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Investigation of Cast Iron Processing to Produce Controlled Dual Graphite Structure in Castings
S.N. Lekakh, J. Qing, V.L. Richards
Missouri University of Science and Technology, Rolla, MO
Copyright 2012 American Foundry Society
ABSTRACT
The change in graphite shape from flake to spheroidal
significantly increases the strength of cast iron while
simultaneously decreases thermal conductivity to about
half of the value in flake graphite iron. In many industrial
applications, such as the cylinder head of diesel engine, a
combination of high strength and thermal conductivity is
essential. A compromise would be to use compacted
graphite iron, while a more effective way would be to
develop controlled distribution of the dual graphite
structure that matches a specific region’s function of the
casting design. In this study, the process of casting such
functionally region-specific material was modeled and
experimentally investigated.
It was shown that four main conditions need to be
satisfied for achieving a controlled dual graphite
structure:
a distinct change in ductile/gray iron melt
composition during mold pouring by melt treatment;
sequential fill of the mold cavity by different melt
compositions using special gating systems;
minimum post filling momentum in the mold and
avoidance of natural convection in the melt before
casting solidification.
Different approaches were tested using an experimental
casting with flat internal cores. The experimental results
achieved thus far were compared to modeling predictions.
The different possible process routes are discussed.
Keywords: ductile iron, gray iron, solidification, dual
graphite structure, modeling.
INTRODUCTION
The shape of graphite in cast irons significantly affects
mechanical and thermo-physical properties of castings.
Table 1 provides the typical mechanical properties of cast
irons with ferritic and pearlitic metal matrix.1, 2
Changing
from the flake-shaped graphite in gray iron (GI) to
spheroidal graphite (SG) in ductile iron (DI) increases
tensile strength by nearly a factor of three without
affecting hardness and provides a higher value of “quality
factor,” which is commonly defined as the ratio of
UTS/HB.
Changing the graphite morphology from flake to
spheroidal also reduces the thermal conductivity at the
same time as strength is increased. In many industrial
applications, such as the cylinder head of a diesel engine,
a combination of high strength and thermal conductivity
is essential. A compromise could be achieved by using
compacted graphite iron. However, for a specific design
application such as the diesel engine head there are
regions of the casting, whose function requires high
strength, and other regions, whose function is primarily
heat conduction between the combustion zone and the
cooling system. Typically, the required metal matrix
consists of pearlite with minimum ferrite and absence of
free carbides.
In many unintentional experiments, different shapes of
graphite have been observed in different parts of the same
casting, while the achievement of controlled,
appropriately located and stable dual graphite structure is
a serious challenge. Stefanescu and coauthors3
investigated the surface layer, so called casting skin, in DI
and compacted graphite (CG) iron. Some metallographic
features were identified and their formation mechanisms
were proposed: graphite degradation due to the fade of
magnesium and graphite depletion due to decarburization
near surface by reaction with oxygen diffused from mold.
Therefore, spheroidal graphite degenerates to CG; CG
degenerates to flake graphite; type A flake graphite
degenerates to type D. For example, flake graphite—fine
flake and compacted to coarse compacted— coarse
compacted and some spheroidal and exploded graphite
layers were identified in the skin. These skin layers
diminish the mechanical properties of the CG iron
casting. These layers are undesirable in conventional
castings while they might be desirable for castings with a
controlled dual graphite structure.
Many researchers have proven the role of sulfur as a
modifier of the graphite morphology. It has been proven
that sulfur is more desirable as denodulizing element than
titanium. Moreover, late sulfur addition favors the
graphite nucleation in ductile iron. Lekakh and Loper8
demonstrated a significant increase in nodule count after
late S-O-additions in magnesium treatment melt. Riposan4
demonstrated that larger sulfur additions facilitated the
transition: SG→ CG → coral graphite→ type B
graphite→ type A graphite, in this order as the sulfur
content increased. Denodulizing magnesium (Mg) treated
Paper 12-024.pdf, Page 1 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
Table 1. Typical Mechanical Properties of Gray (GI), Compacted Graphite (CGI) and Ductile Iron (DI) with Spheroidal Graphite
1, 2
Material Matrix UTS, MPa
Hardness, HB
Elongation, %
Quality factor (YST/HB)
Heat conductivity,
W/mK
GI Pearlitic 220-270 190-230 Less than 1 1.0-1.2 45-55
CGI Ferritic Pearlitic
330-410 420-580
130-190 200-250
5-10 2-5
2.3-2.5 2.0-2.3
35-45
DI Ferritic Pearlitic
420-600 600-800
140-200 240-300
15-25 3-10
2.7-3.0 2.5-2.7
30-35
iron with sulfur additives has been adopted to produce CG
iron of various SG/CG values by adjusting Mg/S ratio. All
studies mentioned previously were done for casting with
mono-shape graphite. However, the achievement of
controlled dual graphite structure in one casting is a
significant challenge.
Melt mixing during mold pouring; natural convection
after the fill and other possible physical phenomena,
which occur during casting filling and solidification
(shrinkage, elemental segregation), could prevent the
formation of the desired dual structure. One of most
important factors preventing the formation of dual
graphite structure is natural convection in the melt, which
results from changes in melt density during cooling and a
significant thermal gradient in the filled casting. In fluid
mechanics, the Rayleigh number (Ra) includes the main
parameters influencing natural convection:
( ) Equation 1
where: Gr is Grashof number which describes the
relationship between buoyancy and viscosity, Pr is
Prandtl number, which describes relationship between
momentum and thermal diffusivities, g is the acceleration
due to gravity, β is thermal expansion coefficient, ϑ is
kinematic viscosity, α is thermal diffusivity, ∆T is
temperature difference and X is characteristic length.
When the Ra number is below a critical value, heat
transfer is primarily controlled by conduction, while heat
transfer is primarily due to convection at higher Ra
values. Factors favoring increased convection are high
thermal expansion, β, and large casting dimensions, X.
The thermal differential, ΔT, is the main driving factor for
convection rate while the total mass transferred by
convection also depends on processing time. High cooling
rate, T,’ increases thermal differential, ΔT, and
simultaneously decreases duration for cooling the poured
melt to near the solidification temperature at which a
dramatic increase of viscosity, ϑ, decreases Ra value.
Therefore, the real situation is very complicated and is an
appropriate application for Computational Fluid
Dynamics (CFD) modeling.
The mixed effect of the natural convection flow and
residual flow from fluid momentum after the completion
of mold filling were numerically investigated5 and the
results clearly showed the necessity to carry out a coupled
filling and solidification analysis. Mampaey6 studied the
influence of mold filling and natural convection on cast
iron solidification using a combination of experimental
tracer technique and CFD modeling. These experiments
indicated that solute as well as precipitated graphite
nodules could be transported by the convection flow in
casting.
The objective of the study reported in this paper was
evaluation of the different possible processing routes for
production of iron castings with a controlled dual graphite
structure. The experimental methods were combined with
CFD modeling with the goal to control the location of
each of the desired structures.
PROCEDURES EXPERIMENTAL The four experimental heats, designated as A, B, C, and
D, were prepared using base pearlitic ductile iron melted
in 200-lb induction furnace from charge materials, which
included industrial ductile iron returns (Table 2). The melt
was heated to 1500C (2732F) and treated in a pocket ladle
by Fe-Si-Mg followed by FeSi base inoculants.
Chemistries of these heats in the ladle after treatment are
given in Table 3.
Table 2. Charge for Melting Ductile Iron
In charge Weight (lbs)
Plain-C steel disks 50
DI return 80
Induction Iron 20
Desulco 9001 3
Cu 0.8
Total induction furnace charge 153.8
Ladle treatment
FSM (46% Si, 5.7% Mg, 1% Ca, 0.4% La, 1% max Al,), 1-10 mm 2.4
Inoculants (75% Si, 4% Al, 1% Ca), 1-0.2 mm 0.8
Part of the magnesium-treated ductile iron melt was
substantially treated with sulfur for the development of a
flake graphite structure. Powder pyrite (FeS2) of size
Paper 12-024.pdf, Page 2 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
Table 3. Chemistry (wt. %) of Experimental DI Metals in Ladle
Heat C Si Mn Cu S P Ni Mg
A 3.62 2.32 0.34 0.63 0.003 0.010 0.08 0.030
B 3.72 2.44 0.40 0.65 0.007 0.014 0.05 0.035
C 3.74 2.60 0.42 0.64 0.003 0.015 0.04 0.043
D 3.66 2.50 0.33 0.64 0.004 0.010 0.04 0.047
Fraction—100 mesh was used for sulfur post—
nodulization-treatment. A list of different experimental
techniques used in these heats is summarized in Table 4,
the reaction tundish design for Heat B, the casting designs
for Heat C with bottom gating system and the casting
design for Heat D with side gating system were illustrated
in Figs.1, 2 and. 3, respectively. The objective and
description of each heat will be discussed in detail in the
following portions of this paper.
Table 4. Experimental Heats Description
Heat Melt treatment
Mold, Casting Gating system
A In ladle 5 stationary molds with 1in. vertical plates
-
B In stream One stationary reaction tundish and 5 moving molds
with 1in. vertical plates
-
C
In stream
Mold with two 1in. horizontal plates and one 11/4in.
internal flat core
Bottom
D In ladle (double pour)
Mold with three 1in. horizontal plates and two 11/4in. internal flat core
Side
a)
b)
Fig. 1. This is the experimental setup of Heat B, used in-stream FeS2 feeding into stationary tundish with a bottom hole (a) and moving molds (b) for collection of treated iron.
Fig. 2. A casting design with a gating system having multiple bottom ingates (Heat C) is illustrated.
Fig. 3. A casting design with a gating system having two side ingates (Heat D) is illustrated.
Thermal analysis of ductile iron treated in the ladle
employed the adaptive thermal analysis system (ATAS).
The thermal cooling curves from castings were obtained
with K-type thermocouples protected by quartz tubes and
connected to a 24-bit National Instrument DAQ. No-bake
molds were placed on an electronic platform scale to
monitor pouring weight on a continuous basis. All heats
were also recorded as video. Chemistry of chilled buttons
was characterized using an arc spectrometer and a Leco
C-S determinator. Castings were cut and microstructure
was evaluated at different locations. Non-metallic
inclusion analysis was done using an Aspex automated
Scanning Electron Microscopy/Energy Dispersive X-ray
(SEM/EDX system.
COMPUTATIONAL Fluent CFD software was used for analysis of the effects
of mold pouring, post-filling momentum, melt natural
convection and casting solidification on sulfur and
magnesium distribution in the casting. The experimental
castings consist of two or three in. thick 10 in.x10 in.
plates with one or two 1-1/4in. thick 8 in.x8 in. internal
flat cores. Taking into consideration the small thickness to
length ratio (1:10), a 2-d computational domain was used
100
50
200
160
50
Ø18
Paper 12-024.pdf, Page 3 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
to simplify the model and economize on computational
requirements. The transient, pressure-based solver with
absolute velocity formulation was used. Heat transfer,
volume of fluid (VOF) with turbulent flow modules were
coupled for modeling of mold pouring. Heat transfer,
laminar flow and solidification modules were coupled for
solving melt cooling with casting solidification. An
additional species transportation module with two
volumetric species including cast iron melt and melt with
magnesium in solution was also coupled for both, mold
filling and casting cooling models. This module was used
for tracing convection and diffusion transport of
magnesium in the melt during processing.
A piece-wise-linear approximation of cast iron density
was used for considering shrinkage in liquid and solid
conditions, as well as the volume increase during graphite
eutectic solidification.7 A simultaneous assumption of
rapid increase in melt viscosity at the moment of graphite
precipitation was applied in the cast iron property data set.
These coupled models allowed considering joint effects of
mold/gating system geometry, heat transfer and cast iron
solidification on transport of dissolved species. Therefore,
it was possible to evaluate the influence of melt mixing in
the mold on formation of dual graphite structure in the
casting.
EXPERIMENTAL RESULTS HEAT A The objective of this heat was to evaluate the effect of
post-treatment sulfur additions to an Mg-treated ductile
iron melt on the shape of graphite in a one-inch wall
thickness casting, solidified in a no-bake mold. Ductile
iron was produced by magnesium treatment and
inoculated in 200-lb ladle. The iron was poured into six
15-lb hand ladles with different amounts of FeS2. The
chemistry and the results of ATAS thermal analysis are
given in Table 5 and Fig.4. Transformation of
microstructure in magnesium treated iron by post-
treatment sulfur additions is shown in Fig. 5. Progressive
sulfur additions produced the expected transformation of
cooling curves in a series typical for iron with spheroidal
graphite, CGI and finally for gray iron with flake
graphite. The specific parameters, which increase, are
recalescence (R) and temperatures of eutectic reaction
(TELOW and TEHIGH). A small amount of post-treatment
sulfur addition (0.02-0.05% FeS2) increased the number
of spheroidal graphite nodules, counted by automated
SEM/EDX analyzer (Table 6). The mechanism of this
effect has been described in literature.8 An addition of
oxygen (CuO will be reduced by C and/or Si in the melt8)
to a large amount of FeS2 had no significant effect on
cooling curve parameters.
A graph of residual sulfur in the casting versus sulfur
additions to the ladle has two distinct regions (Fig. 6a).
First portion of the added sulfur reacted with dissolved
magnesium and part of the reaction product (MgS) floated
to the surface of the melt decreasing the amount of
remaining sulfur in the melt. After full neutralization of
magnesium, the slope of this graph increased because all
thebadded sulfur went into solution in the melt. The
experimental FeS2 additives for forming flake graphite
were close to calculated stoichiometry of MgS reaction. It
was also shown, that post-treatment sulfur addition
produced a relatively wider CGI stability window (from a
full vermicular structure without flake graphite to a
mixture of vermicular with 50% spherical graphite) when
compared to the typical narrow window (Mgresidual =0.015-
0.020) for magnesium treatment (Fig. 6c).
The non-metallic inclusion characteristics are given in
Table 7. In Mg-treated ductile iron, the main inclusions
were different types of complex Si-Al-Mg-Ca oxides with
a limited amount of Mg-Ca-sulfides. Post-treatment sulfur
addition significantly increased the amount of sulfide
inclusions, some of which were attached to flake graphite
(Fig. 7b). It is interesting to observe that sulfides were
mainly present as MgS in the base Mg-treated ductile iron
while complex Mg-Mn-S inclusions were present with
flake graphite in iron produced with post-treatment sulfur
additions (Fig. 7a).
Table 5. Chemistry of Metals and Critical Points on
Thermal Curves from Six Ladles
S Mg TL TELOW TEHIGH R
Base 0.002 0.03 1148.0 1140.1 1141.0 0.9
0.02% FeS2 0.008 0.034 1148.0 1139.8 1141.5 1.7
0.05% FeS2 0.012 0.029 1149.9 1142.9 1147.0 4.1
0.1% FeS2 0.022 0.028 1150.7 1142.9 1149.2 6.3
0.1% FeS2
+0.1%CuO 0.022 0.029 1151.8 1142.4 1148.6 6.2
0.14% FeS2 0.042 0.029 1149.9 1149.9 1156.8 6.9
a)
b)
Fig. 4. Effect of ladle post-treatment sulfur additions into magnesium treated iron (a) on the cooling curves) and (b) on the solidification parameters is graphed.
1120
1130
1140
1150
1160
1170
1180
0 50 100 150 200
Tem
pe
ratu
re,
0C
Time, sec
Base
0.02%FeS2
0.05%FeS2
0.1%FeS2
0.1%feS2+0.1%CuO
0.14%FeS2
0
1
2
3
4
5
6
7
8
1140
1145
1150
1155
1160
0 0.02 0.04 0.06 0.08
Re
ca
lesce
nce
, 0C
Te
mp
era
ture
, 0C
S addition, wt. %
TL
TELOW
TEHIGH
R
Paper 12-024.pdf, Page 4 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
Fig. 5. These photomicrographs show the microstructures of iron castings from six ladles with post-treatment sulfur additions in the ladles.
Table 6. Quantitative Metallography of Graphite Nodules in Mg-Treated Irons with Post-Treatment
Sulfur Addition
Additions N, 1/mm2 D aver, µm
Base 145 24.0
0.02%FeS2 169 22.5
0.05%FeS2 199 20.3
Table 7. Non-Metallic Inclusions in the Castings
Additions Total #/mm2 Average D, µm
sulfides oxides sulfides oxides
Base 83 175 1.37 1.48
0.14 % FeS2
552 56 1.47 1.74
a)
b)
c)
Fig. 6. Effect of post-treatment sulfur addition in the ladle on (a) remaining sulfur and magnesium concentrations in iron, (b) shape of graphite in the castings and (c) typical graphite morphology in magnesium treated iron
10 are graphed.
a) b)
Fig. 7. (a) Composition (b) and backscattered SEM image of sulfide inclusions attached to graphite flake in iron with post-treatment sulfur addition are shown.
HEAT B The objective of this heat was to produce a sharp change
in graphite shape during pouring of Mg-treated melt by
post-treatment with in-stream sulfur injection during mold
0
0.01
0.02
0.03
0.04
0.05
0 0.02 0.04 0.06 0.08
Sin
ca
sti
ng
(w
t. %
)
S added in ladle (wt. %)
Sad [S][S] + [Mg] = {MgS}
Sad [S]
0
0.01
0.02
0.03
0.04
0.05
0 0.02 0.04 0.06 0.08
[S] a
nd
[Mg
] in
iro
n, w
eig
ht %
S addition in melt, weight %
[S][Mg]
FlakeCompactedNodule
Stoichiometry level
Paper 12-024.pdf, Page 5 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
pouring. The percentage of FeS2 additions (> 0.1 wt.%)
for transformation of spheroidal graphite to flake graphite
was defined in the previous Heat A with post-treatment
with sulfur ladle additions. The experimental procedure
for heat B included melt pouring into a stationary tundish
with a bottom nozzle and melt collection into moving
molds as shown in Fig. 1. A graphite nozzle insert in the
stationary tundish with an 18 mm diameter hole was used
for controlling the fill rate near a constant—2 kg/sec. The
first three molds were poured using the base Mg-treated
ductile iron and after that the powder feeder was turned
on for in-stream post-treatment sulfur injection. A powder
feeder injected 2.8 - 3.0 g/s of FeS2 in-stream. In this
experiment, post filling mixing of treated and untreated
melts was prevented and in-stream post-treatment sulfur
injection, provided the desired sharp change in shape of
graphite from spheroidal to flake in sequentially cast
plates (Fig. 8).
Fig. 8. These are the microstructures of casted 1 in. thick plates in different molds: base and in-stream FeS2 injected melt.
HEAT C The objective of this heat was post-treatment in-stream
sulfur addition by injection during pouring the
experimental casting with a bottom gating system having
multiple bottom ingates (Fig. 2). Optimal post-treatment
in-stream sulfur injection parameters were verified in the
previous Heat B. The experimental casting consisted of
two horizontal plates with an internal flat core. The goal
was to produce graphite structure modification in the
bottom plate of the casting. According to this
experimental plan, the last portion of the melt filling the
bottom plate needed to be sulfur injected. Table 8 and Fig.
9 provide experimentally achieved pouring data obtained
from electronic scale and video recording.
Table 8. Pouring Process Data in Heat C
Pouring time, sec
Time to start FeS2, sec
FeS2 to melt, %
Total weight,
kg
Weight of treated part, kg
40.5 21 0.15-0.20 43.8 21
Fig. 9. Pouring weight and pouring rate in Heat C (blue arrow) indicate the starting time of FeS2 injection.
Samples taken from the ladle of the base Mg-treated iron
had spheroidal graphite structure. The last portion poured
into the mold was post-treatment iron with in-stream
sulfur injection. This latter iron was sampled from the
gating and had flake graphite. Thus, in this experiment, a
sharp change of melt chemistry was achieved by using
post-treatment in-stream sulfur injection. However,
entirely mixed CGI microstructures were observed at
multiple locations throughout the casting (Fig. 10). This
result indicated that mold filling mixing, post filling
momentum, natural convection in the mold before casting
solidification or some combination of these could be
responsible for producing mainly CGI structure in casting.
(a) (b)
(c) (d)
(e)
Fig. 10. These are micrographs of the microstructure in sections of experimental casting in Heat C: (a) in ladle, (b) gating system, (c) bottom plate, (d) middle region and (e) casting top.
0
0.2
0.4
0.6
0.8
1
1.2
1.4
1.6
1.8
0
5
10
15
20
25
30
35
40
45
50
0 10 20 30 40 50 60
Po
uri
ng
ra
te,
kg
/se
c
We
igh
t, k
g
Time, sec
Weight
Paper 12-024.pdf, Page 6 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
HEAT D The objective of this heat was experimental verification of
formation of dual graphite structure in the casting when
all of the projected effects, which promoted melt mixing
in the mold were prevented. The procedure for this heat
was designed as an ideal hypothetical case in which
different layers of casting were poured sequentially: first
with gray iron then with ductile iron with an intermediate
time delay for partial cooling of the first layer to avoid
melt flow by natural convection and mixing. The time
delay was designed to allow development of a mushy
zone in the bottom gray iron layer before pouring ductile
iron on the top of gray iron. The casting for this heat had
three 1in. thick, 10 in. x 10 in. horizontal plates with two
1-1/4 in. x 8 in. x 8 in. internal cores each having two core
prints. To use two poured melts, the mold had two side-
type gating systems with gates placed in the parting plane
for one system and above parting plane for the other
gating system (Fig. 3).
Two melts, gray iron and Mg-treated ductile iron, were
prepared simultaneously and sequentially poured into the
mold positioned on electronic platform scale. ATAS
thermal analysis results for these two irons, sampled from
the ladles, are shown in Fig. 11 and chemistries are given
in Table 9. Temperature in the casting was monitored by
four thermocouples, installed on different heights. Gray
iron was poured first into the bottom plate to
approximately one inch thick and the ductile iron melt
which followed was poured with approximately 150 sec
delay to the point that eutectic solidification had started in
the bottom plate (Fig. 12a). Calculation using a computer
assisted thermal analysis technique,9 estimated the
fraction of solid as 0.4-0.5 in the first layer at that time.
Macro observations from multiple vertical cuts (Fig. 13 a)
clearly showed two layers in the casting with a seamless
boundary. The geometry of this boundary had some
shallow areas beneath the gate locations and showed
penetration of the ductile iron from the second pour into
the semisolid gray iron. Micro-observation showed a
straight local boundary layer with a narrow reaction zone.
Vermicular type of graphite particles occur as
protuberances into ductile iron (Fig. 13c), which originate
from the flake graphite. This boundary structure was
formed between the Mg treated iron and the base iron
without Mg and indicated that some amount of residual
Mg is needed to promote growth of vermicular shape
graphite. There were no observations of oxidized films or
indications of cold shut/cold lap. This observation showed
that the narrow boundary layer will have no negative
effect on thermal and mechanical properties on casting
with dual graphite structure.
a)
b)
Fig. 11. These graphs show the ATAS thermal analysis of cooling curves from (a) GI and (b) DI that are used in double poured Heat D.
Table 9. Chemistry of Two Metals Poured In Bottom and Top Parts of Casting (wt. %) in Heat D.
Sample C Si Mn Cu S P Ni Mg
GI melt 3.77 1.44 0.26 0.64 0.009 0.010 0.04 0.004
Casting bottom 3.59 1.46 0.22 0.54 0.007 0.010 0.03 0.006
DI melt 3.66 2.50 0.33 0.64 0.004 0.010 0.04 0.047
Casting top 3.00 2.30 0.16 0.30 0.004 0.010 0.02 0.040
Paper 12-024.pdf, Page 7 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
a)
b)
Fig. 12. (a) The cooling curves are obtained from bottom and top parts of the double poured casting in Heat D and (b) the calculated solid fraction in bottom part of casting is indicated by arrow at the moment of second pouring.
a) b)
c) d)
Fig. 13. (a) This section of casting from Heat D indicates GI in bottom plate and DI in top part of casting. (b) This is a SEM image. (c) Not etched and (d) etched microstructures of the boundary show graphite shape transformation.
DISCUSSION
MODELING PREDICTION AND COMPARISON WITH EXPERIMENTS There are three possible melt mixing mechanisms. The
first mechanism involves melt mixing during mold filling.
Figure 14 illustrates the effect of the gating system on
predicted mixing during mold filling when the chemistry
of the iron was sharply changed during the pouring
process. In these figures, the red color represents the
highest percent Mg in the melt. In the bottom filled gating
system, ductile iron melt was delivered first and gray iron
(blue color) was poured last. The bottom filled casting
with multiple distributed gates provided better results as
compared to a two-gate bottom system (Figs. 14a, 14b).
However, the bottom gating system did not allow gray
iron melt to sequentially and fully replace the ductile iron
melt without intermixing, even at low inlet melt velocity.
The experimental results in Heat C (Fig. 10) confirmed
the intensive melt mixing in the casting when a bottom
gating system was used.
a) b)
c) d)
Fig. 14. These illustrations show the modeling of Mg distribution just after filling molds (a, b) through bottom and (c. d) and side gating systems with variations on ingate location (shown by black arrows).
Better results with less mixing during pouring were
predicted for a symmetrical side gating system (Fig. 14d).
In this case, gray iron was used first for pouring the
bottom part and ductile iron was poured second and
placed above the bottom layer. Experiments (Heat D)
confirmed that some penetration in the bottom layer could
take place by a falling stream from the side gates (Fig.
13a). Also, the gating system design must avoid any
“sloshing” effect from post filling momentum (the second
mixing mechanism) which could take place if the melt did
not enter the mold cavity symmetrically (Fig. 14c).
The third important mixing effect is natural convection
within the melt in the mold cavity. This effect is
illustrated in Fig. 15. It was assumed in this model that
the bottom layer of gray iron was poured ideally into the
lower part of the mold cavity and ductile iron was put on
the top without any post-filling momentum (Fig. 15a).
Two cases were modeled. In the first case, all of the melt
cooled and solidified in the no-bake mold. Modeling
Paper 12-024.pdf, Page 8 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
predicted severe mixing by natural convection when the
casting solidified with complete destruction of the
initially layered structure (Fig. 15b). This result of severe
mixing was observed in Heat C. The mixing due to free
convection could be suppressed significantly if a bottom
chill was applied and the bottom layer of casting
consequently froze rapidly (Fig. 15c). Rapid melt cooling
shortens free convection flow time and concurrently and
quickly develops a mushy zone which increases melt
viscosity. These factors can significantly suppress mixing
due to free convection of the melt.
a)
b) c)
Fig.15. These illustrate the changing Mg-distribution by free melt convection and post filling momentum during melt cooling in mold before casting solidification: (a) after pouring, (b) before solidification in sand mold and (c) using bottom chilling.
PROCESS CLASSIFICATION AND POSSIBLE PROCESS ROUTES Development of dual graphite structure in the same
casting is a serious challenge due to the requirement to
suppress “natural” melt mixing during mold pouring and
casting solidification. A process chart was developed
based on the experiments and the modeling study
discussed, previously (Fig. 16). The process includes
three stages which need to be specifically designed:
development of dual chemistry; sequential mold pouring
using different melt compositions and finally melt cooling
and solidification to prevent melt convection. The bottom
part of the chart indicates the three possible mixing
mechanisms which are deleterious to development of a
controlled dual graphite structure. These mechanisms
include: mixing during pouring, mixing by post-pouring
momentum and mixing by convection. The right side of
this chart contains different possible techniques which can
be exploited for these process stages. For example, dual
chemistries could be produced using two melts (Heat D),
in-steam treatment (Heat C) or so called in-situ techniques
which are the subject of ongoing research. Different
gating system designs included: a two stream system
(Heat D), a bottom-gating system (Heat C) where the
sulfur-treated last liquid metal poured would conceptually
push the first portion of poured metal up to the top; and a
side gating system which allows pouring one liquid metal
on the bottom of casting and placing other metal on the
top. Also, other different combinations can be designed
for specific casting geometry and location in the mold.
Finally, different methods are possible for suppressing
melt mixing in the mold. In Heat C, uncontrolled casting
solidification in no-bake mold created intensive melt
mixing. Freezing of the bottom layer of gray iron to a
mushy condition in Heat D allowed clearer separation of
the dual graphite structure. A more practical way to
control solidification could be by applying internal
or/external chill. A combination of these techniques can
be designed using the suggested process chart.
Fig. 16. Process chart for different routes for producing of dual graphite structure in cast iron casting is illustrated.
CONCLUSIONS
Development of dual graphite structure in the same
casting has potential for future application of cast iron.
Post-treatment of some proportion of the melt by in-
stream sulfur injection is sufficient to achieve controlled
dual graphite structure. This study showed that the main
technological problem of development dual graphite
structure is extensive liquid metal mixing during pouring,
and as a result of post filling momentum and natural
convection. Four different techniques were
experimentally tested and modeled. Both of the
microstructure and macrostructure examinations indicated
the effect of different mixing mechanism on formation of
dual graphite structure in casting.
From the results achieved by so far, mainly CGI structure
was observed throughout the casting poured through a
bottom gating system because of severe mixing. In this
case, pouring mixing, post pouring momentum and
natural convection are all possible mechanisms resulting
in undesired mixing structure. However, “double pour”
and partially freezing of the first poured liquid metal
prevented both pouring and post pouring mixing to some
extent and showed very sharp boundary between DI and
GI.
Paper 12-024.pdf, Page 9 of 10AFS Proceedings 2012 © American Foundry Society, Schaumburg, IL USA
The classification of different process routes with the goal
of suppressing melt mixing was suggested in the process
chart. Based on the process chart, the different ways for
preventing mixing such as directional solidification by
chilling, in-situ sulfur treatment by localized reaction and
a side gating system, which impaired pouring mixing, can
be used for the development of dual graphite structure in
casting.
ACKNOWLEDGMENTS Research was sponsored by Benet Laboratories on behalf
of the US Army Contracting Command Joint Munitions
and Lethality Contracting Center—and was accomplished
under Cooperative Agreement Number W15QKN-11-2-
0001. The views and conclusions contained in this
document are those of the authors and should not be
interpreted as representing the official policies, either
expressed or implied, of AFS or Benet Laboratories or the
U.S. Government. The U.S. Government is authorized to
reproduce and distribute reprints for Government
purposes notwithstanding any copyright notation heron.
The authors would like to thank US Army ARDEC-Benet
laboratories for funding this research. The authors wish to
acknowledge technical support and discussion input of
George Kokos, James Barlow and Zhiping Lin from
Caterpillar. The authors wish to recognize the assistance
of Forrest Huebner, Clinton Ratliff, John Stanek, Jeremy
Robinson, Bradley Bromet, Ian Christian and James
Smoot for heat pouring and sample preparation.
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