A computerized process design system for manufacturing shells and other cup-shaped components

A Computerized Process Design System for Manufacturing Shells and Other Cup-Shaped

Components

J. P. TANG, S. I. OH, and F. M. L E E

A computerized system, called CPDSSM, was developed for design and optimization of the forming operations used in manufacturing artillery shells. The system is capable of simulating the cabbaging, piercing, drawing, and nosing processes, and it can also design streamlined dies for drawing and nosing preforms. This system is augmented by interactive computer graphics and can also be applied in manufacturing cup-shaped products other than shells. For a given set of input data, which includes the billet, die, punch geometries, and the processing conditions such as temperatures, ram speed, and interface conditions, CPDSSM predicts the deformation geometry and the load-displacement relationship in a step-by-step fashion.

The system is validated by comparing the predictions with the experimental measurements. In all cases, the agreements are good. In addition, the use of the computerized system as a design aid for shell manufacturing is demonstrated by a specific shell, which is in production. This example illustrates the capability and the potential benefits of this computerized system.

INTRODUCTION

The manufacture of artillery shells typically involves a series of metalforming, machining, and heat-treatment operations. Depending on the process, the metalforming is done at hot or warm working temperature ranges (conventional, hot-forged method), at room temperature (cold extrusion method), or at both hot working and room temperatures (hot cup, cold draw method). In typical shell-forming operations, the billet is obtained either by nicking and breaking, sawing, or shearing. Then the billet is formed by the cabbaging, piercing, drawing, and nosing processes into a shell, as shown in Figure 1.

It is desirable to increase the productivity of the shell- manufacturing facilities through improvements in process

J. P. TANG and S. I. OH, are with Battelle-Columbus Labora- tories, Columbus, OH 43201. F.M. LEE is with U.S. Army Ar- mament Research & Development Center, Dover, NJ 07801.

J. APPLIED METALWORKING �9 1985 AMERICAN

DRAWING NOSING

PIERCING

CABBAG I NG

I

Fig. 1 - - P r o c e s s sequence of 155-mm MI07 artillery shell manufacturing (max. shell O.D. = 6.16 in., height = 24 in.).

LJ I

design. For this purpose, mathematical models and computer programs have been formulated and validated through confirmation tests under production or near-production environment. ~-~2 These computer programs are capable of de-

SOCIETY FOR METALS VOL. 4 NO. 1 JULY 1985 7

termining and analyzing the metal flow which occurs during the cabbaging, piercing, drawing, and nosing operations. To further enhance the capability and the efficiency, these models were consolidated into a single integrated system for the design of shell-manufacturing operations. ~3.~4

This paper describes briefly the integrated system and demonstrates the use of this system as a design aid in shell manufacturing. The complete description of the processing conditions and detailed explanation of the process modeling of shell-forming operations are discussed by Lahoti et al. ~2

and Tang et al . ~4

THE COMPUTERIZED SYSTEM

As shown in Figure 2, the computerized process design system, CPDSSM, is composed of six functional models, namely CABBAG, PIERCE, CDVEL, DRAWNG, NOS- ING, and VOLUME. These modules are connected by an executive program which is implemented through interactive job control procedures; the intermodule commu- nication is facilitated by an universal data base.

Since the design process is iterative in nature, the system provides the user with a friendly environment, while allowing full control of the design process. Not only can each functional module be executed as a separated module to perform a designated function, but also it can be executed as a part of the total system to serve as a design aid for shell manufacturing. In addition, the computerized system supports dual units (English and SI), to provide easy data preparation and interpretation. The most commonly used material properties, such as flow stress data, density, etc., are included in each module as a user option. Since the flow stress is expressed as a function of temperature, strain, and strain rate, the computerized system is capable of analyzing the shell-forming processes under cold or hot conditions.

CCL .~

5 -I

i CA. AO I I "'ERCE I

I CDVEL I

I I

MATERIAL FILE [

J

I I IORAWNGI I "OS'"G I

I VO'OME I I

1 COMMON

DATABASE

Fig. 2--The flow diagram for the comprehensive system for computerized modeling of shell-forming processes.

The interactive computer graphics capability provided in the system greatly enhances the representation of the results.

CABBAGING PROCESS

Cabbaging is the initial forming operation in the shell- manufacturing process (Figure I). In cabbaging, a guided punch is used to partially pierce and upset one end of the billet, while a boat tail is formed at the other end by an extrusion type of deformation. The pierced recess at the top of the billet provides guidance for the subsequent piercing operation.

As suggested by the FEM simulation results, t2 the deformation modes during the cabbaging process are to occur in the following sequence: (1) forward extrusion, (2) upset t ing/backward extrusion, and (3) backward extrusion.

In order to simplify the analysis of the cabbaging process, the current model assumes that the geometry changes as shown in Figure 3. It is to be noted that upsetting is treated as a separate deformation process in the current model. Nevertheless, a special provision, described later in this section, was made in estimating the forging load when the upsetting process is not separable from the backward extrusion process.

In forward extrusion, the material flows in the same direc- tion as the punch movement. Since the billet is enclosed in the container die, the geometry of the billet after extrusion is determined by the container profile; the material fills up the cavity below the initial contact point. To simulate this deformation mode, the volume of region I in Figure 3(a) is first calculated and is subtracted from the billet volume. The new billet height is then calculated based on the remaining billet volume and original billet plan area. Since the load requirement for the initial forward extrusion is very small when compared with upsetting and backward extrusion, it is ignored in the modeling.

In upsetting, the billet is deformed to fill up the die cavity below the punch tip. To simulate this deformation mode, the new punch-tip position is located by using the secant itera- tion method. ~4 The force required for the simple upsetting of cylindrical billet with maximum friction at the bottom and no friction at the top is given by

F = A~[1 + ( f d / 3 h ) ]

where

F = force required

A = plan area of the billet

-~ = flow stress as a function of strain, strain rate, and temperature

f = friction factor

d = diameter of billet

h = height of the billet

8 VOL. 4, NO. 1, JULY 1985 J. APPLIED METALWORKING

Fig. 3 --Changes of geometry in the cabbaging processes assumed in the model: (a) initial configuration, (b) after tip-forward extrusion, (c) ancr upsetting, (d) backward extrusion steps, and (e) final configuration.

In backward extrusion the undeformed part of the billet remains stationary relative to the container. This, in general, eliminates the large frictional resistance between the container and the billet. The simulation of this deformation mode is performed by dividing punch movements into a finite number of equal steps, based on the punch position at the end of upsetting and the estimated final punch position. The final punch position is calculated by locating the punch position so that the volume of the enclosed region (between punch and container) is equal to that of the billet. For each backward extrusion step, the punch is moved forward one step size and the new billet height is calculated based on the constant workpiece volume requirement. From Kudo, ~5 the force required for the backward extrusion in cabbaging may be obtained by

66 + , 44 In + sin

+ JfSd sin OdA

where

Ap = maximum punch plan area below the billet height in the step

A0 = initial plan area of the billet

A~ = final plan area of the billet

f~ = friction factor between die and billet

f2 = friction factor between punch and billet

0 = angle between tangent to the punch or die surface and radial axis

To predict load-displacement relations, punch displace- ments are calculated based on the relative movements of punch tip with respect to the initial punch-tip position, and load requirements are calculated based on the above formu- lations. In general, the initial forward extrusion is to occur first, since the load requirement for this deformation is very small. Thus, forward extrusion is simulated first. De- pending on the geometries of the container, the die, the punch, and the billet, upsetting and backward extrusion

steps may or may not be separable. If the load requirement for upsetting is less than that of the backward extrusion, upsetting is simulated next. For the case when upsetting and backward extrusion occur simultaneously, the simulation of backward extrusion steps starts at the point where the load required for backward extrusion exceeds the calculated load required for upsetting; in this case, the simulation of the upsetting step is not carried out.

To verify the presented model, the computer program CABBAG was coded and compared with the results obtained from previously performed confirmation tests for the cabbaging of a 155-mm M107 shell. ~2 The material used to

manufacture the 155-mm M107 shell is AISI 1046, and the cabbaging process is done at 1100 ~ by using a hydraulic press with an average ram speed of 140 ipm. From Figure 4, it is seen that the agreement between the measured and predicted load-displacement curves is satisfactory. The slight variation in the shape of the predicted and experimental curves can be largely attributed to the absence of load prediction for the initial forward extrusion step and for simultaneous occurrence of upsetting and extrusion.

J. APPLIED METALWORKING VOL. 4, NO. 1, JULY 1985 9

700

6 0 0

500

c- O

400

155-mm MIO7 Shell. - - - Experiment

Predicted Curve by "CABBAG"

o j _J r / r

o 5 O O - JEI c~ E~

O /

2 0 0 I / / / / 1 / / i /

. . / , / , o o . , Y / "

o . . . . ff I I I 0 2 4 6 8 I0

Punch Displecement, in.

Fig. 4--Comparison of predicted and measured load-displacement curves for cabbaging of M107 shell (max. O.D. = 7 in., height = 13 in.).

P I E R C I N G P R O C E S S

In piercing, a solid billet is extruded backward over a punch to produce a hollow, cup-like part. The shape of the workpiece after piercing is predetermined by the die contour, which allows the material to flow through the gap between the die and the punch-side surface. The pressures involved depend on the billet geometry, lubrication, tool design, ma-

terial, and temperature. The simulation of the piercing process is similar to that of

the backward extrusion in cabbaging. Since the final punch position is predetermined, it is not calculated here. In addition, heat transfer is considered in the modeling of the piercing process. This is necessary because the effect of die chilling is significant, due to the large workpiece-die contact surface and the relatively thin workpiece wall.

In the heat-transfer analysis, it is assumed that temperatures of the workpiece and dies are uniform and that the heat conduction is due only to the temperature gradient across the lubricant film thickness between the dies and workpiece. Then the heat loss, AQ, from the workpiece to

the dies can be approximated by

A Q = hS(Tw - Td)

where Tw and T, are the temperatures of the workpiece and the dies, respectively. Here, S is the die-workpiece contact

area and h is the heat-transfer coefficient. Based on this

assumption, h can be expressed as

h = k /A th

where k is the conductivity of the lubricant and Ath is the film thickness. The temperature drop, AT, of the workpiece

is calculated by

A T = ( A Q / c p ) AT

where c and p are the specific heat and mass density, respectively, and A ~- is the time duration for the temperature drop. In the temperature calculation, it is also assumed that the die temperature does not change during the piercing operation. The temperature estimation is done in a step-by- step manner, since the workpiece temperature and the contact surface area change during the piercing operation. The material constants c and p are provided by the material routines in the program, while the heat-transfer coefficient is specified by the user in the input stage of the simulation.

To verify the presented model, the computer program PIERCE was coded and compared with the results obtained from previously performed confirmation tests for the piercing of a 155-mm M107 shell.12 The workpiece temperature

at the beginning of piercing was 1100 ~ and piercing was done in an 800-ton hydraulic press with an average ram speed of 500 ipm. From Figure 5, it is seen that the agreement between the measured and predicted load- displacement curves is good. The predictions are higher than the measurements at the early stage of the process. The

7~176 /

600 I

5OO

155-mm MI07 Shell

zx Experiment (55) o Prediction by PIERCE

Z 0

- 400 173 < O ..J

o 300 Z

0

14a

~- 200

I00

I I I I I 0 ! 0 2 4 6 8 I0 12

PUNCH DISPLACEMENT, INCH Fig. 5--Comparison of predicted and measured load-displacement curves for piercing of M107 shell (max. O.D. = 7 in., height = 17.75 in.).


difference may be attributed to the approximation made on the initial billet configuration. The predicted load decreases slightly after reaching the peak, while the measured load decreases rapidly during the final steps of deformation. This variation may be due to the constant ram speed assumed in the simulation, while the actual ram speed decreases rapidly at the final stages of deformation. In any event, the overall agreement appears to be satisfactory.

DRAWING PROCESS

In the shell-drawing process, the wall thickness of the shell is reduced while the internal diameter is kept constant. One aspect of this process is that the mandrel and the drawn shell have the same speed at the exit from the die. Consequently, the force necessary to carry out the drawing process is trans- mitted through (a) the draw stress, exerted by mandrel, pushing against the bottom of the shell and (b) the friction stress between the mandrel and the internal surface of the shell. Thus, it is possible to increase the maximum drawing reduction (ratio of initial wall thickness to final wall thickness) without punch through (break in the wall of the drawn shell) by increasing the friction stress at the mandrel- shell interface.

The mathematical model, capable of analyzing drawing through tandem dies, was developed earlier. 1,5-7 The model simulates the actual drawing process by dividing the punch movement into a finite number of discrete steps. The slab method of analysis was used to calculate stresses and drawing loads at each step. This analysis is valid for both conical and streamlined dies, since a complex die profile can be approximated by a series of straight lines.

In calculating the stresses, the flow stress of each element is considered as a function of strain, strain rate, and temperature. The strain in an element is the cumulative strain, and strain rate is calculated from a velocity field describing metal flow.~ The temperature of an element will depend upon the heat generated from plastic deformation, friction at the tool-material interfaces, and the heat conduction to the colder dies and punch.

Based on this stress analysis, a system of computer programs, named DRAWNG, was developed to simulate the drawing operation for artillery shells through multiple dies in tandem. The model was further validated through the confirmation test. 4 Three conical drawing dies in tandem were used. The workpiece of AISI 1046 steel was drawn at a temperature of 1043 ~ by using a hydraulic press with the ram speed of 960 ipm. The predicted and measured load displacement curves are shown in Figure 6. It is seen that peaks and valleys in the theoretical curves are somewhat sharper than measured. This is basically due to the sim- plified heat-generation and heat-transfer analyses used in the model, and due to the complex preform shape used in pro-

O m ~<

o o

E

350

280

210

140 !

7O

0 0

/"-'~, f~ - - Fxperiment01

z L / - ' q

Theroetical ~ ~"-.-J r',~ k,.

Specimen NO. E-9

20 40 60 80

Displocement, in.

Fig. 6--Theoret ica l and experimental load-displacement curves for hot drawing of M107 shell th rough conica l dies ( total area reduction = 36 pet).

duction, compared to that assumed in the analysis. How- ever, this error is not considered significant, since process design and equipment selection in reality is primarily based on the peak loads encountered in the process.

STREAMLINED DRAWING DIE DESIGN

Rationally designed sigmoidal dies have been found to be more efficient than conical dies, and their use produces metallurgically and mechanically superior products. In drawing of shells, conical drawing dies can be replaced by streamlined dies to reduce the drawing load. The deter- mination of an optimum die profile is done by minimizing the total drawing load. The differential equation for the function describing the die profile can be obtained from the application of the principle of virtual work rate.

A general upper-bound velocity field, without discon- tinuities, has been developed for drawing of shells and cups through curved dies. 1-3 In this study, five types of die profiles requiring minimum ram load were examined. The considered classes of die profiles are either with smooth entry and exit, or provided with blending radii in order to approxi- mate continuity in velocities at the die entrance and die exit. The flow stress of the deforming material is considered as a function of the strain, strain rate, and temperature. Thus, both hot and cold drawing of cups and shells are explored. The effects of work hardening and drawing speed upon the die profile, the maximum possible reduction, and the drawing load are estimated. The die profiles considered in the investigation are:

(1) Dies described with two circular curvatures (double- curvature dies)

(2) Dies described with a fourth-order polynominal profile (polynominal dies)

(3) Single-curvature convex dies with blending radius at exit (convex dies)


u

x

. . J

E

3OO

240 f ' ~ ~

180 [

120

60' Theoreticol

Specimen No. N-7

I Experimental

~ ' " ".

\

0 0 20 40 60 80

Displacement, in.

Fig. 7--Theoret ical and experimental load-displacement curves for hot drawing of MI07 shell through streamlined dies (total area reduction = 36 pct).

(4) Single-curvature concave dies with blending radius at entrance (concave dies)

(5) Straight dies with blending radii at entrance and exit (curved-straight-curved dies)

In order to verify the applicability of the presented model, a computer program (CDVEL) was coded, and three double- curvature dies (in tandem) were designed by CDVEL for the drawing of a 155-mm MI07 shell. 4 The workpiece of AISI 1046 steel was drawn at a temperature of 1043 ~ by using a hydraulic press with the ram speed of 960 ipm. The drawing process was then simulated by using the program DRAWNG, and the simulation results were compared with the experimental measurements. From Figure 7, it is seen that the agreement between the measured and predicted load-displacement curves is satisfactory. Again, the dis- crepancy between the load-displacement curves is partially due to the difference between theoretical and actual preform shapes.

NOSING PROCESS

In all modem methods of shell manufacturing, the internal cavity is formed to finish shape, and the machining is re- stricted to the outer surface of the shell. The open end of the rough-machined shell is closed in and the nose is formed. The "closing-in" is accomplished by forcing a contoured die axially over the open end of the shell, while the body of the shell is well supported by a holder.

The flow of metal in nosing is very complex, and a slight variation in the friction or the temperature conditions may result in a misformed shell, due to improper metal flow. In order to determine an optimum combination of the process variables, a number of computerized mathematical models for cold and hot nosing of shells were developed. 8-"

For designing a nosing preform, a method based on the consideration of local strains in the nosed shell was com-

puterized. For predicting time-dependent temperature distri- butions due to induction heating of the preform prior to hot nosing, the uniform heat generation was assumed along the length of the tube inside the coil, and the end effects were neglected. For predicting the metal flow, preforms with uniform wall thicknesses were considered. However, for predicting the load-stroke curve in the nosing operation, preforms with nonuniform wall thicknesses were considered, and the model to simulate the nosing process was developed accordingly. This model simulates the nosing process in a finite number of discrete steps and utilizes Nadai's stress analysis, m It considers the flow stress of the defomaing material as a function of the strain, strain rate, and temperature, and checks for local bulging at the nose base or for Euler's buckling, at each step of simulation.

To validate the mathematical models, the computer program NOSING was coded and tested. 12 A preform designed by NOSING to yield the desired ogive shape for the 155-mm M 107 shell was manufactured and nosed. A hydraulic press with the ram speed of 150 ipm was used for the nosing operation. In addition, the preform was induction heated to a temperature between 450 ~ and 750 ~ prior to the nosing operation. As seen from Figure 8, the preform designed by the program NOSING is as good as that designed by experience.

VOLUME CALCULATION

VOLUME is a computer program to calculate the volume of an axisymmetric part. This program is used to estimate the billet volume required for forming a given shell.

PROCESS DESIGN FOR A 155-MM SHELL WITH THE COMPUTERIZED SYSTEM

To demonstrate the use of the computerized system as a process design aid for shell manufacturing, a 155-mm M107 shell was selected because confirmation test results were available for immediate comparison. A detailed description of this example is given by Tang et al. 14

To simplify the design example, the following factors were considered:

(1) The geometry of the 155-mm MI07 as-nosed shell was used.

(2) AISI 1046 material and its flow stress characteristic were used.

(3) Machining allowance prior to and after nosing were known.

(4) Process sequence for producing the 155-mm M107

shell was known. (5) Workpiece and die interface conditions were known.


Fig. 8 - - N o s e d shells produced from the preforms designed by (a) experience and (b) program NOSING (max. shell O.D. = 6.16 in., height = 24 in.).

It is to be noted here that many uncertainties need to be considered in designing a forming operation. While the interface friction coefficients under various process conditions and lubricant types can be found from published literatures, the heat-transfer coefficients are rarely available. It is often necessary to estimate these coefficients based on the prior production data. Since heat-transfer coefficients may change by an order of magnitude during a forming operation, it is necessary for the designer to take such uncertainties, not only from the analysis but also from the actual production environment, into design considerations.

In Figure 9, the sequence of operations required for producing the 155-ram shell is shown. A billet, either round or round-corner-square (RCS) shape, is first cabbaged, pierced, and drawn. The drawn part is machined to obtain the preform for nosing. The machined preform is then nosed and machined to obtain the final as-nosed shell. The associ- ated modules, which aid the design and the optimization of the designated operations, are also shown in the figure.

Nosing Process Simulation and Preform Design

The design procedure starts with the geometry of the nosed shell. The program NOSING is used to aid the design of the nosing preform and to estimate the punch load required. The preform suggested by the computer program is shown in Figure 10.

Calculation of the Initial Billet Volume

Once the preform shape is determined for the nosing process, the initial billet volume can be determined. This is done by adding a machining allowance to the nosing preform shape. The computer program VOLUME is then used to estimate the required initial billet volume, based on the geometry of the drawn part.

At this stage, the drawn part geometry provides the drawing-punch profile. Since the punch nose shape for piercing is the same as that for drawing, the piercing punch geometry can be determined by allowing a clearance (at forming temperature) between the drawing punch and the inside diameter of the pierced shell, so that the drawing punch can easily slide into the pierced shell. Similarly, the cabbaging punch diameter is determined by allowing a certain clearance between the cabbaged shell and the piercing punch.

Cabbaging Process Simulation

Once the initial billet volume is determined, then the program CABBAG is used to design the cabbaging process. The design of the cabbaging process involves the deter- mination of the billet cross-sectional area, the container die diameter, and the shape of the cabbaging punch. These parameters should be determined in such a way that


C Billet ~ -]

I I I

[" l C Cabbaged ~ _.a

Part _~ -7

[ Piercing ] 'L- I ,

Part -]

I

Drawing I I I

! , P a r t

1 M a c h i n i n g i

1 N o s i n g P r e f o r m ~ --]

1 ' [ ~~ 1 ',

I i

] ' I

1 ' C~-~o~e~ > ; Shell

Fig. 9--Operational sequence for shell.

CABBAG

PIERCE

CDVEL & DRAWNG

NOSING

the manufacturing of 155-mm MI07

(1) cabbaging load should not exceed the equipment capac- ity, (2) the initial positioning of the billet should minimize eccentricity, and (3) the container diameter should allow adequate reduction in the drawing process (note that the piercing container is the same as the cabbaging container).

14 VOL. 4, NO. 1, JULY 1985

1~511n Hie7 SHELL IIOSINQ ( Ws---Fo-s-~ D )

(a)

I S S H n

I

|

ere

I

I !

~ i e 7

IN

SHELL NOSING (PREFORm)

i $.0

IN

IN ns,lo

0

]1 '~;, ~ 4 . , , . , M,. T.

!

I (b) J

Fig. lO--Nosing preform design displayed on CRT: (a) as-nosed (input) and (b) preform (output).

J. APPLIED METALWORKING

TOTRL RAM |.OQIP ! .01:-115 LP

5 0 . ~ .

4@, s~L.

d.O. ~.,

1~.O.

e,. O. - / I

@,0 2 . 0

/ , , , |

I ' I

~ . 0 lO . ( t 4.0 6 .0

DISPLACEMENT X 1.0E~00 IH

Fig. 1 1 - Simulation results obtained from program CABBAG (material

HEIClII" X l . .3E§ 11t

!~;=O

i .

e.ez--Zr-,t- 1 0.0 2.0 4 .~

RADZUS x 1.e~§ zN

= AISI 1045, temperature = 1100 ~

The peak cabbaging load will not vary if the maximum area reduction remains the same. Nevertheless, the shape of the load-displacement curve will change if the punch profile changes. Thus, by varying the punch geometry and/or container geometry when running CARBBAG, the optimal punch and conta iner prof i le can be determined. In Figure 11, the results obtained from the cabbaging simulation are shown.

Pierc ing P r o c e s s S i m u l a t i o n

The container die used in the piercing process is the same as that used in the cabbaging process. As for the punch profile, the shape of the punch nose is the same as that of the shell cavity. The maximum diameter of the piercing punch is determined in such a way that there is a clearance between the pierced inner profile and the container; thus, the contoured drawing punch can slide into the cavity.

The peak piercing load largely depends on the selected punch profile, friction shear factor, and heat-transfer coefficient. It is apparent that the deformation load increases if the maximum area reduction increases. A large surface area of

contact occurs between the billet and the die during piercing; therefore, friction is very significant in determining the peak piercing load. Also, the heat-transfer coefficient affects the amount of heat transfer that occurs between the material and die. This, in turn, affects the temperature distribution in the billet. Since the flow stress of the material is highly dependent upon the temperature, the peak piercing load varies substantially when different heat-transfer conditions are assumed. In view of the effects of the above variables, a sensitivity analysis may be performed by selecting different combinations of punch geometry, friction shear factor, and heat- t ransfer coef f ic ien t when running the program PIERCE. Such an analysis helps to optimize the piercing process. In Figure 12, the results obtained from the piercing simulation are shown.

Drawing Process Simulation

The workpiece geometry for the drawing process is determined from the results of the piercing simulation; the geometry of the drawing punch is determined from an inner profile of the final drawn part. The bottom of the inner


M107 SHELL PIERCING m ~

TOTAL RAt1 LOhD

x I.OE+eS L |

I S . e _

9 . i L

6 . 1 . /J+/ Q , I

I I I I I O.O 2 . 0 4 , 0 6 . 0 IhO I e . e

O l S P L / ~ U E H T X l . O [ + e i I N

HEIGHT x I . O E + H Im

'e.I) ; L I 4 . e

R~DIU$ 11 I ,IIE41'I~ IN

Fig. 12--Simulation results obtained from program PIERCE (material = AISI 1045, initial temperature = 1100 ~

profile of the drawn part is determined by the shell cavity shape and that of the shank profile is determined by the nosing preform shape.

In drawing, the number and the type of dies (conical or streamlined) should be determined first. The number of drawing steps is determined by the required thickness reduction. If streamlined dies are chosen, the program CDVEL can be used to design optimal curved dies, and the program DRAWNG can then be used to simulate the drawing process and to estimate the required load.

The billet cavity height and the die spacing have consid- erable effect on the overall shape of the load-displacement diagram. By proper selection of the spacing and the billet cavity height, with the aid of the program DRAWNG, a better drawing process design can be obtained. It should be mentioned that, due to the simplification of the assumed billet geometry in DRAWNG, process design studies should be carefully planned and interpreted.

For comparison purposes, the actual drawing die setup used in production was selected for simulation. Three

double-curvature dies were also designed for this demon- stration by using the program CDVEL. In either case, no

3 0 0

n~ 24C

% • 18G

o E 12G

6O

|

O0 I00

"" Exper imenta l

" ' " - - . . , " i . . . . . O+s

Streamlined Dies I

I I i I 20 40 60 80

Displ(]cement, in.

Fig. 13--Theoretical and experimental load-displacement curves for hot drawing of M107 shell through conical and streamlined dies (total area reduction = 36 pct).

punch through or wall tear was predicted by the program DRAWNG, since the stresses predicted in the product wall were below the yield strength of the product. Nevertheless, streamlined dies required lower peak punch loads for the operation. This can be seen from Figure 13, where the measured load-displacement relationships are shown with the predicted load-displacement relationship for the conical and

streamlined dies.


SUMMARY AND CONCLUSIONS

The manufacturing of artillery shells typically involves various metalforming operations. Therefore, simulation and optimization of shell manufacturing require the development of mathematical models that simulate these forming operations. Such mathematical models were developed, computerized, and consolidated into a single integrated system. The integrated system now consists of

�9 C A B B A G - - f o r cabbaging simulation �9 P I E R C E - - f o r piercing simulation �9 DRAWNG - - for drawing simulation �9 C D V E L - - f o r streamlined drawing die design �9 V O L U M E - - f o r initial billet volume calculation �9 N O S I N G - - f o r nosing simulation and preform

design

The system was validated by comparing the simulation results with the experimental measurements. The comparison showed very good agreement between predictions and measurements. Through this system, it is expected that a manufacturer of shells or cups can reduce

�9 Die design costs �9 Material waste �9 Extensive trials in die design �9 Lead times

In summary, in this paper the use of an integrated system as a process design aid for shell manufacturing was demonstrated. It may be noted that there is no single optimum route for forming shells and cups. The "optimum" process, in general, depends on equipment available, and it may vary from manufacturer to manufacturer. In practice, a total evaluation of all the operations, considered together, should be made to obtain an overall efficient process sequence. Since all the forming operations involved in shell manufacturing are interrelated, the use of the integrated system for overall process design is quite beneficial.

ACKNOWLEDGMENTS

This paper is based on information developed in a project on computer-aided design and optimization of the shell- manufacturing process. This project was sponsored by the U.S. Army Armament Research Development Center, Dover, New Jersey, with Mr. Fee M. Lee as Program Man- ager. This support is gratefully acknowledged. The authors would also like to acknowledge the technical support and

supervision provided by Dr. T. Altan, Senior Research l~ader, Engineering and Manufacturing Technology De-

partment at Battelle-Columbus Laboratories.

REFERENCES

1. G.D. Lahoti, T.L. Subramanian, and T. Altan: "Development of a Mathematical Model and Computer Programs Capable of Optimizing the Drawing Process for Actual Artillery Shell and Cartridge Cases," Final Report DAAA25-74-00557, Frankfort Arsenal, Philadelphia, PA 19137, May 1975.

2. G.D. Lahoti, P. Gordon, and T. Altan: "Optimum Die Profiles for Drawing of Shells and Cups," presented at NAMRC III, Pittsburgh, PA, May 1975, published in the Proceedings of NAMRC IlL

3. P. E Gordon, E Lee, and G. D. Lahoti: "Heat Transport in Drawing ThroughCurvedDies,"MechanicalResearchCommunications, 1977, vol. 4, pp. 57-62.

4. G.D. Lahoti, P.S. Raghupathi, I". L. Subramanian, and 1". Altan: "Confirmation Tests of Hot and Cold Artillery Shell Drawing Oper- ations," Contract Report ARSCD-DR-79008, U.S. Army Armament Research and Development Command, Dover, NJ 07801, 1979.

5. G.D. Lahoti, P. S. Raghupathi, and E M. Lee: "Applications of Pro- cess Modeling to Shell Drawing Operations Under Actual Production Conditions," paper presented at MTAG Conference, Phoenix, AZ, October 1979, published in Proceedings of MTAG.

6. T.L. Subramanian, G.D. Lahoti, N. Akgerman, and T. Altan: "Computer Aided Simulation and Optimization of Tandem Cup Ironing Processes," Proceedings ofNAMRC-VIII, LaRolla, MO, May 1980, p. 161.

7. T.L. Subramanian, G.D. Lahoti, and T. Altan: "Computer Aided Analyses and Simulation of Some Ironing and Forging Processes," Process Modeling Tools--Proceedings of Process Modeling Sessions of 1980, p. 129, American Society for Metals, Metals Park, OH, 1981.

8. T.L. Subramanian, G.D. Lahoti, and T. Ahan: "Temperature Dis- tributions in Inductively Heated Tubular Components Prior to End Forming," paper presented at NAMRC-V, Amherst, MA, May 1977, published in Proceedings of NAMRC-V.

9. G.D. Lahoti, and T. Altan: "Analysis of Metal Flow in Nosing of Shells," Proceedings of NAMRC-VI, Gainesville, FL, April 1978.

10. G.D. Lahoti, T.L. Subramanian, and T. Altan: "Development of a Computerized Mathematical Model for the Hot/Cold Nosing of Shells," Contract Report ARSCD-CR-78019, U.S. Army Armament Research and Development Command, Dover, NJ 07801, September 1978.

11. G.D. Lahoti and E M. Lee: "Computer Aided Simulation of Hot and Cold Nosing of Shells," Proceedings of NAMRC-VII, Ann Arbor, MI, May 1979.

12. G.D. Lahoti, S. I. Oh, and T. Altan: "Development and Confirmation of a Series of Mathematical Models for the Blocking, Cabbaging, Piercing, and Nosing Operations Involved in Shell Manufacturing," Contract Report ARSCD-CR-91010, U.S. Army Armament Research and Development Command, Dover, NJ 07801, May 1981.

13. S.I. Oh, G.D. Lahoti, E M. Lee, and A. Tang: "Computer Aided Simulation of the Forging Process for Manufacturing Artillery Shells," submitted to American Defense Preparedness Association Meeting, Monterey, CA, November 16-17, 1982.

14. J.P. Tang, S. I. Oh, and T. Altan: "Development of a Computerized System for Computer Aided Design and Optimization of the Shell Manufacturing Process," Contract Report ARSCD-CR-83013, U.S. Army Armament Research and Development Center, Dover, NJ 07801, May 198t.

15. H. Kudo: "Some Analytical and Experimental Studies of Axi- Symmetric Cold Forging and Extrusion--I," Int. J. Mech. Sci., 1960, vol. 2, pp. 102-127.


Documents

A computerized process design system for manufacturing shells and other cup-shaped components