Solidiﬁcation in Spray Forming
Solidiﬁcation in spray forming takes place in two distinct steps: typically half of the alloy latent
heat is removed rapidly from the droplet spray created by gas atomization; the droplets are then
constituted into a billet at deposition where the remaining liquid fraction solidiﬁes relatively
slowly. However, within the droplet spray, individual droplets have diﬀerent thermal and
solidiﬁcation histories and depositing droplets may be solid, mushy, or liquid. Despite many
studies of solidiﬁcation behavior in spray forming, uncertainties and some misconceptions re-
main on how the solidiﬁcation conditions in the spray and billet interact to give rise to the
characteristic spray-formed microstructure comprising reﬁned, polygonal/equiaxed primary
grains with low levels of microsegregation. This article presents a simple numerical model for
the spray-formed grain size arising from the deposition of the various droplets in the spray and
combines insights provided by the model with previous investigations of the phenomena
occurring during and immediately after deposition to propose a comprehensive description of
the important solidiﬁcation behavior during spray forming. Remelting, grain multiplication,
thermal and elemental equilibration, and microstructural coarsening are proposed to play a
critical role in the evolution of the spray-formed microstructure.
Ó The Minerals, Metals & Materials Society and ASM International 2007
THE spray forming process is an advanced casting
process for the manufacture of billet materials and is
shown schematically in Figure 1. Spray forming com-
prises the sequential steps of (1) the continuous gas
atomization of a melt stream to produce a spray of 10- to
500-lm-diameter alloy droplets, (2) droplet cooling at
and acceleration to 50 to
under the action of the atomizing gas, (3) the
deposition of droplets at the growing spray-formed billet
surface, and (4) the relatively slow cooling at 0.1 to
and solidiﬁcation of any residual liquid in
the spray-formed billet.
As well as potential near-net-
shape beneﬁts, the primary advantage of the spray
forming process is the ability to manufacture alloy
compositions that are problematical in conventional
processes such as ingot casting, direct chill casting, and
powder metallurgy. Commercial examples of alloys
manufactured by spray forming include Al-Si based
alloys for cylinder liners,
high speed and speciality
Si-Al alloys for thermal management
and Al-Nd alloys for sputtering targets.
The disadvantages of spray forming include as-sprayed
porosity that requires closing by hot isostatic pressing or
other downstream processes, and losses because not all
droplets created by atomization end up in the spray-
formed billet. These aspects and some of the underlying
process physics have been reviewed in Reference 1.
Despite many investigations of the relationship
between the spray forming conditions, the solidiﬁcation
conditions arising, and the as-sprayed microstructure,[7–16]
some uncertainties remain in how the underlying process
physics and mechanics give rise to the advantageous
spray-formed microstructure that typically comprises
(1) equiaxed/polygonal grains of diameter typically in
the range 20 to 50 lm and the complete absence of
(2) high levels of microstructural homogeneity and
low levels of microsegregation regardless of posi-
tion in the billet, and
(3) the ability to produce this characteristic spray-formed
microstructure in all engineering alloy systems.
For example, Figure 2(a) shows a 26 kg Al-5.3
Mg-1.2 Li-0.28 Zr alloy billet spray formed at Oxford
University, and the corresponding as-spray-formed
microstructure is shown in the electron backscattered
diﬀraction (EBSD) orientation map in Figure 2(b). The
as-spray-formed grain size was ~15 lm and the poros-
ity was <1 area pct. In a further example, Figure 3(a)
shows an optical micrograph of a chill-cast Si-30 wt pct
Al alloy comprising coarse and highly defective primary
Si and interprimary a-Al/Si eutectic with signiﬁcant
shrinkage porosity associated with the wide freezing
range of ~650 °C. In contrast, Figure 3(b) shows the
microstructure of a spray-formed billet of the same
alloy. The microstructure now comprised a more glob-
ular, reﬁned network of primary Si continuously pen-
etrated by a-Al arising from a fully divorced eutectic.
P.S. GRANT, Cookson Professor of Materials, is with the Depart-
ment of Materials, Oxford University, Oxford OX1 3PH, United
Kingdom. Contact e-mail: firstname.lastname@example.org
This article is based on a presentation made in the symposium
entitled ‘‘Solidification Modeling and Microstructure Formation: In
Honor of Prof. John Hunf ’’, which occurred March 13–15, 2006, in the
TMS Spring meeting in San Antonio, TX, under the auspices of the
TMS Materials Processing and Manufacturing Division, Solidification
Article published online February 2, 2007.
1520—VOLUME 38A, JULY 2007 METALLURGICAL AND MATERIALS TRANSACTIONS A
Figure 3(c) shows an EBSD orientation map for the
same microstructure, indicating that the primary Si
network consisted of randomly oriented Si grains with
an average grain size of ~1 lm. These Si-rich alloys oﬀer
an attractive combination of relatively high thermal
conductivity, low thermal expansion, and low density
for thermal management applications.[4,17]
This paper presents a numerical model for the spray-
formed grain size and combines insights provided by the
model with previous investigations of the phenomena
occurring during and immediately after deposition to
propose a comprehensive description of important
solidiﬁcation behavior during spray forming.
II. A MODEL FOR THE SPRAY-FORMED GRAIN
Under a wide range of atomizing conditions for
diﬀerent alloys, the volumetric or mass mean particle
diameter dV has been related to operating variables by
the widely quoted formula
nð1 þ RÞ
where D is the melt delivery nozzle diameter, n is the
ratio of molten metal to gas kinematic viscosity, R is the
ratio of molten metal to atomizing gas mass ﬂow rates,
and b and n are experimental constants. Reported values
of b and n for a wide range of materials and atomizer
geometries are typically 4 to 5 · 10)6
and 0.5, respec-
The probability density Pi of droplet diameters
di about the volumetric mean dV approximates to a log-
normal distribution with a standard deviation r given by
p exp À
ln di À ln dV
where the droplet diameter interval Ddi has a midpoint
di, n is the number of droplet intervals, and
PiDdi ¼ 1 ½3
Assuming dV and r are 100 lm and 0.7 ln (lm),
Figure 4 shows the resulting log-normal
droplet diameter distribution and is typical of those seen
Previous investigations of droplet dynamic and ther-
mal behavior have used numerical models to predict the
solid fraction of the diﬀerent droplet diameters created
Fig. 2—(a) A typical 26-kg Al-5Mg-1.2Li-0.28Zr billet spray formed
at Oxford University and (b) an Al-5Mg-1.2Li-0.28Zr as-sprayed
microstructural orientation map obtained by EBSD showing noncol-
umnar/dendritic equiaxed polygonal grains characteristic of the
spray forming process.
Fig. 1—Schematic of the spray forming process for the manufacture
METALLURGICAL AND MATERIALS TRANSACTIONS A VOLUME 38A, JULY 2007—1521
by atomization as they move away from the point of
atomization under the action of the cold, accelerating
An example of these type of calcu-
lations is shown in Figure 5 for Al-4 wt pct Cu droplet
diameters of 10, 50, 120, and 200 lm. For these droplets
depositing at an axial distance of 0.4 m, diameters less
than 50 lm were predicted to be fully solid; droplets
with a diameter greater than ~200 lm were fully liquid;
and droplets with intermediate diameters were mushy
with a variation in droplet solid fraction dependent
upon droplet diameter. Using these calculations, the
boundaries between droplet diameters that were solid,
mushy, or liquid are shown in Figure 4. While the
details and simplifying assumptions of the various
modeling studies of droplet ﬂight and simultaneous
solidiﬁcation in spray forming may vary, all show a
spread of droplet solid fractions at deposition.
Assuming that the variation of mushy droplet solid
fraction is linear with the logarithm of droplet diameter,
as shown in Figure 6, and ignoring any equilibration
of droplet temperatures or solid fractions (later
discussion), the mass-averaged spray solid fraction fS
Fig. 3—(a) Optical micrograph of chill-cast Si-30 wt pct Al, (b) opti-
cal micrograph of as-spray-formed Si-30 wt pct Al, and (c) EBSD
orientation map of (b).
Fig. 4—A typical log-normal droplet diameter probability density
distribution on a mass or volume basis obtained by gas atomization,
with superimposed assumed boundaries between solid, mushy, and
liquid droplets at the point of deposition during spray forming.
Fig. 5—Calculated variation of Al-4 wt pct Cu droplet solid fraction
as a function of droplet diameter and axial distance from the point
of atomization, indicating that at an axial distance of 0.4 m, the
spray comprises a mixture of solid, mushy, and liquid droplets.
1522—VOLUME 38A, JULY 2007 METALLURGICAL AND MATERIALS TRANSACTIONS A
at the point of deposition is given by the integration of
the distribution of droplet solid fractions in Figure 4.
For the assumed distribution of droplet diameters and
solid, mushy, and liquid droplets in Figures 4 and 6, fS is
0.5, which is in broad agreement with previous studies
(e.g., References 6, 9 and 16).
Figure 7 shows the droplet diameter distribution of
Figure 4 replotted in terms of a normalized number
distribution, together with the same superimposed
boundaries between solid, mushy, and liquid droplets.
Figure 7 shows that the number mean of the distribu-
tion dN was 23 lm and that, while the majority of
droplets by mass or volume arriving at the point of
deposition were either liquid or mushy, in terms of the
number of droplets, the distribution was dominated by
the large number of solid droplets.
The number of solid particles NV,i in the droplet
diameter interval Ddi is given by
NV ;i ¼
for fi > 0 ½4
wherefi is the droplet solid fraction and, since the
probability distribution function in Figure 4 has an
integrated area of unity, the total number of solid and
mushy particles per unit volume NV is simply given by
NV ;i ½5
Assuming in the ﬁrst instance that (1) each solid or
mushy particle in the spray nucleates/becomes one
‘‘embryonic’’ grain in the billet top surface immediately
after deposition and (2) there is no signiﬁcant additional
nucleation in the liquid fraction of the billet top surface
because of little/no undercooling and an excess of nuclei
arising from the deposition of a large number of solid/
then the average embryonic grain
diameter in the billet top surface do can be estimated from
A similar approach has been presented to estimate the
number of subgrains formed in a depositing droplet
during rapid solidiﬁcation.
Table I shows the calculated variation of fS and do
using Eqs.  through  as a function of assumed
deposition conditions described by the particle size
distribution (given by dV and r and where r was varied
independently of dV in this case but recognizing some
interdependency has been suggested
) and the diam-
eters for fully solid (df = 1) and fully liquid (df = 0)
droplets. Two other variables were introduced into the
calculations in Table I:
(1) a critical droplet diameter d*
below which droplets
were assumed to be too small and to have insuﬃ-
cient axial momentum to intersect the billet sur-
face, but instead were seeded in the gas ﬂow, swept
around the billet, and removed as overspray; and
(2) a grain multiplication factor G to account for there
to be more than one nuclei per solid or mushy
droplet, and to account for the breakup of mushy
and remelted droplets on impact with the billet
surface so that droplets may nucleate more than
one embryonic grain in the mushy zone.
Unfortunately, there is little guidance in the litera-
ture as to the possible values for G in the spray
forming process, and so the approach taken here was
to investigate the sensitivity of the calculated embry-
onic grain size to a wide variation in G from 1 to 20. In
this way, the importance of multiple nuclei per droplet
Fig. 6—Assumed variation of droplet solid fraction fi with diameter
di, for fi = 1 at di = 50 lm and fi = 0 at di = 200 lm.
Fig. 7—The log-normal droplet diameter probability density distri-
bution on a number basis for the data in Fig. 4, with the same
superimposed boundaries between solid, mushy, and liquid droplets
at the point of deposition during spray forming.
METALLURGICAL AND MATERIALS TRANSACTIONS A VOLUME 38A, JULY 2007—1523
and grain multiplication in spray forming may be
Table I shows that do was comparatively insensitive to
the signiﬁcant changes in the assumed deposition
conditions and that while some of the model assump-
tions were somewhat arbitrary, they made only a
relatively small diﬀerence to the calculated embryonic
grain diameter. For example, Figure 8 shows the eﬀect
of the grain multiplication factor G on the calculated
embryonic grain diameter. As the multiplication factor
increased from 1 to 5, the calculated grain diameter
decreased from 31 to 18 lm, and reduced further to
11 lm for a grain multiplication factor of 20. Embry-
onic grain diameters cannot be readily compared with
ﬁnal grain diameters from microstructural analysis,
because, as described later, grain coarsening in the
liquid/mushy state is expected. Nonetheless, calculated
embryonic grain diameters in Table I of 11 to 58 lm,
even when signiﬁcant grain multiplication eﬀects or low
spray solid fractions were assumed, are consistent with
ﬁnal spray-formed grain diameters.
may show grain size variations from a few or tens of
microns to millimeters within a billet, the comparatively
low sensitivity of predicted and experimental grain
diameters to strong variations in the condition of
depositing droplets in spray forming arises from two
(1) Figure 7 shows that regardless of speciﬁc spray
forming conditions, it is likely there will always be
a surfeit of small solid particles in the spray,
which, while contributing little to the billet mass,
dominate grain nucleation behavior in the billet
top surface region.
(2) Equation  shows that even when large changes
in the number of solid particles per unit volume in
the spray may be realized, the resulting grain size
has only a relatively weak cube root dependency
on the number of solid particles.
In order to provide experimental support for the grain
size model, the eﬀect of the number of solid particles per
unit volume on the microstructural scale of Si-30 wt pct
Al billets was investigated by the simultaneous coinjec-
tion of solid Si-30 wt pct Al particles during spray
forming. The coinjected solid particles were entrained
continuously into the droplet spray and deposited at the
billet surface as if they had originated from atomization
directly. In this way, the number of solid particles was
controlled independently of the melt mass ﬂow rate and
atomizing conditions. The coinjection of overspray has
become standard commercial practice to improve spray
forming process yield and productivity.
ments were carried out at Sandvik Osprey (Neath,
United Kingdom), and although eﬀorts were made to
keep spray forming conditions constant as the ratio of
particle mass ﬂow rate to liquid alloy mass ﬂow rate
was altered, some small changes had to made to the
proprietary spray forming conditions to control billet
shape, porosity, etc. However, these changes were slight
compared with the changes in coinjection mass ﬂow
rate. The typical microstructure of spray-formed Si-30
wt pct Al has been shown in Figures 3(b) and (c).
Figure 9 shows the variation of the inverse of the
a-Al/Si interface length per unit area obtained by
quantitative image analysis of many digitized optical
micrographs taken from the billet centers as a function
of the ratio of solid Si-30 wt pct Al particle to liquid Si-
30 wt pct Al mass ﬂow rates. As the particle mass ﬂow
rate increased, there was a progressive and marked
microstructural reﬁnement, described by the reduction
in the inverse of interface length per unit area. Consis-
tent with the trends of the calculated embryonic grain
Fig. 8—Variation of calculated average embryonic grain diameter do
as a function of assumed grain multiplication factor G.
Table I. Calculated Variation in fS and do Using Equations
 through  as a Function of Assumed Deposition Condi-
tions Described by the Particle Size Distribution (dV, r); the
Diameters for Fully Solid (df = 1) and Fully Liquid (df = 0)
Droplets; the Critical Droplet Diameter for Deposition d*
the Grain Multiplication Factor G
df = 1
df = 0
(lm) G fS
100 0.7 50 200 0 1 0.5 31
100 0.7 50 200 10 1 0.5 32
100 0.7 50 200 20 1 0.5 37
100 0.6 50 200 10 1 0.5 38
100 0.5 50 200 10 1 0.5 44
100 0.8 50 200 10 1 0.5 28
150 0.7 50 200 10 1 0.31 40
250 0.7 50 300 10 1 0.21 58
80 0.7 50 200 10 1 0.61 28
100 0.7 30 200 10 1 0.39 29
100 0.7 80 200 10 1 0.62 34
100 0.7 50 150 10 1 0.43 30
100 0.7 50 300 10 1 0.59 34
100 0.7 50 200 10 2 0.5 24
100 0.7 50 200 10 5 0.5 18
100 0.7 50 200 10 7.5 0.5 16
100 0.7 50 200 10 10 0.5 14
100 0.7 50 200 10 15 0.5 12
100 0.7 50 200 10 20 0.5 11
1524—VOLUME 38A, JULY 2007 METALLURGICAL AND MATERIALS TRANSACTIONS A
diameters in Table I, Figure 9 suggests that since the Si
volume fraction was measured to be constant for all the
billets investigated, co-injection directly increased the
number and decreased the ﬁnal size of Si particles in the
mushy top surface region of the billet.
Further evidence for the manipulation of spray-
formed grain size by direct control of the number
density of nuclei in the spray is provided in References
22 and 23 where deliberate partial oxidation of Al alloy
droplets was used to create a ﬁne dispersion of highly
numerous but negligible volume oxide particles that
signiﬁcantly reduced the as spray-formed grain size.
III. REMELTING AND EQUILIBRATION
The essentially geometric approach to understanding
spray-formed grain diameter described above does not
account for thermal and compositional equilibration
eﬀects that must occur as droplets with a wide range of
thermal conditions impact onto the billet surface that
has a temperature Teq and solid fraction feq. Outside
thermal transients at the beginning and end of deposi-
tion, Teq and feq should remain approximately constant
as the billet forms.
The variation of Teq and feq with spray forming
parameters has been investigated using a range of
techniques including pyrometry,
Figure 10 shows the ﬁnal grain
size of spray-formed 2000 series Al alloy billets versus
billet top surface solid fraction at that point measured
by plunging a thermocouple into the billet top surface
during spray forming under a range of conditions.
wide range of solid fractions from 0.3 to 0.95 but a
narrow range of grain diameters of 24 to 11 lm were
measured, with grain diameter decreasing progressively
with increasing top surface solid fraction. The study
conﬁrmed that there must be a signiﬁcant liquid fraction
on the billet top surface throughout the spray forming of
good quality billets.
The spray forming conditions
required to ensure that the billet top surface always
contains liquid have been investigated analytically
and, in some studies, optimum conditions for acceptable
porosity, grain size, yield, etc. identiﬁed.
uous presence of a liquid fraction on the billet top
surface distinguishes spray forming as a semicontinuous
casting process distinct from incremental spray deposi-
tion processes such as arc[26,27]
where discrete droplet deposition events are preserved in
the ﬁnal microstructure.
When droplets arrive at the mushy billet top surface,
their velocities are typically in the range 50 to 100 ms)1
depending upon spray forming conditions,[8,29]
following phenomena can occur: solid particles bounce
from the billet surface and are lost as overspray;
substantially liquid droplets splash at the billet surface,
generating a host of smaller secondary droplets that
then either redeposit onto the surface or are lost as
or mushy/liquid droplets rapidly spread
and are incorporated into the billet.
Droplet spreading on impact with a solid or mushy
surface under spray forming conditions is rapid,[32–34]
and lateral spreading of a fully liquid droplet occurs
within a few microseconds.
Solidiﬁcation is relatively
slow and restricts droplet spreading only in the ﬁnal
Once spreading is complete, droplets will
equilibrate to the billet average surface temperature Teq.
The time for thermal equilibration tT
eq can be estimated
where x is a representative equilibration distance and a
is the thermal diﬀusivity. For x of up to several
hundreds of microns for the largest depositing droplets
Fig. 9—Si/a-Al interface length per unit area in spray-formed Si-30
wt pct Al as a function of the ratio of the coinjected Si-30 wt pct Al
particulate ﬂow rate to liquid Si-30 wt pct Al ﬂow rate during spray
forming of billets.
Fig. 10—The ﬁnal grain size of spray-formed 2000 series Al alloy
billets versus billet top surface solid fraction measured by plung-
ing of thermocouples into the billet top surface during spray
METALLURGICAL AND MATERIALS TRANSACTIONS A VOLUME 38A, JULY 2007—1525
and assuming a typical value of a = 10)5
metals, then tT
eq < 0.1 s.[1,19]
Large droplets with a
temperature Ti > Teq will be quenched on deposition;
solid particles with Ti < Teq will be reheated, and
because the billet top surface temperature must be
above the alloy solidus temperature,
Ti < Teq must partially remelt.
For mushy droplets,
an abrupt change in droplet cooling rate from the
spray of typically 102
to that in the billet of
typically 0.1 to 10 Ks)1
will destabilize the dendrite/cell
network growing within the droplet and promote
dendrite/cell network breakup.
Accompanying thermal equilibration will be equili-
bration to the billet solid fraction feq. Figure 11 shows
schematically the changes in solid fraction during the
spray forming process as a function of axial distance
from the point of droplet atomization. The range of
droplet solid fractions depositing at the billet surface
equilibrate in the region of the billet top surface, labeled
the equilibration zone, and the billet solid fraction
(ignoring any radial variations across the billet) then
gradually increases as the billet is slowly withdrawn to
maintain a constant spray distance. Figure 11 also
indicates the smallest solid droplets seeded in the gas
ﬂow, larger solid particles undergoing reheating and
remelting, and mushy particles undergoing thermal
shock and breakup (grain multiplication).
As-spray-formed microstructures such as those shown
in Figure 2(b) indicate that while up to typically half of
the impacting droplet spray mass is solid in the form of a
ﬁne-scale cellular/dendritic morphology within discrete
droplets such as that shown for the 80-lm-diameter Ni
superalloy IN718 droplet in Figure 12, there are no
remnants of the droplet dendritic or cellular microstruc-
ture in the billet. Therefore, the conditions during and
immediately after deposition involving partial remelting,
thermal/mechanical destabilization, and the presence of
both a signiﬁcant liquid fraction and a relatively ﬂat
temperature gradient must facilitate the rapid spheroi-
dization of remnant solid fragments in an attempt to
minimize solid/liquid interfacial area. This process is
shown schematically in Figure 11.
Only thermal and corresponding solid fraction equil-
ibration has been considered. However, impacting
mushy droplets with diﬀerent solid fractions will contain
residual liquid of varying composition. Similarly, par-
tially remelted solid particles will release relatively
concentrated liquid as last to solidify regions remelt.
The time for equilibration of the liquid concentration tC
in the equilibration zone can be estimated from
where DL is the coeﬃcient of diﬀusion for solute atoms
in the liquid. Again, assuming x is up to several
hundred microns and DL = 10)8
, then tC
typically of the order of seconds. The processes of
thermal and compositional equilibration are shown in
the schematic equilibrium phase diagram of Figure 13
for a binary alloy of composition C0, a billet top
surface temperature of Teq, and corresponding solid
and liquid concentrations at equilibrium of Ceq
L , respectively. Impacting mushy droplets 1 and 2 are
assumed to impact the billet surface with temperatures
T1 > Teq and T2 < Teq, respectively. The equilibration
is shown as two distinct thermal and compositional
steps in (a) and (b), respectively, because of the
diﬀering timescales estimated from Eqs.  and .
On thermal equilibration in Figure 13(a) from T1 to
Teq, droplet 1 has solid and liquid compositions that
are too dilute with respect to Teq; while droplet 2
equilibrates from T2 to Teq and has solid and liquid
Fig. 11—Schematic representation of the changes in solid fraction
during the spray forming process and the accompanying change in
microstructure prior to and after droplet deposition.
Fig. 12—Backscatter electron micrograph of a gas-atomized IN718
Ni superalloy droplet cross section showing microsegregation of Nb
in the ﬁne-scale interdendritic regions.
1526—VOLUME 38A, JULY 2007 METALLURGICAL AND MATERIALS TRANSACTIONS A
compositions that are too enriched. Figure 13(b) shows
that solid and liquid compositions will then equilibrate
S and Ceq
L , respectively. The extent to which this
microscale compositional equilibration is achieved
depends upon the local solidiﬁcation time but is
expected in the manufacture of commercial scale
billets. Equilibration of liquid compositions will neces-
sarily involve remelting and homogenization as con-
centrated, relatively hot liquid is diluted by mixing with
cooler, remelted solid of lower solute concentration.
Figure 14 shows a histogram of the maximum to
minimum variation in the concentration of alloying
elements Cr, Fe, Nb, Mo, Ti, and Al in three Ni
superalloy IN718 overspray particles of diameter 25,
100, and 120 lm of the type shown in Figure 12, as well
as the corresponding IN718 ring spray formed at the
same time. Concentration variations were measured
under identical electron probe microanalysis conditions
where the incident electron beam was moved in 2-lm
steps across a number of grain/dendrite boundary
regions, and this procedure was repeated many times
on randomly selected areas of the microstructures. The
Ni superalloy IN718 with nominal composition in
Table II is notoriously susceptible to Nb micro- and
macrosegregation in conventional processing, and the
potential of spray forming to reduce Nb and other
segregation problems has been studied extensively.[37–40]
Figure 14 shows that as the particle diameter decreased
and the cooling rate increased, the extent of alloying
element microsegregation decreased. For all particles,
Nb showed the greatest tendency for microsegregation.
However, all the overspray particles showed greater
Nb microsegregation than the as-spray-formed micro-
structure, despite the much slower cooling rate of the
spray-formed ring. This reduced microsegregation can
now be understood to arise from the large amounts of
reheating, remelting, mixing, and equilibration occur-
ring in the equilibration zone at the ring surface during
V. MACROSEGREGATION AND COARSENING
In the production of billets on a commercial basis,
billet sizes can be up to 600-mm diameter and several
Fig. 13—Equilibration of mushy droplets with temperatures T1 and
T2, impacting a billet top surface with temperature Teq: (a) tempera-
ture and (b) composition.
Fig. 14—Maximum to minimum concentration variations for alloy-
ing elements Cr, Fe, Nb, Mo, Ti, and Al in three Ni superalloy
IN718 overspray particles of diameter 25, 100, and 120 lm and the
corresponding spray-formed Ni superalloy ring manufactured at the
Table II. Nominal Composition of Ni Superalloy IN718
Element Ni C Cr Mo Al Ti B Nb Fe
Wt pct 52.5 0.04 19 3.0 0.5 0.9 0.02 5.1 18.5
METALLURGICAL AND MATERIALS TRANSACTIONS A VOLUME 38A, JULY 2007—1527
meters in length. Consequently, it can be concluded
safely that liquid persists in the spray-formed billet
during and after spray forming for many minutes.
Supporting evidence for the persistence of billet liquid is
given by macrosegregation proﬁles that can develop in
spray-formed billets under some conditions.
erence 12, Cu macrosegregation on the length scale of a
billet was investigated by chemical analysis, and mea-
sured macrosegregation proﬁles conformed to an inverse
segregation, shrinkage-related mechanism dependent
upon the persistence of interconnected liquid through-
out the billet.
Spheroidized solid arising from droplet impact,
remelting, and equilibration described previously was
shown schematically in Figure 11 as isolated islands of
solid surrounded by continuous liquid in a relatively ﬂat
temperature gradient. Although the liquid fraction is
continuously reducing as heat continues to be extracted
from the billet, primarily by convection to the atomizing
with such a high density of solid particles,
diﬀusion-controlled coarsening of the solid may be
expected to occur. The kinetics of coarsening of spray-
formed microstructures has been shown to follow
typical diﬀusion-controlled growth kinetics at solid
fractions of up to ~0.7[42,43]
0 ¼ KðTÞt ½9
where d is the grain diameter at time t, d0 is the starting
grain size at t = 0, and K(T) is the temperature (and
solid fraction) dependent coarsening constant. In a
variety of Al- and Ni-based spray-formed alloy systems,
K(T) increased with solid fraction as interdiﬀusion
distances decreased, until feq ~ 0.7 at which point solid/
solid impingement became signiﬁcant and K(T) re-
Second-phase particles formed during solidi-
ﬁcation have been suggested to play an important role in
inhibiting grain growth at higher solid fractions where
the thickness of liquid ﬁlms delineating grains
approached the second-phase particle size.
ing constants at feq = 0.7 were measured to be in the
range K(T) = 600 to 1400 and 150 to 200 · 10)18
for Al- and Ni-based alloys, respectively.
that on the basis of Table I for an Al alloy that
d0 = 20 lm and K(T) = 600 · 10)18
then Eq.  indicates an increase in grain
diameter to 30 lm after 30 seconds and to 40 lm after
100 seconds. These times are short compared with billet
solidiﬁcation times. Consequently, coarsening can be
expected to contribute signiﬁcantly to the ﬁnal spray-
formed grain size. Similarly to the remelting and
equilibration behavior previously described, because
coarsening involves the transport of solute by diﬀusion,
it can also be expected to play a role in reducing
The developing spheroidal/polygonal spray-formed
grain structure will allow ﬂow of liquid to feed
solidiﬁcation shrinkage in the spray-formed billet much
better than comparable columnar/dendritic morpholog-
ies expected in conventionally cast billets of similar size,
and because approximately half of the alloy latent heat
has been removed prior to deposition, spray forming is
able to produce cast billets in compositions considered
problematical by conventional casting on account of
their large freezing ranges and tendency to shrinkage-
induced macrosegregation and defects. The Si-30 wt pct
Al alloy shown in Figure 3(c) is an excellent example of
a cast structure that cannot be made easily (if at all) by
other casting routes because the alloy has a freezing
range of ~650 °C. More recently, in a modern imple-
mentation of the original idea to use spray forming to
produce thin strip directly,
spray forming is being
used to remove a signiﬁcant fraction of latent heat prior
to twin roll casting in order to make wrought alloys
with relatively wide freezing ranges more ‘‘castable.’’
A similar approach is being explored for the ‘‘nucleated
casting’’ of large diameter superalloy billets in which
spray forming is used to create a high number fraction
of solid particles in diﬃcult to process Ni superalloys,
before the droplet spray that retains a high mass
fraction of liquid is directed continuously into a water-
Because of the broad range of droplet sizes produced
by atomization, there is always a wide range of droplet
thermal conditions within the spray at deposition. A
simple model has been described to relate the distribu-
tion of droplet diameters and solid fraction at deposition
to the embryonic grain diameter in the billet top surface
region. Calculated embryonic grain diameters of 11 to
58 lm are in good agreement with experimental mea-
surements of ﬁnal grain diameters in a range of alloys.
Model predictions are relatively insensitive to assumed
deposition conditions, because the embryonic grain
diameter is determined primarily by the large number
of fully solid particles always present in the spray.
Although these particles contribute little in terms of
volume to the ﬁnal billet, they are critical in determining
the distribution of embryonic grain diameters.
The large thermal mass associated with the larger
liquid droplets must be suﬃcient to partially remelt the
solid component of the spray on deposition so that the
billet top surface is mushy. The reheating and remelting
processes that must occur for a large number of
depositing solid and mushy droplets destabilizes their
dendritic/cellular microstructure, causing it to break up
and providing a grain multiplication eﬀect. The resulting
high number of small solid fragments in the mushy billet
top surface and the corresponding small interdiﬀusion
distances promote rapid thermal and compositional
equilibration that helps to reduce microsegregation. All
remnants of the droplet microstructure are removed and
the solid component in the billet top surface region
spheroidizes in the relatively ﬂat temperature gradient to
reduce solid/liquid interfacial area.
In the subsequent relatively slow cooling of the still
forming billet, the spheroidized solid coarsens quickly in
the early stages due to small interdiﬀusion distances and
highly mobile liquid ﬁlms between embryonic grains.
1528—VOLUME 38A, JULY 2007 METALLURGICAL AND MATERIALS TRANSACTIONS A
Solidiﬁcation drives the further growth of solid grains
until all liquid is consumed.
Despite the reﬁned microstructures produced by
spray forming, it cannot be described as a rapid
solidiﬁcation process because solidiﬁcation occurs in
two distinct steps, of which only the in-ﬂight solidi-
ﬁcation of droplets is rapid and is experienced by only
about half of the spray mass. The key phenomena
that distinguish spray forming from other solidiﬁca-
tion processes are the enormous amounts of remelting,
mixing, and equilibration that occur in the billet top
surface. Spray forming is best regarded as an intrinsic
or self grain reﬁning casting process in which the
purposes of the atomization step are as follows: (1) to
generate a large number density of solid fragments
and particles that will undergo subsequent partial
remelting and grain multiplication; and (2) to remove
a signiﬁcant fraction of the alloy latent heat prior to
deposition so that important but problematical alloys
with wide freeze ranges can be cast successfully into
billets weighing up to several tons.
The author is grateful to past and present members of
the spray processing group at Oxford University for
their many contributions and ideas.
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