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Processing Fiber-Reinforced Polymers: Specific Wear Phenomena Caused by Filler Materials.


The invention of composite polymers has led to an increase of potential applications for plastic materials due to improved mechanical properties, including their use in high-performance structural components. High strength and stiffness by incorporating specific fibers to plastic materials such as glass, carbon and metallic fibers is a well-known concept in materials science. [1-6]. Usually, the fillers are much smaller compared to the polymer pellets (resin), which are used as base material. These fibers are the source of specific wear phenomena during manufacturing. Injection molding represents the most relevant technique in production for structural plastic components [7-10]. Especially during dosing at higher rotational speed, the brittle and heat-resistant fibers are responsible for various wear phenomena within the compression zone of the injection molding screw. Due to their small diameter and high aspect ratio in relation to the polymer pellets, in the hot transition zone, the abrasive fibers protrude from the surface of the compacted solid bed having an abrasive, brush-like behavior [9, 11-13]. As well a certain fraction of cracked fibers migrate toward the melt pool.

Due to this mechanism, the pellets transform into small tribological brushes scrubbing the extruder surfaces and leading to increased component wear. In Fig. 1, the basic principle of an injection molding machine is depicted in relation to this specific mechanism. In general, an injection molding screw can be separated into: (1) feeding zone, (2) compression zone, and (3) melting zone; see Fig. la.

Due to the extreme pressure and temperature in the compression zone, the injected resin starts to melt. Mennig and Lake [9] highlighted that the formation of a liquid phase is driven by friction, thermal conductivity, and heat dissipation within the polymer materials. The reinforcement fibers play a minor role in the course of this phase transition. In particular, during this solid-to-liquid transition, the filler materials tend to move outwards to the surface of the pellets; see Fig. lb. This process strongly depends on the adhesion between the filler and polymer materials [9, 11, 14-16]. The protruding fibers are the source of accelerated abrasive wear, superimposed on the erosion through the melt. On a small scale, both the tribological brushes (protruding fibers from the solid bed) and broken fibers cause abrasion.

Using hard abrasive particles for reinforcement microcutting (Fig. 1d) and micro-plowing (Fig. le), postulated by Zum Gahr [17], are the main wear mechanisms that occur here. In addition, the impact angle of striking fibers is also important for ductile, metallic materials where the maximum wear loss occurs at an impact angle of ~30[degrees], whereas hard materials (>5 GPa) experience maximum erosive wear at an impingement angle of 90[degrees], perpendicular to the surface [18]. Mahler et al. also showed that the interlocking of filler materials, for example, glass fibers in molten polymer, can also affect the predominant wear mechanism [19]. In the context of the melting zone, Bergstrom et al. [20] was able to simulate wear mechanisms observed for the injection unit directly at the entry to the molds, reporting erosion caused by solid fibers in liquid polymer. The authors of the present study also investigated the role of highly filled polymers in totally molten resin condition in a previous paper, where a significant influence of the heat input to the tool steel was found. The dominant wear mechanisms in the metallic matrix areas were microcutting and micro-plowing [21]. In-depth studies on the sliding conditions between carbon-nanotube reinforced polyamide composites and steel disks under dry sliding conditions also suggested plowing and cutting as possible wear mechanisms [22-24]. However, the above-mentioned studies mainly focused either on the sliding conditions, the wear within the mold, or the utilized composite polymer [25-27]. These studies emphasize that the wear rate depends on various parameters such as type, quantity, shape, and mechanical properties of the additives and fillers. Further process parameters, for example, pressure, viscosity, and temperature also influence wear. This wear phenomenon on screw surfaces and barrels caused by protruding fibers from compacted pellets in the solid bed is insufficiently researched in the present literature.

In practical applications, screw surfaces are protected against wear by various approaches such as hardened tool materials and surface modifications, for example, physical vapor-deposited (PVD) coatings [12, 28-30]. Furthermore, welded or sprayed coating materials (e.g., material composites including cemented carbides) are also applied to protect the tool surfaces [31]. However, the wear resistance of such surfaces is estimated empirically, based on laboratory scale testing because of lacking predictive tools.

As no effective methods for the simulation of the previously described wear phenomena exist in any available literature, this work aims to develop such a testing method. A test design should be proposed, capable of testing tool steels or coatings running against real processed polymers. This work should elucidate the main parameters leading to brush-like wear phenomena induced by hard filler materials trapped in the polymer. The applicability of the new test method should be proven for two typically used tool steels (an untreated in addition to a PVD-coated tool steel). Furthermore, differences in the wear mechanisms will be studied in detail.


Pin-on-Disk Tribometer

For the simulation and evaluation of the wear phenomena taking place in the compression zone of an injection molding machine, a pin-on-disk test system was developed. A pin-on-disk tribometer was chosen, as it closely resembles the movement of resin particles against the walls and screws. The steel used for the injection unit is used as test specimen (primary body) in untreated and coated condition. To emulate the resin counterbody as well as possible, glass fiber-reinforced polymer pins need to be used. It was expected that due to the high pressures and temperatures in the test setup, the polymer pins will locally melt in the tribocontact, resulting to the exposure of the fibers (cf. Figure lb) and eventually leading to accelerated wear of the test specimen.

Tribological investigations were performed on a Nanovea pin-on-disk tribometer, under dry sliding conditions against PA 66 pins. To mount the specific samples, a fixing system as shown in Fig. 2 was designed. The specimens were loaded with 40 N at a sliding radius of 7 mm, for which the rotating disk exhibits a constant linear speed of 0.33 m/s against the fixed pin. The pins were manufactured in two different variations in order to allow testing for two contact pressures, as seen in Fig. 2a and b. The chosen load leads to a nominal contact pressure of ~2.04 MPa for the cylindrical shape (Fig. 2a, [empty set]5 mm, low contact pressure--"low cp" tests) and 14.11 MPa for the conical one (Fig. 2b, [empty set]2 mm, high contact pressure--"high cp" tests).

Tests were performed at RT (22 [+ or -] 2[degrees]C, no additional heating) and 200 [+ or -] 10[degrees]C in order to entail melting of the polymer in the contact zone. For the high-temperature tests, the whole tribometer setup was heated up by a standard box furnace, while the temperature was controlled through a thermocouple placed inside the furnace wall. RT tests were done for both the low and high cp condition, while high-temperature tests were only performed for the high cp condition. Additionally, the sliding distances were chosen at 1,000, 5,000, and 10,000 m to obtain measurable wear rates. The tests were repeated twice for each condition.

Pin Material

Polyamide 66 (PA 66), reinforced with 50 wt% glass fibers (PA66-GF50; Ultramid A3WG10, BASF, Ludwigshafen, Germany [32,33]. fiber diameter 10 pm) was chosen for the pins. All PA 66 pins were produced by injection molding using a standard shape--so called "dogbone" (standard tensile samples) with the dimensions stipulated in DIN EN ISO 2740 (MIM-tensile test sample, [empty set]5 mm in the center). For the final pin shape, the dogbone samples were cut in two pieces at the cylindrical part of the bone (see Fig. 2). As described in the previous section, the tip geometry was modified to increase the contact pressure. Both geometries obtained relatively smooth surfaces after preparation for the flat ([R.sub.a] = 0.096 [+ or -] 0.023 [micro]m) and cone ([R.sub.a] = 0.051 [+ or -] 0.014 [micro]m)-shaped pins.

The dogbone shape should help to achieve a homogenous orientation of the glass fibers in the central part of the sample, oriented in the direction of injection and parallel to the main axis. In order to verify the fiber distribution, a computer tomography (CT) analysis was performed using a micro focus-X-ray tube (Phoenixlxs225) with a detector panel manufactured by PerkinElmer (RID 1640).

Disk Materials

As primary body (disk) a well-established steel grade, used for polymer processing--powder-metallurgical (PM) steel (X190CrVMo20 4 PM)--was chosen. The provided PM-steel disks were produced by hot isostatic procedures. After machining into the required disk geometry ([empty set]30 * 8 mm), the samples were quenched and tempered (1,150[degrees]C/520[degrees]C/520[degrees]C/air/[N.sub.2]) and subsequently polished to provide a surface roughness in the range of [R.sub.a] = 0.010 [+ or -] 0.001 [micro]m. The microstructure of the steel is shown in Fig. 3 obtained by a scanning electron microscope (SEM) in back scattered electron mode. It consists of fine dispersed spherical Cr-V-carbides with a size of 1-3 [micro]m (gray) and sulfides smaller than 1 [micro]m (black) in a martensitic matrix. The measured macro-hardness was close to 730 HV10.

As coatings are often used in the injection molding application, a second test series with PVD coating was performed. A Cr/CrN multi-layer system, featuring an additional oxide-top-layer and commonly used in polymer processing industry, was applied onto the PM steel. The CrN coating was prepared in an Oerlikon Balzers deposition system using standard process parameters. The as-deposited CrN coatings had a surface roughness of [R.sub.a] [approximately equal to] 0.017 [+ or -] 0.001 [micro]m.

The hardness (H) and elastic modulus (E) of the untreated and coated steel disks were characterized by nanoindentation with a ultra micro indentation system, equipped with a Berkovich diamond tip applying loads between 3 and 45 mN. The substrate interference on the H and E values was minimized through evaluating the load-displacement curves as described by Oliver and Pharr [34].

Wear Track Analyses

The wear loss of the individual disks (unmodified and coated ones) was quantified by measuring the mass loss and calculating the corresponding volume loss via the material's density. The volume loss was normalized by the normal load (force) and the sliding distance to obtain the wear rate ([mm.sup.3]/Nm).

Investigations on the wear tracks of the untreated steel and coated disks were performed by SEM operated at 5 kV. To investigate material transfer and coating failure mechanisms, elemental analysis was performed via line scan profiles of the original and worn surfaces, using the integrated energy dispersive X-ray spectroscopy (EDS) system (EDAX TEAM[TM]) operated at an acceleration voltage of 20 kV and a working distance of 10 mm.

The resulting wear tracks were quantitatively analyzed by a chromatographic, confocal profilometer (Nanovea PS50), which was also used to determine the roughness of the untreated and coated disks (averaged over three different surface areas of 10,000 [micro][m.sup.2]).


Glass Fiber Distribution in the Pins

To gain an insight into the fiber distribution of the chosen A3WG10 pins, CT micro-focus imaging is shown in Fig. 4. In Fig. 4a, the pin and its analyzed volumes are highlighted. Cross-sectional images emphasize a symmetric, bimodal dispersion of the fibers, see Fig. 4b. In general, the near-surface regions contain a higher amount of oriented fibers than the inner ones: Fig. 4b shows the total pin width (5 mm) while Fig. 4c shows half of the width with higher magnification. This textural behavior is due to the rheological effects of non-Newtonian fluids resulting in higher adhesion to the molding wall, also known as fountain the flow effect [35-37]. In Fig. 4d, a reworked high-contrast image of Fig. 4c confirms the above description, showing aligned fibers next to the mold walls (up to ~2 mm distance) and a random distribution in the middle of the pin (inner ~1 mm diameter). In addition, in Fig. 4e (taken by light microscopy), also round and edge-shaped fibers can be detected, indicating highly oriented fibers on the outermost regions.

Based on this analysis, the conical pin geometry (Fig. 2b) was expected to lead to lower abrasive wear because the regions of highly oriented fibers were removed by the diameter reduction. Nevertheless, when comparing to the real field application with glass fiber-filled resin (random orientation), the configuration used for the present study will lead to a more severe condition. The exacerbated test configuration is beneficial because in injection molding often solid conglomerates occur before the resin liquidizes. This intermediate state is expected to be decisive for the wear mechanisms related to glass fiber acting as brushes. Therefore, a rigid contact area of the pin in combination with the partly oriented glass fibers, represents a proper contact situation for simulating the contact between the filled polymer and the wall/screw surfaces. Increased contact pressure and/or increased temperature in the test should lead to the exposure of fibers and increased wear of the primary body.

Quantitative Wear Results

The wear rates (volume loss normalized by the applied force and sliding distance) of untreated PM steel disks and CrN-coated ones are presented in Fig. 5. As previously mentioned, the tribological tests were carried out for two different contact pressures and temperature combinations. For the larger pin diameters (contact pressure of ~2 MPa) without any additional heating, wear rates in the range of [10.sup.-8] [mm.sup.3]/Nm could be measured after 10,000 m for both untreated and coated disks--see open circles and diamonds, respectively, in Fig. 5. The maximum wear rates of ~[10.sup.-7] [mm.sup.3]/Nm were measured for PM steel disks, loaded with ~14 MPa at a testing temperature of 200[degrees]C, as indicated by full circles. Compared to RT, high cp testing wear rates increase by a factor of ~4-7 because of the high temperatures during testing.

The CrN samples generally show higher wear resistance than the uncoated samples (except for RT, low cp). For the most severe condition (200[degrees]C, high cp), the wear increase for the untreated material compared to the coated samples is almost fivefold for the longest sliding distance. Notably larger wear rates were also observed for higher contact pressures when testing the CrN-coated disks (full diamonds), compared to the low contact pressure tests.

Generally, the normalized wear rates decrease with increasing sliding distance (except for the PM steel at high cp), indicating that running-in wear is more severe than steady-state wear. The transfer of a thin polymer film together with adhesive and mechanical interaction strongly affects the friction conditions as mentioned in Ref. [14]. The tribo-chemical reactions between the broken polymer chains and complexes may promote the formation of a tribofilm on the interface. Therefore, increasing sliding distances and temperatures can also lead to higher amounts of molten polymer, filling worn pits and thereby protecting the surfaces. A similar scenario could be observed in our testing series, as for most of the longer distance tests, reduced wear rates could be measured, for example, CrN at 200[degrees]C and 14 MPa (full diamonds).

Increasing the contact pressure from ~2 to 14 MPa has only a minor influence on the wear rates of CrN compared to an increase in temperature. Generally for the applied setup, the wear rates are relatively small (in order of [10.sup.-6] to [10.sup.-9] [mm.sup.3]/Nm), comparable to studies conducted by Pogacnik et al. [26, 27]. testing PA6 ULTRAMID B3S (unfilled) pins against stainless steel X105CrMol7.

Due to the contact situation involving glass fibers protruding from a polymer matrix, the loading of the steel is locally much more intense with a high abrasive component due to the hard fibers. The contact situation exacerbates especially with increased temperature, as they promote further filler material (glass fibers) protruding to the melting surface in the contact area, brushing, and straining the sample surfaces.

Analyses of Worn Surfaces

Representative line scans acquired by topography measurements after 10,000 m are shown in Fig. 6. In particular, the depth profiles for three different conditions are plotted: low cp at RT; high cp at RT; and high cp at 200[degrees] C and compared to the untested surface condition. The untreated PM steel disk exhibits only minor abrasive wear during RT testing for both contact pressures. To intensify abrasive wear, an increase of the testing temperature to 200[degrees]C was necessary. It is thought that the temperature increase benefits partial melting of the polymer and that new abrasive glass fibers emerge promoting further abrasive wear, which significantly increases the wear loss.

In the case of coated samples, very minor abrasive wear is detectable, as shown in Fig. 6b. The CrN-coated disks experience for all testing conditions slight adhesive wear, which is expected to be transferred polymer.

The observed wear may be due to several different reasons. One major point is the difference in hardness between the CrN-coated disks with values close to 24 GPa compared to 2 GPa for the untreated PM steel. Also, the microstructure of CrN is more uniform than that of the tool steel, with hard phases and softer matrix areas. Furthermore, the difference in thermal conductivity may also play a role: the untreated steel disks exhibit a thermal conductivity of about 14 W/mK (measured at 20[degrees]C [38]) compared to 2-3 W/mK for the CrN-coated ones (also evaluated at 20[degrees]C [39]). This difference suggests that heat induced by the tribological contact is poorly dissipated on the CrN-coated samples; hence, the frictional heat propagates to the counterbody pin. Two counteracting wear phenomena are envisioned for the contact situation: due to frictional heat, molten polymer is partly transferred to the disk surface, but also exposing glass fibers. The polymer melt possibly fills wear pits on the surface. On the other hand, the newly exposed fibers are able to abrade the surface exacerbating the wear regime.

To further investigate the wear phenomena, the wear tracks of the untreated and coated disks were studied by SEM. In Fig. 7a and b, the worn PM steel surfaces tested at 200[degrees]C for low cp and 10,000 m sliding distance are presented. In both images, well-defined wear tracks are visible. An EDS analysis (Fig. 7c) of these wear tracks, as shown in Fig. 7b, highlights an increased Si-content within the grated traces. For this contact situation and testing duration, the abrasive effect of abrading fibers is predominant compared to filling of the worn tracks by melting polymer.

In contrast, the CrN-coated disks only exhibit one single wear path (200[degrees]C, 10,000 m high cp of 14 MPa) as shown in Fig. 7d and e. The observed wear path perfectly matches the diameter of a single fiber (~10 [micro]m). The EDS analysis of the CrN disk reveals (Fig. 7f) that here the small particles are broken fibers sticking on the surface. They are not able to penetrate the surface due to the high hardness of the CrN.

Contrarily to this, for the untreated PM steel material removal takes place. Individual wear traces have widths of ~50 [micro]m (Fig. 7b), likely caused by a bundle of glass fibers abrading the surface. For a more detailed view of the micro wear mechanisms, SEM investigations at higher magnification are shown in Fig. 8. The hard phases are clearly more wear resistant than the surrounding matrix areas, this can be seen in the left part of Fig. 8a, whereas single hard phases are followed by a tail of remaining material, while in front of the hard phases, the matrix is removed. Larger hard phases are partly shifted by the high contact forces, entailing a gap in front of the carbide as visible in the central region of Fig. 8a. Transferred polymer is partly present and visualized as dark areas on the surface. Clearly abrasive wear dominates, and the micro mechanisms of wear were found to be microcutting for the matrix as shown in Fig. 8b. Here, a glass fiber particle has formed a microchip in the wear track. A remnant of the glass fiber (<1 [micro]m) is still visible behind the chip.

In both materials, abrasive marks may be recognized in small areas of the wear track, which were partly filled up with transferred material. The whole load is transferred through several glass fibers on small contact areas, resulting in localized high pressures on the disks' surfaces. These peaks in contact pressure lead to microcutting and microplowing on the disk surfaces. As mentioned by Zum Gahr [17], the ratio of microcutting to microplowing is a function of the ratio of the impact/attack angle to the critical attack angle and depends on the hardness ratio between the abrasive and the bulk material.

The SEM analysis of the pin counterbody is summarized in Fig. 9. The pin surface is relatively uneven, showing clear zones of abrasively scratched areas and islands of molten polymer, suggesting high contact pressures in localized areas, which were responsible for the separated wear tracks shown in Fig. 7a. In Fig. 9b, perfectly round-shaped fibers can be recognized, being an indication for intact fibers, based on the relative small protruding lengths. Due to the results of this analysis, the suggested wear mechanisms (abrasive brush of glass fibers abrading the steel surface) is achieved by the chosen test parameters.


This study deals with the wear mechanism caused by exposed fibers from partly molten glass fiber reinforced polymer resin against tool materials (for barrels or screws of injection molding machines of extruders). The wear phenomena studied typically occur in the compression zone where the melting process starts. To study this regime, a specially designed pin-on-disk setup is introduced, using glass fiber-reinforced polymer pins as counterpart. Two typical tool materials used in plasticizing units were chosen to evaluate the new test setup: a powder metallurgical tool steel and a CrN PVD-coated tool steel. Test parameters (contact pressure, temperature, and sliding distance) were varied in order to simulate known wear mechanisms from the real application.

The obtained wear rates were small due to the high wear resistance of the chosen materials. Wear on a CrN-protected tool steel consisted mostly of adhesive transfer of molten polymer. The wear on the untreated tool steel reproduces the abrasive wear mechanisms known from the application, caused by the exposed fibers for testing at 200[degrees]C with high contact pressures (~14 MPa).

This effect was confirmed by EDS measurements, which showed that the worn areas fit precisely to the positions in the wear track and the magnitude of the glass fiber diameter (10 [micro]m). SEM investigations of the pins clearly exhibit exposed glass fibers and the disk showed microcutting and microplowing due to the abrasive action of the fibers. On the other hand, the more homogeneous and harder nature of the CrN coating allows for a better protection against the relatively hard glass fibers.

Due to the counterbody and test parameters choice, the wear mechanism involving the abrasive role of filler brushes sticking out of still solid compacted pellets in the solid bed could be experimentally simulated. In future, this test setup will improve the screening of materials for barrels and screws of extruders, maximizing the lifetime of these components.


This work was funded by the Austrian COMET Programme (Project K2 XTribology, Grant No. 849109). Bony Vattappillil is acknowledged for English proofreading.


[1.] J.L. Thomason, Compos. Part A Appl. Sci. Manuf., 39, 1732 (2008).

[2.] J.L. Thomason, Compos. Part A Appl. Sci. Manuf., 40, 114 (2009).

[3.] M. Akay and D. Barkley, J. Mater. Sci., 26, 2731 (1991).

[4.] E. Basavaraj, B. Ramaraj, and J.-H. Lee, Polym. Eng. Sci., 53, 1676 (2013).

[5.] N. Sato, T. Kurauchi, S. Sato, and O. Kamigaito, J. Thermoplast. Compos. Mater., 22, 850 (1988).

[6.] G.Y. Lee, C.K. Dharan, and R. Ritchie, Wear, 252, 322 (2002).

[7.] F. Johannaber and W. Michaeli, Handbuch Spritzgiessen, Hanser, Munchen (2001).

[8.] G. Mennig and K. Stoeckhert, Mold-Making Handbook. 3rd ed., Hanser Publications, Munchen (2013).

[9.] G. Mennig and M. Lake, Verschleissminimierung in der Kunststoffverarbeitung, Phanomene und Schutzmassnahmen, Hanser, Munchen (2008).

[10.] A. Shpenev, Model of Composite Wear with Abrasive Particles, Advanced Materials, Proceedings of PHENMA 2017, 459 (2018).

[11.] M. Reinhard, Leistungsbewertung und Einsatzmoglichkeiten von Werkstoffen fur verschleisbeanspruchte Spritzgiessmaschinenund Extruderbauteile, Dissertation, Technische Hochschule Darmstadt, Darmstadt (1987).

[12.] M. Cremer, E. Broszeit, G. Berg, and M. Heinze, Materwiss. Werksttech., 29, 555 (1998).

[13.] M. Heinze, G. Mennig, and G. Palier, Surf. Coat. Technol., 74-75, 658 (1995).

[14.] W. Wieleba, Arch. Civ. Meeh. Eng., 7, 185 (2007).

[15.] W. Friesenbichler, G.R. Berger, J. Perko, and B.E. Gmbh, "Abrasive/corrosive wear on plastic mold steels, measured under practical processing conditions," in the Polym processing Society 23rd Annual Meeting, pp. 3-11 (2007).

[16.] W. Friesenbichler, G.R. Berger, J. Perko, Praxisnahe Prufmethoden fur die Messung des abrasiven und korrosiven Verschleisses, 19. Leobener Kunststoff Kolloquium (2006).

[17.] K.H. Zum Gahr, Microstructure and Wear of Materials, Vol. 10, Elsevier Science Publishers B.V, Siegen (1987).

[18.] K.H. Zum Gahr, Tribol. Int., 31, 587 (1998).

[19.] W.-D. Mahler, Modelluntersuchung zur Verschleiwirkung stromender Schmelzen, Dissertation, Technische Hochschule Darmstadt, Darmstadt (1975).

[20.] J. Bergstrom, F. Thuvander, P. Devos, and C. Boher, Wear, 250, 1511 (2001).

[21.] A. Blutmager, M. Varga, T. Schmidt, A. Pock, and W. Friesenbichler, Polym. Eng. Sci., 59(S1), E302 (2019).

[22.] H. Unal, U. Sen, and A. Mimaroglu, Tribol. Int., 37, 727 (2004).

[23.] H. Unal and A. Mimaroglu, Mater. Des., 24, 183 (2003).

[24.] H. Meng, G.X. Sui, G.Y. Xie, and R. Yang, Compos. Sci. Technol., 69, 606 (2009).

[25.] G. Mennig and G. Palier, Materwiss. Werksttech., 24, 152 (1993).

[26.] A. Pogacnik and M. Kaiin, Wear, 290-291, 140 (2012).

[27.] A. Pogacnik, A. Kupec, and M. Kaiin, Wear, 378-379, 17 (2017).

[28.] K. Bobzin, E. Lugscheider, M. Maes, R. Cremer, and R. Tariq, Tribologie und Schmierungstechnik, 53, 15 (2006).

[29.] R. Rachbauer, A. Blutmager, D. Holec, and P.H. Mayrhofer, Surf. Coat. Technol., 206, 2667 (2012).

[30.] G. Paller, B. Matthes, W. Herr, and E. Broszeit, Mater. Sci. Eng. A, 140, 647 (1991).

[31.] A. Blutmager, M. Varga, U. Cihak-Bayr, W. Friesenbichler, P. H. Mayrhofer, Wear protective effects of tribolayer formation in dry-sliding contacts of hardmetals for rotating machine components, Wear, under review 2019.

[32.] Datasheet Ultramid[R] A3WG10, Wyandotte, MI (2016).

[33.] Datasheet Ultramid[R] (PA), BASF Corporation, Wyandotte, MI (2013)

[34.] W.C. Oliver and G.M. Pharr, J. Mater. Res., 7, 1564 (1992).

[35.] D.J. Coyle, J.W. Blake, and C.W. Macosko, AIChE J., 33, 1168 (1987).

[36.] M.G.H.M. Baltussen, M.A. Hulsen, and G.W.M. Peters, J. Nonnewton Fluid Mech., 165, 631 (2010).

[37.] H. Mavridis, A.N. Hrymak, and J. Vlachopoulos, Polym. Eng. Sci., 26, 449 (1986).

[38.] Datasheet Bohler M390 microclean, Bohler Edelstahl GmbH, Kapfenberg (2006).

[39.] V. Moraes, H. Riedl, R. Rachbauer, S. Kolozsvari, M. Ikeda, L. Prochaska, S. Paschen, and P.H. Mayrhofer, J. Appl. Phys., 119, 225304 (2016).

[40.] A.G. Atkinsi, Interface Focus, 6, 20160019 (2016).

[41.] K. Hokkirigawa and K. Kato, Tribol. Int., 21, 51 (1988).

Andreas Blutmager, (1) Thomas Spahn, (2) Markus Varga (iD), (3) Walter Friesenbichler, (4) Helmut Riedl, (2) Paul Heinz Mayrhofer (2)

(1) Wittmann Battenfeld GmbH, A-2542 Kottingbrunn, Austria

(2) Institute of Materials Science and Technology, TU Wien, A-1060, Vienna, Austria

(3) AC2T research GmbH, A-2700, Wiener Neustadt, Austria

(4) Chair of Injection Moulding of Polymers, Department Polymer Engineering and Science, Montanuniversitat Leoben, A-8700, Leoben, Austria

Correspondence to: M. Varga; e-mail: Contract grant sponsor: Osterreichische Forschungsforderungsgesellschaft; contract grant number: 849109.

DOI 10.1002/pen.25261

Published online in Wiley Online Library (

Caption: FIG. 1. Schematic sketch of an injection molding machine highlighting (a) three different zones of the extruder screw, (b) and (c) the migration of fibers toward the melting surfaces and the possible wear mechanisms--cutting or plowing--are depicted. Microscopical damage caused by these two most important abrasive wear phenomena are schematically depicted in (d) and (c), respectively [40, 41]. [Color figure can be viewed at]

Caption: FIG. 2. Pin-on-disk test arrangement with the different tested pin-tip geometries (a and b). [Color figure can be viewed at]

Caption: FIG. 3. Microstructure of the PM-steel as seen by SEM-BSE (cf. Ref. [21]).

Caption: FIG. 4. CT imaging of the (a) pin, as well as fibers distribution in (b and c), the cross section of the pin for different magnifications, (d) principal fiber texture, and (e) top view of the cut half pin width as seen by light microscopy. [Color figure can be viewed at]

Caption: FIG. 5. Wear rates as obtained for different sliding distances, contact pressures (cp) and disk materials at room temperature and at 200[degrees]C. [Color figure can be viewed at]

Caption: FIG. 6. Topography for (a) PM steel and (b) CrN-coated surface, both at different contact pressures and temperature conditions in comparison to the untested surface. [Color figure can be viewed at]

Caption: FIG. 7. Worn surfaces of (a, b) PM steel and (d, e) PVD-coated surface in different magnifications, with EDS- line scans of wear tracks performed on (c) PM steel and (f) coated disk. [Color figure can be viewed at]

Caption: FIG. 8. Worn surfaces of PM-steel after 200[degrees]C high cp test: (a) wear track and (b) microcutting wear mechanism; the arrows indicate the sliding direction.

Caption: FIG. 9. SEM-pictures of the worn surface of the cone-shaped pin after PM-steel disk testing at high cp, 200[degrees]C and 10,000 m.
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Author:Blutmager, Andreas; Spahn, Thomas; Varga, Markus; Friesenbichler, Walter; Riedl, Helmut; Mayrhofer,
Publication:Polymer Engineering and Science
Geographic Code:4EUAU
Date:Jan 1, 2020
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