Friction, Lubrication, and Wear Technology (1997) Part 11 potx

Perhaps the most direct method is to employ a die steel that is more resistant to wear, that is, one that is harder and that retains its hardness at high die temperatures Ref 54.. Coatin

The results by Thomas (Ref 43) also showed a strong correlation between wear resistance and alloy content (Fig 17) The low-alloy steel (No 5 die steel, which is equivalent to steel number 1 in Fig 15, and 8(a) and Table 8(b)) had relatively poor wear resistance when compared to the 5% Cr hot-work die steel (H12) and the 12% Cr steel This was true irrespective of the hardness level to which the steels were tempered and even, to a certain degree, of the type of workpiece material

Fig 17 Influence of initial die hardness on wear of die steels The wear index is defined as average

cross-sectional area of wear depressions in dies No 5 die steel: 0.6 C, 0.3 Si, 0.6 Mn, 1.5 Ni, 0.6 Cr, 0.25 Mo; 5% Cr steel: 0.33 C, 0.3 Si, 1.0 Mn, 5.0 Cr, 1.5 Mo, 1.5 W, 0.5 V; 12% Cr steel: 0.1 C, 0.25 Si, 0.7 Mn, 2.4 Ni, 12.0

Cr, 1.8 Mo, 0.35 V Source: Ref 43

These findings have been verified and further expanded by other researchers For example, in their carefully controlled forging experiments, Doege, Melching, and Kowallick (Ref 48) and Hecht and Hiller (Ref 50) found low-alloy steels to have far inferior wear resistance as compared to hot-work die steels (Fig 18) because alloying led to higher hardness and the ability to retain strength at high die temperatures The work of Netthöfel (Ref 49) is also in agreement with these observations

Fig 18 Amount of wear of hot-work tool steels as a function of the number of forgings Equivalent steels are in

parentheses Source: Ref 48

Up to now, most of the discussion of alloying has centered on die steels Several workers have also investigated the wear characteristics of nonferrous die materials In Netthöfel's (Ref 49) experiments on forging die wear, it was found that the nickel-base alloy Nimonic 90 had a wear resistance between that of an H12 and H19 steel at a die temperature of 255 °C (490 °F) This is an important finding in view of the fact that the nickel-based alloys are generally many times the cost of

the die steel alloys and are also harder to machine Notthöfel's finding was verified to a certain extent by Ali, Rooks, and Tobias (Ref 42) in their die wear studies in a high-energy-rate forming (HERF) machine (Fig 19) Although the die were depended on whether the top die or bottom die was examined, it was found that Nimonic 90 was only slightly better than

a steel similar to H19 (WEX) Thus, the results reported in Ref 42 and 49 point out the fact that the nickel-base die materials should be reserved for hot-die and isothermal forging applications for which die steels are inappropriate

Fig 19 Total wear volumes for die materials at a mean hardness of 44 HRC Source: Ref 42

Die Hardness. Die hardness is another factor whose influence on abrasive wear is easy to quantify The effect of die hardness is best realized through an understanding of the die wear process itself Misra and Finnie (Ref 51) have summarized a large amount of work on abrasive wear and concluded that two basic processes are involved The first is the formation of plastically deformed grooves that do not involve metal removal, and the second consists of removal of metal

in the form of microscopic chips Because chip formation, as in metal cutting, takes place through a shear process, increased metal hardness could be expected to diminish the amount of metal removal via abrasive wear This trend is exactly what has been observed

The effect of hardness on wear is seen in data on a variety of steels quoted by Kannappan (Ref 40) and by Thomas (Ref 43), which have been discussed previously From examination of Fig 17, it is apparent that the dependence of wear rate

on hardness is greatest for low-alloy die steels such as 6F2 (No 5 die steel in Fig 17) Such a trend has also been reported

by Kannappan (Ref 40) in data on several low-alloy and hot-work die steels

Kannappan (Ref 40) has also discussed the correlation between hardness and wear of die steels with microstructures different from the typical die steel structure of tempered martensite It has been found that the isothermal heat treatment

of steels to produce lower bainite results in better wear resistance (Fig 20) Supposedly, this effect is a result of the fact that isothermal transformation/hardening causes fewer stresses and microscopic cracks (which promote abrasive failure) than does a thermal martensitic transformation

Fig 20 Relative wear resistance with respect to hardness of selected chromium steels with 0.55% C Note the

difference between the effect of quenching followed by tempering (solid lines) and the effect of isothermal treatment/quenching to a lower bainitic region (dashed lines) Relative wear resistance is defined as a number directly proportional to the applied interface pressure and the amount of relative sliding and inversely proportional to the total wear volume Source: Ref 40

Workpiece Temperature. Several researchers have commented on the effect of workpiece temperature on die wear

In his investigation of wear of hammer dies, Thomas (Ref 52) found that in forging of steels, wear increased at first with billet temperature up to 1100 °C (2010 °F) and then decreased with increasing temperature (Fig 21) The initial increase can probably be attributed to the increase in the amount of scale on the billets, which acts as an abrasive during the die wear process However, above 1100 °C (2010 °F), the flow stress drops off rapidly enough to minimize the interface pressure during forging and therefore decrease the effect of scale A similar finding was made by Doege, Melching, and Kowallick (Ref 48), who attributed an increase in die wear as the billet temperature was raised from 800 °C (1470 °F) to

1100 °C (2010 °F) to an increase in the die surface temperature and a simultaneous decrease in wear resistance

Fig 21 Effect of workpiece temperature on wear Source: Ref 52

Lubrication/Die Temperatures. The effects of lubrication and die temperature on die wear have been interpreted in

a variety of often-conflicting ways in the literature This is because lubricants and die temperature influence: lubricity, and hence the amount of metal sliding during forging; the interface pressure during deformation; and the heat transfer characteristics between the dies and workpiece during conventional hot forging The last item is important not only through its influence on heat absorption into the dies, and thus thermal softening and decreased wear resistance of the dies, but also through its effect on the performance of the die and billet lubricants themselves

Investigations into the effect of lubrication on die wear in simple upsetting have shown that wear is greatly increased when the dies are lubricated versus when they are not This effect is shown in the results of Singh, Rooks, and Tobias (Ref 44) from upset tests in a HERF machine (Fig 22) The same phenomenon has been demonstrated by Thomas (Ref 43), who upset successive lots of 1000 samples each on a flat die in a mechanical press In these tests, the amount of wear was greater for the lot involving lubricated compression tests (Fig 23) From these findings, one might conclude that wear increases with lubrication because of increased sliding and that lubrication is detrimental in forging Thomas clarified this point, however, by calculating the amount of wear for equivalent amounts of metal flow past a given point;

he found that lubrication reduces wear by a factor of 3 when compared to forging without lubrication Moreover, he emphasized that in closed-die forging, the amount of metal sliding is fixed by die and preform design and not lubrication Thus, the amount of sliding over the flash land, where wear is usually greatest, depends on the amount of flash that must

be thrown and not on the efficiency of the lubricant employed Because the amount of flash will be roughly the same with

or without lubrication, employing lubricants in closed-die forging should reduce abrasive wear of the flash land and other parts of the die cavity

Fig 22 Effect of lubrication on forging die wear Source: Ref 44

Fig 23 Effect of lubrication on forging die wear Wear index is defined as the average cross-sectional area of

wear depressions in the die Source: Ref 43

The interaction of lubrication and die temperature effects was demonstrated by Rooks (Ref 45) in upset tests on a HERF machine These tests were run with various bulk die temperatures, dwell times, and cycle times Dwell time in the HERF operation includes a short forging phase, a somewhat longer "bouncing" phase, and an extended "after-forging" phase during which the dies and billet are in contact under low pressure Results established that die wear after upsetting of

1000 billets decreased with increasing die temperature This was correlated with decreased amounts of sliding at higher die temperatures due to an increase in the coefficient of friction

The effects of dwell time and cycle time on die wear were also examined by Rooks (Ref 45) Increasing dwell time increases die chilling As a result, metal flow is hindered and die wear is reduced Increased cycle time (time between forgings) tends to have the reverse effect of increasing dwell time (that is, it increases die wear because of lower coefficients of friction and more sliding) However, these effects have been found to be very slight in upset tests, conducted in a HERF machine (Ref 45)

A striking die wear feature that Rooks (Ref 45) and Ali, Rooks, and Tobias (Ref 43) noted concerns the generally higher wear experienced by the top die versus the lower die, which is most noticeable in their lubricated upset tests (Fig 19) This can be attributed to greater chilling on the bottom die because the hot workpiece was placed on it prior to forging

This could, therefore, have been expected to lead to greater friction, less sliding, and thus less abrasive wear than the top die experienced

From a practical standpoint, increased production rate in a forge shop may be expected to lead to lower die life This is almost certainly a result of increased die temperature In forging under production conditions, the die surface temperature observed between two consecutive forging blows seems to remain unchanged throughout a production run (Ref 49) During the actual forging operation, the die surface temperatures increase and reach a maximum peak value and decrease again when the dies are separated and the forging is removed In case the forging "sticks" in one of the dies, the peak surface temperature of that die may increase further and contribute to die wear Therefore, in conducting die wear studies,

it is suggested that an ejector be used to remove the part after forging, so that die temperatures do not increase because a forging sticks in the die In forging of steel at 1200 °C (2190 °F) with dies at about 250 °C (480 °F), surface temperatures will reach approximately 750 °C (1380 °F) if perfect and ideal contact occurs between the forging and the die In reality, however, due to scale and oxidation at the die/material interface, the peak surface temperatures during forging reach 500

to 600 °C (930 to 1110 °F) in mechanical presses and 650 to 700 °C (1200 to 1290 °F) in hammers As an example, die temperatures obtained by Vigor and Hornaday (Ref 53), in forging steel in a mechanical press are given in Fig 24 It can

be seen that the temperature gradient is very large at the vicinity of the die/material interface

Fig 24 Temperatures at the surface and at various depths in forging dies obtained during forging 1040 steel

without lubricant Source: Ref 53

The effects of sliding on die wear are also qualitatively well known in forging practice These effects are taken into account in designing preforms to ensure that more "squeezing" and less lateral flow and sliding action take place during finish forging

Methods of Improving Resistance to Abrasive Wear

From the discussion of the factors that influence abrasive wear, one can deduce methods to improve die performance controlled by this failure mechanism Perhaps the most direct method is to employ a die steel that is more resistant to wear, that is, one that is harder and that retains its hardness at high die temperatures (Ref 54) This could mean changing from a low-alloy die steel to a chromium hot-work die steel The decision to make such a change should be based on the suitability of the new die steel itself in the forging operation and the trade-off between expected increases in die life and increases in material (and machining) costs

Coating, hardfacing, and surface treatment of forging dies often can be employed to improve wear resistance as well Information regarding specific coating and hardfacing alloys (and the methods of their application) and surface treatments such as nitriding and boriding is contained in the following section of this article and will not be reviewed here However, there are numerous instances of such methods increasing die life These include the use of chromium and cobalt-base coatings (Ref 48, 55), weld deposits of higher-alloy steels onto low-alloy steels (Ref 56), weld deposits of nickel and cobalt hardfacing alloys on die steels (Ref 57, 58, 59), ceramic coatings (Ref 60, 61), and surface nitriding (Ref 54, 62, 63)

Another means of reducing wear in the forging of steel involves reducing the scale on heated billets; scale acts as an abrasive during the sliding that occurs between the dies and workpiece Thomas (Ref 52) estimates that poor control of scale can reduce die life as much as 200% Methods of reducing scale are relatively obvious and include the following:

• Using a reducing, or inert, furnace atmosphere

• Using a billet coating to prevent oxidation

• Minimizing time at temperature in the furnace or using induction heating

One final means of decreasing the problem of wear is through improved redesign of the blocker shape This is an important consideration because wear is strongly dependent on the amount of sliding that occurs on a die surface Thus, it

is possible to reduce sliding, thereby reducing wear, by redesigning the blocker shape

Thermal Fatigue

Thermal cycling of the die surfaces during conventional hot forging results in the second most common reason for rejecting dies, namely heat checking Thermal cycling (thermal fatigue) results from the intermittent nature of forging production

The major factors influencing heat checking are:

• Die surface temperatures

• Surface stresses and strains

• Damage accumulation in thermal fatigue

• Microstructural effects of fatigue

Die Surface Temperatures. Information on die temperatures is best obtained from direct measurements Surface temperatures for dies used in a mechanical press have been found to reach about 600 °C (1110 °F) in forging of steel cylinders that were preheated to approximately 1175 °C (2150 °F) and upset to 75% reduction in height (Fig 24) Similar

measurements have been made by Kellow, et al (Ref 64), who upset medium-carbon steel samples in a slow hydraulic

press and a HERF machine Surface thermo-couples were placed at various distances from the axis of the 25 mm (1 in.) diameter billets Experimental results showed that the temperatures obtained along the initial contact area of the workpiece and the die do not differ significantly between low- and high-speed forging However, the temperatures obtained outside the initial contact area, where the billet surface extends during deformation, were significantly higher in high-speed forging (900 °C, or 1650 °F) as compared with low-speed hydraulic press forging (550 °C, or 1020 °F) These results are mainly due to differences in heat generation due to friction, which serves as one means of dissipating the energy produced by the forging machine

Other measurements of die temperatures away from the die surfaces themselves demonstrate that large temperature gradients, as well as high temperatures, are induced in forging dies These measurements include those of Voss (Ref 65), who measured temperatures in low-alloy (6F3) and chromium hot-work steel (H10, H12) radial forging dies preheated to

100 °C (210 °F) before forging (Fig 25) Measurements away from the surface (at 0.5 mm, or 0.02 in., from the surface) show large temperature gradients By comparing die temperatures during forging to those between forging blows (Fig 25), it is apparent that very large temperature changes at the surfaces of forging dies may be expected as well For this reason, large stresses and large strains due to temperature effects are experienced by the surface layers of forging dies

Fig 25 Temperatures in dies with air-water cooling of the dies between blows Initial die temperature: 100 °C

(210 °F) Initial stock temperatures: (1) 1150 °C (2100 °F), (2) 1050 °C (1920 °F), (3) 950 °C (1740 °F) Upper curves are the temperatures achieved during forging; lower curves are the temperatures reached between forging blows Source: Ref 65

Materials with high conductivity are less likely to develop large thermal gradients and fail by thermal fatigue than those with poor thermal conductivity Although conductivity data for the various die materials are scarce, available measurements do show, for instance, that the tungsten hot-work die steels with higher conductivities should be more resistant to heat checking than the chromium hot-work die steels

Surface Stresses and Strains. The stresses and strains that result from the temperature cycles experienced by the forging dies have two main sources: (1) thermal expansion and contraction, and (2) phase changes brought about by temperature cycling The first of these is probably the easiest to quantify This is because the thermal stresses and strains

are approximately proportional to the maximum temperature difference (Tmax - Tmin) experienced by the dies and the thermal expansion coefficient of the die material Most die steels have similar thermal expansion coefficients Therefore,

the thermally induced deformation of the dies is controlled primarily by the magnitude of Tmax - Tmin

As might be expected, the tendency to heat check can be decreased by reducing Tmax - Tmin This can be done in two ways

First, Tmin, or the bulk die temperature, can be increased However, such a change may adversely affect resistance to other

forms of die failure Alternatively, Tmax can be decreased The easiest way to do this is by decreasing the workpiece temperature or by using a lubricant with better thermal insulating properties

Figure 26 shows the effects of increasing Tmin or decreasing Tmax on the fatigue life (in terms of number of cycles to

produce a crack of certain length) of mild steel It is seen that a 100 °C (180 °F) decrease in Tmax is much more beneficial

in extending the fatigue life than a similar increase in Tmin This result is generally true for die steels as well, and can be

attributed to the greater reduction of the strain amplitude by decreasing Tmax

Fig 26 Effects of maximum (a) and minimum (b) temperatures on the fatigue life of En 25 mild steel Source:

Ref 40

The effects of phase changes on thermal fatigue of forging dies has been examined by Rooks, Singh, and Tobias (Ref 66) and Okell and Wolstencroft (Ref 67) Both sets of investigators have concluded that die surface heating and cooling may lead to reversion of the tempered martensite to austenite and subsequent transformation back to martensite Because austenite and martensite have different densities, such phase changes lead to strains and stresses that are imposed by subsurface layers of the dies that do not undergo the transformation

As with the thermally induced strains, transformation-induced strains can be reduce either by keeping the maximum die surface temperature below the Ac1 temperature (the temperature at which austenite forms, which is 800 °C, or 1470 °F),

or by keeping the minimum die surface temperature above the martensite start, Ms, temperature, which depends greatly on alloy composition, typical values being 280 °C ( 535 °F) for H11 and 380 °C ( 715 °F) for H21 Okell and Wolstencroft (Ref 67) suggested the latter possibility, but specified that it should only be used for the more highly alloyed die steels, which have good hot hardness because they resist overtempering

Microstructural Effects on Thermal Fatigue. Because ductility has a large effect on the number of thermal cycles

a forging die can undergo prior to forming cracks, microstructure can have a significant impact on the frequency of heat checking The most important microstructural variables are cleanliness, grain size, and microstructural uniformity Die

steels that are clean resist crack initiation inasmuch as inclusions act as nuclei for crack initiation Thus, the use of a slightly more expensive steel that has been refined to remove inclusions may be a wise investment Grain size can also affect thermal fatigue resistance because grain size has a large influence on crack initiation, with fine-grained material tending to perform better in this respect Lastly, steels whose chemistry and microstructure are uniform (that is, free of segregation) tend to have uniform thermal properties (thermal expansion coefficients) and thus are able to resist thermal stresses and strains that may be developed due to such variations in a uniform temperature field

Methods of Improving Resistance to Thermal Fatigue. The resistance to thermal fatigue can be improved by materials selection, lowering the maximum die temperature variations, or surface treatments

Use of a Steel with Higher Yield Strength. Because thermal fatigue crack growth is controlled by the amplitude of the plastic strain increment, die steels with higher yield strengths, and thus higher elastic limits, are more resistant to thermal fatigue under a given set of process conditions

Lowering the Maximum Die Temperature. Because thermal stresses and strains are related to the temperature changes that the die surfaces experience, decreasing the maximum temperature to which the die surface is exposed is beneficial This can be accomplished by lowering the workpiece temperature or by using lubricants, such as glasses, that act as thermal insulators Lowering the maximum die surface temperature is also helpful in avoiding transformation-induced strains, which, in conjunction with thermal strains, may cause thermal fatigue problems

Raising the Bulk Die Temperature. Thermal stresses and strains and the tendency for heat checking can also be reduced by preheating the dies to higher temperatures Use of this technique should be limited to die steels with good retention of hot hardness, such as the molybdenum hot-work die steels (H41 to H43)

Use of High-Quality Die Steel. Die steels that are clean, of fine grain size, and homogeneous in microstructure resist the initiation of fatigue cracks that are thermally or mechanically induced

Use of Special Surface Finishes or Treatments. By eliminating machining marks, which act as stress concentrators, fatigue crack initiation can sometimes be avoided or delayed Surface treatments such as nitriding or shot peening may also reduce thermal (and mechanical) fatigue problems by inducing residual compressive stresses into the surfaces of forging dies

Mechanical Fatigue

Unlike thermal fatigue, the literature on mechanical fatigue of die steels is sparse Perhaps the largest amount of data on this failure mechanism has been gathered by Thomas in a series of three-point bending experiments (Ref 68) Variables that he investigated included imposed load (of greatest importance in controlling fatigue behavior), hardness, material, position in the die block, strain rate, and temperature

Methods of Improving Resistance to Mechanical Fatigue. Methods of minimizing failure due to mechanical fatigue fall into one of two categories The first of these relates to die design and loading In this area, redesign of dies (flash design) or performs to lower die stresses may totally eliminate problems of mechanical fatigue Also, because of the logarithmic nature of fatigue-failure behavior, often only a slight decrease in applied loading can result in markedly improved fatigue lives Such a reduction in load may be obtained, for example, through better control of the forging energy in hammers and slight modification of the flash configuration in die design

The second major category of methods to improve fatigue resistance comes under the heading of material modification Modifications include treatments (such as shot peening) that put the surface layers of the dies into compression Another alternative is a total change of die material to one with a higher fracture toughness Such a material can support larger fatigue cracks before total fracture occurs

Catastrophic Die Failure/Plastic Deformation

The last two forms of die failure, catastrophic die failure and failure due to plastic deformation, will be discussed only briefly here The first of these, catastrophic die failure, can be considered a special case of mechanical fatigue failure in which the fatigue life is only one cycle It is usually a result of excessive forging stresses or improper die material/heat treatment selection or improper assembly of the die insert in the die holder Forging stresses may be high because of excessive forging energy or because of the general shape of the forging die cavity (Ref 69) With regard to the former,

once the dies have come together in a forging operation, excessive forging energy can only be dissipated by elastic deformation of the dies themselves Therefore, the fracture stress may be reached at points of stress concentration Among the details of the forging die cavity that affect stresses are the draft angles and corner radii

Improper die material selection can also result in catastrophic die failure Dies for hard-to-work materials or in which there are points with high stress concentrations require die steels with good fracture toughness These steels include low-alloy steels (such as 6F3 and 6F7) and some of the chromium hot-work die steels (such as H11) When the more highly alloyed die steels (such as H19 or H21) are employed (because of their wear resistance, for example), they should be tempered to lower-than-normal hardness in order to increase toughness if the dies are susceptible to catastrophic die failure

Failure of forging dies to perform properly because of plastic deformation can be measured by hot hardness or yield strength In general, the yield strengths of steels decrease with increasing temperature However, yield strength is also dependent on the prior heat treatment, composition, and hardness The higher the initial hardness, the greater the yield strength at various temperatures In addition, the yield strengths of different die steels increase with alloy content the tungsten hot-work die steels are harder than the chromium hot-work die steels, which are themselves harder than the low-alloy steels (Fig 2 and 3) Not shown in Fig 2 and 3 is the fact that the molybdenum hot-work die steels are even harder than the tungsten ones and thus manifest the greatest resistance to plastic deformation as far as die steels are concerned

Surface Treatments and Coatings

A variety of die coatings and surface treatments are available to extend the lives of dies limited by wear Among the most common coatings are plated chromium and cobalt-alloy deposits, which adhere to the die surface via a mechanical bond, and weld fusion deposits, which entail a metallurgical bond The latter can be used in rebuilding excessively worn dies as well Recently, the use of ceramic coatings (for example, carbide and nitride coatings) applied by processes such as chemical vapor deposition (CVD) have been found to extend forming die lives

The surface layers of ferrous forming dies can also be hardened by alloying them with nitrogen or boron Nitriding is the most common of the surface treatment techniques and can be accomplished using a gaseous, liquid, or plasma medium The plasma technique (ion nitriding) appears very attractive because the formation of brittle white layers, which are unavoidable in other nitriding processes, can be eliminated or at least minimized Boriding of die surfaces can also be accomplished using a variety of media, and increases in die life of the same magnitude as those obtained with nitriding have been reported

Coatings have been used extensively in net shape forming to reduce friction and wear Coatings can be applied to either the workpiece or the die In the case of the workpiece, the coatings are made of soft material with good adhesion, lubricity, and low shear strength In cold forming, for example, phosphate coatings are used to reduce interface friction and die wear (Ref 70) Resin-bonded coatings containing solid lubricants have also been successfully used (Ref 71, 72, 73) In hot-forming applications, hard coatings are generally used They are applied to the die or mold surface by mechanical, thermal, or chemical means

A hard surface layer reduces the frictional force and the wear rate when sliding against a relatively soft workpiece material if the coating/workpiece material pair is chemically stable and the coating is well bonded and mechanically compatible with the substrate (die material) (Ref 74) The role of the hard layer is to prevent plowing, whereas chemical insolubility is needed to ensure minimal dissolution Hard coatings are especially useful when the dominant wear mechanism is abrasive wear

Typical hard coating materials are oxides, carbides, nitrides, borides, and amorphous glasses One of the main considerations in the choice of a coating material is the quality of bonding between the coating and die material (substrate) The bonding could be chemical, mechanical, or both Chemical bonding is caused by a reaction or diffusion of atoms between the coating and the substrate to form a solid solution at the interface

Alloying Surface Treatments

The lives of steel forming dies limited by wear can often be increased by various surface treatments in which the structure

of the surface is alloyed with, for example, nitrogen, carbon, or boron Other common methods of surface hardening of steel (for example, flame hardening and induction hardening) have not been reported in the literature as having been used

for forming dies This is perhaps due to the large loss in toughness and distortion of the die cavity that these methods may cause

Nitriding is probably the most common treatment for hardening the surface layers of forging dies It is useful for applications in which the surface temperature does not exceed 565 to 595 °C (1050 to 1100 °F) in service As with most

of the surface treatment processes, nitriding finds its greatest application for press-forming dies in which strength and wear resistance are more important than toughness However, there are reports of nitriding being successful in impact applications such as hammer forging (Ref 75) Wear rates have been reduced by as much as 50% using nitriding (Fig 27) (Ref 75, 76)

Fig 27 Relative wear rates of nitrided ad non-nitrided tool steels used in extrusion forging Source: Ref 75

Nitriding processes are performed at temperatures between 495 and 565 °C (925 and 1050 °F) It is important that tempering of the die steel be performed at a temperature exceeding the nitriding temperature prior to nitriding in order to optimize the property combination of the core and the surface of the dies Also, because of the low nitriding temperatures, there is generally little distortion from this heat-treating process

Although the depth and hardness of the nitride case depends a great deal on the nitriding time, these properties (particularly the hardness) are sharply dependent on the composition of the steel as well Die steels containing large amounts of strong nitride formers such as chromium, vanadium, and molybdenum form shallow, very hard surface layers

On the other hand, low-alloy chromium-containing die steels (such as 6G and 6F2) form deeper surface layers that are tougher, but not as hard

Detailed information on specific techniques for nitriding gas nitriding, liquid (salt bath) nitriding, and ion nitriding can

be found in Volume 4, Heat Treating, of the ASM Handbook

Boriding and Carburizing. Boron can be added to surface layers by a diffusion treatment that can be carried out in either gas, molten salt, or pack media at a temperature between 900 to 1100 °C (1650 to 2010 °F), depending upon the process and the material to be borided Extremely hard surface layers with low coefficients of friction are formed, provided the base metal forms borides The process does not require quenching If the base material has to be heat treated, the heat treatment can be done after boriding, although care is required to reduce quenching stresses to prevent spalling of the borided layer

Very little is known about the usefulness of boriding forming dies Vincze (Ref 77) claims a 70% increase in die life with dies surface treated by boriding compared to untreated dies In this study, boriding was carried out by filling the die impression with a mixture of 10% B4C, 40% sodium borate, and 50% hardening salt, and heating the pack at 900 °C (1650 °F) for 3 h This diffusion heat treatment was followed by quenching and tempering Burgreev and Dobnar (Ref 78) also report large increases in hammer forging die life when boriding is used

Boriding of steels is also done electrolytically Boron atoms are electrodeposited onto the metal from a bath of molten salt containing fluorides of lithium, sodium, potassium, and boron The dies are borided in the 800 to 900 °C (1470 to 1650

°F) temperature range in an atmosphere of argon or a mixture of nitrogen and hydrogen Thickness of coating is from 0.013 to 0.05 mm (0.0005 to 0.002 in.), and treatment lasts 15 min to 5 h (Ref 79)

It has been stated that boriding results in undesirable interaction with alloying elements of hot-work die steels (H series) and develops a soft layer (Ref 78) Porosity in the borided layer can develop for steels that require postboriding heat treatment For this reason, it is preferable to limit boriding to those alloys that do not require further high-temperature treatment For example, A6 air-hardening steel can be hardened from the boriding temperature by cooling in air, and only requires tempering This steel, therefore, can be safely borided

Carburizing of hot-work die steels is uncommon and not popular for two main reasons (Ref 80) As with boriding, the high temperatures required for carburizing (815 to 1095 °C, or 1500 to 2000 °F) lead to distortion of the dies on cooling Secondly, the high-carbon surface layer, although it greatly increases hardness, can drastically reduce the toughness of the dies

Additional information on boriding and carburizing can be found in Volume 4, Heat Treating, of the ASM Handbook

Ion implantation is an alloying surface treatment that has been recently applied to forming dies The process was first developed in the late 1960s to introduce electrically active elements into semiconductors of microelectronic devices Now, more than 2000 commercial ion implantation systems are being used worldwide for semi-conductor processing (Ref 81)

Ion implantation research has increased steadily since the first publication of Hartley, et al (Ref 82) The description of

the process, its advantages and shortcomings, and the industrial application of the process have appeared in several review articles (Ref 83, 84, 85, 86, 87, 88, 89) Some ion implantation applications in forming are discussed below

Ion Implantation of Stamping and Cutting Tools. The number of applications involving the reduction of wear in stamping and punching tools probably exceeds all other uses of ion implantation for metals The dominant cause of failure

in high-speed tool steels used in metal stamping is adhesive wear Ion implantation of titanium and carbon has provided the optimum treatment for such tools For example, the life of punches and dies for manufacturing aluminum beverage cans has been improved by 6 to 10 times as compared with that of untreated tooling Similarly, punches for pressing powders into pellets have lasted twice as long after being treated These applications range from forming pharmaceutical pills to more demanding service such as pressing ceramic pellets Selected examples are listed in Table 9

Table 9 Examples of extending tool life via ion implantation

Draw die D2 Drawing 3.5 mm (0.140 in.) hot-rolled steel 22× life before polishing

Flanging ring D2 2 mm (0.080 in.) hot-rolled steel (flywheels) 80× life before polishing

Forming dies W1 1 mm (0.042 in.) cold-rolled steel Pickup reduced, improved finish on

Pilot pins M2 0.08 mm (0.003 in.) 301 stainless steel 5× life

Plastic forming Stainless steel Shaping vinyl siding Wear rate reduced for dies

Sprue bushings P20 (chromium

plated)

Injection molding thermoset (20% glass) At least 4.5× life

Taps M2 (chromium plated Tapping cold-rolled steel nuts 8× life

Ultrasonic electrodes D2 Welding aluminum to steel 2.5× life (mechanical wear)

Wire compacting Inconel, M2 Compacting copper-base composite wire At least 16× life

dies

Source: Ref 90

Ion implantation has also been used on a variety of low-temperature (to 200 °C, or 390 °F) cutting and slitting tools Knife blades for cutting synthetic fibers have been implanted with titanium and carbon to improve wear resistance Other applications include slitting blades for cutting rubber, hay baling cord cutters, plastic bandage cutter tools, corn-husking blades, paper slitters, and paper punches Typical life improvements of 6 to 10 times have been reported for most of these applications

Ion Implantation of Injection Molds. Another of the better-known application areas for the ion implantation process is that of injection molds for plastics and metal powders Simple implantation with nitrogen has provided several-fold improvements in the life of these tool steel molds, along with improved mold-release characteristics and reduced corrosion from chemical additives Molds used for powder metallurgy have also shown less tendency to fracture than untreated molds

Ion Implantation of Extrusion and Wire-Drawing Tools. Ion implantation has proved effective in treating components such as spinnerettes for nylon or rayon fibers, which have numerous tiny orifices The usual failure mechanism in these tools is a combination of corrosion and abrasive wear Implantation of titanium and carbon into the spinnerettes (usually made from stainless steel) has improved the resistance to the abrasive particles (titanium oxide) that are added to the synthetic material The treated spinnerettes have also been more resistant to the corrosive conditions encountered during the cleaning process Ion implantation of wire-drawing dies (up to 25 mm, or 1 in., in diameter) has not only increased the life of the dies, but also improved the surface finish of the drawn wire

References 90, 91, 92, 93, 94, 95 deal specifically with application of ion implantation in forming Additional information can also be found in the article "Ion Implantation" in this Volume

Plasma source ion implantation (PSII) is an ion implantation technique recently developed by J.R Conrad of the University of Wisconsin Madison (Ref 96) Objects to be implanted are placed directly in a plasma source and then pulse-biased to a high negative potential A plasma sheath forms around the target and the ions bombard the entire target simultaneously (Ref 97, 98, 99)

Carbide Coating by Toyota Diffusion Process. Good surface covering and strongly bonding carbide coatings, such as VC, NbC, and Cr7C3, can be formed on die steel surfaces by a coating method developed at Toyota Central Research and Development Laboratory, Inc of Japan (Ref 34)

In the Toyota Diffusion (TD) process, metal dies to be treated are degreased, immersed in a carbide salt bath for a specific time period, quenched for core hardening, tempered, and washed in hot water for the removal of any residual salt The borax salt bath contains compounds (usually ferroalloys) with carbide-forming elements such as vanadium, niobium, and chromium The bath temperature is selected to conform to the hardening temperature of the die steel For example, the borax bath temperature would be between 1000 and 1050 °C (1830 and 1920 °F) for H13 die steel

The carbide layer is formed on the die surface through a chemical reaction between carbide-forming elements dissolved in the fused borax and carbon in the substrate The carbide layer thickens due to reaction between the carbide-forming element atoms in the salt bath and the carbon atoms diffusing into the outside surface layer from the interior of the substrate

The thickness of the carbide layer is varied by controlling the bath temperature and immersion time An immersion time

of 4 to 8 h is needed for H13 steel to produce carbide layers with satisfactory thickness (5 to 10 m) for die-casting applications Dies are then removed from the bath and cooled in oil and salt or air for core hardening followed by tempering

The salt bath furnace consists of a steel pot with heating elements; no protective atmosphere is needed Selective area coating is accomplished by the use of copper or stainless steel masking, plating, thermal spraying, or wrapping with foils The type of carbide coating can be changed easily by using a different bath mixture or more than one pot, each containing different carbide mixtures

The process is applicable to most steels and some nonferrous metals Satisfactory results have been obtained for H12 and H13 steels, which are the most widely used steels for die-casting dies Coated steels exhibit high hardness and excellent resistance to wear, seizure, corrosion, and oxidation In addition, resistance to cracking, flaking, and heat checking is claimed Hardness of the coating depends on layer composition; 3500 HV for vanadium carbide, 2800 HV for niobium carbide, and 1700 HV for chromium carbide

Additional information on the TD process can be found in Ref 100 and in Volume 4, Heat Treating, of the ASM Handbook

MetalLife Treatment

A new surface treatment for dies and molds, MetalLife, utilizes a controlled micropeening method to develop rough texture and beneficial compressive stresses on the surface of dies and molds (Ref 101, 102, 103, 104) Rough texture increases die-lubricant retention characteristics and the residual compressive stresses inhibit the initiation of fatigue microcracks due to the mechanical or thermal cycling of die surfaces during operation

Shotpeening of surfaces is not a new concept, but the conventional process is not very suitable for die and molds because

it results in pitting and stress raisers on the surface that can result in premature tooling failure Uncontrolled use of a smaller blasting media can result in erosive and abrasive wear The key variables that need control are the media size, concentricity, angle of impingement, velocity, and dwell time for each media used Sometimes multiple treatments have

to be performed

Punch and die life increases of 6 to 10 times are claimed using the MetalLife treatment Drawing and forming dies have shown life improvements of 10 to 20 times in certain cases In the field of die casting, MetalLife has been used to close heat-checking cracks and increase the life of dies

Ceramic Coatings

There are a number of ceramic coatings that can be applied by various means to metal parts to improve their service properties; these ceramic materials are electrically nonconductive, have up to 20 times the abrasion resistance of metals, and can withstand temperatures in excess of 2480 °C (4500 °F) (Ref 105) Among the many ceramic wear-resistant materials available for coatings are titanium carbide, titanium nitride, and chromium carbide These materials can be applied to chromium hot-work steels and the air-hardening tool steels by the chemical vapor deposition (CVD) process (Ref 106) In this process, the metal part to be coated is placed in a special reactor vessel after which it is heated and reacted with gas containing coating materials species Selection of suitable coatings and metals depends strongly on the compatibility of the two from a thermal expansion viewpoint If the expansion coefficients are widely different, the coating may crack when the part is cooled to room temperature Because a surface interdiffusion layer is also produced, the possibility of forming soft or brittle compounds must also be considered From previous discussions, it is known that these considerations are also important from the perspective of expected performance during forging A good match of thermal properties is required to prevent heat checking, and tough, hard surface layers are needed to offer resistance to wear and brittle fracture Besides being hard, these coatings generally have good lubricity However, because of oxidation problems, they must be used at temperatures below 650 °C (1200 °F), which is above the typical operating temperature of forming dies made of hot-work die steels To date, CVD coatings have been used for forging dies to a limited extent only

In one application, use of a TiN coating on H26 press-forging dies increased die life from approximately 18,000 (uncoated) to 51,000 pieces, as compared to a die life of 36,000 pieces using chromium-plated dies (Ref 107) The use of these coatings is sure to increase in the future, particularly in applications where life is limited by abrasive wear

Evidence that ceramic surface coatings can extend the life of hot-forging dies has also been demonstrated in die wear trials on H12 (Ref 108, 109) In these trials, die life increases in excess of 10% were reported

Other ceramic coatings that may find hot-forming application include alumina, zirconia, chromium oxide, and magnesium zirconate All of these coatings have excellent wear resistance, especially chromium oxide, which has a diamond pyramid hardness of 1200 kg/mm2 with a 300 g load; aluminum oxide has a slightly lower hardness, 1100 kg/mm2 with a 300 g load (Ref 105) Several ceramic coatings provide excellent thermal resistance in excess of 2480 °C (4500 °F) Zirconia has a melting point close to 2480 °C (4500 °F) and is extremely resistant to thermal shock Magnesium zirconate, with a melting point near 2150 °C (3900 °F), and yttria-stabilized zirconia, with a melting point near 2650 °C (4800 °F), are used as thermal barrier coatings

Ceramic coating application requires the following steps:

• Oxygen acetylene powder

• Oxygen acetylene rod (welding)

• Plasma spraying (torch)

• Detonation gun

In the oxygen acetylene powder method, the powder is fed into a flame at 2760 °C (5000 °F) and sprayed on the substrate

by compressed gas The coatings produced by this method are generally porous with low adhesion This process is of moderate costs

In the oxygen acetylene rod method, fused ceramic material in a rod form is introduced into a 260 °C (500 °F) oxyacetylene torch Molten ceramic is sprayed at speeds up to 170 m/s (550 ft/s) via compressed gas on the target This results in a coating with high cohesive bonding

In plasma spraying, ceramic powder is introduced into a plasma (ionized gas) at temperatures as high as 16,650 °C (30,000 °F) The high-pressure plasma gas accelerates molten particles on the target This method produces well-bonded high-density coatings, but is very expensive

Extremely dense coatings are produced by the detonation gun process This process is preferred when tungsten carbide coatings are to be applied An explosion of oxygen and acetylene gases produces 3315 °C (6000 °F) temperatures, melting the ceramic and producing a molten jet that impinges the target at speeds up to 760 m/s (2500 ft/s)

A novel method for coating metalworking dies with refractory metals has been patented by a group of researchers at United Technology Corporation (Ref 110) In this method, a refractory metal coating is sprayed by a plasma gun and subsequently compacted under conditions of minimum shear stress Refractory metals selected include molybdenum, niobium, tantalum, tungsten, rhenium, and hafnium because they have melting points in excess of 2200 °C (4000 °F) and sufficient ductility for compaction The compaction of the coating is achieved by processing a workpiece, which has previously been formed to the end shape, through the die This pressing of a preformed workpiece reduces the metal flow and shear stresses to a minimum, thereby avoiding shear and spall of the coating Refractory-coated H13 tool steel dies have exhibited significant improvements in wear resistance

Electroplating

Chromium Plating. Chromium is usually applied to metal pieces using electroplating baths composed of chromic acid and some sulfate or fluoride compound (Ref 111) Bath temperatures are between 45 and 65 °C (110 and 145 °F) The kind of plating used for forging dies is called hard chromium plating and results in surface deposits typically between 25 and 500 m (1 and 20 mils) thick After application of the plating, the die cavity is ground to finish dimensions

Chromium plating has been used to a modest extent on industrial forging dies, but there is conflicting evidence as to its value Dies with deep cavities, sharp corners, or projections that show cracking due to thermal or mechanical fatigue should not be chromium plated This may be due to the tendency of hard chromium platings to contain microcracks that

open during cyclic loading On the other hand, dies for thin forgings that must be discarded because of wear (especially at the flash land), and best suited for chromium plating (Ref 112)

Cobalt Plating. As with chromium, various cobalt alloys have been applied to hot-forging die steels, primarily to extend life through reduction of die wear (Ref 113, 114, 115, 116, 117) Typically composed of alloys of cobalt and tungsten or cobalt and molybdenum, these coatings are applied in electroplating baths using so-called electroplating brushes, which allow a small or selected area to be plated Besides offering improvements in wear resistance, it appears that these coatings also possess a low coefficient of friction and can in some cases solve problems involving sticking of the workpiece Many of the criticisms often leveled against coatings, namely loss of adhesion or flaking due to poor resistance to shock loading, appear not to apply to these coatings Hence, it is not surprising that increases of die life up to 100% in press-forging operations are not uncommon with the use of these coatings The kinds of parts for which these improvements have been obtained are varied and include gear levers, turbine blades, and suspension end-sockets for cars

Hardfacing

Hardfacing is a weld fusion process that produces deposits which are metallurgically bonded to the substrate In the early days, hardfacing was used for repair and maintenance of die and molds It is now being used increasingly as an inexpensive means for depositing a hard layer on localized wear-prone die areas

For dies and molds, these deposits have the following applications:

• Deposits of identical material onto a die block to repair it or to allow resinking of it

• Deposits of higher-alloy steels (for example, chromium hot-work steels) onto the die surface of alloy steels to improve the service performance of the dies

low-• Deposits of hard or high-temperature materials (usually cobalt- or nickel-base alloys) onto low-alloy or hot-work steels to improve the service performance of the dies These alloys come under the general heading of hardfacing or hard-surfacing alloys

Hardfacing Processes. Before discussing specific alloys, the processes by which they are deposited will be briefly reviewed The first step in any of the hardfacing processes should be the annealing of the die block into which the rough impression has been sunk (Ref 118) This relieves residual stresses and helps prevent cracking during welding of the surface layer After annealing, the die block should then be reheated to a temperature of 325 to 650 °C (600 to 1200 °F), which is also necessary to minimize cracking due to thermal gradients set up between the surface and the interior during welding The application of the surface layer can then be performed by one of a number of welding processes (Ref 119, 120):

• Gas torch welding (combustible gas welding)

• Manual arc welding

• Submerged arc welding

• Gas shielded arc welding (TIG or MIG)

• Open arc welding

• Thermal spraying

• Fusion treatment

• Plasma spraying (plasma arc welding)

• Transferred arc plasma

• Flame plating

• Deposition process (electroslag welding)

Together with the solidification conditions, the amount of melted base material and base-material dilution is important for wear properties Hardfacing methods differ considerably from each other and also compare with powder spray methods, which show almost no base-material dilution due to mixing

Combustible-gas welding offers many advantages in depositing smooth, precise surfaces of high quality This is done by using a carburizing flame that causes "sweating," or welding of a thin surface layer that spreads freely and prevents metal

buildup For repair of dies, the shielded metal arc method is preferred It allows high productivity and has the advantage

of low heat input and thus minimal distortion of the die cavity

After welding, the die block must be cooled to room temperature to prevent cracking of the weld deposit The die impression is then finished, machined, and ground Heat treatment (austenitizing, quenching, and tempering) of the die block is performed last Once again, differences in thermal properties between the base metal and surface deposit are critical insofar as thermal cracking is concerned Because this is also an important consideration in the performance of forming dies, it is not unusual that welding alloy suppliers are sometimes hesitant about recommending many combinations of die block material and hardfacing alloy (Ref 121)

Hardfacing Alloys. For hardfacing, welding alloys are generally based on iron, cobalt, or nickel Hard phases are formed by addition of carbon (in iron) or boron (in nickel) The volume fraction of hard phase is very important for the wear resistance in the weld deposit Often there is no proportional dependence and the best wear resistance is not achieved by the highest hard-phase concentration

Various ferrous alloys are used to repair steel dies or to lay down deposits with better wear and heat resistance than the substrate These alloys are very similar to the low-alloy and hot-work tool steels in composition Austenitic and austenitic-ferritic materials are preferred for wear resistance under heavy loads

The use of nickel- and cobalt-base alloys in hardfacing offers a considerable cost savings over die blocks made of these alloys In a typical hardfacing operation, a one- or two-alloy layer, each about 0.25 to 1.25 mm (0.010 to 0.050 in.) thick, are deposited on the die If a large amount of buildup is desired or required, however, it is advisable to apply layers of stainless steel or low-alloy filler metal first rather than many layers of the more expensive nickel or cobalt hardfacing materials

Detailed information on the methods for depositing hardfacing alloys and their resistance to wear can be found in the article "Friction and Wear of Hardfacing Alloys" in this Volume

Electro-spark deposition (ESD) is a variation of hard surfacing that has been used extensively in Europe for improving the galling resistance of material (Ref 122) Electrodes of WC, TiC, and Cr3C2 materials have been deposited

on type 316 stainless steel and other substrates The ESD process has been found to be effective in fusing metallurgically bonded coatings to the substrate at low heat with the substrate remaining near the ambient temperature

Hard Coatings for Cold Extrusion

In the cold extrusion of steel, compressive loads up to 3000 MPa (435 ksi) and tensile loads up to 1500 MPa (220 ksi) are not unusual Core hardening (or through hardening) of the tool steel makes the punches brittle, leading to early failures The highly stressed tools, therefore, either should be made from tungsten carbide (a costly material) or they should be hard coated

A very detailed study of chemical and physical treatments for backward can extrusion has been reported by Westheide, et

al (Ref 123) They divided the surface treatments used into reaction- and coating-layer processes (Fig 28) In the

reaction layers, the layer element diffuses into the substrate (die material); in the coating layers, the primary adhesion mechanism is mechanical interlocking Figure 28 also identifies the type of coating that can be deposited on the substrate

Fig 28 Coating processes carried out on die materials used for cold extrusion Source: Ref 123

In this study, billets of case hardening steels (similar to AISI 5120) were used Punches were either made of a working tool steel (similar to AISI D2) or of high-speed tool steel (similar to AISI M2); the former hardened to 62 HRC and the latter to 64 HRC Backward can extrusion experiments were carried out on a 630 kN (70 ton) press at 40 strokes/min The billet height to internal diameter ratio was unity and soap was used as the lubricant

cold-A comparison of treatments for backward can extrusion is given in Fig 29 for 10,000 extrusions Lowest wear rates were obtained for PVD TiN coatings and TD vanadium carbide (vanadized) coatings

Fig 29 Comparison of coatings for backward can extrusion Source: Ref 123

The application of TiN coatings in cold extrusion is reviewed in Ref 123 and 124 Some of the cold extrusion applications where TiN coatings have proved beneficial are hexagonal socket press tools, tools for spur gear teeth, and hydraulic valve stem housings (Ref 123) Vanadium carbide coatings deposited by the TD process have been applied to extrusion dies in rubber forming Die life increases from 30 h for hardened steels to 900 h for coated steels have been found in the production of rubber window seals (Ref 104)

Surface Treatments and Coatings for Cold Upsetting

Westheide (Ref 123) also carried out similar studies on protective wear-resistant coatings for upsetting dies A qualitative comparison of various coatings is provided in Table 10 Hard chromium plating and TiC coatings via CVD were found to have excellent wear resistance

Table 10 Comparison of selected coatings for cold upsetting

Nitriding and nitrocarburizing P G M Chipped off

Note: E = excellent, G = good, M = moderate, and P = poor

Source: Ref 123

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Friction and Wear in the Mining and Mineral Industries

D.L Olson and C.E Cross, Center for Welding and Joining Research, Colorado School of Mines

Introduction

THE PRACTICE of mining and mineral processing, by its nature, involves severe mechanical interactions between metals, as well as between abrasive nonmetallic and metallic materials The abrasive nature of most ores can cause significant wear to both handling and processing equipment

Various wear processes are involved, depending on the nature of the abrasive material, type of loading, and the environment The direct impact of metallic components with the earth requires alloys that have excellent ductility and work-hardening properties The transport of ground material across a surface, such as chutes, requires high surface hardness The crushing or fragmentation of ore requires even higher hardness values Performing mechanical operations

in corrosive environments, such as mineral slurry pumps, requires that the material exhibit both corrosion and wear resistance

To properly select wear-resistant materials, careful analysis of the material application is necessary This selection should consider a very broad range of alloys, including high-strength steels, white irons, austenitic alloys, and high-chromium/high-carbon stainless steels

Types of Damage

Abrasive wear, the most common form of damage in the mineral industry, can be classified as being due to low-stress, high-stress, or impact abrasion (Ref 1, 2) The material-ore interaction is considered a low-stress abrasion application when the wear process does not involve a fracture of the abrasive material

An example is the movement of fragments of ore traveling across a high-strength steel surface (Fig 1a) Low-stress abrasion is not accompanied by significant impact The damage is the result of either scratching or a micrometal removal process Sharp, angular abrasives result in the highest wear rate For this case, materials with high hardness are used to minimize the metal removal rate Abrasion-resistant (AR) steels, alloy cast irons, and ceramic tiles are often used for these applications

Fig 1 Three types of abrasive wear, (a) Low stress (b) High stress (c) Impact Source: Ref 1

High-stress abrasion results when abrasive materials are caught between two normally loaded surfaces The loads are commonly sufficient not only to cause particle penetration of the loading surfaces, but can also fragment the abrasive material This damage is common to ball mills, rock drills, and rock crushers (Fig 1b) Surface damage is due to a combination of scratching and plastic deformation, commonly resulting from fatigue The damage can be minimized by the selection of materials with both high yield strength and hardness

Some sites on mining equipment will be subject to direct impact, which requires those sites to have both hardness and toughness The high localized pressures on impact cause the abrasive materials to cut into the wearing surface, resulting

in very apparent large gouges and scratches This type of damage (Fig 1c) is common to the impact areas of chutes, pulverizing mills, and shovels (Ref 1) The material of choice is often an austenitic alloy that has high toughness and progressively increase in hardness with use through work hardening (Ref 3, 4) High-manganese steels with greater than

14 wt% Mn (known as Hadfield alloys) are examples of traditional impact-resistant austenitic alloys

Corrosive Wear. Wear is primarily the removal of surface metal by mechanical action Often, the mining and mineral processing environment is wet, which cause a chemical action that affects the mechanical wear (Ref 5 , 6, 7, 8, 9) This corrosion effect either increases or decreases the metal removal rate, depending on the natures of both the mechanical

interaction and the corrosion product The synergistic effect of mechanical and chemical interaction causes considerable variation in reported wear rates on mining and mineral processing equipment

Adhesive wear between one metallic part and another is also a common form of damage in the mineral industry It results from the surface tearing of materials that became bonded because of rubbing (Ref 1) The damage involves scoring, galling, or seizure The adhesive wear rate is usually larger when the interacting materials have similar compositions and crystal structures, and it usually occurs under load with little or no lubrication A common location for adhesive wear is at a surface where wire rope rubs against spools and pulleys, for example Figure 2 illustrates adhesive wear on a steel spool that has been interacting with wire rope

Fig 2 Cable wear on large spool, a form of adhesive wear Source: Ref 10

Testing for the types of wear damage described above is often performed according to ASTM test procedures (Ref 11,

12, 13), as well as other special tests designed by specific corporations to evaluate special conditions (Ref 14, 15, 16, 17)

Methods to Improve Wear Resistance

Wear resistance can be increased on mining and mineral processing equipment through proper material selection and equipment design, and through the use of wear plates, as well as hardfacing deposition The method selected will depend

on the type and use of the part and the economics associated with operation of the equipment

Material Selection. A variety of materials that provide wear resistance are commercially available (Ref 18, 19) Some material types are shown in Fig 3 relative to service and environmental needs

Fig 3 Relative placement of various wear-resistant alloy systems when considering abrasion, impact, and

corrosion resistance Source: Ref 20

High impact resistance with low abrasion and corrosion resistance can be achieved with high manganese ferrous alloys A Hadfield alloy (14Mn-1C) is a traditional material for impact resistance In both the as-cast and as-welded conditions, the Hadfield alloy shows good ductility Additionally, the alloy work hardens with impact and deformation, resulting in increased strength and hardness levels (Ref 21) For more corrosive environments, impact resistance can be achieved with the 18-8 (18%Cr + 8%Ni) stainless steels Austenitic stainless steels, such as the 18-8's, also work harden with service, but have the disadvantage of high material cost

Abrasion-resistant plate steels, similar to armor-grade steel, have a unique combination of desirable properties: high hardness, good toughness, and weldability (Ref 22) Abrasion resistance is achieved using a water quenched and tempered martensite, which is often low enough in alloy content (that is, has a low carbon equivalent) to avoid the concerns of weld underbead cracking with proper welding practice Vacuum degassing and desulfurization during steelmaking is required

to minimize inclusions and thus improve toughness in the short transverse direction

As the abrasion resistance increases from a low-stress to high-stress application, the selected material must increase in hardness from martensitic alloy steel to high-carbide materials, including alloyed white irons Applications for martensitic steels of various hardness levels are given in Table 1

Table 1 Materials for surfacing, build-up, and hardfacing

Low hardness

Pearlitic steels Low cost, crack resistant

Mild steels An excellent base for hardfacing

Low-alloy steels Some toughness for build-up of worn areas

A build-up material to restore dimension or as a low-cost base for surfacing applications

Austenitic steels Excellent for heavy impact, tough For metal-to-metal wear under heavy impact conditions

Oxidation and hot wear resistant Furnace parts, red heat frictional wear

14% Mn, Cr-Ni High yield strength austenitic Build-up, crack repair, joining manganese alloys or

manganese to mild steel

14% Mn, 1% Mo Fair abrasion and corrosion resistance, work Railways track build-up

hardens

Semiaustenitic steels Low-cost crack-resistant materials Hardfacing applications

Copper alloys Resistance to seizing under frictional wear

For applications where hot hardness is important

Nickel-chromium Resistance to oxidation For corrosive wear conditions

High compressive strengths

Austenitic alloy irons Better crack-resistant properties than the

martensitic alloy irons

Where erosive wear with or without light impact is present

Chromium-molybdenum

Good crack-resistant qualities Pump and turbine build-up

Nickel-chromium Crack resistant, light impact For erosive wear with light impact

Martensitic steels Abrasion resistance with medium impact

Low carbon (up to

Fair abrasion resistance

For a variety of abrasive wear conditions where there may also be involvement with medium impact

High hardness

High-chromium irons Excellent erosion-resistant properties For general use where hot gases or materials are involved

Austenitic Good abrasion-resistant qualities For farm and earth-moving equipment working in sandy

soils

Martensitic Ability to be rehardened after annealing As a hardfacing in steel mills and refineries where hot

erosion (595 °C, or 1100 °F) is a wear problem

Tungsten-molybdenum alloys

High red hardness For hardfacing of coke oven parts as well as other hot

(425-650 °C, or 800-1200 °F) steel mill applications

Good abrasion resistance, brittle

For most hardfacing applications, but best suited to those applications where hot wear and abrasion (above 650 °C, or

1200 °F) are involved (jet engine turbine blades, gas engine exhaust valve)

Tungsten carbides The ultimate in abrasion-resistant qualities

A variety of materials used for hardfacing to meet an extremely wide range of severe abrasive conditions, especially oil well drill bits, tool joints, rock drill bits

Source: Ref 23

The role of carbon content in steel and iron to promote abrasion resistance is illustrated in Fig 4, which indicates the position of various alloy types relative to their gouging wear rates Figure 5 correlates wear rate to hardness for the various alloys experiencing abrasion wear

Fig 4 Relative placement of various alloys as function of carbon content and abrasion resistance, as measured

by gouging wear ratio Source: Ref 24

Fig 5 Abrasion resistance of various alloy systems, comparing wear rate to hardness Source: Ref 25

Use of Wear Plates. A part that experiences major wear should be designed for easy replacement to prevent excessive processing downtime Figure 6 shows high-manganese steel wear plate castings that are welded into the bucket of a large shovel These castings were made with properly spaced holes, which alloys easy shielded metal are welding of the cast plates to the shovel Similar wear plates are used as working surface in crushers (Ref 26, 27) and pulverizers

Fig 6 High-manganese steel-casted wear plates plug welded to inside of large shovel Source: Ref 10

Figure 7 illustrates the use of wear plates in medium-sized buckets Often, an easily weldable steel plate or strip that is considered expendable is used A typical practice is to incur significant wear in a new bucket to reduce its weight prior to welding these wear plates to it Bucket weight is an important consideration in the efficiency and economics of earth-moving processes

Fig 7 Steel strips welded onto surface of shovel and hardfacing deposition placed on top of steel strips to give

abrasion and impact resistance for a rack consisting of large racks Source: Ref 10

Table 2 lists the various ferrous alloys used for mill liners They include a broad range of high-carbon steel and alloyed white irons

Table 2 Ferrous materials for grinding mill liners

Composition range, %(a) Item Material

Hardness range,

HB(b)

Relative wear rate(c)

1 Martensitic Cr-Mo white iron

2.4-3.2

0.5-1.0

0.5-1.0

23.0

14.0- 3.0

1.0-0-1.5

0-1.2 620-740 88-90

2 Martensitic high-carbon Cr-Mo

steel

1.2

11.0- 1.0

(a) The composition range from an individual supplier will normally be narrower than the ranges given in the

table Production of certain compositions falling within the above ranges may involve proprietary rights covered by patents or trademarks

(b) Hardnesses listed are on the unworn surface of the liners Austenitic and martensitic alloys tend to work

harden on wearing surfaces

(c)

Relative wear rate as determined when wet grinding minus 10mm ( in.) feed to minus 48 mesh on ore containing about 65% quartz, 25-30% feldspar, and 3% pyrite as the principal abrasives, in primary ball mills at Climax, Colorado

(d) A composition containing about 1.0% C, 5-6% Cr, and 1.0% Mo was used as a comparative standard in all

tests and assigned a relative wear rate of 100

(e) Wear rates in the Climax mills could not be determined on this material, because of spalling and breakage

Use of Hardfacing Deposition. Three primary methods are used to deposit hardfacing materials in the mineral industry (Ref 28, 29, 30, 31, 32, 33, 34, 35): arc welding, oxy-fuel gas welding, and thermal spraying Filler metals are available in rod, wire, and powder form

The most traditional process is shielded metal arc (SMA) weld deposition, which involves low capital investment, can be used in all positions, and is an excellent field welding process It is considered a low deposition rate process A large selection of commercial electrode compositions exist

The gas-metal arc welding process is used to make wear-resistant depositions The gas-metal arc wire electrode process increases deposition rate and productivity, but usually requires more equipment setup This gas-metal arc weld deposition process is more flexible than the submerged arc welding process, but it is more difficult than shielded metal arc welding

in terms of moving the welding equipment, including the necessary gas supply for protective cover

The submerged arc welding process deposits in either the flat position or on top of a rotating round part and is capable of high deposition rates It is a fully automatic process that needs to be performed in a stop A wide range of alloy consumables are available, including strip electrodes

Flux-cored electrodes that introduce flux and carbide-forming ingredients to the weld deposit, by additions to the core of the welding wire, are available These flux core additions can also contain generators of plasma and shielding gas which can produce a wire welding process that is self shielding and therefore does not require additional supplied shielding gas

Hardness is commonly used as a measure in the selection of hardfacing materials The abrasion resistance of martensitic alloys is expected to increase with increasing hardness, whereas impact resistance would be obtained from materials of low hardness An extensive list of hardfacing and surfacing materials is given in Table 1, along with their specific advantages The alloy compositions and AWS designations for these various hardfacing materials are given in Table 3

Table 3 Classification and Composition of typical hardfacing alloys

Typical composition, % Group and description

1 Iron-base containing less than 20% alloy additions

Carbon steel bal 0.5 250

Martensitic alloy steel bal 0.1 1 0.7 250

Martensitic alloy steel bal 0.1 3.5 1 350

Martensitic alloy steel bal 0.25 1 1 0.3 350

Martensitic alloy steel bal 0.35 3 1 450

Martensitic alloy steel bal 0.45 5 1 5 0.75 650

Martensitic stainless steel bal 0.1 12 400

Martensitic stainless steel bal 0.25 13 450

High-speed steel bal 0.8 4 0.5 5 2 6 650 Fe 5-A(a)

High-speed steel bal 0.7 4 0.5 7 1 1.5 650 Fe 5-B(a)

High-speed steel bal 0.4 4 0.5 7 1 1.5 600 Fe 5-C(a)

Austenitic Mn steel bal 0.7 0.5 14 4 600 (max) Fe Mn-A(a)

Austenitic Mn steel bal 0.7 0.5 14 1 600 (max) Fe Mn-B(a)

2 Iron-base containing more than 20% alloy additions

Austenitic Cr-Mn steel bal 0.35 14 14 1 0.4 600 (max)

High-speed steel bal 10 15 2.5 750

Austenitic steel bal 0.1 18 3 8 500 (max)

Austenitic iron bal 4 30 6 700 Fe Cr-Al(a)

Martensitic iron bal 2.5 28 1 600

E Ni Mo Cr-1

4 Carbides

AWS A5.21-70 Carbides WC 20/40

>1800 HV WC-40/120 and so forth Arc deposit WC-40/120 and so forth

Tubular rods 50-60% tungsten carbide granules 40-50% Fe

>1000 HV

Sintered rods 50-80% tungsten carbide, up to 10% Cr or Ni, 10-50% Fe >900 HV

Source: Ref 18

(a) Also covered by AWS A5.21-70, "Composite Surfacing Welding Rods and Electrodes."

Use of Design. The geometric pattern of hardfacing deposition also has a role minimizing the wear rate Figures 8, 9,

10, 11, 12, and 13 illustrate various patterns in which hardfacing deposits can be laid Figure 9 suggests that the flow of fine material be perpendicular to stringer bead deposits, whereas with coarse ore, it is common to align the flow of ore with the stringer bead deposits The larger rocks tend to knock off the perpendicular beads Ores with mixed sizes require the spaced weave pattern The general concept is for the interspaces to be filled with fines, which allow the ore to grind upon ore, thereby reducing metal loss The spacing between deposits is directly related to the size distribution of the ore

Fig 8 Typical pattern of hardfacing deposit on lip and teeth of excavator bucket Complete coverage or

hardfacing deposit usually necessary only at edges of lip and teeth Source: Ref 35

Fig 9 Different patterns for hardfacing deposits, depending on size and nature of ore material Source: Ref 32

Fig 10 Hardfacing deposition on lip of shovel that must resist both abrasion and impact Source: Ref 10

Fig 11 Typical pattern of hardfacing deposits onto sides of shovels Source: Ref 10

Fig 12 Effects of various hardfacing coverages on shovel teeth (a) Hardfacing of both top and bottom teeth

(unacceptable) (b) Hardfacing just bottom of tooth (unacceptable) (c) Hardfacing just top of tooth, which is recommended practice that results in self-sharpening of tooth use Source: Ref 32

Fig 13 Worn shovel tooth often rebuilt by welding a new repointer bar, usually made of high-manganese steel,

onto end of tooth prior to depositing hardfacing materials Source: Ref 32

For locations that experience impact loading, it is common to use complete coverage of the hardfacing deposits Scoop buckets are commonly hardfaced on the digging edge using overlapping stringer beads, with the side of the bucket having

a diamond bead pattern, as shown in Fig 10 and 11

Special rules are used for hardfacing deposits on scraper teeth The teeth become blunt, either without hardfacing or with hardfacing on both the top and bottom of the teeth (Fig 12) Shovel teeth should be hardfaced when they are new Hardfacing should be deposited with close spacing, if not overlapped, from the point of the tooth back to 50 mm (2 in.) from the point (Fig 9) The proper method is to deposit hardfacing material only on the top of the tooth This practice will allow the tooth to preferentially wear, resulting in a self-sharpening effect (Fig 12) If the point breaks off, then it is often built up with either steel or 14 wt% manganese steel (Fig 13)

Hardfacing materials commonly have thermal expansion coefficients that are very different from the substrate material on which they are deposited This thermal expansion mismatch causes residual stresses The hardfacing materials commonly have thermal expansion values that are larger than the base structural material, which promotes longitudinal tensile stresses in the deposit

In many hardfacing materials, it is common to have cross-checking, which is a form of transverse cracking that is desirable because it relieves the stresses produced by thermal expansion mismatch between weld deposit and base metal (Fig 14a) If cross-checking does not occur during hardfacing, then the combined residual and externally applied stresses can cause hairline cracks under the weld deposit in the heat-affected zone of the base material (Fig 14b), which can result

in spalling during service

Fig 14 (a) Sufficient cross-checks (b) Insufficient cross-checks Source: Ref 32

In situations where crack-free hard surface deposits are necessary, welding practice that involves a carefully selected welding electrode with preheat and post-heat treatments is necessary (Ref 32) Recommended practices are available from many of the harfacing electrode manufacturers

Ball and Rod Grinding Media

Grinding media, in the form of balls and rods, are made of various alloys (Ref 3, 24, 36, 37, 38, 39, 40, 41) The mining industry consumes cast steel, Ni-hard, and high-chromium white iron grinding balls Grinding rods are commonly hot-

rolled high-carbon steel (AISI 1080 to 1095), as well as AISI 52100 steel The grinding media need to be harder than the ore to be ground The composition and hardness of typical steel and white iron grinding media are compared in Table 4

Table 4 Composition and hardness of typical steel and white iron grinding media

Pearlitic white iron 460 2.75 0.30 0.75

Martensitic white iron 600 3.23 0.62 0.64 1.89 4.26

Cast iron grinding balls are iron-carbon alloys that contain free carbides In unalloyed cast iron balls, iron carbide is the phase that influences hardness With up to 10 wt% chromium, these cast iron balls have mixed carbides, including (Fe,Cr)3C, and hardness values from 800 to 1200 HV At levels above a 10 wt% chromium addition, (Fe,Cr)7C3 carbide is also present, resulting in hardnesses that ranges from 1300 to 1800 HV The nature of these carbides is significant to the wear rate of the grinding media (Ref 38)

Low-alloy pearlitic white iron grinding balls offer a lower-cost alternative to alloyed irons and can be used with many ore types Pearlitic white irons can have up to 2 wt% chromium content and their hardness values are not extremely high, usually ranging from 450 to 600 HV They are available in either as-cast or stress-relieved conditions (Ref 38)

Heat-treated martensitic white iron balls are also available These alloys are usually alloyed with nickel and chromium to allow the formation of a martensitic matrix during the normal cooling process An example of this type of white iron is Ni-hard cast iron, which has from 2.5 to 4.5 wt% nickel and from 1.5 to 2.5% wt% chromium Ni-hard cast iron balls have hardness values ranging from 600 to 750 HV, depending on ball size and carbon content (Ref 38)

Martensitic ductile iron balls with hardness values between 600 and 700 HV offer high hardness with higher toughness than is commonly found with pearlitic and martensitic white iron balls The microstructure consists of a discontinuous carbide network and small graphite nodules in a martensitic matrix The graphite nodules assist in controlling the amount

of free carbides and in promoting the discontinuous carbide network Martensitic ductile irons are cast to yield carbides and then water quenched to produce a martensitic structure The toughness can be improved by tempering (Ref 38)

References

1 V.H Davies and L.A Bolton, The Mechanism of Wear, Weld Surfacing and Hardfacing, The Welding

Institute, Cambridge, UK, 1980, p 4-10

2 D.J Dunn, Metal Removal Mechanisms Comprising Wear in Mineral Processing, Proceedings of

International Conference, Wear of Materials 1985, American Society of Mechanical Engineers, 1985, p

501-508

3 H.A Fabert, Jr., Manganese Steels in Cracking and Grinding Service, Proceedings of Materials for the

Industry, Climax Molybdenum Co., AMAX, 1974, p 163-168

4 J Tasker, Austenitic Manganese Steel Fact and Fallacy, Proceedings of Intermountain Minerals

Symposium, Climax Molybdenum Co., AMAX, 1982 p 3-20

5 W Day, Corrosion and Erosion Effects on Materials for Slurry Pumps, Proceedings of Intermountain

Minerals Symposium, Climax Molybdenum Co., AMAX, 1982, p 79-82

6 I Iwasaki, K.A Natarajan, S.C Riemer, and J.N Orlich, Corrosive and Abrasive Wear in Ore Grinding,

Proceedings of International Conference, Wear of Materials 1985, American Society of Mechanical