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Explosion of a hardening vessel for AAC (autoclaved aerated concrete)
Engineering Failure Analysis 43 (2014) 133–149
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Engineering Failure Analysis journal homepage: www.elsevier.com/locate/engfailanal
Explosion of a hardening vessel for AAC (autoclaved aerated concrete) J.W. Erning ⇑ Federal Institute for Materials Research and Testing (BAM), Germany
a r t i c l e
i n f o
Article history: Received 9 December 2013 Received in revised form 14 March 2014 Accepted 28 May 2014 Available online 25 June 2014 Keywords: Roofing Misalignment Weld imperfection Low cycle fatigue Corrosion fatigue
a b s t r a c t A vessel for the hardening of aerated concrete exploded after modification. The modification was implemented by welding additional barrels to extend the vessel. Two and a half years after the modification (about 500 load cycles) the vessel exploded showing a longitudinal crack starting in one of the new segments. After the modification the dimensions of the vessel were 2000 mm diameter and 31,000 mm length, operational pressures ranged from 0.3 to 16 bar. The material of both the old and new parts of the vessel and the properties of the welds showed no deviations from the specifications. The geometry of the longitudinal weld showed a roofing deviation from circularity. Fractographic investigations showed a crack of 1.6 m length following the longitudinal weld which developed over long time caused by corrosion fatigue. At the time of the explosion, the wall thickness was reduced from 13.5 mm down to about 2 mm leading to longitudinal cracking of the vessel throughout the new parts followed by a circular crack which resulted in a complete separation of the pressurized vessel. The crack development is shown by optical and SEMmicrographs. Calculations of nominal stresses show that the geometrical deviations by linear misalignment and roofing cause a stress (tension load) increase of about 4.8 of the mean tensile load value. Adjacent to the weld this causes a local tension overload; in combination with operation conditions (overheated steam of 203 °C, cyclic operation) this causes corrosion fatigue due to cracking of corrosion layers. The crack formed by this mechanism caused a remaining wall thickness of less than 2 mm followed by a ductile residual fracture. Ó 2014 Elsevier Ltd. All rights reserved.
1. Introduction In 1985 one of a several hardening vessel exploded during standard operation. The originally fabricated vessel was elongated to 31 m about two years ago by adding 2 barrels of 3.85 m length, Fig. 2. The elongation was done by welding the additional barrels into the separated vessel. Standard periodic testing was passed short before the explosion. The explosion caused major property damage as well as injuries. Authorities demanded an investigation after the incident. Technical background: The vessel was operated in the following way: Loading of 5 times 7 m3 of aerated concrete and closing of the vessel, generation of a vacuum of 0.3 bar in about 20 min. Heating with live steam or remaining steam from other hardening vessels according to the following regime: up to 1 bar in 1 h, up to 16 bar in one additional hour, constant at 16 bars for 8–10 h, release of pressure in about 2 h. ⇑ Tel.: +49 30 81041733; fax: +49 30 8104 1737. E-mail address: [email protected] http://dx.doi.org/10.1016/j.engfailanal.2014.05.023 1350-6307/Ó 2014 Elsevier Ltd. All rights reserved.
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1.1. Description of failure event The vessel stayed under pressure of 16 bar for about 5 h when the explosion took place. A part of the vessel of 18 m length has moved through a building and a pine forest to a distance of about 200 m. The defect occurred in the area of a longitudinal weld, Fig. 2, in the part of the vessel that was added two years ago. Crack initiation area in the longitudinal weld was about 1.6 m long. Crack propagation started from both sides of the initial crack in longitudinal direction, one side up to the opening of the vessel, in the other direction about 0.5 m exceeding the assembly weld, completely separating the vessel into two parts in the area of a reinforcement ring attached there. From start of service until the explosion about 500 load cycles as described above have occurred.
1.2. Objectives Objective of the original failure analysis investigation [11] was to determine the cause of the failure of the hardening vessel. In detail the district court Marl, Germany, did put following evidence order [11]: i. What was the cause for the explosion of the hardening vessel? ii. Was the vessel fabricated according to the generally accepted engineering standards? 1. Is the explosion caused by a. An offset between the inner and outer weld seam? b. Flattening of the longitudinal weld causing a lack of roundness in the respective parts of the vessel? c. A crack in the weld between the outer strengthening rings and the vessel? d. The use of wrong filler material? iii. 1. Was it possible to detect the failures or the causes respectively during the acceptance test or the following inspections? 2. Did detectable alterations exist that would require additional technical inspections iv. If these detectable deficiencies would have been addressed by the inspecting people and corrected, could the explosion have been avoided? v. Was the explosion likely due to the possibly determined defects considering the operational use of the owner?
1.3. Description of the vessel The vessel was originally manufactured in 1963 and modified in 1965 and 1983. The material used was a standard pressure vessel material 17Mn4 (1.0481) according to DIN 17155. As Filler material SG 2 with M 21 for the root layer by MAGWelding and S2 with OP100 for the inner and outer weld by UP submerged arc welding was used for the parts added in 1983. Dimensions of the vessel: 2000 mm diameter, 31,000 mm length, 99 m3 volume, max. operating pressure +16 bar, 1 bar; test pressure 20.8 bar; max. operating temperature +203 °C. Inspection period: 10 years for pressure testing, 2 years for inner and outer inspection, last inner inspection 1985–06–03.
1.4. Observations on-site The parts from the explosion site as well as those already secured were inspected. The following observations were made: The vessel tore open longitudinal at the closed side end (right in Fig. 2) over a length of about 13 m (Fig. 2). In part-3-of the older part of the vessel the crack continued following the perimeter causing a separation of the vessel into two independent parts. The 18 m long part towards the closed side end sitting on a floating bearing was moved by the energy of the steam pressure into a nearby forest as already described. In the middle part of the added barrel No. 1 close to the closed end an old crack of about 1.6 m at the edge of the inner weld seam could be detected (Figs. 1 and 2). This old crack, already covered by corrosion products, had propagated to a remaining wall thickness of about 2 mm until the final rupture causing the total loss occurred. The rupture of the vessel started in the already damaged area and continued ductile with high velocity via the base material according to the extent shown in Fig. 2. The initial damage was visible in the second new barrel of the vessel facing the opening side of the vessel without causing a complete rupture of the wall.
2. Experimental The following samples from the exploded vessel were collected for further investigation: Multiple pieces from sheet metal (13.5 mm thickness) (Fig. 1) from the area of the longitudinal welds of the parts that were added during the extension of the vessel (Fig. 2). 10 m of steel profiles from the strengthening rings belonging to the first extension part of the vessel (Fig. 2).
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Fig. 1. Overview of samples from the hardening vessel, left part samples from the original construction, right part samples from the new barrel.
Fig. 2. Design drawing of the hardening vessel with crack position and crack detail.
2.1. Chemical analysis Base metal and filler material were analyzed for the two new barrels of the vessel as well as for the adjacent parts of the original vessel. Analyses were conducted using a spark emission spectrometer ARL Quantovac 31000. Results can be found in Table 1. 2.2. Fractography Fractographic investigations were conducted along the primary fracture and the other already damaged longitudinal welds of the new barrels, see Fig. 2. 2.2.1. Macroscopic observations Cracking always started from the inner surface of the vessel in the notch formed by boiler plates and weld bead. Some samples showed cracks staring on both sides of the weld. The crack started without shear crack formation. The crack propagated almost vertical to the metal sheet surface into the wall of the vessel. One can observe frequently the crack growing into the heat affected zone resulting in branching of the cracks 1–2 mm below the sheet surface and the marginal strips visible from the inside (Fig. 3) in the middle of the thickness of the sheet the surface fracture shows a coarse irregular structure that does not allow any statement concerning the crack propagation. The final depth of the old cracks partially shows a line formation parallel to the sheet surface (Fig. 3) indicating a stepwise propagation of the crack. The crack across the remaining wall thickness leading to the complete failure (remaining wall thickness of the vessel wall of the sample shown in Fig. 3: about 1.5 mm) happened by a ductile overload fracture showing a shear lip. The residual fracture appears bright metallic, the other fracture surface is colored by corrosion products. 2.2.2. SEM investigations Investigations by SEM were used in particular to detect crack propagation features. Three samples were examined taken from sheet sample no. 22 (samples 55 and 57) and 23 (sample 57) according to Fig. 1, i.e. from the area of the primary fracture with large old crack depth (see Fig. 2). The fracture surfaces covered with corrosion products could be recovered partially by treatment with Tri-Norm.
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Table 1 Results of chemical analysis. Element
1.0481
Sample-no. C Si Mn P S Al Cr Cu Mo Nb Ni Ti V Cr + Cu + Mo + Ni Co W
0.14–0.20 60.40 0.90–1.40 60.035 60.030 60.020 60.25 60.30 60.10 60.01 60.30 60.03 60.03 60.70
Base metal
Filler
40
42
61
80
43
60
81
41
44
0.19 0.31 1.10 0.026 0.026 n.n 0.28 0.032 0.006 n.n 0.019 0.002 0.005 0.34 0.006 n.n
0.15 0.20 1.24 0.01 0.005 0.032 0.043 0.131 0.013 n.n 0.066 0.002 0.002 0.26 0.008 n.n
0.15 0.20 1.25 0.01 0.006 0.033 0.043 0.131 0.013 n.n 0.066 0.002 0.003 0.25 0.009 n.n
0.21 0.32 1.09 0.026 0.024 0.004 0.066 0.088 0.012 n.n 0.042 0.002 0.003 0.21 0.013 n.n
0.11 0.43 1.31 0.017 0.012 0.014 0.034 0.12 0.010 n.n 0.055 0.003 0.004 0.22 0.012 n.n
0.11 0.40 1.30 0.019 0.011 0.016 0.032 0.12 0.010 n.n 0.058 0.004 0.004 0.22 0.012 n.n
0.10 0.12 0.72 0.025 0.015 0.003 0.045 0.092 0.343 n.n 0.066 0.002 0.005 0.55 0.017 n.n
0.12 0.40 1.28 0.021 0.017 0.014 0.11 0.083 0.008 n.n 0.043 0.004 0.005 0.24 0.01 n.n
0.094 0.47 1.32 0.016 0.014 0.016 0.032 0.12 0.009 n.n 0.049 0.004 0.004 0.206 0.011 n.n
Sample-no.: 40 old barrel opening side. 43 long. weld, new barrel opening side. 42 new barrel opening side. 60 long. weld, new barrel closed side. 31 new barrel closed side. 81 long. weld, old barrel closed side. 80 old barrel closed side. 41 circ. weld, old/new barrel, opening side. 44 circ. weld, new/new barrel, opening side.
2.3. Metallography For metallographic investigations 10 samples were taken from old and new barrels. Sample no. and place of sampling are connected as follows: Sample no.
Place of sampling
30 31 32 33 51 52/53/54
Circular weld, opening side, connection of old and new barrel, sampling from part 1 Fig. 1 Longitudinal weld, opening side, new barrel, sampling from part 1 Fig. 1 Longitudinal weld, opening side, new barrel, sampling from part 9 Fig. 1 Circular weld, connection of the new barrels, sampling from part 12 Fig. 1 Longitudinal weld, new barrel closed side, sampling from part 12 Fig. 1 Longitudinal weld, new barrel closed side, sampling from part 14 (sample 52) and 17 samples 53 and 54) Fig. 1 Longitudinal weld, new barrel closed side, sampling from part 22 Fig. 1 Longitudinal weld, old barrel bottom side connected to new barrel, sampling from part 27 Fig. 1
56 70
Samples were metallographically prepared and investigated microscopically. Micrographs of the cracks in polished and etched sections were taken. Etching was done by alcoholic sulfuric acid. The microstructure of the welds from the old and new parts of the vessel was compared. 2.4. Mechanical testing 2.4.1. Hardness Vickers hardness HV1 according to DIN 50133-2 was determined using the cross sections from Section 2.3. The testing was conducted with the small load hardness tester Finotest. Table 2 contains the values determined: 2.4.2. Bend test By means of bend tests the deformation- and fracture behavior of the longitudinal welds on the inner surface during quasistatic load was compared. Three bend samples were taken, two from the opening side new barrel (parts 2 and 7
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Fig. 3. fracture surface from the area of the new barrel with long time periodical crack and residual fracture.
Fig. 1) and one from a bottom side old barrel (part 27 Fig. 1). The 50 mm wide samples were used as taken without any surface treatment including the weld seams. 2.4.3. Notched bar impact test The absorbed impact energy was determined for the boiler sheets of the old and new barrels as well as the HAZ of the longitudinal weld (notch layer parallel to the weld, notch ground in the HAZ) of a new and an old barrel on longitudinal and circumferal oriented ISO-V-specimen at 0 °C (partially at 20 °C as well). DIN 17155 (1983) was decisive for the choice of sample design and test temperature. Experiments were conducted according to DIN 50115 using a Losenhausen type 300 J pendulum impact tester. Places of sampling and the results are given in Tables 3 and 4. 2.4.4. Tensile test To investigate the strength and deformation properties given by DIN 17155 for sheets from 1.4081 flat samples were taken from old and new barrels in longitudinal and circumferal directions and tensile tests were conducted. Tests were realized according to DIN 50146 in the 200 kN-range of a DIN 51221 1000 kN Instron type testing machine. Places of sampling and the results are given in Table 5. 2.5. Additional information By assessment of available information the following additional information was gathered: 1. Measurements concerning roofing of the exploded vessel had not been conducted. Measurements taken after the explosion showed roofing to the extent of 5 mm for vessel no. 4, elongated in 1983 and 7 mm in one singular place. Vessel no. 5 is said to have shown roofing up to 10 mm.
Table 2 Results of hardness measurements. Sample no.
30 31 32 33 51b 52 53 54c 56 70 a b c
Hardness HV1 in Base material
HAZa
Filler
195–211 169–185 174–180 164–174 202–213 202–247 175–182 185–187 191–198 167–183 173–193
248 237 226 227 257 267 228 229 247 – 234
210–215
Max hardness value in the heat affected zone. Additional heat influence in the sampling area. Severe plastic deformation (neckdowns) due to the crack progress.
210–220 200–220 235–245 200–220 200–215 210–225 220–240 – 190–200
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Table 3 Results of impact test. Sample no.a
Place of sampling Base metal; Opening side old barrel (sampling part 1 Fig. 1)
11 Q 12 Q 13 Q 14 Q 11 Q 12 Q 13 Q 15 L 16 L 17 L 18 L Base metal; Opening side new barrel (sampling part 12 Fig. 1) 31 Q 32 Q 33 Q 34 Q 35 L 36 L 37 L Notch impact work for sheets from 1.4081 acc. to DIN 17155 (1983) with thickness 660 mm; circumferal samples a b
Temperature (°C)
Impact work
0 0 0 0 0 20 20 0 0 20 20 0 0 0 0 0 0 0 0
12.5 25.5 19.5 22.0 19.0 50.0 33.0 13.8 14.8 50.0 33.0 >143b >143b 216 223 232 254 268 P31
Q: Circumferal sample (sample long. axis parallel to vessel long. axis); L: Longitudinal sample sample long. axis in barrel circ. axis). Samples bent and partially broken.
Table 4 Results of impact test. Place of sampling
Sample no.a
Temperature (°C)
Impact work
Base metal; Closed side new barrel (sampling part 14 Fig. 1)
51 52 53 55 56 57 71 72 73 74 75 76 77 61 62 62 64 65 61 62 62 64 65
0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0
162 209 219 246 199 250 45.0 40.0 59.0 50.0 74.5 64.5 78.0 204 44.0 53.0 32.5 210 76.0 71.0 39.0 36.0 40.0
Base metal; Closed side old barrel (sampling part 27 Fig. 1)
Longitudinal weldb Closed side new barrel (sampling part 17 Fig. 1)
Longitudinal weld
a b
b
Closed side old barrel (sampling part 27 Fig. 1)
Q Q Q L L L Q Q Q Q L L L
Q: Circumferal sample (sample long. axis parallel to vessel long. axis); L: Longitudinal sample sample long. axis in barrel circ. axis). Notch at the inner sheet surface parallel to the weld in the HAZ.
2. The hardening process can result in a pressure increase up to 1.5 bar. The factory itself never observed temperature increase >1 K corresponding to a pressure increase of 0.1 bar. The operator excludes the possibility of pressure increase due to the operation of an open system. 3. Due to the regulation devices including safety valves at the steam generator of the operator (being periodically inspected) exceeding of the maximum operating pressure of the hardening vessel is to be expected in exceptional cases. The maximum pressure will stay below 18 bar in any case. 4. The water used in the steam generator is treated by a hydrogen- and sodium exchanger. The feed water is degassed to remove oxygen, phosphates and a oxygen scavenger are added to remove residual hardness and oxygen. 5. Live steam is provided as saturated steam. Maximum temperature in the stone hardening vessel is 203.4 °C. Steam temperature is measured and registered at the stone hardening vessel. (Temperature reading of 203 °C is existing for the time of failure referring to operators claims.)
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Sample no.a
Yield strength Rp0,2 (N mm2)
Tensile strength RM (N mm2)
Elongation on fracture A5 %
Reduction of area Z%
Opening side new barrel (sampling part 1 Fig. 1)
46 47 48 49 48 49 66 67 68 69 68 69 86 87 88 89
404 396 418 387 352 354 377 387 376 375 400 421 451 467 380 373
556 551 517 514 512 514 519 518 525 522 511 513 546 562 538 533
25.5 25.5 27.5 27.0 30.5 29.0 25.5 26.0 26.5 29.5 27.5 30.5 20.0 21.0 29.5 26.5
62 60 69 66 68 65 63 64 63 63 68 71 58 58 51 53
Opening side new barrel (sampling part 1 Fig. 1 48 L + 49 L and part 12 Fig. 1 48 Q + 49 Q)
Closed side new barrel (sampling part 14 Fig. 1 66 L – 69 Q and part 18 Fig. 1 68 L + 69 L)
Closed side old barrel (sampling part 18 Fig. 1 86 L + 87 L and part 27 Fig. 1 88 Q + 89 Q)
a
L L L L Q Q L L Q Q L L L L Q Q
Q: Circumferal sample (sample long. axis parallel to vessel long. axis); L: Longitudinal sample sample long. axis in barrel circ. axis).
3. Results 3.1. Evaluation of in-service data The evaluation of data provided from periodic inspection gave no reference for any problems in the operational condition of the vessel. The periodic inspection shortly prior to the explosion did not detect any crack or problems with the vessel in question. 3.2. Chemical analysis The boiler plates (samples no. 42 and 61 acc. to Table 1) meet the requirements for the chemical composition of the material 1.4081 (17Mn4) according to DIN 17155 (1983). The values determined meet the values given in the acceptance certificate. The filler material of the inner longitudinal welds (sample no. 43 and 60 acc. to Table 1) show a carbon and aluminum content slightly below the minimal values and a silicon value touching the upper limit or slightly above respectively, if one assumes that base material and filler should be of the same chemical composition. The deviations according to DIN 17155 (1983) and in reference to the base material do not give any indication of a wrong filler material nor would they allow the assumption of a deterioration concerning the fracture properties. 3.3. Fractography SEM investigations showed a honeycomb structure giving evidence of ductile fracture could be detected outside the overload fracture and in the area in front of the shear lip in the middle part of the crack surface as well as in the origin of fracture (Figs. 3 and 6). Line formation in front of the shear lip (Fig. 5) differing clearly from the much finer microstructure layers indicate a stepwise crack propagation in this area (Fig. 4). Line formation in front of the shear lip (Fig. 5) could not be observed in all fracture areas; for the three samples investigated by SEM these were visible in the samples taken from part 22 (Fig. 1). Fatigue striations characteristic for alternate loads differ from the line formation visible in this case. Features of crack propagation allowing conclusions concerning the period of crack formation could not be found. 3.4. Metallography 3.4.1. Microscopic evaluation Microscopic evaluation of the cross-sections confirmed the macroscopic visible finding that all cracks developed from the inner surface of the vessel in the coarse grain area of the heat affected zone directly adjacent to the upper weld layer. Figs. 7 and 8 show two cracks in polished and etched section. The path of the crack in the pictures as well as in Figs. 9 and 10 clearly shows that the cracks started vertically from the surface into the material as well as partial branching. Apart from the crack origin the cracks are independent of the microstructure of the material in the welding area. The propagation of the cracks
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Fig. 4. Coarse line formation near the shear lip (sample 55, Fig. 1).
Fig. 5. Fine line formation in the middle area of the crack surface (sample 57, Fig. 1).
Fig. 6. Structure in the shear lip of sample 55 (Fig. 1).
thus follows a transgranular path. Figs. 11 and 12 show severe corrosive attack at the flanks of the crack into the small branches of the side cracks. The crack in Fig. 7 is heavily corroded as well. In order to determine, whether critical microstructural states influenced the crack initiation, the microstructure of the longitudinal welds of old and new parts of the vessel were compared. The comparison did not show any differences concerning the microstructure. No weld faults such as undercutting, slag inclusions or lack of fusion were detected. Figs. 13 and 14 show the state of the surface in comparison between inner and outer surface. Fig. 13 shows formation of pits due to corrosion at the inner surface whereas Fig. 14 shows the smooth outer surface of the vessel at the same sample.
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Fig. 7. Corroded crack starting from coarse grain zone in the transition weld-sheet, sample 32. Left polished, right etched.
Fig. 8. Crack initiation without branching, sample 52. Left polished, right etched.
3.5. Estimation of loads and stresses due to misalignment and roofing 3.5.1. Geometrical analysis The geometrical relations after the failure in the area of longitudinal welds of the new barrels and the bottom side old barrel are shown in an exemplary manner in Fig. 15 in 1:1 scale. Two geometrical characteristics causing locally tension increase can be observed, namely partially observed misalignment, see Fig. 12, part 4 as well as roofing, being more pronounced in the new barrels, Fig. 12, part 13/21.
3.5.2. Stress analysis The misalignment – the absolute values could not be determined due to the overlaying roofing (angle failures) – leads to a superimposed bending stress and, due to the generally sharper notch formation, to an increased notch stress. The amount of the additional bending stress increases linear with the misalignment as shown in the derivation given in the following
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Fig. 9. Crack with branching, initiation upper left (inner surface of the vessel). Etched, sample 52.
Fig. 10. Crack with branching, initiation upper right (inner surface of the vessel). Etched, sample 56.
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Fig. 11. Crack with branching and shear lip (see Fig. 6). Sample 52.
Fig. 12. Crack branchings with corrosion products (sample 56).
Fig. 13. Pitting corrosion on the inner surface of the vessel (sample 32, new barrel).
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Fig. 14. Outer surface of the vessel without corrosion (sample 32, new barrel).
Fig. 15. Geometric relations in the area of the longitudinal welds of the new barrels (Neuer Kesselschuss) and the adjacent old barrel (Alter Kesselschuss) Inside of the vessel is always at the bottom.
schematic, see Illustration I. Without consideration of the notch, expressed with the stress concentration factor aK, the resulting boundary stress r0 results from the nominal tensile stress rz and the bending stress rb to
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r0 ¼ rz þ rb ¼
145
F 6Fe 6e þ 2 ¼ rz 1 þ bs bs s
For a misalignment e of 1 mm and a wall thickness s of 13.5 mm the boundary stress r0 would result in 1.44-fold of the nominal tensile stress. To determine the roofing that resulted from inadequate smooth curving of the sheet ends in the area of longitudinal welds in the following analytical stress estimation [11] the vessel geometry according to Illustration II was assumed. The amount of roofing h was calculated from the roofing angle formed by the sheet ends adjacent to the weld seam area (see Fig. 12, after failure). The following roofing amounts were determined: (a) Opening side new barrel: h 4.63 mm. (b) Bottom side new barrel: h P 8.63 mm. (c) Bottom side old barrel: h 1.37 mm. Stress estimation according to a roofing amount h at the vessel: Eccentricity k ¼ hr 1 according to the definition in technical codes, i.e. AD-Merkblatt H1
r 1 ¼ rh 1þk qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 2k þ k sin a ¼ 1þk qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 tan a ¼ 2k þ k qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 1 ¼ r 2k þ k cos a ¼
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 P2 ¼ pl ¼ pr 2k þ k : Partial force resulting from inner pressure causing bending stress due to roofing. Components of the partial force P1 resulting from inner pressure, perfect geometry assumed:
P1X ¼
R pa
P1Y ¼
R pa
0 0
p cos urdu ¼ pr sin a ¼ pr
pffiffiffiffiffiffiffiffiffiffi 2kþk2 1þk
2 p sin urdu ¼ prð1 þ cosaÞ ¼ pr 1þk
Equilibrium at partial system III:
F m sin a ¼ P1X ¼ pr sin a F m ð1 þ cos aÞ ¼ P1Y ¼ prð1 þ cos aÞ F m ¼ pr
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Equilibrium at partial system IV:
Shear force
Q ¼ P 2 ¼ P2 ¼ pr
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2k þ k
2
2
Bending moments M ¼ P2 12 ¼ p 12 ¼ 12 pr2 ð2k þ k Þ Highest tensile boundary stress (tensile and bending)
rmax ¼
F m 6M þ 2 s s rh s
r s
rmax ¼ p 1 þ 3 ð2k þ k2
i
For k 1 approximately
rmax ¼ p
r r 1þ6 k s s
mm For rs ¼ 1000 ¼ 74:1 and 13:5 mm
rm p rs (Barlow´s formula, see AD-Merkblatt B1
rmax ¼ 1 þ 444:4k rm In the actual case with r = 1000 mm the following numerical equation results: rmax rm ¼ 1 þ 0:444h if h is inserted in mm.Following the given derivation this results in boundary tension stress rmax for the roofing amounts given above (a) 3.06 (b) 4.84 (c) 1.61 times the calculated mean tensile stress rm. As the bending stresses due to misalignment and roofing add to the mean tensile stress rm from the vessel formula and the resulting stress is additionally enhanced locally by notch effects (transition of sheet to weld, weld seam, undercutting, corrosion marks) one has to assume the stress at the inner surface of the two new barrels in the transition between sheet and weld exceeded the calculated nominal values significantly. 3.6. Mechanical testing 3.6.1. Hardness The determined hardness of the base material is some 7–8% higher as would result from calculation from the tensile strength according to DIN 50150 (08/1976). This is mainly caused by the influence of the low test force for which calculation according to DIN 50150 is not permitted and by the strain hardening of the sample surfaces due to mechanical preparation (grinding). As this is applicable to the HAZ and the filler as well, the relation between the hardness of the three different zones is valid. Due to coarse grain formation and the higher hardness values in the HAZ the weld area reveals a weak spot in the construction, nevertheless the absolute values of hardness are so low that a risk of failure due to lower plastic deformation properties, stress corrosion or hydrogen embrittlement is not existent. Sample no. 51 taken in some 150 mm distance to the circular weld connecting the two new barrels shows a heat treatment independent of the longitudinal welding. This is the probable cause for the elevated hardness levels in base material as well as in the HAZ in comparison to the other samples. Between the sampling area and the circular weld an additional electric hand-weld seam is located on the outside of the vessel. Both facts, heat treatment and hand weld allow the conclusion of additional alignment in the area of the closed side new barrel prior to joining.
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3.6.2. Bend tests The samples were bent under identical test conditions until the onset of microscopic visible cracks. Crack formation always happened by shear crack formation in the notch formed by weld seam and sheet surface, Fig. 16. The deformation until formation of cracks was larger for samples 2 and 7 as for sample 27 with the difference between the first two samples resulting from different notch effect due to different excess weld metal. 3.6.3. Notched bar impact test The minimal notch impact work for the material according to DIN 17155 (1983) at 0 °C is not met by the boiler sheets of the opening side old barrel in longitudinal nor circumferal direction. The 1959 version of DIN 17155 valid until 1983 requires a minimal notch impact work of 34.3 J at room temperature for the DVM sample. As this value is met at 20 °C for the ISO-V specimen, the sheet from the old barrel does meet the 1959 requirements concerning notch impact work, these sheets being of no concern for failure initiation anyway. With clearly differing values for the notch impact work between old and new barrels, showing better values for the new barrel material, the other tested sheets cannot be objected either. The HAZ shows a sufficient notch impact work as well, giving higher results for 0 °C than the base material of the opening side old barrel. 3.6.4. Tensile test The tensile strength values determined meet the requirements of DIN 17155 (1983). 0.2% yield point shows an elevated level. This is caused by prior plastic cold working caused by curving, the failure and the need for straightening of the tensile samples prior to testing.
4. Discussion 4.1. Failure sequence The explosion of the vessel took place after preliminary weakening of the vessel wall by cracks due to a long-term process, i.e. not in one operating cycle generated. The cracks were formed on the inner surface of the added new barrels for the elongation of the vessel starting from the notch between sheet surface and longitudinal weld. Crack propagation went almost perpendicular to the surface into the wall of the vessel. The crack in the starting area of the final defect was some 1.6 m long and resulted in a partial residual wall thickness of 1.5 mm. Formation and propagation of the cracks in the new barrels happened due to low cycle fatigue mechanism, macroscopically without deformation, crack propagation during the explosion took place exclusively after preliminary extensive plastic deformation (ductile).
Fig. 16. Bend tests with 50 mm length samples until visible crack formation.
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4.2. Failure mechanism and root cause The results of all Investigations give strong evidence to corrosion fatigue as the main failure mechanism leading to the explosion. Due to geometrical deviations (misalignment and roofing at the longitudinal weld) local mechanical stresses were significantly higher than designed. This caused low cycle fatigue and crack initiation. Other reasons such as material or welding flaws or a singular overload can be excluded. Corrosion fatigue [1,2] is described in DIN EN ISO 8044 [3] as localized corrosion with formation of cracks as a result of mechanical damage of protecting layers by cyclic critical stress or shrinking of a component. It is always observed in areas of high local stress concentrations, the release of which results in small locally concentrated plastic deformation. This type of corrosion is a fatigue with a low number of load cycles (LCF = low cycle fatigue) combined with high mechanical stress as observed in this case. Crack propagation and appearance of the cracks thus are frequency dependent (resulting in differing amounts of load cycles): Low cycle numbers and low tension amplitude result in a corrosion dominated mechanism, high cycle numbers and high tension amplitudes result in a mechanically dominated mechanism. As the cracking of the protective layers is the major cause for crack initiation, the known operating conditions result in the following failure mechanism: 1. Formation of a protective layer (magnetite) at low pressure by reaction of steam with steel. 2. Cracking of the brittle protective layer in the area of corrosion pits during pressure increase to 16 bar in areas of high local stress or high excess tensile load respectively. 3. Local corrosion at 16 bar. Tensile induced anodic dissolution in the formed notch root can be involved (stress corrosion). 4. Standstill corrosion caused by water in the crack (corrosion of the crack walls without pressure. Necessary for the cracking of the protective layer are at least small plastic deformations of the base material. Corrosion fatigue with crack formation can be observed especially in components with instationary operating conditions resulting in a cyclic repetition of steps 1–4. Literature shows that oxygen containing water at high temperature is sufficient to enable crack formation and propagation under given mechanical conditions. 4.3. Answers to objectives i. The vessel exploded due to LCF caused by mechanical stress caused by geometrical deviations of the new barrels. ii. Roofing as observed here is the cause of a not correct fabrication of the sheets for the vessel barrels. Following the accepted technical regulations the radius of the sheet ends has to follow the radius of the vessel exactly. Roofing caused by the deviation of roundness is not yet included in standardization at the time of the failure. a. An offset between the inner and outer weld seam had no influence on the failure of the vessel. b. Flattening causing a lack of roundness in the area of the longitudinal welds of the new barrels of the vessel is the root cause for the failure. c. A crack in the weld between the outer strengthening rings and the vessel that might have existed prior to the failure could not be detected. d. Use of a wrong filler material can be excluded. iii. 1. It would have been possible to detect the geometrical deviations during inspection. Until the day of the failure a geometrical inspection that would have revealed such flaws was not part of the acceptance test nor the periodical testing. 2. The inner inspection shall reveal damages caused during operation of the device. Whether the cracks in the area of the longitudinal weld were definitely existent and visible during the last inspection is not clear. iv. If these detectable deficiencies would have been addressed by the inspecting people and corrected, the explosion could have been avoided. v. The explosion was caused by geometrical deviations of the components for the new barrels of the hardening vessel. LCF was caused by the geometrically induced stress and the cyclic operation conditions resulted in the formation of cracks leading to the failure of the vessel. Thus the failure was to be expected considering existing manufacturing flaws and operation conditions. 4.4. Reference to similar failure events Some years later, in October 1991, a pressure vessel for hydrogen exploded in Hanau, Germany, causing major damage to almost all buildings of a large facility and other buildings in a distance up to 1000 m [4]. As the explosion took place on a Saturday morning, nobody was working near to the vessel thus causing no personal injuries. The cause for the explosion was the same as in the case described above: roofing of the longitudinal welds of the vessel leading to low cycle fatigue of the weld and subsequent to the failure of the vessel. The general problem of excess tensile stress due to mechanical deviations, especially roofing, is described in the regulations at least since 1995 [5]. Nevertheless, these regulations did
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not take into account roofing below given values for discrete dimension nor the influence of the medium that is stored in the vessel. Thus a miscalculation of the possible load cycles of more than one order of magnitude still was possible. Due to this failure, in 1998 a comprehensive research report dealing with geometrical deviations of pressure vessels was published [6]. Here descriptions of the mechanical deviations are given and mathematical methods for the calculation of the excess loads are presented. The report provides a comprehensive overview of the possible mathematical models and shows solutions for the calculation of the excess stresses caused by the geometrical deviations. 4.5. Reference to A. Martens The process of gathering all data from a reconstruction of construction and failure case subsequently followed by the determination of mechanical data for the materials involved leads consequently to the conclusion of one possible failure cause. The methodology follows strictly the procedures described by Adolf Martens: Reconstruct the failure case, interpret the data and you will find the conclusion. 4.6. Recommendations and preventive actions The results show clearly that even small deviations of roundness result in excess loads on vessel walls. It is necessary to avoid such geometrical deviation in mm-scale for pressure vessels. The problem is described in literature for many years [6–8]. 4.7. Implications for technical regulations The impact of these geometrical deviations were included in German [9] and British standards [10] some years after the incidents. Mathematical methods available nowadays [5] allow exact calculations of the impact of geometrical deviations of pressure vessels. The regulations available today include the determination of such deviations. 5. Conclusions The stone hardening vessel exploded due to a final ductile fracture. Prior to the final fracture, weakening of the vessels wall by LCF took place, starting from a notch between longitudinal weld and wall sheets. Cause for the LCF crack initiation was a geometrical deviation of additional barrels for elongation of the vessel. The roofing of the wall geometry resulted in an excess stress in the respective areas 4.8 times higher than the calculated nominal stress of the vessel wall. Acknowledgements Special acknowledgements to the authors of the original expert report [11] J. Ziebs and K. Naseband and to Amtsgericht Marl for the declassification. References [1] Hickling J. Der Maschinenschaden 1982;55(2):95–105. [2] Hickling J, Blind D. EFC conference on environmental sensitive cracking problems in nuclear installations containing high temperature water Munich; 1984. [3] DIN EN ISO 8044. Corrosion of metals and alloys – basic terms and definitions; 1999. [4] BAM-Expert Report 4-218, 01.10.1992. Causes and cause of events of the explosion of a hydrogen tank on the area of XXX (Ursachen und Hergang des Berstens eines Wasserstofftanks auf dem Gelände der Firma XXX). [5] Weber M. Calculation of structural stress in cylindrical vessels with roofing and internal pressure, (Strukturspannungsberechnung bei zylindrischen Behältern mit Aufdachungen unter Innendruck), in German, BAM Forschungsbericht 224; 1998, ISBN 3-89701-159-X. [6] Haibach E. Fraunhofer Gesellschaft e.V., Laboratorium für Betriebsfestigkeit, FB-77; 1968. [7] Lange K. Lehrbuch der Umformtechnik Band 3 Blechumformung. Springer; 1975. S. 131. [8] Dubbel Taschenbuch für den Maschinenbau, Band II, Springer; 1953. S. 575. [9] AD-Merkblatt HP1. Construction and testing of pressure vessels – design and testing (Herstellung und Prüfung von Druckbehältern – Auslegung und Prüfung), Beuth Verlag; 1995. [10] BS PD 6493 (1991). Guidance on methods for assessing the acceptability of flaws in fusion welded structures, BSI 1991. [11] BAM-Expert Report 1.2/12960. About the causes of the explosion of a stone hardening vessel at XXX company (Über die Ursachen des Berstens (Zerknallens) eines Steinhärtekessels bei der Firma XXX) in German issued 9.6.1986, unpublished.
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Engineering Failure Analysis journal homepage: www.elsevier.com/locate/engfailanal
Explosion of a hardening vessel for AAC (autoclaved aerated concrete) J.W. Erning ⇑ Federal Institute for Materials Research and Testing (BAM), Germany
a r t i c l e
i n f o
Article history: Received 9 December 2013 Received in revised form 14 March 2014 Accepted 28 May 2014 Available online 25 June 2014 Keywords: Roofing Misalignment Weld imperfection Low cycle fatigue Corrosion fatigue
a b s t r a c t A vessel for the hardening of aerated concrete exploded after modification. The modification was implemented by welding additional barrels to extend the vessel. Two and a half years after the modification (about 500 load cycles) the vessel exploded showing a longitudinal crack starting in one of the new segments. After the modification the dimensions of the vessel were 2000 mm diameter and 31,000 mm length, operational pressures ranged from 0.3 to 16 bar. The material of both the old and new parts of the vessel and the properties of the welds showed no deviations from the specifications. The geometry of the longitudinal weld showed a roofing deviation from circularity. Fractographic investigations showed a crack of 1.6 m length following the longitudinal weld which developed over long time caused by corrosion fatigue. At the time of the explosion, the wall thickness was reduced from 13.5 mm down to about 2 mm leading to longitudinal cracking of the vessel throughout the new parts followed by a circular crack which resulted in a complete separation of the pressurized vessel. The crack development is shown by optical and SEMmicrographs. Calculations of nominal stresses show that the geometrical deviations by linear misalignment and roofing cause a stress (tension load) increase of about 4.8 of the mean tensile load value. Adjacent to the weld this causes a local tension overload; in combination with operation conditions (overheated steam of 203 °C, cyclic operation) this causes corrosion fatigue due to cracking of corrosion layers. The crack formed by this mechanism caused a remaining wall thickness of less than 2 mm followed by a ductile residual fracture. Ó 2014 Elsevier Ltd. All rights reserved.
1. Introduction In 1985 one of a several hardening vessel exploded during standard operation. The originally fabricated vessel was elongated to 31 m about two years ago by adding 2 barrels of 3.85 m length, Fig. 2. The elongation was done by welding the additional barrels into the separated vessel. Standard periodic testing was passed short before the explosion. The explosion caused major property damage as well as injuries. Authorities demanded an investigation after the incident. Technical background: The vessel was operated in the following way: Loading of 5 times 7 m3 of aerated concrete and closing of the vessel, generation of a vacuum of 0.3 bar in about 20 min. Heating with live steam or remaining steam from other hardening vessels according to the following regime: up to 1 bar in 1 h, up to 16 bar in one additional hour, constant at 16 bars for 8–10 h, release of pressure in about 2 h. ⇑ Tel.: +49 30 81041733; fax: +49 30 8104 1737. E-mail address: [email protected] http://dx.doi.org/10.1016/j.engfailanal.2014.05.023 1350-6307/Ó 2014 Elsevier Ltd. All rights reserved.
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1.1. Description of failure event The vessel stayed under pressure of 16 bar for about 5 h when the explosion took place. A part of the vessel of 18 m length has moved through a building and a pine forest to a distance of about 200 m. The defect occurred in the area of a longitudinal weld, Fig. 2, in the part of the vessel that was added two years ago. Crack initiation area in the longitudinal weld was about 1.6 m long. Crack propagation started from both sides of the initial crack in longitudinal direction, one side up to the opening of the vessel, in the other direction about 0.5 m exceeding the assembly weld, completely separating the vessel into two parts in the area of a reinforcement ring attached there. From start of service until the explosion about 500 load cycles as described above have occurred.
1.2. Objectives Objective of the original failure analysis investigation [11] was to determine the cause of the failure of the hardening vessel. In detail the district court Marl, Germany, did put following evidence order [11]: i. What was the cause for the explosion of the hardening vessel? ii. Was the vessel fabricated according to the generally accepted engineering standards? 1. Is the explosion caused by a. An offset between the inner and outer weld seam? b. Flattening of the longitudinal weld causing a lack of roundness in the respective parts of the vessel? c. A crack in the weld between the outer strengthening rings and the vessel? d. The use of wrong filler material? iii. 1. Was it possible to detect the failures or the causes respectively during the acceptance test or the following inspections? 2. Did detectable alterations exist that would require additional technical inspections iv. If these detectable deficiencies would have been addressed by the inspecting people and corrected, could the explosion have been avoided? v. Was the explosion likely due to the possibly determined defects considering the operational use of the owner?
1.3. Description of the vessel The vessel was originally manufactured in 1963 and modified in 1965 and 1983. The material used was a standard pressure vessel material 17Mn4 (1.0481) according to DIN 17155. As Filler material SG 2 with M 21 for the root layer by MAGWelding and S2 with OP100 for the inner and outer weld by UP submerged arc welding was used for the parts added in 1983. Dimensions of the vessel: 2000 mm diameter, 31,000 mm length, 99 m3 volume, max. operating pressure +16 bar, 1 bar; test pressure 20.8 bar; max. operating temperature +203 °C. Inspection period: 10 years for pressure testing, 2 years for inner and outer inspection, last inner inspection 1985–06–03.
1.4. Observations on-site The parts from the explosion site as well as those already secured were inspected. The following observations were made: The vessel tore open longitudinal at the closed side end (right in Fig. 2) over a length of about 13 m (Fig. 2). In part-3-of the older part of the vessel the crack continued following the perimeter causing a separation of the vessel into two independent parts. The 18 m long part towards the closed side end sitting on a floating bearing was moved by the energy of the steam pressure into a nearby forest as already described. In the middle part of the added barrel No. 1 close to the closed end an old crack of about 1.6 m at the edge of the inner weld seam could be detected (Figs. 1 and 2). This old crack, already covered by corrosion products, had propagated to a remaining wall thickness of about 2 mm until the final rupture causing the total loss occurred. The rupture of the vessel started in the already damaged area and continued ductile with high velocity via the base material according to the extent shown in Fig. 2. The initial damage was visible in the second new barrel of the vessel facing the opening side of the vessel without causing a complete rupture of the wall.
2. Experimental The following samples from the exploded vessel were collected for further investigation: Multiple pieces from sheet metal (13.5 mm thickness) (Fig. 1) from the area of the longitudinal welds of the parts that were added during the extension of the vessel (Fig. 2). 10 m of steel profiles from the strengthening rings belonging to the first extension part of the vessel (Fig. 2).
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Fig. 1. Overview of samples from the hardening vessel, left part samples from the original construction, right part samples from the new barrel.
Fig. 2. Design drawing of the hardening vessel with crack position and crack detail.
2.1. Chemical analysis Base metal and filler material were analyzed for the two new barrels of the vessel as well as for the adjacent parts of the original vessel. Analyses were conducted using a spark emission spectrometer ARL Quantovac 31000. Results can be found in Table 1. 2.2. Fractography Fractographic investigations were conducted along the primary fracture and the other already damaged longitudinal welds of the new barrels, see Fig. 2. 2.2.1. Macroscopic observations Cracking always started from the inner surface of the vessel in the notch formed by boiler plates and weld bead. Some samples showed cracks staring on both sides of the weld. The crack started without shear crack formation. The crack propagated almost vertical to the metal sheet surface into the wall of the vessel. One can observe frequently the crack growing into the heat affected zone resulting in branching of the cracks 1–2 mm below the sheet surface and the marginal strips visible from the inside (Fig. 3) in the middle of the thickness of the sheet the surface fracture shows a coarse irregular structure that does not allow any statement concerning the crack propagation. The final depth of the old cracks partially shows a line formation parallel to the sheet surface (Fig. 3) indicating a stepwise propagation of the crack. The crack across the remaining wall thickness leading to the complete failure (remaining wall thickness of the vessel wall of the sample shown in Fig. 3: about 1.5 mm) happened by a ductile overload fracture showing a shear lip. The residual fracture appears bright metallic, the other fracture surface is colored by corrosion products. 2.2.2. SEM investigations Investigations by SEM were used in particular to detect crack propagation features. Three samples were examined taken from sheet sample no. 22 (samples 55 and 57) and 23 (sample 57) according to Fig. 1, i.e. from the area of the primary fracture with large old crack depth (see Fig. 2). The fracture surfaces covered with corrosion products could be recovered partially by treatment with Tri-Norm.
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Table 1 Results of chemical analysis. Element
1.0481
Sample-no. C Si Mn P S Al Cr Cu Mo Nb Ni Ti V Cr + Cu + Mo + Ni Co W
0.14–0.20 60.40 0.90–1.40 60.035 60.030 60.020 60.25 60.30 60.10 60.01 60.30 60.03 60.03 60.70
Base metal
Filler
40
42
61
80
43
60
81
41
44
0.19 0.31 1.10 0.026 0.026 n.n 0.28 0.032 0.006 n.n 0.019 0.002 0.005 0.34 0.006 n.n
0.15 0.20 1.24 0.01 0.005 0.032 0.043 0.131 0.013 n.n 0.066 0.002 0.002 0.26 0.008 n.n
0.15 0.20 1.25 0.01 0.006 0.033 0.043 0.131 0.013 n.n 0.066 0.002 0.003 0.25 0.009 n.n
0.21 0.32 1.09 0.026 0.024 0.004 0.066 0.088 0.012 n.n 0.042 0.002 0.003 0.21 0.013 n.n
0.11 0.43 1.31 0.017 0.012 0.014 0.034 0.12 0.010 n.n 0.055 0.003 0.004 0.22 0.012 n.n
0.11 0.40 1.30 0.019 0.011 0.016 0.032 0.12 0.010 n.n 0.058 0.004 0.004 0.22 0.012 n.n
0.10 0.12 0.72 0.025 0.015 0.003 0.045 0.092 0.343 n.n 0.066 0.002 0.005 0.55 0.017 n.n
0.12 0.40 1.28 0.021 0.017 0.014 0.11 0.083 0.008 n.n 0.043 0.004 0.005 0.24 0.01 n.n
0.094 0.47 1.32 0.016 0.014 0.016 0.032 0.12 0.009 n.n 0.049 0.004 0.004 0.206 0.011 n.n
Sample-no.: 40 old barrel opening side. 43 long. weld, new barrel opening side. 42 new barrel opening side. 60 long. weld, new barrel closed side. 31 new barrel closed side. 81 long. weld, old barrel closed side. 80 old barrel closed side. 41 circ. weld, old/new barrel, opening side. 44 circ. weld, new/new barrel, opening side.
2.3. Metallography For metallographic investigations 10 samples were taken from old and new barrels. Sample no. and place of sampling are connected as follows: Sample no.
Place of sampling
30 31 32 33 51 52/53/54
Circular weld, opening side, connection of old and new barrel, sampling from part 1 Fig. 1 Longitudinal weld, opening side, new barrel, sampling from part 1 Fig. 1 Longitudinal weld, opening side, new barrel, sampling from part 9 Fig. 1 Circular weld, connection of the new barrels, sampling from part 12 Fig. 1 Longitudinal weld, new barrel closed side, sampling from part 12 Fig. 1 Longitudinal weld, new barrel closed side, sampling from part 14 (sample 52) and 17 samples 53 and 54) Fig. 1 Longitudinal weld, new barrel closed side, sampling from part 22 Fig. 1 Longitudinal weld, old barrel bottom side connected to new barrel, sampling from part 27 Fig. 1
56 70
Samples were metallographically prepared and investigated microscopically. Micrographs of the cracks in polished and etched sections were taken. Etching was done by alcoholic sulfuric acid. The microstructure of the welds from the old and new parts of the vessel was compared. 2.4. Mechanical testing 2.4.1. Hardness Vickers hardness HV1 according to DIN 50133-2 was determined using the cross sections from Section 2.3. The testing was conducted with the small load hardness tester Finotest. Table 2 contains the values determined: 2.4.2. Bend test By means of bend tests the deformation- and fracture behavior of the longitudinal welds on the inner surface during quasistatic load was compared. Three bend samples were taken, two from the opening side new barrel (parts 2 and 7
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Fig. 3. fracture surface from the area of the new barrel with long time periodical crack and residual fracture.
Fig. 1) and one from a bottom side old barrel (part 27 Fig. 1). The 50 mm wide samples were used as taken without any surface treatment including the weld seams. 2.4.3. Notched bar impact test The absorbed impact energy was determined for the boiler sheets of the old and new barrels as well as the HAZ of the longitudinal weld (notch layer parallel to the weld, notch ground in the HAZ) of a new and an old barrel on longitudinal and circumferal oriented ISO-V-specimen at 0 °C (partially at 20 °C as well). DIN 17155 (1983) was decisive for the choice of sample design and test temperature. Experiments were conducted according to DIN 50115 using a Losenhausen type 300 J pendulum impact tester. Places of sampling and the results are given in Tables 3 and 4. 2.4.4. Tensile test To investigate the strength and deformation properties given by DIN 17155 for sheets from 1.4081 flat samples were taken from old and new barrels in longitudinal and circumferal directions and tensile tests were conducted. Tests were realized according to DIN 50146 in the 200 kN-range of a DIN 51221 1000 kN Instron type testing machine. Places of sampling and the results are given in Table 5. 2.5. Additional information By assessment of available information the following additional information was gathered: 1. Measurements concerning roofing of the exploded vessel had not been conducted. Measurements taken after the explosion showed roofing to the extent of 5 mm for vessel no. 4, elongated in 1983 and 7 mm in one singular place. Vessel no. 5 is said to have shown roofing up to 10 mm.
Table 2 Results of hardness measurements. Sample no.
30 31 32 33 51b 52 53 54c 56 70 a b c
Hardness HV1 in Base material
HAZa
Filler
195–211 169–185 174–180 164–174 202–213 202–247 175–182 185–187 191–198 167–183 173–193
248 237 226 227 257 267 228 229 247 – 234
210–215
Max hardness value in the heat affected zone. Additional heat influence in the sampling area. Severe plastic deformation (neckdowns) due to the crack progress.
210–220 200–220 235–245 200–220 200–215 210–225 220–240 – 190–200
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Table 3 Results of impact test. Sample no.a
Place of sampling Base metal; Opening side old barrel (sampling part 1 Fig. 1)
11 Q 12 Q 13 Q 14 Q 11 Q 12 Q 13 Q 15 L 16 L 17 L 18 L Base metal; Opening side new barrel (sampling part 12 Fig. 1) 31 Q 32 Q 33 Q 34 Q 35 L 36 L 37 L Notch impact work for sheets from 1.4081 acc. to DIN 17155 (1983) with thickness 660 mm; circumferal samples a b
Temperature (°C)
Impact work
0 0 0 0 0 20 20 0 0 20 20 0 0 0 0 0 0 0 0
12.5 25.5 19.5 22.0 19.0 50.0 33.0 13.8 14.8 50.0 33.0 >143b >143b 216 223 232 254 268 P31
Q: Circumferal sample (sample long. axis parallel to vessel long. axis); L: Longitudinal sample sample long. axis in barrel circ. axis). Samples bent and partially broken.
Table 4 Results of impact test. Place of sampling
Sample no.a
Temperature (°C)
Impact work
Base metal; Closed side new barrel (sampling part 14 Fig. 1)
51 52 53 55 56 57 71 72 73 74 75 76 77 61 62 62 64 65 61 62 62 64 65
0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0
162 209 219 246 199 250 45.0 40.0 59.0 50.0 74.5 64.5 78.0 204 44.0 53.0 32.5 210 76.0 71.0 39.0 36.0 40.0
Base metal; Closed side old barrel (sampling part 27 Fig. 1)
Longitudinal weldb Closed side new barrel (sampling part 17 Fig. 1)
Longitudinal weld
a b
b
Closed side old barrel (sampling part 27 Fig. 1)
Q Q Q L L L Q Q Q Q L L L
Q: Circumferal sample (sample long. axis parallel to vessel long. axis); L: Longitudinal sample sample long. axis in barrel circ. axis). Notch at the inner sheet surface parallel to the weld in the HAZ.
2. The hardening process can result in a pressure increase up to 1.5 bar. The factory itself never observed temperature increase >1 K corresponding to a pressure increase of 0.1 bar. The operator excludes the possibility of pressure increase due to the operation of an open system. 3. Due to the regulation devices including safety valves at the steam generator of the operator (being periodically inspected) exceeding of the maximum operating pressure of the hardening vessel is to be expected in exceptional cases. The maximum pressure will stay below 18 bar in any case. 4. The water used in the steam generator is treated by a hydrogen- and sodium exchanger. The feed water is degassed to remove oxygen, phosphates and a oxygen scavenger are added to remove residual hardness and oxygen. 5. Live steam is provided as saturated steam. Maximum temperature in the stone hardening vessel is 203.4 °C. Steam temperature is measured and registered at the stone hardening vessel. (Temperature reading of 203 °C is existing for the time of failure referring to operators claims.)
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Sample no.a
Yield strength Rp0,2 (N mm2)
Tensile strength RM (N mm2)
Elongation on fracture A5 %
Reduction of area Z%
Opening side new barrel (sampling part 1 Fig. 1)
46 47 48 49 48 49 66 67 68 69 68 69 86 87 88 89
404 396 418 387 352 354 377 387 376 375 400 421 451 467 380 373
556 551 517 514 512 514 519 518 525 522 511 513 546 562 538 533
25.5 25.5 27.5 27.0 30.5 29.0 25.5 26.0 26.5 29.5 27.5 30.5 20.0 21.0 29.5 26.5
62 60 69 66 68 65 63 64 63 63 68 71 58 58 51 53
Opening side new barrel (sampling part 1 Fig. 1 48 L + 49 L and part 12 Fig. 1 48 Q + 49 Q)
Closed side new barrel (sampling part 14 Fig. 1 66 L – 69 Q and part 18 Fig. 1 68 L + 69 L)
Closed side old barrel (sampling part 18 Fig. 1 86 L + 87 L and part 27 Fig. 1 88 Q + 89 Q)
a
L L L L Q Q L L Q Q L L L L Q Q
Q: Circumferal sample (sample long. axis parallel to vessel long. axis); L: Longitudinal sample sample long. axis in barrel circ. axis).
3. Results 3.1. Evaluation of in-service data The evaluation of data provided from periodic inspection gave no reference for any problems in the operational condition of the vessel. The periodic inspection shortly prior to the explosion did not detect any crack or problems with the vessel in question. 3.2. Chemical analysis The boiler plates (samples no. 42 and 61 acc. to Table 1) meet the requirements for the chemical composition of the material 1.4081 (17Mn4) according to DIN 17155 (1983). The values determined meet the values given in the acceptance certificate. The filler material of the inner longitudinal welds (sample no. 43 and 60 acc. to Table 1) show a carbon and aluminum content slightly below the minimal values and a silicon value touching the upper limit or slightly above respectively, if one assumes that base material and filler should be of the same chemical composition. The deviations according to DIN 17155 (1983) and in reference to the base material do not give any indication of a wrong filler material nor would they allow the assumption of a deterioration concerning the fracture properties. 3.3. Fractography SEM investigations showed a honeycomb structure giving evidence of ductile fracture could be detected outside the overload fracture and in the area in front of the shear lip in the middle part of the crack surface as well as in the origin of fracture (Figs. 3 and 6). Line formation in front of the shear lip (Fig. 5) differing clearly from the much finer microstructure layers indicate a stepwise crack propagation in this area (Fig. 4). Line formation in front of the shear lip (Fig. 5) could not be observed in all fracture areas; for the three samples investigated by SEM these were visible in the samples taken from part 22 (Fig. 1). Fatigue striations characteristic for alternate loads differ from the line formation visible in this case. Features of crack propagation allowing conclusions concerning the period of crack formation could not be found. 3.4. Metallography 3.4.1. Microscopic evaluation Microscopic evaluation of the cross-sections confirmed the macroscopic visible finding that all cracks developed from the inner surface of the vessel in the coarse grain area of the heat affected zone directly adjacent to the upper weld layer. Figs. 7 and 8 show two cracks in polished and etched section. The path of the crack in the pictures as well as in Figs. 9 and 10 clearly shows that the cracks started vertically from the surface into the material as well as partial branching. Apart from the crack origin the cracks are independent of the microstructure of the material in the welding area. The propagation of the cracks
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Fig. 4. Coarse line formation near the shear lip (sample 55, Fig. 1).
Fig. 5. Fine line formation in the middle area of the crack surface (sample 57, Fig. 1).
Fig. 6. Structure in the shear lip of sample 55 (Fig. 1).
thus follows a transgranular path. Figs. 11 and 12 show severe corrosive attack at the flanks of the crack into the small branches of the side cracks. The crack in Fig. 7 is heavily corroded as well. In order to determine, whether critical microstructural states influenced the crack initiation, the microstructure of the longitudinal welds of old and new parts of the vessel were compared. The comparison did not show any differences concerning the microstructure. No weld faults such as undercutting, slag inclusions or lack of fusion were detected. Figs. 13 and 14 show the state of the surface in comparison between inner and outer surface. Fig. 13 shows formation of pits due to corrosion at the inner surface whereas Fig. 14 shows the smooth outer surface of the vessel at the same sample.
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Fig. 7. Corroded crack starting from coarse grain zone in the transition weld-sheet, sample 32. Left polished, right etched.
Fig. 8. Crack initiation without branching, sample 52. Left polished, right etched.
3.5. Estimation of loads and stresses due to misalignment and roofing 3.5.1. Geometrical analysis The geometrical relations after the failure in the area of longitudinal welds of the new barrels and the bottom side old barrel are shown in an exemplary manner in Fig. 15 in 1:1 scale. Two geometrical characteristics causing locally tension increase can be observed, namely partially observed misalignment, see Fig. 12, part 4 as well as roofing, being more pronounced in the new barrels, Fig. 12, part 13/21.
3.5.2. Stress analysis The misalignment – the absolute values could not be determined due to the overlaying roofing (angle failures) – leads to a superimposed bending stress and, due to the generally sharper notch formation, to an increased notch stress. The amount of the additional bending stress increases linear with the misalignment as shown in the derivation given in the following
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Fig. 9. Crack with branching, initiation upper left (inner surface of the vessel). Etched, sample 52.
Fig. 10. Crack with branching, initiation upper right (inner surface of the vessel). Etched, sample 56.
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Fig. 11. Crack with branching and shear lip (see Fig. 6). Sample 52.
Fig. 12. Crack branchings with corrosion products (sample 56).
Fig. 13. Pitting corrosion on the inner surface of the vessel (sample 32, new barrel).
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Fig. 14. Outer surface of the vessel without corrosion (sample 32, new barrel).
Fig. 15. Geometric relations in the area of the longitudinal welds of the new barrels (Neuer Kesselschuss) and the adjacent old barrel (Alter Kesselschuss) Inside of the vessel is always at the bottom.
schematic, see Illustration I. Without consideration of the notch, expressed with the stress concentration factor aK, the resulting boundary stress r0 results from the nominal tensile stress rz and the bending stress rb to
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r0 ¼ rz þ rb ¼
145
F 6Fe 6e þ 2 ¼ rz 1 þ bs bs s
For a misalignment e of 1 mm and a wall thickness s of 13.5 mm the boundary stress r0 would result in 1.44-fold of the nominal tensile stress. To determine the roofing that resulted from inadequate smooth curving of the sheet ends in the area of longitudinal welds in the following analytical stress estimation [11] the vessel geometry according to Illustration II was assumed. The amount of roofing h was calculated from the roofing angle formed by the sheet ends adjacent to the weld seam area (see Fig. 12, after failure). The following roofing amounts were determined: (a) Opening side new barrel: h 4.63 mm. (b) Bottom side new barrel: h P 8.63 mm. (c) Bottom side old barrel: h 1.37 mm. Stress estimation according to a roofing amount h at the vessel: Eccentricity k ¼ hr 1 according to the definition in technical codes, i.e. AD-Merkblatt H1
r 1 ¼ rh 1þk qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 2k þ k sin a ¼ 1þk qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 tan a ¼ 2k þ k qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 1 ¼ r 2k þ k cos a ¼
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 P2 ¼ pl ¼ pr 2k þ k : Partial force resulting from inner pressure causing bending stress due to roofing. Components of the partial force P1 resulting from inner pressure, perfect geometry assumed:
P1X ¼
R pa
P1Y ¼
R pa
0 0
p cos urdu ¼ pr sin a ¼ pr
pffiffiffiffiffiffiffiffiffiffi 2kþk2 1þk
2 p sin urdu ¼ prð1 þ cosaÞ ¼ pr 1þk
Equilibrium at partial system III:
F m sin a ¼ P1X ¼ pr sin a F m ð1 þ cos aÞ ¼ P1Y ¼ prð1 þ cos aÞ F m ¼ pr
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Equilibrium at partial system IV:
Shear force
Q ¼ P 2 ¼ P2 ¼ pr
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2k þ k
2
2
Bending moments M ¼ P2 12 ¼ p 12 ¼ 12 pr2 ð2k þ k Þ Highest tensile boundary stress (tensile and bending)
rmax ¼
F m 6M þ 2 s s rh s
r s
rmax ¼ p 1 þ 3 ð2k þ k2
i
For k 1 approximately
rmax ¼ p
r r 1þ6 k s s
mm For rs ¼ 1000 ¼ 74:1 and 13:5 mm
rm p rs (Barlow´s formula, see AD-Merkblatt B1
rmax ¼ 1 þ 444:4k rm In the actual case with r = 1000 mm the following numerical equation results: rmax rm ¼ 1 þ 0:444h if h is inserted in mm.Following the given derivation this results in boundary tension stress rmax for the roofing amounts given above (a) 3.06 (b) 4.84 (c) 1.61 times the calculated mean tensile stress rm. As the bending stresses due to misalignment and roofing add to the mean tensile stress rm from the vessel formula and the resulting stress is additionally enhanced locally by notch effects (transition of sheet to weld, weld seam, undercutting, corrosion marks) one has to assume the stress at the inner surface of the two new barrels in the transition between sheet and weld exceeded the calculated nominal values significantly. 3.6. Mechanical testing 3.6.1. Hardness The determined hardness of the base material is some 7–8% higher as would result from calculation from the tensile strength according to DIN 50150 (08/1976). This is mainly caused by the influence of the low test force for which calculation according to DIN 50150 is not permitted and by the strain hardening of the sample surfaces due to mechanical preparation (grinding). As this is applicable to the HAZ and the filler as well, the relation between the hardness of the three different zones is valid. Due to coarse grain formation and the higher hardness values in the HAZ the weld area reveals a weak spot in the construction, nevertheless the absolute values of hardness are so low that a risk of failure due to lower plastic deformation properties, stress corrosion or hydrogen embrittlement is not existent. Sample no. 51 taken in some 150 mm distance to the circular weld connecting the two new barrels shows a heat treatment independent of the longitudinal welding. This is the probable cause for the elevated hardness levels in base material as well as in the HAZ in comparison to the other samples. Between the sampling area and the circular weld an additional electric hand-weld seam is located on the outside of the vessel. Both facts, heat treatment and hand weld allow the conclusion of additional alignment in the area of the closed side new barrel prior to joining.
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3.6.2. Bend tests The samples were bent under identical test conditions until the onset of microscopic visible cracks. Crack formation always happened by shear crack formation in the notch formed by weld seam and sheet surface, Fig. 16. The deformation until formation of cracks was larger for samples 2 and 7 as for sample 27 with the difference between the first two samples resulting from different notch effect due to different excess weld metal. 3.6.3. Notched bar impact test The minimal notch impact work for the material according to DIN 17155 (1983) at 0 °C is not met by the boiler sheets of the opening side old barrel in longitudinal nor circumferal direction. The 1959 version of DIN 17155 valid until 1983 requires a minimal notch impact work of 34.3 J at room temperature for the DVM sample. As this value is met at 20 °C for the ISO-V specimen, the sheet from the old barrel does meet the 1959 requirements concerning notch impact work, these sheets being of no concern for failure initiation anyway. With clearly differing values for the notch impact work between old and new barrels, showing better values for the new barrel material, the other tested sheets cannot be objected either. The HAZ shows a sufficient notch impact work as well, giving higher results for 0 °C than the base material of the opening side old barrel. 3.6.4. Tensile test The tensile strength values determined meet the requirements of DIN 17155 (1983). 0.2% yield point shows an elevated level. This is caused by prior plastic cold working caused by curving, the failure and the need for straightening of the tensile samples prior to testing.
4. Discussion 4.1. Failure sequence The explosion of the vessel took place after preliminary weakening of the vessel wall by cracks due to a long-term process, i.e. not in one operating cycle generated. The cracks were formed on the inner surface of the added new barrels for the elongation of the vessel starting from the notch between sheet surface and longitudinal weld. Crack propagation went almost perpendicular to the surface into the wall of the vessel. The crack in the starting area of the final defect was some 1.6 m long and resulted in a partial residual wall thickness of 1.5 mm. Formation and propagation of the cracks in the new barrels happened due to low cycle fatigue mechanism, macroscopically without deformation, crack propagation during the explosion took place exclusively after preliminary extensive plastic deformation (ductile).
Fig. 16. Bend tests with 50 mm length samples until visible crack formation.
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4.2. Failure mechanism and root cause The results of all Investigations give strong evidence to corrosion fatigue as the main failure mechanism leading to the explosion. Due to geometrical deviations (misalignment and roofing at the longitudinal weld) local mechanical stresses were significantly higher than designed. This caused low cycle fatigue and crack initiation. Other reasons such as material or welding flaws or a singular overload can be excluded. Corrosion fatigue [1,2] is described in DIN EN ISO 8044 [3] as localized corrosion with formation of cracks as a result of mechanical damage of protecting layers by cyclic critical stress or shrinking of a component. It is always observed in areas of high local stress concentrations, the release of which results in small locally concentrated plastic deformation. This type of corrosion is a fatigue with a low number of load cycles (LCF = low cycle fatigue) combined with high mechanical stress as observed in this case. Crack propagation and appearance of the cracks thus are frequency dependent (resulting in differing amounts of load cycles): Low cycle numbers and low tension amplitude result in a corrosion dominated mechanism, high cycle numbers and high tension amplitudes result in a mechanically dominated mechanism. As the cracking of the protective layers is the major cause for crack initiation, the known operating conditions result in the following failure mechanism: 1. Formation of a protective layer (magnetite) at low pressure by reaction of steam with steel. 2. Cracking of the brittle protective layer in the area of corrosion pits during pressure increase to 16 bar in areas of high local stress or high excess tensile load respectively. 3. Local corrosion at 16 bar. Tensile induced anodic dissolution in the formed notch root can be involved (stress corrosion). 4. Standstill corrosion caused by water in the crack (corrosion of the crack walls without pressure. Necessary for the cracking of the protective layer are at least small plastic deformations of the base material. Corrosion fatigue with crack formation can be observed especially in components with instationary operating conditions resulting in a cyclic repetition of steps 1–4. Literature shows that oxygen containing water at high temperature is sufficient to enable crack formation and propagation under given mechanical conditions. 4.3. Answers to objectives i. The vessel exploded due to LCF caused by mechanical stress caused by geometrical deviations of the new barrels. ii. Roofing as observed here is the cause of a not correct fabrication of the sheets for the vessel barrels. Following the accepted technical regulations the radius of the sheet ends has to follow the radius of the vessel exactly. Roofing caused by the deviation of roundness is not yet included in standardization at the time of the failure. a. An offset between the inner and outer weld seam had no influence on the failure of the vessel. b. Flattening causing a lack of roundness in the area of the longitudinal welds of the new barrels of the vessel is the root cause for the failure. c. A crack in the weld between the outer strengthening rings and the vessel that might have existed prior to the failure could not be detected. d. Use of a wrong filler material can be excluded. iii. 1. It would have been possible to detect the geometrical deviations during inspection. Until the day of the failure a geometrical inspection that would have revealed such flaws was not part of the acceptance test nor the periodical testing. 2. The inner inspection shall reveal damages caused during operation of the device. Whether the cracks in the area of the longitudinal weld were definitely existent and visible during the last inspection is not clear. iv. If these detectable deficiencies would have been addressed by the inspecting people and corrected, the explosion could have been avoided. v. The explosion was caused by geometrical deviations of the components for the new barrels of the hardening vessel. LCF was caused by the geometrically induced stress and the cyclic operation conditions resulted in the formation of cracks leading to the failure of the vessel. Thus the failure was to be expected considering existing manufacturing flaws and operation conditions. 4.4. Reference to similar failure events Some years later, in October 1991, a pressure vessel for hydrogen exploded in Hanau, Germany, causing major damage to almost all buildings of a large facility and other buildings in a distance up to 1000 m [4]. As the explosion took place on a Saturday morning, nobody was working near to the vessel thus causing no personal injuries. The cause for the explosion was the same as in the case described above: roofing of the longitudinal welds of the vessel leading to low cycle fatigue of the weld and subsequent to the failure of the vessel. The general problem of excess tensile stress due to mechanical deviations, especially roofing, is described in the regulations at least since 1995 [5]. Nevertheless, these regulations did
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not take into account roofing below given values for discrete dimension nor the influence of the medium that is stored in the vessel. Thus a miscalculation of the possible load cycles of more than one order of magnitude still was possible. Due to this failure, in 1998 a comprehensive research report dealing with geometrical deviations of pressure vessels was published [6]. Here descriptions of the mechanical deviations are given and mathematical methods for the calculation of the excess loads are presented. The report provides a comprehensive overview of the possible mathematical models and shows solutions for the calculation of the excess stresses caused by the geometrical deviations. 4.5. Reference to A. Martens The process of gathering all data from a reconstruction of construction and failure case subsequently followed by the determination of mechanical data for the materials involved leads consequently to the conclusion of one possible failure cause. The methodology follows strictly the procedures described by Adolf Martens: Reconstruct the failure case, interpret the data and you will find the conclusion. 4.6. Recommendations and preventive actions The results show clearly that even small deviations of roundness result in excess loads on vessel walls. It is necessary to avoid such geometrical deviation in mm-scale for pressure vessels. The problem is described in literature for many years [6–8]. 4.7. Implications for technical regulations The impact of these geometrical deviations were included in German [9] and British standards [10] some years after the incidents. Mathematical methods available nowadays [5] allow exact calculations of the impact of geometrical deviations of pressure vessels. The regulations available today include the determination of such deviations. 5. Conclusions The stone hardening vessel exploded due to a final ductile fracture. Prior to the final fracture, weakening of the vessels wall by LCF took place, starting from a notch between longitudinal weld and wall sheets. Cause for the LCF crack initiation was a geometrical deviation of additional barrels for elongation of the vessel. The roofing of the wall geometry resulted in an excess stress in the respective areas 4.8 times higher than the calculated nominal stress of the vessel wall. Acknowledgements Special acknowledgements to the authors of the original expert report [11] J. Ziebs and K. Naseband and to Amtsgericht Marl for the declassification. References [1] Hickling J. Der Maschinenschaden 1982;55(2):95–105. [2] Hickling J, Blind D. EFC conference on environmental sensitive cracking problems in nuclear installations containing high temperature water Munich; 1984. [3] DIN EN ISO 8044. Corrosion of metals and alloys – basic terms and definitions; 1999. [4] BAM-Expert Report 4-218, 01.10.1992. Causes and cause of events of the explosion of a hydrogen tank on the area of XXX (Ursachen und Hergang des Berstens eines Wasserstofftanks auf dem Gelände der Firma XXX). [5] Weber M. Calculation of structural stress in cylindrical vessels with roofing and internal pressure, (Strukturspannungsberechnung bei zylindrischen Behältern mit Aufdachungen unter Innendruck), in German, BAM Forschungsbericht 224; 1998, ISBN 3-89701-159-X. [6] Haibach E. Fraunhofer Gesellschaft e.V., Laboratorium für Betriebsfestigkeit, FB-77; 1968. [7] Lange K. Lehrbuch der Umformtechnik Band 3 Blechumformung. Springer; 1975. S. 131. [8] Dubbel Taschenbuch für den Maschinenbau, Band II, Springer; 1953. S. 575. [9] AD-Merkblatt HP1. Construction and testing of pressure vessels – design and testing (Herstellung und Prüfung von Druckbehältern – Auslegung und Prüfung), Beuth Verlag; 1995. [10] BS PD 6493 (1991). Guidance on methods for assessing the acceptability of flaws in fusion welded structures, BSI 1991. [11] BAM-Expert Report 1.2/12960. About the causes of the explosion of a stone hardening vessel at XXX company (Über die Ursachen des Berstens (Zerknallens) eines Steinhärtekessels bei der Firma XXX) in German issued 9.6.1986, unpublished.