Ch 2 5 Selection of Steel Quality

STEEL CONSTRUCTION: APPLIED METALLURGY __________________________________________________________________________

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STEEL CONSTRUCTION:

APPLIED METALLURGY

Lecture 2.5: Selection of Steel Quality

OBJECTIVE/SCOPE

To describe the selection of steel quality in relation to requirements of toughness.

PRE-REQUISITES

Lecture 2.1: Characteristics of Iron-Carbon Alloys

Lecture 2.3.1: Introduction to the Engineering Properties of Steels

Lecture 2.3.2: Advanced Engineering Properties of steels

RELATED LECTURES

Lecture 2.4: Steel Grades and Qualities

Lecture 2.6: Weldability of Structural Steels

SUMMARY

Selection of the right steel quality for a structure is a matter of major significance as

regards both the safety and the economy of constructional steelwork. This lecture surveys

procedures which have been proposed for this purpose and presents the new rules which

are included in Annex C of Eurocode 3 [1]. All these express, as a function of extreme

service conditions applicable to a structure, a toughness level specified in terms of

performance in the Charpy V test that the selected steel should fulfil, with a transition

temperature at the level of 28J for instance. Numerous comparisons between the output of

different procedures are reported which, on the one hand highlight their consistency, while

on the other hand the possible sources of discrepancies among the various material

requirements determined using these procedures. Such procedures are based on fracture

mechanics concepts such as those of the Stress Intensity Factor, the Crack Tip Opening

Displacement or the Full Yield Criterion. As an introduction the lecture reviews the main

aspects of resistance to brittle failure, with reference to basic documents which the reader

may find it useful to consult for more detail.

1. INTRODUCTION

There are circumstances when the integrity of a structure is governed not by the strength

of the metal but by another property, namely toughness.


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Such situations generally imply the presence of defects in the structure such as cracks or

sharp notches and are favoured by the occurrence of low temperature. The incidence of

dynamic loads is another parameter enhancing the risk of so-called brittle fracture.

Thus the engineer has to consider that the concept of ultimate boundary states and the

fulfilment of the related criteria may apply and lead to a safe design only if the pre-

requisite conditions that prevent brittle failure are met.

In normal circumstances, it is impracticable to undertake a detailed 'fitness-for-purpose'

analysis involving sophisticated fracture mechanics tests either at the design stage or

during fabrication and erection of conventional structures. For such constructions, simple

rules have to be developed and specified in building codes to define which qualities of

steel should be selected to ensure a safe design.

This lecture is divided into sections devoted respectively to:

A brief survey of brittle failure.

A review of different fracture mechanics concepts.

A presentation of the different approaches on which a methodology for steel

selection may be based.

An illustration of the present solution adopted for Eurocode 3 [1] and of the work

planned on this subject in the near future under the auspices of the International

Institute of Welding.

2. THE PHENOMENON OF BRITTLE FAILURE

A material is generally said to be brittle if it cannot be deformed to any appreciable degree

prior to fracture. This behaviour does not imply that the ultimate tensile strength measured

on a smooth specimen during a tensile test is low. On the contrary, the opposite

phenomenon is usually observed. Hardening treatments which aim to increase strength are

usually accompanied by a dramatic degradation of ductility and tend to enhance

brittleness.

Brittleness is neither an absolute nor a simple concept. As a rule, the susceptibility to

brittle behaviour in a given material is increased by:

the lower the temperature to which it is exposed.

the more rapid the loading to which it is submitted.

the more disturbed the stress distribution it experiences.

Brittleness is influenced by ductility, i.e. the capacity of a material to strain plastically, and

by strain-hardenability, i.e. the property of developing a higher strength while undergoing

plastic deformation.

Ductility can easily be appraised in a bend test under strain-controlled conditions where

the material is bent round a mandrel with large plastic strains being induced in the outer

fibres of the specimen (Figure 1). The more ductile material can be bent round a smaller

mondrel without fracture.


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Strain-hardenability is appraised in a tensile test and is quantified by the slope of the

stress-strain curve in the plastic regime. Strain hardening governs the amount of uniform

elongation a material may undergo before necking or fracturing (Figure 2).


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A material showing no ductility is intrinsically brittle and produces, even in a defect-free

situation, negligible plastic elongation during a tensile test and no strain hardening. Such a

material is glass.

Steel usually exhibits ductile behaviour in the tensile test and the presence of some defect

is necessary to induce brittleness. The defect may be a geometrical discontinuity with

sharp edges constituting a stress raiser, or an area in which the mechanical properties are

locally impaired such as the heat affected zone of a weld, or a region of local plastic


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deformation which may undergo subsequent strain ageing. Often, both geometrical and

metallurgical defects are present together; weldments may undergo lack of penetration or

contain cold cracks, while punches or shears may create burrs or notches. A large

embrittling effect may thus be induced.

Should the ductility of the metal at the tip of the notch be very poor, then no possibility

would exist to blunt the notch by plastic deformation. The result would be a brittle failure

occurring at a load which can be calculated from linear elastic theory as a function of the

component dimensions, the notch size and the toughness characteristics of the metal.

In the case where a degree of ductility is available, crack blunting occurs and is reflected

in a degree of crack opening. The fracture behaviour is then significantly influenced by the

strain hardenability of the metal. If little strain hardening is available, then the crack may

propagate through the component at a rather constant stress level, either by ductile tearing

or brittle cracking.

Different fracture behaviour can be observed with a ductile metal which is capable of

sustained strain hardening since propagation of the defect after crack blunting requires an

increasing load to be applied to the component. Such conditions give rise to stable crack

propagation.

3. FRACTURE MECHANICS CONCEPTS AND

TESTING PROCEDURES

Several specific tests have been developed to assess the fracture behaviour of materials in

various loading conditions and defect configurations. It is not the scope of this lecture to

detail the testing procedures or review the methods for derivation of toughness values.

However, it is worth listing the main concepts and types of test as well as the assumptions

on which they rely.

Fracture Mechanics has been approached in a rigorous way on the basis of linear elastic

theory which led to the well-known concept of stress intensity factor, KI [2]. This

parameter defines completely the stress field in the vicinity of a crack. Fracture occurs

when it attains a critical value which is a characteristic of the material. The main

assumption of this theory is the presence of minimal plastic strain at the crack tip. With

steel, such conditions may be met with very high strength products or when the thickness

is large. Table 1 defines the basic equations of Linear Elastic Fracture Mechanics (LEFM)

together with the conditions of applicability.

In many circumstances significant plasticity takes place at the crack front prior to failure

and crack opening may be observed before fracture initiation (Figure 3). The concept of

Crack Opening Displacement (COD) or more precisely (CTOD) defines the amplitude of

the crack tip plasticity under a given stress situation. Fracture initiates when this parameter

reaches a critical value (crit) which is a characteristic of the material and a measurement

of its toughness [3].


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The theory of the J integral is based on the same assumptions as above but computes a

specific fracture energy whose value is independent of the contour of integration and

which is an alternative measurement of the material toughness [4].

It is important to mention that both the CTOD and the J concepts are designed to assess

situations in which fracture occurs in the elasto-plastic region, but imply nevertheless a

relatively small extent of plastic strain at the crack tip. To assess the risk of failure by

brittle fracture, means have been provided to estimate the CTOD or J values in a large

structure containing a defect, as a function of the overall applied loads. These estimated

values are then compared with the critical failure values for the relevant material. A well

known design curve for such fitness-for-purpose analysis is based on the CTOD approach

(Figure 4).

All the above procedures normally require that the product be tested with its full thickness

so as to derive a suitable toughness index. Although this condition may involve test

specimens having rather large dimensions, it can never be considered as appraising the

overall fracture behaviour of a structural component. Therefore, the relevance of the

transferability of the test data for the appraisal of large structures has to be verified by

comparing the computed fracture behaviour to that experimentally observed during tests

on a very large piece whose size is similar to the parts of a real structure.

The Wide Plate test has been designed with this aim. It involves tensile testing, possibly at

lowered temperature, of a wide specimen (1m wide for instance) containing deliberate

through-thickness or surface cracks. A convenient evaluation of performance in the Wide

Plate test is provided by the Yield criterion [5]. When full yield occurs, then all sections of


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the specimen, even those not affected by the defect, develop plastic strains so that the

overall elongation is sufficient to prevent a sudden failure. Full yield also ensures that the

structure can reach its maximum elastic design load, i.e. the product of the material yield

stress and the gross section, as if it were not affected by the defect, which is an obvious


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asset for safety. A critical defect length can be defined above which the criterion can no

longer be fulfilled; then only net section yielding or contained yielding are achieved. Table

2 summarises the main concepts relating to the Wide Plate test.

There are situations in which resistance to failure is not governed by toughness but by the

load bearing capability of the net section in the part affected by a defect. This situation

may occur with quite ductile materials affected by cracks. Plastic collapse corresponds to

the achievement of unlimited displacements in the net section when the applied load

induces a net section stress equal to the material flow stress. A ratio Sr may be defined to

express the safety against plastic collapse for a given loading condition. It is sometimes

more convenient to use the material yield stress as a reference and think in terms of plastic

yield load. A ratio Lr is then considered. Table 3 defines the parameters.

Although the assumptions associated with Linear Elastic Fracture Mechanics and Plastic

Collapse lead to quite opposite fracture mechanisms, both concepts may actually apply to

the same material and even the same structure depending on the defect size. Figure 5

illustrates this situation for the simple example of a wide plate containing a through-

thickness defect. For small defect sizes the lowest fracture resistance is dictated by plastic

collapse concepts, whereas for larger defects the lowest fracture resistance is obtained

from linear elastic theory.

This interactive behaviour supports the so-called two-criteria approach developed by

CEGB in 1976, and known as the R6 procedure, whose latest 3rd revision is now well

developed [6]. A Failure Assessment Diagram (FAD) expresses the risk of failure in a

two-dimensional space using the Sr or Lr parameter as the abscissa and a Kr variable as the

ordinate. Kr is the ratio between the applied and critical KI stress intensity factors. Safe


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and unsafe conditions are discriminated by a curve quantifying the interaction. The main

aspects of this method are summarised in Table 4.

4. METHODOLOGIES FOR STEEL SELECTION

Toughness properties of structural steels are generally classified in material standards such

as the new EN 10025 [7] and EN 10113 [8] in terms of performance in the Charpy V test.

While levels of absorbed energy may be defined for different test temperatures, a simple

and classical characteristic often encountered is the Transition Temperature at the level of

28J: TK28

Any methodology for steel selection applicable to standardised steel grades must involve

the following steps:

i. A definition of extreme service conditions against which to address the resistance to

fracture of a structure, i.e. a size of defect, a mode of loading (static or dynamic) and a

level of internal or external stress.

ii. A method of fracture analysis leading to the derivation of toughness requirements as a

function of the above conditions.

iii. A relationship between the toughness requirements and a transition temperature or

energy in the Charpy V test.

The different methods that are available are reviewed below.

4.1 The French Approach

This procedure was designed by Sanz and co-workers at IRSID in the late seventies,

elaborated by a Working Group of ATS and published in 1980 [9].

It is based on a set of experimental relationships between the transition temperatures in

KIC and Charpy V tests, which are illustrated in Figure 6, which is reproduced from the

original publication.

The analysis of the risk of brittle failure is based on Linear Elastic Fracture Mechanics,

and toughness requirements are derived in agreement with this theory. Some modifications

which were introduced for simplification purposes are highlighted in the document and

result in more conservative assumptions.

A particular feature of the method lies in the fact that defect sizes are defined in relation to

plate thickness so as to match with the conditions of plane strain and thus ensure the

applicability of a LEFM approach.

The main steps involved in the definition and application of the SANZ method are

summarised in Table 5. The final formula leading to the definition of steel quality in terms

of a TK28 index takes account of the service temperature and the strain rate to which the

structure may be exposed, as well as the scatter in the experimental relations.


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Some predictions of the necessary transition temperature at the Charpy V level of 28J are

illustrated in Table 6 for the condition of a structure operating at a minimum service

temperature of 0C and possibly loaded up to the material yield stress at a slow strain rate

(έ = 0,1s-1).

4.2 The British Approach

This method conceived by George in the United Kingdom was first proposed to the

International Institute of Welding in 1979 [10].

It relies on an elasto-plastic analysis of the fracture resistance using the CTOD design

curve for a structure containing a surface flaw of 0,2 times the plate thickness, in a field

having a high level of residual stress and loaded to 0,67 times the yield stress. The steel

quality is defined in terms of the required Charpy V energy at the service temperature

assuming a relationship between this energy and the critical CTOD value. The formula

derived in this way was calibrated against results from wide plate tests and practical

experience.

Table 7 reviews the main assumptions supporting the method and the mathematical

relationships which are derived.

Toughness requirements for a range of steel grades and thicknesses are listed in Table 8.

4.3 The Belgian Approach

Developed in the late 1980's, this method directly links the fracture behaviour of wide

plates containing through-thickness defects which are tensile tested at temperatures

between -10 and -50C, to the following metal characteristics: tensile strength or tensile-

to-yield ratio and Charpy V energy at the same temperature [11].

Net fracture stress of the wide plate is first expressed as a function of defect size in terms

of a linear decreasing function whose ordinate at the origin precisely equals the tensile

strength of the metal and whose slope is inversely proportional to the toughness expressed

by the Charpy V energy. Figure 7 illustrates these relationships. Net fracture stress is then

converted into gross section stress using the specimen width as a correction parameter. At

that stage the critical defect size according to the full yield criterion is computed as the

smaller positive root of a second order equation.

Table 9 lists the principal rules of the model which was originally developed to document

the fracture behaviour of parent metal and later extended to appraise butt welded wide

plates. In the latter case, local values of tensile strength and toughness of the metal at the

crack tip are considered by using Charpy specimens with the notch root positioned in

either the weld metal or the heat affected zone. Hardness measurements are carried out

and converted into a proper Rm when this parameter cannot be directly derived from a

tensile test. For plain or welded plates, the yield stress used in the formulae is always that

of the base metal away from the crack.


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5. METHODOLOGY ADOPTED IN EUROCODE 3

The method for steel selection adopted in Eurocode 3 [1] results from a combination and a

synthesis between the concepts proposed in the French method and the recommendations

of CEGB [6]. The Eurocode procedure is described in recent publications [12, 13] and can

be summarised as follows:


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This approach considers three alternative stressing conditions (S1, S2, S3), two possible

levels of strain rate (R1, R2) and two consequences of failure (C1, C2) which respectively

allow distinction between:

(a) Lower or higher stresses, monolithic elements or welded joints, stress-relieved or as-

welded structures, weaker or stronger effects induced by stress raisers.

(b) Quasi-static or rapid stressing conditions corresponding, on the one hand to permanent

loads, actions of wind, waves or traffic, and on the other hand to impacts, explosions,

collisions.

(c) Either ruptures leading only to localised damage not affecting human life and the

stability of the whole structure, or failure whose local occurrence implies disastrous

consequences for the global resistance of the structure or impairs the safety of people.

The Failure Assessment Diagram according to the R6 Rev. 3 [6] method which is taken

into consideration is that corresponding to the so-called Option 2 which takes account of

the actual tensile stress-strain curve of the steel. Figure 8 illustrates this diagram which in

Eurocode 3 terminates at an abscissa having a value of unity [1].

It is then supposed that the structure may contain semi-elliptical surface flaws whose size

is proportional to the natural logarithm of the thickness (depth=ln t, length = 5ln t).


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Critical Kr values are derived from the failure assessment diagram after computation of the

relevant Lr for the different stress conditions. The necessary toughness, KIC that the

material should display at the relevant service temperature can thus be readily defined.

The conversion of this requirement into a Charpy V TK28 transition temperature is finally

carried out according to the French procedure.

Table 10 summarises the major aspects of the successive derivations. It highlights the fact

that the strict application of this procedure for a given situation would first require a

detailed fracture analysis according to the two-criteria approach so as to derive the

necessary level of toughness required from the material in terms of a critical stress

intensity factor here denoted Kmat. Such an analysis, which must include the respective

contributions of the mechanical and residual stresses, as well as the correction factor for

crack tip plasticity generated by the residual stress, should be carried out carefully

according to a well-documented procedure such as that prescribed in British Standard

PD6493 [14]. The critical stress intensity factor at the minimum design service

temperature then has to be converted to a 28J transition temperature of a Charpy

specimen. Here the rules defined by the French method are followed [9] so as to account

for thickness and strain rate effects and the scatter in the relationship linking TKIC and

TK28.

The procedure presented above should only be applied by specialists in fracture

mechanics. On the other hand, the rules included in Eurocode 3 need to be applied by

design engineers who require a more convenient formulation. With this aim in mind, the

authors of the present Eurocode 3 rules have analysed a limited number of cases

simulating the various loading conditions for different plate thicknesses, so as to derive,

after a statistical analysis, a simplified formulation of the rules. The parameters of thee

rules for practical application are listed in Table 11.

As an illustration of the rules, Tables 12 and 13 list a set of requirements for different

values of yield stress and thickness corresponding to the S1, S2 and S3 load conditions as

well as the C1 and C2 failure consequences.

6. PRESENT STATUS OF THE EUROCODE 3 RULES

Establishing a common basis for steel selection applicable internationally as a substitute

for the existing national rules is a difficult task. Unifying a set of divergent specifications

leads to new rules that are inevitably either less or more constraining than those defined in

one or other of the national codes. On the one hand, concern may be raised against the

reliability of the new rules with regard to a safe design, while on the other hand, criticism

will be expressed towards an excessive degree of conservatism liable to spoil the

economic competitiveness of constructional steelwork.

The present rules in Eurocode 3 [1] have informative status. To encourage as much as

possible the concept of unified rules for steel selection, cooperation has been set up

between the CEN/TC 250/SC3 Committee of Eurocode 3 and the International Institute of

Welding who will study the problem on a multidisciplinary basis coordinated by

Commission X with the support of Commission IX. An introductory article prepared by


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Professor Burdekin, Chairman of Commission X, was presented at the IIW 1992 annual

assembly in this regard [15].

To contribute to the debate covering this important question, comparisons between the

toughness requirements derived from the different approaches mentioned in this paper

have been carried out and are presented below.

7. COMPARISON OF SPECIFICATIONS DERIVED

FROM VARIOUS APPROACHES

Available methods for steel selection differ in several respects: one of the most apparent is

the form of the Charpy requirement which can be expressed either as a transition

temperature or as a level of energy. Suitable conversion formulae are needed so as to

perform comparisons between such methods.

Another factor of divergence is the definition of the flaw size as a function of plate

thickness. Account must be taken of these differences when attempting to compare the

various requirements.

Nevertheless, the main difference between the different approaches may arise from the

assumptions which are made concerning the definition of the fracture model. The French

approach is based on a purely linear elastic analysis, while the British and Belgian ones

assume respectively elasto-plastic and full plastic behaviour of the structure. Another

source of discrepancy could arise from the fact that the available methods were set up at

different times and thus were calibrated on populations of steels which could differ widely

in terms of chemical composition and processing route.

A major result that should emerge from the comparative exercises is therefore verification

of whether the different concepts governing the existing approaches are able to generate

consistent conclusions or, on the contrary, lead to contradictory outputs. It is clear that, if

the latter situation should apply, little confidence could be shown in those models. Their

significance would be restricted to the limited set of conditions and the generation of steels

which were used during their formulation.

If consistent specifications can be derived from several models, another dimension would

be conferred to the problem since, not only the overall reliability of the specifications

would be enhanced, but sources of minor deviations could be better documented.

A significant step in highlighting the possible coherence of Charpy V based specifications

was achieved by de Meester who presented valuable comparisons between the French and

British approaches in 1986 [16]. An illustration of that work is reproduced in Figure 9.

A similar approach can be used to expand this evaluation to other methodologies. It will

be recalled that some procedures express the necessary toughness in terms of a transition

temperature while others require a level of energy at the service temperature. Means of

conversion are, therefore, necessary in such cases. Table 14 reproduces the equations

which were adopted in [16] on the basis of extensive comparisons reported in the 1970's

[17].


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Further information is necessary in order to make comparisons with the Belgian method

since this procedure requires, as an entry parameter, the tensile-to-yield stress ratio of the

relevant steel. This parameter depends, among other things, upon the processing route

undergone by the material. A characteristic value cannot therefore be defined for a given

steel grade. Typical values may, however, be considered on a statistical basis as a function

of the guaranteed yield stress, such as those illustrated in Figure 10 which were proposed

by Dahl and his co-workers [18]. Such a relationship is used in the present analysis to

establish the necessary comparisons. Taking into account that thicker plates generally

display a higher Rm/Re ratio than thinner products, data corresponding to the lower

boundary or those closer to the upper side of the relationship of Figure 10 are selected

depending on the thickness.

Table 15 lists the toughness requirements obtained from the Belgian method applied with

same defect sizes as those considered in the French procedure (cf Table 6). TK28

transition temperatures required by both models are compared in Figure 11, which

highlights the coherence of the respective specifications, especially in the significant field

of negative transition temperatures, which represent the most severe conditions to be

fulfilled by the material.


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A comparison between the French and British methods is provided in Figure 12 using data

which were generated in Tables 6 and 8. Here again specifications from both models are

consistent. It will, however, be noticed that the French procedure considers smaller defect

sizes (as a function of thickness) than the British method, for instance 8mm against 12mm

for a 30mm thick member. This means that, for the same flaw size, the elasto-plastic

fracture analysis developed by George on the basis of the design curve and Charpy COD

relationships would lead to steel toughness requirements that were somewhat less stringent

than those defined by the linear elastic approach of Sanz and the TKIC-TK28 relationships.

A similar conclusion is reached when roughly comparing, on a same defect size basis, the

British rules to the Belgian procedure. Considering, however, that the George model

assumes a design stress of only 2/3 of the yield stress and implementing this stress level in

the CRM model, quite consistent requirements between both methods would be obtained.

This is highlighted in Figure 13 which is drawn with data extracted from Tables 8 and 16.


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Requirements derived from the present Eurocode rules are compared to those from the

French rules using the data listed in Tables 6 with Table 17 (S1 loading and failure

condition C1) and Table 18 (S3 loading and failure condition C2). It is important to note

that for the sake of consistency regarding the effect of strain rate, the same value of 0,1s-1

has been adopted for all procedures. Figures 14 and 15 illustrate that, depending on the

loading conditions and failure consequences, Eurocode 3 rules are either less stringent or,

on the contrary, significantly more constraining than the specifications of the French

standard from which they are partly derived.


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8. DISCUSSION

The comparison exercises reported above have demonstrated that, although the models

were derived from quite different basic assumptions and fracture concepts, and were

validated at different periods on different steel qualities and steel generations, the French,

British and Belgian models lead to consistent Charpy requirements. This agreement is

reached when the procedures are compared to each other on a carefully balanced basis, i.e.

adopting the same defect size as a function of plate thickness and the same stress level.

The strain rate is an important factor, which is an explicit parameter in the French method

but not in either of the others. The Belgian model does, however, take some account of

this effect through the tensile-to-yield ratio which is influenced by the strain rate

sensitivity of the material. All comparisons with the French model were carried out for the

slow strain rate of 0,1s-1 with a view to improving consistency. It is nevertheless clear that

the strain rate sensitivity may vary, not only as a function of the steel grade, but also

according to the applied processing route or chemical composition. This parameter would

certainly be worth being better documented so as to optimise the rules for steel selection in

certain applications involving dynamic effects.

The major sources for possible divergence between the existing rules have been identified.

Many specifications are based on their conventions regarding the evolution of the

admissible defect as a function of member thickness and the prevailing state and level of

stress. This results in the definition of less or more stringent requirements. Such a situation

becomes disturbing and confusing when the procedure that the fabricator has to follow is

not properly documented in those terms, or when the computational steps are complicated

and do not easily provide the possibility of carrying out even a limited parametric analysis.

Simple rules involving clearly defined models and based on rigorous mathematical

treatments, ideally developed into analytical formulae, should constitute a preferred choice

in the establishment of steel selection criteria. On the other hand, complex methodologies

involving advanced concepts of fracture analysis may be misleading since their practical

application would either be too difficult or simplification of the procedure would result in

the infringement of basic fracture mechanics rules.

9. COMMENT ON THE PRESENT EUROCODE 3

RULES

Eurocode 3 rules have been reviewed as well as the philosophy that prevailed in their

development. Comparisons with the French model on which they are partly based, reveal

overall agreement in a large range of transition temperatures, together with the possibility

of shifting widely the requirements depending on the stress conditions (3 levels), the strain

rate (2 conditions), and the consequences of failure (2 conditions): Tables 12 and 13 show

that between S1R1C1 and S3R2C2 conditions, the difference in required TK28

temperatures is equal to or greater than 90C.

These rules still require improvement and need to be discussed within a large forum of

specialists. The cooperation of the International Institute of Welding in this task will bring

a worldwide dimension to the challenge of unifying the rules for steel selection. Initial


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proposals for a coordinated approach to the problem were formulated at the IIW annual

assembly of 1992 [15].

It may be of interest in this regard to quantify roughly how each of the available models

accounts for variations of the plate thickness or of the steel yield stress on the TK28

requirements, all other factors being unchanged. In practice a significant question is to

evaluate the advantages of implementing in a given type of structure, higher strength steels

with thinner gauge as an alternative to conventional grades in thick sections. With this

aim, the TK28 requirements listed in the various Tables 6, 8, 12, 13 and 15 have been

correlated by linear regression analysis to thickness and yield stress so as to derive the

following expression for each model:

TK28 = a - b.e - c.Re

The following b and c coefficients were derived:

French model b = -0,64 c = -0,12

British model b = -0,99 c = -0,12

Belgian model b = -0,85 c = -0,12

Eurocode model b = -0,54 c = -0,15

These values indicate that the French, British and Belgian rules seem more favourable for

the adoption of higher strength steels and thinner gauges than the Eurocode specifications.

10. CONCLUDING SUMMARY

Consistent requirements in terms of a transition temperature in the Charpy V test

may be derived from any one of three national methodologies for steel selection,

each incorporating different concepts of fracture behaviour or analysis, established

at different periods and validated with different grades or generations of steels.

Discrepancies which may appear are more the result of different conventions that

have been adopted for the definition of stress level, safety factors or defect size.

By unifying on a European or wider international basis, criteria for steel selection

would be promoted by a larger analysis of the available models and an exploitation

of their consistency.

11. REFERENCES

[1] Eurocode 3: "Design of Steel Structures": ENV 1993-1-1: Part 1.1, General rules and

rules for buildings, CEN, 1992.

[2] Brock, D., Elementary engineering fracture mechanics, Martinus Nijhorff Publishers,

1987.

[3] Garwood, S. J., A crack tip opening displacement (CTOD) method for the analysis of

ductile materials, ASTM STP 945, June 1985.


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[4] Rice, J. R., A path independent integral and the approximate analysis of strain

concentrations by notches and cracks, Trans. ASME, J. Appl. Mech. 1968, 35 379-

386.

[5] Soete, W. and Denys, R., Evaluation of butt welds based on a strain criterion, Revue

de la Sodure, Lastijd schrift, No. 4, 1975.

[6] Milne, I., Ainsworth, R. A., Dowling, A. R., Stewart, A. T., Assessment of the

integrity of structures containing defects, CEGB document R/H/R6 - Revision 3, May

1986.

[7] EN 10025: Hot Rolled Products of Non-alloy Structural Steels and their Technical

Delivery Conditions, British Standards Institution, 1990.

[8] EN 10113-1: Hot Rolled Products of Weldable Fine Grain Structural Steels Part 1:

General Delivery Conditions, British Standards Institution, 1993.

ENV 10113-2: Hot Rolled Products of Weldable Fine Grain Structural Steels Part 2:

Delivery Conditions for Normalized/Normalized Rolled Steels, British Standards

Institution, 1993.

ENV 10113-3: Hot Rolled Products of Weldable Fine Grain Structural Steels Part 2:

Delivery Conditions for Normalized/Normalized Rolled Steels, British Standards

Institution, 1993.

[9] Sanz, G., Essai de mise au point d'une mthode quantitative de choix des aciers vis-á-

vis du risque de rupture fragile, Revue de Mtallurgie - CIT, Juillet 1980.

[10] George, M., A method for steel selection, Document IIW-IXF.

[11] Defourny, J., D'Haeyer, R., Leroy, V., A metallurgical approach of the parameters

affecting the fracture behaviour of base metal and welded components, IIW

document IX-1607-90/X-1206-90.

[12] Sedlacek, G., Bild, J., Hensen, W., Background document for Chapter 3 of Eurocode

3 "Design Against Brittle Fracture", Aachen 1990.

[13] Brozzetti, J., Sedlacek, G., Hensen, W., Fondements des regles de l'Eurocode 3 en

vue de se garantir du risque de rupture fragile, Construction Mtallique, no. 1, 1991.

[14] Guidance on methods for assessing the acceptability of flaws in fusion welded

structures, PD6493: 1991, BSI.

[15] Burdekin, F. M., Materials selection for welded structural steelwork in Engineering

Design in welded constructions, Pergamon Press, 1992.

[16] de Meester, B., The brittle fracture safe design of welded constructions, Welding in

the world.

[17] Sanz, G., La rupture des aciers, fascicule 2, Collection IRSID OTUA.

[18] Dahl, W., Hesse, W., Krabiell, A., Zur WVerfestigung von Stahl und dessen Einflu

auf die Kennwerte des Zugversuchs; Stahl und Eisen 103 (1983), Heft 2, Seite 87-90.


570

Table 1 Linear Elastic Fracture Mechanics

The three modes of loading at a crack tip:

The stress field around the crack is defined by a parameter K:

u = [KI )].fu) in mode I

Infinite plate:

KI = a)

Finite plate of width W:

KI = a).[sec a/W]

Through thickness crack in Mode I


571

Semi-elliptical crack:

KI = 1,12a/[2 - 0,212(/Re)2]}

= 1 - [(c2-a2)/c2].sin2d - 3/8 + a2/8c2


572

Table 2 Concepts Associated with the Wide Plate Test

Gross section stress: b = F/(W t)

Net section stress: n = F/[(W - 2a) t]

b = n[1 - 2a/W]


573

Full yield : b = Re

General yield : n = Re

Determination of the critical defect length at full yield. Different wide plate tests are

carried out at the same temperature but with different crack sizes.


574

Table 3 Plastic Collapse Concept

According to the plastic collapse concept, the load bearing resistance of the structure

affected by a defect equals the product of the material flow stress by the net section.

The diagram below plots the gross stress at plastic collapse in a plate containing a

through-thickness defect as a function of the 2a/W ratio (defect length to plate width).

Re : material yield stress

Rf : material flow stress

Rm : material tensile strength

Re < Rf Rm

Rf ~ (Rm + Re) / 2

b : gross stress applied to the wide plate

Sr = applied load/plastic collapse load = (b / Rf) (1-2a/W)

Lr = applied load/plastic yield load = (b / Re) (1-2a/W)

At plastic collapse

Sr = Srmax = 1 Lr = Lr

max = Rf/Re


575

Table 4 Main Rules of the R6 Procedure

Box 1:

The R6 Rev. 3 Failure Assessment Diagram

The simplifying concepts behind the R6 procedure

are the two criteria for failure, characterised by:

1. Crack tip failure, where failure occurs when the

applied load equals the LEFM failure load,

(1)

2. Failure by plastic collapse, defined when the load

equals the plastic collapse load and the

displacements in the structure become unlimited. The

plastic collapse load may be formally defined as the

load when the reference stress reaches the material's

flow stress, . The criterion for plastic collapse failure

is given by the ratio Sr.

where Sr =

(2)

and this criterion was used for the initial R6 FAD.

It is often more convenient to think in terms of

plastic yield loads rather than collapse loads as it is

easier to define and calculate these theoretically. In

R6 Rev. 3 the ratio Lr is used as

Lr = (3)

From this, the load for plastic collapse is given as

Lrmax.

With (4)

The simplest FAD requires that Kr < 1 and Lr < Lrmaz.

This ignores any interaction between the two

failure mechanisms. This interaction may be

allowed for by writing the first inequality as Kr <

f(Lr) and choosing an appropriate function for f.

One such function is

f(Lf) = (5)

where Je is the elastic component of J given by

K2/E. Thus when ref is low, J = Je and f(Lr) = 1,

consistent with LEFM failure. When ref is high,

on the other hand, failure is governed by J.

Equation 5 may be used to generate an FAD

directly from the load displacement curves of

specimen tests.

The most general method for calculating equation 5

is to use the reference stress procedures so that

f(Lr) may be calculated directly from the material's

stress strain curve. The general equation for this is

f(Lf) = (6)

where ref is taken from the material's true stress -

true strain curve at respective values of ref. The

figure 6 is a plot of equation 6 as a function of Lr,

using an experimental stress strain curve for a

stainless steel.

Although equation 6 is the preferred option for R6

Rev. 3, it requires a detailed knowledge of the

stress strain curve of the material, especially

around the material's yield point. In some cases,

especially when dealing with hold plant, such

knowledge is not available. Other options of the

FAD have been developed for use in such

circumstances.


576


577

Table 5-1 Main Rules of the French Model

1. BASIC RELATIONSHIP

Between KIC and Charpy V based on Transition Temperature at respectively

100MPa m and 28 J:

(S1)

2. APPLICABILITY

Plane strain condition,

(S2)

where e - is in metres

3. APPLIED STRESS

Material yield stress Re, then

(S3)

where a - is in metres

and from Equation (S2)

(S4)

4. REFERENCE DEFECT SIZE

For thick plates, a semi-elliptic surface defect 50mm long, 25mm deep, which is

equivalent to a 28mm long through thickness defect (a = 14mm).

5. REFERENCE YIELD STRESS

477 MPa is the yield stress level derived from S3 with a = 14mm and KIC = 100

MPa m.

6. REFERENCE THICKNESS

110mm is the thickness satisfying Equation (S4) with a = 14mm.


578


7. REQUIRED KIC VALUES vs YIELD STRESS

For thick products (e 110mm), a 28mm long defect corresponds to 100 MPa m at a

yield stress of 477 MPa. For other values of Re, the following KIC values apply, as

derived from Equation (S3):

Re (MPa) 280 350 410 470 510 640 740

KIC (MPa m) 59 73 86 99 107 134 155

8. DEFECT LENGTH AS A FUNCTION OF THICKNESS

To satisfy Equation (S4) the defect length is related to plate thickness as indicated by the

following table. This table also lists the corresponding KIC values at a reference yield

stress of 477 MPa.

e (mm) 10 20 30 60 80 110 >110

2a (mm) 2,5 5 7,5 15 20 28 28

KIC (MPa m) Re =

477 MPa

30 43 52 74 85 100 100

9. MAJOR RELATIONSHIPS

Taking account of the above tables, it is necessary to consider transition temperatures for

KIC values other than 100, but ranging between 30 and 160MPam. This can be quantified

through the following equation:

(S5)


579


10. RELATION BETWEEN SERVICE TEMPERATURE AND 28J TRANSITION

TEMPERATURE

Equation (S1) is used to relate a minimum service temperature above which the structure

is safe to a required toughness level expressed as a transition temperature of 28J in the

Charpy test.

The minimum service temperature is equal to TKIC = 100 for a structure made of a steel with

reference yield stress 477 MPa, reference thickness 110mm and affected by the reference

defect length a = 14mm.

For other values of yield stress or thickness, other values of KIC apply and corrections are

introduced through Equation (S5).

Corrections are also made to take account of strain rate effects and the scatter in Equation

(S1).

The final relationship is expressed as follows:

(S6)

in which

TS : minimum service temperature for fracture safe design

TK28 : transition temperature of Charpy V at 28J

(Re) : 60 ln rounded to 60 ln

Te : in theory, 60 1n (e in mm); in the model a more

conservative formula is used : 57 ln

TV : (83 - 0,08 Re) 0,17 where is the strain rate in s-1

25 : scatter of Equation (S1)


580


11. DEFINITION OF TOUGHNESS REQUIREMENTS

The necessary steel quality expressed in terms of the 28J transition temperature is derived

as a function of service temperature, strain rate, stress applied to the structure, material

yield stress and thickness from the following formula:


581

Table 6: Some Results of the French Method

Re e 2a Te Tv TK28

@Ts=0C

(MPa) (mm) (mm) (C) (C) (C)@0,1/s (C)

280 10 3 -49 -92 41 53

280 20 5 -49 -73 41 39

280 30 8 -49 -58 41 29

280 60 15 -49 -29 41 8

280 80 20 -49 -16 41 -1

280 110 28 -49 0 41 -12

350 10 3 -27 -92 37 40

350 20 5 -27 -73 37 27

350 30 8 -27 -58 37 16

350 60 15 -27 -29 37 -4

350 80 20 -27 -16 37 -14

350 110 28 -27 0 37 -25

410 10 3 -13 -92 34 33

410 20 5 -13 -73 34 19

410 30 8 -13 -58 34 9

410 60 15 -13 -29 34 -12

410 80 20 -13 -16 34 -22

410 110 28 -13 0 34 -33

480 10 3 0 -92 30 26

480 20 5 0 -73 30 12


582

Table 6 (continued): Some Results of the French Method

480 30 8 0 -58 30 2

480 60 15 0 -29 30 -19

480 80 20 0 -16 30 -28

480 110 28 0 0 30 -39

510 10 3 5 -92 29 24

510 20 5 5 -73 29 10

510 30 8 5 -58 29 0

510 60 15 5 -29 29 -21

510 80 20 5 -16 29 -30

510 110 28 5 0 29 -42

Conditions: loading up to yield stress at a slow strain rate (0,1/s)


583

Table 7 Main Rules of the British Model


Between the critical crack opening displacement ( in mm) of a material at a given

temperature and the Charpy V energy (CV in Joules) at the same temperature:

(G1)

2. APPLICABILITY

Fracture behaviour analysis according to the assumptions that prevail for the definition of

the "Design Curve" in British Standard PD 6493, which for stainless and ferritic steel is

expressed as follows:

(G2)

3. APPLIED STRESS

In the original model, a value of equal to 1,9 Re was adopted as the result of the

superimposition of a design stress equal to 0,67 Re affected by a stress concentration

factor of 1,2 and a residual stress of amplitude equal to Re.

(G3)

4. REFERENCE DEFECT LENGTH

A surface crack size 0,2 times the plate thickness deep and 1 times the thickness long,

which corresponds to an equivalent crack size of 0,2e

(G4)


Combining Equations (G1) to (G4) leads to:

(G5)

equivalent to:

(G6)


584

Table 8 Some Results of the British Method

Re e 2a CV TK28@ Ts=0C

(MPa) (mm) (mm) (J) (C) #

280 10 4 4 106

280 20 8 8 58

280 30 12 12 37

280 60 24 24 6

280 80 32 32 -5

280 110 44 43 -17

350 10 4 5 88

350 20 8 10 46

350 30 12 15 27

350 60 24 30 -3

350 80 32 39 -14

350 110 44 54 -26

410 10 4 6 77

410 20 8 12 38

410 30 12 17 20

410 60 24 35 -9

410 80 32 46 -20

410 110 44 64 -31

480 10 4 7 67

480 20 8 14 31


585

Table 8 (continued): Some Results of the British Method

480 30 12 20 13

480 60 24 41 -15

480 80 32 54 -26

480 110 44 74 -37

510 10 4 7 64

510 20 8 14 28

510 30 12 22 10

510 60 24 43 -17

510 80 32 57 -28

510 110 44 79 -39

#TK28 was derived from CV using formula (ET5) of Table 14


586

Table 9 Main Rules of the Belgian Model


Between net fracture stress in wide plate test containing a through-thickness defect

and Charpy V energy at the same temperature:

(C1)

where CV - is expressed in J and not in J/cm² as in the original model.

2. APPLICABILITY

Fracture appearance transition temperature in the Charpy test is lower than the

temperature of the wide plate test.

3. APPLIED STRESS

Material yield stress on gross-section (the model may nevertheless be applied to

other stresses).

(C2)

4. REFERENCE DEFECT

Length of the through-thickness defect (critical size) satisfying the full yield

behaviour, expressed by the following equation:

(C3)


(C4)


587

Table 10-1 Main Aspects of the Eurocode Rules

Step 1.1 Definition of Stress Levels S1, S2, S3:

S1 - Structural elements containing no weld, or

- stress-relieved welded elements with tensile stresses under 2/3 of yield stress, or

- as-welded element with tensile stresses under 2/10 of yield stress.

S2 - As-welded elements with tensile stresses between 2/10 and 2/3 yield stress, or

- stress-relieved welded elements with stresses below 2 times yield stress.

S3 - As-welded elements with stresses between 2/3 and 2 times yield stress, or

- stress-relieved welded elements with stresses between 2 and 3 times yield

stress.

Note: Stresses here include local stresses concentrated by stress raisers.

(E1)

Step 1.2 Definition of Strain Rates R1, R2

R1 : έ 0,001 s-1 (permanent loads, traffic loads, winds, waves, material handling)

R2 : έ 1 s-1 (impacts, explosions, collisions)

(E2)


588


Step 1.3 Definition of Consequences of Failure C1, C2:

C1 : Localised failure without appreciable consequence on safety of persons and

stability of structure.

C2 : Failure whose local occurrence may cause the global collapse of the structure with

disastrous consequences for persons and economy.

(E3)

Step 2 Definition of Defect Size:

- Semi elliptical surface defect

- Depth (a) equal to natural logarithm of thickness (t)

- Length (2c) equal to 5 times depth:

a = ln (t) (mm)

2c = 5 ln (t) (mm)

(E4)


589


Step 3 Computation of Lr and Kr:

Lr is computed as a function of defect size and stress level.

Kr is computed as a function of Lr from the Fracture Assessment Diagram.

(E5)

Step 4 Computation of Kmat

Kmat is the necessary toughness that must be shown by the material.

Kmat is derived from the KI corresponding to the loading conditions imposed on the

structure and the Kr value derived according to (E5).

A safety factor () is introduced to account for the failure condition:

= 1 for C1

= 1,5 for C2

(E6)

Step 5 Derivation of TK28

The rules of the French method are followed but are translated for KIC values in

N/mm3/2 instead of MPa m

(E7)


590

Table 11-1 Application Parameters of the Eurocode Rules

1. EXPRESSION OF KIC

KIC is the toughness required from the material taking account of its thickness and the

service conditions prevailing for the structure.

(E'1)

Re is the material yield stress guaranteed by the standard for the required thickness (t).

Units: KIC in N/mm3/2

Re in MPa

t in mm

(E'2)

= 1 for condition C1

= 1,5 for condition C2

(E'3)

ln - is natural logarithm

Stress level

S1 S2 S3

ka 0,18 0,18 0,10

kb 0,40 0,15 0,07

kc 0,03 0,03 0,04


591

Table 11-2 Application Parameters of the Eurocode Rules

1. EXPRESSION FOR TK28

(E'4)

TS : Service temperature of the structure


592

Table 12 Some Requirements Derived from the Eurocode Rules

Re

(MPa)

t

(mm)

K1S1

(N/mm3/2)

K1S2

(N/mm3/2)

K1S3

(N/mm3/2)

1

(C)

2

(C)

3

(C)

Tv

(C)@0,001/s

TK28@Ts=0C

(C)@S1

TK28@Ts=0C

(C)@S2

TK28@Ts=0C

(C)@S3

280 10 655 939 1216 -158

-121 -96 19 81 56 37

280 20 813 1185 1514 -

136 -98 -74 19 66 39 21

280 30 933 1364 1726 -

122 -84 -61 19 56 29 12

280 60 1192 1744 2159 -98 -60 -38 19 39 11 -4

280 80 1324 1933 2368 -87 -49 -29 19 31 4 -11

280 110 1489 2167 2622 -75 -38 -19 19 23 -4 -18

350 10 818 1174 1520 -

135 -99 -73 17 67 41 22

350 20 1017 1481 1893 -

114 -76 -51 17 51 24 7

350 30 1166 1705 2157 -

100 -62 -38 17 41 14 -3

350 60 1490 2180 2698 -75 -37 -16 17 24 -3 -19

350 80 1654 2416 2960 -65 -27 -7 17 16 -11 -25

350 110 1861 2709 3277 -53 -16 3 17 8 -19 -32

410 10 958 1375 1781 -

119 -83 -58 16 56 31 12

410 20 1191 1735 2218 -98 -60 -36 16 41 14 -4

410 30 1366 1997 2527 -84 -46 -23 16 31 4 -13

410 60 1745 2553 3161 -60 -21 0 16 14 -14 -29

410 80 1938 2830 3467 -49 -11 9 16 6 -21 -35

410 110 2180 3173 3839 -37 0 19 16 -2 -29 -43

480 10 1122 1610 2085 -

104 -68 -42 14 46 21 2

480 20 1395 2031 2596 -82 -44 -20 14 31 4 -14

480 30 1599 2338 2958 -68 -30 -7 14 21 -6 -23

480 60 2043 2989 3700 -44 -6 16 14 4 -24 -39

480 80 2269 314 4059 -33 5 25 14 -4 -31 -45

480 110 2552 3715 4494 -22 16 35 14 -12 -39 -53

510 10 1192 1711 2215 -98 -62 -36 13 43 17 -2

510 20 1482 2158 2759 -76 -38 -14 13 27 0 -17

510 30 1699 2484 3143 -62 -24 -1 13 17 -10 -27

510 60 2171 3176 3932 -38 0 22 13 0 -27 -43

510 80 2411 3521 4313 -27 11 31 13 -8 -35 -49

510 110 2711 3947 4775 -15 22 41 13 -16 -43 -57

Conditions: S1, S2, S3 loading modes at a R1 strain rate (0,001/s) and C1 failure consequences.


593

Table 13 Some Requirements Derived from The Eurocode Rules

Re

(MPa

)

t

(mm)

K1S1

(N/mm3/2

)

K1S2

(N/mm3/2

)

K1S2

(N/mm3/2

)

1

(C)

2

(C)

3

(C)

Tv

(C)@0,001/

s

TK28@Ts=0C

(C)@S1

TK28@Ts=0C

(C)@S2

TK28@Ts=0C

(C)@S3

280 10 818 1174 1520 -135 -99 -73 61 36 10 -9

280 20 1017 1481 1893 -114 -76 -51 61 20 -7 -24

280 30 1166 1705 2157 -100 -62 -38 61 10 -17 -34

280 60 1490 2179 2698 -75 -37 -16 61 -7 -34 -50

280 80 1654 2416 2959 -65 -27 -7 61 -15 -42 -56

280 110 1860 2708 3277 -53 -16 3 61 -23 -50 -64

350 10 1023 1467 1900 -113 -77 -51 55 24 -2 -21

350 20 1271 1851 2366 -91 -54 -29 55 8 -19 -36

350 30 1457 2131 2696 -78 -40 -16 55 -2 -29 -46

350 60 1862 2724 3372 -53 -15 6 55 -19 -46 -62

350 80 2068 3020 3699 -43 -5 16 55 -27 -54 -68

350 110 2326 3386 4096 -31 7 26 55 -35 -62 -76

410 10 1198 1719 2226 -97 -61 -35 50 16 -10 -29

410 20 1489 2168 2772 -75 -38 -13 50 0 -27 -44

410 30 1707 2496 3158 -62 -24 0 50 -10 -37 -54

410 60 2181 3191 3950 -37 1 22 50 -27 -54 -70

410 80 2422 3538 4333 -27 11 31 50 -35 -62 -76

410 110 2724 3966 4798 -15 23 42 50 -43 -70 -83

480 10 1402 2012 2606 -81 -45 -19 45 8 -17 -36

480 20 1743 2538 3245 -60 -22 2 45 -7 -34 -51

480 30 1999 2922 3697 -46 -8 16 45 -17 -44 -61

480 60 2554 3736 4625 -21 17 38 45 -34 -62 -77

480 80 2836 4142 5073 -11 27 47 45 -42 -69 -83

480 110 3189 4643 5617 1 38 57 45 -50 -77 -91

510 10 1490 2138 2769 -75 -39 -13 42 6 -20 -38

510 20 1852 2697 3448 -54 -16 9 42 -10 -37 -54

510 30 2124 3105 3928 -40 -2 22 42 -19 -47 -63

510 60 2713 3970 4914 -15 23 44 42 -37 -64 -79

510 80 3013 4400 5390 -5 33 53 42 -44 -72 -86

510 110 3389 4933 5968 7 44 63 42 -53 -80 -93

Conditions: S1, S2, S3 loading modes at a R1 strain rate (0,001/s) and C2 failure consequences.


594

Table 14-1 Conversion Between Energy and Transition Temperature in

the Charpy V Test

In the French method, the dependency between KIC and the transition temperature is

defined through a set of experimental results which are the basis of the various

correlations between TK28 and TKIC. This relationship can be expressed by an analytical

formula as follows:

(ET1)

where KIC - is expressed in MPa m .

T - is the difference between the temperatures at which KIC corresponds

respectively to a given value and 100 MPa m

In the same method, it is also considered that the Charpy V energy and the KIC value are

linked by the following relationship:

CV = (ET2)


595

Table 14.2 : Conversion Between Energy and Transition Temperature in

the Charpy V Test

Combining Equations (ET1) and (ET2) leads to:

(ET3)

where T - is the difference between the temperature at which CV must be calculated,

e.g. the service temperature, and TK28.

Thus, it becomes:

(ET4)

where

TS - is the service temperature

TK28 - is the transition temperature at 28 J

For a service temperature of 0C:

(ET5)


596

Table 15: Some Results of the Belgian Method for Comparison with

the French Procedure

Re

(MPa)

e

(mm)

2a

(mm)

Rm/Re CV

(J)

TK28

(C)

280 10 3 1,5 6 77

280 20 5 1,5 10 47

280 30 8 1,5 16 24

280 60 15 1,55 28 0

280 80 20 1,55 37 -12

280 110 28 1,55 53 -25

350 10 3 1,4 7 67

350 20 5 1,4 11 39

350 30 8 1,4 18 17

350 60 15 1,45 32 -6

350 80 20 1,45 43 -17

350 110 28 1,45 62 -30

410 10 3 1,3 8 55

410 20 5 1,3 14 29

410 30 8 1,3 23 8

410 60 15 1,35 39 -13

410 80 20 1,35 52 -24

410 110 28 1,35 75 -37

480 10 3 1,23 10 44

480 20 5 1,23 17 19


597

Table 15: (continued) Some Results of the Belgian Method for Comparison with

the French Procedure

480 30 8 1,23 28 -1

480 60 15 1,28 46 -20

480 80 20 1,28 63 -31

480 110 28 1,28 91 -44

510 10 3 1,2 12 38

510 20 5 1,2 20 14

510 30 8 1,2 32 -6

510 60 15 1,25 51 -23

510 80 2- 1,25 70 -35

510 110 28 1,25 101 -48


598

Table 16: Some Results of the Belgian Method for Comparison with

the British Procedure

Re

(MPa)

e

(mm)

2a

(mm)

Rm/Re CV

(J)#

280 10 4 1,5 5

280 20 8 1,5 9

280 30 12 1,5 14

280 60 24 1,55 27

280 80 32 1,55 37

280 110 44 1,55 51

350 10 4 1,4 5

350 20 8 1,4 10

350 30 12 1,4 15

350 60 24 1,45 29

350 80 32 1,45 39

350 110 44 1,45 54

410 10 4 1,3 5

410 20 8 1,3 11

410 30 12 1,3 16

410 60 24 1,35 31

410 80 32 1,35 42

410 110 44 1,35 58

480 10 4 1,23 6

480 20 8 1,23 11


599

Table 16 (continued): Some Results of the Belgian Method for Comparison with

the British Procedure

480 30 12 1,23 17

480 60 24 1,28 33

480 80 32 1,28 44

480 110 44 1,28 62

510 10 4 1,2 6

510 20 8 1,2 12

510 30 12 1,2 18

510 60 24 1,25 34

510 80 32 1,25 46

510 110 44 1,25 64

# Required Charpy Energy is here computed assuming that the design stress is equal

to 2/3 Re so as to fit with the British model.


600


601


602


603


604


605

Table 17: Requirements from Eurocode 3 at 0,1/s Strain Rate and C1 Failure Consequences

Re

(MPa)

t

(mm)

K1S1

(N/mm3/2)

K1S2

(N/mm3/2)

K1S3

(N/mm3/2)

1

(C)

2

(C)

3

(C)

Tv

(C)@0,1/s

TK28@Ts=0C

(C)@S1

TK28@Ts=0C

(C)@S2

TK28@Ts=0C

(C)@S3

280 10 655 1174 1520 -

135

-99 -73 41 50 24 5

280 20 1017 1481 1893 -

114

-76 -51 41 34 7 -10

280 30 1166 1705 2157 -

100

-62 -38 41 24 -3 -20

280 60 1490 2179 2698 -75 -37 -16 41 7 -20 5

280 80 1654 2416 2959 -65 -27 -7 41 -1 -28 -10

280 110 1860 2708 3277 -53 -16 3 41 -9 -36 -20

350 10 818 1174 1520 -

135

-99 -73 37 52 26 8

350 20 1017 1481 1893 -

114

-76 -51 37 37 10 -8

350 30 1166 1705 2157 -

100

-62 -38 37 27 0 -17

350 60 1490 2180 2698 -75 -37 -16 37 9 -18 -33

Table 17 (continued): Requirements from Eurocode 3 at 0,1/s Strain Rate and C1 Failure Consequences

350 80 1654 2416 2960 -65 -27 -7 37 2 -25 -40

350 110 1861 2709 3277 -53 -16 3 37 -6 -33 -47

410 10 958 1375 1781 -

119

-83 -58 34 43 17 -1

410 20 1191 1735 2218 -98 -60 -36 34 28 1 -17

410 30 1366 1997 2527 -84 -46 -23 34 18 -9 -26

410 60 1745 2553 3161 -60 -21 0 34 0 -27 -42

410 80 1938 2830 3467 -49 -11 9 34 -7 -34 -49

410 110 2180 3173 3839 -37 0 19 34 -15 -42 -56

480 10 1122 1610 2085 -

104

-68 -42 30 35 9 -10

480 20 1395 2031 2596 -82 -44 -20 30 19 -8 -25

480 30 1599 2338 2958 -68 -30 -7 30 9 -18 -35

480 60 2043 2989 3700 -44 -6 16 30 -8 -35 -51

480 80 2269 3314 4059 -33 5 25 30 -16 -43 -57

480 110 2552 3715 4494 -22 16 35 30 -24 -51 -64


510 10 1192 1711 2215 -98 -62 -36 29 32 6 -13

510 20 1482 2158 2759 -76 -38 -14 29 16 -11 -28

510 30 1699 2484 3143 -62 -24 -1 29 6 -21 -38

510 60 2171 3176 3932 -38 0 22 29 -11 -38 -54

510 80 2411 3521 4313 -27 11 31 29 -19 -46 -60

510 110 2711 3947 4775 -15 22 41 29 -27 -54 -68

Conditions: S1, S2, S3 loading modes at a slow strain rate (0,1/s) and C1 failure consequences.

Table 18 Requirements from Eurocode 3 at 0,1/s Strain Rate and C2 Failure Consequences

Re

(MPa)

t

(mm)

K1S1

(N/mm3/2)

K1S2

(N/mm3/2)

K1S3

(N/mm3/2)

1

(C)

2

(C)

3

(C)

Tv

(C)@0,1/s

TK28@Ts=0C

(C)@S1

TK28@Ts=0C

(C)@S2

TK28@Ts=0C

(C)@S3

280 10 818 1174 1520 -

135

-99 -73 41 50 24 5

280 20 1017 1481 1893 -

114

-76 -51 41 34 7 -10

280 30 1166 1705 2157 -

100

-62 -38 41 24 -3 -20

280 60 1490 2179 2698 -75 -37 -16 41 7 -20 5

280 80 1654 2416 2959 -65 -27 -7 41 -1 -28 -10

280 110 1860 2708 3277 -53 -16 3 41 -9 -36 -20

350 10 1023 1467 1900 -

113

-77 -51 37 36 10 -8

350 20 1271 1851 2366 -91 -54 -29 37 21 -6 -24

350 30 1457 2131 2696 -78 -40 -16 37 11 -16 -33

350 60 1862 2724 3372 -53 -15 6 37 -7 -34 -49

350 80 2068 3020 3699 -43 -5 16 37 -14 -41 -56


350 110 2326 3386 4096 -31 7 26 37 -22 -49 -63

410 10 1198 1719 2226 -97 -61 -35 34 27 2 -17

410 20 1489 2168 2772 -75 -38 -13 34 12 -15 -33

410 30 1707 2496 3158 -62 -24 0 34 2 -25 -42

410 60 2181 3191 3950 -37 1 22 34 -16 -43 -58

410 80 2422 3538 4333 -27 11 31 34 -23 -50 -65

410 110 2724 3966 4798 -15 23 42 34 -31 -58 -72

480 10 1402 2012 2606 -81 -45 -19 30 19 -7 -26

480 20 1743 2538 3245 -60 -22 2 30 3 -24 -41

480 30 1999 2922 3697 -46 -8 16 30 -7 -34 -50

480 60 2554 3736 4625 -21 17 38 30 -24 -51 -66

480 80 2836 4142 5073 -11 27 47 30 -32 -59 -73

480 110 3189 4643 5617 1 38 57 30 -40 -67 -80

510 10 1490 2138 2769 -75 -39 -13 29 16 -10 -29

510 20 1852 2697 3448 -54 -16 9 29 0 -27 -44


510 30 2124 3105 3928 -40 -2 22 29 -10 -37 -54

510 60 2713 3970 4914 -15 23 44 29 -27 -54 -70

510 80 3013 4400 5390 -5 33 53 29 -35 -62 -76

510 110 3389 4933 5968 7 44 63 29 -43 -70 -84

Conditions: S1, S2, S3 loading modes at a slow strain rate (0,1/s) and C2 failure consequences.

Ch 2 5 Selection of Steel Quality

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Transcript of Ch 2 5 Selection of Steel Quality