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Fatigue weld failure causes explosion of air receiver
IX)13-:Y44/83/0s040c-I0$03 00/o Pergamon Prey\
I Id.
FATIGUE WELD FAILURE CAUSES EXPLOSION OF AIR RECEIVER LEIGHTON E. SlSSOMt and JOSEPH T. SCARDINAS Tennessee Technological University, Cookeville, TN 38501,U.S.A Abstract-An air receiver tank failed catastrophically with the fracture originating at a fatigue crack located at a leg tack weld. The resulting explosion, equivalent to over 3 lb of TNT, did massive damage to a tire recapping plant and crippled two employees. The fracture origin was the classical semi-elliptical surface crack. The tack weld exhibited incomplete fusion, slag inclusion and excessive porosity. Operation of the controls and safety devices, leak-before-break criteria and weld embrittlement are considered.
1.THE EXPLOSION SCENE SHORTLY after startup of parallel air compressors on a cold morning (well below freezing), an air receiver tank exploded, devastating a tire recapping plant. Two employees, the only ones at the scene when the explosion occurred, were critically injured, suffering head injuries and broken legs. The resulting hole in the building roof is shown in Fig. 1. The wall which was toppled in the foreground was made of 12-in. concrete blocks. The welded structural steel framework of the building was deformed but remained together. Figure 2 is a view from the roof of the rubble inside the building where the air compressors/tanks were located. The major piece of the air receiver which ruptured is shown in Fig. 3. The vessel was 30 in. in outside diameter and 84 in. long with semi-ellipsoidal heads concave to pressure. The shell was 0.250-in. thick, the heads 0.201-in. It was made of SA-455-A, 75,000 psi tensile strength, steel[l]. The vessel had been appropriately inspected and registered by the National Board of Boiler and Pressure Vessel Inspectors on 27 April 1965. Figure 4 is a schematic showing the arrangement of the failed tank along with two other parallel units. Only two of the air compressors-20HP and IOHP-were on at the time of the explosion. The 15 HP unit was off. All these tanks were connected to increase the storage volume. 2. EXPLOSIVEYIELD The total volume of the three interconnected tanks was 65.37 ft3. For a maximum allowable pressure of 200 psi and a peak temperature of 650”F, the tanks contain 34.16 lbm of air. The maximum value for total energy release E when a compressed gas vessel bursts is given by[2] E=
P-Pli V ( r-1 )
where p is the initial absolute pressure in the vessel of volume V, y is the ratio of specific heats of the gas, and pu is the ambient absolute pressure. This gives an upper limit of blast yield E = 4.7 x IO6ft-lbf. This result is based on the assumption that all of the energy which can drive a blast wave does so, depending only on the energy release rate[3,4]. For the vessel in this case, a negligible amount of the energy was absorbed in the fracturing process because the vessel was already stressed to failure. But some of the energy was absorbed in accelerating vessel and compressor fragments[5], which were strewn up to 150ft from the explosion center. The effective blast yield E, is then given by E,=E-2
(fm2)
where u is the mean velocity of mass fragment m. Since all fragments were not found and since most of the fragments were slowed in their trajectories by colliding with building components or other fragments, it is not possible to correct for the kinetic energy of the vessel fragments. tprofessor of Mechanical Engineering and Dean of Engineering SProfessor of Engineering Science and Mechanics. 405
406
L. E. SISSOM and 3. T.
SCARDINA
*SAFETY VAiVE
PILOT VALVE
\,
1120
GAL.
Fig. 4.
We can get a feel, however, for the energy released in this explosion by considering propelling a block weighing 100Ibf from rest to a velocity of 60 mph. The kinetic energy required is
KE = 12,025ft-lbf. The number of such bfocks which could be so propelled, therefore, would be h! = 4.7 x lo6 ft-lbf = 391 ’ 12,025ft-lbf Tremendous destruction can obviously be done by propelling 391 100-lbf blocks from rest to 60 mph! 3. CONTROLS AND SAFETY DEVICES The compressed air system, as shown in Fig. 4, was protected from excessive pressure by a number of over-lapping, independently-activated components. The principal control unit was an automatic pilot valve, shown schematically in Fig. 5, which senses receiver pressure and admits compressed air to the inlet valve regulators when the required pressure has
L. E. SISSOM and I. T.
SCARDINA
6(a).
(b) Fig. 6. (I scale division =&in.)
Fig. 7. (1 scale.division = h in.)
Fatigue weld failure causes explosion of air receiver
(a)
(b) Fig. 8. (1 scale division = kin.)
Fig. 9. (1 scale division = $ in.)
409
411
Fatigueweld failure causes explosion of air receiver CHECK VALVE
Up To Pressure
OPEN
,
COMPRESSOR CYLINDER (Compressor RunningUnloaded Position) DIAGRAM -
INLET VALVE REGULATION
Fig. 5.
been established. The pilot valve in this case was set to unload the compressor at 185psig. (After the it was tested at 175psig.) This permitted the motor and compressor to operate continuously, although under no load when activated. The inlet valve regulators consist of a set of fingers arranged to hold the inlet valve plates open when activated. Air rushing in, on the piston downstroke, is pushed out again, on the piston upstroke, whenever the inlet valve is held open. Since no compression occurs, idling horsepower is very low. Each motor, which powered the 2-cylinder air compressors, was equipped with a pressure switch, set to deactivate the motor at 2OOpsig. Under design conditions, then, the pressure switches would only function if the pilot valve failed. If both the pilot valve and the pressure switches failed, there were safety relief valves throughout the system-one in each cylinder head, one in each tank and one in each intercooler between cylinders. All of these safety relief valves were set at pressures well below the tank’s tested pressure of 300psig. The combined capacity of the relief valves far exceeded the combined capacity of the three compressors. Table 1 shows the results of pressure tests conducted on some of these components which were retrieved after the explosion. Of all the safety valves tested, only one (item 3) could “feel” the receiver pressure. It relieved at 225 psig, 25 psi below the value stamped on its body and approved by the American Society of Mechanical Engineers (ASME). Its rated capacity of 145cfm is well above that which could be produced by both the compressors which were running. Since there was no damage on this safety valve, the maximum pressure which could have existed in the receivers was 225 psig. This would have been the case even if the pilot valve failed and if the pressure switches, set to deactivate the motors at 200 psig, failed to function-a highly improbable set of simultaneous malfunctions. explosion
4. FAILURE OF RECEIVER Most of the air receiver tank was available for fractographic study. A few missing pieces were apparently lost during the initial investigation by others or during the clean-up operation. All of the available fracture surfaces were initially examined macroscopically (1 x to 10 x magnification). Most of the fracture could be characterized as mixed-mode with some regions exhibiting well developed shear lips and others with little or no shear lip. It was also observed that one of the head welds had peeled back over an appreciable portion of its length (approx. 38 in. circumferentially). The peel-back was very straight with obvious separation of the weld bead from the base metal on one side of the weld as if the head weld had “unzipped.” Finally a semi-elliptical region approximately a in. long by i in. wide was found on the fracture surface where one of the legs had been tack welded to the tank. This particular morphology is characteristic of rapid fracture origins. In addition, faint chevron markings were observed on the fracture surface adjacent to the semi-elliptical region. These chevron markings were pointed back toward the semi-elliptical region. This was the only region of this type which was observed over the entire fracture surface. Based on these observations, it was concluded that this was the origin at which the fracture had initiated. The location and macroscopic features of this region are shown in Figs, 6-9.
L. E. SISWM
412
and J. T. SCARDINA Table 1.
. --__l.-.____ Pilot Valve (for constant speed control) Old, discarded valve--used to check test procedure--same as valve on exptoded system Pilot Valve From exploded
system (found
on roof)
Relief Pressure. [psig) 150
175
Safety Vatve--15 HP Receiver ASME: Set 250 psi; Cap. 145 CFM .
225
Safety
200
Valve--l5
HP Head
Safety Valve--l@ HP Head ASME: Seat Set at 100; Cap. Safety
Valve--20
HP Intercooler
Safety
Valve--W
HP Head
90+ 57 CFM * 210’
Pressure Switch (believed to be from 10 HP unit) : Contacts did not open upon pressurizing, but they closed when the pressure fell past 180 psig. They opened when pressurization was stopped at any time past 175 psig. The switch had suffered a substantial deforming blow during the explosion.
‘IWould not reseat. *Did not relieve at 300 psig.
The semi~lliptical region was subsequently examined at magnifications ranging from 10 x to 30 x using optical stereoscopic microscopy. Crack progression marks and ridges approximately perpendicular to these marks were observed in the material beneath the end of one of the tack weld beads. These features, which were present only on the curved, semi-elliptical surface, showed that the fracture origin had been initiated by cyclic loading. At the point where these features ended, a definite curved boundary was formed approximately & in. from the inside surface of the tank, This microscopic examination also revealed welding imperfections at a region of fracture surface a few inches away from the semi-elliptical region. These imperfections included inclusions and pores, some of which exceeded &in. in size, violating universally accepted standards~6~. Tensile tests on a sample taken from the base metal showed an ultimate tensile strength of 78,000 psi. The elongation at fracture in a 2 in. gauge length was 19.5%. These values are within the American Society for Testing Materials (ASTM) specifications for 455 carbon-manganese steel pressure vessel plate. Hardness readings were takan across the leg tack weld a short distance away from the fracture origin. The hardness ranged from 73-81 on the Rockwell B-scale in the base metal and from 90-96 on the Rockwell B-scale in the weld. Two weld coupons were taken from each head weld for standard bend tests. One face bend and one root bend from each head weld were made. Two of the four bend tests failed in that cracks exceeding $ in. developed on the convex surfaces 171. 5. LEAK-BEFORE-BREAK The semi-elliptical fatigue failure region shows that the surface crack nearly penetrated the wall of the air receiver tank. Had such a condition occurred it is quite possible that the tank would have simply leaked rather than fracturing ~~astrophi~lIy. Using the fracture mechanics leak-before-break criteria, it is possible to calculate the required minimum toughness level so that a crack will not become unstable before it penetrates a given thickness material. Rolfe and Barsom[S] list the required KI, values to satisfy the leak-before-break criterion. For a 4 in. thickness (the smallest thickness listed) of a material with a 40,000 psi yield strength, the required K,, value to satisfy this criterion is 35 ksi-in”* if the applied stress equals the yield strength and is 23 ksi-in “* if the applied stress equals one-half the yield strength. It is also possible to use the Irwin solution for a surface flaw in tension{91 to calculate directly the value of K,, for which the specific flaw found in the air receiver tank would become unstable. For a tank pressure of 200 psi and a wall thickness of $in., the nominal stresses in the cylindrical portion of the tank are 8000 psi in the longitudinal direction and 16,000psi in the circumferential direction. Since the flaw is near the head weld, because of the discontinuity due to the attached leg, and the possibility of residual
xxx
---
FRACTURE
WELD
Rm-UII1~
DRAIN
PLUG -\
ml
__.
Fig. 10. Bottom view of tank showing fracture origin and path.
HEAD WELD (“UNZIPPED”)
-
HEAD WELD (INTACT)
FRACTURE ORIGIN AT END OF LEG TACK WELD
/-
P w
414
L. E.
SISSOM
and
J. T. SCARDINA
stresses due to welding, it is reasonable to assume that the applied stress could approach the yield strength of 35,OOOpsifor SA-455-A steel. Using this value for the applied stress, the value of K,,. at which this particular flaw is unstable is 25 ksi-in”‘. Despite the somewhat questionable use of fracture mechanics calculations for a material with this low strength and thickness, these calculations do show that a very low fracture toughness was necessary to explain the catastrophic fracture of the tank. While specific fracture toughness data for SA-455-A steel were not available, a steel of this type would be expected to have a much higher fracture toughness (of the order of 100ksi-in”‘). Since it has been shown that the tank did fracture catastrophically, that an origin typical of such rapid fracture was found, and that the pressure in the tank did not exceed 200 psi; the question then became-what factor or factors could have caused such a low fracture toughness in the steel? This led to a reexamination of the weld attaching the leg to the tank at the region where the fracture originated. As stated earlier this was a tack weld rather than a continuous weld along the total perimeter of the leg. A visual inspection of the weld indicated that it was of poor quality with inadequate fusion between the weld and base metals (this opinion was independently confirmed by two certified welders). It is known that tack welds, particularly poorly applied tack welds, can cause the formation of microstructures which are much more brittle and crack sensitive than those associated with continuous welds[lO]. There are also documented cases of rapid fractures initiating at welds at stress levels significantly lower than would be calculated from fracture mechanics using base metal fracture toughness data[ll]. Thus, it was concluded that the fracture originated at the fatigue crack in an embrittled region adjacent to a tack weld at one of the legs and propagated circumferentially for several inches parallel to the tack weld in the embrittled material. At this point there was a crack large enough to cause rapid fracture to continue in the nonembrittled base metal, and the major fracture then propagated longitudinally along the tank due to the larger circumferential stress. This fracture is shown in Fig. IO. 6. CONCLUSIONS As a result of this investigation the following conclusions were reached. 0 The controls and safety devices were functioning properly at the time of the explosion. 0 The tensile properties of the SA-445-A steel were within the ASTM specifications for this steel. 0 The fracture origin was a semi-elliptical, surface fatigue crack which originated at a leg tack weld. 0 The tack weld, which exhibited incomplete fusion, slag includions and excessive porosity, was a poor quality weld. () Leak-before-break criteria were not met because of weld embrittlement. The fracture propagated circumferentially along the tack weld rather than through the tank wall because of embrittlement of the material surrounding the weld. Once the fracture reached a length of several inches it propagated longitudinally over the length of the pressure vessel and then around the heads of the tank. REFERENCES [I] Another air receiver explosion. National Bourri Bullefin. The National Board of Boiler and Pressure Vessel Inspectors, 35 : 2.Oct. 1911. [2] H. L. Brode, Blast wave from a spherical charge. PhyA. Fluids. 2. 217 (1959). [3] R. A. Strehlow, Accidental explosions. Am. Scienfisf 68(4), 420-428 (1980). [4] R. A. Strehlow, The characterization and evaluation of accidental explosions. Prog. Energy Combusfion Sci. 2, 27-60 (1976). (51 W. E. Baker ef al., Workbook for Estimating E&f\ of Accidenfal Explosions in Propellnnf Ground Handling und Tromporf Systems, NASA Contractor Report 3023, Aug. 1978. [6] ASME Boiler und Pressure Vessel Code, Section V1II. Division 1. p. 265. [7] ASME Boiler und Pressure Vessel Code, Section IX. [8] S. T. Rolfe and J. M. Barsom, Fracture and fatigue control in structures. In Applications of Fracture Mechanics, pp. 394-399 Prentice-Hall, Englewood Cliffs, New Jersey (1977). [9] B. R. Mullinix and D. G. Smith, Fracture mechanics design handbook, p, 194. U.S. Army Missile Research and Development Command Technical Report TL-77-S. Aug. 1977. [lo] M. J. Houle, Problems related to welding qualification, pp. 90-98. Proceedings, 49fh General Meeting, The National Board of Boiler and Pressure Vessel Inspectors, 5-9 May 1980. [II] F. W. Boulger, Fracture toughness comparison5 in steels. In Fracture: VoI. VI, Fracture of Metals (Edited by H. Liebowitz). Academic Press, New York (1969). (Received I I December 1981: received for publication
16 February 1982)
I Id.
FATIGUE WELD FAILURE CAUSES EXPLOSION OF AIR RECEIVER LEIGHTON E. SlSSOMt and JOSEPH T. SCARDINAS Tennessee Technological University, Cookeville, TN 38501,U.S.A Abstract-An air receiver tank failed catastrophically with the fracture originating at a fatigue crack located at a leg tack weld. The resulting explosion, equivalent to over 3 lb of TNT, did massive damage to a tire recapping plant and crippled two employees. The fracture origin was the classical semi-elliptical surface crack. The tack weld exhibited incomplete fusion, slag inclusion and excessive porosity. Operation of the controls and safety devices, leak-before-break criteria and weld embrittlement are considered.
1.THE EXPLOSION SCENE SHORTLY after startup of parallel air compressors on a cold morning (well below freezing), an air receiver tank exploded, devastating a tire recapping plant. Two employees, the only ones at the scene when the explosion occurred, were critically injured, suffering head injuries and broken legs. The resulting hole in the building roof is shown in Fig. 1. The wall which was toppled in the foreground was made of 12-in. concrete blocks. The welded structural steel framework of the building was deformed but remained together. Figure 2 is a view from the roof of the rubble inside the building where the air compressors/tanks were located. The major piece of the air receiver which ruptured is shown in Fig. 3. The vessel was 30 in. in outside diameter and 84 in. long with semi-ellipsoidal heads concave to pressure. The shell was 0.250-in. thick, the heads 0.201-in. It was made of SA-455-A, 75,000 psi tensile strength, steel[l]. The vessel had been appropriately inspected and registered by the National Board of Boiler and Pressure Vessel Inspectors on 27 April 1965. Figure 4 is a schematic showing the arrangement of the failed tank along with two other parallel units. Only two of the air compressors-20HP and IOHP-were on at the time of the explosion. The 15 HP unit was off. All these tanks were connected to increase the storage volume. 2. EXPLOSIVEYIELD The total volume of the three interconnected tanks was 65.37 ft3. For a maximum allowable pressure of 200 psi and a peak temperature of 650”F, the tanks contain 34.16 lbm of air. The maximum value for total energy release E when a compressed gas vessel bursts is given by[2] E=
P-Pli V ( r-1 )
where p is the initial absolute pressure in the vessel of volume V, y is the ratio of specific heats of the gas, and pu is the ambient absolute pressure. This gives an upper limit of blast yield E = 4.7 x IO6ft-lbf. This result is based on the assumption that all of the energy which can drive a blast wave does so, depending only on the energy release rate[3,4]. For the vessel in this case, a negligible amount of the energy was absorbed in the fracturing process because the vessel was already stressed to failure. But some of the energy was absorbed in accelerating vessel and compressor fragments[5], which were strewn up to 150ft from the explosion center. The effective blast yield E, is then given by E,=E-2
(fm2)
where u is the mean velocity of mass fragment m. Since all fragments were not found and since most of the fragments were slowed in their trajectories by colliding with building components or other fragments, it is not possible to correct for the kinetic energy of the vessel fragments. tprofessor of Mechanical Engineering and Dean of Engineering SProfessor of Engineering Science and Mechanics. 405
406
L. E. SISSOM and 3. T.
SCARDINA
*SAFETY VAiVE
PILOT VALVE
\,
1120
GAL.
Fig. 4.
We can get a feel, however, for the energy released in this explosion by considering propelling a block weighing 100Ibf from rest to a velocity of 60 mph. The kinetic energy required is
KE = 12,025ft-lbf. The number of such bfocks which could be so propelled, therefore, would be h! = 4.7 x lo6 ft-lbf = 391 ’ 12,025ft-lbf Tremendous destruction can obviously be done by propelling 391 100-lbf blocks from rest to 60 mph! 3. CONTROLS AND SAFETY DEVICES The compressed air system, as shown in Fig. 4, was protected from excessive pressure by a number of over-lapping, independently-activated components. The principal control unit was an automatic pilot valve, shown schematically in Fig. 5, which senses receiver pressure and admits compressed air to the inlet valve regulators when the required pressure has
L. E. SISSOM and I. T.
SCARDINA
6(a).
(b) Fig. 6. (I scale division =&in.)
Fig. 7. (1 scale.division = h in.)
Fatigue weld failure causes explosion of air receiver
(a)
(b) Fig. 8. (1 scale division = kin.)
Fig. 9. (1 scale division = $ in.)
409
411
Fatigueweld failure causes explosion of air receiver CHECK VALVE
Up To Pressure
OPEN
,
COMPRESSOR CYLINDER (Compressor RunningUnloaded Position) DIAGRAM -
INLET VALVE REGULATION
Fig. 5.
been established. The pilot valve in this case was set to unload the compressor at 185psig. (After the it was tested at 175psig.) This permitted the motor and compressor to operate continuously, although under no load when activated. The inlet valve regulators consist of a set of fingers arranged to hold the inlet valve plates open when activated. Air rushing in, on the piston downstroke, is pushed out again, on the piston upstroke, whenever the inlet valve is held open. Since no compression occurs, idling horsepower is very low. Each motor, which powered the 2-cylinder air compressors, was equipped with a pressure switch, set to deactivate the motor at 2OOpsig. Under design conditions, then, the pressure switches would only function if the pilot valve failed. If both the pilot valve and the pressure switches failed, there were safety relief valves throughout the system-one in each cylinder head, one in each tank and one in each intercooler between cylinders. All of these safety relief valves were set at pressures well below the tank’s tested pressure of 300psig. The combined capacity of the relief valves far exceeded the combined capacity of the three compressors. Table 1 shows the results of pressure tests conducted on some of these components which were retrieved after the explosion. Of all the safety valves tested, only one (item 3) could “feel” the receiver pressure. It relieved at 225 psig, 25 psi below the value stamped on its body and approved by the American Society of Mechanical Engineers (ASME). Its rated capacity of 145cfm is well above that which could be produced by both the compressors which were running. Since there was no damage on this safety valve, the maximum pressure which could have existed in the receivers was 225 psig. This would have been the case even if the pilot valve failed and if the pressure switches, set to deactivate the motors at 200 psig, failed to function-a highly improbable set of simultaneous malfunctions. explosion
4. FAILURE OF RECEIVER Most of the air receiver tank was available for fractographic study. A few missing pieces were apparently lost during the initial investigation by others or during the clean-up operation. All of the available fracture surfaces were initially examined macroscopically (1 x to 10 x magnification). Most of the fracture could be characterized as mixed-mode with some regions exhibiting well developed shear lips and others with little or no shear lip. It was also observed that one of the head welds had peeled back over an appreciable portion of its length (approx. 38 in. circumferentially). The peel-back was very straight with obvious separation of the weld bead from the base metal on one side of the weld as if the head weld had “unzipped.” Finally a semi-elliptical region approximately a in. long by i in. wide was found on the fracture surface where one of the legs had been tack welded to the tank. This particular morphology is characteristic of rapid fracture origins. In addition, faint chevron markings were observed on the fracture surface adjacent to the semi-elliptical region. These chevron markings were pointed back toward the semi-elliptical region. This was the only region of this type which was observed over the entire fracture surface. Based on these observations, it was concluded that this was the origin at which the fracture had initiated. The location and macroscopic features of this region are shown in Figs, 6-9.
L. E. SISWM
412
and J. T. SCARDINA Table 1.
. --__l.-.____ Pilot Valve (for constant speed control) Old, discarded valve--used to check test procedure--same as valve on exptoded system Pilot Valve From exploded
system (found
on roof)
Relief Pressure. [psig) 150
175
Safety Vatve--15 HP Receiver ASME: Set 250 psi; Cap. 145 CFM .
225
Safety
200
Valve--l5
HP Head
Safety Valve--l@ HP Head ASME: Seat Set at 100; Cap. Safety
Valve--20
HP Intercooler
Safety
Valve--W
HP Head
90+ 57 CFM * 210’
Pressure Switch (believed to be from 10 HP unit) : Contacts did not open upon pressurizing, but they closed when the pressure fell past 180 psig. They opened when pressurization was stopped at any time past 175 psig. The switch had suffered a substantial deforming blow during the explosion.
‘IWould not reseat. *Did not relieve at 300 psig.
The semi~lliptical region was subsequently examined at magnifications ranging from 10 x to 30 x using optical stereoscopic microscopy. Crack progression marks and ridges approximately perpendicular to these marks were observed in the material beneath the end of one of the tack weld beads. These features, which were present only on the curved, semi-elliptical surface, showed that the fracture origin had been initiated by cyclic loading. At the point where these features ended, a definite curved boundary was formed approximately & in. from the inside surface of the tank, This microscopic examination also revealed welding imperfections at a region of fracture surface a few inches away from the semi-elliptical region. These imperfections included inclusions and pores, some of which exceeded &in. in size, violating universally accepted standards~6~. Tensile tests on a sample taken from the base metal showed an ultimate tensile strength of 78,000 psi. The elongation at fracture in a 2 in. gauge length was 19.5%. These values are within the American Society for Testing Materials (ASTM) specifications for 455 carbon-manganese steel pressure vessel plate. Hardness readings were takan across the leg tack weld a short distance away from the fracture origin. The hardness ranged from 73-81 on the Rockwell B-scale in the base metal and from 90-96 on the Rockwell B-scale in the weld. Two weld coupons were taken from each head weld for standard bend tests. One face bend and one root bend from each head weld were made. Two of the four bend tests failed in that cracks exceeding $ in. developed on the convex surfaces 171. 5. LEAK-BEFORE-BREAK The semi-elliptical fatigue failure region shows that the surface crack nearly penetrated the wall of the air receiver tank. Had such a condition occurred it is quite possible that the tank would have simply leaked rather than fracturing ~~astrophi~lIy. Using the fracture mechanics leak-before-break criteria, it is possible to calculate the required minimum toughness level so that a crack will not become unstable before it penetrates a given thickness material. Rolfe and Barsom[S] list the required KI, values to satisfy the leak-before-break criterion. For a 4 in. thickness (the smallest thickness listed) of a material with a 40,000 psi yield strength, the required K,, value to satisfy this criterion is 35 ksi-in”* if the applied stress equals the yield strength and is 23 ksi-in “* if the applied stress equals one-half the yield strength. It is also possible to use the Irwin solution for a surface flaw in tension{91 to calculate directly the value of K,, for which the specific flaw found in the air receiver tank would become unstable. For a tank pressure of 200 psi and a wall thickness of $in., the nominal stresses in the cylindrical portion of the tank are 8000 psi in the longitudinal direction and 16,000psi in the circumferential direction. Since the flaw is near the head weld, because of the discontinuity due to the attached leg, and the possibility of residual
xxx
---
FRACTURE
WELD
Rm-UII1~
DRAIN
PLUG -\
ml
__.
Fig. 10. Bottom view of tank showing fracture origin and path.
HEAD WELD (“UNZIPPED”)
-
HEAD WELD (INTACT)
FRACTURE ORIGIN AT END OF LEG TACK WELD
/-
P w
414
L. E.
SISSOM
and
J. T. SCARDINA
stresses due to welding, it is reasonable to assume that the applied stress could approach the yield strength of 35,OOOpsifor SA-455-A steel. Using this value for the applied stress, the value of K,,. at which this particular flaw is unstable is 25 ksi-in”‘. Despite the somewhat questionable use of fracture mechanics calculations for a material with this low strength and thickness, these calculations do show that a very low fracture toughness was necessary to explain the catastrophic fracture of the tank. While specific fracture toughness data for SA-455-A steel were not available, a steel of this type would be expected to have a much higher fracture toughness (of the order of 100ksi-in”‘). Since it has been shown that the tank did fracture catastrophically, that an origin typical of such rapid fracture was found, and that the pressure in the tank did not exceed 200 psi; the question then became-what factor or factors could have caused such a low fracture toughness in the steel? This led to a reexamination of the weld attaching the leg to the tank at the region where the fracture originated. As stated earlier this was a tack weld rather than a continuous weld along the total perimeter of the leg. A visual inspection of the weld indicated that it was of poor quality with inadequate fusion between the weld and base metals (this opinion was independently confirmed by two certified welders). It is known that tack welds, particularly poorly applied tack welds, can cause the formation of microstructures which are much more brittle and crack sensitive than those associated with continuous welds[lO]. There are also documented cases of rapid fractures initiating at welds at stress levels significantly lower than would be calculated from fracture mechanics using base metal fracture toughness data[ll]. Thus, it was concluded that the fracture originated at the fatigue crack in an embrittled region adjacent to a tack weld at one of the legs and propagated circumferentially for several inches parallel to the tack weld in the embrittled material. At this point there was a crack large enough to cause rapid fracture to continue in the nonembrittled base metal, and the major fracture then propagated longitudinally along the tank due to the larger circumferential stress. This fracture is shown in Fig. IO. 6. CONCLUSIONS As a result of this investigation the following conclusions were reached. 0 The controls and safety devices were functioning properly at the time of the explosion. 0 The tensile properties of the SA-445-A steel were within the ASTM specifications for this steel. 0 The fracture origin was a semi-elliptical, surface fatigue crack which originated at a leg tack weld. 0 The tack weld, which exhibited incomplete fusion, slag includions and excessive porosity, was a poor quality weld. () Leak-before-break criteria were not met because of weld embrittlement. The fracture propagated circumferentially along the tack weld rather than through the tank wall because of embrittlement of the material surrounding the weld. Once the fracture reached a length of several inches it propagated longitudinally over the length of the pressure vessel and then around the heads of the tank. REFERENCES [I] Another air receiver explosion. National Bourri Bullefin. The National Board of Boiler and Pressure Vessel Inspectors, 35 : 2.Oct. 1911. [2] H. L. Brode, Blast wave from a spherical charge. PhyA. Fluids. 2. 217 (1959). [3] R. A. Strehlow, Accidental explosions. Am. Scienfisf 68(4), 420-428 (1980). [4] R. A. Strehlow, The characterization and evaluation of accidental explosions. Prog. Energy Combusfion Sci. 2, 27-60 (1976). (51 W. E. Baker ef al., Workbook for Estimating E&f\ of Accidenfal Explosions in Propellnnf Ground Handling und Tromporf Systems, NASA Contractor Report 3023, Aug. 1978. [6] ASME Boiler und Pressure Vessel Code, Section V1II. Division 1. p. 265. [7] ASME Boiler und Pressure Vessel Code, Section IX. [8] S. T. Rolfe and J. M. Barsom, Fracture and fatigue control in structures. In Applications of Fracture Mechanics, pp. 394-399 Prentice-Hall, Englewood Cliffs, New Jersey (1977). [9] B. R. Mullinix and D. G. Smith, Fracture mechanics design handbook, p, 194. U.S. Army Missile Research and Development Command Technical Report TL-77-S. Aug. 1977. [lo] M. J. Houle, Problems related to welding qualification, pp. 90-98. Proceedings, 49fh General Meeting, The National Board of Boiler and Pressure Vessel Inspectors, 5-9 May 1980. [II] F. W. Boulger, Fracture toughness comparison5 in steels. In Fracture: VoI. VI, Fracture of Metals (Edited by H. Liebowitz). Academic Press, New York (1969). (Received I I December 1981: received for publication
16 February 1982)