Minimum Size of Double Continuous Fillet Weld Stainless Steel

Introduction to welding

Klas Weman , in Welding Processes Handbook (Second Edition), 2012

Basic terms

Some basic terms used in welding are defined below.

Weldment. The combined weld, heat affected zone (HAZ – see below) and base metal.

Butt welds. Butt welds join two pieces fitted edge to edge. Full penetration welds are most common and provide a particularly strong weld.

Fillet welds. The alternative to the butt weld, fillet welds join two overlapping pieces (lap joint) or two pieces placed perpendicularly to each other (e.g. a T or L-shaped joint)

Pressure welding. Welding in which sufficient outer force is applied to cause more or less plastic deformation of both the facing surfaces, generally without the addition of filler metal. Usually, but not necessarily, the facing surfaces are heated in order to permit or to facilitate bonding.

Fusion welding. Welding without application of outer force in which the facing surface(s) must be melted. Usually, but not necessarily, molten filler metal is added.

Welding procedure specification (WPS). A document specifying the details of the required variables for a specific application in order to assure repeatability (EN ISO 15609).

Deposition rate. The amount of metal supplied to the joint per unit time during welding.

Parent metal. The metal to be joined, or surfaced, by welding, braze welding or brazing.

Longitudinal direction. The direction along the length of the weldment, parallel to the weld.

Transverse direction. The direction along the width of the weldment, perpendicular to the weld.

Surfacing. Producing a layer of different metal by welding, e.g. with higher corrosion, abrasion or heat resistance than the parent metal.

Heat input. The heat input has great importance for the rate of cooling of the weld. It can be calculated from the formula:

$\begin{array}{cc}\hfill Q=\frac{U.I.60}{V.1000}.\mathrm{Efficiency}\hfill & \hfill \begin{array}{r}\hfill \text{Efficiency}*\hfill \\ \hfill \text{MMA}:0.8\\ \hfill \text{MIG}/\text{MAG}:0.8\\ \hfill \text{SAW}:1.0\\ \hfill \text{TIG}:0.6\end{array}\end{array}$

Where

Q  =   heat input (kJ/mm)

U  =   voltage (v)

I  =   current (A)

V  =   welding speed (mm/min)

(* Efficiencies according to EN 1011-1.)

Heat Affected Zone (HAZ). The heat affected zone, (Figure 1.14), is that area of the base metal not melted during the welding operation but whose physical properties are altered by the heat induced from the weld joint.

Figure 1.14. Fillet weld showing the location of weld toes, weld face, root and heat affected zone.

Throat thickness. Fillet welds are calculated by reference to the throat size. The size required is specified on drawings in terms of throat thickness, t, or the leg length, l, see Figure 1.15.

Figure 1.15. Throat thickness (t) and leg length (l) in a fillet weld.

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Calculating weld size

John Hicks , in Welded Joint Design (Third Edition), 1999

Fillet welds

The fillet weld is the cheapest type of arc welded joint since all that has to be done is to stand one piece of metal against another and run a welding rod or gun where the metals touch. The weld size is not defined by the thickness of the parts being joined as in the case of the butt weld; it can be as small or as large as the design requires or the welder thinks fit but there are limits on the size for other reasons. The minimum size is defined by the need for a minimum heat input to prevent hydrogen cracking, to obtain full fusion and to accommodate any lack of fit between the parts. The maximum size is limited by the economics of welding when above a certain size a butt weld may be more cost effective. Large fillet welds may also set up excessive distortion. The essential simplicity of the fillet weld has caused it to be used extensively in many types of structure. It is incapable of having its size verified by conventionally used non-destructive examination techniques although its external shape and size can be measured; confidence in its internal size and quality must therefore rest with pre-weld attention to fit-up and adherence to qualified welding procedures. The stress distribution in a fillet weld is compounded by the residual stresses due to welding but these are disregarded in most attempts to calculate the stresses in, and the strength of, fillet welds.

The basis of most methods of calculating fillet weld strength is the assumption that the key dimension defining the load carrying capacity of the weld is the throat size. A feature of overload failures of fillet welds is that in many of them the fracture is along their plane of fusion to one or other of the two members being joined and very often within the member material itself. It would therefore appear that the throat size is not the only criterion of strength. One of the other obvious influences is the relative strength and ductility of the weld metal and the parent metal; vary rarely will both properties match.

By convention there are three definitions of face profiles of fillet welds as shown in Fig. 6.2; the mitre fillet which has a flat face, the convex fillet and the concave fillet. Which shape is produced in a single run manual fillet weld depends upon the process, the consumable, the position and the welding conditions as well as the skill of the welder. A multi-run weld can of course be executed to produce any of these profiles. The theoretical static strength of the weld lies in the throat thickness (see Fig. 6.2) and a fillet weld must have full fusion along both legs right down to the root to give the full potential throat thickness. In the mitre weld this is the distance from the root along the normal to the face. In the convex weld it is in effect the same, i.e. the distance along the line through the root normal to the line joining the toes. In the concave weld it is the distance from the root to the tangent to the weld face at the centre of the face. A greater strength can be achieved for the same leg length if a deep penetration fillet weld is made; the throat thickness here is measured from the face to the tangent to the root of the weld as in Fig. 6.2. This can be done with submerged arc welding or MAG/MIG welding using higher currents than for the conventional fillet weld. Although deep penetration manual metal arc electrodes are available there is less confidence in the consistency of penetration along a joint with manual welding than with mechanised welding. The same principles apply to unequal leg fillet welds, see Fig. 6.3.

6.2. Fillet weld throat size.

6.3. Throat sizes for unequal leg fillet welds.

For fillet welds joining parts which are not set at right angles, and for fillet welds between rounded parts such as hollow sections, circular tubes and reinforcing bar various standards and codes give methods of calculating an effective weld throat.

Calculating the shear stress in a fillet weld caused by a load in line with (parallel to) the weld is straightforward; it is the load divided by the weld throat area, see Fig. 6.4.

6.4. Fillet weld with longitudinal shear load.

The weld throat shear stress is

[6.1] ${\tau }_{//}=\frac{P{}_{//}}{Lt}$

This stress is one of two types and two directions of stress which are postulated to exist in a fillet weld. Two are suffixed with the sign for parallel, //, indicating that they result from a load parallel to the length of the weld. Two are suffixed with the sign for perpendicular, ⊥, indicating that they result from a load perpendicular to the length of the weld, see ⊥Fig. 6.5. These are merely symbols describing the type (normal and shear) and direction of stress and do not represent a set of internally balanced stresses.

6.5. Stress notation for fillet weld.

Tests on fillet welds in mild and high yield steels with nominally matching weld metal found the normal stress, σ_//, to have no measurable effect on the strength of the weld. This type of stress is most common in the web to flange weld in an I beam in bending. For design purposes in structural steels it was found that the three other stresses could be related to an allowable stress by a formula of the type

[6.2] $\begin{array}{l}\mathrm{ß}\sqrt{{\sigma }_{\perp }+3\left({{\tau }_{\perp }}^2+{{\tau }_{//}}^2\right)}⩽{\sigma }_{\mathrm{c}}\\ \mathit{and}\\ {\sigma }_{\perp }⩽{\sigma }_{\mathrm{c}}\end{array}$

where σ_c can be the allowable tensile stress or limit state stress. This is used as a basis for fillet weld strength in a number of standards in which values for ß are typically in the region of 0.8–0.9 depending on the strength of the parent metal.

The example above shows how τ_// is calculated, another simple example will show how the other two stresses are derived.

Figure 6.6 shows twin fillets, each with throat thickness t. Then, resolving vertically

6.6. Twin fillet welded joint under load.

[6.3] $P_{\perp }=\frac{2tL}{\sqrt{2}}\left({\sigma }_{\perp }+{\tau }_{\perp }\right)$

and horizontally we can see that

[6.4] ${\sigma }_{\perp }={\tau }_{\perp }$

The stresses so calculated can be put into the equation with the relevant parameters β and σ_c to arrive at a value of t for the design. If there is a load creating a parallel shear stress then this stress can also be entered into the equation.

This is a rather cumbersome procedure for run-of-the-mill work and it is often customary to use just the load divided by the weld throat as a measure of weld stress which in structural steels is then compared with the allowable or limit parent metal shear stress. If a parallel load is present then the two throat stresses are summed as a resultant square root of the sum of the squares.

If τ_t, is the nominal fillet weld throat stress then

[6.6] ${\tau }_t=\frac{\sqrt{{P_{//}}^2+{P_{\perp }}^2}}{2Lt}$

For other materials more complicated routes may be used and the relevant standard or code of practice is followed.

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Manufacturing technology

J. Carvill , in Mechanical Engineer's Data Handbook, 1993

5.12.3 Types of weld

The fillet weld , the most used, is formed in the corner of overlapping plates, etc. In the interests of economy, and to reduce distortion, intermittent welds are often used for long runs, with correct sequencing to minimize distortion. Tack welds are used for temporary holding before final welding.

Plug welds and slot welds are examples of fillet welds used for joining plates. For joining plates end to end, butt welds are used. The plates must have been suitably prepared, e.g. single or double V or U, or single and double bevel or J. To avoid distortion, especially with thick plates, an unequal V weld may be used, the smaller weld being made first.

Resistance welding is used to produce spot welds and stud welding by passing an electric current through the two metal parts via electrodes. In seam welding the electrodes are wheels.

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Design Basis Loads and Qualification

George Antaki , Ramiz Gilada , in Nuclear Power Plant Safety and Mechanical Integrity, 2015

FIV failures have been reported in socket welds to vents or drains; has the industry developed a solution?

Because fillet welds and threaded joints have proven to be prone to vibration-induced fatigue cracking, it is preferable to use butt welds, if practical, in systems subject to vibration. Alternatively, vibration testing of socket welds, sponsored by EPRI (EPRI Report TR-113890) revealed that socket welds with a ratio of length (along the pipe) to width (radial length of the filet) of approximately 2:1 have a much longer fatigue life when subject to high-cycle-low-stress vibration.

One plant experienced four socket weld failures in a 12-month period. The failed welds were downstream of pressure-reducing orifice isolation valves. The orifices themselves showed evidence of wear caused by erosion due to cavitation, with the orifice size growing by wear from 0.2 to 1   inch. This cavitation in turn caused FIV in the downstream piping (IN 98-45).

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The use of fracture mechanics in the fatigue analysis of welded joints

A. Hobbacher , in Fracture and Fatigue of Welded Joints and Structures, 2011

4.4.3 Assessment of a finite element output of a welded joint for fatigue

A fillet weld on a thick-walled steel plate in a special geometrical configuration had to be assessed. No corresponding structural details could be found in the detailed catalogues of codes, so a finite element analysis was performed. The weld toe transition was modelled assuming a toe radius of 1  mm. The stresses have been read from the anticipated crack path; see Fig. 4.15. The stress parts have been separated according to Fig. 4.7 and formulae 4.12. After that, the non-linear stress magnification factors k _t,nl(x) have been determined. With that function, the function M _k(a) could be determined and the stress intensity factors have been calculated, from which the number of live cycles eventually could be derived.

4.15. Stress separation of finite element results.

The nodal stresses of the elements have been used as supporting points for the interpolation of the stresses. The computation starts with the separation of stress parts resulting in σ_m  =   82.55   MPa and σ _b,max  =   22.5   MPa. The remaining stress part is the non-linear peak stress. The integration was done numerically by a stepwise integration of crack and cycle increments and by a parallel computing of the stress intensity factors ∆K and of the function M _k. The step at the beginning was ∆a  =   0.1   mm; 250 steps have been used. The aspect ratio was kept constant to a:c   =   1:10 for conservative estimations. Table 4.4 shows the evolution of the crack depth a.

Table 4.4. Stepwise integration of crack growth (table shows only selected crack lengths)

X (mm)	∆σ (MPa)	a (mm)	Y (a)	M k (a)	∆K (N   mm–3/2)	∆N (cycles)	N (cycles)
0	300	−	–	–	–	–	–
0.1	249.96	0.1	–	–	–	–	–
0.2	219.92	0.2	1.104	2.477	161	47834	47834
0.5	189.8	0.5	1.106	2.093	273	44856	92690
1.0	154.6	1.0	1.109	1.805	354	29488	122178
2.0	129.2	2.0	1.117	1.518	435	31383	153561
5.0	113	5.0	1.144	1.268	598	43699	197260
10	104	10	1.202	1.145	810	29933	227193
25	94	25	1.490	1.057	1466	24865	252058
50	80	50	2.574	1.024	3471	5652	257710
75	64	–	–	–	–	–	–
100	60	–	–	–	–	–	–

The results can be converted in a fatigue class FAT according to the IIW recommendations [8], which is the stress range at 2 million cycles and R  =   + 0.5. This class is FAT 53   MPa, but it must be considered that this fatigue class was derived for a wall thickness of 100   mm. A recalculation to 25   mm, on which the fatigue data of IIW recommendations are based, using a wall thickness correction exponent of n  =   0.3 results in a fatigue class of FAT 80   MPa.

In very thick-walled or voluminous components, the consideration of the stress distribution through the whole section is not necessary. The linearisation and separation of stress needs only to be done in a reasonable surrounding of the estimated final crack. A possible error can be estimated by the function f _w in Table 4.1.

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Connections

Jonathan Ochshorn , in Structural Elements for Architects and Builders, 2010

Fillet weld terminations

In certain cases, fillet welds must be terminated before reaching the edge of the steel elements they are connecting in order to prevent damage (notching, gouging) of the element's edge. Figure 9.33a illustrates a condition where the fillet weld at the underside of a plate (shown as a dotted line) must be interrupted at the corners before turning 90° and being deposited on the opposite side of the same plate. Figure 9.33b illustrates a lap joint that extends beyond a tension element; in such cases, the fillet weld must terminate a distance equal to the weld size, w, from the edge of the tension element.

Figure 9.33. Termination of fillet welds where (a) welds occur on opposite sides of a common plane and (b) a lap joint extends beyond a tension element

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Offshore Structure Platform Design

Mohamed A. El-Reedy Ph.D. , in Offshore Structures, 2012

3.8 Topside Design

In general, major rolled shapes for offshore structure design should be compact sections, as defined by AISC. The minimum thickness of a structural plate or section should be 6 mm. The minimum diameter to thickness ratio of tubular members should not be less than 20, where the diameter is based on the average of the tubular outside and inside diameters.

Connections design to comply with the codes must fulfill the following minimum requirements.

•

The minimum fillet weld should be 6 mm.

•

Wherever possible, joints should be designed as simple joints with no overlap.

•

Tubular joints should be designed in accordance with API RP2A.

The deflection (discussed in Chapter 2) should be matched with the codes and defined in the owner's specification.

•

Deflections should be checked for the actual equipment live loads and casual area live loads. Pattern loading should be considered.

•

Deflection of members supporting deflection-sensitive equipment should be not greater than L/500 for beams and L/250 for cantilevers under live loads.

•

Deflection of beams in the workrooms and living quarters should be not greater than L/360 for beams and L/180 for cantilevers under live loads.

•

Deflection limits for other structures should be L/250 for beams and L/125 for cantilevers under live loads.

In performing structural analysis using software (such as SACS, for example), it is better to define the direction of the model. In most cases, the directions can be:

•

+x axis aligned with platform east

•

+y axis aligned with platform north

•

+z axis aligned vertically upward

•

The datum for the axes should be at chart datum.

In general, only the primary structural steel should be modeled. However, secondary members should be modeled where they are necessary for the structural integrity or to facilitate load input. Deck plates should be modeled as shear panel elements. Joint eccentricities should be modeled using discrete elements rather than using the "offset" facility of SACS. When individual elements are used, joint forces can be more easily extracted from the output.

All the differing analyses (in-place, lift, loadout, etc.) should utilize the same base model. That is, the in-place model should form the basis for all the other analyses to be performed.

The in-service analyses should include a basic model of the jacket structure to ensure the correct stiffness interaction between the jacket and the topside structures. Pile foundations to the jackets need not be considered for topside analysis and design—simple supports are sufficient.

3.8.1 Grating Design

Grating is traditionally member used to cover the platform floor (as in onshore facilities). The check list in Table 3.13 should be filled in, to make certain that the design of the grating matches the client's requirements for its function.

Table 3.13. Grating Design Check List

Item to Be Checked	Yes/No
Grating to be removable
Vibration performance required
Corrosive environment
Grating over stainless steel grating or piping
Panels to be sized to facilitate fabrication
Weight and size limitation was defined
Impact or high local load
Loads meet the design requirement
Special operating loads to be considered
Grating matches the required strength
Grating within the allowable deflection limit
Suitable corrosion protection specified
Anti-slip surface specified
For fiber-reinforced polymer (FRP) the fire performance requirements were specified
For FRP the smoke and toxicity requirements were defined
Grating slope is acceptable
Adequate lateral restraint was provided
Tolerance was checked
Support arrangements provide adequate support at penetrations and cut-outs
There is an isolation between different materials

Grating should have a 1-inch (25-mm) minimum bearing on supporting steel. Where grating areas are shown as removable on the drawings, the weight of fabricated grating sections for such areas should not exceed 160 kg (350 lb).

Most grating and expanded metal needs to be supported in a specific way. The direction in which the load bars run is the important direction for grating and is usually referred to as the span. For expanded metal, the span is in the direction of the strands. Span is always the least dimension given when referencing a panel size. In most cases, the grating has to be supported in the span direction only and it does not require support on all four sides, unlike the floor plate.

The different grating dimensions are shown in Figure 3.41 and the different types of grating are shown in Figure 3.42.

Figure 3.41. Grating dimensions.

Figure 3.42. Types of grating.

There is a relation between the span and the deflection. When you select the grating from the manufacturer, you should refer to his type and calculation. To illustrate a grating sheet based on its type and loads, Table 3.14 has been arranged in increasing strength order. The load, L, is a safe superimposed uniformly distributed load in KPa (100 kg/m² = 0.98 KPa), where D is the deflection (in mm) for L. Loads are calculated in accordance with an allowable bending stress of 171.6 MPa. Note that the load bars are assumed to be simply supported.

Table 3.14. Relation between Grating Dimensions, Maximum Span and Maximum Load

Load bar spacing (mm)	Cross bar pitch	Weight (kg/m2)	Bar size	Load (Kpa)	Spacing between Supports (mm)
					450	600	750	900	1050	1200	1500	1800	2100
40	100	17.5	25 × 3	L	53	30	19	13	10	7	5	3	2
				D	1.4	2.6	4	5.8	7.8	10.2	16	23.1	31.5
60	50	22.3	25 × 5	L	56	32	20	14	10	8	5	3	2
				D	1.4	2.6	4	5.8	7.8	10.2	16	23.1	31
40	100	26.9	25 × 5	L	70	39	24	17	13	9	6	4	3
				D	1.6	2.9	4.5	6.5	8.8	11.5	18.0	25.9	35.3
30	100	22.8	25 × 3	L	70	39	25	17	13	10	6	4	3
				D	1.4	2.6	4	5.8	7.8	10.3	16.0	23.1	31.5
60	50	26.4	32 × 5	L	76	43	27	19	14	11	7	5	3
				D	1.2	2.2	3.4	4.9	6.7	8.7	13.6	19.7	26.8
30	100	34.7	25 × 5	L	91	51	33	23	17	13	8	6	4
				D	1.6	2.9	4.5	6.5	8.8	11.5	18.0	25.9	35.3
40	100	34	32 × 5	L	120	67	43	30	22	17	11	7	5
				D	1.2	2.2	3.4	4.9	6.7	8.7	13.6	19.7	26.8
30	100	28.4	32 × 3	L	114	64	41	28	21	16	10	7	5
				D	1.1	2	3.1	4.5	6.1	8.0	12.5	18.1	24.6
40	100	42.1	40 × 5	L	226	127	81	56	41	31	20	14	10
				D	0.9	1.6	2.5	3.6	4.9	6.4	10.0	14.4	19.7
30	100	62.9	45 × 5	L	377	212	135	94	67	52	33	23	17
				D	0.8	1.4	2.2	3.2	4.3	5.7	8.9	12.8	17.5

Sometimes pipe support will form a concentrated load on the grating, as shown in Figure 3.43; in this situation, the pipe support is usually a base plate resting directly on the grating, and there is no steel under the support.

Figure 3.43. Relation between allowable loads and location of pipe support for different spans.

3.8.2 Handrails, Walkways, Stairways and Ladders

Handrails, walkways, stairways and ladders should be designed in accordance with OSHA 3124.

Handrails should be provided around the perimeter of all open decks and on both sides of stairways. All handrails should be 1.10 m high and made removable in panels no more than 4.5 m long. Handrail posts should be spaced 1.5 m apart. The gap between panels should not exceed 51 mm. Handrails in the wave zone should also be designed to withstand extreme storm maximal wave loadings.

Walkways, stairways and landings should be designed for load combinations:

1.

Dead load plus live loads.

2.

Dead loads plus extreme storm 3-second wind gusts and/or extreme storm maximum waves, whichever is applicable.

Stairways should be of structural steel, using a double runner with serrated bar grating treads and handrails.

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Tension Members and Welds in Thin Cold-Formed Tubes

Xiao-Ling Zhao , ... Gregory Hancock , in Cold-Formed Tubular Members and Connections, 2005

7.4.2 Design Rules

7.4.2.1 AS 4100

Design rules for longitudinal fillet welds in RHS were derived ( Zhao and Hancock 1994, 1995b, Zhao et al 1999) based on both parent metal strength as in AS/NZS 4600 and weld metal strength as in AS 4100. Only those based on weld metal strength are presented here since that is the approach adopted in AS 4100-1998.

The nominal capacity of the connection in tension (N _t) can be expressed as

(7.7) $N_{\text{t}}=\mathrm{0.6·}{f}_{\text{uw}}·a·L$

where f _uw is the weld metal strength, a is the weld throat thickness and L is the weld length. The term 0.6f _uw represents the design shear stress of the weld metal where the value of f _uw is the nominal tensile strength of weld metal given in Table 7.3.

Table 7.3. Table 7.3 Nominal tensile strength of weld metal (from Table 9.7.3.10(1) of AS 4100-1998)

Manual metal arc electrode (AS/NZS 1553.1)	Submerged arc (AS 1858.1) Flux cored arc (AS 2203) Gas metal arc (AS/NZS 2717.1)	Nominal tensile strength of weld metal, f uw N/mm2
E41XX	W40X	410
E48XX	W50X	480

AS 4100 does not specify the requirement of end return welds as shown in Figure 7.1 (b). However a lower capacity factor of ϕ of 0.7 is given for longitudinally welded RHS with a thickness less than 3 mm. This is based on the reliability analysis carried out by Zhao and Hancock (1995b) where tests without end return welds were also included. The design capacity of the connection is given by

(7.8a) $\varphi ·N_{\text{t}}=\varphi \mathrm{·0.6}{f}_{\text{uw}}·a·L$

(7.8b) $\varphi =0.8\text{ when}t\ge 3\text{ mm}$

(7.8c) $\varphi =0.7\text{ when}t<3\text{ mm}$

7.4.2.2 BS 5950 Part 1

Unlike AS 4100, the end return welds are specified as a requirement in BS 5950 Part 1: fillet welds finishing at the ends or sides of parts should be returned continuously around the corners for a distance at least twice the leg length ( s) of the weld.

The longitudinal shear capacity P _L per unit length of weld should be taken as:

(7.9) $P_{\text{L}}=p_{\text{w}}·a$

where p _w is the design strength of a fillet weld and a is the throat thickness of the weld.

The design strength p _w of a fillet weld is given in Table 7.4 (from Table 37 of BS 5950 Part 1), corresponding to the electrode classification and the steel grade.

For other types of electrode and/or other steel grades:

(7.10) $p_{\text{w}}=0.5f_{\text{uw}}\text{ with}{p}_{\text{w}}\le 0.55f_{\text{u}}$

where f _uw is the minimum tensile strength of the electrode as specified in the relevant product standard and f _u is the specified minimum tensile strength of the parent metal.

7.4.2.3 Comment on Weld Throat Failure

In order to have the design capacity (ϕN _t ) of longitudinal fillet welds given in Equation (7.8) less than the parent metal shear resistance V _r given in Equation (7.4), the condition required is:

$\frac{V_r}{\varphi ·N_t}=\frac{0.9\times 0.67\times f_{\text{y}}·t}{0.8\times 0.6\times f_{uw}·a}>1.0$

$\frac{f_y}{f_{uw}}>\frac{0.8\times 0.6\times a}{0.9\times 0.67\times t}$

Assume the weld leg length (s) is the same as the tube thickness (t), i.e. the weld throat thickness (a) becomes $\frac{t}{\sqrt{2}}$ . Therefore

$\frac{f_y}{f_{uw}}>\frac{0.8\times 0.6\times \frac{t}{\sqrt{2}}}{0.9\times 0.67\times t}=0.56$

Similarly when comparing Equation (7.9) and Equation (7.4), the above condition becomes:

$\frac{f_y}{f_{uw}}>0.58$

The nominal yield stress of commonly used cold-formed RHS varies from 228 N/mm² to 460 N/mm² as listed in Table 2.4. The nominal tensile strength of commonly used weld metal is 410 N/mm² and 480 N/mm² as listed in Table 7.3. The smallest ratio of f _y to f _uw is about 0.56 (= 228/410). Therefore for most cold-formed RHS the design capacity given in AS 4100 and BS 5950 Part 1 are always lower than that given by Equation (7.4). In other words the formulae for longitudinal fillet welds in AS 4100 and BS 5950 Part 1 can be used conservatively to design slotted gusset plate connections to RHS in tear-out failure.

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Notch stress intensity approach for seam-welded joints

D. Radaj , ... W. Fricke , in Fatigue Assessment of Welded Joints by Local Approaches (Second Edition), 2006

7.1.2 Principles and variants of the approach

The toe notch of fillet welds or butt welds is a critical region with regard to fatigue crack initiation and propagation. The toe may be a sharp notch (zero notch radius) or a blunt notch (finite notch radius). High notch stresses, actually a stress singularity in the case of sharp notches, occur in this region.These stresses may be described by elastically determined notch stress intensity factors or, in the case of higher loads, by elastic-plastically determined notch stress or notch strain intensity factors. The notch stress intensity factors may be used to define J-integrals, an averaged strain energy density or effective stresses at the corner notch, in order to derive special failure criteria.

The notch stress intensity approach takes account of the notch opening angle (or weld toe angle) and of the notch tip radius (or weld toe radius), proceeding from the nominal or structural stresses in the notched structural member (tensile, bending and shear stresses superimposed). It also comprises the (cyclic) strain-hardening property of the material in the case of elastic-plastic behaviour and a notch support factor when combined with failure criteria.

In its basic version, the approach refers to the elastically determined notch stress intensity factor with minor extensions in the case of a small notch tip radius or minor modifications in the case of small-scale yielding at the notch tip (J-integral or averaged strain energy density at the notch tip).

There is no generally adopted fatigue assessment procedure for welded joints based on notch stress intensity factors available up to now, partly because the method is rather new and partly because only fillet welds have been considered up until recently. Relevant investigations in butt welds have meanwhile been performed with satisfactory results.

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Pressure Vessel Internal Assessment

A. Keith Escoe , in Pressure Vessel and Stacks Field Repair Manual, 2008

Support Clip with Continuous Fillet Weld with an In-Plane Bending Moment, M_zz

A support clip with continuous fillet weld with an in-plane bending moment is shown in Figure 5.14.

Figure 5.14. Support clip with continuous fillet weld with an in-plane bending moment, M_zz ; see Giachino et al. [3].

(5.11) $\sigma =\frac{4.24M_{zz}}{h\left[L^2+3b(L+h)\right]}\text{ psi}(\text{MPa})$

Example 5.2

An example illustrating the design of a support clip with continuous fillet weld with an in-plane bending moment.

A support clip has an in-plane bending moment of 50   kNM. The clip has the following parameters:

L=300   mm; b=120   mm; weld size=t=5   mm

Now, M=50   KNm=50,000   Nm

$h=\frac{t}{cos45°}=\frac{5\text{mm}}{cos45°}=7.071\text{mm}$

Using Eq. 5.11, you can calculate

$\sigma =\frac{4.24(50,000)\text{Nm}\left(\frac{1000\text{mm}}{\text{m}}\right)}{(7.071)\text{mm}[{300}^2+3(120)(300+7.071]{\text{mm}}^2}=149.5\text{MPa}$

Hicks, p. 87−88 [4], has a more comprehensive solution. However, the preceding equation is much simpler and very accurate.

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