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Powder injection moulding of premixed ferritic and austenitic stainless steel powders
Materials Science and Engineering A 528 (2011) 3480–3488
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Powder injection moulding of premixed ferritic and austenitic stainless steel powders M.E. Sotomayor, B. Levenfeld ∗ , A. Várez Materials Science and Engineering Department, Carlos III University of Madrid, Avda. Universidad 30, 28911 Leganés, Spain
a r t i c l e
i n f o
Article history: Received 26 October 2010 Received in revised form 13 January 2011 Accepted 14 January 2011 Available online 19 January 2011 Keywords: Duplex stainless steel Powder injection moulding Sintering Electron microscopy
a b s t r a c t In this experimental work, powder injection moulding (PIM) of premixed 316L and 430L gas-atomized powders was developed to obtain duplex stainless steels. A multicomponent binder constituted of high density polyethylene (HDPE) and paraffin wax (PW) in a volume ratio 50/50 was selected for the process. Feedstocks with powder loadings of 50, 65, 68 and 70 vol.% were prepared. Mixing experiments were carried out at 170 ◦ C according to differential scanning calorimetry (DSC) results. The rheological characterization of feedstocks allowed establishing different rheological parameters as power flow index (n) and activation energy (Ea ) in order to know their suitability for injection moulding. Critical powder volume concentration (CPVC) was determined by means of oil absorption method and a rheological model. The feedstock was injected at 170 ◦ C and three-point bending and tensile parts were obtained. Thermogravimetrical analysis (TGA) of binder allowed us to design the thermal debinding cycle and sintering was carried out in low vacuum at different temperatures. Finally, mechanical properties such as hardness and tensile strength were evaluated. © 2011 Elsevier B.V. All rights reserved.
1. Introduction Duplex stainless steels (DSS) are characterized by a structure consisting of approximately equal amounts of ferrite and austenite. The amount of each phase is obtained by simultaneous control of the chemical composition and heat treatment. Most alloys are designed to contain about equal amounts of each phase as the optimum phase balance. Duplex stainless steels have higher mechanical strength and improved resistance to stress corrosion cracking than austenitic grades. Their toughness is higher than that of the ferritic steels but slightly lower to that of the austenitic ones [1]. As a consequence of this specific combination of properties, duplex stainless steels are an exceptional alternative for several applications. They have found widespread use in the chemical and petrochemical industries and in several medical applications. DSS could be obtained through different ways: using prealloyed powders [2], mixing an austenitic powder with a ferritic stabilizing element powder that during sintering will diffuse in the austenitic phase and will cause its destabilization producing the biphasic microstructure [3,4], or by premixing ferritic and austenitic powders in the adequate proportion [5–12].
∗ Corresponding author. Tel.: +34 916249915; fax: +34 916249430. E-mail addresses: [email protected] (M.E. Sotomayor), [email protected] (B. Levenfeld), [email protected] (A. Várez). 0921-5093/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2011.01.038
Powder injection moulding allows processing of stainless steels powders through a profitable and advantageous method. This technique offers the ability of obtaining pieces with a high shape complexity [13]. This process begins with the selection of powder and binder components. High density polyethylene is a polymer commonly employed for plastic and powder injection moulding. In particular, we have successfully used a binder based on HDPE and paraffin wax as a second component for production of M2 and T15 high speed steels [14,15], bronze [16] and alumina [17] parts. PW is basically added to reduce the viscosity of binder. Usually, multicomponent binders constituted of polymers and other additives, are selected for PIM technology because they improve the debinding process: the binder removal occurs in a wider temperature range and in a more gradual way. The powder and binder are mixed together to obtain a mixture called feedstock. The use of low amounts of binder produces high viscosity feedstock making difficult the moulding process, while large amounts of binder provide low strength and may produce heterogeneous green parts. A successful formulation of a feedstock depends on several rheological considerations. The role of the binder is to impart an adequate flowability to the mixture. This resulting feedstock is injected in order to obtain the green body. After moulding, the binder is removed to obtain the brown part that retains the original shape, and finally it is sintered to shrink in an isotropic way till near full density. After sintering the component has excellent strength, with properties often superior to those available from other processing routes. This technique has been
M.E. Sotomayor et al. / Materials Science and Engineering A 528 (2011) 3480–3488 Table 1 Chemical composition of 316L and 430L gas-atomized powders provided by the supplier (wt.%). 316L Fe Balance
Cr 17.29
Ni 10.83
Mo 2.37
Mn 1.44
Si 0.65
C 0.022
P 0.023
S 0.006
430L Fe Balance
Cr 16.20
Mn 0.71
Si 0.75
C 0.026
P 0.029
S 0.008
employed in recent years for the production of stainless steels components with complex shapes [18–23]. The strength and ductility are similar than obtained in cast and wrought products. In this paper, powder injection moulding of premixed ferritic and austenitic powders to obtain duplex stainless steels was developed. 2. Experimental procedure The starting powder was prepared by premixing 316L and 430L gas-atomized powders in a volume ratio of 50/50 in a Turbula mixer at room temperature for 30 min. The chemical composition of stainless steel powders is summarized in Table 1. The near spherical morphology of the stainless steel powders is revealed in the micrographs of Fig. 1 obtained with a Philips XL 30 scanning electron microscope (SEM). Besides, particle size distributions obtained in a laser scattering particle analyzer Malvern Mastersizer 2000 are shown. A broad particle size distribution provides a higher packing density and as a consequence, a lower amount of binder would be necessary to produce the feedstock. The results obtained from particle size distributions together with density of powders measured in a Micromeritics Accupyc 1330 helium pycnometer are displayed in Table 2. It is extremely necessary to determine particle size and particle size distribution of powder because it strongly affects the different stages of the PIM process [24]. The multicomponent binder system employed in this work was composed of high density polyethylene (HDPE) and paraffin wax (PW) in a volume ratio of 50/50. Differential scanning calorimetry (DSC) analyses were conducted to determine the melting temperatures of binder components and these experiments were performed on a Perkin Elmer Diamond calorimeter with nitrogen as purge gas. In order to eliminate the thermal history samples were firstly heated from 20 to 160 ◦ C at 10 ◦ C/min and then cooled at the same rate. Subsequently, they were heated again at the same conditions and melting temperatures were determined. Thermogravimetric analyses (TGA) were carried out on a Perkin Elmer Pyris 1 thermogravimetric analyzer from 30 to 600 ◦ C at a heating rate of 10 ◦ C/min in order to determine decomposition temperatures of binder and pure components. Some characteristics of binder and pure components are shown in Table 3 (see Refs. [14,16] for further details). The powder-binder mixtures were carried out in a Haake Rheomix 252p mixer with a pair of rotor blades at 170 ◦ C and 40 r.p.m. for 45 min. Mixing temperature was selected taking into
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Table 3 Characteristics of binder and pure components (Tm and Td are melting temperature and decomposition temperature range, respectively). Material
Density (g/cm3 )
Tm (◦ C)
Td (◦ C)
HDPE PW Binder
0.96 0.91 0.93
130.8 53.1 122.4
100 95 320
account DSC experiments. The mixtures were prepared with different powder loadings: 50, 65, 68 and 70 vol.%. Torque evolution was recorded during mixing process. This parameter is proportional to the work required to mix the powder and binder, and the homogeneity of the mixtures is achieved when the torque reaches a steady state value () [25]. A Haake Rheomex CTW100p twin screw extruder was employed to produce enough amount of feedstock in order to be injected. The temperature profile along the barrel was 160/165/170 ◦ C and the extrusion process was conducted at 70 r.p.m. The feedstock was extruded three times to guarantee a high homogeneity, and it was granulated to feed up the injection moulding machine. In order to determine the rheological behaviour of feedstocks, a Haake Rheocap S20 capillary rheometer was employed. The dimensions of the die were 1 mm of diameter and a length of 30 mm in order to keep a L/D ratio of 30. The shear rate was chosen in a range from 100 to 10000 s−1 and a melting time of 10 min was employed for each test. The injection process was performed in an Arburg 220S 250-60 injection machine. The injection moulding parameters were optimized: the barrel temperature profile was adjusted from 155 to 170 ◦ C and mould temperature was 40 ◦ C. The injection pressure was 800 bar with a holding pressure profile in three steps till 25 bar. Three-point bending and tensile parts were obtained. Binder removal was carried out through a thermal debinding in a Goceram GC-DC-50 furnace. The thermal cycle was designed on the basis of the previous thermogravimetric study of the binder. In order to assure the complete removal of binder the carbon content of debound samples was determined using a LECO CS-200 instrument. XRD experiments were carried out in a Philips X’Pert difractometer using Cu K␣ radiation, a voltage of 40 kV and a current of 40 mA. Sintering behaviour of brown parts was studied in low vacuum (<10−2 mbar) at 1100, 1150, 1200, 1250 and 1300 ◦ C. The specimens were heated at a rate of 5 ◦ C/min and held at high temperature for 1 h. Densities of sintered parts were evaluated using Archimedes’ method, and the microstructure was studied by a scanning electron microscope equipped with SE and BSE detectors. A compositional analysis of the alloying elements in different phases was performed by means of energy dispersive analysis of X-rays (EDS). The sintered samples were previously electrochemically etched with KOH 3 M. The tensile tests of sintered parts were carried out in a Microtest testing machine at a strain rate of 1 mm/min, and HV30 hardness of the sintered samples was measured on a Galileo instrument. 3. Results 3.1. Feedstock preparation
Table 2 Characteristics of stainless steel powders. Powder
d90 (m)
d50 (m)
d10 (m)
Pycnometric density (g/cm3 )
316L 430L Premixed 50/50
26.3 16.4 21.3
11.8 8.9 10.2
5.3 4.4 4.9
7.94 7.70 7.82
Fig. 2 shows torque evolution for different feedstocks with different powder loadings. Firstly, the binder components were introduced into the chamber, and subsequently small portions of powder were added until the torque was stabilized. For each feedstock, several additions of powder can be observed with mixing time as an increase in torque value. After the last addition of pow-
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Fig. 1. Scanning electron micrographs and particle size distributions of 316L, 430L and premixed 50/50 powders.
der, the steady state value associated with homogenization of the mixtures is reached. In all the cases the mixing time to achieve the steady state was less than 30 min indicating the good homogeneity of these systems. However, it can be also observed that increasing the powder loading the steady state torque value () slightly increases indicating differences in viscosity of the mixtures. In the case of a powder loading of 70 vol.% the torque curve becomes very noisy and separation between powder and binder, which is detrimental for the PIM process, was discerned. Under these considerations, torque measurements of the mixtures indicated that the optimum amount of metal powder should be less than 70 vol.%. In order to determine the suitability of the formulations for PIM, the homogeneity of each feedstock was evaluated through density measurements of five different amounts of the same batch (Fig. 3). The low deviation from the average value (<0.5%) reveals the good homogeneity of the mixtures. On the other hand and as it was expected, the pycnometric density of feedstock increases as
powder loading increases. These values are also compared with the theoretical density, determined by the rule of mixtures. There is a good agreement between experimental and theoretical values that means that decomposition of paraffin wax did not occur during mixing process. 3.2. Rheological behaviour The rheological behaviour of feedstocks is crucial to evaluate the ability of mixtures to be injected. Fig. 4 shows the variation of the viscosity with the shear rate at 170 ◦ C for the binder and feedstocks with different powder loadings. This shear rate range was selected because similar shear rate values are reached during injection stage. In all the cases, the viscosity decreases as shear rate increases according to a pseudoplastic behaviour which is the most suitable for injection moulding process [26]. Besides, the viscosity of the mixture increases as the powder loading increases. Consid-
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Table 4 Power law index of feedstocks with different powder loadings at 170 ◦ C. Powder loading (vol.%) n r
50 0.72 0.9991
65 0.53 0.998
68 0.50 0.998
70 0.51 0.9990
ering the viscosity values recommended for PIM process, less than 1000 Pa s in the shear rate range between 100 and 1000 s−1 [27], all these feedstocks could be successfully injected. In order to investigate the behaviour of these feedstocks under injection moulding conditions, rheological parameters as power law index (n) and flow activation energy (Ea ) were calculated. Pseudoplastic fluid behaviour fits very well to Ostwald [28] and De Waele [29] power law: = k˙ n
Fig. 2. Torque evolution with mixing time for feedstocks with different powder loadings.
where is the shear stress, ˙ is the shear rate, k is a constant and n is the power law index. In the case of a pseudoplastic behaviour, n < 1. The value of n indicates the degree of sensitivity of viscosity against shear rate which is higher as n is lower. Power law index values for feedstocks with different powder loadings (Table 4) were determined from the slope of logarithmic plots of shear stress versus shear rate in the linear range. As it can be seen, n decreases as powder loading increases. These results are in good agreement with some studies with spherical metallic powder systems published previously by Herranz et al. [14], Aggarwal et al. [30] and Li et al. [31]. A lower flow index for a higher powder loading indicates that there is a better powder-binder ratio for PIM feedstock to get quick powder re-packing and binder molecule orientation during moulding [31]. In general, the influence of temperature on viscosity can be expressed according to an Arrhenius type equation: (T ) = 0 exp
Fig. 3. Theoretical and pycnometric densities of feedstocks with different powder loadings.
E a
RT
Ea is the flow activation energy, R is the gas constant, T is the temperature and 0 is the viscosity at a reference temperature, T0 . All the prepared feedstocks present similar behaviour and, as an example, viscosity values of feedstock with a powder loading of 68 vol.% at different temperatures are shown in Fig. 5(a). Viscosity values decreased as temperature was increased. Linear fits at a shear rate of 1000 s−1 are displayed in Fig. 5(b) and flow activation energy values were determined from the slope of the curves. The values of Ea for different analyzed feedstocks are shown in Table 5. The flow activation energy for the different feedstocks is quite similar, but slightly decreases when increasing the powder loading from 50 to 70 vol.% indicating that the viscosity is not so sensitive to temperature variation. The feedstock with a higher powder loading could be more appropriated to be injection moulded because fluctuations on temperature will not produce an important viscosity change [13]. However, considering that a powder loading of 70 vol.% produced a very noisy torque curve, the most suitable powder loading for injection moulding was 68 vol.%, and this feedstock was selected to go on with the process. This powder loading is even higher than those found in the literature for 316L stainless steel powder. For instance, Omar et al. [19] chose a powder loading of 63 vol.% with a binder constituted of polyethylene, paraffin wax and stearic acid, and Koseski et al. [32] Table 5 Ea values for feedstocks with different powder loadings at 1000 s−1 .
Fig. 4. Viscosity measurements for binder and feedstocks with different powder loadings at 170 ◦ C.
Powder loading (vol.%) Ea (kJ/mol) r
50 21.90 0.9996
65 18.18 0.9998
68 19.72 0.9991
70 15.47 0.9992
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Fig. 5. Viscosity measurements of feedstock with a powder loading of 68 vol.% at different temperatures (a) and linear fits for different powder loadings (b).
employed a 65 vol.% with a binder constituted of paraffin wax and polypropylene. The determination of critical powder volume concentration (CPVC) is important in order to establish the optimum amount of binder in the feedstock. CPVC was experimentally determined by means of oil absorption method and through Janardhana–Reddy theoretical model. The oil absorption method [33,34] has been employed to determine CPVC in the case of ceramic [35–39] and metallic [36] powders. This method consists in loading the mixer chamber with a premeasured volume of dry powder at room temperature with a constant rotation speed of rotor blades. Subsequently, the oil was added stepwise to the camera and the torque is recorded continuously as a function of time. With the addition of liquid, torque increases gradually until a maximum is reached. If the addition of oil continues the torque value decreases abruptly indicating that the critical powder volume concentration is reached. CPVC is defined by the following equation using powder volume Vp and liquid phase volume VL , at the point where torque is a maximum: CPVC =
Vp Vp + VL
The rotation speed was maintained at 50 r.p.m. and an aliquot of 2 ml of oil was added at regular intervals into the chamber. Fig. 6(a) shows the torque and temperature curves against mixing time of the stainless steel powder with oil additions. Initially, when an aliquot of a liquid is added to powder, the torque rises sharply due to the immediate mobilization of the liquid by the powder particles to form a few big clusters. Considerable stresses are therefore
required in order to break these clusters. Thus, during mixing at a constant shear rate, high torque values are encountered initially when only a few big clusters are present. As the addition of liquid continues, the mean equilibrium size of the clusters becomes bigger and bigger as more and more particles join together, until ultimately the whole mass becomes a coherent paste. Corresponding to this, the torque also continuously increases until it reaches a peak at the point of CPVC. From then on, addition of any further liquid serves only to dilate the solid structure and increase the interparticle distances [36]. Therefore, the mixing torque falls as shown in Fig. 6(a). In Fig. 6(b) torque values corresponding to increasing additions of oil are represented. In our case, the maximum is reached with the addition of 12 ml of oil. The temperature will not rise on subsequent oil addition, as there is no further decrease in surface energy, and the capillary forces, also begin to decrease as the excess liquid can only separate the particles. With any further oil addition the temperature falls as the surface forces become ineffective. According to this experimental method, the CPVC for this system is 72 vol.%. Several mathematical models have been proposed to predict the maximum volume fraction of the filler from rheological measurements, which is a very important value for the estimation of optimal filler loading. In the case of PIM, the Janardhana–Reddy model [40] has been successfully applied to evaluate the CPVC. The model equation is given as follows: b = (b )m + b (1 − (b )m ) where is the mixture viscosity; r is the relative viscosity; b is the binder viscosity; r = /b ; is the filler volume fraction; m
Fig. 6. (a) Torque and temperature variation with mixing time for different additions of oil. (b) Torque and temperature values for different oil volumes.
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Fig. 7. Linear fit for Janardhana–Reddy model at 100 s−1 and 170 ◦ C.
is the maximum filler loading; b is the binder volume fraction; (b )m is the critical binder volume fraction. In Fig. 7, the good agreement of experimental data to the Janardhana–Reddy model is presented. The data correspond to those obtained at 100 s−1 and 170 ◦ C. The fit linear coefficient was 0.997 and CPVC was calculated from the slope. A value of 72 vol.% was obtained, similar than previously obtained employing the experimental oil absorption method. Normally, the optimal solids content contains more binder than the critical content and, this optimal loading is approximately 2–5 vol.% lower than critical [26]. As a consequence, an optimum powder loading of 68 vol.% was determined to go on with the injection stage. 3.3. Debinding The binder is only a vehicle that will be burnt out in later steps, and therefore, it will not remain in the final product. The polymeric part was driven off by thermal debinding. The thermal cycle employed is shown in Fig. 8, and it was established on the basis of the previous thermogravimetric study of the binder. The cycle took place under argon atmosphere to prevent the oxidation of metallic particles. The introduction of an air flux during the isothermal step of higher temperature promoted the complete elimination of decomposition products. The first dwell at 190 ◦ C corresponds to paraffin wax degradation. In the next dwell at
450 ◦ C, decomposition of high density polyethylene takes place. The maximum temperature reached is slightly higher than the polyethylene degradation temperature to ensure the full binder removal. This optimized debinding cycle lasted only 4 h. In Fig. 9(a), a scanning electron micrograph of a brown part is shown and no remains of binder can be found. This good distribution of different sized metallic particles allows a good sintering activity. After debinding, free-defects parts were obtained and all of them preserved the shape of green bodies. XRD pattern of the debound part is shown in Fig. 9(b). Only the presence of diffraction peaks corresponding to ferritic and austenitic starting powders was detected. Diffraction maximums of metal oxides were not found; nevertheless small amounts of oxygen were detected on the surface by EDS. After debinding, carbon elemental analysis was carried out to ensure the full polymer elimination. A value of 0.03 wt.% was obtained which is in agreement with the carbon content of the starting powders. Usually, a previous dissolution of paraffin wax from green parts in a suitable solvent allows to reduce the total debinding time. The other constituent is removed later by thermal treatment, which can be carried out rapidly owing to the network of porosity created by the solvent extraction step. Fig. 10 shows the effects of time and temperature on leaching behaviour of paraffin wax in heptane. The amount of removed paraffin wax increases with the time for all temperatures. At 40 ◦ C, complete removal of paraffin wax is not achieved for a leaching time up to 600 min and, even at 50 ◦ C, only the 94% removal is obtained after this time. As it can be seen from the figure, the paraffin wax extraction rate decreases with leaching time indicating that initially removal rate is higher because solvent is in direct contact with wax. The extraction rate becomes smaller as leaching time increases owing to increased length of porous channels through which the polymer must diffuse out [41]. It is known that solvent debinding is a two stage process consisting of dissolution and diffusion. Initially, solvent dissolves one of the components, paraffin in our case, thus forming a porous surface. The solvent then infiltrates into the pores by capillary action. This is followed by diffusion of dissolved paraffin wax out of the green part. The process can be formulated as [42]: ln
1 F
=
De t2 +K 2
where F is the fraction of the remaining soluble paraffin, De is the interdiffusion coefficient of paraffin and solvent, t is time, is the effective length scale and K represents the change in the mechanism controlling the debinding behaviour. The effective length scale is defined as: =
Fig. 8. Thermal debinding cycle.
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HB(2L) 2[HB + B(2L) + (2L)H]
where 2L, H and B are the thickness, length and breadth of the component (in our case, = 1.16). If ln(1/F) is plotted against leaching time, two different slopes can be observed which are associated with dissolution and diffusion stages (Fig. 11). De of paraffin wax in heptane can be determined in both cases. Since diffusion is easier and the solvent is in direct contact with the polymer, the dissolution of the soluble component of binder system is the rate limiting step in the beginning of the debinding process over a leaching time of 250 min [43]. As the process proceeds, a longer diffusion distance through porous channels formed after initial debinding slows down the process, and diffusion becomes the rate determining step. Using the previous equation, De values were estimated to be 2.68 × 10−4 and 1.52 × 10−4 cm2 /s at 40 ◦ C for dissolution and diffusion controlled events, respectively. At this temperature paraffin
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Fig. 9. Scanning electron micrograph (a) and XRD pattern (b) of a debound part. Main diffraction peaks of Fe-␣ and Fe-␥ are indicated.
Fig. 10. Effects of time and temperature on efficiency of heptane leaching process for removing paraffin wax from three-point bending parts.
removal by the dissolution method is ∼1.8 times faster than the removal by diffusion. Moreover, at 50 ◦ C the values calculated were 8.00 × 10−4 and 1.46 × 10−4 cm2 /s for both stages and therefore, De is ∼5.5 times higher in dissolution stage than diffusion stage. As
Fig. 11. Variation of ln(1/F) with leaching time at 40 and 50 ◦ C.
a result, the dissolution stage is more efficient than diffusion one when the temperature increases. In order to determine the influence of the powder in the solvent debinding process, De of paraffin wax in heptane was determined in binder samples. At 40 ◦ C the values calculated in dissolution and diffusion stages were 7.06 × 10−4 and 2.38 × 10−4 cm2 /s, and at 50 ◦ C the values were 1.14 × 10−3 and 2.48 × 10−4 cm2 /s. Values obtained in both dissolution and diffusion stages were higher than values obtained for feedstock. In this way, powder makes difficult paraffin wax extraction. Due to the long time required for solvent debinding, the best choice to remove completely the binder from green parts is only the thermal debinding with a total duration of 4 h previously described. This time is considerably lower than others found in the literature for stainless steels, and which normally take place with a previous solvent debinding. For example, Gülsoy [44] employed a binder constituted of paraffin wax, polypropylene, carnauba wax and stearic acid and, a previous solvent debinding was carried out before a thermal one allowing to employ a total time of 10 h. Koseski et al. [32] employed a previous solvent debinding to remove a binder constituted of polypropylene and wax in a total time of 9 h. 3.4. Sintering Density and hardness reached for the samples after sintering at different temperatures are shown in Fig. 12. As it was expected, sin-
Fig. 12. Density and hardness values of duplex stainless steels sintered in low vacuum at different temperatures.
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Table 6 Tensile strength of sintered duplex stainless steels. Sintering temperature (◦ C) Tensile strength (MPa)
1100 503
1150 629
1200 736
1250 756
1300 682
through different phases. Dark “islands” with a chromium content nearly 20 wt.% are homogeneously distributed in a matrix with a 16 wt.%. Related to nickel amount, the “islands” and the matrix contain a 4 and a 7 wt.%, respectively. This fact clearly indicates the partial Ni-diffusion from austenite to ferrite particles during sintering. However, the duplex microstructure was retained. In general, a heat treatment together with a fast cooling is required after sintering to prevent precipitation of secondary phases as sigma phase [45,46]; however, in all the cases the presence of these phases was not detected. Mechanical properties of sintered samples were also investigated and the results are presented in Table 6. Tensile strength increased from 503 to 756 MPa as sintering temperature increased from 1100 to 1250 ◦ C. However, in the case of samples sintered at 1300 ◦ C, tensile strength decreased because the grain size was increased. These values are considerably higher than obtained through conventional pulvimetallurgy route (above 426 MPa) [12]. 4. Conclusions
Fig. 13. Photograph of green and sintered parts.
tered density was increased as the sintering temperature increased and the highest density obtained was reached even at the temperature of 1250 ◦ C. This density value was 97% of the theoretical one. The evolution of hardness with sintering temperature displays the same trend that density. Vickers hardness values increased from samples sintered at 1100 ◦ C till sintered at 1250 ◦ C. After this sintering temperature, hardness values remain almost constant at 280 HV30. These values are higher than the values obtained for ferritic and austenitic stainless steels processed by PIM. In Fig. 13 green and sintered parts can be seen. The volume contraction obtained was nearly 32% at sintering temperature of 1250 ◦ C. The microstructure of samples sintered at 1250 ◦ C is shown in Fig. 14. Besides to the presence of pores in black contrast, two different contrasts can be observed as a consequence of the presence of two phases with different compositions. EDS microanalysis allowed us to perform a compositional analysis of the alloying elements
The development of powder injection moulding process of duplex stainless steels employing premixed 316L and 430L powders was carried out. The binder used was a multicomponent system constituted by high density polyethylene and paraffin wax. Feedstocks with different powder loadings (50, 65, 68 and 70 vol.%) were prepared. All of them presented a pseudoplastic behaviour which is suitable for injection moulding. CPVC was determined by means of oil absorption method and through Janardhana–Reddy rheological model. In both cases, the value was 72 vol.%. The optimum powder loading was established as 68 vol.%. On the other hand, the elimination of paraffin wax from binder and feedstock in heptane was studied. Two stages, dissolution and diffusion, were clearly distinguished. The interdiffusion coefficients of paraffin in heptane were higher in the case of binder indicating that the presence of powder makes difficult the paraffin wax extraction. Finally, debinding was carried out through an optimized thermal cycle which lasted only 4 h and allowed to obtain freedefects brown parts. After sintering, some Ni diffusion from austenite to ferrite particles was observed, however, these samples displayed a duplex microstructure. Stainless steel parts sintered in low vacuum at 1250 ◦ C presented 97% of the theoretical density and a hardness of 280 HV30. Tensile strength reached 756 MPa which is even higher than P/M duplex stainless steels. Acknowledgements Authors thank the Spanish Agency CICYT (Project MAT200764486-C07 and MAT2010-19837-C06) and the regional Government (Project P2009/PPQ-1629) for financial support. References
Fig. 14. Scanning electron micrograph of a sample sintered in low vacuum at 1250 ◦ C for 1 h.
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Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Powder injection moulding of premixed ferritic and austenitic stainless steel powders M.E. Sotomayor, B. Levenfeld ∗ , A. Várez Materials Science and Engineering Department, Carlos III University of Madrid, Avda. Universidad 30, 28911 Leganés, Spain
a r t i c l e
i n f o
Article history: Received 26 October 2010 Received in revised form 13 January 2011 Accepted 14 January 2011 Available online 19 January 2011 Keywords: Duplex stainless steel Powder injection moulding Sintering Electron microscopy
a b s t r a c t In this experimental work, powder injection moulding (PIM) of premixed 316L and 430L gas-atomized powders was developed to obtain duplex stainless steels. A multicomponent binder constituted of high density polyethylene (HDPE) and paraffin wax (PW) in a volume ratio 50/50 was selected for the process. Feedstocks with powder loadings of 50, 65, 68 and 70 vol.% were prepared. Mixing experiments were carried out at 170 ◦ C according to differential scanning calorimetry (DSC) results. The rheological characterization of feedstocks allowed establishing different rheological parameters as power flow index (n) and activation energy (Ea ) in order to know their suitability for injection moulding. Critical powder volume concentration (CPVC) was determined by means of oil absorption method and a rheological model. The feedstock was injected at 170 ◦ C and three-point bending and tensile parts were obtained. Thermogravimetrical analysis (TGA) of binder allowed us to design the thermal debinding cycle and sintering was carried out in low vacuum at different temperatures. Finally, mechanical properties such as hardness and tensile strength were evaluated. © 2011 Elsevier B.V. All rights reserved.
1. Introduction Duplex stainless steels (DSS) are characterized by a structure consisting of approximately equal amounts of ferrite and austenite. The amount of each phase is obtained by simultaneous control of the chemical composition and heat treatment. Most alloys are designed to contain about equal amounts of each phase as the optimum phase balance. Duplex stainless steels have higher mechanical strength and improved resistance to stress corrosion cracking than austenitic grades. Their toughness is higher than that of the ferritic steels but slightly lower to that of the austenitic ones [1]. As a consequence of this specific combination of properties, duplex stainless steels are an exceptional alternative for several applications. They have found widespread use in the chemical and petrochemical industries and in several medical applications. DSS could be obtained through different ways: using prealloyed powders [2], mixing an austenitic powder with a ferritic stabilizing element powder that during sintering will diffuse in the austenitic phase and will cause its destabilization producing the biphasic microstructure [3,4], or by premixing ferritic and austenitic powders in the adequate proportion [5–12].
∗ Corresponding author. Tel.: +34 916249915; fax: +34 916249430. E-mail addresses: [email protected] (M.E. Sotomayor), [email protected] (B. Levenfeld), [email protected] (A. Várez). 0921-5093/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2011.01.038
Powder injection moulding allows processing of stainless steels powders through a profitable and advantageous method. This technique offers the ability of obtaining pieces with a high shape complexity [13]. This process begins with the selection of powder and binder components. High density polyethylene is a polymer commonly employed for plastic and powder injection moulding. In particular, we have successfully used a binder based on HDPE and paraffin wax as a second component for production of M2 and T15 high speed steels [14,15], bronze [16] and alumina [17] parts. PW is basically added to reduce the viscosity of binder. Usually, multicomponent binders constituted of polymers and other additives, are selected for PIM technology because they improve the debinding process: the binder removal occurs in a wider temperature range and in a more gradual way. The powder and binder are mixed together to obtain a mixture called feedstock. The use of low amounts of binder produces high viscosity feedstock making difficult the moulding process, while large amounts of binder provide low strength and may produce heterogeneous green parts. A successful formulation of a feedstock depends on several rheological considerations. The role of the binder is to impart an adequate flowability to the mixture. This resulting feedstock is injected in order to obtain the green body. After moulding, the binder is removed to obtain the brown part that retains the original shape, and finally it is sintered to shrink in an isotropic way till near full density. After sintering the component has excellent strength, with properties often superior to those available from other processing routes. This technique has been
M.E. Sotomayor et al. / Materials Science and Engineering A 528 (2011) 3480–3488 Table 1 Chemical composition of 316L and 430L gas-atomized powders provided by the supplier (wt.%). 316L Fe Balance
Cr 17.29
Ni 10.83
Mo 2.37
Mn 1.44
Si 0.65
C 0.022
P 0.023
S 0.006
430L Fe Balance
Cr 16.20
Mn 0.71
Si 0.75
C 0.026
P 0.029
S 0.008
employed in recent years for the production of stainless steels components with complex shapes [18–23]. The strength and ductility are similar than obtained in cast and wrought products. In this paper, powder injection moulding of premixed ferritic and austenitic powders to obtain duplex stainless steels was developed. 2. Experimental procedure The starting powder was prepared by premixing 316L and 430L gas-atomized powders in a volume ratio of 50/50 in a Turbula mixer at room temperature for 30 min. The chemical composition of stainless steel powders is summarized in Table 1. The near spherical morphology of the stainless steel powders is revealed in the micrographs of Fig. 1 obtained with a Philips XL 30 scanning electron microscope (SEM). Besides, particle size distributions obtained in a laser scattering particle analyzer Malvern Mastersizer 2000 are shown. A broad particle size distribution provides a higher packing density and as a consequence, a lower amount of binder would be necessary to produce the feedstock. The results obtained from particle size distributions together with density of powders measured in a Micromeritics Accupyc 1330 helium pycnometer are displayed in Table 2. It is extremely necessary to determine particle size and particle size distribution of powder because it strongly affects the different stages of the PIM process [24]. The multicomponent binder system employed in this work was composed of high density polyethylene (HDPE) and paraffin wax (PW) in a volume ratio of 50/50. Differential scanning calorimetry (DSC) analyses were conducted to determine the melting temperatures of binder components and these experiments were performed on a Perkin Elmer Diamond calorimeter with nitrogen as purge gas. In order to eliminate the thermal history samples were firstly heated from 20 to 160 ◦ C at 10 ◦ C/min and then cooled at the same rate. Subsequently, they were heated again at the same conditions and melting temperatures were determined. Thermogravimetric analyses (TGA) were carried out on a Perkin Elmer Pyris 1 thermogravimetric analyzer from 30 to 600 ◦ C at a heating rate of 10 ◦ C/min in order to determine decomposition temperatures of binder and pure components. Some characteristics of binder and pure components are shown in Table 3 (see Refs. [14,16] for further details). The powder-binder mixtures were carried out in a Haake Rheomix 252p mixer with a pair of rotor blades at 170 ◦ C and 40 r.p.m. for 45 min. Mixing temperature was selected taking into
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Table 3 Characteristics of binder and pure components (Tm and Td are melting temperature and decomposition temperature range, respectively). Material
Density (g/cm3 )
Tm (◦ C)
Td (◦ C)
HDPE PW Binder
0.96 0.91 0.93
130.8 53.1 122.4
100 95 320
account DSC experiments. The mixtures were prepared with different powder loadings: 50, 65, 68 and 70 vol.%. Torque evolution was recorded during mixing process. This parameter is proportional to the work required to mix the powder and binder, and the homogeneity of the mixtures is achieved when the torque reaches a steady state value () [25]. A Haake Rheomex CTW100p twin screw extruder was employed to produce enough amount of feedstock in order to be injected. The temperature profile along the barrel was 160/165/170 ◦ C and the extrusion process was conducted at 70 r.p.m. The feedstock was extruded three times to guarantee a high homogeneity, and it was granulated to feed up the injection moulding machine. In order to determine the rheological behaviour of feedstocks, a Haake Rheocap S20 capillary rheometer was employed. The dimensions of the die were 1 mm of diameter and a length of 30 mm in order to keep a L/D ratio of 30. The shear rate was chosen in a range from 100 to 10000 s−1 and a melting time of 10 min was employed for each test. The injection process was performed in an Arburg 220S 250-60 injection machine. The injection moulding parameters were optimized: the barrel temperature profile was adjusted from 155 to 170 ◦ C and mould temperature was 40 ◦ C. The injection pressure was 800 bar with a holding pressure profile in three steps till 25 bar. Three-point bending and tensile parts were obtained. Binder removal was carried out through a thermal debinding in a Goceram GC-DC-50 furnace. The thermal cycle was designed on the basis of the previous thermogravimetric study of the binder. In order to assure the complete removal of binder the carbon content of debound samples was determined using a LECO CS-200 instrument. XRD experiments were carried out in a Philips X’Pert difractometer using Cu K␣ radiation, a voltage of 40 kV and a current of 40 mA. Sintering behaviour of brown parts was studied in low vacuum (<10−2 mbar) at 1100, 1150, 1200, 1250 and 1300 ◦ C. The specimens were heated at a rate of 5 ◦ C/min and held at high temperature for 1 h. Densities of sintered parts were evaluated using Archimedes’ method, and the microstructure was studied by a scanning electron microscope equipped with SE and BSE detectors. A compositional analysis of the alloying elements in different phases was performed by means of energy dispersive analysis of X-rays (EDS). The sintered samples were previously electrochemically etched with KOH 3 M. The tensile tests of sintered parts were carried out in a Microtest testing machine at a strain rate of 1 mm/min, and HV30 hardness of the sintered samples was measured on a Galileo instrument. 3. Results 3.1. Feedstock preparation
Table 2 Characteristics of stainless steel powders. Powder
d90 (m)
d50 (m)
d10 (m)
Pycnometric density (g/cm3 )
316L 430L Premixed 50/50
26.3 16.4 21.3
11.8 8.9 10.2
5.3 4.4 4.9
7.94 7.70 7.82
Fig. 2 shows torque evolution for different feedstocks with different powder loadings. Firstly, the binder components were introduced into the chamber, and subsequently small portions of powder were added until the torque was stabilized. For each feedstock, several additions of powder can be observed with mixing time as an increase in torque value. After the last addition of pow-
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Fig. 1. Scanning electron micrographs and particle size distributions of 316L, 430L and premixed 50/50 powders.
der, the steady state value associated with homogenization of the mixtures is reached. In all the cases the mixing time to achieve the steady state was less than 30 min indicating the good homogeneity of these systems. However, it can be also observed that increasing the powder loading the steady state torque value () slightly increases indicating differences in viscosity of the mixtures. In the case of a powder loading of 70 vol.% the torque curve becomes very noisy and separation between powder and binder, which is detrimental for the PIM process, was discerned. Under these considerations, torque measurements of the mixtures indicated that the optimum amount of metal powder should be less than 70 vol.%. In order to determine the suitability of the formulations for PIM, the homogeneity of each feedstock was evaluated through density measurements of five different amounts of the same batch (Fig. 3). The low deviation from the average value (<0.5%) reveals the good homogeneity of the mixtures. On the other hand and as it was expected, the pycnometric density of feedstock increases as
powder loading increases. These values are also compared with the theoretical density, determined by the rule of mixtures. There is a good agreement between experimental and theoretical values that means that decomposition of paraffin wax did not occur during mixing process. 3.2. Rheological behaviour The rheological behaviour of feedstocks is crucial to evaluate the ability of mixtures to be injected. Fig. 4 shows the variation of the viscosity with the shear rate at 170 ◦ C for the binder and feedstocks with different powder loadings. This shear rate range was selected because similar shear rate values are reached during injection stage. In all the cases, the viscosity decreases as shear rate increases according to a pseudoplastic behaviour which is the most suitable for injection moulding process [26]. Besides, the viscosity of the mixture increases as the powder loading increases. Consid-
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Table 4 Power law index of feedstocks with different powder loadings at 170 ◦ C. Powder loading (vol.%) n r
50 0.72 0.9991
65 0.53 0.998
68 0.50 0.998
70 0.51 0.9990
ering the viscosity values recommended for PIM process, less than 1000 Pa s in the shear rate range between 100 and 1000 s−1 [27], all these feedstocks could be successfully injected. In order to investigate the behaviour of these feedstocks under injection moulding conditions, rheological parameters as power law index (n) and flow activation energy (Ea ) were calculated. Pseudoplastic fluid behaviour fits very well to Ostwald [28] and De Waele [29] power law: = k˙ n
Fig. 2. Torque evolution with mixing time for feedstocks with different powder loadings.
where is the shear stress, ˙ is the shear rate, k is a constant and n is the power law index. In the case of a pseudoplastic behaviour, n < 1. The value of n indicates the degree of sensitivity of viscosity against shear rate which is higher as n is lower. Power law index values for feedstocks with different powder loadings (Table 4) were determined from the slope of logarithmic plots of shear stress versus shear rate in the linear range. As it can be seen, n decreases as powder loading increases. These results are in good agreement with some studies with spherical metallic powder systems published previously by Herranz et al. [14], Aggarwal et al. [30] and Li et al. [31]. A lower flow index for a higher powder loading indicates that there is a better powder-binder ratio for PIM feedstock to get quick powder re-packing and binder molecule orientation during moulding [31]. In general, the influence of temperature on viscosity can be expressed according to an Arrhenius type equation: (T ) = 0 exp
Fig. 3. Theoretical and pycnometric densities of feedstocks with different powder loadings.
E a
RT
Ea is the flow activation energy, R is the gas constant, T is the temperature and 0 is the viscosity at a reference temperature, T0 . All the prepared feedstocks present similar behaviour and, as an example, viscosity values of feedstock with a powder loading of 68 vol.% at different temperatures are shown in Fig. 5(a). Viscosity values decreased as temperature was increased. Linear fits at a shear rate of 1000 s−1 are displayed in Fig. 5(b) and flow activation energy values were determined from the slope of the curves. The values of Ea for different analyzed feedstocks are shown in Table 5. The flow activation energy for the different feedstocks is quite similar, but slightly decreases when increasing the powder loading from 50 to 70 vol.% indicating that the viscosity is not so sensitive to temperature variation. The feedstock with a higher powder loading could be more appropriated to be injection moulded because fluctuations on temperature will not produce an important viscosity change [13]. However, considering that a powder loading of 70 vol.% produced a very noisy torque curve, the most suitable powder loading for injection moulding was 68 vol.%, and this feedstock was selected to go on with the process. This powder loading is even higher than those found in the literature for 316L stainless steel powder. For instance, Omar et al. [19] chose a powder loading of 63 vol.% with a binder constituted of polyethylene, paraffin wax and stearic acid, and Koseski et al. [32] Table 5 Ea values for feedstocks with different powder loadings at 1000 s−1 .
Fig. 4. Viscosity measurements for binder and feedstocks with different powder loadings at 170 ◦ C.
Powder loading (vol.%) Ea (kJ/mol) r
50 21.90 0.9996
65 18.18 0.9998
68 19.72 0.9991
70 15.47 0.9992
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Fig. 5. Viscosity measurements of feedstock with a powder loading of 68 vol.% at different temperatures (a) and linear fits for different powder loadings (b).
employed a 65 vol.% with a binder constituted of paraffin wax and polypropylene. The determination of critical powder volume concentration (CPVC) is important in order to establish the optimum amount of binder in the feedstock. CPVC was experimentally determined by means of oil absorption method and through Janardhana–Reddy theoretical model. The oil absorption method [33,34] has been employed to determine CPVC in the case of ceramic [35–39] and metallic [36] powders. This method consists in loading the mixer chamber with a premeasured volume of dry powder at room temperature with a constant rotation speed of rotor blades. Subsequently, the oil was added stepwise to the camera and the torque is recorded continuously as a function of time. With the addition of liquid, torque increases gradually until a maximum is reached. If the addition of oil continues the torque value decreases abruptly indicating that the critical powder volume concentration is reached. CPVC is defined by the following equation using powder volume Vp and liquid phase volume VL , at the point where torque is a maximum: CPVC =
Vp Vp + VL
The rotation speed was maintained at 50 r.p.m. and an aliquot of 2 ml of oil was added at regular intervals into the chamber. Fig. 6(a) shows the torque and temperature curves against mixing time of the stainless steel powder with oil additions. Initially, when an aliquot of a liquid is added to powder, the torque rises sharply due to the immediate mobilization of the liquid by the powder particles to form a few big clusters. Considerable stresses are therefore
required in order to break these clusters. Thus, during mixing at a constant shear rate, high torque values are encountered initially when only a few big clusters are present. As the addition of liquid continues, the mean equilibrium size of the clusters becomes bigger and bigger as more and more particles join together, until ultimately the whole mass becomes a coherent paste. Corresponding to this, the torque also continuously increases until it reaches a peak at the point of CPVC. From then on, addition of any further liquid serves only to dilate the solid structure and increase the interparticle distances [36]. Therefore, the mixing torque falls as shown in Fig. 6(a). In Fig. 6(b) torque values corresponding to increasing additions of oil are represented. In our case, the maximum is reached with the addition of 12 ml of oil. The temperature will not rise on subsequent oil addition, as there is no further decrease in surface energy, and the capillary forces, also begin to decrease as the excess liquid can only separate the particles. With any further oil addition the temperature falls as the surface forces become ineffective. According to this experimental method, the CPVC for this system is 72 vol.%. Several mathematical models have been proposed to predict the maximum volume fraction of the filler from rheological measurements, which is a very important value for the estimation of optimal filler loading. In the case of PIM, the Janardhana–Reddy model [40] has been successfully applied to evaluate the CPVC. The model equation is given as follows: b = (b )m + b (1 − (b )m ) where is the mixture viscosity; r is the relative viscosity; b is the binder viscosity; r = /b ; is the filler volume fraction; m
Fig. 6. (a) Torque and temperature variation with mixing time for different additions of oil. (b) Torque and temperature values for different oil volumes.
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Fig. 7. Linear fit for Janardhana–Reddy model at 100 s−1 and 170 ◦ C.
is the maximum filler loading; b is the binder volume fraction; (b )m is the critical binder volume fraction. In Fig. 7, the good agreement of experimental data to the Janardhana–Reddy model is presented. The data correspond to those obtained at 100 s−1 and 170 ◦ C. The fit linear coefficient was 0.997 and CPVC was calculated from the slope. A value of 72 vol.% was obtained, similar than previously obtained employing the experimental oil absorption method. Normally, the optimal solids content contains more binder than the critical content and, this optimal loading is approximately 2–5 vol.% lower than critical [26]. As a consequence, an optimum powder loading of 68 vol.% was determined to go on with the injection stage. 3.3. Debinding The binder is only a vehicle that will be burnt out in later steps, and therefore, it will not remain in the final product. The polymeric part was driven off by thermal debinding. The thermal cycle employed is shown in Fig. 8, and it was established on the basis of the previous thermogravimetric study of the binder. The cycle took place under argon atmosphere to prevent the oxidation of metallic particles. The introduction of an air flux during the isothermal step of higher temperature promoted the complete elimination of decomposition products. The first dwell at 190 ◦ C corresponds to paraffin wax degradation. In the next dwell at
450 ◦ C, decomposition of high density polyethylene takes place. The maximum temperature reached is slightly higher than the polyethylene degradation temperature to ensure the full binder removal. This optimized debinding cycle lasted only 4 h. In Fig. 9(a), a scanning electron micrograph of a brown part is shown and no remains of binder can be found. This good distribution of different sized metallic particles allows a good sintering activity. After debinding, free-defects parts were obtained and all of them preserved the shape of green bodies. XRD pattern of the debound part is shown in Fig. 9(b). Only the presence of diffraction peaks corresponding to ferritic and austenitic starting powders was detected. Diffraction maximums of metal oxides were not found; nevertheless small amounts of oxygen were detected on the surface by EDS. After debinding, carbon elemental analysis was carried out to ensure the full polymer elimination. A value of 0.03 wt.% was obtained which is in agreement with the carbon content of the starting powders. Usually, a previous dissolution of paraffin wax from green parts in a suitable solvent allows to reduce the total debinding time. The other constituent is removed later by thermal treatment, which can be carried out rapidly owing to the network of porosity created by the solvent extraction step. Fig. 10 shows the effects of time and temperature on leaching behaviour of paraffin wax in heptane. The amount of removed paraffin wax increases with the time for all temperatures. At 40 ◦ C, complete removal of paraffin wax is not achieved for a leaching time up to 600 min and, even at 50 ◦ C, only the 94% removal is obtained after this time. As it can be seen from the figure, the paraffin wax extraction rate decreases with leaching time indicating that initially removal rate is higher because solvent is in direct contact with wax. The extraction rate becomes smaller as leaching time increases owing to increased length of porous channels through which the polymer must diffuse out [41]. It is known that solvent debinding is a two stage process consisting of dissolution and diffusion. Initially, solvent dissolves one of the components, paraffin in our case, thus forming a porous surface. The solvent then infiltrates into the pores by capillary action. This is followed by diffusion of dissolved paraffin wax out of the green part. The process can be formulated as [42]: ln
1 F
=
De t2 +K 2
where F is the fraction of the remaining soluble paraffin, De is the interdiffusion coefficient of paraffin and solvent, t is time, is the effective length scale and K represents the change in the mechanism controlling the debinding behaviour. The effective length scale is defined as: =
Fig. 8. Thermal debinding cycle.
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HB(2L) 2[HB + B(2L) + (2L)H]
where 2L, H and B are the thickness, length and breadth of the component (in our case, = 1.16). If ln(1/F) is plotted against leaching time, two different slopes can be observed which are associated with dissolution and diffusion stages (Fig. 11). De of paraffin wax in heptane can be determined in both cases. Since diffusion is easier and the solvent is in direct contact with the polymer, the dissolution of the soluble component of binder system is the rate limiting step in the beginning of the debinding process over a leaching time of 250 min [43]. As the process proceeds, a longer diffusion distance through porous channels formed after initial debinding slows down the process, and diffusion becomes the rate determining step. Using the previous equation, De values were estimated to be 2.68 × 10−4 and 1.52 × 10−4 cm2 /s at 40 ◦ C for dissolution and diffusion controlled events, respectively. At this temperature paraffin
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Fig. 9. Scanning electron micrograph (a) and XRD pattern (b) of a debound part. Main diffraction peaks of Fe-␣ and Fe-␥ are indicated.
Fig. 10. Effects of time and temperature on efficiency of heptane leaching process for removing paraffin wax from three-point bending parts.
removal by the dissolution method is ∼1.8 times faster than the removal by diffusion. Moreover, at 50 ◦ C the values calculated were 8.00 × 10−4 and 1.46 × 10−4 cm2 /s for both stages and therefore, De is ∼5.5 times higher in dissolution stage than diffusion stage. As
Fig. 11. Variation of ln(1/F) with leaching time at 40 and 50 ◦ C.
a result, the dissolution stage is more efficient than diffusion one when the temperature increases. In order to determine the influence of the powder in the solvent debinding process, De of paraffin wax in heptane was determined in binder samples. At 40 ◦ C the values calculated in dissolution and diffusion stages were 7.06 × 10−4 and 2.38 × 10−4 cm2 /s, and at 50 ◦ C the values were 1.14 × 10−3 and 2.48 × 10−4 cm2 /s. Values obtained in both dissolution and diffusion stages were higher than values obtained for feedstock. In this way, powder makes difficult paraffin wax extraction. Due to the long time required for solvent debinding, the best choice to remove completely the binder from green parts is only the thermal debinding with a total duration of 4 h previously described. This time is considerably lower than others found in the literature for stainless steels, and which normally take place with a previous solvent debinding. For example, Gülsoy [44] employed a binder constituted of paraffin wax, polypropylene, carnauba wax and stearic acid and, a previous solvent debinding was carried out before a thermal one allowing to employ a total time of 10 h. Koseski et al. [32] employed a previous solvent debinding to remove a binder constituted of polypropylene and wax in a total time of 9 h. 3.4. Sintering Density and hardness reached for the samples after sintering at different temperatures are shown in Fig. 12. As it was expected, sin-
Fig. 12. Density and hardness values of duplex stainless steels sintered in low vacuum at different temperatures.
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Table 6 Tensile strength of sintered duplex stainless steels. Sintering temperature (◦ C) Tensile strength (MPa)
1100 503
1150 629
1200 736
1250 756
1300 682
through different phases. Dark “islands” with a chromium content nearly 20 wt.% are homogeneously distributed in a matrix with a 16 wt.%. Related to nickel amount, the “islands” and the matrix contain a 4 and a 7 wt.%, respectively. This fact clearly indicates the partial Ni-diffusion from austenite to ferrite particles during sintering. However, the duplex microstructure was retained. In general, a heat treatment together with a fast cooling is required after sintering to prevent precipitation of secondary phases as sigma phase [45,46]; however, in all the cases the presence of these phases was not detected. Mechanical properties of sintered samples were also investigated and the results are presented in Table 6. Tensile strength increased from 503 to 756 MPa as sintering temperature increased from 1100 to 1250 ◦ C. However, in the case of samples sintered at 1300 ◦ C, tensile strength decreased because the grain size was increased. These values are considerably higher than obtained through conventional pulvimetallurgy route (above 426 MPa) [12]. 4. Conclusions
Fig. 13. Photograph of green and sintered parts.
tered density was increased as the sintering temperature increased and the highest density obtained was reached even at the temperature of 1250 ◦ C. This density value was 97% of the theoretical one. The evolution of hardness with sintering temperature displays the same trend that density. Vickers hardness values increased from samples sintered at 1100 ◦ C till sintered at 1250 ◦ C. After this sintering temperature, hardness values remain almost constant at 280 HV30. These values are higher than the values obtained for ferritic and austenitic stainless steels processed by PIM. In Fig. 13 green and sintered parts can be seen. The volume contraction obtained was nearly 32% at sintering temperature of 1250 ◦ C. The microstructure of samples sintered at 1250 ◦ C is shown in Fig. 14. Besides to the presence of pores in black contrast, two different contrasts can be observed as a consequence of the presence of two phases with different compositions. EDS microanalysis allowed us to perform a compositional analysis of the alloying elements
The development of powder injection moulding process of duplex stainless steels employing premixed 316L and 430L powders was carried out. The binder used was a multicomponent system constituted by high density polyethylene and paraffin wax. Feedstocks with different powder loadings (50, 65, 68 and 70 vol.%) were prepared. All of them presented a pseudoplastic behaviour which is suitable for injection moulding. CPVC was determined by means of oil absorption method and through Janardhana–Reddy rheological model. In both cases, the value was 72 vol.%. The optimum powder loading was established as 68 vol.%. On the other hand, the elimination of paraffin wax from binder and feedstock in heptane was studied. Two stages, dissolution and diffusion, were clearly distinguished. The interdiffusion coefficients of paraffin in heptane were higher in the case of binder indicating that the presence of powder makes difficult the paraffin wax extraction. Finally, debinding was carried out through an optimized thermal cycle which lasted only 4 h and allowed to obtain freedefects brown parts. After sintering, some Ni diffusion from austenite to ferrite particles was observed, however, these samples displayed a duplex microstructure. Stainless steel parts sintered in low vacuum at 1250 ◦ C presented 97% of the theoretical density and a hardness of 280 HV30. Tensile strength reached 756 MPa which is even higher than P/M duplex stainless steels. Acknowledgements Authors thank the Spanish Agency CICYT (Project MAT200764486-C07 and MAT2010-19837-C06) and the regional Government (Project P2009/PPQ-1629) for financial support. References
Fig. 14. Scanning electron micrograph of a sample sintered in low vacuum at 1250 ◦ C for 1 h.
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