Mn-Containing Steels for High Performance PM Applications
Bruce Lindsley
Hoeganaes Corporation
Cinnaminson, NJ 08077
Suresh Shah, Gilbert Schluterman and Jerry Falleur
Cloyes Gear & Products, Inc.
Subiaco, AR 72865
Presented at PowderMet 2011
San Francisco, CA May 18-21
ABSTRACT
Alloying with manganese provides good hardenability and cost effectiveness in both wrought and PM
steels. Recently introduced ANCORBOND® FLM alloys combine the benefits of Mn with moderate
levels of Mo to produce lean alternatives to Ni- and Cu-containing hybrid alloy steels. The good
hardenability of these Mn-containing steels makes them interesting candidates for sprocket applications,
where martensite formation in components is necessary to impart good wear resistance. The current work
assesses the heat treatment response of these alloys compared with FLN2-4405 under production
conditions.
INTRODUCTION
Powder metallurgy steels prealloyed with Mo have been widely used in the industry for the last two
decades. Molybdenum is a desirable alloying element for PM steel applications for a number of reasons,
including its relatively minor effect on compressibility, high contribution to hardenability, and an easily
reducible oxide. When Fe-Mo steels, such as Ancorsteel® 85 HP, were initially introduced in 1990 [1],
molybdenum enjoyed a relatively low and stable price. Several grades were developed and adopted by
the PM industry including hybrid alloys containing nickel, such as FLN2-4405. This particular alloy
offers a good combination of compressibility and physical properties, leading to its acceptance as one of
the most popular hybrid-alloy PM steels. It is well suited for induction hardening applications as the
prealloyed Mo base alloy provides the necessary hardenability while the Ni addition provides good
toughness.
Unfortunately, the cost of alloying elements has been quite volatile over the last several years. Figure 1
shows the fluctuation in Ni, Mo and Cu price since 2003. These price swings directly impact the cost of
components made with these traditional alloying elements. When the price of these elements is high,
companies within the supply chain may or may not be able to recuperate these charges, but ultimately, the
cost advantage of PM to the end-user is reduced relative to other manufacturing methods.
®
ANCORBOND and Ancorsteel is a registered trademark of Hoeganaes Corporation.
Figure 1. Price of Mo, Ni and Cu alloying elements since January 2003.
A prolonged period of high priced Ni, Mo and Cu could reduce the total PM market due to parts being
made by alternatively alloyed forgings, stampings, machined parts, etc.
In response to these price fluctuations, both powder and part manufacturers have been striving to reduce
the dependence on these volatile elements. One method to achieve this goal is to reduce the alloy content
to the minimum required to meet the part requirements. Powder grades with lower molybdenum content
have been introduced to provide increased flexibility. Additionally, Ni and Cu contents can easily be
tailored in both binder-treated and regular premixes to assist manufacturers in lowering alloy content.
Removal of alloy content and the corresponding reduction in physical properties, however, may not be
allowed based on the mechanical property requirements of the part. In this case, alternative alloying
elements become necessary.
Manganese is a low cost alloying element that is widely used in the wrought industry and imparts both an
increase in strength and an improvement in hardenability. Historically, Mn has not been used extensively
in PM steels for several reasons. It reduces compressibility when prealloyed in the base iron and it is
easily oxidizable. This tendency to oxidize makes atomization of prealloys difficult and limits the
permissible sintering atmospheres to low partial pressures of oxygen. Figure 2 shows the dewpoints
necessary to reduce manganese oxide at different temperatures [2]. At 1120 C, a dew point of -50 C is
required to reduce manganese oxide. While modern sintering furnaces and atmospheres utilizing nitrogen
and hydrogen are able to produce very low dewpoints (-40) in the hot zone of the furnace, Figure 2
suggests that Mn cannot be used with conventional sintering temperatures. Nevertheless, oxidation of
alloys containing Mn can be avoided. Dilution of Mn with other elements reduces the activity of Mn
relative to its pure state and effectively raises the allowable dewpoint. This permits Mn-containing steels
to be sintered at conventional temperatures in nitrogen-hydrogen atmospheres.
The oxygen content of the furnace atmosphere is not the only consideration as oxygen within the compact
is an important variable. Oxygen is present in the gas contained within the porosity, and is also present
on the surface of the metal. Oxygen contents of approximately 0.1 wt% are typical in water atomized iron
VO
(torr)
MoO2
100
10-1
10-2
2O
MnO
Cr2O3
101
PH
Dew point (°C)
FeO
120
0
-20
-40
-60
-80
TiO
0
500
1000
1500
Temperature, °C
10-4
2000
Figure 2. Metal / oxide equilibria in hydrogen [2].
powder. Danninger et al have shown that much of the iron oxide is reduced and removed at temperatures
of 700 C and between 950 C and 1100 C [3]. If an additive is used to introduce manganese, oxygen
should be removed prior to significant Mn diffusion in the part. Researchers have described Mn diffusion
within the compact occurring primarily as a sublimation process where Mn quickly redistributes through
the pore network, followed by deposition of Mn onto the iron particle surfaces and solid state diffusion
into the iron [4]. In this model, the presence of oxygen will inhibit manganese diffusion by forming
MnO. Manganese diffusion from the additive should be limited therefore to temperatures higher than that
required for iron oxide reduction. It was reported that use of pure manganese powder as an additive did
not perform well [5]. The Mn was found to begin diffusing at temperatures below that described above to
reduce iron oxide. Significant MnO formation and limited Mn diffusion into the iron was observed.
Careful selection of the Mn-containing additive must be made if this approach is used. Several additives
have been evaluated with some success, including medium and high carbon ferromanganese [6-9], FeMn-Si master alloys [10], and a specially designed Fe-Mn-C alloy [11].
A newly developed Mn-containing alloy system, ANCORBOND FLM, introduced in 2010, combines the
benefits of manganese and moderate amounts of molybdenum to produce cost-effective alternatives to
high-alloy hybrid and diffusion-alloyed PM steels. This alloy system has been designed for conventional
sintering temperatures and is suitable for as-sintered and sinter-hardened applications. This Cu-free and
Ni-free alloy reduces price volatility and minimizes European recyclability and environmental concerns.
Prior work [12] has shown this alloy system is a good candidate to replace FLN2-4405 with respect to
mechanical properties of sintered test bars. The objective of the current work is to demonstrate the
performance of Mn-containing alloys in a test part compacted and sintered under production conditions.
Additionally, as the chosen part is induction hardened, the response of Mn steels to induction hardening
heat treatment has been evaluated. Mechanical property data and metallography developed from the
processing of a Gen IV crank sprocket are presented in the as-sintered, sinter-hardened, and induction
hardened conditions at densities of 7.0 g/cm3.
EXPERIMENTAL PROCEDURE
Binder-treated premixes of FLN2-4405 and three manganese-containing alloys were used in this study,
Table I. All alloys contained 0.75 wt.% zinc stearate lubricant, 0.25% MnS and Asbury 3203H graphite.
Graphite was limited to 0.6% in mix 4 based on prior results with the higher Mo alloy [12]. Inco 123
nickel powder was used as the Ni source.
Table I. Nominal compositions of the alloys. All mixes contain 0.75 wt.% zinc stearate as a lubricant.
Mix #
Alloy
Fe
wt.%
Mo
wt.%
Ni
wt.%
Mn*
wt.%
Gr
wt.%
1
FLN2-4405
Bal.
0.8
2.0
0.1
0.7
2
0.3Mo1.3Mn
Bal.
0.3
-
1.3
0.7
3
FLM-4005
Bal.
0.5
-
1.3
0.7
4
FLM-4405
Bal.
0.8
-
1.3
0.6
*Mn content does not include that added as MnS
Test specimens were compacted to a density of 7.0 g/cm3 at room temperature. The crank sprockets were
also compacted from the four alloys measuring 38 mm ID, 50 mm OD and 43 mm high. The overall
density of the sprockets was approximately 6.7 g/cm3 and that of the sprocket teeth was 7.1 g/cm3. No
changes were made in the compaction process between the different alloys; all were run with the current
production conditions. The compacts were divided into three thermal processing conditions. All parts
and samples were sintered in a production belt furnace at Cloyes Gear & Products, Inc. for 25 min at
temperature at 1120 °C (2050 °F) in 95N2-5H2 (vol.%). A convective cooling system was used with a
second group to accelerate the cooling rate and partially sinter-harden the compacts. Finally, as-sintered
sprockets were induction hardened by heating for 3 sec on a 60 kW, 450 kHz induction machine and
quenching in oil at 60 °C (140 °F). Heat-treated samples were tempered at 205 °C (400 °F) for 1 h.
Percent dimensional change (from die size), sintered density, and apparent hardness were measured from
the transverse rupture samples using standard MPIF procedures. Tensile testing was performed on dogbone samples using an extensometer, which was left on until failure. Total elongation is reported. Impact
testing was conducted at room temperature on unnotched Charpy samples. The radial crush strength and
the tooth strength of the sprocket were also tested. Figure 3 illustrates how tooth strength and radial crush
strength testing was performed. Tooth strength 1 and radial crush were tested on all parts, whereas tooth
strength 2 testing, in which a pin with a diameter larger than that of the sprocket root is forced between
two teeth until tooth failure, was only performed on hardened teeth. All results presented are averages of
multiple tests.
Samples for metallographic examination were cross-sectioned, mounted in a thermosetting epoxy, ground
and polished using well-established practices. Micro-indentation hardness measurements were made per
ASTM E384-10 on the surface after lightly etching with a 50-50 mixture of 2 vol.% nital – 4 wt.% picral.
The samples were then impregnated with epoxy to ensure accurate representation of the porosity and to
seal off the porosity, thereby avoiding entrapment of polishing compound, water and etchants that can
later lead to staining of the microstructure. The samples were lightly ground, re-polished and etched in a
50-50 mixture of 2% nital – 4% picral to reveal the microstructure. Images were captured on a Leica
MEF4M inverted stage metallograph. The images of induction-hardened teeth in Figure 9 are a montage
of two photographs.
Tooth Strength 1
Tooth Strength 2
Radial Crush Strength
Figure 3. Schematics of tooth testing and radial crush testing.
RESULTS AND DISCUSSION
Figure 4 is a compressibility plot generated from cold compacted samples for the binder-treated premixes.
Mix 1 (FLN2-4405) exhibits excellent compressibility, similar to that found with traditionally alloyed PM
steels (excluding prealloyed Ni grades). The three Mn-containing alloys have lower compressibility due
to the addition of Mn. The compressibility decreases 0.05 to 0.07 g/cm3 compared with FLN2-4405 at
550 MPa. Compressibility loss is one of the major drawbacks of alloying with Mn. Additionally, the Mn
alloys were found to grow upon sintering in a similar fashion to Cu steels. At a density of 7.0 g/cm3, the
dimensional change after sintering was approximately + 0.31% for the Mn alloys versus - 0.03% for the
FLN2-4405 (Table II). Taken together, the Mn alloys require a higher compaction pressure than
traditionally alloyed PM steels to achieve the same sintered density.
In the as-sintered, conventionally cooled condition, alloy 2 (0.3%Mo-1.3%Mn) underperforms
FLN2-4405. While the yield strength and apparent hardness are similar, TRS, UTS and impact are lower
in alloy 2. With the increase in Mo content in alloys 3 and 4, the mechanical properties surpass that of
FLN2-4405. Strength and hardness are higher while the total elongation and impact results are nominally
equivalent. As the hub of the sprocket is used in the as-sintered condition, the 0.5% and 0.8% Mo alloys
impart improved properties relative to FLN2-4405.
Figure 4. Compressibility of alloys studied
ID
Table II. Summary of as-sintered mechanical properties.
SD
DC
TRS
YS
Hard
UTS
3
g/cm
%
MPa
HRA
MPa
MPa
(103 psi)
(103 psi) (103 psi)
Elong
%
Impact
J
(ft lbf)
FLN2-4405
7.01
- 0.03
1033
(150)
54
423
(61.4)
524
(75.9)
1.1
11
(8)
0.3Mo1.3Mn
7.04
+ 0.32
930
(135)
53
419
(60.8)
507
(73.5)
1.0
9
(6)
FLM-4005
7.03
+ 0.31
1116
(162)
58
461
(66.9)
534
(77.4)
0.9
10
(7)
FLM-4405
7.03
+ 0.30
1186
(172)
59
493
(71.4)
603
(87.5)
1.0
12
(9)
The mechanical properties of all alloys increased with accelerated cooling in the sintering furnace Table
III). The dimensional change increased 0.03% to 0.05% for all four alloys, indicating an increase in the
amount of martensite present in the test samples. This is due to the lower density of the martensite phase
relative to a ferrite / carbide microstructure. The faster cooling rate increased hardness and raised yield
strength approximately 100 MPa for alloys 1 and 2 and more than 170 MPa for alloys 3 and 4. Impact
energy did not substantially change with accelerated cooling, and interestingly, the Ni-containing
FLN2-4405 did not show a benefit in impact toughness. Toughness of the sprocket is an important
consideration and the results on the industrially sintered test bars were quite encouraging.
Table III. Summary of mechanical properties for the convectively cooled test specimens.
SD
DC
TRS
YS
Impact
Hard
UTS
Elong
g/cm3
%
MPa
HRA
MPa
MPa
%
J
ID
(103 psi)
(103 psi) (103 psi)
(ft lbf)
FLN2-4405
6.96
+ 0.02
1254
(182)
61
519
(75.2)
615
(89.2)
1.0
9
(7)
0.3Mo1.3Mn
6.99
+ 0.34
1051
(153)
59
525
(76.2)
599
(86.9)
0.9
11
(8)
FLM-4005
6.98
+ 0.35
1316
(191)
64
635
(92.1)
688
(99.8)
0.9
11
(8)
FLM-4405
6.98
+ 0.35
1329
(193)
66
692
(100.4)
763
(110.6)
0.9
11
(8)
Figure 5. Comparison of tensile strength and apparent hardness with conventional cooling and
accelerated cooling. Alloys 1-4 are FLN2-4405, 0.3Mo1.3Mn, FLM-4005 and FLM-4405, respectively.
Crankshaft Sprocket Analysis and Metallography
Microstructure
The Gen IV crank sprocket test pieces were compacted in a 200 ton press using a current production tool
for a crankshaft sprocket used in a high volume V-8 engine. The sprocket has been in production since
1997 and achieves the necessary wear characteristics in the root and flank area of the sprocket through
induction hardening of this region. Because the part and process are mature, it is a good test piece for use
in a comparative study of materials. An example of the sprocket can be seen in Figure 6.
Figure 6. Example of as-sintered test sprocket.
Metallographic sections taken in the hub region of the sprocket revealed an acceptable microstructure for
all four materials. In the as-sintered condition with conventional cooling, the microstructure consists of a
ferrite/carbide microstructure (upper bainite or divorced pearlite) consistent with steels alloyed with
molybdenum, Figure 7. In the FLN2-4405, light etching Ni-rich regions are also present with a small
amount of a martensite (tan colored). The Mn-containing steels have a similar structure with the omission
of the Ni-rich regions. The amount of martensite increases with increasing Mo content, with
approximately 25% martensite in the FLM-4405 alloy. With the use of accelerated cooling, the amount
of martensite in the parts increases substantially, Figure 8. While none of the alloys are fully sinterhardened, significant martensite contents are present in alloys FLM-4005 and FLM-4405. The martensite
content is highest in the FLM-4405, as the increased Mo content overcomes the reduced graphite addition
(0.6%) to boost hardenability relative to the FLM-4005. Less martensite is observed in FLN2-4405 and
the 0.3Mo1.3Mn alloy, as the hardenability of these alloys is reduced. The amount of pearlite / bainite in
alloys 1 and 2 is quite similar. The microstructure observed in the hub is consistent with the mechanical
properties measured in the test bars, where the hardness and strength of alloys 1 and 2 are comparable and
alloys 3 and 4 have higher values.
The case depth and microstructure of the induction hardened teeth is illustrated in Figure 9. No difference
was observed in the ability to harden the sprocket teeth. All four alloys appeared similar, with the
exception of the Ni-regions in the FLN2-4405. The teeth were fully martensitic, and given the moderate
carbon content, no retained austenite was observed. The root section of the sprocket also requires
hardening, and the microstructure of the root is shown in Figure 10. Again, a martensitic case is present
in all four alloys. In the leanest alloy system (0.3Mo1.3Mn), some bainite is present in the case near the
case-core interface, indicating the Mo content may be too low to fully harden in the tooth root for this
application. The sprocket teeth were also evaluated in the as-polished condition. The density of the
FLN2-4405 sample was notably higher in the teeth. This corresponds with the measured densities of the
teeth shown in the following section. The effect of Mn on oxide formation was also a concern with these
materials. No internal oxidation was observed in any of the alloys, and the oxide layer on the surface was
similar between the FLN2-4405 and the Mn-containing alloys, Figure 11. No negative effects of Mn
were noted metallographically in the sprocket teeth.
FLN2-4405
0.3Mo1.3Mn
FLM-4005
FLM-4405
Figure 7. Microstructures of the sprocket hub section for the 4 alloys sintered with conventional cooling.
Samples were etched with 2% picral – 4% nital.
FLN2-4405
0.3Mo1.3Mn
FLM-4005
FLM-4405
Figure 8. Microstructures of the sprocket hub section for the 4 alloys sintered with accelerated cooling
and tempered. Samples were etched with 2% picral – 4% nital.
FLN2-4405
0.3Mo1.3Mn
FLM-4005
FLM-4405
Figure 9. Microstructures of four alloys in the induction hardened teeth. Two photos have been merged
to form the final image. Samples were etched with 2% picral – 4% nital.
A
B
C
D
Figure 10. Induction-hardened tooth root microstructure. A) FLN2-4405, B) 0.3Mo1.3Mn, C) FLM4005, D) FLM-4405 Samples were etched with 2% picral – 4% nital.
a
b
Figure 11. Oxide thickness on the surface of induction-hardened teeth in (a) FLN2-4405 and (b) FLM4005. As-polished
Table IV. Green and sintered density measured in the sprocket.
Green Density (g/cm3)
Alloy
Tooth
Hub
Overall
FLN2-4405
7.21
6.68
6.73
0.3Mo1.3Mn
6.98
6.71
6.73
FLM-4005
7.06
6.69
6.73
FLM-4405
7.06
6.71
6.75
Sintered Density (g/cm3)
Tooth
Hub
Overall
7.07
6.7
6.75
6.9
6.66
6.66
6.96
6.61
6.66
6.85
6.69
6.68
Sectional densities were measured in both the green and sintered condition for the hub region and the
teeth. The average density of the entire part is given in Table IV as ‘Overall’. While there exists some
scatter in the data, trends are present in the part density. The tooth density is higher than that of the hub
and the density of the FLN2-4405 is greater than that of the Mn-containing alloys, especially in the
sprocket teeth. These results are consistent with that observed using metallography. The FLN2-4405
density changes little from green to sintered, as is predicted from the test bars. The drop in sintered
density of the Mn alloys from the green state is also consistent with the test bars. The lower sintered
density in the Mn-containing parts would be expected to have a negative effect on mechanical properties.
The apparent hardness measured in the hub and tooth section of the sprockets is given in Table V. The
hardness values are comparable with all four alloys. The 0.3Mo1.3Mn alloy exhibits the lowest hardness
in both the as-sintered and the convectively cooled condition. The FLM-4005 is quite similar to the
FLN2-4405, while the lower carbon FLM-4405 has higher hardness owing to higher hardenability and
increased martensite content. The apparent hardness of the induction hardened teeth is a function of
density and carbon content only, as the case is fully martensitic. The higher density of the FLN2-4405
results in a higher hardness. Sprockets made with alloys 0.3Mo1.3Mn and FLM-4005 have the same
hardness (same density and same carbon content), while alloy FLM-4405 with 0.53% sintered carbon
results in a lower hardness.
Table V. As-sintered carbon and the apparent hardness measured on the tooth face and at the top of the
hub section of the as-sintered, convectively cooled and induction heat-treated parts.
As-Sintered
Convectively Cooled
Induction
Alloy
Sintered C
Tooth
Hub
Tooth
Hub
Tooth
(wt%)
(HRA)
(HRA)
(HRA)
(HRA)
(HRA)
FLN2-4405
0.65
54
48
57
57
72
0.3Mo1.3Mn
0.63
51
47
57
55
69
FLM-4005
0.64
52
48
58
57
69
FLM-4405
0.53
53
51
60
59
68
The hardness of the martensite in the case hardened teeth is a function of carbon content only; no retained
austenite outside of the Ni-rich regions was observed. Table VI demonstrates that all mixes with 0.7
wt.% graphite had a similar micro-indentation hardness value (715 to 738 HV100 gf) regardless of alloy
content. The difference between 715 and 738 HV100 is approximately one HRC. The lower hardness
nickel-rich regions were avoided during the measurement of the FLN2-4405. Lower micro-indentation
hardness values were obtained with FLM-4405, which contained a lower sintered carbon content. As the
teeth are fully martensitic, micro-indentation hardness is a function of carbon content. The other alloying
elements only serve to improve hardenability. With this in mind, the leanest alloy that has sufficient
hardenability to form a martensitic case under these conditions can be used.
Table VI. Microindentation hardness values for the various mixes, measured in HV100 gf. Converted
values in HRC are given in parentheses.
Processing
FLN2-4405
0.3Mo1.3Mn
FLM-4005
FLM-4405
Induction
Hardening
738
(62)
715
(61)
715
(61)
655
(58)
The mechanical properties of the sprockets were measured by crush strength and tooth strength testing.
The crush strength of the FLN2-4405 sprockets is higher to those made with the Mn alloys. Nevertheless,
the crush strength of the as-sintered sprockets made with 0.3Mo1.3Mn and FLM-4005 is within 2% that
of the FLN2-4405 sprockets. The drop in strength is likely due to the lower density of the Mn-containing
parts. A larger difference in strength was found in parts made with accelerated cooling, and it is thought
that the admixed Ni adds toughness to the hardened components. Higher crush strength was found in the
lower carbon, sinter-hardened FLM-4405 alloy relative to the other Mn-containing alloys, as lower
carbon contents also increase toughness in martensitic materials.
The tooth strength of sprockets made with the FLN2-4405 alloy is higher than that of all three Mncontaining alloys for all conditions tested. The higher tooth density produced in the FLN2-4405 parts
plays a significant role in these results. Within the Mn steels, the FLM-4005 alloy provided the best tooth
strength results for all conditions. The tooth strength values are all within 5% of that measured for the
higher density FLN2-4405. The other two Mn alloys exhibited a more significant reduction in tooth
strength.
Table VII. Crush and tooth strengths measured on the as-sintered, convectively cooled (accelerated) and
induction heat-treated parts.
Tooth
Crush Strength
Tooth Strength 1
Strength 2
Alloy
As-sintered Accelerated As-sintered Accelerated
Induction
Induction
(kgf)
(kgf)
(kgf)
(kgf)
(kgf)
(kgf)
FLN2-4405
4203
4035
1168
1376
1297
1818
0.3Mo1.3Mn
4151
3496
993
1120
1093
1588
FLM-4005
4156
3668
1124
1312
1230
1788
FLM-4405
4037
3988
1008
1235
1181
1537
In the induction-hardened sprockets, all alloys were found to harden adequately during heat treatment.
The sensitivity of the highest Mo-Mn steel (alloy 4) to carbon content as observed in reference 12 limited
the carbon content and therefore the microindentation hardness that was achieved in the sprocket tooth.
The appearance of a small amount of bainite in the tooth root of alloy 2 suggests that 0.3% Mo may be
too low for this application. In combination with the mechanical properties obtained on the sprocket
teeth, Alloy 3 appears to be the most promising alloy to replace the current Mo-Ni alloy. Assuming a
similar density can be obtained in the sprocket teeth, wear is also expected to meet part requirements.
These results indicate that the economically favorable FLM-4005 can be used as a replacement for FLN24405.
CONCLUSIONS
A comparative analysis of the recently developed ANCORBOND FLM Mn-containing steels with the
current FLN2-4405 alloy was performed. An induction-hardened crank sprocket was used evaluate the
different alloys. The Mn-containing alloys have been considered due to their cost-effectiveness and
reduced price volatility relative to the current material. Testing of commercially sintered test bars
revealed the Mn alloys exhibited properties comparable with FLN2-4405. While the test results found
that the 0.3% Mo-1.3% Mn alloy was slightly inferior, the 0.5% and 0.8% Mo versions exceeded the
target properties of the Mo-Ni steel. Sinter hardening was effective in these alloys, and although the
structures were not fully hardened, substantial improvements in strength and hardness were observed.
The desired microstructure was obtained in the induction-hardened sprockets utilizing the Mn-containing
alloys. Martenistic cases were produced and no evidence of preferential oxidation was observed after
hardening. Mechanical testing of the sprockets revealed that the higher density FLN2-4405 teeth
performed better than the Mn-containing steels. Accommodation of the lower compressibility, higher
growth Mn-containing powder would have to be made to separate the role of alloy and density on the
teeth strength and crush strength of the sprocket. Both density and the presence of Ni would be expected
to boost performance in fully hardened sprockets. Alloy 3 (ANCORBOND FLM-4005) is a potential
cost-effective candidate to replace FLN2-4405 in this application.
ACKNOWLEDGMENTS
The authors thank Barry Diamond of Hoeganaes for the extensive metallographic work performed within
the scope of this program.
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