Effect of Pouring Conditions and Gating System Design on

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

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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.


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.


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.


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



29 Constant – 63.5 Cone cup None 1:1:1.7 28.1

Cases are organized by gating system and conditions used


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.


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


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.


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.


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.


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.


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.


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.


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).


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.


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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Effect of Pouring Conditions and Gating System Design on