Development of a Combined Bulging-Piercing Technique to Reduce Forming Load for a Long Semi-Hollow Stepped Part

This paper aims to develop a new forming technique to manufacture a long semi-hollow stepped part. Traditionally, hot backward extrusion is used. This technique is not suitable, because it requires a very high forming load acting on the die and punch especially at the contact between punch and workpiece. As a result, the service life of the punch is very low. Therefore, a new technique to overcome this problem is needed. A combined bulging-piercing technique was proposed and developed in this research. The main concept of this technique is to bulge the part by upsetting the workpiece between the punch and the counter-punch to generate high frictional contact pressure which will help to restrain the material sliding down to the die cavity during the piercing step. In other words, this technique utilizes frictional force at the die-workpiece interface to reduce the forming load of the punch. Finite element modeling was employed to investigate and determine the suitable level of the bulging which can reduce the forming load without generating any signicantly high force to the counter-punch. Only experiments with the minimum forming load were selected and implemented to validate this concept, because other conditions with high load will risk to damage the punch and the machine press of the product line. The results show that this technique can reduce the forming load by almost 40%, and also control a good concentricity of the part and reduce the wall thickness variation.


Introduction
The success of the forging processes requires high tonnage of the press machine and high strength/toughness of tools/dies to deform the materials to the required shape or dimension. These result in high production costs and limits in manufacturing some of the part con gurations.
To reduce the production cost, the reduction in the forming load and the tooling costs is a big and major challenge for engineers and researchers nowadays. In other words, the lower forming load could enhance the service life of the forming tools and lower press machine capacity. Many techniques were developed and introduced to overcome these problems by focusing on the design of billets, toolings, and processes.
Many techniques for the forming load reduction are reviewed and some of them are instanced [1]. Relief hole and axis is a method of the divided ow technique proposed to reduce the forming load in a cold forging process of gears proposed by Kondo et al [2].The material ow is split into 2 parts to ll the tooth tips located at the outer gear teeth and/or ow either radially inwards toward the center axis or axially within a central hole simultaneously [3]. In addition, the optimization of the preform shape and modi cation of the nal forged shape is a technique that can reduce the forming load and also facilitate the die lling, improve material yield, and eliminate forging defects. A modern forging process, which is the combination of the advantages of sheet forming and cold forging, was introduced, namely Flow Control Forming (FCF) [4] or Sheet-bulk metal forming (SBMF) [5]. The upsetting, ironing, and extrusion process are integrated into the conventional sheet metal forming which is typically the sheet forming consisting of the blanking, drawing, and bending processes. Therefore, the sheet material is forced to change shape at all 3-dimensions similar to the material ow in the bulk forming process. It is mentioned that this FCF or SBMF is mainly to reduce the material waste as well as the forming load [6]. Furthermore, a technique for reducing the number of toolings for producing a similar product by only the few dimension differences, namely Common Single-Die Exchange Technique or C-SDET, is proposed [7]. The idea is to determine the representative preform that could be used to produce various product models. In other words, this design aims to develop common generalized preforms by using the same component con guration/shape but varying the dimensions in some places to decrease the number of preforms and tooling installation time (downtime of the machine to install the tooling for each model).
The design of the optimum preforms and tooling shapes requires speci c knowledge and past experience of the engineers which have been done for particular products. The optimum preform shape can lead to defect free parts with minimum required tonnage and waste material [8]. Currently, the computational algorithms are developed and applied to design the optimum preform shape of very complex parts, such as the turbine blade. For example, the backward tracing algorithms together with the advanced nite element simulation were developed to determine the optimum intermediate shape in a shell nosing process to achieve the uniform wall thickness of the nosing part [9]. Furthermore, sensitivity analysis based algorithm of the optimum preform design for single stage forming process was developed. The initial shape of the billet is determined by measuring the sensitivity of design criteria, billet height, and width on the nal shape of the forging to achieve complete die lling [10].
Currently, the tendency to reduce the weight of the machine elements is a big challenge. The weight reduction provides important bene ts, especially for the cost reduction, but the parts must maintain good mechanical properties, lifetime service, and reliability. Therefore, achieving the lightweight components can be practically done in two ways; a) utilizing lightweight materials, such as aluminum alloys or magnesium alloys, and b) optimizing part design, such as hollow design [11].
The utilization of the hollow components often results in material and energy savings throughout the manufacturing process. Metalworking procedures are used to create hollow machine components in which the workpiece is hollowed out over its whole length (with some stock allowed for nishing). This strategy provides practical advantages in addition to the economic bene ts.
In such a case of a semi-hollow stepped shaft, as shown in Figure 1, the hot backward extrusion is conventionally applied to produce such part [12]. This technique provides various advantages in terms of product quality, manufacturing rate, and cost. However, higher forming loads, lubrication costs, limiting shape complexity are the disadvantages of the backward extrusion [13]. Moreover, the backward extrusion is not commonly employed at the high temperature, because the hot backward extrusion process causes the decrease of the tool life especially at the punch corners because of the high contacting stress [14].
The forming load and the uniformity of the wall thickness are also concerned mostly in the backward extrusion. It is mainly controlled by the punch pro le. The in uence of the punch face slope and the punch llet radius on the lateral and axial force was investigated in [15]. The results show that by decreasing the punch llet radius and the punch slope, the lateral force is reduced. In contrast, the smaller punch llet causes the increase of the axial force. Therefore, the proper punch geometry needs to be determined. The uniformity of the wall thickness between straight and circular punch land was investigated by Danckert (2004) [16]. The different length of the punch land causes in changing the contact conditions between the punch land and the cup wall. It results in the punch off-center and variation in the cup wall thickness. Therefore, the utilization of the circular punch land reduced the maximum wall thickness variation by almost 18%. Moreover, the wall thickness uniformity is also affected by the elastic de ection of the die and container including the buckling of the punch [17]. It was recommended that the punch should be short as much as possible, and/or the forming load should be reduced below the buckling threshold of the punch during the backward extrusion to prevent buckling [18]. However, in some forging parts, the internal shape is controlled by the punch pro le which depends on the geometry requirement. It is di cult to avoid or change such cases in some parts, especially for a long semi-hollow part. Therefore, the forming technique to overcome this problem needs to be developed without signi cantly modifying the internal punch shapes.
As seen in Figure 1, this part normally was manufactured by the hot backward extrusion. The preliminary simulation results obtained by the FEM show that the maximum forming load is about 1,500 tons, as seen in Figure 2. This such a high forming load could lead to some damages on the forming tools and the press machine. Therefore, this research aims to propose a new technique for producing the semihollow stepped shafts, namely, a combined bulging-piercing technique. It is composed of two main steps; bulging, and piercing-coining steps. The combined bulging-piercing technique is mainly developed to reduce the forming load in manufacturing of the long semi-hollow stepped shafts ( Figure 1) and still maintain the concentricity of the parts. Three main process parameters, namely the bulging stroke (Sb), the counter-punch lifting displacement (Lc), and the friction value (m), were investigated to determine the effect on the forming load and the die lling by the FEM simulation. Only the case with the minimum forming load was selected to implement experimentally to avoid any damage which might happen in the production press and tools. Then, the die lling, wall thickness measurement, and macro etching ( owline analysis) were performed to validate the design concept and outcomes, outcomes of the load reduction, and accurate geometries.

Mechanism of the combined bulging-piercing technique
The main functions of a combined bulging-piercing technique are to reduce the forming load, control the concentricity of the products, and reduce the side wall thickness variation, especially for forming the long stepped semi-hollow parts.
A schematic of this technique is shown in Figure 3. The main idea of this technique is to employ the counter-punch and the friction at the die-wall interface to control the material ow into the die cavity. Initially, the counter-punch is lift up from the counter-punch reference position by the distance, Lc. The workpiece/billet is placed on the counter-punch, and then the punch is moved down until both are in-contact. The initial height of the billet is de ned as Ti. The stroke that the punch and the workpiece are incontact without deformation is de ned as the punch reference position. Then, the punch is travelled down while the counter-punch is maintained at the reference position to upset the billet to create a bulging. The purpose of the bulging is to generate the contact pressure between the workpiece and the die. The friction at the contact will restrict the material sliding down into the die cavity during piercing. During the bulging step, the punch travels from the punch reference position and then stops at a stroke, Sb, as seen in Figure  3, whereas the counter-punch remains at the initial position. The workpiece is deformed mostly only at the radial direction and forward extrusion with a little or no backward extrusion. Consequently, the workpiece diameter gradually increases and then contacts with the die wall, as seen in Figure 4. With increasing the bulging stroke, the contact area between the workpiece and the die is increased. The forming load of this bulging step slightly increases in the rst period and becomes more prominent, when the workpiece comes into contact with the die, as seen in Figure 5.
As demonstrated in Figure 3, the counter-punch immediately moves down to the counter-punch reference position after the bulging is complete. Then, the punch moves down with the traveling stroke, Sp, to pierce the workpiece as shown in Figure 4. In the piercing step, there are two forming modes; (i) the forward extrusion and (ii) the lateral extrusion, to create the hollow shape and control the concentricity of the workpiece. In Figure 4, with some evidence of the material ow velocity, the material ow during the piercing is aligned with the punch direction. While the workpiece is elongated along the wall, the bottom thickness is also reduced by material owing outward radially. The forming load increases at the beginning and then is roughly constant until the end of the piercing stroke, as seen in Figure 5.
The coining is the last step required to control the nal bottom thickness (Tc) and ful lled the die cavity. It must be noted that even though the nal bottom thickness can be attained with force increasing exponentially, the material may not be completely lled at the bottom die. However, the maximum forming load may occur in one of the two following cases; a) with the de ned nal bottom thickness and b) completely ful ll at the bottom die.
Thus, the total forming stroke, in Eq. 1, is composed of a) the bulging stroke, b) the piercing stroke, and c) the coining stroke, and can be controlled by two conditions; a) a constant stroke which may control the constant bottom thickness and b) a variable stroke for the material is fully lled at the bottom die.
For the constant stroke control, the counter-punch lifting displacement is xed and the bulging stroke is only a variable parameter. The increase of the bulging stroke, not affect the maximum forming load, would mainly increase the piercing stroke but the coining stroke remains constant. However, the nal product with this condition would have the bottom thickness as designed, but the material somehow not be entirely lled at the bottom die.
In contrast, the variable total forming stroke depends on the varied counter-punch lifting displacement, while the bulging stroke is constant. Therefore, the counter-punch lifting displacement variation would results mainly in changing the total forming stroke, as seen in Eq. 2. By this role, the total forming stroke increase corresponds to the increase of the piercing stroke and the decrease of the coining stroke. Hence, the maximum forming load decreases due to the less coining stroke and the more elongated workpiece.
As a result, when the material is completely lled at the bottom die, the nal bottom thickness would be thinner than expected.
For that reason, to achieve the forging parts with the designed thickness and fully lled shape at the bottom die, the impacts of the process parameters need to be studied

Process parameters
The FEM simulations were performed to investigate three main parameters, i.e. (i) the bulging stroke (Sb), (ii) the counter-punch lifting displacement (Cp), and (iii) the friction value (m). All the conditions are summarized in Table 1. In addition, the forming load predictions for each condition were concluded for comparison to nd the applicable condition in the process.
Initially, the workpiece-die contact surface rst touches at the punch stroke of 160 mm. Further, the bulging strokes were varied by 160, 180, 200, 220, and 240 mm, respectively. Then various bulging strokes proceeded continually with the piercing and the coining steps.
The initial position of the counter-punch lifting displacement is at 110 mm above the reference or zerodisplacement of the counter-punch lifting, as shown in Figure 6. Then, the position change was 20 mm for every incremental step. Hence, the lifting displacement of the counter-punch was 90mm, 110 mm, and 130 mm in this numerical study. Eventually, the effect of various bulging strokes and the initial position of the counter-punch on the forming load and the bottom thickness of the workpiece were analyzed.
Here, the friction effect to the forming load was also investigated by using the friction coe cient of 0.5 and 0.7, estimated by the graphite mixture ratio of 5% and 15%, respectively, according to [19]. Speci cally, the bulging stroke was 200 mm, and the counter-punch lifting displacement was 110 mm for studying this friction effect.
After the investigation, one condition that provide possibly lower forming load, would be selected to verify and implement experimentally.

Finite element modeling
The non-isothermal FEM modeling was performed by using a half symmetric model to reduce the computational time, as seen in Figure 7. The 300,000 elements were applied to the workpiece. The forming tools are assumed as the elastic body with the tool steel properties for the hot work, AISI H13. The material property of the workpiece is a function of the temperature range (800 to 1,100 C) and the strain rate range (1.6 to 40 s −1 ) [20], as shown in Figure 8. The chemical compositions of the workpiece are tabulated in Table 2. The heat transfer coe cient was assumed as 7 N/sec/mm/˚C and the initial temperature of the punch and die was assumed at 250 ˚C for this hot forging process simulation. The forming speed was controlled by punch velocity of 100 mm/sec.  Figure 9 shows the load-stroke curve of the counter-punch and the punch during the bulging. During the rst period, the punch load slightly increases before 140 mm of the stroke. After that, the load increases with smoothly changed the slope until the stroke reaches 160 mm. Then, the roughly constant slope of the loading curve can be noticed where the forming load increases linearly due to the higher contact area between the workpiece and the die wall. It is noted that orange and green dots indicate the area of the punch-workpiece contact and the die-workpiece contact, respectively.
The forming loads in each bulging stroke (160, 180, 200, 220, and 240 mm) continued with the piercing and coining steps were explored. The results in Figure 10 shows that the increase in the bulging stroke does not signi cantly affect on the maximum load. It is almost constant, around 1,500 tons. Overall, the piercing load increases almost 20 percent with increasing the bulging stroke, as seen in Figure 11 and Figure 11. Additionally, the increase in the bulging stroke dramatically causes the reduction in the piercing stroke, but the coining stroke remains constant. It is due to the limited total forming stroke. The increase of the forming load within the piercing step is because the di culty of the material to ow in both axial and lateral directions. At 160 mm of the bulging stroke, the forward and the lateral extrusion occur concurrently from the stroke of 160 mm to 350 mm. After that, the lateral ow reduces. The material only ows in the axial direction until the forming stroke is 400 mm, as seen in Figure 13. On the other hand, the forward and lateral ow in the bulging stroke of 240 mm occurs during the stroke from 240 mm to 350 mm, as seen in Figure 14. As a result, the stroke that the material is allowed to ow in the axial and the lateral directions, is shorter than the others. Therefore, it leads to a higher piercing load.
The workpiece bottom thickness of various bulging strokes was measured during the piercing. Figure 15 demonstrates that the increase in the bulging stroke causes a reduction in the bottom thickness. The maximum deviation around 20 mm occurs at the stroke of 250 mm.
The increase of the forming load and the reduction of the nal bottom thickness are plotted in Figure 16. Figure 17 illustrates the material lling into the bottom die cavity in various forming strokes. It shows that when the material is in contact with the bottom die cavity at the forming stroke of 410 mm, the forming load is increased sharply. Then, the bottom die cavity is completely lled at the stroke of 420 mm with the bottom thickness of 63 mm. At last, as designed, the nal bottom thickness is achieved at the forming stroke of 422 mm.
In conclusion, when the total forming stroke is constant, the increase of the piercing stroke relies mainly on the increase of the bulging stroke but it does not lessen the maximum forming load. Therefore, the maximum forming load remains constant. The complete lling at the bottom die cavity is attained before the nal bottom thickness as designed. Therefore, the punch needs to compress the workpiece further to ful ll the bottom die, resulting in a high forming load.

Effect of the counter-punch lifting displacement
The bulging load of various counter-punch lifting displacements is shown in Figure 18. The results show that the deviation of the counter-punch lifting displacement does not cause any signi cant change in the forming load of bulging. Contrarily, it signi cantly impacts the piercing and the coining loads, as seen in Figure 19. The piercing stroke signi cantly decreases for the counter-punch lifting displacement at 90 mm while the coining stroke increases with the extremely high forming load. The piercing stroke extends when the counter-punch lifts up to 110 mm and the coining stroke shortens. However, the maximum forming load remains unchanged. Therefore, the forming part of these two cases seems to be superior, without defects. The most extended piercing stroke is obtained in 130 mm of the counter-punch lifting displacement, and the maximum forming load is minimized from about 1,300 tons to almost 320 tons. Nonetheless, the under lled defect occurs at the bottom of the workpiece.
In such a case of the workpiece bottom thickness, it can be seen that the higher counter-punch lifting displacement dramatically reduces the workpiece bottom thickness, as seen in Figure 20. However, the 130mm of the counter-punch lifting displacement provides the maximum reduction in the workpiece bottom thickness with the minimum forming load. The designed nal bottom thickness and the die lling are not achieved yet.
In conclusion, the maximum forming load increases when the total forming stroke shortens with the total forming stroke varied. To lower the maximum forming load, the total forming stroke increases with increasing the counter-punch lifting displacement. However, a higher total forming stroke would allow some defects; the under lled and the thinner at the nal bottom thickness.

Effect of the friction coe cient
The forming load in the bulging step for various friction coe cients is illustrated in Figure 21. It reveals that the increase of friction coe cient mainly increases the maximum load of the forming punch. The maximum load increases approximately 20 percent, but the counter-punch load is relatively constant. Figure 22 displays the forming load during the piercing and the coining steps for various friction coe cients. The results show that the overall forming load in the piercing increase about 20 percent, but, nearing the end of the stroke, the maximum load dramatically increases from 440 tons to 955 tons, or almost 50 percent.
In conclusion, it is evidence that the friction affects on the forming load, especially at the maximum forming load during the coining. Therefore, the high-performance lubricant is required to reduce the forming load and prevent the overload of the press machine capacity.

Experimental Results
The manufacturing process of the semi-hollow parts is illustrated in Figure 23. Firstly, the initial billet is cut by sawing machine. The square billet was selected in this study because it was available with the required billet size and avoided any initial crack at each corner due to the manufacturing process. The square cross-section of the billet is 140x140 mm and the length is 359 mm, as shown in Figure 24. Further, the rounded edges with radius of 5 mm were treated at each billet corners. Then, the billets were heated up to 1,250 ˚C by the induction heater. Before conveying to the press machine, the heated billets are passed through the high-pressure water tunnel to remove the oxide scale on their surface. For the hot forging process, the 1,250 tons hydraulic press machine is used for this production. This machine also has a cushion with 150 tons load. The water-based graphite with a ratio of 10 percent is utilized for lubrication. The forming tools are made of hot working tool steel (AISI H13) with surface hardening and nitriding surface coating.

An implementation of the combined bulging-piercing technique
The experimental conditions were selected according to the simulation results. The selection criterion was to obtain the superior forging part (complete lling and the nal bottom thickness as designed) with the lower maximum load. The experiment was replicated at least ten times. Those conditions are as follows.; The bulging stroke was xed constant at 200 mm. The counter-punch reference position was de ned to the same position as the FE modeling position and the counter-punch lifting displacement was 110 mm. The counter-punch load was limited to 70-80 tons to avoid any such as buckling and/or bending of the counter-punch. The punch and counter-punch loads were measured while the ram speed was 100 mm/sec.
However, the precise positioning and the constant forming speed in a hydraulic pressing machine are hard to control due to the precision of the control system. In this case, the counter-punch lifting was retracted to be lower than the de ned position during the bulging. Even if using the PLC control, the accurate positioning control is still unsatisfactory. Therefore, to con rm that experimental results were reliable. The experiments with and without the compensated displacement of the counter-punch lifting were conducted. Moreover, the retraction of the counter-punch were monitored for every replication.
In the rst experiment, the loading responses of the punch and the counter-punch (during the bulging) are shown in Figure 25 (red and green lines, respectively). The maximum forming load of the punch and the counter-punch approximately are 140 tons and 68 tons, respectively. The punch and the counter-punch loads obviously increase at the stroke of 850 mm, where the die-workpiece contact occurs. It is jumped up to almost 100 tons in case of the punch. The counter-punch load is also signi cantly increased from about 10 tons to almost 60 tons. Then, the forming load of the punch and the counter-punch is gradually increased until the end of the stroke. Later on, the forming load of the punch is quite constant from the ram stroke of 1,000 to 1,100 mm. After that, it is jumped up rapidly to the maximum forming load which is almost 700 tons.
It was found that the counter-punch could not hold in the position during the bulging. The counter-punch retraction during the bulging step is shown in Figure 26 and the average counter-punch retraction is shown in Figure 27. Initially, the counter-punch is lifted by 110 mm, but, at the end of the bulging stroke, the counter-punch is retracted by almost 6 mm. Therefore, it is lower from 110 mm to almost 104 mm.
This causes the lower de ned counter-punch lifting displacement and the higher maximum forming load. As a result, the compensation of the counter-punch lifting displacement was performed in the second experiment. The compensation of the counter-punch lifting displacement was utilized for preventing the over-retraction by 7 mm. Therefore, it was increased from 110 mm to 117 mm.
In the second experiment, the observed forming load is shown in Figure 25. The bulging load is quite similar to the rst experiment which is almost 140 tons and 67 tons for the punch and the counter-punch. During the bulging step, the counter-punch is still retracted which is about 4 mm, as seen in Figure 26 and Figure 27. It was moved down from 117 mm to 113 mm. However, the last position of the counter-punch is still higher than 110 mm. Therefore, the maximum forming load during the piercing and the coining is dramatically reduced by almost 40% that was lowered from almost 700 tons to 400 tons The owline of the workpieces obtained from the bulging and the piercing and coining are shown in Figure 28 and Figure 29, respectively. It is evident that the workpiece's owline is aligned with the shape of the workpiece without collapsing. As a result, the workpiece is strengthened. The desired nal bottom thickness and the completely lled are also achieved.
Furthermore, the wall thickness of the workpiece was measured in different positions, as seen in Figure  30(a), and the thickness variation is shown in Figure 30(b). It can be seen that by employing this forming technique, the maximum wall thickness variation of the workpiece is about 5.5%. With this wall thickness variation, the workpiece is acceptable for proceeding to the nal machining.

Comparison of the experimental results with the simulation results
The predicted forming loads obtained by the FEM simulation in various friction factors is compared with the forming load the experiment in various counter-punch lifting displacements, as shown in Figure 31. The results show that the predicted forming load during the bulging and piercing region is higher than that of the experiment, and also the predicted maximum forming load is higher than that of the experiment by almost double. The rapid increase of the forming load occurs in different region forming strokes.
The difference between the experimental and predicted forming loads might be due to the forming speed and friction factor. Maintaining a constant forming speed and displacement control for a hydraulic press machine is extremely di cult. The forming speed is directly depended on the forming load. It is slower when the forming load increases. Furthermore, to precisely control the nal bottom thickness of the workpiece, the forming speed was slow down at the end of the forming stroke. As a result, the consistent forming speed was di cult to achieve in this experiment. For the friction, the precise friction factor in this process remained unclear. This method was designed with a constant friction factor which is likewise unattainable. These reasons might cause on the deviation on the predicted and the experimental forming load. However, the trend of the predicted forming load shows good agreement with the experiment. Also, this new developed technique could reduce signi cant forming load as well as provide a good precision of the part wall thickness and concentricity.          Piercing-coining load-stroke curves of various bulging strokes Figure 11 Load-stroke curves of the piercing of various the bulging strokes    Material lling into the bottom die cavity in various strokes Figure 18 Bulging load of the various counter-punch lifting displacements Figure 19 Piercing and coining load of various counter-punch lifting displacement Figure 20 Reduction in the base thickness in case of various counter-punch lifting displacements Bulging load-stroke curves of the punch and the counter-punch in various friction coe cients Figure 22 Piercing load-stroke curves of the punch in case of the various friction coe cients Figure 23 Manufacturing process of the semi-hollow part      Comparison of forming load between the experiment and the FEM simulation