The application of welded, bolted and riveted connections inHSLA steel in structures subjected to high-dynamic loadingCitation for published version (APA):Overbeeke, J. L. (1991). The application of welded, bolted and riveted connections in HSLA steel in structuressubjected to high-dynamic loading. (EUR; Vol. 13626). Luxembourg: Commission of the European Communities.

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Commission of the European Communities

technical steel research

Properties and service performance

The application of welded, bolted and riveted connections

in HSLA steel in structures subjected to high-dynamic loading

5 Commission of the European Communities

technical steel research

Properties and service performance

The application of welded, bolted and riveted connections

in HSLA steel in structures subjected to high-dynamic loading

J. L Overbeeke Eindhoven University of Technology

Laboratory for Structural Fatigue Eindhoven

The Netherlands

Contractor: Stichting Staalcentrum

Rotterdam The Netherlands

Contract No 7210-KD/607 (1.1.1985-30.6.1989)

Final report

Directorate-General Science, Research and Development

1991 PARL EUROP B¡bl¡oth.

N.C. EUR 13626 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, 1991

ISBN 92-826-2819-1 Catalogue number: CD-NA-13626-EN-C

© ECSC-EEC-EAEC, Brussels • Luxembourg, 1991

Printed in Belgium

THE APPLICATION OF WELDED, BOLTED AND RIVETED CONNECTIONS IN HSLA STEEL IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

by J.L. Overbeeke

Summary It is well known that the fatigue resistance of a construction depends heavily upon the quality of its joints. Therefore the application of High Strength Steels, which can lead to a reduction in structural weight and operational costs, is only justified when the quality of the joints are upgraded. This report describes an extensive investigation into the quality of fixed connections in H.S. steel with regard to fatigue. The material used was hot rolled H.S. steel, complying with FeE 560TM. The plate thickness was mainly 6 mm and the types of connections investigated were welded, bolted and rivited joints. From the above it follows that this research applies in the first place to the vehicle industry where weight savings increase the efficiency of transportation considerably. Using the results of our previous research described in EUR 9964 as a firm base, the joining technology was further evaluated and upgraded with regard to the fatigue life, as far as was thought feasable economically. As a follow up S-N curves at R = -1 were produced from constant amplitude and variable amplitude (Spectrum Gauss) fatigue tests. In general endurances up to 2 x 10 were determined. Welded joints It was proved again that overmatching the weld by 40 to 50 HV increases the fatigue strength considerably as does a smooth transition at the weld toe. Pulsed MIG welding with metal cored wire and MIG welding with flux cored wire proved favourable with regard to the latter effect. The improvement by subsequent TIG dressing depends on the Heat Input and on the initial shape of the toe. The better the toe shape the less the improvement can be. For high quality fillet welds the cracks emanate from the weld roots and it follows that full penetration welds become necessary to get the full benefit from the improvements. In addition to the fatigue data given in EUR 9964, S-N curves for CA and VA loading were i.a. generated for non-load carrying longitudinal fillet welds (end welds as welded or TIG-dressed), for load carrying transverse fillet welds, both in 6 mm FeE 560 TM and for butt welds in 9 mm FeE 560TM plate material. The endurances and fatigue strengths obtained were high in comparison with those given in BS 5400. Furthermore spot welded joints in 3 mm FeE 420TM were also investigated and tested in fatigue.

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Friction grip bolted joints As for yield controlled torqueing of HS bolts the friction of the screwthread determines the bolt efficiency, several "lowfriction" coatings were evaluated. The lowest coefficient of friction recorded was u. = 0.09. Furthermore semi-automatic torqueing in yield control proved unreliable. A newly developed high friction primer was applied to friction-grip bolted joints and yielded u » 0.70. Fatigue tests at R = -1 on a heavy duty friction grip joint having this primer on the plate surfaces showed that under V.A. loading with peak stresses up to 0.8RQ2 endurances in excess of 10 are obtained. Riveted joints The joining technique for the recently developed H.S. rivets, see EUR 9964, was further improved with regard to the degree of filling of the hole. As a result of this, the cracks in the riveted joints emanated from just outside the hole by fretting fatigue. Furthermore the minimum dimensions for rivet patterns were established and S-N curves (CA and VA loading) for 3 different joints having up to 4 rivets were generated. Tests on 5 truck frames having a fatigue critical cross member connection have shown that the improvements obtained for small specimens are also realized for complex structural connections under service-type loading. Applications Because of the large amount of fatigue data generated the S-N curves can be used on the one hand as basic design curves for constructions to be built in H.S. Steel and on the other hand for the evaluation of cumulative damage calculation methods applied to these constructions.

IV -

THE APPLICATION OF WELDED. BOLTED AND RIVETED CONNECTIONS IN HSLA STEEL IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

FIGURES OF PART I Fig. 1. Specimen. Fig. 2. Endurances of unnotched specimens, Heat II. Fig. 3. Endurances of unnotched specimens, Heat III. Fig. 4. Endurances of FeE560TM plate materials. Fig. 5. Dimensions of welded specimens, FeE560TM. Fig. 6a. Layout of the welded specimens. Fig. 6b. Clamping device used during welding. Fig. 7. Hardness profile, butt welded joint, 6mm plate

FeE560TM. Fig. 8. Cross section of the welds (typical). Fig. 9. Principle of Pulsed MIG welding. Fig. 10. Edge preparations for the butt welds. Fig. 11. Screening fatigue tests, weldment A (MMA). Fig. 12. Screening fatigue tests, B (MIG). Fig. 13. Screening fatigue tests, weldment C (Pulsed MIG). Fig. 14. Screening fatigue tests, weldment D (Servo Adjusted

MIG) . Fig. 15. Summary of screening fatigue tests,

weldments A through D. Axial loading, R = -1.

Fig. 16. TIG-dressed fillet weld, H.I. = 1.0kJ/mm. Fig. 17. Effect of Heat Input during TIG dressing on the

endurance of fillet welds. Fig. 18. Problems encountered during welding:

a. top side, b. bottom side (torn tack weld), c. distortion.

Fig. 19. Endurances of butt welded joints, C.A. loading, 6mm plate, Heat I [1] .

- V

Fig. 20. Endurances of butt welded joints, V.A. loading, 6 mm plate Heat I [1].

Fig. 21. Endurances of butt welded joints, C.A. loading, 9 mm plate, Heat III.

Fig. 22. Weld profile, 9 mm material, SA welding: a. cross section (V = 4.5), b. weld toe at secondary initiation (V = 500).

Fig. 23. Cross section of cruciform joints: a. Definition of incomplete penetration, b. Series II, 2a » 5mm, c. Series III, 2a =s 1mm.

Fig. 24. Endurances of cruciform joints, C.A. loading. Fig. 25. Endurances of cruciform joints, V.A. loading. Fig. 26. Weld root failures © observed here in comparison

with transition curves after Gurney [11]. Fig. 27. Endurances of attachments, C.A. loading.

Endwelds C. Fig. 28. Endurances of attachments, C.A. loading.

Endwelds A. Fig. 29. Endurances of a longitudinal n.l.c. fillet weld.

FeE560TM, 6mm plate. Axial loading, C.A.

Fig. 30. Endurances of attachments, V.A. loading. Endwelds C.

Fig. 31. Endurances of attachments. V.A. loading. Endwelds A.

Fig. 32. Endurances of a longitudinal n.l.c. fillet weld. FeE560TM, 6mm plate. Axial loading, V.A. Gauss, I = 0.99.

Fig. 33. Specimen FeE420TM plate material, t = 3mm. Fig. 34. Test results C.A., R = -1 on FeE420TM plate

material, t = 3mm. Fig. 35. Nugget size from peel tests versus welding current

3mm FeE420TM. Fig. 36. Spotwelded specimens:

a. Series SI, b. Series S3.

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Fig. 37. Results of CA. tests, spotwelded joints series SI and S3.

Fig. 38. Results of V.A. tests, spotwelded joints Series SI and S3.

Fig. 39. Buttwelded specimens. Fig. 40. Buttwelded joints, details:

a. Angular distortion, b. Distortion in weld direction, c. Profile MIG welding, d. Profile Pulsed-MIG welding.

Fig. 41. Results of C.A. tests, buttwelded joints. Specimens straightened by bending.

Fig. 42. Results of C.A. tests, buttwelded joints. Welding: one sided, Pulsed MIG process. Stretched 2% after welding.

Fig. 43. Results of C.A. tests, buttwelded joints. Welding: one sided MIG process. Stretched 2% after welding.

VII

FIGURES OF PART II Fig. 1. Dimensions of flanged(head) bolts and nuts. Fig. 2. KIscc for high strength steels, reference [3]. Fig. 3. Influence of surface treatment on H-embrittlement

of different bolt-steels [6]. Fig. 4a. Fv as a function of Ms, thread coating T. Fig. 4b. Fv as a function of Ms, for different coatings. Fig. 5. Influence of u for different low-friction coatings

on Fv and Ms. Fig. 6a. Determination of ΜK.

Screwthread : coating M. Bolt/nut faces : dry. Plate surface : high friction primer + M.

Fig. 6b. Determination of μk. Screwthread : coating M. Bolt/nut faces : coating M. Plate surface : high friction primer.

Fig. 7. M versus ø during pretensioning: a. light weight technician, b. heavy weight judoka.

Fig. 8. Record of friction force F versus displacement 5 during 5 cycles, for a frictiongrip bolted joint.

Fig. 9. Joints used for testing frictiongrip joints. Fig. 10. M - ø records:

a. primer WAP 9.4, b. primer WAP 9.5.

Fig. 11. (F a) max related to Mmax during pretensioning. Fig. 12. HRC related to M.„. Fig. 13. Results of C.A. tests, R = -1 for H.D. friction grip

bolted joints. Bolt: M12, 10.9, Coating T, yield control. Plate: primer WAP 9.5.

Fig. 14. Results of V.A. tests, Spectrum Gauss, R = -1 for H.D. friction grip bolted joints.

VIII -

FIGURES OF PART III Fig. 1. Specimen with a central hole.

Kt = 3.07 ref. to gross section stress. Kt = 2.63 ref. to net section stress.

Fig. 2a. Results of C.A. tests, central hole, Series 1. Drilled hole, Heat II material.

Fig. 2b. Results of C.A. tests, central hole, Series 2a. Punched hole, Heat II material.

Fig. 2c. Results of C.A. tests, central hole, Series 2b. Punched hole, Heat I material.

Fig. 2d. Results of C.A. tests, central hole, Series 3. Rejected punch, Heat II material.

Fig. 3. Summary of C.A. tests, central hole. Fig. 4. Crack initiation sites. Fig. 5. Dimensions of the rivet. Fig. 6a. Shape of a punched hole. Fig. 6b. Stapling of the plates. Fig. 6c. Sharpened transition after riveting. Fig. 7. Standardized and modified rivet heads. Fig. 8. Results due to riveting with 2 conical dies:

a. round head die, b. conical die. Fig. 9. Results of V.A. tests, S.S. lapjoint riveted with 2

conical heads. Fig. 10. Stress corrosion tests:

a. Forming matrix, b. Rivet, before (1) and after testing (2), c. String of immersed rivets.

Fig. 11. Dimensions of riveted lapjoints having one rivet: a. single shear, c. exentric single shear, b. double shear, d. test setup, spec. c.

Fig. 12. Results of C.A. tests, S.S. lapjoint.

Fig. 13. Results of C.A. tests, D.S. lapjoints.

Fig. 14. Results of V.A. tests, S.S. lapjoints.

Fig. 15. Results of V.A. tests, D.S. lapjoints.

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Fig. 16. Results of V.A. tests, S.S. lapjoint, e = 9mm. Fig. 17. Specimens with more than one rivet. Fig. 18. Rivet patterns:

a. specimen used for bulge tests, b. rivet pattern dimension.

Fig. 19. Results of C.A. tests, joint with 3 rivets in series.

Fig. 20. Results of V.A. tests, joint with 3 rivets in series.

Fig. 21. Crack initiations from fretting. Fig. 22. Load transmission, connecting elements in series,

a. highly elastic plate, stiff connections. b. stiff plate, highly elastic connections.

Fig. 23. Results of C.A. tests, joint with 3 rivets parallel.

Fig. 24. Results of V.A. tests, joint with 3 rivets parallel.

Fig. 25. Results of C.A. tests, joint with 4 rivets in a square.

Fig. 26. Results of V.A. tests, joint with 4 rivets in a square.

Fig. 27a. Tractor frame. Fig. 27b. H-shaped cross member with H.S. rivets. Fig. 28a. Frame A (reference). Fig. 28b. Frame B. Fig. 28c. Frame C. Fig. 28d. Frame D. Fig. 28e. Frame E.

PREFACE AND GENERAL INTRODUCTION

It is well known that the fatigue resistance of a construction depends upon the quality of the joints of that construction. This implicates also that the total structural weight depends heavily on this quality because mostly it is economically not justified to waist down the (often prismatic) body of the structural elements. Also the application of High Strength Steel, necessary for further reduction in weight and costs, is not justified unless the resistance against fatigue of the joints is upgraded. It is clear that the above is of extreme importance for the vehicle industry, where weight penalties interfere quite heavily with the economics of transportation. Therefore a Dutch research group in which participated the Laboratories or Research Centers of - Eindhoven University of Technology,

DAF-Trucks, Eindhoven, - Hoogovens IJmuiden, IJmuiden, - Smitweld (Lincoln Norweld Group), Nijmegen,

Filare Welding Industries (ESAB Group), Utrecht, NEDSCHROEF Helmond, Helmond, AKZO Coatings Nederland, Sassenheim,

has taken as their subject the improvement of the fatigue resistance of fixed connections in high strength structural steel plate and has carried out an extensive research program.

The large amount of results obtained for different joints and for a variety of joint configurations together with the interconnections that are laid between these results enhance greatly the reliability of the individual test series. Therefore, at least to the opinion of the author, the results obtained are suitable for use in the design office. A pre-requisite is of course, that the load spectra on the construction under design are really known.

Almost all tests on the optimized structural joints were carried out in duplo, viz.

One test series with conventional constant amplitude loading. One test series with variable amplitude loading from the standard spectrum GAUSS.

So on the one hand a significant contribution is made to the art of, or better the use and the avoidance of abuse in, damage calculations from C.A. data.

- XI -

On the other hand the V.A. results obtained under the spectrum GAUSS allow the conversion to other spectra with the help of a relative damage rule with a much higher accuracy than ever can be obtained by damage calculations from constant amplitude data. The joints investigated are:

welded joints - friction grip bolted joints - riveted joints with high strength rivets. It is clear that this investigation builds heavily on the results laid down in report EUR 9964 [1] (Final Report to convention ECSC, 7210.KD/605). In fact this latter investigation was the base from which this investigation started, and therefore it was sometimes necessary to repeat tests already reported in [1] The main results of these two investigations were, with permission of the ECSC, recently presented at the International Conference on "Engineering Integrity through Testing" (EIS*90) [2,3] in order to make the vehicle /construction industry of the European Community aware of our results. Other often used fixed connections for constructions from steel plate material are

spot welded joints An extensive evaluation of the fatigue characteristics of spot welded joints in high strength steel plate was performed bij Eindhoven University in collaboration with DAF Trucks and published a.o. in Welding Research International.[4,5,6] (Part of this research was also performed with financial aid of the CECA under contract 6210-45/6/602) These publications did re-initiate research on spot welded joints on a world wide scale. Some additional tests on spot welds in thin plate material were also carried out within this investigation. Altogether these 3 investigations cover the fatigue behaviour of current fixed connections in high strength steel plate and show how they can be upgraded by a careful choice of production methods. As the production technology of the different joints mentioned above have little in common, this final report is divided into three parts, viz: Part I Welded joints Part II Friction grip bolted joints Part III Riveted joints with high strength rivets.

XII -

These parts are made more or less self-contained. So is e.g. in Part I rather much attention given to the properties of the applied TM steel, but in Part II and III the material properties are - as far as useful -repeated again. The reason for this was to increase the readability of this voluminous report. Acknowledgements As mentioned before, this investigation was carried out in collaboration with 6 dutch companies and would not have been possible without their expertise. Therefore the author wishes to thank the following persons : J.F. Flipsen J.H.M. Martens H.T.M. van Lipzig T.J. Alting N.G.J. Nühn F.W. Schmidt H. Meelker H. Feringa B. Stenneke

of DAF Trucks

of Hoogovens IJmuiden of Filare Welding Ind.

■i II 11

of„Smitweld of NEDSCHROEF Helmond of Akzo Coatings

The author also wishes to thank P.A.M. Jonkers, of our Laboratory for Structural Fatigue, who guided our fatigue tests and carefully analysed the test results. Without external support this investigation into the basic behaviour of structural connections under fatigue loading could not have reached the extent necessary for industrial applicability. Therefore the financial aid provided by the CECA by contract 7210.KD/607 for this research project together with the stimulating discussions in the former Executive Committee F4, Aciers de Construction, are gratefully acknowledged.

Eindhoven, November 1990, Prof.ir.J.L. Overbeeke

XIII

References 1. Overbeeke,J.L. : The fatigue strength of welded, bolted

and riveted joints in high strength low alloy steel. Euroreport EUR 9964 MF EN, 1985.

2. Overbeeke, J.L., Alting, T.J., Jonkers, P.A.M. : Influence of weld parameters on the fatigue strength under C.A. and V.A. loading of welded joints in HSLA steels. Proc. of the Intern.Conf. on Engineering Integrity through Testing, edited by H.G. Morgan. To be published by EMAS (1991).

3. Overbeeke, J.L., Flipsen, J.F., Jonkers, P.A.M. : Improvement of the fatigue strength of riveted joints in HSLA steel by using new high strength rivets. Proc. of the Intern.Conf. on Engineering Integrity through Testing, edited by H.G. Morgan. To be published by EMAS (1991).

4. Overbeeke, J.L. : Fatigue of spotwelded lapjoints. Metal Construction 8 (1976) p.212-215.

Overbeeke, J.L., Draisma, J.: The influence of stress relieving on the fatigue of heavy-duty spotwelded lapjoints. Welding Research Int.,7, 1977, p.241-253. Overbeeke, J.L. : The fatigue behaviour of heavy-duty spot welded lapjoints under random loading conditions. Welding Research Int.,7, 1977, p.254-276.

XV

THE APPLICATION OF WELDED, BOLTED AND RIVETED CONNECTIONS IN HSLA STEELS IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

Part I WELDED CONNECTIONS

Contents Page 1. Introduction 3 2. TM Steels 4

2.1 General 4 2.2 Metallurgy 4 2.3 Welding 5

3. Experimental program for FeE560TM material. 6 3.1 Evaluation program 6 3.2 Design data program 6

4. The FeE560 TM material 7 4.1 Compositions and Mechanical Properties 7 4.2 Endurance data 7

5. Welding the specimens 8 5.1 General 8 5.2 Specimens 9

6. Evaluation of weld parameters and processes 9 6.1 Hardness of the weld and its profile 9 6.2 Influence of the welding process 10 6.2.1 Welding processes used 10 6.2.2 Edge preparations for butt welds type I and II 11 6.2.3 Screening by endurance tests 11 6.2.4 Correction for excentricity of the lap joints 11 6.3 TIG dressing 12

7. Endurance data for welded joints in FeE560TM 13 7.1 Types of fatigue loading 13 7.2 Secondary stresses and misalignments 14 7.3 Butt welded joints 15 7.3.1 Butt welds in 6mm plate material, Heat I 15 7.3.2 Butt welds in 9mm plate material, Heat III 15 7.4 L.c. cruciform joints (Heat II) 17 7.5 Longitudinal n.l.c. fillet welds (attachments) 18

8. Experimental program for the FeE420TM material. 20 8.1 General 20 8.2 The FeE420TM plate material 20 8.2.1 Composition and mechanical properties 20 8.2.2 Endurance data 20 8.3 Spot welded joints 20 8.3.1 Spot welding parameters 20 8.3.2 Endurance data for spot welded joints in

FeE420TM 21 8.4 Butt welded joints 22 8.4.1 Welding processes used 22 8.4.2 Specimens and specimen distortion 22 8.4.3 Endurance data for butt welded joints in

FeE420TM 23

9 . C o n c l u s i o n s 24 1 0 . R e f e r e n c e s 25

TABLES 26

FIGURES 43

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THE APPLICATION OF WELDED, BOLTED AND RIVETED CONNECTIONS IN HSLA STEELS IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

Part I W E L D E D C O N N E C T I O N S

1. INTRODUCTION New generations of High Strength Low Alloy (HSLA) steels as are the micro-alloyed, controlled rolled steels have opened possibilities to build lighter and more efficient constructions. For constructions that are loaded predominantly by static loads,the application of high strength steels does not cause many problems because the paramount design criteria are yield strength, deflection and stability. For constructions that experience life loads which result in fatigue damage, things are different. Apart from the just mentioned static criteria, it is the resistance against fatigue that determines the endurance and it is well known that the fatigue resistance depends at large on the detailed shape of the* structural connections.

So in general the quality of the joints and not the grade of the material, determines the fatigue endurance. This is particularly true for welded constructions, because there is no general relationship between the yield strength of the material and the fatigue strength of welded joints. The reasons for this lack of correlation are: - the unknown geometrical stress concentration factor at

the weld toe. - inclusions in the weld near the fusion line and within

the region of influence of the stress concentration from the toe.

Because of this, and also because the residual welding stresses are usually positive near the surface, the endurance is usually dominated by the growth of macrocracks (defined here as cracks that grow according to LEFM based rules, e.g. Paris law). According to the IIW-Recommendation [2] the growth of macrocracks in structural steels can be described by da/dn = 0.3*10~12 AK3 , independent of the tensile or yield strength of the material. Therefore the Basquin type equations used as design curves for welded joints are generally of the form AS3N = constant (e.g. BS 5400, NEN 2063).

In view of the above, the application of high strength steels in welded constructions for which fatigue is a dominant design criterium is only justified when the

- ̂

resistance against fatigue is upgraded through a careful choice of shape and quality of the welded joints. The investigation described below deals with this aspect. Welded joints in a HSLA steel were optimized with regard to fatigue by evaluating the influence of consumables and welding processes. Furthermore endurance data for constant amplitude (C.A.) loading and variable amplitude (V.A.) loading were generated for several basic types of welded joints. The plate materials used in this investigation are high strength, micro-alloyed steels produced by controlled hot rolling. This type of steel and the plate thicknesses used, are applied extensively in the truck- and heavy vehicle industry. TM STEELS

2.1 General Thermo-mechanically treated steels derive an essential part of their yield strength, their low transition temperature and their cold-forming qualities from a very fine-grained ferritic structure together with precipitation hardening, a low carbon content (C <0.12%) and inclusion control. The differences between the available grades, E275 up to and including E560, depend primarily on differences in manganese content (up to 1.8%) and in the amount of micro-alloying elements. The applicable standard is Euronorm 149-2. Nowadays the most important unit in modern steelworks for producing formable TM steels is the hot strip mill.

2.2 Metallurgy The micrq-aljoying elements Nb, Ti and V are used to form carbo-nitri<|e precipitates, it follows that during the hot rolling under controlled conditions (reduction and temperature) the austenite grain boundaries are pinned by these precipitates so that the grains become elongated and flat. Recryştallisation within these grains refines the structure so that the final grain size is reduced to 5 urn (ASTM size 12) or less. In order to control the shape of the sulfide inclusions, additional elements (Ca, Ce) are added to the melt. This results in improved properties in the transverse direction, which are important for cold-forming. After rolling, the strip is coiled and cooled down in still air. The low cooling rate obtained in this way causes a further precipitation and improves the

. A .

mechanical properties. Before the coiled strip can be used, it is decoiled and straightened. This straightening is performed by repeated rolling in bending, which leaves an equilibrium system of residual stresses in the flat plates. When, in order to manufacture parts, the plates are cut into smaller pieces, the residual stresses are partly released and this results in distortion.(Note however, that flat rolled products from a plate mill also contain considerable residual stresses).

2.3 Welding Because of their low carbon content the TM steels have a good weldability. However, this also implicates that in part of the heat affected zone (HAZ) the drop in hardness (mainly resulting from a loss of precipitation hardening) is not compensated by structural hardening effects. Therefore the higher grades TM steels show a drop in hardness within their HAZ. For the FeE560TM steel as used here, this drop was about 30-40 HV5. An example of a hardness profile is given in fig.7, For the FeE420 TM steel also used here the drop in hardness is negligible. A more detailed description of what happens in the HAZ is given by Suzutki [3]. Furthermore, and this holds for all structural steels, the level of the residual welding stresses and the resulting distortions depend on the yield strength of the material. It follows that for welding high strength steels precautions are necessary to prevent (excessive) distortion, e.g. the use of very heavy tack welds and/or stiff fixtures (see also par. 7.2 and fig. 18).

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3. EXPERIMENTAL PROGRAM for FeE560TM material. The experimental program was split up in 2 parts viz. an evaluation program and a design data program.

3.1 Evaluation program Within the first part of the program, the influence of the consumables and welding processes were evaluated on the base of fatigue tests in order to obtain an optimum in fatigue resistance. The evaluation also included the dressing of welds by the TIG process. The fatigue tests carried out were constant amplitude tests at R = -1 at two or three load levels. Emphasis was on: - The hardness of the weld in comparison with that of the plate [1] .

- The influence of the welding process on the smoothness of the weld and of the weld toe

For the 6 mm plates the common MIG process, 2 more sophisticated MIG processes and Manual Metal Arc welding were evaluated. For the 9 mm plate material the emphasis was on high capacity welding processes, viz. Submerged Arc and Plasma MIG welding. However, for this latter process no stable arc could be obtained due to magnetic blast. Therefore as a second process the Pulsed MIG process was applied.

3.2 Design data program Within the second part of the program, S-N data were generated for the joints made with welding processes that showed the better fatigue performance. The fatigue tests were carried out at R= -1 and - C.A. loading - V.A. loading with the standard spectrum GAUSS (see par

7.1) at I = 0.99 S-N and Srms-N curves were determined for: - plate material, 6 and 9 mm thickness, - butt welded joints in 9 mm plate, fig. 5a, - cruciform joints with load carrying fillet welds in

transverse direction, fig. 5c, - attachments with non-load carrying fillet welds in

longitudinal direction, fig. 5d. The end welds were as welded or TIG-dressed.

Joints with fillet welds were made in 6 mm plate material only. The joints tested are of the more important types in welded constructions, and it is believed that the results presented here will give a designer a firm base for the application of HSLA steels.

6 -

4. THE FeE560TM MATERIAL 4.1 Compositions and Mechanical Properties

6 mm plate material The plates of thickness 6 mm were obtained from Hoogovens IJmuiden B.V. from 2 heats of LD steel. Heat I was an ingot cast steel, micro-alloyed with Nb and V and treated with Ce. This heat was used throughout in our previous investigation (Overbeeke [1]) but was used here only for the evaluation of TIG-dressing, see par 6.3. Heat II was a concast steel, also micro-alloyed with Nb and V but treated with Ca. Both heats were hot rolled. Their composition and mechanical properties are given in table 1. All specimens were taken from one coil (per heat). 9 mm plate material The hot rolled plates of thickness 9 mm were obtained from Thyssentahl A.G. from one coil of a production heat (called III) of LD steel. It was an ingot cast steel, micro-alloyed with Nb and Ti and treated with Ca. Its composition and mechanical properties are also given in table 1.

4.2 Endurance data For determining the S-N curves and the fatigue limit, Sf, of the plate material, specimens according to fig.l were used. The specimens were taken in the longitudinal direction from one or two strips and together they covered the full width of the plate (1000 mm). The machined edges of the specimens were rounded off slightly in order to prevent premature crack initiation from sharp corners. The results for Heat I were reported in [1] The results for Heat II and III are given in table 2 and fig.2 respectively in table 3 and fig.3. The Basquin type equation and the fatigue limits as determined bij the Staircase method are: 6 mm plate, Heat I [1] 8.06 log Sa + log N = 26.208 Sf = 360 N/mm2(50% survival); Sf/Rm = 360/716 = 0.50 (1) 6 mm plate, Heat II. 9.89 log Sa + log N = 30.546 Sf = 304 N/mm2(50% survival); Sf/Rm = 304/694 = 0.44 (2) 9 mm plate, Heat III. 7.64 log Sa + log N = 24.629 (3) Sf = 262 N/mm2(50% survival); Sf/Rm = 262/658 = 0.40 The s.c.f. of the specimen, Kt = 1.07, is not taken into

account in the above figures, because many cracks initiated away from the (rounded) corners. The S-N lines for the 3 Heats together are given in fig.4 As the Sf for Heat I is high as compared to the Sf of Heat II and III, additional tests were carried out. On specimens stretched 3% before machining in order to remove the curvature of and the residual stresses in the specimens. These curvatures, see fig.l, were up to 1.6 mm for 6 mm specimens and up to 2,0 mm for 9 mm specimens taken from the center part of the plate width. The results of CA tests are given in table 4 and from this table it follows that the Sf remained the same. On specimens from which the (thin) millscale was removed by mesh 80 emery paper. The results for 6 mm and 9 mm specimens are given in table 5 and it follows that removing the millscale increases Sf by about 10%. This percentage is so low that a possible difference in "millscale quality," as compared to Heat I, can not be the cause for the lower Sf. So neither the curvature and the corresponding internal stresses, nor the mill scale are the cause of the differences in Sf/Rm ratios.The differences in Sf are apparently due to differences between the materials. It should be remembered, however, that for structural steels the much used ratio Sf/Rm = 0,50-0,55 holds for polished specimens tested in bending and that the ratios for Heat II and III are certainly not low for specimens with millscale tested in push-pull loading.

WELDING THE SPECIMENS 5.1 General

The welded joints were produced at the Laboratories of Filare and Smitweld under supervision of their welding experts. Before the test specimens were produced, welding trials were carried out in order to optimize the weld with regard to the goals to be obtained. These were in general the hardness of the weld metal together with a smooth appearance of the weld toe and a low distortion of the joint. Commercially available consumables suitable at least for IG and 3G position welding were used as far as feasible. In some cases consumables of the cored wire type were modified in order to obtain the required hardness or a better welding performance.

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For every type of weld used, butt welds were produced for static tests. All of these complied with the tensile test according to DIN 50120 and the 180° bend test according to DIN 50121.

5.2 Specimens The shapes and dimensions of the specimens used are given in fig.5 The specimens, except those with longitudinal fillet welds, were produced as follows (see fig.6a): Strips, transverse to the rolling direction of about 1 meter wide (=width of the plate) (& meter for the 9 mm plate material because of the high forces from the residual welding stresses), and 180 to 230 mm long were connected in the desired position with heavy tack welds. Then the strips were clamped in heavy fixtures about every 250 mm after which the welds were made, see fig.6b. The specimens were machined from these welded strips (fig.6a). Parts of the strips with start/stops or tack welds were disregarded. All welds were made in the IG or 2F position with mechanized welding. Before fatigue testing, all machined edges were rounded off, so that the fatigue cracks initiated in general at the centre part of the specimens.

6. Evaluation of weld parameters and processes. 6.1 Hardness of the weld and its profile

In [1] the influence of the hardness of the weld was evaluated. Screening tests with a hardness of the welds ranging from 220 HV up to 290 HV revealed that an overmatched hardness (as compared to the hardness 240 HV of the plate material) results in an increased fatigue resistance and apparently compensates for the dip in hardness to 220 HV within the H.A.Z. A typical hardness profile is given in fig.7 So also in this investigation overmatched welds (260-280 HV as compared to 230 HV for Heat II and 210 HV for Heat III) were applied. It is clear that overmatching is not the one parameter that counts. A second important one is the profile of the weld because of the local stress concentration at the weld toe. However from [1] it appears that overmatching seems more important than the toe profile. In all cases, except the MMA welding used for the end welds of the longitudinal fillet welds, a smooth transition at the toe was obtained. Fig. 8 shows a typical example for a butt weld and for a

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fillet weld, both welded with the Pulsed MIG process, (type C, table 6). Fig. 22 shows an example of a Submerged Arc (table 9) weld and fig.23 for a fillet weld, welded with the MIG proces and flux cored wire, (type B table 6)

6.2 Influence of the welding process, 6mm plate material 6.2.1 Welding processes used

For evaluating the influence of the welding process on fatigue the following processes were used:

A. Shielded Metal Arc (Manual Metal Arc) With this proces the metal rod is covered with a slag forming coating to protect the melt from oxidation. For this project a basic type electrode was used. Alloying elements can be incorporated in the rod or in the coating.

B. MIG This is often used for mechanized welding. The rod is replaced by a wire which is semi automatically fed to the weld. Protection of the weld against oxidation is obtained by shielding the weld by a gas curtain. This gas curtain can be inert (Metal Inert Gas Welding) or a mixture of inert and (surface) active gas (Metal Active Gas Welding) The latter type of shielding is more usual nowadays and it is common practice not to distinguish between MIG and MAG welding but to speak of MIG welding whatever the gas mixture is.

C. Pulsed MIG With this MIG process droplet detachment is caused by pulses of high current. In the time between these pulses the arc is maintained with a low current that is insufficient for melting-off electrode metal. The pulse frequency is regulated so that the melt-off is automatically adjusted. Fig.9 shows the process schematically. Advantages of this process are described by Sol [4].

D. Servo adjusted MIG This process provides a constant Heat Input by keeping the current at a constant level through servo control. Consumables The consumables used, except for MMA welding, were of the cored wire type, either metal cored or metal and flux cored. For this latter type, the core of the wire contains not only alloying elements but also slag forming additions. The more important welding parameters of the 4 processes are given in table 6, together with the measured hardnesses of the welds. The processes used for the 9mm plate material are

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discussed in par.7.3.2. Furthermore it should be noted that in all cases tensile tests and 180° bend tests (d = 3t) showed satisfactory results and that Charpy V-tests (tx8 mm) yielded fracture energies of 40 joules or more at -40°C.

6.2.2 Edge preparations for the butt welds type I and II. The specimens were produced as described in par.5.2. For the butt welds in 6mm plates, 2 types of edge preparations were used viz. Type I: Plates beveled 2 mm on both sides and had an

opening angle of 90°, see fig.10a. Type II Plates fully beveled on one side and 50a opening

angle for the first pass. Grinding out the weld root as shown in fig.10b for the counter pass. This method was also used in [1].

It is clear that type I preparation is much more economic than type II. However, the fatigue resistance of welds with type II preparations is on the average higher than for type I preparations, as is. shown in par. 6.2.3.

6.2.3 Screening by endurance tests. The fatigue resistances of butt welded joints, fig 5a, type I and type II and of fillet welded lapjoints fig.5b, were screened with about 8 specimens per series at 2 or 3 load levels. The results are plotted in fig.11 through 14 and summarized in table 7 and fig.15. The results from [1] (Heat I and hardness of the weld 290 HV) are given too for comparison. It appears from fig. 15 that weldment C yields the better results for the butt welds, while both weldments B and C yield somewhat better results for the fillet welds. Therefore it was decided to use the processes used for weldments B and C, at least as far as possible, for the design data programm.

6.2.4 Correction for excentricity of the lap joints. The low figures for the fillet welded lapjoints (fig.5b.) are due to the excentricity (nominal 6 mm ) of that joint. In [1] it was established that the nominal bending stress (hot spot stress) was 2.42 times the nominal axial stress, Sa. Further that the average bending stresses in the butt welded joints (due to weld distorsion) were ± 0.10 Sa. So for a fillet welded joint with zero excentricity the

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equivalent nominal axial stress is: Sa eq = 2.42/1.10 Sa = 2.2 Sa The values of Saeq are given in table 7.

6.3 TIG dressing It is now well known that remelting the weld toe by the TIG process (without consumables) produces a smooth transition radius and from this an improvement of the fatigue resistance (Haagensen [5], Bignonnet[6]). Improvements from 20% to about 100% in load amplitude have been reported for fillet welds by Minner and Seeger [7]. However, almost all published results are for joints in C-Mn steels with plate thicknesses of 12 mm and above and for these joints heat inputs of at least 1.0 kJ/mm but usually 2: 1.4 kJ/mm are used in order to obtain a smooth radius [5] . Therefore the effect of the heat input on this 6 mm TM steel plate was investigated. The material used was from Heat I. Lap joints according to fig. 5b were used with a hardness of the weld of 290 HV5. Heat Inputs (HI) of 0.7, 1.0 and 1.4 kJ/mm were used for the TIG dressing. Table 8 gives the radii obtained and the hardnesses of the TIG weld and of the HAZ. Fig.16 shows a cross section of a TIG dressed fillet weld. Furthermore it was noted that the surface roughness of the radius increased with decreasing heat input, so with decreasing radius. The results of fatigue tests (CA, axial load, R = -1) are given in fig. 17 together with the S-N curve for the as welded condition of these lapjoints as given in [1] and it follows that: - dressing with a HI of 1,4 kJ/mm does not result in a better endurance, despite the larger radius, because of the lower hardness of the welds.

- dressing with 0.7 or 1.0 kJ/mm results in an appreciable increase (about 50%) in fatigue resistance at the same endurance and also in fatigue strength.

As the results obtained with a HI of 0.7 kJ/mm are equal to or better than those with a HI of 1.0 kJ/mm and because the costs of dressing depend directly on the HI, 0.7 kJ/mm is regarded as the better HI for this 6 mm plate material.

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7. ENDURANCE DATA FOR WELDED JOINTS IN FeE560TM 7.1 Types of fatigue loading.

All fatigue tests were carried out at R = Smin/Smax = -1. Two types of loading were used, viz. -Constant amplitude (C.A.) loading -Variable amplitude (V.A.) loading with the standard loading spectrum GAUSS at a value of the irregularity factor, I = 0.99.

The standard spectrum GAUSS (Haibach [8]) is used for general fatigue investigations (Schütz,[9]) and is based on a Gaussian distribution of the load signal, S(t), in the time domain. Its magnitude is characterized here by Srms, the root mean square value (time average) at zero mean load of the sinusoidal load signal, S(t). A block of this spectrum contains 106 upgoing zero crossings (= 106 cycles) and its maximum peak, Smax

= 5.26 Srms ) .

Its spectral density is characterized by the irregularity factor I, wich is the ratio. 1 = number of zero crossings

number of extremes 2 usefull values of I are indicated, viz. I = 0.99 : characteristic for a resonance type of

response, having a Rayleiqh-type distribution of peaks and amplitudes.

I = 0.70 : characteristic for structures with moderate damping.

Now in [1] tests were carried out with both spectra and it appeared that, on the base of Srms, (or Smax) there is hardly a difference in endurance. This holds for welded and for riveted joints. Therefore it was decided to use only one value of I. I = 0,99 was chosen because the frequencies of testing can be much higher than when I = 0,70. As the distribution of the higher peaks and troughs within a block of 106 cycles is quite regular, it is not necessary to express the endurance in fractional numbers of blocks-to-failure. Therefore the endurance is expressed in cycles, N, the

M When Smax is preferred for presenting results, as iş the case in Germany and is also preferred by Gurney [10] , the Srms values are simply to be multiplied by 5.26.

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number of upgoing zero crossings to failure. The fatigue tests were carried out with servo-hydraulic test equipment (C.A. and V.A. tests) or with a sub-resonance fatigue machine (C.A. tests). The testing frequencies were between 20 and 50 Hz. The signal for the V.A. loading was generated by a specially built pseudo-random generator. This generator produces the desired sequence of peaks and troughs from its ROM and connects each pair of successive extremes with a haversine of pi radians.

7.2 Secondary stresses and misalignments In the preceding paragraph on TM steels it was noted that strips, cut from plates, did not remain flat because of the release of the residual stresses from the production process of the plate. Furthermore there are the axial and angular misalignments due to weld distortion and for the welded lap joints there is a nominal excentricity, e = 6mm, due to the geometry of the joint itself. Therefore the strain amplitudes, appr. 20 mm away from the weld and so out of the influence of the s.c.f.of the toe, were measured with frictiontype-strain gages on several specimens of each series during the C.A. fatigue tests. The accuracy of these gages, as calibrated against bonded strain gages, is +0 / -5%. The averaged values of the measured maximum outer fiber stress, o, (= the hot spot stress) in relation to the nominal axial stress, S, are as follows: - butt joints : oa ~ 1.10 Sa - cruciform joints : Oa » 1.20 Sa - lap joints, e=6 mm,: oa = 2.42 Sa. (see also par 6.2.4) It should be noted that the allowed axial misalignment for specimens of the symmetrical type was restricted to 0.25 mm for the butt welded joints and to 1,0 mm for the cruciform joints. The fact that axial misalignment has a large influence on the endurance is shown in fig. 19 for 2 butt-welded specimens which showed an axial misalignment of 0.75 mm and also in fig.24 for 2 cruciform joints (Series III) having axial misalignments of 2 to 3 mm. That it was difficult to keep the narrow strips in position during welding is shown in fig.18. Fig 18a shows one side after welding one fillet weld. Fig 18b, the reverse side, shows that despite heavy clamps every 250 mm, see fig.6, the heavy tack weld in the middle is broken.

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7.3 Butt welded joints 7.3.1 Butt welds in 6 mm plate material, [1] .

Within this investigation no S - N data, except those given in par.6.2.3, were collected for butt welded joints in Heat II material. Instead the data generated in [1] for Heat I material and a hardness of the weld of nominal 290 HV5 are reproduced here for comparison. C.A.loading.The test results are plotted in fig. 19. The Basquin equation as determined by linear regression was: 5.47 log Sa + log N = 18.39 (4) The fatigue limit (50% survival) is estimated at Sf « 150 N/mm2

In fig. 19 the applicable class D line of BS5400 is also plotted and the improvement obtained needs no comment. V.A.loading. Endurances obtained with the standard spectrum Gauss are plotted in fig. 20. The Basquin equations as determined by linear regressions are: I = 0.99 : 8.90 log Srms + log N = 23.139 (5)

I = 0.70 : 8.22 log Srms + log N = 21.719 (6)

However, from fig. 20 it is clear that the difference between the two types of loading is not significant. In this figure the endurance line for I = 0.99 as calculated from eq. 4 with Miners rule and Sf = 0 is also indicated and it appears that these predictions are unconservative for endurances below N = 107.

7.3.2 Butt welds in 9 mm plate material. For the production of the butt welded specimens, fig.5a, from the 9mm plate material, Heat III, see table 1, two high capacity welding processes were planned, viz. : - Submerged Arc Welding - Plasma- MIG Welding However due to magnetic blast no stable arc could be obtained for the Plasma-MIG proces and therefore it was replaced by Pulsed MIG Welding. Submerged arc welding Details of the welding are given in table 9. No edge preparation was used for the first pass nor for the second (counter) pass. The weld gap was zero. The welds had a very smooth appearance, see fig.22 However the hardness of the weld was as low as 220HV5, because of the extensive mixing of the consumable and the plate material in the melt pool.

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The results of C.A. fatigue tests are given in table 10 and are plotted in fig. 21. The Basquin equation as determined by linear regression is : 2 . 7 3 l o g Sa + l o g N = 1 1 . 1 8 7 (7) and about equal to the class D line of BS5400. The fatigue strength is reached at N ss IO6 and is estimated at: Sf » 70 N/mm2 (7) which is very low indeed. In fig. 21 the results obtained from the evaluation tests on butt welds with hardness 220 HV5 [1] are plotted too and it appears that these results are within the scatterband of the S.A. welds. A cross section of the weld, fig.22a, and an enlargement (V = 500x) at a secondary crack initiation, fig.22b, show that the weld toe profile is flush. This confirms again that', at least for this TM material, the hardness of the weld is of paramount importance for the fatigue resistance. Pulsed MIG welding Details of the welding procedure are given in table 9. The edge preparation was as follows. First pass: 60° opening angle, 4.5 mm deep, zero gap. Second (counter) pass : ground to 60° opening angle, 4,5mm deep. So this edge preparation is comparable to the one used for the evaluation tests, fig. 10, type I. The hardness of the weld was 260 HV5. The results of C.A. fatigue tests are also given in table 10 and plotted in fig. 21. The Basquin equation as determined by linear regression is : 5.88 log Sa + log N = 18.923 (8) Note that the slope is comparable to the one for the C.A. results for the 6 mm butt welds, eq.4. The fatigue limit is estimated at: Sf « 160 N/mm2. (8) which is slightly higher than estimated for the butt welds with comparable hardness in the 6 mm plate material. This difference is, taking into account the low number of specimens having very high endurances, not significant. Note also that the reduction in fatigue strength (eq.3) due to the weld is only 262/160 =1.64

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4 Cruciform .joints Cruciform joints with load carrying transverse fillet welds, see fig. 5c, were fabricated with type B and type C welding from Heat II material. The hardness of the welds was 260-280 HV5. Three series, having different weld penetrations, were tested. Note that 2a, as defined in fig 23a, is an average value and that the penetration varies somewhat along the weld. C.A. loading. The test results are given in table 11 and plotted in fig. 24. Almost all failures were root failures, regardless of the weldment. Note however the detrimental influence of misalignment, see par. 7.2. The Basquin equations for root failures, determined by linear regression, are as follows: Series I: - Weldment C (table 6), partial penetration, 2a » 3 mm 3.02 log Sa + log N = 12.023 (9) Sf: not determined but < 50 N/mm2

Series II:-Weldment B, almost no penetration, 2a « 5 mm. 2.98 log Sa + log N = 11.505. (10) Sf » 35 N/mm2 (based on 1 run-out) Series III:-Weldment B, almost full penetration, 2a « 1mm 3.47 log Sa + log N = 13.300 (11) Sf » 50 N/mm2 (estimate) Comparing Series I with Series II it appears that the inverse slope m = 3.0 for both series, and that the difference is in the constant only. So the endurance is entirely determined by macro-crack propagation. Therefore Series III was fabricated with beveled plates (50°) so that the penetration was as deep as possible (2a a 1 mm) for process B with flux cored wire. This series still showed root failures, but the endurances did increase considerably. The inverse slope remained about the same. In Ch.5 of [11] calculations were made for the transition from weld toe failure to weld root failure for load carrying fillet welds in cruciform joints. The resulting graph is shown in fig.26 together with the actual values for weld failure of the cruciform joints tested here in order to demonstrate the quality of the welds and of the weld toe profiles. These profiles are shown in fig.23a en b. It follows that the S-N lines for toe failures are still unknown, but are anyhow "better" than the ones for root failures.

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It follows also that for high strength material full penetration welds are necessary in order to get the full benefit from high quality welding. V.A. loading. Series I and III were also tested with the standard spectrum Gauss,I = 0.99. The results are given in table 12 and plotted in fig. 25. Again, almost all cracks emanated from the weld roots. The Basquin equations for root failure as determined by linear regresion are as follows: S e r i e s I , (2a » 3 mm) 3 .08 l o g Sr»s + l o g N = 1 1 . 5 8 1 (12)

S e r i e s I I I (2a ss 1 mm) (3 .0 ) l o g Srms + l o g N = (11 .669) (13)

eq. 13 is from a best fit with an inverse slope of 3.0, because of the low number of results suitable for linear regression.

7.5 Longitudinal n.l.c. fillet welds (attachments). Joints with non-load-carrying fillets welds in longitudinal direction and dimensions according to fig. 5d, were fabricated as follows: First the ends of the attachments were welded all around (arc « 180°). Then the fillet welds were made in anti-metric sequence in order to minimize distortion. The following welding procedures were used (see table 6): - Welding C for the end welds and for the fillet welds - Welding A for the end welds and welding B for the fillet welds»because this latter type of welding appeared not suitable for making the end welds because of undercuts due to the necessary low speed welding. As the end welds are the fatigue critical areas of this joint, an increase of the fatigue strength at very low costs is obtained by TIG dressing only these end welds. Therefore specimens from both series were also tested with the end welds TIG dressed with a HI = 0.8 kJ/mm (see par.6.3). C.A. loading. The test results for joints made with welding C (table 6) in the as welded condition and in the TIG dressed condition are given in table 13 and plotted in fig.27. Test results for end welds A (MMA, rod dia 2,5 mm) are given in table 14 and plotted in fig.28 The Basquin equations and the fatigue limits are:

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end welds type C as welded : 4.01 log Sa + log N = 13.438 (14)

Sf = 55 N/mm2 TIG dressed : 4.3 log Sa + log N =14.10 (estimate) (15)

Sf « 70 N/mmz

end welds type A (rod dia. 2.5mm) as welded : 2.71 log Sa + log N = 10.550 (16)

Sf « 25 N/mm2 TIG dressed : 3.97 log Sa + log N = 13.347 (17)

Sf s= 55 N/mm2 (toe cracks) Sf » 40 N/mm2 (root cracks).

In the fig. 27 and 28 the BS5400-class F2 line is plotted for comparison (as 1 = 150 mm, the F-line is formally applicable) and it appears that the results for as welded end welds type A (fig.28) are below this line. TIG dressing improves the endurances considerably (the inverse slope rises from 2.71 to 3.97) and raises the estimated fatigue limit from 25 to 55 N/mm2 (root cracks not taken into account).

However, as shown in fig.29, the results for TIG dressed end welds type A (eq.17) are not better than those for the specimens with end welds type C in the as-welded condition (eq.14). These latter specimens show, see fig.29, a further improvement due to TIG dressing to an estimated inverse slope of 4.3 and Sf » 70 N/mm2. It follows that the initial shape of the toe, which was rather rough for the type A (MMA) welds, is of influence on the improvements obtained from TIG dressing. This was also found by Bignonnet [6] with regard to welds in offshore type joints. V.A. loading. Endurances obtained with the spectrum Gauss, I = 0.99 are given in table 15 and 16 and plotted in fig. 30 and fig. 31. The Basquin equations are as follows: end welds type C as welded : 3 .79 l o g Srms + l o g N = 12 .264 (18) TIG d r e s s e d : 4 . 7 9 l o g Srms + l o g N = 14 .154 (19)

end welds type A a s welded : 3 . 0 5 l o g Srms + l o g N = 10 .752 (20) TIG d r e s s e d : 3 .57 l o g Srms + l o g N = 11 .934 (21)

In fig.32 the 4 S-N lines from V.A. loading are plotted together, and it appears that the evaluation given above for C.A. loading is also valid for the results from V.A. loading.

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8. Experimental program for the FeE420TM material 8.1 General

In par. 3 through 8, fatigue tests series on different welded joints made from 6 and 9 mm hot rolled FeE560TM plate material are reported. Now 9 mm plate material is about the heaviest gage used for heavy vehicles and therefore the tests on 9mm plates were restricted to the material itself (par.4,2) and to butt welded joints (par.7,3) in order to show the similarity with the 6mm plate. Now in the other end of the utilisation scale are hot rolled thin plates and lower material grades. Therefore an additional investigation was carried out for 3mm hot rolled plates in FeE420TM material. As for the 9mm FeE560TM material, the S-N curves for this material and for butt welded joints were determined but, in addition to these, the endurances of spot welded joints were investigated too.

8.2. The FeE420TM plate material 8.2.1 Composition and mechanical properties

The FeE420TM material, hot rolled to a thickness of 3mm, was obtained from Hoogovens IJmuiden BV. The plate surfaces were pickled and oiled. All plates were taken from one coil and degreased before processing. The chemical composition and mechanical properties are given in table 17.

8.2.2 Endurance data Endurances and the fatigue strength for the plain material were determined from specimens fig.33 The results are given in table 18 and plotted in fig.34. The Basquin type equation, based on linear regression is 10.66 log Sa + log N = 31.532 (22) Sf = 239 N/mm2 (st.dev. = 17N/mm2) It follows that, not taking into account Kt = 1,07 of the specimen, the fatigue ratio is: Sf/Rm = 0.43 (22) which is about equal to the ratio found for the 6mm material with millscale.

8.3 Spot welded joints 8.3.1 Spot welding parameters

Spot welds were made with a 315 kVA spot welding machine having current control. Truncated cone (120a) electrodes, RWAA class 2, were used throughout. The tip diameters were 9.2 and 12.3mm. The larger tip face was only used to investigate nugget sizes of 4t. The welding time was 30 cycles and the hold

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time 60 cycles for all welding trials. The nugget diameters (9.2mm tip face), as determined from peel tests, as a function of the welding current are given in fig.35. The results of tensile-shear tests and cross-tension tests are summerized in table 19, together with the ductility ratio's. From these tests it was decided to use nugget diameters

9 < d < 10mm 8.3.2 Endurance data for spot welded joints in FeE420TM.

The specimens used were single shear lap joints having - 1 spot weld, series SI, fig.36a - 3 spot welds parallel, series S3, fig.36b. The sizes of both specimens were kept the same so that the bending stiffnesses would be comparable. The nugget diameter was 9 to 10mm. C.A. loading The results of C.A. tests are given in table 20 and plotted in fig.37. The Basquin type equations are 1 spot weld, series SI 2 . 6 4 l o g Fa + l o g N = 7 . 3 9 5 (23) Ff = 1 .6 kN

3 spot welds, series S3 2 . 8 1 l o g Fa + l o g N = 8 . 5 5 2 (24) Ff = 3 . 7 5 kN

It follows that the strength ratio between the 2 series is rather high, viz.

endurances : 2.35 to 2.50 fatigue limit : 2.35

All failures were plate failures, as usual emanating from the top (bottom) of the spot welds. The S3 specimens showed cracks at every spot weld. V.A. loading The results of V.A. tests, standard spectrum Gauss, I = 0,99, are given in table 21 and plotted in fig.38. The Basquin type equations are: 1 spot weld, series SI 2 . 7 6 l o g Frms + l o g N = 6 . 8 6 9 (25)

3 spot welds, series S3 2 . 9 1 l o g Frms + l o g N = 8 . 0 5 8 (26)

The strength ratio between the 2 series is again 2.35 to 2.50. Note that the exponents of the Basquin equations for C.A. and V.A. loading are essentially the same. Failures were as for C.A. loading.

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8.4. Butt welded joints 8.4.1 Welding processes used

Butt welded joints were produced with the following types of welding processes - MIG welding - Pulsed MIG welding A description of the Pulsed MIG process was given in par. 6.2. The process parameters are given in table 22 and it should be noted that because of the lower strength of the material only commercially available consumables were used. Strips of width 160mm and length 400mm ( = 1/2 platewidth) were welded as described in par.5.2. Despite the heavy fixtures, the distortions after welding were severe. Fig 40a,b show the typical distortions. Typical cross sections for welds from both welding procedures are given in fig.40c and 40d.

8.4.2 Specimens and specimen distortion From these welded strips specimens were produced, of which the dimensions are given in fig.39. As the distortion of the specimens was regarded as too high, a first batch of specimens, (which happened to be Pulsed MIG welded) was straightened by 3-point bending. The results of fatigue tests carried out on these specimens are shown in fig.41 and it appears that:

Sa > 130 N/mm2 N = 2 to 6 * 103 (27) Sa < 130 N/mm2 N > 2.107

Note that at Sa = 140 N/mm2 there is one failure at 3.7 x 10= and another one at 1,4 x 107. Apparently this type of straightening does not yield reproducable (or classic) endurances. So, one can wonder what may be the results of straightening applied after welding in industry. For comparison the class D line from BS5400 is given too, and it appears that the results are still above this line. In view of the above all specimens fig.39a, used to determine the S-N lines were straightened in tension to 3.5mm permanent set before testing. As the total free length was about 180, this permanent set was 2% and just beyond the yield zone.

22

8.4.3 Endurance data for butt welded joints, in FeE420TM. Endurance tests on straightened specimens, see par.8.4.2. were carried out at about 150 Hz, using a servo controlled resonance machine. To prevent (dynamic) buckling, teflon coated guide plates were used. To prevent heating-up, forced air cooling was used so that the measured temperature of the specimens did not exceed 50°C. (It is known that frequency effects in fatigue testing are dominantly from heating up of the specimens). The results of the CA endurance tests are given in table 23 for the MIG-welded specimens and for the Pulsed MIG welded ones. These results are plotted in fig.42 and fig.43 together with the BS5400-D line. The Basquin type equations are: MIG w e l d i n g 5 . 6 3 5 l o g Sa + l o g N = 1 8 . 2 4 1 (28) Sf = 135-N/mm2

Cracks initiated dominantly from the weld root Pulsed MIG welding 6 . 1 8 l o g Sa + l o g N = 2 0 . 0 0 2 (29) Sf = 175 N/mm2

Because of grip failures, most specimens were waisted, see fig.39b. Cracks initiated equally from the weld root and from the weld toe. From the above it appears again that the Pulsed MIG process yields the better results.

23

CONCLUSIONS From this investigation into the fatigue strength of welded joints in HSLA steel, the following conclusions are drawn: FeE560TM, plates 6 and 9 mm: - Due to residual welding stresses the fabrication of welded constructions in high strength (TM) steels needs more precautions with regard to distortions.

- The fatigue strength of welded joints in these steels is considerably upgraded when the weld is overmatched by 40 to 50 HV5 with regard to the plate, provided that the transition at the weldtoe is smooth. A fatigue strength of 160 N/mm2 was obtained for butt welded joints in axial loading at R = -1.

- The welding processes that yielded a high fatigue resistance were : Pulsed MIG welding with metal cored wire MIG welding with (rutile) flux cored wire. Note that metal deposition was by spray transfer for both processes.

- The smooth and large transition radius and the high hardness of the weld toe obtained with TIG dressing yields a high fatigue strength. For 6 mm plate material a heat input of 0.7 kJ/mm is sufficient in this respect.

- Crack growth from weld roots, as occurred in joints with transverse fillet welds and in n.l.c. end welds, are to be avoided because macro-crack growth is almost independent of the grade of structural steel. It follows that full penetration welds are necessary.

- Excentricities and misalignments of welded joints reduce the endurances considerably.

- Fatigue data for 4 basic types of welded joints were generated for C.A. and for V.A. loading up to endurances of N = 2*107.

FeE420TM, plates 3mm - The parameters for spot welding this material were optimized and

- Fatigue data for C.A. and V.A. loading were generated for 2 types of spot welded joints.

- Butt welded joints showed large distortions and it has appeared that the method of straightening influences the endurance considerably.

- Fatigue data for C.A. loading were generated for butt welded joints made with 2 different welding processes and straightened 2% in tension.

General It is anticipated that the reported improvements in production technology together with the generated fatigue data will enable designers to make an efficient use of high strength TM steels in welded constructions.

24 -

10. REFERENCES (1) Overbeeke, J.L.,et al., "The Fatigue Strength of Welded,

Bolted and Riveted Joints in High Strength Low Alloy Steel". EUR 9964 MF EN,1985.

(2) , IIW Recommendation on "The application of an engineering critical assesment in design, fabrication and inspection to assess the fitness for purpose of welded products", Part 4, Fatigue, Doc.IIW XIII-1283-88.

(3) Suzutki,S., Wheatherly, G.C., Houghton, D.C., Acta Metallurgica, Vol.35, 1987, pp.341-352.

(4) Sol, A.M., Joining & Materials, Vol.2, 1989, pp.372-376. (5) Haagensen, P.J., "Improvement of the Fatigue Strength of

Welded Joints", Paper PS 6, Proc. Int.ECSC Conf. on Steel In Marine Structures, Paris 1981, Euroreport EUR 7347 DE EN FR, 1981.

(6) Bignonnet, A., "Improving the fatigue strength of welded steel structures", Paper PS 4, Proc. 3rd Int.ECSC Conf. on Steel In Marine Structures, edited by Noordhoek, C. and de Back, J., Elsevier 1987.

(7) Minner, H.H., and Seeger, T., "Investigations on the fatigue strength of weldable high strength steels St460 and St690 in as welded and TIG-dressed Conditions". Paper 9.5, Proc.Int.ECSC Conf. on Steel In Marine Structures, Paris 1981, Euroreport EUR 7347 DE EN FR, 1981.

8) Haibach,E., Fischer, R., Schütz, W., and Hück, M., "A standard random load sequence of Gaussian type recommended for general application in fatigue testing; its mathematical background and digital generation". Paper 29, Proc.Int.Conf. on "Fatigue Testing and Design", edited by Bathgate, R.G., S.E.E., 1976.

(9) Schütz, W., "Standardized Stress-Time Histories -An Overview", ASTM STP 1006, edited by Potter,J.H. and Watanabe,R.T.,pp.3-16, ASTM, 1989.

(10) Gurney, T.R., Joining and Materials, Vol.2, 1989, pp.320-323 and 390-395.

(11) Gurney, T.R., Fatigue of welded Stuctures, second ed. Cambridge University Press, 1979.

- 25

Table 1. Composition and Properties of the FeE560TM plate materials.

Heat I*) II III

Casting ingot concast ingot Thickness, mm 6 6 9 Chemical composition, %:

C Mn Si P S

micro-alloy, el. Nb, V inel. control grain size,ASTM è 12 Mech.proporties: | -L | X | -L Rm, N/mm2 716 727 694 709 658 682 Re, N/mm2 628 658 611 626 580 626 A % 25 21 30 25 24 20 Charpy V,(tx8) mm2

E at T=-40°C,J 35 29 >40 >30 62 27 Hardness, HV5 240-250 230-240 200-220 *) from [1]

0.09 1.50 0.36 0.016 0.011 Nb, V Ce è 12

0.08 1.44 0.36 0.016 0.006 Nb, V Ca ;> 12

0. 0. 0. 0. 0. Nb Ca 12

07 93 13 014 006 », Ti i 1-14

- 26 -

Table 2. Results of endurance tests at R = -1 for Heat II specimens, t = 6mm

Specimens from STRIP A

Specimens from STRIP B

Sa N/mm2

N * IO" 3

Sa N/mm2

N * IO" 3

400 350 350 320 320 320 320 320 300 300 420 300 450 300 400 280 375

74 211 277 419 508 480 799 882

1,429 >4,712

53 >8,464 ' 16

>11,196 66

>4,063 142

] ] ] ]

320 320 320 320 400 300 300 380 300 400 300 350 280

209 600

1,168 >11,437 -,

58 J 2,244 >5,499 -,

63 J 5,293 -,

266 J >8,392 -,

292 J >4,688

*) retest after runout

- 27

Table 3. Results of endurance tests at R = -1 for Heat III specimens, t = 9 mm

Sa N N/mm2 *10"3 400 74~ 375 97 350 188 320 274 300 472 285 1,314 270 578 270 708 270 1,738 270 >11,351 375 79 255 2,631 255 >9,433 400 44 255 >15,070 320 153 240 >14,854 350 207

*) retested after runout

28

Table 4 Results of C.A. tests at R = -1 of unnotched specimens, from 3% stretched material. Heat II

Specimen 3% stretched before testing

Sa N remarks N/mm2 *10-3

350 350 306 320 320 320 320 300 400 300 375 300 340

102 163

1,757 334 327 626 312

>3,903 50

>2,918 107

>8,181 229

] ] ]

Table 5 Results of C.A. test at R = -1 of unnotched specimens without mill scale.

Heat II (6mm) Heat III (9mm) Sa N remarks Sa N remarks N/mm2 *10-3 N/mm2 *10"3

430 450 360 340 320 400 340 320 375 340

>3.336 23

1.141 684

>10.486 60

1.390 >8.480

509 923

285 350 285 400 300 375 300 300 315

>11,364 177

>25,310 59

>11,856 128

3,217 874 536

] ] ]

29

Table 6 Welding Processes and Consumables, 6mm plate, FeE560TM.

Designation B Welding Process Shielding Gas Ar/CC-2 Consumable type dia, mm AWS-code

SMA (MMA)

n.a. rod basic 4 E11018-M

H.I, kJ/mm (spray transfer Butt welds I 1.0

II 1.1 Fillet welds 1.4 Hardness of the weld, HV5 280(280)

MIG

80/20 wire, flux cored rutile 1.2 E111T1-K3 ) 0.9 1.3 0.9

280(280)

Pulsed MIG

95/5 wire, metal cored n.a. 1.6 E91T-G

0.75 0.7 1.0

260(260)

Servo Adj. MIG

80/20 wire, flux cored basic 1.6 E81T5-K1

0.9 0.9 1.5

280(270) *) Numbers in parenthesis are for the fillet welds.

30

Table 7 Summary of the screening fatigue tests at R = -1 on welded joints made with different processes. Sa eg = eguivalent nominal axial stress after orrection for excentricity after [11.

Weldment Number log N o(log N) of spec.

Remarks

B u t t welded j o i n t s , Sa =

A - I 3 4 .726 A - I I 5 5 .162

B - I B - I I

C - I C - I I

D - I D - I I

Ref. [1]

F i l l e t we

A B C D

Ref. [1]

4 6

3 4

3 4

i lded

3 3 4 3

l a p

5 .505 5 .484

5 .835 6 .032

5 .634 5 .887

5 .723

j o i n t s

5 .415 5 .703 5 .698 5 .357

5 .372

185 N/mm2

0 . 2 3 0 .38

0 .24 0 .32

0 . 2 3 0 .45

0 .10 0 .20

0 .16

Sa = Sa e g . =

0 .15 0 .20 0 .12 0 . 1 1

0 .03

ex r u n o u t s

74 N/mm2

2 . 2 * 74 = 163 N/mm2

31

Table 8 Hardnesses and toe radii after TIG dressing

Heat input, kJ/mm

Hardness (HV5) of the TIG-weld « HAZ (max)1)

Weld toe radius 2),mm max. min.

0.7

320 350

4.3 2.5

1.0

« 320 360

5.0 3.2

1.4

s 270 290

7.2 4.0

1) for weld material and base material about equal 2) as measured from replica's

Table 9 Welding Parameters, 9 mm FeE560TM (Heat III).

Welding process Submerged Arc Pulsed MIG

Weld preparation Consumable AWS-code diameter, mm Flux (DIN 32522)

shielding gas Heat Input kJ/mm Welding speed,m/min Hardnesses, HV5 plate HAZ, min Weld

none

EM3G 4 BAB.l .66.AC 9SKM.HP5 n. a. 1.6 0.7

220 K 195 220

see text

E 91T-G 1.6 n.a.

Ar/C02-95/5 1.5 0.3

220 190-205 260

32

Table 10 Results of C.A. tests at R = -1 on butt welded

Fa kN 210 185 185 150 135 120 105 90 80 65

120 65

150

1°i

S.A. N

*10-3 101 123 57

124 268 296 430

1,150 660

>21,990 510

>14,960 170

nts, FeE560TM,

Welding

* ] ]

remarks

9 mm plate.

Puis« Fa kN 240 210 210 185 165 165 240 150 210 150 210 135 185 120 185 100 150

2d MIG Welding N *1C 40

114 160 323

2,114 >12,153

76 >17,641

374 >19,700

91 >12,030 5,349

>22,284 598

>25,300 712

remarks )-3

] ] ] ] ] ]

*) retested after runout.

- 33

Table 11 Results of C.A. tests on welded cruciform joints Welding : Series I, Pulsed MIG

Series II, Flux cored MIG Series III, Flux cored MIG

Loading : Constant amplitude, R = -1

Series I, 2a = 3mm Series II, 2a = 5mm

Sa N/mm2

160 125 100 100 90 80 70 70 60 .60 60 55 55 50

N *10-3

46 496

1,313 • 1,608 1,687 1,992 2,923 2,102 3,962 2,460 2,473 4,327 7,836

>10,830

remarks

toe failure toe failure

toe failure

runout, but cracked

Sa N/mm2

120 100 100 80 80 60 60 55 50 45 40 35

N *10-3

212 286 425 826

1,212 1,190 1,702 1,739 1,959 3,147 5,453

>23,474

remarks

toe failure

Series III, 2a = 1mm

Sa N/mm2

120 100 100 80 80 60 60

120 50 50

100

N *10-3

74 101

1,404 8,236 5,186 8,150

>48,316 1,723

32,983 >40,130

1,780

* ] ]

remarks

toe toe

fail. fail.

e= e= =3mm =2mm

All failures are root failures unless indicated

*)retested after runout

- 34

Table 12 Results of V.A. tests on welded cruciform joints welding : Series I, Pulsed MIG

Series III, Flux cored MIG loading : Gauss, I = 0.99, R = -1

Series I, Srms N N/mm2 *10-85 438 70 571 70 987 50 1,179 50 2,352 40 4,184 35 4,447 35 7,098 30 11,685 30 >25,039 40 5,854

2a

3

]*

» 3mm remarks

toe fai lure

Series Srms N/mm2

80 80 65 65 50 50 40 40 35 35 30

i III, 2a s

N *10-3

188 825

1,435 1,898 3,203 4,425 5,526 6,232 7,222

>20,569 >22,689

1mm remarks

toe failure toe failure

toe failure

*) retested after runout All failures are root failures unless indicated.

- 35

Table 13 Results of C.A. test, on n.l.c. long, fillet welds

As Sa N/mm2

120 100 90 70 60 60 50 80 50 70

Welding : Loading :

welded N

*10-3 135 216 393 761

1,449 3,171 >8,505 -,

974 J >19,057 -,

655 J

: type C (Pulsed MIG) : Constant

remarks

1

2

failed in

Amplitude, R = -1.

TIG-Sa

N/mm2

90 80 70 70

120 70

100 60 80

grip

-dressed N

*10-3 484

1,455 1,064

>12,287 792

>12,049 322

>15,075 4,342

] ] ]

remarks

2

Almost all cracks initiated at the weld toe. 1) cracks at both end welds. 2) cracks from the root of the end weld.

Table 14 Results of C.A. tests, on n.l.c. long, fillet welds end welds : type A (MMA) fillet welds : type B (Flux cored MIG). Loading : Constant Amplitude, R = -1.

As welded TIG-dressed Sa N remarks Sa N remarks N/mm2 *10~3 N/mm2 *10~3

80 221 60 504 50 956 40 2,016 30 2,184 30 3,227 26.8 6,475 25 >20,741

100 138

80 70 60 60 50 50 45 40 60

100

652 1,396

991 1,207 4,322 6,782

21,357 >24,845 >10,973 -,

276 J

2 2 2

Almost all cracks initiated at the weld toe. 1) cracks at both end welds. 2) cracks from the root of the end weld.

36

Table 15 Results of V.A. tests, on n.l.c. long.fillet welds,

As Srras N/mm2

70 60 50 40 32 25 25 22

Welding: Loading:

welded N r

*10-3

184 517 515

1,243 2,735 8,646 14,366 15,028

type C Gauss,

emarks

1 1

1 1

(Pulsed 1 = 0. .99,

TIG-Srms N/mm2

70 60 50 40 30

MIG) , R = -1

-dressed N

*10-3

174 444

1,259 3,144

10,578

remarks

1 1

All cracks emanated from the weld toe 1) Cracks at both end welds

Table 16 Results of V.A. tests on n.l.c long, fillet welds end welds type A (MMA) fillet welds : type B (Flux cored MIG) Loading Gauss, I = 0.99, R = -1,

As welded TIG-dressed Srms N N/mm2 *10- 3

r e m a r k s Srms N/mm2

N *10-3 remarks

70 60 50 40 30 25 20 20 17

94 215 480 893

2,030 3,429 4,419 9,243 7,210

1 1 1 1 1 1 1

60 50 40 30 26 22 22

363 671

2,501 4,001 5,656

17,073 3,340

1 1

All cracks emanated from the weld toe. 1) cracks at both endwelds

37 -

Table 17 Properties of the 3mm FeE420TM material.

Chemical composition (%) C 0.094 Mn 0.70 Si 0.22 grain size purity inclusion control Mechanical properties Tensile tests on specimens 20 x 3mm2

Nb P S

ol:

0.023 0.010 0.011

ASTM size 13 high Ca

Rm, N/mm2

Re, A5,%

558 464 25

567 488 22

All a-e curves showed a yield zone. Bend test, 180s, radius 3mm : OK

Table 18 Results of C.A. fatigue tests on 3mm plate material FeE420TM.

Sa N remarks Sa N remarks N/mm2 *10~3 N/mm2 *10~3

340 320 300 300 280 280 260 260 250 250 240

1 >9 1

69.3 48.0

317 172 186 121 905 649 ,727 ,696 ,114

240 240 280 230 230 230 320 220 300 220 320

934 >4,338

219 >5,958 3,309 >8,425

59 >7,480

56 >10,532

43

] ] ]

*) retested after runout

38

Table 19 Summary of spot welding trials

nugget size,mm 5/t(=8.7) 4t(=12)

electrode,mm 9.2 12.3

63 plug

Tensile shear test Fmax , kN failure

Cross tension test Tmax , kN failure

Ductility ratio T/F

41 face

31 plug

0.76

40 plug

0.63

- 39

Table 20 Results of C.A. tests at R = -1 on spot welded, single shear lap .joints, FeE420TM, 3mm plate.

single spot weld 3 spot welds parallel Fa N remarks Fa N remarks kN *10-3 kN *10"3

7.0 5.0 3.5 3.5 3.0 2.5 2.0 1.5 4.0 1.5 6.0 1.25 1.75

125 308 740

1,618 1,486 1,460 .3,828

>36,314 601

>64,542 261

>35,743 6,406

20.0 15.0 13.0 90 7.0 5.0 4.0 4.0 3.5

11.0 3.5 5.0

94 171 271 754

1,272 4,117 6,446

10,813 >20,924

389 >50,820

3,226

] ]

*) retested after runout

Table 21 Results of V.A. tests at R = -1 on spot welded, single shear lap joints, FeE420TM, 3mm plate. Loading : Gauss I = 0.99, R = -1

Frms kN

single spot weld

N remarks *10-3

3 spot welds parallel

Frms N remarks kN *10-3

3.0 2.5 2.5 2.0 2.0 1.5 1.5 1.0 1.0 0.8 0.8

298 519 679

1,274 1,386 1,881 2,698 3,323

12,912 9,607

22,698

7.0 6.0 5.0 5.0 4.0 4.0 3.0 3.0 2.5 2.5 2.0 2.0 2.0 7.0 1.7 1.7

306 648

1,448 1,294 2,639 1,545 4,034 5,164 5,073 6,028

23,015 12,706 >38,469

348 21,072 34,604

- 40 -

Table 22. Welding parameters, 3mm FeE420TM

Welding Proces

Weld preparation edge gap, mm no. of runs backing plate

Welding consumable diameter,mm AWS-code availability-Shielding gas Heat input,J/mm Welding speed,m/min

Hardness, HV5 plate HAZ(min) Weld

Bend test 2t-180° (both sides)

MIG (GMAW)

square 1.5 1

no

solid 1.2 ER SOS--G commercial 80Ar/20CO2 0.36 0.48

170 240 OK

180-200

Pulsed MIG

square 0 1

copper plate, groove 1.5mm deep

metal cored 1.6 E71TG-NÌ 1 commercial 95Ar/5C02 1.1 0.40

225 230 OK

41 -

Table 23 Results of C.

Sa N/mm2

220 200 190 170 160 150 140 140 140 220 130 170 130 190

joints,

MIG Welding

N : *10-³

89 229 140 426 688 475 813

"3,820 >21,600

128 >25,200

627 >34,810

405

.A . tests at R = , FeE420TM, 3 mm plate.

remarks

] ] ]

1 2 2 2 2 2 2 1

1

1

2

Pul

Sa N/mm2

220 210 190 190 180 180 170 200 170 240 150

sed MIG

N *10-3

273 452 314 842

1,302 3,101

>21,652 534

>37,664 233

>35,591

-1 on butt welded

Welding

] ]

remarks

1 1 2 2 2 1

2

1

1) Initiation at weld toe 2) Initiation at weld root

42

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20

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LOADING C.A. , R = -1

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LOADING C A . , R = -1

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10 6 107 - * - N - cycles

_ c

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90

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c. CRUCIFORM JOINT

37.5

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d. WELDED-ON ATTACHMENT

Figure 5. Dimensions of welded specimens, FeE560TM.

47

ì i

T

zzz¡ znLznz znnzzzz

T 222 UZ ¡ZIJJ. u

T 1

24

T = TENSILE SPEC.

HEAVY TACK WELD

ROLLING DIRECTION

Figure 6a. Layout of the welded specimens.

Figure 6b. Clamping device used during welding.

- 48 -

hardness HV5

325f

300

275

■ HAZ1 wcldmrtal IHAZ 3 0

°

200-

, X > 325

275

250

225

200 counter pass

H A 2 [ w . l d m . ţ ,

I / A l

5 10 15 20mm 0

>• 1 pass

10 15 20

Figure 7. Hardness prof i le , butt welded jo in t , 6mm plate FeE560TM

Figure 8. Cross section of the welds ( typ ica l ) .

- 49 -

XA XA / / / v XA Z I 1 I I I XA XA 1

LU OC cc ro

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Figure 9. Principle of Pùlsed MIG welding.

TYPE I

50°

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Figure 10. Edge preparations for the but t welds.

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THE APPLICATION OF WELDED, BOLTED AND RIVETED CONNECTIONS IN HSLA STEELS IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

Part II

FRICTION GRIP BOLTED JOINTS

Contents Page 1. Introduction 87 2. Evaluation program 88 3. Plate material 88 4. Bolts and Nuts 88

4.1. Shape, size and grades 88 4.2. Stress corrosion and hydrogen induced cracking 88 4.3. Friction from the screwthread 89

5. Friction grip bolted joints 92 5.1. Friction between nut and plate 92 5.2. Yield controlled torque 92 5.3. High friction primers 93 5.3.1. Primer Development 93 5.3.2. Influence of coating thickness and friction

after cyclic slip 94 6. Endurance tests 95

6.1. Dynamic friction from step tests 95 6.2. Endurance data for a H.D. structural joint 96

7. Conclusions g7 8. References 98

TABLES 99

FIGURES 102

- 85 -

THE APPLICATION OF WELDED, BOLTED AND RIVETED CONNECTIONS IN HSLA STEELS IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

Part II FRICTION GRIP BOLTED JOINTS

Introduction For structural connections quite often use is made of friction grip joints. As the name says, in these joints the load transfer is by frictional forces between the connected elements. The clamping force to attain sufficient grip from friction is usually generated by pretensioned bolts. These joints are e.g. used: When welding is not suitable because of uncertainty in reliability. An example is the assembling of larger parts of a bridge structure in the open, because of tolerances and of the quality of welding in the field under adverse atmospheric conditions . For semi-permanent joints As are joints that are to remain detachable in view of repair or replacement. For joints that as a whole or in part cannot properly be reached by a riveting tool or a welding torch. The latter two cases apply very often when assembling frames of trucks or heavy vehicles. It should be noted that for series production as takes place in the truck- and vehicle industry, the use of friction grip bolted joints is much more expensive than riveting. In [1] tests on friction grip bolted joints were reported with an approved [2] high friction primer on the mating surfaces of the joint and grade 10,9 bolts tightened by yield control. The results obtained were as follows: with regard to torque controlled pretensioning, an increase of 50% in design load is obtained by yield controlled pretensioning of the bolts. Incidental high overloads that cause incidental repeated slip reduce the frictional force to about 60 to 70% of the initial value. Therefore the design load is to be based on the highest service loads anticipated.

87

Evaluation Program In the investigation described below, the emphasis is placed on - the highest allowable grade of the bolts in view of

stress corrosion and hydrogen embrittlement. the coefficient of friction of the screwthread in relation to the clamping force exercised by the bolt, the repeatability of yield-controlled pre-tensioning the static and dynamic frictional properties of 2 high-friction primers suitable for series production, the fatigue endurance of a H.D. bolted joint having 4 bolts in a square.

Plate material. The plate material used complied with FeE560TM according to Euromorm 149-2 and its plate thickness was 6mm. It was obtained from Hoogovens Ymuiden B.V. A description of the general properties of TM steels is given in chapter 2 of Part I. The chemical composition and mechanical properties of the FeE560TM material used here are given in chapter 4 of Part I. For the sake of completeness the mechanical properties that are of interest here are repeated in table 1.

4. Bolts and Nuts 4.1 Shape, size and grades

The size M12 x 1,5 bolts and nuts used were produced by NEDSCHROEF Helmond B.V. They are of the flanged type and comply with DIN 6921 (bolts) and DIN 6923 (nuts), fig.l The nuts were hot formed and heat treated to grade 10 or 12 according to ISO 898-2. The screwthread was cut. The bolts were cold formed and then heat treated to grade 10.9 or 12.9 according to ISO 898-1. Both bolts and nuts were provided with a yellow passivated layer of electrolytic zinc, 5-8 um thick.

4.2 Stress corrosion and hydrogen induced cracking For the possible application of grade 12.9 bolts a short literature survey was made about the risk of stress corrosion cracking in service or hydrogen induced cracking after finishing. Stress corrosion cracking. (Kiscc = the threshold value of the Stress Intensity Factor for crack growth from static load due to corrosion).Kiscc values for a large number of high strength steels in salt water are given in [3]. These values are plotted in fig.2, taken from [3] and this figure shows that high strength materials are susceptible

88

to stress corrosion cracking in brine when the yield strength exceeds 1000-1100 N/mm2, regardless of their composition. Hydrogen induced cracking The most important causes of hydrogen diffusion into the bolt material is the pickling process that preceeds galvanizing and the galvanizing process itself. Furthermore the hydrogen cannot escape afterwards because of the diffusion barrier formed by the zinc [4,5]. Fig.3, from [6], gives a good survey and shows that the maximum allowable Rm for grade 10.9 bolts is the very limit to avoid hazards due to hydrogen induced cracking. Special heat treatments [7] can shift this hazard to a somewhat higher Rm, but these are not yet common in bolt manufacturing industry. As the applied production technology for producing high strength steels and high strength bolts are apparently not yet advanced enough to avoid possible hazards of hydrogen induced cracking or stress corrosion cracking, it is not justified to use bolts with a higher strength than grade 10.9 for structures that are used outdoors. Therefore no tests were carried out with 12.9 bolts.

4.3 Friction from the screwthread When using yield controlled torqueing the influence of the friction from the faces is eliminated but, the clamping force is still dependent on the torque in the shaft of the bolt caused by the friction in the screwthreads. The basic equation to show this is the von Mises Yield cr i ter ium Re2 = a2 + 3 t 2 (1) From this equation one can calculate the influence of the coefficient of friction of the thread on the clamping force. When defining a torque efficiency ñ as

F (Ug = Pl) F(Ug = Ul ) ft = * (2)

F (Ug = 0) F(Ug = 0, $ = 0) the following efficiencies apply for M12 x 1,5 screwthread

Ug % 0,08 90 0,12 85 0,18 76

and it follows that a 10 to 15% increase in clamping forces can be gained by using a true low friction coating.

89

Therefore screening tests were carried out on 5 different coatings that are marketed as low friction coatings,viz : CI : washdip I, 4 years old (used in [1]) C2 : "" I, fresh D : zinc flakes in a binding matrix of zinc Chromate. M : M02S in EP fat P : composition unknown T : washdip II. Whether washdip I or II contains additives is not known. These coatings were tested in 10 different combinations. The bolts and nuts used were M 12 x 1,5 grade 10.9, galvanized (5-8 um) and yellow passivated as described in par.4.1.The mechanical properties of the bolts, as measured in a tensile tests and based on the stress-area As = / 8 * (D2 + d 2 ) = 8 4 , 3 mm2 a r e : Fpo.2 = 80 kN Fm = 9 2 , 8 kN Rpo.2 = 950 N/mm2 Rm = 1 1 0 0 N/mm2

It should be noted that for grade 10.9 bolts 940 <. Rpo.2 <. 1120 N/mm2

So that these bolts were at the lower limit of the RPo,2 specifications (however see also fig.12.). The test were carried out in a servo-hydraulic testing machine and were performed as follows: While the load was increased by a ramp function, the nut was steadily rotated by hand. The tensile load Fv on the bolt and the torqueing moment on the nut Ms were measured simultaneously and provided graphs as given in fig.4. The angle of rotation during the tests was 60s to 90° which, for M12 x 1.5, compares with those for the short clamping lengths, (12 and 18 mm) of the joints used for fatigue testing. All bolts failed by necking except those with Ug = 0,47, which failed in torsion. The results within the elastic range of the bolts are, after correction for the friction of the free rotating piston rod, given in table 2 and plotted in fig.5. The conclusions drawn from these tests are : the friction us of the screwthread, as calculated according to VDI-Richtlinie 2230, (1986), differs by a factor 2 although all coatings used are announced as low friction. So called low friction coatings can loose their properties because of ageing (whatever that is). the coating with the lower friction is dominant, whatever the underlying coating is. the lowest friction as measured here is Ug =0.09 when using coating T. This results is a torque efficiency of ñ = 69/80 = 0,86

90

the torque at yield varies between 21.8 Nm at ñ = 0.86 for coating T to 88 Nm at ñ = 52/80 =0.65 for coating C2. For Coating CI (aged Washdip I) this torque increases to about 107 Nm and bolt failures in torsion resulted.

91

5. Friction grip bolted joints 5.1 Friction between nut and plate

The M-$ relationship for joints was measured for some of the coatings for which ug was measured. From Mmax the friction between the nut and the plate can be determined as follows, see fig.6a and b. In this figure are plotted:

- Fv versus Ms as measured before ( par 4.3) - Fv versus Mk where Fv = pretension in the bolt.

Ms = moment due to the screwthread Mk = moment due to the flanged nut.

and Mk + Ms = M. Now when M = Mmax : dM/d$=0 and dFv/dMs = -Fv/Mk when Uk = constant. From these conditions Fv/Mk can be determined by trial and error. In fig.6 the starting point was A'i and the distance A'i -Ai = Mmax. Drawing tangent a' parallel to line a yields a distance A2 ' - A2 > Mmax. Drawing tangent b' parallel to line b yields a distance B'-B = Mmax. From Fv/Mk the coefficient of friction between nut and plate, Uk, was calculated according

Uk = Mk/(Fv*D/2) (3) were

D = 1/2 (Dl + D2) = 19 mm, Dl = outer diameter of the shiny contact area = 24 mm D2 = inner diameter of the shiny contact area = 14 mm The scatter for Uk was calculated from differences of Mmax (3 tests) and different Fv - Ms measurements. These results are also given in table 2.

5.2 Yield controlled torque For yield controlled torqueing the same apparatus as mentioned in [1] was used initially. This instrument measures M and $ and determines dM/d$ continuously. When an appreciable drop in this gradient (to 2/3 or 1/2) is measured a bell rings indicating that the yield strength of the bolt is exceeded However some dynamic slip tests, comparable to the ones described in par.5.3, did show erratic results and therefore M-$ relationships were measured while assembling the single shear joints used for testing. Two registrations of M-$ are given in fig.7. The test shown in fig.7a was done by a light weight technician (55kg), while the test shown in fig.7b was done by a heavy weight judoka (100kg).

92

Now the semi-automatic torqueing equipment used for yield-controlled torqueing signals when the average slope of the M-$ relationship is decreased by 30% to 50%. From fig.7a it is clear that with a shaky hand due to overstrained muscles (rotation of the nut was in a horizontal plane) such a device cannot work properly. From fig.7b it can be seen that even in the elastic zone the M-4> relationship is not always linear. So in fig.7b the torqueing device will signal at A instead of at B. So the use of semi-automatic equipment to monitor yield-controlled torqueing must be thoroughly checked at the site in order to ensure the reliability of the pre-tensioning of the bolt.

Furthermore it appears from torqueing the bolt by the bolt-head, that the size of the hexagonal part of a head complying with DIN 6921(1983), see fig.l, is too small and impairs the safety of the technician because the wrench has insufficient support from this too low hexagonal part and jumps off. Therefore all bolts were torqued through the nuts.

5.3 High friction primers 5.3.lPrimer Development

In [1] use was made of a commercial available alkali-zinc silicate primer based on zinc powder and alkali-silicate. It is marketed by Akzo-coatings. This is a high friction primer (u > 0,55) that complies with [2]. However> this primer is rather difficult to apply because - mixing zinc powder with a low-viscosity hardener is difficult. - the potlife is shorter than 8 hours Therefore it is not suitable for series production as takes place in truck industry. Investigations within this convention by Akzo in order to make this primer more user friendly were not successful. Therefore experimental high fricton primers were composed by AKZO by modifying an available 2- component organo-zinc silicate primer. For this type of primer the zinc is in the form of a paste and the potlife is in excess of 8 hours. Two promising ones, designated as Wap 9.4 and Wap 9.5 were tested with our program. They comply with the usual tests necessary for qualification (corrosion, scratch)(testing etc.) and their drying time varies from 6 hours at room temperature to 20 minutes at 50°C. Static tests on single shear lap joints coated with 27 ± 3 um primer were carried out. The clamping force was 6Q kN as measured with a washer-type load cell below the bolthead.

- 93

The measured coefficients of static friction, u, where as follows. Primer Wap 9.4: (3 tests)

u = 0.67 ± 0.03 (plateau value) (4a) Primer Wap 9.5: (3 tests)

u = 0.73 ± 0.05 (plateau value) (4b)

5.3.2Influence of coating thickness and friction after cyclic slip Tests under reversed slip conditions were carried out on single shear lap joints (2t = 12 mm) coated with Wap 9.5 primer in 3 thicknesses viz.20,35 and 50um resp. The bolts M12 having low-friction coating M, see par.4.3, were torqued in yield control. The results for 5 succesive cycles are given in table 3. and a typical F-5 record is shown in fig.8 The results are not conclusive, most probably because of variations in the clamping force, but it appears that - there is no significant difference due to the

thickness of the primer at the first plateau value. degradation due to reversed slip is fastest for the 20 urn primer thickness. after 5 reversed cycles the frictional force is roughly 50% of the initial plateau value.

Note that here the clamping length = bolt diameter Note also that for a double shear joint the residual static friction after slip in a stepped fatigue test was 0.6 to 0.7 of the applied load amplitude [1].

94

6. Endurance tests 6.1 Dynamic friction from steptests.

In [1] it was established that, when in a C.A. fatigue test there was no slip after 100,000 cycles the specimen would become a runout. In view of this, the maximium force amplitude from friction was determined by step tests having blocks of 10E5 cycles and load increments per step of 5 kN. The load ratio was R = -1. The joints used were double shear lap joints fig.12b, and the bolts with coating T were torqued by hand. Yielding was determined from a continuous record of M-$ From fig.6. it follows that at dM/d$ = 0 holds: Fv = 85 kN, while the "elastic" range is Fv < 70 kN. (fig.5) So the average clamping forces have been Fv = 80 kN. The results are given in table 4. and it follows that for primer Wap 9.4

(Fa)max = 80-90 kN primer Wap 9.5

(Fa)max = 85~100 kN An analysis of the scatter in results has shown that this scatter is dominantly caused by the differences in torque M, fig.11. However there is no relation between this torque and the hardness of the bolts, fig.12. So the scatter is most probably caused by quality variations of the screwthread or/and its coating.

Fig.10 shows the M-$ records from torqueing Fig. 11 shows a plot of (Fa)max versus Mmax and Fig.12 shows a plot of (Fa)max versus the measured HRC

of the individual bolts. It follows from the above that for this double shear lap joint under fatique conditions holds :

Udyn > 0.50 As for double shear joints the load distribution between the 2 faces is never equal, Udyn > 0.50 can be regarded as a lower limit. Furthermore for bolts at the lower end of the specifications for grade 10.9 (fig.11, 12), the load amplitude (Fa)max > 80 kN. From this it is concluded that it is justified to use for: Design against fatigue Bolts 10.9, Torqued in yield control

Fv 2: 75 kN Primer Wap 9.4 or 9.5, thickness > 30 um

Udyn £ 0.50

95

6.2 Endurance data for a H.D. structural joint As the use of friction-grip bolted joints is common practice (e.g. for bridges) and design rules are available (e.g.[8,9,10]) the value of yield controlled torqueing is demonstrated for only one structural joint. The joint chosen, see fig 9c, is a heavy duty one,viz. a double shear lap joint with 4 bolts in a square, built from 6mm plates grade FeE560TM, coated with WAP 9.5 primer. The pitch and the edge distances were the same as used for the equivalent riveted joint, see Part III. As the 2 gusset plates each have the same thickness as the main plate, the first row of bolts transfers nominally 2/3 of the load to the gusset plates, as follows from elementary stiffness calculations. The results of C.A. tests at R = -1 are given in table 5 and plotted in fig.13. The Basquin type equation is

7,75 log Fa + log N = 22,595 (5) Crack initiation was in the main plate about 6 mm from the holes in axial direction and due to fretting fatigue. Note that the fatigue limit is not determined and is most probably beyond N = 2.107. The results of V.A. tests, spectrum Gauss [11], J = 0.99, are also given in table 5 and plotted in fig.14. As the maximum capacity of our servo controlled hydraulic fatigue machines is 240 kN, which results, with q = 5,26, in Frms ¿ 45 kN, Frms = 45 kN was the maximum load that could be applied. However note that at Fmax = 240 kN, the load per bolt is: 1/2 * 2/3 * 240 = 80 kN, and near the maximum dynamic frictional force as determined in par.6.1. Note furthermore that Smax = 513 N/mm2 which is 84% of the actual yield strength. It follows that the Srms - N curve was not determined also because when Fmax < Fsiip the endurance N £ 107. However, a reliable design curve can be constructed from N = 107 ; Frms = 45 kN. together with 7,75 log Frms + log N = constant (m = 7,75 is taken from the C.A. tests and is in view of the very high endurances involved, certainly on the conservative side). This results in a design curve: Fmax = 5 , 2 6 Frms £ 2 4 0 k N 7 , 7 5 l o g Frms + l o g N = 1 9 . 8 1 2 ( 6 )

96

Conclusions In order to improve the efficiency of friction grip bolted joints in high strength steels, different aspects of joining were investigated. After that fatique tests (C.A. and V.A.) were carried out on a H.D. structural joint. The results are: - Risks of S.C.- and H.I.-cracking limits the usual bolt

quality to grade 10.9 - Flanged (head) bolts and nuts improve the production

efficiency but increases the necessary torque considerably.

- The friction of the screw thread determines Fv. True low-friction coati2ngs on the screw thread yield Ug = 0,09. They also reduce the total torque considerably (uu = 0.10 - 0.12) despite high-friction primer on the plates.

- Yield controlled torqueing results in a reliable pretension, Fv . For M12 x 1,5 bolts grade 10,9 and low friction primer "T", Fv ;> 75 kN.

- 2 new, user-friendly, high friction primers to be used for friction grip joints were tested. Static plateau values of the coefficient of friction between the mating plates were u = 0,67 and 0,73. Under fatique loading the dynamic coefficient of friction Udyn > 0,55 for both primers. Endurance data for a heavy duty structural joint (double shear, 4 bolts in a square) from FeE560TM steelplates of thickness 6mm were determined under C.A and V.A. fatigue loading. For V.A. loading (Gauss, I = 0.99) the minimum endurance was 107 at the maximum allowable design load to avoid slip. At this maximum load the net section stress in the plate was 0.84 RPo,2, which is well above the usual design stress of 0.67 RPo,2

- However, when under variable amplitude loading peak loads occur that cause (cyclic) slip, the friction-grip quality of the joint is seriously reduced.

97 -

8. References 1. Overbeeke,J.L. : The fatigue strength of welded, bolted

and riveted joints in high strength low alloy steel. Euroreport EUR 9964 MF EN, 1985.

2. : Anstrichstoffe auf Alkali-Silikat Grundlage mit Zinkstaub für Reibflächen gleitfester Verbindungen. DGG-TL 918.300 Blatt 85, Mai 1976

3. Wanhill, R.J.H. : Microstructural influences on the fatigue and fracture resistance in high strength structural materials. Eng.Fracture Mechanics 10_ (1978) 337-357

4. Beyer, St. : Sprödbrüche bei Schraubenverbindungen-Ursachen und Abhilfemasznahmen. Mat.-wiss. und Werkstoff technik 19 (1988) p.356-366.

5. Landgrebe R, Kloos, K.H. , Speckhardt,H.: Wasserstoff-inducierte verzögerte Sprödbruchbilding bei hochfesten Schrauben-Abhilfemassnahmen. Mat.-wiss und Werkstofftechnik 19 (1988) p.367-376

6. Kloos, K.H., Landgrebe, R., Speckhart, H.: Untersuchungen zur wasserstof-inducierten Riszbildung bei hochfesten Schrauben aus Vergütungsstählen. VDI-Z Band 122 (1985) Heft 19.

7. Beyer, St.: Einflusz unterschiedener Wärmebehandlungs­verfahren auf die wasserstofinduzierte Sprödbruchbildung bei Vergütungsstählen für die Schraubenfertigung. Z. für Werkstofftechnik 18. (1987) p. 411-422.

8- : Systematische Berechnung hochbeanspruchter Schraubenverbindungen. VDI-Richtunie 2230, VDI-Verlag 1977.

9- : European recommendations for the use of high-strength friction grip bolts in structural steelwork. Report CECM-X-70-8E, European Convention for Constructional Steelwork, 1971.

10. : Anwendung hochfester Schrauben im Stahlbau. Stahlbau Verlag GmbH, Köln, 1974.

11. Schütz, W. : Standardized stress-time histories- An overview. ASTM STP-1006 (1989) p.3-16.

98

Table 1. Mechanical properties, plate material grade thickness mech.prop.

FeE560TM (EUR 149-2) 6 mm

Re , N/mm2 611 Rm, N/mm2 694 A (1=80)% 30 Hardness, HB 230 Fatigue limit Sf, N/mm2,R = -1 304

626 709 25

n.d. *)area 6 x 10 mm.

Table 2. Measured coefficients of friction Ug : between bolt and nut Uk : between flanged nut and plates.

plate condition : shot blast SA 2% -3 + high friction primer, 30um

Surface Coating Ug1) finish (Bolt and nut)

Uk2)

G + P G + P G + P G + P G + P

CI C2 C2 + M C2 + T T

0,47 ± 0,06 0,22 ± 0,02 0,12 ± 0,02 0,09 ± 0,01 0,09 ± 0,01

0,30 ± 0,01 n.d 0,10 ± 0,01 0,12 ± 0,02 n.d.

D D + D +

M T

0,18 ± 0,02 0,12 ± 0,02 0,09 ± 0,01

n.d. n.d 0,12 ± 0,02

B o i t -Nut G + P

B o i t -Nut -

Cl ]

l ]

0 ,18 ± 0 ,02

0 ,18 ± 0 ,02

n . d .

n . d .

1) calculated according to VDI - Richtlinie 2230 (1986) 2) ut calculated with eq.3.

99

Table 3. Max. frictional forces (plateau value) during cyclic tests, Wap 9.5 primer. joint : single shear lap clamping : 10.9 bolt, yield controlled torqueing

Primer thickness

Spec. Plateau frictional force at cycle nr.

nr.

Vim 20 1

2 3

49 53 49

26 26 30

22 23 23

21 21 22

20 20 21

average 50 27 23 21 20 35 4

5 6

48 40 47

38 31 32

34 28 28

31 26 27

28 24 26

average 45 33 30 28 26 50 7

8 9

48 47 45

38 32 30

34 28 23

31 24 20

27 22 18

average 47 33 28 25 22

100

Table 4. Results of step tests at R = -1 on double shear lap .joints having one bolt.

number of cycles *10-3 at Fa = ..kN 70 75 80 85 90 95 100 105

Primer Wap 9.4

Wap 9.5

Spec. 1 2 3 4 5 6 7 8 9 10 11 12

100 100 100 100 100 100 100 100 100 100 100 100

100 100 100 100 100 100 100 100 100 100 100 100

100 100 100 100 100 100 100 100 100 100 100 100

100 100 100 1.2 100 100 100 100 100 100 100 100

100 23.3 95.8 100 100 100 100 59.9 0.2

100 100

100

2.3 2.4 100 100

100 100

100 24.9

100 3.7 51.6

36.2 100 1.1

Table 5. Endurance tests on friction-grip bolted joints, 4 bolts in a sguare, double shear.

C.A. tests R = -1 V.A. tests R = -1 Fa N remarks Frus N remarks kN *10-3 kN *10"3

155 532 155 565 140 670 140 1,015 125 975 45 >13,955 grip failure 125 1,457 125 >18,218 117 2,737 117 4,609 110 13,579 37 >31,920

44 45 45 45 45 45 45 37 37 37

4,437 >19,852 >19,852 10,370 >13,955 10,226 12,700 >30,552 >30,552 >31,920

All failures in the central plate,

101 -

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Figure 1 . Dimensions of flanged(head) bolts and nuts.

200

£ 150

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a. l ight weight technician

b. heavy weight judoka

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Figure 8. Record of friction force F versus displacement 6 during 5 cycles, for a frictiongrip bolted joint.

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THE APPLICATION OF WELDED, BOLTED AND RÏVETED CONNECTIONS IN HSLA STEELS IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

Part III Riveted Joints

Contents Page 1. Introduction 119 2. Research program 120

2.1. Evaluation program 120 2.2. Design data program 120

3. Plate material 120 3.1. the FeE560TM material 120 3.2. quality of the holes 121

4. High strength rivets 122 4.1. General 122 4.2. Preventing "head failures" 122 4.3. Stress corrosion tests 123

5. High friction primers 124 6. Endurance data for riveted joints 125

6.1. General 125 6.2. Joints with one rivet 125 6.2.1. single shear and double shear joints 125 6.2.2. single shear joints with enlarged

excentricity 127 6.3. Joints with more than one rivet. 127 6.3.1. General 127 6.3.2. Rivet pattern dimensions 128 6.3.3. 3 rivets in series 128 6.3.4. 3 rivets parallel 129 6.3.5. 4 rivets in a square 129 6.4. Hole expansion 130

7. Fatigue tests on truck frames 132 7.1. General 132 7.2. Frame specimens 132 7.3. Test stand 132 7.4. Service-load spectrum 133 7.5. Results of service-load tests. 133

8. Conclusions 135. 9. References 735

TABLES 137 FIGURES 142

117

THE APPLICATION OF WELDED, BOLTED AND RIVETED CONNECTIONS IN HSLA STEEL IN STRUCTURES SUBJECTED TO HIGH DYNAMIC LOADING.

Part III R I V E T E D J O I N T S

Introduction Riveting is one of the oldest techniques for building steel structures. Although this technique is now more or less obsolete because of the breakthrough of welding, the number of riveted constructions that are still in use is very large, e.g. railway stations now regarded as monuments, bridges, the Eiffel tower. As the development of welding into a reliable joining technique took place in the decade after 1945, these riveted structures are in use for more than fifty - often for even more than one hundred - years. For these riveted structures when loaded in fatigue, as are bridges, there are no questions left about the endurance with regard to the original design load. The only question now is how much the allowable loads can be increased. So riveting has proved to be a reliable joining technique for steel structures, and it must not be forgotten that it is the major joining technique in aircraft industry. One of the areas where riveting is still in use to build steel structures is the vehicle industry. Most manufacturers of trucks prefer riveted chassis over welded ones because of flexibility and fatigue resistance.

Now especially in the vehicle industry there is a strong tendency to use steels of higher strength because a saving of weight leads directly to an increase of the payload and to a higher efficiency in service. However, riveting technology has had a standstill for decades and no high strength rivets were available in order to utilize the better properties of today's high strength structural steels. To overcome this problem, high strength cold heading rivets with hardnesses up to 350 HB have been developed [1]. Joints made with these rivets from HSLA steel plate, grade FeE560TM, were extensively tested in fatigue and joining techniques were optimized on the base of the test results. After that, endurance curves for constant amplitude and for variable amplitude loading were generated for several basic joints, in order to enable a designer to utilize the results of this development.

119

2. Research program The research program was again split up in 2 parts, viz. an evaluation program and a design data program.

2.1. Evaluation program Within the first part of this program the following aspects were investigated, using the results reported in [1] as a firm base for further development .

- Development of the rivet head and the riveting in order to prevent failures from below the head.

- Stress corrosion tests on H.S. rivets - The application of a high friction primer, designated Wap

9.5, that is much more user-friendly than the one used in .[1] and which proved its quality from the tests on the friction grip bolted joints, reported in part II.

- Determination of the minimum dimensions of rivet patterns.

2.2. Design data program Within the second part of this program S-N data for C.A. loading and Srms -N data for V.A. loading were generated for the following joints: - single shear lap joints having : 1 rivet

3 rivets in series 3 rivets parallel

- double shear lap joints having : 1 rivet 4 rivets in a square

All fatigue tests were carried out at Fraid/Fmax = R = — 1

3. Plate material 3.1. The FeE560TM material

The plate material used throughout this investigation was a hot rolled LD steel that complied with FeE 560TM according to Euronorm 149-2. It was supplied by Hoogovens IJmuiden in two heats, both with a plate thickness of 6mm. This HSLA steel was micro-alloyed with Nb and V, while for inclusion control Ce and Ca respectively were used. More details about these materials are given in part I of this report. The mechanical properties of both heats are given in table 1 and it follows that the differences in R and R 6 in are less than 5% and negligible.

However heat I, which was used in [1] and for parts of this evaluation program, was an ingot cast steel, while

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heat II, which was used for the evaluation program and for fatigue-data generation, was a concast steel. In concast steels the segregations are more concentrated in the mid-plane region than in ingot cast steels. Now for punched holes, the transition from shear to tear is also at the mid-plane region of the plate. It follows that sub-macro lamellar tearing, which could impair the fatigue resistance, could occur due to punching. From observations of the fracture surfaces however, no crack initiations emanating from this midplane were observed.

3.2. Quality of holes Truck frames contain a large number of holes suitable for riveted or bolted connections . The number of holes used for the actual connections is always lower than the total number and depends on the type of truck to be delivered to the customer. So apart from the fatique resistance of connections, the reduction of the fatique resistance due to open holes is also important. As the fatigue resistance of notches, as are holes, also depends on the surface quality, the S-N curves were determined for 6mm plates having 13mm holes made by Series 1 Series 2a,b Series 3

drilling but not deburred punched with a good but used die punched with a rejected (worn) die and an undersized matrix.

The plate material used was from heat II except for series 2b which was from heat I in order to get some comparison between the concast Heat II and the ingot cast Heat I (see par.3.1.). The specimen dimensions are given in fig.l. The specimens were slightly curved (see par.4 of Part I) All cracks emanated from the edge of the hole at the plate surface, irrespective of the heat. The results are plotted in fig.2 through 5. A summary of the results is given in fig.6 and table 2 and from these it appears that

The exponent m of the Basquin-type equation varies no more than 3,56 £ m <. 3,73 while the scatter is very low. So the growth of macrocracks dominates the endurances. the fatigue limit of the punched holes is 40 <. Sfk <. 50 N/mm2 (Where Sfk is the gross section stress as used for frame calculations). which is appr. 50% of that of the drilled hole and appr.150% the fatique limit of holes made with a rejected die.

It follows that the reduction factor, Kf/Kt, due to the surface condition of punched holes is Kf/Kt = 2.5.

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So the fatique resistance of the frame may depend on the surface quality of the open holes instead of on the joints.

4. High strength rivets 4.1. General

During the previous investigation [1] high strength rivets were developed, using a cold heading steel that complied with ISO 4954, class E2. The steel used is 18 MnB4, a micro-alloyed steel used in the Q and T condition. From this material roundhead rivets complying with DIN 124 and having diameters 12mm were produced by NEDSCHROEF Helmond, by standard industrial practice. After Q and T, all rivets were electroplated with 5-8 urn zinc and yellow passivated. From [1] it appears that rivets tempered to a hardness of 280 HB match the hardness of the plates because they

cause a still acceptable deformation of the hole in the FeE560TM plate material. yield a high fatigue resistance of the joints

These 280 HB rivets at a diameter 12mm need a riveting pressure of 400 kN.

4.2. Preventing failures from below the head From [1] it appears that for single shear lap joints, made from plates coated with the high friction primer mentioned in par,5., failure was due to cracks initiated: a. by fretting inside the hole followed by propagation

in the shear plane (fig.4a) when the holes are drilled.

b. at the stress concentration between the fixed head and the shank when the holes are punched. These cracks are caused by tilting of the rivet, see fig.4b, due to incomplete filling of the hole.

In order to prevent this latter type of failure (called "head failure") experiments were carried out with different shapes of the fixed head. The dimensions of the standard rivet are given in fig.5. The stapling of the punched holes, fig.6a, was again [1] as shown in fig,6b and the sharpening of the transition shank-head after riveting due to the constraint by the head is sketched in fig.6c. To avoid this sharpening 2 modified rivet heads, see fig.7, were tried out by riveting tests. Both modifications did not show improvements in degree of filling, while type B was also unacceptable from corrosion point of view because there remained a small gap between head and plate (at the outside) after riveting.

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However, it was established that riveting with a conical die on the roundhead of the rivet transfers enough material from the head into the transition region to prevent sharpening. Fig.8 shows the results.

In order to verify the visual observations, V.A. fatigue tests (GAUSS, I = 0,99) were carried out on 5 single shear lap joints, riveted with type A rivet heads and 2 conical dies. All failures were shear plane failures. The endurances are plotted in fig.9 together with results from [1] for the "standard" joints that produced head failures.

4.3 Stress corrosion tests.

It is known that high strength steels are sensitive to stress corrosion cracking in brine and to hydrogen embrittlement when Re exceeds 1000 N/mm2 (Wanhill(2),Kloos et.al.(3)). Therefore the grade of bolts used in the vehicle industry is usually restricted to grade 10.9, This was already discussed in Part II.

In view of the above, sustained load tests in brine were carried on formed rivets of hardness 320 HB (Rm ~ 1100 N/mm 2). The surface of the rivets was as heat-treated. These rivets, 60 mm long, were riveted in a divisable hardened die having a hole of 13 mm, see fig.10. The radius between the shank and the conical counterhead was r « 0.1 mm in order to make a severe stress concen­tration. 5 rivets were tested for more than 1000 hours in an aerated solution of 3.5% NaCl in distilled water under an axial load of 50 kN (S=380 N/mm2 = 0,35 Rm.) Not any indication of stress corrosion was observed after these tests, so these rivets can be regarded as safe with regard to corrosion from a sea-side atmosphere or from road salt.

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5. High friction primers.

In [1] it was found that the use of common primers often leads to the loosening of rivets, and it was established that the application of a high friction primer prevented this loosening and increased the fatigue resistance of the riveted joints considerably. Furthermore blasting the surface of the plate and the hole to SIS-SA21/2 to SA3 with steel grit (surface roughness Ra = 5um) instead of with corund grit (Ra = 15um) resulted in somewhat higher endurances.

Therefore all plates were shotblasted after punching the holes and coated (thickness 30um) with the newly developed high friction primer, Wap 9.5,(udyn > 0,50) described in Part II.

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6. Endurance data for riveted joints 6.1. General

The following details apply for all joints and tests; Plates material : FeE560TM, heat II thickness : 6mm holes : punched, 13mm Surface treatment (after punching) :

shotblasted to SIS-SA21/2 - 3 coated with 30pm primer Wap 9.5

Rivets type and size : roundhead, 12mm hardness : 280 HB coating : galvanized 5-8 urn thick, yellow

passivated. riveting : conical head, both sides riveting pressure : 400 kN Fatigue tests alternating load, R = -1 Constant Amplitude tests ; Servo-hydraulic or sub-resonance machinery testing frequency 25-40 cps Variable Amplitude tests ; Servo-hydraulic machinery standard load spectrum [4] : Gauss, I = 0.99, see part I testing frequency 15-40 cps.

6.2. Joints with one rivet 6.2.1 Single shear and double shear joints

The dimensions of the specimens are given in fig.11. The results from C.A. loading of single shear (fig.11a) and double shear (fig.lib) lap joints are given in table 3 and 4 and plotted in fig.12 and 13 respectively. The Basquin type equations as determined by linear regression are: single shear : 8.06 log Fa + log N = 17.895 (1)

Ff = 23 kN (estimate) double shear : 5.53 log Fa + log N = 14.887 (2)

Ff = 31 kN The results for V.A. loading with the standard spectrum Gauss and I = 0.99 are also given in table 3 and 4 and plotted in fig.14 and 15 resp. The Basquin type equations are: single shear : 8.68 log Frms + log N = 15.352 (3) double shear : 8.13 log Frms + log N = 16.519 (4) All fractures were in the shear plane except for the double shear lap joint when Frms > 17 kN (Fmax > 90 kN) .

125

These specimens failed from hole elongation The dashed lines in fig.12 through 15 are results reported in [1] for joints having the same geometry but a different high friction primer and a round fixed head. Comparing the above mentioned results with those given in [1], yields the following results: single shear .joints: the endurance for C.A. and V.A. are about equal. double shear joints: the endurances for V.A. loading are about equal, but for C.A. loading the joints here show a reduction in endurance of a factor 2H with regard to the joints tested in [1]. However the fatigue limits are equal. The reason for this reduction in endurance for C.A. loading is not known, but it was observed that for the joints with 2 conical heads more than one crack was formed in both shear planes at the ridges on the rivet shank at the interfaces of the plates, while for joints with one standard roundhead rivet as tested in [1] there usually was only one dominating crack. So this difference in endurance is not attributed to the effect of a different high friction primer but to the effect of riveting with 2 conical headers because of a better filling around the ridges from the punched holes, see fig.6a. Note: From fractographic observations it appears that a number of different crack-initiations in the rivet are possible in riveted joints:

Cracks in joints with drilled holes emanate from fretting in the hole and then grow in the shear plane. Cracks in joints with punched holes and rivets with one round fixed head emanate from the sharp neck below this head, due to a lesser degree of filling just below this fixed head and due to sharpening see par.4,2. The multiple cracks in joints with punched holes and 2 conical heads emanate from the ridges on the rivet shank at the separation planes of the plates (remember the stapling sequence!) due to a better filling of the irregularities of the holes. All cracks emanated at sites on the centre line (load line) of the specimens. No mode III cracks were observed.

Apparently a further increase in endurance of a riveted joint can only be obtained by using drilled holes that are reamed after fixing the position of the plates with regard to one another. This technique was used formerly for pressure vessels, but is not feasible for the truck industry today.

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6.2.2 Single shear .joints with enlarged excentricity, fig.lie In order to obtain more insight into the transverse load carrying capacity of the conical pressed fixed heads, V.A. tests (Gauss, I = 0.99) were carried out on single shear lapjoints with increased excentricity, e = 9mm, see fig.lie, made with standard rivets (type S,) and with rivets type A, see fig.7. Note that for these tests no primer was used on the plates and the holes were drilled. The test setup is shown in fig.lid. The results are given in table 5 and plotted in fig.16, together with the results from [1] from SS lapjoints with normal excentricity (e = 6mm). It follows that: the endurances for joints riveted with type S or type A rivets are the same. All failures were in the shear plane, except for one specimen at Frms = 17 kN. Note that at this load the DS specimens failed from hole elongation. So it appears that the better filling of the hole due to conical pressed roundhead rivets caused a considerable improvement of the capacity of the joint to resist bending of the plates either due to excentricities or due to loads normal to the plates.

6.3. Joints with more than one rivet 6.3.1 General

The experiments described in the preceeding paragraphs were to optimize the quality and the economy of a joint with a single rivet. Investigated were the influence of the hardness of the rivet, the type of primer, the hole quality and the method of riveting. Failures were rivet failures or were due to hole elongation.

In general however, joints consist of a pattern of rivets and cracks in the plates are to occur when the net section stresses in the plates are increased. Therefore the minimum dimensions for rivet patterns were established. After this was done, the following joints were tested, see fig.17. - single shear joints with 3 rivets in series.

// ,, , , with 3 rivets parallel. - double shear joints with 4 rivets in a square. The single shear joint is typical for truck chassis, while the double shear joint is typical for steel constructions. The dimensions of specimens were minimized, see the next paragraph, in order to promote plate failure and so to determine a lower bound of the endurances.

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6.3.2 Rivet pattern dimensions. For rivet patterns in steel constructions the following dimensions are recommended (DIN 15018 T2 and DIN 1050): pitch : a £ 3d edge distance in loading direction : ei £ 2d edge distance in transverse direction : e2 £ 1.5d Because here the rivet is harder than the plate, bulging of the plate may occur more easily than with soft rivets. Therefore the edge distance necessary to prevent bulging for the FeE 560 TM plate material (Heat II) was determined experimentally for rivets of different hardnesses. The specimen used is given in fig. 18a. Criterion was (fl + f2) £ 0.4 mm, which is to be regarded as the beginning of general yield. The results are given in table 6. From table 6 it appears that the minimum edge distance, e, has to be increased from 1.5d to 2d. So for the joints with rivets parallel, the following minimum dimensions were used (see fig. 18b.): d = 12 + 1 = 13 mm a = 4d = 52 mm e2 = 2d = 26 mm ei = 2.67 d = 35 mm Note that the ratio's ei/e2 and a/e2 are kept equal to those in the DIN standards.

6.3.3 3 rivets in series,SS joint,fig.17a. The results are given in table 7. These results are plotted in fig.19 for C.A. loading and in fig.20 for V.A. loading, GAUSS, 1=0,99. In these figures the load F and also the net section stress, S (N/mm2)=2.16F(kN), are given. All fractures except 3 were plate failures. The cracks were initiated near the outer rivets by fretting between the plate surfaces. The initiation sites were a few mm away from the hole at an angle of 10° to 70Q to the tension direction, see fig. 21a. The Basquin type equations are as follows: C.A. loading : 7.80 log Fa + log N=20.392 (5)

or : 7.80 log Sa + log N=23.002 (5a) The fatigue limit for plate failure is estimated at

Ff=65 kN or Sf=140 N/mm2

V.A. loading : 7.74 log Frms + log N=16.965 (6) or : 7.74 log Srms + log N=19.554 (6a)

(3.10a<N<3.107)

Under C.A. loading and very high endurances also 3 rivet failures were observed, so that two S-N lines are to be distinguished in fig.19. The relation between these failures and those for the single rivets, eq.l, is as follows, see fig.22,:

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The loads,L, on the outer rivets are bounded between L=F/2 for rigid (riveted) connections and elastic plates and L=F/3 for rigid plates and elastic (riveted) connections. Therefore in fig. 19 also drawn are the endurance lines which would apply when this joint would have 2 or 3 times the load carrying capacity of the joint with one rivet, eq.l. From the rivet failures it follows that the loads on the (outer) rivets were close to F/3. So the rivet acts as a "soft spring". In fig.20 the lines for 2 and 3 x (eq.3) are plotted too, to allow further comparison with the single shear joint,although no rivet failures were observed under V.A. loading.

6.3.4 3 rivets parallel, SS joint, fig.17b. The results are given in table 8. These results are plotted in fig. 23 for C.A. loading and fig.24 for V.A. loading, GAUSS, 1=0.99. All specimens showed plate failure and the cracks were again initiated by fretting. For C.A. loading the initiation sites were as sketched in fig.21a. For V.A. loading however, the crack in the centre initiated in front of the hole, see fig.21b., apparently from plate bending due to the high peak loads. The Basquin type equations are

C.A. loading : 4.50 log Sa +log N=14.761 (7) St £ 71 N/mm2

V.A. loading : 8.15 log Srms +log N=19.823 (8)

A plot in fig 24 of 3 x (eq.3) shows that for V.A. loading the loadcarrying capacity of this joint is almost 3 times that of the joint with one rivet (however no rivet failures did occur). In view of the above and of the low scatter for V.A. loading where Smax=5.26Srms , i t i s t h o u g h t t h a t t h e s c a t t e r f o r C.A. loading is caused by differences in load transfer between the 3 rivets. Note that the extremely low fatigue limit, Sf £ 71 N/mm2, as compared to the joints with 3 rivets in series, Sf=140 N/mm2, or to the joints with 4 rivets in a square, Sf=171 N/mm2, (described in the next paragraphs) is in agreement with an unequal load transfer by the rivets. Therefore the Sa-N curve, eq.7 is based on the lower endurances measured and is meant as a base for a design curve only.

6.3.5 4 rivets in a square PS joint, fig 17c. The results are given in table 9. These results are plotted in fig.25 for C.A. loading and

129

in fig.26 for V.A. loading, GAUSS, J=0.99. All cracks initiated again from fretting between the plates near the holes, fig.21a, except the one specimen having a low endurance of 2.4 * 106 cycles at Frms = 34 kN. The cracks in this specimen initiated from the edges of the holes in the plate, so most probably the degree of filling of the holes by the rivets has been insufficient, see also par.6.4. The Basquin type equations are

C.A. loading : 4.31 log Sa + log N= 15.727 (9) Note that eq.9 is based, just as eq.7, on the lower endurances in the scatterband. The fatigue limit is estimated at

Sf = 171 N/mm2 which means that, based as usual on the net section stress, Kf = 2 and very low indeed.

V.A. loading : 5.72 log Srms + log N= 17.580 (10) Note that eq. 10 follows from linear regression based on all specimens. From fig.15 it follows that hole elongation (Frms £ 17 kN or Fmax £ 89 kN per rivet) was not to occur, see par.6.3.3. and par.6.2. of Part II. Note also that Frms = 45 kN equals the highest load, viz.240 kN, allowable in our servo controlled fatigue machines, so that all endurances are beyond 106.

6.4. Hole expansion In view of the high fatigue limits obtained for the different joints, together with the observations of crack initiations by fretting on the plate surfaces, fig.21, the condition of the holes after riveting was investigated. Therefore 3 precision riveted, double shear (= 3 x 6mm) joints were made. The holes in the plates were reamed and equal to 13.125 ± 0.001mm The rivet shanks were ground to 12.00 mm diameter. The end planes of the shanks were made flat and square to the shanks in order to prevent a-symmetric riveting. After riveting, the plates were carefully removed from the rivets, and the diameter of the rivet shanks was measured at 9 places ( 3 per plate, in axial and in transverse direction). From comparison of the initial hole diameter (nominal 13mm) with the (unloaded) shank diameter after riveting it was established that, in the central plate, c.q. at the central section of the shank the diameter was 1,0 to 1,2 % larger than the original hole. This means that during riveting the hole is expanded by plastic deformation and after removing the riveting pressure, residual compressive stresses will result.

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The benifical effect of residual compressive stresses has been known for a long time and the effect on fatigue of expanding the holes of lugs by plastic deformation was already shown by Schijve [5] in 1964. The occurence of this phenomenom in riveted joints is, to the authors knowledge never reported in the open literature. It is apparently created by overmatching the hardness of the rivet as compared to the plate. Furthermore it follows from elementary calculations that an interference fit between rivet shank and hole results, because the stresses in the shank are mainly hydrostatic while in the plate the shear stresses are dominating the deformation. Now coming back to the actual riveted joints, it is obvious that the quality of riveting obtained in the laboratory can hardly be met in production. It follows that, at least with punched holes which are partly tapered, fig.6a, and an unfavourable stapling as in fig.6b, it will be difficult to reach a hole expansion across the full thickness of the plates. If this is not the case, than low endurance failures may occur as the one at Frras = 34 kN desc r ibed in p a r . 6 . 3 . 5 .

Therefore it is advised to use drilled holes for those joints in the construction that are most fatigue-critical.

131

7. Fatigue tests on truck frames 7.1. General

In order to demonstrate that the results of tests on small specimens are also applicable for constructions, tests under service loading conditions were carried out on complete vehicle frames. These tests were carried out in the Development Laboratory of DAF Trucks. The chassis chosen for this purpose is that of a tractor with 2 rear axles, fig.27. The H-shaped cross member in the center part of this frame, see fig.27, is heavily loaded and the first prototype tested several years ago did not comply with the desired service life. Failures were in the connections between cross member and coverplates. Therefore this prototype frame was very suited for this purpose, because the standard service loading could also be applied to the "reinforced" version without running into excessive endurances.

"7.2. Frame specimens 5 Specimens were tested viz: Frame A:

H-beam equal to the original first prototype,see fig.28a

plate material : 6mm FeE490TM, pickled before painting rivets : 12mm, A5-A1 (Fe 360) primer : red washprimer

This test is used as a reference Frame g and C:

H-beam geometry- equal to frame A, but: plate material*: 6mm FeE560T^, shot blasted before

painting, rivets : 12mm, hardness 280 HB primer : Wap 9.5 (high friction)

Frame ß and E: H-beam geometry: as shown in fig.28d and 28e. rivet pattern : modified as shown in fig.28d and 28e.

(4 instead of 6 rivets) plate material : aş frames B and C rivets : " primer : "

7.3. Test stand The forces imposed by actuators on the complete frame (without wheels) are in general: 1. Frame torsion (displacement control), introduced at

the front axle. 2. Side loads, imposed on the 2nd rear axle and reacted

132

by the fixed first axle and the front axle. 3. Loads from the trailer imposed on the "fifth" wheel,

a. vertical loads b. side loads c. longitudinal loads (not applied here because

they are not important for the cross member). All actuaters are synchronized by computer control.

7.4. Service-load spectrum From the loads given in par.7.3, torsion and side loads are the most important ones for the cross member. Vertical-and side loads from the "fifth wheel" are less important. Longitudinal loads from this "fifth wheel" do not really strain the cross member, so these loads were omitted. One block of the load spectrum consists of a follow up of 4 sub-blocks, viz:

urban route, fast turns - secondary roads, low speed turns

track on a construction site urban route, low speed turns

Further details are not of interest because all frames experienced the same load blocks and the differences appear as the number of blocks to failure.

7.5. Results of service-load tests. Frame A (reference frame) , see fig.28a. block 457 : first rivet failure block 926 : most rivets have failed

(failed rivets were replaced during the test by pretensioned bolts grade 12.9)

block 2010 : crack in cover plate, (repaired by welding)

block 2692 : crack in cover plate (end of test)

These results compare with the results from tests on the original prototype. Frame B, see fig.28b. block 3277 : crack in cover plate

(repaired by welding) 2 rivets removed deliberately as a try-out for future modifications,

block 4588 : cracks in the cross member from the hole of one of the removed rivets. (end of test)

Frame C, see fig.28c. block 1000 : 2 rivets removed deliberately block 3782 : 2 rivet failures

(replaced by new rivets, riveted in situ) block 4911 : crack in cross member

(end of test)

133

From the above it is clear that the endurances of the H-shaped members of frame B and C, built from FeE560TM, coated with primer Wap 9.5 and joined with H.S. rivets, show a 10-fold increase in endurance. So the technological applicability of this development is proved.

As a follow up to the tests on the frames B and C, two more frames, designated D and E, having a more simple rivet connection (4 instead of 6 rivets) in a different pattern, see fig. 28d and 28e, were tested.

The results were as follows:

Frame D, see fig.28d. block 3271: no rivet failure, end of test.

Frame E, see fig.28e. block 2691: 1 rivet failure, end of test.

After final inspection 7 rivet shanks showed small cracks in the plane between cover plate and H-beam.

From these latter 2 tests it is once more clear that the application of an improved joining technique having a much greater fatigue resistance also allows a more efficient design of structural connections.

So from tests B and C and from tests D and E in comparison with test A, the advantages of the application of this technological development are proved in abundance.

134 -

CONCLUSIONS In order to improve the fatigue strength of riveted joints in high strength steels, high strength rivets were developed.[1] Joints made with these rivets in 6 mm FeE 560TM steel plate were extensively tested. From optimizing the joining techniques on the base of fatigue- and other tests it follows that - Rivets from 18 Mn B4 steel, Q and T to 350 HB or less,

are rivetable by cold pressing.[1] - For FeE 560TM plate material rivets having a hardness of 280 HB are the most suitable.[1]

- Common corrosion resisting primers have a detrimental effect on the endurance. High friction primers are to be used.

- The use of a conical header for the round head of the rivet prevents fatigue failures from the head-shank transition.

- The edge distance of rivets having a hardness á 280 HB to avoid plate bulging in FeE560TM is e £ 2d

- When plate failures occur, as in joints with more than one rivet, cracks initiate from fretting between the plates. The expansion of the holes by plasticity due to the riveting prevents the cracks from emanating from the hole edges. Whether this feature is attainable when joints are made during industrial production is not known.

- No stress corrosion was observed for 320 HB rivets under sustained load tests in brine at S= 380 N/mm2.

- Furthermore it appears that the endurance of open holes depends heavily on the quality of these holes.

Endurance curves were obtained for constant amplitude loading and for variable amplitude loading with the standard spectrum GAUSS, I = 0.99, for the following joints : - single shear joints with 1 rivet

3 rivets in series 3 rivets parallel

- double shear joints with 1 rivet 4 rivets in a square

These endurance curves can be used as (preliminary) design curves.

Tests on truck frames with fatigue-critical cross member connections have shown that the improvements apparent for small specimens are also realized for complex structural elements as are cross beam-coverplate connections in a chassis and can lead to a more economical construction.

- 135 -

9. References 1. Overbeeke,J.L. : The fatigue strength of welded, bolted

and riveted joints in high strength low alloy steel. Euroreport Eur 9964 MF EN, 1985.

2. Wanhill, R.J.H. : Microstuctural influences on the fatigue and fracture resistance in high strength structual materials.' Eng.Fracture Mechanics 10. (1978) 337-357

3. Kloos, K.H., Landgrebe, R., Speckhardt, H. : Untersuchungen zur Wasserstof-induzierten Rissbildung bei hochfesten Schrauben aus Vergütungsstählen. VDI-Z. Band 127 (1985) Heft 19.

4. Schütz, W. : Standardized stress-time histories - An overview. ASTM STP-1006 (1989) p.3-16.

5. Schijve,J.J. : Analysis of fatigue phenomena in Aluminium Alloys . Thesis, Delft Univ. of Techn. 1964.

136 -

Table 1. Mechanical Properties of the 6mm FeE560TM Steel Plates.

FeE 560TM Heat I Heat II casting direction Re, N/mm2 Rm, N/mm2 A (1=80)% Hardness, HB Fatigue limit, Sf ,N/mm2,R =-1

Ingot

» 628 716 25 240 360

cast -L

658 727 21

n.d.

concast

» 611 694 30

230 304

±

626 709 25

n.d.

Table 2.

Series Material Rm N/mm2

Sf N/mm2

Hole

m=log S/l( Sfk N/mm2* Kf/Kt *) gross

Results of fatigue tests 13mm

Dg N

)

central 1

Heat II 694 304 drilled

3.73 90 1.1

hole, Kt= 3,

section stress

2a Heat II

punched

3.56 40 2.5

on plates ,04.

3 Heat II

badly punched 3.61 28**) 3.6**)

with a

2b Heat I 716 360 punched

3.73 50 2.35

**) estimate

137 -

Table 3. Results of endurance tests on single shear riveted .joints, R = -1.

C.A. V.A. Gauss, I = 0.99 Fa N r e m a r k s * ) Frms N r e m a r k s kN * 1 0 - 3 kN * 1 0 - 3

40 40 40 35 35 35 30 30 30 30 27 27 27 23 23 23

56 97

112 203 416 453 387 778

1,387 1,723 1,547 1,579 4,119

493 2,328

>31,249

17 14.5 14.5 12 12 10 10 9 9 ' 8

59 82 451 598 650

9,207 4,051

>21,658 >28,056 >21,432

25 4,401 *) All failures in the shear plane, unless indicated

Table 4. Results of endurance tests on double shear riveted joints, R = -1,

C.A. V.A. Gauss, I = 0.99 Fa N remarks Frms N remarks kN *10-3 kN *10-3

60 329 53 333 53 147 45 295 40 731 35 1,511 35 3,222 33 5,382 30 >26,662 60 61 30 >29,742 45 519

20 20 18 18 17 17 17 15 15 13.5 13.5

1,918 831

1,191 1,810 2,404 1,285 6,984 9,018

12,841 20,736 26,073

1*) 1 1 1

*) All failures in the shear plane, unless indicated 1) failed by hole elongation

138

Table 5.

Rive Frms kN 8.0

10.0 10.5 11.0 11.0 12.5 12.5 14.5 16.0 17.0

Results oi increased rivetinq Loading :

it type "S" N

*10-3 >5,000 4,754 2,179 1,656 2,521

322 515 183 50 11

: endurance t :ests excentricitv, e = Spectrum Gauss,

remarks*)

runout

on lap joi nts with 9mm, double conical

J = 0.99,

Rivet type Frms kN

10.0 10.5 11.0 11.0 12.5 12.5 14.5 16.0 17.0

N *10-3

>9,784 2,057 1,287 1,311

306 446 90 63 39

R = -1.

"A" remarks

runout

head failure

*) All failures in the shear plane except when indicated

Table 6. Maximum rivetinq pressure FB to prevent bulging

Plate HB

240

stripwidth (fi + U

Rivet HB 230 280 313 330 354

<. 2e = 52mm 0,4mm).

FB kN 350 400

400

(fig.18),

F*) kN 350 400 470

530

*) F = necessary riveting pressure [1]

- 139

Table 7. Results of endurance tests at R = -1 on riveted lap-joints, 3 rivets in series, single shear.

C.A. tests V.A. tests, I = 0.99 Fa N remarks Frms N remarks kN *10"3 kN *10"3

100 100 90 90 80 80 70 70 70 65 65 60 60 60

48 58

199 130 466 271 829 907

1,456 7,597 1,325 1,611

14,437 10,354

failed rivet failed rivet rivet

[ in fai [ in fai fai

grip lure grip lure lure

35 30 30 27.5 25 22.5 22.5 20 20

20 317 434 700

1,534 1,934 2,354 7,991

13,175

All failures through the plates unless indicated.

Table 8. Results of endurance tests at R = -1 on riveted lap joints, 3 rivets parallel, single shear.

C.A. tests V.A. tests, I = 0.99 Fa N remarks Frms N remarks kN *10-3 kN *10-3

85 85 75 75 75 70 65 65 60 55 55 50 50

396 649 699

1,247 2,072

580 825

3,774 1,839 1,732 2,930

>14,005 17,138

45 40 40 35 32.5 30 27.5 25

200 211 366 995

1,069 3,358 7,683 19,666

All failures through the plates,

140

Table 9. Results of endurance tests at R = -1 on riveted

Fa kN 120 110 100 100 100 100 120 90 90 85 80 80 80

110

lap joints

C.A. tests N

*10-3 195 353

1,429 7,831

510 >2,889

215 749

1,039 723

>15,574 >12,626 >8,133 1,001

4 rivets

remarks

fail. ]

]

2d in gr

in

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R.A. Frms kN 45 45 40 40 37 37 35 34 34 34 31 31

tests, I N *10-3 1,454 3,726 3,223 3,358 3,420 3,716

>15,475 2,358 7,227 >9,092 19,163 >32,417

.e shear.

= 0.99 remarks

failed in grip

failed in grip

All failures through the plates, unless indicated,

141

Figure 1 . Specimen wi th a central hole.

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K = 2.63 ref. to net section stress

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I

European Communities — Commission

EUR 13626 — The application of welded, bolted and riveted connec­tions in HSLA steel in structures subjected to high-dynamic loading

J. L Overbeeke

Luxembourg: Office for Official Publications of the European Communities

1991 - XV, 174 pp., num. tab., fig. - 21.0 x 29.7 cm

Technical steel research series

ISBN 92-826-2819-1

Catalogue number: CD-NA-13626-EN-C

Price (excluding VAT) in Luxembourg: ECU 15

It is well known that the fatigue resistance of a construction depends heavily upon the quality of its joints. Therefore the application of high-strength steels, which can lead to a reduc­tion in structural weight and operational costs, is only justified when the qua­lity of the joints is upgraded. This report describes an extensive investigation into the quality of fixed connections in HS steel with regard to fatigue. The material used was hot-rolled HS steel, complying with FeE 560TM. The plate thickness was mainly 6 mm and the types of connections investigated were welded, bolted and riveted joints. From the above it follows that this research applies in the first place to the vehicle industry where weight savings increase the efficiency of transporta­tion considerably.

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