JOURNAL OF THE JAPAN WELDING SOCIETY
Online ISSN : 1883-7204
Print ISSN : 0021-4787
ISSN-L : 0021-4787
Volume 36 , Issue 3
Showing 1-14 articles out of 14 articles from the selected issue
  • Shinichi Kaku
    1967 Volume 36 Issue 3 Pages 189-195
    Published: February 25, 1967
    Released: August 05, 2011
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  • Itsuhiko Sejima, Tatsuji Wada
    1967 Volume 36 Issue 3 Pages 196-202
    Published: February 25, 1967
    Released: August 05, 2011
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  • Rei Sakurai, Giichi Nakano
    1967 Volume 36 Issue 3 Pages 203-212
    Published: February 25, 1967
    Released: August 05, 2011
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  • Jyunjiro Muraki, Takayoshi Ishiguro, Hikojiro Yokota
    1967 Volume 36 Issue 3 Pages 213-221
    Published: February 25, 1967
    Released: August 05, 2011
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    Generally speaking, the fatigue strength of welded joints of high tensile strength steels is not so superior to that of ordinary mild steels, but any reasonable explanation has not been reported yet.
    There are many factors reducing the fatigue strength of welded joints, for example, undercutting, geometrical effect of reinforcement, residual stresses and microstructural change in weld heat affected zone, etc.
    Among those many factors, we intended to study the fatigue strength of weld heat affected zone of various kinds of steels range from 60 Kg/mm2 ultimate strength steel to 80 Kg/mm2 ultimate strength steel. In order to reproduce the microstructure of various portions in weld heat affected zone, test specimens were subjected to thermal cycles with synthetic thermal cycle apparatus, seting peak heating ternperatures of these thermal cycles 640, 740, 850, 1000 and 1350°C, corresponding to certain points at varied distances from weld bond, then machined to smooth fatigue specimens.
    Amsler's high frequency Vibrophores fatigue testing machine was used for evaluating the fatigue strength of all specimens. Fatigue limits of each specimens subjected to thermal cycles are nearly equal or higher than that of base metals in every case, and are proportional to hardeness or tensile strength of various portions in weld heat affected zone.
    These results mean that the reduction of fatigue limits of butt-weld joints of high tensile strength, steels is not due to the structural change, hardening or softening in weld heat affected zone.
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  • Zinkichi Tanaka, Tadayoshi Obata
    1967 Volume 36 Issue 3 Pages 222-228
    Published: February 25, 1967
    Released: August 05, 2011
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    The stress relief heat treatment process may be divided into three stages. The first stage is heating process, in which residual stress decreases due to temperature dependence of Young's modulus and yield strength. The second stage is isothermal relaxation process, in which residual stress decreases due to creep. The third stage is cooling process, in which residual stress increases due to temperature dependence of Young's modulus.
    In order to study the relation between isothermal relaxation and creep, a short time creep test and an isothermal relaxation test were carried out using HT60.
    Creep curves might be represented by equation (I). Applying the strain hardening theory, creep rate-stress relation was expressed by equation (5). Combining this with relaxation condition, equation (7) indicating time-residual stress relation was obtained.
    The calculated data by equation (7) were in good agreement with experimental data for test temperatures 500°C and 550°C. But for 600°C and 650°C, the calculated data differed from experimental data. In this case, plastic strain rate in relaxation became independent of initial strain amplitude and dependent only on stress amplitude. So the creep rate theory of steady state creep indicated by equation (8) was adopted. Combining this with relaxation condition, equation (9) indicating time-residual stress relation was obtained for 600°C and 650°C. This time, calculated data coincided sufficiently well with experimental data. For HT60, the estimation of residual stress might be done with sufficient accuracy using equation (7) for 500°C, 550°C and equation (9) for 600°C, 650°C.
    The temperature dependence of creep rate might be expressed by equation (10). Applying this to relaxation phenomena, equations showing time-residual stress relation at various temperature were obtained in the forms of equation (12) and (13). The values of constants in these equations could be obtained only from creep and high temperature tension test data. So isothermal relaxation curves for arbitrary temperature and initial strain could be calculated from creep and tension test data.
    The limiting temperature at which the applicable equation differed from one to another was not precisely confirmed in this study. It might be dependent on steel properties and high temperature properties.
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  • Kunihiko Satoh, Sigetomo Matsui
    1967 Volume 36 Issue 3 Pages 229-237
    Published: February 25, 1967
    Released: August 05, 2011
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    As the factors which affect the contraction process of a weld joint in arc welding, three factors are considered; (1)t welding heat input (Q), (2) plate thickness (h) and (3)gauge length (l) over which the contrac. tion is measured. Besides, effect of transformation expansion of weld during cooling should also be considered.
    In the prior works (parts 1-3), keeping welding heat input constant, effects of the other factors on the contraction were investigated. In the present report, effect of welding heat input on contraction was considered. Knowledge on contraction process of a weld joint has advanced to the following point: The quantity of critical plate thickness (hcr) can be taken as a parameter concerning the mutual effect of heat input (Q) and plate thicness (h), and the critical plate thickness (hcr) is proportionate to the square root of heat input (Table 1). Contraction progresses with different processes according to whether plate thickness (h) is larger than the critical phate thickness (hcr) or not (Fig. 6). In the case of h≥hcr contraction progresses to the final value through the first, the second and the third steps. The final contraction is proportionate to hcr (or to √Q) independently of plate thickness (h). In the case of h<hcr, on the contrary, only the third step appears in contraction process. The final contraction is proportionate to the square of hcr (or to Q) and increases with decrease in plate thickness (Fig. 5).
    Successful agreement was observed between experimental results and calculated ones (Figs. 10-13). However a little difference between them was observed in the effect of moving of heat source which was neglected in the analysis (Fig. 15).
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  • Shigeo Hasebe
    1967 Volume 36 Issue 3 Pages 238-251
    Published: February 25, 1967
    Released: August 05, 2011
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    At first the effect of niobium addition on the mechanical and welding properties of low carbon steel small plate as rolled was investigated in laboratory as compared with vanadium addition. Then the properties of large plate as rolled in works of killed and semikilled steel containing niobium, especially the toughness of the plate were investigated.
    1) To increase the tensile strength of low carbon steel plate as rolled by 5-6 Kg/mm2 the addition of about 0.05% vanadium is a little more effective than the one of niobium equivalent to 0.05% vanadium with respect the impact properties of the plate. But niobium increases the tensile strength of the plate more than vanadium in the range up to about 0.05% individually.
    2) It may be said that we have had for the first time nonheat-treated weldable high tension steel plate with both tensile strength over 60 kg/mm2 and structure consisting of ferrite and pearlite. The addition of both niobium and vanadium to the low alloy steel containing lower content of carbon further improves the impact properties of the plate as rolled.
    3) The high tension steel plate containing niobium has a higher yield ratio as well known and can be welded easily, but the absorbed energy in Charpy impact test of the plate is comparativelly small and this phenomenon is especially controversial in the transverse direction to rolling one. The absorbed energy in Charpy test being small is probably due to the higher tensile strength for the structure consisting of ferrite and pearlite. But the transition temperature of fracture surface in Charpy test is comparativelly good and not worse than the one of low tension steel plate containing no niobium.
    4) The new evaluation criteria for brittle fracture characteristic of structural steels for low temperature application (recently proposed by Japan Welding Engineering Society) adopt the evaluation method by transition temperature of fracture surface in Charpy test, rejecting the one by the absorbed energy. The transition temperature of fracture surface of high tension steel plate containing niobium being good, the small absorbed energy of the plate seems not to be a large defect.
    5) The fracture tests on a larger scale such as press-notch crack starter test (wedge impact test) and double tension test of the plate were also carried out. The results of these fracture tests matched the transition temperature of fracture surface in Charpy test and supported the evaluation results by the above new criteria.
    6) The tensile strength of carbon steel plate containing no niobium coiled immediately after hot-rolling is lowered in general due to the pearlite structure being globularized during slow cooling from coiling temperature, but the one of coiling plate containing niobium is little lowered by precipitation hardening of niobium carbo-nitride. The method to apply globularizing annealing to uncoiled flat plate after hot-rolling may be suitable to improve the impact properties of large plate unable to be coiled.
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  • Kohei Ando
    1967 Volume 36 Issue 3 Pages 252-260
    Published: February 25, 1967
    Released: August 05, 2011
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    Christensen et al. show that the average temperature of the molten pool reaches as high as 1.5-2.0 times the fusion temperature of the base metal when the welding arc is assumed as a moving point heat source.
    In the present paper the molten pool is divided into two parts and the average temperatures are calculated as 1.5-2.95 times and 1.5-1.345 times the fusion temperature for the leading and trailing part of the molten pool respectively as the operational parameter n varies from zero to infinity. The border line of the two parts is MM' in Fig. 2.
    It is well known that the temperatures of the molten pool are estimated as 1800-2200°C for steel from the standpoint of metallugical equilibrium. It is interesting to compare those values to the above calculated temperatures of the trailing part, given as 2300-2060°C for steel.
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  • 1967 Volume 36 Issue 3 Pages 290-303
    Published: February 25, 1967
    Released: August 05, 2011
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  • 1967 Volume 36 Issue 3 Pages 304-317
    Published: February 25, 1967
    Released: August 05, 2011
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  • 1967 Volume 36 Issue 3 Pages 318-329
    Published: February 25, 1967
    Released: August 05, 2011
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  • 1967 Volume 36 Issue 3 Pages 330-346
    Published: February 25, 1967
    Released: August 05, 2011
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  • 1967 Volume 36 Issue 3 Pages 347-363
    Published: February 25, 1967
    Released: August 05, 2011
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  • 1967 Volume 36 Issue 3 Pages 364-380
    Published: February 25, 1967
    Released: August 05, 2011
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