Fracture Toughness Characterization of HSLA Steel WeldmentsExecutive summary Fracture Toughness Characterization of HSLA Steel Weldments Nick Pussegoda; DRDC Atlantic CR 2008-178;

Defence R&D Canada – Atlantic

DEFENCE DÉFENSE&

Fracture Toughness Characterization of

HSLA Steel Weldments

Nick PussegodaBMT Fleet Technology Limited

BMT Fleet Technology Limited311 Legget DriveKanata, ON K2K 1Z8

Contractor Document Number: 6386DFR

Contract Project Manager: Nick Pussegoda, 613-592-2830, ext 205

Contract Project Number: W7707-078025/001/HAL

Contract Scientific Authority: Christopher Bayley, 250-363-4784

The scientific or technical validity of this Contract Report is entirely the responsibility of the Contractor and thecontents do not necessarily have the approval or endorsement of Defence R&D Canada.

Contract Report

DRDC Atlantic CR 2008-178

October 2008

Copy No. _____

Defence Research andDevelopment Canada

Recherche et développementpour la défense Canada

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Fracture Toughness Characterization of HSLA Steel Weldments

Nick Pussegoda BMT Fleet Technology Ltd Prepared By: BMT Fleet Technology Limited 311 Legget Drive Kanata, ON K2K 1Z8 Document Number: 6386DFR Contract Project Manager: Nick Pussegoda, 613-592-2830 ext 205 Contract Project Number: W7707-078025/001/HAL Contract Scientific Authority: Christopher Bayley, 250-363-4784

The scientific or technical validity of this Contract Report is entirely the responsibility of the Contractor and the contents do not necessarily have the approval or endorsement of Defence R&D Canada.

Defence R&D Canada – Atlantic Contract Report DRDC Atlantic CR 2008-178 October 2008

Principal Author

Original signed by Nick Pussegoda

Nick Pussegoda

Approved by

Original signed by Terry Foster

Terry Foster

H/DLP

Approved for release by

Original signed by Ron Kuwahara for

Calvin Hyatt

Chair DRP

© Her Majesty the Queen in Right of Canada, as represented by the Minister of National Defence, 2008

© Sa Majesté la Reine (en droit du Canada), telle que représentée par le ministre de la Défense nationale, 2008

Abstract ……..

High Strength Low Alloy Steel with a minimum specified yield strength (SMYS) of 65 ksi (~450 MPa) (HSLA 65 - ASTM A945 grade 65) has drawn significant interest from Navy and Commercial Shipyards. This is primarily because of weight savings and weldability. It also is intermediate between high strength steels (HSS) with 50 ksi (~450 MPa) SMYS and HSLA 80 with 80 ksi (~550 MPa) SMYS. ASTM A945 grade 65 also specifies a tensile strength of 78-100 ksi (540-690 MPa) and a minimum transversely oriented Charpy impact energy of 70 ft-lb (95 J) at -40oC.

The present study demonstrates the influence of loading rate, temperature and crack location on the fracture toughness. While the specimen geometry of the SENB is known to result in conservative fracture toughness values, it clearly demonstrates the deleterious influence of welding on the fracture behaviour in the coarse grained HAZ. In comparison to the coarse grained HAZ, the two weld metals, 101TM and 71HYN, displayed higher toughness.

Résumé ….....

L’acier faiblement allié très résistant du type HSLA 65 (norme ASTM A945, nuance 65), qui présente une valeur nominale minimum de limite élastique proportionnelle (LEP) de 65 ksi (milliers de lb/po2), soit ~ 450 MPa, constitue un alliage d’un grand intérêt pour les chantiers navals commerciaux et militaires, notamment au chapitre de sa soudabilité et des économies de poids associées à son emploi. L’acier HSLA 65 constitue aussi un alliage intermédiaire qui se situe entre les aciers à haute résistance (AHR), qui présentent une LEP de 50 ksi (~ 450 MPa) et l’acier du type HSLA 80, dont la LEP est de 80 ksi (~ 550 MPa). Selon les exigences établies dans la norme ASTM A945, les aciers de nuance 65 doivent aussi présenter, à -40 ºC, une résistance à la traction se situant entre 78 et 100 ksi (de 540 à 690 MPa), ainsi qu’une valeur minimum de résistance au choc de 70 pi-lb (soit 95 J) lors d’essais de résilience Charpy à orientation transversale.

DRDC Atlantic CR 2008-178 i


ii DRDC Atlantic CR 2008-178

Executive summary

Fracture Toughness Characterization of HSLA Steel Weldments Nick Pussegoda; DRDC Atlantic CR 2008-178; Defence R&D Canada – Atlantic; October 2008.

Introduction or background: High Strength Low Alloy Steel with a minimum specified yield strength (SMYS) of 65 ksi (~450 MPa) (HSLA 65 - ASTM A945 grade 65) has drawn significant interest from Navy and Commercial Shipyards. This is primarily because of weight savings and weldability. It also is intermediate between high strength steels (HSS) with 50 ksi (~450 MPa) SMYS and HSLA 80 with 80 ksi (~550 MPa) SMYS. ASTM A945 grade 65 also specifies a tensile strength of 78-100 ksi (540-690 MPa) and a minimum Charpy impact energy of 70 ft-lb (95 J) at -40 C. o

In a previous study, flux core welding procedures were developed for a matched consumable (Hobart 71HYN) and overmatched consumable (Hobart 101TM) at two weld heat inputs. Welds fabricated from these procedures were tested in order to determine the fracture and mechanical properties of the weld metal and heat affected zone (HAZ). The current study is focused on comparing the fracture toughness of the weldments obtained from the procedures developed for the low heat input weld procedures. This involved both small coupon (ASTM E1820 Single Edge Notched (SENB) testing and test samples that are more representative of structural loading (wide plate tests) with cracks either in the weld metal or HAZ.

This report provides the results from tests carried out on small coupon testing. The fracture toughness tests were performed at both quasi-static (QS) and dynamic loading rates at a representative temperature close to the lowest anticipated service temperature. The dynamic loading rate is representative of slamming under wave loading. Testing was carried out at both the minimum service temperature (-40 C) and at -30 C to assess any change in the fracture toughness behaviour at these two temperatures.

o o

Results: The fracture toughness tests demonstrate the temperature and rate sensitivity of the weld and HAZ. For the fracture tests performed at -30 C the following observations were made: o

• For notches located in the weld metal and tested at the quasi-static loading rate, both welds, made with either the Hobart 71HYN consumable or the Hobart 101TM consumable, resulted in ductile behaviour. Increasing the loading rate for the Hobart 101TM weld resulted in an unstable fracture without any visible ductile extension of the fatigue crack. In contrast, the 71HYN weld metal tested at the higher loading rate remained ductile and did not produce fracture transition.

• For the HAZ region close to the weld fusion line, quasi-static fracture tests displayed transition behaviour while increasing the loading rate generated unstable fracture without any visible ductile extension of the fatigue crack.

A limited number of dynamic and quasi-static fracture tests were performed at -40 C and were restricted to the conditions which had previously generated transition or ductile behaviour.

o

DRDC Atlantic CR 2008-178 iii

• Dynamic rate tests for the 71HYN weld metal indicated that -40 C is the lower limit of ductile behaviour and fracture transition would occur below this temperature. Quasi-static tests were also performed with the fatigue crack located in the HAZ. As expected the fracture toughness values were lower than those tested at -30 C and resulted in unstable fracture without visible extension of the fatigue crack.

o

o

Significance: The present study demonstrates the influence of loading rate, temperature and crack location on the fracture toughness. While the specimen geometry of the SENB is known to result in conservative fracture toughness values, it clearly demonstrates the deleterious influence of welding on the fracture behaviour in the coarse grained HAZ. In comparison to the Coarse grained HAZ, the two weld metals, 101TM and 71HYN, displayed higher toughness.

Future plans: In order to determine the structural significance of the experimentally acquired fracture toughness values, large scale test samples should be tested and compared to the current fracture toughness values.

iv DRDC Atlantic CR 2008-178

Sommaire .....

Caractérisation de la résistance à la rupture de joints soudés effectués avec l’acier de type HSLA

Nick Pussegoda; DRDC Atlantic CR 2008-178; R & D pour la défense Canada – Atlantique; octobre 2008.

Introduction : L’acier faiblement allié très résistant du type HSLA 65 (norme ASTM A945, nuance 65), qui présente une valeur nominale minimum de limite élastique proportionnelle (LEP) de 65 ksi (milliers de lb/po2), soit ~ 450 MPa, constitue un alliage d’un grand intérêt pour les chantiers navals commerciaux et militaires, notamment au chapitre de sa soudabilité et des économies de poids associées à son emploi. L’acier HSLA 65 constitue aussi un alliage intermédiaire qui se situe entre les aciers à haute résistance (AHR), qui présentent une LEP de 50 ksi (~ 450 MPa) et l’acier du type HSLA 80, dont la LEP est de 80 ksi (~ 550 MPa). Selon les exigences établies dans la norme ASTM A945, les aciers de nuance 65 doivent aussi présenter, à -40 ºC, une résistance à la traction se situant entre 78 et 100 ksi (de 540 à 690 MPa), ainsi qu’une valeur minimum de résistance au choc de 70 pi-lb (soit 95 J) lors d’essais de résilience Charpy à orientation transversale.

Dans le cadre d’une étude antérieure, on a élaboré des procédures de soudage avec fil fourré, avec une électrode consommable assortie (Hobart 71HYN) et une autre électrode consommable de nature et de propriétés différentes du joint soudé (Hobart 101TM), à deux valeurs distinctes d’apport de chaleur. Les joints soudés obtenus en suivant ces procédures ont été mis à l’essai afin de déterminer les propriétés mécaniques et la résistance à la rupture du métal de soudure et de la zone thermiquement affectée (ZTA). L’étude dont les résultats sont fournis dans le présent rapport a pour principal objectif de comparer la résistance à la fracture des joints soudés obtenus en suivant les procédures de soudage élaborées pour fabriquer des soudures avec de faibles apports de chaleur. Les travaux exécutés comprenaient entre autres la mise à l’essai, en flexion, d’éprouvettes de petite taille entaillées d’un seul côté (essai de la norme ASTM E1820; Single Edge Notched Bending ou SENB), ainsi que l’examen d’échantillons plus représentatifs de charges structurales (essais avec des plaques larges) présentant des fissures dans le métal de soudure ou dans la ZTA.

Le présent rapport contient les résultats des essais réalisés avec des éprouvettes de petite taille. Les essais de résistance à la rupture ont été exécutés en utilisant une vitesse d’application de charge quasi statique et une vitesse d’application de charge dynamique, et ce, à une température représentative de la plus basse température de service prévue. La vitesse d’application de charge dynamique correspond au tossage que subissent les navires soumis à la charge des vagues. Les essais ont été effectués à la température de service minimum (-40 ºC) ainsi qu’à -30 ºC afin d’évaluer toute variation de la résistance à la rupture des matériaux à ces deux températures.

Résultats : Les résultats des essais de résistance à la rupture démontrent la sensibilité de la soudure et de la ZTA à la température et à la vitesse d’application de la charge. Voici une liste d’observations associées aux essais de cette nature exécutés à -30 ºC :

DRDC Atlantic CR 2008-178 v

• Dans le cas des entailles situées dans le métal de soudure et dont la mise à l’essai a été effectuée à une vitesse d’application de charge quasi statique, les deux joints soudés, soit celui obtenu avec l’électrode consommable Hobart 71HYN et celui obtenu avec l’électrode consommable Hobart 101TM, présentaient un comportement ductile. Dans le cas du joint soudé obtenu avec l’électrode consommable Hobart 101TM, l’augmentation de la vitesse d’application de la charge a entraîné la formation d’une rupture instable sans aucun indice visible d’une propagation ductile de la fissure de fatigue. Par contraste, le métal de soudure du joint soudé obtenu avec l’électrode consommable Hobart 71HYN, qui a été mis à l’essai à une valeur plus élevée de vitesse d’application de la charge, est resté ductile et n’a pas entraîné une transition de comportement en rupture.

• Dans le cas de la section de la ZTA située à proximité de la ligne de fusion du cordon de soudure, les résultats des essais de résistance à la rupture effectués en mode quasi statique indiquent qu’il y a transition de comportement tandis que l’augmentation de la vitesse d’application de la charge a entraîné la formation d’une rupture instable sans aucun indice visible d’une propagation ductile de la fissure de fatigue.

Un nombre restreint d’essais de résistance à la rupture, effectués en modes dynamique et quasi statique, ont été réalisés à -40 ºC et leurs conditions d’exécution ont été restreintes à celles qui avaient entraîné une transition de comportement ou un comportement ductile dans les essais antérieurs.

• Les résultats des essais effectués avec une vitesse d’application de charge dynamique sur le métal de soudure du joint soudé obtenu avec l’électrode consommable Hobart 71HYN indiquent que la température limite inférieure du comportement ductile est de -40 ºC et qu’à une température inférieure, une transition de comportement en rupture se produirait. Des essais en mode quasi statique ont aussi été exécutés, dans le cas d’une fissure de fatigue située dans la ZTA. Tel que prévu, les valeurs de résistance à la rupture étaient inférieures à celles des essais correspondants effectués à -30 ºC et les conditions ont entraîné la formation d’une rupture instable sans aucun indice visible d’une propagation ductile de la fissure de fatigue.

Portée : Les résultats de l’étude qui fait l’objet du présent rapport démontrent les effets de la vitesse d’application de la charge, de la température et de l’emplacement de la fissure sur la résistance à la rupture des éprouvettes mises à l’essai. Bien qu’il ait déjà été établi que la géométrie d’une éprouvette en flexion, entaillée d’un seul côté (SENB), se traduit par des valeurs prudentes de résistance à la rupture, les résultats démontrent clairement les effets nocifs du soudage sur le comportement, au chapitre de la rupture, de la zone thermiquement affectée (ZTA) à grain grossier. Comparativement à cette dernière, les deux métaux de soudure, soit 101TM et 71HYN, présentent une résistance supérieure.

Recherches futures : Afin de pouvoir déterminer la pertinence, en matière de structures, de ces valeurs expérimentales de résistance à la rupture, il faudra effectuer la mise à l’essai à grande échelle d’éprouvettes et comparer les résultats obtenus avec les données actuelles correspondantes.

vi DRDC Atlantic CR 2008-178

Table of contents

................................................................................................................................. i Abstract ……..................................................................................................................................... i Résumé ….....

........................................................................................................................ iii Executive summary................................................................................................................................... v Sommaire .....

........................................................................................................................... vii Table of contents............................................................................................................................... viii List of figures

................................................................................................................................... ix List of tables ............................................................................................................................... 1 1 Introduction ........................................................................................... 2 2 Preperation of Panels and Welding

2.1 ........................................................ 4 Non-Destructive Examination of the Weld Panels .......................................................... 5 3 Fracture Toughness Characterization of the Weldments

3.1 ....................................................... 5 Machining of Fracture Toughness Test Specimens3.2 ...................................................................................................... 6 Fatigue Pre-cracking3.3 .......................................................................................... 7 Fracture Toughness Testing

3.3.1 ...................................................................... 8 Quasi-static loading rate testing3.3.2 .......................................................................... 9 Dynamic loading rate testing

3.4 ................................................................................... 9 Fracture Toughness Test Results3.4.1 .................................................................................. 9 Quasi-static loading rate3.4.2 .................................................................................... 13 Dynamic loading rate

3.5 ...................................................................... 14 Additional Fracture Toughness Testing ............................................................................. 16 4 Discussion of Fracture Toughness Results

4.1 .............................................................................................. 16 Quasi-static loading rate4.2 .................................................................................................. 19 Dynamic loading rate

............................................................................................................. 21 5 Concluding Summary.. ................................................................................................... 23 Annex A Radiography Reports.. ................. 25 Annex B Fracture Toughness Results at Quasi-static (QS) Loading Rate (at -30oC).. .............................. 27 Annex C Fracture Toughness Results at Dynamic Loading Rate (at -30oC).. .......................................... 29 Annex D Fracture Toughness Results for Test Carried out at -40oC

............................................................................................................................... 30 References ................................................................................................................................... 31 Distribution list

DRDC Atlantic CR 2008-178 vii

List of figures

............................................................................................ 2 Figure 1 Prepared Joint Configuration

.................................................................. 2 Figure 2 Preparation Outline for the Four Weld Panels

Figure 3 Macrographs from Welds made using Hobart 101TM (left) and 71HYN.(right) The arrows mark the straight edge side. ............................................................................... 4

Figure 4 Fracture Toughness Specimen Extraction Sketch. (25 mm length of weld was removed from the ends of the welds before saw cutting specimen coupons) ............... 5

Figure 5 Shows the intended fracture plane and specimen location of the machined (B x 2B) specimens. The black broken line indicates the notch location for the HAZ and the white broken line indicates the weld notch location. .................................................... 6

.......................................................................................... 7 Figure 6 Fracture Toughness Specimen

................................................................. 8 Figure 7 Fracture Toughness Testing Cooling Chamber

Figure 8 Load-CMOD curves displaying Jeot behaviour for weld metal specimens 101TM-02 (left) and 101TM-7C (right). The arrow indicates the end of test.............................. 10

Figure 9 Fracture faces of 101TM-02 (left) and 101TM-7C (right). The white arrow show the fatigue crack region and black arrow marks the region of crack extension during the fracture toughness test. .......................................................................................... 10

.......................... 11 Figure 10 Load-CMOD curve display Ju behaviour for test 101TM-01 (HAZ).

Figure 11 Fracture face of 101TM-01. The white arrow shows the fatigue crack region and the broken green line marks the boundary of the crack extension during the fracture toughness test. ................................................................................................ 11

...................... 12 Figure 12 Load-CMOD curve displaying J behaviour for test 101TM-05 (HAZ).c

Figure 13 Fracture face of 101TM-05. The white arrow shows the fatigue crack region and no visible crack extension occurred beyond fatigue crack during the fracture toughness test compared to what is shown in Figure 11. ............................................ 12

........................................ 13 Figure 14 Load-CMOD curve for Test for 101TM-06A is a HAZ test.

Figure 15 Fracture face of 101TM-06A. The white arrow shows the fatigue crack region and no visible crack extension occurred during the fracture toughness test. ..................... 14

Figure 16 Fracture face of the weld centerline specimen 101TM-03 which exhibited Jc. The white arrow shows the fatigue crack region and no visible crack extension occurred prior to the onset of brittle failure during the fracture toughness test. ......... 16

Figure 17 Macrographs taken at the Fatigue Crack in 101TM-05 – Jc ( a) and 101TM-2H J (b). The fusion line is highlighted by the arrow.

eot ........................................................ 19

viii DRDC Atlantic CR 2008-178

List of tables

.................................................................. 3 Table 1 Welding Parameters for Hobart 101TM Welds

................................................................. 3 Table 2 Welding Parameters for Hobart 71HYN Welds

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x DRDC Atlantic CR 2008-178

1 Introduction

BMT Fleet Technology Limited (BMT) was tasked by the Department of National Defence, under PWGSC Contract W 7707-078025/001/HAL to make butt welds in HSLA 65 steels plate using previously developed welding procedures [1]. The welds were made in both “wide plate” and relatively narrow plates using specified welding consumables with the hope of investigating the influence of weld matching on the fracture toughness behaviour. The wide weld panels were shipped to the Dockyard Laboratories (Pacific) after fabrication and the narrow weld panels were used to machine fracture toughness test specimens and perform fracture toughness testing to characterize the properties of the weld and heat affected zone (HAZ). The fracture toughness tests were performed at both quasi-static (QS) and dynamic loading rates.

The report describes the preparation of the weld panels, welding, extraction and machining of the fracture toughness specimens. It also describes the methods used to position the Electric Discharge Machine (EDM) notch in order to sample the weld metal and optimally sample the coarse grain heat affected zone (CGHAZ), fatigue pre-cracking procedure and the subsequent fracture toughness test procedure. Results from the fracture toughness test are presented and discussed with a focus on the effect of weld metal, loading rate and location of the test region, i.e. weld metal and CGHAZ.

DRDC Atlantic CR 2008-178 1

2 Preperation of Panels and Welding

Four panels each 36 inch long (915 mm) were prepared using the HSLA-65 plate supplied to BMT Fleet Technology Limited (BMT) by Defence R&D Canada – Atlantic with the rolling direction (RD) marked. The joint configuration was a single bevel configuration with the bevel side machined at 45° angle (see Figure 1). The matting faces for welding were milled finish in order to remove any heat tinted material from the plasma cutting used to prepare the panels. The single bevel design was chosen in order to provide a nominally straight fronted HAZ.

The preparation outline for the weld panels are presented in Figure 2. The welds were made perpendicular to the rolling direction (RD) that was marked on the supplied plate. The wide panels were cut to approximately 11 inches (280 mm) and when the two panels were welded the width was approximately 22 inches (560 mm). Similarly the narrow panels were cut to approximately 5 inches (125 mm) and when the two panels were welded together the width was approximately 10 inches (250 mm).

Figure 1 Prepared Joint Configuration

Figure 2 Preparation Outline for the Four Weld Panels

2 DRDC Atlantic CR 2008-178

The flux core procedure with the consumables listed in Table 1 and Table 2 were used. The Hobart 101TM electrode was selected as an overmatching consumable and the Hobart 71HYN electrode was selected as a matching consumable [1]. The root opening was set to be between 4 to 5 mm and supported by strong backs. A contoured ceramic backing plate was used to support the molten weld pool, as shown in Figure 1. The strong backs were attached to reduce angular distortion and the set up was successful. For the narrow weld panel made with 71HYN electrode the extreme end strong back was not effective and as a result some distortion was produced at one end of this weld panel. In all, 4 equally spaced strong backs were attached along the length of the weld.

Table 1 Welding Parameters for Hobart 101TM Welds

Table 2 Welding Parameters for Hobart 71HYN Welds

The welding was performed using a mechanized welding gantry in order to ensure consistent and reproducible welds as was made in the previous program [1].

A weld cross section prepared to display the weld profile and pass sequences for the two welding procedures adopted in this program are presented in Figure 3.


Figure 3 Macrographs from Welds made using Hobart 101TM (left) and 71HYN.(right) The arrows mark the straight edge side.

2.1 Non-Destructive Examination of the Weld Panels

The four weld panels were identified as 101TM-N (narrow panel), 101TM-W (wide panel), HYN-N and HYN-W. They were X-rayed at BMT Fleet Technology Ltd by ACUREN to acceptance standard CSA W59, section 125.4.3 (dynamic requirements). The test reports are presented in Appendix A. The test reports indicate that the welds were acceptable.


3 Fracture Toughness Characterization of the Weldments

The two weld panels identified as 101TM-N and HYN-N were used to extract fracture toughness specimens. A schematic for the location of the first specimen for each panel is presented in Figure 4

Specimen locations

Figure 4 Fracture Toughness Specimen Extraction Sketch. (25 mm length of weld was removed

from the ends of the welds before saw cutting specimen coupons)

3.1 Machining of Fracture Toughness Test Specimens

The specimens were machined using the Guidelines in clause 8.1; BS 7448: Part 2: 1997 [2]. The guidelines are practical and allow tolerances for weld misalignment, distortion and specimen blank curvature. Minor weld angular distortion was present in the weld panels from the use of strong backs, while other forms of distortion was not present. The machined specimen dimensions were 12.4 mm (B) by 24.9 mm (2B) for the standard (B x 2B) type fracture toughness specimen geometry. Figure 5 shows the location of the intended fracture plane in the B dimension and specimen location with respect to the weld.


Figure 5 Shows the intended fracture plane and specimen location of the machined (B x 2B)

specimens. The black broken line indicates the notch location for the HAZ and the white broken line indicates the weld notch location.

The machined samples had a surface ground finish on the load line and support surfaces that enabled macro-etching to reveal the weld metal and the HAZ, similar to that shown in Figure 5. This allowed accurate marking of the through thickness notch/fatigue pre-crack locations, i.e. either the weld and HAZ, and followed the guidelines in clauses 6.1 and 8.2 of BS 7448: Part 2. Integrated knife edges were machined into the specimens to allow the use of an MTS clip gauge required to measure the crack mouth opening displacement (CMOD). An EDM notch was then placed to a depth of 10.5 mm using a 0.1 mm wire. The finished specimens measured 12.4 mm (B) x 24.9 mm (2B) x 112 mm (L) which allowed a 99.6 mm loading span.

Prior to pre-cracking, each specimen was laterally compressed by approximately 0.5% of the thickness of the specimen. This lateral compression is known to reduce the variation of weld residual stresses in through-thickness direction (B) of the specimen, and helps promote straight and even fatigue crack-front growth. This procedure was adopted from the guidelines in BS 7448: Part 2, Annex D.

3.2 Fatigue Pre-cracking

Fatigue pre-cracking of specimens was performed following the guidelines of both ASTM E 1820 [3] and BS 7448: Part 2. ASTM E 1820 is for base metal fracture toughness testing while BS 7448 includes specific guidance for weld metal testing. In particular, the BS 7448: Part 2 was used to determine yield and tensile strength that is used to determine the pre-cracking loads. Each of the Bx2B geometry samples were then fatigue pre-cracked to a target of a/W = 0.5.

Fatigue pre-cracking was performed in a servo-hydraulic test machines using an automated schedule developed to ensure that the fatigue loads decrease with crack growth to ensure guidelines in ASTM E 1820 are met. The developed schedule uses elastic compliance


measurements of the load and crack mouth opening displacement (CMOD) during the crack growth process to ensure the fatigue pre-cracking load guidelines are followed. The crack is grown as per the requirements of the standard using both compliance measurement and observation of crack length on the two sides of the specimen surface. Pre-cracking was performed at room temperature. After pre-cracking, the specimen compliance was used to estimate the crack length.

3.3 Fracture Toughness Testing

After fatigue pre-cracking, the specimens were side grooved, to a depth of 1.2 ± 0.1 mm, by careful machining as per clause 7.5 of ASTM E 1820. The groove angle was 60o. The side groove was aligned with the EDM slot, in most cases and in others where the 2 mm (approximate) deep fatigue crack deviated slightly from the plane of the EDM notch as observed on the two surfaces, a “correction” of the alignment of the groove location was adopted with the aim of having the side groove interact with the root of the fatigue crack. Side grooving is beneficial to improve the crack front straightness of crack extension during the fracture toughness tests. Figure 6 shows a fracture toughness specimen after side grooving. The figure also shows the integral knife edge profile and the boundary of the laterally compressed region.

Prior to testing the compliance of a few side grooved specimens was checked to examine the difference in predicted crack lengths. The predicted crack lengths after side grooving was found to be marginally greater (<0.5 mm) than the predicted length before side grooving.

Integral knife edge

Lateral compression Side groove

Figure 6 Fracture Toughness Specimen

The specimen and loading arrangement were enclosed in a cooling chamber (see Figure 7) which provided the test temperature by a solenoid controlled temperature controller. The temperature is controlled using a thermocouple spot welded to the specimen and in this way an accurate control of the specimen temperature is maintained. The solenoid controls the valve that allows the chamber to be cooled by a liquid nitrogen spray aided by a fan for forced circulation which helps achieve a uniform temperature distribution. After the temperature had stabilized for a minimum of 30 minutes, the specimens were then loaded at a quasi-static rate. The load and the clip gauge displacements were digitally acquired for the duration of the test.

During testing the load-CMOD plot was displayed in real time on the PC screen displaying the progress of the test. Later, the acquired data was used to determine the critical J from the input of


specimen dimensions, measured fatigue crack length and elastic properties in accordance with clause A1.4 in ASTM E1820. For the determination of J equations A1.5 and A1.6 were used.

Depending on the nature of the Load-CMOD curve, the specimens were broken open to reveal the fracture surface and allow detailed crack measurements. When a test was terminated with a maximum load plateau, the specimen was first heat tinted in accordance with clause 8.5.1 in ASTM E1820, immersed in liquid nitrogen an then broken open to expose the fracture surface. From the exposed fracture surface, the final crack length was measurement using the procedures in clause 8.5.3 in ASTM E1820. When a test was terminated due to a fracture instability, i.e. brittle crack extension, the specimen was immersed in liquid nitrogen and fractured in order to measure the fatigue crack length.

Thermocouple wire

Fan

Figure 7 Fracture Toughness Testing Cooling Chamber

3.3.1 Quasi-static loading rate testing

For the quasi-static tests the specimens were loaded under displacement control at a loading rate determined according to clause 8.4.2 in ASTM E1820. A loading rate (K rate) check was performed in accordance with clause 8.5 in BS 7448: Part 1-1991, using the linear region of the load-CMOD curve. From these calculations, the K loading rate for the quasi-static loading rate yield around 1−smMPa1 .


3.3.2 Dynamic loading rate testing

In order to achieve the target K-rate of 1100 −smMPa two different control programs were tested. The first control program specified a linear displacement rate while the second approach specified a single fatigue cycle. Using the first approach a ram velocity of 2 mm s-1 was specified and for the initial two tests performed on 71HYN weld center line (WCL), this loading rate produced a K rate of ~ 1−45 smMPa . In order increase the K loading rate a single fatigue cycle (sinusoidal wave) with an amplitude of 2 mm at a rate of 0.5 Hz was specified. This control program produced a K rate of ~ 1−300 smMPa . This control program was selected as the test schedule for the remaining dynamic loading rate tests.

3.4 Fracture Toughness Test Results oThe first series of test were carried out at -30 C. This test temperature coincided with Charpy

Impact Energy tests which reported a ductile to brittle transition behaviour in this range [6] Subsequent, tests at -40oC were performed in order to achieve brittle failure at dynamic conditions for the higher toughness weld metal.

J calculations was performed in accordance with clause A1.4.2.1 (Annex A1) of ASTM E 1820 with a modification of using load-CMOD curve and the Jpl determined using expressions from Proposed Annex to ASTM E 1290 [4].

3.4.1 Quasi-static loading rate

The test results are presented in Appendix B. The “failure type” was categorized as either, Jeot, Ju, and Jc. These categories are based on the shape of the load-CMOD curve and the nature of the crack extension as described below:

• Jeot when a test is terminated after the maximum load plateau. In this case the test is terminated before brittle fracture but after some ductile extension of the crack has occurred. The specimens were usually heat tinted to mark the crack extension and therefore both the fatigue crack length (aave) and final crack length (aavep) have been recorded. It is noted that this categorization is not defined in clause 9 of ASTM E1820, however, it is being currently considered in the ASTM balloting process by the E8 committee. Figure 8 shows typical load-CMOD plots for this category and Figure 9 shows the corresponding fracture face displaying the fatigue crack and the extent of the crack extension at the end of test (eot).

when fracture occurs prior to the maximum load plateau. Figure 10• Ju shows typical load-CMOD plots for this category and Figure 11 shows the corresponding fracture surface displaying the extension beyond the fatigue crack. In general, specimens where the failure type was categorized as Ju, were not heat tinted in order to preserve the metallurgy of the HAZ.

when fracture occurred before any extension beyond the fatigue crack. Figure 12• Jc shows typical load-CMOD plots for this category and Figure 13 shows the corresponding fracture surface and fatigue crack.


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]Figure 8 Load-CMOD curves displaying Jeot behaviour for weld metal specimens 101TM-02

(left) and 101TM-7C (right). The arrow indicates the end of test

Figure 9 Fracture faces of 101TM-02 (left) and 101TM-7C (right). The white arrow shows the fatigue crack region and black arrow marks the region of crack extension during the fracture

toughness test.


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Figure 10 Load-CMOD curve display Ju behaviour for test 101TM-01 (HAZ).

Figure 11 Fracture face of 101TM-01. The white arrow shows the fatigue crack region and the broken green line marks the boundary of the crack extension during the fracture toughness test.


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Figure 12 Load-CMOD curve displaying J behaviour for test 101TM-05 (HAZ). c

Figure 13 Fracture face of 101TM-05. The white arrow shows the fatigue crack region and no

visible crack extension occurred beyond fatigue crack during the fracture toughness test compared to what is shown in Figure 11.


3.4.2 Dynamic loading rate

The test results are presented in Appendix C (note that the initial couple of tests that produced at a loading rate (K rate) of ~ 1−45 smMPa are not included in the results). The “failure type” was categorized in accordance with Section 3.4.1. All of the tests except those for the HYN weld produced unstable fractures before any visible extension from the fatigue crack, i.e. Jc category. The HYN weld test produced Jeot category.

Figure 14 shows typical load-CMOD plots for the J category and Figure 15c shows the corresponding opened up fracture displaying the fatigue crack. The load-CMOD curves were not smooth in the “linear” portion compared to the QS rate tests, however this behaviour is not un-common the dynamic load rate tests [3].

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Figure 14 Load-CMOD curve for Test for 101TM-06A is a HAZ test.


Figure 15 Fracture face of 101TM-06A. The white arrow shows the fatigue crack region and no

visible crack extension occurred during the fracture toughness test.

3.5 Additional Fracture Toughness Testing

As observed in the test results presented in Appendix B for the QS tests, the HAZ toughness values for the 101TM ranged from 1481 (101TM-2H) to 137 (101TM-05) kJ m-2 and reflects the different failure categories ranging from J to Jeot c, respectively. It is possible that fracture toughness in the CGHAZ can result in a “wide” scatter band due to the probabilistic nature of fatigue crack interaction with “brittle metallurgical” phases present in CGHAZ microstructure.


However, the HAZ toughness values of HYN specimens did not display such a wide range and the failure type was J or Ju c, due to unstable fracture. In order to ascertain the relative location of the fatigue crack relative to the fusion line, post test metallography was performed on the previously mentioned specimens. To further examine the HAZ scatter, additional triplicate specimens were tested at -40oC to see if the toughness variation persists at this lower temperature.

For the testing done at dynamic loading rate, the results presented in Appendix C, show that except for HYN weld specimens all of the others displayed Jc behaviour. For HYN weld consistent J o

eot behaviour was observed. For this case it was decided to see if testing at -40 C would result in any change in the fracture behaviour.

The results are presented in Appendix D. Note that specimen 101TM-6 did not meet the BS 7448: Part 1 minimum fatigue crack length/size requirement and therefore is marked as failing the validity criteria requirements. However, this standard does not show an option of a narrow notch, such as EDM, whereas clause 7.4.5.1 in ASTM E1820 allows a minimum length of 0.6 mm for this size of specimen when using a narrow notch. All three tests for the 101TM HAZ produced failure category types Jc. After fatigue pre-cracking, while performing the compliance to estimate the fatigue crack length one of the specimens from HYN weld was overloaded. Therefore only two fracture toughness test were performed at dynamic loading rate at -40oC. These results are also presented in Appendix D. For both tests the fracture toughness behaviour have been categorized as Ju because they fractured during the test compared to the tests carried out on HYN weld at -30oC, however it needs to be noted that the fracture occurred just after the load-CMOD curve reached the maximum load plateau.


4 Discussion of Fracture Toughness Results

4.1 Quasi-static loading rate oThe results are presented in Appendix B for test temperature of -30 C and Appendix D for the

tests at -40oC. At quasi-static testing a range of fracture characteristics were observed depending on the microstructural region sampled by the fatigue crack. Generally the weld centerline specimens displayed consistent Jeot characteristics, while the HAZ samples of both welded panels displayed a range of fracture characteristics.

Except for one weld centerline (WCL) specimen all other tests resulted in ductile crack extension beyond the maximum load plateau at a test temperature of -30oC and were characterized as Jeot. Omitting the specimen which experienced Jc behaviour, the average J for the 101TM consumable was 424.41 kJ m-2 while the average J value for the HYN weld metal was 358 kJ m-2. It needs to be noted that the J values reported for the failure type categorized as Jeot is subjective on the end of test point on the load-CMOD curve (see for example Figure 8).

The fracture surface of the one specimen that produced unstable fracture (101TM-03) categorized as J is shown in Figure 16c can be compared with the other two fracture surface of this set presented in Figure 9. Stereoscopic examination of the fracture at magnifications up to 30X did not reveal any clear location of fracture initiation and therefore cannot be linked to features on the fracture surface.

Figure 16 Fracture face of the weld centerline specimen 101TM-03 which exhibited Jc. The white arrow shows the fatigue crack region and no visible crack extension occurred prior to the onset

of brittle failure during the fracture toughness test.

For the specimens from the HYN weld with the notch in HAZ the variation in the J values from 227 to 75 kJ m-2 is most likely due to the presence of brittle particles and phases, or close to the lower end of the ductile and brittle transition temperature


For the specimens from the 101TM weld with the notch in HAZ the variation in the J values were much larger, ranging from 1481 to 137 kJ m2.

The larger value was associated with a test being terminated after a maximum load plateau was reached, where as lowest value was associated with a test producing an unstable fracture event categorized as J . c

The “intermediate toughness” value was associated with fracture event categorized as J . u

Post test metallography was carried out to determine the fatigue crack location with respect to the weld fusion line for the two specimens that gave the highest value (101TM-2H) and the lowest value (101TM-05). The fatigue crack is closer to the fusion line in the mid-thickness region of specimen 101TM-05 than in specimen 101TM-2H (see Figure 17) and may contribute to the lower toughness. Both weld metal and coarse grain HAZ has non-homogeneous microstructure compared to base metal. The non-homogeneous microstructure results in variability in fracture toughness test results in the fracture transition temperature range [5].

For the test conducted at -40oC (see Appendix D), all three tests for the 101TM HAZ produced failure category types Jc as described in Section 3.5. These tests confirm that the ductile-to-brittle transition temperature for the HAZ is above -40oC.


a)


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Figure 17 Macrographs taken at the Fatigue Crack in 101TM-05 – Jc ( a) and 101TM-2H Jeot (b). The fusion line is highlighted by the arrow.

4.2 Dynamic loading rate

The results are presented in Appendix C for test temperature of -30oC.

When the loading rate (K rate) increased from about 1 to 1300 −smMPa all three specimens from 101TM with the fatigue crack in the weld metal consistently produced low toughness compared to the results at the QS loading rate. By contrast the HYN weld the toughness category did not change and remained as Jeot. However as described in Section 3.5 when the test temperature was reduced to -40oC unstable fracture occurred in the specimens just after the load-CMOD curve reached the maximum load plateau and therefore has a lower J values (see Appendix D) compared to test done at -30oC.


For the specimens from the 101TM weld with the notch in HAZ the tests consistently produced low J values with the failure category type Jc compared to test performed at QS loading rate. Therefore the effect of strain rate on lowering fracture toughness was displayed.

For the specimens from the HYN weld with the notch in HAZ the variation in the J values ranged from 101 to 42 kJ m-2 and all were categorized as Jc values. These ranges of J values are lower than the range obtained at QS rate (see Section 4.1).


5 Concluding Summary

• Both welds, made with either a Hobart 71HYN consumable or the Hobart 101TM consumable, displayed ductile behaviour, except one specimen from 101TM weld at quasi-static (QS) loading rate (K ~ 11 −smMPa ). When the loading rate was increased to the dynamic rate (K ~ 1300 −smMPa ), the 101TM weld clearly displayed fracture transition producing unstable fracture without visible extension of the fatigue crack (i.e. failure category J ) with J values below 100 kJ m-2

c . Whereas, for the 71HYN an increase in the loading rate did not produce fracture transition. However -40oC is likely the lower limit of the ductile behaviour since J characteristics were observed. u

• For the HAZ specimens with a fatigue crack close to the weld fusion line indicated that the fracture transition temperature is around -30oC under quasi-static loading rates. This transition temperature was confirmed by tests performed at -40oC which displayed Jc characteristics and low J energies. It needs to be noted that the microstructure in the CGHAZ does have metallurgical phases that have low fracture toughness and that can also contribute to variability. When the loading rate was increased to a dynamic rate, unstable fracture without visible extension of the fatigue crack (i.e. failure category Jc) was observed with J values below 155 kJ m-2.




Annex A Radiography Reports



Annex B Fracture Toughness Results at Quasi-static (QS) Loading Rate (at -30oC)



Annex C Fracture Toughness Results at Dynamic Loading Rate (at -30oC)



Annex D Fracture Toughness Results for Test Carried out at -40oC


References .....

[1] 6074 BMT FR Welding Procedure Development and Subsequent Fabrication of HSLA 65 Weldments

[2] Fracture Mechanics Toughness Tests, Part 2. Method for determination of KIc, critical CTOD and critical J values of welds in metallic materials, BS 7448: Part 2, BSI, (1997)

[3] Standard Test Methods for Measurement of Fracture Toughness, ASTM E1820-06.

[4] Draft Weld Fracture Test Standard (Proposed Annex to ASTM E1290), ASTM Revision 6, August 10, 1998.

[5] Standard Test Methods of Reference Temperature (To) Ferritic Steels in the Transition Range, ASTM E1921-05

[6] Bayley, C.J. and Mantei A., The Influence of Heat Input on the Fracture and Metallurgical Properties of HSLA-65 Steel Welds Fabrication conditions, Tensile, Impact and Microstructures DRDC Atlantic TM 2008-130. 2008


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1. ORIGINATOR (The name and address of the organization preparing the document. Organizations for whom the document was prepared, e.g. Centre sponsoring a contractor's report, or tasking agency, are entered in section 8.) BMT Fleet Technology Limited 311 Legget Drive Kanata, ON K2K 1Z8

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High Strength Low Alloy Steel with a minimum specified yield strength (SMYS) of 65 ksi (~450 MPa) (HSLA 65 - ASTM A945 grade 65) has drawn significant interest from Navy and Commercial Shipyards. This is primarily because of weight savings and weldability. It also is intermediate between high strength steels (HSS) with 50 ksi (~450 MPa) SMYS and HSLA 80 with 80 ksi (~550 MPa) SMYS. ASTM A945 grade 65 also specifies a tensile strength of 78-100 ksi (540-690 MPa) and a minimum transversely oriented Charpy impact energy of 70 ft-lb (95 J) at -40oC.

The present study demonstrates the influence of loading rate, temperature and crack location onthe fracture toughness. While the specimen geometry of the SENB is known to result inconservative fracture toughness values, it clearly demonstrates the deleterious influence ofwelding on the fracture behaviour in the coarse grained HAZ. In comparison to the coarsegrained HAZ, the two weld metals, 101TM and 71HYN, displayed higher toughness.

14. KEYWORDS, DESCRIPTORS or IDENTIFIERS (Technically meaningful terms or short phrases that characterize a document and could be helpful in cataloguing the document. They should be selected so that no security classification is required. Identifiers, such as equipment model designation, trade name, military project code name, geographic location may also be included. If possible keywords should be selected from a published thesaurus, e.g. Thesaurus of Engineering and Scientific Terms (TEST) and that thesaurus identified. If it is not possible to select indexing terms which are Unclassified, the classification of each should be indicated as with the title.) HSLA Steels, Fracture Toughness, Heat Affected Zone


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Fracture Toughness Characterization of HSLA Steel WeldmentsExecutive summary Fracture Toughness Characterization of HSLA Steel Weldments Nick Pussegoda; DRDC Atlantic CR 2008-178;