Simulation einer Mehrlagenschweißung im Großmotorenbau 2.6.2

Transcrição

2.6.2
Simulation einer Mehrlagenschweißung
im Großmotorenbau
Samuel Sönnichsen, Andreas Junk (CADFEM)
Wärtsilä Switzerland Ltd.
Winterthur, Switzerland
Summary
The crankshaft housing of large two stroke diesel engines (with power up to 80’000 kW) is a welded
structure subjected to a cyclic, constant amplitude loading and designed for infinite life on full load. The
double wall box design, which is standard design since more than 30 years, requires single bevel butt
welds with access only from one side. Geometrical stress concentration on the root side in combination
with additional stress raisers due to weld imperfections, such as lack of penetration, lack of fusion or
cracks, result in high requirements to the fabrication process, since corrective measures from the inside
are not possible. In spite of well known recommendations of welding standards, where favourable effects
of post weld heat treated components are assumed, the main components of the structure of Wärtsilä
engines are not stress relieved. Residual stresses play an important role in the fatigue behaviour.
However, long-term experiences indicate that their effect is rather complex and can be positive as well as
negative.
The present investigation is aimed at a better understanding of the mechanical behaviour of these joints,
particularly regarding residual stress and its effect on the fatigue performance.
Fig. 2: column
Fig.1: Two stroke diesel engine
Fig.3: bedplate
The investigation includes the following steps:
- Stress evaluation by strain gauge measurement and FEM
- Evaluation of residual stress distribution by measurement and weld process simulation (SST)
- Fatigue strength analyses
Keywords
Stress calculation (FEM), Fatigue analysis of welded structure, Residual stress, Weld process simulation
(SST)
th
24 CADFEM Users’ Meeting 2006
International Congress on FEM Technology
with 2006 German ANSYS Conference
October 25-27, 2006 Schwabenlandhalle Stuttgart/Fellbach, Germany
1
0.
Introduction
The crankshaft housings of large two-stroke diesel engines are welded structures with plate thickness up
to 50 mm made of hot rolled carbon steel. The engines operate mainly
on full design load (cyclic forces from combustion and inertia of moving
9
parts) which leads up to 10 cycles after 30 years in service. As fatigue
failure in service requires very expensive repairs and off-hire, a high
survival probability is mandatory. Unfortunately different crack series on
the engine structure accentuate the need for extensive investigations to
improve the situation for future engines. Some cracks occur on few
dynamically loaded areas whereas other highly loaded locations are
running without problems. For some reasons Wärtsilä never applied post
weld heat treatment (PWHT). On one hand, PWHT avoids potential high
tensile residual stress, but on the other beneficial compression stresses
are reduced as well. So the over-all effect of PWHT is not clear a-priori,
and it is essential to know the residual stress distribution and understand
their effect on the fatigue behaviour of the present butt welds. The basic
steps of the procedure are shown in Fig. 5. This investigation is aimed at
clear welding instructions for the production (controlled process).
Fig.4: crack in gear column
Strategy:
Repeatedly increasing the
engine power and
performance requires the
designer to review the rules
continuously
Knowledge
present
Design
quality
s = safety
future
s
k
ac
cr
k
ac
cr
s
Production
1/
Q
action
1/
Q
k
ac
cr
quality
1/
Q
Reliable control of the
welding process has a very
strong influence on the real
fatigue strength and the
safety margin of components
quality
Single side butt weld
(without back weld) is
not well described in
any fatigue design
codes
diameter:
inaccuracies about
allowable fatigue limit
design (stress)
Knowledge
Design
Production
design (stress)
design (stress)
action
Development of technical knowledge
for specific welding application
(single bevel butt weld)
Design improvements in respect of
weld ability, stiffness and prevention of
stress concentration.
result
More precise information
about fatigue limits and
their influencing factors
Detailed instruction for production
welding (qualified process)
Less fluctuation in quality
Increased safety margin
Fig.5: Schematic representation of strategy
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2
Decreased stress level
Increased safety margin
1.
Problem definition
1.1
Design aspects
The double wall box design shown in Fig. 6 leads to single bevel
butt welds.
box
box
crosshead
Fig.6: Crank shaft housing (FEM-Model, right)
and section through the double wall box (left)
Fatigue endurance and reliability are preconditions for an engine structure.
Omitting hot spots and local stiffeners are the most effective measures to avoid damages during life time.
However, using thicker plates at potential critical locations is not always suitable to
increase the safety margin. Due to the fact that the fabrication process itself has a large influence on
fatigue, detailed instructions for the manufacturer help to minimise the variance of welding quality.
Feasibility and production friendliness are further aspects that needs to be considered.
The numerous years of service experiences of the existing double wall structure (without PWHT) can be
regarded and used as long-term fatigue tests, provided they are accompanied by a detailed theoretic
investigations, including stress analysis due to service loads and determination of residual stresses.
1.2
Problem of root welding of a single bevel butt weld
Concerning the fatigue behaviour, the vicinity of the root of a single bevel butt weld is in general the most
critical area. Typical weld defects in this region are shown in Figs. 7-9. Obviously they act like pre-existing
cracks. Since the inside of the box is not accessible there is no possibility for corrections from the inside.
The stress concentration due to the global geometry in combination with the local geometrical
irregularities and uncertainties regarding residual stresses makes a fatigue analysis of such joints difficult.
Possible root imperfections:
Fig.7: Lack of fusion
Fig.8: Hot crack
Fig.9: Lack of penetration
Due to the difficulties to detect root imperfections such as the ones shown above by Non Destructive
Testing (NDT) the structure needs to exhibit a sufficient degree of so-called damage tolerance. This
requires a certain representative crack-like defect to be taken into account in the fatigue analysis and
assessment. According to Linear Elastic Fracture Mechanics (LEFM), crack propagation strongly depends
on residual stress in the affected zone. Compressive residual stresses in the root area can actually
prevent a crack from growing, whereas tensile residual stresses would accelerate it.
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3
1.3
Theoretical assumption of residual stress distribution
In a multi-layer single bevel butt weld the residual stress distribution as well as the deformations depends
mainly on the boundary conditions. If there is no restriction to relative rotation, then each additional weld
layer adds a certain amount of distortion, so a considerable rotational angle χ can accumulate, causing
tensile stresses T’ in the root area. If the relative movement of the plates is prevented, then the
corresponding clamping force gives rise to a bending movement M, which causes compressive stresses
in the root area. In the real case, a relative rotation of the welded plates is prevented by the spatial
arrangement of stiffeners, so it corresponds to Var.2 rather than 1.
M/a
a
T
T
M
χ
T’
C
Variant 2: clamped to
base plate
Variant 1: No kinematical
restriction during welding
Fig.10: Schematic representation of the effect of the kinematical boundary condition /4/
Mean stress sensitivity and residual stress factor according FKM /3/
ƒÐ
AK
2,19
71
R
50
=
1,48
‡
Transverse butt welds
welded from one side
without backing bar,
full penetration
Root controlled by NDT
FAT
0
Description
=
Structural Detail
According to FKM /3/ the influence of
residual stresses can be classified as: 1,0
Low: stress relieved structures
Moderate: thin-walled simple structures
High: thick-walled complex structures
SULZER 85
(fatigue test)
R
1.4
R
29 N/mm2
,5
=0
25
Fig.11: Haigh
diagramm (FAT 71)
acc. to FKM
This ranking of residual stresses
0
represents only their negative impact
-75
-50
-25
0
25
50
75
ƒÐ
on fatigue strength in case of tensile
m
residual stresses. However, if compressive residual stresses prevail in the root region, their effect may be
positive, as shown below.
1000
fatigue test : crack
fatigue test : no crack
S-N-Curve for experimental datas (95% survival probability)
2
stress range (log N/mm )
S-N-Curve for experimental datas (50% survival propability)
stress ratio: R = 0 / no PWHT
100
s (R=0) = ± 46 N/mm2
In 1985, SULZER have performed
fatigue tests. The results of
2
σAK = ± 46N/mm
(specimen not stress relieved) shows
higher fatigue limit than excpected
according FKM /3/.
This specific application utilise the
beneficial influence from residual
stress.
Fig.12: Fatigue test results
10
1.0E+05
1.0E+06
1.0E+07
number of load cycles (log N)
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4
1.0E+08
2.
Stress evaluation
2.1
Strain gauge measurement
Strain gauge mesurements are used for verification of calculated engine loads and to assess the
accuracy of FEM results.
10
ES
0
0°
30°
60°
90°
120°
150°
180°
210°
240°
270°
300°
330°
360°
crank angle (0° = TDC Cyl.1)
stress (N/mm2)
-10
cyl.4
gear
-20
Sx measured FPS
Sx measured ES
-30
Sx calculated
Sx calculated
-40
FPS
-50
Fig.13: 10RTA96C Comparison of mesured and calculated stresses on gear column
100
50
DMS 89
DMS 91
DMS 93
DMS 95
DMS 97
DMS 101
DMS 103
force cyl.2
force cyl.3
4030
30°
0°
91
330°
93
0
0°
strain (mst)
89
30°
60°
90°
120°
150°
180°
210°
240°
270°
300°
330°
60°
360°
95
crank angle (0° = DTC cyl.2)
-50
300°
cyl.3
97
90°
-100
270°
101
cyl.2
120°
-150
side plate
-200
measuring position
Location on top of column less stressed due
to better support of inclined side plate.
240°
103
150°
210°
180°
Inclined side plate
-250
0
600
400
200
0
-200
Fig.14: Exposition of guide shoe force and their impact on stress in side plate
On the engine structure all three normal stress components (transverse,
longitudinal and shear, see Fig. 15) are present. To measure them a triaxial
strain gauge (“rosette”) is required. The evaluation based on nominal stresses
allows a specific assessment of all critical load components. Which one of them
is crucial depends on the weld joint and the load direction. For a single bevel butt
weld with an open root the transverse stress is the most critical load component.
Fig.15: Stress components
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5
-400
2.2
FEM – Calculation
All types of loads acting on the component – stationarly and cyclic ones - must be included in the
FEM – calculation. Often it is suitable to use a shell-element model for such thin-walled components.
Nevertheless a detailed evaluation of the non-linear stress distribution over the plate thickness needs
20-node hexahedron elements. For stress evaluation two different concepts are common:
A) Geometric stress: this concept includes all stress raising effects of a structural detail excluding all
stress concentration due to the weld profile itself.
B) Effective notch stress: this concept calculates the total stress at the root of a reference notch, for
which a radius of 1mm and linear-elastic material behaviour is assumed.
hot spot
geometric
stress
Stress on
surface
Radius 1mm
F
F
Fig.15: geometric stress
Fig.16: effective notch stress
Both these concepts are illustrated in the following by a stress simulation of an actual case of a welded
structure where cracks occurred (Fig. 17).
1
side
plates
Section A - A
A
A
stiffening ribs
bending
tension
compression
longitudinal
Load
1
fuel pump
support
2
2
Fig.17: Overview of crack situation on fuel pump support
Stiffening ribs inside of the crankshaft housing prevent a rotation of the
side plates (caused by the one-sided guide shoe force). This generates
local stress concentrations on the root side of the vertical weld. Crack
propagation as observed (see no.1 inside crankshaft housing / and no.2
outside crankshaft housing) is explained by the residual stresses
(depending on the welding sequence) and the ones due to the service
load (depending on the crank angle offset).
Fig.18: View inside box
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6
SY
SX
50
hot spot
stress
dynamic stress (N/mm2)
40
geometric
stress
30
20
coarse mesh
fine mesh
10
nominal stress
0
0
50
side sloping plate
The geometric stress is mainly
influenced by the design of the edge
relief in the siffening rib. This edge
relief allows for local bending of the
side plate. The bending of the side
sloping plate, which also has an effect
on the stresses in the root area, is not
covered by the geometric stress
concept. For this purpose, the
effective noch concept is preferable.
The use of hot spot stress for
evaluation is only convenient for
comparison of different
countermeasures.The only restriction
is to use the same FE-mesh size.
100
150
distance (mm)
200
R8
0
Fig.19: Notch stress (1 mm radius)
250
side plate
300
stiffening rib
Fig.20: geometric stress (hot spot)
A detailed stress analysis including the defect exhibits a stress concentration in the root area (Fig. 19).
The safety against fatigue is just sufficient if the weld quality is according ISO 5817 (full penetration)
Taking the compressive residual stresses in the root region (see Sect. 3 below) a certain root imperfection
is accepable without any rsik of cracking. In fact, experience - many engines are running without
problems, although quality deficiences in the root are likely to be present - confirm this assumption of a
beneficial influence of the residual stresses on the type of weld considered here.
Stress evaluation with FEM program ANSYS
strain
ANSYS
ANSYS command: prnsol,epto
ε x ε y ε z ε xy ε xz ε yz
y
Evaluation
εa = ε x
ε +ε +ε
ε b = x y xy
stress
ANSYS command: prvect,pdir
Principle
stress
S1
S2
S3
y
vector
S2
dx1
dx2
dx3
dy1
dy2
dy3
2
εc = ε y
x
dz1
dz2
dz3
dx1
α
dy2
dx2
Def: S1 > S2 > S3
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7
εc
εa
α = a tan (dy1 dx1 )
S1
dy1
εb
Selection of both
principle stresses which
are perpendicular to
x
each other
S3 = 0
3.
Residual stress distribution
3.1
Residual stress measurement
Residual stresses, enclosed in the welding seam are not of primary interest. There will always be a
balance between compression and tension. Much more important with regard to the manufacturing
process is the influence to the structure itself. If there are no constraints imposed on the displacements in
any directions, deformations due to the welding distortions occur rather than residual stresses. If
deformations are restricted by geometrical constraints, then much higher residual stresses are produced.
These two conflicting effects have to be optimized for a certain welded structure. An optimised welding
procedure has to allow for free shrinkage to avoid tensile residual stresses on one hand, and to restrict
those deformations that lead to compressive residual stresses (with positive or negative effect on fatigue)
and shape stabillity on the other hand.
low
residual stresses high Fig.21: Schematic representation of interaction
high deformation
low
between residual stresses and
deformation due to degree of restraint
degree of restraint
free
rigid
An efficient method to determine residual stress distributions experimentally is the cut-compliance method
/5/. Measurements performed on a testweld are desribed in /4/. The measurement setup is shown in Fig.
22, and the corresponding results in Fig. 23.
Strain gage F
CC-cut
Fig.22: Experimental set-up
to measure residual
stresses at the weld
Strain gage R
150
cut from root
cut from upper surface
Residual Stress [MPa]
100
50
0
0
5
10
15
20
-50
-100
-150
Root
Distance from root [mm]
Upper
Surface
Fig.23: Measured residual stresses at the weld root section
To cover the root region as well as the upper surface the cutting required by the CC-method was
performed from both sides. As shown in Fig. 23, both measurements are in good agreement with each
other in the overlapping central region, indicating the reliability of the measurement. Apparently, both the
root and the upper surface region are predominantly subjected to compressive stresses. Only in the
central region tensile stresses prevail. Note that this measurement was made after releasing the clamps
that were applied to prevent distortion (see Fig. 10 and Fig. 32). Thus, in the real structure an additional
compressive stress is to be expected in the root region.
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8
3.2
Residual stress calculation
3.2.1
Simulation model
The simulation of the residual stress distribution was
performed by means of a 3-dimensional model of a
representative portion of the weld (Fig. 24). The
model included the cross sections of the material in
pre-weld situation and the 5 layers. The cross
sections were adjusted to the boundaries of the
layers as given in micrograph results. The elements
representing the filler material were activated with
EALIVE according to the process specifications.
Fig. 24: Simulation model
The temperature simulation was done by a transient thermal analysis.
The heat source was simulated with Gauss distribution across the
elements of each layer. The position and direction of the heat source
was adjusted according to the contour of the weld bead.
The structural mechanics boundary conditions were set to represent
planar cross sections at both cut sections of the model. These
constraints are representing the center portion of a very long weld. The
other displacement constraints were chosen according to the inservice
condition of the welded part. They do not relate directly to the
experimental setup described hereafter.
The steel in the weld region and the heat affected zone undergoes
changes in the microstructure. This behavior was simulated by two
different methods, the STAAZ approach and the Leblond approach as
explained below.
3.2.2
Fig. 25: Mesh of weld region
Simulation using the STAAZ method
The simulation was done using the weld
simulation tool SST “Schweißsimulationstool”
licensed by CADFEM. It is based on the
ANSYS program and includes weld specific
simulation algorithms and other features. The
STAAZ approach to simulate the
microstructural changes was proposed by
Ossenbrink, Michailov /1/. This approach
simulates the thermal strain due to
metallurgical phase transformations. To cover
the influence of the thermal history of the
Fig. 26: STAAZ approach
material a triad of significant parameters is
used including the maximum temperature (ST), the time in the austenitic phase (A) and the cool-down
gradient (AZ). This methods directly uses dilatogram data. A set of material test data is required including
dilatogram tests for different triad combinations. During the analysis an interpolation is made at each
location inside the FEM model according to the local triad. This
STAAZ approach avoids the uncertainties of CCT derivation and the
succeeding steps to transfer the CCT data into numerical functions
and parameters. The microstructure contribution to thermal strain is
directly included, there is no calculation of microstructure fractions.
The calculation of the temperature distribution oder time and the
structural mechanics is done in sequence. The transient thermal
analysis uses estimated phase change behavior. The temperature
distribution and function over time were compared to experimental
data. There was a good agreement.
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9
Fig. 27: Temperature as a
function of time
After the thermal analysis is finished the temperatures and
the thermal result triads are transferred to the structural
mechanics simulation. This simulation includes the thermal
strain influenced by microstructure changes. The
mechanical behavior simulates nonlinear stress-strain
relationship, depending on temperature.
Fig. 28 shows the thermal strain contribution during heatup and cool-down cycle of a position of layer 1. It can be
nd
seen that the α-γ change during heat-up and the reverse
Fig. 28: Stress after 2 layer
γ-α change during cool-down occurs at different
temperature levels depending on the local triad values of
maximum temperature (ST), the time in the austenitic phase (A)
and the cool-down gradient (AZ).
Fig. 29 shows the distribution of equivalent stresses after the 2nd
layer is finished. The results were reasonably comparable to test
data. The application of the STAAZ approach has uncertainties
when applied to multilayer welds, since the remelting of material
and its influence on the thermal strain results in unreliable
material data.
nd
Fig. 29: Stress after 2
3.2.3
layer
Simulation using Leblond kinetics
For comparison purposes the simulation was also
done using the Leblond microstructure kinetics
approach /2/. This feature is also available in the
SST Schweißsimulationstool (weld simulation tool).
The Leblond approach approximates the phase
change between the fractions austenite, ferrite,
bainite, martensite, pearlite with analytical functions.
Usually the parameters are derived from CCT
diagram data. Considering the source of CCT
diagrams being dilatograms there are several
processing steps between the data source and the
application. This results in considerable
Fig. 29: Leblond kinetics approach
uncertainties and approximations using this method.
The analysis simulates the change of fractions over time. Thermal and mechanical material values are
combined according to the local fraction relation.
Using the Leblond microstructure kinetics approach in SST the transient thermal field and the structural
behavior are simulated simultaneously in a coupled analysis. The model is meshed with coincident
elements for the thermal distribution (solid70) and the structural mechanics (solid185). The coupling is
done internally in a weak procedure (load vector coupling).
Fig. 30: Microstructure fractions
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10
The initial microstructure conditions are 90% ferrite, 10% pearlite. The resulting distribution of the phase
fractions shows a martensitic microstructure with remaining local austenite spots.
The resulting distributions of
stresses SX (perpendicular to
the weld) and SZ (longitudinal)
are shown in Fig. 31. These
stresses are used to consider
the influence of residual
stresses on service stress
limits.
Fig. 31: Leblond kinetics approach
3.3
Test welding
Laboratory test were carried out to gain temperature and stress data for verification of analytical weld
process simulation. The geometry indicates similar geometry as the real structure.
Position of stress measurement
(at the top and at the bottom)
35°
5
4
3
2
1
20
60
3
100
100
0
2 00
Fig.32: Test configuration
5th layer
300
250
A
60
40
40
B
4th layer
20
D
C
150
3rd layer
Process:
GMAW (CO2), PA-position
2nd layer
position A
position B
position C
position D
1st layer
temperatur (°C)
200
Welding parameters:
1st layer
22V 160A 20cm/min E=1,1KJ/mm
2nd layer
3rd layer
4th layer
5th layer
100
50
pre - heating
14 min
18 min
19 min
20 min
0
0
1000
2000
3000
4000
5000
time (min)
Fig.33: Measured temperatures
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11
6000
7000
8000
9000
10000
after welding (room temperature)
Restraint
190 (-170)
top side (bottom side)
N/mm2
loosen
20 (0) N/mm2
-10 (-130) N/mm2
-70 (-70) N/mm2
210 (-150) N/mm2
40 (30) N/mm2
10 (-140) N/mm2
-60 (-80) N/mm2
Fig.34: Measured stresses
Transverse stresses:
The results of the stress measurement shows, as expected, heavy bending of the plate due to the
2
irregularity of the welding volume. The transverse stresses of about ±180 N/mm have nearly completely
relaxed after the restraints are removed.
Longitudinal stresses:
2
The stresses in longitudinal direction of about - 70 N/mm after loosening the clamps are plausible. These
compressive stresses in the plate indicate high tension in and near the weld. During restraint condition
2
there is an additional bending in longitudinal direction of about ± 60 N/mm . This may be an effect of the
plate dimension and the clamping arrangement and is not likely to appear on the real component.
3.4
Effect of preheating on residual stress
Preheating prior to welding is a simple measure to reduce the shrinkage stresses. The main reason to
utilize preheat is to decrease the cooling rate in the weld and the base metal.
Circular welds
Longitudinal welds
Less deformation after welding
in longitudinal direction
Elongation of inner part
by reason of preheating
Elongation of plate by
reason of preheating
restraint outer part
(prevents free
transverse shrinkage)
thin plate not
restraint
(transverse
shrinkage
possible)
thick
plate
inner part
Preheating reduces longitudinal
shrinkage stresses
Preheating increases transverse
shrinkage stresses
transverse shrinkage
Fig.35: Preheating of high joint restraint
longitudinal shrinkage
Fig. 35 shows schematically representation the effects of preheating. Whether they are positive or
negative depends on the weld joint application. The single bevel butt weld in question corresponds to the
left application. This means that preheating is advantageous, since it reduces longitudinal shrinkage
stresses that are produced by the present boundary conditions.
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12
4.
Fatigue strength analyses
4.1
Stress evaluation
The welding seams of the present structure are subjected to multi-axial stresses. Not only the magnitude
but also the direction of the principle stress changes during one load cycle, i.e. one revolution of the crank
shaft. A variety of different hypothesis for fatigue under such conditions have been published. Wärtsilä
uses the von Mises criterion (GH) with signed principle stress now for many years. This approach results
in high stress values and is often a very conservative interpretation of the load.
The normal stress criterion (NH) enables less conservative predictions of the fatigue strength to be made.
Cracks are formed preferably perpendicular to the maximum tensile stress. According to fracture
mechanics the stress component normal to a possible crack-plane is the crucial one. The normal stress
criterion takes into consideration possible root imperfections in correlation to the critical load component.
Surface:
Root:
20
A
B
12
C
10
angle (0° = TDC Cyl.1)
B
C
4
60°
120°
180°
240°
-10
-20
300°
360°
stress sx (transverse)
stress sy (longitudinal)
60°
120°
-8
-20
G6
A
cyl.6
- +
B
F5
cyl.5
F
5
240°
300°
360°
C
F6
cyl.6
+
-
180°
-4
-16
-40
F5
0°
-12
stress sxy (shear)
-30
F5
angle (0° = TDC Cyl.1)
0
2
ƒÐ
N = }14 N/mm
0°
dyn.stress (N/mm2)
0
2
ƒÐ
N = }27 N/mm
dyn.stress (N/mm2)
A
8
cyl.6
-
cyl.5 F6
cyl.5
G5
F6
A
A
Fig.36 : stress condition on root and surface side (geometric stress concept)
The crucial stress component due to manufacturing deficiencies (imperfections
perpendicular to the stress direction) is the transverse stress σx. According to Fig.
36, the maximum stress range occurs between position A and B.
Normal stress criterion (NH):
σN =
σ x +σ y
2
2
A-A
⎛ σ x −σ y ⎞
+ ⎜⎜
⎝
2
⎟⎟ +τ xy2
⎠
Shear stresses are very low at the relevant positions and can be neglected. If the normal stresses σx and
σy always act proportional or synchronous in phase the values are to be inserted in the formula with the
same (positive) signs /3/.
Regarding stresses on welding surface the transverse and longitudinal stresses act synchronous in
phase. This means that there is no additional influence from the longitudinal stress component. Therefore
normal stress criterion will be generated only by the transverse stress:
σN = σx
2
The decisive root area of the weld is loaded by σx = ± 14 N/mm .
Fatigue resitance is usually derived from constant or variable amplitude tests. The fatigue assessment of
classified structural details (FAT-classes) is based on the nominal stress range. This fatigue values are
based on representative experimental tests and includes the effects of normal fabriction standard.
2
Comparison of σx with the allowable stress range for FAT 71, σAK = ± 29 N/mm , reveals that endurance
is guaranteed with a considerable margin, provided the quality of the weld corresponds to FAT71 (full
penetration without any imperfections). Nevertheless, long term experiences shows that normally a
sytematic defect in the root must be taken into account. In this case the safety drops down dramatically.
th
13
4.2
Fatigue resitance
stress concentration factor kt
As the geometrical stress concept does not apply to root defects, it is necessary to evaluate the fatigue
resistance by use of more refined techniques such as effective notch stress concept or linear elastic
fracture mechanics.
Possible root defects of single bevel butt welds were classified with:
a) stress concentration factor kt for lack of penetration failure (partial penetration)
b) stress intensity factor KI for root crack or lack of fusion failure
3.0
F
root-side (load tension)
root-side (load bending)
M
surface-side (load tension)
surface-side (load bending)
2.5
2.0
1.5
1.0
8%
18%
28%
lack of penetration
38%
Fig.37: stress concentration factor kt for bending and tension/compression
Lack of penetration up to 4 mm (20% in case of a plate thickness of 20 mm) could easily occur if the fit-up
is not done properly. Corresponingly, the stress in the critical area will be increased by a factor 1.6 -1.8
depend on load distribution (see Fig. 37). Therewith, the safety margin of the above example drops just to
zero.
More critical defects such as root cracks or lack of fusion must be assesed with linear elastic fracture
mechanics. The resitance of a material against cyclic crack propagation is characterized by the material
parameters of the “Paris” power law of crack propagation /3/:
da
da
=0
= Co ∗ ΔK m
ΔKth =190−144∗ R
Δ K < Δ K th
if
then
dN
dN
The treshold value ΔKth is influenced by the stress ratio R as idicated in the estimation formula above.
Compressive residual stresses cause a lower R and, therewith, an increased ΔKth, which means that a
higher range ΔK is required to let a crack-like defect grow. This means that the damage tolerance of the
weld (allowance of possible or hypothetical crack-like imperfections, which could not be found by Non
Destructive Testing) is improved by the residual stresses, which is a big advantage concerning the
inspection concept.
4.3
Safety
It is shown theoretically that the residual stress has an important influence on the fatigue performance of
this welded joint. Depending on the defect size, location and orientation on one hand, and the distribution
of the residual stresses on the other, the effect on the fatigue strength can be positive or negative. To
decide about the most beneficial weld procedure and eventual PWHT the critical locations and the
corresponding residual stress distribution should be known.
orientation of
imperfections
orientation of
significant load
F
Utilisation of
residual stresses
+
To fully benefit from the positive effect of residual stress restriction and detailed information are
necessary for production:
- Optimised welding sequence
- Procedure for root repair required
- Flame straightening only in exceptional cases
Only in this special case together with a correct application the decision to not apply PWHT will be
successful.
th
14
5.
Preliminary conclusions and further investigations
Residual stresses and weld defects play a major role in the fatigue behaviour of welded structures. Their
effects should be taken into account in a theoretical analysis.
Compressive residual stresses can increase the damage tolerance of a weld significantly. Therefore, a
PWHT is not necessarily beneficial. To decide about the most adequate procedures one has to be sure of
all influencing factors. This only gives the possibility to restrict the production of welded components
regarding residual stress sensitivity.
With the SST “Schweißsimulationstool” is it now possible to apply analytical investigations for different
applications. However, regarding the variety of required parameters, the results have to be validated by
benchmark measurements.
Examples for further calculations:
A) Variation of heat input: number of layers, excessive parameters
B) Influence of preheating, flame straightening
C) Variation of restraint: circular weld, restricted shrinkage
D) Repair welding
6.
References
/1/
SST Schweißsimulationstool – Werkzeug zur vollständigen numerischen Simulation des
Schmelzschweißens, Schlussbericht zum BMBF-Verbundprojekt 02PD1051, 2005
/2/
J.B.Leblond, J.Deveaux, Acta metall., 32 (1984), 137.
/3/
FKM-Guideline, Analytical strength assessment of components in mechanical engineering
th
5 , revised edition, 2003, Forschungskuratorium Maschinenbau
/4/
Schindler, H.J., Martens, H.J., Sönnichsen, S., “A Fracture Mechanics Approach to Estimate the
Fatigue Endurance of Welded T-Joints including Residual Stress Effects”, to be published in
Fatigue and Fracture of Engineeing Materials and Structures, 2006
/5/
Schindler, H.J.,”Experimental determination of crack closure by the cut compliance method,
“ASTM STP 1343, R. McClung and J.C. Newman, Eds., American Soc. For Testing and
Materials, West Conshohocken, PA. (1999), 175-187
th
15

Simulation einer Mehrlagenschweißung im Großmotorenbau 2.6.2

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