THE USE OF MACHINING ADDITIVES IN FERROUS PM PARTS TO
IMPROVE MACHINABILITY
Bruce Lindsley and Sunil Patel
Hoeganaes Corporation
Cinnaminson, NJ 08077
Suresh Shah, Gregory (Jerry) Falleur and Chanley Chambers
Cloyes Gear & Products, Inc.
Subiaco, AR 72865
ABSTRACT
Although powder metallurgy (PM) alloys are often more difficult to machine than their wrought
counterparts, PM has the advantage of easily admixed machinability additives. The most successful and
widely used additive, manganese sulfide (MnS), can greatly improve tool life in many applications. A
new machining additive (MA) has been introduced to compliment MnS in powder metallurgy PM steels.
This new additive is chemically inert under production processing conditions and does not produce
discoloration and rusting that has been occasionally found with the MnS additive. The MA additive also
reduces tool wear compared with MnS containing mixes in sinter-hardening applications. In this paper,
an industrial scale trial was conducted to determine the preferred additive in an induction hardened
crankshaft sprocket. Tool wear measurements were made throughout the machining process, surface
corrosion was evaluated and the physical properties were tested for each machining additive condition.
INTRODUCTION
Ferrous PM is generally considered a net or near net shape process. Nevertheless, many pressed and
sintered parts require machining prior to final assembly, as certain features must be manufactured by
secondary operations [1]. These machining operations include transverse holes (drilling), undercuts (hard
turning) and other high precision features required for fit and function. PM parts are generally considered
more difficult to machine than their wrought counterparts. The machinability of PM steels differs from
that of wrought steel due to the presence of porosity and the often heterogeneous microstructure [2]. The
porosity in PM steels makes lubrication more difficult and reduces heat transfer from the cutting surface.
Additionally, porosity produces an interrupted cut with the tool, resulting in micro-impacts and a fatigue
condition at the edge of the tool. To achieve similar strengths as wrought products, PM alloys typically
contain higher levels of carbon. As the apparent hardness of a PM specimen approaches that of a wrought
product, the microindentation hardness must be significantly higher to offset the effect of porosity. The
higher microindentation hardness in combination with the porosity is most responsible for the different
machining response of PM compared to wrought steels.
PM, however, has the advantage of being able to admix materials into ferrous alloys, and both powder
producers and part makers have taken advantage of this ability by incorporating machinability additives
such as MnS into the steel. These free-machining additives improve machinability by assisting in chip
formation, lubrication of the tool face and reduction of crater wear [2]. MnS is used extensively in
wrought free-machining grade steels. During the cutting process, MnS deforms along the shear plane,
reduces tool contact time, and forms a lubricating layer on the tool [3]. Several studies have shown the
benefits of MnS additions to the machinability of PM steels [1,2,4-6].
While manganese sulfide has many beneficial attributes, it has some limitations and potentially negative
effects. The use of MnS may damage the sintering furnace through the production of a sulfur containing
gas [4]. High humidity environments can quickly oxidize the MnS and deteriorate the machinability
enhancing properties [5]. In addition, MnS becomes less effective as alloy content increases [6].
A new machining additive (MA) has been introduced that is chemically inert and reduces the tendency for
rusting. This additive has been found to be particularly effective in sinter-hardened steels, especially
when advanced tool materials, such as multi layered coated carbide and boron nitride, are used [7]. In
this study, crankshaft sprockets were produced with differing amounts and types of machining additives
in an FLN2-4405 alloy. Sintering was performed with two different cooling profiles. Quench media
captured in the pore structure was also considered. The mechanical properties of these mixes were tested
and the susceptibility to rusting was also evaluated. More importantly, the machining behavior of the
FLN2-4405 alloy was investigated with no additive, the addition of MnS, or the addition of MA additive.
The benefits of each additive are discussed under several different processing conditions.
EXPERIMENTAL PROCEDURE
The evaluation of different machining additives was conducted with MPIF-designated alloy FLN2-4405.
Mixes were made with Ancorsteel® 85HP, Inco 123 Ni, graphite and lubricant. The ANCORBOND®
process was used for all mixes to ensure uniform powder flow, weight control, and uniform composition
in the parts. Four conditions were tested: no additive, 0.25% MnS, 0.35% MnS and 0.3% MA. The mix
compositions are listed below in Table I. The mixes were compacted into standard mechanical property
specimens (TRS, dogbone tensile, Charpy impact) at nominally 6.7 and 7.1 g/cm3 and into crankshaft
sprockets measuring 38 mm ID, 50 mm OD and 43 mm high. The density of the sprockets was
approximately 6.7 g/cm3 at the ID where the part is machined and 7.1 g/cm3 in the sprocket teeth. The
samples were then sintered in a 95% nitrogen – 5% hydrogen atmosphere at 1120 °C for 25 minutes at
temperature. Two furnace cooling conditions were used; conventional cooling and accelerated cooling
via convection. The rapidly cooled samples were tempered at 205 °C for 1 hour after sintering. The
radial crush strength and the tooth strength of the sprocket were also tested.
The inner diameters of the sprockets were machined. This section sees little heat (~200 °C) from the
induction hardening process performed on the teeth. However, as the entire part is oil quenched after
hardening, samples were tested in two conditions: (1) as-sintered, and (2) heat to 200 °C and oil quench.
This oil quench operation impregnates the sample with oil. The internal boring trials were run on a
Okuma Crown CNC lathe using an uncoated carbide insert (Sandvik H13A) with a DCMT 3(2.5)1 KF
design. The depth of cut was 0.75 mm (0.03 in.) per pass, the feed was 0.08 mm (0.003 in) per revolution
®
Ancorsteel and ANCORBOND are registered trademarks of Hoeganaes Corporation
and the length of cut was 44.5 mm (1.75 in). The cutting speed was 1.5 surface meters per second (300
sfm) and 0.75 m/s (150 sfm) for the conventional and rapidly cooled samples, respectively. No coolant
was used throughout the machining tests. A keyway is present in the ID of the sprocket, so the cut is
interrupted once per revolution. The uncoated carbide insert was used to obtain accelerated wear values
in the samples. It was found to resist catastrophic failure and best demonstrate progressive abrasive wear.
Tool wear was measured after every third cut on the flank of the tool. A new tool edge was used for each
data point generated.
Table I. Nominal compositions (in wt%) of the alloys studied, balance Fe.
Alloy #
Designation
Ni
Mo
Gr
MnS
MA
Acrawax
Zinc
Stearate
1
FLN2-4405
2
0.8
0.7
-
-
0.4
0.35
2
FLN2-4405
2
0.8
0.7
0.25
-
-
0.75
3
FLN2-4405
2
0.8
0.7
0.35
-
0.4
0.35
4
FLN2-4405
2
0.8
0.7
-
0.3
0.4
0.35
The effect of machining additives on surface rusting of the sprockets was studied under three conditions:
(1) ambient plant conditions (winter) for 40 days with as-sintered samples, (2) as-sintered samples in a
humidity chamber, and (3) oil quenched (OQ) samples in a humidity chamber. The conditions in the
humidity chamber were 43 °C (110°F) and 100% humidity for 4 days.
RESULTS
(a)
(b)
Figure 1. Microstructure of the FLN2-4405 sprocket (Alloy 2) at (a) conventional and (b) accelerated
cooling rates.
Microstructure
The microstructures for conventional and accelerated cooling samples are shown in Figure 1. The
microstructure in Figure 1a is a mix of divorced pearlite, martensite and Ni-rich regions. As expected, the
increased cooling rate results in a greater fraction of martensite in the microstructure. The coarse
divorced pearlite structure is completely eliminated, and replaced with either martensite or fine,
unresolved pearlite / bainite. MnS particles (gray) are also evident in Figure 1b.
Mechanical Properties
The mechanical properties were measured at two densities and two cooling conditions. Green strength
was higher in the Acrawax mixes at both densities (Table II). No effect of machining additive was found
on green strength. The dimensional change (DC), hardness, TRS, tensile and impact properties are given
in Table III. MnS causes growth relative to the no additive alloy (roughly +0.05%), whereas the MA
additive causes shrinkage relative to the no additive condition (-0.03%). This is consistent with earlier
work [7]. It can be seen that the addition of machining additives had little effect on mechanical properties
at this density. To determine if there was an effect at different densities, samples were compacted to a 7.1
g/cm3 and tested in the conventionally cooled condition. Again, no effect of the machinability additives
was found on the mechanical properties of this alloy system.
Table II. Green strength of the four alloys at two densities.
Alloy
Green Strength at 6.7 g/cm3
(MPa)
Green Strength at 7.1 g/cm3
(MPa)
1
2
3
4
6.5
5.5
6.5
6.7
10.7
9.4
10.7
10.9
Table III. Sintered properties of the four alloys at different cooling rates and densities.
standard cooling
Alloy
1
2
3
4
Additive
None
0.25 MnS
0.35 MnS
0.3 MA
GD
(g/cm3)
6.67
6.67
6.67
6.67
SD
(g/cm3)
6.65
6.64
6.64
6.67
accelerated cooling, tempered at 400F/1hr
GD
SD
Alloy
Additive (g/cm3) (g/cm3)
1
None
6.67
6.66
2
0.25 MnS
6.67
6.65
3
0.35 MnS
6.66
6.64
6.67
6.68
4
0.3 MA
DC
(%)
-0.05
0.00
0.01
-0.08
TRS
(MPa)
890
855
841
876
Hard
(HRA)
49
49
49
49
YS
(MPa)
412
384
400
404
UTS
(MPa)
487
450
461
481
EL
(%)
1.1
1.1
1.0
1.1
Impact
(J)
8
7
8
8
DC
(%)
-0.06
-0.01
-0.02
-0.09
TRS
(MPa)
1062
1069
1007
1076
Hard
(HRA)
57
58
56
56
YS
(MPa)
521
556
545
527
UTS
(MPa)
578
605
589
599
EL
(%)
0.9
0.9
0.9
1.0
Impact
(J)
9
8
8
9
DC
(%)
0.04
0.08
0.09
0.01
TRS
(MPa)
1276
1200
1172
1228
Hard
(HRA)
56
56
56
56
YS
(MPa)
518
493
510
506
UTS
(MPa)
641
624
636
650
EL
(%)
1.3
1.4
1.3
1.4
Impact
(J)
14
12
12
12
standard cooling
Alloy
1
2
3
4
Additive
None
0.25 MnS
0.35 MnS
0.3 MA
GD
(g/cm3)
7.10
7.08
7.08
7.08
SD
(g/cm3)
7.09
7.05
7.07
7.08
Sprocket samples were tested for both radial crush and tooth strength. Table IV shows the effect of
machining additives on these properties. Again, little effect of the additives was observed. The sample
containing the MA additive (Alloy 4) performed better than the rest in the tooth strength test.
Table IV. Radial crush and tooth strength of the four alloys at two densities.
Alloy
Crush Strength at 6.7 (MPa)
Tooth Strength at 7.1 (kgf)
1
236
1578
2
232
1523
3
224
1534
4
238
1799
Machining
Boring tests were run on all four alloys, processed with both the conventional cooling and accelerated
cooling. For each alloy and cooling method, testing was performed in the as-sintered and the oil
quenched conditions. Tool wear was measured on the tool flank. Figure 2 shows an example of a new
tool and one during testing of Alloy 1. The tool wear is measured from the top of the tool to the lowest
point on the wear scar.
(a)
(b)
Figure 2. Carbide tool (a) before and (b) after machining. Tool in (b) is after 6 passes of the
conventionally cooled sample with no machining additive.
The tool wear was measured every three passes and plotted for each set of conditions. For the
conventionally cooled samples, no additive Alloy 1 caused the highest tool wear in all conditions (Figure
3). In the as-sintered condition, catastrophic tool failure occurred in less than 12 passes. The presence of
machining additives reduced tool wear and the MnS samples resulted in the lowest amounts of wear. The
0.35% MnS curve exhibited a plateau where the wear rate was quite low, after which, wear rates increase
greatly. This behavior has been observed previously [8]. The oil quench (impregnation) operation
greatly benefited machinability and the combination of oil quenchant with the MnS resulted in very low
wear rates. The 0.35% MnS resulted in the least wear of the uncoated carbide tool in all conditions. The
MA additive reduced wear relative to the no additive sample in all conditions; however, greater amounts
of wear were found relative to the MnS samples. In the oil quench samples, after an initial break-in
period, the wear rate is quite similar for the 0.3% MA and 0.25% MnS samples. Overall, the presence of
oil in the sample reduced the tool wear considerably for all four additive conditions.
(a)
(b)
Figure 3. Tool wear for the conventionally cooled samples. (a) as-sintered and (b) oil quenched.
The rapidly cooled set of samples was also tested for machinability. Tool wear data for the samples after
tempering and after oil quench are plotted in Figure 4. The trends are generally similar to the samples
that were cooled more slowly in the sintering furnace. The sinter-hardened material with no machining
additive resulted much higher flank wear and the tool did not survive more than one pass. The carbide
tools used to machine the sinter-hardened, MnS-containing samples had similar wear through 9 passes but
did not exhibit the rapid wear after 12 passes that was present in the conventionally cooled samples.
When the MA-containing sample was turned, greater scatter was found than in the conventionally cooled
samples. Two tools (6 and 12 passes) had flank wear very similar to the conventionally cooled sample in
Figure 3. The other tools (3 and 9 passes) showed higher wear rates. This high scatter is not ideal in a
production environment, where predictable tool life is desired. The combination of accelerated cooling,
the MA additive and the uncoated carbide tool did not work well in this application. The relatively hard
MA additive requires advanced tooling (coated carbide, boron nitride) for the maximum benefit in
machinability. The tool wear of the samples with oil quenchant had a similar behavior to those of the
conventionally cooled samples, with the wear rates being slightly higher in the rapidly cooled condition.
It should be noted that the cutting speed of the rapidly cooled samples was half that of the conventionally
cooled samples to achieve these similar wear values. All additives significantly reduced tool wear
relative to the no additive condition in the oil quenched condition. The no additive sample showed the
plateau behavior described earlier.
Samples were also tested in the rapidly cooled condition without a temper. Tool wear was generally
double that of the tempered samples, with the 0.35% MnS sample again resulting in the least tool wear.
Untempered martensite is known to cause high tool wear as it is quite hard. The combination of
untempered martensite and an interrupted cut is quite challenging for machining and should be avoided if
at all possible. Tool selection is difficult, as many materials that are hard enough for the untempered
martensite are not tough enough to handle the interrupted cut.
(a)
(b)
Figure 4. Tool wear for the sinter-hardened samples. (a) as-sintered & tempered and (b) oil quenched.
Overall, both the MnS and MA additives provided consistently lower tool wear for all conditions
compared to the samples with no additive. Under the current set of machining conditions, manganese
sulfide was found to be the more effective additive. One should keep in mind that an uncoated carbide
tool was used in this testing to produce wear in an accelerated manner. It is entirely possible that an
accelerated test may skew the results relative to a standard industrial practice. Given that the MA additive
is relatively hard, use of a relatively soft insert (compared to a coated carbide, boron nitride or ceramic
insert) may cause increased wear with the MA containing alloy. Given all of the possible machining
conditions present in the PM industry, it is recommended that both additives be evaluated to determine
the best machining response for each application. This assumes that all other factors outside of
machinability are equivalent.
Surface Corrosion
The surface corrosion was measured in two ways. Sintered samples were left exposed to the plant
environment for 40 days. The test occurred during the winter, and although not measured, the
temperature and humidity were lower than that expected during the summer months. Surface corrosion
should therefore be greater in the summer months. Figure 5 shows the surface condition of the parts. It
can be seen that surface staining is apparent in the sprocket teeth with the MnS containing alloys 2,3. The
staining was slight in the 0.25% MnS alloy, but the amount of staining increased greatly with increasing
MnS content. No surface residue was found on the sprocket teeth samples containing no additive or the
MA additive. It is interesting to note that no staining was found in the hub section of any sprockets
tested. Upon further analysis, it was determined that the surface staining was present in the as-sintered
state prior to the 40 day exposure and is likely an interaction between the MnS and the lubricant burnout.
Given that lubricant burnout is a function of density, it is logical that the surface residue changes with part
density and is only located in the high density section of the part. Surface staining in admixed Nicontaining alloys is not unusual as Ni acts as a catalyst for lubricant sooting. It appears that MnS is also a
catalyst for this reaction. Further holding of samples (90 days total) under the plant environment led to no
change in surface appearance. No corrosion (rust) was found on the part surface.
No Add
0.35% MnS
0.25% MnS
0.3% MA
Figure 5. Surface discoloration in sprocket teeth exposed to a manufacturing environment.
Samples were also placed in a humidity chamber for 4 days. Extensive rusting was found over the entire
surface of the as-sintered samples. Figure 6 shows the samples after testing in high humidity. While
rusting was found on all four samples, the two samples containing MnS showed significantly more
corrosion. These samples are total covered with dark red rust, while the no additive and MA additive
samples are roughly 50% covered with light rust. This result is consistent with previous work where MnS
containing alloys increased the tendency for rust formation in high humidity conditions.
No Add
0.25% MnS
0.35% MnS
0.3% MA
Figure 6. Surface corrosion of the as-sintered samples in a humidity chamber.
Quenching the samples in oil had a significant effect on surface corrosion. The amount of surface rust
was dramatically reduced, especially in the MnS containing mixes (Figure 7). Oil quenched alloys 1-3
looked similar to the samples exposed to the plant environment, with some discoloration on the teeth of
the MnS containing alloys. Alloy 4, which contained the MA additive, showed an increased tendency
toward rusting after the quench. Red rust can be found on the teeth of the MA sample. It is unknown
why oil quenching the sample would increase rusting on the teeth of this sample. The bright orange rust
suggests that liquid water was present during the test.
No Add
0.25% MnS
0.35% MnS
0.3% MA
Figure 7. Surface corrosion of the oil-quenched samples in a humidity chamber.
CONCLUSIONS
Machining is an important secondary operation on many PM parts. While PM parts are generally
considered more difficult to machine than their wrought counterparts, different additives are available to
improve the machining response. An evaluation of two machining additives, MnS and MA, was
performed in a crankshaft sprocket made with alloy FLN2-4405. The following conclusions have been
drawn from this study.
•
•
•
•
•
•
No detrimental effect on mechanical properties was found with the addition of either additive at
relatively low levels (0.25% & 0.35% MnS; 0.3% MA).
During machining, both additives reduced tool wear as compared with the no additive alloy.
Under all of the conditions tested, 0.35% MnS provided the best machinability.
The oil quench / impregnation greatly reduced tool wear and was beneficial for all additive and
cooling conditions.
Accelerated cooling in the sintering furnace increased the amount of martensite in the sprocket,
and when untempered, greatly increased tool wear. Tempered martensite also increased tool wear
in the no additive condition. The addition of machining additives minimized the effect of
increased martensite and the tool wear difference between the as-sintered and the sinter-hardened
and tempered samples was similar.
MnS increases the tendency for surface staining related to lubricant burnout and sooting. Parts
containing MA were very clean.
Under the normal plant environment, no rust was found in samples with MnS or MA. In elevated
humidity conditions (43 °C, 100% humidity), MnS accelerates rusting of as-sintered parts, while
the addition of MA into the alloy had no measurable effect. Oil quenchant reduced surface
rusting considerably.
ACKNOWLEDGEMENTS
The authors would like to thank: Paul Kremus of Hoeganaes Corporation for his assistance in sample
preparation and testing; The Sandvik Corporation for technical assistance in insert selection and
measuring technique; and Dewayne Askins, Gilbert Schluterman, Jr., Edward Elsken, Steve Vogt, &
Austin Maddox of Cloyes Gear and Products for their work in sample preparation, evaluation and data
collection.
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