Volume 07 - Powder Metal Technologies and Applications Part 16 doc

Schlieper, Influence of Mechanical Surface Treatments on the Fatigue Properties of Sintered Steels under Constant and Variable Stress Loading, Modern Developments in Powder Metallurgy,

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Fatigue and Fracture Control for Powder Metallurgy Components *

Randall M German and Richard A Queeney, The Pennsylvania State University

Porosity Effects

An inherent physical characteristic of P/M materials is the presence of pores The role of porosity in determining fatigue endurance in powdered metals is akin to that of porosity that is induced through metal solidification in casting or welding However, the porosity that is characteristic of sintered powdered metals, and of those materials subsequently deformation processed, may differ in character and influence from solidification porosity In all cases, porosity catches the attention of the design engineer because it immediately conjures up images of classical stress concentrators In addition, porosity featuring sharp re-entrant corners, a possibility for marginally equilibrium-sintered powder particle boundaries, may be more accurately viewed as crack precursors Pore geometry can be altered by modifications to the sintering cycle, such as a longer hold time or higher temperature, wherein the smoother pores improve strength, fatigue life, and fracture resistance

Ductility is sensitive to the pore structure, but impact, fracture, and fatigue behavior have the greatest sensitivities In general, dynamic strength responses are the most sensitive Even in those materials possessing full density, inferior properties can occur due to microstructural defects Recent applications have pushed P/M into very demanding applications, a good example being the automotive connecting rod Such rods are formed by hot forging a porous P/M preform using an alloy of Fe-2Cu-0.8C They weigh nearly 650 g The ultimate tensile strength is 825 MPa (120 ksi), with a yield strength of 550 MPa (80 ksi) and fatigue endurance limit of 255 MPa (37 ksi) With expansion of P/M fabrication technology into dynamically loaded components, there arise performance limitations associated with fracture and fatigue For static tensile properties, there is a good basis for predicting the effect of residual porosity on strength (Ref 2, 3, 4, 5) Recent research has had an emphasis on fatigue and fracture behavior to fill the database void

Porosity is especially important to the high-cycle fatigue life (Ref 6, 7, 8, 9) Pores play a role in both crack initiation and propagation, typically increasing the threshold stress intensity for crack initiation but lowering the resistance to crack propagation (Ref 10) For an alloy of Fe-2Ni-0.8C the fatigue endurance limit at 107 reverse cycles is between 200 and 250 MPa at a density of 7.1 g/cm3 (approximately 10% porosity) That is approximately 35% of the tensile strength, and many porous injection-molded materials exhibit similar ratios of fatigue strength to tensile strength Closed pores are less detrimental than those that are interconnected and open to the external surface Surface pores act as the preferred site of fatigue crack initiation by acting as stress concentrators (Ref 11) The simplest view of a pore would be that of a spherical hole embedded in a continuous matrix whose response parameters are identical with that of a fully dense form of exactly the same metallurgical state Since the maximum, most positive, principal stress is the controlling load parameter for high-cycle fatigue endurance, the embedded pore in a tensile field is a reasonably approximate model for a stressed porous metal without interacting stress fields (one of less than 10% porosity) The local pore-dominated stress field can be expected to control local events such as crack initiation and threshold stress intensity values ( Kth) For an embedded pore, the local stress fields are related to the far-field (applied) design stresses (Ref 12):

(Eq 1)

Note that, for the average steel, v = 0.30, and the local stresses are magnified by a factor of about 2 The concentrated stress field does not persist far beyond the embedded pore, being reduced to 105% of the design value at a distance of 2a (where a

is the pore radius) from the pore center, thus possibly exerting little influence on far field parameters

If the pore features sharp re-entrant corners, the elastic concentrated stresses are more accurately predicted from an elliptical hole example However, these same pores are unlikely to exhibit smooth ellipsoidal morphologies, and their concentration effects can be more usefully predicted by the proportionality (Ref 13):

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(Eq 2)

Here, a is the major pore length normal to the maximum principal stress direction, and is the radius of curvature at the sharp

corner Suffice to say that the calculated stress concentration value can be large for a nonspheroidized pore

Due to the stress-raising properties of pores, and the fact that most fatigue failures originate at free surfaces, treatments aimed

at surface densification and serendipitous surface strengthening (e.g., coining, shot peening, ausrolling) raise the fatigue endurance limit (Ref 14, 15, 16) As a consequence, the current models for fatigue response in porous sintered materials have

a major dependence on the pore microstructure: the models address the total porosity, alloying homogeneity in near-pore regions, pore size, pore shape, and interpore separation distance (Ref 6, 17, 18, 19, 20, 21, 22, 23, 24) Round pores provide improved resistance to crack propagation Pores act as linkage sites through which cracks can propagate Microstructure-based fatigue models for ferrous alloys have had to address seemingly contradictory porosity effects: round pores retard stable fatigue crack propagation but increase crack extension growth rates by contributing linkage sites While extant theories successfully explain some porosity effects on crack propagation, no total predictive model has been created that embraces microstructure effects (Ref 18, 20, 22) The presence of pores in the reversed plastic zone that is the site of propagating crack damage does not lend itself to facile analysis

Pore structure changes are obtainable through processing and material variations: powder variables (particle size distribution, particle shape); compaction variables (type of lubricant, amount of lubricant, tool motions, maximum pressure); sintering variables (hold time, maximum temperature, atmosphere); and postsintering treatments Figure 1 compares the pore structure

in two sintered stainless steels to emphasize this point Smaller particles result in faster sintering and higher strengths and toughness Associated with the smaller particle sizes are smaller final pores At lower densities (around 6.6 g/cm3 or 18% porosity in steels), a high-relative content of small particles is beneficial to fatigue resistance, while at higher densities (over 7.1 g/cm3 or less than 10% porosity), larger particles prove beneficial The difference relates to the ligament size between pores, which is the determinant of fatigue There are three pore microstructure parameters relevant to the fatigue resistance of porous P/M materials: pore size, pore curvature, and pore spacing These largely reflect the role of stress concentration with respect to the advancing fatigue crack, as noted above (Ref 6) Thus, lower porosity contents, smoother (rounder) pores, and wider interpore separations increase the fatigue endurance strength

Fig 1 Two stainless steels fabricated by P/M, as demonstrations of the microstructure variations possible by

tailoring the powder, compaction, and sintering variables (a) A high-porosity microstructure useful for filtration, formed by press and sinter 1000× (b) A closed-porosity, high-density microstructure useful for mechanical components, formed by injection molding and high-temperature sintering 200×

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Fatigue cracks have been successfully analyzed with regard to their propagation response, and the cyclic growth of a crack can be predicted by the modified Paris growth law:

(Eq 3)

where the material parameters A and n must be experimentally determined for any given material microstate, including

different distributions of porosity The stress analytical variable Keff, the effective stress intensity range, factors out that portion of the total stress range that relieves the stresses holding the crack flanks closed, the opening stress range op, leaving only the stress range component that displaces the crack faces relative to each other In the presence of appreciable levels of mean stress, the relation between load design parameters and fatigue crack propagation rates is given by:

(Eq 4)

Here, the total stress-intensity factor range K is the load variable, but the fatigue ratio R = min/ max and the fracture

toughness Kc, or KIc for a low-ductility sintered member, enter into the fatigue crack propagation response, as do the material constants C and n

Regardless of which of the three fatigue crack propagation relations are relevant in a particular service context, their collective utility lies in predicting member lifetimes, or precise segments of that lifetime Thus, in the case of the Paris law:

(Eq 5)

The calculated endurance cycles to failure, N, can be from an initial flaw size a0 that may be the minimum detectable to final fracture at a = ac, where:

Again, the fracture toughness of the material plays a role in determining structural endurance The fracture resistance KIc, or

KR in the case of more ductile sintered materials (Ref 25), is known to be porosity sensitive (Ref 26) As a first estimate, the sensitivity is about a 100 MPa gain in toughness per percentage point of porosity reduction in quenched and tempered steels Copper-infiltrated steels have toughnesses that run from 40 to 50% those of wrought steels (Ref 27, 28) However, their static strengths are equivalent, reflecting the inability of the steel skeleton to absorb the same level of strain energy release as a fully dense body of the same material

The material parameters in the modified Paris equation (Eq 3) are sensitive to the porosity state, a not unexpected result since

the coefficient and exponent are related to the plastic zone size rp ahead of the advancing crack, the only region of irreversible deformation The coefficient A in Eq 3 increases with increasing porosity fraction (Ref 29), resulting in faster growth rates for

comparable Keff values With higher coefficient measures but constant exponent n values, the threshold stress intensity range Kth, below which no crack growth is thought to occur, also decreases with porosity increase However, Kth values for fully dense materials are sufficiently low that it is not at all clear in what way they could be successfully employed in design practice if service stresses are to be set at appreciable fractions of the yield or tensile strength The reversed plastic zone size is given by (Ref 30):

(Eq 7)

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When the plastic zone size is of the same size as the average pore diameter, it is effectively enlarged by the high strain field in

the vicinity of the pore The net result of the enlarged effective plastic zone is reflected in higher values of the exponent n

(Ref 11), and the enlarged zone is even more pronounced in the response of short cracks driven by the locally raised stress/strain fields associated with design stress concentrators (Ref 15) Copper-infiltrated steels (Ref 27, 28) are as fatigue and fracture resistant as fully dense wrought medium-strength steels Although the performance standards are not up to those

of fully dense martensitic steels, copper infiltration represents a considerable cost savings over forging to full density, while maintaining the cost advantage inherent in press and sinter P/M

Typical fatigue endurance limits (or fatigue strengths at about 107 cycles) are collected in Table 8 for several P/M materials This compilation includes several ferrous alloys, reflecting the high interest in P/M fatigue for automotive applications There are variations in density, alloying, and sintering cycles to show the relative effects on sintered properties Systematic testing

of various alloys has shown that slight changes in the particle size distribution or alloying homogeneity can affect these properties For example, in the Fe-2Ni-0.8C alloy system, shifts in just the iron powder source lead to ±33 MPa (±4.8 ksi) variations in the fatigue endurance strength Accordingly, the values in Table 8 are for relative ranking purposes only and cannot be used as an accurate basis for design of fatigue-sensitive components

Table 8 Representative P/M materials, processing cycles, and fatigue endurance limit

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P+S, pressed and sintered; PIM, powder injection molded and sintered; HIP, hot isostatically pressed; HT, heat treated; SP, shot peened;

DP + DS, double pressed and double sintered; PF, powder forged

Fracture studies of P/M alloys are usually restricted to impact testing, and often this is performed in the unnotched condition because of the low toughness of porous materials Fracture toughness measurements on P/M materials are relatively rare

Tables 9 and 10 summarize some prior findings Table 9 demonstrates the density effect on tensile properties and KIc, for an

Fe-Ni-Mo-C steel Note that as the porosity decreases, the strength essentially doubles, ductility increases substantially (a nearly tenfold gain), the impact energy goes up by a factor of 4, and fracture toughness increases threefold Table 10 collects several examples of the mechanical properties of ferrous alloys and one titanium alloy as representative values obtainable via P/M Clearly, porosity negatively affects the fracture toughness in most materials The fracture toughness of P/M steels is essentially a linear function of density (Ref 18), with sensitivities of about 100 MPa gain in toughness per percentage point of porosity reduction In low-density P/M materials, the fracture crack propagation is rapid because the pore structure amplifies the stress and provides an easy path At low porosity levels, an advancing fracture crack can be blunted by the pores, effectively forming microcracks that improve toughness (Ref 31)

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Table 9 Mechanical properties of pressed and sintered Fe-1.8Ni-0.5Mo-0.5C P/M compacts

Double pressed and double sintered, 1120 °C, h, tested as-sintered

Porosity, % Density,

g/cm3

Elastic modulus,

GPa

Yield strength,

MPa

Ultimate strength,

MPa

Elongation,

%

Notched impact

energy, J

Fracture toughness MPa

Table 10 Representative P/M materials, processing cycles, and fracture toughness

Composition, wt% Processing Density,

g/cm3

Fracture toughness, MPa Fe-4.4Cr-9.2Co-7.2V-3.7Mo-9.2W-2.7C P+S, 1150 °C, 1 h 8.1 13

P+S, pressed and sintered; HIP, hot isostatically pressed; DP + DS, double pressed and double sintered; HF, hot forged

References cited in this section

2 B Karlsson and I Bertilsson, Mechanical Properties of Sintered Steels, Scand J Metall., Vol 11, 1982, p

267-275

3 G.F Bocchini, The Influences of Porosity on the Characteristics of Sintered Materials, Rev Powder Met Phys

Ceram., Vol 2, 1985, p 313-359

4 R Haynes, The Mechanical Behavior of Sintered Metals, Rev Deform Behav Mater., Vol 3, 1981, p 1-101

5 S.H Danninger, G Jangg, B Weiss, and R Stickler, Microstructure and Mechanical Pr operties of Sintered

Iron, Part 1: Basic Considerations and Review of Literature, Powder Met Inter., Vol 25, 1993, p 111-117

6 B Weiss, R Stickler, and H Sychra, High-Cycle Fatigue Behaviour of Iron-Base PM Materials, Metal Powder

Report, Vol 45, 1990, p 187-192

7 R Haynes, Fatigue Behaviour of Sintered Metals and Alloys, Powder Met., Vol 13, 1970, p 465-510

8 S Oki, T Akiyama, and K Shoji, Fatigue Fracture Behavior of Sintered Carbon Steels, J Japan Soc Powder

Met., Vol 30, 1983, p 229-234

9 W.B James and R.C O'Brien, High Performance Ferrous P/M Materials: The Effect of Alloying Method on

Dynamic Properties, Progress in Powder Metallurgy, Vol 42, Metal Powder Industries Federation, 1986, p

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353-372

10 H Danninger, G Jangg, B Weiss, and R Stickler, The Influence of Porosity on Static and Dynamic Properties

of P/M Iron, PM into the 1990's, Vol 1, Proceedings of the World Conference on Powder Metallurgy, Institute

of Materials, London, 1990, p 433-439

11 J Holmes and R.A Queeney, Fatigue Crack Initiation in a Porous Steel, Powder Met., Vol 28, 1985, p 231-235

12 S Timoshenko and J.N Goodier, Theory of Elasticity, McGraw-Hill, 1951, p 359-362

13 F.A McClintock and A.S Argon, Mechanical Behavior of Materials, Addison Wesley, 1966, p 412

14 C.M Sonsino, F Muller, V Arnhold, and G Schlieper, Influence of Mechanical Surface Treatments on the

Fatigue Properties of Sintered Steels under Constant and Variable Stress Loading, Modern Developments in

Powder Metallurgy, Vol 21, Metal Powder Industries Federation, 1988, p 55-66

15 C.M Sonsino, G Schlieper, and W.J Huppmann, How to Improve the Fatigue Properties of Sintered Steels by

Combined Mechanical and Thermal Treatments, Modern Developments in Powder Metallurgy, Vol 16, Metal

Powder Industries Federation, 1985, p 33-48

16 J.H Lange, M.F Amateau, N Sonti, and R.A Queeney, Rolling Contact Fatigue in Ausrolled 1%C 9310 Steel,

Inter J Fatigue, Vol 16, 1994, p 281-286

17 H Kuroki and Y Tokunaga, Effect of Density and Pore Shape on Impact Properties of Sintered Iron, Inter J

Powder Met Powder Tech., Vol 21, 1985, p 131-137

18 F.J Esper and C.M Sonsino, Fatigue Design for PM Components, European Powder Metallurgy Association,

Shrewsbury, UK, 1994

19 K.D Christian and R.M German, Relation between Pore Structure and Fatigue Behavior in Sintered

Iron-Copper-Carbon, Inter J Powder Met., Vol 31, 1995, p 51-61

20 I Bertilsson, B Karlsson, and J Wasen, Fatigue Properties of Sintered Steels, Modern Developments in

21 R.C O'Brien, Impact and Fatigue Characterization of Selected Ferrous P/M Materials, Progress in Powder

Metallurgy, Vol 43, Metal Powder Industries Federation, 1987, p 749-775

22 P.S Dasgupta and R.A Queeney, Fatigue Crack Growth Rates in a Porous Metal, Inter J Fatigue, Vol 3,

1980, p 113-117

23 T Prucher, Fatigue Life as a Function of the Mean Free Path between Inclusions, Modern Developments in

24 K.D Christian, R.M German, and A.S Paulson, Statistical Analysis of Density and Particle Size Influences on

Microstructural and Fatigue Properties of a Ferrous Alloy, Modern Developments in Powder Metallurgy, Vol

21, Metal Powder Industries Federation, 1988, p 23-39

25 I.J Mellanby and J.R Moon, The Fatigue Properties of Heat-Treatable Low Alloy Powder Metallurgy Steels,

Modern Developments in Powder Metallurgy, Vol 18, Metal Powder Industries Federation, 1988, p 183-195

26 J.T Barnby, D.C Ghosh, and K Dinsdale, Fracture Resistance of a Range of Steels, Powder Met., Vol 16,

1973, pp 55-71

27 E Klar, D.F Berry, P.K Samal, J.J Lewandowski, and J.D Rigney, Fracture Toughness and Fatigue Crack

Growth Response of Copper Infiltrated Steels, Inter J Powder Met., Vol 31, 1995, p 316-324

28 R.A Queeney, Fatigue and Fracture Response of Metal-Infiltrated Sintered Powder Metals, Proceedings of

ICM3, Vol 3, Pergamon Press, Oxford, 1979, p 373-381

29 D.A Gerard and D.A Koss, The Influence of Porosity on Short Fatigue Crack Growth at Large Strain

Amplitudes, Inter J Fatigue, Vol 13, 1991, p 345-352

30 P.C Paris, Fatigue An Interdisciplinary Approach, Syracuse University Press, 1964, p 107-117

31 W Pompe, G Leitner, K Wetzig, G Zies, and W Grabner, Crack Propagation and Processes Near Crack Tip

of Metallic Sintered Materials, Powder Met., Vol 27, 1984, p 45-51

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Other Factors Determining Fatigue and Fracture Resistance

It is well established that porosity is the major detriment to fatigue life for P/M materials Beyond porosity, the sintered microstructure is a factor Even in full-density materials fabricated by hot isostatic pressing, microstructure has a role Weak links in the microstructure become evident during fracture For porous structures, these weak links prove to be microstructural inhomogeneities, typically resulting from incomplete diffusional homogenization Often powders (such as iron, nickel, and graphite) are mixed in the compaction stage During heating the intent is for the mixed powders to homogenize to form a uniform microstructure, but this often is inhibited by too short a hold time at the peak temperature Consequently, the alloying elements are poorly distributed and give point-to-point composition and microstructure changes, which are especially evident in postsintering heat treatment response Accordingly, during fatigue or fracture, the weak links become the preferred failure paths Most notable are the negative effects from inadequate homogenization of carbon to ensure uniform strength (Ref 18)

Little is known about the sensitivity of fatigue and fracture to loading conditions for P/M material Table 11 compares the 2 ×

106 fatigue endurance strengths for Fe-1.5Cu-0.6C at 7.1 g/cm3 density using bending and axial fatigue tests The table also

includes a comparison with two loading stress ratios (half-cycle and fully reversed, R = -1) and two notch conditions

(unnotched and notched) (Ref 18) In these cases the unnotched loading shows little sensitivity to axial versus bending fatigue, but a large sensitivity is evident in the presence of notches The notch sensitivity factor is reported to range between 0.32 and 0.43 for many of the common pressed and sintered P/M alloys (Ref 32)

Table 11 Fatigue properties of Fe-1.5Cu-0.6C

Note: Sintered to 7.2 g/cm3 at 1120 °C for 30 min Elastic modulus, 153 GPa; yield strength, 418 MPa; tensile strength, 483 MPa Source:

Full-density materials also suffer from residual microstructure artifacts that degrade the microstructure In hot isostatically compacted powders, the achievement of 100% density is still insufficient to guarantee competitive fracture toughness, fatigue life, or even impact toughness Thermally induced porosity is a subtle problem in many full-density P/M products After consolidation the material is pore free, but it may contain small quantities of adsorbed gas Once the product is put into high-temperature heat treatment or service, this residual gas precipitates to form pores if there is no compressive stress In hot isostatically pressed titanium alloys, gas precipitation reportedly gives a 10 to 20% decrement in fatigue endurance strength (Ref 33)

Another difficulty rests in slight contaminants located on the interfaces that were previously particle surfaces, a feature termed prior particle boundary decorations Figure 2 shows such decorations in a fully densified P/M steel Improper powder

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handling or cleaning prior to consolidation are the primary detriments These contaminants remain on the powder interfaces, even though the structure is fully densified Consequently, a small contamination film runs throughout the structure, providing an easy fracture path that is often traced to a trivial impurity level The fracture path is along the prior particle boundaries and has a characteristic morphology, as shown in Fig 3 In hot isostatically pressed Ti-6Al-4V there is substantial fatigue life improvement due to removal of the contaminant, with a change from 450 MPa endurance strength to 600 MPa due to powder cleaning prior to consolidation (Ref 34)

Fig 2 Prior particle boundary precipitates formed on a hot isostatically pressed steel as the result of contamination

during powder fabrication 500×

Fig 3 A fracture surface showing preferential failure along prior particle boundaries 150×

One option for limiting the detrimental effects from prior particle boundary decorations is to forge the structure after hot isostatic pressing, a process often used in producing aerospace structures to ensure ultimate reliability The forging operation upsets the microstructure and breaks apart the continuous films of contamination The alternative is to resort to clean handling and processing, where the powder is produced by rapid solidification and kept under inert conditions during handling These steps, which minimize segregation and contamination, are employed in the production of aerospace components, microelectronic structures, and high-performance filters

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32 A.F Kravic, The Fatigue Properties of Sintered Iron and Steel, Inter J Powder Met., Vol 3 (No 2), 1967, p

7-13

33 R.L Dreshfield and R.V Miner, Effects of Thermally Induced Porosity on an as- HIP Powder Metallurgy

Superalloy, Powder Met Inter., Vol 12, 1980, p 83-87

34 F.H Froes and C Suryanarayana, Powder Processing of Titanium Alloys, Reviews in Particulate Materials,

Vol 1, A Bose, R.M German, and A Lawley, Ed., Metal Powder Industries Federation, 1993, p 223-276

Steps to Improve Fatigue and Fracture Resistance

Surface pores are particularly detrimental to sintered materials with respect to fatigue life (Ref 35) Accordingly, carbonitriding and other surface strengthening and sealing treatments are most useful Common treatments include shot peening, case hardening, repressing and resintering, coining, sizing, surface ausrolling, and postsintering heat treatments For example, in pressed and sintered ferrous alloys, the endurance limit can be increased on small cross sections (6 by 6 mm) by

at least 20% through shot peening Carbonitriding is even more effective and can double the fatigue endurance limit A typical carbonitride cycle involves heating to 940 °C in a mixture of ammonia and carbon dioxidefor 4 h to form a 0.5 mm deep carbon-rich layer Surface grinding is another approach to improved fatigue strength The larger the cross section of the material, the less benefit possible from surface treatments, because bulk material states will dominate mechanical response The double press and double sinter approach was largely the only viable option for improving density and strength in traditional press and sinter P/M This is more costly and involves extra tooling A newly employed technique for improved fatigue life and fracture toughness in pressed and sintered ferrous alloys is to sinter at higher temperatures The typical sintering temperature for steels is about 1120 °C, largely because of conveyor belt limitations in the furnaces New materials

of construction (ceramic belts) and new conveyor mechanisms (pusher plate and walking beam designs) allow temperature processing regimes Additionally, vacuum sintering usually is not limited in temperature, so it is viable for high-performance components There is more sintering densification at the higher temperatures, so the density gain alone improves properties Induced changes in the pore shape and size also improve fracture and fatigue properties Several examples of the property gains are evident in Table 8 In a comparison of density gains versus sintering temperature effects, it is usually concluded that a change from 1120 to 1280 °C is equivalent to a density gain from 7.1 to 7.4 g/cm3 in terms of both fracture toughness and fatigue endurance strength

higher-For small components the surface treatments are most useful, because the compressive forces extend through a major portion

of the microstructure However, for the porous materials with large cross sections, the need is to sinter at higher temperatures

to improve fatigue and fracture Further, designs that minimize density gradients will assist in minimizing fatigue failure The high-density regions have a higher fatigue strength, and the difference in strength with density often results in failure at the interface between high- and low-density regions For fatigue-sensitive components, the tolerable range of densities is less than 0.05 g/cm3 within the structure As with all fatigue-sensitive components, consideration must be given to surface finishing and processing optimization The keys to improved performance are reduction in the total porosity, elimination of segregation and contamination, and manipulation of the pore microstructure

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There are some unique design opportunities where the microstructure of P/M materials offers a fatigue advantage in spite of the porosity Because many alloys are formed from mixed powders, there is often an enrichment of the alloying addition near pores The resulting higher strength aids local strength and fatigue life Surprisingly, when properly exploited this effect gives higher fatigue and fracture strength to porous structures formed from mixed elemental powders, compared to those formed from prealloyed powders (Ref 36) Unfortunately, pores reduce the elastic modulus, strength, ductility, hardness, and other mechanical properties, so a high-density structure usually proves most successful in fatigue-sensitive applications Indeed, the high sensitivity to porosity mandates that porosity be tightly controlled (density held within 0.05 g/cm3) in regions of high stress concentrations At a given strength level the P/M steels exhibit less notch sensitivity than wrought steels, because pores inherently act to blunt cracks and redistribute the load, especially under complex loading Open pores, which dominate the sintered microstructure for densities below 92% of theoretical, retard fatigue crack propagation Surface densification of P/M steels from shot peening, sizing, coining, surface rolling, or carbonitriding all prove beneficial in improving fatigue strength because of pore closure and surface compressive stresses

35 J.M Wheatley and G.C Smith, The Fatigue Strength of Sintered Iron, Powder Met., Vol 6, 1963, p 141-153

36 U Engstrom, C Lindberg, and I Tengzelius, Powders and Processes for High Performance PM Steels, Powder

Met., Vol 35, 1992, p 67-72

Safety Factors for P/M Materials

Several factors contribute to scatter in the fracture and fatigue properties of P/M materials There is the obvious error in testing, where the range of highest fatigue strength to lowest fatigue strength for a single test condition may be 4% Further, there is typically a notch sensitivity and test error Consequently, a safety factor of 1.4 is often cited as appropriate for sintered P/M alloys (Ref 18) This means that a peak cyclic stress of about 70% of the fatigue endurance strength is the maximum recommended in service

Reference cited in this section

References

1 R.M German, Powder Metallurgy Science, 2nd ed., Metal Powder Industries Federation, 1994

2 B Karlsson and I Bertilsson, Mechanical Properties of Sintered Steels, Scand J Metall., Vol 11, 1982, p

267-275

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3 G.F Bocchini, The Influences of Porosity on the Characteristics of Sintered Materials, Rev Powder Met Phys

Ceram., Vol 2, 1985, p 313-359

4 R Haynes, The Mechanical Behavior of Sintered Metals, Rev Deform Behav Mater., Vol 3, 1981, p 1-101

5 S.H Danninger, G Jangg, B Weiss, and R Stickler, Microstructure and Mechanical Properties of Sintered

Iron, Part 1: Basic Considerations and Review of Literature, Powder Met Inter., Vol 25, 1993, p 111-117

6 B Weiss, R Stickler, and H Sychra, High-Cycle Fatigue Behaviour of Iron-Base PM Materials, Metal Powder

Report, Vol 45, 1990, p 187-192

7 R Haynes, Fatigue Behaviour of Sintered Metals and Alloys, Powder Met., Vol 13, 1970, p 465-510

8 S Oki, T Akiyama, and K Shoji, Fatigue Fracture Behavior of Sintered Carbon Steels, J Japan Soc Powder

Met., Vol 30, 1983, p 229-234

9 W.B James and R.C O'Brien, High Performance Ferrous P/M Materials: The Effect of Alloying Method on

Dynamic Properties, Progress in Powder Metallurgy, Vol 42, Metal Powder Industries Federation, 1986, p

353-372

10 H Danninger, G Jangg, B Weiss, and R Stickler, The Influence of Porosity on Static and Dynamic Properties

of P/M Iron, PM into the 1990's, Vol 1, Proceedings of the World Conference on Powder Metallurgy, Institute

of Materials, London, 1990, p 433-439

11 J Holmes and R.A Queeney, Fatigue Crack Initiation in a Porous Steel, Powder Met., Vol 28, 1985, p 231-235

12 S Timoshenko and J.N Goodier, Theory of Elasticity, McGraw-Hill, 1951, p 359-362

13 F.A McClintock and A.S Argon, Mechanical Behavior of Materials, Addison Wesley, 1966, p 412

14 C.M Sonsino, F Muller, V Arnhold, and G Schlieper, Influence of Mechanical Surface Treatments on the

Fatigue Properties of Sintered Steels under Constant and Variable Stress Loading, Modern Developments in

16 J.H Lange, M.F Amateau, N Sonti, and R.A Queeney, Rolling Contact Fatigue in Ausrolled 1%C 9310 Steel,

Inter J Fatigue, Vol 16, 1994, p 281-286

17 H Kuroki and Y Tokunaga, Effect of Density and Pore Shape on Impact Properties of Sintered Iron, Inter J

Powder Met Powder Tech., Vol 21, 1985, p 131-137

19 K.D Christian and R.M German, Relation between Pore Structure and Fatigue Behavior in Sintered

Iron-Copper-Carbon, Inter J Powder Met., Vol 31, 1995, p 51-61

20 I Bertilsson, B Karlsson, and J Wasen, Fatigue Properties of Sintered Steels, Modern Developments in

21 R.C O'Brien, Impact and Fatigue Characterization of Selected Ferrous P/M Materials, Progress in Powder

Metallurgy, Vol 43, Metal Powder Industries Federation, 1987, p 749-775

22 P.S Dasgupta and R.A Queeney, Fatigue Crack Growth Rates in a Porous Metal, Inter J Fatigue, Vol 3,

1980, p 113-117

23 T Prucher, Fatigue Life as a Function of the Mean Free Path between Inclusions, Modern Developments in

24 K.D Christian, R.M German, and A.S Paulson, Statistical Analysis of Density and Particle Size Influences on

Microstructural and Fatigue Properties of a Ferrous Alloy, Modern Developments in Powder Metallurgy, Vol

21, Metal Powder Industries Federation, 1988, p 23-39

25 I.J Mellanby and J.R Moon, The Fatigue Properties of Heat-Treatable Low Alloy Powder Metallurgy Steels,

Modern Developments in Powder Metallurgy, Vol 18, Metal Powder Industries Federation, 1988, p 183-195

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26 J.T Barnby, D.C Ghosh, and K Dinsdale, Fracture Resistance of a Range of Steels, Powder Met., Vol 16,

1973, pp 55-71

27 E Klar, D.F Berry, P.K Samal, J.J Lewandowski, and J.D Rigney, Fracture Toughness and Fatigue Crack

Growth Response of Copper Infiltrated Steels, Inter J Powder Met., Vol 31, 1995, p 316-324

28 R.A Queeney, Fatigue and Fracture Response of Metal-Infiltrated Sintered Powder Metals, Proceedings of

ICM3, Vol 3, Pergamon Press, Oxford, 1979, p 373-381

29 D.A Gerard and D.A Koss, The Influence of Porosity on Short Fatigue Crack Growth at Large Strain

Amplitudes, Inter J Fatigue, Vol 13, 1991, p 345-352

30 P.C Paris, Fatigue An Interdisciplinary Approach, Syracuse University Press, 1964, p 107-117

31 W Pompe, G Leitner, K Wetzig, G Zies, and W Grabner, Crack Propagation and Processes Near Crack Tip

of Metallic Sintered Materials, Powder Met., Vol 27, 1984, p 45-51

32 A.F Kravic, The Fatigue Properties of Sintered Iron and Steel, Inter J Powder Met., Vol 3 (No 2), 1967, p

7-13

33 R.L Dreshfield and R.V Miner, Effects of Thermally Induced Porosity on an as- HIP Powder Metallurgy

Superalloy, Powder Met Inter., Vol 12, 1980, p 83-87

34 F.H Froes and C Suryanarayana, Powder Processing of Titanium Alloys, Reviews in Particulate Materials,

Vol 1, A Bose, R.M German, and A Lawley, Ed., Metal Powder Industries Federation, 1993, p 223-276

35 J.M Wheatley and G.C Smith, The Fatigue Strength of Sintered Iron, Powder Met., Vol 6, 1963, p 141-153

36 U Engstrom, C Lindberg, and I Tengzelius, Powders and Processes for High Performance PM Steels, Powder

Engineering materials subjected to wear applications include different metal alloys, ceramics, and polymers Powder metallurgy is a class of processing technologies More and more engineering materials and mechanical components are being fabricated by P/M processes Any material processed by P/M that suffers wear loss is the subject of this article However, it is

a formidable task to cover all materials that can be manufactured using P/M techniques and subjected to wear during application Instead, this article focuses on effects on wear of property and microstructure factors that are unique to P/M materials, such as metal alloys and metal matrix composites

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There are two approaches to reduce wear: reduce friction between rubbing objects or increase wear resistance of the materials involved In sliding wear applications, introduction of friction reduction elements into material composition could be the most effective approach in certain specific applications Reduction of friction can be accomplished by lubricant or compositional changes of the material; however, increasing wear resistance is the subject of this article

There are numerous books including handbooks on wear behavior of materials Suggested reading materials listed in Ref 1, 2,

3, 4, and 5 discuss general theory on wear behavior of materials Metallic components fabricated using P/M technology follow the general rules and trends of those theories What P/M technology offers is special composition and microstructure that cannot be fabricated by other conventional processing methods Factors that effect wear properties and are unique to P/M materials are discussed in this article; however, a brief discussion on wear classifications will help in understanding the effects of those factors

Acknowledgement

The section "Full Density Cobalt Alloys" was adapted from the section "Appendix: P/M Cobalt-Base Wear-Resistant

Materials" in the article "Wrought and P/M Superalloys" in Properties and Selection: Irons, Steels, and High-Performance Alloys, Vol 1, ASM Handbook, 1990, p 977-980

References

1 K.-H Zum Gahr, Microstructure and Wear of Materials, Elsevier, 1987

2 I.M Hutchings, Tribology Friction and Wear of Engineering Materials, Edward Arnold, 1992

3 K.C Ludema, Friction, Wear, Lubrication, CRC Press, 1996

4 Friction, Lubrication, and Wear Technology, Vol 18, ASM Handbook, ASM International, 1992

5 K.J.A Brooks, World Directory and Handbook of Hardmetals and Hard Materials, 5th ed., International

Carbide Data, 1992

Wear Resistance of Powder Metallurgy Alloys

Zhigang Fang, Smith Tool, Smith International, Inc

Wear Classifications

Wear can be classified by four main mechanisms: adhesive, abrasion, surface fatigue, and tribochemical reactions Wear can also be classified by different types of relative movements between mating surfaces, such as sliding, rolling, oscillation, impact, and erosive wear

Adhesive wear is defined as the formation and breaking of interfacial adhesive bonds (e.g., cold-welded junctions) The tendency to incur adhesive wear depends on physical and chemical properties of the materials in contact There are five different mechanisms proposed for adhesive wear: (1) mechanical interlocking, (2) diffusion theory, (3) electronic theory, (4) adsorption theory, and (5) chemical adsorption theory Adhesive wear is considered the primary wear mechanism for sliding wear, but it is not equivalent to sliding wear because other mechanisms contribute to wear loss during relative sliding motion between two objects Common terms such as material transfer, galling, scuffing, and scoring are examples of adhesive wear, although scuffing and scoring can mean scratching by abrasive particles

Abrasive wear is the dislodging of material caused by hard particles between or embedded in the surfaces in relative motion, or by hard protuberances on one or both of the relatively moving surfaces Mechanisms responsible for abrasive wear include microploughing, microcutting, and microcracking on the surfaces of metals Abrasion is further divided into two-body or three-body abrasion Abrasion can also be described by high-stress or low-stress abrasion In general, during high-

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stress abrasion, abrasive particles are crushed, while during low-stress abrasion, the particles remain unbroken Erosion is one kind of abrasion which refers to wear caused by impact of particles The hard particles striking the surface are carried by a gas stream or entrained in a flowing liquid The terms "gouging" and "scratching" are often used to describe wear scars left by abrasive action

Surface fatigue results from repeated alternating loading of solid surfaces Most often it occurs during rolling and impact contacts, which causes cyclic surface stressing Mechanisms responsible for surface fatigue wear are undersurface crack formation and flaking material The commonly observed pits are typical of surface fatigue

Tribochemical reactions proceed by continual removal and new formation of reaction layers on the contacting surfaces, such as oxide layers formed in the presence of oxygen The reaction from rubbing surfaces can be caused by gas or liquid in the environment

Factors Affecting Wear Resistance

Root causes of any wear behavior are microstructure and chemical composition, as for any mechanical property However, wear behavior is best understood from three different perspectives: fundamental mechanical properties of the material, microstructure features, and specific wear environments and mechanisms In many cases, wear resistance of a material is largely the function of its intrinsic mechanical properties (i.e., elastic modulus, strength, hardness, and fracture toughness) Depending on wear action, different mechanical properties dominate the rate of wear loss, and different microstructure is desired However, microstructure can also affect wear resistance independently without changing chemical composition or altering basic relationships between mechanical properties and microstructure For example, large differences in wear resistance can occur at the same hardness In other words, changing microstructure features can alter the rate of wear loss of a material with or without concurrent change in common mechanical properties The relationships among wear resistance and mechanical properties and microstructure are also affected by specific wear environments (i.e., specific wear mechanisms) For example, wear resistance against sliding abrasive particles has different requirements for mechanical properties and microstructure than against impact fatigue or some other wear mechanism Wear resistance depends on interrelated mechanical, microstructural and environmental factors (Fig 1)

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Fig 1 Factors affecting wear resistance of materials

Mechanical Properties. Mechanical properties affecting wear resistance of a material include hardness, fracture toughness, elastic modulus, tensile/compressive strength, and impact fatigue strength In many cases, wear resistance is primarily a function of hardness not of toughness, especially for metal alloys which have sufficient fracture toughness The higher the hardness, the higher the resistance to abrasion by hard particles, surface plastic deformation, debonding, and microploughing For wearing systems where plastic deformation is the main mechanism and impact is not prevalent, hardness controls wear rate of the material In cases where fracture toughness of the material is relatively low, the wear application involves impact and repeated loading condition changes, and wear mechanism involves microfracturing, the role of fracture toughness in determining wear rate becomes more important As a general principle, both wear mechanisms plastic flow and fracture, can result in excavation and detachment of material Raising the material hardness will reduce plastic flow but can increase the danger of fracture Hence, it is more often that some compromise between high hardness and high fracture toughness gives the best wear resistance

Microstructure. Materials can be categorized as those having homogeneous microstructures and those having multiphase inhomogeneous materials Homogeneous P/M materials are similar to what can be achieved by traditional processing techniques, such as casting, forging, and heat treatment Homogeneous microstructure indicates that the material is composed

of one single phase or multiple phases that act like a single phase structure with respect to wear Wear resistance is primarily proportional to its hardness and fracture toughness Alloys or composites composed of at least two different phases have microstructures formed either by blending different powders together or one or more phase forms in situ The boundary between a homogeneous microstructure and multiphase structure can be gray For example, cemented tungsten carbide (WC-Co) is a two-phase composite, except for those with extremely fine grains and small mean free path (MFP) Powder metallurgy tool steel is mostly a homogeneous material, although its wear resistance depends on carbide particles in its microstructure The various carbide particles are in situ precipitates The two-phase or multiphase microstructure makes P/M technique unique with regard to tailoring microstructure to phases that cannot be created by casting, forging, or heat treatment

in order to achieve desired wear resistance Microstructure factors that affect wear include volume fraction, particle size, particle size distribution, MFP, and the ratio between MFP and abrasive particle sizes

mechanical properties such as strength and fracture toughness Wear resistance is influenced by volume fraction's influence

on strength and toughness No other direct effects of volume fraction on wear resistance have been reported In general, wear

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resistance is proportional to volume fraction of hard phase particles However, wear resistance levels off when hard phase particles reach certain volume fraction in some alloy systems and applications

mechanical properties At any given volume fraction, finer particle size results in smaller mean free path (MFP), i.e., smaller spacing between hard phase particles When MFP is sufficiently small, the finer the particle size, the more the composite resembles homogeneous material (i.e., the higher the hardness, and the higher the wear resistance) A larger spacing is likely

to cause preferential wear of the softer matrix which leads to uprooting of hard particles In such cases, larger particle size is beneficial for wear resistance Size distribution also affects wear resistance For a given average particle size, wider particle size distribution is likely to give larger MFP However, the effect of different powder size blend is strongly related to wear applications In some cases, multiple wear mechanisms are in effect Therefore, different powder particle sizes may be needed for different roles

= 4 · (1 - VV)/SV

The effects of MFP on wear resistance reflect the combined influence of hard phase volume fractions and particle sizes MFP affects basic mechanical properties including hardness, strength, and fracture toughness As a general rule of thumb, hardness and strength increase with the decrease in MFP Wear resistance is proportional to hardness in most cases; therefore, wear resistance is inversely proportional to MFP The increase of wear resistance as MFP decreases is attributed to the resistance to plastic deformation, which accounts for various wear mechanisms, especially during abrasive and adhesive wear situations However, in cases where microchipping and fracturing is responsible for material loss, an optimum MFP has to be achieved

to maximize wear resistance Microchipping and cracking resistance is proportional to MFP The balance between microchipping and local plastic deformation resistance yields optimum wear resistance Another important factor during abrasive wear is the ratio between MFP and abrasive wear particles When abrasive particle size is smaller than mean free path between hard phase particles, preferential wear of metal phase between them is dominant When abrasive particle size is larger than MFP, overall hardness of the microstructure becomes a dominant factor

borides, nitrides, oxides, and other ceramics Among carbides, tungsten carbide (WC), chromium carbides, and titanium carbide (TiC) are most widely used Vanadium carbide (VC), molybdenum carbides, and tantalum carbides are often used as additives Titanium diboride is known to be extremely wear resistant Usage of TiB2 powder has been on the rise recently Relative hardness of hard phase particles to that of abrasive particles is an important factor attributed to the interaction between the microstructure and wear environment

Wear Environment/Mechanisms. Factors to consider include mechanical loading conditions, wearing mechanisms, harshness of abrasive particles, and hardness and size/shape of abrasive particles In cases of abrasive wear, relative size

between the hard phase particles, dc, and abrasive particle da, is an important factor In general, if the dc < da, rate of wear loss

is greater than it would be if dc > da Wear mechanisms depend on the wear environment and the material In cases of sliding

wear, relative hardness between the two mating surfaces, the chemical affinity between the two mating surfaces, and abrasive particles contained in the materials are all important factors that affect the final outcome Depending on the wearing system, different mechanical properties (e.g., resistance against plastic deformation, crack formation or crack propagation) have to be considered Depending on which wear mechanism is dominant, different material properties and microstructure may be desired Other environment factors, such as corrosive fluid and potential chemical interactions, need also be considered

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Wear-Resistant P/M Alloys

Wear resistant P/M alloys include cemented tungsten carbide, P/M steels, and P/M metal-matrix composites

Cemented Tungsten Carbide

Cemented tungsten carbide and cermet are the most widely used P/M materials for wear applications Cemented tungsten carbide and cermet are used in all types of wear applications including abrasive wear, sliding wear, and erosive wear situations Cemented tungsten is a composite of tungsten carbide, WC, and cobalt metal, Co In most commercially available grades of WC-Co, cobalt contents range from 6.0 to 16.0% There are applications using as low as 3% cobalt or as high as 25% cobalt, although those are not common Tungsten carbide grain size range from submicron to 10 m Commercially available ultrafine WC powder has nanometer size grains such as NanoCarb (Nanodyne, Inc., New Jersey)

Most cemented tungsten carbide manufacturers have their own designations for different grades that are used by suppliers and customers For metal removal cutting tool applications, there are internationally accepted grade designation systems that are used as category reference only For example, the P-M-K system found in Ref 5, and the "C" system in which C-1 through C-

4 are used for abrasion resistant grades and C-5 through C-8 are used for crater and deformation resistant grades C-1 through C-4 are straight WC-Co grades, and C-5 through C-8 contain cubic carbides, such as TiC and TaC for cutting tool applications

A schematic flowchart of processes for cemented tungsten carbide is illustrated in Fig 2 One of the critical factors and an occasional problem in industry today is the carbon control during liquid phase sintering In recent years, low pressure (around

1000 psi) hot isostatic pressing has been used in combination with the standard vacuum liquid phase sintering to improve quality and properties of cemented tungsten carbide products

Fig 2 Schematic flow chart of manufacturing processes for cemented tungsten carbides

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Fundamental mechanical properties of cemented tungsten carbide include hardness, fracture toughness, transverse rupture strength (TRS), compressive strength, and wear resistance Hardness and TRS are widely used in day-to-day engineering activities for selection and quality control, while fracture toughness is mostly used in developing new grades Hardness is proportional to compressive strength and inversely proportional to fracture toughness Transverse rupture strength depends on both strength and toughness of the material When there is significant volume percent of defects such as porosity in microstructure, TRS becomes very defect-sensitive Figure 3 shows the interrelationships between hardness, fracture toughness, and TRS Transverse rupture strength is affected by both crack initiation and crack propagation processes; therefore, it exhibits an apex as hardness increases Wear resistance is more closely related to hardness Figure 4 demonstrates that the higher the hardness, the higher the abrasive wear resistance It also demonstrates that the higher the fracture toughness, the lower the wear resistance The trade-off between wear resistance and toughness has to be made when selecting cemented tungsten grades for specific engineering applications

Fig 3 Mechanical properties of WC-Co alloys

Fig 4 Dependence of wear resistance of WC-Co alloys on their hardness and fracture toughness

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Microstructure of most cemented tungsten carbide belongs to the category of dual-phase composites, in which tungsten carbide grain is the hard phase, and cobalt metal in between WC grains are the ductile phase In submicron super fine grain, the MPF is so small that the composite material behaves more like homogeneous materials Because cobalt content normally ranges from 6 to 20% (8 to 30% vol), it can also be categorized into ductile phase reinforced brittle materials Hardness of the materials decreases and toughness increases as cobalt content increases Grain size of WC also affect properties As grain size increases, hardness decreases and toughness increases Influence of microstructure on properties is reflected in the dependence of hardness, toughness, and wear resistance on MFP The larger the MFP, the lower the hardness and the higher the toughness (Fig 5) In considering wear, different mechanical properties (e.g., resistance against plastic deformation, crack formation or crack propagation) have to be considered

Fig 5 Dependence of hardness and fracture toughness on mean free path of tungsten grains in WC-Co alloys

In the sequence from crack formation to crack propagation and finally fracture, crack propagation plays the determining part

in most metallic alloys In contrast, crack formation seems to be the most important part in ceramics For cemented tungsten carbide, both plastic deformation and microfracturing mechanisms are important However, the dependence of wear resistance on hardness suggests that its resistance to local deformation is dominant at the onset of a wear process Due to limited ductility, initial plastic yield on the wear surface evolves into microfracturing quickly The surface plastic yield is apparently controlled by the cobalt content Preferential wear of cobalt also occurs as cobalt content increases Another important mechanism is the pullout of carbide grains, WC, as a result of preferential wear of cobalt phase The larger the grain size, the less the tendency of carbide grain pullout Therefore, within a certain range of hardness and toughness, larger grain size grades can be more wear resistant than finer grain size grades at an equivalent hardness level

Major categories of applications of cemented tungsten carbide include metal removal cutting tools, rock and earth drilling tools, drawing and punching dies and punches, wear components, and other engineering specialty applications

For metal cutting tools, wear progresses very quickly under high temperature and heavy load Tool life is limited to a very small amount of wear loss, usually indicated by the change in radius of the cutting edge Crater wear on the top surface of a cutting insert is the result of a chemical reaction between metal chips and tungsten carbide under high heat Titanium carbide and tantalum carbide additions are prevalent in most cutting tool grades to combat crater wear Modern cutting tools are often coated by chemical vapor deposition (CVD) with TiN, TiCN, TiC, and in combination with Al2O3 Physical vapor deposition (PVD) is utilized to coat drills, end mills, etc with TiN, TiCN, and CrN

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Cutting tools are categorized based on their specific cutting application, such as general purpose grades, rough turning grades,

or finish turning grades Trade-offs are often made in designing as well as selecting grades between chipping resistance and wear resistance For example, KC-850 (Kennametal, Inc., PA) is a general purpose turning grade used in industry which combines good wear resistance with good chipping resistance GC4015 (Sandvik Coromant Co., Fair Lawn, NJ) is a finishing turning grade with high hardness and crater resistance but moderate chipping resistance It is not suitable for interrupted turning Twist drills often use fine grain (submicron) high hardness grade RTW 2606 (Roger Tool Works, Inc., Rogers, AR)

is a grade for printed circuit board drills with a hardness in the range of 93.5 to 95 Ra

P/M cutting tool materials also include the cermet Cermet is commonly used to describe (TiC, TiCN)-(Ni, Ni/Mo) alloys Cermet has better high temperature hardness than cemented tungsten carbide They are mostly used for finishing turning cutting applications A comprehensive review on cutting tool materials and their wear resistance can be found in the article

"Friction and Wear of Cutting Tools and Cutting Tool Materials" in Friction, Lubrication, and Wear Technology, Vol 18 of the ASM Handbook

Earth drilling tools include petroleum drilling bits, mining bits, coal mining tools, and construction tools that use cemented tungsten carbide as cutting inserts Earth boring bits operate under very heavy load and in extremely abrasive environments Cutting structure of such tools must have very high compressive strength, very high abrasive wear resistance, and sufficient fracture toughness Cemented tungsten carbide offers the best choice In typical grades used for earth boring applications, cobalt content ranges from 6 to 16%, and grain size range from 1 to 10 m Hardness of these grades ranges from 85.0 to 91.0 HRA Corresponding fracture toughness is from 10 to 18 MPa Selection of a specific grade depends on the application For example, 6% cobalt grades with relatively low toughness are often used in applications where mechanical impact is minimum and sliding abrasive wear resistance is the main concern In applications where mechanical impact is unavoidable, grades with higher fracture toughness (by increase cobalt content or grain size) are used

Abrasive wear is considered the primary mechanism for earth boring applications However, depending on different types of rock drilling tools (rotary, percussive, or rotary and percussive), impact and/or impact fatigue can play a major role in causing material loss during drilling Contribution of impact fatigue and abrasion to wear can be separated to a certain extent Figure

6 shows the dependence of impact wear and abrasion wear on cobalt content, respectively The impact wear resistance shows

a maximum of 6 to 8% Co, which is similar to the relationship between compressive strength and transverse rupture strength versus cobalt content For abrasion resistance during rock drilling, the general principle for the effects of grain size, cobalt content, and MFP applies Very hard cemented carbides containing a high volume fraction of carbides exhibit lower abrasive wear loss on fine grain structures Softer, more ductile cemented carbides, containing less volume fraction of carbides, exhibit lower wear loss in the case of coarse grain structures In short, the abrasive wear resistance is generally related to hardness and cobalt content, while the impact wear appears to be closely related to bulk compressive strength and transverse rupture strength Another important factor affecting tool life during earth boring is the size of abrasive particles and its relative ratio

to grain size or mean free path of the microstructure Figure 7 shows that there is optimum /Deff where wear loss is minimum, where Deff is the effective diameter of abrasive particles The optimum values of /Deff for fine grain and coarse

materials are different indicating differences in wear mechanisms

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Fig 6 Contributions of impact wear and abrasive wear in rotary-percussive drilling test Source: Ref 6

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Fig 7 Abrasive wear loss of cemented tungsten carbides in a pin abrasion test vs grain size and the ratio of mean

free path to the effective diameter Deff of abrasive particles Source: Ref 1

Metal forming tools and dies, structural components, and a wide variety of wear parts constitute roughly one-third of total tonnage of cemented tungsten carbide shipped Metal forming tools and dies include hardmetal dies for drawing wire, rod, bar, tube, and special sections For wire drawing dies, crucial properties are abrasive wear resistance and compressive strength High surface finish requirements often dictate selection of fine grain and low cobalt grades However, for larger cross section drawing, such as rods and special sections, toughness is important The biggest market for pressing dies is the carbide industry itself, because nearly all cutting inserts and cemented carbide components are formed by the uniaxial die pressing method Other metal forming tools include pressing dies used for synthetic diamond and cubic boron nitride manufacturing, and rolls for the rolling of steel rod and sections, wire, tape, and foils Foil rolls require perfect mirror surface finish Fine grain and porosity free microstructure, which yields improved rupture strength and fatigue resistance of rolls, is extremely important

Structural components made of cemented carbide include mechanical seals, compressor plungers, lathe jaws, boring bars, grinding spindles, pulverizing pins, and bearing journals Wear parts include nozzles, guides, plungers, balls, slitting wheels and cutters for paper industry, and many other applications

Wear for most metal forming tools, structural, and wear components are classified as "sliding" wear based on their mechanical function Almost all wear mechanisms could occur during sliding wear applications However, depending on specific application, only one or two mechanisms may be dominant For most cemented carbide sliding wear applications, abrasive wear and adhesive wear are the primary mechanisms For metal rolls, obviously, rolling fatigue should be considered

in analyzing wear mechanisms

P/M Steels

There are various P/M tool steels engineered with respect to wear applications The steel compositions vary in carbon and alloying elements such as manganese and phosphorous contents By far the most recognized P/M steels for wear applications, however, are P/M tool steels Compared to conventional tool steels, P/M grades offer distinctive advantages in terms of finer

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grain size and homogeneous microstructure with less segregation The powder metallurgy approach also lends to near-net shape manufacturing important for components made of tool steel from the perspectives of machinability and cost A list of commercially available P/M tool steel grades with their typical composition and characteristics are found in Table 1 and in the article "Particle Metallurgy (PM) Tool Steels" in this Volume A number of P/M high speed steels have been developed that cannot be made by conventional methods because of their high carbon, nitrogen, or alloy contents Examples include CPM Rex 20, CPM Rex 25, CPM Rex 76, and ASP 60 The P/M hot isostatic pressing (HIP) process is the only way to produce such highly alloyed high speed steel (HSS) as Rex 76 and ASP 60

Table 1 Commercial P/M tool steel compositions

Constituent elements, % Trade name AISI

(a) HCHS, high carbon, high sulfur; HS, high sulfur

Figure 8 schematically illustrates P/M processing routes for tool steels The most critical processing steps are manufacturing

of prealloyed powder by atomization and consolidation by HIP There are two competing processes in making powder: water atomization and gas atomization Typical advantages of gas atomization over water atomization include cleanliness and particle shape control Tool steels produced by HIP are normally conducted in argon at 7 to 200 MPa pressure and 1000 to

1200 °C Powder metallurgy tool steels and other ferrous P/M alloys can also be manufactured by nonpressurized sintering techniques, such as solid state, transient, or super-solidus liquid phase sintering Standard sintering techniques have the distinct advantage of low cost and therefore wide application areas; but, in general, substantial compromise is made in terms

of mechanical properties due to presence of significant levels of porosity Lower strength and lower hardness dictate that wear resistance is lower than that of HIP alloys

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Fig 8 Typical processing flow chart of P/M tool steels Source: Ref 7

Powder metallurgy tool steels utilize the same basic heat treatment as their conventional counterparts, but they tend to respond more rapidly and with better predictability to heat treatment because of their more uniform microstructure and finer carbide size Powder metallurgy tool steels also have superior machinability to conventional tool steels due to generally higher sulfur content in P/M alloys

Wear resistance of P/M tool steels is largely associated with its hardness Figure 9 illustrates the relationship between impact toughness and hardness Balance in toughness and wear resistance must be considered in selecting a specific tool steel grade for applications Figure 10 is a representative view of the wear resistance of a range of sintered alloy steels against average hardness Variations in wear resistance versus hardness are due to specific alloys with respect to carbide contents, carbide particle sizes, size distributions, and shapes

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Fig 9 Compilation of fracture toughness of tool steels

Fig 10 Wear rates of a range of sintered alloy steels against average hardness Source: Ref 9

P/M tool steel is essentially a homogeneous material Volume fraction of carbides in a tool steel is <50% (10 to 20% in most cases) and particle size of carbide particles is usually very small, ranging from submicron to a few microns Although wear resistance is proportional to volume fraction of carbides, studies show that the rate of increase in wear resistance diminishes

>30% by volume for specific ferrous-base alloy systems and applications

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On the other hand, the tool steels including high speed steels can be considered as composite materials with a martensitic steel matrix and a harder, particulate reinforcement consisting of primary carbides Wear resistance of P/M tool steels is therefore similar to that of cemented tungsten carbide Microstructure factors controlling wear resistance include carbide volume fraction, carbide type, carbide particle size and size distribution, carbide particle shape, and steel matrix microstructure

Carbide types in tool steels include MC, M2C, and M6C types, where M could be V, W, Mo, or Cr The presence of M2C or

MC type carbides is considered to impart a greater wear resistance to these materials than M6C type carbides Vanadium carbide plays an important role in P/M tool steel PM 10V and PM 15V are primary examples of P/M tools steels with high vanadium content of 9.75 and 14.5 vol% respectively, which is impossible for conventional tool steels The microstructure of

PM 15V contains 22 vol% primary vanadium carbides with the majority having a size of 3 m or finer Table 2 compares high vanadium P/M tools to others It is apparent that wear resistance increases rapidly as vanadium content increases at the expense of impact toughness In general, carbide particle sizes range from submicron to a few microns (<10 m), depending

on processing history including heat treatment Particle size and volume distributions are important in the wear behavior, a high proportion of M2C/MC carbides of large size (6 m) give a greater wear resistance

Table 2 Wear resistance and toughness of high vanadium content steels and other wear resistant steels

Grade Hardness,

HRC

Pin abrasion resistance weight loss,

mg

Cross cylinder wear resistance,

× 1010 lb/in.2

Toughness Charpy C-notch impact

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Table 3 Wear resistance comparison between P/M tool steels and typical WC-Co grades

Materials Wear resistance

under low stress abraison (ASTM G 65),

1000 rev/cc

Wear resistance under high stress abraison (ASTM B 611),

1000 rev/cc WC-15%Co, grain size 5 m 0.435 3.3

WC-16%Co, grain size 6 m 0.238 1.8

Table 4 Qualitative comparison of mechanical properties of P/M tool steel and cemented tungsten carbide

Property P/M tool steel P/M tool steel WC-Co WC-Co

Hardness 50-58 HRC 59-68 HRC >88 HRA <87 HRA

Impact resistance High Medium-low Low Medium

Thermal shock resistance Low Low High Medium

Fig 11 Relative comparison of pin-on-disc wear resistance of PM 24Cr9V, PM 440V, PM 10V, PM D7, and T440B

stainless WC-8Co Source: Ref 10

Thermal fatigue cracking resistance and corrosion resistance are sometimes the primary cause for material loss and fracture Final selection of a material for an application depends on the mechanisms responsible for material loss versus consideration

of combined effects of all properties

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Applications of P/M tool steels and high speed steels include metal removal cutting tools and tooling for metal forming and shaping operations Cutting tools include milling inserts, hole machining reamers, taps, drills, and broaching tools Typical P/M tool steel grades for cutting tools include ASP 30, ASP 60, and the CPM Rex M3 and M4 At similar hardness levels, P/M grades offer superior impact toughness to similar grades processed by conventional means Table 5 compares CPM Rex

20 and CPM Rex M42 to conventional M42 As the result of better basic mechanical properties and more uniform finer microstructure, wear life of P/M tool steel tools are longer than their counterparts Figure 12 is one example of the tool life of milling inserts of ASP 30 and ASP 60 compared to conventional M42

Table 5 Charpy C-notch impact and bend strength of CPM alloys comparing to conventional M42 alloy

Alloy grade Hardness,

HRC

Charpy C-notch impact energy,

J

Bend fracture strength,

MPa

Conventional M42 67.5 7 2565

Fig 12 Comparison of P/M tool steel ASP 30 vs conventional M42 during end mill tests in Ti-6Al 4V Source: Ref 7

Cold- and hot-work rolls are another primary application of P/M tool steels High vanadium P/M tool steel such as CPM 9V and 10V are used for high wear and cold work applications Similar to cutting tools, the more uniform microstructure of P/M cold-work steel yields better toughness Higher vanadium content makes it possible that higher wear resistance is achieved at

an improved toughness simultaneously CPM 10V has proved to be more wear resistant than any commercially available high-alloy tool steel CVP 10V has far superior wear resistance and equivalent impact toughness when compared to conventional D2, M2, and CPM M4 (Fig 13) For hot-work applications, superior thermal fatigue resistance is very important The absence of segregation in P/M tool steels is an advantage over conventional alloys, because premature failure due to thermal fatigue is often attributed to segregation and heterogeneous microstructure H13 and H19 are typical tool steel grades for hot-work applications The P/M versions, P/M H13 and H19 have more uniform properties and equivalent or better toughness Again, high vanadium content can improve wear resistance and toughness

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Fig 13 (a) Wear resistance and (b) impact toughness of CPM10V compared to conventional, P/M, other tool steels

at indicated hardness Source: Ref 7

P/M Metal Matrix Composites

The concept of metal-matrix composite (MMC) refers to a large class of materials with at least one type of reinforcement situated in a metal matrix In general, the scale of the reinforcement is large compared to that of the microstructure The reinforcement can be in the form of continuous fiber, short fiber, or particulate Although it is recognized that MMCs offer considerable potential for enhanced wear resistance, understanding of their wear characteristics is still far from complete

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This is due in part to the inherent complexity of many wear processes, but the problem is compounded by an interplay with microstructure variables in MMC, such as reinforcement content, size, orientation, interface strength, etc

Many MMCs are fabricated by powder metallurgy, in other words, the composites are formed by consolidating constituent powders together Particulate MMC is especially suitable for P/M processing In a narrower sense, MMC refers to modern low-density high-strength alloy systems such as aluminum- or titanium-base composites (e.g., Al-SiCp, Al-Al2O3, Ti-SiCp, and Ti-6Al-4V-SiCp) However, it is clear from the previous discussion that many P/M materials for wear resistant applications bear the same basic microstructure characteristics of MMCs Therefore, many P/M materials can be considered MMC, and many particulate MMCs can be considered P/M materials Hence, general principles governing wear behavior of two or multiphase P/M materials apply to many MMCs

Key factors affecting wear resistance of MMC include volume fraction of reinforcement particles, particle size of the reinforcement, and the diameter of reinforcement relative to the size of the abrading particles Figure 14 illustrates that specific wear rates of aluminum-base composites decrease with the increase of volume fraction of different hard particle reinforcements This behavior is true under either abrasive or adhesive wear situations For a given volume fraction of particles, composites that contain harder particles exhibit a lower wear rate Similar results can be expected for tool steel-base composites such as steel-base TiC particle composites Figure 15 demonstrate the effects of the size of the hard particles Larger ceramic particles and higher contents give greater wear resistance Larger abrasive particles cause significantly more wear to composites because of the increasing chances of fracturing of reinforcement particles under the heavy load exerted by large abrading particles In the case of erosive wear, situations are more complex Factors such as the impact angle play a very important role as well as volume fraction For example, it was found that the minimum erosive wear rate appeared at 34 vol% binder in WC-Co cermets MMC can also contain soft reinforcement Al-Sn alloys, for example, have long been used as bearing alloys In general, when improvements to the wear resistance are obtained in this way, it is normally accompanied by

a reduction in the coefficient of friction

Fig 14 Wear rate of aluminum matrix composites vs volume fraction of reinforcement particles during a sliding

test against steel Source: Ref 11, 12

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Fig 15 Dependence of wear rate of 6061 Al-Safill ( -alumina) composites on volume fraction of reinforcement and

particles size of abrasives Source: Ref 11

1 K.-H Zum Gahr, Microstructure and Wear of Materials, Elsevier, 1987

5 K.J.A Brooks, World Directory and Handbook of Hardmetals and Hard Materials, 5th ed., International

Carbide Data, 1992

6 J Larsen Basse, Powder Metall., Vol 16 (No 31), 1973, p 1-32

7 Properties and Selection: Irons, Steels, and High-Performance Alloys, Vol 1, Metals Handbook, 10th ed., ASM

International, 1990

9 D.S Coleman, J Bates, Q.A Shaikh, and P.R Brewin, Proc 12th Int Plansee Seminar '89, Metallwerk

Plansee GmbH (Reutte, Austria), Vol 2, 1989

10 K.E Pinnow, W Stasko, J.J Hauser, and R.B Dixon, Proc Advances in Powder Metallurgy & Pa rticulate Materials (San Francisco), Vol 6, Metal Powder Industries Federation, June 1992

11 C.T Clyde, Introduction to Metal Matrix Composites, Cambridge Press, 1991

12 Friction, Lubrication, and Wear Technology, Vol 18, ASM Handbook, ASM International, 1992, p 804

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Hardfacing and Thermal Spray

Applications. Metal alloy powders are used for hardfacing and thermal spray coatings for wear-resistant applications Hardfacing is the application of hard, wear-resistant material to the surface of a component by welding, thermal spraying, or a similar process for the main purpose of reducing wear Metal alloy powders used for hardfacing include WC-Co powders, other carbide-metal alloy powders, cobalt-base wear-resistant alloy powders such as Stellite (Deloro Stellite, Inc., Belleville, ON) series of alloys, and some nickel- or iron-base wear-resistant alloy powders

Thermal spray coatings are themselves a major engineering field Numerous different techniques for applying hardmetal powders onto a substrate can be categorized as thermal spray coating Thermal spray techniques include but are not limited to traditional oxyacetylene torch spray, plasma thermal spray, detonation gun (D-gun and super D-Gun) thermal spray, and more recent technologies such as high velocity oxy-fuel (HVOF) thermal spray Figure 16 is a schematic illustration of the plasma spray coating process Common features of thermal spray coatings, differing from other hardfacing processes (such as welding processes), are thickness of thermal spray coating is usually limited to 0.025 to 0.5 mm and bonding between coating and substrate is primarily mechanical Key advantages of thermal coatings over welding are geometry, flexibility, and process temperature of substrate components can be maintained so low that the coating can be applied to finished products

Fig 16 Schematic of plasma spray hardfacing process

Hardfacing materials are often applied by various welding techniques The welding techniques for hardfacing include oxyacetylene torch, gas metal arc, gas tungsten arc, plasma arc, and plasma transfer arc welding Plasma transferred arc process is illustrated in Fig 17 Characteristics of weld-on hardfacing are that coatings are thicker than thermal spray

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coatings, and coatings are metallurgically bonded to the substrate Welded-on wear-resistant hardfacing has wide ranging applications from earth moving equipment to precision bearing and seal surfaces

Fig 17 Schematic of the plasma transferred arc hardfacing process

Another category of hardfacing techniques is cladding Hardfacing materials can be cladded onto substrates by furnace fusing prearranged layers of loosely bonded hardmetal onto substrate (Fig 18) Two layers of metal and metal carbide tapes are stacked over the substrate The tape is made from powder bonded by polytetrafluoroethylene (PTFE) or other polymers One

of the two layers is a carbide powder tape and the other is high temperature brazing alloy powder tape, such as a nickel-base brazing alloys Coating is formed by heating the assembly to brazing temperature and infiltrating the brazing alloy into the carbide layer The final coating of nickel alloy bonded carbide particles is bonded to the substrate during the process Cladding can also be done by HIP Advantages of furnace cladding include better dimensional and thickness control Disadvantages are that the entire components are exposed to high temperature (typically 1000 to 1250 °C) for an extended period of time

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Fig 18 Flow chart of a tape cladding process Source: Ref 13

Metal carbides are the most widely used materials for wear-resistant coatings and hardfacings Among various types of carbides, tungsten carbide is the most popular one Chromium carbide is also very common Other carbides, such as vanadium carbide and molybdenum carbides, are often the choice for additives

welding hardfacing WC-Co powder can be a product of crushing bulk WC-Co alloys or specifically made spherical shaped pellets Tungsten carbide powders for thermal spray coatings usually contain from 6 to 18% cobalt Table 6 lists typical compositions of carbide powders for thermal spray coatings Cr3C2 powder and its alloys are also popular for thermal

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applications Particle sizes of WC-Co and other alloy powders for thermal spray coatings range from 5 to 150 m Powders for torch welding hardfacing ranges from 150 m to 1.2 mm

Table 6 Typical thermal spray alloy powders for wear applications

Composition, % Powders

WC Cr 3 C 2 Co Ni Fe Cr Other WC-12Co bal 12

WC-10Co-4Cr bal 10 4

WC-17Co bal 17

WC-10Ni bal 10

Cr 3 C 2 bal

Cr 3 C 2 -25NiCr bal 24-26NiCr

Cr 3 C 2 -FeCrAlY bal 14-16FeCrAlY

Ni-718 bal 17-21 Mo 2.8-3.3, Nb 4.8-5.5, Fe 14-21, Ti 0.75-1.15

Stellite 6 bal 28 4W-1Si-1C

Stellite 1 bal 31 12.5W-1Si-2.5C

Stellite 12 bal 29 8.5W-1.5Si-1.5C

temperature thermit process during which ore concentrate is converted directly to WC Macrocrystalline tungsten carbide maintains carbon content of 6.13% by weight, the correct stoichiometric amount Macrocrystalline tungsten carbide is grown

in crystals ranging from 1 m to 5 mm Coarse mesh sizes of macrocrystalline carbide are extensively used in abrasion and erosion protection applications The finer size is employed as wear rate modifiers Coarse grain tungsten carbide powder can also be obtained by conventional dry-carburizing tungsten metal method Examples include commercial grades of MAS 2000, MAS 3000, and MAS 5000 by H.C Starck (Newton, MA) Particle sizes of these powders range from 20 to 50 m

eutectic of WC and W2C that can range in carbon content from 3.5 to 4.5% by weight It is manufactured by melting mixtures

of tungsten metal, tungsten carbide, and carbon ingredient at temperatures above 3000 °C The melt is cast into billets, which are then crashed into specified size ranges of powder Cast carbide can be crashed to as coarse as 20 mesh and as fine as -325 mesh powder Cast carbide, mostly used in torch welding hardfacing and infiltrated drag bit matrix body, is often used in combination with sintered tungsten carbide and other carbides to enhance wear resistance Addition of cast carbide tends to degrade toughness of coating

Spherical cast tungsten carbide is a relatively new product in the marketplace with increasing applications Spherical cast carbide has all the basic characteristics of crashed cast carbide Spherical cast carbide has extremely high wear resistance and higher resistance to chipping and cracking than crushed cast carbide due to its spherical shape Spherical cast carbide is manufactured by patented solidification technique (Ref 14), during which molten eutectic carbide droplets are rapidly cooled Spherical cast carbide can be used alone or in combination with other carbides

13 Conforma Clad, Inc., New Albany, IN, technical literature

14 E Findeisen et al., Process of Manufacturing Cast Tungsten Carbide Spheres, U.S Patent 5,089,182, 18 Feb

1992

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Hardfacing Alloy Powders

Cobalt-base hardfacing alloy powders are best known as Stellite powders Typical compositions and wear resistance

of Stellite powders are found in Table 7 Cobalt alloys for hardfacing have advantages for use in combined wear-resistant, corrosion-resistant, and high temperature applications Cost of cobalt-base alloys tends to be higher than nickel- and iron-base alloy powders Cobalt alloys are typically used for overlay applications, deposited readily by plasma transfer arc process

Table 7 Typical cobalt-, nickel-, and iron-base hardfacing alloys

Moderate Good Cutting tools, sheer

blades, reamers, dies, ingot tongs

Excellent Agricultural machinery,

coke shutes, brick-making equipment

5-of carbon steel

Fair Deteriorates

as carbon increases

Excellent for low stress

Poor for high stress

High temperature, high corrosion applications, pump shafts, engine valves

Good for light impact Decreases

as carbon increases

Excellent for low stress

Seal rings, cement pump screws, valves, cams

60 HRC matrix

Superior to all others

Good under light impact

Poor Cutting teeth and edge

holding surfaces of rock drill bits, quarrying, digging, earth moving equipment

Nickel-base hardfacing alloy powders are usually Ni-Cr-B-Si system based Additions of boron and silicon to nickel suppress the melting point so that less heat is required to deposit the nickel-base hardfacing alloy onto substrates However, excessive additions of boron and silicon lower plasticity Most of these alloys contain nickel boride (Ni3B), chromium borides (CrB, Cr5B, and Cr2B), and other complex borides as hard phases Typical alloy compositions and wear resistance are listed

in Table 7 Deposit hardness in these alloys is as high as 60 HRC, depending on the chromium, boron, and silicon contents Nickel-base hardfacing alloys are particularly popular in oil pumping and in glass, ceramic, cement, and plastic industries

Iron-base hardfacing alloys can be classified into Pearlitic steels, Austenitic steels, Martensitic steels, and high-alloy irons The high alloy irons usually contain high carbon and high chromium content More iron-base hardfacing alloys are used in various industries than cobalt- and/or nickel-base hardfacing alloys, primarily due to the lowest cost and highest availability Typical composition and wear resistance comparison is shown in Table 7

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Full Density Cobalt Alloys

Alloys based on cobalt chromium-tungsten-carbon, such as the Stellite family of alloys, are extremely versatile, resistant materials (Ref 15, 16) Full density P/M components also can be manufactured by hot isostatic pressing or extruding operations (Ref 17, 18) for near-net shape processing and refined microstructures

wear-Material Descriptions and Applications. Table 8 gives the nominal chemical analysis for the most popular Stellite P/M materials The microstructure of Stellite alloys contains complex combinations of M7C3, M6C, and M23C6 carbides embedded

in a cobalt-chromium-tungsten superalloy matrix

Table 8 Nominal chemical analysis of most widely used Stellite P/M materials

Composition, % Alloy

Cr W Ni Mo (a) Si (a) Mn (a) Fe (a) C B (a) Other (a)

(a) Maximum; all materials cobalt base

For highly abrasive conditions, alloys such as Stellite 3, 98M2, and Star J are preferable because they contain a large volume fraction of carbides The lower carbon grades are more ductile and are suitable for applications involving shock loading or impact

Typical applications include precision balls, bearing race blanks, valve seat inserts, saw tips, cutters, spacer bushings, and wear pads Stellite 3 is also used in medical instruments Stellite 12 P/M has proved to be exceptional for cutting timber

Material Properties. Tables 9 and 10 show the physical and mechanical properties of Stellite materials The physical properties were measured on castings but are thought to be representative of P/M parts It should be noted that Stellite materials are only weakly magnetic under most conditions

Table 9 Physical properties of Stellite materials

Stellite trade designation Property

2345)

(2215-1260-1357 (2300- 2475)

1255-1341 (2290- 2445)

1239-1299 (2260- 2370)

1340-1396 (2445- 2545)

1224-1275 (2235- 2325)

1215-1299 (2220- 2370)

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(a) P/M materials are typically 97 to 100% dense

Table 10 Mechanical properties of Stellite materials

Stellite P/M trade designation Property

Heat Treatment. Stellite P/M materials are not generally considered heat treatable nor is heat treatment required Stress relief treatment involves a slow heat-up and soaking at 900 °C (1650 °F) for at least 2 h, followed by furnace cooling Stellite

6 can be solution annealed by soaking at 1200 °C (2190 °F) for 2 h followed by rapid cooling This maximizes the corrosion resistance and ductility of Stellite 6 Stellite 31 can be age hardened using a 2 day aging treatment at 800 °C (1470 °F)

Joining. Stellite materials can be joined to other materials by brazing or welding Flame brazing is typically done with silver solder and an oxyacetylene flame Vacuum furnace brazing uses a nickel-base brazing alloy such as AMS 4777

Gas tungsten arc and gas metal arc resistance welding are commonly used for Stellite P/M materials For the fusion processes,

a preheat of 810 °C (1490 °F) is required for the harder grades, such as Stellite 3, 19, 98M2, and Star J The softer grades, such as Stellite 6 and 31, should be preheated to 540 °C (1005 °F) Furnace cooling is required for the harder grades, whereas cooling in still air is satisfactory for the softer grades Stellite 25 or Nistelle W filler metals are recommended for joining Stellite material to mild steel or stainless steel

Machining. With the proper machine setup, Stellite alloys can be rough machined Alloys having hardnesses of <55 HRC can usually be machined, whereas alloys having hardnesses >60 HRC are normally ground For better surface finishes, any of the Stellite wear-resistant alloys should be ground

These alloys are usually machined with tungsten carbide tools Tools for turning should have a 5° end relief, 10° side relief, and a side cutting edge angle of 45° Tools for facing and boring are essentially the same, except for greater clearance where needed For best results in drilling, the drill web should be kept as thin as possible, using carbide-tipped drills In reaming, a

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45° side cutting edge angle should be used The tapping of holes is not recommended for any of the harder alloys, but threads can be produced by electrical discharge machining techniques

Table 11 is a guide to machining these alloys; it has been written to cover the machining of all the alloys and does not necessarily represent the optimum conditions for each In general, the harder alloys should be machined, starting at the lower end of the speed and feed ranges shown in Table 11 Softer alloys or more rigid setups will allow higher speeds and feeds

Table 11 Guide for machining Stellite alloys

(a) C-3 tungsten carbide tools should be used Coolant is water-base fluid diluted 15 parts water to one

part fluid Tools for facing and boring are basically the same as turning tools, except for greater clearances where needed

(b) Depends on bar and size of tool

(c) C-2 tungsten carbide twist drills should be used, with drill web kept as thin as possible Coolant is

same as in (a)

(d) Depends on size of drills

(e) C-3 tungsten carbide tools should be used Reamers should have a 45° side cutting edge angle

Coolant is same as in (a)

Table 12 General recommendations for grinding wheels

Peripheral wheels, Stellite alloy deposits A60JV Emulsifying oil

Form grinding (finer grits recommended for

sharper forms)

A100GV, A120KV, A150GV

Mixed Stellite alloy and steel A46DZB 0.45 kg (1 lb) sal soda to 190 L (50 gal) water

Cup wheels for peripheral wheels (see above) A46JV 3.8 L (1 gal) good grade of water-soluble oil plus 190 L

(50 gal) water

Internal grinding bores greater than 100 mm A54JV, A54MV 3.8 L (1 gal) good grade of water-soluble oil plus 190 L

Tiêu đề	Fatigue and Fracture Control for Powder Metallurgy Components
Tác giả	Randall M. German, Richard A. Queeney
Trường học	The Pennsylvania State University
Chuyên ngành	Powder Metal Technologies and Applications
Thể loại	article

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