Weld World (2017) 61:463–471DOI 10.1007/s40194-017-0422-4

RESEARCH PAPER

Toughness evaluation of EB welds

Christopher Wiednig1 ·Norbert Enzinger1

Received: 3 June 2015 / Accepted: 4 January 2017 / Published online: 15 February 2017© The Author(s) 2017. This article is published with open access at Springerlink.com

Abstract For over a hundred years, Charpy impact testinghas been performed. It is one of the most frequently usedmaterial tests. Due to a very simple test setup, a huge wealthof experience and an enormous database of results, it is stillstate of the art to evaluate toughness of materials and welds.Modern welding technologies like laser welding or electronbeam welding (EBW) are characterized by a low heat input.The high energy density of these technologies results in verynarrow welds. These kind of joints are difficult to analyse bymeans of a standard Charpy test. The combination of weldcross section and properties lead in most cases to a frac-ture path deviation (FPD). Although the crack starts from anotch, it deviates during propagation into the heat-affectedzone or even in the base material. Therefore, the weld itselfis not tested and cannot be characterized. FPD is a well-known issue but so far little attention in standards has beenpaid to this complex topic. Hence, establishing a valid weld-ing procedure specification for beam welding proceduresmay imply difficulties. This work focuses on avoiding FPDin electron beam welds in soft-matensitic steel (1.4313) byusing standard and side notched Charpy impact specimens.

Recommended for publication by Commission IV - PowerBeam Processes

� Christopher Wiednigchristopher.wiednig@voestalpine.com

Norbert Enzingernorbert.enzinger@tugraz.at

1 Institute of Materials Science, Joining and Forming,Graz University of Technology, Graz, Austria

Valid toughness result had to be found for 20- and 100-mm-thick welds at −20 ◦C. Several test with three differentkind of specimen in two different heat-treated conditionswere carried out. It was found that beside the narrow seama major reason for FPD is the significant overmatching ofthe weld. A post weld heat treatment to reduce overmatch-ing facilitates testing but can decrease toughness. Adequateresults for qualification were found and EB welds reachedsufficient toughness at −20 ◦C. Toughness of EB welds wascompared to the toughness of conventional gas metal arcwelds (GMAW). Finally, a recommendation for adaptingthe toughness characterization for narrow and overmatchingseams is proposed.

Keywords (IIW Thesaurus) Ductility · EB welding ·Fracture tests · Notches · Toughness

1 Introduction

In 1884, L. v. Tetmajer introduced an impact test onnotched T-beams to qualify ingot steel (Flusseisen) insteadof wrought iron (Schweisseisen) [1]. He measured theenergy which was consumed by the specimen until breakingand described it as deformation energy. In 1901, AugustinGeorges Albert Charpy reported this work on the 3rd travel-ling exhibition of the International Association for materialstesting. In 1906, the Brussels congress agreed on the impor-tance of a testing procedure with notched bars and the sci-ence of fracture mechanics became an important approachin design [1, 2]. To date, the Charpy impact test is themost convenient procedure regarding ductility testing ofmaterials. The Charpy impact test is the most frequently per-formed mechanical material test next to tensile and hardness

464 Weld World (2017) 61:463–471

testing. Compared to fracture mechanics tests (CT, CTOD),it is easy to carry out but the result cannot be used for dimen-sioning. The measured impact energy is a qualitative valueused for comparison [3]. The execution and evaluation ofthe Charpy test is defined in the ISO standard 148-1:2010and ISO 9016:2011 [4].

Due to the higher probability of flaws in welds, ductilityis of significant importance. Unfortunately, proper tough-ness evaluation of welds is far more difficult than for ahomogeneous material. Because of the very local heat input,a weld offers various zones with different microstructuresand therefore different properties (metallurgical notch). Inthe work of Neves et al. [5] and Elliot [6], the influenceof varying welding parameters on the microstructure andsubsequently the toughness was investigated. They showedthat a change of heat input results in different weld seamtoughness values. Another problem in toughness evaluationof welds is the very local response of weld imperfections.Quintana et al. [7], Satho et al. [8] and Toyoda et al. [9]focus on local brittle zones which can affect crack initiationand growth in a drastic way. Cleavage can be initiated in anotherwise ductile material. This fact is supported by the sta-tistical distribution of ductility values which do not followa bell curve distribution [7] in contrast to other mechanicalproperties like the ultimate tensile strength.

In high energy density welding procedures, further prob-lems occur in addition to those mentioned before:

• Narrow fusion zones• High overmatching situations (UT SFZ � UT SBM)

Fracture path deviation (FPD) occurs as a result of thesecircumstances. FPD is a well-known problem in toughnesstesting of narrow welds. The crack deviates away from thenotch plane in the fusion zone (FZ), into the heat affectedzone (HAZ) and into the base material, providing veryhigh toughness values and leaving the FZ untested (seeFig. 1). FPD occurs in Charpy testing as well as in fracturemechanic testing. So far, this problem has been addressedin several publications [6, 8, 10–17] and different solutionsto overcome this problem have been presented. Crack tipopening displacement tests in different modes were per-formed in some of the previously mentioned publications[10, 11, 14–17]. Research on detecting the ductile-brittletransition temperature was done by Baryraktar et al. in [2,18, 19]. They were carrying out impact tensile testing toeliminate the influence of specimen bending and investi-gate pure crack initiation. Goldak et al. [12] instead wasusing a standard notched specimen but specimens weretaken in the longitudinal direction out of the weld with thenotch transverse to the weld direction (Cross weld Charpytest). Another approach to overcome FPD is to producea wider FZ by placing three welds next to each other

(3-Weld method) [15]. Other studies report testing Sidenotched specimens, where two additional notches are placedin the side planes of standard Charpy specimens [8, 13, 15].

Although a lot of research had been done over the last30 years, little attention has been paid from standardisationinstitutes. In the standard ISO 15614-11:2004 “Specifica-tion and qualification of welding procedures for metallicmaterials-Welding procedure test-Electron and laser beamwelding” [?], the toughness evaluation is left vague and canbe defined by the customer. However, the author is awareof certain cases where an implementation of electron beamwelding was rejected due to the absence of an appropriatelyaccepted toughness qualification of EB seams.

The intention of this paper is to present a solution for acertain case where the toughness evaluation for EB weldswas necessary and a standard Charpy test failed due to FPD(Fig. 1). In this work, toughness was investigated with stan-dard Charpy samples and side notched Charpy samples. Akey issue therefore is that side notched specimen are cheapto produce and can be tested with the same device as thestandard samples. Moreover, earlier studies have shown [13,15] that the absorbed energy of the side notched Charpyimpact test can be related to the absorbed energy of thestandard Charpy test.

2 Experiments

The present paper describes the toughness evaluation ofelectron beam welded soft martensitic steel (1.4313). Thecomposition is listed in Table 1. The ultimate tensilestrength of the base metal (UT SBM ) is 880 MPa, the hard-ness is around 285 HV and the impact energy at−20 ◦Cwasmeasured to be 131 J (arithmetic mean of nine samples).The minimum required toughness for the welded material at−20 ◦C is specified to be 27 J . For investigation, weld sam-ples with thickness of 20 and 100 mm were used. Weldingwas carried out with a probeam universal chamber machine,

Fig. 1 Fracture path deviation

Weld World (2017) 61:463–471 465

Table 1 Composition 1.4313

C Si Mn Cr Mo Ni N

max. 0.029 0.40 0.85 13.00 0.55 4.20 0.04

equipped with a 45 kW EB-gun. The samples were weldedin PA position, using a metal backing. Welding parametersare shown in Table 2.

In a previous test, the seam area with the lowest tough-ness values was estimated by varying the notch position.The lowest toughness values were found in the weld cen-treline. Consequently, all samples tested in this work werenotched in the seam centre. The Charpy samples were takenfrom the top and the middle of the weld. Prior testingshowed that for these two positions, the toughness values areequal and the scattering was less compared to samples takenfrom the root of the weld. The impact tests were carried outwith three different specimen geometries (see Fig. 2):

• Standard Charpy V-notch specimen (Std)• Side notched specimen (SN12) with a 10 x 12 mm cross

section• Side notched specimen (SN) with a 10 x 10 mm cross

section

Notches were machined with a standard Charpy notchmilling tool. The main notches have a depth of 2 mm,whereas the side notches were machined with a depth of1 mm. All notches were machined after etching the milledspecimens to place the notch exactly in the seam centre.Charpy energy in all cases was measured at −20 ◦C. A stan-dard Charpy impact hammer with a maximum energy of

Table 2 Welding Parameters

Thickness (mm) Power (kW) Welding speed (mm/s)

20 9.0 12.0

100 37.5 2.1

300 J was used. Specimens were tested in the as welded(AW) and in post weld heated treated (PWHT) conditions.The post weld heat treatment was performed with a heatingrate of 50 ◦C/h to 580 ◦C. After keeping the specimen 1 hat this temperature, they were cooled down to room temper-ature at 50 ◦C/h. For every testing condition at least threesamples were tested. According to the standard, the meanvalueof the tested samplesgives the absorbed energy. Sampleswhich did not break were excluded from calculation.

To compare toughness of EB welds with toughnessof tungsten inert gas (TIG) and flux-cored metal activegas (FC-MAG) welds, side notched Charpy samples weremachined out of TIG and FC-MAG welded plates. Thesewelds were carried out in 25-mm-thick plates with double-V preparation. In Fig. 3, cross sections of these welds withinscribed Charpy specimen positions are shown. A postweld heat treatment was carried out for the FC-MAG weldsto maximise toughness. TIG welds were not heat treatedafter welding. Charpy test conditions were the same as forthe EBW welds.

The full test matrix is shown in Table 3.

3 Results

The following section presents the results of the investi-gations. Outcomes are divided into three subsections. Inaddition to toughness testing, the microstructure and thehardness of the welds were also investigated. The resultsof the Charpy impact test for the 20-mm samples and the100-mm samples are presented in separate subsections.

3.1 Microstructure

Prior to toughness testing the microstructure of the weldswas investigated. In Fig. 4, half of the FZ and the HAZ isdisplayed. The seam width in the 20-mm weld sample isapproximately 2 mm. A fine grained heat affected zone canbe recognised next to the fusion line with a width of lessthan half a millimetre. The 100-mm sample shows a FZ with

Fig. 2 Specimen geometries,cross-section in the notch plane

(a) (b) (c)

466 Weld World (2017) 61:463–471

Fig. 3 GMAW samples

(a) (b)

Table 3 Charpy test matrix

Std SN12 SN

(a) 20 mm and 100 mm EB welds

AW � � �PWHT � � �BM � – �

(b) TIG and FC-MAG welds

AW TIG TIG TIG

PWHT FC-MAG FC-MAG FC-MAG

BM � – �

approximately 6-mm width. In this case, a fine grained heataffected zone of about 2 mm can be recognised. A coarsegrained zone could not be identified. The dendritic grainsin the FZ in the 20 mm seam are significantly smaller thanthe grains in the 100 mm seam. The average dendrite lengthin the 100 mm FZ is 630 μm, compared to 140μm in the20 mm FZ.

Since Charpy impact tests were carried out on the aswelded and in post weld heat treated conditions, a Vickershardness test was performed inbothconditions. Hardness lines

were measured across the FZ and the HAZ, at half seamdepth. Spacing between the imprints was set to 0.7 mm.Figure 5 shows the impact of the PWHT. Hardness andconsequently the overmatching, between base metal andweld seam, could be decreased. The seam hardness afterthe PWHT was below 350 HV, as required by ISO 15614-1:2004. The missmatch ratio M is calculated as M =HVFZ/HVBM . Average mismatch ratio between the basematerial and the FZ in the as welded condition was MAW =1.31 and for the PWHT condition MPWHT = 1.16.

The microstructures of the TIG and FC-MAG welds didnot show any irregularities and appear like typical multi-pass arc welded microstructures. The hardness of the FZin the TIG welds was measured with 315HV10 and with308HV10 in the FZ of the post weld heat treated FC-MAGwelded cross sections.

3.2 Toughness of 20 mm welds

The average impact toughness of the 20 mm welds is dis-played in Fig. 6. Values which are not valid according thestandard (specimen is not fully broken; fracture deviates outof the FZ) are marked with an asterisk.

For specimens in AW conditions with standard notch(AW-Std), no regular values could bemeasured due to fracturepath deviation (see Fig. 1). Toughness values are extraordinarily

(a) (b)

Fig. 4 Microstructure of FZ and HAZ

Weld World (2017) 61:463–471 467

(a) (b)

Fig. 5 Hardness lines, as welded (yellow) and PWHT (red)

Fig. 6 Absorbed energy of the 20 mm specimen (asterisk-not validvalues)

high and exceed the values of the base material. Lateralexpansion exceeded 1.5 mm for all specimens in this group.

The SN12 specimen with the side notches (AW-SN12)did not return valid results either. Samples did not fullyfracture and inner FPD (fracture plane is curved within thenotches, see Fig. 7) occurred. The fracture path actually

crosses HAZ and base material. The value for the AW-SN12group also exceed the value of the base material.

Specimen with standard 10 x 10 mm geometry andside notches (AW-SN) returned valid values. Some of thespecimen also show this inner FPD but others broke witha flat brittle looking fracture. The absorbed energy reached80 % of the side notched base metal specimens (BM-SN).Lateral expansion did not exceed 0.5 mm.

The situation for the post weld heat-treated specimen ismore obvious. For all tested groups (Std, SN12 and SN),valid results could be measured. The PWHT-Std specimenfractured with a straight fracture plane with 100 % crys-talinity; lateral expansion was 0.5 mm. Both side notchedgeometries (SN and SN12) fractured also in a brittle mannerbut with smaller lateral expansion (<0.5 mm).

3.3 Toughness of 100 mm welds

The average impact toughness of the 100 mm welds is dis-played in Fig. 8. Values which are not valid according to thestandard are marked with an asterisk.

The standard specimen in AW conditions offer no validresult since no sample broke fully. The fracture is clearlyductile with distinctive shear lips and the fracture path

Fig. 7 Inner fracture pathdeviation, AW 20 SN12 andAW 20 SN

(a) (b)

468 Weld World (2017) 61:463–471

Fig. 8 Absorbed energy of the 100-mm specimen (asterisk-not validvalues)

did not leave the FZ (see Fig. 9). Lateral expansion wasmeasured with up to 2 mm.

The specimen with the side notches in AW conditions(AW-SN12 and AW-SN) returned valid results according thestandard; all samples broke fully and no FPD occurred. Thefracture planes are flat with a fully crystalline appearanceand show a lateral expansion of approximately 0.2 mm.

In the PWHT-Std group, two samples did not break fully(values around 180 J, lateral expansion >1.5 mm, notica-ble shear lips). The other PWHT-Std specimen experiencedfracture, with a rough crystalline surface. Small shear lipscan be recognised. The lateral expansion was 0.9 mm.

All side notched specimen (SN12 and SN) in the 100 mmPWHT group behaved comparable to the PWHT sidenotched specimens in the 20 mm group. All specimen returnvalid values. The fracture surface appears brittle and flat withlateral expansion around 0.4 mm. No shear lips were found.

3.4 Comparison EBW to GMAW

For comparison, the results of the 100 mm EB welds wereused, because in this tests the fracture path remained in theFZ (see Fig. 9b). To compare the maximum toughness values

of the different processes, the EB and TIG welds weretested in as welded conditions, whereas the FC-MAG weldswere tested in PWHT conditions. Results are displayed inFig. 10.

The TIG welds offer a superior toughness for all spec-imen geometries. Neither the Std nor the SN and SN12specimens broke fully (values marked with an asterisk).All specimens showed fully ductile behaviour. Shear lipsappeared only on the Std specimens. Lateral expansion wasmeasured on all specimen with at least 0.7 mm.

The absorbed energy of the FC-MAG welds is lowercompared to the EBW welds. The Std FC-MAG specimenbroke fully, showing noticeable shear lips. Lateral expan-sion was below 0.5 mm. The fracture surface appeareddull (ductile) with some crystalline (brittle) areas. Charpybehaviour was as expected from former tests. The SN12and SN specimen did not show shear lips or lateral expan-sion. The fracture surface appeared equal to the Std speci-mens. FC-MAG Charpy values do not show a high decreaseof absorbed energy when using side notched specimens fortesting.

4 Discussion

In several publications, the advantages of EBW, especiallyin thick-walled welding, is mentioned [20–27]. A low over-all heat input and the high cooling rate of this process isa significant characteristic. These circumstances lead to anarrow FZ and a small HAZ with a fine-grained microstruc-ture. A coarse-grained zone was not detectable (see Fig. 4).Generally, fine microstructures provide good toughness val-ues. So in EBW in comparison to most other weldingprocesses, the HAZ is not the critical area with regard toductility [10]. Previous research by Bezensek et al. [10]demonstrated that the lowest ductility in beam welding isthe weld centerline. The high cooling rates also lead to highhardness in the FZ and in the HAZ (see Fig. 5) which isnot favourably regarding ductility. Other studies have shown

Fig. 9 AW-Std 100 mm,Fracture surface and fracturepath

(a) (b)

Weld World (2017) 61:463–471 469

Fig. 10 Absorbed energy of the 100 mm EBW, TIG, and FC-MAGspecimen

that a narrow area with high hardness does not generallylead to toughness problems (see [22]).

However, high hardness in the FZ causes a high mis-match ratio between base metal and FZ [28]. As our findingsshow, this overmatching is essential for FPD; for as weldedspecimens FPD was often observed, whereas in PWHT con-ditions, nearly no FPD occurred. In the literature [6, 11, 12,14], this is explained with the so called Plastic Hinge mech-anism. In narrow welds, the stress field in front of the cracktip extends to the base material. Due to lower strength ofthe BM, UTS in the base metal is reached before the weldmetal reaches its UTS or even yield strength. This leads toa ductile collapse in the base material before a crack can beinitiated in the FZ. In this case, the Charpy value does notrepresent the toughness of the weld, only the energy whichis absorbed by the joint. The author of this paper interviewedseveral material testing institutes were FPD was interpretedas high ductility of the weld.

In 1977, Goldak et al. [12] introduced the crossweld Charpy test (CWCT) were the Charpy specimen ismachined alongside the weld and the notch is perpendicularto the welding direction. In this test, the crack has to pass theFZ but also, depending on the width of the weld, the HAZand the BM. Consequently, the measured Charpy value hasno meaning, with respect to the toughness in the FZ, but theductile brittle transition temperature can be estimated witha microscopic analyse of the fracture surface. Furthermore,because of the anisotropic microstructure of a weld, impacttests along and transverse to the welding direction are notdirectly comparable. Despite these limitations, some test-ing institutes are interpreting the CWCT measurements ofabsorbed energy as toughness values.

The notched Charpy specimen, applied in this work cansuppress FPD in most cases. The side notches cause a

multi axial stress state and fracture is forced within thenotched plane. Thus, absorbed energy becomes significantlower compared to the standard specimen. Shear lips andlateral expansion are suppressed by the side notches andthe specimen fractures in a brittle manner. These results aresupported by Hagihara et al. and Satoh et al. [13, 29]. Toconclude, the side notched Charpy impact test returns veryconservative toughness results.

The quantitative influence of the side notches on thetoughness can be seen from Figs. 6 and 8. The side notchedbase material BM-SN reached about 60 % of the BM-Stdvalue. Taking into account that the SN specimen only hasa cross-section of 64 mm2, the specific toughness value ofthe SN-BM reaches 75 % of BM-Std. These findings corre-late with values found by Hagihara et al. for a laser weldedultra-fine grained steel in the upper shelf area (see Fig. 5 in[13]). This correlation between SN and Std specimen wasalso found for the PWHT samples. As stated in [13], theabsence of shear lips in the side notched specimen is themain reason for the lower toughness value.

Although, all side notched specimens, return a valid resultand fractured in a brittle manner; low lateral expansion,flat fracture surface and crystalline appearance, the impacttoughness, in this case, is above the required minimum of 27 J.Additionally, ductile fractured areas were found on the frac-ture surfaces of all tested EBW specimen. Even the macro-scopic appearance of brittle fractures of the side notchedsamples turned out to be partial ductile (see Fig. 11).

Due to the multi axial stress state, the side notchedspecimens show lower toughness values compared to thestandard specimens (cp. [13, 15]). For that reason, it canbe assumed that a side notched value is a conservativetoughness estimation and a reliable confirmation that thetoughness exceeded the required 27 J at −20 ◦C.

The comparison of side notched EB, TIG and FC-MAGwelds showed that in the case of EBW very conservativevalues are returned. This can be explained by the fact that inthe FZ centre, the dendrites meet during solidification fromboth sides. Impurities and segregation will preferentiallyform in this area. In SN Charpy specimen, the fracture planeis forced to run along this area and therefore the toughnessvalues are strongly decreased. In multi-pass GMAW welds,the fracture path crosses several solidification lines. There-fore, the difference of Std and SN specimens is lowercompared to EB welds. Hence, the “in operation” toughnessof EBW welds can be assumed higher than in this test. Ourfindings clearly show that in this case, EBW joints have suf-ficient toughness and are superior to FC-MAG joints. Thiscomparison shows also that the fracture mode is indepen-dent of the specimen mode. The TIG fracture was ductile for

470 Weld World (2017) 61:463–471

Fig. 11 Fracture surface ofstraight fractured specimen withvalid values

all specimen geometries. Also, the FC-MAG fracture was amixed brittle/ductile fracture in all tests.

A weakness of the side notched Charpy test may be iden-tified when values around 27 J for EB welds are returned.Besides the assumption that this test is conservative, a com-parison to SN specimens machined out of conventionalwelded joints can help to judge fitness for purpose. If this isnot possible, fracture mechanics methods should be used.

5 Conclusion

The present work shows that toughness evaluation of nar-row and overmatching weld seams is a challenge. In orderto be able to obtain valid results, the standard test geometryhas to be adapted. Therefore, the measured results cannot becompared directly with standard Charpy values. Neverthe-less, a clear statement can be given about the utilisation of aseam according to the welding procedure qualification stan-dard. When replacing conventional arc welding procedureswith EBW, Charpy tests with side notched specimen forboth procedures allow a cheap and easy to perform toughnesscomparison.

The results of this work may be summarised as follows:

• Toughness evaluation of 2-mm-wide welds with stan-dard notched Charpy specimen is not possible due tofracture path deviation

• Also with side notched specimens fracture path devia-tion within the fracture plane is possible

• All kinds of FPD indicate considerably higher tough-ness value, than are actually present

• Side notched specimen without FPD show macroscopicbrittle fracture but still return appropriate toughnessvalues due to microscopic ductile fractured areas

• Decreasing the mismatch ratioofthe joint applying a postweld heat treatment leads to a valid testing for standardCharpy specimen as well as for side notched specimens

• Side-notched specimens enable a suitable toughnesscomparison of high-energy-density welding processesto conventional welding processes

• Toughness of EB welds in 1.4313 steel is superior toFC-MAG welds

6 Outlook

This paper provides a recommendation for toughness eval-uation if standard toughness testing fails. However, somequestion are still unanswered and will be investigated inongoing work.

• Fracture surface investigation of side notched Charpyspecimen by means of scanning electron microscope.

• More sophisticated methods (fracture mechanics test:CT, CTOD) are needed to verify and quantify toughnessresults.

• Future research might concentrate on finding a standardbased toughness evaluation method for high energywelding procedures.

Acknowledgments Open access funding provided by Graz Univer-sity of Technology.

Weld World (2017) 61:463–471 471

Open Access This article is distributed under the terms of theCreative Commons Attribution 4.0 International License (http://creativecommons.org/licenses/by/4.0/), which permits unrestricteduse, distribution, and reproduction in any medium, provided you giveappropriate credit to the original author(s) and the source, provide alink to the Creative Commons license, and indicate if changes weremade.

References

1. Blumenauer H (2001) 100 Jahre Kerbschlagbiegeversuch nachCharpy. Materialwissenschaft und Werkstofftechnik 32(6):506–513

2. Bayraktar E, Kaplan D, Schmidt F, Paqueton H, Grumbach M(2008) State of art of impact tensile test (ITT): Its historicaldevelopment as a simulated crash test of industrial materials andpresentation of new ductile/brittle transition diagrams. J MaterProcess Technol 204(1-3):313–326

3. Macherauch E, Zoch H-W (2011) Kerbschlagbiegeversuch.Vieweg+Teubner Verlag, Wiesbaden, pp 281–288

4. ISO. ISO 148-1 - Kerbschalgbiegeversuch nach Charpy (2010)5. Neves J, Loureiro A (2004) Fracture toughness of welds - effect

of brittle zones and strength mismatch. J Mater Process Technol153-154:537–543

6. Elliott S (1984) Electron Beam Welding of C–Mn Steels–Toughness and Fatigue Properties. Weld J:8–16

7. Quintana MA, Babu SS, Major J, Dallan C, James M (2010) WeldMetal Toughness: Sources of Variation. In: 8th ASME Interna-tional Pipeline Conference 2010, Calgary, Canada, pp 559–608

8. Satho K, Tozoda M (1985) Effect of mechanical heterogeneity onfracture toughness evaluation of steel welds, vol 74. Transactionsof the Japan Welding Society

9. Toyoda M, Praunseis Z (2000) The transferability of fracture-mechanics parameters to fracture performance evaluation of weldswith mismatching. MATERIALI IN TEHNOLOGIJE. 34(6)

10. Bezensek B, Hancock JW (2007) The toughness of laser weldedjoints in the ductile–brittle transition. Eng Fract Mech

11. Cam G, Erim S, Yeni C, Kocak M (1999) Determination ofmechanical and fracture properties of laser beam welded steeljoints. Welding Journal-New York

12. Goldak JA, Nguyen DS (1977) A fundamental difficulty inCharpy V-notch testing narrow zones in welds, pp 119–125. WeldJ(Miami, FL, US)

13. Hagihara Y, Tsukamoto S, Otani T (2005) An Evaluation Methodfor Laser Weld Metal Toughness by Side-Notched Charpy Test.Welding in the World 49(7-8):21–27

14. Moltubakk T, Thaulow C, Zhang ZL (1999) Application of localapproach to inhomogeneous welds. Influence of crack positionand strength mismatch. Eng Fract Mech 62:445–462

15. Takashima APY, Minami F (2009) Fracture Toughness Evaluationof Laser Beam-Welded Joints of 780 MPa-Strength Class Steel.Welding in the World 53(11-12):R304–R311

16. Takashima Y, Nishi T, Shoji H, Ohata M, Minami F (2014) Evalu-ation method for Charpy impact toughness of laser welds based onlateral contraction analysis. Welding in the World 58(3):289–295

17. Zhang H, Zhang Y, Li L, Ma X (2002) Influenceofweldmis-match-ing on fatigue crack growth behaviors of electron beam weldedjoints. Materials Science and Engineering: A 334(1-2):141–146

18. Bayraktar E, Hugele D, Jansen JP, Kaplan D (2004) Evaluation ofpipeline laser girth weld properties by Charpy (V) toughness andimpact tensile tests. J Mater Process Technol 147(2):155–162

19. Bayraktar E, Kaplan D, Buirette C, Grumbach M (2004) Applica-tion of impact tensile testing to welded thin sheets. J Mater ProcessTechnol 145(1):27–39

20. Various (2012) International Electron Beam Welding Conference- Lectures of the 2nd IEBW Conference Aachen March 26-30,2012. In: DVS Berichte. Dusseldorf, Germany. DVSMedia GmbH

21. Kihara TH, Minehisa S, Sakabata N, Shibuya T (1984) HighPower Electron Beam Welding of Thick Steel Plates–Method forEliminating Beam Deflection Caused by Residual Magnetism.Weld World 22(5):126–134

22. Wiednig C, Lochbichler C, Enzinger N, Beal C, Sommitsch C(2014) Dissimilar Electron Beam Welding of Nickel Base Alloy625 and 9 % Cr Steel. Procedia Engineering 86:184–194

23. Bohm S (2014) The electron beam as a welding tool: Germanstudy on the state of the art and future requirements. In: Interna-tional Institute of Welding, Commission IV - Power Beam Pro-cesses Annual Assembly 2014. International Institute of Welding,Seoul, Korea

24. Dilthey U, Dobner M (1996) Untersuchung des WerkzeugsElektronenstrahl - Teil 3: Koordination von Strahlstrom undFokussierung. Schweissen und Schneiden 48(10):770–785

25. Sun Z, Karppi R (1996) The application of electron beam weldingfor the joining of dissimilar metals: an overview. J Mater ProcessTechnol 59(3):257–267

26. Dobeneck D (2007) An International History of Electron BeamWelding, 1edn. Pro-beam AG

27. Schultz H (2004) Elektron beam welding. Woodhead PublishingLtd, Cambridge

28. Pavlina EJ, Van Tyne CJ (2008) Correlation of Yield Strength andTensile Strength with Hardness for Steels. Journal of MaterialsEngineering Performance 17:888–893

29. Satoh K, ToyodaM, Nohara K, Takeda S, NayamaM (1982)NotchToughness Evaluation of HT80 Electron Beam Weld Metal byCharpy Test. Journal of the Japan Welding Society 51(8):679–686