The United States armed forces regularly look to researchers for development of alloys and manufacturing methods to produce parts for applications requiring a combination of high strength, ductility, hardness, and impact toughness. Of the existing alloys that can meet or exceed these extreme requirements, steels are significantly cheaper than other alloys commonly used in military applications, such as titanium alloys and specialty steels highly alloyed with high-cost elements such as nickel and cobalt [1] . In particular, high-strength lowalloy (HSLA) steels are desired for their high strength-toweight ratio, relatively low cost, good toughness and ductility, fatigue resistance, wear resistance, good ballistic performance, and relative ease of processing. HSLA steels known as nickelchromium-molybdenum steels, which make up the AISI 43xx series, are often considered the industry standard of comparison for HSLA steels [1] .
The United States Air Force's recent development of Eglin steel and AF 9628 has facilitated the current wave of research into next-generation HSLA steels. These alloys are characterized by both high levels of strength and increased levels of ductility and impact toughness over traditional HSLA alloys such as AISI 4340/4330 [2] . These next-generation HSLA steels contain nickel, chromium, and molybdenum to promote the formation of these desirable properties in the resulting steel [1] . Eglin steel also contains a significant amount of tungsten, which readily forms carbides that result in higher hardness, creep strength, and high-temperature wear resistance [3] .
The first of the next-generation HSLA steels, Eglin steel, was developed by Eglin Air Force Base for use in high strain rate applications such as missile components, penetrating ordnance, and armor plating [4] . The second of the nextgeneration HSLA steels, AF 9628, is similar to Eglin steel but contains no tungsten; it is instead alloyed with additional
The United States armed forces regularly look to researchers for development of alloys and manufacturing methods to produce parts for applications requiring a combination of high strength, ductility, hardness, and impact toughness. Of the existing alloys that can meet or exceed these extreme requirements, steels are significantly cheaper than other alloys commonly used in military applications, such as titanium alloys and specialty steels highly alloyed with high-cost elements such as nickel and cobalt [1] . In particular, high-strength lowalloy (HSLA) steels are desired for their high strength-toweight ratio, relatively low cost, good toughness and ductility, fatigue resistance, wear resistance, good ballistic performance, and relative ease of processing. HSLA steels known as nickelchromium-molybdenum steels, which make up the AISI 43xx series, are often considered the industry standard of comparison for HSLA steels [1] .
The United States Air Force's recent development of Eglin steel and AF 9628 has facilitated the current wave of research into next-generation HSLA steels. These alloys are characterized by both high levels of strength and increased levels of ductility and impact toughness over traditional HSLA alloys such as AISI 4340/4330 [2] . These next-generation HSLA steels contain nickel, chromium, and molybdenum to promote the formation of these desirable properties in the resulting steel [1] . Eglin steel also contains a significant amount of tungsten, which readily forms carbides that result in higher hardness, creep strength, and high-temperature wear resistance [3] .
The first of the next-generation HSLA steels, Eglin steel, was developed by Eglin Air Force Base for use in high strain rate applications such as missile components, penetrating ordnance, and armor plating [4] . The second of the nextgeneration HSLA steels, AF 9628, is similar to Eglin steel but contains no tungsten; it is instead alloyed with additional
The United States armed forces regularly look to researchers for development of alloys and manufacturing methods to produce parts for applications requiring a combination of high strength, ductility, hardness, and impact toughness. Of the existing alloys that can meet or exceed these extreme requirements, steels are significantly cheaper than other alloys commonly used in military applications, such as titanium alloys and specialty steels highly alloyed with high-cost elements such as nickel and cobalt [1] . In particular, high-strength lowalloy (HSLA) steels are desired for their high strength-toweight ratio, relatively low cost, good toughness and ductility, fatigue resistance, wear resistance, good ballistic performance, and relative ease of processing. HSLA steels known as nickelchromium-molybdenum steels, which make up the AISI 43xx series, are often considered the industry standard of comparison for HSLA steels [1] .
The United States Air Force's recent development of Eglin steel and AF 9628 has facilitated the current wave of research into next-generation HSLA steels. These alloys are characterized by both high levels of strength and increased levels of ductility and impact toughness over traditional HSLA alloys such as AISI 4340/4330 [2] . These next-generation HSLA steels contain nickel, chromium, and molybdenum to promote the formation of these desirable properties in the resulting steel [1] . Eglin steel also contains a significant amount of tungsten, which readily forms carbides that result in higher hardness, creep strength, and high-temperature wear resistance [3] .
The first of the next-generation HSLA steels, Eglin steel, was developed by Eglin Air Force Base for use in high strain rate applications such as missile components, penetrating ordnance, and armor plating [4] . The second of the nextgeneration HSLA steels, AF 9628, is similar to Eglin steel but contains no tungsten; it is instead alloyed with additional
The United States armed forces regularly look to researchers for development of alloys and manufacturing methods to produce parts for applications requiring a combination of high strength, ductility, hardness, and impact toughness. Of the existing alloys that can meet or exceed these extreme requirements, steels are significantly cheaper than other alloys commonly used in military applications, such as titanium alloys and specialty steels highly alloyed with high-cost elements such as nickel and cobalt [1] . In particular, high-strength lowalloy (HSLA) steels are desired for their high strength-toweight ratio, relatively low cost, good toughness and ductility, fatigue resistance, wear resistance, good ballistic performance, and relative ease of processing. HSLA steels known as nickelchromium-molybdenum steels, which make up the AISI 43xx series, are often considered the industry standard of comparison for HSLA steels [1] .
The United States Air Force's recent development of Eglin steel and AF 9628 has facilitated the current wave of research into next-generation HSLA steels. These alloys are characterized by both high levels of strength and increased levels of ductility and impact toughness over traditional HSLA alloys such as AISI 4340/4330 [2] . These next-generation HSLA steels contain nickel, chromium, and molybdenum to promote the formation of these desirable properties in the resulting steel [1] . Eglin steel also contains a significant amount of tungsten, which readily forms carbides that result in higher hardness, creep strength, and high-temperature wear resistance [3] .
The first of the next-generation HSLA steels, Eglin steel, was developed by Eglin Air Force Base for use in high strain rate applications such as missile components, penetrating ordnance, and armor plating [4] . The second of the nextgeneration HSLA steels, AF 9628, is similar to Eglin steel but contains no tungsten; it is instead alloyed with additional 48th SME North American Manufacturing Research Conference, NAMRC 48 (Cancelled due to manganese, silicon, nickel, and molybdenum [5] . AF 9628 has better mechanical properties and cost approximately half as much to produce as Eglin steel. Table 1 shows the weight percent elemental compositions of AISI 4330 compared to Eglin steel and AF 9628. While the higher strength and ductility combined with a lower cost is of great benefit to the armed forces, machinability of AF 9628 has been shown to be a problem due to rapid strain hardening of the alloy upon turning. The authors' limited previous experiments of soft turning AF 9628 on a manual lathe with no coolant have shown an increase in hardness of up to 11 HRC after one pass of the tool. Strain hardening is especially problematic when multiple passes of the cutting tool are required. Formation of a white layer, or a very hard surface layer formed in ferrous materials that appears white under the microscope [8] , is detrimental to product performance and must be removed with finishing operations, particularly for applications in the aerospace or automotive industries [9] . The hard white layer formed during machining reduces fatigue strength of the finished part and leads to a brittle surface prone to crack permeation leading to failure of the product in service [9] . In addition to the hardening of the material, strain hardening during turning causes the cutting tool inserts to wear out quickly as the feed rate and depth of cut are increased. Excessive amounts of heat are generated in the turning process in the absence of adequate coolant.
Although a similar study has been performed for milling operations for AF 9628 [10] , this is one of the first studies for lathe operations. The bulk of current applications of AF 9628 involve open die forging a basic cylindrical shape, followed by rapid removal of a large amount of material in the unhardened state using the turning process. It is necessary to minimize strain hardening as much as possible during the soft turning process to avoid excessive machining time and tool wear. Future research is planned to study hard turning of AF 9628. In addition to lessening machining time and tool wear, hard turning of the alloy should minimize strain hardening to avoid post-machining heat treatments meant solely to restore mechanical properties to the surface of an otherwise finished part. While many advanced high strength steels exist for current engineering applications, HSLA steels provide an excellent strength level while maintaining good ductility. For example, although transformation induced plasticity (TRIP) steels have higher strength than HSLA steels, and commercial dual phase (DP) steels have higher ductility than HSLA steels, Ozturk et al. have shown that HSLA 340 has a higher postuniform deformation neck extension than either TRIP800 or DP800 steels [11] . This indicates that HSLA steels can maintain significant ductility after forming without the necessity of costly heat treatment to restore properties.
A wide variety of studies exist for machining HSLA steels.
Many of the turning studies use either coated carbide tools or cubic boron nitride (CBN) tools. While hard turning AISI 4340 with a multilayer coated carbide tool, Suresh et al. found that of the turning parameters, the feed rate has the highest influence on both machining and specific cutting forces, the cutting speed has the highest influence on machining power and tool wear, and the feed rate has the highest influence on surface roughness [12] . In another study by Suresh et al. on turning AISI 4340 steel using a coated carbide insert, it was concluded that a combination of low feed rate, low depth of cut, and low machining time with high cutting speed is beneficial for minimizing the machining force [13] . When investigating minimum quantity of lubrication turning (MQL) of AISI 4140, Hadad and Sadeghi found that the tool-chip interface is approximately 300°C lower in wet turning than dry turning, wet turning reduces cutting forces and helps maintain sharpness of the cutting edge, and surface finish improves due to reduction of tool wear [14] . Other interesting methods for processing HSLA steels include using grind-hardening as an environmentally friendly alternative to heat treating [15] and wire electrical discharge machining (EDM), which introduces a recast layer formation [16] . In many of these studies, tool wear is mentioned as a factor of interest. If tool wear is minimized, tool life can be maximized. This saves money during manufacturing by preventing the purchase of new tools, and by minimizing time associated with tool changes. Maximizing tool life usually comes at the cost of lowering material removal rate. Although tool wear was examined during this study, it was not a primary concern for the end user of these results.
In addition to traditional machining processes, cryogenically-assisted manufacturing processes are emerging as environmentally friendly operations that decrease tool wear and produce functionally superior products [17] . Natasha et al. discovered that the coefficient of friction between the chip and the tool during turning of AISI 4340 steel was reduced up to 73% when using a liquid nitrogen cryogenic coolant compared to dry machining [18] . In a similar study, Natasha et al. saw that the use of -196°C cryogenic coolant during turning of AISI 4340 steel with a carbide tool significantly enhanced the chip breakability during the machining process by increasing the hardness and brittleness of the chips [19] . The current research shows promise for the use of cryogenic turning as a state-ofthe-art machining process for HSLA steels.
The goal of the present research study is to find optimal turning parameters, based on the current processes and current tool inserts, to maximize material removal rate with minimal strain hardening of the alloy. Other areas of interest include minimizing chatter, maximizing tool life, and examining surface finish. This research is intended to serve as a baseline study with quantitative results for strain hardening, microstructure evolution, and tool wear based on maximizing the material removal rate possible with current machining practices for AF 9628; it is not intended to use new, modified, or state-of-the-art practices to reduce these results. The following nomenclature will be used throughout the paper when referring to turning parameters: 
A preliminary baseline study was performed using a manual lathe with no coolant to establish minimum reasonable turning parameters for a CNC lathe with coolant. Both processes required machining a slot in each of the workpieces for a tailstock setup in the lathe. The toolholder used was a Kennametal Kenloc Turn Toolholder catalog number MDJNR165D. The tool inserts were Sumitomo Electric series AC820P. A 0° rake angle was used for all experiments.
Four cylindrical bars were machined on a Bridgeport EZ Pass manual lathe limited to 2000 rpm featuring a 3-jaw chuck. A picture of the manual lathe is shown in Figure 1 .
The goal of the manual lathe tests was to maximize material removal rate (MRR). An initial study was performed with parameters shown in Table 2 to understand material response and determine baseline parameters for the rest of the study. Optimal CNC turning parameters based on current turning practices for AF 9628 were the goal of the study, and the manual turning experiments were solely used to gain a baseline understanding of machinability of the alloy. While increasing the DOC, the heat in the workpiece also increased significantly. Although temperature was not directly measured, the heat produced by the machining operation was assessed by the color of the chips. The chips coming off bright blue or purple in color were known to be significantly hotter than the ones coming off golden or silver in color. Coolant was used between setups to reduce the temperature of the workpiece, but not throughout the parameter testing. The tested parameters did not exceed the power limitations of the lathe or tooling, but the point pressure exceeded the strength of the chuck and limited the amount of possible experiments. An example of the chips discolored from heat is shown in Figure  2 .
After the manual turning experiments were complete, parameters for CNC turning experiments were established. Since no coolant was used in the previous experiments, the parameters used that established the maximum successful material removal rate were used as the baseline parameters for the CNC study. Three surface speeds were used: 550, 650, and 750 SFM; two feed rates: 0.015 and 0.020 in/rev; and three depths of cut: 0.050, 0.100, and 0.150 in. Two cylindrical bars of AF 9628, labeled 5 and 6, were machined for the CNC portion of the study, with nine one-inch travel experimental runs on each. Tables 7 and 8 list the experimental matrices for the CNC study. 
In the manual lathe experiments, tool wear was not measured between runs. The tool insert was replaced upon failure. With increased point pressure, rapid decline in tool life was observed. With about 700 passes already on the tool, a large piece of the cutting edge of the tool chipped off, leading to tool failed after six passes with a DOC of 0.150 inches. An example of a failed tool insert compared to a new insert is shown in Figure 4 .
Between each CNC experimental run, the tool insert was removed and measured for flank wear. The tool was placed with the flank of interest under a Nikon SMZ-800 Stereomicroscope at 6.3x magnification. NIS-Elements D 4.00.03 software was used to take a picture and measure any perpendicular wear marks of interest on the flank. An example of a tool wear measurement is shown in Figure 5 .
For measurements of hardness and surface roughness, flats were cut across the length of the cylinders with a band saw, as shown in Figure 6 .
The experimental runs were then separated with cuts across the diameter, resulting in a half-moon shape for each experimental run. Using a Zygo Zegage optical profilometer, five measurements of surface roughness were taken for each experimental run. A 10X Mirau lens was used with the optical profilometer and was calibrated with a standard silicon carbide reference flat. An example of the surface profile of a part is shown in Figure 7 .
Hardness was measured on both the baseline stock material for all cylinders and each experimental run. Rockwell hardness testing was performed according to ASTM E18-19: Standard Test Methods for Rockwell Hardness of Metallic Materials [20] . The Rockwell C scale was used. Calibration across the C scale was checked with 25.8 ± 1.0 HRC and 63.3 ± 0.5 HRC gauge blocks. Five hardness measurements were taken across the length of each experimental run on the highest crosssectional diameter point, as shown in Figure 8 . 
Samples were taken from experimental runs on Cylinders 3, 5, and 6. A small piece from the end of each half-mood shaped experimental run slice was removed and hot mounted in a phenolic powder. The samples were polished following standard metallurgical procedures.
Microhardness was measured with a Clemex automatic microindenter following ASTM E384-17: Standard Test Method for Microindentation Hardness of Materials [21] . Three lines of five indents were taken across the surface of the sample, ending approximately 200 μm from the machined surface. The minimum allowed load of 25 gf with a dwell time of 15 s was used to produce the smallest possible accurately measurable indents. Microhardness was measured on both the baseline stock material for all cylinders and each experimental run. The indentation diagonal size varied from approximately 12-14 μm. Since the indents were all performed and measured manually, the distance to the machined surface was approximately 35 μm (2.5x diagonal spacing minimum per ASTM standard E384-17), and the spacing between indents varied from about 40-45 μm (> 2.5x diagonal minimum per ASTM standard E384-17). The microhardness sample extraction and indentation pattern are shown in Figure 9 .
Grain refinement was observed with an optical microscope after etching with 4% Nital and by electron backscatter diffraction (EBSD) with a scanning electron microscope (SEM).
Chips were collected from experimental runs on Cylinders 3, 5, and 6 and examined for color, morphology, and thicknesses.
The results for each part of data collection are summarized in the following sections. For all results, it is important to note the new reference to a "Run 10A" on Cylinder 5. Run 4 on Cylinder 5 was accidentally machined over before data could be collected; only chips and tool wear data exist for the original Run 4. Run 10A was created with the same parameters as Run 4, and was machined after Run 9 on Cylinder 5. The data collected for Run 10A serves as a replacement for that which would have been collected from Run 4.
The experiments run for this research study were inadequate in both runtime and quantity to plot a full tool life curve. However, the simple study performed showed that the tool inserts developed flank wear marks quickly, and wear increased significantly with more aggressive cutting conditions, as shown in Figure 10 . Future research may be planned to adequately measure and plot tool life based on each of the viable turning conditions for maximization of material removal rate. 
Surface roughness was expected to be high due to the nature of the experiments run. The goal of the study was to maximize MRR in the unhardened state, so surface finish is anticipated to suffer after such extreme roughing passes. The average surface roughness for the available manual lathe experiments is shown in Table 9 . Surface roughness versus material removal rate for the manual lathe experiments is plotted in Figure 11 . For Cylinders 1, 2, and 4, all of the experimental runs produced an average surface roughness that was better than the baseline stock. Many of the leftover pieces from experiments run on Cylinder 3 were unavailable for analysis, but it looks to have followed similar trends as the other three manually turned cylinders.
The average surface roughness for the CNC lathe experiments is shown in Table 10 . Surface roughness versus material removal rate for the CNC lathe experiments is plotted in Figure 12 . All of the experimental runs produced an average surface roughness that was worse than the baseline stock. The cylinders used for the CNC experiments had a lower baseline roughness than those used in the manual lathe experiments; otherwise, the surface finish is roughly the same for both processes. Depending on the desired surface finish level, a finishing pass will be needed regardless of parameters used to maximize MRR.
In an attempt to quantify strain hardening resulting from the turning process, baseline hardness was compared to hardness after machining. Due to both the limited available sample area for data collection and the microscopically thin hardened layer, it was difficult to prove a statistically significant increase in hardness for any of the runs, so conclusions based on observation of the limited dataset are drawn instead. The average hardness for the available manual lathe experiments is shown in Table 11 . Hardness versus material removal rate for the manual lathe experiments is plotted in Figure 13 . The maximum increase in hardness after manual turning was approximately 5 HRC. It is anticipated that harsher turning parameters, especially without coolant, will result in a greater depth and magnitude of strain hardening. The average hardness for the available manual lathe experiments is shown in Table  12 . Hardness versus material removal rate for the CNC lathe experiments is plotted in Figure 14 . The maximum increase in hardness after CNC turning was again approximately 5 HRC. Although harsher turning conditions were used for CNC testing than manual testing, a coolant flood was also used. The amount of strain hardening between manual and CNC experiments appears to be about the same.
The majority of the microhardness data showed the hardest measurement near the surface of the machined samples; however, the sample did not always get softer as depth from the surface increased until the bulk hardness is reached, as is anticipated. This is likely due to the microscopic nature of the hardened layer, which makes it difficult to pinpoint thickness even with microhardness indents. The small, 25 gf load indents were difficult to measure accurately. A nanoindention machine would make for a better measurement of the thickness of the hardened layer. Although the microhardness data is not displayed in this paper, a reference graph is provided in Figure  15 for the reader to gauge the trends of the data and the magnitude of the measurements taken. A better understanding of the strain hardening of these parts was gained through the microstructural evolution observed through metallographic examination and electron backscatter diffraction. The hardened layer for a sample with DOC 0.145 in, FR 0.020 in/rev, and 650 SFM for a total MRR of 22.62 in 3 /min is visible through Figures 16 and 17 .
Additional micrographs showing the strain hardened layer produced by various machining parameters are shown in Figure  18 . All micrographs are shown at 5000X SEM with a 15 μm scale bar. Sample #3234 had a MRR of 0.018 in 3 /min, sample #3235 had a MRR of 4.68 in 3 /min, sample #3238 had a MRR of 21.06 in 3 /min, and sample #3239 had a MRR of 23.4 in 3 /min.
The strain hardened layer grows with increasing material removal rate for manual turning processes run without coolant. The hardened layer of sample mount #3239 is approximately 20 μm thick; microhardness measurements started at a depth of approximately 35 μm from the machined surface. Thus, future research would suggest hardness measurements be taken using a nanoindenter and additional EBSD analysis to confirm strain hardened layer thickness.
The chips produced by a machining process allow researchers to better understand what is happening to the workpiece. In machining processes, 90 percent of heat produced is removed by the chip [26], so chip evolution and morphology elicits information about the heat input into the workpiece. Hard turning of alloy AISI E52100 high carbon steel led Poulachon and Moisan to conclude that hardness of the material and cutting speed are the two main turning parameters that influence chip morphology [22] . The cutting mechanics of hard steels lead to the formation of saw tooth chips, which are further classified as wavy chip, segmental chip, shear localized chip, or discontinuous chip [23] .
There were generally three chip morphologies seen in these experiments: thin continuous chips, thick continuous chips, and thick discontinuous chips. These morphologies are shown in Figure 19 with a 6-inch scale as reference. Thin continuous chips (Figure 19 (a)) were seen at DOC ≤ 0.050 in when SFM ≤ 650. Thick continuous chips (Figure 19 (b) ) were seen at 0.050 in ≤ DOC ≤ 0.100 in when SFM ≤ 550. Thick discontinuous chips (Figure 19 (c)) were seen at 0.050 in ≤ DOC ≤ 0.100 when SFM ≥ 550 or when DOC ≥ 0.100 in. These results make sense when compared to a study by Maruda et al. on chip formation during turning of austenitic stainless steel 316L [24] . Based on the stainless steel study, turning parameters appear to have the following effects on chip morphology: low temperatures in combination with low cutting speeds produce thin continuous chips, low temperatures in combination with high cutting speeds produce thick continuous chips, and high temperatures in combination with high cutting speeds produce thick discontinuous chips [24] . A higher DOC will produce a higher temperature rise as a thicker piece of material is sheared off of the base part.
Chip thicknesses were measured using a digital caliper and a digital pointed micrometer. The data for chip thicknesses is shown in Figures 20 and 21 . Figure 20 shows data for chip depth thickness, which is the thickness of the chip associated with the direction of the DOC. The depth thickness shows a mostly linear correlation (R 2 = 0.9522) with the DOC. The depth thickness was approximately 66 percent of the intended DOC. Figure 21 shows data for chip feed thickness, which is the thickness of the chip associated with the direction of the FR.
The feed thickness shows a fairly linear correlation (R 2 = 0.9083) with the FR. The depth thickness was approximately 2.2 times the FR.
The following bullets summarize the information gathered from the chips.
• Color: under the best speeds and feeds the chips should be golden to golden blue in color. Silver colored chips mean the machining is performed below the optimal temperature. The purple chips run without coolant are too hot and causing excessive tool wear and may lead to unintentionally annealing the part. The silver chips run with coolant could be run with more aggressive cutting parameters. • Morphology: the goal is a tight curl that is thick enough to break on its own, approximately 2-4 inches long. • Thickness: as expected, SFM does not seem to affect the data, but feed rate and depth of cut are important independent variables.
Manual and CNC turning experiments were carried out on AF 9628 cylindrical bars based on current next-generation HSLA machining practices. Quantitative measurements of tool flank wear, surface roughness, hardness, microhardness, and chip thickness were taken. Qualitative observations were made on microstructural evolution and chip color and morphology. For MRR maximized up to 25.551 in 3 /min, the surface roughness was approximately less than 10 μm, but a finishing pass would likely be necessary to reach a desired surface finish level. Hardening of up to 5 HRC after one pass was observed. The thickness of the hardened layer is small enough that its effects are difficult to see with either microhardness or traditional metallographic inspection, but use of EBSD shows significant grain refinement near the machined surface. The chips tended to be either thin continuous chips, thick continuous chips, or thick discontinuous chips with a color ranging mostly from golden to purple. The thickness of the chips was approximately 66 percent of the desired depth of cut and 2.2 times the given feed rate. Future research is planned to turn this alloy in both the soft and hard states with improved cutting conditions, including the likes of different tooling inserts, cooling practices or fluids, and coatings. Upcoming studies for both soft and hard turning should include a comprehensive tool wear analysis upon selection of final turning parameters. Tool inserts made for hard materials, such as CBN, may be better suited for machining AF 9628. Adjustments to the rake angle and turning parameters may be made in the future to further maximize MRR. Ideal parameters will result in chips of a golden blue color in tight spirals, without significant chatter to optimize MRR and tool life.
The views expressed in this paper are those of the authors and do not necessarily reflect the official policy or position of the U.S. Air Force, the U.S. Department of Defense, or the U.S. Government.