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INTERNAL SHEAR FORGING PROCESSES FOR MISSILE PRIMARY STRUCTURES—ETC (U)
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Internal Shear Forging Processes
for Missile Primary Structures
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S. Rajagopal
S. Kalpakjian
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I I IT Research Instituted
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U.S. Army Missile Command
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Internal shear forging
Shear spinning
Near net shape processing
Primary missile structures
Thermomechanical treatment
Aluminum alloy 2014
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Technology was established for combining thermomechanical treatment
(TMT) with internal shear forging in the manufacture of 2014 aluminum
alloy subshells with integral internal ribs. These structures are cur¬
rently fabricated from 2014-T6 sheet and plate stock by welding together
the cylindrical outer skin of the subshell and a series of internal
stiffening rings. The objective of the project was twofold: to enhance
the properties and performance of these structures by means of TMT and -
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20. Abstract (Continued)
It
'"‘to achieve cost reductions over the current (fabricated) parts as a re¬
sult of single-piece construction and near net shape processing.
Internal shear forging consists of deforming a cylindridtl \ing be¬
tween a rotating external die and an internal roller to result in a
thin-walled tube with internal ribs in the circumferential direction.
This form of processing was conducted, in this project, on a heavy-duty
engine lathe equipped with special-purpose tooling and supplemented with
a heat-treating furnace for the thermal treatments. Several different
TMT cycles were evaluated in terms of the resulting tensile properties.
Production requirements and process economics were also analyzed in
detail.
It was concluded that internal shear forging can successfully re¬
place the current technology; that significant cost savings would accom¬
pany the implementation of the new technology; but that additional work
is necessary to test these parts for residual stresses, fatigue strength
crack propagation, and stress corrosion resistance, before the process
can be considered ready for production.
UNCLASSIFIED _
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FOREWORD
This Final Manufacturing Methods Report covers the work performed by
I IT Research Institute under Contract No. DAAK40-78-C-0264 (MM&T Project
R783204) from 27 September 1978 to 21 March 1981. It is published for
technical information only and does not necessarily represent the recom¬
mendations, conclusions, or approval of the U.S. Army.
This contract was initiated to establish Internal Shear Forging
Processes for Missile Primary Structures, and was conducted under the
direction of Mr. John Honeycutt/DRSMI-RLM, U.S. Army Missile Command,
Redstone Arsenal, Alabama. At IIT Research Institute, Mr. S. Rajagopal ,
Manager, Mechanical Systems and Design, was the principal investigator.
Notable contributions were made by Mr. A. Chakravartty, Mr. C. Hales,
Mr. w. Mitchum, Mr. J. Oni, Mr. A. Omotosho, and Dr. S. Agarwal. Pro¬
fessor S. Kalpakjian of I IT served as consultant to the project team.
Approved by:
/UvaM K,
M. A. H. Howes, Director
Materials and Manufacturing
Technology Division
S. Rajagopal, Manager
Mechanical Systems
and Design Section
CONTENTS
Page
FOREWORD.iii
1 . INTRODUCTION. 1
2. PROGRAM OBJECTIVES AND SCOPE . 2
3. REVIEW OF INTERNAL SHEAR FORGING AND RELATED PROCESSES . 5
3.1 The Basic Process of Shear Forging of Tubes . 6
3.2 Typical Components Produced by the Shear Forging Process
and Their Properties.14
3.3 Theoretical and Experimental Studies in Shear Forging
of Tubes.1?
3.3.1 Spinnabil ity of Materials.18
3.3.2 Material Flow and Resulting Dimensional Changes. . . 24
3.3.3 Forces in Shear Forging of Tubes.31
4. 2014 ALUMINUM ROLLING EXPERIMENTS.36
4.1 Task Objectives.36
4.2 Starting Material . 37
4.3 Rolling Experiments . 41
4.4 Microstructural Observations.47
4.5 Conclusions.52
5. TOOLING DESIGN AND FABRICATION . 53
5.1 Shear Forging Machine . 53
5.2 Basic Considerations in Tooling Design.53
5.2.1 Die Construction.53
5.2.2 Roller Design for Shear Forging.54
5.2.3 Design Parameters Affecting Surface Finish . 55
5.3 Initial Tooling Design and Trials . 57
5.4 Tooling Modifications . 62
6. INTERNAL SHEAR FORGING OF SUBSHELLS.66
6.1 Process Description . ^
6.2 Metal Flow.69
6.3 Shear Forging of 2014 A1 and 2024 A1 Subshells.78
6.4 Thermomechanical Treatment of 2014 A1 Subshells . 81
6.5 Processing of Deliverable Subshells . 93
6.5.1 Deliverable Items.93
6.5.2 Residual Stresses.94
6.5.3 Thin Rib Generation.94
iv
CONTENTS (continued)
Page
7. PRODUCTION REQUIREMENTS AND COSTS. 97
7.1 Equipment and Tooling.97
7.2 Production Process Sequence . 98
7.3 Standard Hours for Subshell Production.98
7.3.1 Heat Treatment Station.99
7.3.2 Spinning Lathe Station . 99
7.3.3 Machining Station.101
7.3.4 Inspection Station . 102
7.4 Subshell Production Cost for
Internal Shear Forging vs. Current Process.102
7.4.1 Production Cost—Internal Shear Forging.103
7.4.2 Production Cost—Current Process . 103
7.4.3 Return on Investment and Payback Period.103
8. CONCLUSIONS AND RECOMMENDATIONS.104
REFERENCES.105
DISTRIBUTION LIST.108
v
TABLES
Number Page
1 Mechanical Properties and Chemical Composition of
As-Received 2014-T651 Aluminum Alloy Rolled Plate . 38
2 Mechanical Properties of 2014-0 Aluminum Alloy Plate
Used as Starting Material In Rolling Experiments.38
3 Rolling Experiments with 2014-0 Aluminum Alloy with
Different Lubricants.44
4 Effect of Rolling Speed and Heat Treatment on Mechanical
Properties of Aluminum Alloy 2014 . 45
5 Effect of Blank Temperature and Rolling Reduction on
Mechanical Properties of Aluminum Alloy 2014.46
6 Mechanical Properties of Internally Shear Forged Samples. . . 79
7 Surface Finish of Internally Shear Forged Parts . 30
8 Tensile Properties of Internally Shear Forged 2014 Aluminum
Alloy after Different Thermal-Mechanical Cycles . 91
Effect of Final Aging Treatment on Tensile Properties of
2014 Aluminum Alloy.92
9
FIGURES
Number Page
1 Subshell target geometry for internal shear forging . 2
2 Program tasks and schedules . 4
3 The shear forging process . 6
4 Jet engine turbine drive shaft roll-formed from AMS 6415
alloy steel forged preform. 8
5 Rear compressor-case layout for flow turning. 8
6 HP9-4 motor case internally rol1-extruded from preform. ... 9
7 Shape capability by shear forming cylinders from rings. ... 10
8 Forces on roller during internal shear forging.11
9 Comparison of grid deformation, photomicrograph, and
isohardness (Vickers) lines in shear forged copper tubing
with 66% reduction, k In. original thickness.12
10 Longitudinal section through the wall of a cylindrical
workpiece showing expansion of the unspun section . 13
11 Distortion of shear forged parts ahead of the deformation
zone.13
12 Schematic diagram of workpiece deformations investigated in
reference 33.13
13 Internal and external roll extrusion with workpiece pulled
through die-roller gap.15
14 Internal roll extrusion of ribs using moving inner mandrel. . 16
15 Test setup to subject tubes to continuously increasing
wall reduction by shear forging . 18
16 Sections of fractured samples from spinnability tests,
indicating degree of forward reduction permissible for
(A) 2024-T4 A1, (B) 6061-T6 A1, (C) annealed copper, and
(D) mild steel.19
17 Maximum reduction in shear forging of various materials as
a function of tensile reduction of area.20
18 Chevron cracking caused by small reductions and large die
angles in extrusion, drawing, and shear forging, shown
here for cold tube extrusion. ..23
19 Buildup in various stages of shear forging and influence
of roller nose radius on buildup formation.25
20 Stages in shear forging with three different methods using
the same roller.26
21 Influence of skew angle on buildup formation.26
22 Influence of contact angle a of roller on buildup formation . 26
vi i
FIGURES (continued)
Number
23
24
25
26
27
28
29
30
31
32
33
34
35
36
37
38
39
40
41
42
43
44
45
Page
Influence of original wall thickness on buildup formation. . . 27
Comparison of calculated angles of buildup with actual
buildup profiles in forward shear forging.27
Influence of roller angle, feed rate, and reduction on
buildup and distortion ahead of roller . 28
Roller geometry and nomenclature . 29
Section through workpiece showing material flow.30
Force coefficients in forward and backward shear forging ... 32
Force factors vs. thickness reduction in backward shear
forging.33
Force factors vs. thickness reduction in forward shear
forging.33
Contact geometry in external and internal shear forging
of tubes.35
Precipitate phases in 2014-T651 as-received plate.39
Finely dispersed precipitate phase observed after
annealing (prior to rolling) . 40
Reduction sequence for rolling of 2014-0 aluminum alloy. ... 42
Summary of rolling experiments for 2014 aluminum alloy ... . 43
Photomicrographs,(a) optical and (b) SEM, of as-rolled
structure after 95 percent reduction by rolling at 149°C . . 48
Photomicrographs, (a) optical and (b) SEM, of as-rolled
structure after 95 percent reduction by rolling at 316°C . . 50
Optical photomicrographs of samples rolled to 95% reduc¬
tion, solution treated and naturally aged to the T4
condition.51
Initial tooling design for internal shear forging.58
Experimental setup for internal shear forging.59
Shear forging roller unit.59
Side view of tooling for internal shear forging.60
Initial tooling setup for internal shear forging . 60
2014 aluminum alloy ring, partially deformed by internal
shear forging.61
Improved tooling design for internal shear forging . 63
viii
FIGURES (continued)
Number Page
46 Modified tooling for internal shear forging, assembled
in LeBlond 2516 engine lathe.64
47 Modified rollers for internal shear forging.65
48 Process elements in internal shear forging of 2014 A1
subshells.66
49 Internal shear forging of aluminum rings into thin-walled,
internally ribbed tubes.67
50 Graphite spray lubrication during initial warm deformation
in internal shear forging.67
51 Adjusting screw being turned for rib forging . 68
52 Intermediate tube wrapped in aluminum foil (to prevent
heating) and loaded in furnace for solution treatment. ... 68
53 Thermomechanically treated, internally ribbed tube being
removed from the die.69
54 Low-magnification photomicrographs showing metal flow in
internal shear forging . 70
55 Fold produced due to inhomogeneous deformation, shown
schematically (a) and on an actual part (b).72
56 Lamination defect, caused on the ID surface during rib
forging by foldover of buildup from previous passes,
shown schematically (a) and on an actual part (b).73
57 Cracking of wall 0D from secondary tensile stresses due
to high radial feed, low temperature, inadequate roller
clearance, and rough machining marks on preform 0D,
shown schematically (a) and on an actual part (b).75
58 Buckling of thick wall ahead of roller and thin wall
behind roller due to heavy reduction and feed rate in
final pass, shown schematically (a) and on an actual
part (b).76
59 Void observed in thin wall in one case, attributable to
hard inclusion or entrapment of wear debris in final
pass, shown schematically (a) and on an actual part (b). . . 77
60 Inside surface finish of part produced by 1SF/RF . 80
61 Thermomechanical treatment (TMT) cycles evaluated during
internal shear forging of 2014 aluminum alloy.83
62 Optical photomicrographs of 2014-0 aluminum specimen
prior to internal shear forging.84
ix
FIGURES (continued)
Number ft,3-
63
Optical photomicrographs of internally shear forged 2014
aluminum specimen after cycle 1.
. . 85
64
Optical photomicrographs of internally shear forged 2014
aluminum specimen after cycle 2.
. . 86
65
Optical photomicrographs of internally shear forged 2014
aluminum specimen after cycle 3.
. . 87
66
Optical photomicrographs of internally shear forged 2014
aluminum specimen after cycle 4.
. . 88
67
Optical photomicrographs of internally shear forged 2 014
aluminum specimen after cycle 5.
. . 89
68
Optical photomicrographs of internally shear forged 2014
aluminum specimen after cycle 6.
. . 90
69
Deliverable subshells, internally shear forged with
thermomechanical treatment .
. . 93
70
Residual stresses from internal shear forging of preaged
. . 95
71
Observation of residual stresses by cutting a thermo-
mechanically processed subshell.
. . 96
1. INTRODUCTION
Missile primary structures are currently fabricated from 2014
aluminum alloy sheet and plate stock by welding together the cylindri¬
cal outer skin of the structure and a series of internal stiffening
rings. Cost reductions are sought for large production quantities by im¬
plementing internal shear forging to produce these structures in monolith¬
ic construction. An additional benefit is the possibility of enhancing
structural performance by incorporating thermomechanical treatments (TMT)
into the internal shear forging process for these structures.
Internal shear forging consists of deforming a cylindrical ring
between a rotating external die and an internal roller, resulting in a
thin-walled tube with integral internal ribs. It is fairly well estab¬
lished as a production process for axisymmetric components and is known
by a variety of names including internal tube spinning and internal roll
extrusion.
The objective of this program was to establish the internal shear
forging process for missile primary structures featuring thermomechanical
treatment. For aluminum alloys, TMT typically involves introducing cold
work into the precipitation-hardening cycle which results, in many cases,
in improved tensile properties, fatigue strength, fracture toughness, and
stress corrosion resistance.
The work performed on this program and covered in this report in¬
cluded tooling design and fabrication, an exploratory study of basic
parameters, internal shear forging experiments, processing of deliver¬
able parts with thermomechanical treatment, tensile property determina¬
tion, and an economic analysis involving considerations of production
requirements and costs for internal shear forging.
2. PROGRAM OBJECTIVES AND SCOPE
The target component for this program (Figure 1) was a simplified
subscale version of a missile primary structure and is referred to here¬
after as the "subshellThe objective of the program was to establish
tooling requirements and manufacturing procedures to shear forge the sub
shell with thermomechanical treatment during processing.
/
Finish Machining
Required
Figure 1. Subshell target geometry for internal shear forging.
The program was organized In two phases. Phase I (Basic Effort)
involved rolling experiments with 2014 aluminum, and tooling design and
fabrication for internal shear forging. In Phase II (Option 1), tooling
trials were made, followed by tooling modifications, final trials.
evaluation of various TMT cycles by tensile property determination,
internal shear forging of deliverable parts, and cost analysis for manu
facture.
Figure 2 shows the tasks and schedules for both phases of the
program.
3
Figure 2. Program tasks and schedules
3. REVIEW OF INTERNAL SHEAR FORGING AND RELATED PROCESSES
This section outlines the state-of-the-art in shear forging of
tubes, both external and internal, whereby the original tubular blank is
subjected to wall thickness variations along its length while maintaining
rotational symmetry. The process is known in industry by a variety of
other names such as shear forming, tube spinning, flow turning, spin
forging, rotary extrusion, roll extrusion, flow forming, hydrospinning,
rotoforming, and floturning. Although all these terms describe the basic
process that is common to all spun parts, there are some differences
among them. These differences generally pertain to the type of rollers,
mandrels, and machines used in making these parts. In this report, the
process will be referred to as shear forging.
The shear forging process in quite versatile in view of the fact
that a great variety of tubular parts can be manufactured with basically
the same tooling. This is done by controlling the path of the rollers
during their longitudinal travel along the length of the workpiece and
by controlling roller geometry and reduction in thickness. Typical com¬
ponents are pressure vessels, automotive, and rocket and missile parts.
The process is suitable not only for prototype or low production runs,
but also for high-production runs on automated equipment. Shear forging
can be quite economical due to relatively low tooling costs, material
utilization, and the ease with which design changes can be made.
Depending on the strength and ductility of the workpiece material
and the capacity of the machine, the shear forging process can be carried
out either at room temperature or at elevated temperatures. The re¬
sulting properties of the product will, of course, depend on the parti¬
cular material and process parameters, such as reduction and temperature.
In general, the mechanical properties are enhanced, thus making the
process attractive for strong but lightweight designs. Various general
surveys of the shear forging process can be found in the literature,
more recent ones being given in References 1 through 4.
In view of its many advantages, this process has been studied quite
extensively during the past thirty years or so. These studies have
ranged from the most analytical to the most applied aspects, the latter
including basically developmental work in tooling and process control in
order to obtain components of certain specified geometries. Contribu¬
tions to the understanding of the process have been concerned with force
and power requirements, spinnability of materials, material flow, sur¬
face deformations at the roller-workpiece interface, properties of spun
parts, and how these factors are related to process parameters and orig¬
inal material properties. Such understanding is essential in order to
make full use of the capabilities of this process and to aid in the de¬
sign of components to be spun. Although the majority of these studies
have been on external shear forging, much of the knowledge gained should
be equally applicable to internal shear forging.
3.1 THE BASIC PROCESS OF SHEAR FORGING OF TUBES
The basic process of shear forging is shown in Figure 3, both for
external and for internal spinning. The original tubular blank is placed
on the mandrel (in the case of internal shear forging, inside the die),
and a roller of a certain geometry travels axially, reducing the thick¬
ness of the blank.
BACKWARD SPINNING
FORWARD SPINNING
ROLLER^
Fr
,1,
u
0 N
^ MANDREL
r
ZZZZZZZZ ' ZZZZ2 ^
Figure 3. The shear forging process (also known as shear forming,
tube spinning, flow turning, spin forging, rotary extrusion,
roll extrusion, flow forming, hydrospinning, rotoforming,
and floturning ).
6
It can be readily recognized that there are a number of independent
variables in this basic process. These are listed below.
a) MandneZ ViameXoM. and fipm
It can be seen that the mandrel does not necessarily have to be of
constant diameter along its length, and that a variety of parts can be
shear forged with shaped mandrels, provided that the mandrels have rota¬
tional symmetry. Figures 4-7.
b) RoiZeA. Size, Ge.omeZx.cf, and OxitnZaZion
A great variety of roller geometries have been developed over the
years. The contact geometry between the roller and the workpiece is
important and could play a critical role in the successful operation of
the process.
Whereas there is a maximum limit to roller diameter in internal
shear forging, there is no such limitation for external shear forging.
It is obvious, of course, that as the roller diameter increases, the
contact area between roller and workpiece increases; this will then in¬
fluence the magnitude of the forces involved, as will be described later.
The power input into this process can be either through the mandrel
(such as by placing it on the headstock of a lathe) or through the roller
(by powering the shaft on which the roller is placed).
c) Feed
The distance traveled by the roller per revolution of the mandrel is
referred to as feed. It is an important parameter as it has influence
on surface finish and dimensional control of the final diameter of the
shear forged product. In situations where more than one roller is used,
the effective feed rate for each roller will be less.
d) RzducZZon pex Pa66
Because of its great influence on a number of parameters, reduction
per pass has been studied extensively. Its effects depend not only on
other variables (such as roller geometry), but also on the mechanical
properties of the workpiece material.
7
I 0.345 min I
0.358 man
Figure 6. HP9-4 motor case (bottom) internally roll-extruded
from preform (top). 12 (All dimensions in inches.)
o 5o">p'*
FlOng* NormoHy
U5#C3 Id K>K5 Dd*^
c CO'TV*** Cy*» n d0'5
Cf 055 «c»or'ond Corstorrt ID
t) Cyt^dftr *»»* *t*Qroi S»«»t*nr»
Ond /£> JD**5
a S«X* PrtXloCSd by PrOpf*tary
Sh®o» Fd^'^O Process
VdV ,r *Q ID ond 00
Figure 7. Shape capability by shear forming cylinders from rings
e) Wokkpl&c.z MateJuaJL
The properties relevant to workpiece materials are mechanical
(strength and ductility), metallurgical (role of impurities, response to
thermomechanical treatment), and physical (specific heat, and coef¬
ficients of thermal conductivity and expansion). The interrelationship
of these properties and other variables of the process is quite complex.
Some details will be discussed later in this section.
The foregoing independent variables will have influence on the
following dependent variables of the process:
a) Foficu > . As can be seen in Figure 8, there are three principal
force components in shear forging of tubes. These are axial, radial,
and tangential. In calculating the individual power components involved
with each of these forces, it can be shown that the most significant
force is the tangential force, F t , which supplies the torque to the
system. The role of the axial force, F a , is rather small by virtue of
the fact that the axial distance traveled by this force is quite small.
The radial force, of course, does not contribute to the power consumed
in the process.
Figure 8. Forces on roller during internal shear forging,
(far F r , and F t are associated with compressive stresses
in the unspun region of the workpiece in the axial,
radial, and tangential directions, respectively ).
11
Knowledge of these forces Is essential not only for determining the
power consumption, but also for the design of machines--particularly
when high rigidity is required for better dimensional control of the
process.
b) McUeAxal Flow . There are two aspects to this parameter. One is
the bulk flow of the material in the deformation zone, Figure 9, and the
other is the distortion, ahead of the roller, of both internal and exter
nal surfaces of the tube being shear forged, Figures 10 to 12.
Figure 9. Comparison of grid deformation, photomicrograph, and
isohardness (Vickers) lines in shear forged copper tubing
with 66 percent reduction, '4 in. original wall thickness . 21
Figure 10. Longitudinal section through the wall of a
cylindrical workpiece showing expansion of the
unspun section. Original wall thickness = 6 mm
material = 0.15 percent C steel. 33
Figure 11. Distortion of shear forged parts ahead of
deformation zone. (a) Cup shear forged to (top) half
and (bottom) almost its full length; (b) transition
between shear forged wall and original wall. 21 *
Figure 12. Schematic diagram of workpiece deformations
investigated in reference 33.
1
Bulk flow characteristics determine the final properties of the
tube (including the role played in thermomechanical treatment), and can
also play a role in defect formation in the product. Surface flow is
important for residual stresses and in dimensional control of all sur¬
faces of the workpiece, although it can also influence the quality of
the product at or near the surfaces. Because they can change the contact
area and geometry at the roller-workpiece interface, surface flow charac¬
teristics also influence forces.
3.2 TYPICAL COMPONENTS PRODUCED BY THE SHEAR FORGING PROCESS
AND THEIR PROPERTIES
A great variety of components have been produced by the shear
forging process of conical, curvilinear, and tubular geometries. In
this section, only tubular shapes will be discussed. In addition to
simple, constant wall thickness tubes, both external and internal geome¬
tries and ribs can be obtained, as shown in Figures 13 and 14, and
earlier in Figures 5 to 7.
External shear forging has been used, thus far, more commonly than
internal shear forging, as reflected in the technical literature.
References 1-15 pertain to external and 16-17 to internal shear forging
(roll extrusion).
Whereas in external shear forging the power to the system is gener¬
ally supplied through the mandrel (with idling rollers), in roll extru¬
sion the power is often supplied through the rollers. The part is
pulled through stationary rollers by grippers (Figure 13) or reduced in
wall thickness by a moving internal roller (Figure 14).
As can be seen from the illustrations in the literature, typical
parts made by these processes are cylindrical shapes of a large variety
of materials, with or without external and internal ribs and flanges.
This process has been proven to be more economical than other processes
to manufacture these same parts. Thus, the process has had wide appli¬
cations in aerospace, military, and nuclear fields, in addition to a
wide range of applications for commercial products.
14
In addition to reducing costs and material savings, the process
enhances the mechanical properties of the material as shown in the pres¬
ent study and also in the technical literature. 3 ’ 18 ’ 19 The increase in
properties will, of course, depend on the workpiece material, reduction,
and temperature. As expected, strength and hardness increase with a
corresponding decrease in ductility. These properties can then be
altered with post-processing heat treatment, such as stress relieving and
annealing.
A significant aspect in property enhancement is the thermomechanical
treatment (ausforming) before and during the shear forging cycle. In
one study on external shear forging of low alloy steel tubes, it was
shown that thermomechanical treatments produced tubes with very high
yield strength combined with good ductility. 20
As it has been shown in the present study, such processing is also
applicable to the internal shear forging of 2014 aluminum tubes with
possible enhancement of their mechanical properties.
3.3 THEORETICAL AND EXPERIMENTAL STUDIES IN SHEAR FORGING OF TUBES
In view of its technological significance, the shear forging process
has been the subject of studies for the past three decades or so. These
studies have ranged from the most analytical to the research and develop¬
ment type work in various industrial organizations. The principal focus
on these studies is mainly toward establishing relationships between
material and process variables and such parameters as forces and spin-
nability of materials.
A general survey of the more technical studies on the subject is
given in Reference 21. Much of the technical literature pertains to
shear forging of conical shapes. The earliest systematic study of tubu¬
lar shapes appears to have been done in 1961. 22 Since then, a number of
publications have appeared on various aspects of shear forging of
tubes. 23-39 None of these studies pertain to internal shear forging.
It is expected, however, that many of the quantitative relationships ob¬
tained for external shear forging would also be applicable to internal
shear forging.
17
The studies conducted thus far have been directed toward an under¬
standing of the following parameters: spinnability of materials, mater¬
ial flow and resulting dimensional changes, and force and power require¬
ments as a function of process parameters and material properties. These
will now be discussed in terms of their engineering relevance.
3.3.1 SnLnnabZtltu of MateAiaZ&
Spinnability of materials is defined as the maximum reduction per
pass that the material can undergo before failure takes place. It is,
of course, desirable to study this process and predict, by analytical or
empirical means, whether or not a given workpiece of certain dimensions
and properties will withstand the stresses imposed upon it. From
Figure 3, it will be noted that in forward shear forging the spun section
is in tension, and in backward shear forging the unspun section is in
compression. Thus, the failure modes will be in tension and by buckling,
respectively.
It appears that the only systematic study of this subject is that
given in Reference 26. With a test setup as shown in Figure 15, a number
ROLLER path
i ^ r
j ROLLER |
Figure 15. Test setup to subject tubes to continuously increasing
wall reduction by shear forging . 26
of metals were spun with increasing reduction in thickness until tensile
failure (or failure in the deformation zone) occurred. Sections of these
metals Illustrating failure are shown in Figure 16. Additional studies
were conducted to determine the influence, if any, of process variables
such as feed, roller corner radius, and roller angle (equivalent to die
angle) on this maximum reduction per pass.
Figure 16. Sections of fractured samples from spinnability tests,
indicating degree of forward reduction permissible for
(A) 2024-T4 Al, (B) 6061-T6 A1 , (C) annealed copper,
(D) mild steel.
Attempts were then made to establish a quantitative relationship
between spinnability and material properties. The results indicated that
the property that was most relevant was the tensile reduction of area of
the original tube material in the longitudinal direction. Figure 17. It
is interesting to note that, up to a tensile reduction in a'ea of about
50%, spinnability was more or less a direct function of the ductility of
the material. The material failed because of lack of ductility as re¬
quired by the thickness variation. Above 50%, a plateau was reached,
indicating that failure was due to tensile stresses in the spun section
of the tube; in this case the ductility of the tube material is of no
primary significance. (This analogy is the same as that used in deter¬
mining maximum theoretical reduction per pass in wire, rod, or tube
drawing.)
l'J
Figure 17. Maximum reduction in shear forging of various materials
as a function of tensile reduction of area. 26
In experiments with mild steel and copper, it was shown that the
maximum reduction per pass decreased with increasing feed. This decrease
could be attributed to an increase in the axial force with increasing
feed, thus increasing the stresses in the spun section of the tube.
Maximum reduction per pass decreased slightly with increasing roller
corner radius, while roller angle had no influence on maximum reduction.
In additional experiments, tubes were reduced in thickness by
spinning, in order to subject them to different degrees of cold work.
These tubes were then shear forged in the same setup as Figure 15 to de¬
termine the maximum reduction per pass. The amount of prior cold work
did not have any significant influence on the maximum reduction. This
means that in multiple-pass operations the same maximum safe reduction
may be employed continuously at each pass.
Although these studies have not been extended to internal shear
forging, there is no particular reason to believe that the external shear
forging results obtained would not be applicable. The two processes are
essentially the same; the wall-thickness-to-mandrel (die)-diameter ratio
is essentially the same, and the deformation zone and the stresses to
which the material is subjected are also the same. The only major dif¬
ference is that, under similar conditions, the contact area between the
roller and the workpiece is larger in internal than in external shear
forging.
As far as backward shear forging is concerned (a process which has
a wider application because of the ease with which it can be carried
out), spinnability becomes somewhat difficult to determine because:
(a) As the thickness reduction per pass increases, reverse buildup under
and ahead of the roller takes place, (b) buckling of the unspun section
of the tube takes place as the longitudinal spinning forces increase due
to increasing reduction, and (c) the buckling will depend on the length
of the unspun tube section.
It would appear, however, that for both the external and internal
case, the ductility of the workpiece material will be important. This is
because the material is subjected to a reduction in thickness and it
21
must have the capacity of undergoing large strains to failure. It should
be pointed out that, as shown in numerous studies, ductility is very
much a function of the state of stress and that a compressive environ¬
ment is desirable for enhanced ductility.
From this point of view, it would appear that internal backward
shear forging is a particularly desirable situation.
The discussion, thus far, has been concerned with maximum reduction
per pass. It should be pointed out that light reductions can also have
adverse effects on the spun product. This is due to the fact that light
reductions will deform mainly the surface layers of the tube with little
deformation of the bulk of the material under the roller.
Whereas light reductions in shear forging may be desirable for work¬
hardening the surface layers of the product (just as in shot peening,
with beneficial increase in the fatigue life), they can have a strong
adverse effect on the product quality. 29 ’ 30 This is due to the fact that,
with light reductions, the plastic zone under the roller does not fully
penetrate the thickness of the tube. This subsequently generates a
hydrostatic tensile stress component which can cause cracks on the
mandrel side of the tube wall (Figure 18). The mechanism is the same as
that obtained in drawing and extrusion of solid tubular parts, the
fracture being called by a variety of names such as centerburst, chevron,
arrowhead, and cuppy core. The generation of these cracks is further
accelerated by the presence of impurities and inclusions, particularly
if they are hard.
To avoid such fractures in shear forging, two important parameters
have to be controlled. One is the percent reduction. Under otherwise
identical conditions, higher reductions will reduce or eliminate fracture
by assuring that the plastic zone penetrates the thickness of the tube.
The other parameter is the roller geometry. The smaller the roller
angle (equivalent to the die angle in drawing), the larger and deeper the
deformation zone. Furthermore, the radius between the roller angle and
its relief angle is important, particularly at small reductions on thin-
walled tubes. This is because the larger this radius, the smaller the
effective roller angle.
22
Cracking resulting from ^o-stage cold e ^ 5 ^ o ; 30 *.
b First stage 1.23, second stage 3.6
(c) First stage 2.0, second stage l.»
(d) First stage 2.0, second stage 3-6
3.3.2 tfatwLaZ Ftou) and Rebutting VimznAionaZ ChangeA
In addition to the bulk deformation of the material between the
roller and the mandrel, deformation of the outer and Inner surfaces of
the tube is important. This aspect of shear forging of tubes has been
the subject of study ever since the systematic work by Thomasett 22 in
1961, although many relevant observations had been made prior to that
time in research studies by commercial organizations.
This subject falls into two categories. One is the diametral con¬
trol of the tube, and the other is the dimensional control of the sur¬
faces of the tube, particularly in generating ribs and flanges of pre¬
scribed geometries.
In external shear forging, two important parameters in diametral
control are the feed and the roller geometry. Diametral growth is akin
to ring rolling and also to spread in plate rolling. In both of these
processes, the geometry of the contact area with respect to the workpiece
dimensions is the significant parameter. Thus, in shear forging, low
feeds, small roller angles, and large nose radius or flat on the roller
will tend to increase the circumferential dimension of the tube—hence
an increase in its diameter. In internal shear forging, this same situa¬
tion will result in the tube becoming tight against the die. It is
obvious that if and when the workpiece has been preheated or if it
develops high temperatures due to the mechanical work input, then, upon
cooling, there will be additional changes in the diameter of the spun
tube. This, of course, will be helpful in the subsequent removal of the
part from the die.
An additional dimensional change that has been observed in external
shear forging is bell mouthing.* In forward shear forging of tubes, the
free end of the tube (i.e., the unspun end) has a tendency to expand
circumferentially even though the section being reduced in thickness is
still quite a distance away (Figures 10 and 19). One solution to this
problem is to use a simple restraining ring on the free end of the tube.
* In internal shear forging, the free end was found to decrease in diam¬
eter, leading to nosing rather than bell mouthing.
£4
Figure 19. Buildup in various stages of shear
forging (top) and influence of roller nose
radius on buildup formation (bottom). 22
However, studies reported in the literature have identified the signifi¬
cant parameters that play a role in this phenomenon. This subject is
closely related to the surface deformations that take place adjacent to
the roller-workpiece contact zone.
These surface deformations have been studied systematically, as de¬
scribed in References 22, 27, 28, 31, and 32 with additional examples
being given in References 24, 33, 38, and 39. Typical examples of buildup
are shown in Figures 19 to 25.
It has been shown that the important parameters in these deforma¬
tions are: reduction in wall thickness, roller angle (or its nose radius
if round), the angle of tilt of the roller axis with respect to the
mandrel axis, and feed. The surface deformation (also called buildup or
Figure 20. Stages in shear forging with three different
methods using the same roller. (a) Roller axis parallel
to mandrel axis, i.e., skew angle = 0°; (b) inclined roller
axis, skew angle = +20°, direction of feed from cup base
to cup flange ; (c) inclined roller axis, skew angle = 20°,
direction of feed from cup flange to cup iase. 22
Figure 21. Influence of skew angle on buildup formation,
(a) 0°, (b) 10°, (c) 20 °. 22
Figure 22. Influence of contact angle <x of roller on
buildup formation. (a) a = 10°, (b) a = 25°. 22
26
Figure 24. Comparison of calculated angles
of buildup with actual buildup profiles
in forward shear forging. 51
I
wave) changes and becomes more prominent as shear forging progresses
along a particular workpiece. It further appears that the ductility of
the material also plays a role to the extent that soft, ductile materi¬
als have a greater tendency to form such buildup ahead of the roller.
As can be seen from the series of samples shown in Figures 19-25,
such buildup increases with increasing reduction, roller angle, and feed
It decreases with increasing roller tilt angle and roller nose radius.
It has been shown that thereis no appreciable difference in buildup be¬
tween forward and backward shear forging under otherwise similar process
ing conditions, as seen in Figure 25.
(b)
Figure 25. Influence of roller angle, feed rate, and ^
reduction on buildup and distortion ahead of roller.
28
In addition to controlling the relevant parameters to reduce build¬
up, a practical method is to have a second flat roller whose cylindrical
surface is tangent to the outer diameter of the unspun section of the
tube. In this way, any tendency for buildup will be continually sup¬
pressed by the cylindrical roller. It is also possible to design the
working roller to have an approach or depressor angle (i.e., flat land
ahead of the contact angle) to serve the same purpose (Figure 26).
By such techniques, a buildup-free deformation zone can be gener¬
ated, thus minimizing surface defects on the spun surfaces (Figure 27).
The disadvantage in the second technique is that a new roller geometry
is needed for each range of heights, as indicated by AS in Figure 26.
Figure 27. Section through workpiece showing material flow.
(a) Using rollers without approach angle,
(b) using rollers with approach angle. 33
3.3.3 FoAcea -in Shea/i Ponging ofa Tubea
Because of their relevance to the design of shear forging equipment
and components, the three principal forces in this process have been
studied both analytically and experimentally. The technical literature
on this subject is given in References 21, 22, 24, 25, and 32-40. It
appears that, in most studies, the experimental data obtained on forces
have been plotted against process parameters and no effort has been made
to obtain analytical or empirical formulas for forces. Furthermore,
there appear to be no studies published on internal shear forging. (Most
of the theoretical studies on forces have been in shear forging of coni¬
cal workpieces, also known as shear spinning.)
It is not within the scope of this review to present a detailed
theoretical analysis of forces in shear forging of tubes; rather, some
results of previous attempts will be presented, with empirical relation¬
ships given for the benefit of design engineers.
a) External Shtax Fonging . There appear to be two studies in the
literature pertaining to expressions for forces. References 22 and 40.
In the former study, the following formulas are given for the forces in
forward shear forging of tubes:
F^ = ct(A t)f
F, = cs(At) [rf tan a) 1 ^
a
F r = a(At) [rf cot a]%
where
o = average flow stress workpiece
At = bite, i.e., (t 0 - tf), where t 0 = original thickness and tf
= final thickness
r = (DmDr)/(Dm + Dr) where Dm and Dr denote the diameter of the
mandrel and rollers, respectively
f = roller feed rate
a = roller angle.
Subscripts t, a, and r refer to tangential, axial, and radial,
respectively.
31
In the second study, 40 the following formulas have been derived:
^t = Kt0't o f
F a = K a o't 0 [D R f tan a]%
F r = K r o't 0 [D R f cot a]*5
where K^, K a , and K r are parameters to be obtained from Figure 26, and
o' = 1.15 ao, where o 0 is the yield stress of the workpiece material.
Figure 28 also shows the results of the two methods of analysis
that have been used--namely, strip solution and slipline solution. It
also shows the results for backward shear forging of tubes.
Figure 28. Force coefficients in forward and
backward shear forging . 40
The two sets of formulas are essentially the same and are in
approximate agreement with each other when reasonable values are substi¬
tuted, using the strip solution for the second set of formulas.
The foregoing studies do not include the friction and redundant
work during deformation. Furthermore, they are based on an Ideal contact
between the roller and workpiece (at an angle, a) but neglect the effects
of the roller nose radius and relief angle. All of these neglected fac¬
tors could play a significant role on the magnitude of forces.
In experimental studies, it has been shown 41 that when shear forging
aluminum, mild steel, and stainless steel, the actual tangential and
radial forces are about twice those predicted by theory, whereas the
axial force agrees well with theory for backward shear forging and is
about 50% higher than the predicted values for forward shear forging.
As a guide for estimating actual forces. Figures 29 and 30 may be
used for backward and forward shear forging of tubes, respectively. They
are based on experimental data and should be used only as guides because
the influence of many process and material parameters can be significant.
Figure 29.
Force factors vs. thickness reduction in
backward shear forging.
Percent Reduction
Figure 30. Force factors vs. thickness reduction in
forward shear forging . 41
33
fa] Internal Shea*. FoAg-cng . As stated earlier In this review, there
appear to be no studies reported on forces in internal shear forging of
tubes in the literature. Because the deviation of forces have, in ex¬
ternal shear forging as well as in many other metalworking processes,
been largely based on the magnitude of the contact area between the tool
and the workpiece, it is possible to obtain an approximate expression
for internal shear forging forces.
The projected axial areas of contact between the roller and the
tube for the two cases are shown in Figure 31. The bite is given by At,
and the significant dimension is the length of the arc ab. Assuming that
At « Dr or Dm, it can be shown that the length ab is given by.
and
3b ext ~ t DRAt/( 1 + Dr/Dm )]h
ab int * [D R At/(l - Dr/D m )]%
If one further assumes that, for the same feed (f), bite (At), and
roller angle (a), the contact area will be directly proportional to the
length ab, then the ratio of forces for internal to external shear
forging will be,
F int /F ext = 1(1 + °R/°M)/(1 -
For the present study where Dr = 216 mm or 8.5 in. and D M = 406 mm
or 16 in., this ratio will then be 1.8. Thus, the forces in internal
shear forging would be nearly twice as much as for the external case
under otherwise identical conditions.
34
4. 2014 ALUMINUM ROLLING EXPERIMENTS
The principal aim of this task (Task 1) of the Basic Effort was to
study the deformation processing characteristics of aluminum alloy 2014,
to establish the starting condition of the alloy, and to predict its
response to deformation during internal shear forging.
4.1 TASK OBJECTIVES
In shear forging, the material is subjected to a reduction in thick¬
ness in an incremental fashion similar to the rolling process. Further¬
more, because the current method of manufacture for these missile structures
utilizes rolled stock, it was decided to study the rolling characteristics
of this alloy and to establish, if possible, the desired processing param¬
eters for internal shear forging. Both material-related variables and
processing variables were considered in the exploratory study. The material
related variables included alloy temper; size and thickness of the work-
piece; starting, intermediate, and final microstructures; and the presence
and distribution of secondary phases. The process-related variables
involved selection of parameters like reduction per mass, rolling temper¬
ature, lubricant, rolling speed, number of rolling passes, heat treatment
between passes, and post-rolling heat treatment. In selecting the range
of these variables, due consideration was given to the interrelationships
between temperature, reduction, and strain rate as these variables directly
influence the reduction behavior and final properties of the rolled material
Initial roll temperature, temperature and thickness of the starting work-
piece, and the heat generated during plastic deformation also influence the
final properties and microstructure of the rolled product.
The alloy 2014 is a heat-treatable age-hardening alloy and contains
A1 with Cu, Mg, and Si as the main alloying elements. Addition of Si
enhances the response to artificial aging (T6 temper), with the final
strength being higher than for the naturally aged (T4) condition. This
36
alloy is widely used in structural applications. For deformation process¬
ing, it is desirable that the alloy be in a fully annealed (0-temper)
condition. The fully annealed temper yields stabilized precipitate phases
in a matrix which has high ductility and can undergo a considerable amount
of plastic deformation.
For the present experiment, the 2014 alloy could not be procured in
the fully annealed condition. Therefore, a 1-inch thick rolled plate was
procured in the T651 temper (solution treated, stretched 0.5-3%, and then
artificially aged). A 150 x 150 x 25 mm (6 x 6 x 1 in.) plate of the 2014-
T651 plate was then annealed at 413°C (775°F) for 2 hours to achieve the
annealed condition (0-temper) before processing.
Microstructures were examined by optical microscopy as well as scann¬
ing electrom microscopy (SEM). Energy-dispersive X-ray (EDX) analysis
was employed for all identification of various constituent phases. All
microscopy specimens were etched with Keller's reagent. Tensile strength
values quoted here were obtained from the longitudinal (parallel to roll¬
ing) direction unless specified otherwise.
4.2 STARTING MATERIAL
Table 1 gives the tensile properties and chemical composition of the
2014-T651 alloy in the as-received condition. The as-received micro¬
structure displayed elongated recrystallized grains interspersed with
particles of precipitate phases. SEM examination at higher magnification
(1500X) revealed the presence of two kinds of precipitate phases, CuA 1 2
(Figure 32a) and (FeMn) 3 SiA1 22 (Figure 32b). Spot EDX analysis confirmed
the presence of the above phases.
The mechanical properties of the 2014-0 alloy (after annealing for
two hours at 413°C from T651 condition) are given in Table 2 . The indi¬
cated value of 192 MPa UTS is close to the specified value (202 MPa) for
37
TABLE 1. MECHANICAL PROPERTIES AND CHEMICAL COMPOSITION
OF AS-RECEIVED 2014-T651 ALUMINUM ALLOY ROLLED PLATE
Item
Measured
Specification
Mechanical
Properties
UTS
295 MPa (42.9
ksi)
483 MPa (70.0 ksi)
Elongation
6% (in 25 mm)
13% (in 50 mm)
Chemical Composition
(wt%)
Chromium
0.01
0.10 max
Copper
4.46
3.90-5.00
Iron
0.52
1.00 max
Magnesium
0.52
0.20-0.80
Manganese
0.72
0.40-1.20
Silicon
0.91
0.50-1.20
Titanium
0.02
0.15 max
Zinc
0.18
0.25 max
TABLE 2. MECHANICAL PROPERTIES OF 2014-0 ALUMINUM ALLOY PLATE
USED AS STARTING MATERIAL IN ROLLING EXPERIMENTS
(Annealed at 413°C-2 hr) i
Item Longitudinal Transverse
Yield strength 118 MPa (17.1 ksi) 94 MPa (13.6 ksi)
(0.2% offset)
UTS 192 MPa (27.9 ksi) 202 MPa (29.3 ksi)
Elongation (in 14% 20%
25 mm gage length)
Reduction in area
43.9%
34.6%
the 2014-0 condition. Detailed microstructural examination by SEM and EDX
revealed that the cooling after annealing treatment was probably not slow
enough, because in addition to remnants of CuA 1 2 and (FeMn) 3 SiA1i 2 phases,
a very finely dispersed phase was also detected in the matrix. Figure 33
shows a high-magnification (5000X) view of this phase. It was identified
to be a precursor of the CuAl 2 phase. Since this phase precipitated at a
lower temperature, its presence should not interfere with deformation
processing at higher temperatures (150-300°C) where it would go back into
solid solution.
SEM No. 4128 5000X
I
Figure 33. Finely dispersed precipitate phase |
observed after annealing (prior to rolling ).
I
I
40
I
4.3 ROLLING EXPERIMENTS
For all rolling experiments, the workpiece was 150 x 150 x 25 mm
in dimensions and was In the annealed condition. Each experiment involved
seven passes in accordance with the rolling schedule given in Figure 34.
A Fenn type 122 rolling mill was used for the present study. The starting
roll temperature was 66°C (150°F) in all cases. Figure 35 presents a sum¬
mary of all the rolling experiments which were carried out in the present
study. Details of the various parameters used in these experiments and
post-rolling heat treatments to obtain T4 and T6 tempers are also given
in Figure 35.
As indicated in Figure 35, samples for mechanical testing were taken
after the second, fourth, and seventh passes for the 149°C (300°F) and
316°C (600°F) rolling experiments. Some of these samples were mounted,
polished, and etched for metallographic examination to study the micro-
structural characteristics of the rolled specimens. Highlights of the
rolling experiments are presented below.
Table 3 surmiarizes the details of the 149°C (300°F) rolling experi¬
ment which was carried out to evaluate the performance of various lubri¬
cants. Performance of all three lubricants was similar in terms of the
as-rolled surface finish. However, the Aquadag lubricant was found to
be marginally better than the others in minimizing the roll-separating
force as determined by thickness reduction allowing for roll deflection.
Therefore, Aquadag was used in subsequent rolling experiments.
In other to test the effect of rolling speed, a rolling experiment
was carried out at 149°C (300°F) with Aquadag lubricant. At rolling '
speeds of 0.4, 1.2, 2.5, and 4.2 m/s (75, 245, 490, and 817 fpm), a
thickness reduction of 95% was obtained after nine passes in each case.
These rolled specimens were mechanically tested in the as-rolled condi¬
tion and also after T4 and T6 aging treatments (Figure 35). Table 4
gives the results of these experiments. As expected, T6 specimens show
higher UTS for all speeds. Over the range of rolling speeds studied,
the rate of deformation, (i.e., rolling speed) seemed to have no signifi¬
cant influence on the mechanical properties.
41
3-38 mm
5th pass
30%
red.
(0.133 in,
Testing
Testing
The above numbers indicate planned reductions only;
actual values differ slightly.
Percentage reduction decreases progressively in accordance with
schedule for internal shear forging, which requires the percentage
reduction to be roughly proportional to the blank thickness.
Figure 34. Reduction sequence for rolling of 2014-0 aluminum alloy.
i
2014-T65T
(Table 1)
Rolling Schedule (Fig. 34)
A) Effect of Blank Temperature
149°C Blank Temperature
B) Effect of Lubricants
1) Aquadag
2) Deltaforge
3) Ultra Graphite Coat
(Table 3)
C) Effect of Rolling Speeds
1) 0.4 m/s (75 fpm)
2) 1.2 m/s (245 fpm)
3) 2.5 m/s (490 fpm)
4) 4.2 m/s (817 fpm)
(Table 4)
D) Effect of Rolling Reduction
1) 2 passes (69%)
2) 4 passes (87%)
3) 7 passes (95%)
(Table 5)
1) 2 passes (69%)
2) 4 passes (87%)
3) 7 passes (95%)
(Table 5)
316°C Blank Temperature
Reduction
Mechanical Testing
1) As-rolled
2) T4 (500°C-45 min,
W.Q., age at R.T.)
3) T6 (500°C-45 min,
W.Q., age 8 hr at 170°C)
Figure 35. Summary of rolling experiments for 2014 aluminum alloy.
TABLE 3. ROLLING EXPERIMENTS WITH 2014-0 ALUMINUM ALLOY WITH DIFFERENT LUBRICANTS
I
TABLE 4. EFFECT OF ROLLING SPEED AND HEAT TREATMENT
ON MECHANICAL PROPERTIES OF ALUMINUM ALLOY 2014
UTS, MPa (ksi)
Elongation
in 25
mm, %
Testing
0.4
1.2
2.5
4.2
0.4
1.2
2.5
Condition
m/s
m/s
m/s
m/s
m/s
m/s
m/s
m/s
As-rolled
405
323
328
369
3.6
7.6
6.0
5.6
(58.7)
(46.8)
(47.5)
(53.5)
T4
356
403
335
339
11.8
12.4
14.0
9.6
(51.6)
(58.4)
(48.6)
(49.1)
T6
396
458
483
379
10.4
14.8
5.6
11.6
(57.4)
(66.4)
(70.0)
(55.0)
Notes:
1. Aquadag was used as lubricant.
2. 95% total rolling reduction (in 9 passes).
3. 149°C (300°F) blank temperature at start of each pass.
4. 66°C (150°F) roll temperature.
5. Test direction was longitudinal, i.e., parallel to rolling
direction.
6. Tensile specimens were 6.4 mm (0.25 in.) width x as-rolled
thickness x (1.00 in.) gage length.
Based on the results presented above, Aquadag lubricant and a rolling
speed of 0.4 m/s (75 fpm) were selected for rolling experiments at 149°C
(300°F) and 316°C (600°F). Figure35 gives the experimental details of
these two rolling experiments, and Table 5 lists the mechanical properties
obtained after rolling at various degrees of reduction and post-rolling
heat treatment. Table 5 shows that thb as-rolled strength increases with
increasing percent reduction at both temperatures. As-rolled specimens
at 149°C exhibit poor ductility, with the elongation decreasing with
increasing reduction. This may be due to an additional strain-hardening
effect because of the presence of a fine precipitate phase in the 2014-0
condition (Figure 33). Reversion of this phase was probably incomplete
at 149°C. Slightly higher UTS values were achieved for 95% reduction in
nine passes at 149°C as shown earlier in Table 4 . After 316°C rolling,
specimens display fairly high ductility. Both for 149°C and 316°C speci¬
mens, solution heat treatment followed by T4 and T6 aging achieved higher
strengths, although the values quoted in Table 5 are slightly lower than
those specified in the literature (UTS: for T4 = 427 MPa or 62 ksi, for
T6 = 483 MPa or 70 ksi).
TABLE 5. EFFECT OF BLANK TEMPERATURE AND ROLLING REDUCTION
ON MECHANICAL PROPERTIES OF ALUMINUM ALLOY 2014
Passes
Total
Reduction in
Thickness,
%
Testing
Condition
UTS, MPa (ksi)
Elongation
in 25 mm, %
149°C a
316°C a
149°C a
316°C a
0
0
Annealed
192
(27.9)
192
(27.9)
14.0
14.0
2
69
As-rolled
229
(33.2)
219
(31.8)
11.7
16.3
T4
328
(47.6)
325
(47.1)
22.2
18.7
T6
372
(53.9)
377
(54.7)
21.8
22.9
4
87
As-rolled
254
(36.9)
332
(48.1)
9.7
13.6
T4
321
(46.5)
345
(50.0)
18.6
18.6
T6
487
(70.6)
392
(56.9)
21.5
20.2
7
95
As-rolled
280
(40.6)
396
(57.4)
3.7
17.9
T4
325
(47.1)
318
(46.1)
12.0
14.0
T6
362
(52.5)
373
(54.1)
—
15.1
Notes :
Roll temperature: 66°C (150°F)
Rolling speed: 0.4 m/s (75 fpm)
Test direction: longitudinal (L)
Specimen geometry: 6.4 mm (0.25 in.) width x as-rolled thickness x
25.4 mm (1.00 in.) gage length
a Blank temperature at commencement of each pass.
46
The mechanical properties achieved in heat-treatable alloys are
determined by microstructural changes which accompany processing and
heat treatment. A study of microstructures is therefore essential at
various stages of the processing sequence. The following section sum¬
marizes microstructural observations made during the course of the
present experiments.
4.4 MICROSTRUCTURAL OBSERVATIONS
Macroscopic samples of annealed and rolled specimens were taken from
the longitudinal, long-transverse, and short-transverse directions and were
examined, after etching, at a low magnification (25X) for visible flaws
like cracks, folds, and inclusions. Metallographic specimens were pre¬
pared from 149°C and 316°C rolled specimens from all three rolling direc¬
tions after 69, 87, and 95% reductions. These specimens were examined by
optical microscopy (100 to 250X) and SEM (up to 1500X) to reveal the grain
structure, material flow, grain fragmentation, and fragmentation and redis¬
tribution of the precipitate phases CuA 1 2 and (FeMn) 3 SiAl 12 which were
present in the 2014-0 condition. EDX analysis was used on SEM specimens
to identify various phases.
Specimens rolled at 149°C displayed elongated and fragmented grains.
The grains were elongated in the direction of working and flattened in the
thickness direction. No pronounced difference was noted in grain structures
after various reduction levels. This may be due to the intermediate anneal¬
ing that was used between passes to restore the temperature of the workpiece
to 149°C. Figure 36a shows the as-rolled structure, in the long-transverse
direction, after 95% reduction at 149°C. With increasing reduction, the
banded precipitate structure became more random, the precipitate size grew
slightly; however, no visible difference was detected in precipitate density.
Both CuA 1 2 (with some Mg) and (FeMn) 3 SiAl 12 precipitates were present.
Figure 36bshows an SEM micrograph at 500X for the same specimen. White
precipitates are the CuAl 2 phase, and the gray angular precipitates belong
to the (FeMn) 3 SiAl 12 phase.
47
. V
Neg. No. 50090
(a)
200X
Figure 36. Photomicrographs, (a) optical and (b) SEM, of as-rolled
structure after 95 percent reduction by rolling at 149°C.
For 316°C rolling experiments, the grain structure tended to coarsen
slightly with increasing reduction. Since 316°C is probably in the re¬
crystallization regime for this alloy, and rolled specimens were reheated
to 316°C between passes, some recrystallization and grain size changes are
expected to occur. Figure 37a shows a typical long-transverse section show¬
ing precipitate and grain structures after 95% reduction at 316°C. EDX
analysis on SEM specimens showed the presence of only the CuA 1 2 phase.
A very fine precipitate was found to develop after 316°C rolling, and it
could be detected only at higher magnifications (Figure 37tJ. As mentioned
before, a similar phase was also detected after annealing. This phase
probably disappears at the 316°C rolling temperature, but reappears upon
cooling from the rolling temperature.
The main purpose for conducting post-rolling heat treatment to obtain
T4 and T6 conditions was to check whether the final rolled product could be
obtained in these high-strength tempers irrespective of the rolling temper¬
ature and percent reduction. The solution treatment for all specimens was
500°C (930°F), and the time at temperature was 45 min. After solution
treatment, specimens were water-quenched and then aged at room temperature
to obtain the T4 temper. For the T6 temper, quenched specimens were arti
ficially aged at 170°C (340°F) for 8 hr.
Microstructural examination of solution treated and aged specimens
from the 149°C experiment revealed that the actual solution treatment
temperature was probably higher than 500°C. This resulted in melting of
the eutectic and solid solution constituents at grain boundaries. Conse¬
quently, upon subsequent aging to T4 and T6 tempers, this brittle grain
boundary phase persisted. Figure 38ashows the worst case of this grain
boundary melting phenomenon in the case of 95% reduction followed by
solution treatment and natural aging to achieve T4 temper. By contrast,
solution treated and aged specimens from 316°C rolling showed normal
recrystallized grains with strengthening phases that were precipitated
49
Neg. 49879
ZUUX
Figure 37. Photomicrographs, (a) optical and (b) SBM, of as-rolled
structure after 95 percent reduction by rolling at 316°C.
Neg. 49959
JOOX
(b)
Figure 38. Optical photomicrographs of samples rolled to 95 percent
reduction, solution treated and naturally aged to the T4 condition,
(a) 149°C rolling temperature (evidence of grain boundary melting
during solution annealing); (b) 316°C rolling temperature (equiaxed
recrystallized grains after T4 treatment).
r
during aging to achieve T4 and T6 tempers. An example of such a micro¬
structure is shown in Figure 38b for the case of 316°C rolling with 95%
reduction and in the T4 condition. Figure 38b shows a long-transverse
section. The presence of equiaxed recrystallized grains is to be noted.
4.5 CONCLUSIONS
The rolling experiments described above lead to the following
important conclusions:
1) It is essential that the starting condition
of the 2014 alloy for bulk deformation
processing be in a completely annealed condi¬
tion in order to achieve maximum ductility
of the matrix with stabilized precipitates.
In the present study, this condition was not
achieved to the desired extent, perhaps
because of the T651 temper of the starting
material.
2) The present experiments indicate at higher
temperatures, deformation rates (by changing
the rolling speed) do not produce a pronounced
change in the appearance of the grain structure.
Shear forging of the proposed part may involve
localized differences in deformation rates, but
the present results indicate that this should
not pose any serious problem.
3) Post-rolling heat treatment experiments demon¬
strate that the utmost care must be exercised
during solution treatment in order to achieve
the desired strength and ductility in the
final product. Grain boundary melting can be
avoided if the solution treatment is preceded
by a homogenizing anneal at a lower tempera¬
ture. Precise temperature control is, however,
extremely important.
4
, \
1
1
i :
52
I
5. TOOLING DESIGN AND FABRICATION
5.1 SHEAR FORGING MACHINE
A heavy-duty engine lathe (LeBlond Model 2516) was converted into
an experimental shear forging machine using the tooling described later
in this section. The lathe specifications are as follows:
• swing over bed and carriage wings 635 mm (25 in.)
• face plate diameter 597 mm (Z3h in.)
■ available number of spindle speeds 36
• feed range 0.114-6.604 mm/rev
(0.0045-0.2600 in/rev)
• spindle speed range 10-1300 rpm
• maximum horsepower 40
5.2 BASIC CONSIDERATIONS IN TOOLING DESIGN
Some of the important aspects of tooling design for internal shear
forging, based on the metalworking requirements imposed on the tooling
components, are discussed below.
5.2 .1 Via ConSiAuction
The high radial force in internal shear forging, particularly when a
roller of small contact angle is used, causes diametral expansion of the
shear forged part inside the die. While this problem could be alleviated
somewhat by using a roller of adequate contact angle, it is doubtful that
it can be avoided altogether. It was necessary, therefore, to design the
die as a split construction that could be separated to enable part removal
and yet retain sufficient strength and rigidity to withstand the forces
generated during shear forging.
It is important to have a smooth surface on the inside of the die to
allow unrestricted axial movement of the shear forged tube during the form¬
ing process and to attain a high degree of surface finish on the outside of
the part. Further, the inside of the die must possess adequate hardness
(HRC 55-60) to withstand the localized contact/indentation type of loading
imparted by the roller in internal shear forging.
5.2.2 RolteJi Design joK Shea*. fogging
The nomenclature used for the rollers was illustrated earlier in
Figure 26. The "nose radius" R is usually equal to the thickness of the
spun part for most workpiece materials, but may be increased as the soft¬
ness of the material increases.
The "contact angle" or "deformation angle" a is responsible for the
relative magnitudes of the radial and axial components of force during
shear forging and thus influences the selection of the size of roller
bearings. Smaller values of a lead to higher radial components of force,
increasing the tightness between the shear forged part and the external
die. Too large a value of a could prove detrimental to surface finish
and smooth flow of the material under the roller.
The "approach angle" or "depressor angle" 6 prevents the workpiece
material from rising ahead of the roller in backward shear forging. The
axial "extrusion" force, which causes the material to bulge out, increases
with increasing a. Therefore, proper selection of the angle B will permit
a greater angle a, resulting in a more efficient operation.
The "relief angle" y and the "land" l are largely responsible for the
surface appearance on the inside of the internally shear forged part.
Small values of y and/or large £ values cause a burnishing effect, result¬
ing in a bright surface finish. However, these may also increase the
pressure between the tube and the die and cause local buckling in the
walls o'f the tube.
The "clearance angle" 6 is provided in order to keep the length of
land small and, at the same time, to have enough material of roller behind
the deformation zone to withstand the shear forging forces.
54
The value AS determines the maximum possible reduction in thickness
per pass. The actual reduction in thickness (t Q - t^) should be just
marginally smaller than AS in order to utilize the angle 8 effectively
(Figure 26).
5.2.3 VeAi.gn PaAamcteAA Affecting SuA^acc Fin.l&h.
The following factors have been shown to influence the overall sur¬
face finish of the end product:
a) contact angle, a
b) nose radius, R
c) relief angle, y
d) surface finish of roller
e) feed rate
f) speed of shear forging.
These factors are discussed below.
a) Contact angle, a . A roller with a very small contact angle will
cause high radial forces to develop that tend to expand the tube radially.
Being prevented from doing so by the presence of the rigid external die,
the shear forged tube could buckle circumferentially in the thin-walled
region, particularly during the final pass, and cause a wrinkled surface
on the final part.
Too large a value for a would increase the severity of distortion
ahead of the roller in the deformation zone and could damage the inside
surface of the tube.
Contact angles of about 30° have been found to result in good per¬
formance during shear forging.
b) Note siadiu&, R . It has been found that leaving the nose as a sharp
corner causes a rough finish on the end product. On the other extreme,
very large values of R tend to reduce the effective contact angle, increas¬
ing radial forces and the likelihood of circumferential wrinkling. As a
rule of thumb, nose radii are selected to be equal to the thickness of
the tube (after shear forging) for most materials, with sharper radii for
stronger materials and somewhat larger radii for softer materials.
55
c) Relief angle., y . This angle influences the inside surface finish
of the shear forged part more than any other roller parameter. The sur¬
face left behind by the roller is a replica not of the leading edge but
rather the trailing edge of the roller. There is a small amount of
elastic recovery of the tube after deformation, and a small relief angle
of 1-2° allows this recovery to take place in contact with the trailing
edge. A value of y = 0° implies that this recovery will take place after
the material passes the trailing edge with no contact with the roller it¬
self. The same is true if y is larger than 2°, and in such cases the sur¬
face of the tube is poor. A y value of 1.5° would be near optimum, and
this value was used in designing the shear forging roller.
By similar reasoning, the length of land £ should be about 3/16 in.
for optimum results. The length of the approach zone and the magnitude
of the approach angle 8 depend on the tendency of the material to build
up ahead of the roller, and they, too, govern surface finish. Optimum
choice of these variables depends on the values of a and R.
d ) Rolled AoArface jjjuAh , A fine microfinish on the trailing edge of
the roller is necessary in order to have smooth surface finish on the in¬
side of the tube. While an ordinary grinding operation may prove adequate
for the rest of the roller, the length of land on the trailing edge requires
additional polishing with diamond paste to provide the necessary degree of
finish (2-3 yin. rms).
£) Feed naZz . The role of feed rate on surface finish is more predom¬
inant in the case of shear spinning of cones where the roller consists of
a true radius throughout the deformation zone. For shear forging rollers
of the type shown in Figure 26, higher feeds would not result in wavier
surfaces. Even at high feeds, these rollers should provide satisfactory
finish, all other parameters being satisfactory.
$) Speed o& ih&aA faxging . The volume rate of deformation (dV/dt,
in 3 /min) can be expressed in terms of the die ID (D M , in.), the original
thickness of the tube (t Q , in.), the feed rate (f, in/rev), and the spindle
rpm (N) as:
56
dV
dt
= 7rD M t fN
M o
The product fN is the axial velocity of the roller. For the same feed
rate, a higher rpm means a greater rate of deformation of the material.
The upper bound N for the process is governed by the horsepower available
and the maximum permissible temperature rise of the workpiece.
5.3 INITIAL TOOLING DESIGN AND TRIALS
The tooling designed initially for shear forging the subshells is
shown schematically in Figure 39. The die is chucked to the lathe spin¬
dle, and the overhang is supported by the steady rest. The workpiece,
in the form of a 2014-0 aluminum ring, is located inside the die, rotated
in conjunction with the die, and deformed by the axial traversal of the
shear forging roller.
Assembly procedures were established to result in accurate alignment
of the rotating parts with the axis of the lathe. The experimental setup
is shown in Figure 40. Initial trials were restricted to observations of
the principal motions involved in shear forging, i.e., the die rotation,
roller rotation, roller axial feed, and roller radial feed.
The shear forging roller is shown assembled to the roller arm in
Figure 41. The fine surface finish on the outside of the roller is to
be noted. Figure 42 shows a side view of the die holder (extreme left),
the die (the five peripheral holes are for shoulder screws holding the
two halves of the die together), the steady rest (center) which supports
the die, and the shear forging roller (extreme right).
During operation (Figure 43), the aluminum alloy workpiece (1) is
first preheated to the forming temperature by an oxyacetylene torch (2)
and is then internally shear forged by the axial movement of the shear
forging roller (3) at a preset reduction (allowing for deflection). The
loau is transmitted through the workpiece to the rotating external die
(4), whose deflection is restrained by a three-roller steady rest (5).
The oxyacetylene torch is kept on during the operation to maintain the
forming temperature during roller traversal.
57
die from high~pr>
oilers (5)■
It was seen during the initial trials that adequate shear forging
reductions could not be taken on account of excessive deflection of the
tooling away from the preset reduction. Further, even when this deflec¬
tion could be contained, the steady-rest rollers indented the die quite
deeply (see arrow in Figure 43), preventing significant reductions from
being taken. The maximum deformation that was permitted by the original
tooling is seen in Figure 44.
Neg. No. 5184!
Figure 44. 2014 aluminum alloy ring, partially deformed by
internal shear foraing. The experiment was interrupted
due to excessive deflection and die indentation by
steady rest rollers.
61
5.4 TOOLING MODIFICATIONS
To minimize deflections and enable the lathe to perform capably under
the high forces generated during shear forging, the loading and support
members of the tooling were "boot-strapped" together (Figure 45). The
loading path thus generated would go from the roller arm directly to the
nterconnected support arm, with load transfer to the lathe possible only
in the event of gross misalignment.
The modified tooling assembled in IITRI's LeBlond 2516 engine lathe
is shown in Figure 46. The principal components are the die (A); a sup¬
port roller (B) mounted on the support roller arm (C); a shear forging
roller (D) mounted on the shear forging roller arm (E) which pivots with
respect to the support roller arm about the swivel pin (F). Wall thickness
reductions can be set by tightening or loosening the adjustment screw (G)
at the aft end of the roller assembly. A set of oxyacetylene torches (H)
provides workpiece heating during shear forging, while the carriage arm (I)
takes up the benidng moment due to tangential force and prevents carriage
lift-off during operation.
With the above system, the lathe experiences only the axial and tan¬
gential components of force (both relatively small). The large, radial
component of force is taken up by the tooling assembly itself.
This experimental setup was successful in accommodating the forming
loads without deflecting significantly. The shear forging and support
roller arm could each be expected to deflect radially away from the work-
piece and die by approximately 0.25mm (0.010 in.), under maximum operating
conditions, for a total deviation, from the preset reduction, of 0.50 mm
(0.020 in.). The large deflections of up to 6 mm {k in.) previously en¬
countered, due to the lack of support rigidity between the steady rest and
the lathe bed, were thus largely eliminated.
In addition, a complete set of rollers was fabricated (Fig. 47) for
shear forging, rib forging, and for blending the rib-forged section to
the shear-forged section.
62
Improved tooling des
6. INTERNAL SHEAR FORGING OF SUBSHELLS
6.1 PROCESS DESCRIPTION
Using the modified tooling described in the foregoing section,
thick-walled 2014-0 rings were successfully shear forged into thin-
walled internally ribbed tubes. Thermomechanical treatments were evalu¬
ated by introducing cold work at various stages of the precipitation
hardening cycle for the workpiece material. The processing sequence is
shown for one TMT cycle in Figure 48.
1) Sa + Wq + aa 2) Shear forge 3) Rib forge
" fwd end fwd end
4) Shear forge
aft end
rH
> - \
1 i
i
T
l
. -j-
!
-1
n
i
j
H
i — —\
. _-
i
1
i
1 —. —
1
U
i
1_U_ I I_ U- 1 I--- ; -l
5) Rib forge 6) Trim and 7) aa^ 8) inspect
aft end undercut rib
Figure 48. Process elements in internal shear forging
of 2014 A1 subshells.
The deformation processes involved are illustrated schematically in
Figure 49. Lubrication during the initial passes--which were made at
approximately 177°C (350°F)--was accomplished using a graphite (Aquadag)
spray on the workpiece, Figure 50. Figure 51 shows the process of ribs
forging to orthogonalize the ribs. Solution treatment was accomplished
in a Globar-heated electric furnace (Figure 52) at 55°C (930°F). The
66
Figure 51. Adjusting screw being turned for rib forging
final passes following solution treatment and different preaging treat¬
ments were made at room temperature (to prevent overaging) using a light
oil for lubrication. Figure 53 shows the final, thermomechainically
Figure 53. Thermomechanically treated, internally ribbed tube
being removed from the die.
6.2 METAL FLOW
The various types of metal flow encountered in internal shear
forging are shown in Figure 54. Figures 54a and 54b illustrate a near'
uniform flow of metal under the roller (shown cross-ruled above the
workpiece section) with no folding of metal ahead of the roller. This
type of metal flow typically occurred after the first two or three
passes, with gradual heating and increased plasticity of the workpiece
enabling higher percent reductions to be taken. In the initial passes,
the metal flow was usually less uniform because of loading limitations
and insufficient percent reduction in the wall. An extreme case of
nonuniform flow is shown in Figure 54c, with intense shearing of the
surface layers of metal, and formation of dead-metal zones, folds, and
laminations. In contrast, the rib forging operation that was performed
to orthogonalize the rib section produced uniform flow of metal under
the roller (Figure 54d) despite some shaving of the vertical face of the
rib due to friction.
Numerous structural defects were observed in the early trials, which
had to be studied and resolved. Their causes and remedies are discussed
below, in decreasing order of the frequency of occurrence of such
defects.
a) Foldi on OutAlde. Solace . The compressive stress field in the
deformation zone is concentrated directly under the shear forging roller
since the area of contact is limited in the case of the roller-workpiece
interface. At the workpiece-die interface, the same compressive load is
distributed over a larger contact area. The end result is that the
majority of plastic deformation is confined to the workpiece layers
immediately below the workpiece, especially for low initial reductions.
Consequently, the axial elongation of the part is more pronounced,
in the early passes, in the inside surface layers of the workpiece, re¬
sulting in a bulging out of ID material. Figure 55. During subsequent
passes, the bulge produced on the ID is folded back to the OD (die sur¬
face) resulting in a discontinuity on the outside surface of the shear
forged part. Fortunately, this fold is generally formed at the tail end
of the final skin and is removed by trimming about 50 mm (2 in.) off
either end of the finished component.
b) Lamim£ionA on Insj.dc SuAfiace . Another effect of inhomogeneous
deformation is the buildup observed ahead of the roller. This in itself
is usually not detrimental to the part. However, in extreme cases or
when rib forging is performed to square the rib section, previously
built-up layers become flattened to produce foil-like laminations on the
ID surface, Figure 56.
71
to inhomogeneous deformation,
nd on an actual part (b).
i
I
Lamination
Roller
Bu i 1 dup
orkpiece
Neg. No. 52614
Figure 56. Lamination defect, caused on the ID surface during
rib forging by foldover of buildup from previous passes,
shown schematically (a) and on an actual part (b).
This can be avoided (a) before the fact by undertaking a cleanup
cut on the rib area prior to rib forging, or (b) after the fact by re¬
moving the laminations with emery cloth and following up with a final
shear forging pass.
c) Cracking . Cracking of the outside surface of the wall, Figure 57
is due to secondary tensile stresses in the axial direction generated by
the wedge action under the roller. This phenomenon is usually evident
when the roller is plunged into the material radially, when the material
is formed cold or in the T6 temper, when the clearance angle on the
roller land is low, or when deep impressions are left by the turning
operation during preform preparation which later act as stress raisers.
d) Buckling . In backward shear forging, buckling ahead of the rol¬
ler is the result of high axial load which, in turn, is caused by high
reduction and feed rate. Figure 58. In addition, for the particular
case of internal shear forging, buckling was also observed in the reduced
section behind the roller (Figure 58). This is attributed to diametral
expansion of the reduced section within a rigid die, which sets up a
compressive hoop stress in the thin skin and causes the latter to buckle.
e.) Void* . In one instance, examination of a finished part revealed
a through-hole in the reduced wall section. Figure 59. The cause of
this void formation is not clear. Two reasons proposed are the presence
of a hard inclusion in the preform of greater size than the final thick¬
ness of the part, and the entrapment--ahead of the roller or on the die--
of shaved particles of buildup material or other forms of wear debris.
Shaving of metal by the wiping action of the roller is not discussed
separately since it is not believed to produce any structural damage by
itself. Nevertheless, it can be detrimental to the ID surface unless
the shaved particles are periodically removed from the vicinity of con¬
tact between the roller and the workpiece.
74
(a)
Neg. No. 52613
(b)
Figure 57. Cracking of wall OD from secondary tensile stresses due to
high radial feed, low temperature, inadequate roller clearance, and
rough machining marks on preform OD, shown schematically (a) and on
an actual part (b).
75
6.3 SHEAR FORGING OF 2014 A1 AND 2024 A1 SUBSHELLS
In addition to 2014 A1 (program alloy), a few tests were conducted
with 2024 workpieces fabricated from plate by roll-bending and welding.
In contrast, all the 2014 A1 workpieces were ring-rolled to provide a
weld-free structure.
The strength, elongation, and hardness values for 2014 and 2024
aluminum alloys after internal shear forging are given in Table 6. The
as-formed skin showed decreasing strength with increasing wall reduction.
This is not the expected result of softening during mechanical working,
since the preform was fully annealed to begin with and the processing
was conducted below the recrystallization temperature. However, it may
be an evidence of the Bauschinger effect, since the tensile tests re¬
versed the polarity of the stress field present during forming.
The strength levels after heat treatment to the T6 condition showed
no significant effect of deformation behavior, the yield and tensile
strengths being in accordance with handbook data for these materials.
Surface roughness measurements on parts produced by internal shear
forging (Table 7) indicate that the surface finish on the ID (20-65 yin.
AA) is consistent with standard cold rolling practice. Figure 60.
The ID finish is dictated by the finish on the roller (3 yin.), and
by the feed rate and roller geometry. For the case of forward shear
forging (samples 9 and 10), no land was provided on the roller, resulting
in a rough surface (150 yin. AA).
The finish on the 00 is generally a replica of the die contact sur¬
face. In the present case, the die was finish-turned to approximately
16-32 yin. AA, but subsequent use resulted in localized roughening to
32-63 yin. Consequently, the OD surface finish is seen in Table 7 to
vary from 15 to 80 yin. AA.
In addition, entrapment of graphite lubricant on both surfaces of
the workpiece could be responsible for considerable scatter in surface
roughness measurements, as seen in Table 7.
78
TABLE 6. MECHANICAL PROPERTIES OF INTERNALLY SHEAR FORGED SAMPLES
Identi-
fication
No.
ISF
Reduction,
%
Yield
Strength,
MPa (ksi)
UTS,
MPa (ksi)
Elongation
in 25 mm,
%
Hard¬
ness,
HRB
2014-F a
l c
0
261
(37.9)
352
(51.1)
8.0
33
2 d
90
208
(30.2)
270
(39.1)
10.0
28
2014-T6 b
3 C
0
439
(63.6)
492
(71.3)
11.0
84
4 d
90
419
(60.8)
521
(75.5)
10.0
82
2024-F a
5 e
0
197
(28.5)
304
(44.1)
11.0
45
6 d
93
188
(27.3)
263
(38.1)
12.0
30
7 d
95
179
(25.9)
230
(33.4)
5.0
22
2024-T6 b
8 e
0
385
(55.8)
481
(69.8)
11.0
85
9 d
93
350
(50.8)
439
(63.6)
14.0
78
10 d
95
377
(54.7)
456
(66.1)
8.0
78
Notes:
1. Handbook properties are as follows:
2014-T6: 414 MPa (60 ksi) yield; 483 MPa (70 ksi) UTS;
13% elongation
2024-T6: 393 MPa (57 ksi) yield; 476 MPa (69 ksi) UTS;
10% elongation
2. Forming temperature was 150 o -200 o C in all cases.
a Tested in as-internally shear forged condition.
b Tested after solution annealing, water quenching, and artificial aging
to the T6 temper.
c Test direction was along the circumference of the internally shear
forged tube; sample dimensions were 6.4 mm (k in.) dia. x 25 mm (1 in.)
gage length.
d Test direction was along the length of the internally shear forged
tube; sample dimensions were (as-forged thickness) x 6.4 mm width x
25 mm gage length.
e Test direction was along the length of the internally shear forged
tube; sample dimensions were 6.4 mm dia. x 25 mm gage length.
79
TABLE 7. SURFACE FINISH OF INTERNALLY SHEAR FORGED PARTS
Feed
AA Surface Roughness,
um (pin.)
Spacing,
ID
0D
0D
No.
ISF Mode
mm
(in.)
(axial)
(axial)
(circumf.)
1
Backward
2.54
(0.100)
0.5
(20)
0.8 (30)
—
2
Backward
1.91
(0.075)
1.7
(65)
1.3 (50)
—
3
Backward
1.52
(0.060)
1.0
(40)
0.9 (35)
—
4
Backward
2.03
(0.080)
0.5
(20)
1.1 (45)
2.0 (80)
5
Backward
1.52
(0.060)
1.1
(45)
1.4 (55)
--
6
Backward
1.27
(0.050)
0.5
(20)
1.5 (60)
1.3 (50)
7
Backward
2.54
(0.100)
0.9
(35)
1.1 (45)
1.5 (60)
8
Backward
2.16
(0.085)
0.5
(20)
1.8 (70)
--
9
Forward
0.89
(0.035)
3.8
(150)
0.5 (20)
0.4 (15)
10
Forward
0.64
(0.025)
3.8
(150)
1.0 (40)
1.9 (75)
Note: In all cases, the workpiece material was 2014 aluminum,
internally shear forged at 150°-20Q°C, 50 rpm (1.0 m/s) speed, 0.13 mm/
rev feed rate for final pass, and 1 mm final thickness.
\
/
/
Neg. 52616
Figure 60. Inside surface finish of part produced by ISF/RF.
80
6.4 THERMOMECHANICAL TREATMENT OF 2014 AT SUBSHELLS
Thermomechanical treatment (TMT) is a means of improving the
properties of certain alloys by introducing plastic deformation into the
heat-treatment cycle. The microstructural changes accompanying the pro¬
cess are different from those in conventional processing. Depending on
the alloy and the nature of the TMT, this can result in improved tensile
strength and elongation, better fatigue, creep, and wear resistance, and
higher levels of fracture toughness and stress corrosion resistance.
Reviewing Soviet progress in this area, Azrin et al . 40 detailed the
following findings for precipitation-hardenable aluminum alloys:
1) Plastic deformation accelerates the precipitation reaction
in these alloys, due to the hereditary effect of prior de¬
formation of subsequent precipitation, and also the dynamic
effect of precipitation accompanying deformation. This
leads to finer, more homogeneous precipitation than possible
in conventional processing (without deformation).
2) In high-strength aluminum alloys, stress corrosion resis¬
tance can be enhanced by partially aging (artificially) the
solutionized material prior to deformation, followed by
final aging. Homogeneous precipitation occurs within the
grains, by this form of double aging, at the expense of
grain-boundary precipitation. In addition, continuous
boundary precipitates are fragmented, resulting, overall,
in improved resistance to grain-boundary fracture from stress
corrosion. .
3) In many aluminum alloy systems, TMT can lead to reduced
strength from overaging as a result of the accelerated kinet¬
ics of the final aging operation. Higher overall proper¬
ties have been obtained with Al-Cu-Mg, Al-Mg-Si, and Al-Mg-Zn
systems when using a TMT sequence of solutionize, cold work,
and age. The double-aging sequence of solutionize, age, cold
work, and age has been applied successfully to the above
alloy systems and also to Al-Cu-Li alloys. The double-aging
sequence generally leads to better resistance to stress cor¬
rosion cracking. In many cases, double aging also produces
an enhanced rate of work hardening during deformation to
result in dramatic increases in strength after final aging.
4) In some alloys, fracture toughness is enhanced due to a
distorted grain-boundary structure resulting from TMT; this
structure changes the mode of fracture from intergranular
to transgranular.
81
Six different TMT cycles were evaluated in this program (Figure 61).
They differed from one another primarily in the stage of the precipitation¬
hardening sequence in which cold work was introduced. In addition, the
time and temperature of preaging (after solution treatment, before cold
work) was varied, wherever possible, to investigate the precipitation
kinetics to a limited extent.
In all cases, the starting material was a ring-rolled and fully an¬
nealed ring of 2014 aluminum (Figure 62). The microstructures resulting
from the various TMT cycles are shown in Figures 63-68. In general, cold
work following solution treatment is seen to fragment the grains, with the
most dramatic fragmentation occurring when cold work is imposed on a fully
aged, T6 structure.
The tensile properties measured under these conditions are given in
Tables 8 and 9. On the basis of the data shown in Table 8, cycle 6--
which is a double-aging TMT cycle--was selected for production of the de¬
liverable subshells. Two other cycles--cycle 2 (T6) and cycle 4 (T9)--
were selected for preparation of samples for further testing by MIRADCOM.
In general, cold work appears to hasten the precipitation-hardening
kinetics of 2014 aluminum (Table 9). The final aging temperature thus has
to be lower than for conventional artificial aging to the T6 condition to
avoid overaging.
The tensile test data presented in Tables 8 and 9 indicate that the
response of 2014 aluminum to thermomechanical treatment is only marginal
insofar as tensile properties are concerned. Although higher strengths
have been attained here than in conventional T6 heat-treatment, it has
been at the expense of ductility (tensile elongation). The effect on
fatigue strength, fracture toughness, and stress corrosion cracking would
be evaluated by MIRADCOM, if warranted, external to this program.
82
2014-0 RING
Figure 61. Thermomechanical treatment (TMT) cycles evaluated
during internal shear forging of 2014 aluminum.
Neg. Nos. (A) 53047
(T) 53046
(R) 53045
t
R
Figure 62. Optical photomicrographs of 2014-0 aluminum specimen
prior to internal shear forging.
84
Neg. Nos. (A) 53052
(T) 53061
(R) 53060
Fig. 63. Optical photomicrographs of internally shear forged 2014 aluminum
specimen after cycle 1: Warm working, solutionizing and quenching, natural
aging to T4 condition, cold working 40%, and artificial aging at 150°C-10 hr
AD-A102 848
UNCLASSIFIED
I IT RESEARCH INST CHICAGO IL F / 6 20/11
INTERNAL SHEAR FORGING PROCESSES FOR MISSILE PRIMARY STRUCTURES—ETC<U)
JUL 81 S RAJAGOPAL* S KALPAKJIAN DAAK40-78-C-0264
IITRI-M06013-33 NL
i
it
l
1
\
100 pm
— i
> *
/
/ 0
. / %
/
*
A
•f
• * ® *
\j0
/
/
• ✓
/
/
i
i
!■'
I ^
Neg. Nos. (A) 53036
(T) 53037
(R) 53038
R
<
1
Figure 64. Optical photomicrographs of internally shear forged 2014 aluminum
specimen after cycle 2: Warm working, solutionizing and quenching, and
artificial aging at 170°C-10 hr to T6 condition.
' 4
1
i :
1
86
Neg. Nos. (A) 53042
(T) 53043
(R) 53044
T
Figure 65. Optical photomicrographs of internally shear forged 2014 aluminum
specimen after cycle 2: Warm working, solutionizing and quenching, artifi¬
cial aging at 170°C-10 hr (T6), cold working 20%, and artificial aging at
150°C-10 hr.
Neg. Nos. (A) 53065 '
(T) 53064
(R) 53063
Figure 66. Optical photomicrographs of internally shear forged 2014 aluminum
specimen after cycle 4: Warm working, solutionizing and quenching, artifi¬
cial aging at 170°C-10 hr (T6), and cold working 20% to T9 condition.
Neg. Nos. (A) 53041
(T) 53039 v .
(R) 53040
T
R
Figure 67. Optical photomicrographs of internally shear forged 2014 aluminum
specimen after cycle 5: Warm working, solutionizing and quenching, cold
working 30%, and artificial aging at 150°C-10 hr to T8 condition.
fc
¥
100 nm |
r
*
Neg. Nos. (A) 53059
(T) 53058
(R) 53057
♦
T -
A
Figure 68. Optical photomicrographs of internally shear forged 2014 aluminum
specimen after cycle 6: Warm working, solutionizing and quenching, artificial
aging at 170°C-1 hr, cold working 10%, and artificial aging at 150°C-10 hr.
90
TABLE 8. TENSILE PROPERTIES OF INTERNALLY SHEAR FORGED 2014
ALUMINUM ALLOY AFTER DIFFERENT THERMAL-MECHANICAL CYCLES
Test
Cycle* Direction**
Yield
Strenqth
MPa ksi
Tensile
Strenqth
MPa ksi
Elongation
in 25 mm, %
1 A
394
57.1
418
60.6
11
T
404
58.6
434
62.9
9
2 A
442
64.1
458
66.4
8
T
423
61.4
451
65.4
7
3 A
513
74.4
518
75.1
5
T
505
73.3
510
74.0
4
4 A
536
77.7
540
78.3
3
T
506
73.4
507
73.5
2
5 A
359
52.0
424
61.5
8
T
413
59.9
445
64.5
10
6 A
427
62.0
442
64.1
10
T
445
64.5
456
66.2
6
Specimen Geometry: 6
mm
width x
1 mm thickness x
25 mm gage length.
* Warm work 94%, solutionize (500°C-4 hr)
and quench, followed by
1 - Natural age (T4),
cold work (40%),
and artificial age at 150°C-10 hr
2 - Artificial age
at
170°C-10 hr (T6)
•
3 - Artificial age
at
170°C-10 hr (T6)
, cold work (20%)
, and artificial
age at 150°C-10 hr.
4 - Artificial age at 170°C-10 hr (T6), and cold work 20%, (T9).
5 - Cold work (30%) and artificial age at 150°C-10 hr (T8).
6 - Artificial age at 170°C-1 hr, cold work (10%), and artificial age at
150°C-10 hr.
** A - Axial direction of tube.
T - Tangential direction of tube.
91
--
TABLE 9
1
1
. EFFECT OF FINAL AGING TREATMENT ON TENSILE PROPERTIES |
OF 2014 ALUMINUM ALLOY
Tensile
Cycl e*
Final Aging
Test
Direction**
Strenqth,
MPa ks1
Strenqth,
MPa ks1
Elongation
In 25 mm, %
1
150°C-2 hr
A
363
52.7
419
60.8
12
T
394
57.1
430
62.4
8
150°C-10 hr
A
394
57.1
418
60.6
11
T
404
58.6
434
62.9
9
170°C-2 hr
A
352
51.1
393
57.0
6
T
397
57.6
425
61.7
7
170°C-10 hr
A
343
49.8
387
56.1
9
T
422
61.2
434
62.9
5
3
150*0-2 hr
A
436
63.2
492
71.3
3
T
473
68.6
484
70.2
6
150°C-10 hr
A
513
74.4
518
75.1
5
T
505
73.3
510
74.0
4
170°C-2 hr
A
493
71.5
503
72.9
3
T
499
72.3
509
73.8
5
170°C-10 hr
A
483
70.0
496
72.0
4
T
474
68.8
483
70.1
4
5
150*C-2 hr
A
357
51.8
423
61.4
9
T
369
53.5
428
62.1
12
150-C-10 hr
A
359
52.0
424
61.5
8
T
413
59.9
445
64.5
10
170°C-2 hr
A
345
50.0
432
62.6
9
T
414
60.0
445
64.6
9
170°C-10 hr
A
419
60.7
425
61.6
4
T
447
64.8
451
65.4
6
6
150®C-2 hr
A
309
44.8
424
61.5
13
T
372
53.9
421
61.0
8
150"C-10 hr
A
427
62.0
442
64.1
10
T
445
64.5
456
66.2
6
170°C-2 hr
A
356
51.7
430
62.3
12
T
400
58.0
422
61.2
7
170*C-10 hr
A
363
52.6
439
63.6
7
T
446
64.7
456
66.2
6
Specimen Geometry: 6 mm width x 1 mm thickness x 25 mm gage length.
* Warm work 94%, solutlonize (500°C-4 hr) and quench, followed by
1. Natural age (T4), cold work (40%), and artificial age at 150'C-
150*C-10 hr.
3. Artificial age at 170°C-10 hr (T6), cold work (20%), and
artificial age at 150‘C-IO hr.
5. Cold work (30%) and artificial age at 150°C-10 hr (T8).
6. Artificial age at 170*C-1 hr, cold work (10%), and artificial
age at 150*C-10 hr.
92
litfailMMa
G.5 PROCESSING OF DELIVERABLE SUBSHELLS
6.5.1 VeZive.n.ablti Item
Three deliverable subshells were processed according to the schematic
shown earlier in Figure 48. This involved shear forging (170°C), rib
forging (170°C), solutionizing and quenching, artificial preaging at
170°C-1 hr, shear forging (cold) by 10%, final artificial aging at 150°C-
10 hr, and trimming to size. In addition, the skins of an additional sub¬
shell section were processed by T6 and T9 treatments and delivered to
MIRADCOM for metallurgical and mechanical testing.
The three intact subshells are shown in Figure 69.
Neg. No. 53314
0.08X
Figure 69. Deliverable subshells, internally shear
forged with thermomechanical treatment.
93
6.5.2 Residual SVlu&cm
It was observed that internal shear forging with TMT--with the final
passes performed cold on a preaged structure--resulted in considerable re¬
sidual stresses. These stresses were either nonexistent or not apparent
when shear forging annealed parts at warm temperatures.
Figures 70 shows one manifestation of the residual stresses remaining
in the subshell after processing: departure from circular to hexagonal
profile at some distance away from the reinforcing presence of the central
rib. In one case, when an axial slit was made in the thermomechanically
processed subshell, it coiled up into a spiral (Figure 71)--further evi¬
dence of residual stresses being present.
6.5.3 Then R lb G&n&uitlon
The target component depicted earlier in Figure 1 was intended to be
internally shear forged leaving the central region as a single, thick rib.
Grooving the thick rib in a lathe setup would then generate two separate,
thin ribs. However, in producing the deliverable subshells, it was not
possible to machine the rib to the desired specifications, for the reason
given below.
The shear forging die was designed as a split (two-piece) construction
held together by tangential bolts, to facilitate part removal after forming
Under load, the two-piece assembly was progressively distorted (and the
bolts stretched) to result in an oval die and, hence, subshells with ap¬
proximately 1.6 mm (1/16 in.) ovality.
It was decided not to machine the ribs so as to avoid the risk of
breaking through the skin and rendering the subshells unusable.
★
The experiments showed that a split die is not necessary and that by
reverse feeding the roller, the part can be forcibly extracted from the
die. A strong, rigid, single-piece die could thus be used to eliminate
ovality and rib machining problems.
7. PRODUCTION REQUIREMENTS AND COSTS
This section details the equipment and tooling requirements for
internal shear forging of thermomechanically treated 2024 aluminum sub-
shells in production, the cycle time and production rates anticipated,
and the cost benefits of implementing this technology in production.
7.1 EQUIPMENT AND TOOLING
The capital equipment for internal shear forging consists primarily
of one or more heat-treating furnaces and one or more heavy-duty engine
lathes or, preferably, spinning lathes. The quantity of each of these
items would be determined by the cycle time and the rate of production
desired. Tooling for internal shear forging would be as described earlier
in Section 5, modified as necessary to suit the particular spinning machine
used. Additional items include a lubricant spray system, and oxyacetylene
heating sytem (optional) and material handling equipment.
An annual requirement of 2000 parts calls for one spinning lathe
($200,000) and three heat-treating furnaces ($300,000) for solutionizing,
preaging, and final aging. The total investment in implementing the new
technology in place of the current process is estimated to be $1,000,000.
As seen later, the annual return on this investment is expected to be
$6.86 per dollar invested, for a payback period of 1.7 months.
The shear forging die* rollers, support members, and other items of
tooling are expected to cost $100,000 including rework and replacement
parts for one year's production. Amortization over 2000 parts will result
in a tooling add-on-cost of $50 per part.
The cycle time, standard labor hours per part, and the total cost per
part are discussed below.
As discussed earlier in Section 6.5.3, the production die would be built as
a single-piece unit to eliminate the problem of the distortion and resulting
part ovality. All other items of tooling will be, in principle, similar to
the internal shear forging tooling described in Section 5.4. Engineering
drawings for these components are available at IITR1 (Ref. IITRI-M06013).
97
7.2 PRODUCTION PROCESS SEQUENCE
Internal shear forging of 2014 aluminum subshells with thermo¬
mechanical treatment can be divided into eight process elements
(Figure 48):
1. solution anneal, water-quench, and
partial artifical age (sa + wq + aa p )
2. shear forge forward end
3. rib forge forward end
4. shear forge aft end
5. rib forge aft end
6. trim and undercut ribs
7. final artificial age (aa^)
8. inspect.
7.3 STANDARD HOURS FOR SUBSHELL PRODUCTION
The standard hours per subshell will be computed on the basis of a
throughput time of two hours per subshell, i.e., a production rate of
0.5 subshells per hour.* The eight operations listed above would be con¬
ducted in four stations:
1. heat treatment station (2-man crew)
2. spinning lathe station (2-man crew)
3. machining station (1-man crew)
4. inspection station (1-man crew).
The standard hours per subshell will thus be 2 hours x 6 men
= 12 man-hours. This number must now be verified against the actual
throughput capability of each station.
*0n a two-shift basis, this would satisfy an annual requirement of
2000 subshells.
98
7.3 .1 Hexut Tne/UmeM. Station
The estimated standard hours for heat treatment are 2 hours x 2 men
= 4 man-hours. This is a conservative estimate, as demonstrated below:
Furnace
Time
at
Temp.,
hr
Through¬
put
Time,
hr
Minimum
Furnace
Capacity,
subshells
Actual
Standard Hours
for Loading-
Unloading
1 (Solutionizing)
4
2
2
0.50
2 (Preaging)
1
2
1
0.25
3 (Final aging)
10
2
5
1.25
2.00
In other words, a 2-nan crew can easily handle a throughput of 0.5
subshells/hr at the heat-treatment station for solutionizing, preaging,
and final aging operations.
7.3.2 Spinning Lathe. Station
At this station, a 2-man crew must shear forge and rib forge the
forward and aft ends of the subshell. The standard hours estimated ear¬
lier were 2 hours x 2 men = 4 man-hours. Let us now verify that this is
within the manpower capability.
a.) ¥omaAd. End Shea*, fogging- ~l\\e operations involved in shear forg¬
ing one end of the subshell are as follows:
Pass No.
Thickness, mm Travel, mm Time,
Start Finish Start Finish min
Change roller - 5 min
Flatten taper - 10 min
Change roller - 5 min
Shear forge - 10 min
Forward end rib forging, total time = 30 min/part
Standard hours per part with a 2-man crew * 1.00
100
c) A jt End Shcan. Fogging- -The partially forged subshell Is now
reversed inside the die to expose the hitherto undeformed wall to the
shear forging roller. This Is identical to forward and shear forging
(item a) except that an additional one minute will be spent In removing
the subshell from the shear forging die and reversing its position in
the die. Thus,
Aft end shear forging, total time = 26.87 min/part
Standard hours per part with a 2-man crew = 0.90
d ) A jt End Rib Tosiging - -This is identical to forward end rib forg¬
ing (item b) except that an additional one minute will be required for
subshell removal after the operation. Thus,
Aft end rib forging, total time = 31 min/part
Standard hours per part with a 2-man crew =1.03
c) Actual Standard Houte joK Spinning Lathe Station --As a total of
the time consumed in each of the preceding four operations, the floor-to-
floor time at the spinning lathe station is 113.74 min or 1.9 hours per
subshell, which is within the estimated 2 hours per subshell. The actual
standard hours at this station will be 3.8 per subshell, which justifies
the use of four standard hours per subshell in the economic analysis.
7.3.3 Machining Station
Having shear forged and rib forged the ring into a thin-walled shell,
the ends must now be trimmed to size and a groove machined in the middle
of the thick rib to generate two thin ribs. These operations could con¬
veniently be conducted on the spinning machine setup Itself or, alterna¬
tively, in a separate machining station requiring:
10 min/part for trimming both ends
20 min/part for rib undercutting
for a total of 0.5 standard hours per subshell with a one-man crew. Thus,
the earlier estimate of 1 standard hour per subshell is well within the
manpower capacity for this operation.
101
*«**••• v f \ #>
7.3.4 ln&pzc£Lon Station
The In-plant Inspection tasks would consist of:
Dimensional test:
10 min/part
Hardness test:
10 min/part
Ultrasonic test:
10 min/part
X-ray test:
20 min/part
All others:
10 min/part
Inspection of the subshells thus accounts for one standard hour per
subshell, as estimated previously.
7.4 SUBSHELL PRODUCTION COST OF
INTERNAL SHEAR FORGING vs. CURRENT PROCESS
The labor requirement for internal shear forging of 2014 aluminum
subshells with thermomechanical treatment has been estimated at 12 stan¬
dard hours per part, whereas the current process of fabricating the sub¬
shell as a welded structure was estimated to consume 130 standard hours
per part. 41 The production costs derived from these figures show that
shear forged subshells would cost about one-eighth the cost of a compar¬
able fabricated structure.
The production costs for internal shear forging and the current
process are estimated below, along with the projected annual return on
investment and payback period from the $1 million cost of implementing
the new technology in production.
102
7.4.1 Production Co it--Internal Skeen fogging
Direct labor ($8/hr x 12 hr/part x 2000 parts) $192,000
Manufacturing overhead (175% of direct labor) 336,000
Raw material ($3/1b x 10 Ib/part x 2000 parts) 60,000
Tooling (including rework and replacement parts) 100,000
Consumables (lubricants, cutting tools, and supplies) 10,000
G&A, profit, etc. (35% of above costs) 244,300
Total cost of 2000 shear forged subshells $942,300
Cost per subshell 471.15
7.4.2 Production Coit--Current Pro C.&6A
Direct labor ($8/hr x 130 hr/part x 2000 parts) $2,080,000
Manufacturing overhead (175% of direct labor) 3,640,000
Raw material ($3/1 b x 4 Ib/part x 2000 parts) 30,000
Tooling (welding fixtures) 20,000
Consumables (electrodes, cutting tools, and supplies) 10,000
G&A, profit, etc. (35% of above costs) 2,023,000
Total cost of 2000 fabricated subshells $7,803,000
Cost per subshell 3,901.50
7 .4.3 Return on Investment and Payback Period
The anticipated savings in producing 2000 subshells by internal shear
forging, as opposed to current fabrication, is $6,860,700. Since the 2000
parts would represent one year's production, the return on investment on
the $1,000,000 invested in new equipment and facilities would be $6.86 in
the first year on each dollar invested.
The payback period for the investment would be 1.7 months correspond¬
ing to the production of 392 subshells by the new method.
103
8. CONCLUSIONS AND RECOMMENDATIONS
Using shear forging tooling in conjunction with an engine lathe,
internal shear forging experiments were conducted to establish procedures
for producing missile primary structures as monolithic construction with
integral ribs. This Manufacturing Methods and Technology Program success¬
fully established guidelines for large-volume production of these struc¬
tures and showed the cost savings resulting from implementation of the
internal shear forging processes.
The major conclusions from this program are that:
• Internal shear forging can be implemented to produce
missile primary structures to near-net shape.
• The cost savings accompanying implementation of this
process are substantial.
• Although the strengthening effect of thermomechani¬
cal treatment seems to be minimal for 2014 aluminum
on the basis of tensile data, further testing for
fatigue, fracture, and stress corrosion resistance
is necessary before the effects of TMT can be
confirmed.
• Dimensional stability and residual stresses result¬
ing from TMT must be fully characterized.
Before this cost-effective, near net shape technology is implemented,
however, it is recommended that the thermomechanical aspects of the pro¬
cess be carefully reviewed and evaluated, in terms of the following:
• trade-off between higher property levels (if existent)
and increased processing cost due to multiplicity of
steps,
• influence of process variables on the reproductivity
of end properties and product appearance,
• magnitude, polarity, and distribution of residual
stresses resulting from TMT and their effect on the
dimensions and mechanical properties of shear forged
subshells.
104
REFERENCES
1. C. L. Packham, "Metal Spinning and Shear and Flow Forming,"
Metallurgies and Metal Forming , June 1976, pp. 168-70; July 1976,
pp. 203-206; August 1976, pp. 250-252; September 1976, pp. 281-285.
2. C. Wick, "Metal Spinning: A Review and Update," Manufacturing
Engineering , January 1978, pp. 73-77.
3. Metal Deformation Processing, Vol. II, Defense Metals Information
Center, Battelle Memorial Institute, Columbus, Ohio, Report No.
226, 1966, pp. 49-66.
4. Rotary Metalworking Processes , Proceedings, First Int. Conf., IFS
(Conferences) Ltd., Bedford, U.K., 1979.
5. K. W. Stalker and K. W. Moore, "Cold Power Spinning Saves Material,
Cuts Costs," American Machinist , May 9, 1955.
6. "Power Spinning Conical and Tubular Parts," Product Engineering,
August, 1956.
7. J. Genis and W. Mallindine, "Rotary Extrusion Reduces Costs and
Saves Material," Machinery (NY), April 1958, pp. 115-121.
8. J. H. Peters, "Flow Turning to Increase Strength, Save Weight,
and Reduce Costs," ASME Paper No. 59-A-277, 1959.
9. L. E. Zwissler, "Spinning Makes Stronger Rocket Cases," Metal
Progress, December 1960.
10. "Heavy Flow-Forming," Aircraft Production, November 1962, pp. 374-
376.
11. D. J. Campion, "Spinning and Related Forming Techniques with Par¬
ticular Reference to Maraging Steel," Sheet Metal Industries,
March 1967, pp. 160-166.
12. G. E. Gott, J. M. Lynch, and S. M. Jacobs, "Are Shear Spinning and
Roll Extrusion Production Processes for Large Parts?" Metal Progress,
March 1968, pp. 95-99.
13. "Case Histories Demonstrate Metal Spinning's Virtues," Modern Metals,
June 1978, pp. 46-49.
14. D. H. Pollitt, "The Practice and Potential of Flow Forming Processes,"
in Rotary Metalworking Processes, Proceedings, First Int. Conf., IFS
(Conferences) Ltd., Bedford, U.K., 1979.
15. "Shear Forming of Thin-Wall Seamless Tubes," Machinery, January 1964,
pp. 120-121.
105
16. A. W. Ernestus, "Roll Extrusion, A New Metal-Forming Technique,"
American Machinist , June 29, 1959, pp. 84-86.
17. D. L. Corn, "Roll Extruding Precision Seamless Pipe and Tubing,"
Metal Progress , June 1977, pp. 28-31. See also "Recent Advances
in Roll Extrusion," in Rotary Metalworking Processes , Proceedings,
First Int. Conf., IFS (Conferences) Ltd., Bedford, U.K., 1979,
pp. 243-250.
18. J. M. Steichen and R. L. Knecht, "Mechanical Properties of Roll
Extruded Nuclear Reactor Piping," ASME Paper No. 75-PVP-41, 1975.
19. "Shear Forming: How It Affects Properties," American Machinist,
July 6, 1965, pp. 57-59.
20. E. S. Jones, "Aus-Shear Forming of Low Alloy Steel Cylinders,"
General Electric Co., Cincinnati, Ohio, ca. 1960.
21. S. Kalpakjian, "An Application of Theory to an Engineering Prob¬
lem: Power Spinning," in Fundamentals of Deformation Processing,
Syracuse University Press, 1964, Chapter IV, pp. 71-103.
22. E. Thomasett, "Krafte und GrenzformSnderungen beim AbstreckdrUcken
zyl indrischer, rotationssymmetrisches Hohlkb’rper aus Aluminium,"
Doctoral Dissertation, University of Stuttgart, 1961.
23. S. Kalpakjian, "An Experimental Study of Plastic Deformation in
Power Spinning," C.I.R.P. Annalen, Vol . 10, 1962, pp. 58-64.
24. H. Jacob, "Erfahrungen beim FIiessdrilcken zylindrischer WerkstUcke,"
Fertigungstechnik and Betrieb, Vol. 12, March 1962, pp. 169-178.
25. E. Siebel and K. A. Droge, "Forces and Material Flow in Spinning,"
Werkstattstechnik und Maschinenbau, Vol. 45, January 1955, pp. 6-9.
(See also The Engineers’ Digest , Vol. 16, No. 5, May 1955, pp. 193—
195.)
26. S. Kalpakjian, "Maximum Reduction in Power Spinning of Tubes,"
Trans. ASME , J. Eng. Ind., Vol. 86, 1964, pp. 49-54.
27. C. H. Wells, "The Control of Buildup and Diametral Growth in Shear
Forming," Trans. ASME, J. Eng. Ind., Vol. 90, 1968, pp. 63-70.
28. H. J. Dreikandt, "Untersuchung iiber das Driickwalzen zylindrisher
Hohlkorper und Beitrag zur Berechnung der gedruckten FISche und
der Krafte," University of Stuttgart, 1973.
29. S. Kalpakjian, "Chevron Fracture in Tube Reduction by Spinning,"
in Advances in Research on the Strength and Fracture of Materials,
Vol. 2A, Pergamon Press, NY, 1977.
106
30. A. K. Cruden, Report No. 341, National Engineering Laboratory, 1968;
also M. G. Cockroft in Ductility, ASM, 1968, p. 218.
31. M. Hayama and H. Kudo, "Analysis of Diametral Growth and Working
Forces in Tube Spinning," Bulletin of the Japan Society of Mechanical
E>iginecvs, Vol. 22, 1979, pp. 776-784.
32. M. Hayama and H. Kudo, "Experimental Study of Tube Spinning," Bidle-
tin of the Japan Society of Mechanical Engineers, Vol. 22, 1979,
pp. 769-775.
33. H. Jacob and F. Garreis, "Rollenanordnung und Rollenform beim
FI iessdriicken kreiszylindrischer Korper," Fertigungstechnik und
Bctrieb, Vol. 15, No. 5, May 1965, pp. 279-283,
34. H. Jacob and F. Garreis, "Berechnung der auftretenden Krafte beim
FI iessdriicken zyl indrischer Hohlkb'rper," Fertigungstechnik und
Betrieb, Vol. 14, No. 8, August 1964, pp. 493-497.
35. H. Jacob and F. Garreis, "FIiessdriicken mit Schragstellung der
Rollen und deren Auswirkung und den Werkstofffluss und Umformkrafte
(Dreh-Umformmaschine)," Fertigungstechnik und Betrieb, Vol. 16,
No. 1, January 1966, pp. 42-45.
36. V. I. Eliseev and E. I. Isachenkov, "Selection of Force Parameters
When Flow-Turning Cylindrical Components," Russian Engineering
Journal (Translation of Vestnik Mashinosirceniia) , Vol. 49, No. 11,
1969, pp. 68-71.
37. P. Bennich, "Tube Spinning," Int. Journal Prod. Res., Vol. 14, No. 1,
1976, pp. 11-21.
38. J. A. Bennaton and E. Appleton, "An Experimental Metalforming
Machine and Its Application to Cylindrical Flowturning," Rotary
Metalworking Processes, Ref. 4.
39. T. Rammohan and R. Mishra, "Studies on Power Spinning of Tubes,"
Int. Journal Prod. Res., Vol. 10, No. 4, 1972, pp. 351-364.
40. M. Azrin, G. B. Olson, E. B. Kula, and W. F. Marley, Jr., "Soviet
Progress in Thennomechanical Treatment of Metals," J. Appl. Metal¬
working, Vol. 1, No. 2, 1980, pp. 5-34.
41. J. Long, MIRA0C0M, private communication, 28 January 1981.
107
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