IWSD Module 4 -Design of Welded Joints

International Welded

Structure Designer----IWSD

Module 4. Design of welded joints

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IWSD Version1.0 Date 18.10.2010

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4.1 Categories of welded joints ..................................................................................................... 5

4.1.1 Classification of welded joints ........................................................................................................................... 5

4.1.2 Definitions ............................................................................................................................................................... 14

4.1.3. Correlation of loading and control of welds ............................................................................................ 17

4.1.4. Welded joints realized on actual metallic structures ........................................................................... 20

4.2 Design of welded joints with predominantly static loading .................................................... 23

4.2.1. Scope ........................................................................................................................................................................ 23

4.2.2. Basis of design ...................................................................................................................................................... 25

General requirements ........................................................................................................................................................ 25

4.2.3. Welded connections ........................................................................................................................................... 28

General ...................................................................................................................................................................................... 28

Global analysis ...................................................................................................................................................................... 28

Loading actions .................................................................................................................................................................... 29

4.2.4. Basic principles .................................................................................................................................................... 30

Calculation of welded joints ............................................................................................................................................ 30

Directional method ............................................................................................................................................................. 33

Simplified method ................................................................................................................................................................ 34

Resistance calculation of welds ..................................................................................................................................... 37

4.2.5. Types of stress raisers and notch effects .................................................................................................. 43

4.2.6. Determination of stress and stress intensity factors ........................................................................... 50

Definition of Stress Components ................................................................................................................................... 50

Nominal stress ....................................................................................................................................................................... 50

Calculation of nominal stress ......................................................................................................................................... 52

Measurement of nominal stress .................................................................................................................................... 53

4.2.7 Structural hot spot stress ................................................................................................................................. 53

General ...................................................................................................................................................................................... 53

Determination of structural hot spot stress ............................................................................................................ 55

Calculation of structural hot spot Stress................................................................................................................... 56

Measurement of structural hot spot stress .............................................................................................................. 60

Determination of stress ..................................................................................................................................................... 61

Structural hot spot stress concentration factors and parametric formulae ............................................ 61

4.2.8 Effective notch stress ......................................................................................................................................... 62

Calculation of effective notch stress ............................................................................................................................ 62

Stress intensity factors ...................................................................................................................................................... 63

Calculation of stress intensity factors by parametric formulae ..................................................................... 63

4.3 Design of welded joints with predominantly fatigue loading .................................................. 64

4.3.1 Basic principles ..................................................................................................................................................... 65

Increasing accuracy and efficiency of mechanical characteristics .............................................................. 68

Distribution function of durability at the action of variable loading .......................................................... 69

Statistical processing method ........................................................................................................................................ 70

4.3.2 S – N Diagram......................................................................................................................................................... 70

4.3.3 Collective applications of voltage .................................................................................................................. 71

4.3.4 Fatigue resistance ................................................................................................................................................ 73

4.3.5 The average voltage effect ................................................................................................................................ 75

4.3.6 Fatigue resistance of classified structural details .................................................................................. 77

4.3.7 Linear Damage Calculation by "Palmgren-Miner" ................................................................................. 80


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4.3.8 Nonlinear Damage Calculation ....................................................................................................................... 83

4.3.9 Fatigue resistance against structural hot spot stress ........................................................................... 83

A.Fatigue Resistance using Reference S-N Curve .................................................................................................. 83

B. Fatigue resistance using a reference detail ........................................................................................................ 84

4.3.10 Fatigue resistance against effective notch stress ................................................................................. 86

4.3.11 Fatigue strength modifications .................................................................................................................... 86

4.3.12 Wall Thickness.................................................................................................................................................... 87

4.3.13 Improvement techniques ............................................................................................................................... 88

Applicability of improvement methods ..................................................................................................................... 89

Burr Grinding ......................................................................................................................................................................... 90

TIG dressing ............................................................................................................................................................................ 91

Hammer peening .................................................................................................................................................................. 91

Needle peening ...................................................................................................................................................................... 92

4.3.14 Effect of elevated temperatures .................................................................................................................. 92

4.3.15 Effect of corrosion ............................................................................................................................................. 93

4.3.16 Fatigue resistance against crack propagation ....................................................................................... 93

4.3.17 Fatigue assessment by crack propagation calculation ...................................................................... 95

4.3.18 Fatigue assessment by service testing ...................................................................................................... 96

A. General ................................................................................................................................................................................. 96

B. Acceptance criteria ........................................................................................................................................................ 98

C. Safe life verification ....................................................................................................................................................... 98

D. Fail safe verification ...................................................................................................................................................... 99

E. Damage tolerant verification .................................................................................................................................... 99

4.3.19 Fatigue resistance of joints with weld imperfections ........................................................................ 99

A.Types of Imperfections .................................................................................................................................................. 99

B. Effects and assessment of imperfections ........................................................................................................... 100

C. Misalignment ................................................................................................................................................................. 101

D. Undercut .......................................................................................................................................................................... 102

E. Porosity and inclusions ............................................................................................................................................. 103

4.3.20 Fatigue resistance values for structural details in steel and aluminium assessed on the

basis of nominal stresses .......................................................................................................................................... 105

4.4 Design against brittle fracture .............................................................................................. 124

4.4.1. General ................................................................................................................................................................. 124

4.4.2. Mechanical behaviour under tensile loads ............................................................................................ 125

4.4.3. Impact testing .................................................................................................................................................... 127

A. Notched-bar impact tests ......................................................................................................................................... 127

B. Instrumented Charpy test ........................................................................................................................................ 130

C. High rate impact test ................................................................................................................................................. 132

C1. Explosion bulge test ................................................................................................................................................. 132

C2. Drop weight test. ....................................................................................................................................................... 134

C3. Robertson crack-arrest test .................................................................................................................................. 135

C4. Fracture analysis diagram .................................................................................................................................... 135

4.4.3. Fatigue testing ................................................................................................................................................... 138

4.4.4. Fracture mechanics approach .................................................................................................................... 140

A. General .............................................................................................................................................................................. 140

B. Linear-elastic fracture toughness testing ........................................................................................................ 144

C. Nonlinear fracture toughness testing ................................................................................................................ 145


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4.4.5. New standards for fracture mechanics testing of metallic materials ......................................... 146

List of figures ............................................................................................................................. 148

List of tables .............................................................................................................................. 150


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4.1 Categories of welded joints

Objective

The students will understand the differences between functional weld categories and how the

design requirements will depend on the categories

Scope

Weld categories

Primary load carrying joints

Connecting joints

Binding joints; accessory joints

Expected results

Identify various classes of welded joints based on their function.

Explain the load-bearing requirement of various weld categories.

Explain the need to avoid the under- and over-size of the throat thickness.

Illustrate the role of joint preparation and weld penetration for load-carrying joints.

Identify joint categories from an engineering structure.

4.1.1 Classification of welded joints

The welding operation must be understood as the realization of a non-detachable joint

between two or more parts, named components, by heating and/or applying a pressure with

or without using filler material. In the welding area, material of components can be in melting

or plastic sate assuring the continuity of materials the components are made out of. In

technical literature, standards and norms, inclusively welded joints are classified according to

the welder’s position against the joint, the way the parts to be welded are situated one

against the other and the way edges are processed, inclusively when the thickness of jointed

parts exceeds 8-10 mm.

The classification of welded joints takes in account the international terminology (Figure 1).

a) Considering the welding process there are:

a. Welding by melting

b. Pressure welding

b) According to the purpose:

a. Resistance joining

b. Sealing up joining

c. Hardening joining


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d. Surfacing joining

c) Position of components in the joining process:

a. Butt welding, when components are in the same plane (1, 4)

b. Fillet welding, with constructive variants

i. T, when components form an angle by joining (2)

ii. Overlapped, when components are in contact on a certain area

(3)

d) According the direction of loading:

a. Frontal joining, when the loading is transversal against the

longitudinal axis

b. Longitudinal joining, when the loading is on the direction of the

longitudinal axis

e) Welding position:

a. horizontal (5),

b. flat weld (6),

c. vertical up, vertical down (7a and 7b, 7c),

d. horizontal vertical weld (8),

e. overhead (9).

Welding positions differentiate them according to the circular scale disks accepting the

horizontal line as reference, so:


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Figure 1 Classification of electric arc welded joints

.

Horizontal position, in the range 45 ÷ 135 (7b)

Position in vertical plane, in the range 135 ÷ 225 şi 315 ÷ 45 (7c)

Position overhead, in the range 225 ÷ 135 (7c)


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f) Continuity of deposited welds:

a. Continuous joining, when the length of the joining is identical with that of

components to be welded

b. Discontinuous joining, when the joining length sum is more reduced than that

of components to be welded

c. Weld spots joining, when the components joining is locally assured

g) Number of cooling ways:

a. bimetallic (10)

b. multimetallic (11, 12)

h) Accessibility when welding:

a. one side joining (13, 14, 15)

b. both sides joining (16, 17a, b, c)

i) Number of weld metal passings:

a. one passing (18)

b. more passing’s (19)

j) Thickness uniformity:

a. equal (20)

b. unequal (21)

k) Shape and geometry of the groove:

a. butt welds: in I(13), in V(14), in double V or X(16), K

b. fillet welds: with non-processed web (17 b), with processed web (17 c)

l) Metallurgical group of materials to be welded:

a. Homogeneous

b. heterogeneous

Homogeneous joints are realized with base and filler materials belonging to the same

metallurgical group. The heterogeneous ones, have one or both components, and the filler

material, respectively from different metallurgical groups.

m) The mechanization degree can be:


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a. manual welding

b. semi-mechanized welding

c. automated welding

Analyzing the way welded joints are formed according to EC 3-1-8 the following types of

welded joints are defined:

1. Fillet welds, which can be continuous or intermittent fillet welds,

2. Fillet welds all round, in fact fillet welds on the contour of holes made in one of the

overlapped components,

3. Butt welds,

4. Plug welds and

5. Flare groove welds.

Table 1 presents the classification criteria and the type of welds. Butt welds can be realized

with full or partial penetration.

In the category of fillet welds are framed all welds between components making between

them an α angle in the range 60° and 120°. Besides the common fillet welds, which thickness

„a” is considered equal with the height of the inscriptible triangle in the cross section of the

weld, descended from its root on the exterior side, EC 3 also stipulates fillet welds with full

penetration, which thickness depends on the technology and equipment used. The design

codes foresee the obligation to check by preliminary test probes.

Table 1 Classification criteria and weld type according to EC 3-1-8.

No.

Classifica

tion

criteria

Weld type Representation

1 Butt

welds

Butt welds

with full

penetration

and V, 2V, U,

2U groove


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2 Butt

welds

Butt welds

with

penetration

and V and U

groove

3 Fillet

welds

Continuous

welds

4 Fillet

welds

Fillet weld

with 2x½V

5 Fillet

welds

Fillet welds

with V, J, K

and 2J

groove


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6 Fillet

welds

Intermittent

welds:

-alternative,

-bilateral,

7 Fillet

welds

Fillet welds

with deep

penetration

8 Fillet

welds

Fillet welds

with partial

penetration

completed

with

deposition

9

Overlapp

ed welds

Continuous

fillet welds:

- lateral,

- frontal


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10 Overlapp

ed welds

Fillet welds:

- all around,

- oblong

11 Overlapp

ed welds Plug welds

12

Flare

groove

welds

Welds

between

oblong

groove welds

Accepting this type of weld intervene following the improvement of welding technologies,

which at present allow the significant penetration of fillet welds in the material of the welded

components.

So, it is possible to realize actual weld thickness “a” bigger than those considered in common

fillet welds, where the penetration is more reduced and not taken into account.

Obviously, here appears as necessary the direct designer- executants relation, which have to

collaborate during the design stage having as objective the possibility to realize deep

penetrated fillet welds, relation that is not a problem for firms realizing the design

documentation, execution and the assemblage of metallic structures.

As regards the full penetrated welds, both the butt welds and the T welds, changes in

designation appear. For example, X weld is designated as double V weld, the K one is

named double J weld, and the ½ V and ½ U welds are named semi V, and J, respectively.

A significant difference consists in accepting the partial penetration welds, both for butt welds

and for T welds; they are named double V and double U welds, semi double V welds,

respectively.

EC 3 also stipulates for T welds the possibility to use butt welds with partial penetration

completed with fillet welds, which thickness is established according to specifications of

design codes.


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4.1.2 Definitions It is necessary to introduce main vocabulary notions under the form of technical terms,

weblated with components of single sided joint, double sided joint, respectively (Figure 2).

The following definitions are used:

basic component of a joint: specific parts of a joint that has an identified contribution

on structural characteristics;

Figure 2 Components of single sided joint, double sided joint, respectively

Connection: a place where two components and inter-connection means are

interconnected;

Connected member: element that is supported by the element it is connected to;

Joint: assembly of basic components which make possible the connection of

elements so that relevant forces and internal moments can be transferred form one to

another. As for example, a beam- column joint consists in a web type cassette in a

connection (single-sided joint) or two connections (double-sided joint),

Joint configuration: type or location of joint or joints in an area where two or more

inter-connected elements meet (Figure 2);

Structural properties of a joint: resistance to internal forces and moments in

interconnected elements, rigidity and its rotation capacity;

Uniplanar joint: in a lattice structure a uniplanar joint connects elements situated in a

single plane.


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Figure 3 presents images of above classified welds, to be identified and commented by

students.

Electric arc welds are classified according to different criteria:

1. According to the joint type:

a. butt welds

b. fillet welds

2. According to the position the welds are made, butt welds can be:

a. horizontal welds, in horizontal plane

b. horizontal welds, in vertical plane

c. vertical welds (can be performed up-down and down-up)


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d. overhead welds which are the most difficult to be performed

3. Fillet welds can be performed:

a. horizontal fillet weld

b. horizontal weld


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c. vertical weld, performed by descending or ascending the electrode

d. overhead weld

Figure 3 Types of welded joints

4.1.3. Correlation of loading and control of welds

Design codes stipulate checking of stresses in welds with relations: as < Rs = γ R,

respectively Ts < Rs = γ R, where γ is a coefficient depending on the loading nature, which

values are presented in table 2.


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Table 2 Provisions regarding the correlation of loading and control of welds

Joint type

Weld type

Loading and

calculus

relations

Performed

welds

Control of

welds γ Rs = γR

1 2 3 4 5 6

BUTT WELD

with deep penetration

bxa

xaabsA 2

6

2baWs

compression

s

c

s

s

c RA

N

automated

semiautomat

ed manual

Common

means 1 R

tensile

i

s

s

s

i RA

N

automated - || - 1 R

semiautomat

ed manual - || - 0.8 0.8 R

semiautomat

ed manual

With X

or γ rays 1 R

shearing

s

f

s

s RA

T

automate

semiautomat

e manual

Common

means 0.6 0.6 R

bending

s

inc

s

s RW

M

automated - || - 1 R

semiautomat

ed manual

Common

means 0.8 0.8 R

OVERLAPPED

Filet welds

3 mm ≤ a ≤ 0,7 · tmin

sis laA ; sis ll

Tensile-

compression

s

s

s RA

T

automated

semiautomat

ed manual

Common

means 0.7 0.7 R

T Tensile automated Common 0.7 0.7 R


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As it results from table 2 coefficient γ and finally the calculus resistance of welds, depend on

the calculus resistance of the material to be welded, loading mode of the weld (γ = 1 for

compression, tensile, respectively, for controlled but welds with performance procedures, γ =

0,8 for tensile loading in butt welds, if the weld control is made with less performing

procedures, γ = 0.6 for butt welds shear loaded and γ = 0.7 for fillet welds, where only

tangential stresses, t are checked.

In the calculation of weld sizes according to EC 3-1-8, limits of geometric sizes are

stipulated.

For example, for the fillet welds thickness “a” the following condition has to be respected:

3 mm <a <0,7 tmin

The “a” values have to be checked by measurements and preliminary probes, in the

case of fillet welds with deep penetration, respectively of butt welds with partial

penetration completed with fillet welds.

For the minimum weld length, some norms foresee 40 mm, while EC 3 foresees only 30 mm,

but keeps the prescription: lmin - 6a.

It must be retained that EC 3 foresee the acceptance of fillet welds with constant thickness

on the whole length, if it can be practically realized, not considering the existence of final

craters form the ends of the welds. Contrary, the reduction of weld with “2a” is maintained.

Additionally the return of welds is accepted, in the same plane, after the corner of the

overlapped components. The return of weld is considered when calculating the weld length,

if it has the same thickness “a” as the its rest.

Fillet welds

3 mm ≤ a ≤ 0,7 · tmin

62

2baWs

; ls=2b

compression

shearing

s

s

s RA

N

baAs 2

semiautomat

ed manual

means

bending

s

s

s RW

M


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When the distribution of stresses along the fillet welds is significantly influenced by the

rigidity of elements or of jointed components, the non-uniformity of this distribution is

considered by using a reduced effective length: bef. When the length of welds exceeds 150a,

resistance of the weld is reduced with a factor JLw < 1.

C 3-1-8 also provides special restrictions to use single sided fillet welds and the butt welds

with single sided partial penetration, in case of bending and tensile stresses.

4.1.4. Welded joints realized on actual metallic structures Figure 3 presents welded joints realized on actual metallic structures.

a) b) c)

d) e)


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f) g)

i) h)


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j)

k) l)

m)

Figure 4 Types of welded joints on technological equipment


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4.2 Design of welded joints with predominantly static loading

Objective

The students will understand how the throat thickness of weld will be defined in

predominantly static loaded joints.

Scope

Throat thickness

Elastic and plastic design

Deformation capacity Stress components in a fillet weld

Correlation factor for weld strength

Design strength

As appropriate, a suitable design guidance document, e.g., EN 1993 Eurocode 3-part 1-8:

Design of Steel Structures: Design of Joints, may be used.

Expected result at comprehensive level

Explain the assumptions involved in the design of predominantly static loaded joints.

Identify relevant stress values from a type stress-time history for a structural

component. Calculate the design strength of end welds based on weld stress

components.

Calculate the design strength of side welds based on weld stress components.

Calculate the strength reduction factor for long side welds or transverse stiffeners.

Calculate the needed throat thickness for a full strength primary load carrying weld.

Calculate the throat thickness for a binding welded joint.

4.2.1. Scope Present chapter deals with the design rules of joints and is prepared in accordance with the

provisions of EN 1993, Part 1-8. Consequently, the main attention is focused on design

methods for joints subjected to predominantly static loads. In conjunction with the provisions

of EN 1993, Part 1-8, the methods described may also apply for applications generating

dynamic loads, particularly from wind action, unless otherwise noted. Popular brands of steel

to be used in conjunction with the design methods presented are S 235, S 275, S 366, S 420

and S 460. Joints fatigue design, not subject to this chapter.

Components used in modern engineering usually have to bear high mechanical loads.

Because mechanical equipment is often used at or near design limitations, great care must

be employed in selecting the proper materials to use for a particular design application. The

need for high-performance materials in such industries as aerospace and power generation

has advanced the use of design parameters in the evaluation of material behaviour. The term

"mechanical behaviour" encompasses the response of materials to external forces.


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The successful employment of metals in engineering applications relies on the ability of the

metal to meet design and service requirements and to be fabricated to the proper

dimensions. The capability of metallical structures to meet these requirements is determined

by the mechanical, physical, chemical and fabrication properties of the metal components

with welding joinings (Figure 5).

Figure 5 The overview of engineering properties of materials.

Various tests have been devised to reveal the mechanical properties of materials, related

with structural integrity, with two main types of loading conditions, namely static loading and

dynamic loading. Assessment only at static behaviour is almost an idealization.

A static load is applied only once; it induces strain in the material very slowly and gradually

and remains constant throughout the service life of the component. Tension, compression,

hardness, and creep tests are used to reveal mechanical properties under a static loading

condition.

Dynamic loads can be classified into impact loads and fatigue loads. An impact load

resembles a static load in that it is applied only once. However, it differs from a static load

because it introduces strain in the material very rapidly. Charpy impact test is devised to

measure the behaviour in these circumstances of materials.

Design for structural and mechanical functions is based on the useful strength or allowable

stress of engineering materials. Usually, in such applications, materials are selected to

operate within their elastic range. Sometimes, however, machine parts and structures are

operated at stresses exceeding their elastic limit. Also, to guard against catastrophic failure,

it is taken into account that the material should plastically deform rather than fracture in case

of a sudden overload condition. During service, engineering products are usually subjected

to complex systems of stresses.

Tension, hardness, creep, impact toughness, and fatigue tests have long been used to

evaluate the mechanical properties of engineering materials in mechanical welding joining

structures. More recently, the fracture toughness test has emerged as another important test.

Compression is a less common mechanical test. Another test rarely used to specify the

mechanical properties of materials is the torsion test. As described below, the uniaxial stress-


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strain relationship determined from the tension test reveals a number of important

mechanical properties of the material, usable for engineering calculations.

Application of all simulation and testing programs is routine in many design groups in

worldwide. Set design calculations are based on results of a very broad palette of testing and

practical experiments. However, test range for particular engineering component requires

special attention. Reliability and functionality are two of the most prized qualities required

from engineered components. These are not achieved by accident. Indeed, considerable

scientific and technological endeavour is expended to help achieve them, because without

them the functionality of our whole society would be seriously jeopardized. Individual

structures are, of course, designed and manufactured to perform an individual specified

function, be they large or small. For example, a turbine should generate and transmit power,

a bridge should carry traffic, and a pressure vessel should contain a liquid or gas under

pressure. These constitute large structures, many of which are hidden from the general

public, but whose function is taken for granted by them. Other structures and components

can involve the public at a very personal level, like a mechanical heart valve or a

replacement hip, or relatively mundane domestic appliances. Yet more are hidden in

instruments and service systems, like computers, banking systems, telecommunication

systems. Loss of functionality in any one of these components can, therefore, have

consequences which far exceed the immediate damage to the component in question. Many

of welding joining metallic structures and components are required to operate under tight

controllable operating conditions, while others operate under unpredictable and

uncontrollable regimes. The environment may also be variable, regardless of the operating

regime. All the structures must be capable of operating to their design function for the period

for which that function is required, in terms of reliability and safety requirements. For a heart

valve, this may be the remaining lifetime of a patient, let say decades, while for a building or

a bridge it may be several hundred years. Additionally, operating conditions may change

throughout life: on bridges loads may increase as traffic becomes heavier and more frequent,

storage vessels may be required to store heavier charges as technology changes, electricity

generating plants may be required to switch from operating continuously at base load to two,

shift operation for peak lopping, and rail tracks may have to carry higher speed and heavier

trains. One way in which a structure may fail to meet its engineering function by mechanical

failure. This occurs when the structure, or part of it, loses its mechanical integrity to such an

extent that it ceases to perform as designed.

The mechanical integrity required to function as designed is what is meant by the term

"structural integrity", and they are all dedicated to the various methods that are inherent in

the "assurance" of structural integrity. These methods involve activity at all stages of life,

during conception, design, manufacture, operation, and decommissioning of a structure, and

the disciplines required to ensure structural integrity are all embracing.

4.2.2. Basis of design

General requirements

The design methods taken from EN 1993 assume that the standard of construction is

as specified in the execution standards set designer and that the construction materials and

products used are those specified in EN 1993 or in the relevant material and product

specifications.


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All joints shall have a design resistance such that the structure is capable of satisfying all the

basic design requirements provided by the designer according to specific codes, including in

EN 1993 parts 1-1, 1-8.

The partial safety factors γM for joints are given in table 3.

Table 3 The partial safety factors γM for joints

Resistance of members and cross-sections Mo , 1M and

2M see EN

1993 -1-1

Resistance of bolts

2M

Resistance of rivets

Resistance of pins

Resistance of welds

Resistance of plates in bearing

Slip resistance

- for hybrid connections or connections under fatigue

loading

- for other design situations

3M

3M

Bearing resistance of an injection bolt 4M

Resistance of joints in hollow section lattice girder 5M

Resistance of pins at serviceability limit state 6M

Preload of high strength bolts 7M

Resistance of concrete 0M see EN 1992

Recommended values are as follows: γM2 = 1,25; γM3 = 1,25 for hybrid connections or

connections under fatigue loading and γM3 = 1,1 for other design situations; γM4 = 1,0;

γM5=1,0 ; γM6 = 1,0 ; γM7 = 1,1.

Joints subject to fatigue should also satisfy the principles given in EN 1993-1-9.

The forces and moments applied to joints at the ultimate limit state shall be determined according to the principles in EN 1993-1-1.The resistance of a joint shall be determined on the basis of the resistances of its basic components. In terms of tensile strength, breaking combining should take place outside the typical areas.

Frequently, in the design of joints, linear-elastic or elastic-plastic analysis may be used.

Where fasteners with different stiffnesses are used to carry a shear load the fasteners with the highest stiffness should be designed to carry the design load. However there may be some


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cases.

Joints shall be designed on the basis of a realistic assumption of the distribution of internal

forces and moments. The following assumptions should be used to determine the distribution

of forces:

a) the internal forces and moments assumed in the analysis are in equilibrium with the

forces and moments applied to the joints,

b) each element in the joint is capable of resisting the internal forces and moments,

c) the deformations implied by this distribution do not exceed the deformation

capacity of the fasteners or welds and the connected parts,

d) the assumed distribution of internal forces shall be realistic with regard to relative

stiffnesses within the joint,

e) the deformations assumed in any design model based on elastic-plastic analysis are

based on rigid body rotations and/or in-plane deformations which are physically

possible, and

f) any model used is in compliance with the evaluation of test results (see EN 1990).

Where a joint loaded in shear is subject to impact or significant vibration one of the following

jointing methods should be used:

welding

bolts with locking device

preloaded bolts

injection bolts

other types of bolt which effectively prevent movement of the connected parts

rivets

Where slip is not acceptable in a joint (because it is subject to reversal of shear load or for

any other reason), preloaded bolts in a Category B or C connection, fit bolts rivets or

welding should be used.

For wind and/or stability bracings, bolts in Category A connections may be used.

Where there is eccentricity at intersections, the joints and members should be

designed for the resulting moments and forces, except in the case of particular types of

structures where it has been demonstrated that it is not necessary.

In the case of joints of angles or tees attached by either a single line of bolts or two lines of bolts any possible eccentricity should be taken into account in accordance with set design. In-plane and out-of-plane eccentricities should be determined by considering the relative positions of the centroidal axis of the member and of the setting out line in the plane of the connection (Figure 6). For a single angle in tension connected by bolts on one leg the simplified design method given in set design, may be used.


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The effect of eccentricity on angles used as web members in compression is given in

EN 1993-1-1, Annex BB 1.2, special attention is to be treated as.

Figure 6 Setting out lines

4.2.3. Welded connections

General

Conforming to EN 1993-1-1, apply to weldable structural steels and to material

thicknesses of 4 mm and over. Also apply to joints in which the mechanical properties of

the weld metal are compatible with those of the parent metal.

For welds in thinner material reference should be made to EN 1993 part 1.3 and for welds in

structural hollow sections in material thicknesses of 2.5 mm and over guidance is

given section 7 of EN 1993.

For stud welding reference should be made to EN 1994-1-1. Further guidance on stud

welding can be found in EN ISO 14555 and EN ISO 13918.

Quality level C according to EN ISO 5817 is usually required, if not otherwise specified. The

frequency of inspection of welds should be specified in accordance with the rules in set

design.

Lamellar tearing shall be avoided. Guidance on lamellar tearing is given in EN 1993-1-10.

The specified yield strength, ultimate tensile strength, elongation at failure and minimum

Charpy V-notch energy value of the filler metal, should be equivalent to, or better than that

specified for the parent material. Generally, it is safe to use electrodes that are

overmatched related to the steel grades being used.

Global analysis

The effects of the behaviour of the joints on the distribution of internal forces and moments

within a structure, and on the overall deformations of the structure, should generally be taken

into account, but where these effects are sufficiently small they may be neglected.


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To identify whether the effects of joint behaviour on the analysis need be taken into account,

a distinction may be made between three simplified joint models as follows:

Simple, in which the joint may be assumed not to transmit bending moments

Continuous, in which the behaviour of the joint may be assumed to have no effect on

the analysis

Semi-continuous, in which the behaviour of the joint needs to be taken into account in

the analysis

The appropriate type of joint model should be determined from table 4, depending on the

classification of the joint and on the chosen method of analysis.

The design moment-rotation characteristic of a joint used in the analysis may be simplified by

adopting any appropriate curve, including a linearised approximation (e.g. bi-linear or tri-

linear), provided that the approximate curve lies wholly below the design moment-rotation

characteristic.

Table 4 Type of joint model

Method of global analysis

Classification of joint

Elastic Nominally pinned Rigid Semi – rigid

Rigid - Plastic

Nominally pinned Full - strength Partial - strength

Elastic – Plastic

Nominally pinned Rigid and full - strength

Semi – rigid and partial – strength

Semi – rigid and full – strength

Rigid and partial - strength

Type of joint model

Simple Continuous Semi - continuous

Loading actions

All types of fluctuating load acting on the component and the resulting stresses at potential

sites for static and variable loading have to be considered. Stresses or stress intensity

factors then have to be determined according to the assessment procedure applied.

Frequently, a fatigue load is a more common type of load, and it is applied several times in a

cyclic manner. Fatigue test is exclusively used to determine mechanical properties under

cyclic loading condition. As important as is the fracture toughness.

The actions originate from live loads, dead weights, snow, wind, waves, pressure,

accelerations, dynamic response, etc. Actions due to transient temperature changes should


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be considered. Improper knowledge of fatigue actions is one of the major sources of fatigue

damage.

4.2.4. Basic principles

Calculation of welded joints

Weld load capacity is affected by:

Joint geometry

Cross section effective area

Fracture resistance of used materials

Fracture resistance depends on:

Structural heterogeneity of weld zones (BM, HAZ, WM)

Biaxiality effect of the stress state

In the absence of defects, ability to weld butt load applied perpendicularly on the seam axes

is:

when fracture occurs in base metal FrMB = RrMB . Ao (4.2.1)

when fracture occurs in weld FrSUD = RrSUD . AS (4.2.2)

where Ao , AS are cross section areas, with pot defects in BM, WELD, respectively, and RrMB

. RrSUD fracture resistances of the BM, WELD, respectively /N/mm2/.

Load capacity of the material deposited when welding, with defects is expressed by relation:

FrdSUD = RrdSUD . AS = RrdSUDef( AS- Ad ) (4.2.3)

where RrdSUD is the fracture nominal resistance of the deposited material with defects,

RrdSUDef – effective fracture resistance relative to net area ( AS - Ad ), Ad - defects affected

area.

The global resistance of the weld depends on the effective fracture resistance of the weld

containing defects and a linear variation factor relative to the size of the defect:

RrdSUD = RrdSUDef.[1 - (Ad/ AS )] (4.2.4)


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Defects induce the change of the stress state by its concentration in the defect section

expressed by the concentration coefficient:

kS = RrdSUDef / RrSUD > 1 (4.2.5)

Figure 7 presents the evolution of the bearing capacity of the weld with the defect area.

Figure 7 Change of the bearing capacity of weld with defect area.

The weld must provide superior bearing capacity to the base material:

FrSUD = FrMB sau FrdSUD = FrMB (4.2.6)

In the previous Figure there are two domains:

I. where FrdSUD = RrdSUD . AS > FrMB = RrMB . Ao (4.2.7)

the bearing capacity is attributed to RrMB , and fracture produces in BM,

II. where FrdSUD = RrdSUD . AS < FrMB = RrMB . Ao (4.2.8)

the bearing capacity is attributed to RrdSUD , and fracture produces in SUD

Switching between the two areas is defined by the relation:


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RrMB . Ao = kS . RrdSUDef (AS - Ād) (4.2.9)

Where it is explained:

Ād = Ao [ (AS/ Ao) - (RrMB / kS RrSUD)] (4.2.10)

Admitting that (AS/ Ao) = 1, there results:

Ād = Ao [ 1 - (RrMB / kS RrSUD)] (4.2.11)

The previous relation is valid when the selection of the base material is made on the criterion

RrdSUD > RrMB. If this criterion refers to the yield limit, the ratio R0.2 / Rr is considered. This ratio

is statistically situated at:

0.60 for non-alloy steel base materials heat resistant alloy

0.80 for non-alloy filler materials

0.85 for alloy filler materials

Defects with round shapes (sulphurs, inclusions, cavities) respect the mentioned

considerations. Defects with great acuity, such as cracks, lack of penetration, are not

subjected to the mentioned considerations. The weld behaviour is controlled by the material

capacity to inhibit the propagation of the defect.

As regards the calculus dimensions for welds, in EC3-1-8 limits are stipulated that are also to

be found in other norms, but different limits, too. For example, for the thickness of fillet welds

the condition: 3 mm ≤ a ≤ 0.7 tmin

has to be respected and values “a” checked by preliminary

probes, in the case of deep penetration fillet welds, of partial penetration deep welds

completed with fillet welds, respectively.

For the minimum weld length, EC 3 stipulates 30 mm, but keeps the prescription: lmin

≥ 6a.

In EC 3 is provided the acceptance of fillet welds with constant thickness on their whole

length, if this can be practically accomplished, not taking into account the existence of final

craters from th end of welds.

Otherwise is maintained the requirement related to the reduction of the weld length with 2a.

In addition, the return of welds is acceptable, in the same plane, after the corner of the

overlapping parts, a return to be taken into account in calculating the length of weld, if the

thickness is the same.

When stress distribution along the weld angle is significantly influenced by the rigidity of

components or joined parts, uniformity of this distribution is taken into account by using a

reduced effective length “beff

“ and when the weld length exceeds 150 a, the weld strength is

reduced with a factor βLw

< 1.

EC 3-1-8 also provides special restrictions to use one side fillet welds and one side partial

penetration deep welds, when subjected to bending and tensile stresses. Calculation of weld


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strength is determined according to EC 3 as function of fracture tensile nominal strength

tensile of the steel used in joining fu and not as a function of its yield limit f

y.

The design resistance of a fillet weld should be determined using:

Directional method

Simplified method

Directional method

In directional method, the forces transmitted by a unit length of weld are resolved into

components parallel and transverse to the longitudinal axis of the weld and normal and

transverse to the plane of its throat.

The design throat area Aw should be taken as Aw =Σ a. leff.

The location of the design throat area should be assumed to be concentrated in the root. A

uniform distribution of stress is assumed on the throat section of the weld, leading to the

normal stresses and shear stresses (Figure 8), as follows:

• ζ⊥ - is the normal stress perpendicular to the throat

• ζ|| - is the normal stress parallel to the axis of the weld

• ч⊥ - is the shear stress (in the plane of the throat) perpendicular to the axis of the weld

• ч || - is the shear stress (in the plane of the throat) parallel to the axis of the weld.

Figure 8 Stresses on the throat section of a fillet weld

The normal stress parallel to the axis is not considered when verifying the design resistance

of the weld.

The design resistance of the fillet weld will be sufficient if the following are both satisfied:

2

5,0222 /3 MWuf and 2/ Muf (4.2.12)


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where:

- fu is the nominal ultimate tensile strength of the weaker part joined;

- β w is the appropriate correlation factor taken from table 5.

Welds between parts with different material strength grades should be designed using the

properties of the material with the lower strength grade.

Table 5 Correlation factor β w for fillet welds.

Standard and steel grade Correlation

factor βw EN 10025 EN 10210 EN 10219

S 235

S 235 W S 235 H S 235 H 0.8

S 275

S 275 N/NL

S 275 M/ML

S 275 H

S 275 NH/NLH

S 275 H

S 275 NH/NLH

S 275 MH/MLH

0.85

S 355

S 355 N/NL

S 355 M/ML

S 355 W

S 355 H

S 355 NH/NLH

S 355 H

S 355 NH/NLH

S 355 MH/MLH

0.9

S 420 N/NL

S 420 M/ML S 420 MH/MLH 1.0

S 460 N/NL

S 420 M/ML

S 420 Q/Ql/QL1

S 460 NH/NLH S 460 NH/NLH

S 460 MH/MLH 1.0

Simplified method

In the simplified method, the design resistance of a fillet weld may be assumed to be

adequate if, at every point along its length, the resultant of all the forces per unit length

transmitted by the weld satisfy the following criterion:

F.w,Ed ≤ Fw,Rd (4.2.13)


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where:

F.w,Ed is the design value of the weld force per unit length;

F.w,Rd is the design weld resistance per unit length.

Independent of the orientation of the weld throat plane to the applied force, the design

resistance per unit length Fw,Rd should be determined from:

Fw,Rd = fvw.d a (4.2.14)

where:

fvw.d is the design shear strength of the weld.

The design shear strength fvw.d of the weld should be determined from:

2

,

3/

Mw

u

dvw

ff

(4.2.15)

where:

fu and βw are defined previous.

The design resistance of a full penetration butt weld should be taken as equal to the design

resistance of the weaker of the parts connected, provided that the weld is made with a

suitable consumable which will produce all-weld tensile specimens having both a minimum

yield strength and a minimum tensile strength not less than those specified for the parent

metal.

The design resistance of a partial penetration butt weld should be determined using the

method for a deep penetration fillet weld. The throat thickness of a partial penetration butt

weld should not be greater than the depth of penetration that can be consistently achieved.

The design resistance of a T-butt joint, consisting of a pair of partial penetration butt welds

reinforced by superimposed fillet welds, may be determined as for a full penetration butt weld

if the total nominal throat thickness, exclusive of the unwelded gap, is not less than the

thickness „t” of the part forming the stem of the tee joint, provided that the unwelded gap is

not more than (t / 5) or 3 mm, whichever is less (Figure 9).

The design resistance of a T-butt joint which does not meet the requirements should be

determined using the method for a fillet weld or a deep penetration fillet weld, depending on

the amount of penetration. The throat thickness should be determined in conformity with the

provisions for both fillet welds and partial penetration butt welds.


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Figure 9 Effective penetration of T-butt welds.

The design resistance Fw,Rd of a plug weld should be taken as:

Fw,Rd = fvw.d .Aw (4.2.16)

where

fvw.d is the design shear strength of a weld,

Aw is the design throat area and should be taken as the area of the hole.

The distribution of forces in a welded connection may be calculated on the assumption of

either elastic or plastic behaviour. It is acceptable to assume a simplified load distribution

within the welds.

Residual stresses and stresses not subjected to transfer of load need not be included when

checking the resistance of a weld. This applies specifically to the normal stress parallel to the

axis of a weld.

Welded joints should be designed to have adequate deformation capacity. However, ductility

of the welds should not be relied upon.

In joints where plastic hinges may form, the welds should be designed to provide at least the

same design resistance as the weakest of the connected parts.

In other joints where deformation capacity for joint rotation is required due to the possibility of

excessive straining, the welds require sufficient strength not to rupture before general

yielding in the adjacent parent material.

If the design resistance of an intermittent weld is determined by using the total length ltot, the

weld shear force per unit length Fw,Ed should be multiplied by the factor (e + l/l) (Figure 10).


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Figure 10 Calculation of the weld forces for intermittent welds

Resistance calculation of welds

a) with full penetration

Resistance calculation of deep full penetrated welds is taken as equal with the resistance of

the weakest joined part, provided that welding is done by filler materials that will ensure in all

tensile tests, yield limit (fy) and fracture resistance (f

u) greater than or equal to the basic

material. As for deep welds, the calculation area of weld is equal with the cross section area

of the base material, as accepting the equality of the weld resistance calculation with that of

the base material, practically the weld verification is identical with that of the base material

and effectively it is not necessary any more.

b) with partial penetration

Proceed as for fillet welds with deep penetration. Thicknesses of welds with partial

penetration "a" that can effectively be determined by preliminary tests, within the certification

action of the welding technology.

c) with partial penetration completed with fillet welds

The procedure is similar with that for deep welds with full penetration provided that

requirements in correlation between characteristics, limits and geometrical conditions are

met. When the aforementioned conditions are not met, proceed as for fillet welds or deep

penetration welds.

Plug welds may be used:

To transmit shear

To prevent the buckling or separation of lapped parts, and


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To inter-connect the components of built-up members but should not be used to resist

externally applied tension.

The diameter of a circular hole or width of an elongated hole, for a plug weld should be at

least 8 mm more than the thickness of the part containing it.

The ends of elongated holes should either be semi-circular or else should have corners

which are rounded to a radius of not less than the thickness of the part containing the slot,

except for those ends which extend to the edge of the part concerned.

The thickness of a plug weld in parent material up to 16 mm thick should be equal to the

thickness of the parent material. The thickness of a plug weld in parent material over 16 mm

thick should be at least half the thickness of the parent material and not less than 16 mm.

In the case of welds with packing, the packing should be trimmed flush with the edge of the

part that is to be welded.

Where two parts connected by welding are separated by packing having a thickness less

than the leg length of weld necessary to transmit the force, the required leg length should

be increased by the thickness of the packing.

Where two parts connected by welding are separated by packing having a thickness equal to, or greater than, the leg length of weld necessary to transmit the force, each of the parts should be connected to the packing by a weld capable of transmitting the design force.

The effective length of a fillet weld should be taken as the length over which the fillet is

full-size. This may be taken as the overall length of the weld reduced by twice the

effective throat thickness “a”. Provided that the weld is full size throughout its length

including starts and terminations, no reduction in effective length need be made for either

the start or the termination of the weld. A fillet weld with an effective length less than 30

mm or less than 6 times its throat thickness, whichever is larger, should not be designed

to carry load.

The effective throat thickness, a, of a fillet weld should be taken as the height of the

largest triangle (with equal or unequal legs) that can be inscribed within the fusion

faces and the weld surface, measured perpendicular to the outer side of this

triangle(Figure 11). The effective throat thickness of a fillet weld should not be less than 3

mm.

Figure 11 Throat thickness of a fillet weld.


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In determining the design resistance of a deep penetration fillet weld, account may be

taken of its additional throat thickness (Figure 12), provided that preliminary tests show

that the required penetration can consistently be achieved.

Figure 12 Throat thickness of a deep penetration fillet weld.

For solid bars the design throat thickness of flare groove welds, when fitted flush to the

surface of the solid section of the bars, is defined in Figure 13. The definition of the design

throat thickness of flare groove welds in rectangular hollow sections.

Figure 13 Effective throat thickness of flare groove welds in solid sections.

Where a transverse plate (or beam flange) is welded to a supporting unstiffened flange of an

I, H or other section, Figure 14, and provided that the design condition given is met, the

applied force perpendicular to the unstiffened flange should not exceed any of the relevant

design resistances as follows:

That of the web of the supporting member of I or H sections ,

Those for a transverse plate on a RHS member,


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That of the supporting flange as given by formulas, calculated assuming the applied

force is concentrated over an effective width, beff, of the flange as given as relevant.

Figure 14 Effective width of an unstiffened T – joint

For an unstiffened I or H section the effective width beff should be obtained from:

beff = tw s k.tf (4.2.17)

where:

k = (tf/tp ) ( fy, f/f y,p ) for k ≤1 (4.2.18)

f y,f is the yield strength of the flange of the I or H section;

f y,p is the yield strength of the plate welded to the I or H section.

The dimension s should be obtained from:

– for a rolled I or H section: s= r

– for a welded I or H section: s= √2 . a

In lap joints the design resistance of a fillet weld should be reduced by multiplying it by a

reduction factor βLw to allow for the effects of non-uniform distribution of stress along its

length. The provisions do not apply when the stress distribution along the weld corresponds

to the stress distribution in the adjacent base metal, as, for example, in the case of a weld

connecting the flange and the web of a plate girder.

Generally in lap joints longer than 150a the reduction factor βLw should be taken as βLw.1

given by:


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βLw.1 = 1,2 Lj /(150a) but βLw.1 ≤1 (4.2.19)

where:

L j is the overall length of the lap in the direction of the force transfer.

For fillet welds longer than 1,7 metres connecting transverse stiffeners in plated members,

the reduction factor βLw may be taken as βLw.2 given by:

β Lw.2 = 1,1 βw /17 but 0,6≤ βLw.2 ≤1 (4.2.20)

where:

β w is the length of the weld (in metres).

Local eccentricity should be avoided whenever it is possible.

Local eccentricity (relative to the line of action of the force to be resisted) should be taken into

account in the following cases:

- where a bending moment transmitted about the longitudinal axis of the weld produces

tension at the root of the weld (Figure 15 a),

- where a tensile force transmitted perpendicular to the longitudinal axis of the weld

produces a bending moment, resulting in a tension force at the root of the weld (Figure 15 b).

Local eccentricity need not be taken into account if a weld is used as part of a weld group around the perimeter of a structural hollow section.

a) Bending moment produces tension at the b) Tensile force produces tension at the root of the weld root of the weld

Figure 15 Local eccentricity

Local eccentricity need not be taken into account if a weld is used as part of a weld group around the perimeter of a structural hollow section.

In angles connected by one leg, the eccentricity of welded lap joint end connections may

be allowed for by adopting an effective cross-sectional area and then treating the

member as concentrically loaded.


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For an equal-leg angle, or an unequal-leg angle connected by its larger leg, the effective area may be taken as equal to the gross area.

For an unequal-leg angle connected by its smaller leg, the effective area should be taken as equal to the gross cross-sectional area of an equivalent equal-leg angle of leg size equal to that of the smaller leg, when determining the design resistance of the cross-section, see EN 1993-1-1. When determining the design buckling resistance of a compression member, the actual gross cross-sectional area should be used.

In angles connected by one leg, the eccentricity of welded lap joint end connections may be

allowed for by adopting an effective cross-sectional area and then treating the member

as concentrically loaded.

Welding may be carried out within a length 5t either side of a cold-formed zone ( table

6), provided that one of the following conditions is fulfilled:

– the cold-formed zones are normalized after cold-forming but before welding;

– the r/t -ratio satisfy the relevant value obtained from table 6.

Table 6 Conditions for welding cold-formed zone and adiacent material

r/t

Strain due

to cold forming

(%)

Maximum thickness (mm)

Generally Fully killed

Aluminium

– killed steel (Al

≥ 0,02%)

Predominan

tly static loading

Where

fatigue

predominates

≥ 25

≥ 10

≥ 3.0

≥ 2.0

≥ 1.5

≥ 1.0

≥ 2

≥ 5

≥ 14

≥ 20

≥ 25

≥ 33

any

any

24

12

8

4

any

16

12

10

8

4

any

any

24

12

10

6

B. Calculating resistance of welds in filled holes

Calculating resistance of a filled hole is taken equal to:

Fw,Rd = fvw,d · Aw (4.2.21)


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where fvw.d

– is shear calculating resistance of the weld,

Aw – hole area where the weld is performed (Circular or elongated).

In conclusion, calculation of welds is made reducing the effect of loading in relation to the

centre weight of the weld area calculation. In simple loading, this leads to one type of stress

(ζ or η) in this area, stresses that must not exceed the calculating resistance of welds

In the case of fillet welds it is acceptable to rebate the calculating area of weld in the cathetes

plan and carrying out the verification in relation to the rebated area. In the case of compound

loading an equivalent stress is determined on the bases of the Huber – Mises concept

Rech 22 3 (4.2.22)

where α has the value 1,1, and R is the calculating resistance of the base material.

As it results from the EC 3 norm, analytical relations are expressly provided to check the

weld strength only for fillet welds and welds in filled holes and two methods to check fillet

welds.

4.2.5. Types of stress raisers and notch effects Different types of stress raisers and notch effects lead to the calculation of different types of

stress. The choice of stress depends on the fatigue assessment procedure used (table 7,

Figure 16, 17).

Table 7 Stress raisers and notch effects

Type Stress raisers Stress determined Assessment procedure

A General analysis of sectional forces

using general theories e.g. beam

theory, no stress risers considered

Gross average

stress from

sectional forces

not applicable for fatigue

analysis, only component

testing

B A + macrogeometrical effects due to

the design of the component, but

excluding stress risers due to the

welded joint itself.

Range of nominal

stress (also modi-

fied or local nomi-

nal stress)

Nominal stress approach


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C A + B + structural discontinuities due

to the structural detail of the welded

joint, but excluding the notch effect of

the weld toe transition

Range of structural

Structural Stress

(hot spot stress)

Structural Stress (hot spot

stress) approach

D A + B + C + notch stress

concentration due to the weld bead

notches a) actual notch stress b)

effective notch stress

Range of elastic

notch stress (total

stress)

a) Fracture mechanics

approach b) effective

notch stress approach

Figure 16 Modified or local nominal stress Figure 17 Notch stress and structural stress

Besides the usual corner welds, the thickness "a" is considered equal to the height of the

triangle in cross section of weld recordable, lowered from its roots on the outer side, EC May

3 provides deep penetration welds corner with a thickness depends on technology and

equipment required for execution and check the preliminary tests (table 8).


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Table 8 Characteristics, limitations and conditions related to the type of welding.

Joint

type Weld type

Characteristics, limitations

and conditions

0 1 2

in T

, in

an

gle

FILLET WELDS

1. continuous 60° ≤ α ≤ 120°

α < 60° are considered to

be deep welds with partial

penetration

α < 120° their strength is

determined by tests

The return of welds is

imposed to the ends with 2a

and notation on drawings

SIS ll returns (for a

= constant) min

Sl min (30 mm

or 6a); 150max Sl a

For > 150a weld strength

is reduced with βLW

3 mm ≤ a ≤ 0.7 · tmin

effw laA

in T

, in

an

gle

2. interrupted

Not to be used in

corrosive environments.

At the ends of parts both

side welds are used.

max. Lwe ≥ 0.75b and

0.75b1

min. L1 ≤ 16t and 16t1 or

200 mm

min. L2 ≤ 12t and 16t1 and

0.25b sau 200 mm

Standard EN 1993, part 1-8, covers the design of fillet welds, fillet welds all round, butt

welds, plug welds and flare groove welds. Butt welds may be either full penetration butt

welds or partial penetration butt welds.

Both fillet welds all round and plug welds may be either in circular holes or in elongated

holes.


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The most common types of joints and welds are illustrated in EN 12345.

Fillet welds may be used for connecting parts where the fusion faces form an angle of

between 60° and 120°.

Angles smaller than 60° are also permitted. However, in such cases the weld should be

considered to be a partial penetration butt weld.

For angles greater than 120° the resistance of fillet welds should be determined by testing in

accordance with EN 1990 Annex D: Design by testing.

Fillet welds finishing at the ends or sides of parts should be returned continuously, full size,

around the corner for a distance of at least twice the leg length of the weld, unless access or

the configuration of the joint renders this impracticable. In the case of intermittent welds this

rule applies only to the last intermittent fillet weld at corners.

End returns should be indicated on the drawings.

Intermittent fillet welds shall not be used in corrosive conditions.

In an intermittent fillet weld, the gaps (L1 or L2) between the ends of each length of weld Lw

should fulfil the requirement given in Figure 18. In an intermittent fillet weld, the gap (L1 or

L2) should be taken as the smaller of the distances between the ends of the welds on

opposite sides and the distance between the ends of the welds on the same side. Correlated

with previous Figure, to remember:


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Figure 18 Geometric elements of intermittent fillet weld

The larger of Lwe ≥ 0.75 b and 0.75 b1

For build-up members in tension:

The smallest of L1 ≤ 16 t and 16 t1 and 200 mm

For build-up members in compression or shear:

The smallest of L2 ≤ 12 t and 12 t1 and 0.25 b and 200 mm

In any run of intermittent fillet weld there should always be a length of weld at each end of

the part connected.


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In a built-up member where plates are connected by means of intermittent fillet welds, a

continuous fillet weld should be provided on each side of the plate for a length at each end

equal to at least three-quarters of the width of the narrower plate concerned (Figure 18).

Fillet welds all round, comprising fillet welds in circular or elongated holes, may be used only

to transmit shear or to prevent the buckling or separation of lapped parts. The diameter of a

circular hole, or width of an elongated hole, for a fillet weld all round should not be less than

four times the thickness of the part containing it. The ends of elongated holes should be

semi-circular, except for those ends which extend to the edge of the part concerned.

The centre to centre spacing of fillet welds all round should not exceed the value necessary

to prevent local buckling, show in table 9.

A full penetration butt weld is defined as a weld that has complete penetration and fusion of

weld and parent metal throughout the thickness of the joint.

A partial penetration butt weld is defined as a weld that has joint penetration which is less

than the full thickness of the parent material.

Intermittent butt welds should not be used.

Table 9 The centre to centre spacing of fillet welds all round

Distances

and spacings,

see Figure 3.1

Minimu

m

Maximum1) 2) 3)

Structures made from steels

conforming to EN 10025 except

steels conforming to EN 10025-5

Structures

made from steels

conforming to EN

10025

Steel

exposed to the

weather or other

corrosive

influences

Steel not

exposed to the

weather or other

corrosive

influences

Steel used

unprotected

End

distance e1

1.2 do 4t+40 mm The larger of

8t or 125 mm

Edge

distance e2 1.2 do 4t+40 mm

The larger of

8t or 125 mm

Distance e3

in slotted holes

1.5 do 4)

Distance e4

in slotted holes

1.5 do 4)

Spacing p1 2.2 do The smaller

of 14t or 200 mm

The smaller

of 14t or 200 mm

The smaller

of 14tmin or 175

mm


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Spacing p1,0 The smaller

of 14t or 200 mm

Spacing p1,i The smaller

of 28t or 400 mm

Spacing p2 5) 2.4 do

The smaller

of 14t or 200 mm

The smaller

of 14t or 200 mm

The smaller

of 14tmin or 175

mm

1) Maximum values for spacing, edge and end distances are unlimited, except in the following

cases:

– for compression members in order to avoid local buckling and to prevent corrosion in

exposed members and;

– for exposed tension members to prevent corrosion.

2) The local buckling resistance of the plate in compression between the fasteners should be

calculated according to EN 1993-1-1 using 0.6 pi as buckling length. Local buckling between

the fasteners need not to be checked if p1/t is smaller than 9ε. The edge distance should not

exceed the local buckling requirements for an outstand element in the compression

members; see EN 1993-1-1. The end distance is not affected by this requirement.

3) t is the thickness of the thinner outer connected part.

4) The dimensional limits for slotted holes are given in 2.8 Reference Standards: Group 7.

5) For staggered rows of fasteners a minimum line spacing of p2 = 1.2d0 may be used,

provided that the minimum distance, L, between any two fasteners is greater than 2.4d0,


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4.2.6. Determination of stress and stress intensity factors

Definition of Stress Components

The stress distribution over the plate thickness is non-linear in the vicinity of notches. The

stress components of the notch stress ζln

are (Figure 19):

ζmem

membrane stress,

ζben

shell bending stress,

ζnlp

non-linear stress peak

Figure 19 The stress distribution over the plate thickness.

If a refined stress analysis method is used, which gives a non-linear stress distribution, the

stress components can be separated by the following method:

the membrane stress ζmem

is equal to the average stress calculated through the

thickness of the plate, and it is constant through the thickness,

the shell bending stress ζben

is linearly distributed through the thickness of the plate,

and tt is found by drawing a straight line through the point “0” where the membrane

stress intersects the mid-plane of the plate. The gradient of the shell bending stress is

chosen such that the remaining non-linearly distributed component is in equilibrium.

the non-linear stress peak ζnlp

is the remaining component of the stress.

The stress components can be separated analytically for a given stress distribution ζ (x) for x=0 at surface to x=t at through thickness.

Nominal stress

Nominal stress is the stress calculated in the sectional area under consideration,

disregarding the local stress raising effects of the welded joint, but including the stress raising

effects of the macrogeometric shape of the component in the vicinity of the joint, such as e.g.

large cut outs (Figure 20). Overall elastic behaviour is assumed.


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Figure 20 Nominal stress in a beam-like component

The nominal stress may vary over the section under consideration. E.g. at a beam-like

component, the modified (also local) nominal stress and the variation over the section can be

calculated using simple beam theory. Here, the effect of a welded on attachment is ignored.

The effects of macrogeometric features of the component as well as stress fields in the

vicinity of concentrated loads must be included in the nominal stress. Consequently,

macrogeometric effects may cause a significant redistribution of the membrane stresses

across the section. Similar effects occur in the vicinity of concentrated loads or reaction

forces. Significant shell bending stress may also be generated, as in curling of a flange, or

distortion of a box section (Figure 21, 22).

The secondary bending stress caused by axial or angular misalignment needs to be

considered if the misalignment exceeds the amount which is already covered by fatigue

resistance S-N curves for the structural detail (Figure 23). This is done by the application of

an additional stress raising factor km,eff

.

Figure 21 Examples of macrogeometric effects


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Figure 22 Modified (local) nominal stress near concentrated loads

Figure 23 Axial and angular misalignment

Intentional misalignment (e.g. allowable misalignment specified in the design stage) is

considered when assessing the fatigue actions (stress) by multiplying by the factor. If it is

non-intentional, it is regarded as a weld imperfection which affects the fatigue resistance and

has to be considered by dividing the fatigue resistance (stress) by the factor.

Calculation of nominal stress

In simple components the nominal stress can be determined using elementary theories of

structural mechanics based on linear-elastic behaviour. In other cases, finite element method

(FEM) modelling may be used. This is primarily the case in:

a) complicated statically over-determined (hyperstatic) structures,

b) structural components incorporating macrogeometric discontinuities, for which no

analytical solutions are available.

Using FEM, meshing can be simple and coarse. Care must be taken to ensure that all stress

raising effects of the structural detail of the welded joint are excluded when calculating the

modified (local) nominal stress.

If nominal stresses are calculated in fillet welds by a coarse finite element mesh, nodal forces

should be used in a section through the weld instead of element stresses in order to avoid

stress underestimation.


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Measurement of nominal stress

The fatigue resistance S-N curves of classified structural details are based on nominal

stress, disregarding the stress concentrations due to the welded joint. Therefore the

measured nominal stress must exclude the stress or strain concentration due to the

corresponding discontinuity in the structural component. Thus, strain gauges must be placed

outside of the stress concentration field of the welded joint. In practice, it may be necessary

firstly to evaluate the extension and the stress gradient of the field of stress concentration

due to the welded joint. For further measurements, simple strain gauge application outside

this field is sufficient.

4.2.7 Structural hot spot stress

General

The structural or geometric stress Fhs

at the hot spot includes all stress raising effects of a

structural detail excluding all stress concentrations due to the local weld profile itself. So, the

non-linear peak stress Fnlp

caused by the local notch, i.e. the weld toe, is excluded from the

structural stress. The structural stress is dependent on the global dimensional and loading

parameters of the component in the vicinity of the joint. It is determined on the surface at the

hot spot of the component which is to be assessed. Structural hot spot stresses Fhs

are

generally defined at plate, shell and tubular structures. Figure 24 shows examples of

structural discontinuities and details together with the structural stress distribution.

Figure 24 Structural details and structural stress

The structural hot spot stress approach is recommended for welded joints where there is no

clearly defined nominal stress due to complicated geometric effects, and where the structural

discontinuity is not comparable to a classified structural detail. Definition of structural hot spot

stress show in Figure 25.


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Figure 25 Definition of structural hot spot stress

The structural hot spot stress can be determined using reference points and extrapolation to

the weld toe at the considered hot spot. The method as defined here is limited to the

assessment of the weld toe, i.e. cases “a” to “e” in Figure 4.2.22. It is not applicable in cases

where crack will grow from the weld root and propagate through the weld metal, i.e. cases “f”

to “I” in Figure 26.

Figure 26 Various locations of crack propagation in welded joints

The method of structural hot spot stress may be extended to the assessment of spots of the

welded joint susceptible to fatigue cracking other than on plate surface, e.g. on a fillet weld

root. In this case, structural hot spot stress on surface is used as an indication and estimation

of the stress for the spot in consideration. The S-N curves or structural hot spot stress

concentration factors used for verification in this case depend largely on geometric and

dimensional parameters and are only valid within the range of these parameters.


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In case of a biaxial stress state at the plate surface, it is recommended to use the principal

stress which is approximately in line with the perpendicular to the weld toe, i.e. within a

deviation of ±60º (Figure 27).

Figure 27 Biaxial stress at weld toe

The other principal stress may be analysed, if necessary, using the fatigue class for parallel

welds in the nominal stress approach.

Besides the definitions of structural hot spot stress as given above, two types of hot spots

have to be distinguished according to their location on the plate and their orientation to the

weld toe (table 10).

Determination of structural hot spot stress

Determination of structural hot spot stress can be done either by measurement or by

calculation. Here the non-linear peak stress is eliminated by linearization of the stress

through the plate thickness or by extrapolation of the stress at the surface to the weld toe.

The following considerations focus on extrapolation procedures of the surface stress, which

are nearly the same in measurement and calculation.

Firstly the stresses at the reference points, i.e. extrapolation points, have to be determined;

secondly the structural hot spot stress has to be determined by extrapolation to the weld toe.

Table 10 Types of hot spots

Type Description Determination

a Structural hot spot stress transverse to

weld toe on plate surface

Special FEA procedure or

measurement and extrapolation

b Structural hot spot stress transverse to

weld toe at plate edge

Special FEA procedure or

measurement and extrapolation


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The structural hot spot stress may be determined using two or three stress or strain values at

particular reference points apart from the weld toe in direction of stress. The closest position

to the weld toe must be chosen to avoid any influence of the notch due to the weld itself

(which leads to a non-linear stress peak). This is practically the case at a distance of 0.4 t (t =

plate thickness) from the weld toe. The structural hot spot stress at the weld toe is then

obtained by extrapolation. Identification of the critical points (hot spots) can be made by:

a) measuring several different points,

b) analysing the results of a prior FEM analysis,

c) experience of existing components, which failed.

Calculation of structural hot spot Stress

In general, analysis of structural discontinuities and details to obtain the structural hot spot

stress is not possible using analytical methods. Parametric formulae are rarely available.

Thus, finite element (FEM) analysis is mostly applied.

Usually, structural hot spot stress is calculated on the basis of an idealized, perfectly aligned

welded joint. Consequently, any possible misalignment has to be taken explicitly into

consideration by the FEA model or by an appropriate stress magnification factor km. This

applies particularly to butt welds, cruciform joints and one-sided transverse fillet welds at

free, unsupported plates (Figure 28).

Figure 28 Types of hot spots

The extent of the finite element model has to be chosen such that constraining boundary

effects of the structural detail analysed are comparable to the actual structure.

Models with thin plate or shell elements or alternatively with solid elements may be used. It

should be noted that on the one hand the arrangement and the type of the elements have to

allow for steep stress gradients as well as for the formation of plate bending, and on the

other hand, only the linear stress distribution in the plate thickness direction needs to be

evaluated with respect to the definition of the structural hot spot stress. The stresses should


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be determined at the specified reference points.

For FEM analysis, sufficient expertise of the analyst is required. Guidance is given in [2-3]. In

the following, only some roughure (Figure 29.a), the elements have to be arranged in the

mid-plane of the structural components. 8-noded elements are recommended particularly in

case of steep stress gradients. In simplified models, the welds are not modelled, except for

cases where the results are affected by local bending, e. g. due to an offset between plates

or due to the small distance between adjacent welds. Here, the welds may be included by

vertical or inclined plate elements having appropriate stiffness or by introducing constraint

equations or rigid links to couple node displacements.

a) b)

Figure 29 Typical meshes and stress evaluation path for a welded detail

An alternative particularly for complex cases is recommended using prismatic solid elements

which have a displacement function allowing steep stress gradients as well as plate bending

with linear stress distribution in the plate thickness direction. This is offered, e. g., by

isoparametric 20 node elements with mid-side nodes at the edges, which allow only one

element to be arranged in the plate thickness direction due to the quadratic displacement

function and the linear stress distribution. At a reduced integration, the linear part of the

stresses can be directly evaluated. Modelling of welds is generally recommended (Figure

29.b).

The element lengths are determined by the reference points for the subsequent

extrapolation. In order to avoid an influence of the stress singularity, the stress closest to the

hot spot is usually evaluated at the first or second nodal point. Therefore, the length of the

element at the hot spot has to correspond at least to its distance from the first reference

point. Coarser meshes are possible with higher-order elements and fixed lengths, as further

explained below.

Appropriate element widths are important particularly in cases with steep stress gradients.


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The width of the solid element or the two shell elements in front of the attachment should not

exceed the attachment width “w”, i. e. the attachment thickness plus two weld leg lengths.

Usually, the structural hot spot stress components are evaluated on the plate surface or

edge. Typical extrapolation paths are shown by arrows in Figure 25. If the weld is not

modelled, it is recommended to extrapolate the stress to the structural intersection point in

order to avoid stress underestimation due to the missing stiffness of the weld.

Type “a” hot spots

The structural hot spot stress ζhs

is determined using the reference points and extrapolation

equations as given below (Figure 30).

Figure 30 Reference points at different types of meshing

1. Fine mesh with element length not more than 0.4 t at the hot spot: Evaluation of nodal

stresses at two reference points 0.4 t and 1.0 t, and linear extrapolation.

2. Fine mesh as defined above: Evaluation of nodal stresses at three reference points

0.4 t, 0.9 t and 1.4 t, and quadratic extrapolation. This method is recommended in

cases with pronounced non-linear structural stress increase to the hot spot.

3. Coarse mesh with higher-order elements having lengths equal to plate thickness at

the hot spot: Evaluation of stresses at mid-side points or surface centres respectively,

i.e. at two reference points 0.5 t and 1.5 t, and linear extrapolation.

tths 0,14,0 67,067,1 (4.2.23)

ttths 4,19,04,0 72,024,252,2 (4.2.24)


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tths 5,15,0 50.050,1 (4.2.25)

Type “b” hot spots

The stress distribution is not dependent of plate thickness. So, the reference points are given

at absolute distances from the weld toe or from the weld end if the weld does not continue

around the end of the attached plate.

4. Fine mesh with element length of not more than 4 mm at the hot spot: Evaluation of

nodal stresses at three reference points 4 mm, 8 mm and 12 mm and quadratic

extrapolation (eq. 4).

5. Coarse mesh with higher-order elements having length of 10 mm at the hot spot:

Evaluation of stresses at the mid-side points of the first two elements and linear extra-

polation (eq. 5).

mmmmmmhs 1284 33 (4.2.26)

mmmmhs 155 5,05,1 (4.2.27)

Correlation between relatively coase and fine models, to type of model and weld toe it is in

table 11.

Table 11 Correlation between relatively coase and fine models, to type of model and weld toe

Type of model

and weld toe

Relatively coase models Relatively fine models

Type a Type b Type a Type b

Element

size

Shells t x t max t x

w/2*)

10 x 10 mm ≤0.4 t x t or

≤0.4 t x w/2

≤ 4 x 4 mm

Solids t x t max t x w

10 x 10 mm ≤0.4 t x t or

≤0.4 t x w/2

≤ 4 x 4 mm

Extrapo-

lation

points

Shells 0.5 t and 1.5 t

mid-side

points**)

5 and 15 mm

mid-side points

0.4 t and 1.0 t

nodal points

4. 8 and 12 mm

nodal points

Solids 0.5 and 1.5 t

surface centre

5 and 15 mm

surface centre

0.4 t and 1.0 t

nodal points

4. 8 and 12 mm

nodal points

*)

w = longitudinal attachment thickness + 2 weld leg lengths **)

surface centre at


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transversal welds, if the weld below the plate is not modelled (see Figure 28.a).

Measurement of structural hot spot stress

The recommended placement and number of strain gauges is dependent of the presence of

higher shell bending stresses, the wall thickness and the type of structural stress (Figure 31).

Figure 31 Examples of strain gauges in plate structures

The centre point of the first gauge should be placed at a distance of 0.4 t from the weld toe.

The gauge length should not exceed 0.2 t. If this is not possible due to a small plate

thickness, the leading edge of the gauge should be placed at a distance 0.3 t from the weld

toe. The following extrapolation procedure and number of gauges are recommended:

Type “a” hot spots

a) Two gauges at reference points 0.4 t and 1.0 t and linear extrapolation (eq. 6).

tths 0,14,0 67.067,1 (4.2.28)

b) Three gauges at reference points 0.4 t, 0.9 t and 1.4 t, and quadratic extrapolation in

cases of pronounced non-linear structural stress increase to the hot spot (eq. 7).

tttht 4,19,04,0 72,024,252,2 (4.2.29)


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Often multi-grid strip gauges are used with fixed distances between the gauges. Then the

gauges may not be located as recommended above. Then it is recommended to use e.g.

four gauges and fit a curve through the results.

Type “b” hot spots

Strain gauges are attached at the plate edge at 4, 8 and 12 mm distant from the weld toe.

The hot spot strain is determined by quadratic extrapolation to the weld toe (eq. 8).

mmmmmmhs 1284 33 (4.2.30)

Tubular joints

For tubular joints, there exist recommendations which allow the use of linear extrapolation

using two strain gauges. Here, the measurement of simple uniaxial stress is sufficient.

Determination of stress

If the stress state is close to uniaxial, the structural hot spot stress is obtained approximately

from eqn. (9).

hshs E (4.2.31)

At biaxial stress states, the actual stress may be up to 10% higher than obtained from eqn. (3). In this case, use of rosette strain gauges is recommended. If FEA results are available giving the ratio between longitudinal and transverse strains εy/εx , the structural hot spot stress ζ

hs can then be resolved assuming that this principal stress is about perpenticular to

the weld toe.

21

1

v

v

E x

y

xhs

(4.2.32)

Instead of absolute strains, strain ranges ∆ε = εmax − εmin are usually measured and

substituted in the above equations, producing the range of structural hot spot stress ∆ζhs.

Structural hot spot stress concentration factors and parametric formulae

For many joints between circular section tubes parametric formulae have been established

for the stress concentration factor khs

in terms of structural structural stress at the critical

points (hot spots). Hence the structural hot spot stress ζhs

becomes:

nomhshs k (4.2.33)


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where ζnom

is the nominal axial membrane stress in the braces, calculated by elementary

stress analysis.

4.2.8 Effective notch stress Effective notch stress is the total stress at the root of a notch, obtained assuming linear-

elastic material behaviour. To take account of the statistical nature and scatter of weld shape

parameters, as well as of the non-linear material behaviour at the notch root, the real weld

contour is replaced by an effective one. For structural steels and aluminium an effective

notch root radius of r = 1 mm has been verified to give consistent results.

The method is restricted to welded joints which are expected to fail from the weld toe or weld

root. Other causes of fatigue, e.g. from surface roughness or embedded defects, are not

covered. Also it is also not applicable where considerable stress components parallel to the

weld or parallel to the root gap exist.

The method is also restricted to assessment of naturally formed weld toes and roots. At

machined or ground welds, toes or roots shall be assessed using the notch stress and the

fatigue resistance value of a butt weld groud flush to plate.

The method is well suited to the comparison of alternative weld geometries. Unless

otherwise specified, flank angles of 30° for butt welds and 45° for fillet welds are suggested.

In cases where a mean geometrical notch root radius can be defined, e.g. after certain post

weld improvement procedures, this geometrical radius plus 1 mm may be used in the

effective notch stress analysis. The method is limited to thicknesses t ≥ 5 mm. For smaller

wall thicknesses, the method has not yet been verified.

Calculation of effective notch stress

Effective notch stresses or stress concentration factors can be calculated by parametric

formulae, taken from diagrams or calculated from finite element or boundary element

models. The effective notch radius is introduced such that the tip of the radius touches the

root of the real notch, e.g. the end of an unwelded root gap (Figure 32).

Figure 32 Effective notch stress concentration factors


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Possible misalignment has to be considered in the calculations.

Because the effective notch radius is an idealization, the effective notch stress cannot be measured directly in the welded component. In contrast, the simple definition of the effective notch can be used for photo-elastic stress measurements in resin models.

Stress intensity factors

Fracture mechanics assumes the existence of an initial crack ai. It can be used to predict the

growth of the crack to a final size af. Since for welds in structural metals, crack initiation

occupies only a small portion of the life, this method is suitable for assessment of fatigue life,

inspection intervals, crack-like weld imperfections and the effect of variable amplitude

loading.

The parameter which describes the fatigue action at a crack tip in terms of crack propagation

is the stress intensity factor (SIF) “K”.

Fracture mechanics calculations generally have to be based on total stress at the notch root, e.g. at the weld toe. For a variety of welded structural details, correction functions for the local notch effect and the nonlinear stress peak of the structural detail have been established. Using these correction functions, fracture mechanics analysis can be based on structural hot spot stress or even on nominal stress. The correction function formulae may be based on different stress types. The correction function and the stress type have to correspond.

Stress intensity factor determination methods are usually based on FEM analyses. They may be directly calculated as described in the literature, or indirectly using the weight function approach.

Calculation of stress intensity factors by parametric formulae

First, the local nominal stress or the structural Structural hot spot stress at the location of the

crack has to be determined, assuming that no crack is present. The stress should be

separated into membrane and shell bending stresses. The stress intensity factor (SIF) “K”

results as a superposition of the effects of both stress components. The effect of the

remaining stress raising discontinuity or notch (non-linear peak stress) has to be covered by

additional factors “M”k.

benkbenbenmemkmemmem MYMYaK ,, (4.2.34)

where

ζmem - membrane stress

ζben -shell bending stress,


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Ymem - correction function for membrane stress intensity factor,

Yben - correction function for shell bending stress intensity factor,

Mk, mem - correction for non-linear stress peak in terms of membrane action,

Mk, ben - correction for non-linear stress peak in terms of shell bending.

The correction functions Ymem

and Yben

, the formulae for stress intensity factors, Mk-factors

can be found in the literature.

4.3 Design of welded joints with predominantly fatigue loading

Objective

The students will understand how the fatigue behaviour of welded joints and be able to

perform relevant fatigue life calculations.

Scope

Fatigue of welded joints:

o stress concentrations


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o residual stresses

o initial defects

Constant and variable amplitude loading

Cumulative damage FAT class

Overview of fatigue calculation methods in a relevant design guidance document, e.g.,

IIW Doc. XIII-1965-03/XV-1127-03 “Recommendations for fatigue design of welded

joints and components”

Expected result at comprehensive level:

Explain the assumptions involved in the design of predominantly static loaded joints.

Identify relevant stress values from a type stress-time history for a structural

component. Calculate the design strength of end welds based on weld stress

components.

Calculate the design strength of side welds based on weld stress components.

Calculate the strength reduction factor for long side welds or transverse stiffeners.

Calculate the needed throat thickness for a full strength primary load carrying weld.

Calculate the throat thickness for a binding welded joint.

4.3.1 Basic principles Fatigue resistance is usually derived from constant or variable amplitude tests. The fatigue

resistance data given here are based on published results from constant amplitude tests.

The fatigue resistance data must be expressed in terms of the same stress as that controlled

or determined the generation of those data.

In fatigue assessment, the fatigue actions and the fatigue resistance are related by means of

an appropriate assessment procedure. It must be ensured that all three elements (actions,

resistance and assessment procedure) correspond. Three procedures may be distinguished:

a) Procedures based on S-N curves, such as nominal stress approach structural hot

spot stress approach effective notch stress approach.

b) Procedures based on crack propagation considerations.

c) Direct experimental approach by fatigue testing of components or entire structures.

If normal and shear stress occur simultaneously, their combined effect shall be considered.

Three cases may be distinguished:

a) If the equivalent nominal shear stress range is less than 15% of the equivalent normal

stress range or if the damage sum due to shear stress range is lower than 10% of

that due to normal stress range, the effect of shear stress may be neglected.

b) If the normal and shear stress vary simultaneously in phase, or if the plane of


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maximum principal stress is not changed significantly, the maximum principal stress

range may be used.

c) If normal and shear stress vary independently out of phase, in damage calculation the

damage sums shall be calculated separately and finally added. The usage of 1/2 of

the calculated life cycles is recommended.

Fracture mechanics crack propagation calculations should be based on maximum principal stress range.

The S-N curve represents a material characteristic, experimentally considered for any

loading type or body. It is known as the material base curve. ConFigureation must be plain to

assure homogenous tensile / compression mechanical loading. Results obtained on bodies

subjected to cyclic bending, or bodies with geometric concentrators, reflect the effect of

certain influence factors.

Figure 33 presents two curves:

- curve „a” defines the asimptotic level of stress SR under which fracture does not happen

any more indifferently of the loading cycles,

- curve „b” defines a level of the loading at which the material can failure for a defined

number of cycles.

Distinct zones on the curve S-N (Figure 34): quasistatic fracture, oligocyclic, and polycyclic

fatigue, respectively

Figure 33 S-N Diagram Figure 34 Specific zones of the S-N diagram

Very high stress leads to quasistatic fracture. The oligocyclic fatigue is localized in the range

102-105 cycles, and for a greater number of cycles the polycyclic fatigue works.

Stress is a time periodical function. Assembly of stress values during a period is called cycle.

ConFigureation of cycles is presented in Figure 1.

Parameters:

- stress (S): maximum / minimum, (Smax / Smin) (Figure 35),

- period (T),


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Figure 35 Loading parameters

- average stress Sm = (Smax + Smin) / 2 (4.3.1)

- cycle amplitude Sa = (Smax – Smin) / 2 (4.3.2)

- cycle asymmetry coefficient r = Smin / Smax (4.3.3)

- a cycle characteristic = Sv / | Sm | (4.3.4)

So, Smax = Smed + Ra/2, respectively Smin = Sm- Ra/2. (4.3.5)

Variation of cycle asymmetry coefficient with amplitude (Sa) and average stress (Smed) is

represented in Figure 36.

Figure 36 Correlation of the cycle asymmetry coefficient, amplitude and average stress

Loading complex aleatory spectra reflect, in the nearest way the in service situations.

Mathematical presentation is possible by the Fourier series decomposition, as it is presented

later in this material.

In conventional endurance testing, there are different definitions of failure. In general, small

specimens are tested to complete rupture, while in large components the observation of a

through wall crack is taken as a failure criterion. In fracture mechanics crack propagation

testing, the crack growth rate data are derived from crack propagation monitoring.

All fatigue resistance data are given as characteristic values, which are assumed to have a

survival probability of at least 95%, calculated from a mean value of a two-sided 75%

confidence level, unless otherwise stated.


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The (nominal) stress range should be within the limits of the elastic properties of the material.

The range of the design values of the stress range shall not exceed 1.5 fy for nominal normal

stresses or 1.5 fy / √3 for nominal shear stresses. fy – actual or specified yield strength of the

material.

Evaluation to variable loading supposes to define form the beginning the using requirements

of characteristics for components, subassemblies, products. Requirement is generated by

the balance technical efficiency and involved costs. The establishment of fatigue

characteristics is influenced by the dispersion of primary results: structural macro/micron

homogenities of used materials, surface state, effect of stress constructive concentrators,

change in time of testing conditions (temperature, environment, etc.), testing technical

systems state, personnel qualification. Costs are generated by the volume of probes,

duration, cost for the exploitation of experimental technical systems, etc, of personnel,

respectively.

Accordingly, experiments are being planned.: Volume of probes, forming series, their

optimum distribution on loading parameter packages: type of loading, stresses, strains,

frequency, environment conditions, etc. Statistical processing of results is important to define

fatigue limit curves.

Increasing accuracy and efficiency of mechanical characteristics

This is made possible by planning the volume of experiments and adopting a more rational

method for statistical processing of results.

If the aim of testing is to assess the mathematically the mechanical characteristics, selection

„n” in considering the normal distribution is determined by relation:

n = (2 / a2) Z 2 1-/2 (4.3.6)

or n = Z 2 1-/2 / a2 (4.3.7)

where is the variation coefficient of determined mechanical characteristics; a –

mathematical error (tolerated) relative to the assessment of the average value, Z1-/2 –

quintile level; P = 1-/2 – statistical reliability, representing the probability of effective error

non-changed by assessing the average value of maximum error characteristics a or a; a –

maximum relative error (tolerated) when estimating the average value in sizes of the mean

square deviation of the analyzed mechanical characteristic.

The testing volume is corrected with values of variation coefficient of selection methods using

the relation:

N = (v2 /a2). t2, k (4.3.8)


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where: t, k is the value of the statistic quintile value for the level P = 1-/2 for the degree of

freedom number k = n – 1. Usually is adopted = 0.1 or 0.05, more rarely 0.01.

The size of the maximum value a and a is adopted depending on the necessary accuracy

of average values of determinations:

- reduced accuracy , a , and a = 1,

- average accuracy: a = (0,4-0,5) , and a = 0,4-0,5,

- high accuracy: a = (0,2-0,3) , and a = 0,2-0,3.

Distribution function of durability at the action of variable loading

Inevitable dispersion of results necessary to define a higher limit, inferior one, respectively,

for durability curves.

Analysis of results evinces that the aleatory size x = log (N – No) is distributed according to

a normal law. The main difficulty to use the normal distribution law of the magnitude „x” in

order to estimate the resource of construction elements in natural size is the complexity of

determining the quieting sensitivity. In order to test products or models there exists an

economic limit. In case a great number of samples are used the sensitivity of determinations

is valid for high stresses in correlation with the limited duration of the experiment. For law

stresses, in the case of real components, the assessment on samples takes long times.

Here from the necessity to adopt more simple distributions for which parameters are

easier to determine.

By using the independence assumption on the coefficients of fatigue life variation „R -1” of

the base cycle number and the fatigue curve equation for the fracture probability P = 0.5 is:

Ra = R-1 +a(log N) - (4.3.9)

or Ra = R-1 +b(log N) - (4.3.10)

where and depend on the analyzed material.

Square mean deviation „” for durability N, for symmetric cycle can be determined by the

relation:

log N = (R -1 / )[alog N + (R-1 /a) a+1 log N)] (4.3.11)

or log N = (R -1 /2,3) [ 1 + (R-1 /b) e 2,3 a log N] (4.3.12)

as to the previous relations.


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Statistical processing method

The experimental results are tabulated positions after increasing duration of durability,

including non-fractured specimens. Stress level at which all specimens were broken series,

determine media selection, dispersion, standard deviation of the logarithm of durability,

confidence intervals and the overall dispersion. Plot the empirical distribution of PN durability

in logarithmic coordinates for several levels of stress amplitude.

Estimation of quintiles related to the durability values Np is possible with the relation:

log Np = a + zp .

(4.3.13)

where a and are estimations of mathematical expectations, the square mean deviation of

the size x = log N, respectively zp - quintile of the normal aleatory size level.

Estimated sensitivity threshold based on the layout where the law of normal distribution of

random size x = log (N - No). Estimate, which is covering and is considered the threshold of

sensitivity, is determined by the relationship:

No = (Nmin . Nmax – N2 0,5 ) / (Nmax + Nmin – 2 N0,5) (4.3.14)

where Nmax , Nmin , N0,5 reprezintă valorile durabilităţi maxime, minime, mediane

determintate.

Relationship is valid for n ≥ 20 and (Nmax + Nmin) ≥ 5 N0, 5). Such calculations made

produced results similar to the graphical determination.

4.3.2 S – N Diagram SN curve is a characteristic of material, experimental high for any kind of request, or body.

It is recognized as the basic curve of the material. ConFigureation must be plain to ensure

uniform application of mechanical traction / compression.

Results obtained by cyclic bending required bodies, or bodies containing geometric

concentrators, reflecting the effect of certain influencing factors.

Figure 37 presents two curves:

- curve „a” defines the asymptotic voltage of SR below which no fracture occurs regardless of

the number of cycles of application,


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- curve „b” defines a level of application at which the material can fail at a specified number

of cycles.

Distinct areas on the SN curve (Figure 38): quasi-static fracture, cycle fatigue life, polycyclic,

respectively.

Figure 37 S-N Diagrams Figure 38 Specific zones of the S – N diagram

Application with very high voltage leads to quasi-static fracture.

Cycle fatigue life is located within 102-105 cycles, and for greater number of fatigue cycles

works the multicyclic fatigue.

4.3.3 Collective applications of voltage Aleatory complex loading spectra decomposed in Fourier series, under the form:

∞

S(t) = Smed + ∑ (αi.cos 2 π i . f. t + βi. sin2 π i . f. t ) (4.3.15)

i = 1 T

where: Smed = (1/T) . ∫ S. dt 0 T

αi = (2/T) ∫ Sa . cos (2 π i . f. t) dt

0 variable components of the stress (4.3.16)

T

βi = (2/T) ∫ Sa . sin (2 π i . f. t) d 0


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Sai = (αi 2 + βi

2 )

½ amplitude of i order harmonic (4.3.17)

Variable loadings unperiodical and transitrory are described by relation:

S(t) = Sa . e –α t . cos β . t (4.3.18)

It reflects loading situations, in the service, as real as possible (Figure 39).

There are a number of loading classes:

Sai = Sai+1 ± ∆ S (4.3.19)

In the representation Sai+1 / Sa as a function of N, the cumulated frequency curve is realized

(Figure 40).

Figure 39 Spectra of aleatory loading.

Figure 40 Curve of cumulated frequencies.

Number of loading cycles N = 5.105 ÷ 106, for a representative situation.

Fullness coefficient of spectrum:


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Cp = (Sa)N- / Smax. (4.3.20)

where: (SA)N- - intensity of loading corresponding to the maximum frequency,

Smax – maximum intensity od loading.

0 < Cp < 1 : 1 – harmonic loading spectrum,

0 - loading spectrum with normal log distribution.

Conventional number of loading cycles: N = 2.105 ÷ 5.106

Left side of the spectrum: slow loading low frequency, high amplitude.

Right side of the spectrum: fast loading: high frequency, low amplitude.

Depending on reproductive possibilities of loading spectrum testing program types are:

- blocks with monotonic load decrease,

- blocks with monotonic increase/ decrease of load,

- blocks that schematize aleatory loadings,

- complex programs on the bases of modelling the Markov superior oder process.

4.3.4 Fatigue resistance When a part is subjected to repeated cyclic loadings fracture can occur without observing

the degradation during the loading. The applied loading can be reduced so that evident

degradation is not noticed. Failure is the result of changes at micro/ sub microscopic level by

cumulative and irreversible degradation. The degradation process is correlated and

determined by the cyclic plastic deformation. The elastic deformation is reversible and does

not generate material degradation. Only the cyclic plastic deformation generates irreversible

changes, mainly in the dislocation substructure.

The fatigue resistance Sr is defined as the highest value of the maximum stress the

specimen, the material does not fails, indifferently of the number of loading cycles. It

represents the value corresponding to the unlimited durability of the material at the

respective loading.

Frequently is this necessary to determine the number of cycles at which material fails under

a prescribed loading over the level of fatigue resistance.

In order to determine materials characteristics at variable loadings the following methods are

used:

direct: Wőhler’s classical method, steps, Probit, with progressive loading (Prot, Locati)

indirect: based on the change of physical constants during the application variables

(elastic modulus, work absorption, magnetic permeability)

based on dependencies between fatigue strength and mechanical properties (R0, 2,

Rm, A, Z.


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Wöhler method call the dependence of the maximum tension the material yield (Rmax,,

Smax) and corresponding number of cycles to fracture (N). Testing continues until at least

one probe does not break. Include consecrated values (N0):

steels 2.106 – 2. 107 cycles,

light alloys 2. 107 – 108 cycles

With pairs of values: stress fracture (S) - number of cycles (N) the diagram in Figure 41 is

made: fatigue life curve (Wöhler), for determined loading conditions. The voltage coefficient

is the asymmetry index (R).

In semi-logarithmic coordinates, there are three areas (Figure 42): static requests or few

cycles (I) limited durability (II), non-limited durability or fatigue resistance (III).

Figure 42 Fatigue resistance curve

In logarithmic coordinates, the curve appears as two lines connected.

Dispersion results are affected by the homogeneity of the material, preparing samples, test

conditions, etc. In the plane S – N the curve can be replaced by an izoprobability network of

curves generated by the equation:

.,0

constpdNSNfNp

(4.3.21)

where p is probability (0-1). At the S3 loading level point 'a' is the probability that no

specimen fracture, and point "b" – the probability of survival of 0%. Significance points are

treated alike 'c', 'd' for a specified number of cycles.

The main factors affecting the durability of variable loadings, including weld fatigue limit are:

- asymmetry degree of the cycle,

- coefficient of fullness,

- overloading,

- geometry of the joint,

Figure 41 Fatigue curve domains


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- level of residual stresses,

- technological factors.

The asymmetry degree of the cycle (R) is defined by the dependence between the maximum

and the minimum stress.

Coefficient of fullness (cp) expresses the influence of the intensity of the aleatory loading

spectrum.

With decreasing degree of fullness of the spectrum and removal application spectrum with

constant amplitude at which cp = 1, the fatigue resistance and durability increase

continuously.

The cause is weight decrease of active processes of degradation and increase the effect of

structural strength. The range of use cp = 0-1. Other factors are subsequently treated in

terms of improving the reaction to variable loading.

4.3.5 The average voltage effect Fatigue curve is obtained by highlighting current-stress dependence on the number of cycles

(S-N). An alternative is the strain-stress dependence. It defines:

Testing with "soft loading cycles' when the stress is the evaluated one

Testing with "hard loading cycles” when the strain is the evaluate one

The difference between the two regimes is the most obvious in asymmetric cycles.

"Hard cyclic" loading with average tensile stress (Smed), leads to cyclic creep. "Soft cyclic"

with average tensile strain (εmed) leads to the stress-relieving of the creep stress. Figures 43

and 44 present the stress-relieving of average stress in case of pulsed loading, respectively

to the stress-relieving of creep stress for asymmetric cyclic loading. They represent typical

behaviours of materials at oligocyclic loading.

Figure 43 Stress-relieving cat pulsed constant

Figure 44 Creep stress-relieving asymmetric cyclic loading and controlled stress


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strain

The average stress (Smed) is stress –relieved after several loading cycles. The hystereses

loop becomes completely symmetric as a function of the stress. The stress-relieving rate

depends on the material, amplitude of deformation and its average value (εa, εmed). The final

value of the average stress, after stress-relieving must not become nule. Cyclic creep can

develop only in the case of the soft asymetric loading cycle (Controlled amplitude of stress).

The example is suggestively expressed in the previous Figure: Smed, Sa =const. Material

reacts so that εa, εmed ≠ constant. εmed increases with the number of cycles(N). Depending on

the material, Smed, Sa and T creep stops after N cycles or continues to fracture.

In the first case, fracture is base don fatigue, and in the second case it is base don plastic

instability.

The effect of average strain depends on the existence of average stress which is not zero

anbd can be understood only on the bases of stress change. So, the effect of average strain

is not significant when the average stress is rapidly relaxed during the controlled strain cycle,

but can be very important if the stress-relieving is a slow one. The stress-relieving rate

depends on material and strain. When the strain is higher the stress-relieving is more

reduced, and the effect on the average strain is more reduced.

Overlapping of the cyclic and average stress components in controlled conditions is to be

found out by cyclic creep. The tensile average stress shortens the fatigue life, while the

average compression stress make it longer. Figure 45 presents the effect of average stress

son fracture mechanisms under controlled stress conditions. There are four intervals:

Figure 45 Effect of average stress son fracture mechanisms for controlled stress testing.

a) For reduced levels of stress amplitude, under the fatigue limit there is no failure

indifferently of the average stress level

b) For levels of stress amplitude over the fatigue limit (median area of the diagram)

failure is fatigue typical, by initiation and propagation of cracks, preceded by creep

strain afferent to the average stress which is ≠ 0.


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c) For high levels of average stress the cyclic creep and ductile fracture prevail, by the

contraction of area.

d) For cyclic compression loading and high average stress the loss of stability occurs by

buckling.

For a prescribed amplitude of stress, it is found out that durability decreases, but not

monotonously, when average stress increases.

The effect of average stress on fatigue durability is suggestively expressed in Figure 46. The

maximum stress (Smax) depends on the average one (Sm) and the number of cycles (N). So,

a set of parametric curves (parameter N), is obtained and the experimental determination of

the diagram S-N is compulsory.

Figure 46 Influence of average stress son fatigue resistance for different Nf values.

4.3.6 Fatigue resistance of classified structural details The fatigue assessment of classified structural details and welded joints is based on the

nominal stress range. In most cases structural details are assessed on the basis of the

maximum principal stress range in the section where potential fatigue cracking is considered.

However, guidance is also given for the assessment of shear loaded details, based on the

maximum shear stress range. Separate S-N curves are provided for consideration of normal

or shear stress ranges, as illustrated in Figures 47 and 48 respectively.


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Figure 47 Fatigue resistance S-N curves for steel, normal stress

Figure 48 Fatigue resistance S-N curves for aluminium, normal stress

Care must be taken to ensure that the stress used for the fatigue assessment is the same as

that given in the tables of the classified structural details. Macro-structural hot spot stress

concentrations not covered by the structural detail of the joint itself, e.g. large cut-outs in the

vicinity of the joint, have to be accounted for by the use of a detailed stress analysis, e.g.

finite element analysis, or appropriate stress concentration factors.

The fatigue curves are based on representative experimental investigations and thus include

the effects of:

structural hot spot stress concentrations due to the detail shown,

local stress concentrations due to the weld geometry,

weld imperfections consistent with normal fabrication standards,

stress direction,

welding residual stresses,

metallurgical conditions,

welding process (fusion welding, unless otherwise stated),

inspection procedure (NDT), if specified,

postweld treatment, if specified.

Furthermore, within the limits imposed by static strength considerations, the fatigue curves of


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welded joints are independent of the tensile strength of the material.

Each fatigue strength curve is identified by the characteristic fatigue strength of the detail at 2

million cycles. This value is the fatigue class (FAT).

The slope of the fatigue strength curves for details assessed on the basis of normal stresses

is m=3.00. The constant amplitude knee point is at 1. 107

cycles. The slope at higher number

of cycles is m=22.

The slope of the fatigue strength curves for detailed assessments on the basis of shear

stresses is m=5.00, but in this case the knee point is at 108

cycles. The slope at higher

number of cycles is m=22.

The descriptions of the structural details only partially include information about the weld

size, shape and quality. The data refer to a standard quality as given in codes and standard

welding procedures. For higher or lower qualities, conditions of welding may be specified and

verified by test.

The fatigue classes given in table 4.3.1 shall be modified as given in chapter 4.3.5. The

limitations of weld imperfections shall be considered.

All butt welds shall be full penetration welds without lack of fusion, unless otherwise stated.

All S-N curves of details are limited by the material S-N curve, which may vary due to

different strengths of the materials.

Disregarding major weld defects, fatigue cracks originate from the weld toe, and then propagate through the base material, or from the weld root, and then propagate through the weld throat. For potential toe cracks, the nominal stress in the base material has to be calculated and compared with the fatigue resistance given in the tables. For potential root cracks, the nominal stress in the weld throat has to be calculated. If both failure modes are possible, e.g. at cruciform joints with fillet welds, both potential failure modes have to be assessed.

Fatigue verification is carried out using the design spectrum of fatigue actions in terms of

stress ranges Δζi,S,d

, in which the stresses of the characteristic spectrum Δζi, S, k have been

multiplied by the partial safety factor γF for fatigue actions.

The design resistance S-N curve based on design resistance stresses ΔζR,d

, in which the

characteristic resistance stress ranges ΔζR,k

have been divided by the partial safety factor γM

for fatigue resistance.

The design resistance S-N curve may be modified further according to the needs of the

damage calculation procedure.

For constant amplitude loading, the characteristic stress range ΔζR,k

at the required

number of stress cycles is firstly determined. Secondly the fatigue criterion is checked:


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M

kR

FkSdS

,

.,

(4.3.22)

At variable amplitude loading, cumulative damage calculation procedure is applied. Usually

a modified "Palmgren-Miner"-rule, is appropriate. For load spectra which are sensitive to the

position of the fatigue limit or cut-off limit, or in which the spectrum changes during the

service time, additional assessment using the nonlinear damage calculation method is

recommended.

In fields of application, where no test data or service experience exist and the shape of the

stress spectrum is not close to constant amplitude, it is recommended to proceed according

to the damage calculation.

4.3.7 Linear Damage Calculation by "Palmgren-Miner" For fatigue verification it has to be shown that the calculated usable cycles are larger than

the anticipated number of cycles occurring in service of the structure:

f

t

t

dN

nD

1

1

2

varNNN const

usable

(4.3.23)

where ΣDd damage by summation.

“i” index for block number in load spectrum of required design life ni number of cycles of

stress range Δζi,S,d

in load spectrum block i Ni number of cycles at which design stress range

Δζi,S,d

causes failure in the modified design fatigue resistance S-N curve. Nvar

number of

cycles calculated at variable amplitude load by use of damage summation ΣDd

Nconst

number

of cycles calculated at constant amplitude load of maximum stress range in spectrum Nusable

number of calculative usable cycles.

The order of sequence of the blocks has no effect on the results of this calculation.

In some cases it might be convenient to calculate an equivalent constant amplitude stress

range ΔζE and to compare it directly to the constant amplitude resistance S-N curve

neglecting the constant amplitude fatigue limit.

For the grid of fatigue resistance classes and an initial slope of m=3 predominantly used, the values of the modified characteristic fatigue resistance S-N curves have been calculated. Stepping down one class corresponds to a division by 1.12. So different levels of safety γ

M of

SN curve may be achieved (Figure 49, 50 and table 12).


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Figure 49 Modified resistance S-N curves of steel for Palmgren-Mine summation

Figure 50 Modified resustance S-N curves of aluminium for Palmgren-Miner summation

Table 12 FAT data, stress at knee-point of S-N curve, constants of tentative S-N curves and constants for Palmgren-Miner summation

FAT class

[MPa]

stress at knee

point [MPa]

# of cycles lower

than knee point of

S-N curve

# of cycles higher than knee point of S-N

curve Constant C: N=C/ Δζ m

Δζ at 2e6c. Δζ at 1e7 c. m=3 constant ampl. m=5 varable ampl. m=22

125 73.1 3.906E+12 2.0440E+47 2.091E+16

112 65.5 2.810E+12 1.8250E+46 1.207E+16

100 58.5 2.000E+12 1.5082E+45 6.851E+15


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90 52.7 1.458E+12 1.4852E+44 4.046E+15

80 46.8 1.024E+12 1.1129E+43 2.245E+15

71 41.5 7.158E+11 8.0564E+41 1.236E+15

63 36.9 5.001E+11 5.8070E+40 6.800E+14

56 32.8 3.512E+11 4.3511E+39 3.773E+14

50 29.3 2.500E+11 3.5958E+38 2.141E+14

45 26.3 1.823E+11 3.5411E+37 1.264E+14

40 23.4 1.280E+11 2.6533E+36 7.016E+13

36 21.1 9.331E+10 2.6128E+35 4.143E+13

32 18.7 6.554E+10 1.9578E+34 2.299E+13

28 16.4 4.390E+10 1.0374E+33 1.179E+13

25 14.6 3.125E+10 8.5731E+31 6.691E+12

22 12.9 2.130E+10 5.1494E+30 3.531E+12

20 11.7 1.600E+10 6.3259E+29 2.192E+12

18 10.5 1.166E+10 6.2295E+28 1.295E+12

16 9.4 8.192E+09 4.6677E+27 7.184E+11

14 8.2 5.488E+09 2.4733E+26 3.685E+11

12 7.0 3.456E+09 8.3262E+24 1.705E+11

m=5

160 116.0 2.097E+17 5.2373E+51 2.100E+17

80 58.0 6.554E+15 1.2487E+45 6.564E+15

70 50.8 3.361E+15 6.6164E+43 3.367E+15

Δη at 2e6 c. Δη at 1e8 c.

100 45.7 2.000E+16 3.2973E+44 2.000E+16

80 36.6 3.277E+15 2.4922E+42 3.277E+15

36 16.5 1.209E+14 6.0904E+34 1.209E+14

28 12.8 3.442E+13 2.2836E+32 3.442E+13


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4.3.8 Nonlinear Damage Calculation A nonlinear fracture mechanics damage calculation is recommended in cases, where:

1. The Miner summation is sensitive to the exact location of the knee point of the fatigue

resistance S-N curve

2. The spectrum of fatigue actions (loads) varies in service or is changed, and so the

sequence of loads becomes significant

3. The resistance S-N curve of a pre-damaged component has to be estimated.

4.3.9 Fatigue resistance against structural hot spot stress

A.Fatigue Resistance using Reference S-N Curve

The S-N curves for fatigue resistance against structural hot spot stress are given in the table

13 for steel and aluminium, where the definition of the FAT class is given in chapter 4.2. The

resistance values refer to the as-welded condition unless stated otherwise. The effects of

welding residual stress are included. Effects of misalignment are not included.

The design value of the structural hot spot stress range shall not exceed ΔFhs

< 2.fy.

Table 13 Fatigue resistance against structural hot spot stress

No

.

Structural detail Description Requirements FAT

Steel

FAT

Alu.

1

Butt joint As welded, NDT 100 40

2

Cruciform or T-

joint with full

penetration K-

butt welds

K-butt welds, no

lamellar tearing

100 40

3

Non load-

carrying fillet

welds

Transverse non-load

carrying attachment,

not thicker than main

plate, as welded

100 40

4

Bracket ends,

ends of

longitudinal stif-

feners

Fillet welds welded

around or not, as

welded

100 40

5

Cover plate ends

and similar joints

As welded 100 40


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6

Cruciform joints

with load-

carrying fillet

welds

Fillet welds, as welded 90 36

No Structural detail Description Requirements FAT FAT

. Steel Alu.

7

Lap joint with

load carrying fillt

welds

Fillet welds, as welded 90 36

8

Type “b” joint

with short

attachment

Fillet or full penetration

weld, as welded

100 40

9

Type “b” joint

with long

attachment

Fillet or full penetration

weld, as welded

90 36

Note: Table does not cover effects of misalignment. They have to be considered explicitely in determination of stress range.

For hollow section joints, special hot-spot stress design S-N curves have been

recommended by the IIW. The tubular joint design curves should not be applied to other

types of structure.

B. Fatigue resistance using a reference detail

The tables of the fatigue resistance of structural details, or fatigue data from other sources

which refer to a comparable detail, may be used. The reference detail should be chosen as

similar as possible to the detail to be assessed. Thus the procedure will be:

a) Select a reference detail with known fatigue resistance, which is as similar as

possible to the detail being assessed with respect to geometric and loading

parameters.

b) Identify the type of stress in which the fatigue resistance is expressed. This is usually

nominal stress.

c) Establish a FEM model of the reference detail and the detail to be assessed with the

same type of meshing and elements.

d) Load the reference detail and the detail to be assessed with the stress identified in b).


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e) Determine the structural hot spot stress ζhs, ref

of the reference detail and the

Structural hot spot stress ζhs, assess

of the detail to be assessed.

f) The fatigue resistance for 2 million cyles of the detail to be assessed FATassess

is then

calculated from fatigue class of the reference detail FATref

by:

ref

assesshs

refhs

assess FATFAT ,

,

(4.3.24)


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4.3.10 Fatigue resistance against effective notch stress A. Steel

The effective notch stress fatigue resistance against fatigue actions, as determined for steel

and aluminium, is given in table 14. The fatigue resistance value refers to the as-welded

condition. The effect of welding residual stresses is included. Possible misalignment is not

included.

Table 14 Effective notch fatigue resistance for steel

No. Quality of weld notch Material Description FAT

1 Effective notch radius equalling

1 mm replacing weld toe and

weld root notch

Steel

Notch as-welded, normal

welding quality m=3

225

2 Effective notch radius equalling

1 mm replacing weld toe and

weld root notch

Aluminium

Notch as-welded, normal

welding quality m=3

75

4.3.11 Fatigue strength modifications A. Stress ratio

A1. Steel

For stress ratios R<0.5 a fatigue enhancement factor f(R) may be considered by multiplying

the fatigue class of classified details by f(R). The fatigue enhancement factor depends on the

level and direction of residual stresses (Figure 51). It should only be used if reliable

information or estimation of the residual stress level was present. The following cases are to

be distinguished:

I. Unwelded base material and wrought products with negligible residual stresses

(<0.2fy), stress relieved welded components, in which the effects of constraints or

secondary stresses have been considered in analysis. No constraints in assembly.

f(R) = 1.6 for R < -1

f(R) = -0.4 .R + 1.2 for -1 ≤ R ≤0.5

f(R) = 1 for R > 0.5

II. Small scale thin-walled simple structural elements containing short welds. Parts or

components containing thermally cut edges. No constraints in assembly.

f(R) = 1.3 for R < -1

f(R) = -0.4 . R + 0.9 for -1 ≤ R ≤ -0.25

f(R) = 1 for R > -0.25


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III. Complex two or three-dimensional welded components, componentswith

globalresidual stresses, thickwalled components.

f(R) = 1 no enhancement

Figure 51 Enhancement factor f(R)

The ranking in categories I, II or III should be done and documented by the design office. If

no reliable information on residual stress is available, f(R)=1.

It has to be noted in this respect that stress relief in welded joints is unlikely to be fully

effective and long range residual stresses may be introduced during assembly of

prefabricated welded components. For such reasons, it is recommended that values of

f(R)>1 should only be adopted for welded components in very special circumstances.

A2. Aluminium

The same regulations as for steel are recommended.

4.3.12 Wall Thickness A. Steel

The influence of plate thickness on fatigue strength should be taken into account in cases

where cracks start from the weld toe. The fatigue resistance values here given refer to a wall

thickness of 25 mm at steel. The reduced strength is taken in consideration by multiplying the

fatigue class of the structural detail by the thickness reduction factor f(t). The thickness

correction exponent “n” is dependent on the effective thickness teff

and the joint category

(table 15 and Figure 52). The same way a benign thinness effect might be considered, but

should be verified by component test.

Table 15 Thickness correction exponents

Joint category Condition n

Cruciform joints, transverse T-joints, plates with transverse attachments as-welded 0.3

Cruciform joints, transverse T-joints, plates with transverse attachments toe ground 0.2


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Transverse butt welds as-welded 0.2

Butt welds ground flush, base material, longitudinal welds or attachements any 0.1

n

effttf

25 where t >25 mm

If L/t ≤ 2 then teff = 0.5·L

else teff = t (4.3.25)

The plate thickness correction factor is not required in the case of assessment based on

effective notch stress procedure or fracture mechanics.

Figure 52 Toe distance

B. Aluminium

The same regulations as for steel are recommended.

4.3.13 Improvement techniques A. General

Post weld improvement techniques may raise the fatigue resistance. These techniques

improve the weld profile, the residual stress conditions or the environmental conditions of the

welded joint. The improvements methods are:

1. Methods of improvement of weld profile:

a. machining or grinding of weld seam flush to surface,

b. machining or grinding of the weld transition at the toe,

c. remelting of the weld toe by TIG-, plasma or laser dressing.

2. Methods for improvement of residual stress conditions:


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a. peening (hammer-, needle-, shot- or brush-peening),

b. coining,

c. overstressing,

d. stress relieving thermal treatment

3. Methods for improvement of environmental conditions:

a. Painting

b. resin coating

The effects of all improvement techniques are sensitive to the method of application and the

applied loading, being most effective in the low stress high cycle regime. They may also

depend on the material, structural detail and dimensions of the welded joint. Consequently,

fatigue tests for the verification of the procedure in the endurance range of interest are

recommended in lot of references.

For some post welding improvement procedures, direct recommendations are given below.

They may be used under the following circumstances:

a) increasing the fatigue strength of new structures,

b) a verification by test is recommended,

c) repair or upgrading of existing structures.

The recommendations apply to nominal stress and structural hot spot stress method; they do not apply to effective notch stress and fracture mechanics method.

Applicability of improvement methods

Examples of joint suitable for improvement show in Figure 53.

Figure 53 Examples of joint suitable for improuvement

The recommendations apply to all arc welded steel or aluminium components subjected to

fluctuating or cyclic stress and designed to fatigue limit state criterion. They are limited to

structural steels up to a specified yield strength of 900 MPa and to structural aluminium

alloys commonly used in welded structures, primarily of the AA 5000 and AA 6000 series.


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The recommendations apply to welded joints of plates, of sections built up of plates or similar

rolled or extruded shapes, and hollow sections. If not specified else, the plate thickness

range for steel is from 6 to 150 mm, for aluminium from 4 to 50 mm.

The application is limited to joints operating at temperatures below the creep range. In general, the recommendations do not apply at low cycle fatigue conditions, so the nominal stress range is limited to

For the special improvement procedures additional restrictions may be given.

The improvement procedures described below, apply solely to the weld toe and to cracks

starting from this point. All other points of a possible start of fatigue cracks therefore should

be carefully considered as e.g. the weld root or weld imperfections. (Figure 54).

The recommendations do not apply to joints operating under free corrosion.

Figure 54 Examples of joints, at which an improvement might be limited by a possible root crack

Burr Grinding

Fatigue cracks initiate usually at the weld toe at points of cold fusion or other sharp crack-like

defects of a few tenth of a millimetre. The grinding has firstly to remove these defects and

secondly to create a smooth weld transition and thus to reduce the stress concentration. All

embedded imperfection which emerge to the surface at grinding must be repaired. The

benefit of burr grinding is given as a factor on the stress range of the fatigue class of a non-

improved joint (table 16).

Table 16 Benefit factors on stress of burr grinding and TIG dressing

Area of application Mild steel fy < 355

MPa

Steel fy > 355 MPa &

aluminium alloys

All structural details leading to a IIW fatigue

class of 90 at steel or 32 at aluminium or lower

as applicable

1.3 1.5 (1.3*)


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All S-N curves and fatigue classes for

assessment by structural hot spot stress, but

no higher class than 100 at steel or 40 at

aluminium. Butt joints to be assessed by the

nominal stress fatigue class.

1.3 1.5 (1.3*)

For transverse fillet welds at continuous plates, corresponding to FAT 80 at steel or FAT 28 at aluminium in the catalogue of details

TIG dressing

By TIG (tungsten inert gas) dressing, the weld toe is remolten in order to remove the weld

toe undercut or other irregularities and to smoothen the stress concentration of the weld

transition (table 17). The recommendations apply to partial or full penetration arc welded fillet

welds in steels with a specified yield strength up to 900 MPa and to wall thicknesses >= 10

mm operating non-corrosive environment or under conditions of corrosion protection. The

details of the procedure are described in references.

Hammer peening

By hammer peening, the material is plastically deformed at the weld toe in order to introduce

beneficial compressive residual stresses. The recommendation is restricted to steels with

specified yield strength up to 900 MPa and structural aluminium alloys, both operating non-

corrosive environment or under conditions of corrosion protection. The recommendations

apply for plate thicknesses from 10 to 50 mm at steel and 5 to 25 mm at aluminium and to

arc welded fillet welds with a minimum weld leg length of 0.1.t, where t is the wall thickness

of the stressed plate. The details of the procedure are described in references (table 17).

Table 17 Benefit on stress of hammer peening (nominal stress)

Area of application Benefit Requirements

All structural details leading

to a IIW fatigue class of 90 at

steel or 40 at aluminium or

lower

Upgrade for steel to FAT 125

for aluminium to FAT 56

Max. amount of nominal

compressive stress in load

spectrum < 0.25 · fy ,

including proof loading

if R < 0 then use Δζ

if R >=0 then use max ζ

instead of Δζ (for aluminium

use fy of HAZ !)

For structural hot spot stress see recommendations for needle peening.


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Needle peening

By needle peening, the material is plastically deformed at the weld toe in order to introduce

beneficial compressive residual stresses. Before any application, it is recommended to grind

the weld toe in order to remove undercut and weld toe irregularities and subsequently to

finish with a sandpaper tool for a glossy surface. The details of the procedure are described

in in references (table 18).

Table 18 Benefit on stress of needle peening (nominal stress)

Area of application Benefit Requirements

All structural details

leading to a IIW fatigue

class of 90 at steel or 40

at aluminium or lower

Upgrade for steel to

FAT 125 for aluminium

to FAT 56

Max. amount of nominal compressive

stress in load spectrum if R<0 then use

if R>=0 then use instead of (for aluminium

use fy of HAZ)

At all peening techniques, the structural hot spot stress approach should be applied only to

joints with fillet welds (with any penetration) and not to butt joints. The structural hot spot

stress, which includes the stress increase due to the structural geometry and possible

misalignments can be assessed by the corresponding material S-N curve, e.g. FAT 160 for

steel and FAT 71 for aluminium alloys in conjunction with the slope exponent m=5.0 . In this

way, the base metal at the weld toe is assumed to have a lower fatigue strength than the

peened weld.

4.3.14 Effect of elevated temperatures A. Steel

For higher temperatures, the fatigue resistance data may be modified with a reduction factor given in Figure 55. The fatigue reduction factor is a conservative approach and might be raised according to test evidence or application codes.

Figure 55 Fatigue strength reduction factor for steel at elevated temperatures


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B. Aluminium

The fatigue data given here refer to operation temperatures lower than 70 °C. This value is a

conservative approach. It may be raised according to test evidence or an applicable code.

4.3.15 Effect of corrosion The fatigue resistance data given here refer to non-corrosive environments. Normal

protection against atmospheric corrosion is assumed. A corrosive environment or

unprotected exposure to atmospheric conditions may reduce the fatigue class. The fatigue

limit may also be reduced considerably. The effect depends on the spectrum of fatigue

actions and on the time of exposure.

For steel, except stainless steel, in marine environment not more than 70% of the fatigue

resistance values in terms of stress range shall be applied and no fatigue limit be considered.

In fracture mechanics crack propagation calculations the constant C0 of the Paris Power Law

shall be multiplied by a factor of 3.0 . A threshold value shall not be considered.

No further specific recommendations are given for corrosion fatigue assessment.

4.3.16 Fatigue resistance against crack propagation The resistance of a material against cyclic crack propagation is characterized by the material

parameters of the "Paris" power law of crack propagation:

mKCdN

da 0 if thKK then 0

dN

da (4.3.25)

where the material parameters are C0 - constant of the power law, m - exponent of the power

law, ΔK range of cyclic stress intensity factor, ΔKth

- threshold value of stress intensity, under

which no crack propagation is assumed R, ratio Kmin

/Kmax

, taking all stresses including

residual stresses into account.

In the absence of specified or measured material parameters, the values given below are

recommended. They are characteristic values.

B. Steel

C0 = 1.58 .10

-11

(units in MPa%m and m) or

C0 = 5.0.10

-13

(units in N*mm-3/2

and mm)

m = 3

ΔKth = 6.0 - 4.56.R but not lower than 2 (units in MPa%m), or

ΔKth = 190 - 144 .R but not lower than 62 (units in N*mm

-3/2

)

C. Aluminium


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C0 = 1.27.10

-9

(units in MPa%m and m) or

C0 = 4.00.10

-11

(units in N*mm-3/2

and mm)

m = 3

ΔKth = 2.0 - 1.5.R but not lower than 0.7 (units in MPa%m), or

ΔKth = 63 – 48.R but not lower than 21 (units in N*mm

-3/2

)

Where the parameters for a fracture mechanics fatigue assessment are not known and only

the resistance S-N curve is known, the S-N curve can be used to derive dimensionless

fracture mechanics parameters, which allow a damage calculation. The procedure is based

on the "Paris" power law of crack propagation:

mKCdN

da 0 if thKK then 0

dN

da (4.3.26)

where “a” crack parameter, damage parameter (dimensionless); N Number of cycles; ΔK

range of stress intensity factor; ΔKth

threshold value of stress intensity factor range; C0, m

material constants.

The characteristic stress intensity factor range ΔKS,k

of the fatigue action is calculated with

the stresses of the spectrum ΔFi,S,k

and the crack parameter “a”:

aK kSkS ,, (4.3.27)

The characteristic resistance parameters can be derived from the characteristic constant

amplitude fatigue resistance S-N curve: The threshold value corresponds to the fatigue limit,

ΔKth,k

=ΔFL,R,k

, m equals the slope of the S-N curve, and the constant C0,k

can be calculated

from a data point (ΔFS-N

and NS-N

) on the S-N curve, preferably from the fatigue class at 2

x106

cycles:

m

NSNS

kNm

C

)2(

2,0 (4.3.28)

The fatigue verification is executed according to 4.4, using an initial crack parameter ai=1

and a final one af=4 or a large number e.g. a

f=10

9

. The restrictions on life cycles given in 4.3

are to be considered. The actual fatigue class of a pre-damaged component is FATact.

=

FAT/a.


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4.3.17 Fatigue assessment by crack propagation calculation The fatigue action represented by the design spectrum of stress intensity factor ranges:

FkSidSi KK ,,,, (4.3.27)

is verified by the material resistance design parameters against crack propagation

Mk

m

Mkd CCC ,0,0,0

M

kth

dth

KK

,

,

(4.3.28)

using the "Paris" power law

mKCdN

da 0 if thKK then 0

dN

da (4.3.29)

where a - crack parameter, damage parameter; N Number of cycles; ΔK range of stress

intensity factor; ΔKth

threshold value of stress intensity factor range; C0, m material constants.

At stress intensity factors which are high compared with the fracture toughness of the

material, Kc, an acceleration of crack propagation will occur. In these cases, the following

extension of the "Paris" power law of crack propagation is recommended. In the absence of

an accurate value of the fracture toughness, a conservative estimate should be made.

c

m

K

KR

KC

dN

da

1

0 (4.3.30)

where Kc fracture toughness, R stress ratio.

The number of life cycles N is determined by integration starting from an initial crack

parameter ai to a final one a

f. The calculated number of life cycles N has to be greater or

equal to the required number of cycles.

In general, the integration has to be carried out numerically. The increment for one cycle is:

m

dd KCda ,0 , if dthd KK , then da = 0 (4.3.31)

It is recommended that a continuous spectrum is subdivided to an adequate number of

stress range blocks, e.g. 8 or 10 blocks, and the integration performed block wise by

summing the increments of “a” and the number of cycles of the blocks. The entire size of the

spectrum in terms of cycles should be adjusted by multiplying the block cycles by an

appropriate factor in order to ensure at least 20 loops over the whole spectrum in the


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integration procedure.

4.3.18 Fatigue assessment by service testing

A. General

Components or structures may be tested or verified in respect to fatigue for different reasons:

a) Existence of a new design with no or not sufficient knowledge or experience of fatigue

behaviour.

b) Verification of a component or structure for a specified survival probability under a

special fatigue action (stress) history.

c) Optimization of design and/or fabrication in respect of weight, safety and economy

after pre-dimensioning. Pre-dimensioning may be done by the use of higher fatigue

resistance data, according to a lower survival probability in comparison with the

resistance data given here. Then the verification is achieved by a subsequent

component testing (Figure 56).

Figure 56 Example of scatter of test data

A pre-dimensioning leading to the mean values of the resistance data may be done by multiplying the resistance values in terms of stress by a factor of 1.5, which is based on a standard deviation of log cycles of 0.25 and an exponent of m=3.00 .

The verification or assessment depends of the safety strategy considered. Safe life, fail safe

and damage tolerant strategy have to be distinguished.

The fatigue tests should be performed using the data of the fatigue action history, factored by

the partial safety factors γF and γ

M, i.e., the stress levels of the action history have to be

multiplied by γF . γ

M for testing (table 19).


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Table 19 Testing approaches

No. Testing procedure Approach

1 all specimens of the samples are tested until failure all failed

2 testing is stopped at failure of first specimen of the sample first to fail

3 testing is stopped when „p” specimens of the „n” samples fail „p” to fail

The all failed approach is the normal way of testing at small size samples of which each

specimen represents the same weld details. The statistical analysis uses the data of the

failed specimens disregarding the non-failed ones.

The first to fail approach may be used at a large scale sample with the same weld details

and loading. The test is stopped at the first failure of a specimen.

The “n” to fail approach is used in similar conditions as the “first to fail” one, when repairs of

crack details can be performed during the test. Each time when a detail fails, the test is

stopped and the failed detail is repaired. Repairs are stopped depending of test conditions.

At the end possibbly all details have failed and thus the “all failed” approach is applied. If only

“p” specimens out of the “n” size of the sample failed, the “p to fail” approach is used.

This chapter considers the all failed and first to fail approaches.

The following test result data should be documented according to the selected approach:

The mean of the log of number of cycles at failure of all “n” failed samples or details.

The number of cycles of the first failed detail within “n” tested details.

The number of cycles of the first “p” failed details within “n” tested details.

The tests should be performed according to well established and appropriate procedures or

standards.

For the evaluation of service tests, an estimate of the standard deviation of logN has to be

made, taking into account that the standard deviation varies with the life cycle of the

component to be assessed.

For the number of test results being n>10, the standard deviation has to be calculated. For the number of test results being n<10, or if the procedure of first failure or “p” failures in

“n” specimens is used, the standard deviation can be estimated as follows:

0.178 for geometrically simple structures at a number of cycles between 104

and 105

0.25 for complex structures at cycles up to 106

-----no estimate can be given for higher cycles near the endurance limit. Here special verification procedures are recommended.


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B. Acceptance criteria

The number of design life cycles of the component or structure should be less than the

minimum probable number of the test life cycles:

F

NN T

d (4.3.32)

where NT

number of test life cycles of the test specimens corresponding to the log mean

value or number of cycles of the first test specimen to fail, whichever is applicable. F factor

dependent of the number of test results available as defined in tables 4.3.8, 4.3.9. The F-

factors refer to a 95% survival probability at a two sided confidence level of 75% of the mean

(see also 6.4) Nd

number of design life cycles, up to which the component or structure may

be used in service.

If all components or test specimens are tested to failure, table 20, shall be used.

Table 20 F-factors for failure of all test specimens

Stdv. \ n 2 4 6 8 10

0.178 3.93 2.64 2.45 2.36 2.30

0.200 4.67 2.97 2.73 2.55 2.52

0.250 6.86 3.90 3.52 3.23 3.18

If the tests are carried out until failure of the first test specimen, table 21 shall be used, the

factor F may be further modified according to safety requirements.

Table 21 F-factors for the first test specimen to fail

Stdv. \ n 2 4 6 8 10

0.178 2.72 2.07 1.83 1.69 1.55

0.200 3.08 2.26 1.98 1.80 1.64

0.250 4.07 2.77 2.34 2.09 1.85

C. Safe life verification

Safe life verification considers each structural element and detail as independent. Each

element has to fulfill the acceptance criteria previous as defined.

The partial safety factors γF applied to fatigue actions (loads) and γ

M applied to fatigue

resistance may be selected.


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D. Fail safe verification

Fatigue life verification of failsafe structures depends largely on the design and operation

parameters of a structure. The effectiveness of statically over-determined (hyperstatic)

behaviour or redundancy of structural components, the possibility of detection of failures in

individual structural parts and the possibility of repair determine the level of safety required in

the individual structural parts. So, no general recommendation can be given.




E. Damage tolerant verification

The verification is based on crack growth measurements, starting from a crack size, which

can be detected in inspection up to a critical crack size, at which the limit state of critical

safety against brittle or plastic fracture or other modes of failure of the remaining sectional

area is attained.

The criteria for factoring the observed life cycles for the test depend of the application. It is

recommended to establish agreement on the factor F.




4.3.19 Fatigue resistance of joints with weld imperfections

A.Types of Imperfections

The types of imperfections covered in this document are listed below. Other imperfections,

not yet covered, may be assessed by assuming similar imperfections with comparable notch

effect.

a. Imperfect shape

All types of misalignment including centre-line mismatch (linear misalignment) and angular

misalignment (angular distortions, roofing, peaking).

b. Volumetric discontinuities

Gas pores and cavities of any shape. Solid inclusions, such as isolated slag, slag lines, flux, oxides and metallic inclusions.

c. Planar discontinuities.

All types of cracks or cracklike imperfections, such as lack of fusion or lack of penetration (Note that for certain structural details intentional lack of penetration is already covered, e.g. at partial penetration butt welds or cruciform joints with fillet welds)

If a volumetric discontinuity is surface breaking or near the surface, or if there is any doubt

about the type of an embedded discontinuity, it shall be assessed like a planar discontinuity.


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B. Effects and assessment of imperfections

At geometrical imperfections, three effects affecting fatigue resistance can be distiguished, as summarized in table 22.

a. Increase of general stress level

This is the effect of all types of misalignment due to secondary bending. The additional

effective stress concentration factor can be calculated by appropriate formulae. The fatigue

resistance of the structural detail under consideration is to be lowered by division by this

factor.

Table 22 Categorisation and assessment procedure for weld imperfections

Effect of imperfection Type of imperfection Assessment

Rise of general stress le-

vel Misalignment

Formulae for effective

stress concentration

Local notch

effect additive

Weld shape

imperfections, undercut Tables given

competitive Porosity and inclusions

not near the surface Tables given

Cracklike imperfection Cracks, lack of fusion and

penetration, all types of

imperfections other than

given here

Fracture mechanics

b. Local notch effect

Here, interaction with other notches present in the welded joint is decisive. Two cases are to

be distinguished:

b1. Additive notch effect

If the location of the notch due to the the weld imperfection coincides with a structural

discontinuity associated with the geometry of the weld shape (e.g. weld toe), then the fatigue

resistance of the welded joint is decreased by the additive notch effect. This may be the case

at weld shape imperfections.

b2. Competitive notch effect

If the location of the notch due to the weld imperfection does not coincide with a structural

geometry associated with the shape geometry of the weld, the notches are in competition.

Both notches are assessed separately. The notch giving the lowest fatigue resistance is

governing.


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c. Cracklike imperfections

Planar discontinuities, such as cracks or cracklike imperfections, which require only a short

period for crack initiation, are assessed using fracture mechanics on the basis that their

fatigue lives consist entirely of crack propagation.

After inspection and detection of a weld imperfection, the first step of the assessment

procedure is to determine the type and the effect of the imperfection as given here.

If a weld imperfection cannot be clearly associated to a type or an effect of imperfections listed here, it is recommended that it is assumed to be cracklike.

C. Misalignment

Misalignment in axially loaded joints leads to an increase of stress in the welded joint due to

the occurrence of secondary shell bending stresses. The resulting stress is calculated by

stress analysis or by using the formulae for the stress magnification factor km.

Secondary shell bending stresses do not occur in continuous welds longitudinally loaded or in joints loaded in pure bending, and so misalignment will not reduce the fatigue resistance. However, misalignment in components, e.g. beams, subject to overall bending may cause secondary bending stresses in parts of the component, where the through thickness stress gradient is small, e.g. in a flange of a beam, where the stress is effectively axial. Such cases should be assessed.

Some allowance for misalignment is already included in the tables of classified structural

details (3.2). In particular, the data for transverse butt welds are directly applicable for

misalignment which results in an increase of stress up to 30%, while for the cruciform joints

the increase can be up to 45%. In local concepts of fatigue analysis, a small but unevitable

amount of misalignment according to a stress manification factor of km

=1.05 is already

included in the fatigue resistance S-N curves.

In special joints, i.e. all listed in table 23, the effect of a larger misalignment has to be addi-

tionally considered in the local stress (structural hot spot stress or effective notch stress. The

misalignement effect may be present even in the vicinity of supporting structures. A

corresponding stress increase has to be taken into account also in crack propagation

analyses. In all those cases, where the stress magnification factor is directly calculated, the

effective stress magnification factor km, eff

should be calculated as given below:

eredalreadym

calculatedm

effmk

kK

cov,

,

, (4.3.33)

For the simultaneous occurrence of linear and angular misalignment, both stress

magnification:

111 ,, angularmaxialmm kkk (4.3.34)

As misalignment reduces the fatigue resistance, the fatigue resistance of the classified

structural detail (3.2) has to be divided by the effective stress magnification factor.


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Table 23 Consideration of stress magnification factors due to misalignment

Type of km analysis

Nominal stress

approach Structural hot spot and effective notch approach

Type of welded joint km already cover-

ed in FAT class

km already

covered in SN

curves

Default value of effective km to

be considered in stress

Butt joint made in shop

in flat position

1.15 1.05 1.10*

Other butt joints 1.30 1.05 1.25*

cruciform joints 1.45 1.05 1.40*

One-sided fillet welds 1.25 1.05 1.20**

*) but not more than (0.95 + 3. emax /t), where emax = permissible misalignment and t = wall

thickness of loaded plate,

**) but not more than (0.95 + 0.3. tref/t), where tref = reference wall thickness

D. Undercut

The basis for the assessment of undercut is the ratio u/t, i.e. depth of undercut to plate

thickness. Though undercut is an additive notch, it is already considered to a limited extent in

the tables of fatigue resistance of classified structural details for steels and aluminium (table

24, 25).

Undercut does not reduce fatigue resistance of welds which are only longitudinally loaded. Table 24 Acceptance levels for weld toe undercut in steel

Fatigue class Allowable undercut u/t

butt welds fillet welds

100

90

80

71

63

56 and lower

0.025

0.05

0.075

0.10

0.10

0.10

not applicable

not applicable

0.05

0.075

0.10

0.10

Notes: a) undercut deeper than 1 mm shall be assessed as a crack-like imperfection. b) the table

is only applicable for plate thicknesses from 10 to 20 mm


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Table 25 Acceptance levels for weld toe undercut in aluminium

Fatigue class Allowable undercut u/t

butt welds fillet welds

50

45

40

36

32

28 and lower

0.025

0.05

0.075

0.10

0.10

0.10

not applicable

not applicable

0.05

0.075

0.10

0.10

Notes: a) undercut deeper than 1 mm shall be assessed as a crack-like imperfection. b) the table

is only applicable for plate thicknesses from 10 to 20 mm

E. Porosity and inclusions

Embedded volumetric discontinuities, such as porosity and inclusions, are considered as

competitive weld imperfections which can provide alternative sites for fatigue crack initiation

than those covered by the fatigue resistance tables of classified structural details steels and

aluminium (table 26, 27).

Before assessing the imperfections with respect to fatigue, it should be verified that the

conditions apply for competitive notches, i.e. that the anticipated sites of crack initiation in the

fatigue resistance tables do not coincide with the porosity and inclusions to be assessed and

no interaction is expected.

It is important to ensure that there is no interaction between multiple weld imperfections, be it

from the same or different type. Combined porosity or inclusions shall be treated as a single

large one. The defect interaction criteria must well know, for the assessments of cracks also

apply for adjacent inclusions. Worm holes shall be assessed as slag inclusions.

If there is any doubt about the coalescence of porosity or inclusions in the wall thickness

direction or about the distance from the surface, the imperfections shall be assessed as

cracks. It has to be verified by NDT that the porosity or inclusions are embedded and

volumetric. If there is any doubt, they are to be treated as cracks.

The parameter for assessing porosity is the maximum percentage of projected area of

porosity in the radiograph; for inclusions, it is the maximum length.


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Table 26 Acceptance levels for porosity and inclusions in welds in steel

Fatigue class Max. length of an inclusion in mm Limits of

porosity in %

of area * ** as-welded stress relieved +

100

90

80

71

63

56 and lower

1.5

2.5

4

10

35

no limit

7.5

19

58

no limit

no limit

no limit

3

3

3

5

5

5

* Area of radiograph used is length of weld affected by porosity multiplied by width of

weld

** Maximum pore diameter or width of an inclusion less than 1/4 plate thickness or 6

mm

+ Stress relieved by post weld heat treatment.

Table 27 Acceptance levels for porosity and inclusions in welds in aluminium

Fatigue class Max. length of an

inclusion in mm **

Limits of porosity in % of

area * **

as-welded

40 and higher

36

32

28

25

15 and lower

1.5

2.5

4

10

35

no limit

0 +)

3

3

5

5

5

* Area of radiograph used is length of weld affected by porosity multiplied by width of weld

** Maximum pore diameter or width of an inclusion less than 1/4 plate thickness or 6 mm

+) Single pores up to 1.5 mm allowed

***Tungsten inclusions have no effect on fatigue behaviour and therefore do not need to be

assessed.


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4.3.20 Fatigue resistance values for structural details in steel and

aluminium assessed on the basis of nominal stresses (table 28)

Table 28 Fatigue resistance values for structural details in steel and aluminium assessed on the basis of nominal stresses

No. Structural Detail

Description

(St.= steel;

Al.=

aluminium)

FAT

St.

FAT

Al.

Requirements and

Remarks

1

Rolled or

extruded

products,

components

with mashined

edges,

seamless

hollow

sections.

m = 5

St.: For high

strength steels

a higher FAT

class may be

used if verified

by test.

Al.: AA

5000/6000

alloys AA 7000

alloys

160

70

80

No fatigue resistance of

any detail to be higher at

any number of cycles!

Sharp edges, surface and

rolling flaws to be removed

by grinding. Any machining

lines or groves to be par-

allel to stresses! For high

strength steels a higher

FAT class may be used if

verified by test.

2

Machine gas

cut or sheared

material with

subsequent

dressing, no

cracks by

inspection, no

visible

imperfections

m = 3

140

All visible signs of edge

imperfections to be

removed. The cut surfaces

to be mashined or ground,

all burrs to be removed. No

repair by welding refill!

Notcheffects due to shape

of edges have to be

considered.


106 Of 150

3

Machine

thermally cut

edges, corners

removed, no

cracks by

inspection m =

3

125 40


of edges have to be

considered.

4

Manually

thermally cut

edges, free

from cracks

and severe

notches m = 3

100


of edges have to be

considered.

5

Manually

thermally cut

edges, uncon-

trolled, no

notch deeper

than .5 mm m =

3

80


of edges have to be

considered.

6

Transverse

loaded butt

weld (X-groove

or V-groove)

ground flush to

plate, 100%

NDT

100 40

All welds ground flush to

surface, grinding paralell to

direction of stress. Weld

run-on and run-off pieces to

be used and subsequently

removed. Plate edges to be

ground flush in direction of

stress. Welded from both

sides. No misalignement.

Required quality cannot be

inspected by NDT !

7

Transverse butt

weld made in

shop in flat

position, toe

angle # 30°,

NDT

90 36

Weld run-on and run-off

pieces to be used and

subsequently removed.

Plate edges to be ground

flush in direction of stress.

Welded from both sides.

Misalignment <5%

8

Transverse butt

weld not

satisfying

conditions of

212, NDT Al.:

Butt weld with

80

32

25








107 Of 150

toe angle #50°

Butt welds with

toe angle >50/

Misalignment <10%

9

Transverse butt

weld, welded

on ceramic

backing, root

crack

80 28

Backing removed, root

visually inspected.

Misalignment <10%

10

Transverse butt

weld on

permanent

backing bar

terminating >10

mm from plate

edge, else

71

63

25

22 Misalignment <10%

11

Transverse butt

welds welded

from one side

without backing

bar, full

penetration

root controlled

by NDT no

NDT

71

36

28

12

Misalignment <10%

12

Transverse

partial

penetration butt

weld, analysis

based on

stress in weld

throat sectional

area, weld

overfill not to

be taken into

account.

36 12

The detail is not

recommended for fatigue

loaded members.

Assessment by notch

stress or fracture

mechanics is preferred.

13

Transverse butt

weld ground

flush, NDT,

with transition

in thickness

and width slope

1:5 slope 1:3

100

90

80

40

32

28


surface, grinding paralell to

direction of stress. Weld

run-on and run-off pieces to

be used and subsequently



stress. Misalignment <10%


108 Of 150

slope 1:2 Exceeding misalignment

due to thickness step to be

considered.

14

Transverse butt

weld made in

shop, welded in

flat position,

weld profile

controlled,

NDT, with

transition in

thickness and

width: slope 1:5

slope 1:3 slope

1:2

90

80

72

32

28

25






Misalignment <10%

Exceeding misalignment

due to thickness step to be

considered.

15

Transverse butt

weld, NDT, with

transition on

thickness and

width slope 1:5

slope 1:3 slope

1:2

80

71

63

25

22

20


pieces

to be used and subse-

quently



stress.

Misalignment <10%

Exceeding

misalignments due to

thickness

step to be considered.


109 Of 150

16

Transverse butt

weld, different

thicknesses

without

transition,

centres

aligned. In

cases, where

weld profile is

equivalent to a

moderate slope

transition, see

no. 222

71 22

Misalignment <10% of

smaller

plate thickness

17

Three plate

connection,

root crack

71 22 Arc welds: Misalignment

<10%

18

Transverse butt

weld flange

splice in built-

up section

welded prior to

the assembly,

ground flush,

with radius

transition, NDT

100 40


surface,

grinding paralell to direction

of

stress. Weld run-on and

run-off


subsequently


ground


19

Transverse butt

weld splice in

rolled section

or bar besides

flats, ground

flush, NDT

80 28


surface,


of


run-off


subsequently


ground


110 Of 150


20

Transverse butt

weld splice in

circular hollow

section, welded

from one side,

full penetration,

root crack root

inspected by

NDT no NDT

71

36

28

12 Welded in flat position.

21

Tubular joint

with permanent

backing

71 28 Welded in flat position.

22

Transverse butt

weld splice in

rectangular

hollow section,

welded from

one side, full

penetration,

root crack root

inspected by

NDT no NDT

56

36

25

12 Welded in flat position.

23

Transverse butt

weld ground

flush, weld

ends and

radius ground,

100% NDT at

crossing

flanges, radius

transition.

100 40


surface,


of


run-off


subsequently


ground



No


111 Of 150

misalignment. Required

weld quality

cannot be inspected by

NDT

24

Transverse butt

weld made in

shop at flat

position, weld

profile

controlled,

NDT, at

crossing

flanges, radius

transition

90 36


pieces

to be used and subse-

quently


ground


Welded

from both

sides.Misalignment <5%

25

Transverse butt

weld ground

flush, NDT, at

crossing

flanges with

welded

triangular

transition

plates, weld

ends ground.

Crack starting

at butt weld.

For crack of

throughgoing

flange see

details 525 and

526.

80 32


surface,


of

stress. Plate edges to be

ground flush

in direction of stress.

Welded

from both sides.

Misalignment <10%

26

Transverse butt

weld, NDT, at

crossing

flanges, with

welded

triangular

transition

plates, weld

ends ground.

Crack starting

at butt weld.

71 28


flush in

direction of stress. Welded

from

both sides.Misalignment

<10%


112 Of 150

For crack of

throughgoing

flange see

details 525 and

526!

27

Transverse butt

weld at

crossing flan-

ges. Crack

starting at butt

weld. For crack

of throughgoing

flange see

details 525 and

526!

50 20 Welded from both sides.

Misalignment <10%

28

Automatic

longitudinal

seam welds

without

stop/start

positions in

hollow sections

with stop/start

positions

125

90

50

36

29

Longitudinal

butt weld, both

sides ground

flush parallel to

load direction

125 50

30

Longitudinal

butt weld,

without

stop/start

positions, NDT

with stop/start

positions

125

90

50

36

31

Continuous

automatic

longitudinal ful-

ly penetrated

K-butt weld

without

stop/start

positions

125 50

No start-Stop position is

permitted

except when the repair is

performed by a specialist

and

inspection is carried out to


113 Of 150

(based on

stress range in

flange) NDT

verify

the proper execution of the

weld.

32

Continuous

automatic

longitudinal

double sided

fillet weld

without stop-

/start positions

(based on

stress range in

flange)

100 40 Discussion: EC3 has 112

33

Continuous

manual

longitudinal

fillet or butt

weld (based on

stress range in

flange)

90 36

34

Intermittent

longitudinal

fillet weld

(based on

normal stress

in flange ζ and

shear stress in

web ч at weld

ends). ч / ζ = 0

0.0 - 0.2

0.2 - 0.3

0.3 - 0.4

0.4 - 0.5

80

71

63

56

32

28

25

22

Analysis based on normal

stress

in flange and shear stress

in web

at weld ends.

representation by

formula:

-steel 80.[1- (Δ ζ/ Δ ч) but

>=36

-alum. 36.[1- (Δ ζ/ Δ ч) but

>=14

35

Joint at

stiffened

knuckle of a

flange to be

assessed

according to

no. 411 - 414,

depending on

type of joint.


114 Of 150

Stress in

stiffener plate:

Af = area of

flange ASt =

area of stiffener

Stress in weld

throat: Aw =

area of weld

throat

36

Unstiffened

curved flange

to web joint, to

be assessed

according to

no. 411 414,

depending on

type of joint.

Stress in web

plate: Stress in

weld throat: Ff

axial force in

flange t

thickness of

web plate a

weld throat

The resulting force of Ff-left

and

Ff-right will bend the flange

perpenticular to the plane

of main

loading. In oder to minimize

this

additional stressing of the

welds,

it is recommended to

minimize

the width and to maximize

the

thickness of the flange.

Stress

longitudinally to the weld is

to be

considered. At additional

shear,

principle stress in web is to

be

consired (see 321 to 323)

37

Cruciform joint

or T-joint, K-

butt welds, full

penetration, no

lamellar

tearing,

misalignment

e<0.15At, weld

80 28

Material quality of

intermediate plate has to

be checked against

susceptibility of lamellar

tearing. Misalignment

<15% of primary plate.


115 Of 150

toes ground,

toe crack

38

Cruciform joint

or T-joint,

single-sided arc

welded fillet or

partial

penetration Y-

butt weld, no

lamellar

tearing,

misalignment of

plates e<

0.15At, stress

at weld root.

Penetration

verified.

Penetration not

verified.

71

36

25

12

Analysis based on stress in

weld root. Excentricity e of

plate t and weld throat a

midpoints to be considered

in analysis. Stress at weld

root: )F w, root = )F w,

nom A (1+6e/a) An

analysis by effective notch

procedure is recommended

39

Splice of rolled

section with

intermediate

plate, fillet

welds, weld

root crack.

Analysis base

on stress in

weld throat.

36 12

40

Splice of

circular hollow

section with

intermediate

plate,

singlesided butt

weld, toe crack

wall thickness

> 8 mm wall

thickness < 8

mm

56

50

22

20

Welds NDT inspected in

order to ensure full root

penetration.


116 Of 150

41

Splice of

circular hollow

section with

intermediate

plate, fillet

weld, root

crack. Analysis

based on

stress in weld

throat. wall

thickness > 8

mm wall

thickness < 8

mm

45

40

16

14

42

Splice of

rectangular

hollow section,

single-sided

butt weld, toe

crack wall

thickness > 8

mm wall

thickness < 8

mm

50

45

20

18

Welds NDT inspected in

order to ensure full root

penetration.

43

Splice of

rectangular

hollow section

with

intermediate

plate, fillet

welds, root

crack wall

thickness > 8

mm wall

thickness < 8

mm

40

36

16

14

44

Weld

connecting web

and flange, loa-

ded by a

concentrated

force in web

plane

perpendicular

to weld. Force

distributed on

Full penetration butt weld.


117 Of 150

width b = 2Ah

+ 50 mm.

Assessment

according to

no. 411 - 414.

A local bending

due to

eccentric load

should be

considered.

45

Transverse

non-load-

carrying attach-

ment, not

thicker than

main plate K-

butt weld, toe

ground Two-

sided fillets, toe

ground Fillet

weld(s), as

welded thicker

than main plate

100

100

80

71

36

36

28

25

Grinding parallel to stress

At one sided fillet welds, an

angular misalignment

corresponding to km = 1.2

is already covered

46

Transverse

stiffener welded

on girder web

or flange, not

thicker than

main plate. K-

butt weld, toe

ground Two-

sided fillets, toe

ground fillet

weld(s): as

welded thicker

than main plate

100

100

80

71

36

36

28

25

For weld ends on web

principle stress to be used

47

Non-

loadcarrying

stud as welded

80 28

48

Trapezoidal

stiffener to

deck plate, full

penetration butt

weld,

71 25


118 Of 150

calculated on

basis of

stiffener

thickness, out

of plane

bending

49

Trapezoidal

stiffener to

deck plate, fillet

or partial

penetration

weld, out of

plane bending

50 16

Calculation on basis of

stiffener thickness and weld

throat, whichever is smaller

50

Longitudinal

fillet welded

gusset at

length l l < 50

mm l < 150 mm

l < 300 mm l >

300 mm

80

71

63

50

28

25

20

18

For gusset near edge: see

525 "flat side gusset" If

attachement thickness <

1/2 of base plat thickness,

then one step higher

allowed (not for welded on

profiles!)

51

Longitudinal

fillet welded

gusset with

radius

transition, end

of fillet weld

reinforced and

ground, c < 2 t,

max 25 mm r >

150 mm

90 32 t = thickness of attachment

52

Longitudinal

fillet welded

gusset with

smooth

transition

(sniped end or

radius) welded

on beam flange

or plate. c < 2 t,

max 25 mm r >

0.5 h r < 0.5 h

or n < 20°

71

63

25

20

t = thickness of attachment

If attachement thickness <




profiles!)


119 Of 150

53

Longitudinal

flat side gusset

welded on

plate edge or

beam flange

edge, with

smooth

transition

(sniped end or

radius). c < 2t2,

max. 25 mm r >

0.5 h r < 0.5 h

or n < 20°

50

45

18

16


For t2 < 0.7 t1, FAT rises

12%

54

Longitudinal

fillet welded

gusset with

radius

transition, end

of fillet weld

reinforced and

ground, c < 2 t,

max 25 mm r >

150 mm

90 32 t = thickness of attachment

55

Longitudinal

fillet welded

gusset with

smooth

transition

(sniped end or

radius) welded

on beam flange

or plate. c < 2 t,

max 25 mm r >

0.5 h r < 0.5 h

or n < 20°

71

63

25

20


If attachement thickness <




profiles!)

56

Longitudinal

flat side gusset

welded on

plate edge or

beam flange

edge, with

smooth

transition

(sniped end or

radius). c < 2t2,

max. 25 mm r >

50

45

18

16


For t2 < 0.7 t1, FAT rises

12%


120 Of 150

0.5 h r < 0.5 h

or n < 20°

57

Longitudinally

loaded lap joint

with side fillet

welds Fatigue

of parent metal

Fatigue of weld

(calc. on max.

weld length of

40 times the

throat of the

weld

50

50

18

18

Weld terminations more

than 10 mm from plate

edge. Buckling avoided by

loadin or design!

58

Lap joint

gusset, fillet

welded, non-

load-carrying,

with smooth

transition

(sniped end

with n<20° or

radius), welded

to loaded

element c<2At,

but c <= 25 mm

to flat bar to

bulb section to

angle section

63

56

50

22

20

18

t = thickness of gusset

plate

59

Transverse

loaded overlap

joint with fillet

welds. Stress in

plate at weld

toe (toe crack)

Stress in weld

throat (root

crack)

63

36

22

12

Stresses to be calculated

using a plate width

equalling the weld length.

For stress in plate,

excenticity to be

considered, as given in

chapters 3.8.2 and 6.3.

Both failure modes have to

be assessed separately.

60

End of long

doubling plate

on I-beam,

welded ends

(based on

stress range in

56

50

45

20

18

16

End zones of single or

multiple welded cover

plates, with or without

frontal welds. If the cover

plate is wider than the

flange, a frontal weld is


121 Of 150

flange at weld

toe) tD # 0.8 t

0.8 t < tD # 1.5

t tD > 1.5 t

needed. No undercut at

frontal welds!

61

End of long

doubling plate

on beam,

reinforced

welded ends

ground (based

on stress range

in flange at

weld toe) tD #

0.8 t 0.8 t < tD

# 1.5 t tD > 1.5

t

71

63

56

28

25

22

Grinding parallel to stress

direction.

62

End of

reinforcement

plate on rectan-

gular hollow

section. wall

thickness: t <

25 mm

50 20 No undercut at frontal weld!

63

Reinforcements

welded on with

fillet welds, toe

ground Toe as

welded

80

71

32

25

Grinding in direction of

stress! Analysis based on

modified nominal stress,

however, structural stress

approach recommended.

64

Stiff block

flange, full

penetration

weld

71 25

65

Stiff block

flange, partial

penetration or

fillet weld toe

crack in plate

root crack in

weld throat

63

36

22

12


122 Of 150

66

Flat flange with

> 80% full

penetration butt

welds, modified

nominal stress

in pipe, toe

crack

71 25

67

Flat flange with

fillet welds,

modified

nominal stress

in pipe, toe

crack.

63 22

68

Tubular branch

or pipe

penetrating a

plate, K-butt

welds.

80 28

If diameter > 50 mm, stress

concentration of cutout has

to be considered

Assessment by structural

hot spot is recommended.

69

Tubular branch

or pipe

penetrating a

plate, fillet

welds. Toe

cracks. Root

cracks

(analysis based

on stress in

weld throat)

71

36

25

12



to be considered



70

Nozzle welded

on plate, root

pass removed

by drilling.

71 25



to be considered



71

Nozzle welded

on pipe, root

pass as

welded.

63 22



to be considered



72

Circular hollow

section butt

joint to massive

bar, as welded

63 22

Root of weld has to

penetrate into the massive

bar in order to avoid a gap

perpenticular to the stress


123 Of 150

direction.

73

Circular hollow

section welded

to component

with single side

butt weld, bac-

king provided.

Root crack.

63 22

Root of weld has to

penetrate into the backing

area in order to avoid a gap

perpenticular to the stress

direction.

74

Circular hollow

section welded

to component

single sided

butt weld or

double fillet

welds. Root

crack.

50 18

Impairment of inspection of

root cracks by NDT may be

compensated by adequate

safety considerations (see

chapter 5) or by

downgrading down to 2

FAT classes.

75

Circular hollow

section with

welded on disk

K-butt weld, toe

ground Fillet

weld, toe

ground Fillet

welds, as

welded

90

90

71

32

32

25

Non load-carrying weld.

76

Tube-plate

joint, tubes

flattened, butt

weld (X-

groove) Tube

diameter < 200

mm and plate

thickness < 20

mm

71 25

77

Tube-plate

joint, tube

slitted and wel-

ded to plate

tube diameter <

200 mm and

plate thickness

63

45

18

14


124 Of 150

< 20 mm tube

diameter > 200

mm or plate

thickness > 20

mm

4.4 Design against brittle fracture

Objective:

The students will be acquainted with the brittle fracture analysis based on linear elastic

fracture mechanics.

Scope:

Fracture toughness

Critical stress intensity

Critical crack size Temperature and material toughness

Overview of calculation methods in a relevant design guidance document, e.g., EN

1993 Eurocode 3-part 1-10: Design of Steel Structures: Selection of materials for

fracture toughness and through thickness properties

Expected result at comprehensive level:

Review theory of fracture mechanics and brittle fracture.

Explain relationship between material fracture toughness and temperature.

Review calculation procedures in a relevant design guidance document.

Compute critical crack size for structural element with typical material properties.

Compute stress intensity factor for a welded connection.

4.4.1. General Toughness of steel structures is treated exhaustively in design codes, thanks to a long

history of technical events due to factors that degrade materials after shorter or longer is

used. Toughness characterizes the behaviour of the steel structure damaged by mechanical

characteristics apparently not affecting them. For this reason the current presentation aims at

evaluating the behaviour of materials at the request of traction, bending the shock in different

ways, to understand the significance of defects in material harm to the intensity and tension

associated. This problem is highlighted in the context of variable demands, which promotes

germination and growth of cracks.


125 Of 150

4.4.2. Mechanical behaviour under tensile loads The mechanical behavior of metals is described by their deformation and fracture

characteristics under applied tensile, compressive, or multiaxial stresses. Determination of

this mechanical behavior is influenced by several factors that include metalllurgicall material

variables, test methods, and the nature of the applied stresses. In welded joinings, external

loads are distributed over structurally heterogeneous areas, which include residual tensions.

The basic data on the mechanical properties of a material are obtained from a tension test, in

which a suitably designed specimen is subjected to increasing axialload until it fractures. The

main emphasis is on mechanical behavior during the engineering tension test, which is

wide1y used to provide basic design information on the strength of materials and as an

acceptance test for the specification of materials. In this test procedure, a specimen is

subjected to a continualIy increasing uniaxialload (force), while simultaneous observations

are made of the elongation of the specimen. An engineering stress-strain curve is

constructed from load-elongation measurements (Figure 57). The stress used in this stress-

strain curve is the average longitudinal stress in the specimen, obtained by dividing the load

P, by the original specimen cross section area, Ao.

Figure 57 Comparison of engineering and true stree-strain curve

For homogen material, the strain in the engineering stress-strain curve is the average linear

strain, obtained by dividing the elongation of the specimen gauge length, δ, by its original

length, Lo. Since both the stress and the strain are obtained by dividing the load and

elongation by constant factors, the load-elongation curve will have the same shape as the

engineering stress-strain curve. The two curves are frequently used interchangeably. The

shape of the curve and magnitudes of stress and strain of the material will depend on its

composition, heat treatment, prior history of plastic deformation, and the strain rate,

temperature, and state of stress imposed during testing. The basic parameters used to

describe the stress-strain curve of a metal are the tensile strength, yield strength or yield


126 Of 150

point, percent elongation, and reduction of area. The first two are strength parameters; the

last two indicate ductility.

For welded joints are determined only resistance characteristics. In addition it is possible to

analyze the location and appearance of fracture section.

The general shape of the engineering stress-strain curve requires further explanation.

In the elastic region stress is linearly proportional to strain. When the load exceeds a value

corresponding to the yield strength, the specimen undergoes gross plastic deformation. It is

permanent1y deformed if the load is re1eased to zero. The stress producing continued

plastic deformation increases with increasing plastic strain, i.e., the metal strain-hardens. The

volume of the specimen remains constant during plastic deformation, AL = AaLo, and as the

specimen elongates, it decreases uniformly along the gauge length in cross-section area.

Initially, strain hardening more than compensates for this decrease in area and the

engineering stress (proportional to load P) continues to rise with increasing strain. Eventually

a point is reached where the decrease in specimen cross-sectional area is gre ater than the

increase in deformation load, arising from strain hardening. This condition wilI be reached

first at some point in the specimen that is slightly weaker than the rest. All further plastic

deformation is concentrated in this region, and the specimen begins to neck or thin down

local.

Ductility is a qualitative, subjective property of a material. In general, measurements of

ductility are of interest in two ways:

1. To indicate the extent to which a metal can be deformed without fracture in

metalworking operations such as rolling and extrusion.

2. An indication to the designer, in a general way, of the ability of the metal to flow

plastically before fracture. A high ductility indicates that the material is likely to deform

locally without fracture.

The slope of the initial linear portion of the stress-strain curve is the modulus of elasticity, or

the Young' s modulus. The modulus of elasticity is a measure of stiffness of the material, for

computing deflections of beams and other members. However, an increase in temperature

decreases the modulus of elasticity.

The ability of a material to absorb energy when deformed elasticalIy and to retuffi it when

unloaded is called resilience. This is usually measured by the modulus of resilience, which is

the strain energy per unit volume required to stress the material from zero stress to yield

stress So.

The toughness of a material is its ability to absorb energy in the plastic range. The ability to

withstand occasional stresses above yield stress without fracturing is particularly desirable in

parts such are freight-car couplings, gears, chains, and crane hooks. Toughness is a

commonly used concept which is difficult to pin down and define. One way of looking at

toughness is to consider that it is the total area under the stresssstrain curve.

The engineering stress-strain curve for homogen materials, does not give a true indication of

deformation characteristics of a metal because it is based entirely on original dimensions of

the specimen, and these dimensions change continuously during the test. Also, ductile metal


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pulled in tension becomes unstable and necks down during the course of the test. Since the

cross-section area of the specimen decreases rapidly at this stage in the test, the load

required to continue deformation falls off. The average stress based on original are a

decreases likewise, and this produces the fall-off in the stress-strain curve beyond the point

of maximum load. Actually, the metal continues to strain-harden all the way up to fracture, so

that stress required to produce further deformation should also increase. If true stress is

used, based on the actual cross-section area of the specimen, it is found that the stress-

strain curve increases continuously up to fracture. If the strain is measured also

instantaneously, the curve which is obtained is known as a true-stress-true-strain curve, or a

flow curve. Any point on the flow curve can be considered as the yield stress for a metal

strained in tension by the amount shown off the curve. Thus, if the load is removed at this

point and then reapplied, the material will behave elastically throughout the entire range of

reloading.

4.4.3. Impact testing Three basic factors contribute to a brittle-cleavage type of fracture:

a triaxial state of stress,

a low temperature,

a high strain rate or rapid rate of loading.

All three of these factors do not have to be present at the same time to produce brittle

fracture. A triaxial state of stress, such existing at a notch, and low temperature are

responsible for most service failures of the brittle type. However, since these effects are

accentuated at a high rate of loading, many types of impact tests have been used to

determine the susceptibility of materials to brittle behaviour. Steels which have identical

properties when tested in tension or tors ion at slow strain rates can show pronounced

differences in their tendency for brittle frac ture when tested in a notched-impact test.

The situation becomes more complex when heterogeneous metallurgical materials, as if

welded joints. Since the ship failures occurred primarily in structures of welded construction,

it was considered for a time that this method of fabrication was not suitable for service where

brittle fracture might be encountered. A great deal of research has since demonstrated that

welding, per se, is not inferior in this respect to other types of construction. However, strict

quality control is needed to prevent weld defects which can act as stress raisers or notches.

New electrodes have been developed for a weld with better properties than the mild-steel

plate. The design of a welded structure is more critical than the design of an equivalent

riveted structure. It is important to eliminate stress raisers and reduce rigidity.

A. Notched-bar impact tests

Various types of notched-bar impact tests are used to determine the tendency of a material

to behave in a brittle manner. The results obtained from notched-bar tests are not convenient

for design, since it is not possible to measure the components of the triaxial stress condition

at the notch. Furthermore, there is no general agreement on the interpretation or significance

of results obtained with this type of test. Nowadays, Charpy specimen is supported as a


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beam in a horizontal position and loaded behind the notch by the impact of a heavy swinging

pendulum (Figure 58.) with the high impact velocity.

Figure 58 Dimensions of Charpy V notch standard specimen

Plastic constraint at the notch produces a triaxial state of stress. The maximum plastic stress

concentration is given by:

Kζ = [1 + (π/2) – (ω/2)]

(4.4.1)

where ω is the included flank angle of the notch.

The relative values of the three principal stresses depend strongly on the dimensions of the

bar and the details of the notch. The standard Charpy V specimen is thick enough to ensure

a high degree of planc-strain and triaxiality across almost all of the notched cross section,

and provides a severe condition for brittle fracture. Therefore, nonstandard specimens

should be used with great care. The principal measurement from the impact test is the

energy absorbed in fracturing the specimen. After breaking the test bar, the pendulum

rebounds to a height which decreases as the energy absorbed in fracture increases.

The energy absorbed for fracture, in joules (J), of ten designated KV, is read directly from a

calibrated dial on the impact tester. Sometimes impact test results are expressed ut energy

absorbed per unit cross-sectional area of the specimen (notch or impact toughness).

Fracture energy measured by the Charpy test is only a relative energy and cannot be use,

direct1y in design equations. Another common result obtained from the Charpy test is based

on examination of the fracture surface. The fracture is fibrous (shear fracture) granular

(cleavage fracture), or a mixture of both. These different modes of failure are readily

distinguishable even without magnification. The flat facets of cleavage fracture provide a high

reflectivity and bright appearance, while the dimpled surface of a ductile fibrous fracture

provides a lightabsorptive surface and dull appearance. Usually an estimate is made of the

percentage of the fracture surface that is cleavage (or fibrous) fracture.


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Figure 59 shows how the fracture appearance of steel changes from 100 percent flat

cleavage (Ieft) to 100 percent fibrous fracture (right) as the test temperature is increased.

The fibrous fracture appears first around the outer surface of the specimen (shear lip) where

the triaxial constraint is the least.

Gradual decrease in the granular region and increase in lateral contraction at the notch with

increasing temperature is visible. Sometimes in the Charpy test the ductility is measured as

indicated by the percent contraction of the specimen at the notch.

The notched-bar impact test is most meaningful when conducted over a range of

temperature so that the temperature for ductile-to-brittle transition can be determined (Figure

60). The energy absorbed decreases with decreasing temperature but for most cases the

decrease is not sharp at a certain temperature, and it is difficult to determine accurately the

transition temperature. In selecting a material from the standpoint of impact toughness or

tendency to brittle failure, the important factor is the transition temperature. Steel A, on

Figure 60.a, shows higher impact toughness at room temperature; yet its transition

temperature is higher than that of steel B. The material with the lowest transition temperature

is to be preferred. Notched-bar impact tests are subject to considerable scatter, particularly in

the region of the transition temperature.

5 oC 38 oC 100 oC

Figure 59 Fracture surfaces of Charpy specimens of mild steel, tested at different temperatures


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Figure 60 Transition-temperature curves for a) two steels, b) transition temperature criterion

Most of this scatter is due to local variations in the properties of the steel, but also notch

shape and depth are critical variables, which can not be perfectly reproduced. Proper

placement of the specimen in the anvil is also important.

The principal advantage of the Charpy V -notch impact test is that can readily be carried out

over a range of subambient temperatures. Moreover, the design of the test specimen is well

suited for measuring differences in notch toughness in low-strength materials such as

structural steels. The test is used for comparing the influence of alloy studies and heat

treatment on notch toughness. It frequently is used for quality control and material

acceptance. By collecting and testing samples of welded joints areas, obtain information

about tenacity is located.

The major difficulty is that the results of the Charpy test are not directly applicable in design.

B. Instrumented Charpy test

The conventional Charpy test measures the total energy absorbed in fracturing the

specimen. Additional information can be obtained if the impact tester is instrumented to

provide a load-time history of the specimen during the test. It is possible to determine the

energy required for initiating fracture (crack) and the energy required for propagating

fracture. It also yields information ob the load for general yielding, the maximum load and the

fracture load - information very important for designer. The area under diagram is

proportional to absorbed energy.

For same absorbed energy of two different steels tested at different temperatures, the ratio

between energies for crack initiation and propagation can be different at low temperatures,

although is the same at room temperature. Such a behavior can help as an additional

criterion of material weldability. It can be noticed, without a deep analysis, that higher crack


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propogation energy is convenient for welded joint, having in mind that crack-like defects

cannot be excluded in welded structures.

Because the root of the notch in a Charpy specimen is not as sharp as in fracture mechanics

tests with precracked specimens, there has been a trend toward using standard Charpy

specimens which are precracked by the introduction of a fatigue crack at the tip of the V

notch. These precracked specimens have been used in the instrumented Charpy test to

measure dynamic fracture toughness values (KId).

The significance of impact testing is illustrated by test results performed for two high strength

steels (table 29, Figure 61).

Table 29 Chemical composition and tensile characteristics of tested steels

Steel Chemical composition (%) Tensile characteristics

C Si Mn Cr Ni Mo V Al Yield

strength

YS

[MPa]

Ultimate

tensile

strength

UTS

[MPa]

Elon-

gation

A

[%]

Reductin

of cross

section

area

Z [%]

A

B

0,1

0,3

0,27

0,28

0,35

0,73

1,11

2,05

2,65

1,87

0,26

0,30

0,1

-

0,05

-

780

940

825

1015

18,0

16,7

68,0

58,2

Figure 61 Instrumented impact test resu1ts obtained with Charpy V specimen for steels A and B

L-notch in cross-rolling direction; C-notch in rolling direction. 1-crack initiation energy, 2-crack

propagation energy, 3-total energy.


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The difference in strength and ductility of these steels is not expressed in the same level as

in the case with impact toughness properties. Steel A, with low carbon content, exhibited

high impact energy at low temperatures (down to - 100 °C) for crack propagation and also

crack initiation. However, there a significant effect of rolling direction in that case, which

should be taken into account. For steel B, with 0.3% C, the impact energy is low, and nil

ductility transition temperature can be determined (between -40°C and - 60°C).

Typical curve obtained on instrumented Charpy pendulum is presented in Figure 62. In this

Figure all oscillations of Charpy specimen during testing on instrumented pendulum

indicated.

Figure 62 Typical load vs. time record showing fracture phases of Charpy specimen

C. High rate impact test

Probably the chief deficiency of the Charpy impact test is that the small specimen is not

always a realistic model of the actual situation. Not only does the small specimen lead to

considerable scatter, but a specimen with thickness of 10 mm cannot provide the same

constraint as would be found in a structure with much greater thickness.

At a particular service temperature the standard Charpy specimen shows a high shelf

energy, while actually the same material in a thick-section structure has low toughness at the

same temperature. The most logical approach to this problem is the development of tests

that are capable of handling specimens of extended thickness (e.g. explosion bulge test,

drop weight test).

C1. Explosion bulge test

The basic need for large specimens resulted from the inability to produce fracture in small

laboratory specimens at stresses below gross yield stress, whereas brittle fractures in ship

structures occur at service temperatures at elastic stress levels, as experieneed with Liberty

ships.


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Die support (rig) with the base allows bulging of properly positioned test plate (specimen).

Cast explosive charge of specified mass and power should be applied at properly determined

distance, obtained by cardboard box over the test plate. High rate of explosion loading

contributes to brittle fracture of test plate. Tests can be carried out over a range of

temperatures and then the appearance of the fracture determines the transition temperatures

(Figure 63). Below the nil-ductility-transition (NDT) temperature the fracture is a flat (elastic)

fracture running completely to the edges of the test plate. Above the nil ductility temperaature

a plastic bulge forms in the center of the plate, but the fracture is still a flat elastic fracture out

to the plate edge. At a still higher temperature the fracture does not propagate outside of the

bulged region. The temperature at which elastic fracture no longer propagates to the edge of

the plate is called the fracture transition elastic (FTE). The FTE marks the highest

temperature of fracture propagation by purely elastic stresses. At yet higher temperature the

extensive plasticity results in a helmet-type bulge. The temperature above which this fully

ductile tearing occurs is the fracture transition plastic (FTP).

The explosion bulge test makes use of a large plate specimen that incorporates novel

features in its preparation and testing procedure. However, the application of explosion in the

test introduced inconveniences and a new loading type had been proposed.

As an illustration, the results of explosion bulge test with the plates (BM) of steels A aud B

(table 4.4.1) are presented in Figure 64. After each shot, the reduction of thickness ΔR and

bulge extension B were measured. Again the effect of rolling direction of stee1 A is

significant, and steel B exhibited linear behaviour.

Figure 63 Fracture appearance vs. temperature for explosion-crack-starter

test NDT - Nil Ductility Transition; FTE-Fracture Transition Elastic; FTP-Fracture Transition

Plastic


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Figure 64 Typical results of explosion bulge test for steels A and B

The graphs show reduction of thickness ΔR and bulge development B vs. number of

explosions L-notch in hard bead in cross rolling direction; C-notch in rolling direction.

C2. Drop weight test. Experience gathered with the explosion bulge test in NRL has led to

the development of drop-weight test (DWT), in order to avoid the explosion. The DWT energy

is obtained from potential energy of falling mass (weight). Due to significant weight of the tup

and height of device, more energy can be obtained compared to Charpy pendulum.

The drop weight test was developed specifically for the determination of the NDT

temperature on full thickness plates. The simplicity of the drop-weight specimen, the

apparatus for applying load and the interpretation of results, contributed to wide use of this

test. The stress applied to the specimen during the impact loading is limited to the yield point

by a stopping block attached to the anvil below the specimen (Figure 65). This is the practical

device for evaluating the ability of the steel to withstand yie1d point loading in the presence

of a small flaw.

Figure 65 Drop weight test conFigureation the anvil stop.


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C3. Robertson crack-arrest test

This method provides a relationnship between the stress level and the ability of the material

to arrest a rapidly propagating crack (Figure 66). A uniform elastic tensile stress is applied to

a plate specimen 150 mm wide. A rapid brittle fracture is initiated by impact loading at a

starter crack on the cold side of the specimen. The crack propagates up a temperature

gradient toward the hot side. The point across the specimen width at which the temperature

is high enough to give sufficient ductility to blunt the crack is called the crack-arrest

temperature (CAT). In a test alternative form, the temperature across the specimen is

constant and tests are carried out with successive specimens at increasing temperature until

the CAT is reached. Crack-arrest tests on mild steel below NDT show that the CAT is

independent of temperature but the stress level for crack arrest is very low. If the stress is

greater than 35 to 55 MPa, brittle fracture will occur. Obviously, this stress level is too low for

practical engineering design, so that steels cannot be used below the NDT. While crack-

arrest tests are among the most quantitative ofbrittle fracture tests, they are not used

extensively due to required large testing machines and large specimens.

C4. Fracture analysis diagram

Nil-ductility transition temperature as determined by the drop weight test is regarded the most

important reference point on the fracture analysis diagram because of the simplicity with

which it is determined, and because a steel is characterized by a single NDT. Fracture

analysis system introduces considerable promise for guiding engineering design and

selection of steel for fracture-safe weldments and structures. More detailed consideration is

necessary before use of transition points by the fracture analysis diagram, through reference

to basic properties of the tension test. The sub-ambient temperature dependences of yield

strength ζo and ultimate tensile strength ζu in a metal (Figure 69).

For an unnotched specimen, the material is ductile until a very low temperaature, point A,

where ζo= ζu. Point A represents the NDT temperature for a flaw-free material. The curve

BCD represents the fracture strength of a specimen with a small flaw (a<0.1 mm). The

temperature corresponding to point C is the highest temperature where the fracture strength

ζf » ζu. Point C represents the NDT for a specimen with a small crack or flaw. The presence

of a small flaw raises the NDT of a steel by about 90°C.

Increasing the flaw size decreases the fracture stress curve, as in curve EF, until with

increasing flaw size a limiting curve of fracture stress HJKL is reached. Below the NDT, the

limiting safe stress is 35 to 55 MPa. Above NDT the stress required for unstable propagation

of a long flaw (JKL) rises sharply with increasing temperature. This is the crack-arrest

temperature curve (CAT). The CAT defines the highest temperature at which unstable crack

propagation can occur at any stress level. Fracture will not occur for any point to the right of

the CAT curve. The temperature above which elastic stresses cannot propagate a crack is

the fracture transition elastic (FTE). This is defined by the tempeerature when the CAT curve

crosses the yield-strength curve (point K). The fracture transition plastic (FTP) is the

temperature where the CAT curve crosses the tensile estrength curve (point L). Above this

temperature the material behaves as if it was flaw-free, for any crack, no matter how large,

cannot propagate as an unstable fracture.


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Figure 66 Robertson crack-arrest test Figure 67 Temperature dependence of yield strength

(σo), tensile strength (σu)

Data obtained from the DWT and other large-scale fracture tests have been assembled into

a useful design procedure called the fracture analysis diagram (FAD). The NDT as

determined by the DWT provides a key data point to start construction of the fracture

analysis diagram and transition temperature features of steel (Figure 4.4.12).

Figure 68 Fracture-analysis diagram showing influence of various initial flaw sizes

For mild steel below NDT the CAT curve is flat. A stress level in excess of 35 to 55 MPa

causes brittle fracture, regardless of the size of the initial flaw. Extensive correlation between

NDT and Robertson CAT tests for a variety of structural steels has shown that the CAT curve

bears a fixed relationship to the NDT temperature. Thus, the NDT -1°C provides a

conservative estimate of the CAT curve at stress of ζo /2, NDT + 15°C provides an estimate

ofthe CAT at ζ = ζo, and the FTE and NDT +50°C provides an estimate of the FTP. So, once

NDT for structural steels is determined, the entire scope of the CAT curve can be established

well enough for engineering design. The curve traced out represents the worse possible case


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for large flaws in excess of 600 mrn. One can imagine a spectrum of curves translated

upward and to the left for smaller, less severe flaws. Correlation with service failures and

other tests has allowed the approximate determination of curves for a row of initial flaw sizes.

Thus, the FAD provides a generalized relationship of flaw size, stress, and temperaature for

low-carbon structural steels of the type used in ship construction. The fracture analysis

diagram can be used in several ways for design (Figure 69). One simple approach would be

to use the FAD to select a steel which had an FTE that was lower than the lowest expected

service temperature. With this criterion the worst expected flaw would not propagate so long

as the stress remained elastic. Since the assumption of elastic behaviour is basic in

structural design, this design philosophy would be tantamount to being able ta ignore the

presence of flaws and brittle fracture. However, this procedure may prove to be too

expensive and overconservative. A slightly less conservative design against brittle fracture,

but still a practical approach, would be to design on the basis of an allowable stress level not

exceeding ζo/2. From Figure is visible that any crack will not propagate under this stress so

long as the temperature is not below NDT -1°C.

The dynamic tear test (DT) can also be used to construct the FAD (Figure 70), using NDT as

base (dashed line). Below the NDT the fracture is brittle and has aflat, featureless surface

devoid of any shear lips. At temperatures above NDT there is a sharp rise in energy for

fracture and the fracture surfaces begin to develop shear lips. The shear lips become

progressively more prominent as the temperature is increased to the FTE. Above FTE the

fracture is ductile, void coalescence-type fracture. The fracture surface is a fibrous slant

fracture. The upper shelf of energy represents the FTP. The lower half of the DT energy

curve traces the temperature course of the CAT curve from NDT to FTE.

Figure 69 Fracture analysis diagram for steel A Figure 70 Application of DT test result for fracture analysis


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4.4.3. Fatigue testing

Fatigue is the progressive, localized, and permanent structural damage that occurs when a

material is subjected to cyclic or fluctuating strains at nominal stresses that have maximum

values less than (and of ten much less than) the static yield strength of the material. This

process of fatigue failure can be divided into different stages, which, from the standpoint of

metallurgical processes, can be divided into five stages, defined by the characterization of

the underlying fatigue damage of a material (Figure 71):

cyclic plastic deformation prior to fatigue crack initiation,

initiation of one or more microcracks,

propagation or coalescence of microcracks to form one or more microcracks,

propagation of one or more macrocracks,

final failure.

Figure 71 Different phases of fatigue life and relevant factors

It also clearly defines the requirement of plastic deformation for the onset of crack initiation.

In general, three simultaneous conditions are required for the occurrence of fatigue damage:

cyclic stress,

tensile stress,

plastic strain.

If any one of these three conditions is not present, a fatigue crack will not initiate and

propagate. The plastic strain resulting from cyclic stress initiates the crack, and the tensile

stress (which may be localized tensile stresses caused by compressive loads) promotes

crack propagation.

The stages of fatigue can also be defined in more general terms from the perspective of

mechanical behavior of crack growth. For example, another division of the fatigue process is

defined as follows:


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nucleation (initiation of fatigue cracks),

structurally dependent crack growth rates (often called the "short crack" or "small

crack" phase),

crack growth rates that can be characterized by either linear elastic fracture

mechanics, elastic-plastic fracture mechanics, or fully plastic fracture mechanics,

final instability.

This definition of the stages in the fatigue process is roughly equivalent to the first, except

that crack propagation is expressed in terms of crack growth rates, and nucleation is meant

to include all processes leading up to crack initiation. In general, the fatigue process consists

of a crack initiation and a crack propagation phase. There is, however, no general agreement

when (or at what crack size) the crack initiation process ends, and when the crack growth

process begins. Nonetheless, the separation of the fatigue process into initiation and

propagation phases has been an important and useful advance in engineering. Another

important engineering advance is the transfer of the multistage fatigue process from the field

to the laboratory. In order to study, explain, and qualify component designs, or to conduct

failure analyses, a key engineering step is often the simulation of the problem in the

laboratory. Any simulation is, of course, a compromise of what is practical to quantify, but the

study of the multistage fatigue process has been greatly advanced by the combined methods

of strain-control testing and the development fracture mechanics of fatigue crack growth

rates. This combined approach (Figure 72) is a key advance that allows better understanding

and simulation of both crack nucleation in regions of localized strain and the subsequent

crack growth mechanisms outside the plastic zone. This integration of fatigue and fracture

mechanics has had important implications in many industrial applications for mechanical and

materials engineering.

Figure 72 Laboratory simulation of the multistage fatigue process.

The three basic types of fatigue properties, are in table 30 presented.


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Table 30 The three basic types of fatigue properties

Design philosophy Design methodology Principal testing data desctription

Safe-life, infinite-life Stress-life S-N

Safe-life, finite-life Strain-life ε-N

Damage tolerant Fracture mechanics da/dN-ΔK

The S-N and ε-N techniques are usually appropriate for situations where a component or

structure can be considered a continuum (i.e., those meeting the "no cracks" assummption).

In the event of a crack-like discontinuity, the S-N or ε-N methods (except through residuallife

testing) offer little or quantitative basis for assessment of fatigue life.

Another limitation of the S-N and ε -N methods is the inability of the controlling quantities to

make sense of the presence of a crack. A brief review of basic elasticity calculations shows

that both stress and strain become astronomical at a discontinuity such as a crack, far

exceeding any recognized property levels that might offer some sort of limitation. Even

invoking plasticity still leaves inordinately large numbers or, conversely, extremely low

tolerable loads.

The solution to this situation is the characterization and quantification of the stress field at the

crack tip in terms of stress intensity in linear elastic fracture mechanics. It recognizes the

singularity of stress at the tip and provides a tractable controlling quantity and measurable

material property. The use of the stress intensity as a controlling quantity for crack extension

under cyclic loading thus enhances the engineering analysis of the fatigue process.

4.4.4. Fracture mechanics approach

A. General

The concepts of fracture mechanics are basic ideas for developing methods of predicting the

load-carrying capabilities of structures and components containing cracks. Though virtualIy

alI design and standard specifications require the definition of tensile properties for a

material, these data are only partly indicative of inherent mechanical resistance to failure in

service. Except for those situations where gross yielding or highly ductile fracture represents

limiting failure conditions, tensile strength and yield strength are often insufficient

requirements for design of failure-resistant structures. Brittle fracture can also occur if

toughness, resistance to corosion, stress corosion, or fatigue resistance is reduced too much

in achieving high strength.

Fracture of structural and equipment components as a result of cyclic loading has long been

a major design problem and the subject of numerous investigations. Although a considerable

amount of fatigue data are available, the majority have been concemed with the nominal

stress required to cause failure in a given number of cycles-namely, S-N curves. Usually,

such data are obtained by testing smooth specimens which, although of some qualitative use

for guiding material selection, are subject to limitations caused primarily by the failure to


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adequately distinguish between fatigue-crack-initiation life and fatigue-crack-propagation life

(Figure 73).

The existence of surface irregularities and crack type imperfections reduces and may

eliminate the crack-initiation portion of the fatigue life of the component. Fracture emechanics

methodology offers considerable promise for improved understanding of the initiation and

propagation of fatigue cracks and problem resolution in designing to prevent failures by

fatigue.

Figure 73 Different scenarios for fatigue crack growth

Figure 74 Schematic illustration of variation of fatigue-crack -growth rate, da/dN, with alternating stress intensity, ΔK, in steels, showing regions of primary crack-growth mechanisms.


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Initiation of fatigue cracks in structural and equipment components occurs in regions of

stress concentrations, such as notches, as a result of gradients and fluctuation stresses. The

material element at the tip of a notch in a cyclically loaded component is subjected to the

maximum stress range, Δζmax.

Consequently, this material element is most susceptible to fatigue damage and is, in general,

the origin of fatigue-crack initiation. It can be shown that, for sharp notches, the maximum-

stress range on this element can be related to the stress intensity-factor range, ΔKI, as

follows:

Δζmax =(2/ √ π). (ΔKI / √ρ ) = ρζ(kt) (4.4.2)

where ρ is the notch-tip radius, Δζ is the range of applied nominal stress, and kt is stress-

concentration factor.

Extensive data have shown that the fatigue-crack-propagation behaviour of metals is

controlled primarily by the stress intensity factor range, ΔKI. The fatigue-crack propagation

behaviour of metals can be divided into regions (Figure 74):

- the behaviour in region I exhibits a fatigue-crack propagation threshold, ΔKth, wich

corresponds to the stress-intensity-factor range, below which cracks do not propagate under

cyclic-stress fluctuations.

- the behaviour in region II represents the fatigue-crack-propagation behaviour above ΔKth,

which can be represented by the power-law relationship:

da/dN = A (ΔKth)n (4.4.3)

Extensive fatigue crack growth rate data for various steels show that the primary parameter

affecting growth rate in region II is the stress intensity factor range, and that the mechanical

and metallurgical properties of these steels have negligible effects on the fatigue crack

growth rate in a room temperature air environment. The stress ratio and mean stress have

negligible effects on the rate of crack growth in region II. Also, the frequency of cyclic loading

and the wave form (sinusoidal, triangular, square, or trapezoidal) do not affect the rate of

crack propagation per cycle of load for steels in benign environments. The acceleration of

fatigue-crack-growth rates that determines the transition from region II to region III appears to

be caused by the superposition of a brittle or a ductile-tearing mechanism onto the

mechanism of cyclic subcritical crack extension, which leaves fatigue striations on the

fracture surface. These mechanisms occur when the strain at the tip of the crack reaches a

critical value. Thus, the fatigue-rate transition from region II to region III depends on the

maximum stress-intensity factor, on the stress ratio, and on the fracture properties ofthe

material.


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The fracture mechanics approach is based on a mathematical description of the

characteristic stress field that surrounds any crack in a loaded body. When the region of

plastic deformation around a crack is small compared to the size of the crack (as is of ten

true for large structures and high-strength materials), the magnitude of the stress field around

the crack is related to the stress-intensity factor, K, with:

K=ζ(√a).Y(a/W) (4.4.4)

where: ζ is remotely applied stress, a - characteristic flaw size dimension, Y - geometry

factor that depends on the ratio of the crack length a, to the width W, determined from linear

elastic stress analysis.

The stress-intensity factor K, thus represents a single parameter that includes both the effect

of the stress applied to a sample and the effect of a crack of a given size in a sample. The

stress-intensity factor can have a simple relation to applied stress and crack length, or the

relation can involve complex geometry fac tors for complex loading, various conFigureations

of real structural components, or variations in crack shapes.

These concepts provide a basis for defining a critical stress-intensity factor (Kc) for the onset

of crack growth, as a material property independent of specimen size and geometry for many

conditions of loading and environment. In general, when the specimen thickness and thein-

plane dimensions near the crack are large enough relative to the size of the plastic zone,

then the value of K at which growth begins is a constant and, generally, minimum value

called the plane-strain fracture toughness factor, KIc, of the material. The parameter KIc is a

true material property in the same sense as is the yield strength of a material. The value of

KIc determined for a given material is unaffected by specimen dimensions or type of loading,

provided that the specimen dimensions are large enough relative to the plastic zone to

ensure plane-strain conditions around the crack tip (strain is zero in the through-thickness or

z-direction). Therefore, plane-strain fracture toughness, KIc, is particularly pertinent in

materials selection because, unlike other measures of toughness, it is independent of

specimen configuration.

Whether the fracture is ductile or brittle does not directly influence the deformation process

that a component or specimen might undergo during the measurement of toughness. The

deformation process is generally described as being linear-elastic or nonlinear. This

determines which parameter is used in the fracture toughness test characterization. All

loading begins as linear-elastic. For this, the primary fracture parameter is the well-known

crack-tip stress-intensity factor, K.

If the toughness is relatively high, the loading may progress from linear-elastic to nonlinear

during the toughness measurement, and a nonlinear parameter is needed. The nonlinear

parameters that are most often used in toughness testing are the J-integral, labeled J, and

the crack tip opening displacement (CTOD), labeled δ. Because all loading starts as linear-

elastic, the nonlinear parameters are all written as a sum of a linear component and a

nonlinear component.


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The test methods covered include linear-elastic and nonlinear loading, slow and rapid

loading, crack initiation, and crack arrest. The development of the test methods followed a

chronological pattern; that is, a standard was written for a particular technology soon after

that technology was developed. Standards written in this manner tend to become exclusive

to a particular procedure or parameter. Because most fracture toughness tests use the same

specimens and procedures, this exclusive nature of each new standard did not allow much

flexibility in the determination of a fracture toughness value.

B. Linear-elastic fracture toughness testing

Fracture mechanics and fracture toughness testing began with a strictly linear-elastic

methodology using the crack-tip stress-intensity factor, K. The linear-elastic methods of frac

ture toughness testing are used to measure a single point fracture toughness value. For

fracture by a brittle mechanism, this is no problem. Fracture occurs at a distinct point, and

the fracture toughness measurement is taken as a value of the fracture parameter at that

point. For fracture by a ductile mechanism, the fracture is a process, and the fracture

toughness measurement is an R-curve. To get a single value for this fracture toughness, a

point on the R-curve must be chosen. This usually involves a construction procedure.

The first fracture toughness test that was written as a standard was the KIc test method,

ASTM E 399 or BS 5447. This test measures fracture toughness that develops under

predominantly linear-elastic loading with the crack-tip region subjected to near strain

constraint conditions through the thickness. The test was developed for essentially ductile

fracture conditions, but can also be used for brittle fracture. As a ductile fracture test, a single

point to define the fracture toughness is desired. To accomplish this, a point where the

ductile crack extension equals 2% of the original crack length is identified. This criterion is

illustrated schematically with a K based R curve (Figure 75). This criterion gives a somewhat

size-dependent measurement, and so validity criteria are chosen to minimize the size effects

as well as to restrict the loading to essentialIy the linear-elastic regime. In this way, the KIc

test can serve as a model for the other discussions.

The first element of the test is the selection of a test specimen. Five different specimen

geometries are recommended. These are the single edge-notched bend specimen, SE(B),

compact tension specimen, C(T), arc-shaped specimen, A(T), disk-shaped compact

specimen, DC(T), and the arc-shaped bend specimen, A(B). Many of these specimen

geometries are used in the other standards as well. The SE(B) and C(T) specimens are

traditional fracture toughness specimens used in every fracture toughness test method. The

other three are special geometries that represent structural component forms. Most fracture

toughness tests are conducted with either the edge-notched bend or compact specimens.

The choice between the bend and compact tension specimen is based on:


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Figure 75 Schematic of K-based crack resistance, R curve with position of KIc

the amount of material available (the bend takes more),

machining capabilities (the compact has more detail and costs more to machine),

the loading equipment available for testing.

Specimens for the KIc test must be precracked in fatigue before testing. This means that a

sharp crack is developed at the end of a notch by repeated loading and unloading of the

specimen, that is, fatigue loading.

C. Nonlinear fracture toughness testing

Linear-elastic parameters are used to measure fracture toughness for relatively low

toughness materials, which fracture under or near the linear-loading portion of the test. For

many materials used in structures, it is desirable to have high toughness, a value at least

high enough so that the structure would not reach fracture toughness before signifiicant

yielding occurs. For these materials, it is necessary to use the nonlinear fracture parameters

to measure fracture toughness properties. The two leading nonlinear fracture parameters are

J and δ. For many of the nonlinear fracture toughness measurements, the fracture mode is a

ductile one. In this case the fracture toughness is measured by an R curve, that is, a plot of

the fracture-characterizing parameter as a function of the ductile crack advance. The

evaluation of R -curve toughness requires three measurements during the test: load,

displacement, and crack length. In the standards, the crack length has been measured

visually and the fracture surface by an elastic unloading compliance method that uses the

elastic properties of the specimen geometry to evaluate crack length. Methods that have also

been used are an electric al potential drop method and a key curve, or normalization,

method. The measured values are:


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crack-tip opening displacement (CTOD) δ: the crack displacement due to elastic and

plastic deformation at variously defined locations near the original (prior to an

application of load) crack tip.

J-integral, J: a mathematical expression, a line or surface integral that encloses the

crack front from one crack surface to the other, used to characterize the local stress-

strain field around the crack front.

4.4.5. New standards for fracture mechanics testing of metallic materials The development of standard fracture toughness test methods is present a permanent

problem. They are the common test method, a new fracture toughness standard that

combines most of the standard test methods into a single standard, and the transition

fracture toughness standard. Corresponding standard for testing of weldments is still being

developed.

a) Common fracture toughness test method. Because the JJc and J-R curve test

standards are similar in many respects, they have been combined into a single test

standard, ASTM E 1820. This standard method also allows a measurementof fracture

toughness using the linear elastic parameter, K, and the nonlinear parameters, J and

δ. The idea of a common method is that most of the fracture toughness tests use the

same specimens, instrumentation, and test procedures. The way individual methods

were written in the past allows for the likelihood that a test can produce an invalid or

unqualified result with no way to use the analysis procedure of another test method to

try to obtain an acceptable result. The common method combines all measurements

of fracture toughness into a single standard instead of many specialized standards.

Therefore, after the test has been completed, the behavior of the material can dictate

the nature of the analysis used, and a satisfactory fracture toughness result can be

achieved for most tests. The analysis can use a linear elastic or an elastic-plastic

parameter; it can use a single point fracture measurement or an R-curve toughness

measurement.

b) Transition fracture toughness testing for ferritic steels has long been a problem. The

fracture behavior is usualIy brittle sometimes after an initial period of ductile crack

extension. The toughness values show extensive scatter and size dependency that

cause difficulty in the characterization of toughness for the evaluation of structures.

The scatter and size dependency has been attributed to statistical influences and

constraint differences. Characterization of the toughness relies mainly on the

statistical handling of the data. Test method ASTM E 1921 has been deve1oped

recently to handle the problems of transition fracture toughness testing. The

specimens, fixtures, instrumentation, test procedures, and calculation of toughness

parameters folIow existing standards, for example, ASTM E 1820. The evaluations of

the statistical aspects are handled with weakest-link weibull statistical distribution.

In final, without mechanical testing, there is no true structural integrity. In many procedures of

structural integrity assessment, precise evaluation of mechanical properties, especially

tensile properties, is of crucial importance. This is even more pronounced in the case of

weldment integrity assessment because of its heterogeneity.


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Although the conservative approach can accept the lowest strength as the relevant

parameter for weldment integrity assessment (e.g. if FAD is applied), in the case of more

precise analysis it is necessary to evaluate strength properties for alI weldment regions with

different microstructure, either directly, using tensile microspecimens, or indirectly, e.g. by

measuring the microhardness. One should notice, that even more sophisticated integrity

procedures like J-R vs Crack Driving Force analysis, strongly depend on tensile properties,

i.e. flow strength.

Toughness is often used in structural integrity assessment procedures, e.g. in Fracture

Analysis Diagrammes (FAD), at least indirectly, as the nil ductility temperature (NDT). It is

well known that NDT temperature is often significantly different for different microstructures in

a weldment, indicating once again the need for precise evaluation of a mechanical property

in order to get the reliable structural integrity estimate of a material and weldment.

Present damage tolerant design philosophy doesn't address issue about flaw existence.

Moreover, it claims that there is no component and structure without flaws, which shift

fracture mechanic design and testing methodology on a new level, lead to real and

comprehensive structural integrity.


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List of figures Figure 1 Classification of electric arc welded joints _______________________________________________________________ 8 Figure 2 Components of single sided joint, double sided joint, respectively _____________________________________14 Figure 3 Types of welded joints ___________________________________________________________________________________17 Figure 4 Types of welded joints on technological equipment ____________________________________________________22 Figure 5 The overview of engineering properties of materials. __________________________________________________24 Figure 6 Setting out lines __________________________________________________________________________________________28 Figure 7 Change of the bearing capacity of weld with defect area. _____________________________________________31 Figure 8 Stresses on the throat section of a fillet weld ___________________________________________________________33 Figure 9 Effective penetration of T-butt welds. ___________________________________________________________________36 Figure 10 Calculation of the weld forces for intermittent welds _________________________________________________37 Figure 11 Throat thickness of a fillet weld. _______________________________________________________________________38 Figure 12 Throat thickness of a deep penetration fillet weld. ___________________________________________________39 Figure 13 Effective throat thickness of flare groove welds in solid sections. ____________________________________39 Figure 14 Effective width of an unstiffened T – joint _____________________________________________________________40 Figure 15 Local eccentricity _______________________________________________________________________________________41 Figure 16 Modified or local nominal stress _______________________________________________________________________44 Figure 17 Notch stress and structural stress _____________________________________________________________________44 Figure 18 Geometric elements of intermittent fillet weld ________________________________________________________47 Figure 19 The stress distribution over the plate thickness. ______________________________________________________50 Figure 20 Nominal stress in a beam-like component _____________________________________________________________51 Figure 21 Examples of macrogeometric effects __________________________________________________________________51 Figure 22 Modified (local) nominal stress near concentrated loads ____________________________________________52 Figure 23 Axial and angular misalignment _______________________________________________________________________52 Figure 24 Structural details and structural stress _______________________________________________________________53 Figure 25 Definition of structural hot spot stress ________________________________________________________________54 Figure 26 Various locations of crack propagation in welded joints _____________________________________________54 Figure 27 Biaxial stress at weld toe _______________________________________________________________________________55 Figure 28 Types of hot spots _______________________________________________________________________________________56 Figure 29 Typical meshes and stress evaluation path for a welded detail ______________________________________57 Figure 30 Reference points at different types of meshing ________________________________________________________58 Figure 31 Examples of strain gauges in plate structures ________________________________________________________60 Figure 32 Effective notch stress concentration factors __________________________________________________________62 Figure 33 S-N Diagram Figure 34 Specific zones of the S-N diagram ________________________________________66 Figure 35 Loading parameters ____________________________________________________________________________________67 Figure 36 Correlation of the cycle asymmetry coefficient, amplitude and average stress ______________________67 Figure 37 S-N Diagrams ___________________________________________________________________________________________71 Figure 38 Specific zones of the S – N diagram ____________________________________________________________________71 Figure 39 Spectra of aleatory loading. ____________________________________________________________________________72 Figure 40 Curve of cumulated frequencies. _______________________________________________________________________72 Figure 41 Fatigue resistance curve _______________________________________________________________________________74 Figure 42 Fatigue curve domains _________________________________________________________________________________74 Figure 43 Stress-relieving cat pulsed constant strain ____________________________________________________________75 Figure 44 Creep stress-relieving asymmetric cyclic loading and controlled stress _____________________________75 Figure 45 Effect of average stress son fracture mechanisms for controlled stress testing. _____________________76 Figure 46 Influence of average stress son fatigue resistance for different Nf values. ___________________________77 Figure 47 Fatigue resistance S-N curves for steel, normal stress ________________________________________________78 Figure 48 Fatigue resistance S-N curves for aluminium, normal stress _________________________________________78 Figure 49 Modified resistance S-N curves of steel for Palmgren-Mine summation _____________________________81


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Figure 50 Modified resustance S-N curves of aluminium for Palmgren-Miner summation ____________________81 Figure 51 Enhancement factor f(R) _______________________________________________________________________________87 Figure 52 Toe distance ____________________________________________________________________________________________88 Figure 53 Examples of joint suitable for improuvement _________________________________________________________89 Figure 54 Examples of joints, at which an improvement might be limited by a possible root crack ___________90 Figure 55 Fatigue strength reduction factor for steel at elevated temperatures _______________________________92 Figure 56 Example of scatter of test data _________________________________________________________________________96 Figure 57 Comparison of engineering and true stree-strain curve ____________________________________________ 125 Figure 58 Dimensions of Charpy V notch standard specimen __________________________________________________ 128 Figure 59 Fracture surfaces of Charpy specimens of mild steel, tested at different temperatures ___________ 129 Figure 60 Transition-temperature curves for a) two steels, b) transition temperature criterion ___________ 130 Figure 61 Instrumented impact test resu1ts obtained with Charpy V specimen for steels A and B __________ 131 Figure 62 Typical load vs. time record showing fracture phases of Charpy specimen ________________________ 132 Figure 63 Fracture appearance vs. temperature for explosion-crack-starter ________________________________ 133 Figure 64 Typical results of explosion bulge test for steels A and B ___________________________________________ 134 Figure 65 Drop weight test conFigureation the anvil stop. ____________________________________________________ 134 Figure 66 Robertson crack-arrest test __________________________________________________________________________ 136 Figure 67 Temperature dependence of yield strength (σo), tensile strength (σu) _____________________________ 136 Figure 68 Fracture-analysis diagram showing influence of various initial flaw sizes ________________________ 136 Figure 69 Fracture analysis diagram for steel A _______________________________________________________________ 137 Figure 70 Application of DT test result for fracture analysis ___________________________________________________ 137 Figure 71 Different phases of fatigue life and relevant factors ________________________________________________ 138 Figure 72 Laboratory simulation of the multistage fatigue process. __________________________________________ 139 Figure 73 Different scenarios for fatigue crack growth ________________________________________________________ 141 Figure 74 Schematic illustration of variation of fatigue-crack -growth rate, da/dN, with alternating stress

intensity, ΔK, in steels, showing regions of primary crack-growth mechanisms. ______________________________ 141 Figure 75 Schematic of K-based crack resistance, R curve with position of KIc _______________________________ 145


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List of tables Table 1 Classification criteria and weld type according to EC 3-1-8. _______________________________________10 Table 2 Provisions regarding the correlation of loading and control of welds ______________________________18 Table 3 The partial safety factors γM for joints _________________________________________________________________26 Table 4 Type of joint model ______________________________________________________________________________________29 Table 5 Correlation factor β w for fillet welds. __________________________________________________________________34 Table 6 Conditions for welding cold-formed zone and adiacent material ___________________________________42 Table 7 Stress raisers and notch effects _______________________________________________________________________43 Table 8 Characteristics, limitations and conditions related to the type of welding. ________________________45 Table 9 The centre to centre spacing of fillet welds all round ________________________________________________48 Table 10 Types of hot spots ______________________________________________________________________________________55 Table 11 Correlation between relatively coase and fine models, to type of model and weld toe ________59 Table 12 FAT data, stress at knee-point of S-N curve, constants of tentative S-N curves and constants

for Palmgren-Miner summation __________________________________________________________________________________81 Table 13 Fatigue resistance against structural hot spot stress ______________________________________________83 Table 14 Effective notch fatigue resistance for steel __________________________________________________________86 Table 15 Thickness correction exponents ______________________________________________________________________87 Table 16 Benefit factors on stress of burr grinding and TIG dressing ______________________________________90 Table 17 Benefit on stress of hammer peening (nominal stress) ____________________________________________91 Table 18 Benefit on stress of needle peening (nominal stress) ______________________________________________92 Table 19 Testing approaches ____________________________________________________________________________________97 Table 20 F-factors for failure of all test specimens ____________________________________________________________98 Table 21 F-factors for the first test specimen to fail ___________________________________________________________98 Table 22 Categorisation and assessment procedure for weld imperfections ____________________________ 100 Table 23 Consideration of stress magnification factors due to misalignment ____________________________ 102 Table 24 Acceptance levels for weld toe undercut in steel _________________________________________________ 102 Table 25 Acceptance levels for weld toe undercut in aluminium __________________________________________ 103 Table 26 Acceptance levels for porosity and inclusions in welds in steel ________________________________ 104 Table 27 Acceptance levels for porosity and inclusions in welds in aluminium __________________________ 104 Table 28 Fatigue resistance values for structural details in steel and aluminium assessed on the basis

of nominal stresses _____________________________________________________________________________________________ 105 Table 29 Chemical composition and tensile characteristics of tested steels _____________________________ 131 Table 30 The three basic types of fatigue properties _______________________________________________________ 140

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IWSD Module 4 -Design of Welded Joints