Materials and Design 31 (2010) 1096–1104

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Materials and Design

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Influence of mould design on the solidification of heavy forging ingots of lowalloy steels by numerical simulation

A. Kermanpur a,*, M. Eskandari a, H. Purmohamad a, M.A. Soltani c, R. Shateri b

a Department of Materials Engineering, Isfahan University of Technology, Isfahan 84156-83111, Iranb Iron and Steel Society of Iran, Isfahan 84156-83111, Iranc Department of Materials Engineering, Islamic Azad University Majlessi Branch, Isfahan, Iran

a r t i c l e i n f o a b s t r a c t

Article history:Received 26 June 2009Accepted 23 September 2009Available online 26 September 2009

Keywords:(G) Numerical simulation(C) Ingot casting(A) Low alloy steel(C) Solidification(C) Hot forging

0261-3069/$ - see front matter � 2009 Elsevier Ltd. Adoi:10.1016/j.matdes.2009.09.045

* Corresponding author. Tel.: +98 311 3915738; faxE-mail address: ahmad_k@cc.iut.ac.ir (A. Kermanp

Ingot casting of a 6-ton, heat-treatable Cr–Mo low alloy steel was simulated using finite element methodin three dimensions. Effects of casting parameters including bottom pouring rate, mould slendernessratio, mould slope, and height and shape of the hot top isolate on solidification behaviour and crack sus-ceptibility during subsequent hot forging of the ingot were investigated. The simulation model was val-idated against experimental data of two different ingot mould designs. Influences of the castingparameters on the riser efficiency and possible crack formation in the intersection of hot top and ingotbody during subsequent open-die forging of the cast steel ingots were discussed. Results showed thatpouring the melt under a constant rate, reducing the mould slenderness ratio, and using a proper designfor the hot top isolate would all improve the riser efficiency and thereby possibly reduce crack suscepti-bility during subsequent hot forging.

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1. Introduction

Today’s forging industry requires a wide range of raw materials,all of which must meet certain standards that limit the quality ofthe semi-finished products. In addition to imparting a certainshape and geometric dimensions, the forging process eliminatesdefects in the initial semi-finished product as it breaks upcoarse-grained dendritic structures and nonmetallic inclusions[1]. Thus, the final product is characterized both by the inheritedmacrostructural nonuniformity of the ingot and by the nonunifor-mity which results from plastic deformation. However, crackingmay occur during hot forging of steel ingots originating from thecast microstructure or unsuitable forging conditions. The intersec-tion between the hot top (riser) and ingot is a critical region inwhich circumferential cracks could form during the primary stagesof forging. The crack then propagates into the ingot and leads tohigh crap formation. Fig. 1 shows a typical circumferential crackthat is formed during the open-die forging of a low alloy steelingot.

The experimental investigation is not always possible andappropriate because it leads to experiments with a lot of parame-ters at complicated circumstances and as a result to great materialexpenditures. That is why a combination of mathematical model-ling and experimental investigations is nowadays acquiring greatersignificance. Attempts have been made by many researchers to

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understand temperature distribution and solidification of large in-gots through computer simulation of ingot casting process usingthe finite element method (FEM). Chernogorova and Vabishchevich[2] investigated the process of the solidification of a binary alloy ina cylindrical metal mould. Tashiro et al. [3] investigated the influ-ence of hot top and mould design on the formation of centralporosities and loose structure in heavy forging ingot (100 and135 ton ingots) by FEM. Gu and Beckermann [4] numerically sim-ulated melt convection and macro-segregation in the casting of alarge steel ingot. Their simulation was based on model for multi-component steel solidification with melt convection and involvesthe solution of fully coupled conservation equations for the trans-port phenomena in the liquid, mush, and solid. Radovic and Lalovic[5] developed two-dimensional (2D) numerical model of ingotsolidification based on the Fourier’s differential equation as wellas energy of lattice defects. On the basis of their numerical modelresults, they calculated temperature distribution, temperature gra-dient, distribution solid and liquid phase and increment of solidfraction. Recently, the casting and solidification processes of large,tool steel ingots were modelled numerically and the ingot shapewas optimized with respect to the real solidification conditions,suppressing the ingot’s internal discontinuities and obtaining anacceptable level of structural and chemical homogeneousness [6].Besides numerical modelling, artificial intelligence methods arealso under development to design and study ingot manufacturingprocesses. A prediction model based on data mining roadmapincluding dynamic polynomial neural network and bootstrapmethod is recently developed by Bae et al. [7]. They collected trace

Fig. 1. A typical circumferential crack formed during the open-die forging of a 6-tonlow alloy steel ingot.

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parameters on-line and measured measurement parameters bysampling inspection. First, statistical methods were used for datageneration, and then modelling was performed, using the gener-ated data, to improve the performance of the models.

The present paper reports development of a 3D numerical mod-el of filling and solidification of a 6-ton low alloy steel ingot usingthe commercial finite element software ProCASTTM [8]. The modelwas validated against experimental data of a real casting shop.The validated model was then used to evaluate the effects of differ-ent casting parameters aiming at eliminating cracking during hotforging of steel ingots.

2. Experimental procedure

The ingot casting operations were carried out in hexahedroncast iron moulds in an industrial alloy steel company. Each mouldwas consisted of three parts including stool (mould entrance), kokil(mould body) and ring (mould riser), as shown in Fig. 2. The insidesurface of the ring was covered by insulation material with 25 mmthickness. Chemical compositions of the low alloy steel and thecast iron mould are listed in Table 1. The steel melt was bottompoured at temperature 1600 �C and flow rate of about 7 kg s�1 inthe 200 �C preheated mould through stool. When the melt level

Fig. 2. Top view of the components of a 6-ton steel mould (P64).

was reached to the hot top, exothermic material was distributedover the melt free surface. Different height to diameter (slender-ness) ratios of the mould, riser shapes and pouring regimes wereinvestigated. Two ingot moulds were cast under different condi-tions: P59 with the slenderness ratio of 1.5 and melt capacity of5900 kg, and P64 with the slenderness ratio of 2.3 and melt capac-ity of 6400 kg. Table 2 shows the experimental conditions investi-gated in the simulations.

3. Numerical simulation

The filling and solidification stages of the ingot casting weresimulated by the finite element software ProCAST in 3D. A coupledflow-thermal–mechanical model was set up in the software basedon the Navier–Stokes equations of fluid flow, Fourier’s equation ofheat flow and stress calculations for gap formation [9]. The k–emodel was used for the turbulent flow calculations. The enthalpymethod was used for applying the latent heat release during solid-ification. The stress module was activated in the software in orderto consider the effect of gap formation in the metal/mould inter-face on solidification. The stress calculations were started as soonas the fraction of solid was larger than a critical fraction definedby the user (say, 0.5).

Effect of gap formation in the melt/mould and melt/insulatorinterfaces is very crucial on solidification calculations. Initial valuesof the interfacial heat transfer coefficients between melt/mouldand melt/insulator were assigned as 600 and 100 Wm�2 K�1,respectively [8]. This heat transfer coefficient is a function of ferro-static pressure of the melt given by the following equation:

h ¼ h0 1þ PA

� �ð1Þ

where h0 is the initial value of the heat transfer coefficient, P is thepressure of melt and A is the empirical constant to account for con-tact pressure. During a thermo-mechanical calculation, gaps mayform between the different domains (e.g. between the casting andthe mould). ProCASTTM automatically accounts for the modificationof the interface heat transfer coefficient when gaps are formingwhich is given by equation:

h ¼ 11

h0þ Rgap

; Rgap ¼1

kgapþ hrad

ð2Þ

where k is the conductivity of air, ‘gap’ is air gap width and hrad isradiative equivalent heat transfer coefficient [8].

The finite element mesh of mould parts and ingot is shown inFig. 3, consisting of 22,690 nodes and 103,516 tetrahedral ele-ments. This mesh was selected based on the mesh sensitivity anal-ysis performed for several mesh refinements. Fig. 4 shows theboundary conditions applied to the model. As it can be seen, an in-let condition was assigned at the mould bottom for entering themelt into the mould; an isolation condition was set over the meltsurface at the riser as the melt was covered by the insulation mate-rials; a natural convection was assigned over the mould body withthe environment around the mould. Using the simulation model,fourteen different simulation runs were conducted to investigatethe effects of processing parameters as listed in Table 2.

4. Results

4.1. Effect of slenderness ratio

The filling and solidification sequences of the P64 and P59 ingotmoulds with the melt flow rate of 7 kg s�1 are shown in Figs. 5 and6, respectively (e.g. experiments # 1 and 2 in Table 2). Based on thetechnology used in the steel plant, the melt flow rate was de-

Table 1Chemical compositions of ingot and mould.

Material Code Alloys elements

C (%) Mn (%) Si (%) Cr (%) Mo (%) Cu (%) Mg (%) S (%) P (%)

Melt Alloy steel 0.4 0.7 0.3 1.10 0.2 – – 0.03 0.035Mould GGG60 3.53 0.34 2.88 – – 0.64 0.05 0.06 0.02

Table 2Experimental conditions used for the simulations.

Experiment # Ingot mould Mould slenderness ratio Mould slope (�) Pouring rate (kg s�1) Hot top

Mould Hot top Insulation height Shape

1 P64 2.3 2.8 7 3.5 Full Circular2 P59 1.5 2.8 7 3.5 Full Circular3 P59L 1.1 2.8 7 3.5 Full Circular4 P64H 2.8 2.8 7 3.5 Full Circular5 P64 2.3 2.8 7 7 Full Circular6 P64 2.3 2.8 7 21 Full Circular7 P59 1.5 2.8 7 7 Full Circular8 P59 1.5 2.8 7 21 Full Circular9 P64 2.3 2.8 7 7 Partial Circular

10 P59 1.5 2.8 7 7 Partial Circular11 P64 2.3 2.8 7 3.5 Full Polygonal12 P59 1.5 2.8 7 3.5 Full Polygonal13 P64 2.3 10 7 7 Full Circular14 P59 1.5 10 7 7 Full Circular

Fig. 3. Finite element mesh of the mould parts and ingot.

Fig. 4. The boundary conditions of the model.

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creased from 7 kg s�1 to half (e.g. 3.5 kg s�1) when the melt freesurface touched the hot top (see Figs. 5c and 6c). This was initiallythought to be useful for improving the performance of the mouldriser.

Table 3 shows a reasonable agreement between the measuredpouring and solidification times with the simulated ones. It canbe seen that the filling time for the two ingots is different (e.g.1064 s for P64 vs. 1014 s for P59). It can be noticed that more fillingtime was achieved for the larger mould slenderness ratio. Further-more, comparing Fig. 5e with Fig. 6e reveals that more solidifica-tion was taken place in the riser for the mould P64, when

pouring is completed. Table 3 also shows that the total solidifica-tion time for the mould P64 is more than that of mould P59 (e.g.12,555 vs. 11,285 s, respectively). Therefore, the solidification timewas increased by increasing the mould slenderness ratio.

In order to investigate the effect of slenderness ratio on direc-tionality of the melt solidification, the solidified shell thickness atthe mid-height of the mould for four different slenderness ratioswere evaluated. These include the P59 mould, one mould with aslenderness ratio lower than P59 (e.g. 1.1 vs. 1.5) called P59L, theP64 mould, and one mould with a slenderness ratio higher thanP64 (e.g. 2.8 vs. 2.3) called P64H. The experimental conditions ofthe new moulds are shown in Table 2 as experiments # 3 and 4,

Fig. 5. Distribution of simulated solid fraction during filling and solidification of P64 ingot at: (a) 35 s, (b) 320 s, (c) 760 s, (d) 910 s, (e) 1065 s, (f) 3320 s, (g) 4560 s, (h) 7775 s,and (i) 12,555 s.

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respectively. Fig. 7 shows variation of the solidified shell thicknessat the mid-height of the moulds versus time. As diameter of themoulds was different, the normalized values of the shell thicknesswere processed. Also, as the filling time of the moulds was not thesame, the reference time was the time at which each mould wasjust filled. It can be seen that the transverse solidification at themid-height of the moulds are approximately similar for the fourmoulds at the beginning, but it becomes different with the progres-sion of solidification. Less transverse solidification is taken placewith a lower slenderness ratio.

4.2. Effect of pouring rate

Effect of the melt pouring rate on the solid fraction distributionin the ingot P64 is shown in Fig. 8 (e.g. experiments # 1, 5 and 6 inTable 2). Three different pouring rate regimes were used for thefilling of the hot top: (a) 7 kg s�1 in the mould and 3.5 kg s�1 inthe hot top; (b) constant pouring rate of 7 kg s�1, during fillingthe whole mould; and (c) 7 kg s�1 in the mould and 21 kg s�1 inthe hot top. The simulation results show that increasing the meltpouring rate in the hot top resulted in decreasing the amount of

Fig. 6. Distribution of simulated solid fraction during filling and solidification of P59 ingot at: (a) 35 s, (b) 287 s, (c) 703 s, (d) 859 s, (e) 1014 s, (f) 2630 s, (g) 4770 s, (h) 7830 s,and (i) 11,285 s.

Table 3Experimental observations of the filling and solidification times (in seconds) versusthe simulated ones for two moulds P64 and P59.

Mould P64 P59

Data Experimental Simulated Experimental Simulated

Filling time (s) 1080–1140 1064 960–1080 1014Solidification time

(s)18,000 12,554 15,600 11,284

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solid formed in the hot top during the pouring operation. This wasclearly opposite to the technology used in the steel plant based onwhich the pouring rate was decreased to half in the hot top. Fig. 9shows cooling curves of two different locations in the hot top while

using both constant and reduced pouring rates. It can be seen thatincreasing the pouring rate in hot top led to slower cooling.

4.3. Effect of height of isolate in the hot top

Distribution of the simulated solid fraction during solidificationof P64 ingot using a complete and a partial isolation material in thehot top is shown in Fig. 10 (e.g. experiments # 5 and 9 in Table 2).In Fig. 10a, the hot top is fully covered by the isolating material,whereas one third of the isolating material is removed from thebottom part of the hot top in Fig. 10b. As expected, it can be seenthat decreasing the isolate height resulted in decreasing the solid-ification time. Moreover, thickness of the solid formed on the bot-tom part of the hot top is the same as the mould region for the hottop with the partial isolating material.

Fig. 7. Normalized solidified shell thickness at mid-height of four moulds withdifferent slenderness ratios in terms of time after pouring is completed.

Fig. 9. Simulated cooling curves for two different locations in the hot top of P64ingot poured at constant and reduced rates.

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4.4. Effect of shape of isolate in the hot top

The solid fraction distribution in the P64 ingot during solidifica-tion with two different isolate shapes is shown in Fig. 11 (e.g.experiments # 1 and 11 in Table 2). It can be seen that the ingotwith the polygonal hot top is cooled faster than the one with thecircular one due to the more contact area in the polygonal shape.In addition, the polygonal hot top is characterized with a deepershrinkage at the ingot top.

4.5. Effect of mould slope

Effect of the mould slope on the solidification time of P59 in-got with different slopes is shown in Table 4 (e.g. experiments# 7 and 14 in Table 2). As shown, increasing the mould slopeleads to increasing solidification time. It seems that, increasingtop diameter of the ingot with the higher slope, leads to decreas-ing heat transfer in radial direction and increasing the solidifica-tion time. In addition, a higher mould slope led to a more verticalsolidification (Fig. 12). It should be noted that increasing themould slope is not practically interested, as it increases the num-ber of forging passes during the initial stages of the open-dieforging operation.

Fig. 8. Distribution of simulated solid fraction when pouring is completed for the mould P(b) constant 7 kg s�1, (c) 7 kg s�1 in the mould and 21 kg s�1 in hot top.

5. Discussion

From the point of view of casting efficiency, it is necessary todesign a suitable hot top such that melt solidification is preferablytaken place directionally from the mould bottom to the hot toprather than transversely from the wall. Therefore, vertical solidifi-cation of the melt in the mould should be facilitated, whereastransverse solidification from the walls of hot top towards the in-got centre must be eliminated [10]. On the other hand, as the ther-mal conductivity of the hot top materials is much lower than thatof the mould material, the cooling effect is much lower in the hottop and consequently, a much coarser dendritic structure will beformed in the hot top. It is well known that mechanical propertiesand high temperature workability of the cast microstructures arecrucially depended upon the dendrite arm spacing of the dendriticstructure [11]. It can therefore be concluded that workability of thesolid formed in the hot top would be lower than that of the mouldregion.

The above-mentioned thermal simulation results clearlyshowed that among the different process parameters of the ingotcasting of low alloy steels, the slenderness ratio, pouring rate andisolate material had significant effect on solidification of the lowalloy steel in hot top. Simulation results of the filling and solidifi-cation times of the P64 and P59 moulds with different pouringrates are shown in Table 5. As shown, increasing the pouring ratein hot top resulted in increasing solidification time of the ingotand improving directional solidification pattern (Fig. 8). In fact,

64 under different pouring rates: (a) 7 kg s�1 in the mould and 3.5 kg s�1 in hot top,

Fig. 10. Distribution of simulated solid fraction during solidification of the P64 ingot using (a) complete and (b) partial isolate in the hot top.

Fig. 11. Distribution of simulated solid fraction during solidification of the P64 ingot using (a) circular and (b) polygonal hot top.

1102 A. Kermanpur et al. / Materials and Design 31 (2010) 1096–1104

increasing the pouring rate in the hot top resulted in less time forheat transfer during filling hot top and thereby less solid is formed

in the hot top. This would increase the total solidification time asexpected. The simulation results showed that axial solidification

Table 4Effect of mould slope on solidification time of P59 ingot.

Mould slope (�) 2.8 10Filling time (s) 833 828Solidification time (s) 11,360 14,397

Fig. 13. Photograph of a 6-ton low alloy steel ingot during the open-die forging thesimilar ingot. Note that no crack is formed in the intersection of the hot top (right)and ingot (left).

A. Kermanpur et al. / Materials and Design 31 (2010) 1096–1104 1103

is improved by increasing the pouring rate at hot top, whiles thelateral solidification is minimised. In other words, directional solid-ification will be improved by increasing the pouring rate. Sincecooling rate is much higher in the mould region compared to hottop, the cast microstructure would be finer with increasing thepouring rate. It is expected that this microstructure possesses bet-ter workability during the subsequent open-die forging of the in-gots. Fig. 13 shows the photograph of the same ingot shown inFig. 1 that is successfully hot forged and no crack is formed duringforging. This ingot was cast under constant pouring rate of 7 kg s�1

(i.e. the pouring rate in hot top is double compared to the one inFig. 1). It confirms that increasing pouring rate in hot top of P64 in-got resulted in reducing crack susceptibility in the critical intersec-tion of hot top and mould ingot.

According to the simulation results shown in Fig. 7 it was foundthat lowering the slenderness ratio of the mould could improvedirectional solidification pattern in the mould (e.g. lower trans-verse shell thickness at the mid-height of the mould as shown inFig. 7) and therefore it may reduce the crack susceptibility. Itshould be however noted that, on the other hand, as the cast ingotsare subsequently subjected to hot forging operation, more forgingswould be needed for the ingots with a lower slenderness ratio. Infact, although a lower slenderness ration is preferable in terms of

Fig. 12. Distribution of simulated solid fraction during solidificati

directional solidification, but the ratio should not be such low thatthe subsequent hot forging become huge and uneconomical. Thismay also increase the crack susceptibility of the ingots. Therefore,the P59L design is not practically interested, as it increases thenumber of forging passes during the initial stages of the open-dieforging operation, and the P59 mould is more preferable.

In practice it was found that when the 200,000 kg ingot (P200)was cast with a reduced height of hot top isolate, no crack was

on of the P59 ingot with mould slope of (a) 2.8� and (b) 10�.

Table 5Simulation results of filling and solidification times of P64 and P59 ingots.

Experiment # Ingot Pouring rate(kg s�1)

Time (s)

Mould Hot top Mould filling till hot top Hot top filling Total pouring Solidification after ingot filling Total ingot solidification

1 P64 7 3.5 759 305 1064 11,490 12,5542 P59 7 3.5 703 311 1014 10,270 11,2845 P64 7 7 759 149 908 11,902 12,8106 P64 7 21 759 49 808 12,077 12,8858 P59 7 21 703 43 746 10,754 11,500

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formed in the intersection of hot top with the ingot during forging.In other word, removing the insulating material from the bottom ofhot top made it possible for the melt in hot top to solidify similar tothe mould region. Therefore, solid formed in the intersection of hottop and ingot represented similar workability and no crack wasformed there. It was noticed that under this condition, the crackformation was shifted up to where the insulation wall was in-stalled over the hot top wall.

This work clearly showed that numerical simulation of steel in-got casting can be a valuable design tool to investigate the effectsof casting parameters on the solidification behaviour and sound-ness of the heavy cast ingots that are subsequently subjected tohot forging. In order to produce sound ingots after hot forging, itis recommended to adjust casting parameters such that constantpouring rate during casting of the whole ingot, lower slendernessratio of the mould, and proper hot top riser are used, all of whichwould enhance vertical solidification rather than transversesolidification.

6. Conclusions

1. Pouring the melt under constant rate in the mould with a lowerslenderness ratio is preferred to enhance vertical solidificationand to reduce transverse solidification in the hot top region.This would lead to the higher riser efficiency.

2. Height and shape of the insulating material in the hot top influ-enced the solidification behaviour. The circular cross section forthe hot top is preferred.

3. Increasing the mould slope led to increase solidification timeand better directional solidification, but it was not practicallyinterested.

4. Circumferential cracks in the intersection of the mould hot topand the ingot body may be formed during the primary stages ofopen-die forging of cast steel ingots due to the selection ofunsuitable parameters of the previous casting operation.Improving vertical solidification in the mould may reduce cracksusceptibility during hot forging.

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