Daniel J. Raiser*, Wojciech Z. Misiolek*, Bruce Lindsley**
*Institute for Metal Forming
Lehigh University
Bethlehem, PA, USA
**Hoeganaes Corporation
Cinnaminson, NJ, USA
The effect of post sintering cooling rate on the microstructure of a PM sintered steel was investigated by assigning three different cooling conditions to the same alloy. The various microstructures produced by these different cooling conditions were evaluated and quantitatively analyzed for volume fraction of bainite, martensite, and retained austenite. Machinability tests were also performed on samples from all three cooling conditions to determine how these microstructures affected tool life. This was done using a turning operation and the tool life machinability criteria. It was found that the volume fraction of martensite was the most important factor in determining machinability. Lower martensite contents produced much higher machinability and vice versa. Lower martensite contents also resulted in decreased hardness. Bainite was found to be a much more favorable phase from a machinability standpoint. It was also observed that the trajectory of the chips during the turning operation was an indicator of the condition of the tool edge.
INTRODUCTION
Ferrous PM parts have long been known to provide good hardness and strength properties with near net shape processing at low cost. These qualities have led to the replacement of many parts that were traditionally cast or wrought. Unfortunately, PM parts are thought to be worse in terms of machinability than the cast and wrought materials. Although the amount of secondary machining (such as drilling and hard turning) on PM parts is typically lower than other methods, the tool life for performing machining operations on PM parts can be extremely low. This is due to many factors, including non-optimized cutting conditions for PM, but the largest difference is the presence of porosity in PM parts. This leads to interrupted cutting and a low thermal conductivity in the cutting zone [1]. Porosity in the PM part can also interfere with proper lubrication during cutting. These effects result in a higher cutting temperature that is detrimental to tool life.

Many methods of improving the machinability of PM parts are currently available. Most of these methods involve admixing machining additives into the initial alloy mix. These additives typically increase the machinability by acting as chip breakers during machining which helps to reduce thrust force.
These additives, however, can be costly and if too much is added to the mixture, they can have an adverse effect on the strength and dimensional change of the final part.
The microstructures seen in PM parts are very different from those found in similar cast or wrought parts.
Wrought parts are generally machined in a normalized or annealed condition and heat-treated afterwards, whereas some PM components are partially or fully hardened after sintering. In addition, many PM parts have heterogeneous microstructures, due to elemental alloying additives. Given these different microstructures, the specific effect of the different phases found in a PM microstructure on machinability is not well known. The goal of this research is to investigate the relationship between microstructure and machinability as well as the corresponding effect to the mechanical properties of the material. Through better understanding of this relationship it is shown that a compromise can be made which results in a significant improvement in machinability without a large decrease in mechanical properties.
MATERIAL SELECTION
The material used in this experiment was Ancorsteel
®
SH 737, a water-atomized, pre-alloyed powder,
supplied by Hoeganaes Corporation. The pre-alloyed composition is presented in Table I.
Table I – Composition of Ancorsteel 737 SH powder
Composition (weight %)
C Mn Mo Ni O
<0.01 0.42 1.25 1.40 0.15
Additive copper and carbon were also added to the pre-sintered mixture in the amounts of 2.0 wt % and
0.9 wt % respectively. The properties of this PM steel have shown potential in many applications in which high wear resistance, tensile strength, and apparent hardness are desired. These applications include automotive, home appliance, and industrial applications in parts such as brushings, gearings, and sprockets [2]. All samples contained 0.75 wt% EBS wax lubricant and were pressed to a nominal 7.0 g/cm
3
density. Samples were sintered in a continuous belt furnace at 1120 °C in a 90% nitrogen - 10% hydrogen atmosphere. Consequently the samples were also tempered at 205 °C for one hour.
The study involved three cooling conditions. The first cooling condition (cooling condition #1) is the slowest of the three and involves a post-sintering cooling rate of 0.5 ºC/sec. The middle cooling condition (cooling condition #2) involved a post-sintering cooling rate of 0.6 ºC/sec. The fastest of the three cooling conditions (cooling condition #3) had a post-sintering cooling rate of 1.5 ºC/sec. These cooling rates represent the average change in temperature while the sample was cooling from 650ºC to
315ºC. The post-sintering cooling rate was altered by utilizing a conventive cooling system and by changing the belt speed at which the samples went through the cooling zone in the furnace. These cooling conditions were determined as a result of a prior cooling rate screening study and were shown to display a satisfactory range of microstructures for this experiment.
®
Ancorsteel is a registered trademark of Hoeganaes Corporation

EXPERIMENTAL PROCEDURE
Machinability is a parameter that has many different definitions, and thus, many different types of tests.
Among the most popular test criteria are tool wear, surface finish, power used, and cutting speed. The testing done for this paper used the tool wear criteria. Basically, the more machining operations you are able to perform with one tool, the higher the machinability of the material.
The machining process selected was a turning operation performed using a Lodge & Shipley Powerturn
20hp Engine Lathe. Turning is a popular secondary machining operation due to its ability to generate good part concentricity, dimensional accuracy, and sealing face properties. The samples used in the machinability testing were PM steel cylindrical sleeves measuring 32 mm in height. The outer diameter for the samples was 44.4 mm and the inner diameter was 25.4 mm, giving a sleeve wall thickness of
9.5 mm. The turning operation consisted of rotating the sample to achieve an average cutting speed of
1.5 m/s, and dragging a tool along the outer edge of the sample to remove material and perform a uniform ring thickness reduction. The feed rate of the tool was kept constant at 3 mm/rev. Each machining operation consisted of a 2.54 mm outer diameter reduction which translates to a 1.27 mm ring thickness reduction. After five cuts, the sample was removed and replaced with a new cylinder. At this time, the removed tool was examined under a light optical microscope (LOM) to measure flank wear. The wear progression was measured using the MeasureGraph 1-2-3 computer program that is accurate within 2.0
µm. After the wear progression was measured and recorded, the tool was placed back on the lathe and the next sample was turned down for five cuts. This process, with wear progression being checked after every five cuts with the tool, was repeated until either the tool failed, or 150 cuts were completed. Tool failure could be determined by either visual failure (burning of sample and large increase in sparks during turning), audible failure (loud crack during turning), or wear progression failure (no visible edge left on tool).
Before each new sample was placed on the lathe, the cylinder edge on one of the flat sides of the cylinder was ground down to make it smoother. In this way, the initial contact between the tool and sample during the first cut would be with a rounded edge rather than a sharp one. This was done to prolong tool life so that more data would be available for the harder compositions.
Figure 1 – Schematic of the lathe setup for this experiment. The white piece represents the tool while the gray area represents the workpiece. The rate at which the tool is brought across the workpiece is known as the feed rate (f) and the speed at which the workpiece spins is called the cutting speed (v). The depth cut (d) represents the degree to which the radius of the workpiece is reduced. [3]

The tool used in this experiment was a TX-150 carboloy cemented carbide square insert. The insert size number was 433 and the tool parameters used in the machinability testing were a rake angle of 5 degrees, a clearance angle of 0 degrees, and a cutting edge angle of 15 degrees. A schematic of the lathe setup can be seen in Figure 1.
Samples prepared for microstructural study were taken out of the cross sections of the un-machined parts.
The samples were etched using a 50-50 mix of 2% nitol and 4% picral. Staining was also performed on the samples to distinguish the martensite phase from the retained austenite background. This was done using a 25% sodium bi-sulphide stain. The microstructures produced by the different chemical compositions and cooling conditions were analyzed quantitatively to determine the volume fraction of the various phases. This was done using a point count method. The volume fractions of the various phases were found by averaging the results of 10 fields for each condition at 600x magnification. A 100 pt. square test grid was used for the point count analysis. It is important to note that due to the fineness of the microstructures produced, martensite and retained austenite were counted as one phase. Also, porosity was counted during this analysis, but these counts have been subtracted out. In this way, the volume fractions of the phases add up to 100%. The porosity for all conditions was typically around
12%.
RESULTS
The summarized results of the machinability testing are shown in Table II. The results are displayed as the flank wear after 150 cuts if the tool did not fail, and average number of cuts to failure if the tool failed prior to 150 cuts. In the case of cooling condition #3, multiple tests were performed since each test consumed fewer samples. Therefore, the result listed in the table is an average of three tests, along with the standard deviation. Since the standard deviation is less than five, which is the number of cuts in between flank wear measurements, the average value is considered representative.
Table II - Results of machinability testing for all cooling conditions along with their average apparent hardness values.
Cooling Condition
#1 (0.5 ºC/sec)
Flank Wear after 150 cuts
(µm)
408
Average # of Cuts to
Failure n/a
Average Apparent
Hardness (HRA)
63.9
#2 (0.6 ºC/sec) n/a 130 66.9
#3 (1.5 ºC/sec) n/a 38 std. 4.7 71.7
Figures 2 through 4 also display the results of the machinability testing for the alloy at all three cooling conditions. The graphs show the progression of the flank wear throughout the testing, which was measured as the width of the flank wear band. It should be noted that Figure 4 shows a representative machinability test results since many were done at this particular cooling condition and the test results averaged.

1000
900
800
700
600
500
400
300
200
100
0
0 20 40 60 80
Number of Cuts
100 120 140 160
Figure 2 – Machinability testing for cooling condition #1. The graph shows the progression flank wear.
1000
900
800
700
600
500
400
300
200
100
0
0 20 40 60 80 100 120 140 160
Number of Cuts
Figure 3 – Machinability testing for cooling condition #2. The graph shows the progression flank wear.

1000
900
800
700
600
500
Failure Region
400
300
200
Steady State Wear
Break-in Period
100
0
0 20 40 60 80 100 120 140 160
Number of Cuts
Figure 4 – Machinability testing for cooling condition #3. The graph shows the progression flank wear.
This progression shows three distinct stage with different wear rate. They are labeled on the graph.
The results of the microstructure characterization showing percentage contribution of different phases present, as well as average apparent hardness values, are summarized in Table III.
Table III - Results of point count analysis for all cooling conditions with their respective average apparent hardness values.
Cooling Condition % Bainite % Martensite &
Retained Austenite
Average Apparent
Hardness (HRA)
#3 (1.5 ºC/sec) n/a 100 71.7
Representative microstructures for the studied alloy at all three post sintering cooling conditions can be seen in Figures 5 through 10. The admixed Cu diffuses along pore boundaries and changes the hardenability of the alloy in the near pore regions. The center of the large particles have little Cu and the lower hardenability results in bainitic structures at the slower cooling rates. Copper also lowers the martensite finish temperature, so that more retained austenite is present in the alloy adjacent to the pores.

Figure 5 - Low magnification image of an etched sample at cooling condition #1. The microstructure mostly consists of bainite and martensite mixed w/ retained austenite. The black areas are porosity.
Martensite/
Retained
Austenite
Bainite
Figure 6 - Low magnification image of an etched sample at cooling condition #1. The microstructure mostly consists of bainite and martensite mixed w/ retained austenite. The black areas are porosity.

Figure 7 - Low magnification image of an etched sample at cooling condition #2. The microstructure consists of mostly martensite mixed w/ retained austenite with 21% bainite. The black areas are porosity.
Martensite/
Retained
Austenite
Bainite
Figure 8 - High magnification image of an etched sample at cooling condition #2. The microstructure consists of mostly martensite mixed w/ retained austenite and some bainite. The black areas are porosity.

Figure 9 - Low magnification image of an etched sample at cooling condition #3. The microstructure consists entirely martensite mixed w/ retained austenite. The black areas are porosity.
Martensite/
Retained
Austenite
Figure 10 - High magnification image of an etched sample at cooling condition #3. The microstructure consists entirely martensite mixed w/ retained austenite. The black areas are porosity.

DISCUSSION
For the tests that resulted in tool failure, three distinct stages were found for the flank wear progression.
The first stage, known as the break-in period, showed a fast wear rate that typically took place within the first five machining operations. From there, a steady state wear progression set in which would continue until the flank wear approached 650-800 µm. At this point, the tool wear rate would ramp up again and continue to increase until catastrophic failure occurred. This stage is known as the failure region [3].
The images of the microstructures seen in Figures 5-10 correspond well to the results of the point count analysis listed in Table III. The martensite and retained austenite content definitely appears to be increasing in the micrographs as the post sintering cooling rate is increased. Also, it can be seen that the amount of bainite in the microstructure decreases as the post sintering cooling rate is increased.
When comparing the microstructure data in Table III with the machinability data in Table II, a clear distinction can be seen between in the volume fraction of martensite and the resulting machinability. A larger amount of martensite in the microstructure results in a lower tool life and therefore lower machinability. The relationship is logical since martensite is a much harder phase than bainite.
Therefore, machining through this phase is much more detrimental to the tool and will result in a larger amount of abrasive wear. It is also shown in the results that the effect of the volume fraction of martensite on machinability is very large. When the martensite in the microstructure is reduced by 20%, as it can be seen between cooling condition #3 to cooling condition #2, the machinability more than triples. Four tools would have been needed to perform 130 machining operations (the lifetime of the tool which machined the samples which underwent cooling condition #2) on samples that underwent cooling condition #3. This represents a considerable increase in production cost. A correspondence of the volume fraction of martensite to the apparent hardness is also seen in Tables II and III. Since traditionally, machinability tends to decrease as hardness increases, this trend also provides very useful information.
The microstructure of the sample cooled at 1.5
°
C/s also contained a trace amount of free copper. An example of free copper in the microstructure can be seen in Figure 11. The presence of free copper has been linked to improved machinability [4].
Cu
Figure 11 - High magnification micrograph of composition F with cooling condition #3. The image shows a particle of free copper among the porosity in the as polished sample (arrow).

Perhaps more important to machinability is the presence of retained austenite in this martensitic microstructure. Retained austenite contents of 10% have been found in this alloy system. The austenite in much softer than martensite and should be a much easier phase to machine than martensite. It is expected that if the amount of retained austenite was reduced, either by cryogenic thermal treatments or by a reduction in Cu and / or carbon, the machinability would be even poorer for this high cooling rate condition.
Lastly, a general trend concerning the trajectory of the chips was observed during the turning process.
The general progression of the chip directory throughout the machinability tests is shown in Figure 12.
1
3
2
4
5
Figure 12 - Drawing illustrating the progression of the chip trajectory during the machinability tests. This progression is representative of all tests that resulted in an eventual tool failure.
The trend shown in this figure represents a progression of chip trajectory and associated tool wear that resulted in eventual tool failure. If tool failure did not occur, the progression of the chip trajectory ended up somewhere in the middle. It should be noted that radical change from stage three (chips flying down and in the cutting direction) to stage four (chips flying up and opposite the cutting direction) in the chip trajectory progression was observed to always occur shortly before the tool edge failed in the testing.
This trend was found to be consistent for all mixes in that 5-10 cuts after the change in chip direction, the tool would fail. This radical change in chip trajectory was also heralded by a marking on the top of the tool itself. This “comet tail” marking was also observed for all mixes and signified that the tool edge was about to fail. This marking is shown on a used tool in Figure 13.

Figure 13 - Image of the “comet tail” feature on the tool (arrow). This feature appeared when the chip trajectory changed from down and in the cutting direction, to up and opposite the cutting direction. The appearance of this feature was a consistent indicator that the tool edge was about to fail.
Since both of these features always occurred consistently during a certain stage in the tool life, it is possible that they could be used as a new criteria for determining when a tool needs to be changed. The most likely of these two criteria to be used in any kind of industrial capacity is the chip trajectory since the markings on the tool would not always be able to be observed by an operator or detected by a piece of machinery. The usefulness of tool life criteria would, of course depend on the tolerances of the part since the tool failure criteria used in this experiment (audible change or visible change on surface finish).
However, since the easily observed change in the chip trajectory always occurred close to the end of the tool life by this experiment’s failure criteria, it would be an effective way to get as much use out of a tool before it had to be replaced. In this way, the chip trajectory has the potential to be used to give insight into the condition of the tool edge without having to stop the machining process.
CONCLUSIONS
Based on the performed tests the following conclusions can be formulated:
1.
The wear progression of the cemented carbide tool when turning PM steel has three stages. A high wear rate is experienced in the break-in and failure stages of wear, and a lower wear rate is present during the steady wear condition.
2.
The volume fraction of martensite in a sample is the most important parameter in predicting how well a particular material will machine. As the volume fraction of martensite in the microstructure increases, the machinability severely decreases.
3.
The machinability can be maximized by optimizing the martensite and bainite volume fractions and the associated hardness within the part specifications. A post sintering cooling rate can be

chosen to produce a hardness closest to the minimum since bainite is a much more favorable phase from a machinability standpoint. If this is not feasible, other avenues such as advanced tool materials and machinability additives can be utilized.
4.
The trajectory of the chips during the turning operation was observed to be an indicator of the condition of the tool edge for this set of machining conditions. A radical change in chip trajectory was indication that the tool edge would fail within the next 5-10 cuts.
ACKNOWLEDGEMENTS
The authors would like to thank Hoeganaes Corporation for the support of this research. Partial support of Wojciech Z. Misiolek is provided by the Loewy Family Foundation through the Loewy Professorship at Lehigh University. Additionally, the authors are grateful to Thomas Murphy and Arlan Benscoter for help with metallograpohy analysis and Ed Force for his help with machinability experiments.
REFERENCES
1. A. Davala, R.J. Causton, S.H. Luk, F.G. Hanejko, “Productivity Enhancement
Through Improved Machinability Materials,” SAE Transactions: Journal
of Materials & Manufacturing , Vol. 108, pgs. 280-292. 1999.
ISSN: 0096-736X.
2. T.E. Haberberger, T.J. Cornelio, “Field Experience on a New Sinter-Hardening
Material,” International Conference on Powder Metallurgy & Particulate
Materials , 2000.
3. M.P. Groover, Fundamentals of Modern Manufacturing, Second Edition, 2002,
John Wiley and Sons, Inc., Hoboken, NJ.
4. T.F. Murphy, M.C. Baran, “An Investigation into the Effect of Copper and
Graphite Additions to Sinter-Hardening Steels,” Hoeganaes Technical
Paper.