To recap some of what we saw last month, we set out the basic concepts and began reviewing the experimental procedure and results. These results are discussed and summarized here.
The mechanical-property evaluation shows that the pre-alloyed material system has higher ultimate tensile strength (~3%), yield strength (~5%) and apparent hardness (~6%) at all tempering temperatures compared to the diffusion-bonded material system. The diffusion-bonded material has slightly higher impact energy compared to the pre-alloyed material. These results are in line with the MPIF standard 35 published data.
The tensile strength of both materials stays relatively constant at tempering temperatures of 160-220°C (300-425°F). A sharp decline in tensile strength occurs as tempering temperatures reach beyond 220°C (425°F), from where both materials measured a 7% drop in tensile strength from 220-275°C (425-525°F). In contrast, the yield strength steadily increases as tempering temperatures approach 230°C (450°F) before declining again as temperatures go beyond 250°C (475°F).
The impact energy shows a steady decrease at tempering temperatures above 200°C (400°F) for both material systems. Both material systems show a linear decrease in apparent hardness and microhardness over the full tempering-temperature range. Based on this study, the optimum tempering temperatures range between 220-250°C (425-475°F), where both materials achieve high tensile strengths while maintaining good yield strength, impact energy, apparent hardness and microhardness. A summary of the properties for each material system is shown in Table 4.
The microstructures of the tensile specimens tempered at 200°C (400°F) are shown in Figure 9. The FLN2-4405 microstructures at each tempering temperature consist of martensite with nickel-rich austenite where the nickel was present. The FD-0205 tensile specimens also consist of martensite with nickel-rich austenite where the nickel was diffusion-bonded.
In order to understand how the mass of a component affects the microstructure at each tempering temperature, a study was conducted using two different-size puck specimens manufactured from the FLN2-4405 material. For this investigation the pucks were heat treated and tempered using the same parameters as the specimens submitted for mechanical-property testing. The microstructure of the surface and core was evaluated on the pucks and compared to the tensile specimens. The measurements of each specimen are shown in Table 5.
The quenching process during heat treatment relies on the transfer of heat to the quenching medium in order to rapidly cool the components. The mass effect of a component will determine how quickly it is cooled, and various cooling rates may be seen throughout different sections of the component. A component with small cross sections will cool more rapidly compared to large cross sections. In the same respect, the outer surface layer of larger components will cool faster than the core, resulting in variations in hardness and microstructure the closer to the center of the part.
The microstructures of the core of the 40-mm and 100-mm pucks at different tempering temperatures after heat treatment are shown in Figure 10. This study shows that the 40-mm pucks were able to achieve a fully martensitic microstructure with no bainite observed in the core. The 100-mm pucks have a martensitic surface, but bainite was observed within the core after heat treatment (~20%). Tempering has no effect on the microstructures.
Comparatively, the microstructures of the tensile specimens with the smaller cross section also developed a fully martensitic microstructure for both material systems. All specimens showed a similar drop in microhardness from the surface to the core due to stress relief as a result of tempering. This study shows that there is no mass effect on components with a cross-sectional area less than 1,000 mm2, but a mass effect is observed on components with a cross-sectional area of 2,500 mm2. The mass effect is due primarily to heat treatment and not tempering temperature.
Figure 11 shows the phase-mapping percentage of martensite at incremental distances from the part surface into the core on a 100-mm puck specimen tempered at 200°C (400°F). The component is 100% martensite at the part surface and 1 mm below the surface with no bainite observed. A small amount of bainite is observed and the martensite percentage slightly decreases at 2 mm below the surface. A large amount of bainite is present and the martensite percentage drops significantly at 3 mm below the surface. The core of the component measures approximately 80% martensite.
Figure 12 shows images taken by a scanning electron microscope to compare the differences between untempered martensite and the tempered martensite. The untempered martensite consists of thick, plate-like needles throughout its structure. This structure is a result of carbides being trapped in the crystal lattice during heat treatment (quenching), resulting in high internal stresses throughout the structure.
As the tempering temperature increases, alterations occur within the microstructure as the carbon is precipitated out of the crystal lattice. The atoms rearrange and form dispersed, spherical carbides in the martensite. This alteration forms a new structure called “tempered martensite” and has lower internal stresses compared to the untempered martensite. The tempered-martensite structure supports the mechanical properties obtained so that as the tempering temperatures increase, the internal stresses decrease and the strength, apparent hardness, impact energy and microhardness also decrease.
- The pre-alloyed material system has higher ultimate tensile strength (~3%), yield strength (~5%) and apparent hardness (~6%) at all tempering temperatures compared to the diffusion-bonded material system.
- The diffusion-bonded material has slightly higher impact energy compared to the pre-alloyed material.
- A sharp decline in tensile strength occurs as the tempering temperature goes beyond 220°C. Both material systems saw a 7% drop in tensile strength above this temperature.
- The yield strength increases as the tempering temperatures approach 250°C (475°F), but it decreases above this temperature.
- The impact energy of both material systems is similar within the tempering range of 160-200°C (325-400°F). As the tempering temperature increased above 200°C (400°F), the impact energy of FLN2-4405 decreased 24%, while the impact energy of FD-0205 material decreased 29%.
- The apparent hardness levels decrease linearly as the tempering temperature increases. The FLN2-4405 material exhibited higher apparent hardness levels at all tempering temperatures compared to the diffusion-bonded FD-0205 material
- The microhardness of the tensile specimens decreases linearly as the tempering temperature increases. The microhardness levels of both FLN2-4405 and FD-0205 are similar.
- Based on this study, optimum tempering temperatures to obtain the highest ultimate tensile strengths while retaining good yield strength, impact energy and hardness levels are between 220°C and 250°C.
- The mass-effect study of the FLN2-4405 specimens shows the microhardness exhibits a linear decrease as the tempering temperature increases. The microhardness also decreases at each 1-mm increment below the surface.
- The 40-mm puck specimens were able to achieve fully martensitic microstructures, while the 100-mm puck specimens have a mixture of bainite and martensite within the core.
- No mass effect can be observed for tempering. The mass effects are primarily seen during heat treatment (quenching).
- The surface of the 100-mm puck specimens tempered at 200°C is 100% martensite. A significant drop in martensite content is observed at 3 mm below the surface. The core of the component is approximately 80% martensite.
- The microstructure of the untempered puck specimens consists of plate-like martensite needles. The martensite contains dispersed spherical carbides typical of tempered martensite at the highest tempering temperature.
- The microstructure arrangements are consistent with the mechanical properties obtained. As the tempering temperature increases, the internal stresses decrease and the strength, apparent hardness, impact energy and microhardness also decrease.
For more information: Contact Amber Tims, technical service engineer, North American Höganäs Co. 111 Hoganas Way, Hollsopple, PA 15935; tel: 814-479-3528; e-mail: firstname.lastname@example.org; web: www.hoganas.com. Both parts of this article came from Proceedings of the 2018 International Conference on Powder Metallurgy & Particulate Materials, June 17-20, 2018; Category 10 Material Properties – Page 755, Paper 122; Metal Powder Industries Federation Princeton, N.J.
- F. Fillari, T. Murphy, I. Gabrielov; “Effect of Case Carburizing on Mechanical Properties and Fatigue Endurance Limits of P/M Steels,” Hoeganaes Corporation, USA, Borg Warner Automotive, USA
- S. Saritas, R.Causton, B. James, A. Lawley; “Effect of Microstructural InHomogeneities on the Fatigue Crack Growth Response of a Prealloyed and Two Hybrid P/M Steels,” Proceedings for PM2002, Hoeganaes Corporation, USA, Gazi University, Turkey, Drexel University, USA 2002
- “Heat Treatment of Plain Carbon and Low-Alloy Steels: Effects on Macroscopic Mechanical Properties,” Massachusetts Institute of Technology Department of Mechanical Engineering, Cambridge, MA 2004
- T. Digges, S. Rosenberg; “Heat Treatment and Properties of Iron and Steel,” U.S. Department of Commerce National Bureau of Standards Monograph 18 1960 p. 10-18
- S. Ropar, R.Warzel III, B. Hu; “Martensitic PM Materials,” North American Höganäs Proceedings for PM2016
- Dr. H. K. Khaira; “Hardenability” Manit, Bhopal; Https://www.slideshare.net/RakeshSingh125/f46b-hardenability, Nov. 2013
- “Höganäs AB Handbook for Metallography,” No. 6, Höganäs AB 2015
- Lindskog, P.; “Controlling the Hardenability of Sintered Steels,” Höganäs AB, Höganäs, Sweden
- MPIF Standard 35 Material Standards for PM Structural Parts. (n.d), MPIF