A Numerical Investigation of Gas Flow Effects on High-Pressure Gas Atomization Due to Melt Tip...

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METALLURGICAL AND MATERIALS TRANSACTIONS B VOLUME 28B, OCTOBER 1997935

A Numerical Investigation of Gas Flow Effects on High-Pressure Gas Atomization Due to Melt Tip GeometryVariation

J. MI, R.S. FIGLIOLA, and I.E. ANDERSON

A parametric numerical study is presented on how melt feed tube geometry influences the gas flowfield of a high-pressure gas atomizer (HPGA). The axisymmetric, turbulent, compressible NavierStokes equations are solved for the gas-only flow in the vicinity of the melt tip for a confined-feed,annular-slit atomizer. The numerical results indicate that a fully retracted melt tip can develop a highoverambient base pressure over a wide range of operating pressures, while an extended melt tip candevelop a subambient pressure, called an aspirating effect, which encourages melt to flow. Addinga taper angle to an extended melt tip decreases the aspirating effect. However, a melt tip can developa radial pressure gradient along its base that forces the melt to move radially outward into a high-shear region of the flow, thus encouraging droplet formation. A fully retracted melt tip develops thehighest such radial force, but this tends to decrease in magnitude with tip extension. Hence, acompromise must be established whereby the good aspiration character of a longer tip is balancedagainst the good pressure-driving force along the base of a shorter tip. The results provide detailsof the flow field and the effects due to variation of melt tip taper and extension.

I. INTRODUCTION

High-pressure gas atomization (HPGA) can be an effi-cient method for producing high yields of ultrafine metaland alloy powders.[1,2] Currently, many commercial metaland alloy powders are being produced using HPGA, in-cluding high-alloy tool steels, nickel-based superalloys, andaluminum alloys.[39]

Earlier studies[913] have shown evidence that both thesystems gas operating pressure and the melt feed tube tipgeometry effectively control the gas dynamics within theatomization zone of the HPGA process. The design of theHPGA melt tip has been proven to be an essential featureof the process efficiency. Some melt tip geometric param-eters can be altered, such as the melt tip extension length,

s, and melt tip taper angle, , shown in Figure 1. It isexpected that the gas flow characteristics and melt aspira-tion will change with different melt tip designs.

Operation of an HPGA system within the aspiration re-gime results in a stable atomization condition and has beenassociated with a fine powder production.[39] Under this fa-vorable condition, the subambient melt tip base pressure per-mits the liquid metal to accelerate into contact with the gasstream. As was illustrated in the previous publications,[1014]

the liquid metal exiting the pour tube is seen to flow fromthe melt tip base center area to the tip edge along the melt

tip base. Therefore, the pressure gradient from the melt tipbase center to the melt tip base edge plays an essential rolein driving this radial movement.

J. MI, formerly Research Assistant, Department of MechanicalEngineering, Clemson University, is Research Associate, Metals andCeramics Division, Ames Laboratory, Ames, IA 50011. R.S. FIGLIOLA,Professor and Chair, is with the Department of Mechanical Engineering,Clemson University, Clemson, SC 29634-0921. I.E. ANDERSON, SeniorMetallurgist, is with the Metals and Ceramics Division, Ames Laboratory,Ames, IA 50011.

Manuscript submitted April 30, 1996.

Anderson and Rath[3] reported that efforts to use a fullyretracted melt tip design at high operating pressures resultedin unstable atomization with subsequent melt flow freeze-up. Extending the melt tip into the flow was found to sta-

bilize the melt flow, and adding taper to the tip improvedpowder refinement. Meanwhile, the extended tip was foundto develop a subambient pressure over its base for a widerange of operating pressures.[11] In an experimental study

by Anderson and Figliola,[10] three different melt tip taperangles63, 45, and 0 degwere used to measure pressureat a single point on the melt tip base. Schlieren imagingdocumented very different flow fields. For a wide range of

operating pressures, the 63 deg tip failed to providean aspirating effect, whereas the 45 deg tip and theextended straight 0 deg tip did. These results suggestedthat taper angle can control the aspiration capability of amelt feed tip. Indeed, commercial operators often point outthat a straight tip extension will stabilize the atomizer meltflow, preventing freeze-up caused by gas flowing back upthe tube. However, a straight extended tip design is alsoknown to be rather inefficient for producing very fine pow-ders, especially at high operating pressures.[10] The straighttip is also prone to erosion owing to the high impact energydissipated by the gas jet at the tube wall.[7,8,10]

The effects of gas atomizing pressure on the HPGA gasflow field were examined numerically by Mi et al.[14] The

study effectively documented for one melt tip extension andtaper geometry that both a subambient pressure and a radial

pressure gradient were developed along the melt tip base.These act to force the melt to flow radially outward fromthe tip orifice, thereby directing the melt to make contactwith the high velocity gas found along the edge of the at-omization zone. This mechanistic explanation is consistentwith observations.[6,10]

This article reports on a numerical investigation of theinfluence of the melt tip extension length and melt tip taperangle on the gas flow field during HPGA operation. Such


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936VOLUME 28B, OCTOBER 1997 METALLURGICAL AND MATERIALS TRANSACTIONS B

Fig. 1The cross-sectional view of the HPGA melt feed tube tipgeometry.

Table I. Different HPGA Melt Tip Geometry Designs

Tip

1 2 3 4 5 6 7Tip length, s

(mm) 0 1.93 2.34 3.86 1.93 1.93 1.93Tip taper angle,

(deg) 45 45 45 45 0 30 63Tip base diameter,

db (mm) 10.03 8.43 8.09 6.83 10.03 9.00 8.43

Melt tip orifice diameter, do 4.76 mm; gas jet diameter, dj 0.813mm; and gas jet apex angle, 45 deg.

information is valuable for the design of an efficient at-omizer and its operation, for an understanding of the pro-cess dynamics, as well as for the intelligent control ofHPGA through real-time modification of the geometric

parameters.

II. NUMERICAL MODEL

Based on the assumptions of steady state, axisymmetric,compressible, and turbulent flow, the conservation equa-tions for the present problem are defined by the continuitymomentum, and energy equations. Mi et al.[14] have out-lined these equations, which are written subsequently usingstandard index notation with subscripts i and j referring tothe axial and radial components of the variable, as appro-

priate.Continuity equation:

(u )i 0 [1]

xi

Momentum equation:

(u u ) upi j j ( ) [2]eff

x x x xi i i i

Energy equation:

(u h) p hi u (K ) [3]j eff

x x x xi j i i

In these equations, xi is the spatial coordinate, ui is thetime-averaged velocity component, p is the time-averaged

pressure, n is the time-averaged static enthalpy, is thetime-averaged density, and is the viscous dissipationfunction. The terms eff and Keff are the effective viscosityand conductivity, which include laminar and turbulent ef-fects, respectively. The thermodynamic properties of the

gas are related through the equation of state for a perfectgas.

The previous equations can be transformed to the form ofa generalized transport equation, which includes diffusionterm, convection term, and source term. Then, a general so-lution procedure based on the SIMPLEST algorithm[15] isapplied to solve all the governing equations. The differentialequations are formulated using the control volume ap-

proximation. A structured grid was created using high re-finement in regions containing large pressure gradients. Theturbulence model closure, computational domain, gridmesh, code validation, boundary conditions, and their spe-cific handling were discussed in detail previously.[14]

The effect of the melt tip geometry on the HPGA gasflow field and melt aspiration condition was studied for fourextension lengths and four tip taper angles, as shown inTable I. The values selected bracket a range found in theliterature.[1,3,5,6,814] The general melt tip design remains sim-ilar to the generic design of previous studies[1014] andreplicated by others.[5,9] References[5,9,1014] cite using theTip2 design. In the geometry used here, the melt tip frustumsurface coincides with the innermost edge of the gas jetring. The total apex angle of the gas jet is 45 deg.The gas properties used are those of argon.

Referring to Table I, the melt tip extension length s rep-resents the tip length extending beyond the gas jet orifice

plane. In particular, a fully retracted melt tip, common toseveral gas atomizer processes, has a tip length of s 0.0mm. The tip taper angle is defined as twice the angle be-tween the tip frustum surface and the nozzle axis. Here, an 45 deg tip exactly matches the 45 deg gas jet apexangle. An 63 deg tip turns the extended tip frustumsurface and the gas flow toward the nozzle axis, while an 30 deg tip turns the gas flow away from the nozzleaxis. An 0 deg tip (Tip5) is simply an extended straighttip, and forms a limit for taper angle. An example of anozzle having an extended straight melt tip was presented

by Unal.[6]

III. RESULTS AND DISCUSSION

Results are expressed in terms of Mach number contoursin the gas flow field and static pressures along the melt tip

base. Pressures external to the flow domain are held at pa 1 atm. Each geometry is studied over a range of stag-nation pressures. The stagnation pressure po is set at the gasmanifold just upstream of the onset of the straight gas jet

passage. This pressure differs from the operating (stagna-tion) pressure measured well upstream, usually at an at-omizer regulator. The relation between operating stagnation

pressure and gas jet stagnation pressure depends on the par-


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ticular system used, and the two will differ considerablydue to internal shocks and losses in the transmission lines.

A. Tip Extension Length Variation

Figure 2 illustrates the calculated HPGA gas flow Machnumber contours at the stagnation pressure Po 4 MPa forthe four melt tip lengths. Please note that the rough orstepwise solid surfaces seen in Figure 2 are a manifestationof the plotting program and not of the grid used. For pre-

sentation purposes, the flow chart is rotated 90 deg, suchthat the gas flow is from left to right, and because of theimposed axisymmetric assumption, the flow is cut at thecenterline. Subsonic areas are dark and supersonic areas aregray to bright in these figures. Shock structures are exhib-ited by bright or gray to dark transitions.

For all tip lengths used, the presence of a subsonic re-circulating flow (dark area) attached to the tip base andshock structures downstream of this recirculation regioncan be discerned. On exciting a gas jet orifice, the gas ex-

pands through a large expansion wave (bright areas). Withincreased tip length, this expansion wave becomes largerand moves downstream. The maximum Mach number ob-served in the bulk gas flow field always occurs at the center

of the expansion wave pattern (brightest area). For s 0.0mm and s 1.93 mm, a normal shock is seen to form

perpendicular to the nozzle centerline downstream of themelt tip, approximately halfway down the flow domain. Be-cause of the axisymmetric condition, the shock is a diskwith a radius that increases with tip extension. For s 2.34mm ands 3.86 mm, the shock along the nozzle centerline

becomes concave.The computed values for pressure at the melt tip base

centerline are presented as Pb/Po for four melt tip lengthsin Figure 3. Over a wide range of low operating pressures,the base centerline pressure Pb remains approximately equalto the ambient pressure Pa independent of tip extension. Athigher pressures, as indicated by small values of Pa/Po, themelt tip base pressure ratio, Pb/Po, becomes independent ofstagnation pressure. As noted in earlier studies ofHPGA,[1014] this behavior is associated with the open orclosed character of the recirculation region. In an openwake, the recirculation region is not surrounded by shockstructures and the melt tip has a pressure remains sensitiveto ambient flow conditions. At some critical pressures, theinternal shock structures elongate to enclose the recircula-tion zone, leading to a closed wake, and the Pb/Po valuethat is essentially independent of changes in ambient flowconditions. For the extended tips tested, the wake closesnear Pa/Po 0.025 (Po 4 MPa), while for the fully re-tracted melt tip, the wake closes near Pa/Po 0.06 (1.7MPa). Within the closed wake operation, the P

b/P

ovalues

are lower for a longer tip than those for a shorter tip. InFigure 4, it is clear that an extended melt tip with 45deg taper angle develops a subambient to slightly over-ambient centerline pressure at all operating pressures cited,with base pressure decreasing with tip length. The fullyretracted tip shows overambient pressures.

Morton[13] studied the effect of the melt tip length on theHPGA melt tip centerline base pressure using the hydraulicanalogy. The results are shown in Figure 5. Although thehydraulic analogy is not exact, its results provide a good

indication of the relative trend in pressure in high-speedgas flows. The measured pressure results of Figure 5 arequite consistent with the numerical results of Figure 4: thefully retracted tip develops a much higher base centerline

pressure at all operating pressures relative to the extendedtips; a longer tip exhibits a lower base pressure, especiallywithin the high pressure regime; and the extended tips alldisplay a minimum base centerline pressure for the oper-ating pressures tested. This minimum value is found to oc-cur at a higher operating pressure with a longer extensionfor a given taper angle.

Figure 6 illustrates the HPGA melt tip base pressure dis-tributions for the four tip lengths at Po 4 MPa. The non-dimensional radial distance from the centerline is based onmelt tip radius, db/2. As seen in the figure, a very large

portion of the base for a fully retracted tip experiences ahigh overambient pressure. Among the three extended tipstested, the s 3.86 mm tip experiences the lowest base

pressure. Base pressure decreases systematically with in-creasing tip length. Still consistently, the base pressureshows a maximum at the base centerline and a minimumnear the base corner, this minimum due to flow expansionaround the corner.

It was suggested

[10]

that the driving-pressure force alongthe melt tip base is important in the HPGA process, as thispressure gradient would tend to drive the melt radially out-ward from the base centerline to the base corner. Becauseof the high kinetic energy in the high-shear layer betweenthe recirculation region and the expansion region, an effi-cient atomization would take place there. As shown in Fig-ure 6, a longer tip, with its more uniform base pressure,develops a smaller pressure gradient along the base betweenthe centerline and the corner, relative to a shorter tip. How-ever, as manifested in Figure 4, a longer tip encourages alower base centerline pressure. This provides a designtrade-off between the pressure driving force available totransport melt to the shear layer and aspiration for a stable

melt flow.

B. Tip Taper Angle Variation

Figure 7 illustrates the calculated HPGA gas flow Machnumber contours for the four taper angles at Po 4 MPa.All tips are found to demonstrate a large expansion wave

pattern downstream of the melt tip, with the straight tiphaving the largest expansion and correspondingly the high-est maximum Mach number observed. The shock structuresare well defined for all tips studied. For the three taperedtips, normal shocks are seen to form perpendicular to thecenterline about midway down the flow domain, with the 63 deg tip developing the smallest radius. The recir-culating flow region found in the tip taper angle study re-mains subsonic for all tips tested.

The HPGA melt tip centerline base pressure ratios forthe four tip taper angles are shown in Figure 8. The resultsreveal open wake and closed wake behavior similar to thetip extension studies. However, the actual Pb/Po values in-crease with taper angle. Similar results have been reportedfor the case of the axisymmetric truncated plug nozzle flowfield. In particular, a smaller plug angle, equivalent to taperangle here, led to a lower Pb/Po value.

[16]

The melt tip centerline base pressure, Pb, experiences a


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Fig. 2The calculated HPGA gas flow Mach number contours at Po 4 MPa for the four melt tip lengths.

minimum with stagnation pressure (Figure 9). The 0deg straight tip experiences the lowest value of Pb, whilethe 63 deg tip experiences an overambient conditionfor all Po applied.

Figure 10 presents the experimental results based on thehydraulic analogy of HPGA melt tip base centerline pres-sure measurements for three model tips.[13] The measuredresults are consistent with the calculated results shown inFigure 9: the 63 deg tip experiences the highest pres-sure, and the straight 0 deg tip experiences the lowest

pressure. Centerline pressure measurement curves also dis-play a minimum for all three model tips, with a smallertaper angle at a higher stagnation pressure. With a similargeometric melt tip, Veistinen et al.[8] reported that an over-ambient base pressure developed and increased as the meltfeed tube taper angle was increased.

Figure 11 shows the calculated HPGA melt tip base pres-sure distributions at Po 4 MPa for the four tip taperangles used. The 63 deg tip experiences an overam-

bient pressure over half of the tip radius. With a decrease


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Fig. 3The calculated HPGA melt tip base centerline pressure ratios dueto different melt tip lengths.

Fig. 4The calculated HPGA melt tip base centerline pressures due todifferent melt tip lengths.

Fig. 5HPGA melt tip base center pressure measurement results for fourmelt tip lengths using the hydraulic analogy.[13]

Fig. 6The calculated HPGA melt tip base pressure at Po 4 MPa forthe four melt tip lengths.

in taper angle, the base region experiencing this overam-bient pressure decreases in area. The straight 0 deg tipexhibits the lowest base pressure and, as such, develops the

best aspiration effect of all taper angles tested. On the otherhand, it is noticed that the straight tip develops a moreuniform pressure distribution over its tip base comparedwith the other three tapered angle tips. This demonstratesthat a tip with a larger taper angle develops a higher driving

pressure gradient along its base. But a smaller taper angletip develops a lower melt tip base centerline pressure.Therefore, this trade-off clearly suggests that when consid-ering both atomization stability and atomization effective-ness, the HPGA melt tip taper angle needs to be chosencarefully to balance the driving force and aspiration needed.

Both the numerical results presented herein and support-ing experimental observations demonstrate that the gas flowcharacteristics and melt tip aspiration in the HPGA processare strongly affected by the melt tip extension length andtip taper angle. Increasing extension length or decreasingtaper angle tend to increase aspiration, while the oppositegeometry changes are favorable for a radial melt flow.

Hence, these geometric variables are important designparameters and can be expected to influence the gas at-omization process.

IV. CONCLUSIONS

The effects of melt feed tube tip extension length andmelt tip taper angle on the HPGA gas flow field and meltaspiration are examined numerically. The main gas flowfeatures characterized in the numerical solutions preserve astrong recirculating region, a shear layer, and shock struc-tures in qualitative conformity with experimental results.The calculations indicate that a fully retracted melt tip de-velops a high overambient pressure along the melt tip baseand may experience problems for stable atomization, whilea longer melt tip extension develops a better melt aspiratingeffect, providing for a stable melt flow. However, a strong

pressure gradient along the melt tip base that drives meltradially outward toward a strong shear layer responsible fordroplet size refinement diminishes with melt tip extension.The results also show that a small tip taper angle encour-


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Fig. 7The calculated HPGA gas flow Mach number contours at Po 4 MPa for the four melt tip taper angles.

ages a desirable aspirating condition. But a tip with a largertaper angle develops a higher driving-pressure gradientalong its base. Therefore, this trade-off clearly suggests thatwhen considering both atomization stability and effective-ness, the HPGA melt tip taper angle and extension lengthneed to be chosen carefully to balance between aspirationand the driving force that provides size refinement.

NOMENCLATUREdb melt tip base diameterdj gas jet diameterdo melt tip orifice diameterh time-averaged enthalpy

Keff effective conductivityp time-averaged pressurePa ambient pressurePb melt tip base center pressure


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Fig. 8The calculated HPGA melt tip base centerline pressure ratios dueto different melt tip taper angles.

Fig. 9The calculated HPGA melt tip base centerline pressures due todifferent melt tip taper angles.

Fig. 10HPGA melt tip base center pressure measurement results forthree melt tip taper angles using the hydraulic analogy.[13]

Fig. 11The calculated HPGA melt tip base pressure at Po 4 MPa forthe four melt tip taper angles.

Po stagnation pressure assumed to exist at the gasmanifold

s melt tip extension lengthui time-averaged velocity component

xi coordinate melt tip taper angle gas jet apex angleeff effective viscosity time-averaged density viscous dissipation function

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1. J.D. Ayers and I.E. Anderson: J. Met., 1985, vol. 37, pp. 16-21.2. N.J. Grant: J. Met., 1983, vol. 35, pp. 20-27.3. I.E. Anderson and B.B. Rath: Proc. Symp. on Rapidly Solidified

Crystalline Alloys, S.K. Das, B.H. Kear, and C.M. Adam, eds.TMS-AIME, Warrendale, PA, 1985, pp. 219-45.

4. R.A. Ricks and T.W. Clyne: J. Mater. Sci. Lett., 1985, vol. 4, pp.814-17.

5. S.D. Ridder and F.S. Biancaniello: Mater. Sci. Eng., 1988, vol. 98,pp. 47-51.

6. A. Unal: Metall. Trans. B, 1989, vol. 20B, pp. 833-43.7. U. Backmark, N. Backstrom, and L. Arnberg: Technical Bulletin No.

IM-2051, Swedish Institute for Metals Research, Stockholm, 1985.8. M.K. Veistinen, E.J. Lavernia, J.C. Baram, and N.J. Grant: Int. J.

Powder Metall., 1989, vol. 25, pp. 89-92.9. P.I. Espina, S.A. Osella, S.D. Ridder, F.S. Biancaniello, and G.E.

Mattingly: in P/M in Aerospace, Defense, and DemandingApplications, APMI, Princeton, NJ, 1993, pp. 101-07.

10. I.E. Anderson and R.S. Figliola: in Modern Development in PowderMetallurgy, P.U. Gummeson and D.A. Gustafson, eds., APMI,Princeton, NJ, 1988, pp. 205-17.

11. I.E. Anderson, R.S. Figliola, and H. Morton: Mater. Sci. Eng., 1991,vol. 148, pp. 101-13.

12. I.E. Anderson, H. Morton, and R.S. Figliola: in Physical Chemistryof Powder Metals Production and Processing, W.M. Small andD.G.C. Roberson, eds., TMS, Warrendale, PA, 1989, pp. 229-49.

13. H. Morton: Masters Thesis, Clemson University, Clemson, SC, 1990.14. J. Mi, R.S. Figliola, and I.E. Anderson: Mater. Sci. Eng., 1996, vol.

208, pp. 20-29.15. S.V. Patankar: Numerical Heat Transfer and Fluid Flow, McGraw-

Hill Hemisphere, New York, NY, 1980.16. C.R. Hall and T.J. Mueller: J. Spacecraft, 1972, vol. 9, pp. 337-42.

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