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Mechanisms of high temperature sliding abrasion of metallic materials
151
Wear, 152 (1992) 1.5-159
mechanisms of high temperature metallic materials*
sliding abrasion of
A. Fischer Znstitut fiir WerhTtoj%, Lehrstuhl
(Received
Wer~to~techn~~
April 18, 1991; revised and accepted
Ru~r-~~~ye~~t~t,
W-4630
&chum
1 (FRG)
July 24, 1991)
Abstract Abstract The results of sliding abrasion tests up to 750 “C are presented and discussed. The wear mechanisms acting at elevated temperatures are completely different from those observed at 25 “C. Fragments of flint particles penetrate into the softer surfaces. In addition, layers are generated which adhere strongly to the abraded areas of the specimens. These fragments and layers protect the surfaces against abrasion. The wear rates therefore decrease between 25 and 650 “C. The materials benefit from these effects as far as the metal matrix is capable of supporting the hard constituents (hard phases, layers, fragments). Thus, above
650 “C the hot strength of the metal matrix influences
the wear behaviour
distinctly.
1. Introduction
The wear behaviour of materials can be described by means of the acting wear mechanisms. These are governed by the structure of the tribosystem, the tribological stresses and the material properties. DIN 50320 [l] shows four major wear mechanisms, i.e. abrasion, adhesion, surface fatigue and tribochemical reaction. These mechanisms may appear separately or combined depending on certain types of wear. Considering additional submechanisms such as microploughing, microcutting and microcracking, for abrasion a close correlation between the microst~cture and the tribological properties of materials can be derived. However, this is possible only for simple tribological systems in which just one wear mechanism appears separately or is most pronounced. The acting wear mechanisms can be determined by microscopical investigations of the worn surfaces, the subsurface regions and the wear debris. For room temperature wear this has been carried out by many researchers for the mechanisms of abrasion, adhesion, surface fatigue and t~bo-oxidation l2-83. Wear at elevated temperatures, unfortunately, has been investigated little [9-G]. However, for the development of materials subjected to high temperature wear the wear mechanisms have to be known too. This becomes most important because at elevated temperatures the wear mechanisms are completely different from those observed at room temperature 116, 171. In this paper the wear mechanisms acting in sliding abrasion of metallic materials with and without hard phases above 500 “C are described and discussed. *Paper presented April 7-11, 1991.
at the International
Conference
on Wear of Materials,
Orlando,
FL, USA,
0 1992 - Elsevier Sequoia. All rights reserved
152
2. Experimental
procedure
2.1. Materiab tested For the wear tests different
materials
with well-known
microstructures
are chosen.
50CrV4 (BS 735M50, AISI 6150) in various heat-treated states (w, annealed; g, hardened; g+ a, hardened and tempered) and NiCr20AlTi (Nimonic 80A) contain precipitates smaller than 1 Frn which do not affect the wear behaviour even though they influence the macrohardness. Eutectic hard phases of M7C3 type appear in the cold-worked tool steel X155CrVMo12 1 (BS BD2, AISI D2) and the cobalt-base hardfacing alloy CoCd9W (Stellite 6) (Fig. 1). The average chemicai compositions and the hardness values prior to testing are listed in Table 1. 2.2. Tribological system and test rig According to the German wear terminology a tribological system has four major components [l]: body, counterbody, interfacial medium and surrounding medium. With
CoCr29W
as
cast
Fig. 1. Microstructure TABLE Hardness hardened
70pm I
of an alloy containing
eutectic
hard phases (cast CoCr29W, Stellite 6).
1 and average chemical composition and tempered)
of materials tested (w, annealed; g, hardened;
g C a,
Material
Wardness (W 30)
Average chemical composition (wt.%)
NiCr2OAlTi
280
Ni, 20%Cr, 2.5%Al, l%Ti
50CrV4 w SOCrV4 g 50CrV4 gi-a
270 680 480
Fe, OS%C, l%Cr
190 700 620
Fe, 1.6%C, 12%Cr, l%V
430
Co, 1.2%C, 29%Cr, 4%W
X155CrVMo12 X155CrVMo12 X155CrVMolZ CoCr29W
1 w 1 g 1 gta
153
respect to this the disc is the body, the ring the counterbody, the layer of abrasive particles and wear debris the interfacial medium and air (25 “C) or dry argon (above 500 “C) the surrounding medium (Fig. 2). Loose flint particles are filled into the hollow pusher beam and supplied to the interface between body and counterbody through notches in the ring via the relative motion. The ratio between the abraded areas of the disc (Ad) and ring (A,) is 0.88. The weight losses of the disc (AG,) and ring (AG,), which are made of the same material, are added and related to their density (p), A,,, A, and the length of the wear path (L). This brings about the dimensionless wear rate for sliding abrasion (VVSa):
(1) The test parameters were as follows. The size of the abrasive particles ranged from 63 to 100 Frn, the mean value being 80 pm. The rotational speed of the body was 28 mm s-r, with no measurable heating of the specimens near the worn surfaces. With respect to the complex flow properties of loose abrasive particles and the need for a stationary flow the nominal load was limited to 0.82 MPa. A changeable recipient, which is heated by a furnace, contains the specimens, drive shaft and pusher beam. The tests were carried out at 25, 550, 6.50 and 750 “C. 2.3. Microscopical analyses The wear appearances were investigated on the worn surfaces as well as in the subsurface regions using a scanning electron microscope with an attached energydispersive X-ray (EDX) analyser. All specimens were cleaned intensively by an ultrasonic cleaning process to remove the loose wear debris before scanning electron microscopy 1 (SEM) analyses. To investigate the subsurface regions, the specimens were prepared as follows. After the electroless deposition of a nickel layer the discs were cut with a low speed diamond cutting saw tangentially in the middle of the abraded area. The taper sections Test
Force
F
Envirorent (Air.Argon.etc)
nterfacial abr.
Rotation
particles)
U
Fig. 2. Structure of the ring-on-disc tribologicai system.
154
were ground by lo-mush Sic paper and polished with 1 pm diamond paste. Afterwards the iron-base materials were etched for 5 s in 10% nitric acid at 25 “C. The nickelbase alloy was etched for 50 s in V2A etchant (100 mt distilled water, 100 ml hydro&h~ori~ acid, 10 ml nitric acid, 0.3 ml pickling inhibitor) and the cobait-base alloy for 5 s in Murak~~ (100 ml distjiied water, 10 g potassium bexaeyanoferrate(III), 10 g ~otassjum hydroxide) at 25 “C. To avoid blooming by non-metallic constituents, the taper sections were covered with a gold layer of about 3 nm thickness by sputtering.
3. Results
and discussion
Figure 3 shows the wear rates measured between 25 and 750 OC using flint (300 HV 0.05) as abrasive particles. Obviously, the wear rates at 550 and 650 “C are lower than those measured at room temperature. This is more pronounced when the materials are softer. The arrows indicate tests which had no constant wear rates during the entire testing time. These wear tests had to be interrupted because of the low strength of the rneta~l~cmatrices at these temperatures bringing about chip and groove formation of a few millimetres in size. This severe damage of the surfaces resulted in failure of the specimens (Fig. 4). The wear rates measured at room temperature depend on the microstructures and properties of the constituents in terms of the microIlardness of the metal matrix and the volume fraction of the hard phases. The harder the metal matrix of a material is, the less it is scratched by relatively soft abrasive particles such as flint (Fig. 5). Eutectic hard phases, which are harder than flint, cannot be scratched by these abrasive particles. The hard phases have therefore to be removed from the surfaces by two different ways of micr~ra~king. First, owing to the progressive wear of the metal matrix the hard phases protrude from the surfaces and are easily torn ofI. Secondly, the eutectic hard phases do not withstand the locally acting high stresses and are fractured, losing their support within the metal matrix. Thus the fear-reducing effect
Fig. 3. Sliding abrasion rates of metallic materials up to 750 “C. Fig. 4. Wear appearance
on an
iron-based material without hard phases at 750 “C.
Flint,
80 urn, XP
C, Air
680
Fig. 5. Influence of bulk hardness 4 w; bottom, 50 CrV 4 g. Fig. 6. Appearance
i-h’30
Flint,
80 pm, 650°C,
on the groove formation
of sliding abrasion
Argon
of metallic materialts: top, 50 CrV
on NiCrZOAi~i at 25 and 650 “C.
of eutectic hard phases is more pronounced for those materials which have a metal matrix of sufficient strength. If the hard constituents are embedded in a soft metal matrix (Fig. 3, X155CrVMo12 1 w), their influence on the wear resistance is not as pronounced as it is for materials with a hard metal matrix (Fig. 3, X155CrVMo12 1 g}. A distinct capability of the metal matrix to work harden renders a similar beneficial effect on the wear behaviour (Fig. 3, CoCr29W). However, these apparently simple correlations between wear behaviour and the microstructure of metallic materials, with abrasion being the predominant wear mechanism, are in fact much more complicated 123. Comparing the worn surfaces of the nickel-based alloy after the 25 and 650 “C wear tests, it is obvious that the formation of a layer has taken place during the wear process {Fig. 6). This layer consists of fractured or deformed flint particles and sticks strongly to the worn surface. Even an intensive treatment of the specimens in an ultrasonic cleaning bath did not tear off these layers. EDX analyses of the worn areas showed an increasing silicon content with increasing temperature, e.g. for NiCdOAlTi: at 25 “C, 7% f 5%, at 550 “C, 40% ct lS%, at 650 “C, 50% rt 15%, at 750 “C, 65% rt 15%. Even though the measured results scatter, the tendency is unequivocal and similar for all other materials investigated (Fig. 7). The unworn materials contain almost no silicon. Thus the increased silicon contents after room temperature tests can be explained by embedded tine fragments of abrasive particles which are pressed into the metal matrices and remain there.
_ 100
200
300
Testing
Fig. 7. Scattering
400
500
Temperature
SO0
700
800
in “C
bands of the silicon content
NiCr20AITi,750~.Ar,Flint,80 measured
um,0.82
MPa
on the worn surfaces.
Fig. 8. Layer on the worn surface of NiCROAITi containing
silicon (taper section).
The higher the temperature, the softer are the metal matrices and the more abrasive particles penetrate the worn surfaces. Therefore the deformed flint particles generating layers (Fig. 8) and their fragments sticking rigidly to the surface act like hard phases. They protect the surfaces against abrasion, overbalancing the effect of metal matrix softening with increasing temperature. For materials whose microstructure consists of eutectic hard phases embedded in a sufficiently hard metal matrix the circumstances are similar but the effect is less pronounced (Fig. 3, CoCr29W, X155CrVMo12 1 g). One reason for this is that these alloys already contain constituents impeding abrasion. Therefore the layers and fragments sticking to the surfaces just increase the volume fraction of uncuttable constituents (Fig. 9). Another reason is the different kind of motion of the abrasive particles. Owing to sufficiently supported hard phases, the abrasive particles change from a predominant sliding motion to a rolling motion. This can be seen by the pronounced indentations in the worn surface of an alloy containing hard phases (Fig. 10). Materials without eutectic hard phases, such as NiCr20AlTi, show grooves (Fig. 6). Owing to the change in the type of motion from sliding to rolling, the fraction of layers on the surfaces decreases. This can be verified by comparing the average silicon contents in the worn areas for the materials with and without eutectic hard phases (Fig. 7). The microscopically acting mechanical stresses of a rolling abrasive particle are lower than those of a particle that sticks within the body and is driven through the surface of the counterbody. Therefore the wear rates are lower. Above 650 “C the hot strength of the metal matrix becomes more important for the wear behaviour because the hard phases as well as the layers have to be sufficiently supported within the surfaces to impede wear efficiently. Therefore the iron-base materials fail above 650 “C. Owing to the fact that the depth of the grooves produced by the abrasive particles is larger than the thickness of the layers and the sizes of the flint fragments and eutectic hard phases, they are removed from the surface. Investigation of these surface layers shows a laminated structure consisting of base
ic .bides
X155CrVMo
12 l,Flint,BO
Um,750”C,Ar.O.82
MI Ja
CoCr29W,
Fig. 9. Flint fragment
embedded
Fig. 10. Indentations
from rolling abrasive particles
Fig. 11. Laminated structure tool steel at 750 “C. material particles
650°C.Ar,Flint,80
in the WOI rn surface of a cold-worked
MPa
on the worn surface of CoCr29V.J.
of the surface layer produced
(metal matrix and eutectic hard (Fig. II). Thus, these uncuttable
pm,033
tool steel (taper sect :ion).
by sliding abrasion of a cold-worked
phases) and deformed or fractured abrasive constituents do not provide any protection
against sliding abrasion. Metal matrices with a sufficient hot strength, as is the case for the nick& and cobalt-base alloys, are deformed less. Thus they support the uncuttable constituents
Cr-Distribution Fig.
12. Twin
layer
Si-Distribution on the worn
surface
of NiCdOAlTi
containing
silicon
and
Si-Cr.
even at higher temperatures, providing a steady low wear rate (Fig. 3, NiCr20AlTi, CoCr29W). The bonding mechanism between the deformed flint particles and the worn surfaces is unknown, but there is evidence for chemical bonding by the generation of an Si-Cr solid solution. Figure 12 shows a twin layer on the worn surface of the nickel-base alloy consisting of a silicon- and an Si-Cr-containing layer. Owing to the fact that flint contains no chromium, it must stem from the base material. The structure of this Si-Cr layer has not been examined yet. However, it shows that some kind of tribochemical reaction must have taken place. Even though a temperature of more than 1000 “C is required for such a solid-solid diffusion-controlled reaction, it may take place at distinctly lower temperatures under tribological stresses [18].
4. Conclusions Sliding abrasion tests of metallic materials with and without eutectic hard phases have been carried out between 25 and 750 “C. Investigation of the wear appearances and acting wear mechanisms brought about the following results. (1) Up to 650 “C the wear rates decrease with increasing temperature. This is more pronounced for materials with a softer metal matrix. Above 500 “C layers are generated which stem from deformed flint particles and adhere strongly to the abraded areas. In addition, flint fragments penetrate the softer surfaces. Both effects render hard constituents impeding abrasion. (2) The strong bonding between the layers and the worn surfaces is achieved by mechanical interlocking and adhesion. (3) It is not clear how the observed diffusion processes were possible at the relatively low temperatures. In addition, the influence of work hardening above 550 “C on the wear behaviour is unknown.
159 Acknowledgments
The author would like to thank cordially Professor Dr.-Ing. I-I. Berns for the initiation and further discussion of this work. In addition, the author would like to thank Ms. C. Rademacher and Mrs. M. Boettcher for their parts in wear testing and microstructural analyses. This investigation was sponsored by the Deutsche Forschungsgemeinschaft under contract SFB 316/C2.
References 1 DlN 50320: Wear, ~e~n~i~~s and Te~inolo~~ Beuth, Berlin, 1980. 2 K.-H. Zum Gahr, Microstructure and Wear of Murerials, Ttib&gy Series, Vol. IO, Elsevier, Amsterdam, 1987. 3 H. Czichos, Tribology - A Systems Approach to the Science and Technology of Friction, ~ub~cution and Wear, T~boIo~ Series, Vol. 1, Elsevier, Amsterdam, 1977. 4 K.-H. Habig, Verschieiss und Hiirte von Werkstoffen, Carl Hanser, Munich, 1980. 5 J. Hailing, Principles of Tribology, Macmillan, London, 1979. 6 D. A. Rigney, Sliding wear of metals, Ann. Rev. Mater. Sci., 18 (1988) 141-163. 7 N. P. Suh, The delamination theory of wear, Wear, 25 (1973) 111-124. 8 T. F. J. Quinn, A review of oxidational wear, NASA Contractor Rep. 3686, 1983. 9 M. F. Amateau and W. A. Glaeser, Survey of materials for high temperature bearing and sliding applications, Wear, 7 (1964) 385-418. 10 Yu. G. Tkachenko, V. N. Klimenko, I. N. Gorbatov, V. A. Maslyuk and D. 2. Yurchenko, Friction and wear of Cr.-carbide Ni-alloys at temperatures of 20-1000 “C, Poroshk. Metall., 11 (1978) 60-63. 11 M. Imai and E. Rabinowicz, Lubrication by low-melting-point metals and elevated temperatures, ASLE Trans., 6 (1963) 286-294. 12 J. Saverna, B. Vigneron and J. M. Schissler, Development of a hot abrasive testing machine. Effects of initial structure before testing. Application to steel Z3OC, Mater. Corms. Znd., 55 (11-12) (1979) 651-652. 13 K. Wellinger and H. Uetz, Gleit-, Spiil- und Strahlverschleissprtifung, Schweiz. Arch. Angew. Wiss. Techn., 24 (1958) l-11. 14 R. Zeihe, H. E. Schweite and L. Zagar, Heissabrieb feuerfester Baustoffe, Toni&. Ztg., 98 (3) (I974) 53-58. 15 S. Soemantri, A. C. McGee and I. Finnie, Some aspects of abrasive wear at elevated temperatures, in K. C. Ludema (ed.), Wear of Matetials 1985, Vancouver, April 1985, ASME, New York, 1985, pp. 338-344. 16 A. Fischer, Korngleit- und Gleitverschleiss metallischer Werkstoffe bei erhbhten Temperaturen, in K.-H. Zum Gahr (ed.), Reibung und Verschleiss bei Me~alI~sche~ und ~ich~~etuZl~chen Werkstoflen, Syrup. der DGM, Bad Nauheinr, October 1989, Deutsche Gesellschaft fiir Materialkunde, Oberursel, 1990, pp. 265-277. 17 M. Gienau, M. Woydt and K.-H. Habig, Hochtemperatur Tribometer fur Reibungs- und Verschleisspr~fungen bis 1000 “C, Ma~e~ulp~~n~ 29 (7-8) (1987) 197-199. 18 G. Heinicke, Tribochemtitq, Carl Hanser, Munich/Vienna, 1984, pp. 405-406.
Wear, 152 (1992) 1.5-159
mechanisms of high temperature metallic materials*
sliding abrasion of
A. Fischer Znstitut fiir WerhTtoj%, Lehrstuhl
(Received
Wer~to~techn~~
April 18, 1991; revised and accepted
Ru~r-~~~ye~~t~t,
W-4630
&chum
1 (FRG)
July 24, 1991)
Abstract Abstract The results of sliding abrasion tests up to 750 “C are presented and discussed. The wear mechanisms acting at elevated temperatures are completely different from those observed at 25 “C. Fragments of flint particles penetrate into the softer surfaces. In addition, layers are generated which adhere strongly to the abraded areas of the specimens. These fragments and layers protect the surfaces against abrasion. The wear rates therefore decrease between 25 and 650 “C. The materials benefit from these effects as far as the metal matrix is capable of supporting the hard constituents (hard phases, layers, fragments). Thus, above
650 “C the hot strength of the metal matrix influences
the wear behaviour
distinctly.
1. Introduction
The wear behaviour of materials can be described by means of the acting wear mechanisms. These are governed by the structure of the tribosystem, the tribological stresses and the material properties. DIN 50320 [l] shows four major wear mechanisms, i.e. abrasion, adhesion, surface fatigue and tribochemical reaction. These mechanisms may appear separately or combined depending on certain types of wear. Considering additional submechanisms such as microploughing, microcutting and microcracking, for abrasion a close correlation between the microst~cture and the tribological properties of materials can be derived. However, this is possible only for simple tribological systems in which just one wear mechanism appears separately or is most pronounced. The acting wear mechanisms can be determined by microscopical investigations of the worn surfaces, the subsurface regions and the wear debris. For room temperature wear this has been carried out by many researchers for the mechanisms of abrasion, adhesion, surface fatigue and t~bo-oxidation l2-83. Wear at elevated temperatures, unfortunately, has been investigated little [9-G]. However, for the development of materials subjected to high temperature wear the wear mechanisms have to be known too. This becomes most important because at elevated temperatures the wear mechanisms are completely different from those observed at room temperature 116, 171. In this paper the wear mechanisms acting in sliding abrasion of metallic materials with and without hard phases above 500 “C are described and discussed. *Paper presented April 7-11, 1991.
at the International
Conference
on Wear of Materials,
Orlando,
FL, USA,
0 1992 - Elsevier Sequoia. All rights reserved
152
2. Experimental
procedure
2.1. Materiab tested For the wear tests different
materials
with well-known
microstructures
are chosen.
50CrV4 (BS 735M50, AISI 6150) in various heat-treated states (w, annealed; g, hardened; g+ a, hardened and tempered) and NiCr20AlTi (Nimonic 80A) contain precipitates smaller than 1 Frn which do not affect the wear behaviour even though they influence the macrohardness. Eutectic hard phases of M7C3 type appear in the cold-worked tool steel X155CrVMo12 1 (BS BD2, AISI D2) and the cobalt-base hardfacing alloy CoCd9W (Stellite 6) (Fig. 1). The average chemicai compositions and the hardness values prior to testing are listed in Table 1. 2.2. Tribological system and test rig According to the German wear terminology a tribological system has four major components [l]: body, counterbody, interfacial medium and surrounding medium. With
CoCr29W
as
cast
Fig. 1. Microstructure TABLE Hardness hardened
70pm I
of an alloy containing
eutectic
hard phases (cast CoCr29W, Stellite 6).
1 and average chemical composition and tempered)
of materials tested (w, annealed; g, hardened;
g C a,
Material
Wardness (W 30)
Average chemical composition (wt.%)
NiCr2OAlTi
280
Ni, 20%Cr, 2.5%Al, l%Ti
50CrV4 w SOCrV4 g 50CrV4 gi-a
270 680 480
Fe, OS%C, l%Cr
190 700 620
Fe, 1.6%C, 12%Cr, l%V
430
Co, 1.2%C, 29%Cr, 4%W
X155CrVMo12 X155CrVMo12 X155CrVMolZ CoCr29W
1 w 1 g 1 gta
153
respect to this the disc is the body, the ring the counterbody, the layer of abrasive particles and wear debris the interfacial medium and air (25 “C) or dry argon (above 500 “C) the surrounding medium (Fig. 2). Loose flint particles are filled into the hollow pusher beam and supplied to the interface between body and counterbody through notches in the ring via the relative motion. The ratio between the abraded areas of the disc (Ad) and ring (A,) is 0.88. The weight losses of the disc (AG,) and ring (AG,), which are made of the same material, are added and related to their density (p), A,,, A, and the length of the wear path (L). This brings about the dimensionless wear rate for sliding abrasion (VVSa):
(1) The test parameters were as follows. The size of the abrasive particles ranged from 63 to 100 Frn, the mean value being 80 pm. The rotational speed of the body was 28 mm s-r, with no measurable heating of the specimens near the worn surfaces. With respect to the complex flow properties of loose abrasive particles and the need for a stationary flow the nominal load was limited to 0.82 MPa. A changeable recipient, which is heated by a furnace, contains the specimens, drive shaft and pusher beam. The tests were carried out at 25, 550, 6.50 and 750 “C. 2.3. Microscopical analyses The wear appearances were investigated on the worn surfaces as well as in the subsurface regions using a scanning electron microscope with an attached energydispersive X-ray (EDX) analyser. All specimens were cleaned intensively by an ultrasonic cleaning process to remove the loose wear debris before scanning electron microscopy 1 (SEM) analyses. To investigate the subsurface regions, the specimens were prepared as follows. After the electroless deposition of a nickel layer the discs were cut with a low speed diamond cutting saw tangentially in the middle of the abraded area. The taper sections Test
Force
F
Envirorent (Air.Argon.etc)
nterfacial abr.
Rotation
particles)
U
Fig. 2. Structure of the ring-on-disc tribologicai system.
154
were ground by lo-mush Sic paper and polished with 1 pm diamond paste. Afterwards the iron-base materials were etched for 5 s in 10% nitric acid at 25 “C. The nickelbase alloy was etched for 50 s in V2A etchant (100 mt distilled water, 100 ml hydro&h~ori~ acid, 10 ml nitric acid, 0.3 ml pickling inhibitor) and the cobait-base alloy for 5 s in Murak~~ (100 ml distjiied water, 10 g potassium bexaeyanoferrate(III), 10 g ~otassjum hydroxide) at 25 “C. To avoid blooming by non-metallic constituents, the taper sections were covered with a gold layer of about 3 nm thickness by sputtering.
3. Results
and discussion
Figure 3 shows the wear rates measured between 25 and 750 OC using flint (300 HV 0.05) as abrasive particles. Obviously, the wear rates at 550 and 650 “C are lower than those measured at room temperature. This is more pronounced when the materials are softer. The arrows indicate tests which had no constant wear rates during the entire testing time. These wear tests had to be interrupted because of the low strength of the rneta~l~cmatrices at these temperatures bringing about chip and groove formation of a few millimetres in size. This severe damage of the surfaces resulted in failure of the specimens (Fig. 4). The wear rates measured at room temperature depend on the microstructures and properties of the constituents in terms of the microIlardness of the metal matrix and the volume fraction of the hard phases. The harder the metal matrix of a material is, the less it is scratched by relatively soft abrasive particles such as flint (Fig. 5). Eutectic hard phases, which are harder than flint, cannot be scratched by these abrasive particles. The hard phases have therefore to be removed from the surfaces by two different ways of micr~ra~king. First, owing to the progressive wear of the metal matrix the hard phases protrude from the surfaces and are easily torn ofI. Secondly, the eutectic hard phases do not withstand the locally acting high stresses and are fractured, losing their support within the metal matrix. Thus the fear-reducing effect
Fig. 3. Sliding abrasion rates of metallic materials up to 750 “C. Fig. 4. Wear appearance
on an
iron-based material without hard phases at 750 “C.
Flint,
80 urn, XP
C, Air
680
Fig. 5. Influence of bulk hardness 4 w; bottom, 50 CrV 4 g. Fig. 6. Appearance
i-h’30
Flint,
80 pm, 650°C,
on the groove formation
of sliding abrasion
Argon
of metallic materialts: top, 50 CrV
on NiCrZOAi~i at 25 and 650 “C.
of eutectic hard phases is more pronounced for those materials which have a metal matrix of sufficient strength. If the hard constituents are embedded in a soft metal matrix (Fig. 3, X155CrVMo12 1 w), their influence on the wear resistance is not as pronounced as it is for materials with a hard metal matrix (Fig. 3, X155CrVMo12 1 g}. A distinct capability of the metal matrix to work harden renders a similar beneficial effect on the wear behaviour (Fig. 3, CoCr29W). However, these apparently simple correlations between wear behaviour and the microstructure of metallic materials, with abrasion being the predominant wear mechanism, are in fact much more complicated 123. Comparing the worn surfaces of the nickel-based alloy after the 25 and 650 “C wear tests, it is obvious that the formation of a layer has taken place during the wear process {Fig. 6). This layer consists of fractured or deformed flint particles and sticks strongly to the worn surface. Even an intensive treatment of the specimens in an ultrasonic cleaning bath did not tear off these layers. EDX analyses of the worn areas showed an increasing silicon content with increasing temperature, e.g. for NiCdOAlTi: at 25 “C, 7% f 5%, at 550 “C, 40% ct lS%, at 650 “C, 50% rt 15%, at 750 “C, 65% rt 15%. Even though the measured results scatter, the tendency is unequivocal and similar for all other materials investigated (Fig. 7). The unworn materials contain almost no silicon. Thus the increased silicon contents after room temperature tests can be explained by embedded tine fragments of abrasive particles which are pressed into the metal matrices and remain there.
_ 100
200
300
Testing
Fig. 7. Scattering
400
500
Temperature
SO0
700
800
in “C
bands of the silicon content
NiCr20AITi,750~.Ar,Flint,80 measured
um,0.82
MPa
on the worn surfaces.
Fig. 8. Layer on the worn surface of NiCROAITi containing
silicon (taper section).
The higher the temperature, the softer are the metal matrices and the more abrasive particles penetrate the worn surfaces. Therefore the deformed flint particles generating layers (Fig. 8) and their fragments sticking rigidly to the surface act like hard phases. They protect the surfaces against abrasion, overbalancing the effect of metal matrix softening with increasing temperature. For materials whose microstructure consists of eutectic hard phases embedded in a sufficiently hard metal matrix the circumstances are similar but the effect is less pronounced (Fig. 3, CoCr29W, X155CrVMo12 1 g). One reason for this is that these alloys already contain constituents impeding abrasion. Therefore the layers and fragments sticking to the surfaces just increase the volume fraction of uncuttable constituents (Fig. 9). Another reason is the different kind of motion of the abrasive particles. Owing to sufficiently supported hard phases, the abrasive particles change from a predominant sliding motion to a rolling motion. This can be seen by the pronounced indentations in the worn surface of an alloy containing hard phases (Fig. 10). Materials without eutectic hard phases, such as NiCr20AlTi, show grooves (Fig. 6). Owing to the change in the type of motion from sliding to rolling, the fraction of layers on the surfaces decreases. This can be verified by comparing the average silicon contents in the worn areas for the materials with and without eutectic hard phases (Fig. 7). The microscopically acting mechanical stresses of a rolling abrasive particle are lower than those of a particle that sticks within the body and is driven through the surface of the counterbody. Therefore the wear rates are lower. Above 650 “C the hot strength of the metal matrix becomes more important for the wear behaviour because the hard phases as well as the layers have to be sufficiently supported within the surfaces to impede wear efficiently. Therefore the iron-base materials fail above 650 “C. Owing to the fact that the depth of the grooves produced by the abrasive particles is larger than the thickness of the layers and the sizes of the flint fragments and eutectic hard phases, they are removed from the surface. Investigation of these surface layers shows a laminated structure consisting of base
ic .bides
X155CrVMo
12 l,Flint,BO
Um,750”C,Ar.O.82
MI Ja
CoCr29W,
Fig. 9. Flint fragment
embedded
Fig. 10. Indentations
from rolling abrasive particles
Fig. 11. Laminated structure tool steel at 750 “C. material particles
650°C.Ar,Flint,80
in the WOI rn surface of a cold-worked
MPa
on the worn surface of CoCr29V.J.
of the surface layer produced
(metal matrix and eutectic hard (Fig. II). Thus, these uncuttable
pm,033
tool steel (taper sect :ion).
by sliding abrasion of a cold-worked
phases) and deformed or fractured abrasive constituents do not provide any protection
against sliding abrasion. Metal matrices with a sufficient hot strength, as is the case for the nick& and cobalt-base alloys, are deformed less. Thus they support the uncuttable constituents
Cr-Distribution Fig.
12. Twin
layer
Si-Distribution on the worn
surface
of NiCdOAlTi
containing
silicon
and
Si-Cr.
even at higher temperatures, providing a steady low wear rate (Fig. 3, NiCr20AlTi, CoCr29W). The bonding mechanism between the deformed flint particles and the worn surfaces is unknown, but there is evidence for chemical bonding by the generation of an Si-Cr solid solution. Figure 12 shows a twin layer on the worn surface of the nickel-base alloy consisting of a silicon- and an Si-Cr-containing layer. Owing to the fact that flint contains no chromium, it must stem from the base material. The structure of this Si-Cr layer has not been examined yet. However, it shows that some kind of tribochemical reaction must have taken place. Even though a temperature of more than 1000 “C is required for such a solid-solid diffusion-controlled reaction, it may take place at distinctly lower temperatures under tribological stresses [18].
4. Conclusions Sliding abrasion tests of metallic materials with and without eutectic hard phases have been carried out between 25 and 750 “C. Investigation of the wear appearances and acting wear mechanisms brought about the following results. (1) Up to 650 “C the wear rates decrease with increasing temperature. This is more pronounced for materials with a softer metal matrix. Above 500 “C layers are generated which stem from deformed flint particles and adhere strongly to the abraded areas. In addition, flint fragments penetrate the softer surfaces. Both effects render hard constituents impeding abrasion. (2) The strong bonding between the layers and the worn surfaces is achieved by mechanical interlocking and adhesion. (3) It is not clear how the observed diffusion processes were possible at the relatively low temperatures. In addition, the influence of work hardening above 550 “C on the wear behaviour is unknown.
159 Acknowledgments
The author would like to thank cordially Professor Dr.-Ing. I-I. Berns for the initiation and further discussion of this work. In addition, the author would like to thank Ms. C. Rademacher and Mrs. M. Boettcher for their parts in wear testing and microstructural analyses. This investigation was sponsored by the Deutsche Forschungsgemeinschaft under contract SFB 316/C2.
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