Volume 01 - Properties and Selection Irons, Steels, and High-Performance Alloys Part 13 pps

Pease, Mechanical Properties of Steam Blackened P/M Materials, in Modern Developments in Powder Metallurgy, Vol 18-21, Metal Powder Industries Federation, 1988 Ferrous Powder Metallurg

(a) None, as sintered; Light blackening, 2 h exposure in 538 °C (1000 °F) steam; heavy blackening, 4 h exposure in 538 °C (1000 °F) steam

(b) Unnotched Charpy test at room temperature

References cited in this section

8 H Ferguson, Heat Treatment of P/M Parts, Met Prog., Vol 107 (No 6), June 1975, p 81-83; Vol 108 (No

2), July 1975, p 66-69

9 L.F Pease III, J.P Collette, and D.A Pease, Mechanical Properties of Steam Blackened P/M Materials, in

Modern Developments in Powder Metallurgy, Vol 18-21, Metal Powder Industries Federation, 1988

Ferrous Powder Metallurgy Materials

Revised by Leander F Pease III, Powder-Tech Associates, Inc

Re-pressing

As a secondary mechanical forming operation performed at room temperature, repressing is done primarily to increase density, which increases mechanical and physical properties and hardness Improvements in part dimensions can also be achieved by re-pressing The amount of material deformation achieved with repressing is greater than in sizing because the forces used are greater than the sizing forces The reduction in height of a ferrous part generally ranges from 3 to 5%

As with sizing, part tolerance after re-pressing depends on material type and part size

Re-pressing generally refers to the application of high pressures on a sintered part at room temperature, while powder forging refers to processes in which a P/M preformed part is kept at an elevated temperature during the application of high pressure (see the section "Powder Forging" ) At room temperature and at pressures as high as or higher than the compacting pressure, re-pressing increases the strength of a sintered P/M part by decreasing its porosity and by cold working the metal The part is considerably strengthened, but at the expense of ductility Resintering after re-pressing

increases the ductility and toughness of the part without diminishing its strength Those materials that are difficult to press after sintering usually can be re-pressed if the sintering is done at a low temperature at which alloying cannot take place; this low-temperature sintering is called presintering For iron alloys, presintering is done at 845 °C (1550 °F)

The effect of pressing on the density of ferrous P/M materials is shown in Fig 14 The density that is achieved by pressing depends on the density of the sintered or presintered compact, the re-pressing pressure and lubricant, and whether the powder used was prealloyed or mixed from elemental powders

re-Fig 14 Effect of re-pressing on density of powder metallurgy compacts Alloy steel powders (4640

composition) were compacted at various pressures, then sintered, re-pressed, and resintered For each specimen, the final density is indicated by the intersection between the curve that indicates the re-pressing pressure and the grid line that indicates the green compacting pressure (a) Prealloyed powder (b) Diffusional alloy made from elemental powders

Ferrous Powder Metallurgy Materials

Revised by Leander F Pease III, Powder-Tech Associates, Inc

Powder forging is a process in which unsintered, presintered, or sintered powder metal preforms are hot formed in confined dies The process is sometimes called P/M forging or P/M hot forming, or is simply referred to by the acronym P/F

Powder forging is a natural extension of the conventional press and sinter (P/M) process, which has long been recognized

as an effective technology for producing a great variety of parts to net or near-net shape Figure 15 shows the powder forging process In essence, a porous preform is densified by hot forging with a single blow Forging is carried out in heated, totally enclosed dies, and virtually no flash is generated

Fig 15 The powder forging process

The shape, quantity, and distribution of porosity in P/M and P/F parts strongly influence their mechanical performance Powder forging is a deformation processing technology aimed at increasing the density of P/M parts and thus their performance characteristics

There are two basic forms of powder forging:

• Hot upsetting, in which the perform experiences a significant amount of lateral material flow

• Hot re-pressing, in which material flow during densification is mainly in the direction of pressing The form of densification is sometimes referred to as hot restriking, or hot coining

While P/F parts are primarily used in automotive applications where they compete with cast and wrought products, parts have also been developed for military and off-road equipment

The economics of powder forging have been reviewed by a number of authors (Ref 10, 11, 12, 13, 14, 15) Some of the case histories included in the section "Applications of Powder Forged Parts" in this article compare the cost of powder forging with that of alternative forming technologies

Material Considerations

The initial production steps of powder forging (performing and sintering) are identical to those of the conventional press and sinter P/M process Certain defined physical characteristics and properties are required in the powders used in these processes In P/M parts, surface finish is related to the particle size distribution of the powder In powder forging, however, the surface finish is directly related to the finish of the forging tools

Typical pressing grades are -80 mesh with a median particle size of about 75 μm The apparent density and flow are important for maintaining fast and accurate die filling The chemistry affects the final alloy produced, as well as the compressibility

Green strength and compressibility are more critical in P/M than they are in P/F applications Although there is a need to maintain edge integrity in P/F preforms, there are rarely thin, delicate sections that require high green strength Because P/F preforms do not require high densities (typically 6.2 to 6.8 g/cm3), the compressibility obtainable with prealloyed powders is sufficient However, carbon is not prealloyed because it has an extremely detrimental effect on compressibility

The two principal requirements for P/F materials are a capability to develop an appropriate hardenability that will guarantee strength and to control fatigue performance by microstructural features such as inclusions

Hardenability. Nickel and molybdenum have the advantage that their oxides are reduced at conventional sintering temperatures Alloy design is therefore a compromise, and the majority of atomized prealloyed powders in commercial use are nickel/molybdenum based, with manganese present in limited quantities The compositions of three commercial P/M steels are:

(a) All compositions contain balance of iron

The higher cost of nickel and molybdenum, along with the higher cost of powder, compared with conventional wrought materials, is often offset by the higher material utilization inherent in the P/F process

More recently, P/F parts have been produced from iron powders (0.10 to 0.25% Mn) with copper and/or graphite additions for parts that do not require the heat-treating response or high-strength properties achieved through the use of the low-alloy steels

Inclusion Assessment. Because the properties of material powder forged to near-full density are strongly influenced

by the composition, size distribution, and location of nonmetallic inclusions (Ref 16, 17, 18), a method has been developed for assessing the inclusion content of powders intended for P/F applications (Ref 19 , 20, 21 , 22) Samples of powders intended for forging applications are re-press powder forged under closely controlled laboratory conditions The resulting compacts are sectioned and prepared for metallographic examination The inclusion assessment technique involves the use of automatic image analysis equipment The compact used for inclusion assessment may also be used to measure the amount of unalloyed iron powder particles present

Process Considerations

The development of a viable powder forging system requires the consideration of many process parameters The mechanical, metallurgical, and economic outcomes depend to a large extent on operating conditions, such as temperature, pressure, flow/feed rates, atmospheres, and lubrication systems Equally important consideration must be given to the types of processing equipment, such as presses, furnaces, dies, and robotics, and to secondary operations, in order to obtain the process conditions that are most efficient This efficiency is maintained by optimizing the process line layout

Examples of effective equipment layouts for performing, sintering, reheating, forging, and controlled cooling have been reviewed in the literature (Ref 10) Figure 16 shows a few of the many possible operational layouts Each of these process stages is reviewed in the following sections

Fig 16 A powder forging process line Source: Ref 23

Preforming. Preforms are manufactured from admixtures of metal powders, lubricants, and graphite Compaction is predominantly accomplished in conventional P/M presses that use closed dies

The control of weight distribution within preforms is essential to produce full density and thus maximize performance in the critical regions of the forged component Excessive weight in any region of the preform may cause overload stresses that could lead to tool breakage at forging

Successful preform designs have been developed by an iterative trial-and-error procedure, using prior experience to determine the initial shape More recently, computer-aided design (CAD) has been used for preform design

Preform design is intimately related to the design and dimensions of the forging tooling, the type of forging press, and the forging process parameters Among the variables to be considered for the preforming tools are:

• Temperature, that is, preform temperature, die temperature, and, when applicable, core rod temperature

• Ejection temperature of the forged part

• Lubrication conditions, that is, influence on compaction/ejection forces and tooling temperatures

• Transfer time and handling of the preform from the preheat furnace to the forging die cavity

Correct preform design not only entails having the right amount of material in the various regions of the preform, but also

is concerned with material flow between the regions and prevention of potential fractures and defects (Fig 17)

Fig 17 Configuration for the ring preform (a) for forging the part shown in (b) (b) Cross section of the part

under consideration for powder forging

Sintering and Reheating. Preforms may be forged directly from the sintering furnace; sintered, reheated, and forged;

or sintered after the forging process The basic requirements for sintering in a ferrous P/F system are:

• Lubricant removal

• Oxide reduction

• Carbon diffusion

• Development of particle contacts

• Heat for hot densification

Oxide reduction and carbon diffusion are the most important aspects of the sintering operations For most ferrous powder forging alloys, sintering takes place at about 1120 °C (2050 °F) in a protective reducing atmosphere with a carbon potential to prevent decarburization Typical P/M sintering has been performed at 1120 °C (2050 °F) for 20 to 30 min Increases in temperature will reduce the time required for sintering by improving oxide reduction and increasing carbon

diffusion Chromium-manganese steels have been limited in their use because of the higher temperatures required to reduce their oxides and the greater care needed to prevent reoxidation

Any of the furnaces used for sintering P/M parts, such as vacuum, pusher, belt, rotary hearth, walking beam, roller hearth, and batch/box, may be used for sintering or reheating P/F preforms The sintered preforms may be forged directly from the sintering furnace; stabilized at lower temperatures and forged; or cooled to room temperature, reheated, and forged All cooling, temperature stabilization, and reheating must occur under protective atmosphere to prevent oxidation

Induction furnaces are often used to reheat axisymmetric preforms to the forging temperature because of the short time required to heat the material Difficulties may be encountered in obtaining uniform heating throughout asymmetric shapes because of the variation in section thickness

Powder forging involves removing heated preforms from a furnace, usually by robotic manipulators, and locating them

in the die cavity for forging at high pressures (690 to 965 MPa, or 100 to 140 ksi) Preforms may be graphite coated to prevent oxidation during reheating and transfer to the forging die Lubrication of the die and punches is usually accomplished by spraying a water-graphite suspension into the cavity

The forging presses commonly used in conventional forging, including hammers, high energy rate forming machines, mechanical presses, hydraulic presses, and screw presses, have been evaluated for use in powder forging The essential characteristics that differentiate presses are contact time, stroke velocity, available energy and load, stiffness, and guide accuracy

Metal Flow in Powder Forging. Draft angles, which facilitate forging and ejection in conventional forging, are eliminated in P/F parts This means that greater ejection forces on the order of 15 to 20% of press capacity as a minimum are required for the powder forging of simple shapes However, the elimination of draft angles permits P/F parts to be forged closer to net shape

Tool Design. In order to produce sound forged components, the forging tooling must be designed to take into account:

• Preform temperature

• Die temperature

• Forging pressure

• The elastic strain of the die

• The elastic/plastic strain of the forging

• The temperature of the part upon ejection

• The elastic strain of the forging upon ejection

• The contraction of the forging during cooling

• Tool wear

Secondary Operations. In general, the secondary operations applied to conventional components, such as plating and peening, may be applied to P/F components The most commonly used secondary operations involve deburring, heat treating, and machining

The heat treatment of P/M products is the same as that required for conventionally processed materials of similar composition The most common heat-treating practices involve treatments such as carburizing, quench and temper cycles, and continuous-cooling transformation

The amount of machining required for P/F components is less than the amount required for conventional forgings because

of the improved dimensional tolerances Standard machining operations may be used to achieve final dimensions and surface finish One of the main economic benefits of powder forging is the reduced amount of machining required Improved machinability can be accomplished by the addition of solid lubricants such as manganese sulfide

Mechanical Properties

Wrought steel bar stock undergoes extensive deformation during cogging and rolling of the original ingot This creates inclusion stringers and leads to planes of weakness, which affect the ductile failure of the material The mechanical

properties of wrought steels vary considerably with the direction test pieces are cut from the wrought billet Powder forged materials, on the other hand, undergo relatively little material deformation, and their mechanical properties have been shown to be relatively isotropic

The mechanical properties of P/F materials are usually intermediate to the transverse and longitudinal properties of wrought steels The rotating-bending fatigue properties of P/F material have also been shown to fall between the longitudinal and transverse properties of wrought steel of the same tensile strength

While the performance of machined laboratory test pieces follows the intermediate trend described above, in the case of actual components, P/F parts have been shown to have superior fatigue resistance This has generally been attributed not only to the relative mechanical property isotropy of powder forgings, but also to their better surface finish and finer grain size

This section reviews the mechanical properties of P/F materials The data presented represent results obtained on machined standard laboratory test pieces Data are reported for four primary materials The first two material systems are based on prealloyed powders (P/F-4600 and P/F-4200; see the section "Hardenability" in this article) The third material, based on an iron-copper-carbon alloy, was used by Toyota in 1981 to make P/F connecting rods; Ford Motor Company introduced powder forged rods with a similar chemistry in 1986 Mechanical property data are therefore presented for copper and graphite powders mixed with an iron powder base to produce materials that generally contain 2% Cu Some powder forged components are made from plain carbon steel This is the fourth and final material for which mechanical property data are presented

Forging Mode. It is well known that the forging mode has a major effect on the mechanical properties of components With this in mind, the mechanical property data reported in this section were obtained on specimens that were either hot upset or hot re-press forged

Heat Treatments. There were three heat treatments used in developing the properties of the prealloyed powder forged materials: case carburizing, blank carburizing, and through-hardening (quenching and tempering)

Hardenability. Jominy hardenability curves are presented in Fig 18 for the P/F-4600, P/F-4200, and carbon alloys Testing was carried out according to ASTM A 255 Specimens were machined from upset forged billets that had been sintered at 1120 °C (2050 °F) in associated ammonia

iron-copper-Fig 18 Jominy hardenability curves for (a) P/F-4600, (b) P/F-4200, and (c) iron-copper-carbon materials at

various forged-carbon levels Vickers hardness was determined at a 30 kgf load

Tensile, Impact, and Fatigue Properties. Tensile properties were determined on test pieces with a gage length of

25 mm (1 in.) and a gage diameter of 6.35 mm (0.25 in.) Testing was carried out according to ASTM E 8 using a crosshead speed of 0.5 mm/min (0.02 in./min) Room-temperature impact testing was carried out on standard Charpy V-notch specimens according to ASTM E 23 Rotating-bending fatigue (RBF) testing was performed using single-load, cantilever, rotating fatigue testers The tensile, impact, and fatigue data for the various materials are summarized in Tables

9, 10, and 11

Table 9 Mechanical property and fatigue data for P/F-4600 materials

Sintered at 1120 °C (2050 °F) in dissociated ammonia unless otherwise noted

Ultimate tensile strength

0.2% offset yield strength

Room-temperature Charpy V-notch impact energy

Fatigue endurance limit

Forging mode Carbon,

%

Oxygen, ppm

MPa ksi MPa ksi

Elongation

in 25 mm (1 in.), %

Reduction

of area, %

J ft · lbf

Core hardness, HV30

MPa ksi

Ratio of fatigue endurance to

Upset 0.39 260 825 120 745 108 21.0 57.0 62.4 46.0 269

Upset(g) 0.58 280 860 125 760 110 20.0 50.0 44.0 32.5 270

Upset(h) 0.80 360 850 123 600 87 19.5 46.0 24.4 18.0 253

(a) Sintered at 1260 °C (2300 °F) in dissociated ammonia

(b) Sintered at 1120 °C (2050 °F) in endothermic gas atmosphere

Table 10 Mechanical property data for P/F-4200 materials

Ultimate tensile strength

0.2% offset yield strength

Forging mode Carbon,

%

Oxygen, ppm

MPa ksi MPa ksi

Elongation in 25

mm (1 in.), %

Reduction of area, %

Quenched and tempered

(b) Sintered in dissociated ammonia at 1120 °C (2050 °F)

(c) Sintered in dissociated ammonia at 1260 °C (2300 °F)

(j) Tempered at 660 °C (1225 °F)

(k) Tempered at 675 °C (1250 °F)

Table 11 Mechanical property and fatigue data for iron-copper-carbon alloys

Sintered at 1120 °C (2050 °F) in dissociated ammonia, reheated to 980 °C (1800 °F) in dissociated ammonia, and forged

Ultimate tensile strength

0.2% offset yield strength

Room-temperature Charpy V-notch impact energy

Fatigue endurance limit

Forging mode Carbon,

%

Oxygen, ppm

MPa ksi MPa ksi

Elongation

in 25 mm (1 in.), %

Reduction

of area, %

J ft · lbf

Core hardness, HV30

MPa ksi

Ratio of fatigue endurance to

Re-press 0.82 220 1170 170 745 108 10 12.8 2.7 2.0 368 475 69 0.41

(a) Still-air cooled

(b) Forced-air cooled

The iron-copper-carbon alloys were either still-air cooled or forced-air cooled from the austenitizing temperature of 845

°C (1550 °F) The austenitizing temperature influences core hardness These iron-copper-carbon alloys are often used with manganese sulfide additions for enhanced machinability The tensile, impact, and fatigue properties for a sample with a 0.35% manganese sulfide addition are compared with a material without sulfide additions in Table 12 Data from the samples with manganese sulfide and sulfurized powders are included for comparison The manganese sulfide addition had little influence on tensile strength, whereas the sulfurization process degraded tensile properties

Table 12 Mechanical property and fatigue data for iron-copper-carbon alloys with sulfur additions

Sintered at 1120 °F (2050 °F) in dissociated ammonia, reheated to 980 °C (1800 °F) in dissociated ammonia, and forged

Ultimate tensile strength

0.2% offset yield strength

Room-temperature Charpy V-notch impact energy

Fatigue endurance limit

%

Oxygen, ppm

Reduction

of area, %

J ft · lbf

Core hardness, HV30

MPa ksi

Ratio of fatigue endurance to

tensile strength

Manganese sulfide 0.59 270 0.13 915 133 620 90 11 23.2 6.8 5.0 290 430 62 0.47

Sulfur 0.63 160 0.14 840 122 560 81 12 21.4 6.8 5.0 267 415 60 0.50

Compressive Yield Strength. The 0.2% offset compressive yield strengths for P/F-4600 at various forged carbon levels after different heat treatments are summarized in Table 13

Table 13 Compressive yield strengths of P/F-4600 materials

Sintered at 1120 °C (2050 °F) in dissociated ammonia

0.2% offset compressive yield

strength

Forged carbon

content, %

Forged oxygen content, ppm

0.77 410 Tempered at 695 °C (1280 °F) 700 101

Effect of Porosity on Mechanical Properties. The mechanical property data summarized in the previous sections are related to either hot re-press or hot upset forged pore-free material The general effect of density on mechanical properties is presented in Table 14

Table 14 Tensile and impact properties of P/F-4600 hot re-pressed at two temperatures

Re-pressing

temperature

pressing stress

Re-Re-pressed density

0.2% offset yield strength

Ultimate tensile strength

Charpy notch impact

Quality Assurance for P/F Parts

Many of the quality assurance tests applied to wrought parts are similar to those used for powder forged parts Among the parameters specified are part dimensions, surface finish, magnetic particle inspection, composition, density, metallographic analysis, and nondestructive testing

Part Dimensions and Surface Finish. Typical tolerances for P/F parts are summarized in Table 15 The as-forged surface finish of a P/F parts is directly related to the surface finish of the forging tool Surface finish is generally better than 0.8 m (32 in.) which is better than that obtained on wrought forged parts This good surface finish is beneficial to the fatigue performance of P/F parts

Table 15 Typical tolerances for P/F parts

Magnetic particle inspection is used to detect surface blemishes such as cracks and laps

Density. Sectional density measurements are taken to ensure that sufficient densification has been achieved in critical areas Displacement density checks are generally supplemented by microstructural examination to assess the residual porosity level For a given level of porosity, the measured density will depend on the exact chemistry, thermomechanical condition, and microstructure of the sample Parts may be specified to have a higher density in particular regions than is necessary in less critical sections of the same component

Metallographic Analysis. Powder forged parts are subjected to extensive metallographic evaluation The primary parameters of interest include those discussed below

The extent of surface decarburization permitted in a forged part will generally be specified The depth of decarburization may be estimated by metallographic examination, but it is best quantified using microhardness measurements as described in ASTM E 1077

Surface finger oxides are defined as oxides that follow prior particle boundaries into the forged part from the surface

and cannot be removed by physical means such as rotary tumbling An example of surface finger oxides is shown in Fig

19 Metallographic techniques are used to determine the maximum depth of surface finger oxide penetration

Fig 19 Surface finger oxides (arrows at upper right) and interparticle oxide networks (arrow near lower left) in

a powder forged material

Interparticle oxides follow prior particle boundaries They may sometimes form a continuous three-dimensional network, but more often will, in a two-dimensional plane of polish, appear to be discontinuous An example is presented

in Fig 19

Unalloyed iron powder contamination in low-alloy powder forged parts can be quantified by means of the etching procedure described in the earlier section "Material Considerations."

The nonmetallic inclusion level in a P/F part may also be quantified using the image analysis technique described in the section "Material Considerations." However, if the section of a component selected for inclusion assessment is not pore-free, image analysis procedure are not applicable

Nondestructive Testing. Although metallographic assessment of P/F parts is common, it is also useful to have a nondestructive method for evaluating the microstructural integrity of components It has been demonstrated that this can

be achieved with a magnetic bridge comparator

Applications of Powder Forged Parts

Previous sections compared powder forging and drop forging and illustrated the range of mechanical property performance that can be achieved in powder forged material The present section concentrates on two examples of P/F components and highlights some of the reasons for selecting P/F parts over those made by competing forming methods

Example 1: Converter Clutch Cam

The automotive industry is the principal user of P/F parts, and components for automatic transmissions represent the major areas of application One of the earliest powder forgings used in such an application is the converter clutch cam (Fig 20) The primary reason that powder forging was chosen over competitive processes was that it reduced manufacturing costs by 58%, compared with the conventional process of machining a forged gear blank

Fig 20 Powder forged converter clutch cam used in an automotive automatic transmission Courtesy of

Precision Forged Products Division, Federal Mogul Corporation

Powder forged cams are made from a water-atomized steel powder (P/F-4200) containing 0.6% Mo, 0.5% Ni, 0.3% Mn, and 0.3% graphite Preforms weighing 0.33 kg (0.73 lb) are compacted to a density of 6.8 g/cm3 The preforms are sintered at 1120 °C (2050 °F) in an endothermic gas atmosphere with a 2 °C (35 °F) dewpoint The sintered preforms are graphite coated before being induction heated and forged to near-full density (less than 0.2% porosity) using both axial and lateral flow After forging, the face of the converter clutch cam is ground, carburized to a depth of 1.75 mm (0.070 in.), and surface hardened by means of induction The part requires a high density to withstand the high Hertzian stress experienced by the inner cam surface in service Machining requires only one step on the P/F cam; seven machining operations were required for the conventionally processed cam

Example 2: Powder Forged Connecting Rods

Connecting rods were among the components selected for a number of P/F development programs in the 1960s However,

it was not until 1976 that the first P/F connecting rod was produced commercially This was the connecting rod for the Porsche 928 V-8 engine

The connecting rod for the Porsche 928 engine was made from a water-atomized low-alloy steel powder (0.3 to 0.4% Mn, 0.1 to 0.25% Cr, 0.2 to 0.3% Ni, and 0.25 to 0.35% Mo) to which graphite was added to give a forged carbon content of

0.35 to 0.45% The forgings were oil quenched and tempered to a core hardness of 28 HRC (ultimate tensile strength of

835 to 960 MPa, or 121 to 139 ksi), followed by shot peening to a surface finish of 11 to 13 on the Almen scale

The preform was designed so that the P/F component had less than 0.2% porosity in the critical web region The P/F connecting rod had considerably better fatigue properties than did conventional drop forged rods Its weight control was good enough to allow a reduction in the size of the balance pads, resulting in a weight savings of about 10% (it weighed

~1 kg, or 2 lb) Powder forged connecting rods are currently used in both the Porsche 928 and 944 engines

The first high-volume commercialization of P/F connecting rods was in the 1.9 L Toyota Camry engine In this design, the balance pads were completely eliminated Despite the publication of the results of development trials in 1972, it was not until the summer of 1981 that production rods were introduced

Toyota selected a copper steel (Fe-0.55C-2Cu) based on a water-atomized iron powder to replace conventional forgings, which had been made from a quenched and tempered 10L55 free-machining steel The preform, which has a preshaped partial I-beam web section, has an average green density of 6.5 g/cm3 The preform shape is such that forging is predominantly in the re-pressing mode However, some lateral flow does take place where required in critical regions, such as the web

Preforms are sintered for 20 min at 1150 °C (2100 °F) in an endothermic gas atmosphere in a specially designed rotary hearth furnace During sintering, the preforms are supported on flat, ceramic plates The preforms are allowed to stabilize

at about 1010 °C (1850 °F) before closed-die forging

Exposure of the preform to the atmosphere during transfer to the forging dies is limited to 4 to 5 s An ion nitriding treatment is applied to the punches and dies in the regions at which forging deformation occurs The connecting rods are forged at the rate of 10 per minute, and tool lives of over 100,000 pieces have been reported

The forged rods are subjected to a thermal treatment after forging This results in a ferrite/pearlite microstructure with a core hardness of 240 to 300 HV (30 kgf load) Subsequent operations include burr removal, shot peening, straightening, sizing, magnetic particle inspection, and finish machining

Savings in material and energy are substantial for the P/F rods The billet weight for a conventional forging is 1.2 kg (2.65 lb); the P/F preform weighs 0.7 kg (1.54 lb)and requires little machining In addition to the benefits in process economics, the variability in fatigue performance for the P/F rods is reported to be half that of conventionally forged parts

Ford Motor Company has introduced P/F connecting rods in the 1.9 L four-cylinder engine used in the Ford Escort model Five million connecting rods were in service by June 1989 Figure 21 illustrates the Ford P/F connecting rod Ford has also announced plans to use P/F rods in its modular engine, which is scheduled for production in 1992

Fig 21 Powder forged connecting rod for 1.9 L automobile engine A similar rod will be used in the modular

engine

References cited in this section

10 G Bockstiegel, Powder Forging Development of the Technology and Its Acceptance in North America,

Japan, and West Europe, in Powder Metallurgy 1986 State of the Art, Vol 2, Powder Metallurgy in Science

and Practical Technology series, Verlag Schmid, 1986, p 239

11 P.K Jones, The Technical and Economic Advantages of Powder Forged Products, Powder Metall., Vol 13

(No 26), 1970, p 114

12 Economic Aspects of P/M-Hot-Forming, Mod Dev Powder Metall., Vol 7, 1974, p 91

13 J.W Wisker and P.K Jones, The Economics of Powder Forging Relative to Competing Processes Present

and Future, Mod Dev Powder Metall., Vol 7, 1974, p 33

14 W.J Huppmann and M Hirschvogel, Powder Forging, Review 233, Int Met Rev., No 5, 1978, p 209

15 C Tsumuti and I Nagare, Application of Powder Forging to Automotive Parts, Met Powder Rep., Vol 39

(No 11), 1984, p 629

16 R Koos and G Bockstiegel, The Influence of Heat Treatment, Inclusions and Porosity on the Machinability

of Powder Forged Steel, Prog Powder Metall., Vol 37, 1981, p 145

17 B.L Ferguson, H.A Kuhn, and A Lawley, Fatigue of Iron Base P/M Forgings, Mod Dev Powder Metall.,

Vol 9, 1977, p 51

18 G.T Brown and J.A Steed, The Fatigue Performance of Some Connecting Rods Made by Powder Forging,

Powder Metall., Vol 16 (No 32), 1973, p 405

19 W.B James, The Use of Image Analysis for Assessing the Inclusion Content of Low Alloy Steel Powders

for Forging Applications, in Practical Applications of Quantitative Metallography, STP 839, American

Society for Testing and Materials, 1984, p 132

20 R Causton, T.F Murphy, C.-A Blande, and H Soderhjelm, Non-Metallic Inclusion Measurement of

Powder Forged Steels Using an Automatic Image Analysis System, in Horizons of Powder Metallurgy, Part

II, Verlag Schmid, 1986, p 727

21 W.B James, "Quality Assurance Procedures for Powder Forged Materials," Technical Paper 830364, Society of Automotive Engineers, 1983

22 W.B James, Automated Counting of Inclusions in Powder Forged Steels, Mod Dev Powder Metall., Vol

14, 1981, p 541

23 W.B James, New Shaping Methods for P/M Components, in Powder Metallurgy 1986 State of the Art,

Vol 2, Powder Metallurgy in Science and Practical Technology series, Verlag Schmid GmbH, 1986, p 71

Note cited in this section

* The text, photographs, and tables in this section on powder forging were adapted from the article "Powder

Forging" in Forming and Forging, Volume 14 of ASM Handbook, formerly 9th Edition Metals Handbook

Ferrous Powder Metallurgy Materials

Revised by Leander F Pease III, Powder-Tech Associates, Inc

Metal Injection Molding (MIM) Technology

When 60 vol% of fine (10 m, or 400 in.) metal powder is blended with 40 vol% lubricant and binder, the resulting mixture can be injection molded, much like a conventional plastic Any shape that can be molded in plastic can be molded

in metal powder The use of paraffin waxes and thermoplastics such as polyethylene or polypropylene provides the rheological basis for allowing the mixture to flow around corners and into undercuts in a way that is impossible in the uniaxial pressing of binder-free metal powders The process leads to a whole new class of P/M geometries

In the original 1971 Wiech process, the wax is removed with a liquid solvent The remaining plastic binder is thermally removed during a very careful heat-up and sintering The latter process takes 3 to 5 days During the long debindering and sintering cycle at 1200 to 1300 °C (2200 to 2400 °F), near-theoretical density is reached Following further development

at Witec, the process (Wiech II) now consists of removal of the wax in air, at about 175 °C (350 °F), over a 1 to 3 day

period This oxidizes the part, resulting in a brownish color, while the oxide imparts extra strength for handling Sintering takes place in a closed reaction vessel in a controlled atmosphere of argon and hydrogen The 60% dense as-molded parts shrink 14 to 20% and achieve near-full density (>95% dense) The parts have very fine, noninterconnected porosity and much better elongation, toughness, and dynamic properties than conventionally pressed and sintered materials Such parts are limited to section sizes of 9.5 to 13 mm (0.375 to 0.5 in.) and have tolerances of ±0.003 mm/mm (±0.003 in./in.) Because of the oxidation that occurs during wax removal, the process may be fundamentally limited to smaller cross sections Also, the oxidation of sensitive elements such as chromium and manganese is only reversed at later sintering, with difficulty Modification of the early Wiech process using solvent to debinder the wax, is now the most commonly used debindering technique The residual thermoplastic can be handled in a vacuum furnace or atmosphere pusher furnace

The Rivers process, licensed by Haynes International, uses methylcellulose as a binder, along with small amounts of water, glycerine, and boric acid The methylcellulose dissolves in cold water to form a binder Upon injection into a warm mold, it gels to form a fairly rigid part that can be removed from the tooling The part is dried at 120 °C (250 °F) to remove water and then sintered by any convenient sintering process This can be done in belt furnaces, pushers, or vacuum furnaces, at nearly any heat-up rate Depending on the fineness of the starting powder and the sintering temperature, densities of 90 to 99% of full density are obtained An injection molded and sintered 6.8 kg (15 lb) Stellite block has demonstrated the capability to make parts with a 75 mm (3 in.) thick section This degree of binder removal is possible because the methyl cellulose breaks down easily upon heat-up and follows the channels created by the evaporated water

Although there are some large-scale users of the Wiech-type batch sintering reactors in Japan (340 kg, or 750 lb, in a 60 h period), the trend in the United States is toward more continuous furnaces The use of a 1315 °C (2400 °F) molybdenum wound pusher furnace has become widespread A single-stage vacuum furnace to evaporate and thermally remove all binders and lubricants from MIM parts has been developed It is an outgrowth of vacuum furnaces, which have long been used to dewax and liquid phase sinter carbide parts The MIM parts are then heated to sintering temperatures of 1100 to

1315 °C (200 to 2400 °F) Because the parts are not oxidized, the process works well with sensitive chromium- and manganese-containing materials, as well as low-alloy steel A 45 kg (100 lb) load has a floor-to-floor time of 12 h or less and runs under microprocessor control

Advantages of the MIM Process. The tolerance and shape factor advantages were noted above The design engineer should think of using injection molded parts in mechanisms or as small enclosures Some of the MIM fabricators produce runs of 2000 to 5000 pieces, particularly on more expensive parts They can do a short run such as these because there is

no danger to the tooling at setup, while in conventional P/M there is more risk As the process becomes ever better understood and better controlled, small batches can be grouped together for sintering Sintering to near-full density seems

to reach an end point for shrinkage This contributes to product uniformity, as long as the original molded part has uniform metal powder density

Because the process uses fine powders (5 to 20 μm), alloys have been formed in situ by the diffusion of mixed elements

or master alloys Some stainless steel alloys have been made in this way, using an Fe-30Cr master alloy It is convenient

to make 4100 series alloys without the need for prealloyed powders The capability to blend an alloy allows a flexibility that does not exist in conventional powder metallurgy There are only a limited number of prealloyed powders available

Sintering to near-full density gives excellent toughness, elongation, and other dynamic properties This is aided by the presence of fine spheroidized porosity versus the sharp, stress raiser porosity of conventional powder metallurgy (see Fig

22 for a MIM microstructure) Without open pores, the parts can be plated, used in pressure-controlled environments, or used in food handling applications The stainless steels thus produced have very low carbon contents and are more corrosion resistant than typical P/M materials

Fig 22 Typical MIM microstructure with rounded, isolated areas of porosity BASF grade OM carbonyl iron

sintered for 1 h at 1315 °C (2400 °F) in vacuum 94% of full density Unetched 180×

The capital equipment (presses) for injection molding is more economical than that for large-scale P/M presses Tool life

is at least 300,000 pieces These factors help offset the added short-term material cost

Factors Impeding Growth of MIM Technology. In conventional press and sinter P/M, dimensional tolerances and quality control are very significant Re-pressed or coined P/M parts are made to a tolerance of ±0.013 mm (±0.0005 in.)

on a 25 mm (1 in.) diam part As-sintered tolerances on well-behaved alloys are ±0.001 mm/mm (±0.001 in./in.) Most MIM producers are achieving ±0.003 mm/mm (±0.003 in./in.) with some companies offering 0.001 mm/mm (0.001 in./in.) on selected dimensions The industry as a whole needs to reach the tolerance levels being offered by conventional powder metallurgy Some MIM producers do re-press their parts to straighten them and achieve tolerances like those of conventional powder metallurgy

The raw materials are expensive: in 1989, carbonyl Fe, reduced, is $8.60/kg ($3.90/lb) With binder added, ferrous material costs may be $7 to 10/kg ($3.20 to 4.50/lb) The total cost for low-alloy steel MIM parts is $30 to 40/kg ($14 to 18/lb) This explains why parts of less than 20 g (0.7 oz) have been the initial focus of sales activity Work is underway

on ways to liquid phase sinter mixtures of $0.70/kg ($0.32/lb) iron and a special graphite (Ref 24) At 1150 °C (2100 °F), some regions of the iron-graphite parts do sinter to full density, but the process is not yet successful The development of sintering techniques for the coarser, less expensive powders will enormously impact the economics of the process There are several Japanese and U.S companies at work developing a way to make less expensive fine Fe powders or alloys

MIM green parts can exhibit defects as they exit the press Variations in metal loading or the segregation of metal and plastic will result in dimensional variations as full density is approached Real-time x-ray equipment has disclosed voids and low density areas in injection molded ceramic parts with diameters of 0.1 mm (0.004 in.) It also may be possible to measure density variation and to look for cracks and pores in MIM parts with x-ray equipment

Mechanical Properties. Typical mechanical properties for 95% dense MIM materials are given in Table 16 The Metal Injection Molding Association and ASTM Subcommittee B09.11 are currently developing the standardized minimum properties for the MIM materials that are in current use These will be available in 1990 Figure 23 shows the new standard tensile test bar for metal injection molded materials It is pulled by inserting pins through the holes in each end and is self-aligning

Table 16 Mechanical properties of metal injection molded materials

tensile strength

0.2% offset yield strength

MPa ksi MPa ksi

%

Sintered 470 68 255 37 21 74 HRB Sinter

Heat treated

1450 210 1095 159 4.0 38 HRC Sinter, heat treat, then

hold 1 h at 232 °C (450 °F)

Heat treated

1125 163 980 142 5.5 31 HRC Sinter, heat treat, then

hold 1 h at 343 °C (650 °F)

Fe-2Ni-0.43C

Heat treated

880 128 710 103 11.0 21 HRC Sinter, heat treat, then

hold 1 h at 427 °C (800 °F)

Sintered 605 88 345 50 11.5 79 HRB Sinter

Heat treated

1620 235 1350 196 2.5 44 HRC Sinter, heat treat, then

hold 1 h at 232 °C (450 °F)

Heat treated

1345 195 1115 162 4.5 78 HRC Sinter, heat treat, then

hold 1 h at 343 °C (650 °F)

4600 + 0.45% C

Heat treated

1090 158 938 136 6.0 31 HRC Sinter, heat treat, then

Source: Brunswick Technetics Division of Brunswick Corporation

Fig 23 Tensile bars specially developed for testing MIM Courtesy of Omark Industries, Advanced Forming

Technology, and Metal Injection Molding Association

Applications. MIM parts have been used in a variety of production parts for automobiles Other applications of MIM materials include the molding of threads on pressure manifolds and the fabrication of gun sight parts with a special nongalling cam locking mechanism that could not be machined on the parts

The first two MIM production parts in automobiles were a part for an ignition lock (Fig 24) and a single-part replacement (Fig 25a) for a two-part turn signal lever assembly (Fig 25b) Both have been in service since July 1988

Fig 24 MIM part (upper left) for an automobile ignition lock The key forces the MIM part into contact with a

security switch Courtesy of SSI Technologies

Fig 25 Single-piece MIM part (a) that replaced a two-piece automobile turn signal lever assembly (b) The

smaller MIM part in (a) was the first version, while the larger MIM part is the finished version that replaced the two-part assembly shown in (b) Courtesy of the Remington Arms Division of E.I Du Pont de Nemours & Company, Inc

Figure 24 shows the entire ignition lock and the MIM subcomponent As the key is inserted in the lock, the cam-shaped MIM part moves away and depresses an electrical switch, which is part of the security system The initial design of the part was too small and complicated for the model shop to make and it was prototyped from the MIM tooling

The turn signal indicator lever is an example of the replacement of a two-piece assembly (Fig 25b) with a single MIM part The lower portion of Fig 25(a) shows the first version of the MIM part, and the upper view shows the final 19.0 g (0.670 oz) MIM part that replaced the assembly The MIM material is iron with 2% Ni, sintered and then case hardened

It replaced AISI 4037 and SAE 1018 case hardened The MIM part succeeded because of its superior strength compared

to the two-piece assembly The core properties of the materials are 415 MPa (60 ksi) tensile strength and 15% elongation

at 60 HRB

Reference cited in this section

24 L.F Pease, An Approach to Near Full Density P/M Automotive Components Liquid Phase Sintering of Iron Graphite Materials, Technical Paper 870131, Society of Automotive Engineers, 1987

Ferrous Powder Metallurgy Materials

Revised by Leander F Pease III, Powder-Tech Associates, Inc

References

1 R.M German, Powder Metallurgy Science, Metal Powder Industries Federation, 1984

2 F.V Lenel, Powder Metallurgy Principles and Applications, Metal Powder Industries Federation, 1980

3 P/M Materials Standards for P/M Structural Parts, MPIF Standard 35, Metal Powder Industries Federation,

1988

4 C Durdaller, The Effect of Additions of Copper, Nickel and Graphite on the Sintered Properties of

Iron-Base Sintered P/M Parts, Progress in Powder Metallurgy, Metal Powder Industries Federation, Vol 25,

1969, p 71-100

5 L.G Roy and L.F Pease III, Through Hardening of P/M Materials, in Progress in Powder Metallurgy, Vol

42, Metal Powder Industries Federation, 1986

6 A.F deRege, G l'Espérance, and L.F Pease III, Prealloyed MnS Powders for Improved Machinability in

P/M Parts, in Near Net Shaping Manufacturing, P.W Lee and B.L Ferguson, Ed., ASM

INTERNATIONAL, 1988, p 57-67

7 C Durdaller, Copper Infiltration of Iron-Base P/M Parts, Hoeganaes Corporation, 1969

8 H Ferguson, Heat Treatment of P/M Parts, Met Prog., Vol 107 (No 6), June 1975, p 81-83; Vol 108 (No

2), July 1975, p 66-69

9 L.F Pease III, J.P Collette, and D.A Pease, Mechanical Properties of Steam Blackened P/M Materials, in

Modern Developments in Powder Metallurgy, Vol 18-21, Metal Powder Industries Federation, 1988

10 G Bockstiegel, Powder Forging Development of the Technology and Its Acceptance in North America,

Japan, and West Europe, in Powder Metallurgy 1986 State of the Art, Vol 2, Powder Metallurgy in

Science and Practical Technology series, Verlag Schmid, 1986, p 239

11 P.K Jones, The Technical and Economic Advantages of Powder Forged Products, Powder Metall., Vol 13

(No 26), 1970, p 114

12 Economic Aspects of P/M-Hot-Forming, Mod Dev Powder Metall., Vol 7, 1974, p 91

13 J.W Wisker and P.K Jones, The Economics of Powder Forging Relative to Competing Processes Present

and Future, Mod Dev Powder Metall., Vol 7, 1974, p 33

14 W.J Huppmann and M Hirschvogel, Powder Forging, Review 233, Int Met Rev., No 5, 1978, p 209

15 C Tsumuti and I Nagare, Application of Powder Forging to Automotive Parts, Met Powder Rep., Vol 39

(No 11), 1984, p 629

16 R Koos and G Bockstiegel, The Influence of Heat Treatment, Inclusions and Porosity on the

Machinability of Powder Forged Steel, Prog Powder Metall., Vol 37, 1981, p 145

17 B.L Ferguson, H.A Kuhn, and A Lawley, Fatigue of Iron Base P/M Forgings, Mod Dev Powder Metall.,

Vol 9, 1977, p 51

18 G.T Brown and J.A Steed, The Fatigue Performance of Some Connecting Rods Made by Powder Forging,

Powder Metall., Vol 16 (No 32), 1973, p 405

19 W.B James, The Use of Image Analysis for Assessing the Inclusion Content of Low Alloy Steel Powders

for Forging Applications, in Practical Applications of Quantitative Metallography, STP 839, American

Society for Testing and Materials, 1984, p 132

20 R Causton, T.F Murphy, C.-A Blande, and H Soderhjelm, Non-Metallic Inclusion Measurement of

Powder Forged Steels Using an Automatic Image Analysis System, in Horizons of Powder Metallurgy, Part

II, Verlag Schmid, 1986, p 727

21 W.B James, "Quality Assurance Procedures for Powder Forged Materials," Technical Paper 830364, Society of Automotive Engineers, 1983

22 W.B James, Automated Counting of Inclusions in Powder Forged Steels, Mod Dev Powder Metall., Vol

14, 1981, p 541

23 W.B James, New Shaping Methods for P/M Components, in Powder Metallurgy 1986 State of the Art,

Vol 2, Powder Metallurgy in Science and Practical Technology series, Verlag Schmid GmbH, 1986, p 71

24 L.F Pease, An Approach to Near Full Density P/M Automotive Components Liquid Phase Sintering of Iron Graphite Materials, Technical Paper 870131, Society of Automotive Engineers, 1987

Austenitic Manganese Steels

Revised by D.K Subramanyam,* Ergenics Inc.; A.E Swansiger, ABC Rail Corporation; and H.S Avery, Consultant

Introduction

THE ORIGINAL AUSTENITIC manganese steel, containing about 1.2% C and 12% Mn, was invented by Sir Robert Hadfield in 1882 Hadfield's steel was unique in that it combined high toughness and ductility with high work-hardening capacity and, usually, good resistance to wear Consequently, it rapidly gained acceptance as a very useful engineering material

Hadfield's austenitic manganese steel is still used extensively, with minor modifications in composition and heat treatment, primarily in the fields of earthmoving, mining, quarrying, oil well drilling, steelmaking, railroading, dredging, lumbering, and in the manufacture of cement and clay products Austenitic manganese steel is used in equipment for handling and processing earthen materials (such as rock crushers, grinding mills, dredge buckets, power shovel buckets and teeth, and pumps for handling gravel and rocks) Other applications include fragmentizer hammers and grates for automobile recycling and military applications such as tank track pads Another important use is in railway trackwork at frogs, switches, and crossings, where wheel impacts at intersections are especially severe Because austenitic manganese steel resists metal-to-metal wear, it is used in sprockets, pinions, gears, wheels, conveyor chains, wear plates, and shoes

Austenitic manganese steel has certain properties that tend to restrict its use It is difficult to machine and usually has a yield strength of only 345 to 415 MPa (50 to 60 ksi) Consequently, it is not well suited for parts that require close-tolerance machining or that must resist plastic deformation when highly stressed in service However, hammering, pressing, cold rolling, or explosion shocking of the surface raises the yield strength to provide a hard surface on a tough core structure

Note

* Formerly with Abex Corporation

Austenitic Manganese Steels

Revised by D.K Subramanyam,* Ergenics Inc.; A.E Swansiger, ABC Rail Corporation; and H.S Avery, Consultant

Composition

Many variations of the original austenitic manganese steel have been proposed, often in unexploited patents, but only a few have been adopted as significant improvements These usually involve variations of carbon and manganese, with or without additional alloys such as chromium, nickel, molybdenum, vanadium, titanium, and bismuth The most common of these compositions, as listed in ASTM A 128, are given in Table 1

Table 1 Standard composition ranges for austenitic manganese steel castings

Carbon and Manganese. The ASTM A 128 compositions in Table 1 do not permit any austenite transformation when the alloys are water quenched from above the Acm (that is, the temperature that corresponds to the boundary between the cementite-austenite and the austenite fields) However, this does not preclude lower ductility in heavy sections because of slower quenching rates The effect is due to the formation of carbides along grain boundaries and other interdendritic areas and to some degree affects nearly all commercial castings except the very smallest Figure 1 shows Acm

temperatures for 13% Mn steels containing between 0.6 and 1.4% C Figure 2 shows the effects of carbon and manganese content on the Ms temperature, that is, the temperature at which martensite starts to form from austenite upon cooling, of a homogeneous austenite with all carbon and manganese in solid solution

Fig 1 Solubility of carbon in 13% Mn steels Source: Ref 1

Fig 2 Variation of Ms temperature with carbon and manganese contents Source: Ref 2

The mechanical properties of austenitic manganese steel vary with both carbon and manganese content Figure 3 indicates that carbon increases strength up to the range of ASTM A 128, grade A A pleateau is indicated at 1.05 to 1.35% C content Any departure from this curve can be attributed to grain size unless good statistical evidence is found The plateau at 827 MPa (120 ksi) is based on the 97-heat, 270-test scatter graph shown in Fig 4 The data points on Fig 4 were used to calculate the standard deviation, σ, data in Fig 3 As carbon is increased it becomes increasingly difficult to retain all of the carbon in solid solution, which may account for reductions in tensile strength and ductility Nevertheless, because abrasion resistance tends to increase with carbon, carbon content higher than the 1.20% midrange of grade A may

be preferred even when ductility is lowered Carbon content above 1.4% is seldom used because of the difficulty of obtaining an austenitic structure sufficiently free of grain boundary carbides, which are detrimental to strength and

ductility The effect can also be observed in 13% Mn steels containing less than 1.4% C because segregation may result in local variations of ±17% (±0.2%C) from the average carbon level determined by chemical analysis

Fig 3 Variation of properties with carbon content for austenitic manganese steel containing 12.2 to 13.8% Mn

Data are for castings weighing 3.6 to 4.5 kg (8 to 10 lb) and about 25 mm (1 in.) in section size that were water quenched from 1040 to 1095 °C (1900 to 2000 °F) Flow under impact is the total reduction in height sustained by a cylindrical specimen 25 mm (1 in.) in both diameter and length after absorbing 20 blows of 680

J (500 ft · lbf) each Source: Abex Research Center

Fig 4 Distribution of tensile strength and ductility values for 97 heats of manganese steel The chemical

compositions indicated are average for the specific data points plotted Test specimens were 25 mm (1 in.) diam bars, austenitized and quench-annealed from 1010 °C (1850 °F) or above Source: Ref 3

The 0.7% C (min) of grades D and E-1 may be used to minimize carbide precipitation in heavy castings or in weldments, and similar low carbon contents are specified for welding filler metal Carbides form in castings that are cooled slowly in the molds In fact, carbides form in practically all ascast grades containing more than 1.0% C, regardless of mold cooling rates They form in heavy-section castings during heat treatment if quenching is ineffective in producing rapid cooling throughout the entire section thickness Carbides can form during welding or during service at temperatures above about

275 °C (530 °F)

If carbon and manganese are lowered together, for instance to 0.53% C with 8.3% Mn or 0.62% C with 8.1% Mn, the workhardening rate is increased because of the formation of strain-induced α (body-centered cubic, or bcc) martensite However, this does not provide enhanced abrasion resistance (at least to high-stress grinding abrasion) as is often hoped (Ref 4)

Manganese contributes the vital austenite-stabilizing effect of delaying transformation (but not eliminating it.) Thus, in a simple steel that contains 1.1% Mn, isothermal transformation at 370 °C (700 °F) begins about 15 s after the steel is quenched to that temperature, whereas in a 13% Mn steel, transformation at the same temperature does not begin until after 48 h (Ref 1) Below 260 °C (500 °F), phase changes and carbide precipitation are so sluggish that for all practical purposes they may be neglected, in the absence of deformation, if manganese content exceeds 10%

Figure 5 shows the influence of manganese content on the strength and ductility of cast austenitic steel that has been solution treated and water quenched It confirms the observations of many investigators, including Sir Robert Hadfield (Ref 5), who studied the influence of manganese content up to about 22% Manganese content has little effect on yield

strength In tensile testing, ultimate strength and ductility increase fairly rapidly with increasing manganese content up to about 12% and then tend to level off, although small improvements normally continue up to about 13% Mn

Fig 5 Variation of properties with manganese content for austenitic manganese steel containing 1.15% C Data

are for castings weighing 3.6 to 4.5 kg (8 to 10 lb) and about 25 mm (1 in.) in section size that were water quenched from 1040 to 1095 °C (1900 to 2000 °F) Flow under impact is the total reduction in height sustained

by a cylindrical specimen 25 mm (1 in.) in both diameter and length after absorbing 20 blows of 680 J (500 ft · lbf) each Source: Abex Research Center

Silicon and Phosphorus. As noted in Table 1, silicon and phosphorus are present in all ASTM A 128 grades of austenitic manganese steel Silicon is seldom added except for steelmaking purposes Silicon content exceeding 1% is uncommon, because foundries do not like to have the silicon pyramid in melts containing returned scrap A silicon content of 1 to 2% might be used to increase yield strength to a moderate degree, but other elements are preferred for this

effect Loss of strength is abrupt above 2.2% Si, and Mn steel containing more than 2.3% Si may be worthless On the other hand, silicon levels below 0.10% show decreased fluidity during casting

The availability of low-phosphorus ferro-manganese since about 1960 has enabled steelmakers to reduce phosphorus levels in manganese steel to a large extent The preferred practice is to hold the phosphorus content below 0.04% even though 0.07% is permitted by ASTM A 128 Levels above 0.06%, which formerly were prevalent, contribute to hot shortness and low elongation at very high temperatures and frequently are the cause of hot tears in castings and underbead cracking in weldments It is particularly advantageous to keep phosphorus at the lowest possible level in the grades that are welded, and in manganese steel welding electrodes, and in heavy section castings

Common Alloy Modifications. The most common alloying elements are chromium, molybdenum, and nickel (see Table 1) Added to the usual carbon level of about 1.15%, both chromium and molybdenum increase yield strength (Fig 6) and flow resistance under impact

Fig 6 Effects of (a) chromium, (b) molybdenum, and (c) nickel contents on the tensile properties of cast

manganese steel Steel was cast in 25 mm (1 in.) diam test bars, reheated to 1095 °C (2000 °F), and water quenched Source: Ref 6

Chromium additions are less expensive for a given increase, and chromium grades (ASTM A 128, grade C, for instance) are probably the most common modifications ASTM A 128, grade B, often contains some chromium also The 2% chromium addition in grade C does not significantly lower toughness in light sections However, in heavier sections, its effect is similar to that of raising the carbon level; the result is a decrease in ductility due to an increase in the volume fraction of carbides in the microstructure Chromium additions have been used up to 6% for some applications, sometimes

in combination with copper, but these grades no longer receive much attention Chromium enhances resistance to both atmospheric corrosion and abrasion, although the latter effect is not always consistent and depends on the individual application It is also used up to 18% in low-carbon electrodes for welding manganese steel Because of the stabilizing effect of chromium on iron carbide, higher heat-treatment (solutionizing) temperatures are often necessary prior to water quenching

Molybdenum additions, usually 0.5 to 2%, are made to improve the toughness and resistance to cracking of castings in the as-cast condition and to raise the yield strength (and possibly toughness) of heavy-section castings in the solution-treated and quenched condition These effects occur because molybdenum in manganese steel is distributed partly in solution in the austenite and partly in primary carbides formed during solidification of the steel The molybdenum in solution effectively suppresses the formation of both embrittling carbide precipitates and pearlite, even when the austenite

is exposed to temperatures above 275 °C (530 °F) during welding or in service The molybdenum in primary carbides tends to change the morphology from continuous envelopes around austenite dendrites to a less harmful nodular form, especially when the molybdenum content exceeds 1.5%

The 1% molybdenum grades (ASTM A 128, grade E-1, and AWS A5.13, grade EFeMn-B) are resistant to the reheating effect that limits the usefulness of the standard B-2, B-3, and B-4 grades Grade E-1 is adapted to heavy-section castings used in roll and impact crushers that are frequently reheated during weld buildup and overlays

Grade E-2, which contains about 2% Mo, may be given a special heat treatment to develop a structure of finely dispersed carbides in austenite This heat treatment entails a partial grain refinement (U.S Patent 1,975,746) by pearlitizing near

595 °C (1105 °F) for 12 h and water quenching from 980 °C (1800 °F) This type of microstructure has been found to enhance abrasion resistance in crusher applications The tensile properties of specimens removed from worn cone crusher parts ranged from 440 to 485 MPa (64 to 70 ksi) in yield strength, 695 to 850 MPa (100 to 125 ksi) in tensile strength, and

15 to 25% in elongation

The addition of molybdenum in amounts greater than 1% can increase the susceptibility of the manganese steel to incipient fusion during heat treatment Incipient melting refers to a liquation phenomenon that occurs because of the presence of low-melting constituents in interdendritic areas, both within individual grains and along grain boundaries This tendency is aggravated by higher P levels (>0.05%), higher pouring temperatures (which promote segregation in the casting), and higher carbon levels (>1.3%) in the steel

As a further use, molybdenum is added to the lean manganese steel grade F partly to suppress embrittlement in both cast and heat-treated conditions

as-Nickel, in amounts up to 4%, stabilizes the austenite because it remains in solid solution It is particularly effective for suppressing precipitates of carbide platelets, which can form between about 300 to 550 °C (570 to 1020 °F) Therefore, the presence of nickel helps retain nonmagnetic qualities in the steel, especially in the decarburized surface layers Nickel additions increase ductility, decrease yield strength slightly, and lower the abrasion resistance of manganese steel Nickel

is used primarily in the lower-carbon or weldable grades of cast manganese steel and in wrought manganese steel products (including welding electrodes) In wrought products, nickel is sometimes used in conjunction with molybdenum The stability of nickel-containing manganese steels when reheated is shown in Table 2 Table 3 contains compositional and tensile data for nickel manganese steels used in naval applications

Table 2 Mechanical properties of three reheated austenitic manganese steels

impact strength(a) Condition

Elongation,

% (a)

Reduction in area, % (a)

(a) Average of two determinations

(b) 1000 kg load; average of six determinations

Table 3 Compositions and tensile properties of 14Mn-Ni steels meeting MIL-S-17758 (Ships)

Alloy(a) Element or property

Type 1 Type 2

Composition limits