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4.2 Design of welded joints with predominantly static loading 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 behavior. The term "mechanical behavior" encompasses the response of materials to external forces. 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 4.2.1).

Figure 4.2.1 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 behavior 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 behavior 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 usualIy 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 mecanical 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

15

materials is the torsion test. As described below, the uniaxial stress-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 funnctionality are two of the most prized qualities required from engineered components. Theese are not achieved by accident. Indeed, considerable scientific and technological endeavor is expended to help achieve them, because without them the functionalitv 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 commputers, 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 metalic 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 structutural 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. 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 4.2.1.

Table 4.2.1 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

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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 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 devic – 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 4.2.2). For a single angle in tension connected by bolts on one leg the simplified design method given in set design, may be used. The effect of eccentricity on angles used as web members in compression is given in EN

17

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

Figure 4.2.2 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. 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 4.2.2, 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.

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Table 4.2.2 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. Frecvently, 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 impoprtant 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 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, - corss section effective aria - 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 cros section areas, withpot defects in BM, WELD, respectively, and RrMB . RrSUD fracture resistences 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 fracfture 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) Defects induce the change of the stress state by its concentration in the defect section expressed by trhe concentration coefficient: kS = RrdSUDef / RrSUD > 1 (4.2.5) Figure 4.2.3 presents the evolution of the bearing capacity of the weld with the defect area.

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Figure 4.2.3 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: RrMB . Ao = kS . RrdSUDef (AS - Ād) (4.2.9) Where it is explained:

Ād = Ao [ (AS/ Ao) - (RrMB / kS RrSUD)] (4.2.10) Addmitting 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 (suphlurs, inclusionsi, 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 t

min 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: l

min≥ 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. Otherwize 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.

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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 “b

eff “ 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 strength is determined according to EC 3 as function of fracture tensile nominal strength tensile of the steel used in joining f

u and not as a function of its yeald limit f

y.

The design resistance of a fillet weld should be determined using: - directional method, - simplified method.

a). 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 4.2.4), as follows: • σ - is the normal stress perpendicular to the throat • σ

|| - 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 4.2.4 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:

25,0222 /3 MWuf and 2/ Muf (4.2.12)

where: - fu is the nominal ultimate tensile strength of the weaker part joined; - β w is the appropriate correlation factor taken from table 4.2.3. Welds between parts with different material strength grades should be designed using the properties of the material with the lower strength grade. Table 4.2.3 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 H S 275 NH/NLH

S 275 H S 275 NH/NLH 0.85

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S 275 M/ML S 275 MH/MLH 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

b) 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) 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

udvw

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 4.2.5). 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 4.2.5 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 4.2.6).

Figure 4.2.6 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 resistancxe of the weakest joined part, provided that welding is done by filler materials that will ensure in all tensile tests, yeald limit (f

y) 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.

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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 corelation 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 – 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 4.2.7). The effective throat thickness of a fillet weld should not be less than 3 mm.

Figure 4.2.7 Throat thickness of a fillet weld.

In determining the design resistance of a deep penetration fillet weld, account may be taken of its additional throat thickness (figure 4.2.8), provided that preliminary tests show that the required penetration can consistently be achieved.

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Figure 4.2.8 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 4.2.9. The definition of the design throat thickness of flare groove welds in rectangular hollow sections.

Figure 4.2.9 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 4.2.10, 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, – that of the supporting flange as given by formules, calculated assuming the applied force is concentrated over an effective width, beff, of the flange as given as relevant.

Figure 4.2.10 Effective width of an unstiffened T – joint

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

beff = tw 2s 7k.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:

25

– 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:

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 4.2.11.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 4.2.11.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 4.2.11 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. 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 4.2.4), provided that one of the following conditions is fulfilled: – the cold-formed zones are normalized after cold-forming but before welding;

26

– the r/t -ratio satisfy the relevant value obtained from table 4.2.4. Table 4.2.4 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%)

Predominantly 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)

where fvw.d

– is shear calculating resistance of the weld, A

w – 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 (σ or τ) 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 equivqlent 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 4.2.5, figure 4.2.12, 4.2.13).

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Table 4.2.5 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

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 4.2.12 Modified or local nominal stress Figure 4.2.13 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 4.2.6).

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Table 4.2.6 Characteristics, limitations and conditions related to the type of welding.

Joint type Weld type Characteristics, limitations and

conditions 0 1 2

in T

, in

angl

e

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) minSl 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

angl

e

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. 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.

29

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 4.2.14. 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. Corelated with previous figure, to remember:

Figure 4.2.14 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

30

In any run of intermittent fillet weld there should always be a length of weld at each end of the part connected. 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 4.2.14). 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 4.2.7. 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 4.2.7 The centre to centre spacing of fillet welds all round

Distances and spacings, see

Figure 3.1 Minimum

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

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 spacings, 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.

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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, 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 σln are (figure 4.2.15): σmem membrane stress, σben shell bending stress, σnlp non-linear stress peak

Figure 4.2.15 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 σmem is equal to the average stress calculated through the thickness of the plate, and it is constant through the thickness, - the shell bending stress σben 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 σnlp is the remaining component of the stress. The stress components can be separated analytically for a given stress distribution σ (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 4.2.16). Overall elastic behaviour is assumed.

Figure 4.2.16 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

32

(figure 4.2.17, 4.2.18). 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 4.2.19). This is done by the application of an additional stress raising factor km,eff.

Figure 4.2.17 Examples of macrogeometric effects

Figure 4.2.18 Modified (local) nominal stress near concentrated loads

Figure 4.2.19 Axial and angular misalignement

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.

33

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. 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 para-meters 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 4.2.20 shows examples of structural discontinuities and details together with the structural stress distribution.

Figure 4.2.20 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 4.2.21.

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Figure 4.2.21 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 4.2.22.

Figure 4.2.22 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 suceptible 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. In case of a biaxial stress state at the plate surface, it is recommeded 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 4.2.23).

35

Figure 4.2.23 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 distiguished according to their location on the plate and their orientation to the weld toe (table 4.2.8). 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 eleminated by linearisation of the stress throughth 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 4.2.8 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

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

36

welded joint. Consequently, any possible misalignment has to be taken explicitely 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 4.2.24).

Figure 4.2.24 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 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 4.2.25.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 4.2.25 Typical meshes and stress evaluation parth 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

37

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 4.2.25.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. 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 4.2.21. 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 σhs is determined using the reference points and extrapolation equations as given below (figure 4.2.26).

Figure 4.2.26 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 centers 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)

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

38

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 extrapolation (eq. 5).

mmmmmmhs 1284 33 (4.2.26)

mmmmhs 155 5,05,1 (4.2.27) Corellation between relatively coase and fine models, to type of model and weld toe it is in table 4.2.9. Table 4.2.9 Corellation 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 center

5 and 15 mm surface center

0.4 t and 1.0 t nodal points

4. 8 and 12 mm nodal points

*)

w = longitudinal attachment thichness + 2 weld leg lenths **)

surface center at transversal welds, if the weld below the plate is not modelled (see figure 4.2.24.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 4.2.27).

Figure 4.2.27 Examples of strain gauges in plate structures

39

The center 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)

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 σ

hs can then

be resolved assuming that this principal stress is about perpenticular to the weld toe.

21

1

v

vE 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 ∆σhs. 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 σhs becomes:

40

nomhshs k (4.2.33)

where σnom 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 4.2.28).

Figure 4.2.28 Effective notch stress concentration factors

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”.

41

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 σmem - membrane stress σben -shell bending stress, 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.