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Investigation of the laser–powder–atmosphere interaction zone during the selective laser melting process
Accepted Manuscript Title: Investigation of the laser-powder-atmosphere interaction zone during the selective laser melting process Author: Amal Masmoudi Rodolphe Bolot Christian Coddet PII: DOI: Reference:
S0924-0136(15)00208-3 http://dx.doi.org/doi:10.1016/j.jmatprotec.2015.05.008 PROTEC 14413
To appear in:
Journal of Materials Processing Technology
Received date: Revised date: Accepted date:
20-2-2015 14-4-2015 11-5-2015
Please cite this article as: MASMOUDI, A., BOLOT, R., CODDET, C.,Investigation of the laser-powder-atmosphere interaction zone during the selective laser melting process, Journal of Materials Processing Technology (2015), http://dx.doi.org/10.1016/j.jmatprotec.2015.05.008 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Investigation of the laser-powder-atmosphere interaction zone during the selective laser melting process
LERMPS -University of Technology Belfort – Montbeliard, Sevenans, 90010 Belfort, France
*[email protected]
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Abstract
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a
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Amal MASMOUDIa,* , Rodolphe BOLOTa and Christian CODDETa
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Selective Laser Melting (SLM) consists in manufacturing parts in a chamber filled with a protective gas, by melting successive thin powder layers using a laser. This study
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aims, at first, to understand the laser–material–atmosphere interactions that involve various complex and simultaneous physical phenomena such as heat transfer in the
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target, melting and vaporization of the material, and expansion of the generated vapor.
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A modelling approach of the process was developed to investigate the effect of some process parameters, such as the exposure time (texpo) of the laser beam on each point
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along its trajectory, the pressure of the protective gas (Argon) and the spatial thermal energy distribution on the dimensional characteristics of the melted track. Keywords: Laser melting process, Rapid prototyping, Additive manufacturing, Numerical modelling, CFD model, Metallic vapour. 1. Introduction
Additive manufacturing processes, also known as rapid manufacturing or rapid prototyping (Edson et al., 2006) allow manufacturing complex shape parts directly from 3D CAD data. Among those processes, Selective Laser Melting (SLM) was developed
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to produce fully functional metal parts presenting high mechanical properties directly from powder (Yadroitsev et al., 2011). It consists in manufacturing parts layer-by-layer using a local melting of a powder bed provided by laser irradiation, followed by
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immediate consolidation (Yadroitsev et al., 2006). The process is carried out in a
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chamber filled with a protective gas such as Argon to avoid the metal oxidation. SLM is a complex process involving several physical and chemical phenomena
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especially during the laser-powder interaction and it is yet not sufficiently understood. Thus, further researches are required and several numerical simulations are developed
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particularly focusing on the parts properties and on the laser-powder interaction zone under various processing parameters (Ehsan et al., 2011). A selection of those
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researches is briefly discussed in the following section.
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Gusarov et al. (2009a, 2009b, 2010) studied the phenomena related to the laser-powder
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interaction by modelling the penetration of heat and radiations in the powder bed by taking multiple reflections into account. They numerically calculated the temperature
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distribution in the laser-powder interaction region and investigated the influence of the powder layer thickness and of the laser power on the melt pool shape. They also studied the effects of the scanning velocity. Hussein et al. (2013) developed a three-dimensional model to predict the temperature distribution and stress field inside a single metallic layer formed on the powder bed without support. The simulation results revealed the magnitude of the temperature and stresses at different locations of the layer. The temperature gradients were the highest at the start of the first track scan and drops subsequently for all scanning speeds.
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Verhaeghe et al. (2009) presented a model allowing determining the effect of evaporation. They found that evaporation becomes a very significant phenomenon at high beam powers and energy densities. They also investigated thermal gradients and
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fluid dynamics within the melt pool.
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Donghua et al. (2014) developed a numerical simulation regarding the influence of the energy density of the laser beam on the melt pool dynamics and densification
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mechanisms.
Khairallah et al. (2014) developed a 3D mesoscopic simulation model to study the laser-
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induced melting of a random bed of stainless steel 316L particles on a solid substrate, and the solidification of the melt pool into either a continuous or a discontinuous track
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as a result of the Plateau–Rayleigh instability. Their study demonstrates the importance
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of simulation as a tool to determine the underlying physics governing the SLM process.
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Moreover, simulation may be helpful to study the interplay between the build parameters in order to optimize the process.
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Kurian et al. (2014) presented a numerical investigation to understand the effect of some process parameters during the laser melting of grade 316L stainless steel powder on an AISI 316L substrate using a pulsed Nd-YAG laser. They presented the temperature distribution in a single powder layer. They studied the effect of some process parameters such as the laser power, the scanning speed and the size of the beam on the dimensional characteristics of the melt region from the temperature solution. The results confirmed found that the laser power and scanning speed of the laser beam have a noticeable effect on the track characteristics.
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Nevertheless, few studies are concerned with the role of the atmosphere on the process and parts properties. Zhang et al. (2013a) studied pure Ti parts produced under vacuum and Zhang et al. (2013b) analyzed the influence of the protective gas nature on 316L
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parts. They showed that a decrease of the scan speed may provide a denser material
under vacuum conditions. However, a decrease of the chamber pressure gave also rise
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to a lower cross-section of the track, resulting certainly from an increase of the
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vaporization level.
Therefore, a 3D numerical model was developed in the present work to study the laser-
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powder-atmosphere interaction under different conditions. This model is able to consider the fundamental parameters involved in the SLM process such as the laser
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beam properties, the thermal energy distribution, the powder layer characteristics (material nature, particle size distribution, thickness, etc.), the nature and pressure of the
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environment gas, and other process parameters. This model is applied to investigate the
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material behavior.
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thermal and dimensional characteristics of the melt pool zone and the evaporated
Furthermore, the results obtained by numerical simulation are used to understand the effective properties of SLM parts manufactured during experimental works, in order to check the accuracy of the developed simulation model. 2. Modelling approach
A numerical simulation of the selective laser melting process was developed to study the laser-powder-atmosphere interaction by considering the different process parameters.
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In practice, a single line (Fig. 1c) scanned by the laser beam on a thin (50 μm) powder layer placed over a dense substrate (1 mm) of the same material (i.e., 316L stainless steel) and exposed to a protective Argon atmosphere was simulated using FLUENT
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software as shown in Fig. 1.
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The mesh size was adapted (Fig. 1b) to provide thin cells in the interaction zone of the
laser with the powder bed. On the contrary the mesh size was increased in other regions
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in order to reduce the computational time.
Fig. 1 Modelling steps of a 316L stainless steel powder bed irradiation by a laser beam
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(a) different regions of the calculation domain, (b) mesh, (c) laser irradiation of the powder bed.
2.1. Governing equations
The transient heat conduction equation was considered in solid regions as: h . T x, y, z t
(W .m 3 )
(1)
in which stands for density (kg.m-3), h for enthalpy (J.kg-1), for thermal conductivity (W.m-1.K-1), T for temperature (K) and represents the volumic heat source corresponding to the laser.
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Additionally, transient convective movements in the Argon atmosphere were taken into account by considering the Navier-Stokes equations as: -
Continuity:
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.v 0 t
(2)
Momentum:
v . v v p . g t
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-
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in which v stands for the velocity vector.
(3)
-
Energy:
Dp Dh . T v Dt Dt
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in which p is the pressure, is the stress tensor and g is the gravitational acceleration.
(4)
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dissipation function.
t
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Dt
d
with by definition De e v. e whatever e, and v corresponds to the viscous
2.2. Boundary and initial conditions A uniform temperature throughout the powder bed was considered before the laser irradiation:
T (t 0) T0 288 K
(5)
in which T0 is the ambient temperature. SLM layers are then built over a thick solidified material block (i.e., made of powder melted and solidified) presenting thermo-physical properties of the dense material (as first approximation). 6
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The thermal flux on the opposite face of the material block (the rear face of the plate) was neglected, corresponding to an adiabatic condition. This assumption was made due to the relatively high thickness of the block and to the fast laser irradiation that do not
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allow heating the rear face up to a significant value. Concerning the limits of the domain, the sides and the top face of the fluid domain were considered as "free
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surfaces" at which only the pressure was fixed (i.e., a relative pressure of 0 Pa).
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The laser source of the SLM machine considered in this study (from Realizer) is an YLR-100-SM single mode CW Ytterbium fiber laser (1064–1100 nm). This laser is
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used to melt the powder bed layer-by-layer (Ruidi et al., 2009). According to the calculation domain including both solid and fluid regions, the heating of the powder bed
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by the laser takes place at the gas/solid interface (i.e., corresponding to an internal face of the calculation domain). In practice, this fluid/solid interface was represented by the
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way of a “coupled wall”, meaning that the continuity of the thermal flux on both sides
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of this interface is considered. In order to keep this possibility, the laser heat source was
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thus considered as a volumic heat source concentrated in the first layer of the mesh in the solid side. For this, a Gaussian distribution of the power density is assumed (Fig. 2). It is the most common model of the power radial distribution (Mumtez et al., 2010) and is defined by:
r2 ( r ) 0 exp 2 2
(6)
in which 0 P2 , P is the power of the laser beam which was fixed during the 2 simulation (100 W), α is the absorption coefficient of dense 316L stainless steel and
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d where d is the apparent spot diameter of the laser beam (d = 68 µm so that σ = 4
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17 µm).
Fig. 2 Gaussian distribution of the thermal flux
(7)
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2r 2 (r ) 0 exp 2
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defined as:
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In some other studies (Hussein et al., 2013), the Gaussian distribution is sometimes
in which 2
According to Gusarov et al. (2009b), the penetration of the laser energy into the powder bed involves multi-reflection phenomena, providing a much higher effective absorption rate (about 0.8 for 316L) than the absorptivity of a flat dense surface (about 0.3 for 316L). This model would be suitable for the case of a laser beam impacting on loose particles without melting. However, after our results and considering our parameters, the laser impact point during the SLM process is mainly positioned on the melt pool
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(and not on unmelted particles). Thus, in the present study, the effective absorption coefficient of the liquid melt pool was considered as α = 0.3 for 316L stainless steel
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(Rai1 et al. 2007). 2.3. Effective properties of the solid materials
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Basic properties of 316L stainless steel such as melting point (about 1750 K) and boiling point (3090 K) were taken into account during the numerical simulations
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(Cédric, 2007; IAEA, 2008).
During the SLM process, the powder bed is molten and rapidly solidified into a dense
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solid state. Thus, the latent heat of fusion (270 kJ/kg) was considered and equations shown in Table 1 were used to define the 316L stainless steel properties in the liquid
Heat capacity -1
-1
Dense solid state
Powder state
462 + 0.134 T
462 + 0.134 T
7433 + 0.0393× T -
8084 + 0.4209 × T -
8084 + 0.4209 × T -
-4
-5
755
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Сp (J.kg .K )
Liquid state
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Properties of 316L
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Table 1. Thermo-physical properties of 316L
Density ρ (kg.m-3)
Thermal conductivity λ (W.m-1.K-1)
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and solid (dense and powder) states (IAEA, 2008).
1.801.10 × T
2
12.41 + 0.003279 × T
3.894. 10 × T
2
9.248 + 0.01571 × T
3.894.10-5 × T 2 (1 )
9.248 + 0.01571 × T (1 - )
in which ε = 0.44 represents the porosity level of the powder bed and T is the temperature in [K].
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Many correlations have been proposed in the literature in order to estimate the thermal conductivity of powder materials (Filali, 2006; Deissler, 1952). However, these models give rise to highly variable results.
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In the present study, an experimental procedure was developed to estimate the thermal conductivity of powder materials. Finally, a corrective factor of 0.56 (i.e.,
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corresponding to 1-) was deduced and the thermal conductivity was defined as shown
2.4. Modelling of the surrounding atmosphere
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in Table 1. Details concerning this method shall be given in another specific paper.
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Processing with the laser beam was carried out in a closed chamber filled with a protective gas (Argon) to avoid oxidation of the metallic material.
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During this work, the Argon environment was considered as a domain of compressible ideal gas. Moreover, the Lennard-Jones potential was considered to describe
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interactions between the gas molecules (for estimation of the transport properties). This
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potential is the most used for simple gases and is expressed as (Cuadros et al., 1996):
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12 6 E ( r ) 4 r r
(8)
in which is the depth of the potential energy well and is the distance at which the potential is zero. Table 2 lists the Lennard-Jones parameters for Iron (Cuadros et al., 1996) and Argon (Hirschfelder et al., 1954) taken from the literature.
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Table 2. Lennard-Jones parameters for Fe and Ar
(Angstrom Ǻ)
/κ (K)
Fe
2.321
6104
Ar
3.418
124
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Element
During the laser-powder interaction, the thermal power transferred is high enough to
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provide a melting of the powder bed in the region directly exposed to the laser beam.
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Depending on the scanning conditions, the boiling point may be reached, thus providing a partial vaporization of the material. So, the latent heat of evaporation of 316L stainless
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steel (745 kJ/kg) was considered (IAEA, 2008). During each numerical time step, the fraction of evaporation of cells at the material/Argon interface is calculated, providing a
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mass source to the continuity equation in the gas side. The expansion of the metallic
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vapor in the Argon atmosphere depending on the pressure is thus investigated.
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3. Experimental procedure
A specific selective laser melting machine developed at LERMPS laboratory allowing working under vacuum or in a controlled atmosphere was used in this work (Fig. 3). A special feature of this machine is its ability to vary the Argon gas pressure between 10 and 105 Pa. It is equipped with an Nd-YAG fiber laser with a maximum power of 120 W.
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Fig. 3 Vacuum selective laser melting machine developed at LERMPS
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laboratory
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The scanning speed of the laser is controlled thanks to oscillating mirrors driven by stepper motors. In that case, the laser movement on the powder bed is discontinuous: the laser stops during a short exposure time (texpo) on a point and then moves to the next point at a distance Pdist along the trajectory. Thus, the effective scanning speed depends on the two parameters texpo and Pdist according to the following expression:
veff
Pdist
(9)
Pdist t exp o v1
in which ν1 is a characteristic velocity depending on the stepper motors.
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This characteristic velocity corresponds to the maximum velocity possible with the stepper motors and can be estimated by considering a very large point to point distance. In our case, this characteristic velocity was thus estimated to 2.5 m/s. For industrial
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applications, this velocity is sometimes considered as sufficiently high to neglect its effect on the average scanning speed, which is thus sometimes defined as Pdist/texpo.
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A two parameter DOE (Design Of Experiments) was used to quantify the effect of the
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effective scan speed (changed with texpo), and that of the chamber pressure, on the
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characteristics of single tracks (Fig. 4 and Tables 3 and 4).
Fig. 4 316L stainless steel tracks manufactured with several scanning speeds (1 to 4) at several pressure levels (A to F) of Argon gas
Tables 3 and 4 present the four values of the effective scan speed and six values of the chamber pressure that were considered for this DOE. Three different tracks were made for each parameter set (Fig. 4). The average of each result such as line width and depth was considered for comparisons with the
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corresponding simulation. The process parameters were of course those used for the numerical modelling (laser power of 100 W, spot diameter of 68 µm, powder bed thickness of 50 µm and distance of 40 µm between two successive points).
1
400
40
2
200
40
3
120
40
4
100
40
Scan speed veff (m/s)
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Pdist (µm)
0.1 0.2
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texpo (µs)
0.3 0.4
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Parameter No.
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Table 3. Scan speeds applied to elaborate the single lines
A
B
C
Pressure (mbar)
995
500
250
D
E
F
100
10
1
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Parts
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Table 4. Argon gas pressure selected to elaborate the single lines
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4. Experimental results and discussion
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Fig. 5 displays more precisely the effect of the Argon gas pressure on the tracks obtained with an exposure time of 400 µs and a distance of 40 µm between two successive points (i.e., an effective scan speed of 0.1 m/s). At high Argon pressure, a continuous solid track is formed following the powder melting and solidification (Fig. 5 (a – d)). At low pressures (i.e., below 100 mbar), the solid track disappears and only a slight remelting of the solid surface just under the powder bed is observed. In this situation, the difference between the melting and vaporization temperatures becomes very low and it may thus be supposed that the liquid phase is vaporized as soon as it is formed, thus explaining the absence of any solid track
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on the surface. For still lower pressures, the liquid phase would not even be formed and a sublimation of the solid phase could take place (i.e., direct vaporization of the solid
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particles).
Fig. 5 Effect of the Argon gas pressure on SS 316L tracks prepared by the SLM process with veff = 0.1 m/s (a) 995 mbar (b) 500 mbar (c) 250 mbar (d) 100 mbar (e) 10 mbar (f) 1 mbar
In order to confirm these assumptions, the Clausius-Clapeyron equation (Takamichi et al., 1993) provides the evolution of the vaporization temperature versus pressure as: d( ln p) ΔH vap d(1/T) R
(10)
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in which R = 8.3145 J.mol-1.K-1 is the universal constant of gases and ΔHvap is the enthalpy of evaporation of the liquid metal. More particularly, the vapor pressure of liquid 316L steel may be estimated from
18868 T
cr
log10 p 11.1183
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(IAEA, 2008):
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with p (Pa) and T (K).
(11)
18868 11.1183 log10 ( p )
(12)
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Tv
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The change of the vaporization temperature with the chamber pressure is thus given by:
Based on this equation, the equilibrium temperature values at the different Argon-gas
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pressure levels considered in the DOE are given in Table 5.
Parameters
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Table 5. Vaporization temperature (Tv) of 316L for different pressure levels values
2000
1000
500
250
100
10
1
Tv (K)
3243
3084
2940
2807
2650
1865
1697
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p (mbar)
Tv represents the temperature at which gas and liquid phases are in equilibrium. According to Table 5, the vaporization temperature decreases strongly with the lowering of the pressure. One may also notice that the conventional boiling temperature of stainless steel (i.e., about 3093 K) is well predicted at 1 bar, and that there is no more difference between the melting and boiling points (about 1750 K) at 2.2 Pa. Below this pressure, a sublimation of the material occurs.
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5. Numerical results 5.1. Temperature distribution
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As mentioned previously, many numerical researches were developed to study the effect of the processing parameters on the temperature distribution of the powder bed. For
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example, according to Kurian et al. (2014), the temperature of the melt pool rises
rapidly and reaches about 3500 K for a 316L stainless steel powder bed with a high
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energy density and /or a low scan speed. However, in most cases, the environment properties (nature and pressure of the gas) within the chamber of the SLM machine
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were not considered.
After what was said above, the maximum temperature of the molten material should not
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exceed the vaporization temperature that depends on the liquid / gas equilibrium (see
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equation 12).
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One may hence assume that a decrease of the environment pressure should imply a decrease of the laser power or an increase of the scan speed in order to limit the material
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vaporization. Nevertheless, it becomes very sensitive at low pressure levels due to the decrease of the difference between the melting and vaporization temperatures. In order to take this phenomenon into account, the temperature of the molten pool was thus limited in the model by the equilibrium vaporization temperature provided by equation 12. In counterpart, the additional heat transferred once the vaporization temperature is reached contributes to vaporize a fraction of the material. Within the program, lines of C coding were written in order to represent that:
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The effect of vaporization is first omitted in the thermodynamic properties of the material. Then after each time step, the mass fraction of the vaporized material is calculated from the temperature excess beyond the vaporization point as: c p (T Tv ) H vap
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y
(13)
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Once this fraction is calculated, the local temperature T in the cell is reset to Tv for cells
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with a higher temperature. This method allows considering the vaporization
phenomenon by introducing a mass source term to the continuity equation in the gas
y mcell dt
(14)
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side. At each time step dt, this mass source term is set as:
in which y is the mass fraction of the liquid cell vaporized, mcell is the mass of material
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within this cell (i.e..Vcell) and dt corresponds to the time step.
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Fig. 6 displays the temperature distribution in the scanned zone for an effective scan speed of 0.1 m.s-1 (i.e., texpo = 400 µs and Pdist = 40 µm) in the case of a pressure of
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1 bar, a laser power of 100 W and a laser spot diameter of 68 µm. The laser trace on the powder bed is clearly observed and the maximum temperature is limited by the calculated vaporization temperature.
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Fig. 6 Temperature distribution (K) for a single track scanned linearly (veff = 0.1 m/s, p = 1 bar)
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For these process conditions, the temperature range of the scanned zone is comprised between 288 K (the initial temperature) and the vaporization temperature of about
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3080 K. The temperature of the powder bed in the region directly exposed to the laser
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beam largely exceeds the melting temperature of 316L stainless steel (≈1700 K) causing
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the formation of a molten pool. From this molten pool, the heat propagates in the surrounding loose powder. The power of the source is also high enough to cause a local evaporation of the material. A fast cooling down of the molten zone is then observed when the laser beam leaves the zone so that temperatures below the melting point (solid phase) are quickly obtained behind the hot spot (Fig. 6). 5.2. Melt pool dimensions and shape The morphology of the melt pool region predicted with this model can directly be compared to that of single tracks manufactured (see Fig. 4 and Fig. 5).The melt pool width and depth penetration are the most important characteristics to obtain fully dense 19
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parts. In particular, the depth of penetration and the width of the track have to be sufficient to allow overlapping and welding of the successive layers and tracks (Stoffregen et al., 2011).
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Fig. 7 and Fig. 8 show respectively the morphology of the melt pool (thanks to the
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temperature field) predicted by the simulation and that of the track obtained
experimentally after a single scan on the powder bed. In the simulation, the melt pool
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dimensions were determined from the temperature distribution along the laser path (Fig. 7). The width of the molten pool is defined as the distance of the molten material along
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the direction perpendicular to the scan direction (Y-axis). The depth of the melt pool was measured from the top of the substrate to the molten depth inside the mater (Z-axis)
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so that it does not include the initial thickness of the powder bed. A typical image of a cross-section of the track obtained experimentally is shown in Fig.
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8. The experimental cross-section of the track (right figure in Fig. 8) shows a convex
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top surface, resulting from the flow of the liquid metal. Nevertheless, the corresponding
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numerical result (left figure in Fig. 8) doesn’t show this characteristic feature. In practice, the use of a model allowing following the liquid/gas interface (e.g., such as the VOF model for example) would be required to provide such a result. However, the main model developments performed in this work were focused on another critical point (i.e., the effect of the surrounding atmosphere and the contribution of the convective movements in this atmosphere). The depth and width of the melt pool estimated with these methods were investigated for different sets of parameters and are presented in Fig. 9. The figure shows the range and average experimental melt pool dimensions, as well as the corresponding simulated result.
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Fig. 7 Measurement method of the melt pool dimensions in the simulation of the melt
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pool morphology
Fig. 8 Cross section of a SS 316L track obtained with a scanning speed of 0.1ms-1 at a pressure of 995 mbar 21
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The range of variation of experimental data results from the origin characteristic of the laser power which varies above and below the specified power (Loh et al., 2015) and from the anisotropic nature of the initial powder bed. As shown in Fig. 9, by comparing
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numerical and experimental results, a quite good correlation was found and the
predicted melt pool width and depth fall within the corresponding experimental range.
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The present analysis method of the numerical results seems thus suitable to provide a
Ac ce p
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good prediction of the melt pool dimensions.
Fig. 9 Numerical and experimental results of the melt pool depth and width
As also shown in Table 6, the melt pool depth and width are directly proportional to the effective scan speed of the laser. 22
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Table 6. Dimensions of tracks obtained by numerical modelling – Case of a chamber pressure of 1 bar texpo (µs)
Pdist (µm)
Melt pool width (µm)
Melt pool depth (µm)
400
40
168
28
0.2
200
40
143
16
0.3
120
40
127
8
0.4
100
40
122
3
cr
0.1
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Scan speed (m/s)
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The melt pool dimensions decrease with the increase of the scan velocity (i.e. lowering
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of the exposure time). The exposure time of the laser beam affects the energy density transferred into the material and is hence an important factor to provide a full melting of
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the material. Below a certain limit, a decrease of the exposure time can give rise to an increase of the porosity level.
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5.3. Surrounding gas motion around the laser-powder-atmosphere interaction zone
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The material evaporation in the interaction zone provides a flux of metal vapor, causing
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convection movements in the gas just over the irradiated surface. This phenomenon seems to become essential for low pressure conditions. Fig. 10 displays the velocity vectors of the evaporated material in the laser-powder-atmosphere interaction zone for an effective scan speed of 0.1 m/s (i.e., exposure time of 400 µs and distance of 40 µm between two successive points) and a chamber pressure of 100 mbar.
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Fig. 10 Velocity vectors (m/s) of the evaporated material for an Argon-gas pressure of
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100 mbar, in the near surface plane.
Fig. 11 shows the evolution of the velocity field in the gas side just over the laser impact point for different chamber pressures in the range 100-2000 mbar. The color scale is the same for the 5 pictures. For high pressure levels, the velocity is low. However, a decrease of the chamber pressure provides a strong increase of the velocity and an expansion of the affected region.
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Fig. 11 Evolution of the velocity field provided by the evaporated material for several
Argon gas pressures, in the near surface plane and viewed from the cross section of the melt pool.
The maximum velocity corresponding to the different pressure levels of Fig. 11 is plotted versus pressure on Fig. 12. As mentioned above, for high pressure levels, the maximum velocity remains quite low (i.e., below 100 m/s for a pressure above 800
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mbar). On the contrary, the maximum velocity increases strongly with the lowering of the pressure in the chamber (i.e., up to 300 m/s for the lowest pressure of 100 mbar). As first approximation, the maximum velocity seems almost inversely proportional with the
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pressure in the chamber.
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Fig. 12 Evolution of the maximum velocity of the evaporated material with the Argon gas pressure
The corresponding concentration of the metal vapor is shown in Fig. 13. An expansion of the vapor in the chamber is observed with the decrease of the ambient pressure. The evaporated material expands more freely in the rarefied atmosphere (100 mbar). On the contrary, at high pressure, the region of the high concentration of the evaporated material remains small and is characterized by a higher density and a slower propagation speed.
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ip t cr us an M d te
Ac ce p
Fig. 13 Concentration of the evaporated material (mass fraction) for several Argon-gas pressures
Thus, Argon acts not only as a shielding gas in the process chamber, but it also squeezes the evaporated material into a smaller volume. For low pressures, the amount of mater vaporized increases and the generated vapor expands significantly in the atmosphere. Interactions between the laser beam and the metal vapor requires complementary investigations in order to quantify the effect of the vapor on the heat transferred to the powder bed by the laser. Concerning this point, a test was performed with metal vapors
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Page 27 of 34
generated by a plasma jet, a laser beam oriented through the generated vapor and a calorimeter to measure the power transferred by the laser. In the considered conditions, the power transferred was not affected by the presence of the metallic vapor and the
ip t
Argon plasma jet, so that the absence of interactions between the laser and the
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vaporized material can be assumed.
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6. Conclusion
A study of the powder-laser-atmosphere interaction during the SLM process was
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presented in this work. A 3D modeling approach was developed and validated progressively by experimental results. The model provides a first approach of the role of
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the metal vapor generated in the interaction region of the laser beam and the powder bed. An effect of the pressure within the chamber was already shown: a decrease of the
d
pressure gives rise to a strong expansion of the metal vapor in the interaction region.
A high pressure environment may contribute to decrease the convective movements
Ac ce p
-
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Based on these results, the following conclusion can be drawn:
in the gas side in the laser impact region.
-
A low pressure environment contributes to decrease the difference between the melting point and the vaporization point and can thus increase the rate of evaporation of the material. This high vaporization can also contribute to blow the powder particles. The combination of these two effects prevent building tracks at low pressure levels (i.e., below 100 mbar in the present case).
-
An adaptation of the parameters, such as a change of the scan speed, seems thus to be required in order to work under low pressure conditions. Nevertheless, it can
28
Page 28 of 34
become very sensitive due to the lowering of the difference between the melting and vaporization temperatures of the material with the decrease of the pressure in the chamber.
ip t
Finally, the use of a high pressure environment could contribute to decrease the
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convective movements in the atmosphere and to lower the vaporization rate of the
material. The use of a surrounding pressure above 1 bar could thus be tested in a future
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work.
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References:
Cédric André, 2007, modélisation quantitative du procédé de frittage sélectif par laser:
lausanne, Lausanne, EPFL.
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relation paramètres/microstructure, PhD thèse, école polytechnique fédérale de
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Cuadros F., I. Cachadina, W. Ahumada, 1996, Determination of Lennard-Jones
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Interaction Parameters using a new procedure, Molecular Engineering: 319-325.
Ac ce p
Deissler R.G., Eian C.S., 1952, Investigation of effective thermal conductivities of powders, National advisory committee for aeronautics, Washington.
Donghua Dai, Dongdong Gu, 2014, Thermal behavior and densification mechanism during selective laser melting of copper matrix composites: Simulation and experiments, Materials & Design 55, 482-491.
Edson Costa Santos, Masanari Shiomi, Kozo Osakada, Tahar Laoui., 2006, Rapid manufacturing of metal components by laser forming. International Journal of Machine Tools and Manufacture 46, no 12-13,1459-1468.
29
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Ehsan,Ebrahimnia Bajestan,Hamid Niazmand, Weerapun Duangthongsuk, Somchai Wongwises, 2011, Numerical investigation of effective parameters in convective
Journal of Heat and Mass Transfer 54, no 19-20, 4376-88.
ip t
heat transfer of nanofluids flowing under a laminar flow regime,International
cr
Filali Mohamed, 2006, Conductivité thermique apparente des milieux granulaires
soumis à des contraintes mécaniques : modélisation et mesures, thèse, chapitre2.
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Gusarov A. V., I. Smurov, 2009b,Two-dimensional numerical modelling of radiation transfer in powder beds at selective laser melting ». Applied Surface Science 255,
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no 10, 5595-5599.
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Gusarov A. V., I.Yadroitsev, Ph.Bertrand, I.Smurov, 2009a, Model of Radiation and Heat Transfer in Laser-Powder Interaction Zone at Selective Laser Melting,
d
Journal of Heat Transfer 131, no 7.
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Gusarov A. V.,I. Smurov,2010,Modelling the interaction of laser radiation with powder
Ac ce p
bed at selective laser melting, Physics Procedia 5, Part B 381-394. Hirschfelder J. O. , C.F. Curtiss, R. B. Bird, 1954, Molecular Theory of gases and liquids, pp 1110-1112, Wiley, New York.
Hussein Ahmed, Liang Hao, Chunze Yan, Richard Everson, 2013, Finite element simulation of the temperature and stress fields in single layers built withoutsupport in selective laser melting, Materials & Design 52,638-647.
IAEA (International Atomic Energy Agency), 2008, Thermophysical Properties of Materials for Nuclear Engineering: A Tutorial and Collection of Data, Vienna, pages 168-170.
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Khairallah Saad A., Andy Anderson, 2014, Mesoscopic simulation model of selective laser melting of stainless steel powder, Journal of Materials Processing Technology 214, no 11,2627-36. doi:10.1016/j.jmatprotec.2014.06.001.
ip t
Kurian Antony, N. Arivazhagan, K. Senthilkumaran, 2014, Numerical and experimental
cr
investigations on laser melting of stainless steel 316L metal powders ,Journal of Manufacturing Processes 16, no 3,345-55.
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Loh, Loong-Ee, Chee-Kai Chua, Wai-Yee Yeong, Jie Song, MahtaMapar, Swee-Leong Sing, Zhong-Hong Liu and Dan-Qing Zhang , 2015, Numerical investigation and
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an effective modelling on the Selective Laser Melting (SLM) process with aluminum alloy 6061,International Journal of Heat and Mass Transfer 80,
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288-300.
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Mumtaz K.A., N. Hopkinson, 2010, Selective Laser Melting of thin wall parts using
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pulse shaping, Journal of Materials Processing Technology, 210, no 2, 279-287. Rai1 R., JW Elmer, T. A. Palmer T. DebRoy, 2007, Heat transfer and fluid flow during
Ac ce p
keyhole mode laser welding of tantalum,Ti–6Al–4V, 304L stainless steel and vanadium, Journal of Physics D: Applied Physics.
Ruidi Li, YushengShi, Jinhui Liu, Huashan Yao, Wenxian Zhang, 2009, Effects of processing parameters on the temperature field of selective laser melting metal powder, Powder Metallurgy and Metal Ceramics 48, no 3-4, 186-195.
Stoffregen H., Fischer J., Siedelhofer C., Abele E., 2011, Selective Laser Melting of porous Structures , Proceedings of the 22nd Annual International Solid Free form Fabrication (SFF) Symposium, University of Texas, Austin, p. 680.
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Takamichi Iida, Roderick I.L. Guthrie, 1993, The physical properties of liquid Metal, Clarendon Press, Oxford. Verhaeghe F., T. Craeghs, J. Heulens, L. Pandelaers, 2009,A pragmatic model for
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selective laser melting with evaporation , Acta Materialia 57, no 20 6006-6012.
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Yadroitsev I., I.Smurov, 2011, Surface Morphology in Selective Laser Melting of Metal
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Powders, Physics Procedia, 264-270.
Yadroitsev I., Ph. Bertrand, I. Smurov, 2007, Parametric analysis of the selective laser
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melting process, Applied Surface Science 253 Elsevier , pp 8064-8069. Zhang Baicheng, Lucas Dembinski, Christian Coddet, 2013b, The study of the laser
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parameters and environment variables effect on mechanical properties of high compact parts elaborated by selective laser melting 316L powder, Materials
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Science and Engineering: A 584,21-31. doi:10.1016/j.msea.2013.06.055.
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Zhang Baicheng, Hanlin Liao, Christian Coddet, 2013a, Microstructure evolution and
Ac ce p
density behavior of CP Ti parts elaborated by Self-developed vacuum selective laser melting system , Applied Surface Science.doi:10.1016/j.apsusc.2013.04.090.
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Figure and Table captions Figure 1. Modelling steps of a 316L stainless steel powder bed irradiation by a laser beam (a) different regions of the calculation domain, (b) mesh, (c) laser irradiation of
ip t
the powder bed. Figure 2. Gaussian distribution of the thermal flux.
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Figure 3.Vacuum selective laser melting machine developed at LERMPS laboratory.
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Figure 4. 316L stainless steel tracks manufactured with several scanning speeds (1-4) at several pressure levels (A to F) of Argon gas.
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Figure 5. Effect of the Argon gas pressure on SS 316L tracks prepared by the SLM process (a) 995 mbar (b) 500 mbar (c) 250 mbar (d) 100 mbar (e) 10 mbar (f)1 mbar –
M
veff=0.1 m/s.
Figure 6. Temperature distribution (K) for a single track scanned linearly (veff =
d
0.1 m/s, p= 1 bar).
Ac ce p
melt pool morphology.
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Figure 7. Measurement method of the melt pool dimensions in the simulation of the
Figure 8. Cross section of a SS 316L track obtained with a scanning speed of 0.1 m.s-1 at a pressure of 995 mbar.
Figure 9. Numerical and experimental results of the melt pool depth and width. Figure 10. Velocity vectors (m/s) of the evaporated material for an Argon-gas pressure of 100mbar, in the near surface plane. Figure 11. Evolution of the velocity field provided by the evaporated material for several Argon gas pressures, in the near surface plane and viewed from the cross section of the melt pool.
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Figure 12. Evolution of the maximum velocity of the evaporated material with the Argon gas pressure. Figure 13. Concentration of the evaporated material (mass fraction) for several Argon-
ip t
gas pressures.
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Table 2. Lennard-Jones parameters for Fe and Ar
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Table 1. Thermo-physical properties of 316L
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Table 3. Scan speeds applied to elaborate the single lines
Table 4. Argon gas pressure selected to elaborate the single lines
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Table 5. Vaporization temperature (Tv) of 316L for different pressure levels
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pressure of 1 bar.
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Table 6. Dimensions of tracks obtained by numerical modelling – Case of a chamber
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S0924-0136(15)00208-3 http://dx.doi.org/doi:10.1016/j.jmatprotec.2015.05.008 PROTEC 14413
To appear in:
Journal of Materials Processing Technology
Received date: Revised date: Accepted date:
20-2-2015 14-4-2015 11-5-2015
Please cite this article as: MASMOUDI, A., BOLOT, R., CODDET, C.,Investigation of the laser-powder-atmosphere interaction zone during the selective laser melting process, Journal of Materials Processing Technology (2015), http://dx.doi.org/10.1016/j.jmatprotec.2015.05.008 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Investigation of the laser-powder-atmosphere interaction zone during the selective laser melting process
LERMPS -University of Technology Belfort – Montbeliard, Sevenans, 90010 Belfort, France
*[email protected]
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Abstract
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a
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Amal MASMOUDIa,* , Rodolphe BOLOTa and Christian CODDETa
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Selective Laser Melting (SLM) consists in manufacturing parts in a chamber filled with a protective gas, by melting successive thin powder layers using a laser. This study
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aims, at first, to understand the laser–material–atmosphere interactions that involve various complex and simultaneous physical phenomena such as heat transfer in the
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target, melting and vaporization of the material, and expansion of the generated vapor.
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A modelling approach of the process was developed to investigate the effect of some process parameters, such as the exposure time (texpo) of the laser beam on each point
Ac ce p
along its trajectory, the pressure of the protective gas (Argon) and the spatial thermal energy distribution on the dimensional characteristics of the melted track. Keywords: Laser melting process, Rapid prototyping, Additive manufacturing, Numerical modelling, CFD model, Metallic vapour. 1. Introduction
Additive manufacturing processes, also known as rapid manufacturing or rapid prototyping (Edson et al., 2006) allow manufacturing complex shape parts directly from 3D CAD data. Among those processes, Selective Laser Melting (SLM) was developed
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Page 1 of 34
to produce fully functional metal parts presenting high mechanical properties directly from powder (Yadroitsev et al., 2011). It consists in manufacturing parts layer-by-layer using a local melting of a powder bed provided by laser irradiation, followed by
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immediate consolidation (Yadroitsev et al., 2006). The process is carried out in a
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chamber filled with a protective gas such as Argon to avoid the metal oxidation. SLM is a complex process involving several physical and chemical phenomena
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especially during the laser-powder interaction and it is yet not sufficiently understood. Thus, further researches are required and several numerical simulations are developed
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particularly focusing on the parts properties and on the laser-powder interaction zone under various processing parameters (Ehsan et al., 2011). A selection of those
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researches is briefly discussed in the following section.
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Gusarov et al. (2009a, 2009b, 2010) studied the phenomena related to the laser-powder
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interaction by modelling the penetration of heat and radiations in the powder bed by taking multiple reflections into account. They numerically calculated the temperature
Ac ce p
distribution in the laser-powder interaction region and investigated the influence of the powder layer thickness and of the laser power on the melt pool shape. They also studied the effects of the scanning velocity. Hussein et al. (2013) developed a three-dimensional model to predict the temperature distribution and stress field inside a single metallic layer formed on the powder bed without support. The simulation results revealed the magnitude of the temperature and stresses at different locations of the layer. The temperature gradients were the highest at the start of the first track scan and drops subsequently for all scanning speeds.
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Verhaeghe et al. (2009) presented a model allowing determining the effect of evaporation. They found that evaporation becomes a very significant phenomenon at high beam powers and energy densities. They also investigated thermal gradients and
ip t
fluid dynamics within the melt pool.
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Donghua et al. (2014) developed a numerical simulation regarding the influence of the energy density of the laser beam on the melt pool dynamics and densification
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mechanisms.
Khairallah et al. (2014) developed a 3D mesoscopic simulation model to study the laser-
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induced melting of a random bed of stainless steel 316L particles on a solid substrate, and the solidification of the melt pool into either a continuous or a discontinuous track
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as a result of the Plateau–Rayleigh instability. Their study demonstrates the importance
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of simulation as a tool to determine the underlying physics governing the SLM process.
te
Moreover, simulation may be helpful to study the interplay between the build parameters in order to optimize the process.
Ac ce p
Kurian et al. (2014) presented a numerical investigation to understand the effect of some process parameters during the laser melting of grade 316L stainless steel powder on an AISI 316L substrate using a pulsed Nd-YAG laser. They presented the temperature distribution in a single powder layer. They studied the effect of some process parameters such as the laser power, the scanning speed and the size of the beam on the dimensional characteristics of the melt region from the temperature solution. The results confirmed found that the laser power and scanning speed of the laser beam have a noticeable effect on the track characteristics.
3
Page 3 of 34
Nevertheless, few studies are concerned with the role of the atmosphere on the process and parts properties. Zhang et al. (2013a) studied pure Ti parts produced under vacuum and Zhang et al. (2013b) analyzed the influence of the protective gas nature on 316L
ip t
parts. They showed that a decrease of the scan speed may provide a denser material
under vacuum conditions. However, a decrease of the chamber pressure gave also rise
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to a lower cross-section of the track, resulting certainly from an increase of the
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vaporization level.
Therefore, a 3D numerical model was developed in the present work to study the laser-
an
powder-atmosphere interaction under different conditions. This model is able to consider the fundamental parameters involved in the SLM process such as the laser
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beam properties, the thermal energy distribution, the powder layer characteristics (material nature, particle size distribution, thickness, etc.), the nature and pressure of the
d
environment gas, and other process parameters. This model is applied to investigate the
Ac ce p
material behavior.
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thermal and dimensional characteristics of the melt pool zone and the evaporated
Furthermore, the results obtained by numerical simulation are used to understand the effective properties of SLM parts manufactured during experimental works, in order to check the accuracy of the developed simulation model. 2. Modelling approach
A numerical simulation of the selective laser melting process was developed to study the laser-powder-atmosphere interaction by considering the different process parameters.
4
Page 4 of 34
In practice, a single line (Fig. 1c) scanned by the laser beam on a thin (50 μm) powder layer placed over a dense substrate (1 mm) of the same material (i.e., 316L stainless steel) and exposed to a protective Argon atmosphere was simulated using FLUENT
ip t
software as shown in Fig. 1.
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The mesh size was adapted (Fig. 1b) to provide thin cells in the interaction zone of the
laser with the powder bed. On the contrary the mesh size was increased in other regions
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d
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an
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in order to reduce the computational time.
Fig. 1 Modelling steps of a 316L stainless steel powder bed irradiation by a laser beam
Ac ce p
(a) different regions of the calculation domain, (b) mesh, (c) laser irradiation of the powder bed.
2.1. Governing equations
The transient heat conduction equation was considered in solid regions as: h . T x, y, z t
(W .m 3 )
(1)
in which stands for density (kg.m-3), h for enthalpy (J.kg-1), for thermal conductivity (W.m-1.K-1), T for temperature (K) and represents the volumic heat source corresponding to the laser.
5
Page 5 of 34
Additionally, transient convective movements in the Argon atmosphere were taken into account by considering the Navier-Stokes equations as: -
Continuity:
ip t
.v 0 t
(2)
Momentum:
v . v v p . g t
us
-
cr
in which v stands for the velocity vector.
(3)
-
Energy:
Dp Dh . T v Dt Dt
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an
in which p is the pressure, is the stress tensor and g is the gravitational acceleration.
(4)
Ac ce p
dissipation function.
t
te
Dt
d
with by definition De e v. e whatever e, and v corresponds to the viscous
2.2. Boundary and initial conditions A uniform temperature throughout the powder bed was considered before the laser irradiation:
T (t 0) T0 288 K
(5)
in which T0 is the ambient temperature. SLM layers are then built over a thick solidified material block (i.e., made of powder melted and solidified) presenting thermo-physical properties of the dense material (as first approximation). 6
Page 6 of 34
The thermal flux on the opposite face of the material block (the rear face of the plate) was neglected, corresponding to an adiabatic condition. This assumption was made due to the relatively high thickness of the block and to the fast laser irradiation that do not
ip t
allow heating the rear face up to a significant value. Concerning the limits of the domain, the sides and the top face of the fluid domain were considered as "free
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surfaces" at which only the pressure was fixed (i.e., a relative pressure of 0 Pa).
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The laser source of the SLM machine considered in this study (from Realizer) is an YLR-100-SM single mode CW Ytterbium fiber laser (1064–1100 nm). This laser is
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used to melt the powder bed layer-by-layer (Ruidi et al., 2009). According to the calculation domain including both solid and fluid regions, the heating of the powder bed
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by the laser takes place at the gas/solid interface (i.e., corresponding to an internal face of the calculation domain). In practice, this fluid/solid interface was represented by the
d
way of a “coupled wall”, meaning that the continuity of the thermal flux on both sides
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of this interface is considered. In order to keep this possibility, the laser heat source was
Ac ce p
thus considered as a volumic heat source concentrated in the first layer of the mesh in the solid side. For this, a Gaussian distribution of the power density is assumed (Fig. 2). It is the most common model of the power radial distribution (Mumtez et al., 2010) and is defined by:
r2 ( r ) 0 exp 2 2
(6)
in which 0 P2 , P is the power of the laser beam which was fixed during the 2 simulation (100 W), α is the absorption coefficient of dense 316L stainless steel and
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Page 7 of 34
d where d is the apparent spot diameter of the laser beam (d = 68 µm so that σ = 4
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an
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cr
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17 µm).
Fig. 2 Gaussian distribution of the thermal flux
(7)
Ac ce p
2r 2 (r ) 0 exp 2
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defined as:
d
In some other studies (Hussein et al., 2013), the Gaussian distribution is sometimes
in which 2
According to Gusarov et al. (2009b), the penetration of the laser energy into the powder bed involves multi-reflection phenomena, providing a much higher effective absorption rate (about 0.8 for 316L) than the absorptivity of a flat dense surface (about 0.3 for 316L). This model would be suitable for the case of a laser beam impacting on loose particles without melting. However, after our results and considering our parameters, the laser impact point during the SLM process is mainly positioned on the melt pool
8
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(and not on unmelted particles). Thus, in the present study, the effective absorption coefficient of the liquid melt pool was considered as α = 0.3 for 316L stainless steel
ip t
(Rai1 et al. 2007). 2.3. Effective properties of the solid materials
cr
Basic properties of 316L stainless steel such as melting point (about 1750 K) and boiling point (3090 K) were taken into account during the numerical simulations
us
(Cédric, 2007; IAEA, 2008).
During the SLM process, the powder bed is molten and rapidly solidified into a dense
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solid state. Thus, the latent heat of fusion (270 kJ/kg) was considered and equations shown in Table 1 were used to define the 316L stainless steel properties in the liquid
Heat capacity -1
-1
Dense solid state
Powder state
462 + 0.134 T
462 + 0.134 T
7433 + 0.0393× T -
8084 + 0.4209 × T -
8084 + 0.4209 × T -
-4
-5
755
Ac ce p
Сp (J.kg .K )
Liquid state
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Properties of 316L
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Table 1. Thermo-physical properties of 316L
Density ρ (kg.m-3)
Thermal conductivity λ (W.m-1.K-1)
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and solid (dense and powder) states (IAEA, 2008).
1.801.10 × T
2
12.41 + 0.003279 × T
3.894. 10 × T
2
9.248 + 0.01571 × T
3.894.10-5 × T 2 (1 )
9.248 + 0.01571 × T (1 - )
in which ε = 0.44 represents the porosity level of the powder bed and T is the temperature in [K].
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Page 9 of 34
Many correlations have been proposed in the literature in order to estimate the thermal conductivity of powder materials (Filali, 2006; Deissler, 1952). However, these models give rise to highly variable results.
ip t
In the present study, an experimental procedure was developed to estimate the thermal conductivity of powder materials. Finally, a corrective factor of 0.56 (i.e.,
cr
corresponding to 1-) was deduced and the thermal conductivity was defined as shown
2.4. Modelling of the surrounding atmosphere
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in Table 1. Details concerning this method shall be given in another specific paper.
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Processing with the laser beam was carried out in a closed chamber filled with a protective gas (Argon) to avoid oxidation of the metallic material.
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During this work, the Argon environment was considered as a domain of compressible ideal gas. Moreover, the Lennard-Jones potential was considered to describe
d
interactions between the gas molecules (for estimation of the transport properties). This
te
potential is the most used for simple gases and is expressed as (Cuadros et al., 1996):
Ac ce p
12 6 E ( r ) 4 r r
(8)
in which is the depth of the potential energy well and is the distance at which the potential is zero. Table 2 lists the Lennard-Jones parameters for Iron (Cuadros et al., 1996) and Argon (Hirschfelder et al., 1954) taken from the literature.
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Page 10 of 34
Table 2. Lennard-Jones parameters for Fe and Ar
(Angstrom Ǻ)
/κ (K)
Fe
2.321
6104
Ar
3.418
124
cr
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Element
During the laser-powder interaction, the thermal power transferred is high enough to
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provide a melting of the powder bed in the region directly exposed to the laser beam.
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Depending on the scanning conditions, the boiling point may be reached, thus providing a partial vaporization of the material. So, the latent heat of evaporation of 316L stainless
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steel (745 kJ/kg) was considered (IAEA, 2008). During each numerical time step, the fraction of evaporation of cells at the material/Argon interface is calculated, providing a
d
mass source to the continuity equation in the gas side. The expansion of the metallic
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vapor in the Argon atmosphere depending on the pressure is thus investigated.
Ac ce p
3. Experimental procedure
A specific selective laser melting machine developed at LERMPS laboratory allowing working under vacuum or in a controlled atmosphere was used in this work (Fig. 3). A special feature of this machine is its ability to vary the Argon gas pressure between 10 and 105 Pa. It is equipped with an Nd-YAG fiber laser with a maximum power of 120 W.
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Page 11 of 34
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Fig. 3 Vacuum selective laser melting machine developed at LERMPS
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laboratory
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The scanning speed of the laser is controlled thanks to oscillating mirrors driven by stepper motors. In that case, the laser movement on the powder bed is discontinuous: the laser stops during a short exposure time (texpo) on a point and then moves to the next point at a distance Pdist along the trajectory. Thus, the effective scanning speed depends on the two parameters texpo and Pdist according to the following expression:
veff
Pdist
(9)
Pdist t exp o v1
in which ν1 is a characteristic velocity depending on the stepper motors.
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Page 12 of 34
This characteristic velocity corresponds to the maximum velocity possible with the stepper motors and can be estimated by considering a very large point to point distance. In our case, this characteristic velocity was thus estimated to 2.5 m/s. For industrial
ip t
applications, this velocity is sometimes considered as sufficiently high to neglect its effect on the average scanning speed, which is thus sometimes defined as Pdist/texpo.
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A two parameter DOE (Design Of Experiments) was used to quantify the effect of the
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effective scan speed (changed with texpo), and that of the chamber pressure, on the
Ac ce p
te
d
M
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characteristics of single tracks (Fig. 4 and Tables 3 and 4).
Fig. 4 316L stainless steel tracks manufactured with several scanning speeds (1 to 4) at several pressure levels (A to F) of Argon gas
Tables 3 and 4 present the four values of the effective scan speed and six values of the chamber pressure that were considered for this DOE. Three different tracks were made for each parameter set (Fig. 4). The average of each result such as line width and depth was considered for comparisons with the
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Page 13 of 34
corresponding simulation. The process parameters were of course those used for the numerical modelling (laser power of 100 W, spot diameter of 68 µm, powder bed thickness of 50 µm and distance of 40 µm between two successive points).
1
400
40
2
200
40
3
120
40
4
100
40
Scan speed veff (m/s)
cr
Pdist (µm)
0.1 0.2
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texpo (µs)
0.3 0.4
an
Parameter No.
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Table 3. Scan speeds applied to elaborate the single lines
A
B
C
Pressure (mbar)
995
500
250
D
E
F
100
10
1
d
Parts
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Table 4. Argon gas pressure selected to elaborate the single lines
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4. Experimental results and discussion
Ac ce p
Fig. 5 displays more precisely the effect of the Argon gas pressure on the tracks obtained with an exposure time of 400 µs and a distance of 40 µm between two successive points (i.e., an effective scan speed of 0.1 m/s). At high Argon pressure, a continuous solid track is formed following the powder melting and solidification (Fig. 5 (a – d)). At low pressures (i.e., below 100 mbar), the solid track disappears and only a slight remelting of the solid surface just under the powder bed is observed. In this situation, the difference between the melting and vaporization temperatures becomes very low and it may thus be supposed that the liquid phase is vaporized as soon as it is formed, thus explaining the absence of any solid track
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Page 14 of 34
on the surface. For still lower pressures, the liquid phase would not even be formed and a sublimation of the solid phase could take place (i.e., direct vaporization of the solid
Ac ce p
te
d
M
an
us
cr
ip t
particles).
Fig. 5 Effect of the Argon gas pressure on SS 316L tracks prepared by the SLM process with veff = 0.1 m/s (a) 995 mbar (b) 500 mbar (c) 250 mbar (d) 100 mbar (e) 10 mbar (f) 1 mbar
In order to confirm these assumptions, the Clausius-Clapeyron equation (Takamichi et al., 1993) provides the evolution of the vaporization temperature versus pressure as: d( ln p) ΔH vap d(1/T) R
(10)
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Page 15 of 34
in which R = 8.3145 J.mol-1.K-1 is the universal constant of gases and ΔHvap is the enthalpy of evaporation of the liquid metal. More particularly, the vapor pressure of liquid 316L steel may be estimated from
18868 T
cr
log10 p 11.1183
ip t
(IAEA, 2008):
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with p (Pa) and T (K).
(11)
18868 11.1183 log10 ( p )
(12)
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Tv
an
The change of the vaporization temperature with the chamber pressure is thus given by:
Based on this equation, the equilibrium temperature values at the different Argon-gas
d
pressure levels considered in the DOE are given in Table 5.
Parameters
te
Table 5. Vaporization temperature (Tv) of 316L for different pressure levels values
2000
1000
500
250
100
10
1
Tv (K)
3243
3084
2940
2807
2650
1865
1697
Ac ce p
p (mbar)
Tv represents the temperature at which gas and liquid phases are in equilibrium. According to Table 5, the vaporization temperature decreases strongly with the lowering of the pressure. One may also notice that the conventional boiling temperature of stainless steel (i.e., about 3093 K) is well predicted at 1 bar, and that there is no more difference between the melting and boiling points (about 1750 K) at 2.2 Pa. Below this pressure, a sublimation of the material occurs.
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Page 16 of 34
5. Numerical results 5.1. Temperature distribution
ip t
As mentioned previously, many numerical researches were developed to study the effect of the processing parameters on the temperature distribution of the powder bed. For
cr
example, according to Kurian et al. (2014), the temperature of the melt pool rises
rapidly and reaches about 3500 K for a 316L stainless steel powder bed with a high
us
energy density and /or a low scan speed. However, in most cases, the environment properties (nature and pressure of the gas) within the chamber of the SLM machine
an
were not considered.
After what was said above, the maximum temperature of the molten material should not
M
exceed the vaporization temperature that depends on the liquid / gas equilibrium (see
d
equation 12).
te
One may hence assume that a decrease of the environment pressure should imply a decrease of the laser power or an increase of the scan speed in order to limit the material
Ac ce p
vaporization. Nevertheless, it becomes very sensitive at low pressure levels due to the decrease of the difference between the melting and vaporization temperatures. In order to take this phenomenon into account, the temperature of the molten pool was thus limited in the model by the equilibrium vaporization temperature provided by equation 12. In counterpart, the additional heat transferred once the vaporization temperature is reached contributes to vaporize a fraction of the material. Within the program, lines of C coding were written in order to represent that:
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Page 17 of 34
The effect of vaporization is first omitted in the thermodynamic properties of the material. Then after each time step, the mass fraction of the vaporized material is calculated from the temperature excess beyond the vaporization point as: c p (T Tv ) H vap
ip t
y
(13)
cr
Once this fraction is calculated, the local temperature T in the cell is reset to Tv for cells
us
with a higher temperature. This method allows considering the vaporization
phenomenon by introducing a mass source term to the continuity equation in the gas
y mcell dt
(14)
M
m
an
side. At each time step dt, this mass source term is set as:
in which y is the mass fraction of the liquid cell vaporized, mcell is the mass of material
d
within this cell (i.e..Vcell) and dt corresponds to the time step.
te
Fig. 6 displays the temperature distribution in the scanned zone for an effective scan speed of 0.1 m.s-1 (i.e., texpo = 400 µs and Pdist = 40 µm) in the case of a pressure of
Ac ce p
1 bar, a laser power of 100 W and a laser spot diameter of 68 µm. The laser trace on the powder bed is clearly observed and the maximum temperature is limited by the calculated vaporization temperature.
18
Page 18 of 34
ip t cr us
an
Fig. 6 Temperature distribution (K) for a single track scanned linearly (veff = 0.1 m/s, p = 1 bar)
M
For these process conditions, the temperature range of the scanned zone is comprised between 288 K (the initial temperature) and the vaporization temperature of about
d
3080 K. The temperature of the powder bed in the region directly exposed to the laser
te
beam largely exceeds the melting temperature of 316L stainless steel (≈1700 K) causing
Ac ce p
the formation of a molten pool. From this molten pool, the heat propagates in the surrounding loose powder. The power of the source is also high enough to cause a local evaporation of the material. A fast cooling down of the molten zone is then observed when the laser beam leaves the zone so that temperatures below the melting point (solid phase) are quickly obtained behind the hot spot (Fig. 6). 5.2. Melt pool dimensions and shape The morphology of the melt pool region predicted with this model can directly be compared to that of single tracks manufactured (see Fig. 4 and Fig. 5).The melt pool width and depth penetration are the most important characteristics to obtain fully dense 19
Page 19 of 34
parts. In particular, the depth of penetration and the width of the track have to be sufficient to allow overlapping and welding of the successive layers and tracks (Stoffregen et al., 2011).
ip t
Fig. 7 and Fig. 8 show respectively the morphology of the melt pool (thanks to the
cr
temperature field) predicted by the simulation and that of the track obtained
experimentally after a single scan on the powder bed. In the simulation, the melt pool
us
dimensions were determined from the temperature distribution along the laser path (Fig. 7). The width of the molten pool is defined as the distance of the molten material along
an
the direction perpendicular to the scan direction (Y-axis). The depth of the melt pool was measured from the top of the substrate to the molten depth inside the mater (Z-axis)
M
so that it does not include the initial thickness of the powder bed. A typical image of a cross-section of the track obtained experimentally is shown in Fig.
d
8. The experimental cross-section of the track (right figure in Fig. 8) shows a convex
te
top surface, resulting from the flow of the liquid metal. Nevertheless, the corresponding
Ac ce p
numerical result (left figure in Fig. 8) doesn’t show this characteristic feature. In practice, the use of a model allowing following the liquid/gas interface (e.g., such as the VOF model for example) would be required to provide such a result. However, the main model developments performed in this work were focused on another critical point (i.e., the effect of the surrounding atmosphere and the contribution of the convective movements in this atmosphere). The depth and width of the melt pool estimated with these methods were investigated for different sets of parameters and are presented in Fig. 9. The figure shows the range and average experimental melt pool dimensions, as well as the corresponding simulated result.
20
Page 20 of 34
ip t cr us an
M
Fig. 7 Measurement method of the melt pool dimensions in the simulation of the melt
Ac ce p
te
d
pool morphology
Fig. 8 Cross section of a SS 316L track obtained with a scanning speed of 0.1ms-1 at a pressure of 995 mbar 21
Page 21 of 34
The range of variation of experimental data results from the origin characteristic of the laser power which varies above and below the specified power (Loh et al., 2015) and from the anisotropic nature of the initial powder bed. As shown in Fig. 9, by comparing
ip t
numerical and experimental results, a quite good correlation was found and the
predicted melt pool width and depth fall within the corresponding experimental range.
cr
The present analysis method of the numerical results seems thus suitable to provide a
Ac ce p
te
d
M
an
us
good prediction of the melt pool dimensions.
Fig. 9 Numerical and experimental results of the melt pool depth and width
As also shown in Table 6, the melt pool depth and width are directly proportional to the effective scan speed of the laser. 22
Page 22 of 34
Table 6. Dimensions of tracks obtained by numerical modelling – Case of a chamber pressure of 1 bar texpo (µs)
Pdist (µm)
Melt pool width (µm)
Melt pool depth (µm)
400
40
168
28
0.2
200
40
143
16
0.3
120
40
127
8
0.4
100
40
122
3
cr
0.1
ip t
Scan speed (m/s)
us
The melt pool dimensions decrease with the increase of the scan velocity (i.e. lowering
an
of the exposure time). The exposure time of the laser beam affects the energy density transferred into the material and is hence an important factor to provide a full melting of
M
the material. Below a certain limit, a decrease of the exposure time can give rise to an increase of the porosity level.
d
5.3. Surrounding gas motion around the laser-powder-atmosphere interaction zone
te
The material evaporation in the interaction zone provides a flux of metal vapor, causing
Ac ce p
convection movements in the gas just over the irradiated surface. This phenomenon seems to become essential for low pressure conditions. Fig. 10 displays the velocity vectors of the evaporated material in the laser-powder-atmosphere interaction zone for an effective scan speed of 0.1 m/s (i.e., exposure time of 400 µs and distance of 40 µm between two successive points) and a chamber pressure of 100 mbar.
23
Page 23 of 34
ip t cr us an M d
te
Fig. 10 Velocity vectors (m/s) of the evaporated material for an Argon-gas pressure of
Ac ce p
100 mbar, in the near surface plane.
Fig. 11 shows the evolution of the velocity field in the gas side just over the laser impact point for different chamber pressures in the range 100-2000 mbar. The color scale is the same for the 5 pictures. For high pressure levels, the velocity is low. However, a decrease of the chamber pressure provides a strong increase of the velocity and an expansion of the affected region.
24
Page 24 of 34
ip t cr us an M d te Ac ce p
Fig. 11 Evolution of the velocity field provided by the evaporated material for several
Argon gas pressures, in the near surface plane and viewed from the cross section of the melt pool.
The maximum velocity corresponding to the different pressure levels of Fig. 11 is plotted versus pressure on Fig. 12. As mentioned above, for high pressure levels, the maximum velocity remains quite low (i.e., below 100 m/s for a pressure above 800
25
Page 25 of 34
mbar). On the contrary, the maximum velocity increases strongly with the lowering of the pressure in the chamber (i.e., up to 300 m/s for the lowest pressure of 100 mbar). As first approximation, the maximum velocity seems almost inversely proportional with the
te
d
M
an
us
cr
ip t
pressure in the chamber.
Ac ce p
Fig. 12 Evolution of the maximum velocity of the evaporated material with the Argon gas pressure
The corresponding concentration of the metal vapor is shown in Fig. 13. An expansion of the vapor in the chamber is observed with the decrease of the ambient pressure. The evaporated material expands more freely in the rarefied atmosphere (100 mbar). On the contrary, at high pressure, the region of the high concentration of the evaporated material remains small and is characterized by a higher density and a slower propagation speed.
26
Page 26 of 34
ip t cr us an M d te
Ac ce p
Fig. 13 Concentration of the evaporated material (mass fraction) for several Argon-gas pressures
Thus, Argon acts not only as a shielding gas in the process chamber, but it also squeezes the evaporated material into a smaller volume. For low pressures, the amount of mater vaporized increases and the generated vapor expands significantly in the atmosphere. Interactions between the laser beam and the metal vapor requires complementary investigations in order to quantify the effect of the vapor on the heat transferred to the powder bed by the laser. Concerning this point, a test was performed with metal vapors
27
Page 27 of 34
generated by a plasma jet, a laser beam oriented through the generated vapor and a calorimeter to measure the power transferred by the laser. In the considered conditions, the power transferred was not affected by the presence of the metallic vapor and the
ip t
Argon plasma jet, so that the absence of interactions between the laser and the
cr
vaporized material can be assumed.
us
6. Conclusion
A study of the powder-laser-atmosphere interaction during the SLM process was
an
presented in this work. A 3D modeling approach was developed and validated progressively by experimental results. The model provides a first approach of the role of
M
the metal vapor generated in the interaction region of the laser beam and the powder bed. An effect of the pressure within the chamber was already shown: a decrease of the
d
pressure gives rise to a strong expansion of the metal vapor in the interaction region.
A high pressure environment may contribute to decrease the convective movements
Ac ce p
-
te
Based on these results, the following conclusion can be drawn:
in the gas side in the laser impact region.
-
A low pressure environment contributes to decrease the difference between the melting point and the vaporization point and can thus increase the rate of evaporation of the material. This high vaporization can also contribute to blow the powder particles. The combination of these two effects prevent building tracks at low pressure levels (i.e., below 100 mbar in the present case).
-
An adaptation of the parameters, such as a change of the scan speed, seems thus to be required in order to work under low pressure conditions. Nevertheless, it can
28
Page 28 of 34
become very sensitive due to the lowering of the difference between the melting and vaporization temperatures of the material with the decrease of the pressure in the chamber.
ip t
Finally, the use of a high pressure environment could contribute to decrease the
cr
convective movements in the atmosphere and to lower the vaporization rate of the
material. The use of a surrounding pressure above 1 bar could thus be tested in a future
us
work.
an
References:
Cédric André, 2007, modélisation quantitative du procédé de frittage sélectif par laser:
lausanne, Lausanne, EPFL.
M
relation paramètres/microstructure, PhD thèse, école polytechnique fédérale de
d
Cuadros F., I. Cachadina, W. Ahumada, 1996, Determination of Lennard-Jones
te
Interaction Parameters using a new procedure, Molecular Engineering: 319-325.
Ac ce p
Deissler R.G., Eian C.S., 1952, Investigation of effective thermal conductivities of powders, National advisory committee for aeronautics, Washington.
Donghua Dai, Dongdong Gu, 2014, Thermal behavior and densification mechanism during selective laser melting of copper matrix composites: Simulation and experiments, Materials & Design 55, 482-491.
Edson Costa Santos, Masanari Shiomi, Kozo Osakada, Tahar Laoui., 2006, Rapid manufacturing of metal components by laser forming. International Journal of Machine Tools and Manufacture 46, no 12-13,1459-1468.
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Ehsan,Ebrahimnia Bajestan,Hamid Niazmand, Weerapun Duangthongsuk, Somchai Wongwises, 2011, Numerical investigation of effective parameters in convective
Journal of Heat and Mass Transfer 54, no 19-20, 4376-88.
ip t
heat transfer of nanofluids flowing under a laminar flow regime,International
cr
Filali Mohamed, 2006, Conductivité thermique apparente des milieux granulaires
soumis à des contraintes mécaniques : modélisation et mesures, thèse, chapitre2.
us
Gusarov A. V., I. Smurov, 2009b,Two-dimensional numerical modelling of radiation transfer in powder beds at selective laser melting ». Applied Surface Science 255,
an
no 10, 5595-5599.
M
Gusarov A. V., I.Yadroitsev, Ph.Bertrand, I.Smurov, 2009a, Model of Radiation and Heat Transfer in Laser-Powder Interaction Zone at Selective Laser Melting,
d
Journal of Heat Transfer 131, no 7.
te
Gusarov A. V.,I. Smurov,2010,Modelling the interaction of laser radiation with powder
Ac ce p
bed at selective laser melting, Physics Procedia 5, Part B 381-394. Hirschfelder J. O. , C.F. Curtiss, R. B. Bird, 1954, Molecular Theory of gases and liquids, pp 1110-1112, Wiley, New York.
Hussein Ahmed, Liang Hao, Chunze Yan, Richard Everson, 2013, Finite element simulation of the temperature and stress fields in single layers built withoutsupport in selective laser melting, Materials & Design 52,638-647.
IAEA (International Atomic Energy Agency), 2008, Thermophysical Properties of Materials for Nuclear Engineering: A Tutorial and Collection of Data, Vienna, pages 168-170.
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Khairallah Saad A., Andy Anderson, 2014, Mesoscopic simulation model of selective laser melting of stainless steel powder, Journal of Materials Processing Technology 214, no 11,2627-36. doi:10.1016/j.jmatprotec.2014.06.001.
ip t
Kurian Antony, N. Arivazhagan, K. Senthilkumaran, 2014, Numerical and experimental
cr
investigations on laser melting of stainless steel 316L metal powders ,Journal of Manufacturing Processes 16, no 3,345-55.
us
Loh, Loong-Ee, Chee-Kai Chua, Wai-Yee Yeong, Jie Song, MahtaMapar, Swee-Leong Sing, Zhong-Hong Liu and Dan-Qing Zhang , 2015, Numerical investigation and
an
an effective modelling on the Selective Laser Melting (SLM) process with aluminum alloy 6061,International Journal of Heat and Mass Transfer 80,
M
288-300.
d
Mumtaz K.A., N. Hopkinson, 2010, Selective Laser Melting of thin wall parts using
te
pulse shaping, Journal of Materials Processing Technology, 210, no 2, 279-287. Rai1 R., JW Elmer, T. A. Palmer T. DebRoy, 2007, Heat transfer and fluid flow during
Ac ce p
keyhole mode laser welding of tantalum,Ti–6Al–4V, 304L stainless steel and vanadium, Journal of Physics D: Applied Physics.
Ruidi Li, YushengShi, Jinhui Liu, Huashan Yao, Wenxian Zhang, 2009, Effects of processing parameters on the temperature field of selective laser melting metal powder, Powder Metallurgy and Metal Ceramics 48, no 3-4, 186-195.
Stoffregen H., Fischer J., Siedelhofer C., Abele E., 2011, Selective Laser Melting of porous Structures , Proceedings of the 22nd Annual International Solid Free form Fabrication (SFF) Symposium, University of Texas, Austin, p. 680.
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Takamichi Iida, Roderick I.L. Guthrie, 1993, The physical properties of liquid Metal, Clarendon Press, Oxford. Verhaeghe F., T. Craeghs, J. Heulens, L. Pandelaers, 2009,A pragmatic model for
ip t
selective laser melting with evaporation , Acta Materialia 57, no 20 6006-6012.
cr
Yadroitsev I., I.Smurov, 2011, Surface Morphology in Selective Laser Melting of Metal
us
Powders, Physics Procedia, 264-270.
Yadroitsev I., Ph. Bertrand, I. Smurov, 2007, Parametric analysis of the selective laser
an
melting process, Applied Surface Science 253 Elsevier , pp 8064-8069. Zhang Baicheng, Lucas Dembinski, Christian Coddet, 2013b, The study of the laser
M
parameters and environment variables effect on mechanical properties of high compact parts elaborated by selective laser melting 316L powder, Materials
d
Science and Engineering: A 584,21-31. doi:10.1016/j.msea.2013.06.055.
te
Zhang Baicheng, Hanlin Liao, Christian Coddet, 2013a, Microstructure evolution and
Ac ce p
density behavior of CP Ti parts elaborated by Self-developed vacuum selective laser melting system , Applied Surface Science.doi:10.1016/j.apsusc.2013.04.090.
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Figure and Table captions Figure 1. Modelling steps of a 316L stainless steel powder bed irradiation by a laser beam (a) different regions of the calculation domain, (b) mesh, (c) laser irradiation of
ip t
the powder bed. Figure 2. Gaussian distribution of the thermal flux.
cr
Figure 3.Vacuum selective laser melting machine developed at LERMPS laboratory.
us
Figure 4. 316L stainless steel tracks manufactured with several scanning speeds (1-4) at several pressure levels (A to F) of Argon gas.
an
Figure 5. Effect of the Argon gas pressure on SS 316L tracks prepared by the SLM process (a) 995 mbar (b) 500 mbar (c) 250 mbar (d) 100 mbar (e) 10 mbar (f)1 mbar –
M
veff=0.1 m/s.
Figure 6. Temperature distribution (K) for a single track scanned linearly (veff =
d
0.1 m/s, p= 1 bar).
Ac ce p
melt pool morphology.
te
Figure 7. Measurement method of the melt pool dimensions in the simulation of the
Figure 8. Cross section of a SS 316L track obtained with a scanning speed of 0.1 m.s-1 at a pressure of 995 mbar.
Figure 9. Numerical and experimental results of the melt pool depth and width. Figure 10. Velocity vectors (m/s) of the evaporated material for an Argon-gas pressure of 100mbar, in the near surface plane. Figure 11. Evolution of the velocity field provided by the evaporated material for several Argon gas pressures, in the near surface plane and viewed from the cross section of the melt pool.
33
Page 33 of 34
Figure 12. Evolution of the maximum velocity of the evaporated material with the Argon gas pressure. Figure 13. Concentration of the evaporated material (mass fraction) for several Argon-
ip t
gas pressures.
us
Table 2. Lennard-Jones parameters for Fe and Ar
cr
Table 1. Thermo-physical properties of 316L
an
Table 3. Scan speeds applied to elaborate the single lines
Table 4. Argon gas pressure selected to elaborate the single lines
M
Table 5. Vaporization temperature (Tv) of 316L for different pressure levels
Ac ce p
te
pressure of 1 bar.
d
Table 6. Dimensions of tracks obtained by numerical modelling – Case of a chamber
34
Page 34 of 34