3 ,',& ,
Chapter 3 Flow
3.3 Fluidity (maximum fluidity length) Lf
3.3.4 Effect of surface tension
However, for investment casting the ceramic shell allows a complete range of temperatures to be chosen without difficulty. From Equation 3.3 it is seen that the freezing time is proportional to the difference between the freezing point of the melt and the temperature of the mould. The few tests of this prediction are reasonably well confirmed (for instance, Campbell and Olliff 1971).
One important prediction is that when the mould temperature is raised to the melting point of the alloy, the fluidity becomes infinite; i.e. the melt will run for ever! Actually, of course, this self- evident conclusion needs to be tempered by the realization that the melt will run until stopped by some other force, such as gravity, surface tension or the mould wall! All this corresponds to common sense. Even so, this elimination of fluidity limitations is an important feature widely used in the casting of thin-walled aluminium alloy investment castings, where it is easy t o cast into moulds held at temperatures in excess of the freezing point of the alloys at approximately 600°C. Single crystal turbine blades in nickel-based alloys are also cast into moulds heated to 1450°C or more, again well above the freezing point of the alloy.
Any problems of fluidity are thereby avoided.
Having this one concern removed, the founder is then left with only the dozens of additional important factors that are specified for the casting. Solving one problem completely is a help, but still leaves plenty of challenges for the casting engineer!
Flow 89
1000
8oo-
E
600-s-
ul r
W Q
-
'E 400- 5
200
0
6 0 0 - 8 I I I I I I I I t t I t i I i i
Casting temp.
"C
1570 1
o 1620 I 10 500
E E 400
$ 300 5 200
0 2 4 6 8
Strip thickness, mm
a LM2511.5 mm 0 LM2512.5 mm A LM2513.5 mm A LM2516.5 mm 0 LM2516.5 mm GR + A17Si/l.5 mm
A17Sil3.5 mm
+ A17SiI6.5 mm
-
3.5 mm
-
2.5 mm
-
-0 50 100 150 200
o 1520
100: -
0 E O ' L
,
, , ,,
, ,, :
0 2 4 6 a 10
Strip thickness, mm Figure 3.14 ( a ) Fluidity data.for a low alloy steel, and ( b ) , f o r a stainless steel poured in a straight channel. ,furan bonded sand mould (Boutorabi et al. 1990).
Figure 3.15 Fluidity o f u variety of AI-7Si and Al-7Si-O.4Mg alloys, one grain refined GR, yhowing linear behaviour with section thickness and casting temperature (Boutorabi et ul.
1990).
were removed, giving considerably improved reproducibility of the fluidity test.)
Typical results for a vacuum-cast nickel-based superalloy are given in Figure 3.19 (Campbell and Olliff 197 1). Clearly, the 1.2 mm section fills more fully than the 0.6 mm section. However, it is also clear that at low casting temperature the filling of both sections is limited by the ability of the metal to flow prior to freezing. At these low casting temperatures the fluidity improves as temperature increases, as expected.
However, above a metal casting temperature of approximately 1500°C further increases of temperature do not further improve the filling. As the metal attempts to enter the diminishing sections of the mould, the geometry of the liquid front is closely defined as a simple cylindrical surface. Thus it is not difficult to calculate the thickness of the mould at any point. Half of this thickness is taken as the radius of curvature of the liquid metal meniscus (Figure 3.20). It is possible to predict, therefore, that the degree of filling is dictated by
0 x, 1 2 3 4 5 6 7 8 9 1 0 Thicknesdmm
Figure 3.16 Fluiditj of a varietj of grey and ductile cast irons showing linear behaviour with section thickness and casting temperature (Boutorabi et al.
I 990).
Figure 3.17 Fluidity results re- 7.5 alloys.
0 0 + I 2 3 4 5 6 7 presented from Figure 3.15 for Al-
Section thickness x (mm)
Flow 91 0
5
\
i-
I- P
-
/
7
Figure 3.18 Aerofoil fluidity test mould. The outlines of the ca.st shape are computed f o r increasing values of yl pgh, units in rnillimetres (Campbell and Olliff 1971).
100 c 80
& 60
-
2 40 20 0 1 .z
8
Q
m
1
.- c
m
1.2 mm
I I I I
00 1400 1500 1600 1700
Casting temperature ("C) Figure 3.19 Results from the aerofoilfluidity test (Campbell and Olliff 1971) (lines denote theoretical predictions; points are experimental data).
the local balance at every point around the perimeter of the meniscus between the filling pressure due to
i
LI
i-
93
.
Figure 3.20 Geometry ofthe aerofoi1,fluidity test (Campbell and Olliff 1971).
the metal head and the effective back pressure due to the local curvature of the metal surface. In fact, if momentarily overfilled because of the momentum of the metal as it flowed into the mould, the repulsion effect of surface tension would cause the metal to 'bounce back', oscillating either side of its equilibrium filling position, finally settling at its balanced, equilibrium state of fullness.
The authors of this work emphasize the twin aspects of filling such thin sections; flowability limited by heat transfer, and fillability limited by surface tension.
At low mould and/or metal temperatures, the first type of filling, flowability, turns out to be simply classical fluidity as we have discussed above.
Metallographic examination of the structures of aerofoils cast at lower temperatures showed columnar grains grown at an angle into the direction of flow, typical of solidification occurring while the metal was flowing. The flow length was controlled by solidification, and thus observed to be a function of superheat and other thermal factors.
as we have seen.
The second type of filling, fillability, occurs at higher mould and/or metal temperatures where the heat content of the system is sufficiently high that solidification is delayed until after filling has come to a stop. Studies of the microstructure of the castings confirm that the grains are large and randomly oriented, as would be expected if the metal were stationary during freezing. Filling is then controlled by a mechanical balance of forces. The mode of solidification and further increases of temperature of the metal and the mould play no part in this phase of filling.
In a fluidity test of simpler geometry consisting of straight strips of various thickness, the linear plots of fluidity Lf versus thickness x and superheat ATs are illustrated in Figures 3.15 and 3.16 for Al- 7Si alloy and cast iron in sand moulds. It is easy to combine these plots giving the resultant three- dimensional pyramid plot shown in Figure 3.17.
The plot is based on the data for the A1 alloy in Figure 3.15. In terms of the pressure head h, and the intercepts ATo and xo defined on fluidity plots 3.15 and 3.16, the equation describing the slightly skewed surface of the pyramid is
(3.15) Where C i s a constant with dimensions of reciprocal temperature. For the A1 alloy, C i s found from Figure 3.15 to have a value of about 1.3 f 0.1 K-I, ATo = 30 f 5 , y = 2 Nm-' allowing for contribution of oxide film to the surface tension, p = 2500 kgm-3, g = 10 msK2 and h = 0.10 m. We can then write an explicit equation for fluidity (mm) in terms of superheat (degrees Celsius) and section thickness (mm):
Lf = C(ATs
+
ATO)(X - (2y/pgh))Lf = 1.3(ATs
+
3 0 ) ( ~ - 1.6)For a superheat ATs = 100°C and section thickness x = 2 mm we can achieve a flow distance Z+ = 68 mm for AI-7Si in a sand mould. If the head h were increased, fluidity would be higher, as indicated by Equation 3.15 (but noting the limitations discussed in section 3.3.2).
As we have seen, in these thin section moulds both heat transfer and surface tension contribute to limit the filling of the mould, their relative effects differ in different circumstances. This action of both effects causes the tests to be complicated, but, as we have seen, not impossible to interpret. Further practical examples of the simultaneous action of heat transfer and surface tension will be considered in the next section.