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Fine spherical powder production during gas atomization of pressurized melts through melt nozzles with a small inner diameter
Powder Technology 356 (2019) 759–768
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Fine spherical powder production during gas atomization of pressurized melts through melt nozzles with a small inner diameter Xing-gang Li a,b,c,⁎, Qiang Zhu b,c, Shi Shu d, Jian-zhong Fan d, Shao-ming Zhang d a
Academy for Advanced Interdisciplinary Studies, Southern University of Science and Technology (SUSTech), Shenzhen 518055, China Department of Mechanical and Energy Engineering, Southern University of Science and Technology (SUSTech), Shenzhen 518055, China Shenzhen Key Laboratory for Additive Manufacturing of High-performance Materials, Shenzhen 518055, China d National Engineering & Technology Research Center for Nonferrous Metal Matrix Composites, General Research Institute for Nonferrous Metals (GRINM), Beijing 100088, China b c
a r t i c l e
i n f o
Article history: Received 4 November 2018 Received in revised form 27 June 2019 Accepted 7 September 2019 Available online 09 September 2019 Keywords: Additive manufacture Metal powder Gas atomization Melt nozzle Mass median diameter Satellite particle
a b s t r a c t A pressure-gas atomizer was developed, in which the melts were pressurized through melt nozzles with a small inner diameter, aiming for a small mass median diameter (MMD, d50,3) and high productivity of fine spherical powders. The maximum melt flow resistance in a melt nozzle, defined as the sum of the capillary resistance and viscous pressure drop, was analyzed by varying the inner diameter of the melt nozzle (D0). The calculation results indicate that the maximum melt flow resistance increases quickly with the decrease of D0, and varies in an order of 100–102 kPa for different metal melts when D0 reduces from 4.0 mm to 0.5 mm. Atomization runs with three kinds of aluminium (Al) alloys were accomplished using melt nozzles with different inner diameters in a pilot plant whereby an over-pressure in a range of Δpl = 30–45 kPa can be maintained on the melts to enhance the melt flowing in the melt nozzle. The experimental results indicate that the atomization efficiency can be well improved by reducing the inner diameter of the melt nozzle, which resulted in a small MMD, narrow particle size distribution and high fine powder yield. For Al-I alloy powders, when the inner diameter of the melt nozzle reduces from D0 = 3 mm to D0 = 1 mm, the particle MMD reduces from d50,3 = 86.13 μm to d50,3 = 40.42 μm, and the powder yield b53 μm increases from 27.60% to 62.57%. For Al-III alloy powders, when the inner diameter of the melt nozzle reduces from D0 = 4 mm to D0 = 2 mm, the particle MMD reduces from d50,3 = 120.10 μm to d50,3 = 54.82 μm and the powder yield b53 μm increases from 20.70% to 48.20%. Moreover, the satellite particles and lamellae sticking on the particle surface were reduced when a melt nozzle with a small inner diameter was employed in a gas atomization process. © 2019 Elsevier B.V. All rights reserved.
1. Introduction The emerging technological field of Additive Manufacturing (AM) typically needs specifically tailored materials and particles, especially metal powders with precisely optimized size, shape and morphology. For example, fine spherical metal powders in a size range of b53 μm are usually preferred during selective-laser-melting (SLM) processing of metal parts. Moreover, powder agglomeration and hollow particles should be avoided to guarantee a good processability and the final product property. At present fine spherical powders for metal AM are produced primarily by gas atomization technology. Two most widely used gas atomization systems in industry are free-fall type and close-coupled type, in both of which a hot molten metal stream, usually in a form of a round jet, is disintegrated into droplets by high pressure cold gas jets ⁎ Corresponding author at: Academy for Advanced Interdisciplinary Studies, Southern University of Science and Technology (SUSTech), Shenzhen 518055, China. E-mail address: [email protected] (X. Li).
https://doi.org/10.1016/j.powtec.2019.09.023 0032-5910/© 2019 Elsevier B.V. All rights reserved.
surrounding the molten metal stream, and subsequently the atomized metallic droplets are subjected to high cooling rates and deep undercooling in a high speed atomization gas flow. In a free-fall configuration, the molten metal falls freely for some distance in the direction of gravity before being hit by the atomization gas, and this configuration generally produces relatively coarse particles. In a close-coupled configuration, gas jets hit the molten metal immediately after it comes out of the melt nozzle, and this configuration is frequently used for the production of metal powders with mass median diameter (MMD, d50,3) ranging from 10 to 100 μm. The above gas atomization techniques are often characterized with wide particle size distribution and high specific gas consumption. Larger or smaller particle size ranges are usually separated which, in turn, reduces the productivity of the process. A wide size distribution of the raw metal powder means a low productivity. For example, as for the AlSi10Mg powders for the SLM process, the powder yield in a size range b 53 μm is usually only 30 vol% in a conventional gas atomization process. Therefore, metal powder costs currently form a high fraction of additively manufactured metal parts. Research on gas atomization process has put emphasis on spherical powders with a
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small mass median diameter and high productivity. Atomization efficiency or fine powder yield can be improved mainly by three kinds of approaches: 1) optimization of gas nozzle design and the arrangement between gas nozzle and melt nozzle, 2) augmenting atomizing gas properties, and 3) manipulating the melt properties. Most of research work has been focused on the first two aspects. Significant development efforts based on gas flow dynamics have resulted in the advances in gas nozzle design and the improvement in the arrangement between gas nozzle and melt nozzle. Gas nozzles with a convergent-divergent (C-D) [1] or De Laval nozzle [2] contour in gas flow channels, as in rocket exhaust nozzle design, were widely studied to increase atomization efficiency and uniformity and to reduce the required gas supply pressures. Based on a close-coupled configuration, Anderson & Terpstra [1] developed high pressure gas atomization (HPGA) nozzles with discrete convergent-divergent jets, which exhibited higher atomization efficiency compared with those with straight bore jets. The wetting by the melt (lick back), which is a cause of gas nozzle failure, can be avoided by adjusting the angle matching or the relative placement between gas nozzle and melt nozzle to eliminate or restrict the local flow separation around the tip of the melt nozzle [3,4]. The atomization efficiency can be well improved by augmenting the properties of the atomization medium, such as gas pressure and temperature. HJE Co, Inc. [5] and PSI Ltd. [6] developed hot gas atomization techniques. The sonic velocity of a gas is directly proportional to the square root of the absolute temperature of the gas. Thus, increasing atomization gas temperature can lead to substantial kinetic energy available to disrupt the melt stream into a finer distribution of droplets, resulting in improved atomization efficiency. The efficiency of an atomization process can be alternatively optimized by increasing the specific liquid surface energy before atomization [7]. Within twin fluid (gas) atomization, the fragmentation must take place in a region where the velocity difference between the atomization gas and the liquid is highest. Therefore, the liquid must be transformed and transported into an area where the maximum velocity difference occurs. These aspects can be realized in a prefilming hybrid atomizer where the disintegration process can be subdivided into two steps: a prefilming and a gas atomization step [8]. During the prefilming step, a large specific liquid surface area is generated to prepare an efficient secondary disintegration step by gas atomization. Czisch and Fritsching [9] developed a prefilming hybrid atomizer, which was based on a combination of a single-fluid rotary prefilming concept and an external mixing twin-fluid atomizer. Within this specific design, the feed material is first spread out by a spinning disc operated in sheet formation mode, forming a free-flowing liquid film, which was subsequently conveyed into the most effective atomization zone by the local gas flow and atomized by a high speed gas flow from the external mixing gas atomizer. A different hybrid atomization process for particle production was developed to hybridize effectively gas atomization with centrifugal atomization in Minagawa et al. [10]. In this hybrid atomizer configuration the melt spray from an external mixing twin fluid nozzle impinges onto a rotating disc, resulting in a controlled film that finally acts as centrifugal atomizer for fine droplets with narrow particle size distribution within metal particle production. Another prefilming hybrid atomizer concept, designated as pressure-swirl-gas-atomization (PSGA), was developed based on a combination of pressure-swirl-atomization (PSA) as a prefilming step and gas atomization [11]. Within a PSGA process, the pressurized melt initially enters a pressure swirl nozzle and leaves the nozzle exit as a conical hollow cone film due to its swirl and centrifugal forces; as the second step, an external mixing gas nozzle is employed to disintegrate the liquid film, ligaments, and any large droplets. The PSGA atomizer design, especially the pressure swirl nozzle, is sophisticated and has been used to produce fine spherical powders of low melting point metals with small median diameter and narrow size distribution
[12,13], as well as particulate reinforced metal-matrix-composite (MMC) particles [14]. Atomization of pressurized non-metallic liquids has been widely investigated, for example, in the field of fuel combustion in a engine [15,16], where the liquid fuel is injected into a combustion chamber under a high injection pressure and then atomized into very tiny droplets, thus enhancing the fuel combustion efficiency. As for the atomization of pressurized metal melts, besides the pressure-swirl-gasatomization (PSGA) technique discussed above, a periodic overpressure is often imposed on the melt in a Drop-on-Demand (DOD) process [17,18], whereby the melt is periodically pushed through a melt nozzle, forming a discontinuous chain of mono-sized droplets, which subsequently solidify as powder particles. Based on the authors' knowledge, Raman and Surette [19] firstly proposed the method, i.e., gas atomization of pressurized metal melts in a form of round jet for powder production in their invention patent, and presented preliminary experimental results, which indicated that good yields of fine particles were obtained by imposing over-pressure on the melt (Sn-5%Pb) when it went through a melt nozzle with an inner diameter of D0 = 2 mm. However, since then, few studies have been reported on this method. Although, in the practice of metal powder production by gas atomization, it may be very common to impose a slight over-pressure on the melt to enhance the melt flow, few studies have been performed to systematically investigate the influence of over-pressure and nozzle inner diameter on the melt flow behavior in the melt nozzle. In this paper, the pressure swirl nozzle in a PSGA process is replaced by a simple melt nozzle with a cylindrical channel, through which the pressurized melt goes into the spray chamber as a round liquid jet. Attempts to reduce the particle size of metal powders concentrate on reducing the inner diameter of the melt nozzle (D0). With this approach, a very thin liquid jet is expected to form at the melt nozzle orifice, which can enhance the subsequent gas atomization efficiency, thus resulting in an improved yield of fine particles. Firstly, the melt flow resistance in a melt nozzle is analyzed by varying the inner diameter of the melt nozzle. Next, an introduction is given to the powder plant and the powder production process, in which an over-pressure can be maintained on the melts as a driving force to enhance the melt flowing in the melt nozzle. Finally, three kinds of Al-alloys are gas-atomized using melt nozzles with different inner diameters, and the asatomized powders are characterized with emphasis on the properties such as mass median diameter, particle size distribution span, fine powder yield and powder morphology. 2. Theory/calculation In a conventional gas atomization process, the driving forces for the melts to go through a melt nozzle include the aspiration pressure formed in front of the melt nozzle by high speed gas jets and the gravitational force. In this case, a melt nozzle with a large diameter is often used to ensure that the melt can go through the melt nozzle by overcoming the capillary force and the viscous force between the melts and the inner wall of the melt nozzle, thus avoiding clogging in the melt nozzle but resulting in coarse powder particles. Lubanska [20] has established the correlation with mass median diameter (MMD, d50,3) of gas-atomized powders that demonstrates the benefits of reducing the inner diameter of the melt nozzle (D0) for the production of fine powders: d50;3 ¼ D0 K lub
0:5 νl 1 1 1þ GMR ν g We
ð1Þ
where
We ¼
ðΔU Þ2 ρl D0 σ
ð2Þ
X. Li et al. / Powder Technology 356 (2019) 759–768
is the Weber number, which represents the ratio of inertial force including the density of the melt (ρl) to the surface tension of the melt (σ), Klub is a constant that must be determined for a particular atomizer, νl and νg are the kinematic viscosities of the melt and the gas, respectively, GMR is the ratio of gas to melt mass flow rate, and ΔU is the relative velocity between the melt and the gas. Take Eq. (2) into Eq. (1), which can be rearranged into. 0:5 1 d50;3 ¼ K lub Q D0 1 þ GMR
ð3Þ
where Q is related to the properties of atomized melts and atomizing gas, as. Q¼
νl σ ν g ðΔU Þ2 ρl
!0:5 ð4Þ
For specific atomizer and material systems, Eq. (3) indicates that the variation of the particle mass median diameter (MMD, d50,3) is only related to the melt nozzle diameter (D0) and the GMR. The MMD is directly proportional to the square root of D0, which provides a theoretical basis to reduce the particle size of metal powders using a melt nozzle with a small inner diameter in a gas atomization process. However, with the decrease of D0, the melt flow resistance in the melt nozzle may increase, which cause a reduction in melt mass flow rate and even melt clogging in the channel. The next will give a quantitative description of the melt flow resistance related to the inner diameter of the melt nozzle. 2.1. Capillary resistance Fig. 1 (a) and (b) depict two cases of the clogged inner-channels of the melt nozzle by melt columns, respectively. The shape of the innerchannel is cylindrical. Assuming that the melt is in a full liquid state in the clogged melt nozzle, a balance is reached between the capillary force (Fs), viscous force (Fv) and gravity. (See Table 1.) The capillary pressure (ps) can be obtained by the Young-Laplace equation [23] as: ps ¼
4σ cosθ D0
ð5Þ
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Table 1 Surface tension and viscosity at melting point [21,22]. Metal
Sn
Zn
Mg
Al
Cu
Ni
Fe
Melting point (°C) Surface tension (N/m) Viscosity (mPa•s)
232 0.550 1.85
420 0.789 3.85
650 0.577 1.25
660 0.871 1.30
1083 1.330 4.0
1455 1.796 4.90
1530 1.862 5.5
the melt flowing in the melt nozzle, as shown in Fig. 1 (a); if the contact angle θ N 90°, the value of the capillary pressure (ps) is negative, which means that the capillary force (Fs) will prevent the melt flowing in the melt nozzle, as shown in Fig. 1 (b). For gas atomization of metal melts, the melt nozzles are usually made from refractory materials, such as graphite, alumina, magnesia, zirconia and boron nitride. In most cases, the surfaces made from these materials are poorly wetted by metal melts, and the contact angles (θ) between metal melts and solid surfaces are larger than 90°. This means that the capillary force (Fs) will act in an opposite direction against the melt flow in a melt nozzle. According to Eq. (5), the capillary resistance (ps) is calculated for the melts of several common metals. The calculated results are plotted against the inner diameter of the melt nozzle (D0) in Fig. 2 and against the liquid/solid contact angle (θ) in Fig. 3, respectively. Fig. 2 indicates that the capillary resistance increases quickly with the decrease of the inner diameter of the melt nozzle (D0). In the case of a constant liquid/solid contact angle θ = 150°, the capillary resistance (ps) varies in an order of 100-101 kPa when the inner diameter of the melt nozzle (D0) reduces from 4.0 mm to 0.5 mm. Fig. 3 indicates that if the solid surface is non-wetted by the melts, the capillary resistance increases with the increase of the liquid/ solid contact angle (θ). 2.2. Viscous pressure drop Assuming that the melts are incompressible and Newtonian fluids in a laminar flow flowing through a melt nozzle, the viscous pressure drop (pv) can be described using the Hagen-Poiseuille law [23] as: pv ¼
8μLQ πR40
ð6Þ
where μ is the dynamic viscosity, L is the inner-channel length of the melt nozzle, R0 is the inner radius of the melt nozzle, and Q is the melt
which relates the capillary pressure (ps) at the melt/gas interface to the inner diameter of the melt nozzle (D0), the surface tension of the melt (σ), and the liquid/solid contact angle (θ). Eq. (5) indicates that the capillary pressure (ps) is directly proportional to the surface tension of the melt (σ), and inversely related to the inner diameter of the melt nozzle (D0). If the contact angle θ b 90°, the value of the capillary pressure (ps) is positive, which means that the capillary force (Fs) is a driven force for
Melt
Melt
D0
Fs
θ Fs
D0
Fs
θ Fs
Fig. 1. Illustration of melt clogging inner-channel in melt nozzle, liquid/solid contact angle (a) θ b 90°, (b) θ N 90°.
Fig. 2. Capillary resistance vs. melt nozzle inner diameter (D0) in the case of a constant liquid/solid contact angle (θ) for different metals.
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viscous pressure drop varies in an order of 100-102 kPa when the inner diameter of the melt nozzle (D0) reduces from 4.0 mm to 0.5 mm. 2.3. Melt flow resistance
Fig. 3. Capillary resistance vs. liquid/solid contact angle (θ) in the case of a constant melt nozzle inner diameter (D0) for different metals.
volumetric flow rate, which is defined as: Q ¼ Ul • πR20
ð7Þ
where Ul is the melt normal discharge velocity at the nozzle exit. Combine Eqs. (6) and (7) and substitute R0 with D0/2 to obtain:
The capillary resistance is usually important in the start phase of the process when the melt nozzle is not completely wetted. During a spray run the melt nozzle should be wetted completely. However, in many cases the melt clogging happens at the very beginning of the process before the melt nozzle is not completely wetted, especially when a melt nozzle with a small inner diameter is employed. In some cases, when the melt clogging happens during a spray run, the solidified melt column in a melt nozzle shows the solidification front with a curved surface, which implies that the capillary force plays a role. Therefore, the maximum resistance that impedes the melt flow in a melt nozzle could be defined as the sum of the capillary resistance (ps) and the viscous pressure drop (pv). As shown in Fig. 5, the maximum melt flow resistance is plotted against the inner diameter of the melt nozzle (D0). The melt nozzle with a smaller inner diameter has larger flow resistance. For a constant liquid/solid contact angle of θ = 150° and a normal discharge velocity of Ul = 5 m/s, the maximum resistance that impedes the melt flow in a melt nozzle, with an inner-channel length of L = 40 mm, varies in an order of 100-102 kPa when the inner diameter of the melt nozzle reduces from 4.0 mm to 0.5 mm. Therefore, an additional driving force is necessary for the melt to flow well through a nozzle with a small inner diameter. In this paper, an over-pressure is imposed on the melt in a tundish to enhance the melt flowing in a melt nozzle by filling gas into the furnace and then building a pressure difference between the furnace and the spray chamber. 3. Powder plant
pv ¼
32μLU l D20
ð8Þ
which indicates that the viscous pressure drop is directly proportional to the dynamic viscosity, the nozzle inner-channel length and the melt normal discharge velocity, and inversely proportional to the square of the inner diameter of the melt nozzle. Fig. 4 shows the calculated results of the viscous pressure drop (pv) plotted against the inner diameter of the melt nozzle (D0), which indicates that the viscous pressure drop increases quickly with the decrease of the inner diameter of the melt nozzle (D0). For a normal discharge velocity of Ul = 5 m/s and an inner-channel length of L = 40 mm, the
Fig. 4. Viscous drop pressure (pv) vs. melt nozzle inner diameter (D0) for different metals, L - inner-channel length of melt nozzle, Ul - melt discharge velocity.
The powder production facility used in this study mainly consists of melting furnace, spray chamber, vacuum system and powder collection system. The melting furnace is a pressure vessel which can withstand an over-pressure (Δpl) of at least 100 kPa in the present experiments. The furnace cover and body are locked and sealed through a design as in a pressure cooker. The furnace and the spray chamber are two separate containers and connect with each other only through a melt nozzle at the bottom of the tundish. Both of the furnace and the spray chamber are firstly evacuated to a vacuum level of 10−2-10−1 Pa, and then refilled with nitrogen gas (N2) to a pressure of 1 atm prior to melting
Fig. 5. Melt flow resistance vs. melt nozzle inner diameter (D0) for different metals, L inner-channel length of melt nozzle, Ul - melt normal discharge velocity, θ - liquid/solid contact angle.
X. Li et al. / Powder Technology 356 (2019) 759–768
to reduce the oxygen content. When the preset superheat of the melt in the crucible is achieved, refill the nitrogen gas into the furnace to the preset over-pressure, and then transfer the melt from the crucible to the tundish. At last the pressurized melt is pushed through the melt nozzle into the spray chamber, where the melt jet is disintegrated into droplets by the high speed gas jets. The gas nozzle consists of 20 straight bore jets with a diameter of 1.1 mm at each outlet. The diameter of the circumference, along which the 20 gas-nozzle outlets are uniformly distributed, is DG = 28 mm, and an atomization angle of ɑ = 45° was designed for the gas nozzle (see Fig. 6 (b)). The powder collection system consists of cyclone separator and powder collector. Most of the coarse particles are deposited into the powder collector at the bottom of the spray chamber, while the rest of the particles are carried into the cyclone separator by the gas flow and further deposited into the powder collector at the bottom of the cyclone separator. The exhaust from the cyclone separator is rigorously filtered and at last emitted into the atmosphere. The powders in different powder collectors are mixed in a powder mixer. 4. Experimental design and procedure Three kinds of Al-alloys, designated as Al-I, Al-II and Al-III, respectively, were atomized with a discrete jet gas nozzle using commercial purity nitrogen at 2.0 MPa. The nominal composition of the alloys is listed in Table 2. In contrast with the Al-II alloy, the Al-III alloy contains extra 18-20 wt% Si and 5-7 wt% Fe, which implies that the Al-III melt would be more viscous than the Al-II melt. As indicated in Fig. 6 (a), the crucible and the tundish are made from commercial purity graphite, while the melt nozzle is made from high purity graphite. Each 7 kg alloy charge was induction superheated by 150 K over the melting point in the crucible. The melt nozzles were tested by varying the inner diameter from 4 mm to 1 mm. It should be mentioned that a higher over-pressure would result in a higher melt mass flow rate, thereby reducing the fine powder yield. Therefore, the over-pressure imposed on the top of the melt in this study lies in a range of 3045 kPa, which is on the same level as the calculated results of the melt flow resistance in melt nozzles (see Fig. 5). In case of melt freezing in the melt nozzle, the tundish is also superheated by 150 K over the melting point temperature for each alloy by resistance heating; moreover, a separate heating device was employed to heat the melt nozzle to about 500 °C during atomization. The vertical distance between the melt nozzle exit and the gas nozzle outlet is H = 5-10 mm (see Fig. 6 (b)). Table 3 summarizes the experimental parameters of the gas atomization experiments of three Al-alloys. The melt atomization duration mainly depends on the inner diameter of the melt nozzle and the over-pressure imposed on the melt in the furnace, and ranges between 2 min - 6 min. The melt mass flow rate (Ml), which is defined as the ratio of the powder weight finally obtained
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Table 2 Nominal composition (wt%) of Al-alloys tested in experiments.
Al-I (AlSi10Mg) Al-II (2009 alloy) Al-III
Si
Mg
Cu
Fe
Al
9.0–11.0 – 18.0–20.0
0.2–0.5 1.0–1.6 1.0–1.6
– 3.2–4.4 3.2–4.4
– – 5.0–7.0
Bal. Bal. Bal.
in powder collectors to the melt atomization duration, generally varies from 80 kg/h to 210 kg/h. The gas flow measurements were performed using the method developed in Anderson & Terpstra [1], in which both the initial gas cylinder pressure and the final gas cylinder pressure after a gas flow period were recorded at an ambient-temperature equilibration state of the supply cylinder. In this study, sixteen gas cylinders were assembled in series with each other to provide gas for melt atomization. For all the tests here, the atomization gas pressure was set at 2.0 MPa, which resulted in a gas mass flow rate of about 200 kg/h. Each run of atomized Al-alloy powders were pre-screened to b500 μm, eliminating irregular particulates and atomization debris. For characterization samples were taken from the powders of each run and analyzed by laser diffraction and scanning electron microscopy (SEM). 5. Experimental results and discussion 5.1. Particle size distribution The Al-alloy powders were produced by gas atomization of pressurized melt jets through the melt nozzles of different inner diameters. Fig. 7 shows the particle size distributions (cumulative volume or mass) of the powders of Al\\I and Al-III alloys. The powder samples were wetdispersed and the measurements were carried out with a laser light diffraction instrument (Beckman Coulter Particle Size Analyzer LS 13320). As the inner diameter of the melt nozzle decreases, the particle size distribution curve shrinks and moves towards the region of small particle size. The results indicate that the particle size distribution is greatly influenced by the inner diameter of the melt nozzle, especially when it reduces to b3 mm. 5.2. Mass median diameter The values of particle mass median diameter (MMD, d50,3) were extracted from the particle size distributions and plotted against the inner diameter of the melt nozzle (D0), as shown in Fig. 8. For the three runs of the atomized Al-I alloy powders, the particle MMD is 86.13 μm for D0 = 3 mm, 60.45 μm for D0 = 2 mm and 40.42 μm for D0 = 1 mm, respectively, indicating a decrease of ~53% in MMD when the inner diameter of the melt nozzle reduces from D0 = 3 mm to D0 = 1 mm. For the three runs of the atomized Al-III alloy powders, the MMD is 120.10 μm
Fig. 6. (a) Illustration of melt transferring, pressurizing, guiding and gas atomization process; (b) Sketch of nozzle design under operation.
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Table 3 Experimental parameters in gas atomization of different Al-alloys. Alloy
D0 (mm)
Δpl (kPa)
Ml (kg/h)
pg (MPa)
Mg (kg/h)
GMR (−)
ΔTl (K)
Al-I Al-I Al-I Al-II Al-III Al-III Al-III
3 2 1 3 4 3 2
30 35 40 30 35 40 45
139.2 102.27 87.97 120.89 203.69 184.05 104.60
2.0 2.0 2.0 2.0 2.0 2.0 2.0
200 200 200 200 200 200 200
1.44 1.96 2.27 1.65 0.99 1.09 1.91
150 150 150 150 150 150 150
Note: D0 - melt nozzle inner diameter, Δpl - over-pressure on the top of the melt, Ml - melt mass flow rate, pg - atomization gas pressure, Mg - gas mass flow rate, GMR - gas-to-metal mass flow rate ratio, ΔTl - melt superheat.
for D0 = 4 mm, 102.20 μm for D0 = 3 mm and 54.82 μm for D0 = 2 mm, respectively, indicating a decrease of ~54% in MMD when the inner diameter of the melt nozzle reduces from D0 = 4 mm to D0 = 2 mm. As described in Eq.(3), the decrease of the particle MMD in this study could be mainly attributed to the nozzle geometrical effect and the GMR increase, both of which are induced by the reduction in D0. The Al-III alloy was atomized under similar GMRs by employing the melt nozzle of D0 = 4 mm and D0 = 3 mm, respectively (see Table 3). The former resulted in a ~18% larger particle MMD than the latter, which indicates the significant nozzle geometrical effect on the variation of the particle MMD. In the case of D0 = 3 mm, the atomization of the Al-III alloy resulted in a ~16% larger particle MMD than the Al-II alloy, possibly due to the higher viscosity of the Al-III alloy containing excessive Si and Fe. The results of the particle MMD have also been collected and plotted as a function of the GMR in Fig. 9, which indicates that the particle MMD decreases with the increase of the GMR. In the present work, the reduction in the melt nozzle diameter led to the decrease of melt mass flow rate, while the over-pressure imposed on the melt was selected as low as possible to maintain a low melt mass flow rate, thus raising the GMR. Fig. 10 shows the results of the particle MMD as a function of the factor [D0(1 + 1/GMR)]^0.5 in Eq. (3), which represents the comprehensive effect of the simultaneous changes of melt nozzle diameter and the GMR on the variation of the particle MMD. Given constant values of KlubQ, the variation of the particle MMD with the factor [D0(1 + 1/ GMR)]^0.5 according to Eq. (3) is shown in Fig. 10, which indicates that the values of the particle MMD obtained from the present work are located between the two curves of KlubQ = 35 and KlubQ = 45
Fig. 8. Mass median diameter (MMD, d50,3) vs. melt nozzle inner diameter (D0), measurements conducted by laser light diffraction (wet dispersion).
based on Eq. (3). Here the variation in KlubQ value reflects the change in material properties. 5.3. Particle size distribution span The size distribution span of the powder particles is defined as the diameter ratio. span ¼
d90;3 −d10;3 d50;3
ð9Þ
A narrower distribution means a higher yield at a specific mass median diameter. The span values were derived from the particle size distribution and plotted against the inner diameter of the melt nozzle (D0), as shown in Fig. 11. The results indicate that the particle size distribution span generally reduces with the decrease of D0, especially when D0 b 3 mm. For example, the size distribution span of the atomized Al\\I alloy powder decreases from span = 2.16 to span = 1.52 when the inner diameter of the melt nozzle varies from D0 = 3 mm to
Fig. 7. Cumulative mass distribution for different melt nozzle inner diameters (D0), measurements conducted by laser light diffraction (wet dispersion).
X. Li et al. / Powder Technology 356 (2019) 759–768
Fig. 9. Mass median diameter (MMD, d50,3) vs. the ratio of gas to melt mass flow rate (GMR), measurements conducted by laser light diffraction (wet dispersion).
D0 = 1 mm; the size distribution span of the atomized Al-III alloy powder decreases from span = 2.75 to span = 2.05 when the inner diameter of the melt nozzle varies from D0 = 4 mm to D0 = 2 mm. In the case of D0 = 3 mm, the span values are 2.16, 1.73 and 2.74, respectively, for Al-I, Al-II and Al-III alloys.
5.4. Fine powder yield The powder yield b53 μm is defined here as the cumulative mass or volume fraction at the size of 53 μm in the particle size distribution curve, which is measured by laser light diffraction. Fig. 12 is a plot of the powder yield b53 μm as a function of the inner diameter of the melt nozzle (D0). An obvious increase in the fine powder yield could be observed with the decrease of D0. For the atomized Al-I alloy powders, the powder yields b53 μm are 27.60% for D0 = 3 mm, 41.50% for D0 = 2 mm and 62.57% for D0 = 1 mm, respectively; for the atomized Al-III alloy powders, the powder yields b53 μm are 20.70% for D0 = 4 mm, 25.90% for D0 = 3 mm and 48.2% for D0 = 2 mm, respectively. For the Al-II alloy, a 26.40% powder yield b53 μm is obtained by use of a melt nozzle with D0 = 3 mm, which is similar to the results of the Al-I and Al-III alloys obtained by use of the same melt nozzle.
Fig. 10. Mass median diameter (MMD, d50,3) vs. the factor [D0(1 + 1/GMR)]^0.5 in Eq. (3), measurements conducted by laser light diffraction (wet dispersion)，curves of KlubQ = 45 and KlubQ = 35 calculated according to Eq. (3).
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Fig. 11. Powder size distribution span vs. melt nozzle inner diameter (D0), measurements conducted by laser light diffraction (wet dispersion).
5.5. Powder morphology SEM images of the Al-I alloy powder samples in a size range of 38 and 53 μm are presented in Fig. 13. It can be observed that the powders generated with a melt nozzle of D0 = 3 mm exhibit a lot of satellites (small particles) sticking on the surface of larger particles compared with those generated with a melt nozzle of D0 = 1 mm. Moreover, the micrographs of the Al-I alloy powder samples in a size range of b38 μm, as shown in Fig. 14, indicate that the particles generated with a melt nozzle of D0 = 1 mm are more spherical and smoother. SEM images of the Al-III alloy powder samples, generated with melt nozzles of D0 = 4 mm and D0 = 2 mm, respectively, are presented in Fig. 15. A few of satellites sticking on the particle surface are observed in both of the powder samples. But the particle surface of the latter seems more spherical and smoother. SEM images of the Al-II and Al-III alloy powder samples, generated with a melt nozzle of D0 = 3 mm, are presented for comparison in Fig. 16, which indicates satellites and lamellae sticking on the particle surface in both of the powder samples. The gas-atomized powders in this study are spherical in general. However, the phenomena of satellites and lamellae sticking on the particle surface, especially in the powders generated using a melt nozzle
Fig. 12. Powder yield b53 μm vs. melt nozzle inner diameter (D0), measurements conducted by laser light diffraction (wet dispersion).
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Fig. 13. SEM images of Al-I alloy powders with a size range 38 μm b d b 53 μm: (a) melt nozzle inner diameter D0 = 3 mm, (b) melt nozzle inner diameter D0 = 1 mm.
with a larger inner diameter, are often observed, which reduce the powder sphericity and smoothness, and further reduce the powder flowability and apparent density. The formation of those satellites and lamellae on the particle surface should be ascribed to the frequent particle/droplet collisions in the atomization spray process, where the gas/droplet momentum exchange and heat transfer are mainly influenced by the droplet size. Small droplets, which would firstly solidify and are more easily accelerated by the high-speed atomizing gas, collide with large droplets that are not fully solidified. The collision may cause adhesion, partial penetration or fusion between droplets, leading to the formation of satellite particles sticking on the surface of the large
particles; or cause droplet deformation, leading to the formation of lamellae covering the particle surface. In this study, the powders, generated using a melt nozzle with a smaller inner diameter, usually exhibit a narrower or more uniform size distribution, which implies a lower particle/droplet collision frequency in the atomization spray process because of similar particle/droplet inertia; in addition, the reduction in the inner diameter of the melt nozzle leads to a higher GMR (see Table 3), which enhances the droplet solidification. Therefore, the powders generated using a melt nozzle with a smaller inner diameter are more spherical and smoother with fewer satellites and lamellae sticking on the particle surface.
Fig. 14. SEM images of Al-I alloy powders with a size range d b 38 μm: (a) melt nozzle inner diameter D0 = 3 mm, (b) melt nozzle inner diameter D0 = 1 mm.
Fig. 15. SEM images of Al-III alloy powders with a size range d b 53 μm: (a) melt nozzle inner diameter D0 = 4 mm, (b) melt nozzle inner diameter D0 = 2 mm.
X. Li et al. / Powder Technology 356 (2019) 759–768
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Fig. 16. SEM images of powders with a size range d b 53 μm, melt nozzle inner diameter D0 = 3 mm: (a) Al-II alloy powders, (b) Al-III alloy powders.
5.6. Concerned issues 5.6.1. Powder production rate For powder production by gas atomization, the melt mass flow rate (kg/h) is an important process parameter, which determines the powder throughput / production rate. The reduction in the melt nozzle inner diameter will lead to the increase of the melt flowing resistance in the nozzle, thus lowering the melt mass flow rate. However, in this paper, this deficiency has been made up for by imposing overpressure to enhance the melt flowing. According to Eslamian and Ashgriz [24], the liquid discharge velocity at a pressure nozzle exit is proportional to the square root of the injection pressure difference, which means that the over-pressure imposed on the melt can play a decisive role in manipulating the melt mass flow rate in a nozzle with a very small inner diameter. In Gao et al. [25], gas atomized AlSi10Mg powders were produced by varying the melt nozzle inner diameter from 4.2 mm to 5.0 mm, and the melt was driven through the nozzle under the gravitational force, resulting in the melt mass flow rate ranging between 72.6 kg/h and 208.8 kg/h. In this paper, the melt nozzles with inner diameters 1 mm - 4 mm were employed, and with an over-pressure 30 kPa–45 kPa imposed on the melt, the melt mass flow rate varied in a range between 88.0 kg/h and 203.7 kg/h. The above comparison indicates that the reduction in the powder production rate caused by the reduction in the melt nozzle inner diameter can be counterbalanced by raising the over-pressure on the melt to enhance the melt flowing. 5.6.2. Gas consumption The over-pressure imposed on the melt in the tundish was built by filling inert gas like N2 into the furnace. Given the furnace with an inner diameter Φ = 1.2 m and a height H = 1 mm, when filling N2 into the furnace to a pressure level Δpl = 100 kPa, the gas consumption is b20% of the storage capacity of a standard cylinder (Volume: 0.040m3, max. Pressure: 15 MPa), which is insignificant compared with the gas consumption for atomization. Moreover, the gas consumption for atomization is greatly reduced by using a small inner diameter melt nozzle. For example, at atomization pressure pg = 2.0 MPa, the AlSi10Mg powder (Al-I) yield for particle size b53 μm can achieve 40% by use of a melt nozzle with an inner diameter D0 = 2 mm; however, in a conventional process employing a melt nozzle with a larger inner diameter, e.g., D0 = 3 mm - 4 mm, atomization pressure pg ≥ 3.0 MPa is usually needed to achieve the same fine powder yield. 5.6.3. Melt clogging in nozzle The over-pressure imposed on the melt in the tundish will enhance the melt flowing, and thus reduce the probability of the melt clogging in the nozzle. However, with the atomization process, the melt nozzle, and
even its heating device, will be gradually cooled by the high speed cold atomizing gas flow. In this case, the reduction in the melt nozzle inner diameter will enhance the cooling and solidification of the melt in the nozzle, increasing the risk of melt clogging. When the melt solidifies in the nozzle, the over-pressure imposed on the melt will lose its role in enhancing melt flowing. In the Drop-on-Demand (DoD) process, since there is no cooling from the cold atomizing gas flow, the CuSn melt can go through a melt nozzle with an inner diameter D0 = 0.35 mm [17], while the Cu melt can go through a melt nozzle with an inner diameter D0 = 0.30 mm [18]. In the present work, there are often a few of solidified melt remnants at the bottom of the tundish when an atomization run was over. With the decrease of the melt nozzle inner diameter, the volume of those remnants increased, especially for the Al-III alloy. This made it difficult and time-consuming to clean the tundish and thus would reduce the life span of the tundish. The melt nozzle used in the present work was changed after each atomization run. 6. Conclusions and future work The production of fine spherical metal powders with a small mass median diameter and high productivity was realized by manipulating the inner diameter of the melt nozzle during gas atomization of pressurized melts. (1) The maximum melt flow resistance in a melt nozzle, consisting of capillary resistance (ps) and viscous pressure drop (pv), was derived as a function of the inner diameter of the melt nozzle (D0) for different metal melts. The calculation results indicate that the maximum melt flow resistance, as well as the capillary resistance and the viscous pressure drop, increases quickly with the decrease of D0, especially in a range of D0 b 3 mm. When D0 reduces from 4.0 mm to 0.5 mm, the maximum melt flow resistance varies in an order of 100-102 kPa, the capillary resistance 100-101 kPa, and the viscous pressure drop 100-102 kPa. Therefore, an additional driving force is necessary for the melt to flow well through a nozzle with a small inner diameter. This has been achieved in the present work by imposing an overpressure on the melt in a tundish to enhance the melt flowing in the melt nozzle. (2) A novel pressure-gas atomization device was proposed for the production of fine metal powders, whose specific nozzle geometry in combination with an over-pressure unit and a nozzle heating unit allows for smaller nozzle inner diameters and in turn for the production of smaller powder size distributions as conventional gas atomization devices. Atomization runs with three kinds of Al-alloys were accomplished by use of melt nozzles of different inner diameters, together with a range of Δpl =
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30-45 kPa imposed on the melts. The experimental results indicate that the atomization efficiency can be well improved by reducing the inner diameter of the melt nozzle (D0), which leads to a smaller MMD, narrower particle size distribution, and higher fine powder yield. Moreover, the particle surface quality can be improved when melt nozzles with a small inner diameter are employed during gas atomization process. (3) It should be mentioned that the over-pressure imposed on the melts works only when the melt in the melt nozzle is still in a liquid or semi-liquid state. With the decrease of D0，the heat transfer between the melt and the inner wall of the melt nozzle will be enhanced, which may lead to melt solidification in the melt nozzle. In this case, the melt nozzle will be clogged despite a high over-pressure imposed on the melt. In this study, a separate heating device has been employed to warm the melt nozzle, but the high speed cold gas jets from the gas nozzle can gradually cool the melt nozzle and even the heating device during atomization. Therefore, hot gas atomization techniques as developed in [5,6] will be integrated with the method developed in this paper in future. By this way, the preheated atomizing gas can slow down the heat transfer between the melt and the melt nozzle, which is helpful in forming a thinner melt jet through a melt nozzle with a smaller inner diameter. (4) It should be declared that some key process parameters, such as melt nozzle diameter (D0), over-pressure on the melt (Δpl), and melt mass flow rate, correlate with each other. This makes it complex to discuss the results just with changing of one parameter at a time based on our understanding of the relationship between these parameters. For example, the melt mass flow rate, which determines the GMR in the present work, is influenced by both of the melt nozzle diameter and the over-pressure imposed on the melt. If the results are investigated just with changing of the melt nozzle diameter and taking into account only the nozzle geometrical effect, a fixed melt mass flow rate must be obtained by adjusting the over-pressure on the melt in each case of melt nozzle diameter. When we discuss the dependence of the particle size and its distribution on the melt properties, a fixed melt mass flow rate must also be obtained by adjusting the over-pressure on the melt when a same melt nozzle of D0 is employed. Because the exact relationship between melt nozzle diameter, over-pressure on melt and melt mass flow rate has not been determined in the present work, this makes it very expensive to perform a systematical study of the particle properties just with changing of one parameter at a time. In future, the numerical simulations will be performed and combined with experimental validation to determine the exact relationship between melt nozzle diameter, over-pressure on melt and melt mass flow rate. Acknowledgements The work was supported by Beijing Nova Program (No. xx2018036), Shenzhen Key Laboratory for Additive Manufacturing of Highperformance Materials (No. ZDSYS201703031748354), and Shenzhen Science and Technology Plan (No. JSGG20180508152608855).
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Powder Technology journal homepage: www.elsevier.com/locate/powtec
Fine spherical powder production during gas atomization of pressurized melts through melt nozzles with a small inner diameter Xing-gang Li a,b,c,⁎, Qiang Zhu b,c, Shi Shu d, Jian-zhong Fan d, Shao-ming Zhang d a
Academy for Advanced Interdisciplinary Studies, Southern University of Science and Technology (SUSTech), Shenzhen 518055, China Department of Mechanical and Energy Engineering, Southern University of Science and Technology (SUSTech), Shenzhen 518055, China Shenzhen Key Laboratory for Additive Manufacturing of High-performance Materials, Shenzhen 518055, China d National Engineering & Technology Research Center for Nonferrous Metal Matrix Composites, General Research Institute for Nonferrous Metals (GRINM), Beijing 100088, China b c
a r t i c l e
i n f o
Article history: Received 4 November 2018 Received in revised form 27 June 2019 Accepted 7 September 2019 Available online 09 September 2019 Keywords: Additive manufacture Metal powder Gas atomization Melt nozzle Mass median diameter Satellite particle
a b s t r a c t A pressure-gas atomizer was developed, in which the melts were pressurized through melt nozzles with a small inner diameter, aiming for a small mass median diameter (MMD, d50,3) and high productivity of fine spherical powders. The maximum melt flow resistance in a melt nozzle, defined as the sum of the capillary resistance and viscous pressure drop, was analyzed by varying the inner diameter of the melt nozzle (D0). The calculation results indicate that the maximum melt flow resistance increases quickly with the decrease of D0, and varies in an order of 100–102 kPa for different metal melts when D0 reduces from 4.0 mm to 0.5 mm. Atomization runs with three kinds of aluminium (Al) alloys were accomplished using melt nozzles with different inner diameters in a pilot plant whereby an over-pressure in a range of Δpl = 30–45 kPa can be maintained on the melts to enhance the melt flowing in the melt nozzle. The experimental results indicate that the atomization efficiency can be well improved by reducing the inner diameter of the melt nozzle, which resulted in a small MMD, narrow particle size distribution and high fine powder yield. For Al-I alloy powders, when the inner diameter of the melt nozzle reduces from D0 = 3 mm to D0 = 1 mm, the particle MMD reduces from d50,3 = 86.13 μm to d50,3 = 40.42 μm, and the powder yield b53 μm increases from 27.60% to 62.57%. For Al-III alloy powders, when the inner diameter of the melt nozzle reduces from D0 = 4 mm to D0 = 2 mm, the particle MMD reduces from d50,3 = 120.10 μm to d50,3 = 54.82 μm and the powder yield b53 μm increases from 20.70% to 48.20%. Moreover, the satellite particles and lamellae sticking on the particle surface were reduced when a melt nozzle with a small inner diameter was employed in a gas atomization process. © 2019 Elsevier B.V. All rights reserved.
1. Introduction The emerging technological field of Additive Manufacturing (AM) typically needs specifically tailored materials and particles, especially metal powders with precisely optimized size, shape and morphology. For example, fine spherical metal powders in a size range of b53 μm are usually preferred during selective-laser-melting (SLM) processing of metal parts. Moreover, powder agglomeration and hollow particles should be avoided to guarantee a good processability and the final product property. At present fine spherical powders for metal AM are produced primarily by gas atomization technology. Two most widely used gas atomization systems in industry are free-fall type and close-coupled type, in both of which a hot molten metal stream, usually in a form of a round jet, is disintegrated into droplets by high pressure cold gas jets ⁎ Corresponding author at: Academy for Advanced Interdisciplinary Studies, Southern University of Science and Technology (SUSTech), Shenzhen 518055, China. E-mail address: [email protected] (X. Li).
https://doi.org/10.1016/j.powtec.2019.09.023 0032-5910/© 2019 Elsevier B.V. All rights reserved.
surrounding the molten metal stream, and subsequently the atomized metallic droplets are subjected to high cooling rates and deep undercooling in a high speed atomization gas flow. In a free-fall configuration, the molten metal falls freely for some distance in the direction of gravity before being hit by the atomization gas, and this configuration generally produces relatively coarse particles. In a close-coupled configuration, gas jets hit the molten metal immediately after it comes out of the melt nozzle, and this configuration is frequently used for the production of metal powders with mass median diameter (MMD, d50,3) ranging from 10 to 100 μm. The above gas atomization techniques are often characterized with wide particle size distribution and high specific gas consumption. Larger or smaller particle size ranges are usually separated which, in turn, reduces the productivity of the process. A wide size distribution of the raw metal powder means a low productivity. For example, as for the AlSi10Mg powders for the SLM process, the powder yield in a size range b 53 μm is usually only 30 vol% in a conventional gas atomization process. Therefore, metal powder costs currently form a high fraction of additively manufactured metal parts. Research on gas atomization process has put emphasis on spherical powders with a
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small mass median diameter and high productivity. Atomization efficiency or fine powder yield can be improved mainly by three kinds of approaches: 1) optimization of gas nozzle design and the arrangement between gas nozzle and melt nozzle, 2) augmenting atomizing gas properties, and 3) manipulating the melt properties. Most of research work has been focused on the first two aspects. Significant development efforts based on gas flow dynamics have resulted in the advances in gas nozzle design and the improvement in the arrangement between gas nozzle and melt nozzle. Gas nozzles with a convergent-divergent (C-D) [1] or De Laval nozzle [2] contour in gas flow channels, as in rocket exhaust nozzle design, were widely studied to increase atomization efficiency and uniformity and to reduce the required gas supply pressures. Based on a close-coupled configuration, Anderson & Terpstra [1] developed high pressure gas atomization (HPGA) nozzles with discrete convergent-divergent jets, which exhibited higher atomization efficiency compared with those with straight bore jets. The wetting by the melt (lick back), which is a cause of gas nozzle failure, can be avoided by adjusting the angle matching or the relative placement between gas nozzle and melt nozzle to eliminate or restrict the local flow separation around the tip of the melt nozzle [3,4]. The atomization efficiency can be well improved by augmenting the properties of the atomization medium, such as gas pressure and temperature. HJE Co, Inc. [5] and PSI Ltd. [6] developed hot gas atomization techniques. The sonic velocity of a gas is directly proportional to the square root of the absolute temperature of the gas. Thus, increasing atomization gas temperature can lead to substantial kinetic energy available to disrupt the melt stream into a finer distribution of droplets, resulting in improved atomization efficiency. The efficiency of an atomization process can be alternatively optimized by increasing the specific liquid surface energy before atomization [7]. Within twin fluid (gas) atomization, the fragmentation must take place in a region where the velocity difference between the atomization gas and the liquid is highest. Therefore, the liquid must be transformed and transported into an area where the maximum velocity difference occurs. These aspects can be realized in a prefilming hybrid atomizer where the disintegration process can be subdivided into two steps: a prefilming and a gas atomization step [8]. During the prefilming step, a large specific liquid surface area is generated to prepare an efficient secondary disintegration step by gas atomization. Czisch and Fritsching [9] developed a prefilming hybrid atomizer, which was based on a combination of a single-fluid rotary prefilming concept and an external mixing twin-fluid atomizer. Within this specific design, the feed material is first spread out by a spinning disc operated in sheet formation mode, forming a free-flowing liquid film, which was subsequently conveyed into the most effective atomization zone by the local gas flow and atomized by a high speed gas flow from the external mixing gas atomizer. A different hybrid atomization process for particle production was developed to hybridize effectively gas atomization with centrifugal atomization in Minagawa et al. [10]. In this hybrid atomizer configuration the melt spray from an external mixing twin fluid nozzle impinges onto a rotating disc, resulting in a controlled film that finally acts as centrifugal atomizer for fine droplets with narrow particle size distribution within metal particle production. Another prefilming hybrid atomizer concept, designated as pressure-swirl-gas-atomization (PSGA), was developed based on a combination of pressure-swirl-atomization (PSA) as a prefilming step and gas atomization [11]. Within a PSGA process, the pressurized melt initially enters a pressure swirl nozzle and leaves the nozzle exit as a conical hollow cone film due to its swirl and centrifugal forces; as the second step, an external mixing gas nozzle is employed to disintegrate the liquid film, ligaments, and any large droplets. The PSGA atomizer design, especially the pressure swirl nozzle, is sophisticated and has been used to produce fine spherical powders of low melting point metals with small median diameter and narrow size distribution
[12,13], as well as particulate reinforced metal-matrix-composite (MMC) particles [14]. Atomization of pressurized non-metallic liquids has been widely investigated, for example, in the field of fuel combustion in a engine [15,16], where the liquid fuel is injected into a combustion chamber under a high injection pressure and then atomized into very tiny droplets, thus enhancing the fuel combustion efficiency. As for the atomization of pressurized metal melts, besides the pressure-swirl-gasatomization (PSGA) technique discussed above, a periodic overpressure is often imposed on the melt in a Drop-on-Demand (DOD) process [17,18], whereby the melt is periodically pushed through a melt nozzle, forming a discontinuous chain of mono-sized droplets, which subsequently solidify as powder particles. Based on the authors' knowledge, Raman and Surette [19] firstly proposed the method, i.e., gas atomization of pressurized metal melts in a form of round jet for powder production in their invention patent, and presented preliminary experimental results, which indicated that good yields of fine particles were obtained by imposing over-pressure on the melt (Sn-5%Pb) when it went through a melt nozzle with an inner diameter of D0 = 2 mm. However, since then, few studies have been reported on this method. Although, in the practice of metal powder production by gas atomization, it may be very common to impose a slight over-pressure on the melt to enhance the melt flow, few studies have been performed to systematically investigate the influence of over-pressure and nozzle inner diameter on the melt flow behavior in the melt nozzle. In this paper, the pressure swirl nozzle in a PSGA process is replaced by a simple melt nozzle with a cylindrical channel, through which the pressurized melt goes into the spray chamber as a round liquid jet. Attempts to reduce the particle size of metal powders concentrate on reducing the inner diameter of the melt nozzle (D0). With this approach, a very thin liquid jet is expected to form at the melt nozzle orifice, which can enhance the subsequent gas atomization efficiency, thus resulting in an improved yield of fine particles. Firstly, the melt flow resistance in a melt nozzle is analyzed by varying the inner diameter of the melt nozzle. Next, an introduction is given to the powder plant and the powder production process, in which an over-pressure can be maintained on the melts as a driving force to enhance the melt flowing in the melt nozzle. Finally, three kinds of Al-alloys are gas-atomized using melt nozzles with different inner diameters, and the asatomized powders are characterized with emphasis on the properties such as mass median diameter, particle size distribution span, fine powder yield and powder morphology. 2. Theory/calculation In a conventional gas atomization process, the driving forces for the melts to go through a melt nozzle include the aspiration pressure formed in front of the melt nozzle by high speed gas jets and the gravitational force. In this case, a melt nozzle with a large diameter is often used to ensure that the melt can go through the melt nozzle by overcoming the capillary force and the viscous force between the melts and the inner wall of the melt nozzle, thus avoiding clogging in the melt nozzle but resulting in coarse powder particles. Lubanska [20] has established the correlation with mass median diameter (MMD, d50,3) of gas-atomized powders that demonstrates the benefits of reducing the inner diameter of the melt nozzle (D0) for the production of fine powders: d50;3 ¼ D0 K lub
0:5 νl 1 1 1þ GMR ν g We
ð1Þ
where
We ¼
ðΔU Þ2 ρl D0 σ
ð2Þ
X. Li et al. / Powder Technology 356 (2019) 759–768
is the Weber number, which represents the ratio of inertial force including the density of the melt (ρl) to the surface tension of the melt (σ), Klub is a constant that must be determined for a particular atomizer, νl and νg are the kinematic viscosities of the melt and the gas, respectively, GMR is the ratio of gas to melt mass flow rate, and ΔU is the relative velocity between the melt and the gas. Take Eq. (2) into Eq. (1), which can be rearranged into. 0:5 1 d50;3 ¼ K lub Q D0 1 þ GMR
ð3Þ
where Q is related to the properties of atomized melts and atomizing gas, as. Q¼
νl σ ν g ðΔU Þ2 ρl
!0:5 ð4Þ
For specific atomizer and material systems, Eq. (3) indicates that the variation of the particle mass median diameter (MMD, d50,3) is only related to the melt nozzle diameter (D0) and the GMR. The MMD is directly proportional to the square root of D0, which provides a theoretical basis to reduce the particle size of metal powders using a melt nozzle with a small inner diameter in a gas atomization process. However, with the decrease of D0, the melt flow resistance in the melt nozzle may increase, which cause a reduction in melt mass flow rate and even melt clogging in the channel. The next will give a quantitative description of the melt flow resistance related to the inner diameter of the melt nozzle. 2.1. Capillary resistance Fig. 1 (a) and (b) depict two cases of the clogged inner-channels of the melt nozzle by melt columns, respectively. The shape of the innerchannel is cylindrical. Assuming that the melt is in a full liquid state in the clogged melt nozzle, a balance is reached between the capillary force (Fs), viscous force (Fv) and gravity. (See Table 1.) The capillary pressure (ps) can be obtained by the Young-Laplace equation [23] as: ps ¼
4σ cosθ D0
ð5Þ
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Table 1 Surface tension and viscosity at melting point [21,22]. Metal
Sn
Zn
Mg
Al
Cu
Ni
Fe
Melting point (°C) Surface tension (N/m) Viscosity (mPa•s)
232 0.550 1.85
420 0.789 3.85
650 0.577 1.25
660 0.871 1.30
1083 1.330 4.0
1455 1.796 4.90
1530 1.862 5.5
the melt flowing in the melt nozzle, as shown in Fig. 1 (a); if the contact angle θ N 90°, the value of the capillary pressure (ps) is negative, which means that the capillary force (Fs) will prevent the melt flowing in the melt nozzle, as shown in Fig. 1 (b). For gas atomization of metal melts, the melt nozzles are usually made from refractory materials, such as graphite, alumina, magnesia, zirconia and boron nitride. In most cases, the surfaces made from these materials are poorly wetted by metal melts, and the contact angles (θ) between metal melts and solid surfaces are larger than 90°. This means that the capillary force (Fs) will act in an opposite direction against the melt flow in a melt nozzle. According to Eq. (5), the capillary resistance (ps) is calculated for the melts of several common metals. The calculated results are plotted against the inner diameter of the melt nozzle (D0) in Fig. 2 and against the liquid/solid contact angle (θ) in Fig. 3, respectively. Fig. 2 indicates that the capillary resistance increases quickly with the decrease of the inner diameter of the melt nozzle (D0). In the case of a constant liquid/solid contact angle θ = 150°, the capillary resistance (ps) varies in an order of 100-101 kPa when the inner diameter of the melt nozzle (D0) reduces from 4.0 mm to 0.5 mm. Fig. 3 indicates that if the solid surface is non-wetted by the melts, the capillary resistance increases with the increase of the liquid/ solid contact angle (θ). 2.2. Viscous pressure drop Assuming that the melts are incompressible and Newtonian fluids in a laminar flow flowing through a melt nozzle, the viscous pressure drop (pv) can be described using the Hagen-Poiseuille law [23] as: pv ¼
8μLQ πR40
ð6Þ
where μ is the dynamic viscosity, L is the inner-channel length of the melt nozzle, R0 is the inner radius of the melt nozzle, and Q is the melt
which relates the capillary pressure (ps) at the melt/gas interface to the inner diameter of the melt nozzle (D0), the surface tension of the melt (σ), and the liquid/solid contact angle (θ). Eq. (5) indicates that the capillary pressure (ps) is directly proportional to the surface tension of the melt (σ), and inversely related to the inner diameter of the melt nozzle (D0). If the contact angle θ b 90°, the value of the capillary pressure (ps) is positive, which means that the capillary force (Fs) is a driven force for
Melt
Melt
D0
Fs
θ Fs
D0
Fs
θ Fs
Fig. 1. Illustration of melt clogging inner-channel in melt nozzle, liquid/solid contact angle (a) θ b 90°, (b) θ N 90°.
Fig. 2. Capillary resistance vs. melt nozzle inner diameter (D0) in the case of a constant liquid/solid contact angle (θ) for different metals.
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viscous pressure drop varies in an order of 100-102 kPa when the inner diameter of the melt nozzle (D0) reduces from 4.0 mm to 0.5 mm. 2.3. Melt flow resistance
Fig. 3. Capillary resistance vs. liquid/solid contact angle (θ) in the case of a constant melt nozzle inner diameter (D0) for different metals.
volumetric flow rate, which is defined as: Q ¼ Ul • πR20
ð7Þ
where Ul is the melt normal discharge velocity at the nozzle exit. Combine Eqs. (6) and (7) and substitute R0 with D0/2 to obtain:
The capillary resistance is usually important in the start phase of the process when the melt nozzle is not completely wetted. During a spray run the melt nozzle should be wetted completely. However, in many cases the melt clogging happens at the very beginning of the process before the melt nozzle is not completely wetted, especially when a melt nozzle with a small inner diameter is employed. In some cases, when the melt clogging happens during a spray run, the solidified melt column in a melt nozzle shows the solidification front with a curved surface, which implies that the capillary force plays a role. Therefore, the maximum resistance that impedes the melt flow in a melt nozzle could be defined as the sum of the capillary resistance (ps) and the viscous pressure drop (pv). As shown in Fig. 5, the maximum melt flow resistance is plotted against the inner diameter of the melt nozzle (D0). The melt nozzle with a smaller inner diameter has larger flow resistance. For a constant liquid/solid contact angle of θ = 150° and a normal discharge velocity of Ul = 5 m/s, the maximum resistance that impedes the melt flow in a melt nozzle, with an inner-channel length of L = 40 mm, varies in an order of 100-102 kPa when the inner diameter of the melt nozzle reduces from 4.0 mm to 0.5 mm. Therefore, an additional driving force is necessary for the melt to flow well through a nozzle with a small inner diameter. In this paper, an over-pressure is imposed on the melt in a tundish to enhance the melt flowing in a melt nozzle by filling gas into the furnace and then building a pressure difference between the furnace and the spray chamber. 3. Powder plant
pv ¼
32μLU l D20
ð8Þ
which indicates that the viscous pressure drop is directly proportional to the dynamic viscosity, the nozzle inner-channel length and the melt normal discharge velocity, and inversely proportional to the square of the inner diameter of the melt nozzle. Fig. 4 shows the calculated results of the viscous pressure drop (pv) plotted against the inner diameter of the melt nozzle (D0), which indicates that the viscous pressure drop increases quickly with the decrease of the inner diameter of the melt nozzle (D0). For a normal discharge velocity of Ul = 5 m/s and an inner-channel length of L = 40 mm, the
Fig. 4. Viscous drop pressure (pv) vs. melt nozzle inner diameter (D0) for different metals, L - inner-channel length of melt nozzle, Ul - melt discharge velocity.
The powder production facility used in this study mainly consists of melting furnace, spray chamber, vacuum system and powder collection system. The melting furnace is a pressure vessel which can withstand an over-pressure (Δpl) of at least 100 kPa in the present experiments. The furnace cover and body are locked and sealed through a design as in a pressure cooker. The furnace and the spray chamber are two separate containers and connect with each other only through a melt nozzle at the bottom of the tundish. Both of the furnace and the spray chamber are firstly evacuated to a vacuum level of 10−2-10−1 Pa, and then refilled with nitrogen gas (N2) to a pressure of 1 atm prior to melting
Fig. 5. Melt flow resistance vs. melt nozzle inner diameter (D0) for different metals, L inner-channel length of melt nozzle, Ul - melt normal discharge velocity, θ - liquid/solid contact angle.
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to reduce the oxygen content. When the preset superheat of the melt in the crucible is achieved, refill the nitrogen gas into the furnace to the preset over-pressure, and then transfer the melt from the crucible to the tundish. At last the pressurized melt is pushed through the melt nozzle into the spray chamber, where the melt jet is disintegrated into droplets by the high speed gas jets. The gas nozzle consists of 20 straight bore jets with a diameter of 1.1 mm at each outlet. The diameter of the circumference, along which the 20 gas-nozzle outlets are uniformly distributed, is DG = 28 mm, and an atomization angle of ɑ = 45° was designed for the gas nozzle (see Fig. 6 (b)). The powder collection system consists of cyclone separator and powder collector. Most of the coarse particles are deposited into the powder collector at the bottom of the spray chamber, while the rest of the particles are carried into the cyclone separator by the gas flow and further deposited into the powder collector at the bottom of the cyclone separator. The exhaust from the cyclone separator is rigorously filtered and at last emitted into the atmosphere. The powders in different powder collectors are mixed in a powder mixer. 4. Experimental design and procedure Three kinds of Al-alloys, designated as Al-I, Al-II and Al-III, respectively, were atomized with a discrete jet gas nozzle using commercial purity nitrogen at 2.0 MPa. The nominal composition of the alloys is listed in Table 2. In contrast with the Al-II alloy, the Al-III alloy contains extra 18-20 wt% Si and 5-7 wt% Fe, which implies that the Al-III melt would be more viscous than the Al-II melt. As indicated in Fig. 6 (a), the crucible and the tundish are made from commercial purity graphite, while the melt nozzle is made from high purity graphite. Each 7 kg alloy charge was induction superheated by 150 K over the melting point in the crucible. The melt nozzles were tested by varying the inner diameter from 4 mm to 1 mm. It should be mentioned that a higher over-pressure would result in a higher melt mass flow rate, thereby reducing the fine powder yield. Therefore, the over-pressure imposed on the top of the melt in this study lies in a range of 3045 kPa, which is on the same level as the calculated results of the melt flow resistance in melt nozzles (see Fig. 5). In case of melt freezing in the melt nozzle, the tundish is also superheated by 150 K over the melting point temperature for each alloy by resistance heating; moreover, a separate heating device was employed to heat the melt nozzle to about 500 °C during atomization. The vertical distance between the melt nozzle exit and the gas nozzle outlet is H = 5-10 mm (see Fig. 6 (b)). Table 3 summarizes the experimental parameters of the gas atomization experiments of three Al-alloys. The melt atomization duration mainly depends on the inner diameter of the melt nozzle and the over-pressure imposed on the melt in the furnace, and ranges between 2 min - 6 min. The melt mass flow rate (Ml), which is defined as the ratio of the powder weight finally obtained
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Table 2 Nominal composition (wt%) of Al-alloys tested in experiments.
Al-I (AlSi10Mg) Al-II (2009 alloy) Al-III
Si
Mg
Cu
Fe
Al
9.0–11.0 – 18.0–20.0
0.2–0.5 1.0–1.6 1.0–1.6
– 3.2–4.4 3.2–4.4
– – 5.0–7.0
Bal. Bal. Bal.
in powder collectors to the melt atomization duration, generally varies from 80 kg/h to 210 kg/h. The gas flow measurements were performed using the method developed in Anderson & Terpstra [1], in which both the initial gas cylinder pressure and the final gas cylinder pressure after a gas flow period were recorded at an ambient-temperature equilibration state of the supply cylinder. In this study, sixteen gas cylinders were assembled in series with each other to provide gas for melt atomization. For all the tests here, the atomization gas pressure was set at 2.0 MPa, which resulted in a gas mass flow rate of about 200 kg/h. Each run of atomized Al-alloy powders were pre-screened to b500 μm, eliminating irregular particulates and atomization debris. For characterization samples were taken from the powders of each run and analyzed by laser diffraction and scanning electron microscopy (SEM). 5. Experimental results and discussion 5.1. Particle size distribution The Al-alloy powders were produced by gas atomization of pressurized melt jets through the melt nozzles of different inner diameters. Fig. 7 shows the particle size distributions (cumulative volume or mass) of the powders of Al\\I and Al-III alloys. The powder samples were wetdispersed and the measurements were carried out with a laser light diffraction instrument (Beckman Coulter Particle Size Analyzer LS 13320). As the inner diameter of the melt nozzle decreases, the particle size distribution curve shrinks and moves towards the region of small particle size. The results indicate that the particle size distribution is greatly influenced by the inner diameter of the melt nozzle, especially when it reduces to b3 mm. 5.2. Mass median diameter The values of particle mass median diameter (MMD, d50,3) were extracted from the particle size distributions and plotted against the inner diameter of the melt nozzle (D0), as shown in Fig. 8. For the three runs of the atomized Al-I alloy powders, the particle MMD is 86.13 μm for D0 = 3 mm, 60.45 μm for D0 = 2 mm and 40.42 μm for D0 = 1 mm, respectively, indicating a decrease of ~53% in MMD when the inner diameter of the melt nozzle reduces from D0 = 3 mm to D0 = 1 mm. For the three runs of the atomized Al-III alloy powders, the MMD is 120.10 μm
Fig. 6. (a) Illustration of melt transferring, pressurizing, guiding and gas atomization process; (b) Sketch of nozzle design under operation.
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Table 3 Experimental parameters in gas atomization of different Al-alloys. Alloy
D0 (mm)
Δpl (kPa)
Ml (kg/h)
pg (MPa)
Mg (kg/h)
GMR (−)
ΔTl (K)
Al-I Al-I Al-I Al-II Al-III Al-III Al-III
3 2 1 3 4 3 2
30 35 40 30 35 40 45
139.2 102.27 87.97 120.89 203.69 184.05 104.60
2.0 2.0 2.0 2.0 2.0 2.0 2.0
200 200 200 200 200 200 200
1.44 1.96 2.27 1.65 0.99 1.09 1.91
150 150 150 150 150 150 150
Note: D0 - melt nozzle inner diameter, Δpl - over-pressure on the top of the melt, Ml - melt mass flow rate, pg - atomization gas pressure, Mg - gas mass flow rate, GMR - gas-to-metal mass flow rate ratio, ΔTl - melt superheat.
for D0 = 4 mm, 102.20 μm for D0 = 3 mm and 54.82 μm for D0 = 2 mm, respectively, indicating a decrease of ~54% in MMD when the inner diameter of the melt nozzle reduces from D0 = 4 mm to D0 = 2 mm. As described in Eq.(3), the decrease of the particle MMD in this study could be mainly attributed to the nozzle geometrical effect and the GMR increase, both of which are induced by the reduction in D0. The Al-III alloy was atomized under similar GMRs by employing the melt nozzle of D0 = 4 mm and D0 = 3 mm, respectively (see Table 3). The former resulted in a ~18% larger particle MMD than the latter, which indicates the significant nozzle geometrical effect on the variation of the particle MMD. In the case of D0 = 3 mm, the atomization of the Al-III alloy resulted in a ~16% larger particle MMD than the Al-II alloy, possibly due to the higher viscosity of the Al-III alloy containing excessive Si and Fe. The results of the particle MMD have also been collected and plotted as a function of the GMR in Fig. 9, which indicates that the particle MMD decreases with the increase of the GMR. In the present work, the reduction in the melt nozzle diameter led to the decrease of melt mass flow rate, while the over-pressure imposed on the melt was selected as low as possible to maintain a low melt mass flow rate, thus raising the GMR. Fig. 10 shows the results of the particle MMD as a function of the factor [D0(1 + 1/GMR)]^0.5 in Eq. (3), which represents the comprehensive effect of the simultaneous changes of melt nozzle diameter and the GMR on the variation of the particle MMD. Given constant values of KlubQ, the variation of the particle MMD with the factor [D0(1 + 1/ GMR)]^0.5 according to Eq. (3) is shown in Fig. 10, which indicates that the values of the particle MMD obtained from the present work are located between the two curves of KlubQ = 35 and KlubQ = 45
Fig. 8. Mass median diameter (MMD, d50,3) vs. melt nozzle inner diameter (D0), measurements conducted by laser light diffraction (wet dispersion).
based on Eq. (3). Here the variation in KlubQ value reflects the change in material properties. 5.3. Particle size distribution span The size distribution span of the powder particles is defined as the diameter ratio. span ¼
d90;3 −d10;3 d50;3
ð9Þ
A narrower distribution means a higher yield at a specific mass median diameter. The span values were derived from the particle size distribution and plotted against the inner diameter of the melt nozzle (D0), as shown in Fig. 11. The results indicate that the particle size distribution span generally reduces with the decrease of D0, especially when D0 b 3 mm. For example, the size distribution span of the atomized Al\\I alloy powder decreases from span = 2.16 to span = 1.52 when the inner diameter of the melt nozzle varies from D0 = 3 mm to
Fig. 7. Cumulative mass distribution for different melt nozzle inner diameters (D0), measurements conducted by laser light diffraction (wet dispersion).
X. Li et al. / Powder Technology 356 (2019) 759–768
Fig. 9. Mass median diameter (MMD, d50,3) vs. the ratio of gas to melt mass flow rate (GMR), measurements conducted by laser light diffraction (wet dispersion).
D0 = 1 mm; the size distribution span of the atomized Al-III alloy powder decreases from span = 2.75 to span = 2.05 when the inner diameter of the melt nozzle varies from D0 = 4 mm to D0 = 2 mm. In the case of D0 = 3 mm, the span values are 2.16, 1.73 and 2.74, respectively, for Al-I, Al-II and Al-III alloys.
5.4. Fine powder yield The powder yield b53 μm is defined here as the cumulative mass or volume fraction at the size of 53 μm in the particle size distribution curve, which is measured by laser light diffraction. Fig. 12 is a plot of the powder yield b53 μm as a function of the inner diameter of the melt nozzle (D0). An obvious increase in the fine powder yield could be observed with the decrease of D0. For the atomized Al-I alloy powders, the powder yields b53 μm are 27.60% for D0 = 3 mm, 41.50% for D0 = 2 mm and 62.57% for D0 = 1 mm, respectively; for the atomized Al-III alloy powders, the powder yields b53 μm are 20.70% for D0 = 4 mm, 25.90% for D0 = 3 mm and 48.2% for D0 = 2 mm, respectively. For the Al-II alloy, a 26.40% powder yield b53 μm is obtained by use of a melt nozzle with D0 = 3 mm, which is similar to the results of the Al-I and Al-III alloys obtained by use of the same melt nozzle.
Fig. 10. Mass median diameter (MMD, d50,3) vs. the factor [D0(1 + 1/GMR)]^0.5 in Eq. (3), measurements conducted by laser light diffraction (wet dispersion)，curves of KlubQ = 45 and KlubQ = 35 calculated according to Eq. (3).
765
Fig. 11. Powder size distribution span vs. melt nozzle inner diameter (D0), measurements conducted by laser light diffraction (wet dispersion).
5.5. Powder morphology SEM images of the Al-I alloy powder samples in a size range of 38 and 53 μm are presented in Fig. 13. It can be observed that the powders generated with a melt nozzle of D0 = 3 mm exhibit a lot of satellites (small particles) sticking on the surface of larger particles compared with those generated with a melt nozzle of D0 = 1 mm. Moreover, the micrographs of the Al-I alloy powder samples in a size range of b38 μm, as shown in Fig. 14, indicate that the particles generated with a melt nozzle of D0 = 1 mm are more spherical and smoother. SEM images of the Al-III alloy powder samples, generated with melt nozzles of D0 = 4 mm and D0 = 2 mm, respectively, are presented in Fig. 15. A few of satellites sticking on the particle surface are observed in both of the powder samples. But the particle surface of the latter seems more spherical and smoother. SEM images of the Al-II and Al-III alloy powder samples, generated with a melt nozzle of D0 = 3 mm, are presented for comparison in Fig. 16, which indicates satellites and lamellae sticking on the particle surface in both of the powder samples. The gas-atomized powders in this study are spherical in general. However, the phenomena of satellites and lamellae sticking on the particle surface, especially in the powders generated using a melt nozzle
Fig. 12. Powder yield b53 μm vs. melt nozzle inner diameter (D0), measurements conducted by laser light diffraction (wet dispersion).
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Fig. 13. SEM images of Al-I alloy powders with a size range 38 μm b d b 53 μm: (a) melt nozzle inner diameter D0 = 3 mm, (b) melt nozzle inner diameter D0 = 1 mm.
with a larger inner diameter, are often observed, which reduce the powder sphericity and smoothness, and further reduce the powder flowability and apparent density. The formation of those satellites and lamellae on the particle surface should be ascribed to the frequent particle/droplet collisions in the atomization spray process, where the gas/droplet momentum exchange and heat transfer are mainly influenced by the droplet size. Small droplets, which would firstly solidify and are more easily accelerated by the high-speed atomizing gas, collide with large droplets that are not fully solidified. The collision may cause adhesion, partial penetration or fusion between droplets, leading to the formation of satellite particles sticking on the surface of the large
particles; or cause droplet deformation, leading to the formation of lamellae covering the particle surface. In this study, the powders, generated using a melt nozzle with a smaller inner diameter, usually exhibit a narrower or more uniform size distribution, which implies a lower particle/droplet collision frequency in the atomization spray process because of similar particle/droplet inertia; in addition, the reduction in the inner diameter of the melt nozzle leads to a higher GMR (see Table 3), which enhances the droplet solidification. Therefore, the powders generated using a melt nozzle with a smaller inner diameter are more spherical and smoother with fewer satellites and lamellae sticking on the particle surface.
Fig. 14. SEM images of Al-I alloy powders with a size range d b 38 μm: (a) melt nozzle inner diameter D0 = 3 mm, (b) melt nozzle inner diameter D0 = 1 mm.
Fig. 15. SEM images of Al-III alloy powders with a size range d b 53 μm: (a) melt nozzle inner diameter D0 = 4 mm, (b) melt nozzle inner diameter D0 = 2 mm.
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Fig. 16. SEM images of powders with a size range d b 53 μm, melt nozzle inner diameter D0 = 3 mm: (a) Al-II alloy powders, (b) Al-III alloy powders.
5.6. Concerned issues 5.6.1. Powder production rate For powder production by gas atomization, the melt mass flow rate (kg/h) is an important process parameter, which determines the powder throughput / production rate. The reduction in the melt nozzle inner diameter will lead to the increase of the melt flowing resistance in the nozzle, thus lowering the melt mass flow rate. However, in this paper, this deficiency has been made up for by imposing overpressure to enhance the melt flowing. According to Eslamian and Ashgriz [24], the liquid discharge velocity at a pressure nozzle exit is proportional to the square root of the injection pressure difference, which means that the over-pressure imposed on the melt can play a decisive role in manipulating the melt mass flow rate in a nozzle with a very small inner diameter. In Gao et al. [25], gas atomized AlSi10Mg powders were produced by varying the melt nozzle inner diameter from 4.2 mm to 5.0 mm, and the melt was driven through the nozzle under the gravitational force, resulting in the melt mass flow rate ranging between 72.6 kg/h and 208.8 kg/h. In this paper, the melt nozzles with inner diameters 1 mm - 4 mm were employed, and with an over-pressure 30 kPa–45 kPa imposed on the melt, the melt mass flow rate varied in a range between 88.0 kg/h and 203.7 kg/h. The above comparison indicates that the reduction in the powder production rate caused by the reduction in the melt nozzle inner diameter can be counterbalanced by raising the over-pressure on the melt to enhance the melt flowing. 5.6.2. Gas consumption The over-pressure imposed on the melt in the tundish was built by filling inert gas like N2 into the furnace. Given the furnace with an inner diameter Φ = 1.2 m and a height H = 1 mm, when filling N2 into the furnace to a pressure level Δpl = 100 kPa, the gas consumption is b20% of the storage capacity of a standard cylinder (Volume: 0.040m3, max. Pressure: 15 MPa), which is insignificant compared with the gas consumption for atomization. Moreover, the gas consumption for atomization is greatly reduced by using a small inner diameter melt nozzle. For example, at atomization pressure pg = 2.0 MPa, the AlSi10Mg powder (Al-I) yield for particle size b53 μm can achieve 40% by use of a melt nozzle with an inner diameter D0 = 2 mm; however, in a conventional process employing a melt nozzle with a larger inner diameter, e.g., D0 = 3 mm - 4 mm, atomization pressure pg ≥ 3.0 MPa is usually needed to achieve the same fine powder yield. 5.6.3. Melt clogging in nozzle The over-pressure imposed on the melt in the tundish will enhance the melt flowing, and thus reduce the probability of the melt clogging in the nozzle. However, with the atomization process, the melt nozzle, and
even its heating device, will be gradually cooled by the high speed cold atomizing gas flow. In this case, the reduction in the melt nozzle inner diameter will enhance the cooling and solidification of the melt in the nozzle, increasing the risk of melt clogging. When the melt solidifies in the nozzle, the over-pressure imposed on the melt will lose its role in enhancing melt flowing. In the Drop-on-Demand (DoD) process, since there is no cooling from the cold atomizing gas flow, the CuSn melt can go through a melt nozzle with an inner diameter D0 = 0.35 mm [17], while the Cu melt can go through a melt nozzle with an inner diameter D0 = 0.30 mm [18]. In the present work, there are often a few of solidified melt remnants at the bottom of the tundish when an atomization run was over. With the decrease of the melt nozzle inner diameter, the volume of those remnants increased, especially for the Al-III alloy. This made it difficult and time-consuming to clean the tundish and thus would reduce the life span of the tundish. The melt nozzle used in the present work was changed after each atomization run. 6. Conclusions and future work The production of fine spherical metal powders with a small mass median diameter and high productivity was realized by manipulating the inner diameter of the melt nozzle during gas atomization of pressurized melts. (1) The maximum melt flow resistance in a melt nozzle, consisting of capillary resistance (ps) and viscous pressure drop (pv), was derived as a function of the inner diameter of the melt nozzle (D0) for different metal melts. The calculation results indicate that the maximum melt flow resistance, as well as the capillary resistance and the viscous pressure drop, increases quickly with the decrease of D0, especially in a range of D0 b 3 mm. When D0 reduces from 4.0 mm to 0.5 mm, the maximum melt flow resistance varies in an order of 100-102 kPa, the capillary resistance 100-101 kPa, and the viscous pressure drop 100-102 kPa. Therefore, an additional driving force is necessary for the melt to flow well through a nozzle with a small inner diameter. This has been achieved in the present work by imposing an overpressure on the melt in a tundish to enhance the melt flowing in the melt nozzle. (2) A novel pressure-gas atomization device was proposed for the production of fine metal powders, whose specific nozzle geometry in combination with an over-pressure unit and a nozzle heating unit allows for smaller nozzle inner diameters and in turn for the production of smaller powder size distributions as conventional gas atomization devices. Atomization runs with three kinds of Al-alloys were accomplished by use of melt nozzles of different inner diameters, together with a range of Δpl =
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30-45 kPa imposed on the melts. The experimental results indicate that the atomization efficiency can be well improved by reducing the inner diameter of the melt nozzle (D0), which leads to a smaller MMD, narrower particle size distribution, and higher fine powder yield. Moreover, the particle surface quality can be improved when melt nozzles with a small inner diameter are employed during gas atomization process. (3) It should be mentioned that the over-pressure imposed on the melts works only when the melt in the melt nozzle is still in a liquid or semi-liquid state. With the decrease of D0，the heat transfer between the melt and the inner wall of the melt nozzle will be enhanced, which may lead to melt solidification in the melt nozzle. In this case, the melt nozzle will be clogged despite a high over-pressure imposed on the melt. In this study, a separate heating device has been employed to warm the melt nozzle, but the high speed cold gas jets from the gas nozzle can gradually cool the melt nozzle and even the heating device during atomization. Therefore, hot gas atomization techniques as developed in [5,6] will be integrated with the method developed in this paper in future. By this way, the preheated atomizing gas can slow down the heat transfer between the melt and the melt nozzle, which is helpful in forming a thinner melt jet through a melt nozzle with a smaller inner diameter. (4) It should be declared that some key process parameters, such as melt nozzle diameter (D0), over-pressure on the melt (Δpl), and melt mass flow rate, correlate with each other. This makes it complex to discuss the results just with changing of one parameter at a time based on our understanding of the relationship between these parameters. For example, the melt mass flow rate, which determines the GMR in the present work, is influenced by both of the melt nozzle diameter and the over-pressure imposed on the melt. If the results are investigated just with changing of the melt nozzle diameter and taking into account only the nozzle geometrical effect, a fixed melt mass flow rate must be obtained by adjusting the over-pressure on the melt in each case of melt nozzle diameter. When we discuss the dependence of the particle size and its distribution on the melt properties, a fixed melt mass flow rate must also be obtained by adjusting the over-pressure on the melt when a same melt nozzle of D0 is employed. Because the exact relationship between melt nozzle diameter, over-pressure on melt and melt mass flow rate has not been determined in the present work, this makes it very expensive to perform a systematical study of the particle properties just with changing of one parameter at a time. In future, the numerical simulations will be performed and combined with experimental validation to determine the exact relationship between melt nozzle diameter, over-pressure on melt and melt mass flow rate. Acknowledgements The work was supported by Beijing Nova Program (No. xx2018036), Shenzhen Key Laboratory for Additive Manufacturing of Highperformance Materials (No. ZDSYS201703031748354), and Shenzhen Science and Technology Plan (No. JSGG20180508152608855).
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