Castings 2 Episode 11 ppt

238 Castings binder, and so allowing the core to collapse earlier. 2. Tie bars can be connected across the open side of the box, thereby holding the walls in place, and balancing the effect of the contraction of the closed side. The tie bar need not be a separate device. It can be the running system of the casting, carefully sized so as to carry out its two jobs effectively. f/ 4- + This raises the important issue of the influence of the running and feeding system. Unfortunately these appendages to the casting cannot be neglected. They can be used positively to resist casting distortion as above. Alternatively they can cause distortion and even tensile failure as shown in the simple case in Figure 8.8a. Alternatively, if this casting is fed at the flange end, leaving the plate free to contract along its length as shown in Figure 8.8b, the problem is solved. However, note the important point that whether or not casting ‘a’ has suffered any tensile failure, it will be somewhat longer than casting ‘b’ . Thus different pattern contraction allowances are appropriate for these two different constraint modes. Free contraction Figure 8.8 (a) Effect of the running and feeding systems imposing constraint on the contraction of the casting. (b) Applying the running and feeding io the opposite end of the casting removes the problem. of casting (b) In more complex castings the effect of geometry can be hard to predict, and harder to rectify if the casting is particularly badly out of shape. Especially for large, thin-walled castings requiring close dimensional tolerance it may be wise to include for a straightening jig in the tooling price. 8.2.2 Casting constraint Even if the casting were subjected to no constraint at all from the mould, it would certainly suffer internally generated constraints as a result of uneven cooling. The famous example of this effect is the mixed-section casting shown in Figure 8.9a. If a failure occurs it always happens in the thicker section. This may at first sight be surprising. The explanation of this behaviour requires careful reasoning, as follows. (b) Figure 8.9 (a) Thickhhin-section casiing showing iensile stress in the thick section: (b) an even-walled casting showing internal tensile stress. First, the thin section solidifies and cools. Its contraction along its length is easily accommodated by the heavier section, which simply contracts under the compressive load since it is hot, and therefore plastic, if not actually still molten. Later, however, when the thin section has practically finished contracting, the heavier section starts to contract. It is now unable to squash the thin section significantly, which has by now become rigid and strong. The result is the bending of the thin section, or the failure in tension of the thick section, which, depending on its temperature, will experience a tensile load. It can therefore stretch plastically, or hot tear, or cold crack (these defect modes are discussed later). The example shown in Figure 8.9b is another common failure mode. The internal walls of a casting remain hot for longest even though the casting may have been designed with even wall sections. This is, of course, simply the result of the internal sections being surrounded by other hot sections. The reasoning is therefore the same as that for the thick- Linear contraction 239 reinforced by heavy-section ribs. It is often seen in thin-section boxes that have reinforcing ribs around the edges of the box faces. The general argument is the same as before: the thin, flat faces cool first, and the subsequent contraction of the heavier ribs causes the face to buckle, springing inwards or outwards. This is known as ‘oil-can distortion’; an apt name which describes the exasperating nature of this defect, as any attempts to straighten the face cause it to buckle in the opposite direction, taking up its new reversed curvature. It can be flipped backwards and forwards indefinitely, but not straightened permanently. Once a casting exhibits oil-can distortion it is practically impossible to cure. The effect is more often seen after quenching from heat treatment, where, of course, the rate of cooling is greater than in the mould, and where the casting does not have the benefit of the support of the mould. Oil-can distortion may be preventable by careful design of ribs, to ensure that their geometrical modulus (Le. their cooling rate) is similar to. or less than, that of the thinner flat face. Alternatively, the cooling rate from the quench needs to be equalized better, possibly by the use of polymer quenchants and/or the masking of the more rapidly cooling areas. In ductile iron castings that exhibit expansion on solidification due to graphite precipitation, the expansion can be used to good effect to reduce or eliminate the necessity for feeders, particularly if the casting cools uniformly. Tafazzoli and Kondic (1977) draw attention to the problem created if the cooling is not uniform. The freezing of sections that freeze first leads to mould dilation in those regions that solidify last. Although these authors attribute this behaviour to the mismatch between the timing of the graphite expansion and the austenite contraction, it seems more likely to be the result of the pressure within the casting being less easily withstood by those portions of the mould that contain the heavier sections of the casting. This follows from the effect of casting modulus on mould dilation; the lighter sections are cooler and stronger, and the thicker sections are hotter and thus more plastic. Any internal pressure will therefore transfer material from those sections able to withstand the pressure to those that cannot. The thinner sections will retain their size while the thicker sections will swell. Tafazzoli and Kondic recommend the use of chills or other devices to encourage uniformity. In a classic series of papers Longden (193 1-32, 1939-40, 1947-48, 1948) published the results of measurements that he carried out on grey iron lathe beds and other machine beds. The curvature of a casting up to 10 m long could result in a maximum out-of-line deviation (camber) of 50 mm or more. Longden summarized his findings in a nomogram that allowed him to make a prediction of the camber I I I I I I I I I I I I I I I I I I I I I /thin-section casting above. The internal walls of the casting suffer tension at a late stage of cooling. This tension may be retained as a residual stress in the finished casting, or may be sufficiently high to cause catastrophic failure by tearing or cracking. The same reasoning applies to the case of a single-component heavy-section casting such as an ingot, billet or slab, and especially when these are cast in steel, because of its poor thermal conductivity. The inner parts of the casting solidify and contract last, putting the internal parts of the casting into tension (notice it is always the inside of the casting that suffers the tension; the outside being in compression). Because of the low yield point of the hot metal, extensive plastic yielding occurs at high temperature. However, as the temperature falls, the stress cannot be relieved by plastic flow, so that increasing amounts of stress are built up and retained. There has been much experimental and theoretical work in this area. Some of this work will be touched on in section 8.5. An example shown in Figure 8.10 shows the kind of distortion to be expected from a box section casting with uneven walls. The late contraction of the thicker walls collapses the box asymmetrically (the casting is at risk from tensile failure in the thicker walls, but we shall assume that neither tearing nor cracking occurs in this case). There is clearly some strong additional effect from mould constraint. If the central core were less rigid, then the casting would contract more evenly, remaining more square. I I I I I I I I I I I I I I I I II 11 II II II II II II I I I_-___ ___-___ ____-__ There is an important kind of distortion seen in plate-shaped castings which have heavy ribs adjoining the edges of the plate, or whose faces are 240 Castings to be expected on any new casting. The reverse camber was then constructed into the mould to give a straight casting. Although Longden’s nomogram is probably somewhat specific to his type of machine tool bases, it is presented in Figure 8.1 1 as an example of what can be achieved in the prediction of casting distortion. It is to be expected that castings of other types will exhibit a similar relationship. We can convert Longden’s qualitative summary into a quantitative form, where length L and depth D are in metres, and wall section thickness w and camber c are in mm. Simplifying Longden’s nomogram within the limits of accuracy of the original data, and making the further approximation that the slight curved lines on the left-hand side are straight lines through the value L = 1.5 m, then with perhaps about 10 per cent accuracy for wall thickness w from 10 to 40 mm the camber is given by: c = (I. - 1.5)(7.62 - 1.073~) - 6600 + 310 and for wall thickness w from 40 to 70 mm: c = (L - 1.5)(144 - 2.03~)( 1 - D) We now move on to a further type of internal constraint that appears to be universal in castings of all sizes and shapes, and which is rarely recognized, but was investigated by Weiner and Boley (1963) in a theoretical study of a simple slab casting. They assumed elastic-plastic behaviour of the solid, and that the yield point of the solid was zero at the melting point (not quite true, but a reasonable working approximation) and increased as the casting cooled. They found that plastic flow of the solid occurs at the very beginning of solidification. The stress history of a given particle was found to be as follows. On freezing, the particle is subject to tension, and since the yield stress is initially zero, its behaviour is at first plastic. As it cools, the tensile stress on it increases and remains equal to the yield stress corresponding to its temperature until such time as the rate of increase of stress upon it is less than the rate of increase of its yield stress. It then starts to behave elastically. Soon after, unloading begins, the stress on the particle decreasing rapidly, becoming compressive, and finally reaching the yield stress in the opposite direction. Its behaviour remains plastic thereafter. Weiner and Boley’s analytical predictions have been accurately confirmed in a later numerical study by Thomas and Parkman (1 998). To sum up their findings, in a solidifying material there will be various deformation regimes. These are (i) a plastic zone in tension at the solidification front, since the strength of the solid is low; (ii) a central region where the stresses are in the elastic range; and (iii) a zone at the surface of the casting where there is plastic flow in compression. The overall scheme is illustrated in Figure 8.12. The propagation of the tensile plastic region, the central elastic zone, and the compressive plastic zone are reminiscent of the propagation of the various transformation zones through the sand mould. These are waves of the stresslstrain regime that spread through the newly solidified casting, 600 ]500 Length of casting L (rn) Depth of casting D (rn) Figure 8.1 1 Camber- nomogram by Longden (1948) for machine tool bed castings in grey iron. (During conversion to metric units rhe figure has been smoothed somewhat, and broken lines show minor extrapolation.) A 10 m long casting with 25 mm section side wdls, and 750 mm depth will show approximately 210 mm camber. Linear contraction 2-1 I Mould Elastic compression * c:; :: * Plastic I Elastic Plastic Liquid + 4 t Solid Compression Tension Distance - I I V I I ii I I I 0 0.2 0.4 0.6 0.8 1.0 Distance through the solid remaining parallel to the solidification front and remaining at the same relative distances, as illustrated in Figure 8.12. If the yield stress were not assumed to be zero at the freezing point, but were to be given some small finite value, then the analysis would be expected to be modified only very slightly, with a narrow elastic zone appearing at the solidification front on the right-hand side of Figure 8.12. The analysis will be fundamentally modified for materials that undergo certain phase changes during cooling. If the crystallographic rearrangement involves a large enough shear strain, or change of volume, as is common in steels cooling through the y to atransition, for instance, then the material will be locally strained above its yield point, adding an additional plastic front which will propagate through the material. The phenomenon is known as trans@rmation induced plasticity (TRIP). This additional opportunity for the plastic relief of stress will fundamentally alter the distribution of stress as predicted in Figure 8.12. However, Figure 8.12 Figure 8.12 Ela.stic/p/tr.stic reginies in (I simple tltih custing (after Weiner and Bole! 1963). is expected to be reasonably accurate for many other metals such as zinc-, aluminium-, magnesium-, copper- and nickel-based alloys, and for thoqe steels that remain single phase from solidification to room temperature. The high internal tensions predicted by this analysis will be independent of, and will be superimposed on, stresses that arise as a result of other mould and/or casting constraints as we have discussed above. It is not surprising, therefore, to note that on occasions castings fail while cooling after solidification, as will be digcussed in section 8.4. Drezet and co-workers (2000) showed that the elastidplastic model would not be fundamentally altered if creep flow behaviour were assumed instead of the elastic-plastic flow behaviour with yield stress a function of temperature. They found that such simulations were insensitive to the rheological model employed, but that the deformation was mainly a simple function of the thermal contraction and the conditions for continuity. 242 Castings Richmond and Tien (1971) and Tien and Richmond (1982) criticize the model by Weiner and Boley on the grounds that they do not take account of the friction at the casting/mould interface. When Richmond and Tien include this they find that the casting/mould interface is no longer in compression but in tension. This is almost certainly true for large castings in metal moulds such as steel ingots in cast iron ingot moulds, where the pressure between the mould and casting is high, the friction is high, and the mould is rigid. These authors explain the occurrence of surface cracks in steel ingots in this way. However, Weiner and Boley are likely to be more nearly correct for smaller castings in sand moulds. Here the interfacial pressure will be less, and the surface more accommodating, and the air gap ensuring that the casting and mould are not in contact at all in some places. All these factors will reduce the restraint due to friction. Thus their analysis remains probably the most appropriate for medium-sized shaped castings. 8.3 Hot tearing 8.3.1 General A hot tear is one of the most serious defects that a casting can suffer. Although it has been widely researched, and is understood in a general way, it has remained a major problem in the foundries, particularly with certain hot-tear-prone alloys. It has to be admitted that the most important insights into hot tearing behaviour have recently emerged not from scientific experiments in the laboratory, but from experience on the shop floor of foundries. Thus these important findings have not been published in the scientific journals. Briefly, for those who wish to read no further, the key result that has recently emerged is that hot tearing can usually be eliminated in most castings and most alloys by simply improving the filling system of a casting. The reader is referred to section 8.3.10.1. In the meantime, there is much useful background that has been clarified by careful, systematic research over the years. This is reported here. It will become clear that it is consistent with the view that bifilms introduced by the running system are implicated in a major way. However, the reader will find it illuminating to review the experimental data, keeping in mind the fact that bifilms are almost certainly the major underlying cause, originating with the poor pouring technique. The defect is easily recognized from one or more of a number of characteristics: 1. Its form is that of a ragged, branching crack. 2. The main tear and its numerous minor offshoots 3. 4. 5. 6. 7. generally follow intergranular paths. This is particularly clear on a polished section viewed under the microscope. The failure surface reveals a dendritic morphology (Figure 8.13a). The failure surface is often heavily oxidized (prior, of course, to any subsequent heat treatment). This is more particularly true of higher temperature alloys such as steels. Its location is often at a hot spot, and where contraction strain from adjoining extensive thinner sections may be concentrated. It does not always appear under apparently identical conditions; in fact it seems subject to a considerable degree of randomness in relation to its appearance or non-appearance, and to its extent. The defect is highly specific to certain alloys. Other alloys are virtually free from this problem. Before we go on to discuss the reasons for all this behaviour, it is worth bearing in mind the most simple and basic observation: The defect has the characteristics of a tear. This disarmingly obvious characteristic immediately gives us a powerful clue about its nature and its origin. We can conclude that: A hot tear is almost certainly a uniaxial tensile failure in a weak material. This may appear at first sight to be a trivial conclusion. However, it is fundamental. For instance, it allows us to make some important deductions immediately: 1. Those theories that are based on feeding difficulties can almost certainly be dismissed instantly. This is because feeding problems result in hydrostatic (i.e. a triaxial) stress in the residual liquid, causing pores or even layer porosity in the liquid phase. If the triaxial stress does increase to a level at which a defect nucleates, then the liquid separates and expands (triaxially) to create a pore among the dendrites. The dendrites themselves are not affected and are not pulled apart. They continue to interlace and bridge the newly formed volume defect, as was discussed for layer porosity in particular (section 7.7.2). This is in contrast to the hot tear, where it is clear from micrographs and X-ray radiographs (Figure 8.13b) that the dendrites open up a pathway first. The opened gap drains free of liquid later. Because the liquid is in open contact with the already drained parts of the tear, it clearly cannot be under any significant hydrostatic [...]... D: x = (Dt)1 '2 (8 .11) 8.5 .2 Quenching stress where D = KIpC and K is the thermal conductivity, about 20 8 Wm-' K-' for aluminium, the density p is 27 00 kgm-' and the specific heat C, is approximately 1000 Jkg-' K-' This gives t h e thermal diffusivity D as about IO4 m' s-' For an aluminium bar of only 20 mm diameter, Figure 8 .25 shows that quenching in water will cool the bar from 500 to 25 0°C within... range Hydrostatic tension in dendrite mesh Binary alloy composition (4 - 2. 57 Binary alloy composition c (b) Figure 8 .22 Sutnmary of ( a ) the eBect qf grain size and presence u f bi$lrns on hot tearing, and ( h ) the relation hetwmv the conditions ,for the incidence of porosity or hot tearing 25 8 Castings illustrated in Figure 8 .22 The continuing low level of porosity at higher solute contents approaching... times from annealing temperatures, especially for large castings Even so, steel castings are not at the same risk of failure from a quenching stress as aluminium alloy castings This is because steels in general enjoy an elongation to failure usually in the range 20 to 40 per cent Thus 1 or 2 per cent quenching 1000 Figure 8 .25 Rates of cooling qf a 20 mm diameter aluminium bar 10 000 when quenched by... developed by the flow of feed metal through the 25 2 Castings Non-equilibrium freezing range 100 90 ' - _ 80 I? c P S c _ - 1 - 70 -c crack length for cone I 9 , / /' _ - n frpp7inn range / v c // 60 , / 0 c Q 8 50 u) u) m _ c 2 p icmg iceptibility ICOEtfficient 40 a , c I 30 Hydrostatic tension I I I J I 20 ' I / , I /\' 10 / 1 ' \ / \ / / 0 0 \ 0 1 I 1 \ '4 2 3 4 5 Copper (wt per cent) 6 7 dendrite... therefore that the temperature differences and cooling rates applying below the gamma-alpha-transformation Linear contraction 26 1 90 - 80 - 70 - a 60- 2 m ? c 50- - a -, m - 40- m, E m oz 3 020 - io - I Oi 0.5 time I 1 I I I 1 2 4 8 16 32 Time in mould before stripping (min) I 64 Figure 8 .24 Residiiul stress in uluminium alloy and grey iron custings as rr ,fiitictioti o/ stripping time Dura,from Dodd ( 1950)... in Figure 8.13a The fracture follows the film, the fracture surface exhibiting steps at integral numbers of dendrite arms, as explained in Figure 2. 41 Naturally, the hot tear morphology is also seen in the room temperature fracture seen in Figures 2. 42 and 2. 43 Clearly, after being flattened out by the growth of dendrites, the bifilm is simply a hot tear waiting to be opened and so be revealed If it... ends of the cast bars If the restraint was not released before the castings cooled to 20 0°C then a loud crack was heard, corresponding to the complete fracture of the bar Similar failures during the cooling of steel castings are also well known, as was, for instance, reported by Steiger as long ago as 1913 During the cooling of steel castings there is a succession of particularly vulnerable temperature... coefficient of thermal expansion of the alloy a We have : E = AL/L = aAT and so from Equation 8.8: CT = a E A T Pouring temp 730°C Stripping time 60 min Mould hardness 25 /35 AI alloy RR 59 lot 0 2 4 6 8 Water in sand (per cent) 10 12 Figure 8 .23 Residual stress in the centre member of a three-bar frame as a function of water content of the greensand mould (Dodd 1950) (8.9) (8.10) The stress therefore depends... cent Sn alloys by Chakrabarti (20 00) Surfaces of hot torn alloys illustrate the brittle nature of the failure in this alloy (Figure 8.17) The alloy has such a large freezing range, close to 430°C (extending from close to pure A1 at 660°C down to nearly pure Sn at 23 2°C) that in the hot tear ring test the alloy appears even more susceptible to failure by hot tearing than A201 alloy When subjected to the... lengths of beam (see below) 8.3.9 Methods of testing PV AI-Zn u Y 9 V 3 c e 50 0 2 4 6 8 10 Alloy addition (wt per cent) 12 Figure 8 .20 Hut-tearing behaviour of variouA alloj s rubjected to the ring die test Data from Dodd (1955), Dodd et a1 (1957), Punzphrel and Ljonp (1948) and Piimphre\ and Moore (1949) Clyne et al (19 82) later extended the model to the cracking of steels Here, as a consequence of . an A1-4.2Cu alloy has a strength of over 20 0 MPa at 20 "C, that falls to 12 MPa at 500"C, 2 MPa at the solidus temperature, and finally to zero at a liquid fraction of about 20 per. metal through the 25 2 Castings Non-equilibrium freezing range 100 90 - ._ I? 80 P S 70 9 c 60 0 8 50 m c c ._ v - ._ c Q u) u) ._ c 40 2 p 30 20 a, c 10 0. of the thermal contraction and the conditions for continuity. 24 2 Castings Richmond and Tien (1971) and Tien and Richmond (19 82) criticize the model by Weiner and Boley on the grounds that