Steel-Based Infiltration – A Method to Achieve Full Density HigherPerformance Powder Metal Parts
F. J. Semel and K. S. Narasimhan
Hoeganaes Corporation, Cinnaminson, NJ 08077
ABSTRACT
This paper presents the current state of the art of the iron base infiltration process. The several papers
detailing the efforts made in the last few years to develop the process and assess its mechanical property
potential and dimensional change characteristics are reviewed and presented in summary form.
INTRODUCTION
Recent efforts to develop iron base infiltration, as a viable parts making process were first reported in two
papers in 2004 [1, 2]. A survey of the open literature, at the time, indicated that past efforts to develop the
process predated the present ones by about twenty years or more. The survey disclosed six articles on the
subject [3 - 8]. All were published before 1985 and were apparently the output of independent research in
three laboratories. They reported studies on three different alloy systems. However, none of the associated
findings were sufficiently compelling to justify continued efforts and subsequent interest in the process
waned. A common difficulty at that time, which may have been central to this outcome, was the general
inability to prevent the adverse effects of admixed alloy segregation during processing after mixing. Thus,
it may well be that the key element underlying the relatively greater success of the present efforts is the
intervening development of the binder treatment process [9, 10, 11]. Certainly, there is absolutely no
doubt in the case of the present process that it would not have even survived the preliminary assessment
stage had it not been for binder treatment.
Subsequent to the 2004 papers, which detailed the metallurgical fundamentals of the process, studies were
reported in each of the following two years that provided further information as to the properties to be
expected of the infiltrated parts. Alloy effects on tensile and transverse rupture properties were reported in
2005, and the effects of thermal processing on dimensional change and infiltrated surface appearance
were reported in 2006 [12, 13]. Minus a few results that have patent potential, these four papers represent
the present state of the art of the process. However, they still remain to be reviewed and consolidated as
such. Thus, the objective of this paper is to indicate the present state of the art of the process by
summarizing and, in certain instances, updating earlier findings in the form of a review.
Development of the Basic Infiltration Process
The focus of the earliest work [1, 2], as presented in 2004, was to establish a basic understanding of the
process and to define the conditions needed to implement it as a practical matter. It was shown that
infiltration to near theoretical densities was possible in several Fe-C based alloy systems. The necessary
processing conditions generally included temperatures below 1200 oC (2190 oF), times of less than ½ hour
and the use of standard hydrogen-nitrogen atmospheres with modest methane additions to control the
carbon potential (typically less than 0.5 v/o). The required infiltrant composition was at or near the
corresponding eutectic liquidus value of the selected alloy system and the required base compact
composition was likewise at or near the eutectic solidus value. The possibility to combine an infiltrant
composition of one alloy system with a base compact composition of another system was also
demonstrated. In addition, it was shown that base compact densities of 6.8 g/cm3 or less were sufficient to
obtain a virtually pore free density after infiltration. The optimum base compact density in terms of
maximizing the infiltration rate and minimizing the infiltrant weight to full density was determined to be
about 6.7 g/cm3.
The as-infiltrated carbon contents of the alloy systems that were studied typically ranged from about 2 %
to about 2.35 %. In an otherwise un-alloyed Fe-C composition, carbon contents in this range normally
result in an as-infiltrated microstructure which consists of pearlite in a network of hyper-eutectoid grain
boundary carbides. Such microstructures are inherently brittle and, unless submitted to a so-called
‘malleablizing’ anneal have limited potential for structural applications [14]. Consequently, the possibility
to use alloy additions to graphitize the hyper-eutectoid carbon and produce cast-iron like microstructures
was investigated. It was found that modest additions of either silicon or nickel were effective in this
regard. However, based on what was generally known of their alloying effects as well as the results of a
series of trials with pre-alloyed nickel compositions, it was concluded that the silicon was the better
choice for future development. Consequently, subsequent work was directed towards obtaining a
sufficient understanding of its effects on both the infiltration and graphitization processes to facilitate the
design of silicon containing infiltrant and base compact compositions. The ultimate objective was to
normalize the development of the technology in terms of one such alloy in each case as standard iron
base infiltration compositions.
Standard Compositions
As it turned out, the defining studies indicated silicon contents that nominally averaged 0.18 % for the
infiltrant composition and 0.75 % for the base compact composition. Based on the eutectic equilibrium of
the Fe-C-Si system as indicated by the ThermoCalc program [15], the corresponding carbon contents
were 4.28 % for the infiltrant and 1.91 % for the base compact. The infiltrant weight to full density in this
alloy system is about 15% of the base compact weight. Thus, based on the infiltrant and base compact
compositions in each case, its easily shown that the respective carbon and silicon contents to be expected
after infiltration are about 2.27 % and 0.66 %. Likewise, assuming that all of the hyper-eutectoid carbon
is graphitized during the process and that the balance forms pearlite, it can also be shown that the final
density to be expected is about 7.53 g/cm3. The attendant studies of the systems reasonably confirmed
each of these expectations.
Distortion Effect
The distortion effect, which may be unique to iron base infiltration, is a result of density gradations in the
infiltrated compact that are manifest as a disparity in the lateral dimensions of the infiltrated and opposing
un-infiltrated surfaces. The greatest variations always occur immediately under the infiltrated surface to a
depth of a few millimeters but may occur elsewhere as well. As a consequence, the magnitude of the
effect is measured simply as the difference in the lengths of the infiltrated and opposing un-infiltrated
surfaces. Typically, the effect is large enough that if not otherwise mitigated, the resultant parts will
require a machining step before they can be put into service in all but the least demanding applications.
Preliminary study of the effect indicated that there were two different causes: a primary one; and an,
intermittently occurring, secondary one. The primary cause was determined to be liquid penetration and
dissolution during infiltration of the sinter bonds existing in and just below the surface of the base
compact followed by lateral expansion of the affected area under the influence of the surface tension
forces that act on the as yet un-infiltrated liquid. The secondary cause is incomplete graphitization of the
hypereutectoid carbon, which is typically limited to the lower regions of the infiltrated compact.
Theoretical considerations suggested several possibilities to mitigate the primary cause of the effect.
However, the simplest way found to prevent it is to use an infiltrant compact with a smaller cross section
than that of the base compact and to position it on the base compact so that the centerlines of the two
roughly coincide with each other. The secondary cause of the effect was determined to be amendable to
either or both of the alloy content of the graphitizing element and the cooling rate after infiltration. Given
the silicon contents of the standard compositions, the maximum permissible cooling rate needed to
prevent distortion due to this cause is about 20 oC/min. (0.6 oF/sec.).
Mechanical Property Studies
The study reported in 2005 was conducted to survey the transverse rupture and tensile properties of the
standard composition as well as to determine the effects on tensile properties of minor alloy
modifications of the base compact composition [12]. The latter included one or more of copper, nickel,
manganese and molybdenum. In addition to the as-infiltrated properties, the study also surveyed the
effects of various heat treatments. These included stress relieving, sub-critical (or so-called light)
annealing, and normalizing.
The study was conducted in three phases, each with slightly differing objectives in terms of the alloy
modifications and the heat treatments that were examined. The sintering and infiltration conditions,
however, were the same in all three cases: the infiltrant weight was nominally 13.5% of the base compact
weight; thermal processing was in a production belt furnace; the atmosphere was 90 v/o N2 – 10 v/o H2
with 0.25 v/o CH4 added to control the carbon potential; infiltration was at 1185 oC (2165 oF) for ½ hour
at temperature; and, cooling subsequent to infiltration averaged about 10 oC/min. (0.3 oF/sec.).
Phase I
The tensile properties of the standard infiltrant and base compact compositions and the alloy effects of
modest additions to the base compact composition of copper, nickel, and manganese were determined in
each of the as-infiltrated, stressed relived, and sub-critically annealed conditions. The as-infiltrated results
are shown below in Table 1.
The final carbon contents of the specimens in this case ranged from ~ 2.0 to 2.1 %. Assuming complete
graphitization of the hyper-eutectoid carbon contents involved, the corresponding pore free densities of
the specimens would be expected to range from ~ 7.52 to 7.54 g/cm3. Thus, the densities of the five
compositions listed in the table were all upwards of 96 % of the pore free value.
Table 1 – As-Infiltrated Properties of the Si Base and Cu, Ni and Mn Modified Mixes
Yield Strength
Ultimate Strength Elongation Hardness
Alloy ID
3
MPa
(103 psi) % in 2.5 cm
HRA
MPa
(10 psi)
353.7
(51.3)
468.2
(70.0)
1.4
56
Base (0.75%Si)
422.0
(61.2)
597.1
(86.6)
2.2
60
Base + 1 Cu
376.5
(54.6)
504.0
(73.1)
1.7
55
Base + 1 Ni
(62.5)
582.6
(84.5)
1.8
60
Base + 1 Cu + 1 Ni 430.9
384.7
(55.8)
515.7
(74.8)
1.5
57
Base + 0.5 Mn
Density
g/cm3
7.47
7.43
7.46
7.25
7.34
As expected, the tensile and hardness results in the table showed that the Si Base composition had the
lowest overall properties. In comparison, each of the alloy compositions exhibited substantially higher
strength values and in most cases, modestly higher ductility and hardness values as well. The greatest
increases in all four properties were in the specimens of the Base + 1Cu and the Base + 1Cu + 1Ni
compositions. In view of the relatively low density of the latter, the general implication of the findings
was that copper was the single most effective alloy addition.
The stress relief anneal generally led to very modest improvements in the ultimate strength and ductility
values but to little or no change in the yield strength, hardness or density values. In contrast, the subcritical anneal led to fairly substantial ductility increases but to equally significant decreases in strength
and hardness. The density, in most cases, was unaffected by this treatment. The results of the sub-critical
anneal are shown below in Table 2.
Table 2 – Sub-Critically Annealed Properties of the Si Base and Cu, Ni and Mn Modified Mixes
Yield Strength
Ultimate Strength
Elongation Hardness Density
Alloy ID
MPa
(103 psi)
% in 2.5 cm
HRA
g/cm3
MPa
(103 psi)
315.8
(45.8)
475.8
(69.0)
2.1
55
Si Base
7.47
363.4
(52.7)
561.9
(81.5)
2.7
57
Base + 1 Cu
7.39
326.8
(47.4)
490.2
(71.1)
2.5
55
Base + 1 Ni
7.46
(48.5)
506.8
(73.4)
3.4
53
Base + 1 Cu + 1 Ni 334.4
7.25
324.1
(47.0)
475.8
(69.0)
2.1
54
Base + 0.5 Mn
7.34
The presence of free graphite in these compositions makes them similar in many respects to the cast irons.
Apart from alloy content and the microstructure of the iron base matrix, it is the morphology of the
graphite precipitates that largely determines the properties of the cast irons. In general, there are four
different morphologies that comprise the predominant types in each of the four principal cast iron grades.
In order of decreasing symmetry and correspondingly, of decreasing potential in terms of mechanical
properties, these include: 1) the nodular or spheroidal type of the Ductile cast irons; 2) the temper carbon
type of the Malleable cast irons; 3) the vermicular or compacted type of the Compacted Graphite cast
irons; and, 4) the flake type of the Grey cast irons [14]. The predominant graphite morphology of the
present compositions including those that have yet to be discussed is the vermicular or compacted type.
Thus, it is of interest to compare the properties of the present compositions with those of the Compacted
Graphite (or CG) cast irons. These are shown below in Table 3 [16].
A brief review of the data in Table 3 will show that the highest strength and hardness properties are
associated with the normalized condition and the highest ductilities are those of the ferritized (full
annealed) condition. Predictably, the strength and hardness values of the nickel-containing grade were
generally better than those of the un-alloyed grade.
Table 3 - Typical Tensile Properties Of Compacted Graphite Cast Irons
Yield
Tensile
Elongation
Iron
Strength
Strength
Condition
(a)
Matrix
MPa
(ksi)
MPa
(ksi)
%
60% F
263
(38.1) 325
(47.1)
2.8
As-Cast
(b)
100% F
231
(33.5) 294
(42.6)
5.5
Ferritized
(c)
90% P
307
(44.5) 423
(61.3)
2.5
Normalized
328
(46.7) 427
(61.9)
2.3
As-Cast
…
(b)
100% F
287
(41.6) 333
(48.3)
6.0
Ferritized
(c)
90% P
375
(54.4) 503
(73.0)
2.0
Normalized
Hardness
HRA
48
47
52
53
49
56
(e)
Nickel
%
1.5
1.5
1.5
(a) F, ferrite; P, pearlite. (b) Annealed, 2 hr. at 900 oC (1650 oF), furnace cooled to 690 oC (1275 oF), held 12 hr.,
cooled in air. (c) Austenitized 2 hr. at 900 oC (1650 oF), cooled in air. (e) Converted from Brinell values.
Comparison of the data in this table with those in the earlier Table 1 will show that the strength and
hardness of all of the infiltrated compositions were significantly better than those of both of the CG irons
in the as cast condition and of the plain CG iron in the normalized condition. Otherwise, the strength and
hardness of the Si Base composition closely approached those of the nickel-containing grade in the
normalized condition while those of the remaining four compositions were either equivalent or
substantially better than those of the nickel grade in the normalized condition.
The fact that the strengths and hardnesses of the infiltrated compositions were generally better than those
of the CG cast irons is thought to be attributable to the inherently lower densities of the cast irons. For
example, compared with the present compositions, the cast irons generally have both higher carbon and
higher silicon contents and each lead to significantly lower pore free densities. In the case of the CG
irons, quantitative estimates indicate that their pore free density ranges from about 7.17 g/cm3 in the
ferritized condition to about 7.25 g/cm3 in the normalized condition.
It is also of interest to compare the properties of the present compositions with those of the standard P/M
grades. Since the microstructures of the infiltrated compositions are predominantly pearlitic, the most
realistic comparison is with the predominantly pearlitic P/M grades or, in effect, with the 0.6 to 0.9 %
carbon containing grades in the as-sintered condition. MPIF Standard 35 lists three such grades, which are
otherwise plausibly comparable in terms of their total ‘second’ alloy contents. These include the
following: F-0008, FC-0208 and FN-0208. The highest tensile properties that Standard 35 lists for each
are shown below in Table 4.
Table 4 – Tensile Properties of Comparable Standard P/M Grades
Yield Strength Ultimate Strength Elongation
MPIF
Grade Designation
MPa
(103 psi)
% in 2.5 cm
MPa
(103 psi)
275.8
(40.0)
393.0
(57.0)
1.0
F-0008-35
448.2
(65.0)
517.1
(75.0)
<1.0
FC-0208-60
379.2
(55.0)
620.6
(90.0)
3.0
FN-0208-50
Hardness
HRA *
44
52
54
Density
g/cm3
7.0
7.2
7.4
* Converted from HRB values.
Limiting the considerations to the Si Base and Base + 1Cu compositions, comparison of the data in this
table with those in the earlier Table 1 will show that the properties of the Si Base were far superior to
those of the F-0008 grade and otherwise approached those of the FC-0208 grade. Similarly, the properties
of the Base + 1Cu composition were generally superior to those of the FC-0208 grade and rivaled those of
the FN-0208 grade. Here again, the higher density of the infiltrated compositions is almost certainly the
major underlying cause of the indicated differences.
Phase II
The second phase of the study had the objective to determine the effects on tensile properties of modest
additions of molybdenum and molybdenum plus copper. The additions were made to the standard Si Base
composition and the properties were again determined in the as-infiltrated, stress relieved, and subcritically annealed conditions.
As in Phase I, the properties in the stress-relieved condition were in most cases very similar to those of
the as-infiltrated condition. However, in the case of the Base + 0.5 Mo + 2 Cu composition, the stress
relief effected significant increases in both the yield and ultimate strength values of the order of 70 Mpa
(10,000 psi) each. Thus, in this instance, it is appropriate to present and discuss the stress relieved
properties rather than the as-infiltrated ones. These are shown overleaf in Table 5.
Table 5 -Tensile Properties of the Molybdenum Containing Alloys in the Stress Relieved Condition
Elastic
Yield
Ultimate
Elongation Hard. Den.
Modulus
Strength
Strength
Alloy ID
GPa (106 psi) MPa (103 psi) MPa (103 psi) % in 2.5 cm RHA g/cm3
162.0 (23.5) 450.2 (65.3) 648.1 (94.0)
2.0
56
Base + 0.3 Mo
7.49
162.7 (23.6) 570.2 (82.7) 685.4 (99.4)
1.3
58
Base + 0.5 Mo
7.53
1.7
61
Base + 0.3 Mo + 1 Cu 171.0 (24.8) 588.1 (85.3) 748.1 (108.5)
7.50
1.3
69
Base + 0.5 Mo + 2 Cu 160.7 (23.3) 557.8 (80.9) 753.6 (109.3)
7.53
In addition to the strength and elongation values, the table also lists the elastic modulus values that were
observed in the tests. These data were included because this property is chiefly affected by density and the
high densities of the present compositions were expected to manifest as increased modulus values. In fact,
the values shown in the table are generally higher than those quoted in MPIF Standard 35 for the majority
of P/M grades which, of course, typically involve lower densities.
Compared with the properties of the Si Base and Base + 1Cu compositions, as listed in the earlier Table 1,
the balance of the data in this table show very substantial strength and hardness improvements but modest
decreases in ductility. The best overall properties in the table are those of the Base + 0.3 Mo + 1 Cu
composition. The fact that these somewhat exceeded those of the higher alloyed Base + 0.5 Mo + 2 Cu
composition suggested that the leaner of the two may actually be closer to the optimum in terms of alloy
content. However, as will be seen, the properties in the sub-critically annealed condition did not support
this idea. Nevertheless, they did show an unusually high ductility value in the case of the leaner alloy.
These findings are shown below in Table 6.
Table 6 - Sub-Critically Annealed Properties of the Molybdenum Containing Compositions
Elastic
Yield
Ultimate
Elong.
Hard. Den.
Modulus
Strength
Strength
Alloy ID
3
3
MPa (10
MPa (10
% in 2.5 cm RHA
g/cm3
GPa (106 psi)
psi)
psi)
172.4 (25.0) 344.1 (49.9) 568.1 (82.4)
3.3
51
Base + 0.3 Mo
7.49
171.0 (24.8) 383.4 (55.6) 588.8 (85.4)
3.7
55
Base + 0.5 Mo
7.53
5.1
52
Base + 0.3 Mo + 1 Cu 186.2 (27.0) 373.0 (54.1) 588.1 (85.3)
7.50
2.1
60
Base + 0.5 Mo + 2 Cu 162.7 (23.6) 490.9 (71.2) 608.8 (88.3)
7.53
Copper and molybdenum are commonly added to the Compacted Graphite cast irons and here again,
where reasonably direct comparisons were possible, the indications were that the properties of the present
iron base infiltrated compositions were generally superior in strengths and hardness and reasonably
comparable in ductility [17, 18].
More interesting, perhaps, is a comparison of the properties of the present compositions with those of the
well-known P/M grades that contain molybdenum. The properties of three such grades as selected from
the Low Alloy Steel and Diffusion Alloyed Steel categories of MPIF Standard 35 are shown overleaf in
Table 7. In each case, the data correspond to steels in the as-sintered condition with carbon contents in the
eutectoid range from 0.4 to 0.9 %. Significantly, the strength, hardness and modulus values that are
shown in the table are essentially the highest values listed in the Standard for steels in the as-sintered
condition.
As those familiar with the MPIF system of grade designation will appreciate, the alloy contents of the
indicated grades are generally both different and, in particular, higher than those of the present infiltrated
compositions. All three of the P/M grades contain at least 0.5 % molybdenum and either 1.75 or 4.0 %
Table 7 – Properties of Selected Molybdenum Containing P/M Steels
Elastic
Yield
Ultimate
MPIF
Elong.
Modulus
Strength
Strength
Grade Designation
GPa (106 psi) MPa (103 psi) MPa (103 psi) % in 2.5 cm
162.0 (23.5) 482.7 (70.0) 689.5 (100.0)
2.0
FLN2-4405-60
158.6 (23.0) 503.3 (73.0) 710.2 (103.0)
1.0
FD-0208-65
168.9 (24.5) 489.5 (71.0) 861.9 (125.0)
2.0
FD-0408-65
Hard.
Den.
RHA
56
56
58
g/cm3
7.30
7.25
7.40
nickel. In addition, each of the diffusion alloyed (FD pre-fixed) grades nominally contain 1.5 % copper.
However, despite these differences, comparison of the data in this table with those in the earlier Table 6
will show that the strength, hardness and modulus values of each of the two copper containing variants of
the infiltrated compositions were superior to those of the FLN2-4405 and FD-0208 grades and were either
superior to or closely approached those of the FD-0408 grade. Otherwise, the ductility values of the two
data sets were reasonably comparable.
Phase III
The objective of the third and last of the three smaller studies mentioned was to examine the effects on
properties of copper modifications of up to 2 % of the Si Base composition in four conditions as follows:
as-infiltrated, stress relieved, sub-critically annealed, and normalized. The study also included
determinations of the TRS properties in the as-infiltrated condition. These results are shown below in
Table 8.
Table 8 - TRS Properties of the Si Base and Copper Modified Base in the As-infiltrated Condition
Transverse
Dimensional
Hardness
Density
Rupture Strength
Change
Alloy ID
3
%
RHA
g/cm3
MPa
(10 psi)
1051.5
(152.5)
0.59
58
Si Base
7.44
1100.4
(159.6)
0.77
60
Base + 1 Cu
7.40
1148.7
(166.6)
0.63
62
Base + 2 Cu
7.43
The TRS properties were chiefly of interest because target applications include Grey cast iron
components and rupture strength is a commonly cited property of the Grey irons. However, the latter is
typically quoted in terms of the breaking load of a standard test specimen and is not directly comparable
with the usual P/M values. For comparison, standard P/M specimens were prepared from a Grey iron
component of interest. They exhibited an average TRS value of 465 MPa (67,500 psi), a hardness of
53 RHA, and a density of 7.23 g/cm3.
The tensile properties of the subject compositions in each of the four process conditions mentioned are
shown overleaf in Table 9.
A review of these data will show that the densities of the Si Base and Base + 1 Cu compositions were a
little lower than earlier (Tables 1 & 2) and that the density of the Base + 2 Cu composition was essentially
intermediate of these two. All of the values were upwards of 98 % of the pore free value (~7.53 g/cm3).
Comparison of the tensile properties of the Si Base and Base + 1 Cu compositions with those of the
earlier study will show that the two data sets were generally similar. In the case of the Base + 2 Cu
composition, the major effects of the additional copper appeared to be to increase the yield strength and
hardness and decrease the elongation values relative to both the Si Base and Base + 1Cu compositions.
The indicated increases in the yield strength were in all cases fairly substantial (≥ 70 MPa ≅ 10,000 psi)
whereas the hardness increases were generally marginal.
Table 9 - Tensile Properties of the Si Base Composition and Two Copper Containing Varian in the
As-Infiltrated and various conditions of Heat treatment
Yield Strength
Ultimate Strength
Elongation Hardness Density
Alloy ID
MPa
(103 psi)
% in 2.5 cm
HRA
g/cm3
MPa
(103 psi)
As-Infiltrated
366.7
(53.1)
502.6
(72.9)
1.9
59
Si Base
7.46
410.3
(59.5)
613.7
(89.0)
2.4
60
Base + 1 Cu
7.39
481.3
(69.8)
604.7
(87.7)
1.5
60
Base + 2 Cu
7.44
Stress Relieved
354.4
(51.4)
528.8
(76.7)
2.3
59
Si Base
7.46
413.0
(59.5)
606.1
(87.9)
2.2
59
Base + 1 Cu
7.39
484.0
(70.2)
580.6
(84.2)
1.1
61
Base + 2 Cu
7.42
Sub-critically Annealed
298.6
(43.3)
524.7
(76.1)
3.4
56
Si Base
7.46
322.0
(46.7)
535.1
(77.6)
4.0
54
Base + 1 Cu
7.42
382.0
(55.4)
535.7
(77.7)
3.1
56
Base + 2 Cu
7.43
Normalized
584.0
(84.7)
781.2
(113.3)
2.2
65
Si Base
7.44
638.5
(92.6)
832.9
(120.8)
2.1
65
Base + 1 Cu
7.40
732.9
(106.3)
835.7
(121.2)
1.1
67
Base + 2 Cu
7.41
A review of the properties in the normalized condition will show that this heat treatment had virtually no
effect on density or elongation but very substantial effects on strength and hardness. For example, a
general comparison of these properties with those in the balance of the paper will show that in most cases
they surpassed the best of the latter including those of the molybdenum modified compositions (Table 5).
Finally, its interesting to make a general comparison of the properties in Table 9 with the those of the
Ductile irons in the plain or essentially un-alloyed condition (i.e. containing from 3 to 4 % C, 0.1 to 1.0 %
Mn, and 1.8 to 2.8 % Si) [19]. Remarkably, in fact, although marginal in ductility, the strength and
hardness values indicated by the present findings significantly exceed those of the latter in most instances.
For example, the minimum requirements of the three highest strength grades of plain Ductile iron in
accordance with ASTM A 536 are shown below in Table 10.
Table 10 – Minimum Tensile Property Requirements of Ductile Iron According to ASTM A 536
Yield
Ultimate
ASTM
Elongation Typically Recommended
Strength
Strength
Grade
Process Condition
3
3
Designation
MPa
(10 psi) % in 2.5 cm
MPa (10 psi)
413.7 (60.0)
551.6
(80.0)
3.0
As Cast
80-60-03
482.7 (70.0)
689.5
(100.0)
3.0
Normalized
100-70-03
620.6 (90.0)
827.4
(120.0)
2.0
Oil Quenched & Tempered
120-90-02
A cursory comparison of the present findings with these data will show the following: 1) the requirements
of the 80-60-03 grade were closely approached by the properties of the Base + 1 Cu composition in the
as-infiltrated and stress relieved conditions; 2) the strength and hardness requirements of the 100-70-03
grade were significantly exceeded by the properties of the Si Base composition in the normalized
condition; and, 3) the requirements of the oil quenched and tempered 120-90-02 grade were met by the
properties of the Base + 1 Cu composition in the normalized condition.
To be fair, what is being called normalizing here may be more akin to sinter hardening. As is generally
well known, the normalizing heat treatment derives its name from the fact that it consists of austenitizing,
usually at a low temperature above the upper critical, followed by normal cooling in still air. As applied
to cast irons, the general aims are to eliminate hyper-eutectoid carbides if they exist, refine the grain size
and produce an iron base matrix that has a predominantly pearlitic microstructure. The cooling rate of the
process naturally depends on the mass of the casting and may be less than ~10 oC/min (0.3 oF/sec).
In comparison, the findings in the present case derive from specimens that were ‘normally’ cooled in the
cooling zone of a P/M furnace. The average cooling rate was estimated to be ~100 oC/min (3 oF/sec) in
the temperature range from 845 to 315 oC (1550 to 600 oF). This, of course, is essentially an intermediate
cooling rate in the sinter hardening range and, at the very least, would be expected to produce a finer
grain size and a smaller interlamillar spacing than would be typical of a normally cooled large casting.
Dimensional Change and Surface Appearance
The study reported in 2006 examined the factors effecting the dimensional change characteristics and
surface quality of the resulting parts. The dimensional change characteristics included part uniformity as
well as part-to-part consistency. The surface quality related to the disposition and effects of infiltrant
residues on the structural integrity and appearance of the infiltrated surfaces. The objective was
essentially to define the conditions needed to implement the process as a practical matter.
Dimensional Change Results
It was explained in the introductory remarks of the paper that the infiltrated compositions according to the
present technology offer a wide choice of dimensional change possibilities ranging from a high of
upwards of 1 % to a low of 0 % or less. However, it was uncertain what effects the processing required to
produce a particular value in this range would have on the dimensional consistency of the resulting parts.
Thus, the primary objective of this phase of the study was to determine which of the extremes of the
indicated range was associated with the best overall dimensional control in terms of part-to-part
variability and within part uniformity.
Two trials were conducted to evaluate the indicated dimensional change effects. The trials were designed
to indicate statistically significant standard deviation differences at the 80% confidence level. The total
number of specimens in each case was 48. The base compact geometry in the trials was a cylindrical disc
measuring about 4.5 cm (1.75 ins.) in diameter and 1.4 cm (0.55 ins.) in height. In addition to the
infiltrated density and dimensional change properties, diametric parameters that indicated the
concentricity and the top to bottom uniformity of the specimens were also evaluated. The values of
selected statistics of four subsets of the specimens that reflected the order of processing in the trial were
compared to assess the operational stability of the processing.
In Trial 1 of the study, the specimens were processed at a low temperature (1165 oC ≈ 2125 oF) and using
a high infiltrant weight (i.e. 16.5% of the base compact weight). The resulting infiltrated density averaged
7.40 g/cm3 versus an expected value of ~ 7.47 g/cm3 and a pore free value of 7.54 g/cm3.The shortfall in
density was traced to relatively high infiltrant residue weights. The standard deviation of the density was
0.018 g/cm3 and the corresponding range of variation in the data was about 1% of the average. The
average diametric dimensional change was 1.29 % from green and 1.37 % from die. The standard
deviation, in this case, was 0.051 % and the corresponding 4σ value of 0.204% indicated a maximum
diametric variation about the average of ± 0.0046cm (0.0018 ins.). The diameter data of the trial
confirmed this value. The concentricity parameter indicated that the specimens were out of round by an
average of 0.0043 cm (0.0017 ins.). The top to bottom uniformity parameter indicated that the bottom
diameters were larger than the top ones in 90% of the cases. The average difference was 0.0031 cm
(0.0012 ins). The general indications of the data of the four subsets of the specimens was that the process
was stable during the course of the trial.
In Trial 2, the specimens were initially processed at the same low temperature as in Trial 1 but using a
relatively lower infiltrant weight ( i.e. 12.5 % of the base compact weight) and then submitted to a high
temperature liquid phase sintering step at 1190 oC (2180 oF). The infiltrated density averaged 7.47 g/cm3
versus an expected value of 7.54 g/cm3 (i.e. the pore free value). Here again, the shortfall was traced to
relatively high infiltrant residue weights. The standard deviation of the density was 0.022 g/cm3 and the
corresponding range of variation in the data was about 1.5 % of the average, 50% greater than in Trial 1.
The average diametric dimensional change in this case was -0.01 % from green and 0.07 % from die. The
standard deviation was 0.060 %, about 20 % greater than in the first trial. The corresponding 4σ value of
0.24% indicated a maximum diametric variation about the average of ± 0.0053cm (0.0021 ins.). The
actual diameter data of the trial indicated a slightly lower variation of ± 0.0048cm (0.0019 ins.). Here
again, the concentricity parameter indicated that the specimens were out of round. The average was
0.0057 cm (0.0022 ins.), about 30 % greater than in the first trial. The top to bottom uniformity parameter
indicated an average difference of 0.0131 cm (0.0051 ins.), a little over 4 times the value of the first trial.
In this case, the bottom diameters of the specimens were uniformly larger than the top ones. Frictional
effects between the specimens and the sintering trays were apparently responsible for these differences.
As earlier, the general indications of the data of the four subsets of the specimens was that the process
was stable during the course of the trial.
In view of these findings, it was concluded that at least in the near term, there is essentially nothing to be
gained in terms of dimensional consistency from using the sinterability inherent in the compositions of
the technology to produce a low average dimensional change value. The potential advantage that a near
zero dimensional change offers to use the compaction die to size the parts would almost certainly be
negated in this case by the very large top to bottom differences that were observed. In addition, the results
of earlier studies suggested that the slightly higher infiltrated densities that the associated processing is
likely to effect will generally not lead to significantly higher tensile properties. In part, this is because the
pores that remain after infiltration are both closed and generally well rounded and in part, because at
contents of 1 or 2 %, they are not the dominant stress raising feature of the microstructure.
Surface Appearance Results
The infiltrated surface appearances of the specimens of both trials were problematic. As expected, the
infiltrant slugs had left thin, centrally located residues which were roughly circular in shape and ~ 2 cm
(0.75 ins) in diameter. Unexpectedly, however, the residues were somewhat adherent to the underlying
surfaces and had to be mechanically removed. This was done manually using a knife-edge as a pry. In the
case of the specimens of the first trial , the residues normally came off the surfaces intact and without too
much difficulty. In the case of the specimens of the second trial, there was a greater tendency for the
residues to breakup during the process and they were generally more difficult to remove. These changes
evidently reflected the effects of the generally higher process temperature in the second trial.
The general condition of the un-infiltrated surfaces of the specimens as well as of the infiltrated surfaces
to within a few millimeters of the residues was excellent. In comparison, the condition of the surfaces
under and just adjacent to the residues was poor. In the general case, these areas were either flawed with
erosion pits or covered with residual particles including partially melted and/or un-melted iron and
residual oxides, sulfides and possibly, un-reacted graphite. In some places, the pitting and particles
occurred together. In the case of the specimens of the first trial, low power microscopic examinations
suggested that both defects were more cosmetic than structural. However, this was not true of the
specimens of the second trial. The erosion pits in particular, in this case, appeared to be deep enough in
many places to threaten the structural integrity of the surfaces.
Neither these findings nor the aforementioned high residue weights were expected. Earlier, albeit more
limited, studies which were also based on a slightly different infiltrant composition, had generally
produced very different results. The observed infiltrant residue weights (as a percentage of the starting
weights) normally varied somewhat but were never seen to be as high as the present ones. Similarly,
although the resultant infiltrated surfaces were seldom flawless, they were generally better than the
present ones.
These indications led to the third and final trial of the study. The results of this trial largely confirmed the
importance of the infiltrant composition and, more generally, showed that material selection in the design
of the infiltrant was far more important than hitherto suspected.
DISCUSSION AND CONCLUSIONS
Experience to date with iron base infiltration suggests that thermal processing will be the difficult step for
parts makers. Compared with traditional P/M processing, the required temperatures are of the order of
70 to 85 oC (125 to 150 oF) higher and it’s likely that the process will prove to be more troublesome in
terms of establishing and managing the thermal conditions that are needed to produce dimensionally
stable parts. Much will also depend on the ability to supply quality mixes to implement the process. In
fact, it is precisely the difficulty which the infiltrant mix represents in this respect that has so far been the
main sticking point in commercializing the process. Five different approaches to overcoming the
associated problems have been identified, so it’s reasonable to expect that a solution exists. However, it
still remains to develop and evaluate these approaches in order to decide how best to proceed.
Otherwise, the process appears to offer considerable potential versus both regular and high density P/M
processing. Its three most obvious advantages in comparison with traditional processing include the
following:
1) Base compact densities of 6.8 g/cm3 or less are sufficient to obtain a virtually pore free density
after infiltration. Thus, a particular advantage of the technology as compared with the traditional
high pressure / high density processes [20, 21] is that the low starting densities essentially provide
the potential to press larger parts that can later be infiltrated to the same or even higher final
densities.
2) Infiltrated densities within 98% of the pore free value were demonstrated to be routinely possible.
In contrast with traditional high density processing as exemplified by the double press and sinter
and warm compaction technologies, the pores remaining after infiltration are completely isolated.
Hence, infiltrated parts are likely to exhibit better dynamic properties (e.g. fatigue and impact
properties) along with better machining and plating properties as well as greater applicability to
the manufacture of hydraulic parts.
3) And, finally, the generally higher densities obtainable by iron base infiltration hold out the
promise of higher or equivalent properties at significantly lower total alloy contents and thus, the
likelihood of greater overall economy in parts production, especially, in view of present alloy
costs.
ACKNOWLEDGMENTS
Special thanks are due to the Ben Franklin Technology Partners of Pennsylvania for funding a part of this
research, and (in memoriam) to Mister W. B. Bentcliff of the Hoeganaes Laboratory for his help in
obtaining the data used in preparing the manuscript.
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