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EFFECT OF POURING CONDITIONS AND GATING SYSTEM DESIGN ON AIRENTRAINMENT DURING MOLD FILLING
Seyyed Hojjat Majidi and Christoph BeckermannDepartment of Mechanical Engineering, University of Iowa, Iowa City, IA, USA
Copyright � 2018 American Foundry Society
https://doi.org/10.1007/s40962-018-0272-x
Abstract
Air entrainment during mold filling is a major source of
oxide inclusion formation in metal casting. A model was
recently developed by the authors to predict the volumetric
air entrainment during pouring of metal castings. In the
course of validating the model with experimental data for
plunging liquid jets, it was shown that the air entrainment
rate during mold filling depends fundamentally on the
velocity and diameter of the jet formed by the pouring
stream. In this study, the effect of more complex pouring
conditions and gating system design on air entrainment is
examined. Simulations are performed investigating the air
entrainment characteristics of castings filled without a
gating system, and with bottom-gated and side-gated filling
systems. Results indicate that reducing the head height and
pouring time, and the addition of a nozzle extension sig-
nificantly reduces the air entrainment. In addition, using an
offset pouring basin with a stopper and pressurizing the
gating system further reduces the volume of entrained air.
Simulation results also show that the generation of vortex
flows inside the filling system is beneficial in reducing free
surface turbulence, which results in less air entrainment
and oxide inclusion formation during mold filling.
Keywords: oxide inclusions, air entrainment model,
gating system, free surface turbulence, vortex flow,
casting simulation
Introduction
Defects in cast parts affect the material properties and
machinability of cast metals, resulting in scrapped castings
or their premature failure. Oxide inclusions are among
casting defects that are troublesome and damaging to
casting performance. Oxygen reacts with the chemical
constituents in the liquid metal during mold filling to form
oxide inclusions which can be transported into the final cast
part. In aluminum alloy castings, as the liquid metal
experiences free surface turbulence, the dry side of the
oxide covering the melt becomes in contact with the dry
side of the other oxide and forms double oxide films or
bifilms1 leading to inclusions entrapped in the casting. In
carbon and low-alloy steel castings, the reaction between
oxygen in the atmosphere and the most reactive elements in
deoxidized steel results in reoxidation inclusions.2 In duc-
tile iron casting, the reaction of magnesium oxide and silica
during magnesium treatment and further during mold fill-
ing is responsible for the formation of dross inclusions.3
Oxide inclusions can form during mold filling as a result of
air entrainment. In free surface flows, air is entrained at
flow discontinuities when the liquid surface experiences
turbulence. Important examples of air entraining flows are:
a liquid jet plunging into a pool, a breaking wave, and a
hydraulic jump, where a fast moving liquid discharges into
a low-velocity atmosphere. Air entrainment manifests itself
in the form of bubbles. For a plunging liquid jet, experi-
ments have shown that the amount of air entrainment
depends on the velocity, diameter, and turbulence level of
the liquid jet.4–7 During filling of metal castings, the
pouring conditions and gating system design affect the air
entrainment, and hence the extent of oxide inclusion for-
mation. Examples of air entraining flows in casting gating
systems include: liquid metal impingement inside sprues,
breaking waves in runners, falling jets inside of the mold
cavity, and rising vertical jets or fountains formed in bot-
tom-gated castings. Once air is entrained, the oxygen in the
entrained air reacts with liquid metal constituents to form
oxide inclusions. Upon formation, these inclusions are
transported with the liquid metal and can ultimately end up
as nonmetallic inclusions in the solidified casting.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 255
Therefore, pouring liquid metals with less air entrainment
is a practical and effective method for the production of
clean and inclusion free, metal castings. Air entrainment
should not be confused with air entrapment. Air entrapment
refers to the creation of air pockets when liquid metal
fronts converge upon each other or on a mold wall. It can
occur due to poorly designed ingates and in addition when
the mold is not properly vented or the mold permeability is
insufficient. Air entrainment, on the other hand, is always
associated with the formation of small bubbles at flow
discontinuities.
Several experimental and numerical studies have been
conducted to investigate the effects of pouring conditions
and filling system design on air/oxide entrainment. Studies
have recommended that reducing the velocity of liquid
metal at ingates has a significant effect on minimizing air
entrainment and oxide defect formation. Several criteria
have been proposed for a critical ingate velocity, below
which free surface turbulence in the flow entering the
casting is avoided. The balance between inertial and sur-
face tension forces acting on the liquid metal has been
suggested to predict the critical velocity at the ingate for
the onset of entrainment.1,8,9 The critical entrainment onset
velocity is calculated to be 0.45 and 0.5 m s-1 for pure
liquid aluminum and iron, respectively. The model devel-
oped by Lai et al.10 compares the simulated instantaneous
free surface area of the melt to the instantaneous free
surface area assuming the liquid metal fills the mold qui-
escently. The difference between these two free surface
areas is then used to determine the oxide entrainment
magnitude. The results of this study found the instanta-
neous free surface area increases significantly above ingate
velocities of 0.4–0.5 m s-1, and compared favorably with
the previous findings.1 Hsu et al.11 proposed using a vortex
gate to reduce the free surface turbulence at the entrance
into the casting. Creating vortex flows inside the gating
systems has been championed in Campbell’s work.1
Campbell argues that vortex wells and gates efficiently
absorb the kinetic energy of the liquid metal, and describes
how these gating features have been successfully imple-
mented in steel casting.
Several models have been developed to predict entraining
events in aluminum casting based on velocity vectors, free
surface normals, and liquid volume fractions in free surface
cells. As described earlier, oxide entrainment in aluminum
castings requires dry sides of oxide films to be in contact
with each other. For this to occur, the velocity directions of
the two overlapping flow streams must be opposite.1,8
Using this criterion, Yang et al.12 developed a model to
predict the entrainment of oxides in aluminum castings.
Applying this model, it was shown that a vortex runner is
beneficial in reducing the free surface turbulence inside the
runner, consequently providing smooth liquid metal flows
at ingates. The predicted results were validated through
comparison with experimental measurements. The effect of
runner design on the formation of oxide inclusions was
previously studied experimentally by Latona et al.13 Mea-
surements showed that a pressurized gating system with an
area ratio of 4:8:3 (where the lowest melt velocity occurs in
the runner) is beneficial in reducing dross inclusions in
ductile iron castings. The model developed by Reilly
et al.14 uses the velocity vectors, free surface normals, and
liquid volume fractions in free surface cells to define a
series of Boolean logic criteria for predicting the occur-
rence of entraining events. The model was applied to
several gating systems, and results showed the majority of
oxides are entrained at the surface of the pouring basin and
at the sprue base. Yue15 modeled three gating orientations:
direct pour, vertical bottom gate, and horizontal side gate.
Results showed that for a constant head height, the filling
system with the plunging jet (direct pour) entrained more
air compared to the two other filling systems. Castings
produced by the plunging jet flow type (direct pour) were
less reliable than castings produced by the rising jet flow
(vertical bottom gate) and returning wave flow (horizontal
side gate) as demonstrated by bending strength test results
for the three casting filling methods.
Few attempts have been made to simulate air entrainment
and compare predictions to water modeling experimental
results under varying pouring parameters and conditions.
Such results can be used to guide the best practices of
pouring liquid metals. Experimental water modeling stud-
ies have shown that shorter falling heights and filling times
reduce the volume of entrained air.16,17 Additionally, it has
been shown through water modeling that use of several
gating system features significantly reduces air entrain-
ment: an offset pouring basin with a step before the sprue
entrance, a nozzle extension submerged into the pouring
basin, and a whirl gate.18,19
A model for predicting the entrained air during mold filling
allows casting process engineers to evaluate pouring con-
ditions and gating systems. Based on the work by Ma
et al.,20 the authors have recently developed a model21 for
predicting the local air entrainment rate resulting from
disturbances at the free surface of the liquid metal flow.
This sub-grid air entrainment model was implemented in a
casting filling simulation software to calculate local air
entrainment rates during filling. While the authors’ previ-
ous work described the model’s development21 and vali-
dation by comparison with water modeling experiments,22
the goal of these developments is to apply the model to
liquid metals and predict the air entrainment during mold
filling. In the present study, the capabilities of the previ-
ously developed air entrainment model are demonstrated
by simulating liquid metal flow in typical casting filling
systems and calculating air entrainment rates. Results
presented here provide quantitative comparisons of the
effects of various pouring conditions on air entrainment,
and hence oxide inclusion formation. The metal casting
pouring conditions investigated here are process variables
256 International Journal of Metalcasting/Volume 13, Issue 2, 2019
which can be modified to minimize inclusion defects:
nozzle diameter, filling time, gating components, and gat-
ing ratios.
Air Entrainment Model
The model described here is implemented as part of a
standard casting filling simulation23 by performing calcu-
lations on flow velocities at the free surface to predict the
local air entrainment rate. The casting filling simulation
calculates the melt velocity and geometry of the free sur-
face by solving the Navier–Stokes equations with a volume
tracking method (VOF method) progressively through time
as the casting fills.
Disturbances exist at the surface of free surface flows, for
example on the periphery of the liquid jets (Figure 1).
These disturbances make the liquid–air interface rough,
and air pockets are trapped inside these disturbances as
indicated in Figure 1. For a plunging liquid jet, once the
liquid jet impinges on the surface of a quiescent pool, the
trapped air pockets are entrained into the bulk liquid where
they are broken into smaller air bubbles and carried away
with the liquid flow. In the model applied here, the local air
entrainment rate is a function of the turbulent kinetic
energy, k, and the normal derivative of the normal com-
ponent of the liquid velocity at the free surface interface,
oun=on20:
q ¼ Cent
k
g
oun
onEqn: 1
where q is the volumetric air entrainment rate per unit
interfacial area, Cent is an entrainment coefficient, and g
(m s-2) is the gravitational acceleration. For any free
surface flow, air is entrained only if the gradient term in
Eqn. 1 is positive, oun=on[ 0. The turbulent kinetic
energy in the model is estimated from the sum of the
squares of the fluctuating velocity components relative to a
spatially averaged mean velocity. Integrating the air
entrainment rate per unit interfacial area q (m s-1) over
the interfacial area As(m2), the volumetric air entrainment
rate Qa (m3 s-1) is calculated as
Qa ¼ZZ
As
qdA Eqn: 2
The entrainment coefficient used in the model was
Cent ¼ 0:039, and it was determined by calibrating the
predicted relative air entrainment rates to experimental
measurements reported in Reference 6 for plunging water
jets that have variable liquid jet velocities and diameters at
low turbulence intensities. A previously published paper21
explains the details of the model including the calculations
of the turbulent kinetic energy and normal derivative of the
normal component of the liquid velocity at the interface,
and the calibration of the entrainment coefficient. Using
this air entrainment coefficient value, good agreement
between measurements and predictions was achieved for
plunging water jets for a range of liquid jet velocities and
diameters. An example velocity field for a plunging/free
falling liquid jet is shown in Figure 2a with falling height,
hj, liquid jet velocity, uj, and jet diameters at the nozzle, dN,
and the impact location, dj. Air entrainment results for a
liquid jet with constant turbulence intensity are given in
Figure 1. Illustration of a plunging liquid jet depicting air bubbles entrained atthe impact location. Air pockets trapped inside disturbances along theperiphery of the free surface of the jet and important variables are shown.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 257
Figure 2b, c. The air entrainment is presented throughout
this paper using the relative air entrainment, defined as the
ratio of volumetric air entrainment rate, Qa, to the
volumetric flow rate of liquid, Qw. As shown in
Figure 2b, with good experimental and model agreement,
increasing the liquid jet velocity increases the relative air
entrainment rate, since increasing the liquid jet velocity
increases the turbulent kinetic energy of the liquid jet.
While as shown in Figure 2c, increasing the liquid jet
diameter reduces the relative air entrainment rate. The
explanation for the effect of jet diameter on air entrainment
requires more thought. Increasing the liquid jet diameter
increases the perimeter of the liquid jet, which increases
the volume of air pockets on the periphery of the liquid jet.
This effect increases the volumetric air entrainment rate in
proportion to the perimeter of the jet Qa / dj. However, the
relative air entrainment is the volumetric air entrainment
rate divided by the volumetric flow rate. Since the flow rate
for a circular jet that has a given velocity is proportional to
the inverse of the jet area, Qw / 1=d2j , the relative air
entrainment rate is therefore Qa=Qw / 1=dj. Therefore, fora given liquid jet velocity, increasing the jet diameter
reduces the relative air entrainment as shown for both
model and experimental results in Figure 2c.
Simulation of Air Entrainment During Mold Filling
The effect of pouring conditions and gating system designs
on the air entrainment during casting filling processes was
investigated using the model and simulation. For all cases
presented here, the cast part is represented by a rectangular
block of 406.4 mm (1600) length and width, and 304.8 mm
(1200) height. The filling simulations were performed using
a commercial casting simulation software,23 and the air
entrainment model was implemented based on the soft-
ware’s velocity and free surface geometry calculations. The
material properties for low-alloy steel and furan sand mold
were used from the software database. The pouring tem-
perature was 1600 �C (2912 �F) for all cases given here. A
uniform mesh with grid spacing of 4.5 mm was used in all
cases. Even though the simulations performed here are for
low-alloy steel and constant grid spacing, the relative
comparisons for air entrainment results between pouring
parameters and gating systems should not be affected by
the metal type and grid spacing. The readers are advised
that the values for relative air entrainment presented here
are for relative comparison. As part of ongoing work,
experiments are being developed using liquid metals to
provide air entrainment measurements to compare with
predicted model results.
In order to evaluate three distinct mold filling methods,
castings with direct pour (no gating), bottom gating, and
side gating were simulated. Geometries and dimensions of
the simulation cases without and with gating systems are
given in Figure 3. As shown in Figure 3a, for configura-
tions with no gating, the liquid steel is directly poured into
the mold cavity. Table 1 lists all cases studied here to
predict air entrainment in steel casting. In all of these
configurations, the distance between the ladle nozzle exit
Figure 2. (a) Velocity field for a free falling liquid jetshowing: falling height, hj, liquid jet velocity, uj, and jetdiameters at the nozzle, dN, and the impact location, dj,and relative volumetric air entrainment rate as a functionof (b) liquid jet velocity and (c) liquid jet diameter at theimpact location.
258 International Journal of Metalcasting/Volume 13, Issue 2, 2019
and the mold inlet was 50.8 mm (200). The effect of tur-
bulence intensity was neglected in these simulations.
The effects of nozzle diameter, dN, flow rate type (constant
vs variable flow rate), nozzle extension, and fill time (or
volumetric flow rate), tfill, were studied for the no gating
(direct pour) configurations. To study the effect of nozzle
diameter on the air entrainment, six nozzle diameters were
simulated, with smallest and largest nozzle diameters being
25.4 mm and 76.2 mm, respectively. In addition, for each
nozzle diameter, a constant and a variable flow rate was
examined. For configurations with variable flow rate, the
volumetric flow rate is not constant during mold filling.
The flow rate profile is kept the same for all nozzle
diameters. To calculate the variable flow rate profile, two
assumptions were made. First, it is assumed that the
maximum velocity of the liquid jet does not exceed
uj ¼ 6 m s-1. This velocity is reported as a critical tran-
sition velocity where the mechanism of air entrainment
changes and the disturbances on the periphery of the liquid
jet are not solely responsible for all of the air entrain-
ment.4–7 A maximum volumetric flow rate was selected
based on the maximum velocity for the smallest nozzle
diameter dN ¼ 25:4 mm. Second, an average filling time of
tfill ¼ 25:4 s is assumed corresponding to a ‘‘base’’ case
that has a constant volumetric flow rate of
Qs ¼ 2 9 10-3 m3 s-1. Moreover, the effect of nozzle
extension was studied for two nozzle diameters,
dN ¼ 47:6 mm and 63.5 mm. For configurations with the
nozzle extension, an extension with the same diameter as
the nozzle was added to the nozzle exit. Nozzle extension
with lengths of 203.2 mm (nozzle extended halfway) and
330.2 mm (nozzle extended all way down to the bottom of
the casting) were simulated. Additionally, four filling times
(or constant/average flow rates) were studied to examine
the effect of the filling time on the air entrainment.
The effects of pouring cup, pouring basin, sprue, and well
designs on air entrainment were investigated. Results for
simulated cases presented here include: conical pouring
cup, offset pouring basins with and without stopper, use of
vortex sprue, and use of vortex well. A vortex gate was
simulated with the bottom-gated configuration, and for the
side-gated configuration, several gating ratios’ were
examined. The geometries and dimensions of the pouring
cup/offset basin and vortex components are shown in
Figures 4 and 5, respectively. In all configurations with
gating systems, the nozzle diameter and volumetric flow
rate of liquid steel are constant with dN ¼ 63:5 mm and
Qs ¼ 2 9 10-3 m3 s-1, respectively, and the pouring cup/
basin height is hcup ¼ 152:4 mm. The slight variation in
filling times, indicated in Table 1, is due to the difference
in gating system volume.
A single case was simulated to study the effect of stopper
in offset pouring basins. The stopper is placed at the
sprue entrance, sealing the entrance of the sprue. The
liquid is poured from the ladle to the basin, and the
stopper is removed when the basin is filled to a certain
level. Once the stopper is removed, the liquid fills the
sprue and the gating system. After removing the stopper,
liquid metal is continuously provided from the ladle to
maintain a constant falling height from the ladle lip to
impact location inside basin. Since moving objects can-
not be defined in the software used, a porous filter with
large pressure loss coefficient was defined at the sprue
Figure 3. Casting geometry and gating systems used inair entrainment simulations: (a) direct pour (no gatingsystem), (b) bottom-gated system, and (c) side-gatedsystem. Dimensions are in mm.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 259
entrance to apply the ‘‘stopper’’ effect on the offset basin,
and a script file was used to define the time and condi-
tions for removing the stopper. While the flow is blocked
at the filter location, liquid steel is allowed to fill the
basin for a time of 2 s. Then, the stopper is removed
accordingly, and liquid steel is allowed to fill the gating
system and casting while continue filling the basin. For
all the bottom-gated configurations, the sprue and runner
diameters are ds ¼ dr ¼ 76:2 mm, and the ingate diame-
ter is din ¼ 177:8 mm. Details of the gating system
components for the side-gated filling systems are shown
in Table 2.
For all the cases, instead of modeling the ladle, the volu-
metric flow rate of liquid steel was used as the input for the
simulations.
Results and Discussion
For each simulation case, results for the velocity field, the
relative air entrainment rate variation over time during the
pouring event, and the final (total) relative air entrainment
volume are presented. To calculate the final relative
entrained air volume, Va=Vs, first, the volume of entrained
Table 1. Summary of Air Entrainment Simulation Cases Presented
Trial Flow ratetype
Nozzleextension
Nozzle diameter,dN (mm)
Pouring cup/basin Vortex component Gating ratio,Abs:Ar:Ain
Fillingtime, t (s)
No gating—direct pour
1 Constant No 25.4 – – – 25.4
2 Constant Halfway 25.4 – – – 25.4
3 Constant All way 25.4 – – – 25.4
4 Constant No 31.75 – – – 25.4
5 Constant No 38.1 – – – 25.4
6 Constant No 50.8 – – – 25.4
7 Constant No 63.5 – – – 12.6
8 Constant No 63.5 – – – 16.8
9 Constant No 63.5 – – – 25.4
10 Constant Halfway 63.5 – – – 25.4
11 Constant All way 63.5 – – – 25.4
12 Constant No 63.5 – – – 50.5
13 Constant No 76.2 – – – 25.4
14 Variable No 25.4 – – – 25.4
15 Variable No 31.75 – – – 25.4
16 Variable No 38.1 – – – 25.4
17 Variable No 50.8 – – – 25.4
18 Variable No 63.5 – – – 25.4
19 Variable No 76.2 – – – 25.4
Bottom gated
20 Constant – 63.5 Cone cup None 1:1:5.4 30.6
21 Constant – 63.5 Offset basin None 1:1:5.4 31.3
22 Constant – 63.5 Offset basin with stopper None 1:1:5.4 31.4
23 Constant – 63.5 Offset basin Vortex sprue 1:1:5.4 31.2
24 Constant – 63.5 Cone cup Vortex well 1:1:5.4 32.3
25 Constant – 63.5 Cone cup Vortex well and gate 1:1:5.4 34.5
Side gated
26 Constant – 63.5 Cone cup None 1:2:2 30.3
27 Constant – 63.5 Cone cup None 4:8:3 30.1
28 Constant – 63.5 Cone cup None 1:1:1 29.1
29 Constant – 63.5 Cone cup None 1:1:1.7 28.1
Cases are organized by gating system and conditions used
260 International Journal of Metalcasting/Volume 13, Issue 2, 2019
air, Va, is determined by integrating the volumetric air
entrainment rate over time, and then, this value is divided
by the volume of liquid steel poured, Vs. In the results that
follow, first simulation results for filling the casting without
a gating system (direct pour) are given, and the effects of
nozzle diameter, flow rate constancy, nozzle extension, and
filling time are shown. In the second results subsection, the
effects of gating system components on air entrainment are
presented for castings with bottom-gated filling system and
compared with the direct pour case. Finally, the results for
the side gating configuration are given where several gating
ratios are compared with each other and with the direct
poured case.
Direct Pour: No Gating
The velocity and local air entrainment contours are shown
in Figure 6a at t ¼ 8 s for the ‘‘base’’ case using direct
pour (no gating), and nozzle diameter and fill time of
dN ¼ 63:5 mm and tfill ¼ 25:4 s, respectively. Air bubbles
are entrained at the periphery of the liquid jet where it
impinges on the surface of the liquid pool (Figure 6b). In
Figure 6c, the total relative air entrainment (local air
entrainment summed over the system volume) over time is
shown. The largest spike in the relative air entrainment plot
corresponds to the liquid jet initial impingement on the
pool surface. After the initial impact, starting from around
3 s into the filling, the relative air entrainment decreases
significantly to a smaller value around 0.1 and it reduces
gradually until the end of filling.
The effect of nozzle diameter on the air entrainment is
shown in Figure 7. Using a constant filling time of
tfill ¼ 25:4 s, six nozzle diameters were simulated. No
nozzle extension was used for this case. For a constant
filling time (constant volumetric liquid flow rate),
increasing the nozzle diameter reduces the liquid jet
velocity at the nozzle exit (Figure 7b), and the liquid jet
diameter at impact, and hence the air entrainment.
Increasing the nozzle diameter by a factor of three reduces
the liquid jet velocity at nozzle exit uN by a factor of more
than 6 (Figure 7b), while looking at the velocity contours
(Figure 7a), the liquid jet at impact, uj, reduces by a factor
of approximately 2. In addition, though the liquid steel
pouring stream contracts more for cases with larger nozzle
diameters, the liquid jet diameter at impact is still larger for
large nozzle diameters (Figure 7a). The reduction in liquid
jet velocity at impact, uj, along with the increase in liquid
jet diameter at impact, dj, significantly reduces the relative
air entrainment rate (Figure 2). As shown in Figure 7c, the
initial spike is largest for the smallest nozzle diameter
configuration dN ¼ 25:4 mm, and this case entrains sig-
nificantly more air than other configurations throughout the
pouring event. From Figure 7d, it is shown that for a
constant filling time tfill ¼ 25:4 s increasing the nozzle
diameter by a factor of three reduces the relative entrained
air volume approximately 5 times, and more modest
increases in dN also result in substantially less air
entrainment.
In bottom pour ladles, as the ladle is emptied, the liquid jet
velocity at nozzle exit reduces with time; hence, the liquid
jet velocity at impact reduces as filling proceeds. To study
the effect of this variable flow rate on air entrainment,
constant and variable flow rate configurations were simu-
lated for each nozzle diameter. In Figure 8, the relative air
Figure 4. Geometries and dimensions (in mm) for (a) conical pouring cup and(b) offset pouring basin used in air entrainment simulations.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 261
entrainment volume is compared for constant and variable
flow rates for the nozzle diameters used in Figure 7 where
the filling time of tfill ¼ 25:4 s. The volumetric steel flow
rate used for all nozzle diameters and velocity profiles for
three of the nozzle diameters are shown in Figure 8a, b,
respectively. Note that since the filling time for variable
Figure 5. Geometries and dimensions (in mm) of vortex generating gating compo-nents: (a) offset basin for creating vortex sprue, (b) vortex gate, and (c) vortex well.
Table 2. Overview of the Sprue, Runner, and Ingate Geometry for Side-Gated Filling System Simulation Cases Pre-sented in Paper
Gating ratio,Abs:Ar:Ain
Sprue base diameter, dbs(mm)
Runner width, wr
(mm)Runner height, hr(mm)
Ingate width, win
(mm)Ingate height, hin(mm)
1:2:2 63.5 114.3 55.4 127.0 49.9
4:8:3 63.5 114.3 55.4 93.5 25.4
1:1:1 63.5 88.9 35.6 127.0 24.9
1:1:1.7 47.0 76.2 22.8 127.0 22.8
All cases used a nozzle diameter and volumetric flow rate of dN ¼ 63:5 mm and Qs ¼ 2� 10�3 m3 s-1, respectively
262 International Journal of Metalcasting/Volume 13, Issue 2, 2019
flow rate is the same as the constant flow rate configuration
(tfill ¼ 25:4 s), the velocity plots intersect halfway during
filling (tfill ¼ 12:7 s) in Figure 8b. The velocity (left) and
relative air entrainment rate development during filling
(right) are compared for the smallest nozzle diameter
configuration in Figure 8c. The velocity contours are
shown at two times, at 4 and 20 s into the filling. The larger
initial velocity at 4 s in Figure 8c is clear for the variable
flow rate case. For the constant volumetric flow rate, the
liquid velocity at nozzle exit remains constant during fill-
ing, while for the variable case, the velocity decreases as
filling proceeds. From Figure 8c, initially the relative air
entrainment rate is significantly larger for the variable flow
rate. At the beginning due to high liquid jet velocity for the
variable flow rate, significant free surface turbulence
occurs once the liquid steel impinges on the bottom of the
mold box. Additionally, the breaking waves resulted from
splashing, contributing to more air entrainment. However,
after tfill ¼ 12:7 s, as the liquid velocity at the nozzle exit
reduces below uN ¼ 3:95 m s-1 (liquid velocity at the
nozzle exit for the constant pouring rate), the air entrain-
ment drops below the air entrainment rate of the constant
pouring rate case. Comparison between the constant and
variable flow rates (Figure 8d) indicates that increasing the
nozzle diameter reduces the difference of the relative
entrained air volume between constant and variable flow
rates. For smaller nozzle diameters, air entrainment is
significantly larger due to the much larger velocity at the
beginning of filling. However, as the nozzle diameter is
increased, the velocity difference between the constant and
variable flow rate configurations is less (Figure 8b), and the
air entrainment difference between the constant and vari-
able flow rates becomes negligible (Figure 8d).
The effect of using a nozzle extension (or shroud) on air
entrainment is shown in Figure 9 for two nozzle diameters
Figure 6. Plots of (a) velocity field and (b) local volumetric air entrainment rate, Qa,at t ¼ 8 s, and (c) total relative air entrainment rate as a function of time for thesimulation case with no gating (direct pour), nozzle diameter of dN ¼ 63:5 mm, andfilling time of tfill ¼ 25:4 s (Qs ¼ 2� 10�3 m3 s-1) with no nozzle extension.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 263
and a constant filling time of tfill ¼ 25:4 s. The addition of
a nozzle extension applies friction to the liquid flow, which
reduces the liquid steel velocity exiting the nozzle exten-
sion, and therefore reduces air entrainment. Also, once the
nozzle extension becomes submerged inside the liquid steel
pool, the relative air entrainment rate drops to a small value
as shown in Figure 9b. For a submerged nozzle extension,
the nozzle extension exit is located at the impact location,
and the liquid steel has virtually no interaction with the
surrounding air and no air pocket is trapped in the
periphery of the liquid jet. Therefore, once the nozzle
extension becomes submerged in the liquid, air entrainment
decreases significantly.
The effect of fill time (or alternatively the flow rate) on the
air entrainment, and oxide inclusion formation, has been
debated among foundry engineers. The effect of fill time on
the air entrainment is demonstrated by the results in Fig-
ure 10 for the no nozzle extension and nozzle diameter of
dN ¼ 25:4 mm case. Longer filling time implies longer
interaction of liquid metal with the air, which implies that
more air entrainment occurs. For configurations with short
filling time, the liquid steel fills the mold cavity fast, and
this results in the free falling height (distance from the
nozzle exit to the impact location) and liquid jet velocity at
impact reducing over a short period of time as shown in
Figure 10a for the four cases at 8 s into the filling process.
This is the primary reason that less air is entrained for
faster fill times. As indicated in Figure 10b, the relative air
entrainment drops drastically for the shortest filling time
(high volumetric flow rate), while the reduction in air
entrainment is a slower process for the longest filling time
case. Clearly, for a pouring process using a constant head
height and nozzle diameter, reducing the filling time
reduces the air entrainment.
Bottom-Gated Castings
The effect of using various gating system components on
air entrainment in filling bottom-gated castings was simu-
lated. The effect of pouring cup/basin design and the use of
a stopper for the offset basin on the air entrainment are
Figure 7. Plots showing the effect of the nozzle diameter on variables and the relative air entrainment rate for theno gating case and a constant fill time of tfill ¼ 25:4 s (Qs ¼ 2� 10�3 m3 s-1): (a) the velocity contours at t ¼ 8 s forfour of the nozzles, (b) liquid velocities at the nozzle exit, (c) relative volumetric air entrainment rates for the nozzlesin (a), and (d) total relative entrained air volume at the end of filling.
264 International Journal of Metalcasting/Volume 13, Issue 2, 2019
shown in Figure 11. In Figure 11a, the velocity contours at
8 s from start of filling are shown for the five cases con-
sidered using a nozzle diameter of dN ¼ 63:5 mm and a
flow rate of Qs ¼ 2� 10�3 m3 s-1. The final air
entrainment results show that the casting with direct pour
(no gating) entrains the least amount of air (Figure 11c).
This might be surprising at first glance. However, it must
be pointed out that the main reason for this outcome is the
Figure 8. Plots showing the effect of variable and constant flow rates on the relative air entrainmentrates and key variables for the no gating case using a fill time of tfill ¼ 25:4 s (Qs ¼ 2� 10�3 m3 s-1):(a) variable and constant volumetric flow rates used as a function of time for all nozzle diameters, (b)nozzle velocities for variable and constant flow rates for three of the nozzle diameters simulatedversus time, and (c) the velocity fields at times t ¼ 4 s and t ¼ 20 s during filling, and the relativevolumetric air entrainment rates for the nozzle diameter dN ¼ 25:4 mm case, and (d) the final relativeentrained air volume for cases having with and constant flow rates for all nozzle diameterssimulated.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 265
difference in falling heights between the direct pour con-
figuration and the bottom-gated configurations (see Fig-
ure 3). The initial falling height for the direct pour case is
355.6 mm (Figure 3a), compared to 711.2 mm for bottom-
gated case (Figure 3b), and this has resulted in lower air
entrainment for the direct pour configuration. As men-
tioned earlier, reducing the falling height reduces the liquid
jet velocity at impact, and therefore the air entrainment.
The velocity and air entrainment rates for the conical cup
and offset basin are also compared in Figure 11. The gating
system with conical cup entrains more air than the one with
offset basin. Early during filling the offset basin
configuration entrains more air, but the conical pouring cup
case entrains larger amounts throughout most of the later
pouring process. In an offset basin, the liquid steel first
impinges on the bottom of the basin surface, and later it
plunges to the base of the sprue. For the pouring basin, the
flow velocity at the sprue entrance is greatly reduced from
the conical pouring cup case due to the presence of the
cylindrical step feature (or wier), which results in a smaller
liquid velocity at the sprue base. For a conical cup, the
liquid steel exiting the ladle nozzle directly impinges on
the sprue base and results in significantly larger liquid jet
velocities at the sprue base. Also for the offset basin con-
figuration, once the liquid steel passes the circular step/weir
Figure 9. Plots showing the effect of nozzle extension and nozzle diameter on thevelocity field and relative air entrainment for the case with no gating system withnozzle diameters of 25.4 and 63.5 mm and a fill time of tfill ¼ 25:4 s(Qs ¼ 2� 10�3 m3 s-1): (a) the velocity fields at t ¼ 8 s, (b) relative volumetric airentrainment rate during pouring, and (c) final relative entrained air volume for thetwo nozzle sizes and three extension conditions.
266 International Journal of Metalcasting/Volume 13, Issue 2, 2019
of the offset basin occurs inside the sprue (large spike in
Figure 11b). However, after approximately t ¼ 4 s, air
entrainment significantly reduces to a small value and
continues to drop as the filling proceeds. Overall, the offset
basin shows improvement over the conical cup. In addition,
the effect of stopper for an offset basin was examined.
Once the stopper is removed, the liquid steel fills the sprue
in a short period of time, reducing the interaction with
surrounding air; therefore, the stopper reduces the effect of
initial liquid metal impingement inside the sprue. The use
of stopper slightly reduces the air entrainment as seen by
the results in Figure 11c.
Another case of interest shown in Figure 11 is the effect on
air entrainment for the offset basin with vortex sprue gating
system. Creation of a vortex flow inside the sprue requires
careful design of the offset basin. The design shown in
Figure 5a is patterned after Reference 1. The flow from the
ladle impinges on the basin surface, and as it enters the
offset part of the basin, it follows the circular path of the
design (see rightmost image in Figure 5a). The offset part
of the basin is designed to generate a tangential flow to the
sprue wall (see Figure 12a top view). The liquid steel
swirls down the sprue and does not directly impinge on the
bottom of sprue as shown in Figure 12a. In this flow pro-
cess, the sprue walls generate additional frictional losses in
the liquid steel flow and markedly reduce the velocity.
Note that this vortex flow reduces the initial spike of the air
entrainment to a great extent for the offset basin. These
results appear to contradict an often held belief among
foundry engineers that such vortex flows draw air into the
gating system and increase oxide inclusions.1
The relative effectiveness of the vortex well and vortex
gate to reduce air entrainment is given by the results in
Figure 13. In the case of a vortex well, after the liquid steel
impinges on the well surface and fills the sprue well, the
free surface turbulence at the sprue base significantly
reduces. Comparison between the velocity contours in
Figure 13a indicates that for configurations with the vortex
well, most of the sprue fills up early during the filling,
which reduces the falling height from the ladle exit into the
sprue, and consequently reduces air entrainment. The aim
of the vortex gate shown in Figure 13 is to reduce the
liquid metal velocity at the ingate and to reduce the free
surface turbulence at the entrance to casting cavity, and
Figure 10. Plots showing the effect of filling time (pouring rate) on the velocity field and relative airentrainment for the no gating system case with nozzle diameter dN ¼ 63:5 mm and filling times of12.7, 16.9, 25.4 and 50.8 s: (a) the velocity fields at t ¼ 8 s from the start of filling, (b) relativevolumetric air entrainment rate during filling, and (c) final relative entrained air volume for each filltime.
International Journal of Metalcasting/Volume 13, Issue 2, 2019 267
therefore air entrainment. In a vortex gate, the flow from
the runner tangentially enters the gate, and swirls inside the
gate and fills the gate and the casting quiescently. The
combination of vortex well and vortex gate is shown to
reduce the air entrainment to almost half of the conical
configuration in Figure 13c. The simulation results agree
with Campbell’s findings; however, as stated by Campbell,
further research is required to confirm the advantages of
these vortex generating gating components.1
Side-Gated Castings
Non-pressurized and pressurized gating systems have long
been used for eliminating or reducing oxide inclusions in
metal castings. In pressurized gating systems, the minimum
cross-sectional area (choke) of the gating system is located
at the ingate. Therefore, these gating systems are beneficial
in filling the gating system fast, hence reducing the liquid
metal interaction with the surrounding air. Unfortunately,
the ingate velocity can be small in these filling systems,
which increases the possibility of free surface turbulence at
the casting entrance. Conversely for a non-pressurized
gating system, the choke for the system is located at the
sprue base. The liquid metal velocity in non-pressurized
gating systems is usually low, and the liquid metal has
more time to interact with air. A newer gating system
design approach proposed by Campbell has been widely
accepted in metal castings, the ‘‘Naturally Pressurized’’
gating system. In this gating system, local chokes are
avoided and the liquid metal is pressurized throughout the
entire filling system. As a result, the cross-sectional areas
of the filling system components are small enough such
that surface turbulence is avoided to a great extent.
Air entrainment results for side-gated casting filling sys-
tems with several gating ratios are shown in Figure 14. The
gating ratios produce non-pressurized, pressurized, and
naturally pressurized filling systems, and their results are
compared to the direct pour no gating configuration. Sim-
ilar to bottom-gated castings, air entrainment comparisons
are made for a nozzle diameter and volumetric flow rate of
dN ¼ 63:5 mm and Qs ¼ 2 9 10-3 m3 s-1, respectively.
Figure 11. Plots showing the effect of the pouring cup, basin, vortex sprue and stopper conditionson the velocity field and relative air entrainment rate for nozzle diameter dN ¼ 63:5 mm and flow rateof Qs ¼ 2� 10�3 m3 s-1: (a) the velocity fields at t ¼ 8 s from start of filling, (b) relative volumetric airentrainment rate during filling, and (c) final relative entrained air volume for the five simulationcases.
268 International Journal of Metalcasting/Volume 13, Issue 2, 2019
In Figure 14a, filling velocity fields are shown at t ¼ 8 s
from the start of pouring. Note that at this time the sprue is
almost filled with liquid steel for the naturally pressurized
gating system configuration using the small sprue, while
for the non-pressurized gating system, the liquid level is
close to the sprue base. As a result, the liquid jet velocity at
impact, and the air entrainment, is larger for non-pressur-
ized gating system. The results of relative air entrainment
in Figure 14b indicate the pressurized and naturally pres-
surized gating systems show a secondary peak. This sec-
ondary peak corresponds to the time when the liquid jet
enters the casting. A smaller ingate produces a larger liquid
velocity entering the casting, and a significant amount of
air is entrained when the liquid steel emanates from the
smaller ingates. However, after the liquid level inside the
casting reaches the top of the ingate, at approximately
t ¼ 4 s, the air entrainment decreases drastically. As filling
proceeds, air entrainment continues dropping until it
reaches zero at the end of filling. In summary, the faster the
liquid steel fills the gating, the less interaction occurs
between the liquid metal and surrounding air, resulting in
less air entrainment. Clearly, these results show that natu-
rally pressurized gating systems are beneficial in reducing
the total amount of entrained air. Considering the gating
dimensions in Figure 3, it is important for the reader to
appreciate that even though the initial free falling height of
the naturally pressurized gating system with small sprue
(h ¼ 584:2 mm) is significantly larger than that of the
direct pour case (h ¼ 355:6 mm), the gating with the
greater falling height entrains less air.
Conclusion
Casting manufacturing processes have many variables
which can be difficult to control and can lead to defects in
cast components. If foundry engineers can design reliable
processes for filling castings, they can reduce air entrain-
ment and the oxide inclusion defects it causes, and elimi-
nate a major source of quality variability and rework. Here
air entrainment modeling is applied to demonstrate the
effect of pouring conditions and gating system design on
the air entrainment during mold filling. The results
demonstrate that quantitative comparisons between filling
methods, based on physical mechanisms, are possible. As
mentioned several times here, there are many beliefs,
opinions, and anecdotal experiences in the foundry industry
surrounding the best ways to fill castings, and now these
can be quantitatively tested.
Modeling results presented here show that reducing the
total head height and adding a nozzle extension to the end
of the nozzle reduces the air entrainment during mold
filling with a bottom pour ladle. In addition, for a given
Figure 12. Velocity field results for the simulation configuration using the vortexsprue and bottom filled gating system with nozzle diameter dN ¼ 63:5 mm and flowrate of Qs ¼ 2� 10�3 m3 s-1: (a) two views of the flow field development inside theoffset basin at early fill times, and (b) views of the velocity flow field at 8 s from thestart of filling for the entire model domain (3D view) and on the mid-plane of themodel (side view).
International Journal of Metalcasting/Volume 13, Issue 2, 2019 269
filling time, increasing the nozzle diameter reduces the
liquid jet velocity, which significantly reduces the
entrained air volume. Using an offset basin with a stopper
and gating components which create vortex flows tangen-
tial to gating system walls markedly reduce the entrained
air volume. Results also show that producing clean castings
requires pressurizing the gating system. This study
produced promising results and demonstrates the possibil-
ities of future applications of the present air entrainment
model. With further experimental validation, the air
entrainment model will be a powerful tool for the evalua-
tion of filling and gating systems. Ongoing and future
model development work will link the present air entrain-
ment model to an inclusion generation and transport model,
Figure 13. Plots showing the effect of using vortex components in gating systems on the velocityfield and relative air entrainment rate for nozzle diameter dN ¼ 63:5 mm and flow rate ofQs ¼ 2� 10�3 m3 s-1: (a) the velocity contours for the five simulation cases at 8 s from the start offilling, (b) relative volumetric air entrainment rate for the cases during filling, and (c) the final relativeentrained air volume at the end of filling.
270 International Journal of Metalcasting/Volume 13, Issue 2, 2019
where the final oxide inclusion size and location can be
predicted. Real-world casting trials should be conducted to
further validate the model predictions.
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