft material will endure the majority of the damage. The highest standard deviation observed during the measurements was equal to 11%, meaning that the hardness of both materials remains fairly constant throughout the extension of their surfaces. Moreover, the dif ference between the results provided by the micro-indentation and the standard hardness test were small. For instance, the R3 steel shows a micro-hardness only 7% higher than the micro-hardness while the dif ference for the PC/ABS is 14%. This variation is likely related to a size effect phenomenon and a small density difference between top layer and substrate. Nevertheless, as this variation was rather low, it indicates that the solids are mostly homogeneous over depth. Therefore, the particles produced over the wear cycles should have a fairly constant hardness.
3. Results and discussion In this section, three aspects are discussed: i) hardness; ii) scar morphology; iii) wear analysis; and iv) life estimations. 3.1. Hardness The hardness test results, listed in Table 2, revealed the R3 steel as a �77% harder surface when compared to the PC/ABS 60:40. When 4
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preserved, see Fig’s. 7d-7f. Similar results were obtained by Wang [27], whose research on torsional fretting wear of 1045 steel against a monomer casting nylon filled with glass fiber resulted in analogous profiles: a central region with almost no indentation and an annular area with severe wear grooves. Nevertheless, in all tests in the reference configuration, erosion marks produced by a chipping effect are observed inside the adhesion zone. This pattern is evident in both sides of the interface and becomes more prominent at 2.0 kN (see Fig.s 7j-7l) since under this condition the adhesion zone is 18–22% bigger. In most cases, chipping marks with this aspect are created by the smallest debris pro duced in the mixed-slip zone which, due to the angular movement of the pad, are brought into the adhesion and crushed to even smaller particles. Three trends can be noticed when analyzing these marks: i. as the number of cycles increases the chipping marks become more slender; ii. as one moves away from the first point of contact (center of the adhesion zone) the chipping marks get larger; and iii. the depth of these chipping marks remained fairly constant as the number of cycles increased. These three features corroborate the assumption that the debris decrease in size as the number of cycles increase. Moreover, a dark red powder (typical of oxidised debris) was found inside the adhesion zone. Since the chipping marks were rather mild it is also reasonable to assume that their hardness was not significantly altered by the oxidation process, rendering the debris almost harmless. It is important to emphasize that, if any oxide particles were present in the wear mark after test, they were completely removed during the cleaning process (compressed dry air) as no trace of oxidation (oxide debris/oxidation spots) was found on the samples during the microscopy analysis. On the other, the partial-slip zone shows severe damage. The scars contain protrusions, grooves and lumps unevenly spread indicating an intricate exchange of material. Although the overall shape of the scars is the expected circular pattern, in some cases a rather elliptical mark is found (see Fig’s. 7b and 7e). The exact opposite behavior is found in the PC/ABS marks, as can be seen in Fig. 8. Overall, the wear marks observed in the PC/ABS samples have a smooth bowl-shape (crater). Therefore, a continuous transition from the adhesion zone to the mixed-slip zone is observed. Although a few dispersed dragging scars (marked by black arrows) are found, the damage promoted by the debris is mostly small scratches. Nevertheless, cracks and micro-fractures are found in the corners of all samples (as indicated by the white circles such as the one in Fig. 8a). We believe that these scratch marks were formed at the early stage of the wear process, when the debris produced in the mixed-slip zone would damage the borders of the contact zone. This condition is only evident in the PC/ABS samples, as the material is rather softer and weaker when compared to the R3 steel. As the load increases, the central adhesion area, which used to pre sent several scratches, starts to show a more smooth pattern, see Fig’s 8b-8e. This behavior is notable at the 200k cycles condition depicted in Fig. 8f. Another peculiar feature is the polishing effect created at the borders of the contact zone (denoted by the black circles). These ob servations, once again, indicate a decrease of the debris over load and over number of cycles as observed in the reference condition.
Fig. 6. Computer-aided volume measurement of the material removed (dark yellow: top view) and deposited (yellow: top view and pink: 2D view). (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.) Table 2 Hardness of the flat counterpart batch. HV0.1
HBW 25/300
R3 steel
PC/ABS 60:40
198�5
122�14
199�2
120�8
198�4
115�3
183�7
103�2
188�12
105�3
186�2
104�2
3.2. Scar morphology The images of the scars provide a great level of detail and revealed interesting aspects of the morphology of the worn surfaces. Three important features of the configurations studied are observed right after cleaning the samples: 1. For compression test (i.e. zero cycles) a regular smooth hemispher ical mark is obtained, which is barely visible due to the low inden tation depth produced. Therefore, these results are only detailed in subsection 3.3; 2. In all tests against the PC/ABS flat counterparts and all compression tests the hemispherical pad remained intact; 3. The fretting wear tests in the reference configuration lead to signif icant damage in both pad and flat counterpart.
3.3. Wear analysis As previously mentioned, the analysis of the 3D profiles reveals that material transfer occurred during the tests in the reference configura tion. In some cases, one side of the interface is basically unscathed and only material deposition occurred, as seen in Fig. 9. However, in most cases a complex and nonlinear exchange of material is observed. Therefore, in order to quantify the net wear volume of the fretting wear tests, each side of the interface is evaluated independently and the measurements are taken with respect to a reference surface (e.g. a plane for the flat counterpart or a spherical cap for the hemispherical pad). Results for the volume above the reference surface, Vþ , void below the reference surface, V , and net wear volume of the interface, Vw , are
Representative top views of the profiles at reference condition wear are shown in Fig. 7, we warn the reader that (due to changes in shape and overall dimensional of the wear marks) the scales used were not kept constant. The wear marks on R3 steel samples show, as expected, two distinct zones: an adhesion zone located at the center of the contact region; and a thick outer annular region representing the partial-slip zone. The former shows a shallow indentation and minute signs of damage while the latter presents severe fretting wear scars. On one hand, results with the reference configuration at 1 kN reveal an adhesion zone (inner region of the white dashed circle) well 5
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Fig. 7. Top view of the R3 steel samples.
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Fig. 8. Top view of PC/ABS 60:40 flat counterpart samples.
listed in Table 3. As already stated, the pads used in the compression tests did not presented any sign of damage. Moreover, despite the fact that wear marks are found on the pads used in the 50k cycles tests none of them actually lost material. Furthermore, at 50k cycles the material removed from the flat counterpart can be found deposited on itself and at the center of the pad’s tip. At 100k and 200k cycles both surfaces presented wear damage and a considerable amount of material deposited on their surfaces. Although it is not trivial to discern the origin of the material deposited in each side, it is reasonable to assume that they were removed from their respective surfaces. A graphical representation of the reference configuration net wear volume over number of cycles is provided in Fig. 10. An important observation is that the initial mark created by the compression phase of the tests with 2 kN is 6x bigger in volume than ones produced in the tests with 1 kN. Moreover, the running-in phase is observed and it is likely to have lasted less than 50k cycles. Following [28], this non-linear phase of the tests will be disregarded for a coherent analysis of the wear coefficient after its stabilization. The wear rate observed on the 2 kN tests (1.73E-4mm3/kilocycle) is approximately 2x the one observed in the 1 kN tests (8.00E-5mm3/kilocycle). The high R-squared values obtained for the linear regressions (0.83 for 1 kN and 0.90 for 2 kN) indicate that, despite the complex material exchange observed in the R3 steel pairing, a predictive behavior (from 50k to 200k cycles) is achieved. The measurements of the wear craters produced on the PC/ABS 60:40 flat counterparts are listed in Table 4. The wear volumes observed in this second configuration reached values two orders of magnitude larger than the reference configuration. This behavior was expected as a polymer has a rather softer and weaker structure when compared to a metal, which is evident in the graphical representation of the data dis played in Fig. 11.
The wear rate observed (during the stable phase) was equal to 2.22E2mm3/kilocycle for the 2 kN tests and 1.74E-3mm3/kilocycle for the 1 kN tests, a difference of over 10 times. Therefore, since the R3 steel shown a �2x increase on its wear rate at the same loads, one could state that the PC/ABS is 5x more susceptible to load variations. However, results also show two paramount advantages: a smoother transition from the running-in phase to the stable phase and an even better linear fitting. Although in this analysis it is also assumed that the stable wear phase of the PC/ABS samples started after 50k cycles, we believe that the required number of cycles for achieving this condition might be lower. The R-squared values of 0.86 and 0.91, for 1 kN tests and for 2 kN tests respectively, indicate a reliable linear trend. Furthermore, as no damage is induced on the pads during the PC/ABS tests, this same predictable behavior is observed on the evolution of wear depth of the flat coun terparts, see Fig. 12. 3.4. Life estimations In the guidelines for failure analysis of mooring components [29] it is stated that the dimension of surface defects must be inferior to 5% of the component’s nominal diameter. For instance, hemispherical defects on standard offshore chain links (nominal diameter of 120 mm) must have less than 6 mm in diameter or a perfect hemispherical scar with a radius of 3 mm and a 56 mm3 in volume. In this setting, the wear marks observed at 200k cycles (0.04 mm3 for 1 kN and 0.09 mm3 for 2 kN) would represent 0.07% and 0.16% of the tolerance for surface defects in mooring chains. Considering the most optimistic scenario observed in this study, i.e. a one-sided linear evolution of the wear volume at 1 kN load, at approximately 285.7 M cycles a standard chain link would be flag for replacement. However, given the complex material exchange and non-linear behavior observed in this study, such high lives are never attained. In order to address this issue a rather complicated and costly 7
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Fig. 10. Net wear volume versus number of cycles: reference configuration. Table 4 Volume estimations: PC/ABS 60:40 flat counterpart. Condition
dx [mm]
dy [mm]
hx [mm]
hy [mm]
V[mm3]
1 kN, 0 cycles
2.879 3.186 3.869 4.053 4.159 4.058 4.229 4.307 4.480 4.192 5.579 5.773 6.252 6.194 6.785 6.149
3.102 3.207 3.978 3.996 4.107 4.039 4.271 4.299 4.271 4.335 5.625 5.809 6.297 6.244 6.874 6.242
0.013 0.013 0.053 0.055 0.059 0.056 0.073 0.084 0.052 0.031 0.149 0.203 0.261 0.222 0.345 0.309
0.011 0.016 0.048 0.050 0.053 0.051 0.076 0.085 0.054 0.033 0.142 0.202 0.252 0.221 0.361 0.322
0.042 0.059 0.307 0.333 0.379 0.347 0.529 0.615 0.404 0.229 1.799 2.670 3.970 3.370 6.485 4.773
1 kN, 50k cycles 1 kN, 100k cycles
Fig. 9. Height profiles of the wear marks on R3 steel samples: (a) Hemi spherical pad with severe deterioration at the borders of the contact zone; (b) Flat counterpart with unscathed surface and material deposition.
1 kN, 200k cycles 2 kN, 0 cycles 2 kN, 50k cycles
Table 3 Volume measurements: reference configuration. Flat
2 kN, 100k cycles
Pad
Interface
Condition
Vþ [mm3]
V [mm3]
Vþ [mm3]
V [mm3]
1 kN, 0 cycles
– – 0.020 0.025 0.031 0.042 0.035 0.030 – – 0.036 0.023 0.047 0.061 0.077 0.062
– –
– – 0.020 0.026 0.018 0.036 0.029 0.025 – – 0.071 0.049 0.082 0.055 0.032 0.066
– – 0.000 0.000 0.000 0.028 0.014 0.005 – – 0.000 0.000 0.040 0.037 0.002 0.011
1 kN, 50k cycles 1 kN, 100k cycles 1 kN, 200k cycles 2 kN, 0 cycles 2 kN, 50k cycles 2 kN, 100k cycles 2 kN, 200k cycles a b
– –
0.012 0.019 0.012 0.011 0.008 0.007 0.046 0.013 0.013 0.012 0.017 0.034
2 kN, 200k cycles
Vw [mm3] 0.003 0.003 0.028 0.032 0.037 0.039 0.042 0.043 0.018 0.018 0.061 0.059 0.076 0.067 0.090 0.083
(a) (a)
more than 4.6 M cycles would be required. The simplicity of this assessment is a decisive feature as it greatly facilitates maintenance planing and the prediction of safe limits to avoid unexpected failures. 4. Conclusions
(b) (b)
This study presents a sphere-to-flat torsional fretting wear analysis of two material pairings: a R3 steel pad against a R3 steel counterpart; and a R3 steel pad against a PC/ABS 60:40 counterpart. Analysis of hardness, wear morphology and wear rate are provided. The main conclusions drawn from the observations made in this study can be summarized as follows:
Spherical indentation volume for the 1 kN load. Spherical indentation volume for the 2 kN load.
� In both configurations addressed, a direct relation between wear volume, applied load and number of cycles is observed. After a stable condition is reached (50k cycles), this relation is linear in both configurations; � R3 steel shows higher hardness and lower wear rate when compared to the PC/ABS 60:40. However, the former shows a rather unpre dictable behavior as complex material exchange are observed. This condition becomes even more evident as the load is increased.
maintenance technical procedure is employed. A much more reliable scenario is observed in the proposed solution. The linear relation between wear depth and number of cycles (observed in Fig. 12) enables an easy evaluation of the lifespan of a PC/ABS 60:40 coating. For instance, at 1 kN, over 25.4 M cycles would be needed to completely remove a 5 mm thick protective layer, where as for 2 kN, 8
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� The torsional fretting wear test have shown that the PC/ABS 60:40 flat counterparts are able to protect the interface for a substantial number of cycles while displaying a predictable wear behavior, which is essential for a coating material; � Finally, the linear relation between wear depth and number of cycles on the polymer tested enables a straightforward estimation of wear life of a mooring component. The findings of this study are promising. However, more research is needed. Analysis considering different blend ratios and types of poly mers will follow. Acknowledgments This work was financially supported by the Brazilian National Council for Scientific and Technological Development (CNPq) under the project [406724/2016-4], by the Foundation for the Support of Research (FAPDF) under the project [0193.001521/2016] and by the Coordenaç~ ao de Aperfeiçoamento de Pessoal de Nível Superior (CAPES) [Finance Code 001]. The authors gratefully acknowledge this support. Our thanks are also extended to Eng. Marcelo Fonseca dos Santos and Petrobras/CENPES for the offshore components supplied. References
Fig. 11. Net wear volume versus number of cycles: PC/ABS 60:40 flat counterpart.
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Fig. 12. Wear depth versus number of cycles: PC/ABS 60:40 flat counterpart.
Moreover, the wear volume of the reference condition required a more costly and time consuming evaluation method; � Both configurations have shown scars with two distinct zones: a central circular adhesion zone and an annular mixed-slip zone. The size of these zones were proportional to the applied load; � In the reference configuration, the majority of the damage is found in the mixed-slip zone whereas in the PC/ABS 60:40 flat counterpart the damage is more evenly distributed; � Interesting scratching marks were found in both configurations and they evolved as the number of cycles increased. A dark red powder (oxidised debris) and a black powder (polished particles) were found on the wear scars of the reference configuration and the PC/ABS 60:40 flat counterpart, respectively; 9
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[26] Steier V, Pires M, Doca T. The influence of diamond-like carbon and anodised aluminium oxide coatings on the surface properties of the SAE 305 aluminium alloy. J Braz Soc Mech Sci Eng 2018;40:41. [27] Wang S. Torsional wear behavior of monomer casting nylon composites reinforced with GF: effect of content of glass fiber. Tribol Trans 2013;56(2):178–86.
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