DUCTILE IRON DATA FOR DESIGN ENGINEERS
SECTION IV. AUSTEMPERED DUCTILE IRON
What material offers the design engineer the best combination of low cost, design flexibility, good machinability, high strength-to-weight ratio and good toughness, wear resistance and fatigue strength? Austempered Ductile Iron (ADI) may be the answer to that question. ADI offers this superior combination of properties because it can be cast like any other member of the Ductile Iron family, thus offering all the production advantages of a conventional Ductile Iron casting. Subsequently it is subjected to the austempering process to produce mechanical properties that are superior to conventional ductile iron, cast and forged aluminum and many cast and forged steels.
Figures 4.1 and 4.2 compare the mechanical properties of ADI to those of conventional Ductile Irons. Figure 4.1 also provides a comparison between the tensile strength-elongation relationships for the ASTM A897 ADI specification and that of the ASTM A536 specification for conventional Ductile Iron. These, and other Ductile Iron specifications are discussed in further detail in Section XII. Compared to the conventional grades of Ductile Iron, ADI delivers twice the strength for a given level of elongation. In addition, ADI offers exceptional wear resistance and fatigue strength.
Figure 4.3a shows the strength of Ductile Iron and ADI compared to cast and forged steels. Ductile Iron has commercially replaced as cast and forged steels in the lower strength region, now ADI is finding applications in the higher strength regions. As shown in Figure 4.3b, the yield strength of ADI is over three times that of the best cast or forged aluminum. In addition ADI weighs only 2.4 times more than aluminum and is 2.3 times stiffer. ADI is also 10% less dense than steel. Therefore, when you compare the relative weight per unit of yield strength of ADI with that of various aluminums and steels (Figure 4.4) it is easy to see the engineering and design advantages inherent in ADI.
For a typical component, ADI costs 20% less per unit weight than steel and half that of aluminum. When we now analyze the cost-per-unit-strength of ADI vs. various materials (Figure 4.5) the economic advantages of ADI become apparent.
The mechanical properties of Ductile Iron and ADI are primarily determined by the metal matrix. The matrix in conventional Ductile Iron is a controlled mixture of pearlite and ferrite. (Tempered martensitic matrices may be developed for wear resistance, but lack the ductility of either as-cast Ductile Iron or ADI). The properties of ADI are due to its unique matrix of acicular ferrite and carbon stabilized austenite; called Ausferrite. The austempering process is neither new or novel and has been utilized since the 1930’s on cast and wrought steels. The austempering process was first commercially applied to Ductile Iron in 1972 and by 1998 world-wide production was approaching 100,000 tonnes annually.
The preponderance of information on the austempering of steel and the superficial similarities between the austempering heat treatments applied to steels and ADI, have resulted in comparisons which are incorrect and damaging to the understanding of the structure and properties of ADI. ADI is sometimes referred to as "bainitic Ductile Iron", but correctly heat treated ADI contains little or no bainite. Bainite consists of a matrix of acicular (plate-like) ferrite and carbide. ADI’s ausferrite matrix is a mix of acicular ferrite and carbon stabilized austenite. This ausferrite may resemble bainite metallographically, however it is not because it contains few or none of the fine carbides characteristic in bainite. An ausferrite matrix will only convert to bainite if it is over tempered.
The presence of austenite in ADI also leads to a harmful misconception. It is "retained" in the sense that it has persisted from the austenitizing treatment, but it is not the "retained austenite" that designers and metallurgists equate with unstable, incorrectly heat treated steel. The austenite in ADI has been stabilized with carbon during heat treatment and will not transform to brittle martensite even at sub-zero temperatures.
The presence of stable, carbon enriched austenite also accounts for another inadequately understood property of ADI. While thermodynamically stable, the enriched austenite can undergo a strain-induced transformation when exposed to high, normal forces. This transformation, which gives ADI its remarkable wear resistance, is more than mere "work hardening". In addition to a significant increase in flow stress and hardness (typical in most metallic materials), this strain induced transformation also produces a localized increase in volume and creates high compressive stresses in the "transformed" areas. These compressive stresses inhibit crack formation and growth and produce significant improvements in the fatigue properties of ADI when it is machined after heat treatment or subjected to surface treatments such as shot peening, grinding or rolling. (See Section IX).
ADI is a group of materials whose mechanical properties can be varied over a wide range by a suitable choice of heat treatment. Figure 4.6 illustrates the strong correlation between austempering temperature and tensile properties. A high austempering temperature, 750F (400C), produces ADI with high ductility, a yield strength in the range of 500 MPa (72 ksi) with good fatigue and impact strength. These grades of ADI also respond well to the surface strain transformation previously discussed which greatly increases their bending fatigue strength. A lower transformation temperature, 500F (260C), results in ADI with very high yield strength (1400 MPa (200 MPa)), high hardness, excellent wear resistance and contact fatigue strength. This high strength ADI has lower fatigue strength as-austempered but it can be greatly improved with the proper rolling or grinding regimen. Thus, through relatively simple control of the austempering conditions ADI can be given a range of properties unequaled by any other material.
Like other Ductile Iron specifications presented in Chapter XII, ASTM A897 defines the minimum tensile properties for different grades of ADI. Figure 4.1 indicates that the ranges of properties exhibited by ADI exceed these minima, but does not offer quantitative evidence on which materials selection decisions can be made with confidence. Competent producers of both conventional Ductile Iron and ADI recognize that they must not only provide statistically significant mechanical property data, but also give evidence of SPC and their commitment to continual improvement. Figure 4.7 provided by CMI International offers statistical evidence that grade 125/80/10 ADI can be produced with mechanical properties significantly in excess of those required by the specification. (The data shown represent 18 months of foundry production and over 600 heat treat lots).
The modulus of elasticity in tension for ADI lies in the range of 22.5-23.6 X 106 (155-163 GPa). Figure 4.8 shows the relationship of ADI’s Young’s Modulus to that of other materials.
Traditionally, components have been designed on the basis of preventing failure by plastic deformation. As a result, design codes used either the 0.2% yield stress or the ultimate tensile stress when specifying material properties, and designers then applied a safety factor when determining the acceptable working stress level in the component. Both designers and metal casters have recognized that structures also fail by brittle fracture and fatigue, especially in the presence of a crack-like defect. As a result, fracture toughness, the intrinsic resistance of the material to crack propagation, is becoming an essential part of the package of material properties used by designers to select materials for critical applications.
There is a dearth of fracture toughness data for ADI, for two very valid reasons. First, being a relatively new material, efforts to define a mechanical property database have concentrated on the more conventional and easily acquired tensile properties, with early efforts at defining toughness being confined largely to an extension of the notched and un-notched Charpy test used to characterize conventional Ductile Irons. Second, the more ductile grades of ADI, for which fracture toughness is a critical property, do not behave in a linear elastic manner when subjected to standard LEFM tests, and toughness must be determined by yielding fracture mechanics techniques such as the J integral crack opening displacement (COD) tests. Nevertheless, when the combined knowledge of the toughness properties is assembled, they reveal three important facts:
Figure 4.2 shows that ADI heat treated to produce high strength has a static fracture toughness of 55-70 MPa(m)1/2, which is greater than that of Ductile Iron with a matrix of tempered martensite or pearlite. Furthermore, it shows that ADI heat treated to a lower strength (higher ductility) grade has a fracture toughness in excess of 100 MPa(m)1/2, twice as tough as pearlitic Ductile Iron. When compared to forged steel and conventional Ductile Iron, both with equal or inferior mechanical properties, ADI exhibits un-notched Charpy impact values at room temperature that are less than those of forged steel, but three times higher than conventional Ductile Iron (Table 4.1). Figure 4.9 shows that the room temperature un-notched Charpy values of ADI are substantially higher than those required by ASTM A897-90 at all levels of strength.
Table 4.1 Comparison of the mechanical properties of forged steel, pearlitic Ductile Iron and Grade 150/100/7 ADI.
** Un-notched charpy at room temperature.
Table 4.2 provides a comparison of yield strength and fracture toughness, K1C, between ADI , conventional austenitic Ductile Irons, and quenched and tempered AISI 4140 and 4340 steels. With K1C values in the range of 59-86 MPa m1/2, ADI had a fracture toughness which was superior to all other Ductile Irons, except Ni-Resist, and equal to or higher than most of the quenched and tempered steels. This table also compares the ratio of K1C to yield strength for these materials. This ratio, which is proportional to the size of flaw that can be tolerated when materials are stressed to a constant fraction of their yield strength, indicates that ADI has equal or greater flaw tolerance than pearlitic Ductile Iron and quenched and tempered steels.
Table 4.2 Comparison of yield strength, fracture toughness and flaw tolerance between ADI, conventional and austenitic Ductile Irons, and quenched and tempered steels.
* Isothermal Transformation Time: 1 hour.
The ASME Gear Research Institute used both the ASTM Short Rod Fracture Test (SRFT) and the non-standard Single Tooth Impact (STI) test to evaluate various ADI materials for gear applications and compare their fracture toughness with carburized 8620 steel. Figure 4.10 shows ASTM Short Rod Fracture Toughness for ADIs ranging from 55 to 105 MPa (m)1/2 at room temperature. This compares favorably with a reported room temperature fracture toughness of 22 to 33 MPa (m)1/2 for carburized and hardened 8620 steel. The ADI toughness levels increase strongly with increasing austempering temperature. (The austempering condition is indicated in degrees F and hours; i.e. ADI 750(.75 hr) was austempered at 750F for 45 minutes).
The Single Tooth Impact test results shown in Figure 4.11 are consistent with Short Rod Fracture Toughness data, with ADI superior to carburized steel for both peak load and fracture energy criteria at both room temperature and -40 degrees.
Figure 4.12 illustrates ADI dynamic fracture toughness data which were produced by instrumented impact tests performed on Charpy-size specimens. The results, in the range of 40 to 80 MPa(m)1/2, show that toughness decreases with higher levels of manganese.
Regardless of the type of toughness test, ADI results were superior to those of conventional Ductile Iron, and were equal to, or better than competitive steels. Additionally, all toughness tests revealed that the toughness of ADI increases with austempering temperature to a maximum around 650-700F (340-370C). Figure 4.13 confirms that this relationship is a further manifestation of the influence of the volume fraction of stabilized austenite on the ductility and toughness of ADI.
An extensive work done at National University of Mar del Plata (Argentina) compared the properties of ADI to those of 4140 steel. It concluded that while a standard notched Charpy test indicated that the properties of ADI were inferior to those of 4140 steel, fracture toughness tests indicate "a much less significant difference". In fact, it was found that the strain rate of the fracture test had a much more significant effect on the steel than on the ADI. It concluded that "the comparison of toughness of ADI and steels, should not be based on the impact energy measurements. Fracture mechanics properties, such as K1C, should be used for design purposes".
As shown in Figure 4.14, ADI has fatigue properties equal or superior to those of forged steels. When subjected to surface treatments such as rolling, peening or machining after heat treatment, the fatigue strength of ADI is increased significantly. (See Figures 3.34, 3.35 and Table 3.3 in Section III, and Figures 4.35 and 4.36. Figures 4.14 and 4.15 also indicate that ADI is moderately notch sensitive in fatigue, with a notch sensitivity ratio (ratio of notched to un-notched endurance limits) ranging from 1.2 to 1.6 (see Figure 4.16) for the notch geometry tested. Conventional ferritic and pearlitic Ductile Irons have a notch sensitivity of about 1.6 and steels with fatigue strengths similar to ADI exhibit notch sensitivity ratios as high as 2.2-2.4. To avoid problems caused by notch sensitivity, components with sharp corners should be redesigned to provide generous fillets and radii. When required, fillet rolling or shot peening can be employed to further increase resistance to fatigue failure.
Figure 4.15 relates the fatigue strengths of notched and un-notched ADI to tensile strength and austempering temperature. Comparison of this figure with Figure 4.13 reveals several interesting facts. Unlike conventional Ductile Iron, the un-notched fatigue limit of ADI does not follow the tensile properties, demonstrates a maxima at the condition of lower tensile strength and maximum stabilized austenite content in the metal matrix. These relationships result in an endurance ratio (ratio of fatigue strength to tensile strength) that is 0.5 for lower strength ADI and decreases to 0.3 as the tensile strength increases to its maximum. (See Figure 4.16). The notched ADI fatigue strengths shown in Figure 4.15 increase rapidly with tensile strength and, as a result, high strength ADI is the least notch sensitive.
Figure 4.17 shows an important relationship between austempering temperature and the endurance limit of shot peened ADI. The dramatic rise in endurance limit in ADIs austempered above 600F (315C) is related to the increased response to peening resulting from the higher austenite contents characteristic of higher austempering temperatures.
Figures 4.18 and 4.19 indicate that for gear applications, shot peened ADI has single tooth bending fatigue and contact fatigue superior to as cast and conventionally heat treated Ductile Irons, and cast and through hardened steels. They also show that peened ADI is competitive with gas nitrided and case carburized steels.
To this point we have discussed fatigue strength in terms that assume infinite life below a certain load. In fact, as the number of loading cycles are increased all materials undergo changes in their ability to withstand further loading. Figure 4.20 shows the relationship between number of cycles and the allowable single tooth bending stress (in ksi). Grade 2 ADI represents an ADI austempered at 675F (357C) and Grade 5 represents an ADI austempered at 500F (260C). Figure 4.21 shows the relationship between the number of cycles and the allowable contact stress for those same ADIs. Figure 4.22 shows the relationship of number of cycles to allowable rotating bending stress for a Grade 1050 ADI (austempered at 675F (357C)) in the as-austempered condition.
To accurately model the finite element behavior of materials that are dynamically loaded the design engineer uses certain coefficients and exponents to predict a component’s fatigue behavior. As of this writing much work is being done to develop those numbers for Ductile Irons and ADI. Table 4.3 shows the typical fatigue coefficients and exponents for a 300 BHN ADI. (Courtesy of Ford Motor Company and Meritor Heavy Vehicle Systems).
Table 4.3 Typical fatigue coefficients and exponents for 300 BHN ADI
Austempered Ductile Iron offers the design engineer abrasion resistance that is superior to competitive materials over a wide range of hardness. Generally, ADI will outwear competitive materials at a given hardness level. For example (from Figure 4.23) an ADI component at 30 to 40 Rc will wear comparably to a quenched and tempered steel component at nearly 60 Rc in an abrasive wear environment. This property, shown in Figures 4.23 and 4.24, allows the designer to select the combination of strength, ductility and abrasion resistance that will provide the best component performance in a particular application.
The superior abrasion resistance, and the low sensitivity of abrasion resistance to bulk hardness are related to the strain-induced transformation of stabilized austenite which occurs when the surface of an ADI component is subjected to deformation. The result of this transformation is a significant increase in surface hardness shown in Figure 4.25. This increase in surface hardness, and its relationship to microstructure, are responsible for the reduced sensitivity of abrasion resistance to hardness. As the bulk hardness of ADI is reduced by the austempering temperature, the amount of stabilized austenite increases (see Figure 4.26). This increase in austenite content increases the hardness increment produced by surface deformation. As a result, a Ductile Iron component austempered to produce a lower hardness displays an abrasion resistance greater than that predicted by its bulk hardness, provided that the abrasion mechanism involves sufficient deformation to transform the surface layers to martensite.
Through variations in austempering conditions, the designer can optimize the abrasion resistance and related mechanical properties of an ADI component. For a combination of high toughness and abrasion resistance an austempering temperature in the range of 650-700F (350-375C) should be used. When a combination of high strength and abrasion resistance are required, an austempering temperature of 500F (260C) will yield the best results.
Compared to competitive materials, the machinability of Ductile Iron has been one of its major advantages. When the substantial increases in strength and wear resistance offered by ADI are considered, it would be logical to assume that ADI could present machining problems. However, cost savings in machining are frequently mentioned as reasons for converting to ADI. The reasons for this surprising combination of mechanical properties and machinability are two-fold. First, the machinability of the softer grades of ADI is equal or superior to that of steels with equivalent strength, and second, the predictable growth characteristics of ADI during austempering allow, in many cases, for it to be machined complete in the soft as-cast or annealed state before heat treatment. This allows for faster machine feeds and speeds and greatly increased tool life. (See Figure 4.27).
As discussed in a later section on surface deformation treatments of gears, the processing of ADI parts should could follow one of several paths, (Figure 4.34), depending primarily on the grade of ADI, but also on the surface treatment benefits that may be gained from machining after austempering. As shown in Figures 4.28 and 4.29, cutting tool life decreases substantially as the hardness of ADI increases and the cutting speed and metal removal rate increase. For these reasons, only the 125/80/10 and 150/100/07 grades of ADI should be machined after austempering. Parts processed to the higher strength grades should receive the following processing sequence:
While annealing adds cost to the casting the benefits of more predictable dimensional change during austempering and greatly improved machinability often more than offset the cost of annealing. In order to obtain the benefits of the excellent machinability of annealed Ductile Iron, the designer must have confidence in the reproducibility of the growth of the machined casting during heat treatment. Papers presented at the 1st International Conference on ADI indicated that dimensional changes during heat treatment varied from a slight contraction to a growth of approximately 0.4%, (Figure 4.30). It was stressed that as long as the prior Ductile Iron microstructure was consistent in pearlite/ferrite ratio, predictable growth occurred and close tolerance ADI parts could be produced successfully by machining prior to heat treatment.
The production of ADI ring and pinion gear sets for General Motors rear wheel drive cars (model years 1977-1979) provides a good example of the reduction in machining costs offered by ADI. As shown in Table 4.4, tool life improvement for different machining operations performed on the ferritized ADI blanks ranged from 20% to over 900% compared to similar operations required to produced the forged, carburized and hardened 8620 steel gears that they replaced. The overall cost savings to GM of converting from carburized steel to austempered ring and pinion gears was approximately 20%.
Tables 4.5 and 4.6 provide guidelines for the machining of the lower strength grades of ADI. In general, reduced surface speeds and increased feed rates provide the best metal removal rate in ADI.
The mechanical properties offered by ADI make it an attractive material for demanding applications in which strict specifications must be met consistently. While greatly enhancing the properties of a conventional Ductile Iron casting, the ADI process cannot compensate for casting defects that would impair mechanical properties. ADI castings should, therefore, be produced free from surface defects, and with the following microstructural parameters.
These requirements are essentially the same as those required to produce good quality Ductile Iron. To assure quality, ADI should be purchased from casting and heat treatment suppliers that have well developed process control systems who can demonstrate that they are consistently capable of producing high quality castings and heat treatments.
The production of a high quality casting is essential but, by itself, not a sufficient condition to ensure optimum properties in ADI. The casting must be heat treated properly by a supplier capable of taking into account the interaction between casting dimensions, composition, microstructure and the desired properties in the austempered casting. For this reason, there should be close cooperation between the designer, metal caster, heat treater and machine source from conception of the design to delivery of the castings.
In many cases, the composition of an ADI casting differs little from that of a conventional Ductile Iron casting. When selecting the composition, and hence the raw materials, for both conventional Ductile Iron and ADI, consideration should be given first to limiting elements which adversely affect casting quality through the production of non-spheroidal graphite, or the formation of carbides and inclusions, or the promotion of shrinkage. The second consideration is the control of carbon, silicon and the major alloying elements (See Table 4.7) that control the hardenability of the iron and the properties of the transformed microstructure. When determining the alloying requirements both the section size and type and the severity (or speed) of the austempering quench must be considered.
For a typical salt quench with agitation section sizes up to about 3/8 inch (10 mm) can be successfully through hardened without pearlite with even unalloyed Ductile Iron. For a highly agitated austemper quench with water saturation section sizes of up to ¾ inch (20 mm) can be through hardened with no additional alloying. For castings of heavier section size selective alloying is required to through harden the parts and avoid pearlite in the heat treated microstructure.
Figure 4.31 summarizes the hardenability of copper, nickel and molybdenum by relating the levels of these elements to the maximum diameter bar that can be satisfactorily austempered. (The BCIRA data utilizes a relatively fast quench while the Dorazil data reflects an austempering bath with a lower quench severity). The relative hardenability contribution of manganese is between that of molybdenum and copper.
For economic reasons, or to avoid metallurgical problems, combinations of alloys are often used to achieve the desired hardenability in ADI. To avoid micro-segregation and the resultant degradation of mechanical properties associated with higher levels of manganese and molybdenum, their levels should be carefully controlled with the desired hardenability obtained by supplementary additions of first copper (up to about 0.8%), then nickel.
The primary purpose of adding copper, nickel or molybdenum to ADI is to increase the hardenability of the matrix sufficiently to ensure that the formation of pearlite is avoided during the austempering process. These elements have only marginal effect on the mechanical properties of ADI that is properly austempered. (The austempering process determines the properties after austempering; not the alloying). Only the minimum amount of alloys required to through harden the part should be employed. Excessive alloying only increases the cost and difficulty of producing the good quality Ductile Iron necessary for ADI. Ultimately, the amount of alloying required will be a function of the metal caster’s base composition, the casting section size and type and the characteristics of the chosen heat treatment process.
A typical iron composition (and control range) that can be used is shown below:
*(Carbon and silicon should be controlled to produce the desired carbon equivalent for the section size being produced).
The aforementioned composition does not guarantee ADI properties, nor is it mandatory. However, this composition is a typical, industrially successful ADI composition. A good controlled chemistry like this one combined with a consistently high nodule count and nodularity and a consistent pearlite/ferrite ratio in clean, shrink-free Ductile Iron will provide the most robust process for the production of ADI.
ADI is produced by an isothermal heat treatment known as austempering. Austempering consists of the following steps as shown in Figure 4.32:
The austenitizing may be accomplished by using a high temperature salt bath, an atmosphere furnace or (in special cases) a localized method such as flame or induction heating. The austempering is most typically carried out in a nitrite/nitrate salt bath but in special cases it can be accomplished in hot oil (up to 470F (243C)), or molten lead or tin.
The critical characteristics are:
The austenitizing temperature controls the carbon content of the austenite which, in turn, affects the structure and properties of the austempered casting. High austenitizing temperatures increase the carbon content of the austenite, increasing its hardenability, but making transformation during austempering more problematic and potentially reducing mechanical properties after austempering. (The higher carbon austenite requires a longer time to transform to ausferrite). Reduced austempering temperatures generally produce ADI with the best properties but this requires close control of the silicon content, which has a significant effect on the upper critical temperature of the Ductile Iron.
Austenitizing time should be the minimum required to heat the entire part to the desired austenitizing temperature and to saturate the austenite with the equilibrium level of carbon, (typically about 1.1-1.3%). In addition to the casting section size and type, the austenitizing time is affected by the chemical composition, the austenitizing temperature and the nodule count.
Austempering is fully effective only when the cooling rate of the quenching apparatus is sufficient for the section size and hardenability of the component. The minimum rate of cooling is that required avoid the formation of pearlite in the part during quenching to the austempering temperature. The critical characteristics are as follows:
The use of a correctly designed austempering system with a suitably high quench severity, and the correct loading of castings, can minimize hardenability requirements of the casting resulting in significant savings in alloy costs.
As illustrated earlier in Figure 4.6, austempering temperature is one of the major determinants of the mechanical properties of ADI castings. To produce ADI with lower strength and hardness but higher elongation and fracture toughness, a higher austempering temperature (650-750F (350-400C)) should be selected to produce a coarse ausferrite matrix with higher amounts of carbon stabilized austenite (20-40%). Grades 125/80/10 and 150/100/07 would be typical of these conditions. To produce ADI with higher strength and greater wear resistance, but lower fracture toughness, austempering temperatures below 650F (350C) should be used.
Once the austempering temperature has been selected, the austempering time must be chosen to optimize properties through the formation of a stable structure of ausferrite. Figure 4.33 schematically illustrates the influence of austempering time on the stabilization of austenite, and shows the hardness of the resultant matrix. At short austempering times, there is insufficient diffusion of carbon to the austenite to stabilize it, and martensite may form during cooling to room temperature. The resultant microstructure would have a higher hardness but lower ductility and fracture toughness (especially at low temperatures). Excessive austempering times can result in the decomposition of ausferrite into ferrite and carbide (bainite) which will exhibit lower strength, ductility and fracture toughness. At the highest austempering temperature (750F (400C)) as little as 30 minutes may be required to produce ausferrite. At 450F (230C) as much as four hours may be required to produce the optimum properties; Figure 4.32.
Note: a strength level maxima is achieved in ADI at an austempering temperature of about 475-525F (250-275C). At temperatures below that range the hardness may increase but the strength may decrease due to the presence of martensite mixed in with the ausferritic matrix. (In other words, as the austempering temperature is incrementally decreased below 475F (250C) the material behaves increasingly like a quenched and tempered Ductile Iron).
The austempering treatment used to produce ADI can result in small residual tensile surface stresses. Even so, ADI has excellent fracture toughness and fatigue strength. However, ADI’s ability to resist the initiation of fatigue cracks (especially in cyclic bending) may be greatly enhanced by inducing compressive stresses to the surface after the austemper heat treatment.
These small tensile stresses can be easily replaced with rather substantial compressive stresses if the ADI is subjected to any surface treatment involving the sufficient surface deformation to cause a strain-induced transformation of the stabilized austenite. Such treatments could include conventional machining operations such as turning, grinding, milling, hobbing or special treatments such as shot peening or surface rolling.
For parts subjected to fatigue failure, performance can be enhanced significantly if machining operations can be performed after austempering. However, this processing sequence is limited by the machinability of the different grades of ADI. Figure 4.34 shows the two processing flow charts used by the ASME Gear Research institute to produce ADI test gears. Using machinability as the main criterion, the GRI selected process #2 for blanks with heat treated hardnesses in the range of 30-34 HRC and process #1 for blanks with hardness greater than 37 HRC. Figure 4.35 reveals that hobbing after austempering resulted in a 20% increase in fatigue strength of gear teeth.
Shot peening offers a controllable means of selectively hardening certain parts of a finished casting to produce significant improvements in fatigue properties. Figure 4.35 shows that shot peening a hobbed gear increased fatigue strength by 60%. Single tooth fatigue tests conducted by the GRI on ADI gears indicated that shot peening doubled the fatigue strength. The GRI work also identified a significant correlation between residual compressive stress produced by peening and the endurance limit (Figure 4.36). In addition to being able to be applied to selected areas of a part, peening also offers the advantage of increasing surface hardness without detrimentally affecting the ductility and strength of the remainder of the component. Post peening surface treatments such as honing may be required to reduce surface roughness in parts subjected to rolling contact fatigue. (For more information see Section IX).
Surface rolling or burnishing produced with a hardened roller, or mating part in the case of gear hardening, may be used independently or as a post-peening operation to improve the fatigue properties of ADI. Gear Research Institute tests show that roll burnishing of a peened gear with a carburized steel mate produced a six-fold increase in fatigue life. Fillet rolling has been used very effectively to offset reductions in fatigue life produced by stress concentrations at changes in cross section in castings such as crankshafts and gear clusters (see Table 3.3 and Section IX).
There are many properties needed for specific design applications that are not currently published. Without these important properties engineers must make assumptions about a material’s behavior and mistakes can be made, (as discussed in some of the referenced papers listed in this chapter). Table 4.9 summarizes some of those properties for the five ASTM897 grades.
The ADI market represents nearly all segments of manufacturing. Figure 4.37 shows the approximate breakdown of the North American ADI market.
Heavy Truck and Bus Components
Light Auto and Truck Components
Truck spring support in ADI which replaced cast steel (Muhlberger GGG100B/A)
Construction and Mining Components
Assorted ADI conveyor components
In North America, ADI achieved a major breakthrough in 1977, when General Motors converted a forged and case hardened steel ring gear and pinion to ADI for Pontiac rear drive cars and station wagons. The decision came after nine years of development work and six years of field testing. The automaker was able to gain both significant cost savings and product improvement by changing to ADI.
ADI Timing Gears for Cummins B-Series diesel engines. Replaced forged and case carburised 1022 steel with 30% cost saving.
In 1983, the Cummins Engine Co. began to use ADI timing gears, produced to AGMA class 8 standards, in its B and C series diesel engines. These gears were machined and hobbed from annealed Ductile Iron castings. A crown shaving operation was carried out on the gear teeth prior to austempering, and the only operations performed after austempering were the grinding of the bore diameter and shot peening. Annual production exceeds 30,000 sets and the cost savings are estimated at 30% compared to the forged and carburized 1022 steel gears previously used.
Table 4.8 Comparison of energy requirements for the production of ADI and forged and carburized steel gears.
Table 4.8 describes the energy savings of almost 50% resulting from the conversion to ADI gears. In addition to savings in energy and overall production costs, ADI gears offer the following advantages:
Austempered Ductile Iron gears to patented specifications K9805.
Austempered Ductile Iron Hypoid Axle Gears: Conversion to Cast Ductile Iron from Forged Steel gave:
Crankshafts are another potentially significant application for Austempered Ductile Iron. The first commercially produced ADI part was a small crankshaft for a hermetically sealed refrigerator compressor. It was cast by Wagner Castings Company (US) for Tecumseh Products. Production of that part was initiated in 1972 and since that time, millions of those crankshafts have been produced.
Engines being developed by the automotive industry require weight reduction in parts that will be required to handle increased power. Automotive design engineers have evaluated ADI as a candidate for both the replacement of forged steel crankshafts and the upgrading of existing Ductile Iron crankshafts. The Ford Motor Company made an exhaustive, three year study of ADI crankshafts and concluded that they met all design criteria. During this study, the importance of fatigue testing was identified, and the following results were obtained:
A thorough, joint Motor Industry Research Association / Cast Metals Development Laboratories study on ADI crankshafts concluded that properly fillet rolled ADI crankshafts exhibited fatigue properties comparable to, or better than, the best forged and heat treated steel crankshafts.
In another documented crankshaft study conducted at the Manchester (England) Materials Science Center, the authors demonstrated the performance capability of ADI crankshafts in one cylinder commercial and four cylinder automotive engines. They noted a 10% rotating weight reduction and an estimated 30% cost savings.
As of this writing ADI crankshafts are employed in high volume commercial applications and low volume automotive applications. As the specific power requirements for automotive engines are increased, ADI will become a more viable alternative to the heavier, more expensive forged steel crankshaft.
An assortment of ADI "ground engaging" parts. Courtesy of Applied Process Inc.
ADI crawler type track shoe. Courtesy of Applied Process Inc.
ADI crankshaft for a hermetically sealed compressor (first produced in 1972).
As design engineers become more familiar with ADI’s strength, toughness, wear resistance and noise damping properties and learn about the impressive cost and weight savings reported in successful ADI conversions from steel castings, weldments and forgings and aluminum castings and forgings, ADI will continue its remarkable growth.
E. Dorazil, "Mechanical Properties of Austempered Ductile Iron", Foundry Management & Technology, July, 1986, 36-45.
T. Luyendijk and H. Nieswaag, "The Influence of Silicon on the Toughness of Bainitic Ductile Iron." Proceedings of the 50th International Foundry Conference, Cairo, 1983.
A Design Engineer's Digest of Ductile Iron, 5th Edition, Qrr-Fer et Titane Inc., Montreal, Quebec, Canada, 1983.
"Ductile Iron," Foundry Management & Technology, September, 1989.
R. A. Harding, "Effects of metallurgical process variables on austempered ductile irons," Metallurgical Materials, 1986.
D. J. Moore, K. B. Rundman, and T. N. Rouns, "The Effect of Thermomechanical Processing on Bainitic Formation in Several Austempered Ductile Irons," Proc. lst Conf . on Austempered Ductile Irons, Chicago, IL, 1984.
T. Shiokawa, "Austempering of nodular cast irons, their mechanical properties and some practical applications," 59th Japan Ductile Cast Iron Association Conference, 1978.
M. Johansson, "Austenitic-Bainitic Ductile Iron," Transactions, American Foundrymen's Society, 1977.
P.A. Blackmore and R.A. Harding, "The effects of metallurgical process variables on the properties of austempered ductile irons.", presented at the lst International Conference on Austempered Ductile Iron, April 2-4, Rosemount, Chicago, Illinois, USA.
R. D. Forest, "The challenge and opportunity presented to the SG iron industry by the development of austempered Ductile Iron," BCIRA International Conference, "SG Iron the next 40 years", 1987.
G. F. Ruff, "Comparison of Austempered Ductile Iron Versus Steel," American Foundrymen's Society Congress, Hartford, Connecticut, 1988.
S.C. Lee and C.C. Lee, "The Effects of Heat Treatment and Alloying Elements of Fracture Toughness of Bainitic Ductile Cast Iron.", AFS Transactions, vol 96, pp 827-838, 1988.
S. Shepperson and C. Allen, "The Abrasive Wear Behaviour of Austempered Spheroidal Cast Irons," International Conference on Wear of Materials, Houston, Texas, 1987.
M. Kurkinen, "Wear-resistant components in nodular cast iron.", Werkstatt und Betrieb, 1981, vol 114, No 3, pp 165-168.
H. Muhlberger, "Progress in the production, understanding, and application of Germanite austempered Ductile Iron.", 2nd International Conference on Austempered Ductile Iron, March 17-19, 1986, Ann Arbor, Michigan, USA.
T. E. Prucha and G. F. Ruff, "Implementing Advanced Materials Through Integrated Casting Design and Manufacturing," High Integrity Castings, ASM International, 1988.
J. Janowak and R. Gundlach, "Development of a Ductile Iron for Commercial Austempering," Transactions, American Foundrymen's Society, 1983.
R. D. Forest, "Austempered Ductile Iron for Both Strength and Toughness," Machine Design, September, 1985.
M. Grech and J. M. Young, "Impact Properties of a Cu-Ni Austempered Ductile Iron," Cast Metals, 1988, vol 1 (2), 98-103.
M. Grech and J. M. Young, "Influence of Austempering Temperature on the Characteristics of Austempered Ductile Iron Alloyed with Cu and Ni.", Paper No. 90-160, presented at the American Foundrymen's Society Congress, April 20-24, 1990, Detroit Michigan, USA.
M. Gagne, "The Influence of Manganese and Silicon on the Microstructure and Tensile Properties of Austempered Ductile Iron," Transactions, American Foundrymen's Society, 1985.
J. F. Janowak, "Iron Foundries Facing Tough Questions on ADI" journal of Heat Treating, May, 1985.
AMAX Inc., "Austempered Ductile Iron," AMAX Publication M-626, 1985.
R. A. Harding, "Austempered Ductile Iron Gears," Materials and Design, August/September, 1985.
N. M. Lottridge and R. B. Grindahl, "Nodular iron hypoid gears.", Proceedings of the SAE Conference on Fatigue, Dearborn, Michigan, USA, April 14-16, 1982.
A. G. Fuller, "Austempered Ductile Irons - Present Applications," Materials and Design, June/July, 1985.
I. C. H. Hughes, "Austempered Ductile Irons - Their Properties and Significance," Materials and Design, June/July, 1985.
Y. J. Park, P. A. Morton, M. Gagne, R. Gollar, "Continuous Cooling Transformation Diagrams and Austempering study of Cu-Mo Ductile irons," Transactions, American Foundrymen's Society, 1984.
W. E Gruver, "Air Hardening Cast Iron," U. S. Patent 3,334,120, 1970.
R. Christ and Cynthia Krist, et al, "Model Certification Specification-Austempered Ductile Iron (ADI)", Ductile Iron Society Department of Army Contract DAAA08-93-C-0066 (1995).
J. Keough, "An ADI Market Primer", Foundry Management & Technology, October/November 1995.
B. Kovacs, "Development of Austempered Ductile Iron for Crankshafts", ASME 2nd International Conference on Austempered Ductile Iron, Ann Arbor, Michigan, USA 1986.
K. Hayrynen, "Another Avenue for Ductile Iron Foundries", Modern Casting Magazine, August 1995.
J. Keough, "ADI-A Designer Gear Material", Gear Technology-The Journal of Gear Manufacturing, 1995.
R. Martinez, R. Boeri and J. Sikora, "Impact and Fracture Properties of ADI, Comparison with 4140 Steel", American Foundrymens Society Congress, 1998. Paper No. 98-008.
Keough, "ADI vs. Aluminum - No Contest", International Pig Iron Secretariat Competition, 1995.
B. Kovacs, J. Keough, D. Pramstaller, "Austempered Ductile Iron Process Development", Gas Research Institute Contract 5085-235-1183, (Dec. 1988).
M. Bahmani, R. Elliot and N. Varahram, "Austempered Ductile Iron: a Competitive Alternative for Forged Induction-Hardened Steel Crankshafts", International Journal of Cast Metals Research, 1997, 9,249-257.
T. Chatterley, P. Murrell (MIRA/CDC), "ADI Crankshafts-An Appraisal of Their Production Potential, SAE Technical Paper Series, No. 980686. SAE Congress, Detroit, Michigan, USA, February 1998. (Reprinted from: Advancements in Fatigue Research and Applications SP-1341).