EFFECT OF MICROSTRUCTURAL INHOMOGENEITIES
ON THE FATIGUE CRACK GROWTH RESPONSE OF
A PREALLOYED AND TWO HYBRID P/M STEELS
Suleyman Saritas*, Robert Causton**, W. Brian James*** and Alan Lawley****
*Professor, Department of Mechanical Engineering, Gazi University,
Maltepe, Ankara, 60570, Turkey
**Senior Scientist, Hoeganaes Corporation, Cinnaminson, NJ 08077, USA
***Manager, Application Engineering, Hoeganaes Corporation, Cinnaminson, NJ 08077, USA
****Grosvenor Professor, Department of Materials Engineering, Drexel University,
Philadelphia, PA 19104, USA
Presented at PM2TEC2002
World Congress on Powder Metallurgy & Particular Material
June 16-21, 2002 Orlando, Florida
ABSTRACT
In the first phase of this study, the effect of microstructural inhomogeneities on the tensile and impact response of
a prealloyed (FL-4405) and two hybrid (FLC2-4405 and FLN2-4405) P/M steels based on prealloyed
Ancorsteel 85 HP was evaluated. In phase two we assess crack propagation response. The base powder and
additions were mixed with 0.75 w/o Lonza Acrawax in 227 kg (500 lb) batches. A density of 7.4 g/cm3 was
obtained by double pressing (550/550 MPa). Sintering temperatures of 2050 °F(1120 °C) and 2300 °F (1260 °C )
were utilized. A group of sintered compacts of each alloy was heat treated by quenching from 1650 °F (900 °C)
into warm oil at 160 °F (70 °C) followed by tempering at 375 °F (190 °C) for 1 h. Two groups of sintered
compacts of the FLC2-4405 and FLN2-4405 alloys were sinter hardened and tempered at 375 °F (190 °C) for 1 h.
Experimental data showed that the P/M steels exhibit comparable fatigue crack growth rates (1.1207E-4 to
3.0185E-4 mm/cycle) at a stress intensity range of 1000 MPa (mm)1/2. Quenched and tempered microstructures
resulted in the highest fatigue crack growth rate. Sinter hardening of FLC2-4405 and FLN2-4405 lowered the
fatigue crack growth rate. High temperature sintering reduced the fatigue crack growth rate in FL-4405 but
increased it in FLC2-4405 and FLN2-4405.
INRODUCTION
Fatigue is a complex phenomenon influenced by numerous factors. For powder metallurgy (P/M) steels;
pores, microstructure, surface finish, residual stresses and external notches are of primary importance [121]. Surface pores act as preferred sites for fatigue crack initiation by acting as stress concentrators [1113]. Since the size, shape and orientation of pores vary from pore to pore, individual stress concentration
factors are unknown. The fact that most fatigue failures originate at free surfaces suggests that any surface
treatment which increases the density or the hardness of the surface region will retard fatigue crack
initiation and thus improve fatigue life[16-20]. The shape and distribution of pores may be altered by the
powder characteristics and processing route (e.g. powder size distribution, alloying and lubricant
additives, and compaction pressure) and the propensity for the pores to act as crack precursors may be
reduced [15, 22-28]. The microstructures of P/M steels depend on alloying technique and the production
procedures. There are four major and distinct alloying techniques used for formulating P/M steels and
these establish a classification for the steels: admixed, diffusion alloyed, prealloyed and hybrid [29]. The
resulting microstructures may be homogeneous or heterogeneous with respect to chemical composition
and to the constituent phases. The effect of homogeneity / heterogeneity on fatigue properties is subject to
differences of opinion. From the classical metallurgical view homogenous structures are desirable and
prealloyed steels are preferred in this regard. On the other hand, many investigators [30-35] note that the
constituents of heterogeneous microstructures reinforce each other (as in composite materials) and thus
they result in improved performance. In order to improve further the mechanical properties of sintered
steels, through hardening or surface hardening can be performed, depending on the application of the part.
If the final component is exposed primarily to axial stresses, through hardening is preferred, whereas
surface hardening is applied to components subjected to high bending stresses and/or high surface
pressures such as in gear applications. The highest fatigue endurance limit is always obtained after
quench-hardening and tempering [36]. Sinter-hardening, which is not a secondary operation, is becoming
attractive since it eliminates problems of distortion and oil retention when quenching porous steels [37,
38]. There is little information available on the fatigue behavior of sinter-hardened steels [39, 40].
The effect of pores on the fatigue crack propagation rate depends on their morphology. Round pores are
known to decrease the propagation rate by blunting the crack tip [22] whereas pores with sharp edges act
as linkage sites which assist crack propagation [13]. While existing theories explain successfully some of
the effects of porosity on crack propagation, no complete predictive model has been developed that
embraces microstructural effects [8,41-43]. In addition to the total amount of porosity, pore curvature, and
pore separation distance influence the response of P/M steels to the fatigue environment [44].
Fatigue crack propagation rates have been analyzed, and the cyclic growth of a fatigue crack can be
predicted by the modified Paris-Erdogan law:
da
= A(∆K) n
dN
(1)
where da/dN is the fatigue crack propagation rate, A and n are material constants determined
experimentally, and ∆K is the stress intensity range. da/dN vs ∆K curves of P/M steels are identical in
shape to those of wrought steels. However, they are one order of magnitude higher. The material
parameters in Eq.1 are sensitive to porosity. The coefficient A increases with increasing porosity, resulting
in higher crack growth rates than in wrought steels at comparable values of ∆K [8]. The threshold stress
intensity range ∆Kth, below which no crack growth occurs, is independent of porosity for 0 to 8 v/o
porosity. At higher levels of porosity, ∆Kth decreases as the amount of porosity increases [44]. This
transition is attributed to a change from interconnected pores to isolated pores at a density of about
7.2 g/cm3 (92% pore-free density).
In the present study, the fatigue crack growth behavior of one prealloyed (FL-4405) and two hybrid
(FLC2-4405 and FLN2-4405) P/M steels based on Ancorsteel 85 HP are reported. In the first phase of this
study, the effect of microstructural inhomogeneities on the tensile, impact and hardenability response of
these steels was evaluated [45-47]. Three compositions, two sintering temperatures and three heat
treatment conditions are under investigation.
EXPERIMENTAL PROCEDURE
Materials and Processing
Three P/M steels based on Hoeganaes Ancorsteel 85 HP were examined. The compositions and properties
of the premixes (FL-4405, FLC2-4405 and FLN2-4405) are given in Table I. After sintering, the carbon
level was 0.57 w/o in each steel; oxygen levels were in the range 0.009 – 0.055 w/o.
Table I. Compositions and Properties of Premixes
Premix*
Flow
(s/50g)
Apparent
Density
(g/cm3)
Composition (w/o)
Ancorsteel 85 HP
Copper
FL-4405
3.11
30
Balance
0
FLC2-4405
3.11
30
Balance
2
FLN2-4405
3.13
29
Balance
0
* Metal Powder Industries Federation, Princeton, NJ, Standard 35
Nickel
Graphite
Lubricant
0
0
2
0.6
0.6
0.6
0.75
0.75
0.75
A batch size of 227 kg (500 lb) was mixed for each of the three alloys in the Hoeganaes pilot plant using
the ANCORBOND process. To achieve a density of about 7.4 g/cm3 the powder was double pressed at
550/550 MPa (80,000 psi) pressure and double sintered (DPDS). The dimensions of the test blanks were
12.7x12.7x100 mm.
Sintering was carried out in an Abbott ceramic belt high temperature furnace in an atmosphere of 75 v/o
hydrogen - 25 v/o nitrogen for 30min. Compacts were presintered at 1450 °F (790 °C) after first pressing.
The furnace was equipped with a Varicool cooling zone to provide accelerated cooling from the sintering
temperature. The sintering and heat treating conditions employed in the present study are summarized in
Table II. For quench hardening and tempering, test pieces were austenitized in a Lindberg sealed quench
furnace at 1650 ºF (900 ºC) in an atmosphere of 75 v/o hydrogen - 25 v/o nitrogen for 30 min. The blanks
were quenched into oil preheated to 160 ºF(70 ºC). Excess oil was wiped from the quenched blanks before
tempering in a Blue M oven in air. Tempering was carried out at a temperature of 375 ºF (190 ºC) for 60
min. The sinter hardened test pieces were tempered under the same conditions as the heat-treated test
pieces.
Table II. Sintering and Heat Treating Conditions
Process
Condition
Sintered
Sinter-Hardened
Quenched +
Tempered
Sintered
Sintering
Temperature
ºF(ºC)
2050 (1120)
2050 (1120)
2050 (1120)
2300 (1260)
Cooling After
Sintering
Standard
100% Varicool
Standard
Austenitization
Temperature
ºF(ºC)
1650(900)
Tempering
Temperature
ºF(ºC)
375 (190)
375 (190)
Standard
-
-
The fatigue crack growth test pieces with the dimensions given in Figure 1 were machined from the
blanks. 45º notches were machined by milling using with a cutter with a nose radius of 0.25 mm. The
machined test pieces were stress relieved at 375 ºF (190 ºC) to remove any residual stresses from the
machining and grinding operations. Densities of the green and sintered steels were determined by the
water displacement method outlined in MPIF, Standard 42 [48].
12.7
50
50
3
6.35
Figure 1. Fatigue crack growth test piece (ASTM E 647 & E 399)
Fatigue Crack Growth Testing
Fatigue crack propagation tests were performed using a Dartec servo-hydraulic testing machine. A standard bend
specimen geometry (Figure 1) was used according to ASTM standards E 399 [49] and E 647 [50]. The dimension
(B) was chosen as half of the dimension W. Specimens were loaded in a three-point bending configuration with a
support span (S) of four times W. The load ratio (R) was 0.1. The bend test fixture was manufactured according to
ASTM standard E 399. A sinusoidal waveform and a testing frequency of 1 Hz were selected. Specimens were
fatigue precracked to obtain an initial crack length (a0) of 4 mm (1.0 mm fatigue precrack) in stress cycles of not
less than 104, as specified in ASTM standards E 399 and E 647. Crack length was measured with a traveling
microscope (accurate to 0.01 mm). Readings were taken at 0.1 mm intervals. Shedding of the load was inevitable,
and the minimum crack extension between shedding was 0.5 mm (except in the quenched and tempered
specimens). The stress intensity range ∆K was calculated from the equation given in ASTM standard E 399:
∆K =
∆PS
f (α )
BW 3 / 2
(2)
where ∆P is the load range (Pmax – Pmin) in N, S is the span (the distance between the axes of the support cylinders;
50.8 mm), B is 6.35 mm, W is 12.7 mm and f(α) is a calibration factor that was calculated from the equation given
in ASTM Standard E 399:
f (α ) =
[
3α 1/ 2 1.99 − α (1 − α )(2.15 − 3.93α + 2.7α 2
2(1 + α )(1 − α )3 / 2
]
(3)
where α = a/W, and a is the crack length.
Fatigue crack growth rates (da/dN), were calculated from the experimental data by using an incremental
polynomial method, as suggested in ASTM standard E 647. A graph of crack length vs number of stress
reversals was plotted and a second order polynomial (parabola) was fitted to the data points. The equation
of the parabola was obtained from the software. The derivative of the equation was taken to represent
da/dN and the values were calculated at each experimental point. Three specimens were tested for each
steel under the same processing conditions.
Metallography
Metallographic specimens of all test materials were analyzed by optical microscopy in the polished and
etched conditions. Polished metallographic samples were also utilized to evaluate the stereological
parameters of the pores (shape, size and distribution). This was accomplished by means of a Clemex 1024
automated image analysis system. Analysis of the pore shape was conducted to determine the average
degree of circularity, where circularity was determined from the equation [51]:
Circularity = 4πA/(P)2
(4)
where A = pore area, and P = pore circumference.
A form factor of unity represents a circular pore in the plane of analysis; as the number decreases from
unity, the degree of irregularity increases.
EXPERIMENTAL RESULTS
Microstructures
(a)
(b)
(c)
Figure 2. Microstructures of FL-4405: (a) as sintered at 2050 ºF (1120 oC)
(b) sintered at 2050 ºF (1120 oC) and quenched and tempered, (c) as sintered at 2300 ºF (1260 oC)
(a)
(b)
(c)
(d)
Figure 3. Microstructures of FLC2-4405: (a) as sintered at 2050 ºF (1120 oC),
(b) sintered at 2050 ºF (1120 oC) and quenched and tempered, (c) as sintered at 2300 ºF (1260 oC)
(d) sinter hardened from 2050 ºF (1120 oC)
Figures 2 to 4 show representative microstructures of the three alloys for each of the processing
conditions. As-sintered microstructures consist primarily of divorced pearlite in the three steels. FLC24405 and FLN2-4405 exhibit some martensitic areas. In addition, some nickel-rich areas are seen in
FLN2-4405. Quenching produced a structure of more than 90% martensite in the steels. Sinter hardening
produced about 50% martensite in FLC2-4405 and FLN2-4405. Figure 5 shows the distribution of the
circularity of the pores. The numbers in the legend in each figure refer to sintering temperature (ºF).
Fatigue Crack Growth
A sample fatigue crack growth curve is given in Figure 6. The graph corresponds to one of the FLC2-4405
steel specimens in the sinter hardened condition. A parabolic trend line was added to the data points and
the software gave the equation representing the parabola. The derivative of the equation was taken and the
da/dN values were calculated at each data point by using the derivative. Figures 7 to 9 show fatigue crack
growth graphs for the three steels. The fatigue crack growth graphs are replotted in Figures 10 to 13 for
specific processing conditions. Table III summarizes the mechanical properties and fatigue crack growth
data for the steels.
(a)
(b)
(c)
(d)
Figure 4. Microstructures of FLN2-4405: (a) as sintered at 2050 ºF (1120 oC),
(b) sintered at 2050 ºF (1120 oC) and quenched and tempered, (c) as sintered at 2300 ºF (1260 oC),
(d) sinter hardened from 2050 ºF (1120 oC)
30
(a)
DPDS (2050)
25
Frequency (%)
DPDS (2300)
20
15
10
5
0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
1
Circularity
30
DPDS (2050)
(b)
Frequency (%)
25
DPDS (2300)
20
15
10
5
0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
1
Circularity
30
DPDS (2050)
(c)
frequency (%)
25
DPDS (2300)
20
15
10
5
0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
Circularity
Figure 5. Effect of sintering temperature (oF) on the circularity of pores:
(a) FL-4405, (b) FLC2-4405, (c) FLN2-4405
1
8,5
y = 1,9611E-08x2 - 8,1869E-04x + 1,2730E+01
8
Crack Length, a (mm)
7,5
7
6,5
6
5,5
5
4,5
4
3,5
15000
17000
19000
21000
23000
25000
27000
29000
31000
33000
35000
Stress Reversal, N (cycle)
Figure 6. Representative fatigue crack growth curve; FL-4405 sintered at 2300 ºF (1120 oC)
1,00E-03
FL-4405-2050 F
da/dN (mm/cycle)
FL-4405-2300 F
FL-4405-QT
1,00E-04
1,00E-05
100
1000
∆Κ (MPa
(mm)
(
10000
1/2
)
Figure 7. Fatigue crack growth rate vs stress intensity range for FL-4405.
1,00E-03
FLC2-4405-2050 F
FLC2-4405-2300 F
da/dN (mm/cycle)
FLC2-4405-QT
FLC2-4405-SH
1,00E-04
1,00E-05
100
1000
∆K (MPa (mm)1/2 )
10000
Figure 8. Fatigue crack growth rate vs stress intensity range for FLC2-4405.
1,00E-03
FLN2-4405-SH
FLN2-4405-2050 F
da/dN (mm/cycle)
FLN2-4405-QT
FLN2-4405-2300 F
1,00E-04
1,00E-05
100
1000
10000
∆K (MPa (mm)
1/2
)
Figure 9. Fatigue crack growth rate vs stress intensity range for FLN2-4405.
1,00E-03
FL-4405
FLC2-4405
da/dN (mm/cycle)
FLN2-4405
1,00E-04
1,00E-05
100
1000
10000
∆K (MPa (mm)
1/2
)
Figure 10. Fatigue crack growth rate vs stress intensity range for steels sintered at 2050 ºF (1120 oC)
1,00E-03
FL-4405
FLC2-4405
da/dN (mm/cycle)
FLN2-4405
1,00E-04
1,00E-05
100
1000
10000
∆K (MPa (mm)1/2 )
Figure 11. Fatigue crack growth rate vs stress intensity range for steels sintered at 2300 ºF (1260 oC)
0,001
FL-4405
FLC2-4405
da/dN (mm/cycle)
FLN2-4405
0,0001
0,00001
100
1000
10000
∆K (MPa (mm)1/2 )
Figure 12. Fatigue crack growth rate vs stress intensity range for steels in QT condition
0,001
FLC2-4405
da/dN (mm/cycle)
FLN2-4405
0,0001
0,00001
100
1000
∆K (MPa (mm)1/2 )
Figure 13 . Fatigue crack growth rate vs stress intensity range for steels in SH condition
10000
Table III. Mechanical Properties as a Function of Composition and Processing History
Material
Sintering
Temp.
°F (°C)
Density
(g/cm3)
Heat
Treatment
Tensile
Strength
(MPa)
FL-4405
2050
(1120)
2300
(1260)
2050
(1120)
7.39
7.35
7.35
AS*
QT**
AS
647
1486
659
da/dN
(mm/cycle)
at
1000 MPa
(mm)1/2
2.3953E-4
2.6854E-4
1.9785E-4
7.28
7.33
7.33
7.25
AS
QT
SH***
AS
788
1432
822
745
7.44
7.39
7.39
7.39
AS
QT
SH
AS
826
1644
982
770
FLC2-4405
FLN2-4405
2300
(1260)
2050
(1120)
2300
(1260)
∆Kth
Paris-Erdogan Constants
MPa(mm)1/2
A
n
820
845
826
6.5655E-15
2.3454E-16
4.2712E-16
3.5207
4.0196
3.8886
2.5580E-4
3.0185E-4
2.1851E-4
2.6944E-4
836
828
826
832
1.1612E-9
8.8693E-17
3.9626E-12
4.1234E-13
1.7810
4.1773
2.5805
2.9384
1.5473E-4
2.8297E-4
1.1207E-4
1.9923E-4
810
836
830
837
2.6521E-13
2.3364E-12
3.1132E-13
5.3526E-12
2.9220
2.6944
2.8521
2.5236
*AS: As sintered,
**QT: Quenched + tempered,
***SH: Sinter-hardened
DISCUSSION
Mechanical properties reported previously [45] are summarized in Table III, and show that the P/M steels
investigated demonstrate strength levels suitable for high performance applications (tensile strength >
1000 MPa). At a nominal sintered density of 7.4 g/cm3 high temperature sintering did not produce any
change in tensile strength for FL-4405 but it did reduce tensile strength slightly in the FLC2-4405 and
FLN2-4405 steels. Sinter hardening of the FLC2-4405 and FLN2-4405 steels from 2050 ºF(1120 ºC)
resulted in tensile strength levels that are higher than the as-sintered values, but much lower than the
quenched and tempered values. These strength levels, coupled with the microstructures, confirm that the
cooling rate of the furnace in 100% Varicool is not sufficient for the complete transformation of austenite
to martensite.
Fatigue crack growth data are summarized in Table III for the steels following the different processing
conditions. The threshold stress intensity range (∆Kth) for crack initiation is similar for all the steels. The
average value is about 830 MPa (mm)1/2 . The fatigue crack growth rate at a stress intensity range of
1000 MPa (mm)1/2 is from 1.1207E-4 to 3.0185E-4 for all the steels. Most of the data points overlapped.
As seen in Figures 7 to 9, the quenched and tempered microstructure resulted in a higher fatigue crack
growth rate, followed by the 2050 oF (1120 oC) sinter, 2300 oF (1260 oC) sinter and sinter hardening.
Thus, sinter hardening resulted in the lowest fatigue crack growth rates. When the steels are compared for
the same processing condition (Figures 10 to 13), the FLN2-4405 steel exhibited the lowest fatigue crack
growth rate, except for the quenched and tempered case where FL-4405 had the lowest fatigue crack
growth rate. It is believed that the inhomogeneous microstructure (a mixture of soft and hard phases)
produced by sinter hardening decreased the fatigue crack growth rate.
To obtain the Paris-Erdogan constants, given in Table III, power curves were fitted to the experimental
data set. These appear as straight lines on the log-log graph. Regression coefficients (R2) for the fitting of
these lines were about 0.70 which indicates that the fit is not very good. The low values of the regression
coefficients may be explained partly by the parabolic trends near the threshold region (region I) and near
the final fracture region (region III). It appears that the main reason for the low values of the regression
coefficients is an insufficient number of data points . Additional specimens should be tested to define
more clearly the fatigue crack growth curves.
The microstructures of the FL-4405 steel in the as-sintered conditions (Figures 2(a) and 2(c)) are similar.
The microstructure consists primarily of divorced pearlite and some ferrite. The quenched and tempered
microstructure consists of almost 100% tempered martensite (Figure 2(b)). The microstructures of the
FLC2-4405 steel in the as-sintered conditions (Figures 3(a) and 3(c)) are comparable. The brownish
etching areas consist of fine martensite. The presence of more than 50% cleavage facets on the tensile
fracture surfaces confirms this observation [45]. The microstructure consists primarily of divorced
pearlite. The quenched and tempered microstructure is almost 100% tempered martensite (Figure 3(b)).
The sinter hardened microstructure is about 50% martensite, 45% pearlite and some retained austenite
(Figure 3(d)). The microstructure of the FLN2-4405 steel in the as-sintered condition (Figures 4(a) and
4(c)) is comparable. The brownish etching areas consist of fine martensite. The fracture surfaces again
exhibit more than 50% cleavage facets [45]. The microstructure is composed primarily of divorced
pearlite with white etching nickel-rich areas. The quenched and tempered microstructure is composed of
90% tempered martensite (Figure 4(b)). The remaining 10% is composed of nickel-rich areas and
retained-austenite. The sinter-hardened microstructure consists of about 50% martensite, 40% pearlite
with the balance comprising nickel-rich areas and retained-austenite (Figure 4(d)).
As reported in the previous work [45], most of the pores (>60%) in the P/M steels investigated are in the
size range (by area) <100 µm2 which corresponds to pores < 11 µm dia. The frequency of pores in the 100
to 200 µm2 range (corresponding to 11 to 16 µm dia) is about 10% and is much less for larger pores. In
contrast, there are few pores > 1000 µm2 corresponding to a pore dia > 35 µm dia in the steels, which is
true primarily for the copper steel.
High temperature sintering is effective in increasing the circularity of the pores (Figure 5). The effect of
this is more dominant for pores which are near spherical initially with a circularity or shape factor of 0.8
to 0.9. For pores with a shape factor < 0.6, high temperature sintering is not as effective as it is for pores
with a shape factor of 0.9.
CONCLUSIONS
1. In the sintered, sinter hardened, and quenched and tempered conditions, FL-4405,
FLC2-4405 and FLN2-4405 P/M steels at 7.4 g/cm3 nominal density exhibit comparable fatigue
crack growth rates (1.1207E-4 to 3.0185E-4 mm/cycle) at a stress intensity range of 1000 MPa
(mm)1/2 .
2. Quenched and tempered microstructures resulted in the highest fatigue crack growth rates.
3. Sinter hardening of FLC2-4405 and FLN2-4405 lowered the fatigue crack growth rates.
4. High temperature sintering reduced the fatigue crack growth rates in FL-4405, but a reverse trend
was observed in FLC2-4405 and FLN2-4405.
AKNOWLEDGMENTS
Professor Saritas is indebted to the Hoeganaes Corporation for financial support during a sabbatical leave
(2000/2001) at Drexel University, Philadelphia, PA. The authors express their gratitude to Tom Murphy
for assistance with the pore and microstructural analysis.
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