Thus, slow strain rate tests. (SSRT) were performed in a simulated degraded concrete pore solution (pH = 8) with high chlorides content ([Cl. -. ]=17,5g.L. -1. ).
Stress corrosion cracking and hydrogen embrittlement behaviour of high strength duplex and austenitic stainless steels in simulated concrete pore solution B. Grimault a,b, L. Gaillet a, M. Drissi-Habti a, T. Chaussadent a, E. Chauveau b and M. Mantel b a
Institut Français des Sciences et Technologies des Transports, de l’Aménagement et des Réseaux, France b
Ugitech, France
Summary Stainless steels are playing an increasingly role in civil engineering, especially in coastal and marine constructions, where the concentration of chloride ions is higher. Their use as concrete reinforcement rebar is an interesting example of that idea. However, although some studies about the use of stainless steels in prestressing are emerging, no realistic realization has been performed. This might be due to technical lock as the high mechanical strength needed for prestressing seems to be difficult to achieve with stainless steels. Moreover, even if these mechanical characteristics could be obtain, the stress corrosion cracking behavior of this kind of grade in simulated concrete pore solution is unknown. In this context, the aim of this work is to study, regarding hydrogen embrittlement (HE) and chloride stress corrosion cracking (SCC), two stainless steel grades (austenitic and duplex) which were formed to present high strength. Hydrogen embrittlement resistance of the two grades was studied thanks to the ISO 15630-3 standard stress corrosion cracking test where a corrosive solution of ammonium thiocyanate NH4SCN was used. A more realistic environment was used to evaluate their susceptibility to SCC. Thus, slow strain rate tests (SSRT) were performed in a simulated degraded concrete pore solution (pH = 8) with high chlorides content ([Cl-]=17,5g.L-1). In the two cases, the results are presented and discussed with respect to the microstructure of the two stainless steel grades.
1
Introduction
Generally, steels used for concrete prestressing are carbon steels due to their low cost. The alkalinity of concrete and the chloride protection barriers able the passivity of the prestressing reinforcements and thus their corrosion protection. However, in some places (marine environment or places subjected to de-icing salt or road pollution), two phenomena can lead to prestressing reinforcement depassivation: (i) the diffusion of carbon dioxide in the pores of the concrete resulting in a decrease in pH of the pore solution, initially close to 13,2 [1], by carbonation, (ii) the presence of chloride ions coming adsorbed on the metal surface and destroy the passivity. This corrosion is considered as the main cause of deterioration of concrete structures such as bridges or buildings. In the case of prestressing, as the reinforcements are 1
under tension (= 80%Rm), corrosion is characterized by two phenomena: chloride induced stress corrosion cracking and hydrogen embrittlement [2]. For an application in prestressing, high strength stainless steels, with an ultimate tensile stress of about 1860MPa, are needed [3] and could be obtained by cold working. During this step, for austenitic and duplex stainless steel, strain induced martensitic transformation could appeared [4] and this body centered cubic (bcc) phase is known to be associated with a corrosion resistance reduction [12,5]. Ghosh and al. [6,7] show that cold worked 304 stainless steel is lower resistant to chloride induced stress corrosion cracking than solution annealed one due to the presence of strain induced martensite. According to the authors, it presence results in higher SCC susceptibility as dissolution rate of martensite is much higher than that of austenite. Some authors [8] show that this higher dissolution rate could be linked to the dislocation pile-ups associated with the presence of martensite laths. In this way, strain induced martensite could have an indirect effect on stress corrosion cracking resistance. In the case of hydrogen embrittlement, the effect of martensite is clearer. It presence favors the hydrogen diffusion and the area embrittled [9,10]. In this context, some studies about the use of stainless steels as prestressing reinforcement are emerging [11,12]. The stainless steel grades belong the austenitic family and present ultimate strength close to 1500MPa, which is below the specifications needed for prestressing use. In these studies, the adverse effect of strain induced martensite on SCC and HE resistance is enhanced. The work describes in this paper compare the stress corrosion cracking and hydrogen embrittlement behavior of a high strength austenitic and duplex stainless steel with mechanical characteristic compatible with prestressing. The effect of microstructure is studied. Furthermore, the effect of strain induced martensite is discussed. 2
Experimental procedures
2.1 Materials Two stainless steel grades were considered in this study. The “lean duplex” 1.4362 and the austenitic 1.4310. The composition of these two grades is given in table 1. After hot rolling, the two stainless steels were cold worked to a diameter of 3.25 mm. The reduction of area associated with drawing is about 90% for the lean duplex 1.4362 and 80% for the austenitic 1.4310. To obtain the high mechanical characteristics needed for prestressing application, thermo-mechanical stressrelieving is then applied. During this heat treatment, the wires were both under stress and heated by Joule effect. It lower the residual stresses induced by cold working and increased the yield stress by micro plasticity [13,14]
2
Table 1: Chemical composition and mechanical characteristics of the two stainless steels grades studied Elements (wt%) Mechanical characteristics Grade Cr Ni Mo N Mn S C Rm* Rp0,1%* A%** Z%*** 1.4362 22,3 4,26 0,274 0,122 1,09 0,0004 0,017 1960 1850 1,15 47 1.4310 17,7 7,76 0,38 0,033 0,9 0,0013 0,098 2035 1700 1,65 46 *
MPa Ultimate elongation *** Necking
**
Magnetic measurements before and after drawing were performed on two grades to determine the percentage of strain induced martensite. In the case of the duplex grade, after drawing these measurements indicate 47% of ferrite, 15% of austenite and 38% of martensite. For austenitic and again after drawing, the percentage of martensite is 30%. The wire obtained presented high mechanical characteristics (see table 1). 2.2 Hydrogen embrittlement - FIP 1 test This test is derived from the ISO 15630-3 standard. It’s use for the hydrogen embrittlement characterization of the carbon prestressing steels. The principle of this test is to maintain the sample under a stress near to 80% ± 2% of its ultimate strength as it is immersed in a corrosive solution. The solution used contains 250g.L-1 of ammonium thiocyanate NH4SCN heated at 50°C. The pH of the solution obtained is close to 4,5. The test is considered complete when the wire is broken. Therefore, the parameter characterizing the material is the time to failure tf. During this test, a very large amount of hydrogen is absorbed in the material, making it very severe and far from realistic conditions. This severity is reinforced by the low pH of the solution. The reactions (1) to (3) have to be taking into account for this test. The dissociation of the ammonium ion (reaction 1) allows the formation of hydrogen which can then be adsorbed to the surface of the material. The presence of thiocyanate ions SCN-, which is a poison for the recombination reaction of hydrogen (reaction 3), promotes the hydrogen absorption into the metal.
NH 4+ ↔ NH 3 + H +
(1) −
SCN H + + e − → H ads → H ads + M → MH abs → H ads + H ads → H 2
(2) (3)
2.3 Chloride induced stress corrosion cracking - Slow strain rate test Traction frame is a rigid structure provided with a cylinder at the top. The three selected low strain rates (ε = 10-6s-1, 2.10-7s-1, 4.10-8s-1) are made by injection, using a micro-pump Eldex, of a small amount of oil in the cylinder. The force is measured by a load cell of 100kN and elongation by an extensometer. During the test, part of 1
Fédération Internationale de la Précontrainte
3
the wire is immersed in a corrosive solution thanks to a corrosion cell equipped with a three-electrode assembly (figure 2). The sample is used as working electrode whereas the counter and reference electrode are platinum and saturated calomel electrode (SCE), respectively. This three-electrode cell is connected to an SP-300 potentiostat BIOLOGIC used in floating mode, which allows electrochemical measures for non-insulated samples. The heating of the solution at 50°C or 80°C is carried in a remote tank. Then, circulation of the solution is required during the tests. The solution used simulates a chlorinated carbonated concrete with a pH close to 8 and a concentration of chloride ions of 17,5g.L-1. Before each tests, the polarization curve of the sample is determined at a scan rate of 100mV/min from the corrosion potential Ecorr to the pitting potential Epit. This potential is defined as the potential for which the current density equals to 100µA.cm-2. Then, the slow strain rate test starts under potential control which matches with Epit. Prior to the test, specimens were grounded with SiC paper to 1200-grit and then degreased with ethanol/acetone. They are then aged for 24 hours.
Working electrode = studying material
SCE
Platinum counter electrode
Flowing solution
Figure 2: Corrosion cell used for the slow strain rate tests
3
Results
3.1 Hydrogen embrittlement The hydrogen embrittlement behavior of the two grades was examined with a maximum duration of the test set at one week. This time is higher than the one associated with the failure of the carbon prestressing steel (C-steel in the table 2), generally close to 20 hours. The results obtained show a difference in the time to failure of the two stainless steel grades (see table 2). Thus, in the case of the austenitic stainless steel, the time to failure is close to 130 hours whereas the duplex one didn’t break. Complementary test in the same conditions shows that the time to failure of the lean duplex 1.4362 exceeds 650 hours. 4
Table 2: Time to failure associated with the stainless steel grades studied after the FIP-Test
tf (h)
1.4362 1.4310 C-steel >650 128±30 17±7
The metallographic analysis provides some information. In the case of the austenitic stainless steel grade, the crack propagates initially in mode Ι and over a length of 180µm (figure 3). Then, the strongly oriented microstructure of the wire after drawing generates propagation in mode ΙΙ before the shear area, characteristics of the final rupture. This kind of cracking is typical of what Toribio et al. [15] observed in the case of carbon prestressing steel. The hydrogen impact on the austenitic stainless steel 1.4310 can be characterized by the value of the stress intensity factor. An approximation of this factor could be made considering the wire as a cylinder with a diameter D of 3,25mm. The Valiente method [16] suggests calculating KI-HE by the following expression:
a K I − HE = M ( )σ πa D
σ is the applied stress, a the crack depth and M(a/D) a dimensionless function: −1
1
2 2 4 2 a a a a a M ( ) = 0,473 − 3,286 + 14,797 − D D D D D With this method, the stress intensity factor is estimated at 45MPa m , which is in the same order as the one find for prestressing carbon steel in the same conditions [17]. Thus, hydrogen seems to generate the same effect in the two steels, from the stress intensity factor point of view, but in a higher time for the austenitic stainless steel.
180µm Initiation crack
Secondary crack Shear area
Figure 3: Fracture surface of the austenitic stainless steel grade after FIP-Test
In the case of the duplex grade, there was no spontaneous rupture at the end of the test duration set (1 week) but several corroded areas can be observed on the surface of the sample. Longitudinal sections were performed on these areas (figure 4). We can see the presence of stress corrosion cracks which could reach the size of 200µm. The morphology of these cracks, which is dependent on the crack length, is 5
shown in figure 5. Initially, the crack seems to propagate in both phases equally (figure 5). The ferritic phase seems to be preferentially dissolved, which is in accordance with the cathodic protection of austenite by ferrite when the two phases are under tension [18]. Thus, in this first part of crack, a stress corrosion cracking mechanism which included a dissolution step couldn’t be excluded. Then, when the size of the crack is such that the oxidizing species are no longer present at the crack tip (figure 6.a. and b.), the phenomenon of hydrogen embrittlement seems to be preponderant. The crack propagates preferentially in the austenitic phase and at α/γ grain boundaries. The hydrogen trapping at α/γ grain boundaries [19], which generates a local concentration of hydrogen, could explains the preferential propagation of the cracks at these sites. Therefore, the presence of strain induced martensite agreed with the fact that cracks aren’t arrested by the austenite phase, contrary to what could be seen in the literature [20,21]. However, these two points can’t explain the apparent resistance of the ferrite phase to hydrogen cracking. Indeed, the hydrogen diffusion in ferrite and martensite, two bcc phases, is similar and close to 9.10-5cm2.s-1 [22].
200µm
Figure 4: Optical microscope observation of the cracks observed on duplex grades
1μm
10µm
α
Stress
α’+γ Dissolution of α
Figure 5: Back-scattered figure of the morphology of stress corrosion cracks observed in the part near to wire surface
6
5µm
10µm
α
α
Stress
Crack in α’+γ phase
α’+γ α’+γ Cracks at α/γ grain boundaries
a.
b.
Figure 6: Back-scattered figure of the morphology of stress corrosion cracks far from the surface, a. longitudinal section, b. transversal section
3.2 Chloride induced stress corrosion cracking 3.2.1 Strain rate effect on SCC susceptibility The tests describe in this first was made at a temperature of 50°C for the austenitic grade 1.4310. In these conditions, the influence of the imposed strain rate on the tensile behaviour (characterise by Rε 2) of the austenitic grade is reported in figure 6. For a strain rate ε =10-6s-1, the ultimate elongation is similar in air and in the corrosive solution (Rε=1). However, for strain rates less than ε=2.10-7s-1, a marked reduction of the ultimate elongation occurs in the corrosive media (Rε10-6s-1), the pitting corrosion kinetic is too slow regarding to the material elongation: the failure is nearly due to mechanical deformation. For a strain rate close to ε=4.10-8s-1, there is an interaction between corrosion and deformation which able the SCC to occurs. The comparison of the two stainless steels grades will be realized at a strain rate where SCC is susceptible to appear, i.e. close to ε=4.10-8s-1.
2
Rε =
ε Cl ε air
−
with εCl- and εair the ultimate elongation of the material in the corrosive media and in air respec-
tively.
7
Figure 7: Strain rate effect on the ratio Rε
2mm 3.2.2 Stainless steel grades comparison- Temperature effect Figure 8 shows the temperature effect on the tensile behaviour of the two grades. The adverse effect of temperature on the SCC resistance of the austenitic grade is shown in this figure as both Rε and Rσ3 decreased with temperature. Generally, an increased temperature increases the risk for SCC to occur [25]. Moreover, in the range between 20°C and 80°C, the chloride ions activity increases with increasing temperature. Thereby, at 80°C, the passive film of the austenitic might be less resistant than at 50°C and the crack is more prone to propagate. In the case of the lean duplex grade, even if the corrosive media deteriorates his mechanical resistance, no effect of temperature is observed (Rε and Rσ values are similar for 50°C and 80°C) in the range studied.
3
Rσ =
σ Cl σ air
−
with σCl- and σair the ultimate strength of the material in the corrosive media and in air.
8
Figure 8: Temperature effect on the ratio Rε and Rσ for the two stainless steel grades studied
Figure 9 and 10 show the crack morphology for the austenitic or for the duplex stainless steel. In the case of austenitic 1.4310, it seems that dissolution occurs near to the principal crack. The orientation of this dissolution is the same as the martensite lath. According to Ghosh and Kain [8,9], martensite lath would be a preferential path for SCC propagation as their dissolution rate is higher than the one of austenitic matrix. In our case, the SCC resistance might be linked with the high density of martensite laths, which is associated with high density of dislocation.
In the duplex 1.4362, selective dissolution of one phase participated in the SCC process. According to the back-scattered longitudinal observation of the fracture surface (figure 10.b.), we could conclude that ferrite (in dark) is dissolve preferentially during the SCC crack propagation. These cracks are close to the one observed after FIP-test, which could be interpreted by the occurrence of a similar mechanism.
Crack
Stress
Dissolution
a. b. Figure 9: Morphology of the cracks observed in the case of the austenitic grade (T°C=80°C), a. longitudinal section, b. details of the crack
In figure 10.a., we could observed that the crack propagation in the austenite+martensite phase seems to be realized by the dissolution of martensite lath. The surface associated to this phase also shows preferential attack.
9
Stress
α Selective dissolution
α α’+γ
α’+γ
a. b. Figure 10: Morphology of the cracks observed in the case of the duplex grade (T°C=80°C), a. fracture surface, b. Longitudinal section of the fracture surface
4
Conclusion
In summary, it seems that strain induced martensite have a strong influence on the hydrogen embrittlement and stress corrosion cracking resistance of both austenitic and duplex grade. Martensite lath are associated with the propagation of cracks: (i) it favour the diffusion of hydrogen or, (ii) the dissolution of martensite laths increase the SCC development. However, although our work doesn’t include this kind of study, the martensite lath dissolution seems to be related with dislocation pile-ups. The comparison of the two grades was made with FIP-test and slow strain rate test in simulated concrete solution at a strain rate close to 4.10-8s-1. The lean duplex stainless steel presents greater resistance in the two cases, which is consistent with literature data. 5
Acknowledgments
The authors wish to thank Ugitech company for the supply of stainless steels studied and for the financial and technical supports of this project.
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References
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