Materials Selection and Design (2010) Part 17 docx

In the second case, P is increased from 0 to 534 kN (120 kips) and then brought back to zero. The entire history of
stresses in the three bars is shown in Fig. 1(b). When P exceeds 400 kN (90 kips), the two outer bars deform plastically
and, because of the reduced modulus, begin to share less load. The stress in the two outer bars follows the path ABCD,
whereas that in the middle bar follows the path ABEF. It can be seen that when P is again zero (unloading), the stresses in
the three bars do not go back to zero. Instead, the middle bar has a residual tensile stress of 78.8 MPa (11.4 ksi), and each
of the two outer bars has a residual compressive stress of 39.4 MPa (5.7 ksi). Because there is no external load on the
assembly, the residual stresses in the three bars are in self-equilibrium. A comparison of the two loading histories
indicates that the presence of inhomogeneous plastic deformation in the three bars is responsible for the generation of
residual stresses. Similarly, mechanical residual stresses occur in any component when the distribution of plastic
deformation in the material is inhomogeneous, such as the surface deformation in shot-peening operation.
Thermal Loads. A similar three-bar model explaining the generation of residual stresses due to inhomogeneous plastic
deformation caused by thermal loads is discussed by Masubuchi (Ref 6, presumably adopted from Ref 7). In this model,
three carbon-steel bars of equal length and cross-sectional area are connected to two rigid blocks at their ends. The middle
bar is heated to 593 °C (1100 °F) and then cooled to room temperature, while the two outer bars are kept at room
temperature. Some of the details are not clearly explained in Ref 6, but the problem is very similar to the previous
example. When the temperature in the middle bar is raised, the requirements of compatibility and equilibrium imply that a
compressive stress be generated in the middle bar and tensile stresses in the two outer bars; the stress in each of the two
outer bars being half of that in the middle bar. If the temperature in the middle bar is so high that its stress exceeds yield
but in the two outer bars the stresses are still below yield, residual stresses will occur in the three bars when the
temperature of the middle bar is brought back to room temperature (i.e., on unloading). Similarly, if the stresses in all
three bars exceed yield but by different amounts, residual stresses will still occur when the temperature of the middle bar
is brought back to room temperature. Indeed, this case is very similar to that of a cylinder immersed vertically in a
quenchant where, during the initial stages of quenching, the temperature in the outer layer is much lower than that in the
inner core.
The three-bar model can be further utilized to explain the generation of residual stresses due to the mismatch in
coefficients of thermal expansion. For example, suppose the two outer bars represent the layers of matrix in a composite
lamina and the inner bar represents a layer of fibers. The coefficient of thermal expansion of the two outer bars is equal
but, in general, different from that of the middle bar. It is assumed that the initial temperature of all the three bars is equal,
which corresponds to a certain processing temperature much higher than room temperature. When the assembly is
brought to room temperature, the requirements of compatibility and equilibrium will be satisfied if a system of forces
(residual stresses) is established such that the sum of the forces in the two outer bars is equal and opposite to that in the

middle bar. In this case, the presence of unequal plastic deformation is not a prerequisite for the generation of residual
stresses. This explains why, while selecting the constituent materials for a composite or for a coating, the designers try to
minimize the mismatch between their coefficients of thermal expansion.
Solid-State Transformation. In quenching, welding, and casting processes, many metals such as steels undergo one
or more solid-state transformations. These transformations are accompanied by a release of latent heat, a change in
volume, and a pseudoplasticity effect (transformation plasticity). All of these affect the state of residual stresses in the
part. The release of latent heat during solid-state transformation is similar to that during the liquid-to-solid transformation,
albeit of a smaller amount. The change (increase) in volume occurs due to the difference in mass densities of the parent
phase (e.g., austenite) and the decomposed phases (pearlite, ferrite, bainite, and martensite). In steels, the volumetric
change due to phase transformation is in contrast to the normal contraction or shrinkage during cooling (Ref 8).
A simple example of transformation plasticity is shown in Fig. 2, which is based on the results of a constrained
dilatometry experiment (Ref 9). The figure shows that during cooling in the phase transformation regime, the presence of
even a very low stress may result in residual plastic strains. Two widely accepted mechanisms for transformation
plasticity were developed by Greenwood and Johnson (Ref 10) and Magee (Ref 11). According to the former, the
difference in volume between two coexisting phases in the presence of an external load generates microscopic plasticity
in the weaker phase. This leads to macroscopic plastic flow, even if the external load is insufficient to cause plasticity on
its own. According to the Magee mechanism, if martensite transformation occurs under an external load, martensitic
plates are formed with a preferred orientation affecting the overall shape of the body.

Fig. 2 Transformation plasticity. Source: Ref 9
Material Removal. A fact that is often overlooked in discussing residual stresses caused by various manufacturing
processes is the effect of material removal on the state of stresses in the product. Consider, for example, that a casting
mold must be finally broken and removed, or a forging die must be retracted. Likewise, in making a machined part some
of the material has to be removed. All of these operations change the state of stress in the part. In order to fully understand
this concept, three examples discussed in Ref 5 should be considered.
The first example entails an assembly of two concentric springs of slightly different lengths, L
i
and L
o
, as shown in Fig.

3(a); the subscripts i and o refer to inner and outer springs, respectively. The bottom ends of the two springs are fixed.
Then, the upper ends are tied to a rigid block that is free to move only in the vertical direction. The two springs adopt a
compromise length, L, which is in between L
i
and L
o
, as shown in Fig. 3(b). As a result, the two springs develop equal and
opposite forces: compressive in the longer inner spring and tensile in the outer shorter spring. The assembly of the two
springs may be viewed as analogous to the assembly of a cast part and its mold or to the assembly of the forged part and
the die, or to a machined part before some portion of it is removed. Then, the removal of the outer spring becomes
analogous to removal of material during machining (Ref 5, 12), of the casting mold (Ref 13, 14, 15, and 16), or of the
forging die (Ref 17). Two cases are considered. In the first case, the stresses in both springs are assumed to be within their
elastic limits. When the outer spring is removed, the force acting on it is transferred to the inner spring in order to satisfy
equilibrium and the inner spring returns to its original length. In the second case, it is assumed that the inner spring has
undergone a certain amount of plastic deformation. When the outer spring is removed, the inner spring does not return to
its original length, L
i
. In either case, because the two springs and, therefore, the forces, are concentric, the residual stress
in the inner spring becomes zero when the outer spring is removed.

Fig. 3 Residual stresses in an assembly of two springs with unequal initial lengths. Source: Ref 5

For the second example, reconsider the three-bar model from the section "Mechanical Loads" in this article. After
creating residual stresses in the three bars by loading and unloading the assembly, bar 3 is removed, by (for example)
machining. As shown in Fig. 4(a) and 4(b), a redistribution of stresses in the remaining two bars takes place. The resultant
stresses at the centroids of the two bars become -14.8 MPa (-2.14 ksi) in bar 1, and 14.8 MPa (2.14 ksi) in bar 2. Also, the
assembly rotates (distorts) by an angle of 4.3 × 10
-3
radians.

Fig. 4 Effect of asymmetric material removal in the three-bar model of Fig. 1. Source: Ref 5
The third example in Ref 5 is of a thick-walled cylinder with an internal diameter of 101.6 mm (4 in.) and an outer
diameter of 203.2 mm (8 in.) as shown in Fig. 5(a). Both ends of the cylinder are restrained axially, and the cylinder is
subjected to an internal pressure. A 25.4 mm (1 in.) thick (along the axis) slice of the cylinder is analyzed by subdividing
it into 10 equal finite elements (5.08 mm, or 0.2 in., thick each) in the radial direction (Fig. 5b). The residual stresses are
created by increasing the pressure from zero to 345 MPa (50 ksi), and then back to zero. The elements 1 and 2 are
removed successively. The variation of the three stress components along the radius is shown in Fig. 6, before material
removal (i.e., the residual stresses) and after removing the two layers. It may be noted that in an overall sense, the level of
residual stresses goes down as the material is removed. However, this is not necessarily true in a local sense. Consider, for
example, the circumferential stress at the centroids of elements 3 and 4 in Fig. 6(b); it increases as the material is
removed.

Fig. 5 Cylinder with internal pressure and its finite element mesh. Source: Ref 5

Fig. 6 Effect of removing layers of elements of material from the inside of the cylinder. Source: Ref 5

Important conclusions from the three examples discussed above can be summarized as follows:
• When the material removal is symmetric with respect to the stress distribution (Fig. 3
), the residual
stresses in the remainder of the assembly or part are very small or even zero.
• When the material removal is not symmetric with respect to the stress distribution (Fig. 4, 6
), the
residual stresses in the remainder of the assembly or part are not necessarily small.
• Material removal may result in an increase in stresses at some locations of the assembly or the part (
Fig.
6).

References cited in this section
5.
U. Chandra, Validation of Finite Element Codes for Prediction of Machining Distortions in Forgings,

Commun. Numer. Meth. Eng., Vol 9, 1993, p 463-473
6. K. Masubuchi, Analysis of Welded Structures, Pergamon Press, 1980, p 94-96
7. W.M. Wilson and C.C. Hao, Residual Stresses in Welded Structures, Weld. J.,
Vol 26 (No. 5), Research
Supplement, 1974, p 295s-320s
8. W.K.C. Jones and P.J. Alberry, "The Role of Phase Transformation in the Developmen
t of Residual
Welding Stresses," Central Electricity Generating Board, London, 1977
9. J B. Leblond, G. Mottet, J. Devaux, and J
C. Devaux, "Mathematical Models of Anisothermal Phase
Transformations in Steels, and Predicted Plastic Behavior," Mater. Sci. Technol., Vol 1, 1985, p 815-822
10.

G.W. Greenwood and R.H. Johnson, The Deformation of Metals under Small Stresses During Phase
Transformation, Proc. Royal Soc., Vol 283, 1965, p 403-422
11.

C.L. Magee, "Transformation Kinetics, Microplasticity and A
ging of Martensite in FE31 Ni," Ph.D. Thesis,
Carnegie Institute of Technology, 1966
12.

U. Chandra, S. Rachakonda, and S. Chandrasekharan, Total Quality Management of Forged Products
through Finite Element Simulation, Proc. Third International SAMPE Metals and Metals Processing Conf.,

Vol 3, F.H. Froes, W. Wallace, R.A. Cull, and E. Struckholt, Ed., SAMPE International, 1992, p M379-
M393
13.

U. Chandra, Computer Prediction of Hot Tears, Hot Cracks, Residual Stresses and Distortions in Precision

Castings: Basic Concepts and Approach, Proc. Light Metals, J. Evans, Ed., TMS, 1995, p 1107-1117
14.

U. Chandra, Computer Simulation of Manufacturing Processes Casting and Welding, Comput.
Model.
Simul. Eng., Vol 1, 1996, p 127-174
15.

U. Chandra, R. Thomas, and S. Cheng, Shrinkage, Residual Stresses, and Distortions in Castings,
Comput.
Model. Simul. Eng., Vol 1, 1996, p 369-383
16.

A. Ahmed and U. Chandra, Prediction of Hot Tears, Residual Stresses and Distortions in Castings Including
the Effect of Creep, Comput. Model. Simul. Eng., to be published
17.

U. Chandra, S. Chandrasekharan, and R. Thomas, "Finite Element Analysis of the Thread Rolling Process,"
Concurrent Technologies Corporation, submitted to Knolls Atomic Power Lab, Schenectady, NY, 1995
Control of Residual Stresses
U. Chandra, Concurrent Technologies Corporation

Computer Prediction of Residual Stresses
In recent years, the finite element method has become the preeminent method for computer prediction of residual stresses
caused by various manufacturing processes. A transient, nonlinear, thermomechanical analysis software is generally
employed for that purpose. Some of the mathematics that form the basis of such software is common for all
manufacturing processes. Such common mathematics is summarized by this section. However, because every process is
unique, some mathematical requirements are, in turn, dependent on the process. Also, for the simulation of certain
processes a sequential thermomechanical analysis is adequate, whereas for others a coupled analysis may be preferred or
even essential. Such subtleties are pointed out later when individual processes are discussed.

Ignoring convection, the following conduction heat-transfer equation is solved with appropriate initial and boundary
conditions:


(Eq 1)
where T is the temperature at an arbitrary location in the workpiece at time t, k is the thermal conductivity of the material,
c
is the rate of heat generated per unit volume, is the density, C
p
is the specific heat, and is the differential operator;
all material properties are assumed to vary with temperature. The term
c
accounts for the release of latent heat during
liquid-to-solid transformation in casting and welding processes or during solid-state phase transformation in quenching,
welding, or casting processes. It also accounts for the heat of plastic deformation in forging and other bulk deformation
processes. The initial and boundary conditions are process dependent. Details of converting Eq 1 into its finite element
form and of numerical solution are available in a number of technical papers and textbooks and are not repeated here. For
a general treatment of the subject, the reader is referred to Ref 18, 19, 20, and 21.
The transient temperatures computed above are used as loading for the subsequent transient stress/displacement analysis.
Using the incremental theory, the total strain increment { } at time t can be divided into various components (Ref 22,
23, 24, 25, and 26):
{ } = {
e
} + {
t
} + {
p
} + {
cr
} + {

v
} + {
tr
}


(Eq 2)
where superscripts e, t, p, cr, v, and tr refer to elastic, thermal, plastic, creep, volumetric change, and transformation
plasticity components, respectively. The first three strain terms are needed in the simulation of every manufacturing
process discussed here, whereas the use of the other three terms is dependent on the process and are pointed out as
appropriate. Also, mathematical details for the first four strain terms are discussed in most standard references (Ref 22,
23), whereas the details for the last two terms are discussed often in the context of the simulation of quenching and
welding processes (Ref 24, 25, 26).
In forging and other large deformation processes, the term
c
in Eq 1 represents the heat of plastic deformation and leads
to a coupling between Eq 1 and 2.
At present, no single computer code is capable of predicting residual stresses caused by all manufacturing processes.
However, several general-purpose finite element codes are capable of predicting these stresses to a reasonable degree of
accuracy for at least some of the manufacturing processes (Ref 27, 28, 29). In addition, some of these codes permit
customized enhancements leading to more reliable results for a specific process. Before attempting to predict residual
stresses due to a manufacturing process, it is advisable to compare the capabilities of two or three leading codes and use
the one most suited for the simulation of the process in consideration. Examples of such comparisons are given in Ref 12
and 13 for forging, quenching, and casting processes. It must be noted that, due to continuous enhancement in these
codes, it is always advisable to compare the capabilities of their latest versions.

References cited in this section
12.

U. Chandra, S. Rachakonda,

and S. Chandrasekharan, Total Quality Management of Forged Products
through Finite Element Simulation, Proc. Third International SAMPE Metals and Metals Processing Conf.,

Vol 3, F.H. Froes, W. Wallace, R.A. Cull, and E. Struckholt, Ed., SAMPE International, 1992, p M379-
M393
13.

U. Chandra, Computer Prediction of Hot Tears, Hot Cracks, Residual Stresses and Distortions in Precision
Castings: Basic Concepts and Approach, Proc. Light Metals, J. Evans, Ed., TMS, 1995, p 1107-1117
18.

G. Comini, S. Del Guidice, R.W. Lewis, and O.C. Zienkiewicz, Finite Element Solution of Non-
linear Heat
Conduction Problems with Special Reference to Phase Change, Int. J. Numer. Methods Eng.,
Vol 8, 1974, p
613-624
19.

R.W. Lewis, K. Morgan, and R.H. Gallagher, Finite Element
Analysis of Solidification and Welding
Processes, ASME Numerical Modeling of Manufacturing Processes, PVP-PB-
025, R.F. Jones, Jr., H.
Armen, and J.T. Fong, Ed., American Society of Mechanical Engineers, 1977, p 67-80
20.

A.B. Shapiro, "TOPAZ A Finite Element Heat Conduction Code for Analyzing 2-
D Solids," Lawrence
Livermore National Laboratory, Livermore, CA, 1984
21.

B.G. Thomas, I. Samarasekera, and J.K. Brimacombe, Comparison of Numerical Modeling Techniques for
Complex, Two-Dimensional, Transient Heat-Conduction Problems, Metall. Trans., Vol 15B, 1984, p 307-
318
22.

A. Levy and A.B. Pifko, On Computational Strategies for Problems Involving Plasticity and Creep,
Int. J.
Numer. Methods Eng., Vol 17, 1981, p 747-771
23.

M.D. Snyder and K J. Bathe, A Solution Procedure for Thermo-Elastic-Plastic and Creep Problems,
Nuc.
Eng. Des., Vol 64, 1981, p 49-80
24.

S. Sjöström, Interactions and Constitutive Models for Calculating Quench Stresses in Steel, Mater.
Sci.
Technol, Vol 1, 1985, p 823-829
25.

S. Das, G. Upadhya, and U. Chandra, Prediction of Macro- and Micro-
Residual Stress in Quenching Using
Phase Transformation Kinetics, Proc. First International Conf. Quenching and Control of Distortion,
G.E.
Totten, Ed., ASM International, 1992, p 229-234
26.

S. Das, G. Upadhya, U. Chandra, M.J. Kleinosky, and M.L. Tims, Finite Element Modeling of a Single-
Pass GMA Weldment, Proc. Engineering Foundation Conference on
Modeling of Casting, Welding and

Advanced Solidification Processes VI, T.S. Piwonka, V. Voller, and L. Katgerman, Ed., TMS, 1993, p 593-
600
27.

"ABAQUS," Version 5.5, Hibbitt, Karlsson and Sorenson, Pawtucket, RI, 1995
28.

"ANSYS," Release 5.3, ANSYS, Inc., Houston, PA, 1996
29.

"MARC," Version K 6.2, MARC Analysis Research Corporation, Palo Alto, CA, 1996
Control of Residual Stresses
U. Chandra, Concurrent Technologies Corporation

Measurement of Residual Stresses
It is generally not possible to measure residual stresses in a product during its manufacture; instead, they are measured
after the manufacturing process is complete. Smith et al. (Ref 30) have divided the residual stress measurement methods
into two broad categories: mechanical and physical. The mechanical category includes the stress-relaxation methods of
layer removal, cutting, hole drilling, and trepanning, whereas the physical category includes x-ray diffraction (XRD),
neutron diffraction, acoustic, and magnetic. The layer-removal technique as originally proposed by Mesnager and Sachs
(Ref 4) is only applicable to simple geometries such as a cylinder with no stress variation along its axis or circumference,
or to a plate with no variation along its length or width. Thus, whereas it could be used to measure quench-induced
residual stresses in a cylinder or a plate, it is not suitable for measuring complex stress patterns such as those caused by
welding. The layer removal and cutting techniques, however, have been applied to pipe welds in combination with
conventional strain gages and XRD measurements. The layer-removal technique is also used to measure residual stresses
in coatings.
Hole-drilling and trepanning techniques can be used in situations where the stress variation is nonuniform, but they are
generally restricted to stress levels of less than one-third of the material yield strength. Also, these two techniques can be
unreliable in areas of steep stress gradients and require extreme care while drilling a hole or ring in terms of its alignment
as well as the heat and stress generation during drilling (Ref 31). For such reasons, and others, these two techniques have

found little application in the measurement of weld-induced residual stresses.
Of the methods in the physical category, XRD is probably the most widely used method, the neutron diffraction method
being relatively new. These two methods measure changes in the dimensions of the lattice of the crystals, and from these
measurements the components of strains and stress are computed. The XRD technique has undergone many
improvements in recent years. With the development of small portable x-ray diffractometers, the technique can be used
for on-site measurement of residual stresses. It should be noted, however, that this technique is capable of measuring
strains in only a shallow layer (approximately 0.0127 mm, or 0.0005 in., thick) at the specimen surface. To measure
subsurface residual stresses in a workpiece, thin layers of materials are successively removed and XRD measurements are
made at each exposed layer. For reasons discussed in the section "Material Removal" in this article, the measurements at
an inner layer should be corrected to account for the material removed in all the previous layers. Reference 32 gives
analytical expressions for such corrections in cases of simple geometries and stress distributions. For more complex cases,
it still remains difficult to determine subsurface residual stresses accurately.
In contrast to the x-rays, neutrons can penetrate deeper into the metals. For example, in iron the relative depth of
penetration at the 50% absorption thickness is about 2000 times greater for neutrons than for x-rays. Only a few materials,
such as cadmium and boron, absorb neutrons strongly. However, to gain the advantage of greater penetration of neutrons
requires the component to be transported to a high flux neutron source (Ref 30), which limits the use of the technique.

References cited in this section
4. R.G. Treuting, J.J. Lynch, H.B. Wishart, and D.G. Richards, Residual Stress Measurements,
American
Society of Metals, 1952
30.

D.J. Smith, G.A. Webster, and P.J. Webster,
Measurement of Residual Stress and the Effects of Prior
Deformation Using the Neutron Diffraction Technique, The Welding Institute, Cambridge, UK, 1987
31.

C.O. Ruud, A Review of Nondestructive Methods for Residual Stress Measurement, J. Met.,
Vol 33 (No.

7), 1981, p 35-40
32.

M.E. Hilley, J.A. Larson, C.F. Jatczak, and R.E. Ricklefs, Ed., "Residual Stress Measurement by X-
Ray
Diffraction," SAE J784a, Society of Automotive Engineers, 1971
Control of Residual Stresses
U. Chandra, Concurrent Technologies Corporation

Residual Stresses Caused by Various Manufacturing Processes
Casting. In the past, little attention has been paid to the control of residual stresses in casting; much of the interest was
focused on the prediction and control of porosity, misruns, and segregation. A review of the transactions of the American
Foundrymen's Society or of earlier textbooks on casting (e.g., Ref 33) reveals practically no information on the subject;
even the ASM Handbook on casting (Ref 34) provides little insight. In a recent book, Campbell (Ref 35) has included a
brief discussion of residual stresses summarizing the work done by Dodd (Ref 36) with simple sand-mold castings. Dodd
studied the effect of two process parameters: mold strength, by changing water content of sand or by ramming to different
levels, and casting temperature. The conclusions of these costly experiments could have been more economically and
easily arrived at by using the basic concepts discussed in the section "Fundamental Sources of Residual Stresses" and
further amplified in the following paragraphs.
When a casting is still in its mold, the stresses are caused by a combination of the mechanical constraints imposed by the
mold, thermal gradients, and solid-state phase transformation. Also, creep at elevated temperature affects these stresses.
Finally, when the casting is taken out of its mold, it experiences springback that modifies the residual stresses.
As discussed in Ref 13, 14, 15, and 16, the computer prediction of residual stresses in castings requires a software that is
capable of performing coupled transient nonlinear thermomechanical analysis (see the section "Computer Prediction of
Residual Stresses" in this article). In addition, it should be able to account for the following:
• Release of latent heat during liquid-to-solid transformation, that is, in the mushy region
• Mechanical behavior of the cast metal in the mushy region
• Transfer of heat and forces at the mold-metal interface
• Creep at elevated temperatures under condition of varying stress
• Enclosure radiation at the mold surface to model the investment-casting process

• Mold withdrawal to model directional solidification
• Mold (material) removal
The author and his coworkers have recently modified a commercial finite element code and have analyzed simple sand-
mold castings (Ref 15, 16). Computer simulation of these castings indicates that: (1) for an accurate prediction of
transient and residual stresses, consideration of creep is important; creep is also found to make the stress distribution more
uniform; and (2) just prior to mold removal the stresses in the casting can be extremely high, but after the mold removal
they become very small (owing to the springback discussed in the section "Material Removal" ) except in the areas of
stress concentration. The residual stresses after the mold removal will not necessarily be small if the casting is complex
and the mold removal is asymmetric with respect to the stress distribution. Also, small variations in mold rigidity are not
found to have any noticeable effect on residual stresses, which confirms the observations based on trial and error using
green-sand molds with various water contents (Ref 35).
Although very little work is published thus far on the subject of control of residual stresses in castings, finite element
simulation methodology is now sufficiently advanced to enable the study of the effect of various process and design
parameters on the residual stresses in castings, for example, superheat, stiffness and design of the mold, design of the
feeding system and risers, and the design of the part itself. Also, residual stresses caused by different casting practices
such as sand-mold, permanent-mold, investment casting, and so forth, can be determined. As the manufacturers and end-
users of cast products become more aware of the status and benefits of the computer-simulation methodology, it can be
expected to play a very important role in controlling residual stresses in complex industrial castings. At present, the
biggest limiting factor in the use of simulation is the lack of thermophysical and mechanical properties data for the cast
metal and the mold materials.
Forging. As with the casting process, little attention has been paid in the past to the control of residual stresses caused by
forging; most of the interest was in predicting the filling and the direction of material flow. Now, due to recent advances
in computer-simulation techniques, it is possible to predict and control the residual stresses in forged parts.
Large plastic flow of the workpiece material is inherent in the forging process. The material flow is influenced by a
number of factors including the die shape and material, forging temperature, die speed, and lubrication at the
die/workpiece interface. Therefore, finite element simulation software used to predict and control residual stresses in the
part should be capable of accounting for these factors. Because a significant amount of energy is dissipated during forging
in the form of heat due to plastic deformation, a coupled thermomechanical analysis becomes necessary especially for
nonisothermal forging. Other factors contributing to the complexity of the finite element simulation of this class of
problems are: temperature-dependent thermal and mechanical properties of the materials (especially for a nonisothermal

forging); the choice of solution algorithm and remeshing due to large plastic deformation in the workpiece; and
mathematical treatment of the die/workpiece interface that includes heat transfer, lubrication, and contact. The last two
terms in Eq 2 need not be considered in the simulation of the forging process.
Finite element simulation of the forging process with simple geometries and of a two-dimensional idealization of the
thread-rolling process (Ref 17) showed that, although the stresses in the workpiece are high during the deformation stage,
the stresses after retraction of the die (residual stresses) are no longer high except in the regions of stress concentration.
Again, similar to the simulation of the casting processes, it is premature to generalize this conclusion but it is clear that
the technique of computer simulation of forging and many other bulk deformation processes has advanced to a stage
where it can assist in controlling the residual stresses in the part by performing a detailed parametric study with much less
investment of time and capital than trial and error on the shop floor.
Quenching involves heating of the workpiece to the heat treatment temperature followed by rapid cooling in a
quenchant (e.g., air, water, oil, or salt bath) in order to impart the desired metallurgical and mechanical properties. The
choice of a quench medium is the key element; it should be such that it removes the heat fast enough to produce the
desired microstructure, but not too fast to cause transient and residual stresses of excessive magnitude or of an adverse
nature (e.g., tensile instead of compressive). The heat removal characteristic of a quenchant is known to be affected by a
number of factors including the size, shape, orientation of the workpiece (even for simple shapes such as plates and
cylinders, the heat removal is different at the bottom, top, and side surfaces); the use of trays and fixtures to hold the
workpiece in the quenchant; composition of the quenchant; size of the pool and its stirring, and so forth (Ref 37, 38, 39).
Additional difficulties arise when, due to economic reasons, quenching is performed in a batch process.
In the past, using trial and error, shop-floor personnel have come up with some interesting strategies to control the
residual stresses (and warpage), for example, air delay or an intentional delay while transporting the workpiece from the
heating furnace to the quenchant, and time quenching or performing the quenching operation in two steps. In the first
step, the part is quenched in a medium such as a salt bath until the part has cooled below the nose of time-temperature
transformation curve, followed by quenching in second medium such as air to slow the cooling rate. Obviously,
perfecting the quenching operation by trial and error can be an extremely time-consuming task.
At first glance, computer simulation of the quenching process may appear to be simple. It involves an uncoupled transient
nonlinear small deformation thermomechanical analysis (as outlined in the section "Computer Prediction of Residual
Stresses" in this article), with due consideration to solid-state transformation effects (Ref 9, 24, 25); creep is generally
ignored. However, the major difficulty lies (for reasons discussed in the preceding paragraph) in a lack of knowledge of
the heat removal characteristic of various quenchants, which is mathematically represented as the convective heat transfer

coefficient at the outer boundary of the workpiece. Other difficulties arise due to the lack of thermophysical and
mechanical properties of the workpiece material at elevated temperatures. Still, at least in the United States, major aircraft
engine manufacturers and their forging vendors have been using computer simulation to control quench-related cracking
and residual stresses for some time. One such example involving a turbine disk is discussed in Ref 40. The reported work
was performed without the benefit of sophisticated simulation software that could account for solid-state transformation
effects. For proprietary reasons, few such cases are published in the open literature.
Machining. Many complex parts in aerospace and other key industries are made by machining forgings, castings, bars,
or plates to their net shapes. The presence of residual stresses in the workpiece affects its machinability and, on the other
hand, the machining process also creates residual stresses and undesired distortions in the part and alters the already
existing stress state. In order to minimize or eliminate these adverse effects, machine-shop personnel often experiment
with a number of process parameters, for example, depth of cut, speed of the cutting tool, and coolant. For single-point
turning, they frequently flip the workpiece in order to balance the distortions and stresses evenly on the two sides. This
trial and error is frequently combined with statistical process control.
A serious problem associated with machining and residual stresses is often manifested in the form of part distortion. For
example, consider the example in Table 1 (Ref 12). The table shows the results of a dimensional check on 30 samples of
an aircraft engine part that was made by machining heat treated forgings procured from three different vendors (10
samples each). The location at which the dimensional check was performed is identified on the figure included in the
table. It was found that: (1) for all forgings from any one vendor, the drop was almost identical; (2) the drop in forgings
from vendor B was within the specifications, but not so in the case of the other two vendors; and (3) the drop in forgings
from vendors A and C was on the two opposite sides of that from vendor B. It was recognized that all heat treated
forgings contained residual stresses. When these forgings were machined to net shapes, distortions occurred for two
reasons: the release of residual stresses from the removed portion of the workpiece and the machining process itself. The
former can be easily modeled using the finite element method if the magnitude of residual stresses in the forgings prior to
machining is known (Ref 5). However, to this author's knowledge, no serious attempt has so far been made to predict
residual stresses in a workpiece due to the machining process itself. If an attempt of this type is to provide reliable results,
it must take into account such factors as: depth of cut, speed of the cutting tool, interaction between the tool and the
workpiece (heat and force), coolant, and clamping/unclamping of the workpiece. It must also recognize the fact that the
location of contact between the tool and the workpiece moves as the machining process progresses. If such technology
could be developed, it would become possible to predict and control the overall distortions and residual stresses in a part
after machining, thereby reducing scrap.

Table 1 Dimensional check of a machined part

Drop No.

mm in.
Vendor A
1 2.692

0.1060

2 2.769

0.1090

3 2.725

0.1073

4 2.756

0.1085

5 2.738

0.1078

6 2.667

0.1050

7 2.769

0.1090

8 2.743

0.1080

9 2.680

0.1055

10 2.743

0.1080

Vendor B
1 2.921

0.1150

2 2.906

0.1144

3 2.997

0.1180

4 2.941

0.1158

5 2.936

0.1156

6 2.870

0.1130

7 2.972

0.1170

8 2.954

0.1163

9 2.926

0.1152

10 2.954

0.1163

Vendor C
1 3.294

0.1297

2 3.124

0.1230

3 3.200

0.1260

4 3.195

0.1258

5 3.213

0.1265

6 3.251

0.1280

7 3.294

0.1297

8 3.127

0.1231

9 3.251

0.1280

10 3.277

0.1290

Source: Ref 12
Welding. The residual stress distribution in welded joints depends on a number of process and design parameters such as
the heat input, speed of the welding arc, preheat, thickness of the welded part, groove geometry, and weld schedule.
Welding engineers have long used trial and error to obtain a suitable combination of these parameters in order to control
the residual stresses.
The role of computer simulation in the prediction of residual stresses in weldments is the subject of a recent review (Ref
14). Major elements of computer simulation of the process are:
• Mathematical representation of the heat input from the welding source
• A transient thermal analysis
• A tran
sient stress/displacement analysis; the flow of molten metal and thermal convection in the weld
pool are generally ignored
Following Rosenthal (Ref 41), a semisteady state approach is often used, although some attempts at full three-
dimensional analysis have also been made. As mentioned earlier, it is now possible to account for volumetric change and
transformation plasticity effects. Because of the short time periods involved, creep is ignored. In the case of a single-pass
weld or a weld with few passes (e.g., four or five), it is now possible to predict residual stresses with reasonable accuracy.
But, as the number of passes increases (e.g., 20 or 30), it becomes computationally intractable to model each pass. The
scheme of lumping several passes into one layer has been employed with less than satisfactory results. In addition to
excessive computation time, other major difficulties with the simulation of a multipass weld are: the numerical errors tend
to accumulate with each pass, and the changes in metallurgical and mechanical properties of material in previously
deposited layers during deposition of a subsequent layer are difficult to quantify and to account for in the finite element
analysis. The technique of lumping several layers together aggravates these problems.

Between the mid 1970s and early 1980s, the Electric Power Research Institute (EPRI) in the United States sponsored a
program to systematically study the effects of various process and geometric parameters such as the heat input, welding
method (gas tungsten arc welding, submerged arc welding, laser, and plasma), speed of the welding arc, diameter and
thickness of the pipe, and groove geometry, on residual stresses in pipe welds (Ref 42, 43, 44, 45, 46, 47, 48, 49, 50, and
51). In addition, various thermal processes such as heat-sink welding, backlay welding, and induction heat treatment were
investigated to verify if the residual stresses on the inner surface of the pipe could be changed from tensile to compressive
to avoid intergranular stress-corrosion cracking. Both experimental and finite element methods were used in the study.
The results of this effort are summarized in Ref 52.
A very interesting effort related to in-process control and reduction of residual stresses and distortions in weldments is
being pursued (Ref 53, 54). The effort aims at moving beyond mere analysis of residual stresses and distortions to
aggressively controlling and reducing them. To accomplish this objective, the effort is subdivided into the development of
the following three primary capabilities: prediction, sensing, and control. For prediction purposes, a series of computer
programs have been developed, include simple but fast one-dimensional programs that analyze only the most important
stress component, that is, the one parallel to the weld line. Sensing capability refers to a set of devices including a laser
interferometer to measure minute amounts of distortions, a laser vision system to measure large amounts of distortions,
and a mechanical system to measure radi of curvature. Finally, to control the residual stresses, various techniques
including changes in heating pattern and application of additional forces have been attempted. References 53 and 54
provide further examples of the application of this methodology in reduction of residual stresses in weldments in high-
strength steels and girth-welded pipes.
Coating. Coatings are being used extensively in aerospace, marine, automobile, biomedical, electronics, and other
industries. For example, in modern jet aircraft engines, approximately 75% of all components are coated. Some of the
reasons for the application of coatings are: thermal barrier, wear resistance, corrosion resistance, oxidation protection,
electrical resistance, and repair or dimensional restoration of worn parts. A variety of methods are used for the deposition
of coatings on a substrate; the following discussion is limited primarily to the thermal spray process.
The prediction of residual stresses in a coating/substrate system is in its infancy. These stresses result from the difference
in the coefficient of thermal expansion of the coating and substrate materials and from plastic deformation of the substrate
material. The limited number of numerical studies conducted thus far have been related to small button-type specimens
where the coating material was assumed to be fully molten and deposited instantaneously. These efforts have ignored
several important factors, for example:
• The presence of partially molten particles in the spray

• A nonuniform deposition of coating material normal to the axis of the plasma jet
• A liquid-to-solid and solid-state transformation
• Imperfect bond between the coating and the substrate
• The relative motion between the spray and the substrate
Also, because a layer of coating consists of several successive passes, the effect of any new pass on its adjacent
previously deposited pass in terms of partial remelting, additional material buildup, solute diffusion, and redistribution of
residual stresses could be important and should be accounted for. If more than one layer is involved, for example, in
functionally graded coatings, modeling the effect of a whole new layer of material on the previously deposited layer
would be computationally prohibitive. Also, due to the morphology of the coating material on deposition, its thermal and
mechanical properties are extremely difficult to measure and, thus, are generally unavailable for simulation purposes. Due
to such reasons, end-users of coated products still rely on the methods of trial and error and statistical process control for
the selection of an optimal combination of process parameters in order to control the residual stresses in coated parts (Ref
55).

References cited in this section
5.
U. Chandra, Validation of Finite Element Codes for Prediction of Machining Distortions in Forgings,
Commun. Numer. Meth. Eng., Vol 9, 1993, p 463-473
9. J B. Leblond, G. Mottet, J. Devaux, and J C. Devaux, "Mathematical Models of Anisothermal Pha
se
Transformations in Steels, and Predicted Plastic Behavior," Mater. Sci. Technol., Vol 1, 1985, p 815-822
12.

U. Chandra, S. Rachakonda, and S. Chandrasekharan, Total Quality Management of Forged Products
through Finite Element Simulation, Proc. Third International SAMPE Metals and Metals Processing Conf.,

Vol 3, F.H. Froes, W. Wallace, R.A. Cull, and E. Struckholt, Ed., SAMPE International, 1992, p M379-
M393
13.

U. Chandra, Computer Prediction of Hot Tears, Hot Cracks, Residual Stresses and Distortions
in Precision
Castings: Basic Concepts and Approach, Proc. Light Metals, J. Evans, Ed., TMS, 1995, p 1107-1117
14.

U. Chandra, Computer Simulation of Manufacturing Processes Casting and Welding, Comput.
Model.
Simul. Eng., Vol 1, 1996, p 127-174
15.

U. Chandra, R. Thomas, and S. Cheng, Shrinkage, Residual Stresses, and Distortions in Castings,
Comput.
Model. Simul. Eng., Vol 1, 1996, p 369-383
16.

A. Ahmed and U. Chandra, Prediction of Hot Tears, Residual Stresses and Distortions in Castings Including
the Effect of Creep, Comput. Model. Simul. Eng., to be published
17.

U. Chandra, S. Chandrasekharan, and R. Thomas, "Finite Element Analysis of the Thread Rolling Process,"
Concurrent Technologies Corporation, submitted to Knolls Atomic Power Lab, Schenectady, NY, 1995
24.

S. Sjöström, Interactions and Constitutive Models for Calculating Quench Stresses in Steel, Mater.
Sci.
Technol, Vol 1, 1985, p 823-829
25.

S. Das, G. Upadhya, and U. Chandra, Prediction of Macro- and Micro-Residual Stress in Quenching
Using

Phase Transformation Kinetics, Proc. First International Conf. Quenching and Control of Distortion,
G.E.
Totten, Ed., ASM International, 1992, p 229-234
33.

R.W. Heine, C.R. Loper, and P.C. Rosenthal, Principles of Metal Casting, Tata McGraw-Hill, N
ew Delhi,
India, 1976
34.

Casting, Vol 15, ASM Handbook, (formerly Metals Handbook, 9th ed.), ASM International, 1988
35.

J. Campbell, Casting, Butterworth Heinmann, Oxford, U.K., 1991
36.

R.A. Dodd, Ph.D. Thesis, Department of Industrial Metallurgy, University of Birmingham, U.K., 1950
37.

H.E. Boyer and P.R. Cary, Quenching and Control of Distortion, ASM International, 1988, p 11
38.

T.V. Rajan, C.P. Sharma, and A. Sharma, Heat Treatment Principles and Techniques, Prentice-
Hall of
India Private Ltd., 1988
39.

S. Segerberg and J. Bodin, Variation in the Heat Transfer Coefficient around Components of Different
Shapes During Quenching, Proc. First International Conf. Quenching and Control of Distortion,
G.E.

Totten, Ed., ASM International, 1992, p 165-170
40.

R.A. Wallis, N.M. Bhathena, P.R. Bhowal, and E.L. Raymond, Application of Process Modelling to Heat
Treatment of Superalloys, Ind. Heat., 1988, p 30-33
41.

D. Rosenthal, The Theory of Moving Sources of Heat and Its Application to Metal Treatments,
Trans.
ASME, Nov 1946, p 849-866
42.

R.M. Chrenko, "Residual Stress Studies of Austenitic and Ferritic Steels," Conference on Residual Stresses
in Welded Construction and Their Effects, London, Nov 1977
43.

R.M. Chrenko, "Weld Residual Stress Measureme
nts on Austenitic Stainless Steel Pipes," Lake George
Conference, General Electric, 1978, p 195-205
44.

W.A. Ellingson and W.J. Shack, Residual Stress Measurements on Multi-
Pass Weldments of Stainless Steel
Piping, Exp. Mech., Vol 19 (No. 9), 1979, p 317-323
45.

W.J. Shack, W.A. Ellingson, and L.E. Pahis, "Measurement of Residual Stresses in Type-
303 Stainless
Steel Piping Butt Weldments," Report NP-1413, Electric Power Research Institute, June 1980
46.

E.F. Rybicki and P.A. McGuire, "A Computational Mod
el for Improving Weld Residual Stresses in Small
Diameter Pipes by Induction Heating," 80-C2-PVP-
152, Century 2 Pressure Vessels and Piping Conference,
San Francisco, CA, Aug 1980, American Society of Mechanical Engineers
47.

A.F. Bush and F.J. Kromer, Re
sidual Stresses in a Shaft after Weld Repair and Subsequent Stress Relief,
Exp. Tech., Vol 5 (No. 2), 1981, p 6-12
48.

F.W. Brust and R.W. Stonesifer, "Effect of Weld Parameters on Residual Stresses in BWR Piping
Systems," Report NP-1743, Electric Power Research Institute, March 1981
49.

F.W. Brust and E.F. Rybicki, A Computational Model of Backlay Welding for Controlling Residual
Stresses in Welded Pipes, J. Pressure Vessel Technol. (Trans. ASME), Vol 103, 1981, p 226-232
50.

R.M. Chrenko, Thermal Modification of Welding Residual Stresses, Residual Stress and Stress Relaxation,

E. Kula and V. Weiss, Ed., Plenum Publishing, 1982, p 61-70
51.

E.F. Rybicki, P.A. McGuire, E. Merrick, and J. Wert, The Effect of Pipe Thickness on Residual Stresses
due to Girth Welds, J. Pressure Vessel Technol. (Trans. ASME), Vol 104, 1982, p 204-209
52.

U. Chandra, Determination of Residual Stresses due to Girth-Butt Welds in Pipes,
J. Pressure Vessel
Technol. (Trans. ASME), Vol 107, 1985, p 178-184
53.

K. Masubuchi, In-Process Control and Reduction of Residual Stresses and Distortion in Weldments,
Proc.
Practical Applications of Residual Stress Technology, C.O. Ruud, Ed., ASM International, 1991, p 95-101
54.

K. Masubuchi, Research Activities Examine Residual Stresses and Distortion in Welded Structures,
Weld.
J., Dec 1991, p 41-47
55.

R.V. Hillery, Coatings Producibility, The Leading Edge,
GE Aircraft Engines, Cincinnati, Ohio, Fall 1989,
p 4-9
Control of Residual Stresses
U. Chandra, Concurrent Technologies Corporation

Stress-Relief Methods
The basic premise of a stress-relief method is to produce rearrangement of atoms or molecules from their momentary
equilibrium position (higher residual stress state) to more stable positions associated with lower potential energy or stress
state. These methods can be classified into three broad categories: thermal, mechanical, and chemical (Ref 4, p 134). The
following concern the methods in the first two categories.
Thermal stress-relief methods include annealing, aging, reheat treatment (e.g., postweld heat treatment), and others. In
general, a stress-relief operation involves heating the part to a certain temperature, holding at the elevated temperature for
a specified length of time, followed by cooling to room temperature. Primary reduction in residual stresses takes place
during the holding period due to creep and relaxation. Thus, computer simulation of a thermal stress-relief method

generally entails a thermal-elastic-plastic-creep analysis of the part. A simple, one-dimensional computer analysis of
residual stresses in thin plates along with experimental verification is discussed by Agapakis and Masubuchi (Ref 56).
More sophisticated thermal-elastic-plastic-creep simulations of the annealing of single pass and multipass girth-butt welds
in pipes are presented in Ref 57 and 58.
A number of subcategories of mechanical stress-relief methods are listed in Ref 4. Of these, the methods in the static-
stressing subcategory such as stretching, upsetting, bending and straightening, and autofrettage are common, and these
should not pose much difficulty in simulation by the finite element method. Similarly, in the mechanical surface treatment
subcategory, it should be possible to model the surface-rolling method. However, within the same subcategory, shot
peening (a frequently used stress-relief method) is likely to be difficult to simulate; and to this author's knowledge, no
realistic attempt has yet been made to do so. The obvious reason is that, whereas it should be possible to model a single
impact, modeling multiple impacts will be difficult, just as it is for modeling multipass welding.
In recent years, the method of vibratory stress relief (especially in the subresonant region) has received considerable
attention (Ref 59, 60). The basic premise of this method is that the presence of residual stresses in a part changes
(increases) its natural resonant frequency. When the part is subjected to vibrations below its new frequency, the metal
absorbs energy. During this process, the stresses redistribute gradually and the resonant frequency shifts back to the point
corresponding to a residual stress-free (or almost free) state. The process does not change the metallurgical or mechanical
properties of the material. The technique has been found successful in relieving residual stresses induced by thermal
processes such as welding and casting, but not those induced by cold working. It has also been applied to reduce residual
stresses in parts prior to machining in order to minimize distortions. It has been found particularly beneficial in low- and
medium-carbon steels, stainless steels, and aluminum alloys, but not in copper alloys. In view of the fact that the
technique is much simpler, quicker, and more inexpensive than the thermal-relief methods, it merits further study.

References cited in this section
4. R.G. Treuting, J.J. Lynch, H.B. Wishart, and D.G. Richards, Residual Stress Measurements,
American
Society of Metals, 1952
56.

J.E. Agapakis and K. Masubuchi, Analytical Modeling of Thermal Stress Relieving in Stainless and High
Strength Steel Weldments, Weld. J. Res. Suppl., 1984, p 187s-196s

57.

B.L. Josefson, Residual Stresses and Their Redistribution During Annealing of a Girth-Butt Welded Thin-
Walled Pipe, J. Pressure Vessel Technol. (Trans. ASME), Vol 104, 1982, p 245-250
58.

B.L. Josefson, Stress Redistribution During Annealing of a Multi-Pass Butt-Welded Pipe,
J. Pressure
Vessel Technol. (Trans. ASME), Vol 105, 1983, p 165-170
59.

R.A. Claxton, Vibratory Stress Relieving
Its Advantages and Limitations as an Alternative to Thermal
Treatments, Heat. Treat. Met., 1974, p 131-137
60.

A.G. Hebel, Jr., Subresonant Vibrations Relieve Residual Stress, Met. Prog., Nov 1985, p 51-55
Control of Residual Stresses
U. Chandra, Concurrent Technologies Corporation

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Institute of Metals, 1948, p 47-59
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Materials, 1949

4. R.G. Treuting, J.J. Lynch, H.B. Wishart, and D.G. Richards, Residual Stress Measurements,
American
Society of Metals, 1952
5. U. Chandra, Validation of Fin
ite Element Codes for Prediction of Machining Distortions in Forgings,
Commun. Numer. Meth. Eng., Vol 9, 1993, p 463-473
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W.K.C. Jones and P.J. Alberry, "The Role of Phase Transformation in the Development of Residual
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C. Devaux, "Mathematical Models of Anisothermal Phase
Transformations in Steels, and Predicted Plastic Behavior," Mater. Sci. Technol., Vol 1, 1985, p 815-822
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C.L. Magee, "Transformation Kinetics, Microplasticity and Aging of Martensite in FE31 Ni," Ph.D. Thesis,
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U
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through Finite Element Simulation, Proc. Third International SAMPE Metals and Metals Processing Conf.,

Vol 3, F.H. Froes, W. Wallace, R.A. Cull, and E. Struckholt, Ed., SAMPE International, 1992, p M379-
M393
13.

U. Chandra, Computer Prediction of Hot Tears, Hot Cracks, Residual Stresses and Distortions in Precision
Castings: Basic Concepts and Approach, Proc. Light Metals, J. Evans, Ed., TMS, 1995, p 1107-1117
14.

U. Chandra, Computer Simulation of Manufacturing Processes Casting and Welding, Comput.
Model.
Simul. Eng., Vol 1, 1996, p 127-174
15.

U. Chandra, R. Thomas, and S. Cheng, Shrinkage, Residual Stresses, and Distortions in Castings,
Comput.
Model. Simul. Eng., Vol 1, 1996, p 369-383
16.

A. Ahmed and U. Chandra, Prediction of Hot Tears, Residual Stresses and Distortions in Castings Including
the Effect of Creep, Comput. Model. Simul. Eng., to be published
17.

U. Chandra, S. Chandrasekharan, and R. Thomas
, "Finite Element Analysis of the Thread Rolling Process,"
Concurrent Technologies Corporation, submitted to Knolls Atomic Power Lab, Schenectady, NY, 1995
18.

G. Comini, S. Del Guidice, R.W. Lewis, and O.C. Zienkiewicz, Finite Element Solution of Non-lin
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S. Sjöström, Interactions and Constitutive Models for Calculating Quench Stresses in Steel, Mater.
Sci.
Technol, Vol 1, 1985, p 823-829
25.

S. Das, G. Upadhya, and U. Chandra, Prediction of Macro- and Micro-Residual Str
ess in Quenching Using
Phase Transformation Kinetics, Proc. First International Conf. Quenching and Control of Distortion,
G.E.
Totten, Ed., ASM International, 1992, p 229-234
26.

S. Das, G. Upadhya, U. Chandra, M.J. Kleinosky, and M.L. Tims, Finite Element Modeling of a Single-
Pass GMA Weldment, Proc. Engineering Foundation Conference on
Modeling of Casting, Welding and
Advanced Solidification Processes VI, T.S. Piwonka, V. Voller, and L. Katgerman, Ed., TMS, 1993, p 593-
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C.O. Ruud, A Review of Nondestructive Methods for Residual Stress Measurement, J. Met.,
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M.E. Hilley, J.A. Larson, C.F. Jatczak, and R.E. Ricklefs, Ed., "Residual Stress Measurement by X-
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R.W. Heine, C.R. Loper, and P.C. Rosenthal, Principles of Metal Casting, Tata McGraw-
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Casting, Vol 15, ASM Handbook, (formerly Metals Handbook, 9th ed.), ASM International, 1988
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J. Campbell, Casting, Butterworth Heinmann, Oxford, U.K., 1991
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R.A. Dodd, Ph.D. Thesis, Department of Industrial Metallurgy, University of Birmingham, U.K., 1950
37.

H.E. Boyer and P.R. Cary, Quenching and Control of Distortion, ASM International, 1988, p 11
38.

T.V. Rajan, C.P. Sharma, and A. Sharma, Heat Treatment Principles and Techniques, Prentice-
Hall of
India Private Ltd., 1988
39.

S. Segerberg and J. Bodin, Variation in the Heat Transfer Coefficient around Components of Different
Shapes During Quenching, Proc. First International Conf. Quenching and Control of Distortion,
G.E.
Totten, Ed., ASM International, 1992, p 165-170
40.

R.A. Wallis, N.M. Bhathena, P.R. Bhowal, and E.L. Raymond, Application of Process Modelling to Heat
Treatment of Superalloys, Ind. Heat., 1988, p 30-33
41.

D. Rosenthal, The Theory of Moving Sources of Heat and Its Application to Metal Treatments,
Trans.
ASME, Nov 1946, p 849-866
42.

R.M. Chrenko, "Residual Stress Studies of Austenitic and Ferritic Steels," Conference on Residual Stresses
in Welded Construction and Their Effects, London, Nov 1977
43.

R.M. Chrenko, "Weld Residual S
tress Measurements on Austenitic Stainless Steel Pipes," Lake George
Conference, General Electric, 1978, p 195-205
44.

W.A. Ellingson and W.J. Shack, Residual Stress Measurements on Multi-
Pass Weldments of Stainless Steel
Piping, Exp. Mech., Vol 19 (No. 9), 1979, p 317-323
45.

W.J. Shack, W.A. Ellingson, and L.E. Pahis, "Measurement of Residual Stresses in Type-
303 Stainless
Steel Piping Butt Weldments," Report NP-1413, Electric Power Research Institute, June 1980
46.

E.F. Rybicki and P.A. McGuire, "A Co
mputational Model for Improving Weld Residual Stresses in Small
Diameter Pipes by Induction Heating," 80-C2-PVP-
152, Century 2 Pressure Vessels and Piping Conference,
San Francisco, CA, Aug 1980, American Society of Mechanical Engineers
47.

A.F. Bush and
F.J. Kromer, Residual Stresses in a Shaft after Weld Repair and Subsequent Stress Relief,
Exp. Tech., Vol 5 (No. 2), 1981, p 6-12
48.

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Stresses in Welded Pipes, J. Pressure Vessel Technol. (Trans. ASME), Vol 103, 1981, p 226-232
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E. Kula and V. Weiss, Ed., Plenum Publishing, 1982, p 61-70
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Technol. (Trans. ASME), Vol 107, 1985, p 178-184
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Proc.
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Design for Surface Finishing
Eric W. Brooman, Concurrent Technologies Corporation

Introduction
AN AXIOM among surface-finishing industry professionals is that the quality of a finish is only as good as the quality of
the substrate and its pretreatment. That is, to obtain a finish of high quality, meeting all performance specifications and
customer expectations, great care has to be taken in identifying and using the appropriate pretreatment. The latter could be
grinding, heating, buffing, cleaning, or a number of other processes to prepare the surface properly. Although not
articulated in quite the same way, the surface-finishing industry also recognizes that the design of the part (component or
assembly) can have a significant influence on the ability to use satisfactory pretreatments and obtain quality finishes. The
overall part design, and the design of surface features and their size, can have an impact not only on the choice of
pretreatments and finishes, but also on the efficacy of these processes and the results obtained. This article provides some

guidelines about general design principles for different types of surface-finishing processes, which include cleaning,
organic coatings, and inorganic coatings applied by a variety of techniques.
Many of the guidelines discussed here apply equally as well to other articles in this section and vice versa. Therefore,
although what is presented here is a fairly comprehensive summary of the topic of design for surface finishing, useful
information can be found in other articles such as "Design for Machining" and "Design for Heat Treatment." As is
stressed elsewhere and discussed in this article, design must be considered an integral part of the overall manufacturing
process and cannot be considered in isolation.
Definitions and more detailed descriptions of the processes discussed in this article can be found in Surface Engineering,
Volume 5 of ASM Handbook.
Design for Surface Finishing
Eric W. Brooman, Concurrent Technologies Corporation

Design as an Integral Part of Manufacturing
In recent years, manufacturing processes have been evaluated in terms of life-cycle costs and their impact on the
environment, and even ecology in general. Several scenarios have been proposed for the life cycle of materials, which of
necessity incorporates manufacturing processes. In the article "Introduction to Manufacturing and Design" in this
Volume, an example of one such scheme is given where a material flows through the extraction, refining (preparation),
shapemaking and structural treatments (manufacturing), and surfacing (surface-finishing) stages before being assembled
and placed in use. Manufacturing also can be viewed as part of an "integrated product- and process-development"
process, also described in that article. This approach places emphasis on "product design" and "process design" in a
concurrent-engineering environment and forces product developers to consider both simultaneously, rather than
sequentially and separately (Ref 1). If these concepts are extrapolated, with a focus on design in relation to surface
finishing, the iterative process shown in Fig. 1 results. The scheme presented in this figure is entirely consistent with
another modern manufacturing concept, namely that of continuous improvement.

Fig. 1 Flow diagram for incorporating design principles into surface-finishing operations

Having stressed the importance of considering design precepts as an integral part of the manufacturing process and
product improvement, the following guidelines should be considered as just that guidelines. Each application should be
treated on an individual basis, and surface-finishing processes should be flexible enough to permit design improvements,

and vice versa. In addition, it should be obvious from Fig. 1 that the earlier in the product-development cycle
manufacturing-related design considerations are addressed, the more efficient will be the manufacturing process and the
better the quality of the surface finish will be. Communication with and input from the manufacturing and surface-
finishing staff are very important in establishing a satisfactory product design.
Design of the part or component and pretreatment selections are important, but in keeping with the concept of design
being an integral part of manufacturing and surface finishing, process equipment design restrictions and fixturing design
also are very important. Both can influence the quality of the resulting finish. Figure 2 shows their interrelated roles
schematically. In this overview, emphasis is placed on issues relating to part design and process equipment. Fixturing
must be designed and tailored for each individual application and is beyond the scope of this article.

Fig. 2 Interrelation between part design, equipment limitations, and fixturing

Reference cited in this section
1.

H.A. Kuhn, Concurrent Technologies Corp., Johnstown, PA, personal communication, 1996

Design for Surface Finishing
Eric W. Brooman, Concurrent Technologies Corporation

General Design Principles Related to Surface Finishing
There are a number of general design principles that apply to a variety of finishing processes, while others are specific to
individual finishing techniques. These general principles are discussed in this section, and the following three sections
discuss design aspects relating to: (1) surface-preparation techniques, including cleaning; (2) organic finishing techniques;
and (3) inorganic finishing techniques. References 2 and 3 provide some background material on the various finishing
techniques discussed, while Tables 1, 2, and 3 summarize the important design limitations for these three categories. The
subject of materials selection is covered elsewhere in this Volume. The choice of materials can limit the choices for
surface finishing. Where appropriate, significant limitations are described.
Table 1 Summary of design limitations for selected surface-preparation processes
Process Design limitations

Avoid recesses, holes, channels, and similar features (such as closely spaced ribs) that could trap blasting
media
Avoid thin cross sections (such as fins, louvers, walls) that could be distorted by the blasting media
Blasting/deburring
Avoid intricate designs and surface features
Broaching/honing
Typically used for inside diameters of tubes and other cylindrical parts, or for grooves, large holes, and
other cavities
Surfaces must be accessible to tools and withstand the local pressure and heat buildup
Avoid very thin cross sections/wall thickness
Surfaces must be accessible to tools and withstand the local pressure and heat buildup
Avoid very thin cross sections/wall thickness that could deflect
Avoid sharp corners and edges
Brushing/burnishing
Avoid intricate designs and surface features
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air or evolved gases and prevent chemical action from occurring or cause
uneven attack
Chemical milling
Mask areas not to be attacked
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent surface chemical reactions from occurring or cause staining
Conversion coating
Mask areas not to be attacked
Allow for electrical contact to be made on nonsignificant surfaces
Avoid features that would trap process chemicals or prevent satisfactory rinsing

Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Electrocleaning
Avoid features that could trap air or evolved gases and would prevent cleaning from occurring or cause
staining
Allow for electrical contact to be made on nonsignificant surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air or evolved gases and prevent polishing action from occurring or cause
staining
Electropolishing
Mask areas not to be attacked
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent etching action from occurring
Avoid sharp corners and edges
Avoid shallow intricate designs and surface features
Etching
Mask areas not to be attacked
Surfaces must be accessible to tools and withstand the local pressure and heat buildup
Avoid very thin cross sections/wall thickness
Avoid sharp corners, edges, and protuberances
Grinding
Avoid intricate designs and surface features
Surfaces must be accessible to tools (preferably flat or simple, curved contours
Avoid very thin cross sections/wall thickness that cannot withstand the local pressure and heat buildup
Avoid sharp corners and edges
Lapping/buffing
Avoid intricate designs and surface features that would trap the lapping/buffing compounds

Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent pickling action
Pickling
Avoid flat surfaces on small parts that could stick together, exclude the acid, and prevent the pickling
action
Surfaces must be accessible to tools and withstand the local pressure and heat buildup
Avoid very thin cross sections/wall thickness
Avoid sharp corners, edges, and protuberances
Polishing
Avoid intricate designs and surface features that could trap the polishing compound
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent cleaning from occurring
Solvent cleaning,
immersion
Avoid flat or curved surfaces on small parts that could stick together during immersion and prevent
cleaning of those surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or prevent
satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent cleaning from occurring
Solvent cleaning,
ultrasonic
Avoid thin cross sections that could be damaged by the energy released during cavitation
Avoid features (e.g., small recesses, blind holes, cavities) that would trap smut and process chemicals or
prevent satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover

Avoid features that could trap air and prevent coating removal from occurring
Stripping, chemical
Mask areas not to be attacked
Avoid recesses, holes, channels, and similar features that could trap blasting media
Avoid thin cross sections or intricate designs that could be damaged by the stripping media
Stripping, mechanical
Mask areas not to be attacked
Avoid thin cross sections or intricate designs that could be distorted by the thermal cycling
Stripping, thermal
Try to provide uniform cross-sectional mass throughout the part to help provide a uniform temperature
distribution during heating

Table 2 Summary of design limitations for selected organic finishing processes
Process Design limitations
Allow for electrical contact to be made on nonsignificant surfaces
Avoid features that could trap air and prevent wetting by process solutions
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Electrocoating
Avoid thin cross sections or intricate designs that could become distorted during drying/curing cycle
Allow for electrical contact to be made on nonsignificant surfaces
Avoid features that could trap air and prevent wetting by process solutions
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Electropolymerization
Avoid thin cross sections or intricate designs that could become distorted during drying/curing cycle
Surfaces must be accessible to application tools (preferably flat or simple, curved contours)
Avoid features that would trap excess paint
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent coating from occurring
Painting, brushing or
dipping

Avoid thin cross sections or intricate designs that could become distorted during drying/curing cycle
Surfaces must be accessible to application tools (preferably flat or simple, curved contours)
Allow for fixturing/racking on nonsignificant surfaces
Avoid features that would trap excess paint
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent coating from occurring
Painting, solvent spraying
Avoid thin cross sections or intricate designs that could become distorted during drying/curing cycle
Allow for fixturing/racking on nonsignificant surfaces
Allow for electrical contact to be made on nonsignificant surfaces
Avoid deep recesses and blind holes that cause the "Faraday cage" effect
Powder coating
Avoid thin cross sections or intricate designs that could become distorted during drying/curing cycle
Allow for fixturing/racking on nonsignificant surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals
Avoid thin cross sections or intricate designs that could be distorted by the thermal cycling
Sol-gel coating
Try to provide uniform cross-sectional mass throughout the part to help provide a uniform temperature
distribution during heating cycle
Allow for fixturing/racking on nonsignificant surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize carryover
Avoid features that could trap air and prevent coating from occurring
Solution coating
Avoid thin cross sections or intricate designs that could become distorted during drying/curing cycle

Table 3 Summary of design limitations for selected inorganic finishing processes
Process Design limitations
Allow for electrical contact to be made on nonsignificant surfaces
Avoid, if possible, sharp edges and corners, ridges, blind holes, etc., that would prevent uniform

density distribution
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover
Avoid features that could trap air and prevent electrochemical reactions from occurring
Avoid features that could trap evolved gases and cause staining
Anodizing
Mask areas not to be anodized
Cementation/diffusion
Surfaces must be thoroughly deburred and cleaned before cladding, so design principles for these
processes also apply
Avoid thin cross sections or intricate designs that could become distorted during thermal cycling
Mask areas not to be coated
Only for relatively simple shapes, especially with flat surfaces
Cladding
Surfaces must be thoroughly cleaned before cladding, so design principles for cleaning also apply
Allow for fixturing/racking on nonsignificant surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or
prevent satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover
Avoid features that could trap air and prevent chemical reactions from occurring or cause staining
Electroless plating
Mask areas not to be coated
Allow for electrical contact to be made on nonsignificant surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or
prevent satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover
Avoid features that could trap air and prevent surface chemical reactions from occurring or cause

staining
Electrophoretic plating
Mask areas not to be coated
Allow for electrical contact to be made on nonsignificant surfaces
Avoid, if possible, sharp edges and corners, ridges, blind holes, etc., that would prevent uniform
current density distribution; or use current robbers and/or shields
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or
prevent satisfactory rinsing
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover
Avoid features that could trap air and prevent deposition from occurring
Avoid features that could trap evolved gases and cause staining
Avoid thin cross sections (such as fins, louvers, walls) that could be distorted by internal stress in
the coating
Electroplating (plating,
electrodeposition)
Mask areas not to be coated
Allow for fixturing/racking on nonsignificant surfaces for discrete, small parts
Best for relatively simple shapes (e.g., tubing) and flat surfaces
Allow for excess coating material to drain quickly
Allow for doctor blades or air knives to be used to obtain uniform coating thickness
Hot dipping, galvanizing
Avoid thin cross sections that could become distorted during thermal cycling
Surfaces must be accessible (preferably flat or simple, curved contours)
Allow for fixturing/racking on nonsignificant surfaces
Avoid features that would trap excess paint
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover
Avoid features that could trap air and prevent coating from occurring
Avoid thin cross sections or intricate designs that could become distorted during drying/fusing

cycle
Inorganic painting, slurry
coating
Mask areas not to be coated
Allow for electrical contact to be made on nonsignificant surfaces or use a conductive screen
Avoid features that would shield the surface from the beam (line-of-sight limited) unless multiple
beams are used or part is rotated/translated in beam
Avoid high aspect ratio holes and recesses, grooves, etc., that would not allow the beam to reach
the bottom surfaces
Ion implantation
Mask areas not to be coated
Allow for electrical contact to be made on nonsignificant surfaces or use a conductive screen
Avoid features that would shield the surface from the beam (line-of-sight limited) unless multiple
beams are used or part is rotated/translated in beam
Avoid high aspect ratio holes and recesses, grooves, etc., that would not allow the beam to reach
the bottom surfaces
Ion plating
Mask areas not to be coated
Allow for fixturing/racking on nonsignificant surfaces
Avoid features that would shield the surface from the laser beam (line-of-sight limited) unless
multiple beams are used or part is rotated/translated in beam
Avoid high aspect ratio holes and recesses, grooves, etc., that would not allow the beam to reach
the bottom surfaces
Avoid thin cross sections or intricate designs that could be damaged by local heating during glazing

Laser glazing
Mask areas not to be treated
Mechanical (peen) plating
Allow for fixturing/racking on nonsignificant surfaces on large parts
Avoid features that could trap air and prevent activation by the process chemicals from occurring

Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover of activating solutions
Avoid recesses, holes, channels, and similar features that could trap peening media
Avoid thin cross sections (such as fins, louvers, walls) that could be distorted by the peening action
Avoid sharp edges and corners that could be damaged by the peening media
Avoid intricate designs and small surface features that cannot be reached by the peening media
Provide good natural drainage or use drainage holes on nonsignificant surfaces to minimize
carryover
Mask areas not to be coated
Allow for fixturing/racking on nonsignificant surfaces
Avoid features (e.g., small recesses, blind holes, cavities) that would trap process chemicals or
prevent satisfactory rinsing
Avoid features that could trap air and prevent surface chemical reactions from occurring or cause
staining
Passivation
Mask areas not to be attacked
Allow for fixturing/racking on nonsignificant surfaces
Design should allow for surface roughening to promote adhesion, so blasting design precepts also
apply
Avoid features that would shield the surface from the spray (line-of-sight limited) unless multiple
sprays are used or part is rotated/translated in spray plume
Avoid high aspect ratio holes and recesses, grooves, etc., that would not allow the spray to reach
the bottom surfaces
Avoid thin cross sections (such as fins, louvers, walls) that could be distorted by the local heating
and kinetic energy
Thermal spraying
Mask areas not to be coated
Allow for fixturing/racking on nonsignificant surfaces
Avoid thin cross sections (such as fins, louvers, walls) that could be distorted by heating, if needed
prior to coating deposition

Vacuum processes are line-of-sight limited, so similar design precepts to those for ion plating will
apply
Vapor deposition (CVD, PVD,
RVD)
Mask areas not to be coated

(a)
CVD, chemical vapor deposition; PVD, physical vapor deposition; RVD, reactive vapor deposition

Fabrication Processes. Some methods of fabrication such as the forging, extrusion, molding, and casting of metals
and ceramics can lead to surface defects that must be removed by subsequent surface-finishing techniques, such as
grinding, lapping, and polishing or electropolishing, or hidden by techniques such as applying a leveling copper deposit
before a decorative plated finish. Defects include laps, tears, cracks, pores, shrinkage cavities, gating and venting
residues, ejection marks, and parting lines. Careful design of the casting or molding operation including the dies, gates,
vents, and overflows will minimize finishing problems by ensuring such defects are avoided, occur on nonsignificant
surfaces, or are hidden by specially incorporated design features, such as steps or ridges at parting lines.
When polymeric materials are being cast, molded, extruded, or formed, it is especially important that the design and
tooling lend themselves to producing an acceptable surface finish because any finishing operation that removes surface
layers could expose porosity and other defects and remove aesthetic qualities such as smoothness and luster. Also, the
selection of finishing tools and conditions is much more critical because of the physical properties (e.g., softness,
plasticity) of the polymeric materials and the potential for damage (e.g., heat distortion) caused by heat generated during
the finishing operation. This topic is discussed more fully in Ref 4. As with metals and ceramics, some undesirable
attributes of the fabrication process, such as parting lines, can be hidden by added design features.
Whatever the type of material being cast or molded, dimensional and warpage allowances must be made in the design of
the tooling (i.e., dies) to accommodate shrinkage and distortion during solidification and cooling. Otherwise, parts may be
undersized or require excessive machining to obtain the specified dimensional tolerances.