,

Report EUR 13083 EN

Commission of the European Communities

nuclear science and technology Mechanical behaviour

of dissimilar metal welds

C. Escaravage Novatome

10, rue Juliette Recamier F-69001 Lyon

Contract No RAP-090-F

Final report

This work was performed under the Commission of the European Communities for the Working Group 'Codes and standards', Activity Group 3:

'Materials within the fast reactor coordinating committee'

Directorate-General Science, Research and Development

1990 EUR 13083 EN

Published by the COMMISSION OF THE EUROPEAN COMMUNITIES

Directorate-General Telecommunications, Information Industries and Innovation

L-2920 Luxembourg

LEGAL NOTICE Neither the Commission of the European Communities nor any person acting on behalf of the Commission is responsible for the use which might be made of the

following information

Cataloguing data can be found at the end of this publication

Luxembourg: Office for Official Publications of the European Communities, 1990

ISBN 92-826-1895-1 Catalogue number: CD-NA-13083-EN-C

© ECSC-EEC-EAEC, Brussels · Luxembourg, 1990

Printed in Belgium

FOREWORD AND EXECUTIVE SUMMARY

The Commission of the European Communities is assisted in its actions regarding fast breeder reactors by the Fast Reactor Coordinating Committee which has set up the Safety Working Group and the Working Group Codes and Standards (WGCS). The latter's mandate is to harmonize the codes, standards and regulations used in the EC member countries for the design, material selection, construction and inspection of LMFBR components.

The present report is the revised final report of CEC Study Contract N° RA1-0091-F performed under WGCS/Activity Group 3 : Materials. As most AG3 work, this study concerned the evaluation of material data to be used in LMFBR design codes. The main contractor was CEA (France) with UKAEA (UK) and INTERATOM (FRG) as participants.

The report concerns the evaluation of high cycle fatigue properties of three austenitic stainless steels : type AISI 316 (UKAEA tests), type AISI 316L (CEA tests) and type AISI 304 (INTERATOM tests). The data on these steels comprised some 550 data points from 14 casts. This data set covered a wide range of testing parameters : temperature from 20 to 625°C, frequency from 1 to 20,000 Hz, constant amplitude and random fatigue loading, with and without mean stress etc. However, the testing conditions chosen by the three partners differed considerably because they had been fixed independently and not harmonised prior to the tests. This created considerable difficulties when the results were pooled for the evaluations to be performed in the present study.

Experimental procedures and statistical treatments used for the three sub-sets of data are first described and discussed. Results are presented in tables and graphs. Although it is often difficult to single out the influence of each parameter due to the different testing conditions, several interesting conclusions can be drawn :

- The HCF properties of the three steels are consistent with the 0.2 % proof stress, the fatigue limit being larger than, the latter at temperatures above 550°C. The type 304 steel has lower tensile properties than, the two other steels and hence also lower HCF properties.

- Parameters which clearly have a significant effect on HCF behaviour are mean stress or R-ratio (less in the non endurance region than in the endurance region), temperature, cast or product.

- Other parameters have probably a weak or no effect but it is difficult to conclude due to insufficient data : environment, specimen orientation, frequency, specimen geometry.

Recommendations for future work conclude the report. One of these is that partners of collaborative actions should agree on a common test matrix and experimental procedures prior to any testing.

L.H. Larsson

III —

C O N T E N T S

Page

1 - INTRODUCTION

TYPES OF JOINTS INVESTIGATED 2 2.1. Ferritic steel types 2 2.2. Austenitic steel types 3 2.3. Geometry of the welded joints 4 2.4. Weld metals 5 2.5. Welding procedures 6

3 - EXAMINATIONS OF DISSIMILAR METAL WELDS AFTER ELEVATED TEMPERATURE SERVICE 8 3.1. Main observations 8 3.2. Unexpected failures II

4 - LONG DURATION TESTS AND LABORATORY EXAMINATIONS 11 4.1. Laboratory tests for reproducing service failures 11 4.2. Creep tests 12 4.3. Unexpected failure mechanisms proposed from

observed microstructures 13 4.3.1. Austenitic type of welded joints 14 4.3.2. Nickel base type of welded joints 16

5 - MECHANICAL PROPERTIES OF DISSIMILAR METAL WELDED JOINTS WHEN CREEP IS NON SIGNIFICANT 17 5.1. Fatigue behaviour of dissimilar metal welded

joints 17 5.1.1. Fatigue life tests - Initiation locations 17 5.1.2. Crack propagation 18

5.2. Impact toughness 19

6 - LIFE AND RESIDUAL LIFE EVALUATION 19 6.1. Life evaluation : Extrapolation and acceleration

of the failure mechanism 19 6.2. Residual life assessment : Prediction methods 20

V —

7 - CONCLUSIONS 22 7.1. Evaluation and improvement of dissimilar metal

weld lives 22 7.1.1. Low temperatures service 22 7.1.2. Elevated temperature service 23

7.2. Concluding recommendations 24

8 - REFERENCES 27

TABLES AND FIGURES 37

VI

1 - INTRODUCTION

In liquid metal fast breeder reactors, two types of dissimilar metal welds can be found. The first one contains the welds joining the austenitic stainless steel vessels with the roof in carbon or low alloy steels. It is possible to locate these welds in low temperature areas (kept below 200 °C) ; they are not likely to be highly loaded. In this case, defect free welds rather than high strength welds are wished. If necessary, fatigue behaviour of these joints shall be checked. The second type can be found in sodium loops and water or steam pipes connected to the steam generator. During service, they may reach high temperatures in the creep range. Depending on the types of steels selected for the exchange tube and steam generator shell, these dissimilar metal welds are used :

* On large diameter sodium pipes (in this case the dissimilar metal welds can be far from the steam generator if part of the sodium loop or intermediate heat exchanger is made of a ferritic steel).

* On large diameter steam pipes.

* On small tubes of the size of steam generating tubes if the dissimilar metal welds are located on each tube of the bundle before the collecting pipes.

As regards this type of dissimilar metal welds related to the steam generating system and concerned by creep, the experience gained on conventional generating plants may be of interest.

2 - TÏPES OF JOINTS INVESTIGATED

Dissimilar metal welds mentioned in this note are obtained with materials quoted in Tables 1 to 3. They have been submitted to : * post service operations, * laboratory testing programmes.

2.1 - FERRITIC (*) STEEL TYPES Table 1 shows the chemical analysis of the ferritic steels : * Low temperature types as 25 MN 4 or 16 MND 5 for which the weld defects have been studied.

* Creep resisting and oxydation resistant steel types which contain molybdenum and mainly chromium up to 9 %. In the examinations and tests with this second type of ferritic steels, emphasis is on cracking probably due to high temperature (> 500 °C) service and to cycles from service to lower temperature or to room temperature.

Comparison of the service behaviour of the different kinds of ferritic steels is very difficult. The failure mechanism is indeed a long term one. However, low alloy steels with chromium content less than 2 % are no longer used because the effects of oxydation has been recognized to contribute to the failure of some dissimilar welds. Besides, the use of stabilized 2 1/4 Cr 1 Mo Nb to prevent carbon migration does not seem to improve significantly life duration at elevated temperatures (62).

(*) The ferritic steel designation means carbon, low alloy and alloy steels which are neither stainless (Cr < 13 %) nor austenitic whatever the microstructure obtained by heat treatements.

2 -

The cold cracking tendency of low alloy steels 25 MN 4, 20 M 5 M and 16 MND 5 during the welding of dissimilar metals must be taken into account. Butt welding experience in laboratories and workshops ort the other ferritic steels is not sufficient to draw a conclusion. Consequently, comparison of the different types of steels from the point of view of cold cracking is not possible.

Two areas of dissimilar metal welds are concerned by hydrogen assisted cold cracking :

* The martensitic border always present in dissimilar metal welds designed for low temperature service, whatever the ferritic steel and weld metal types (austenitic or nickel base alloy) ; and often existent in dissimilar metal welds of creep resisting steels (par. 2.5).

* The heat affected zone (HAZ) where temperature had risen above AC 1.

The risks of disbonding of the martensitic area and of cracking in the HAZ perpendicular to the weld line are dependent upon the residual stresses, the geometry and hydrogen content of the welds. These points are mainly dependent on the welding procedure.

In addition, Reference 8 connects this risk with the segregations and as a consequence, rather with the fabrication methods than with the chemical composition. So this risk is to be considered mainly for heavy forgings ; more attention should be paid to the cold cracking of the ferritic steels in comparison with the case of homogeneous welding of these steels.

2.2 - AUSTENITIC STEEL TYPES

Their chemical analyses are given in Table 2, but the behaviour of dissimilar welds does not seem to be dependent on the kind of austenitic stainless steel. Alloy 800 is a particular case for its thermal expansion coefficient which approaches that of ferritic steel (see following table..)

The use of a 800 or 800 H alloy transition piece is a means to decrease the calculated stresses resulting from thermal cycling. The weld between the 800 alloy and the ferritic steel brings about as many metallurgical problems as the weld between a ferritic steel and an austenitic one with a nickel base alloy as filler metal.

MEAN THERMAL EXPANSION COEFFICIENT (20 °C - 550 °C)

FERRITIC STEEL 2 1/4 Cr 1 Mo

14 χ 10~6 °C-1

AUSTENITIC STEEL

18.5 x 10"6 °C-1

ALLOY 800

17 χ 10"6 0C _ 1

NICKEL BASE FILLER METALS

14.8 x 10"6 °C_1

2.3 - GEOMETRY OF THE WELDED JOINTS

The joints mentioned in the references, which report examinations after several failures, are butt welded pipes. Two sizes are concerned :

* Steam pipes with thickness greater than 40 mm and diameters ranging from 140 to more than 350 mm.

* Tubes of conventional plants of wall thicknesses between 8 and 20 mm and diameters ranging from 50 to 75 mm.

There are also some collecting pipes, the size of which lies btween the aformentioned ranges. The problems due to weld defects and unexpected service failures can be found to the same extent on both sizes of pipes.

In the case of laboratory tests, butt welded pipes were used for experiments reproducing the type of cracks observed after service failures (References 7 to 9, 20). For more fundamental characterization tests, butt welded plates were used (References 19, 21, 22, 27).

As regards the influence of the welded joint geometry, several authors mentioned that cracks described in Paragraph 3.2, in the HAZ of ferritic steels, are more numerous when the fusion line is perpendicular to the tube axis (Reference 35).

2.4 - WELD METALS

The use of ferritic weld metals (References 3, 37 in Table 3 for instance) as filler metal for dissimilar welds did not last long. Ferritic weld metal with niobum as carbide stabilizing element was also used for buttering layers on the ferritic steel side.

The two main types of weld metals presently used are : * austenitic stainless steels, * nickel base alloy.

The chemical analyses are given in Table 3. As the extensive use of the second type of filler metals began more recently (very often to repair dissimilar welds initially made with an austenitic stainless steel) and because service life of dissimilar metal welds using nickel base alloy weld metals is longer, it appeared only in 1980 that this type of heterogeneous joints was also affected by service failures.

The use of nickel base alloy as filler metal is no longer considered as the final solution for unexpected failures of dissimilar metal welds in cyclically shifted temperatures. However, an increase of service life in comparison with austenitic stainless steel has been observed : relative values of expansion coefficient provide a first explanation of these findings (Cf. Paragraph 2.2). Other effects of a more metallurgical origin are supposed to also contribute to the longer service life of dissimilar metal welds made with nickel base alloy in pure creep conditions without temperature cycles.

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Besides failures due to thermal cyclic conditions analyzed in Paragraph 4.3, few failures were attributed to weld defects, particularly to hot cracking in the weld metal. Solutions to prevent this type of cracks are available in the case of austenitic stainless steels : mainly ferrite content control. A great experience has been accumulated over many years with systematic experiments to clarify the influence of each element of chemical composition. As regards nickel base alloys, the knowledge of the parameters likely to induce hot cracking is quite limited. The influence of some elements (Nb and Si) was pointed out but precise recommendations to prevent hot cracking (as in the case of austenitic stainless steel) are not yet available.

2.5 - WELDING PROCEDURES

Welding procedures play a role in the behaviour of welded joints i.e. they can prevent weld defects. In the case of dissimilar metal welds two kinds of cracking are to be feared : * hot cracking in the weld metal, * cold cracking in the ferrite base metal or in the narrow martensitic border.

For dissimilar metal welds in operation at low temperatures, a good fatigue behaviour is wished ; in this case, defect free welds are obtained with controlled welding procedures. To prevent the second kind of cracking, the fusion line should be smooth without variation in the thickness of the martensitic narrow border and any anomaly in the dilution zone should be avoided. A welding procedure limiting the stress is also beneficial. Welding procedures with two buttering layers made with automatic submerged arc welding with preheating and postheating are chosen to achieve this aim (Reference 84).

For dissimilar welds at elevated temperatures during operation (in heavy thickness), the currently used procedures are the shielded metal manual arc

welding (SMAW) and the inert gaz tungsten arc welding (GTAW). One or several buttering layers have been sometimes used with these procedures. These layers were intended to slow down transfers of chemical elements (mainly of carbon) during welding and afterwards in service. For the failures discussed in Paragraph 4.3, the buttering technique does not provide a full solution. Owing to the lack of programmed comparative tests, it is difficult to quantitatively evaluate the advantage in life duration obtained by buttering.

In case of elevated temperature operating welds with smaller sizes and often made at the plant location, welding procedures are not extensively documented. The examined bibliography only points out the angle between fusion line and tube axis as an acting parameter (Reference 35). In the case of creep resistant ferritic steel welded with a nickel base alloy, welding conditions leading or not to a martensitic narrow border are not known.

As regards the effect of postwelding heat treatment, taking into account analysis of the failure mechanisms (Paragraph 4.3) some authors think that a stress relieving heat treatment has not to be applied or has to be as reduced in temperature and duration as possible for welding without cold cracking (Reference 4). In fact the stress relieving treatment has the same metallurgical effect on the weld microstructure than the elevated temperature hold time during creep in service. But, of course, the thicker the weld, the more difficult it is to justify reduction or suppression of the treatment.

Particular welding procedures have been used for dissimilar metal welds in conventional plants : * In Great Britain, flash butt weld procedures without filler metal : the procedure seems to be no longer used due to the numerous defects of lack of fusion type (Reference 22).

* In the United States, induction pressure welding, a procedure also without filler metal, was also used. References show that this type of weld has a better life performance than dissimilar metal welds with austenitic steel

as filler metal, but that the type of failure described hereinafter (Paragraph 4.3) also occurs. Furthermore, no systematic comparison with nickel alloy dissimilar metal welds has been found.

Recently, heterogeneous welds made by friction have been tested in laboratory (thermal cycles and creep - Reference 81). This procedure also results in creep ruptures in the carbon depleted area although extrapolated 5 times to rupture at 540 °C seem to exceed 10 hours.

3 - EXAMINATIONS OF DISSIMILAR METAL WELDS AFTER ELEVATED TEMPERATURE SERVICE

3.1 - MAIN OBSERVATIONS Number of papers (References 9, 10, 13, 15, 16, 18, 30, 35, 37, 43, 45, 47, 52, 64, 69, 70, 74, 75, 77, 82, 83) report examinations and observations on welds taken from plant after significant service time. Samples were taken from : * failed welds examined, * pieces chosen at random or in the vicinity of failed-welds (for comparative examinations).

In all cases, the failures reported in the references are related to pieces cyclically heated with a maximum temperature higher than 500 °C. In the Reference 83, it is indicated that there is no dissimilar metal weld unexpected failure on plant components working with a maximum temperature of 400 to 475 °C. Thus, the first point to be noticed is that creep is involved in the failure mechanism. Moreover, the intergranular character of cracks and fracture surfaces confirms that creep is present : several authors have reported that typical creep cavities were observed.

The influence of thermal cycling is more difficult to clear out. It can : * initiate cracks by fatigue, * set up the thermal stresses necessary for the creep damage.

In both cases, using a nickel base alloy as filler metal reduces the expansion coefficient gap, decreases the stress level and increases life as it has been statistically noted (Reference 37).

The influence of the oxydation of the ferritic steel is noted in many reports, especially in the area in which the main crack has been initiated. It appears to be blocked by oxyde which can act as a stress concentrator in fatigue. It seems that the oxydation appearance is different in cyclically fatigued zones (oxyde notch) and in less cycled zones (stepped oxydation).

The failure takes place normally in the ferritic steel zone between the fusion line and the non affected zone of the ferritic base metal. The different microstructures observed and their influence will be discussed later. The failure location can be explained : * By a fatigue mechanism with an oxydation effect because reinforced oxydation occurs there.

* By a creep damage mechanism related to the microstructural evolution due to the diffusion transfer of elements between the weld metal and the base metal (Paragraph 4.3).

On several points, many authors are in agreement : * the narrow martensitic border due to dilution near the fusion line on the weld metal side,

* the increase of carbon in the weld metal beyond this narrow border, * the grain growth in part of the HAZ of the ferritic steel.

Other points are still discussed as e.g. the exact origin of phases seen near the fusion line in the ferritic steel and the carbon depletion of the part of the HAZ nearer the fusion line. These points can be dependent on : * the weld metal used (austenitic stainless steel or a nickel base alloy), * the post weld heat treatment.

They will be discussed later (Paragraph 4.3).

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3.2 - UNEXPECTED FAILURES

The words "unexpected failures" are often applied to the failures described above. This can have several different meanings :

* Rupture without plastic deformation detected during operation or visible after failure ; this type of rupture can occur by creep when the creep ductility, at least locally, is low.

* Early failure by comparison with the design made for the elevated temperature service in relation to the creep strength of the less resistant (i.e. the ferritic) steel ; the effects due to the expansion coefficient gap are taken into account only by simplified computation ; the stress concentration due to the creep strain rate gap is not considered ; no attempt to quantitatively check the discrepancy between calculated time to failure and service life to failure is reported in the referenced papers.

The initiation of the cracks can be influenced by other factors which explain the scatter of observed service life :

- fatigue enhanced by oxydation, - initial weld defect in weld metal or on the surface.

Numerous are the expert reports mentioning possible initiation of the main crack leading to failure at weld defects, as :

* lack of fusion, * lack of penetration, * undercut, * shrinkage groove, * cold cracks, * reheat cracks.

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4 - LONG DURATION TESTS AND LABORATORY EXAMINATIONS

4.1- LABORATORY TESTS FOR REPRODUCING SERVICE FAILURES

Number of experimental devices were used for producing, in laboratory environment, cracks similar to those discussed in Paragraph 3.1. Further details are given in the References 7, 8, 15, 17, 31, 32, 46, 62, 69.

In addition to constant load creep tests to be analysed in the following paragraph, other tests were developped i.e. :

* temperature cycling, * temperature and pressure cycling (maximum temperature and maximum pressure in phase),

* temperature and pressure cycling with bending, * temperature and pressure cycling, with, during high temperature hold time, a temperature gradient through the thickness of the tube,

* temperature cycling of a prebent sample, (Reference 47) * various devices of temperature cycling with inhibited expansion or contraction.

These tests are an attempt to reproduce some or all of the load components likely to be present in service and to accelerate material damage.

Some of the tests (References 15,17) have led to failures and cracks different from service ones (References 13-15). These cases are similar to short term creep tests where the cracks observed are not the typical ones found after unexpected service failure. Two reasons can explain this difference :

* The time spell applied to accelerated tests for obtaining quick results, is too short to produce the creep damage mechanisms of Paragraph 4.3.

* Tests failed to reproduce an important load component occurring in the service failure process (probably during initiation).

11 -

The sources of quick temperature cycling with bending favoured the second hypothesis : in addition to the bending stress, there are also stresses due to the expansion coefficient gap and periodically restored by temperature cycling. But the tests are short time ones (1 000 h - 2 000 h at 600 °C).

On the other hand, examination after the tests of Reference 47, indicate that the bending effect is more important than temperature cycling in order to simulate cracks observed after service. With the progress of observation, typical features of unexpected failure cracks have been precised. It is no longer sure that accelerated tests under complex loading reproduce all these features.

4.2 - CREEP TESTS

Creep in an important part of the damage observed in dissimilar metal welds. It was obvious to perform creep tests with cross welds sampling in dissimilar metal weldments. In a first step, experimental conditions chosen to accelerate the creep damage have resulted in ruptures in the ferritic steel far away from the HAZ or in the weld metal, different from the ones observed in service. Creep tests resulting in rupture very near the fusion line (as in the case of service failure) have been obtained under the following conditions :

2 * With large section (20 χ 15 mm ) and accelerating the creep damage by increasing the test temperature above service temperatures (the stress being limited to 100 - 120 MPa) (Reference 19).

* By limiting the stress to 150 - 135 MPa and 115 MPa respectively for the austenitic stainless steel welds and the nickel base alloy ones (References 21-22).

The dissimilar metal cross weld creep tests have shown strain concentration : * in the ferritic base metal whatever the stress, * in a region nearer to the fusion line as the stress is lower.

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This latter point has been obtained by painting a grid on the specimen surface. This grid has also shown that the strain concentration near to the fusion line does not occur from the beginning of creep but is set up step by step during creep hold time.

One can relate the creep elongation to the length of the ferritic steel part of the gauge length and compare the obtained curve to the creep curve of the base metal 2 1/4 Cr 1 Mo. (References 19, 21). It appears that the difference between the two curves occurs only when the dissimilar metal cross weld specimen goes into the tertiary creep. This finding is explained if the strain concentration in the zone near to the fusion line occurs at a late stage of creep time.

Indentation creep tests (Reference 56) have also shown that the zone near to the fusion line is weaker in creep strength. However, the region where the creep strain rate is the largest is located at 1 mm from the fusion line. The region where unexpected failure cracks occur is nearer to the fusion line (less than 150 microns) and corresponds to the area where the creep strain rate gradient is the steepest.

4.3 - UNEXPECTED FAILURE MECHANISMS PROPOSED FROM OBSERVED MICROSTRUCTURES

The prime importance of creep in the unexpected failure mechanisms is now confirmed by the observation of cross weld specimen creep strained to rupture under moderate stresses. Ruptures and cracks very near the fusion line, in the ferritic heat affected zone, have been obtained for dissimilar metal joints welded with both types of filler metals. In the present paragraph, it will be discussed how much observations can explain the creep weakness of the fusion line nearest part of the ferritic steel HAZ.

13

4.3.1 - Austenitic type of welded joints

After welding, from weld metal towards the ferritic steel, one can see :

* The filler metal with dilution free chemical analysis. * The etching resistant zone (light etching zone) very hard and enriched in chromium and nickel by dilution : it is a martensitic zone (lath or acicular martensites) and has often a narrow border shape.

* The interface or the fusion line. * The highly affected part of the ferritic steel HAZ with increased grain size and modified grain boundary orientations : the proportion of boundaries parallel and perpendicular to the fusion line (pavement like structure) is greater than in the random initial orientation.

* The true HAZ with a hardness peak (1 mm from the fusion line). * The non affected ferritic base metal.

Reference 47 proposes a mechanism of solution of the carbides of the grain boundaries near the fusion line because of the carbon enrichment of the weld metal. As carbides disappear, the grain of the austenite can grow before cooling.

The observations of weldments after service or long term tests at elevated temperature show the following points :

* A carbon peak appears in the weld metal. * The tempered martensite is less hard ; precipitation of carbides occurs in the martensite.

* The boundaries of the grown grain area are also with carbides. * In the true HAZ, the hardness peak becomes smooth.

This picture is schematic ; welding conditions and post weld heat treatment for stress relieving can modify the dilution zone microstructure details (residual austenite between the martensite and the interface) and initiate the carbide precipitation.

14

Deeper and deeper progressive localisation of creep damage in the grown grain area at 1 or 2 grain diameters from the fusion line (< 150 micron) is the source of unexpected failures. The first reason for this localisation was thought to be the carbon depleted zone due to the carbon transfer towards the weld metal. Though carbon transfer is thought to be true, the formation of the carbon depleted zone has not yet been confirmed even after more recent investigations (Reference 4). The careful examination of the heat affected zones 2 1/4 Cr 1 Mo steel homogeneously welded or welded to 9 Cr steels have detected a grown grain and, with some experimental difficulties, carbon depleted zone (Reference 55), after post weld heat treatment at 725 °C. The width of this zone is increased above 100 microns by aging. Therefore the carbon depletion hypothesis does not explain such a strong localisation of the damage : the carbon depleted zone would be larger than 150 micron. Another explanation is proposed for a deeper localisation : the early initiation of creep cavities by intergranular carbides in the grown grain zone, carbides playing a role either mechanical or by carbon depletion of the matrix, but very locally. Reference 21 points out fracture surfaces even nearer to the interface (10 micron - 20 micron) which follow the shape of weld seams. Interface (*) cracks participate in this rupture (type I). It occurs at slightly shorter rupture times than the fracture in the grown grain area described above (type II) : stress for type I failure being 100 to 135 MPa. But this remark is not in agreement with the observation that the damaged zone is narrower as the stress is lower (Reference 22). More likely, the two mechanisms are competitive and slight changes in the welded joint microstructure produce one or the other type of cracks. Figure 1 shows the ranges of temperatures and of times to rupture where type I and type II creep ruptures are observed.

(*) Interface cracks are attributed to multiaxial stresses which come from creep strain compatibility in weld metal (near zero creep stain rate) and in the ferritic steel.

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4.3.2 - Nickel base type of welded joints

Many aspects are similar with the case discussed above. However, some differences must be pointed out : the reported observations are less coherent and some detailed aspects are more dependent on the local welding conditions. In Reference 21, two types of fusion line zone on the ferritic steel side are observed with the following features : * Type I : Narrow fusion line without clear indications of martensite,

with, on the weld metal side, high Ni and Cr contents and carbides formed during welding giving a peak of carbon at 20 micron or less from the interface.

* Type II : A thicker fusion line with a hardness peak, near the interface in the weld metal ; this zone is the martensitic zone pointed out by many metallographic examinations.

Transverse joint creep tests (References 19-21) give type II creep rupture : the initial crack appears close to the interface and seems to go away in the HAZ and to branch. Cavitation indications are rare.

The results of the examinations of nickel base welded joints, after service cracking, are relatively different (Reference 18) ; M-^Cg carbides lie nearly continuously on the interface ; creep cavities are associated with these carbides and promote the rupture by growth and coalescence. In Reference 43, for a case of in service rupture, several indications show that the propagation of cracks through the carbides and along grain boundaries, one grain from the interface, are two competitive mechanisms for cracking. It seems that the transverse creep tests have yet been of too short duration to observe rupture from interface carbides ; but different features of the fusion line can also result in different types of rupture. The use of such results for extrapolation and dissimilar metal creep life prediction requires some cautions in the case of two mechanisms in competition. Figure 2 shows the ranges of temperatures and times to rupture where type II ruptures are observed.

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5 - MECHANICAL PROPERTIES OF DISSIMILAR METAL WELDED JOINTS WHEN CREEP IS NON SIGNIFICANT

5.1 - FATIGUE BEHAVIOUR OF DISSIMILAR METAL WELDED JOINTS

5.1.1 - Fatigue life tests - Initiation locations Fatigue tests on transverse specimens from dissimilar metal welded joints (References 71, 72, 73) show fatigue lives similar to ferritic base metal lives when Δε < 2 %. For very high strain ranges Δε > 2 %, a slightly smaller fatigue life as compared with ferritic base metal is observed : this reduction is thought to be due to heterogeneous strain arising in the fatigue specimen near the fusion line.

In Reference 42, a fatigue specimen with diametral extensometry, is used with maximum strain in the heat affected zone ; such a specimen does not give a number of cycles to rupture different from that of the ferritic base metal.

In Reference 72, fatigue tests were made on specimens taken from K-shaped and cruciform welded joints, with as welded or flush ground surface of the weld seam. A first series of fatigue tests were stress controlled ; several

5 points are to be noted for high number of cycles (N > 10 ) :

* Flush grinding improves the life over 2 χ 10 cycles by a factor which exceeds 2.

* A first improvement (60 %) is obtained by grinding the welded seam on the ferritic steel side, but grinding of the welded seam on the austenitic steel side results in a second improvement of fatigue life (20 % ) .

* In the as-welded state, fatigue initiation takes place at the welded seam edge, on the ferritic steel side ; it is shifted on the edge of the welded seam on the austenitic side when the other weld seam is ground and takes place in the austenitic base metal when all the welded seams are ground.

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* When a buttered layer is used, the initiation takes place in various areas in the as-welded condition (welded seam edge on both sides) and in the ground condition (austenitic base metal or weld metal). The latter type of initiation is due to inclusions in the weld metal not detected by non destructive examinations.

* Using a buttering improves the behaviour in the as-welded condition ; in the completely ground condition, the presence of inclusions is the dominating factor for the fatigue life.

* Grinding suppresses the stress concentration coming from the geometry of the welded seam edge surface, but also modifies the residual stress field near the surface.

When fatigue tests are strain controlled as it is the case in the second type of tests reported in Reference 72 most of the initiations and of the ruptures take place in the weld metal, the initiations being located at slag inclusions.

These results are interesting about the effects of the surface condition and of slag inclusions in weld metal ; but the effects are not specific to a dissimilar welded joint between a ferritic steel and an austenitic stainless steel.

5.1.2 - Crack propagation When the initiation occurs in the decarburized HAZ (Reference 71), the propagation can be observed in this area but :

* The propagation rate is not greater than in the ferritic base metal. * A deviation of the crack in the austenitic weld metal is sometimes observed.

* The martensitic border does not influence the crack path.

Reference 7 shows that crack growth rates in the weld metal of dissimilar joints can be : * reduced by compressive residual stresses, * increases, at least macroscopically, by the presence of inclusions.

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5.2 - IMPACT TOUGHNESS

The Charpy tests of Reference 68 on specimens taken from dissimilar metal welds, give greater impact values for weld metal containing more nickel (40 Ni 16 Cr weld metal can be compared with 24 Cr 20 Ni weld metal). This effect is thought to be due to the thickness of the martensitic border which decreases the impact level when exceeding 7.5 micron.

The Charpy tests (Reference 60) also show the effect of heat treatment of various duration (10,30 and 60 h at 550 °C, 600 °C and 650 °C). The impact toughness decreases much less in the weld metal with 70 % of nickel than in the weld metal with 23 % of chromium and 12 % of nickel. In the latter case, the impact toughness drop increases with ferrite content.

6 - LIFE AND RESIDUAL LIFE EVALUATION

6.1 - LIFE EVALUATION : EXTRAPOLATION AND ACCELERATION OF THE FAILURE MECHANISM

To get creep results usable for the prediction of life time of dissimilar metal welds, it is necessary to accelerate tests on cross weld specimens in comparison with service conditions.

In this connection, two important points appear from the reported works :

1/ It is better to use an increased temperature because an increase of the stress results in other failure mechanisms : a curvature is observed in the isothermal logarithmic stress versus time plot ; as a consequence, extrapolations on such a plot are doubtful. The temperature must yet be moderately above the service temperature in order to as little as possible modify the metallurgical evolution. Tests are made at a given stress at the highest temperature to obtain short results, then going down step by step near the service temperature with the longer tests. In this case, linear extrapolations to the service temperature seem to be possible.

19

2/ Uncertainty bound to the necessary extrapolation in life assessment, is greater in the case of nickel weld metal than in the case of austenitic weld metal. This is the counterpart of longer creep life but also the observation of two different competitive failure locations does increase the doubt about the validity of the extrapolations.

Figures 3 and 4 show the same time to rupture, temperature logarithmic diagram, the results of creep tests and the life durations observed after service or after more sophisticated laboratory tests. In the case of austenitic weld metal (Figure 3), the domain of test results joins the domain of the observed service life. In the case of nickel base weld metal, the available tests results are yet : * too short with a factor of about four in comparison with observed lifes,

* at a temperature increased by 25 or 30 °C above the service temperatures.

6.2 - RESIDUAL LIFE ASSESSMENT : PREDICTION METHODS

They are under development. Indeed it is difficult to evaluate lives and residual lives, taking into account the different creep laws of the base metal, the ferritic heat affected zone and the weld metal and the calculation of the related stresses by methods other than the simplified ones used for design. The stress analysis has been improved in the case of a weld metal and a base metal of different creep behaviours. But the problem is more complicated when the heat affected zone is expected to have also a particular behaviour. The development of experimental or quasi experimental methods for total or residual life assessment has been undertaken. The first efforts have studied (Reference 30) the carbon evolution : * width of the carburized zone, * size of Mo-C carbides. The latter measurement is relatively well correlated with the temperature and time conditions seen by the material, but to our knowledge, the relation of this measurement with the creep residual life was not tested.

-20

Then the observations related more directly to the creep damage were tentatively used : * maximum length of microcracks, * proportion of cracked grain boundaries, * file of typical micrographies corresponding to known steps of the damage

(medium time or end of secondary creep, medium time of tertiary creep).

The development of these methods will be long because it needs numerous quantitative examinations of out of service samples and specimens of isothermal or thermally cycled creep tests both interrupted and conducted to rupture. These examinations are quite necessary to quantify the damage accumulation which is presently known only schematically. As a counterpart, it can be noted that these methods can become non destructive with the parallel development of replica techniques.

The PODIS method, under development in the United States, is a mixed method of calculation and of observation of damaged samples. Its principle is to separate the damage in three parts (instead of two in current design evaluation) : * Primary damage proportional to the time at temperature : it is a creep damage in t/t_ (t_ being the rupture time under primary stresses).

* Secondary damage proportional to the number of cycles n/N : it is a fatigue damage.

* "Intrinsic" damage proportional to the number of cycles and to the temperature variation during the cycle : η χ ΔΤ.

The observations and tests on samples, after damage accumulation, are used in relation with pure creep tests, to determine the five coefficients needed to calculate the total damage.

21

7 - CONCLUSIONS

7.1 - EVALUATION AND IMPROVEMENT OF DISSIMILAR METAL WELD LEVES

7.1.1 - Low températures service

In this case, the improvement of dissimilar metal welds will result from the development, qualification and control of the welding procedure to obtain safe welded joints :

* Without defects i.e. : - free of hydrogen assisted cold cracks in the HAZ and in the martensitic border zone,

- free of hot cracks in the weld metal.

* With a martensitic border zone of regular and as small as possible thickness obtained by controlling the dilution of the weld metal and the base metal.

In these conditions, the fatigue life of dissimilar metal welds does not seem to be shorter than the life of the base metals to be joined. In the case of dissimilar metal welds as well as in the case of homogeneous welds, fatigue life depends on surface finish, in particular at the edges of the welded seams.

The thickness of the martensitic border zone has an influence on the impact toughness of the weld and, probably on its toughness. It is not possible to quantify toughness and to compare with the toughness of the base metal because the analysis of fracture mechanics tests in heterogeneous welds is difficult. Nickel base weld metal is to be preferred for loaded joints from the impact toughness point of view. But with nickel base weld metal, the welding procedure is more difficult due to the hot cracking risk. Furthermore, the hot cracking risk depends on the non ferritic base metal by the effect of the dilution on the non ferritic side of the weld.

22 -

7.1.2 - Elevated temperature service

To obtain for the dissimilar metal welds a life as long as that of the weaker of the base metals, it is not enough to produce defect free welded joints when the creep damage mechanism is likely to operate. Creep reduction factors seem to be applied to dissimilar metal welds.

Three effects can lower the creep strength : * The additional strains and stresses due to the expansion coefficient gap between the weld metal and the ferritic base metal.

* The difference in creep strain rates between the weld metal and the ferritic base metal.

* The microstructure of the zone where the creep damage is localised i.e. either the interface between the ferritic base metal and the weld metal, or the heat affected zone near the fusion line at a distance of one or two grains.

The joints with a nickel base alloy as weld metal, are in a better situation, at least for the first effect. Cross weld creep tests (Paragraph 4.2), cyclic tests (Paragraph 4.1) and service experience (Paragraph 3.1) confirm this point.

For the second effect, welded joints of the more creep resistant ferritic steels should have a better behaviour. But we have not yet found any systematic study of dissimilar metal welds of high creep resistant steel (like 9-12 % chromium steel) and the comparison with 2 1/4 Cr Mo metal welds is not possible.

The localisation of the creep damage has been observed but not completely explained by a metallurgical mechanism. But both localised damage zones have been produced in the low stress creep tests and the corresponding mechanisms seem to be competitive with nickel base as well as with austenitic steel weld metal.

- 23

Considering the results available, for the evaluation of the reduction factors to be applied to dissimilar metal welds in time and temperature service conditions, the necessary extrapolations are more difficult in case of nickel base welds than in the case of austenitic ones.

High stress tests are not convenient for extrapolation ; mechanisms must be accelerated by higher temperature. There are no clear reasons to consider separately the two types of fracture.

7.2 - CONCLUDING RECOMMENDATIONS

1 For all service conditions considerable thought should be given to selection and qualification of welding procedures which is not as easy as in the homogeneous welded joint case for the following reasons :

* On the ferritic steel side - cold cracking prevention, - control of dilution for the band or martensite.

* On the austenitic steel side - hot cracking prevention specially when the weld metal is a nickel base alloy,

* In the weld metal - no inclusion nor porosity (at least no large inclusion) in the weld deposit.

2 For high or frequent cyclic loading care should be taken with the surface finishing of the weld seam. Flush grinding or machining result in fatigue strength improvement and failure locations shift from the weld reinforcement to the base metals. The buttering technique improves the fatigue strength only if the buttered layers are as inclusion free as the filling of the weld.

24

3 As good fatigue strength as in the base metals can be expected of dissimilar metal welds if they are free from defects mentioned in 1 and 2 above.

4 When brittle fracture must be considered, nickel base alloy is to be preferred as weld metal, for the higher impact strength and toughness of the welded joint.

5 When creep strength is important, the use of nickel base alloy should be preferred to austenitic steel as weld metal.

6 Even in the case of nickel base weld metal, reduction factors should be established to take into account the reduced creep strength of the dissimilar welded joint as compared to the ferritic steel (the weaker of the two base metals), a weak creep strength zone cannot be avoided by current welding practices.

7 The post weld heat treatment must be limited to the minimum level required for welding procedure qualification (to prevent cold cracking) ; it can initiate early metallurgical degradation leading to weakened zones (fusion line and/or nearest HAZ) where creep damage will localise.

8 After thermal cycling, the state of stress is complex due to the expansion coefficient gap but also to the accommodation of the different strain rates ; in life prediction it can be useful to add an intrinsic cyclic damage variying with η χ ΔΤ to the current creep and fatigue damages due to the calculated stress and strain range.

25

9 Creep tests vinder moderate stress are needed, more specially in the nickel base weld metal case, in order to :

* improve extrapolation methods,

* experimentally define parameters used in life prediction,

* explain more completely the mechanism responsible for creep weak zones by micrographie examination and analysis.

First tests must be performed at temperatures higher than expected service temperatures, going down afterwards to temperatures closer to the latter. Stresses are to be limited to the 0.2 % yield stress of the austenitic base metal (120-150 MPa). Tests to failure and tests stopped after noticeable damage accumulation are expected to provide complementary information.

10 Thermal cyclic creep fatigue tests are needed if intrinsic damage is used in life prediction to adjust the parameters involved.

11 Ferritic steels, more creep resistant than the 2 1/4 Cr 1 Mo one, are expected to give improved dissimilar welded joints for creep due to the reduced gap in creep strain rates with the austenitic stainless steels. But this experimental evidence is not yet available.

12 Localised oxydation has not been quantitatively checked but can act in the failure mechanism.

26

8. REFERENCES

1 JF. KING Behaviour and properties of welded transition joints between austenitic stainless steels and ferritic steels - A litterature review. Oak Ridge National Laboratory Report. ORNL-TM - 5163, nov. 1975.

2 CD. LUNDIN Dissimilar metal welds - Transition Joints Litterature Review. Weld. Jour. Res. Sup p. 58S, feb. 1982.

3 G. FABER, T. GOOCH Assemblages soudés entre aciers inoxydables et aciers faiblement alliés - Etat de la question. Doc. IIS/IIW 703-82, le soudage dans le monde, Volume N° 5/6, 1982.

4 R. VLSWANATHAN, RI. JAFFEE, J. DIMMER Dissimilar metal welds in power plants. Conf. Joining Dissimilar Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

5 PL. KLUEH Dissimilar metal weld failures in boiler tubing. Power Engineering, feb. 1984.

6 RH. RYDER, C.C. LI., R. VISWANATHAN, J. DIMMER Dissimilar Metal Weld Failures in Power Plants - Causes and Remedies Trends in Electric Utility Research - Conference on Recent Developments in Electric Utility Research. Midwest University - Chicago april 1984.

7 M.DC. MOLES, MJ. TINKLER, HJ. WESTWOOD Testing Austenitic - Ferritic Superheater Tube - Transition Welds Under Combined "Two Shifting" and Bending Stresses. J. Materials for Energy Systems V.3, p. 19, june 1981.

8 M.DC. MOLES, MJ. TINKLER, HJ. WESTWOOD Accelerated testing of three austenitic - ferritic superheater tube transition welds under simulated "two shift" conditions. Metals Technology - p. 121, april 1981.

9 RB. DOOLEY, GG. STEPHENSON, MJ. TINKLER, M.DC. MOLES, HJ. WESTWOOD Ontario Hydro Experience With Dissimilar Metal Welds in Boiler Tubing Welding Journal, Weld. Res. Sup. p. 45S, feb. 1982.

10 V. WRIGHT Service experience with austenitic ferritic superheater transition welds within CEGB Midlands. Conference on Welding Dissimilar Metals. Droitwich, may 1969.

11 LM. WYATT Dissimilar metals joints used in the steam circuit of electrical generating plant Conference on Welding Dissimilar Metals. Droitwich, may 1969.

- 27

12 T. ROWLEY, T. ROWBERRY, CA. ALDRIDGE Problems associated with the Design Inspection and use of large diameter ferritic/ austenitic joints in power plants. Conference on Welding Dissimilar Metals. Droitwich, may 1969.

13 J. BARFORD, KS. PROBERT Interfacial effects in dissimilar steel joints. Int. Conf. on Welding Research Related to Power Plant. Southampton. Marchwood 17-21, sept. 1972.

14 AK. HARDY, T. ROWLEY, JA. WILLIAMS Austenitic ferritic welded transition joints for high temperature applications. Conference on ferritic steels for Fast Reactor Steam Generators. BNES - London, may 1977.

15 TR. ROWBERRY, BI. BAGNALL, JA. WILLIAMS Analysis of service experience with large austenitic/ferritic steampipe joints in CEGB Midland Region plant. Conference Welding and Fabrication in the Nuclear Industry. BNES - London 1979.

16 AT. PRICE CEGB experience with small diameter dissimilar metal welds in coal fired boilers. Conf. Joining Dissimilar Metals. AWS EPRI Pittsburg, 24 et 25 august 1982.

17 JD. WILLIAMS, RD. NICHOLSON Failure modes in heavy section dissimilar joints subjected to accelerated cycle creep loading conditions. Joint Engineering Material Research Seminar. University College of Swansea, april 1982.

18 RD. NICHOLSON Interfacial structures in nickel based transition joint after long term service. CEGB Report N° RD/M/N 1131 - CEGB Marchwood Eng. Lab. ; Marchwood 1980.

19 RD. NICHOLSON The creep rupture properties of austenitic and nickel-based transition joints. CEGB R e p o r t N° RD/M/N 1176 N. 81 CEGB Marchwood Eng. Lab . ; Marchwood 1981. Metals Technology 9 p . 305 - 311, 1982.

28

20 TR. ROWBERRY, PB. BATES Service simulation tests on austenitic/ferritic transition joints for steam pipe applications. Par. II : Metallurgical examination of four Joints. CEGB Report N° SSD MIS R231. 70 Midlands Region, dec. 1970.

21 Β. ΝΑΤΉ Creep rupture and creep crack growth behaviour of transition joints. Int. Conf. on Welding Technology for Energy Application ; Gatlinburg, may 1982.

22 IJ. CHILTON, AT. PRICE, B. WILSHIRE Creep deformation and local strain distribution in dissimilar metal weld between AISI type 316 and 2.25 Cr-1 Mo steels made with 17 Cr 8 Ni 2 Mo weld metal. Metals Technology V. 11 p. 383, sept. 1984.

23 PJ. ALBERRY, B. CHEW, WKC. JONES Prior austenitic grain growth in heat effected zone of a 0.5 Cr MoV steel. Metals Technology V. 4 p. 317-315, june 1977.

24 PJ. ALBERRY, WKC. JONES Structure and hardness of 0.5 Cr MoV and 2 Cr Mo simulated heat affected zone. Metals Technology V. 4 p. 557-566, dec. 1977.

25 PJ. ALBERRY, J. MYERS, B. CHEW An improved welding technique for heat affected zone refinement. Welding and Metal Fabrication V. 45 N° 9 p. 549-553, nov. 1977.

26 PJ. ALBERRY Simple test reveals level of two layer refinement. Welding and Metal Fabrication p. 543-547, nov. 1981.

27 Β. ΝΑΤΉ Strengthening of dissimilar metal welds by post weld heat treatment. 2d Internat. Conf. Creep and Fracture of Engineering Materials and Structures. Swansea 1984.

28 WKC. JONES Heat treatment effect on 2 Cr Mo joints welded with a nickel base Electrode. Weld. Jour. Welding Res. Sup. p. 225S, may 1974.

29 DW. JAMES, M. NEUMANN, J. SOO The integrity of 9 Cr 1 Mo to stainless steel transition joints in AGR steam generator. Joint Engineering Material Research Seminar, Swansea, april 1982.

29

30 JD. PARKER Failure examination of a thick section AISI 347/2 1/4 Cr 1 Mo transition joints. CEGB RD/M/1169 R 81, July 1981.

31 JT. TUCKER, F. EBERLE Development of a ferritic-austenitic weld joint for steam plant application. Welding Journal, Welding Research Sup. p. 529S, nov. 1956.

32 RW. EMERSON, RW. JACKSON, CA. DAUBER Transition joints between austenitic and ferritic. Steel piping for high temperature steam service. Welding Journal, Welding Research Sup. p. 385S, sept. 1962.

33 PE. HASS Results of industry wide survey dissimilar metal weld perfomance. Conf. Joining Dissimilar Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

34 S. IBARRA Experiences with dissimilar welds in the petroleum industry. Conf. Joining Dissimilar Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

35 S.O. HILTON, S.N. CATO, F.V. ELLIS Metallurgical characteristics of austenitic weld metal transition joint failures. Conf. Joining Dissimilar Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

36 RF. GURNEA Radiographic technique for detecting distress in dissimilar weld joints. Conf. Joining Dissimilar Weld Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

37 GL. MESSER TVA's experience and repair procedures of dissimilar welds. Conf. Joining Dissimilar Weld Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

38 G.H. HARTH Alternatives for new construction. Conf. Joining Dissimilar Weld Metals. AWS EPRI Pittsburgh, 24-25 august 1982.

-30

39 CR. BRINKMAN, JF. KING, JP. STRIZAK, RL. KLUEH, MK. BOOKER Mechanical properties of transition joint materials in support of LMFBR steam generator design. Int. Conf. on Ferritic Steel for Fast Reactor. Steam Generators BNES London, may 1977.

40 JF. KING, GM. SLAUGHTER Transition joint welding development for LMFBR steam generators. Int. Conf. on Ferritic Steel for Fast Reactor. Steam Generators BNES London, may 1977.

41 JF. KING, MD. SULLIVAN, GM. SLAUGHTER Development of an improved stainless to ferritic steel transition joint. Welding Journal, Welding Research Sup. p. 354S, nov 1977.

42 CR. BRINKMAN, JP. STRIZAK, JF. KING Elevated temperature fatigue characterisation of transition joint weld metal and heat affected zone in support of breeder steam generator development. ASTM STP 648, Fatigue Testing of Weldments p. 218, 1978.

43 RJ. GRAY, JF. KING, JM. LEITNAKER, GM. SLAUGHTER Examination of a failed transition weld joint and the associated base metals. Oak Ridge National Laboratory Report. ORNL 5223, January 1977.

44 GM. GOODWIN, JF. KING Bridging the gap between dissimilar materials. Welding Design and Fabrication p. 90-91, may 1979.

45 RL. KLUEH, JF. KING Austenitic stainless steel - ferritic steel weld joint failures. Welding Journal, Welding Research Sup. p. 302S, sept. 1982.

46 RL. KLUEH, JF. KING, JL. GRIFFITH A simple test for dissimilar - metal welds. Welding Journal, Welding Research Sup. p. 154S, January 1983.

47 KH. HOLKO, CC. LI Dissimilar weld failure analysis and development program. Int. Conf. on Welding Technologuy for Energy Application. Gatlinburg, may 1982.

48 CC. LI, R. VISWANATHAN, RH. RYDER The microstructure and remaining l i f e of d iss imi lar metal weldments a f te r service in foss i l f i red b o i l e r s . ASME I n t . Conferences on Advances in L i f e . Pred ic t ion Methods Albany, 18-20 ap r i l 1983.

31

49 AW. DALCHER, TM. YANG, CL. CHU High temperature thermal elastic analysis of dissimilar metal transition joints. Jour. Eng. Materials and Technology p. 65, January 1977.

50 TM. YANG, AW. DALCHER Inelastic behaviour of a dissimilar metal welded pipe transition joint-Comparison of experimental measurements and analytical prediction. Inst. Mech. Eng. 1980.

51 CC. YANG, AW. DALCHER An inelastic analysis of dissimilar metallic pipe joints for CRBRP applications 1982 Pressure Vessel piping conference. Orlando 27 june - 2 July 1982.

52 DI. ROBERTS, RH. RYDER, R. VISWANATHAN Performance of dissimilar welds in service. Journ. Pressure Vessel Technology V. 107, p. 247, august 1985.

53 R. VISWANATHAN Dissimilar metal weld and boiler creep damage evaluation for plant life extension. Jour. Pressure Vessel Technology V. 107, p. 219, august 1985.

54 FRENCH. N. DAVID High nickel joints unite dissimilar steels. Welding Design and Fabrication, V. 54, p. 92-93, may 1981.

55 AA. TAVASSOLI, A. LECLOU Migration du carbone et évolution structurale de liaisons hétérogènes soudées entre aciers "ferritiques" 2,25 Cr Mo et 9,5 Cr Mo Nb V. Jour. Nucl. Mat. 96 p. 329, 1981.

56 WS. GIBBS, SH. WANG, DK. MATLOCK, DL. OLSON High temperature impression creep testing of weldments.J Welding Journal, Welding Research Sup. p. 153S, june 1985.

57 W. JAHN Schweissverbidungen zwischen ferritischen und austenitischen Stählen in Versuch and Praxis. Mitteilungen der VGB N° 7 p. 3, 10 april 1962.

58 H. WIRTZ Bedentung and Anwendung von Pufferchichten beim Schweissen und Metallspritzen. Schweissen und Schneiden V. 22 N° 10, p. 417-421, 1970.

32

59 R. LISON Zur Problematik der Schweissenverbindungen zwischen underschiedliche Wertstoffen unter besonderer Berücksichtigung des Schmelzschweissens Schweissen und Schneiden V. 28 N° 3, p. 89-92, 1978.

60 H. WIDEOWITHZ, E. FOLKHARD Schweisszusatzwerkstoffe fur Austenit. Ferrit - Verbindungen. Schweisstechnik V. 32 N° 7, p. 101-107, 1977.

61 K. HENNEMANN Schmelzchweissverbindungen zwischen verschiedenartigen Stahlen. V.D.I.Z.V. 119, N° 12 p. 615-620, 1977.

62 E. JAHN Temperaturwechselversuche an Schweissverbindungen zwischen austenitisehen und ferritischen Stählen. VGB Kraftwerkstechnick V. 61, p. 509, 1981.

63 A. OHTOMO, F. OKADA, Y. SAIGA A failure analysis of butt weld joints between 1 1/4 Cr 1/2 Mo steel and HK 40. Journal of the Society of Materials Science Japan V. 24 N° 258, p. 204-209, march 1975.

64 A. OHTOMO, S. OHISHIBASHI, Y. SAIGA Analysis on the failure of butt weld joint between 2 1/4 Cr 1 Mo steel tube and stainless steel tube. IHI Engineering review V. 10 N° 2, p. 17, april 1977.

65 O. AKIBO, A. YAMAMURA Dissimilar welds in thermal power plants. ASTM-ASME-MPC Joint Comittee Dissimilar Weld Task Force Meeting. La Jolla, California, 8 October 1981.

66 H. HANEDA, M. ARAOKA, K. SETOGUCHI, JD. FOX, WF. SIDDAL Prediction of creep fatigue damage in major steam generator components. 38 th American Power Conference 1976.

67 K. SETOGUCHI, T. DAIKIKU, F. MASUYAMA, H. HANEDA, K. MURAISHI Life prediction of dissimilar welds joints for boiler tubing . ASME Int. Conf. on Advances in Life Predictions Methods. Albany, p. 179, 18-20 april 1983.

68 M. WATANABE, I. WATANABE Brittleness at bonded part of deposit metal by austenitic stainless steel electrode to the ferritic steel mother metal. Doc. IIW No X 511-69, July 1969.

33

69 F. MASUYAMA, K. SETOGUCHI, H. HANEDA, F. NANJO Findings on creep - fatigue damage in pressure parts of long-term service - exposed thermal power plant. Jour, of Pressure Vessel Technology V. 107, p. 260, august 1985.

70 JTW. PASTOORS, J. VAN LIEVE KEMA Research on Dissimilar Joints. Seminar on Dissimilar Welds in Fossil Fired Boilers. New Orleans, 23-24 feb. 1984.

71 C. AMZALLAG, JL. BERNARD, E. BOLLINGER Amorçage et propagation des fissures de fatigue dans des joints d'assemblages hétérogènes en aciers inoxydables et faiblement allié. Congrès IIS Dublin, juillet 1978.

72 C. MAILLARD-SALIN, JP. GEREY Comportement en fatigue de soudures hétérogènes. Recherche CECA N° 7210. KD312 et 313 F47-80, octobre 1984.

73 C. AMZALLAG, JL. BERNARD, P. RABBE, G. SLAMA Fatigue behaviour of welded joints between dissimilar metals (austenitic stainless steel and low alloy steels). 5 th Inter. Conf. on fracture, Cannes, 29 march - 30 april 1981.

74 ARRIZZOLI, A. GAUCHET EDF-GDL Rapp d'expertise D5002/AZI/R83-285, October 1983.

75 A. GAUCHET, PERRET, CHATELAIN, PITARD EDF-GDL, Rapport d'expertise D5002/GHT/R82-808, février 1983.

76 BJL. DARLASTON, DS. WALTERS Some considerations of the application of cyclic data to the design of welded structure. ASTM-STP 520 - Fatigue at Elevated Temperatures p. 688-698, 1973.

77 R. ROWBERRY, BI BAGNALL Austenitic ferritic - steampipe transition joints in Midlands Region plants. Service experience to december 1977. CEGB Report N° SSD/MID/R66/78.

78 AK. HARDY Elastic stress distributions in welded transition joints. CEGB Report N° B/M/M.657 - CEGB Marchwood Eng. Lab. Marchwood 1972.

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79 A.A. TAVASSOLI, H. TOURON, M. WEISZ Effect of decarburization on structural and mechanical properties of ferritic steels. Nuclear Technology V. 55 p. 302, nov. 1981.

80 A.A. TAVASSOLI, A. LECLOU, P. PETREQUIN Conséquences de migration du carbone dans les liaisons hétérogènes soudées. Colloque de Métallurgie INSTIN, juin 1986.

81 K.G.K. MURTY, S. SUNDERESAN Thermal behaviour of austenitic-ferritic transition joints made by friction welding. Welding Research Sup. p. 327S-334S, december 1986.

82 D.M. HADDRILL, J.D. RUSSEL Cracking associated with dissimilar metal welds at the top ends of reformer tube. Conference on welding of dissimilar metals Droitwich may 1969

83 F.P. FROST Service experience with some dissimilar metal welds in the process industry Conference on welding of dissimilar metals Droitwich may 1969

84 M. FONDEVIOLE, A. VIGNES Technologie des liaisons bimétalliques acier inoxydable austénitique ou alliage NiCrFe sur acier 20 M5 M ou 16 MND5 DVS 75 pl82 1975

35-

T A B L E S A N D F I G U R E S

37

TABLE 1

FERRITIC STEELS INVOLVED IN DISSIMILAR METAL WELDED JOINTS

TYPICAL ANALYSIS ON PRODUCTS

GRADE

00 CT)

SA 516 25 MN 4 M 20 M 5 M 16 MND5 (13 MND5) 0 ,5 % Mo

1 Mo 0 ,5 Cr MoV 1 Cr 1/2 Mo SA 213 T i l 1 1/4 Cr 1 Mo 2 Cr Mo SA 213 T22

2 1/4 Cr 1 Mo (10 Cr Mo 9-10)

10 CD 9-10

10 Cr Mo Nb 9-10 SA 213 T21 3 Cr 1 1/2 Mo 3/4 V 3 Cr 1/2 Mo 3/4 V 1/2 SA 213 T5 25 NCD 7-5

< 0 ,21 « 0,25 < 0,25

« 0 ,22

(0 ,12) (0 ,12) Í 0,15

(0,13) 4 0,15 (C,13) (0 ,14) (0,12) (0 ,10) (0,066) (0,081)

( 0 , 0 8 )

< 0,15

« 0,15

Mn

0 ,60-0 ,90 « 1,20 < 1,20

1,15-1,60

(0,62) (0 ,6) 0 , 3 -0 ,6

(0 ,56) 0 ,3 -0 ,6 (0 ,47) ( 0 , 4 0 ) ( 0 , 8 7 ) ( 0 , 5 8 )

( 0 , 4 3 ) ( 0 , 4 7 )

(0 ,61)

0 , 3 -0 ,6

0 ,3 -0 ,6

Si

0 ,15-0 ,40 « 0,55 Í 0,50

0 ,10-0 ,30

(0,28) (0,25) 0 ,5 -1

( 0 , 2 2 )

< 0 ,5

(29)

( 0 , 2 2 ) ( 0 , 3 0 ) ( 0 , 3 5 )

( 0 , 3 3 ) ( 0 , 2 8 )

( 0 , 4 0 )

< 0 ,5

« 0 ,5

Cr

Í 0 ,25

« 0 ,25

(0 ,45)

(1) 1 ,0-1 ,5

(1 ,89) 1 ,9-2,6 (2 ,27) (2 ,21) (2 ,07) (2 ,39) (2 ,12) (2 ,34)

(2 ,26)

2 ,65-3 ,35

4 , 0 - 6 , 0

« 0 ,25

0 , 4 3 - 0 , 5 7

(0 ,54) (0 ,5) 0 ,44-0 ,65

13

( 0 , 8 8 )

0 ,87 -1 , ( 0 , 9 8 )

(0 ,90) (0 ,78) (1 ,04) (1 ,08) (1 ,08)

(1 ,12)

0 ,8 -1 ,06

0 ,45-0 ,65

« 0,02

Í 0,01

( 0 ,25 )

( 0 ,05 )

( 0 ,02 ) ( 0 ,04 )

N i

« 0,50

0 ,50-0 ,80

( 0 , 2 0 )

(0 ,19)

(0 ,17) (0 ,23) (0 ,71)

Cu Nb

« 0 ,25

« 0 ,20

(0 ,12)

(0 ,11)

(0 ,09) (0 ,27) (0 ,15)

(1 ,05)

REFERENCES

34 71-83 83

72-83 3 11 3-24 10-11-4-47 34-35

16

4 à 9, 36, 37, 43, 44, 47, 59, 53, 54

13 , 19, 20, 22, 30, 32, 35 39, 41 , 42, 45, 56, 64, 70

74-75

62

4-47

12

12-15

4-37

72

TABLE 2

AUSTENITIC STAINLESS STEELS INVOLVED IN DISSIMILAR METAL WELDED JOINTS

TYPICAL ANALYSIS ON PRODUCTS

-P» O

GRADE

304 (Z6Cn 18-10) SA 213 TP 304 H SA 351 CF8 321

SA 213 TP 321H Z6 CNT 18-12

SA 213 TP 347 347 SA 213 TP 347H X8 Cr Ni Nb 16-13 316

SA 213 TP 316H Esshete 1250 X 8 Cr Ni Mo Nb 16-13 800 H Sanicro 71

C

« 0,08 0,04-0,10 « 0,08 (0,06)

0,04-0,10 0,04-0,08

< 0,08 (0,06) 0,04-0,10

(0,08)

(0,06)

0,04-0,10 (0,055) (0,1)

(0,09)

< 0,10 (0,03)

Mn

« 2 < 2 « 1,5 (1)

« 2 « 2

« 2 (1) « 2

(1,40)

(1,5)

« 2 (1,7) (6)

(1,30)

« 1,5 (0,49)

Si

< 0,75 Í 0,75 S 2,0 (0,5)

< 0,75 « 1

< 0,75 (0,5) ΐ 0,75

(0,60)

(0,5)

« 0,75 (0,6) (0,5)

(0,44)

Í 1,0 (0,5)

Ni

8-11 8-11 8-11 (10)

9-13 11-13

9-13 (12) 9-13

(13,23)

(13)

11-14 (13,5) (15) (12,67)

30-35 (71,8)

Cr

18-20 18-20 18-21 (18)

17-20 17-19

17-20 (18) 17-20

(16,47)

(17)

16-18 (16,5) (10) (16,98)

19-23 (17,1)

MO

« 0,5

(2,5)

2-3 (2,3) (1) (2,32)

Ti

(0,5)

4C-0.6 4C-0.6

(0,25)

0,15-0,60 (0,39)

Nb

10C-1,0 (0,8) 8C-1,0 (1)

(0,96)

B

0,0015-0,0060

Al

0,15-0,60

Fe

(8,1)

REFERENCES

72 9-35-47 34 5, 10, 16, 37, 39, 40, 43, 44, 45, 75 7, 8, 9, 35, 47, 52

10, 11, 12, 15, 16, 31,45 35 35, 36, 47, 70 62 10, 11, 12, 15, 16, 18, 21, 22, 32 ,39 ,56, 64 29-47 11 10, 11, 16, 28

62 39, 41, 42 29

TABLE 3

WELD METALS INVOLVED IN DISSIMILAR METAL WELDED JOINTS

GRADE

121 Cr 2 LC (0,5 * Mo) 2 1/4 Ce Mo Nb-t

306L

E 300L E 308 Ho F O K Cn 16-13

309L

E 309 ER 309 25 Cr 12 NI

E 309 Cb 1« Cr « NI 1 Nb E 347 (1« Cr 12 Ni 1 Nb) E 310 316 Ι­Ε 31« E 16-1-2 Araex QT ( 17 Cr β NI 2 Mol Inconel 62/ER NI Cr 3 I n e ™ · ! 92/ER NI Cr 6 Incori.1 132 E NI Cr Fel Incon.l 182 E NI Cr Fe 3 Inconel λ E NI er r. 2 Fox 50 Ni Nb NCF Cb

PROC.

Electrodes Buttar Ing SAW-Buttering Kl.et rod.Β El.et rod.B Elect rod·« SAW-Buttering El.et rod.η TIO El.et rod.o Buttering Electrodes Electrodes

Electrodes

Electrodes

Electrodes Electrodes

Electrodes

TIO

TIO

Electrodes

Electrodes

Electrodes

Electrodes

Electrodes

C

ί 0,05

< 0,030

ί 0,04 < 0,06 (0,141

ί 0,04

< 0,15 ί 0,12

< 0,12

< 0,0« (0,07) < 0,20 (0,031 ί 0,0t ί 0,10

(0,07)

< 0,1

ί 0, OB

ί 0,06 (0,053)

ί 0,1

ί 0,1

(0,07)

(0,07)

Μη

< 0,9

< 2

0,5-2,5 0,5-2,5 12,60)

« 2

0,5-2,5 1-2,5

è,5-2,5

0,5-2,5 (2,5) < 1,25

0,5-2,5 0,5-2,5

(2,5)

2,5-3,5

2,0-2,7

« 3,5 (0,97)

5-9

1-3

(1,911

(1,14)

Si

ί 0,6

< 1,50

ί 0,9 ί 0,9 (0,22)

< 1,5

< 0,9 0,3-7

ί 0,90

< 0,90 (0,75) < 0,75

ί 0,90 ί 0,60

(0,61

< 0,5

< 0,5

< 0,75 (0,24)

« 1

< 0,75

(0,36)

(0,31)

Ρ

ί 0,04

ί 0,025

ί 0,04 ί 0,04

< 0,025

< 0,04 ί 0,03

$ 0,04

< 0,04 (0,024) < 0,04

< 0,04 < 0,03

(0,02)

< 0,03

< 0,03

< 0,03 (0,000)

«ί 0,03

ί 0,03

S

ί 0,03

ί 0,025

< 0,03 < 0,03

< 0,025

ί 0,03 ΐ 0,03

< 0,03

ί 0,03 10,011) $ 0,03

ί 0,03 ί 0,03

10,02)

< 0,015

ί 0,015

ί 0,02 10,006)

ί 0,015

ί 0,02

Cr

19-21

16-21 16-21 (15,971

22-26

22-25 23-25

22-25

16-21 (19,1) 25-26 117,7) 17-20 14,5-16,5

(17)

16-22

14-17

13-17 (16,33)

13-17

13-17

110,66)

(14,64)

Ni

9,5-11,5

9-11 9-12 (12,57)

11,5-13,5

12-14 12-14

12-14

9-11 (9,9) 12-14 (10) 11-14 7,5-9,5

161

ι 61

ι 6Ί

> 62 (70,64)

ι 5' ι 62

(49,411

69,14)

Mo

0,4-0,65

< 0,75 2-3 (0,24) ί 0,5

< 0,75 < 0,75

ί 0,75

< 0,75

< 0,5 (2,29) 2-3

1,0-2,0 (21

0,5-2,9

Nb

(1)

0,7-1

6XC-1 11)

KO,02) < 0,75

2-3

1,5-4 (1,931

1-2,5

0,5-0

(1,24 1

10,65)

Ti

KO,05)

< 0,75

2,5-3,5

(0,11)

< 1

re

< 3

< 6

< 11 (9,241

< 10

< 12

(34,5)

112,·)

Cu

< 0,75 < 0,75

< 0,75 < 0,75

< 0,75

< 0,75 (0,16) < 0,5

< 0,75

< 0,75

*J 0,5

3 0,5 < 0,5 (0,06)

*j 0,5

< 0,5

STANDARD

RCC-M

SFA 5-4 SFA 5-4

RCC-M

SFA 5-4 SFA 5-9

SFA 5-9

SFA 5-4

SFA 5-4

SFA 5-4

SFA 5-4

SFA 5-14

SFA 5-14

SFA 5-11

SFA 5-11

SFA 5-11

REFERENCES

3, 6 4, 37 15, 20

71, 72, 73

72 62 71, 73

5, 6, 7, 6 16, 37, 45, 46, 64, 72

4 10

16, 15, 30 4, 56, 70 21 10, 12, 13, 16

19, 22

4, 5, 6, 40 43, 44, 45

4, 16, 47

4, 5, 6, 40 43, 44 13, 16, 17 19, 21, 26 41, 45

4, 5, 13, 16 17, 26, 32 47, 54

62

62

Λ ^errperätu/LSL '¿L • -hyC)í2.X \JLi_pÌTÌAAi2-

-f* ho

650

€00 Η · o · ·

55Ö-- •^iriftjL -fco OLi_prtu>ia.

tkl Olirsi

Λ01

40' AtH /lo-

FIGURE 1

DISSIMILAR METAL WELDS WITH STAINLESS STEEL AS FILLER METAL

CREEP TESTS TO RUPTURE

A Te/Y.p2VAf~tur-fì °c

S5D

5Ό0.

· · ·

• ·

• ·

£50-

Λο*- Ab' faH

(fej Λ O*

FIGURE 2 DISSIMILAR METAL WELDS WITH NICKEL BASED FILLER METAL

CREEP TESTS TO RUPTURE

* ■

*·

A T^pu^Q-WeX. °(L.

ε so

6Ό0

550

o

o

·— ·

· · • · · <4<6

• · ·

• ·

cxX>

ο Φ

* ·

o

Φ e«v>û5S. NAJSJ-C/ dypQ ρ ~fceJLl-g.

/s ioJLOOLQ ¿OVy/ -{rr'tlhyL. -Le_£f-i

Ö O

O .o o Γ "o , l

/lo 3 X /o

FIGURE 3

Ti οίε ih)

DISSIMILAR METAL WELDS WITH AUSTENITIC FILLER METALS

αϊ

A ~Τοι* Κην./θ7.ι„ o . ° C_

Ç5o

Sbo-

· · ·

• 4»

· · • ·

0 0

m · m · · ·

O ìk\ -k £i>-h"e_úL< - / ^ . i Piltra

o o

550. TÍIWA Γ (Λ)

/ I D ' λ o k

FIGURE 4

λ*1

DISSIMILAR METAL WELDS WITH NICKEL BASE FILLER METALS

European Communities — Commission

EUR 13083 — Mechanical behaviour of dissimilar metal welds

C. Escara vage

Luxembourg: Office for Official Publications of the European Communities

1990 — VI, 45 pp., tab., fig. — 21.0 x 29.7 cm

Nuclear science and technology series

EN

ISBN 92-826-1895-1

Catalogue number: CD-NA-13083-EN-C

Price (excluding VAT) in Luxembourg: ECU 5

This report addresses the problems of dissimilar metal welds connecting an austenitic stainless steel component to a ferritic steel component. In LMFBRs such welds appear at the junction of the austenitic stainless steel vessel with the ferritic steel roof and in sodium and water or steam pipes. The latter are exposed to high temperatures in the creep range. A wide range of austenitic stainless steels and ferritic steels (carbon steels, low allow steels and alloy steels) are covered; the study encompasses more than 20 different weld metals (austenitic stainless steels and nickel base alloys). The report begins with a presentation of the materials, geometries and welding procedures treated in the study, followed by a review of service experience from examinations of dissimilar metal welds after elevated tem­perature service, in particular failed welds. Results of laboratory tests performed for reproducing service failures are then discussed. A further section is devoted to a review of test results on fatigue behaviour and impact toughness for dissimilar metal welded joints when creep is not significant. Finally, the problem of residual life assessment is addressed. A set of recommendations concludes the report. They concern the material selection, welding procedure, life prediction and testing of dissimilar metal welds.

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