Wear, 105 (1985) 1 - 17 1

~FLUENCE OF P~EC~ITATION AND SOLUTION STRENGTHEN~G ON ABRASIVE WEAR RESISTANCE*

STAFFAN SeDERBERG’, ULF BRYGGMAN$ and ALFRED0 CANALES**

Uniuersity of Houston, Houston, TX 77004 (U.S.A.)

(Received April 26,1985; accepted May 9,1985)

Summary

The influence of pr~ipi~tion and solution s~en~hen~g on gouging abrasion resistance has been studied using a 6061 aluminium alloy heat treated to obtain four different microstructures. Single-event abrasive grooves were generated using a modified Charpy impact tester and the specific energy consumption in grooving was recorded and used as a measure of wear resistance. In addition, the me~han~ms of chip formation and material removal were studied by met~lo~aphic analysis of “quick-sap” specimens.

The results of the grooving tests show that for small grooves precipita- tion strengthening yields higher energy values than solution strengthening, the wear resistance increasing with decreasing precipitate size. The ranking of the different heat ~eatments with respect to specific grooving energy depends, however, on the considered damage depth and the opposite order of ranking is found at the largest groove depths. On the basis of the results of the grooving tests, a discussion on the influence of groove depth and microstructure on chip formation and specific grooving energy is presented. In particular, the inhomogeneous nature of the plastic deformation involved in chip formation is discussed, and the resulting importance of the thermal stability of the material in determining wear resistance is emphasized.

1. Introduction

Abrasive wear refers to the displacement and subsequent removal of material from a surface by the grooving action of hard elements. The domi- nant mechanism of material removal is microcutting by which chip-like fragments are detached from the surface. Grooves can also be formed solely

*Paper presented at the International Conference on Wear of Materials, Vancouver, Canada, April 14 - 18,198s.

‘Present address: Uppsala University, Institute of Technology, Box 534, S-761 21 Uppsala, Sweden.

*Present address: AB Bofors hers, S-160 31 hers Styckebruk, Sweden. **Present address: Celanese Engineering Resins, Bishop, TX 78343, U.S.A.

~943-~64%1%6f $3.36 @ Elsevier ~quoia/~in~d in The Netherlan~

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by plastic deformation without any material actually being removed. How- ever, the combined cumulative action of several grooving events will still cause material removal as a result of surface fatigue and delamination. In addition, wear fragments may form by direct spalling of surface layers in hard and brittle materials.

The abrasive wear process depends on the physical conditions defining the tribosystem. The importance of shape, size and mechanical properties of the abrading elements on wear rate is well established [l - 41. The contact conditions also strongly influence the wear process and three different classes can be distinguished.

(i) Two-body abrasion, in which the abrading elements are attached to or part of one of the two mating surfaces (as in the case of grinding with bonded abrasives).

(ii) Three-body abrasion, in which the abrading elements are loose but entrapped between the mating surfaces (as in the case of polishing using an abrasive paste).

(iii) Impact abrasion, in which the abrading elements impinge on the surface being abraded (as in the case of sandblasting). This case is referred to as either gouging abrasion or erosion, depending on the momentum of the abrading elements.

The classification above is important since it cannot be assumed that a material exhibiting a high wear resistance within one class will perform equally well at conditions corresponding to another class. Of these classes gouging abrasion generally causes the most severe material damage, especially since impact and corrosion are frequently concurrent and interactive damage mechanisms. Since abrasion is not amenable to lubrication, wear damage must be minimized by careful material selection and, wherever possible, design considerations. Consequently, much research has been devoted to the study of abrasive wear and the development of abrasion-resistant mate- rials, as reviewed by Moore [l] and Zum Gahr [ 51.

Several attempts to correlate the abrasive wear rate with experimental and material parameters are found in the literature [l, 6 - 91. Theoretical modelling is, however, extremely difficult since the abrasive events are localized to small volumes in the surface layer and are subjected to extreme conditions of stress, strain, strain rate and temperature. The resulting prop- erties of the affected material will thus be drastically different from those of the bulk [ 10 - 121. Therefore material selection for abrasion resistance must be mainly based on wear data from comparative material testing, performed at conditions as close as possible to the intended application [13,14]. In spite of the large amount of experimental data available, it has proven dif- ficult to draw significant conclusions concerning the influence of micro- structural parameters on abrasive wear resistance [ 15,161. This is in a large part due to the microstructural complexity of the steels and cast irons usually included in these investigations. Therefore more basic research ef- forts are required to determine the influence of different strengthening mechanisms on abrasive wear.

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In the research described in this paper, a 6061 aluminium alloy is used as a model material. The material was heat treated to obtain maximum solu- tion and precipitation strengthening and, in addition, heat treatments producing coarser precipitates were included. The gouging abrasion resis- tance was tested using a single-pass pendulum technique [12,17,18]. The pendulum test utilizes the concept that the energy required to remove a specific amount of material by grooving can be used as a quantitative mea- sure of gouging abrasion resistance. This method is found to give very good correlation with results from field testing of excavator teeth and wear data are available for a large number of materials [19]. In addition, specific energy data are also available from earlier work on grindability-machin- ability where a similar pendulum dynamometer technique has frequently been used [20 - 231. Another advantage of pendulum grooving is the ability to perform quick stops of the abrasive action. This allows detailed studies of the mechanisms of chip formation and material removal in gouging abrasion.

2. Experimental details

2.1. Test equipment The gouging abrasion experiments were performed in a modified

impact testing machine, applying the same principles as in the “Uppsala pendulum” [ 12,17 - 191. Instead of letting the falling pendulum hammer fracture the specimen, an abrasive tip is mounted on the pendulum to form an arcuate groove on the specimen surface (Fig. 1). The energy consumed in the grooving event is recorded (as in a standard impact test) as a func- tion of specimen mass loss. The specific grooving energy (energy consumed per unit mass of material removed) is then computed and plotted as a function of groove depth or mass loss.

The pendulum used has in impact energy of 174 J, a radius of 0.43 m and an entrance velocity of 4.3 m s- ‘. The abrasive tip is made of cemented

Pendulum

1.

fi

x

Abrasive Tip’

r!

,/’ \

‘\ ./ ’ I

‘\. I Energy Consumed

1 ] ‘.

‘.

\ Y

\ I \ I \ \ d /’ \ \ .

69 Specimen

Fig. 1. Principle of the single-pass pendulum grooving test,

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carbide and ground to the geometry of a square-based pyramid with 45” apex angle. The base area is 11 mm X 8 mm and the area of the truncated flank is 1.0 mm X 1.0 mm. The tip is mounted on the pendulum with one of its faces perpendicular to the grooving direction. Rectangular bar speci- mens (100 mm X 13 mm X 10 mm) are positioned in a specimen holder which is designed to allow vertical adjustments for setting of the desired groove depth.

In the quick-stop experiments, a stud is mounted on the pendulum behind the abrasive tip. The stud hits the back of the specimen when the tip is at its maximum groove depth. By using breaker pins to hold the speci- men in position at the front, the impact of the stud releases the specimen from the holder and ejects it tangentially away from the moving tip.

2.2. Materiuls and testing procedure A 6061 aluminium alloy was tested in the grooving experiments. The

material was heat treated to four different conditions, corresponding to (i) solution strengthening, (ii) precipitation strengthening (to maximum hard- ness), (iii) over-aging and (iv) annealing.

The heat treatment sequence used in each case is summarized in Table 1. A series of introductory heat treatment experiments were performed to ensure that the desired microstructures were produced. The average grain size of the matrix phase was measured using light optical microscopy (LOM) and was found to be 100 pm for all four materials. In Table 2 values of measured hardness and estimated values of yield strength and fracture strain obtained from handbook data [ 241 are given.

TABLE 1

Heat treatment procedures

Material Heat treatment

Solution strengthened

Precipitation strengthened

Overaged

Annealed

2 h at 550 “C, quenched to 0 “C

2 h at 550 “C, quenched to 0 “C 4 h at 185 “C, quenched to 0 “C

2 h at 550 “C, quenched to 0 “C 70 h at 230 “C, quenched to 0 “C

2 h at 415 OC, furnace cooled to 20 “C

Directly on completion of each heat treatment, the specimens were wet ground and polished to remove the surface oxide film produced by the heat treatment. The hardness and mass of the specimen were recorded and the grooving experiment was performed. The structure of the solution- strengthened material is unstable and gradually decays into two phases even at room temperature. Therefore only one specimen was heat treated at a time in this case to minimize the time lapse between quenching and grooving.

TABLE 2

Microstructural and mechanical properties

Material Hardness (HRF)

Yield Fracture strength strain (MPa) (%I

Solution strengthened 39 f 3 150 23 Precipitation strengthened 97 + 1 275 12 Overaged 69 + 1 175 12 Annealed 115 2 55 25

Between 20 and 30 grooving experiments were performed for each heat treatment with a range of groove -depths corresponding to mass losses from 0.5 to 150 mg. In addition, two to four quick-stop experiments were performed for each heat treatment. The groove depths were selected to correspond to the different modes of chip formation observed in the spe- cific grooving energy measurements. The quick-stop specimens were sub- jected to careful metallographic analysis using scanning electron microscopy (SEM), LOM and microhardness measurements.

3. Results

3.1. Specific grooving energy measurements The results of the grooving experiments are summarized in Fig. 2 for

all four heat treatments by plotting the specific grooving energy e as a func- tion of the mass loss W. A common characteristic of the different materials (and of gouging abrasion in general [12,18]) is that the specific energy

a E 3 . l,

‘\ 6 ‘.

.I

w’ .\

i $dy_- E ---___ ---_ . ii ---___

8 --.___ -.__ 0.5 I

1.0 10 100

Mass Loss (mg)

Fig. 2. Specific grooving energy e plotted us. mass loss W based on the computed k and q values of Table 3: - - -, solution strengthened; - - - -, precipitation strengthened; -._ , overaged; -, annealed.

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decreases with increasing groove depth. It is seen from the figure that this can be expressed as an empirical relationship:

The lines plotted in Fig. 2 are obtained from a least-squares fit of the experimental data and the corresponding computed iz and q values are given in Table .3. As seen from the figure, a well-defined transition point exists at W = 10 mg in the precipitation-strengthened material, at which both the k and the q v .lues change. It is interesting to note that the present results agree well with earlier results from specific energy measurements in grinding with respect to both specific energy values and the existence of a transition point in the precipitation-strengthened material [20, 251. For all but the largest groove depths, the specific grooving energy values exceed the specific melting energy of aluminium which is approximately 1 J mg-l .

TABLE 3

k and q values from grooving tests

Material k (J mKq 1

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Solution strengthened 1.63 0.94 Precipitation strengthened (W 6 10 mg) 2.82 0.47 Precipitation strengthened (W > 10 mg) 1.08 0.88 Overaged 2.08 0.81 Annealed 1.64 0.88

The intersecting curves in Fig. 2 show that the ranking of the four heat treatments with respect to wear resistance is strongly dependent on the damage size considered. At small groove depths the precipitation- strengthened material yields the highest specific grooving energy values followed by the overaged, the annealed and the solution-strengthened ma- terials. At the other end of the scale, i.e. at large groove depths, the ranking of the materials is just the opposite with solution strengthening giving the highest specific grooving energy values. These results show the importance of considering the damage size range in materials selection against gouging abrasion.

3.2. Quick-stop experiments and metallography Throughout the grooving experiments all chips removed were collected

and examined. By weighing it was found that the mass loss of the specimen coincides with the mass of the chip in practically all cases. This observation emphasizes the importance of studying the mechanisms of chip formation in basic investigations of gouging abrasion. From examination of chips and quick-stop specimens produced in the present experiments, it is found that the morphology of the chips depends on material and groove depth. The

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solution-strengthened, overaged and annealed materials all display similar chip formation characteristics, while the precipitation-strengthened mate- rial exhibits a radically different behaviour.

For small groove depths, the precipitation-strengthened material yields long continuous chips, similar to those usually obtained in machining (Fig. 3). The lamellar inhomogeneous deformation structure of the chip is seen from the SEM micrograph, although the etched cross section shows the individual shear bands to be thin and very closely spaced. The effect of increasing the groove depth is to cause localization of the shear bands into zones of high strain density, leaving the material between the zones less deformed. The resulting chip has a serrated appearance as seen from Fig. 4.

The transition from a fine lamellar to a serrated chip morphology is gradual, but at a groove depth corresponding to a mass loss of approximately 10 mg a drastic change in the mode of chip formation takes place. At these large groove sizes, a completely segmented chip is formed with the individual segments attached to each other only by a thin isthmus of material (Fig. 5).

‘.“;,y’, -;.> , mn

(a) (‘b .“.4?Ty$y&y,+ , mm

Fig. 3. Fine lamellar machining type of chip formation in the precipitation-strengthened material at a small groove depth: (a) SEM photograph; (b) detail of (a).

Fig. 4. Serrated chip formation in the precipitation-strengthened material at a medium groove depth: (a) SEM photograph; (b) LOM photograph.

Fig. 5. Segmented chip formation in the precipitation-strengthened material at a large groove depth: (a) SEM photograph; (b) LOM photograph.

From the etched cross section in the figure it is seen that the deformation is localized to a fairly wide shear zone, eventually resulting in the thin isthmus seen in the SEM micrograph. These zones are characterized by an extremely high strain density while the material within each segment ap- pears little deformed. From a comparison with the results of the specific energy measurements, it is seen that the incipient formation of segmented chips, coincides with the change of slope observed in Fig. 2 for the precipi- tation-strengthened material, It should be noted that, because of the varia- tion in groove depth during each grooving event, all three modes of chip formation described above are observed in chips corresponding to large mass losses.

In the three other materials a completely different type of chip is produced, the material typically being “piled up” in front of the tip. In Fig. 6 an example of chips formed

(a) Fig. 6. Pile-up type of chip formation in the solution-strengthened material at a small groove depth: (a) SEM photograph; (b) LOM photograph.

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lamellae observed by SEM are seen in cross section to be separated by cracks instead of shear zones. The cracks appear to have been nucleated at the surface, subsequently propagating towards the tip, leaving the indi- vidual lamellae only attached along the thin heavily sheared zone. The propagation length of these cracks is little affected by variations in groove depth and for larger grooves chip formation takes place by a combination of crack propagation and shear along closely spaced bands (Fig. 7). The ratio of crack length to shear length decreases rapidly with increasing groove depth, causing the chips from large grooves to appear almost homoge- neously sheared (Fig. 8).

The solution-strengthened material deviates somewhat in behaviour compared with the overaged and annealed materials in two respects. For the very largest groove sizes produced, partial and irregular segmentation is observed (see Fig. 8(a)). At the other end of the scale, this is the only

Fig. 7. Pile-up type of chip formation caused by combined crack formation and shear in the solution-strengthened material at a medium groove depth: (a) SEM photograph; (b)

Fig. 8. Pile-up type of chip formation caused predominantly by shear in the solution- strengthened material at a large groove depth: (a) SEM photograph; (b) LOM photograph.

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material in which shallow grooves can be formed without any observable chip formation.

Significant differences are also observed when comparing the groove bottom topographies of the different materials. The precipitation-strength- ened, overaged and annealed materials are generally characterized by a smooth groove bottom (Fig. 9(a)), although some scale formation is usually found on the entrance side of the groove (see Fig. 9(b)). In the solution- strengthened material, scale formation is observed along the whole groove and the topography is much rougher in comparison with the other three materials (Fig. 10(a)). This type of surface structure is usually considered as an indication of built-up edge formation in machining [ 26, 271, but the scales can also be the result of sliding adhesive interaction at the interface of the truncated flank of the abrasive tip. The scale topography becomes extremely rough in the extreme cases mation in this material (see Fig. lO( b)).

of grooves formed without chip for-

(a) (b) Fig. 9. Surface topography of groove bottoms: (a) centre of groove, annealed material; (b) beginning of groove, precipitation-strengthened material.

(4 (b) Fig. 10. Surface topography of groove bottoms: (a) centre of groove, solution-strength- ened material; (b) centre of small groove with no chip formation, solution-strengthened material,

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4. Discussion

The strong influence of heat treatment on gouging abrasion resistance is clearly demonstrated by the present experiments. It is also found that the relative ranking of materials with respect to specific grooving energy depends on the damage depth considered. For example, the well-known rule-of- thumb that expresses abrasion resistance as being proportional to hardness applies in the studied alloy only to groove depths corresponding to a mass loss of approximately 3 mg. Another interesting result is that the mode of chip formation is strongly influenced by material microstructure and groove depth. In the literature, similar changes in chip characteristics are reported to take place when the rake angle of the abrading tip is varied [ 281. Ac- cording to the generally accepted nomenclature in abrasion, the precipi- tation-strengthened material displays a cutting-type chip formation (i.e. a critical attack angle less than 45”) at small groove depths while chip forma- tion in the other three materials would be classified as plowing (i.e. a critical attack angle greater than 45’) [ 29, 301. However, with the present test configuration material removal is obtained in a single-pass experiment also in the case of plowing. In this section the results of the grooving experiments are discussed on the basis of the observed modes of chip formation with special emphasis on microstructural effects.

4.1. The three zones of deformation The plastic work expended in chip formation is localized to three

deformation zones, usually referred to as the primary, secondary and tertiary zones [31] (Fig. 11). If no friction occurred at the interface between the abrading tip and the material, plastic deformation would be confined to the primary zone only. In this case, the chip thickness will equal the groove depth which, using the present geometry with a rake angle of -45”, cor- responds to a shear plane angle of 22.5’. (The shear plane angle is a param- eter commonly used in metal cutting studies and refers to the angle that a vector extending from the tool tip to the free surface at the chip root makes with the direction of motion.) In practice, friction in the secondary zone causes the chip to slow down, thereby reducing the shear plane angle. The frictional work in the tertiary zone does not affect the mode of chip forma- tion but will contribute to the energy consumption in grooving.

Fig. 11. The primary (I), secondary (II) and tertiary (III) deformation zones in chip for- mation.

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4.2. Influence of groove depth A common characteristic of the four materials tested is that the specific

grooving energy decreases with increasing groove depth, a phenomenon known as the “size effect” in abrasion [32]. The recorded specific energy is the sum of the components in the grooving direction of the work expended in the three deformation zones. In a first approximation, it can be assumed that the shear flow stresses rP, 7, and rt in each zone are independent of groove depth. If the contact length in the secondary zone is assumed to be proportional to the groove depth, the resulting force F in the grooving direction can be expressed as

F = F, + FId + F,d2 (2)

where d is the groove depth and F,, F, and F2 are constants. The second term will be dominant since F1 includes the work performed in both the pri- mary and the secondary zones. F, refers to the deformation in the tertiary zone at the truncated flat, while the work expended at the sides of the tip will contribute to both F, and F2. For all but the largest groove depths, the parabolic term can be neglected and the specific grooving energy is seen to decrease with increasing groove depth according to

eo e=e,+ -

W

where W is the mass loss and ei and ee are constants. Although this equation is not of the same form as eqn. (l), it predicts a similar size effect. The strong influence on the work expended in the tertiary zone on specific energy readings has also been experimentally demonstrated by Malkin [ 331.

In reality, the shear flow stress values will not stay constant with changing groove depth, As a first approximation it can be assumed that the effect of a change in groove depth is only to change the chip dimensions with a scaling factor. If this is true both the length and the thickness of the primary and secondary zones will be proportional to the groove depth. This means that the volume of the two zones will vary as d*, while the area through which the heat generated in these zones is conducted is proportional to the length of the zones, i.e. to d. Heat conduction from the deformation zones thus becomes less effective with increasing groove depth, resulting in higher temperatures within the zones, The resulting thermal softening is an additional factor contributing to the size effect. This thermal effect can be indirectly observed experimentally in the case of grooving in the precipi- tation-strengthened material. Careful microhardness measurements reveal that the hardness in the shear zones is lower than in the surrounding material which is probably due to thermally assisted coalescence during chip forma- tion. The temperature effect is most pronounced in the secondary zone, where the highest temperatures are generated [34]. Consequently, rS drops faster than 7, with increasing groove depth and the shear plane angle will increase. This was experimentally confirmed by measuring the shear plane angle on metallographic cross sections of the quick-stop specimens as a

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Groove Depth (mm)

Fig. 12. Shear plane angle VS. groove depth for the precipitation-strengthened ( -) and the solution-strengthened (- - -) materials.

function of groove depth for the precipitation- and solution-strengthened materials (Fig. 12). The change in shear plane angle will in itself contribute to the size effect, since it is only the components of the work in the grooving direction that make up the recorded energy consumption in the pendulum test.

4.3. Chip formation The metallographic examination of quick-stop specimens demonstrates

that chip formation is always an inhomogeneous deformation process. A constitutive equation for the plastic flow stress r which includes the effect of strain 7, strain rate + and temperature T has been proposed by Sundara- rajan [35]:

r = Ci(1 - C,T)y” + Cs In Q

+ C4nlm+ exp r 0

(4)

where all other parameters appearing in the equation are constants. The first term of the equation corresponds to the athermal component of the flow stress, the second term represents the thermally activated motion of disloca- tions past short-range obstacles and the third term describes the viscous drag experienced by a dislocation when moving through the lattice.

Using newly developed equipment for high strain rate testing it has been demonstrated that the constants C,, C,, C3 and n can be experimen- tally determined [36,37]. Although no data are yet available for the alumin- ium alloy studied in the present investigation, it is generally observed that the second term in eqn. (4) above becomes practically constant at f values above 10’ s-i. The third term, in contrast, is only important at very high strain rates (f > lo3 s-l) [38]. Consequently, the strain rate dependence is small at intermediate strain rates (10’ s-l < $ < lo3 s-l). In this region the

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flow stress is a linearly decreasing function of temperature while the strain dependence is described by a power function with an exponent less than unity. This suggests that the effect of thermal softening is stronger than that of strain hardening and, once flow is initiated, the concomitant tem- perature rise will cause localized softening. The material thus continues to shear in the same region, the deformation spreading sideways to form a shear band. Eventually, the stress within the band will relax owing to the continuous motion of the chip up along the rake face of the tip. At this point, shear will be activated on a new plane starting from the edge of the tip and the process will repeat itself. The characteristic lamellar structure of all chips can thus be explained as arising from a thermal instability phe- nomenon. A more detailed discussion on the basic processes in chip forma- tion can be found in the metal cutting literature [ 39, 401.

As already discussed, the shear plane angle is determined by the ratio of 7, to 7,. Consequently, a material with pronounced thermal softening should be characterized by a high shear plane angle because of the reduced resistance to plastic flow in the secondary zone. Transmission electron microscopy studies of quick-stop specimens have demonstrated that dynamic recrystallization takes place in the secondary and tertiary zones in machining [ 11, 411 and the same is expected to apply in abrasion. The precipitation- strengthened material, having the highest yield strength, will be the most susceptible to thermal softening, especially since dynamic recrystallization may be accompanied by dynamic coalescence of precipitates in this material. This observation is in good agreement with the shear plane angle measure- ments of Fig. 9.

The alternative to plastic shear as a means of relaxing the stresses in the primary shear zone is crack nucleation and propagation [ 421. In the present experiments the observation of cracks nucleated at the chip root and propa- gating along the primary shear zone can also be explained as a temperature effect. The two end points of the primary zone both represent points of high stress concentration. At the edge of the tip the temperature is high and shear proceeds by plastic deformation. At the chip root on the surface the temperature is much lower and the high stress is instead relaxed by crack nucleation. The effect of increasing groove depth is to increase the tern- perature in the primary zone and thereby to reduce the extent of crack propagation. The fact that very little crack formation is observed in the precipitation-strengthened material may be explained by the high shear plane angle. From geometry only it is seen that this will result in a higher tem- perature at the chip root.

In general, the lamellar spacing is expected to increase with increasing groove depth, since the higher temperature in the primary zone will in- crease the thickness of the shear zone separating the individual lamellae. This effect is strongest in the precipitation-strengthened material owing to its high sensitivity to thermal softening, which explains the observed gradual transition from fine lamellar to serrated chip formation in this material.

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The transition to segmented chip formation, in contrast, is sudden and must correspond to a drastic change in material behaviour within the deformation zones. The observed extrusion of each segment along two well-defined shear zones of extremely high strain density suggests adiabatic shear conditions (Fig. 13(a)). The fact that segmented chip formation is only obtained in the precipitation-strengthened material can be explained by the sudden thermal softening that occurs when the solution temperature is reached in the primary zone. Dynamic dissolution is not unlikely in spite of the very short duration of the shearing event since material transport is provided by plastic flow and diffusion is not required. This hypothesis that dynamic dissolution occurs is supported by the results of microhardness measurements on cross sections of quick-stop specimens. These measure- ments show that the microhardness within the shear zones coincides with that of the annealed material while the microhardness of the segments is the same as that of the unaffected material.

One effect of the extremely localized softening taking place (see Fig. 13(b)) is that very high strain rates are generated and the viscous drag com- ponent of eqn. (4) becomes important. As discussed by Sullivan et al. [43], the sudden increase in 7, when q exceeds lo4 s’-’ in the secondary zone will cause stick on the rake face and the shear in the primary zone will be halted. Stick conditions will continue to prevail until r, has decreased and a new segment can be formed. Consequently, segmented chip formation cor- responds to a stick-slip type of contact.

(b) Fig. 13. Localized deformation in the primary deformation zone at large groove depths in the precipitation-strengthened material: (a) in the centre of the primary zone; (b) at the intersection with the outer surface.

4.4. Gouging abmsion resistance The recorded specific grooving energy is composed of three compo-

nents, each one corresponding to one of the three zones of deformation. The component from the primary zone is, however, the dominant com- ponent at all but the smallest groove sizes. At small groove depths the temperature rise in the primary zone is comparatively small and it is not

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surprising that the materials approximately rank according to room tem- perature yield strength in the small damage size range. With increasing groove depth dynamic recrystallization and dynamic coalescence are initiated. As previously discussed, the sensitivity to these thermal softening mechanisms increases with decreasing precipitate size in the studied alloy. The more finely dispersed the precipitates are, the more localized the shear becomes and the energy consumption is reduced. Consequently, the highest specific energy values are recorded for the solution-strengthened material at large groove depths, followed by the precipitation-treated material in order of decreasing precipitate size.

The precipitation-strengthened material represents an extreme case of localized softening at large groove depths. It is interesting to observe from the discussion above that the solution- and precipitation-strengthened mate- rials are sheared in the same microstructural condition. A lower specific energy is, however, recorded in the precipitation-strengthened material since higher temperatures are generated in the primary shear zones owing to the extreme shear localization in this material.

4.5. Conclusions The experiments performed clearly illustrate the importance of con-

sidering the damage size when comparing the abrasion resistance of mate- rials. The results suggest that materials, in the first approximation, rank according to yield strength at small damage sizes. With increasing groove size, the thermal stability of the material becomes increasingly important and solution strengthening is more effective in providing wear resistance than precipitation strengthening.

Great care must be taken in extrapolating wear data from one alloy system to another. Some of the trends observed seem, however, likely to apply also in more microstructurally complex materials. The steel grades usually used for abrasion resistance are characterized by segmented chip formation at large groove depths [19]. This is in agreement with the above discussion since these materials are precipitation treated by tempering or controlled rolling. The groove size corresponding to the transition to seg- mented chip formation will depend on the thermal stability of the precip- itates as well as the thermal properties of the material.

One characteristic of the pendulum test is that only the work per- formed in the direction of grooving is recorded. The work expended in the perpendicular direction will in a real case affect the penetration depth of the abrasive tip. A material with a high shear angle will thus be somewhat underestimated as to abrasion resistance, but the effect is usually small because of the low values of the shear plane angles in abrasion.

References

1 M. A. Moore, Abrasive wear. In D. A. Rigney (ed.), Fundamentals of Friction and Wear of Materials, American Society of Mechanical Engineers, New York, 1980, p. 73.

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2 G. R. Davies,Proc., Inst. Mech. Eng., London, 180 (1965 - 1966) 237. 3 R. C. D. Richardson, Wear, 11 (1968) 245. 4 R. C. D. Richardson, J. Agric. Eng. Res., 12 (1967) 22. 5 K.-H. Zum Gahr, Abrasive wear on metallic materials. In E. Hornbogen and K.-H.

Zum Gahr (eds.), Metallurgical Aspects of Wear, Deutsche Gesellschafte fir Metall- kunde, Stuttgart, 1981, p. 73.

6 E. Rabinowicz, Friction and Wear of Materials, Wiley, London, 1966. 7 K.-H. Zum Gahr, 2. Metallkd., 69 (1978) 312. 8 K.-H. Zum Gahr, 2. Metallkd., 69 (1978) 643. 9 A. G. Evans and T. R. Wilshaw, Acta Metall., 24 (1976) 939.

10 M. A. Moore, Wear, 17 (1971) 51. 11 S. Hogmark, S. Soderberg and 0. Vingsbo, Proc. 3rd Znt. Conf. on Mechanical Behau-

iour of Materials, Cambridge, August 20 - 24, 1979, p. 621. 12 U. Bryggman, S. Hogmark and 0. Vingsbo, Proc. 2nd Znt. Conf. on Wear of Materials,

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