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Scientific Reports volume 14, Article number: 20643 (2024 ) Cite this article conductive coating
To reduce the number of casualties in explosion accidents, blast-resistant shelters can be used to protect personnel in high-risk areas of petrochemical processing plants. In this work, the deformation behaviours of uncoated and polyurea-coated blast-resistant plates were studied through gas explosion tests. An ANSYS/LS-DYNA model of a polyurea-coated shelter was established, and the dynamic responses of the shelter under various explosion loads were analysed. A series of fuel–air explosion tests were carried out to investigate the explosion resistance of the full-scale shelter. The results showed that compared with the uncoated blast-resistant plate, the deformation of the polyurea-coated blast-resistant plate was significantly reduced. The overall deformation of the shelter was the central depression of the wall and the inward bending of the frame. The damage effect of a typical high-overpressure, low-duration load was greater than that of typical low-overpressure, long-duration load. The shelter remained intact under three repeated explosive loads, with cracks appearing on the inner wall but no collapse or debris splashing. The shock wave attenuation rate of the shelter reached over 90%, which could significantly reduce the number of indoor casualties.
Most of the materials involved in petrochemical processing plants are flammable and explosive hydrocarbon liquids and gases. The leakage of dangerous materials could lead to fire and explosion accidents, which often caused heavy casualties and economic losses1,2. Related statistics also show that 75% of the casualties in petrochemical plant accidents are related to the destruction of buildings caused by vapor cloud explosions3. However, a considerable number of existing buildings in petrochemical enterprises were only designed considering fire protection requirements but not to withstand blast loading. When it is difficult to implement explosion-resistant renovation for existing buildings, new blast-resistant shelters could also be used to replace the existing buildings4. Compared with cast-in-place reinforced-concrete shelters, prefabricated steel shelters involve short construction periods, which significantly reduces the cost of interrupted production. In addition, prefabricated conventional container-sized steel shelters also have the advantages of mobility, reusability, and the ability to be assembled into large explosion-resistant shelters.
Polyurea is a specific chemically cross-linked elastomeric copolymer by a rapid chemical reaction between a short chain amine and an isocyanate curative5. Due to its superior physical and mechanical properties, it is extensively used as a protective coating in many industrial applications, such as vehicles, ships, and buildings6,7,8,9,10. Wang et al.11 investigated the peak pressure for damage, failure modes, and failure mechanisms of spray-on polyurea-reinforced clay brick masonry unit walls subjected to blasts. A polyurea layer can significantly improve the blast resistance of a wall and turn the collapse of unreinforced walls into local mortar joint separation and the development of flexure in the walls. Bahei-El-Din et al.12,13 studied the strain reduction effect of a polyurea interlayer within a sandwich plate. The inserted polyurea layer could alter the propagation of stress waves into the foam core, leading to a significant reduction in the face sheet strain. Gardner et al.14 investigated the dynamic behaviours of E-glass vinyl ester (EVE) composite panels. Applying a polyurea layer between the foam core and the back facesheet improved the overall blast performance of the EVE composite panel and maintained the structural integrity. Hou et al.15 evaluated the effect of the structural configuration on the air blast resistance of polyurea-coated composite steel (PCCS) plates. Rear-side sprayed and sandwiched PCCS plates had superior air blast resistances to their monolithic steel counterparts. Chen et al.16 conducted air blast experiments to identify the deformation and failure of polyurea-coated steel (PCS) plates with different steel strengths. The increase in the polyurea strength led to a dramatic decrease in the energy absorption proportion of the kinetic energy of polyurea fragments. Guo et al.17 conducted contact explosion experiments on uncoated and rear-side-coated reinforced-concrete slabs to analyse the dynamic response and explosion resistance mechanism of the polyurea. It was found that the polyurea coating dissipated explosive energy through elastic deformation, plastic deformation, fracture, and debonding. Spraying polyurea on the protected substrate could effectively reduce the deflection of the substrate and mitigate substrate damage. Once the brittle substrate was broken under a high explosive load, the polyurea coating could also act as a capture system6,18, reducing the number of fragments and effectively mitigating the threat of flying debris.
Steel shelters consist of a steel frame and blast-resistant walls, and the walls are usually composed of explosion-resistant plates. The application of polyurea-coated blast-resistant plates has good potential for further improving the explosion resistance of shelters. The above literature focused on the study of polyurea-coated plates/walls with high explosives as the explosion source and analysed the blast resistance performances of the polyurea-coated components. There have been few studies on the gas explosion resistance of polyurea-coated plates. Furthermore, due to the limitation of gas explosive devices and testing sites, the previous gas explosion tests mainly evaluated single structural components19,20,21, and full-scale analysis of the explosion resistance of a whole shelter has been rarely conducted. Liquid fuel can mix with air to form a flammable cloud, which is commonly referred to as a fuel–air explosive (FAE)22. FAE cloud explosion can simultaneously generate multiple damage effects, such as shock waves, seismic waves, and fireballs. Compared with trinitrotoluene (TNT) explosions, FAE cloud explosions have lower peak overpressures and longer durations23,24, and their damage effect is more similar to that of an industrial gas cloud explosion. FAE cloud explosions can be used to test shelters in the petrochemical industry.
In this study, gas explosion tests of uncoated and polyurea-coated blast-resistant plates were carried out, and a polyurea-coated shelter was designed based on the explosion tests and numerical simulations. With a liquefied hydrocarbon station of a certain refining enterprise as an example, the explosion load of the shelter was calculated using the parameters of the incident shock wave, and the dynamic behaviour of the shelter under different explosion loads was analysed. A special FAE was proposed, and a series of field tests were carried out on the full-scale shelter. The explosion overpressure and the status of the markers in the shelter were recorded and analysed, and the explosion resistance performance of the full-scale shelter was verified. Therefore, this study could provide guidance for the design and application of explosion-resistant shelters in the petrochemical industry and other process industries.
The specimens used in the test included rectangular steel frames and blast-resistant plates. The blast-resistant plate was a composite sandwich structure with a fibre-reinforced concrete core covered by galvanized steel sheets on both sides. The thickness of the fibre-reinforced cement core was 8.5 mm, and the thickness of the single-layer galvanized steel sheet was 0.5 mm. The plane size of the blast-resistant plate was 2400 × 1200 mm, with a total thickness of 9.5 mm. The rectangular steel frame was made from a hollow steel tube with a cross section of 100 × 100 mm and a 4 mm thickness.
In the blast resistant engineering projects of petrochemical enterprises, the blast resistant plate was usually fixed on the steel frame with self-tapping screws. Therefore, in this study, in order to be consistent with engineering conditions, a series of screws were used to fix the blast resistant plate (Fig. 1). The size of self-tapping screws was M5.5 × 35 mm, and the average screw spacing was about 240 mm. Two specimens with the same steel frame were prepared. The blast-resistant plate of one specimen was sprayed with 3 mm of polyurea, and the blast-resistant plate of the other specimen was uncoated.
The schematic diagram and a photograph of the gas explosive device is shown in Figs. 2 and 3, respectively. The gas explosion test device was composed of a test plate structure, gas distribution system, gas storage pipeline, and measuring system. The test plate structure was composed of a blast-resistant plate, rectangular steel frame, and fixed frame, as illustrated in Fig. 1. The rectangular steel frame was anchored to the fixed frame through steel rods. In this study, the gas storage pipeline was DN500 in diameter and 36 m in length. The combustible gas used was propane with a concentration of 4.2%. The data measurement system included a synchronous trigger, a data acquisition instrument, pressure sensors, and a high-speed camera (FastCAM Mini Ax, Phototron). The arrangement of pressure sensors is shown in Fig. 3. The pressure sensor (HTP 400, The 401 Research Institute of China Aerospace Science and Technology Corporation)) was used to record the pressure time history of the explosion shock wave, and the measuring range of the sensor was 0–5 MPa. After the test plate was installed, the square flaring (2 × 2 m) was sealed with plastic film. The propane was pumped from the gas cylinder into the storage pipeline by the gas distribution system, and then the spark plug was remotely controlled for ignition.
Schematic diagram of the test device.
The structural response process of the uncoated blast-resistant plate is shown in Fig. 4. The uncoated blast-resistant plate underwent large deformation under an explosion load, but the appearance of the blast-resistant plate remained intact without high-speed debris.
Structural response process of the uncoated blast-resistant plate.
The overpressure time-history curves are shown in Fig. 5. P1 was the upper sensor of the blast-resistant plate, and P2 was located at the bottom. The peak overpressure of P1 was 224 kPa and the positive duration was 13 ms. The peak overpressure of P2 was 801 kPa, and the positive duration was 11.4 ms. The front face of the uncoated blast-resistant plate after the explosion impact is shown in Fig. 6. The plate protruded outward to form two inverted trapezoidal platform shapes, and the maximum displacement reached 12.2 cm. The intermediate square steel component played a key supporting role to prevent further deformation and damage of the blast-resistant plate. However, the four sides of the plate were close to the centre due to the explosion impact, causing some of the fixed bolts to undergo shear damage. However, due to the large convex deformation of the plate under an explosion impact, some fixed bolts were damaged.
Front face of blast-resistant plate.
Under the same test conditions, an explosion test was carried out on the blast-resistant plate sprayed with 3 mm polyurea on the rear face. The structural response process of the polyurea-coated blast-resistant plate is shown in Fig. 7. During the test, there were no visible cracks in the coating, the blast-resistant plate was relatively flat, and the overall structure did not undergo large deformation.
Structural response process of the blast-resistant plate coated with 3 mm polyurea.
It can be seen from Fig. 8 that the peak overpressure of the blast wave measured by P1 was 211 kPa, and the positive duration was 12 ms. Due to the damage of sensor P2, the data of the lower part of the blast-resistant plate was not collected. However, the load of sensor P1 (211 kPa, 12 ms) was basically consistent with that of the uncoated plate (224 kPa, 13 ms). Therefore, the repeatability of the explosion load in the two tests was acceptable. The front face of the coating-reinforced blast-resistant plate is shown in Fig. 9. After the explosion impact load, the blast-resistant plate also protruded outward. However, compared with the uncoated blast-resistant plate, the deformation of the coated plate was significantly reduced, and the maximum deformation was 3.4 cm. With the reduction of the plate deformation, the fixed bolts around the coated plate were basically not damaged, reducing the risk of the plate falling off.
Front face of blast-resistant plate.
According to theoretical calculations in previous research, the wave impedance of steel was 4.02 × 107 (N∙s/m3), and the wave impedance of polyurea was 1.10 × 106 (N∙s/m3)25. There was a wave impedance mismatch between polyurea and steel. Since polyurea was coated on the rear face, and the stress wave propagated from the high-impedance steel to the low-impedance polyurea. The reflected stress wave formed an unloading attenuation effect on the incident stress wave, alleviating the damage degree of steel sheet.
Under the combined effect of polyurea deformation energy absorption and stress wave attenuation, the deformation of polyurea coated plate was significantly reduced compared to uncoated plate, and the explosion resistance of coated plate was improved. Polyurea had the microscopic characteristic of microphase separation, and its microscopic energy absorption mechanism also included shock wave induced rearrangement of soft and hard segments26,27, hydrogen bond breaking and recombination28, and viscous dissipation29.
Referring to previous studies18,30,31,32,33, the elastic–plastic material model *MAT_PIECEWISE_LINEAR_ PLATICITY was used to simulate the explosive mechanical behaviour of the polyurea. The stress–strain curves of polyurea at different strain rates (Fig. 10) were obtained using an MTS material testing machine and split Hopkinson tensile bar34. The stress–strain curves of polyurea exhibited significant nonlinearity and strain rate sensitivity. To characterize the strain rate effect of polyurea, the dynamic increase factor (DIF) of the stress was calculated and is plotted in Fig. 11. The DIF is the ratio of the stress at a specific strain rate to the stress at the reference strain rate (0.001 s−1) under the same strain (0.02).
Stress–strain curves of polyurea.
Dynamic increase factor of polyurea.
The Cowper–Symonds model can be used to describe the strain rate effect of polyurea under extreme loads35:
where \(\dot{\varepsilon }\) is the strain rate. Calculated by the least squares method, the constants C and p were 650.18 s−1 and 0.98, respectively.
The stress–strain curve of polyurea at a strain rate of 0.001 s−1 was input into *MAT_PIECEWISE_ LINEAR_ PLATICITY model, and the Cowper-Symonds model was used to define the strain rate effect of polyurea. The main parameters of polyurea, including the parameters of the Cowper-Symonds model, are presented in Table 1. Fiber-reinforced concrete was described by the *MAT_BRITTLE_DAMAGE model, and the parameters of the fibre-reinforced cement are shown in Table 236. The galvanized steel plate and frame square steel were characterized by the *MAT_PLASTIC_KINEMATIC model. The material parameters of the galvanized steel plate and the frame square steel are shown in Table 337,38.
Numerical simulations of the gas experiment were performed using the finite element software ANSYS/LS-DYNA. The finite element model is shown in Fig. 12, and the load curve method is used to directly apply explosive loads on the blast-resistant plate. During the experiment, there was a significant difference in the explosion overpressures between the upper and lower parts of the blast-resistant plate. In the simulation model, the explosion load was applied to the blast-resistant plate in the upper and lower zones. The pressure curve measured by sensor P1 was adopted for the upper load of the blast-resistant plate, and the pressure curve measured by sensor P2 was adopted for the lower load of the blast-resistant plate. Due to the failure of sensor P2 in the coated plate test, no valid data was collected. However, as mentioned above, the repeatability of the gas explosion test was acceptable. Therefore, the data of sensor P2 in the simulation was the same as that of the uncoated plate test.
Finite element model of blast-resistant plate and steel frame.
The galvanized steel plate and the square steel frame were meshed by shell elements, while the other structures were meshed using solid elements. The overall grid size of the model was 2.5 cm, and the total number of grid elements was 30,301. The square steel frame was anchored to the fixed frame through steel rods. Therefore, the square steel frame adopted fixed boundary conditions. The fixed bolts of the polyurea-coated blast-resistant plate were not damaged, and thus, bond contact was adopted between the polyurea-coated plate and the square steel frame.
The structural response process of the polyurea-coated blast-resistant plate is shown in Fig. 13. The blast-resistant plate exhibited significant deformation under the explosive load, followed by a certain degree of deformation rebound. The experimental and numerical simulation comparison of the plastic deformation of the polyurea-coated blast-resistant plate is shown in Fig. 14. Due to the supporting effect of the central square steel frame, the blast resistance plate formed two concave deformations between the central square steel component and the boundary square steel.
Simulation of structural response process of polyurea-coated plate.
Comparison of plastic deformation of polyurea-coated blast-resistant plate.
The maximum elastic–plastic deformation of the blast-resistant plate was 5.8 cm at 6.1 ms (Fig. 15). After a period of vibration, the residual deformation of the blast-resistant plate was 3.6 cm. The plastic deformation of the resistance plate measured in the experiment was 3.4 cm, and the relative error between the simulation and experimental values was 5.88%. The experimental results were in good agreement with the numerical simulation results.
The control room of a petrochemical enterprise was adjacent to the liquefied hydrocarbon loading and unloading platform. There was a risk of accidental leakage of liquefied hydrocarbons causing the control room to be exposed to the explosive wave. According to the architectural drawings, the building was a brick–concrete structure with a low resistance to progressive collapse. To reduce the risk of casualties in explosion accidents, a new blast-resistant steel structure shelter was proposed.
Based on our previous simulation of liquefied hydrocarbon leakage and explosion39, it was determined that the peak incident overpressure (\(P_{so}\) ) of the shock wave at the new shelter was 14.56 kPa, and the positive duration (\(t_{d}\) ) was 174.56 ms. The design appearance size of the shelter was 6.2 × 2.8 × 2.7 m. The blasting load of each face of the shelter (Fig. 16) could be calculated by substituting the shock wave and building parameters into the following Formulas (2), (3), (4), (5), (6), (7), (8), (9), (10) 4.
Blast loading curves of shelter.
The peak reflected overpressure (\(P_{r}\) ) and the stagnation pressure (\(P_{s}\) ) were calculated as follows:
where Cd is the drag coefficient, and \(q_{o}\) is the peak dynamic wind pressure. For closed rectangular buildings, the Cd of the front wall was set as 1.0, and the values for the side wall, roof, and back wall were − 0.4. The effective duration (te) was calculated as follows:
where S is the clearing distance (m), U is the shock front velocity (m/s), and \({\text{t}}_{{\text{c}}}\) is the reflected overpressure clearing time (s).
The equivalent peak overpressure (\(P_{a}\) ) and rise time (\(t_{r}\) ) were calculated as follows:
where \(C_{e}\) is the equivalent peak overpressure, and L1 is the length of the structural member in the direction of the shock wave (m).
The equivalent peak overpressure (\(P_{{\text{b}}}\) ) and time of arrival (\(t_{a}\) ) were calculated as follows:
where \({\text{L}}\) is the building width (m).
The rise time (\(t_{r}\) ) is calculated as follows:
The shelter consisted of a steel frame and polyurea-coated blast-resistant plates. The main beam and column of the frame were 100 × 100 × 4 mm square steel, and the secondary beam and column were 100 × 50 × 4 mm square steel. The steel grade of the frame beam column was Q345. To facilitate the installation of 1.2 m-long blast-resistant plates, the maximum spacing between the beams and columns was 0.6 m (Fig. 17a). Consistent with the gas explosion testing, 3 mm polyurea was sprayed on the rear face of the blast-resistant plate. The full-scale shelter model is shown in Fig. 17. Galvanized steel plates and steel frames were modelled by shell elements, while other components were modelled using solid elements. For the structural simulation of a mobile refuge chamber under an explosion load37, the mesh size of the shelter model was set to 2.5 cm, and the total number of grid elements was 515,120. The shelter was fixed to the precast concrete foundation through anchoring, and thus, the bottom of the shelter was set as a fixed boundary.
Full-scale shelter model. (a) Steel frame. (b) Steel frame and panel.
The deformation process of the shelter under an explosion load is shown in Fig. 18. The overall deformation of the shelter was the central depression of the wall, and the shelter wall remained intact without cracking. The displacement time-history curves of the centre of each wall are shown in Fig. 19, and the maximum deformation values of the walls are summarized in Table 4. The shelter steel frame showed inward bending deformation, and there were no fractures in the square steel beams and columns (Fig. 20). The displacement curves of the maximum deformation points on each side of the frame are shown in Fig. 21, and the maximum deformation values are listed in Table 4.
Displacement time-history curves of the centre of each wall.
Deformation process of steel frame.
Displacement curves of the maximum deformation points on each side of the frame.
During the deformation process of the shelter, the blast-resistant plate did not detach from the steel frame. Therefore, as presented in Table 4, the maximum displacement of the wall was basically the same as the maximum displacement of the frame (except for the right side). The number of horizontal square steel components on the right side was less than that on the left side, and the maximum displacement of the right-side square steel components was basically the same as that of the left-side square steel components. However, the displacement of the right-side wall increased significantly (Table 4), which indicated that the horizontal square steel component mainly affected the deformation of the wall.
Process of framework stress variation.
Stress curves of the front column with the largest deformation.
There are two typical blast shock waves in the explosions of the blast-resistant structures in the petrochemical industry. One was the peak overpressure of 21 kPa with a duration of 100 ms, and the other was the peak overpressure of 69 kPa with a duration of 20 ms4. Based on the peak overpressure and its duration, the wall load of the shelter was calculated by Formulas (2), (3), (4), (5), (6), (7), (8), (9), (10).
Under the action of the explosion wave (69 kPa, 20 ms), the cloud map at the moment when the shelter displacement and stress reached the maximum values is shown in Fig. 24. The blast face of the shelter was subjected to the greatest reflected overpressure, and the maximum displacement and stress occurred on the front side. The maximum displacement time-history curve of the shelter is shown in Fig. 25, and the maximum displacement reached 24.1 cm at 19 ms. The maximum stress time-history curve of the shelter is shown in Fig. 26, and the maximum stress reached 783 MPa at 12 ms.
Shelter structure response (69 kPa, 20 ms). (a) Maximum displacement (19 ms). (b) Maximum stress (12 ms).
Maximum displacement curves of shelters under different loads.
Maximum stress curves of shelters under different loads.
Under the action of the explosion wave (21 kPa, 100 ms), the cloud map at the moment when the shelter displacement and stress reached the maximum value is shown in Fig. 27. Different from the high explosion load (69 kPa, 20 ms), the maximum displacement of the shelter appeared on the right wall, but the wall was not damaged. The blast-resistant wall load would be transferred to the steel frame, which was the main bearing member of the shelter and directly determined the safety of the shelter. In order to compare the responses of shelter structures under different loads, the maximum displacement and stress of the shelter were uniformly characterized by the maximum displacement and stress of the steel frame. The maximum displacement time-history curve of the shelter is shown in Fig. 25, and the maximum displacement reached 3.72 cm at 15 ms. The maximum stress time-history curve of the shelter is shown in Fig. 26, and the stress reached its maximum value of 634 MPa at 12 ms.
Shelter structural response (21 kPa, 100 ms). (a) Maximum displacement (15 ms). (b) Maximum stress (12 ms).
The simulation results showed that the shelter was not damaged under the action of two typical explosion waves. The deformation and stress of the shelter under a typical high explosive overpressure (69 kPa, 20 ms) were higher than those under a typical low explosive overpressure (21 kPa, 100 ms). Therefore, the dynamic response analysis of the shelter under a high explosion load (69 kPa) and different durations (5, 10, and 20 ms) was carried out. To compare with the results of subsequent FAE explosion tests, the simulated load also included the maximum load of the FAE explosions. The maximum displacement and maximum stress time-history curves of the shelter under different loads are shown in Figs. 25 and 26, respectively, and the maximum displacement and stress are summarized in Table 5.
According to Table 5, the structural response of the shelter was closely related to the incident overpressure and duration. The maximum displacement and stress of the shelter under a high incident overpressure significantly decreased with decreasing duration. When the duration decreased from 30 to 5 ms, the maximum displacement decreased from 28.8 to 4.81 cm, and the maximum stress was reduced from 814 to 679 MPa. However, the maximum displacement and stress of the shelter under a high explosion overpressure for a low duration (69 kPa, 5 ms) were still higher than those under a low explosion overpressure for a high duration (21 kPa, 100 ms).
Based on the steel frame and polyurea-coated blast-resistant plate described above, a full-scale shelter engineering prototype was processed. Considering the aesthetic appearance and internal insulation, an engineering shelter wall was designed with a multi-layer composite structure. The outer layer of the shelter wall was a corrugated steel plate, the middle layer was a polyurea-coated blast-resistant plate and thermal insulation material, and the inner layer was a gypsum board. A layer of fireproof coating was added on the surface of polyurea to improve its fire resistance. The full-size shelter for the FAE test is shown in Fig. 28.
To obtain damage effects similar to those of a gas explosion, a cylindrical vessel filled with liquid oil and a solid energetic material was designed as the FAE source. As shown in Fig. 29, the central tube inside the cylindrical vessel was filled with a TNT explosive, and direct-current transformer oil was filled in the space between the central tube and the vessel wall. In each explosion test, the fuel oil used was 10 L and the TNT explosive mass was 575 g. To fully disperse the fuel oil after detonation, as shown in Fig. 30, the FAE source vessel was placed on a steel frame with a height of 1.6 m. The detonator was remotely controlled to detonate the TNT explosive, which caused the fuel oil to be scattered, and the fuel oil deflagrated during the dispersal process.
Schematic of fuel–air explosion test.
To systematically evaluate the protective effect of a full-size shelter on the internal personnel and critical equipment, as shown in Fig. 30, pressure sensors were placed inside and outside the shelter, and a table and a chair were placed inside the shelter. The distances between the external pressure sensors and the shelter were 1 m (P1), 2 m (P2), and 3 m (P3). Two pressure sensors (P13 and P14) were arranged on the ground along the central axis inside the shelter, and the distances from the explosion face of the shelter were 1 and 2 m, respectively. A total of nine pressure sensors (P4–P12) were installed on the blasting face of the shelter. The type of pressure sensor was 113B (PCB Piezotronics Inc), with a range of 0–3.4 MPa and a sensitivity of 1.45 mV/kPa. A high-speed camera (Photon mini ux100) was used to record the dynamic response of the shelter. The recording frequency of the high-speed camera was 2000 fps.
The FAE test of the full-size shelter was conducted in the field (Fig. 31). In this study, three groups of repeated explosion tests were carried out for the same shelter. The distances between the explosion source and the shelter were 5, 3, and 2 m successively.
Fuel–air explosion test setup.
Figure 32 shows the typical process of the FAE test (test 1). After the detonator ignited, the shell of the explosive device ruptured (378 ms), and then the liquid fuel was rapidly scattered around to form a flat spherical cloud. The fuel cloud expanded to contact the shelter at 420 ms. At 450 ms, a flame spreading phenomenon occurred, and the cloud ignited and exploded from inside to outside, forming a fireball. The fireball radius reached the maximum at about 650 ms.
Typical process of fuel–air explosion (test 1).
Due to the overall high-speed vibrations of the shelter and the damage of the sensor data line caused by the explosion wave, only some sensors measured effective explosion overpressure data (Fig. 33). The peak overpressures measured by the ground sensors P1, P2 and P3 were 118.56, 83.62, and 48.94 kPa, respectively. The positive pressure durations of the ground sensors P1, P2, and P3 were 2.41, 2.59, and 5.13 ms, respectively. The overpressure of the FAE gradually decreased with increasing distance from the explosion source, while the positive duration gradually increased with increasing distance from the explosion source.
Overpressure time-history curves of test 1.
Sensor P6 at the bottom of the shelter measured a peak overpressure of 35.17 kPa and a positive duration of 5.43 ms. The overpressures of sensors P13 and P14 inside the shelter were 0.81 and 0.44 kPa respectively, with an average value of 0.63 kPa. The positive pressure durations of sensors P13 and P14 were 4.18 and 5.88 ms, respectively. Due to the shielding effect of the shelter on the explosion shock wave, the overpressure inside the shelter was significantly lower than the overpressure outside the shelter. The sensors inside the shelter were arranged on the ground, and thus, the measured value (35.17 kPa) of the sensor at the bottom of the shelter was taken as the external shock wave value. The attenuation rate of the shelter to the blast overpressure was calculated to be about 98.2%.
Figure 34 shows the shelter after test 1. The internal table and chair were displaced but had not toppled (Fig. 34a), and the shelter internal wall was basically undamaged (Fig. 34b). As shown in Fig. 34c, residual oil and burning marks remained on the outer wall of the shelter, but there were no cracks or evident deformation on the outer wall. The shelter effectively blocked the shock wave overpressure and high-temperature fireball in the explosive accident, and it could have significantly alleviated the damage to personnel and equipment in the shelter. The shelter explosion-resistant window cracked (Fig. 34d) due to the impact of the explosion splash, but no debris splashed. In order to analyse the ability of the main structure of the shelter to withstand repeated explosion impacts, the explosion-resistant window was replaced with a steel plate in the subsequent explosion tests.
Shelter after the first explosion. (a) Internal table and chair. (b) Interior wall. (c) Outside surface. (d) Explosion-resistant window.
In test 1, there was almost no visible damage to the shelter wall. Thus, the blast distance was reduced to 3 m in test 2. Because it was too close to the explosion source, data transmission failures occurred, and most sensors were unable to collect effective explosion pressure data. The effective explosion overpressure curves are shown in Fig. 35. The shock wave overpressure measured by the lower sensor P5 of the shelter was 159.6 kPa, and the positive pressure duration was 2.06 ms.
Overpressure time-history curves of test 2.
Due to the reduction of the distance between the explosion source and the shelter, the peak overpressure on the lower part of the shelter increased significantly compared with that of 35.17 kPa in test 1. The overpressure of sensor P13 inside the shelter was 14.31 kPa. With the peak overpressure (159.6 kPa) of the sensor at the lower part of the shelter as the reference value, the overpressure attenuation rate of the shelter was calculated to be 91.0%.
Figure 36 shows the shelter after test 2. The table and chair inside the shelter were not overturned (Fig. 36a). A slight vertical crack appeared on the inner wall of the shelter (Fig. 36b). There were burning marks on the exterior wall, but no visible deformation occurred (Fig. 36c).
Shelter after the second explosion. (a) Internal table and chair. (b) Interior wall. (c) Outside surface.
The explosion distance in test 3 was further reduced to 2 m. The overpressure time-history curve of test 3 is shown in Fig. 37. The pressure sensors were too close to the explosion source, resulting in significant multi-peak characteristics of the overpressure curves. To conservatively evaluate the performance of the shelter to resist explosions, the first wave peak with a lower value was taken as the peak pressure. The peak overpressure measured by ground sensor P2 and sensor P6 at the lower part of the shelter were 299.6 and 38.62 kPa, respectively.
Overpressure time-history curves of test 3.
Since the TNT explosive was located directly above sensor P2 in test 3, the overpressure value measured by sensor P2 was higher than that in test 1 (83.62 kPa). The explosion distance in test 3 was further reduced, but the smaller distance between the explosion source and the shelter also led to the insufficient dispersal of fuel oil. Thus, the overpressure value of the shelter wall was between the values of test 1 and test 2. The inside and outside structures of the shelter after three repeated explosions are shown in Fig. 38. There was no visible damage to the table or chair inside the shelter (Fig. 38a). Lateral and vertical cracks appeared on the inner wall of the shelter, and plasterboard peeled away slightly at the intersection of the cracks (Fig. 38b). There was no significant deformation on the exterior wall (Fig. 38c).
Shelter after the third explosion. (a) Internal table and chair. (b) Interior wall. (c) Outside surface.
The results of the three FAE explosion tests are summarized in Table 6. For the lower peak overpressure (35.17 kPa), the shock wave attenuation rate of the shelter reached 98.2%, and for the higher peak overpressure (159.61 kPa), the shock wave attenuation rate was still greater than 90%. After three repeated temperature–pressure coupled loads, transverse and vertical cracks appeared on the inner wall of the shelter, but no overall collapse or splashing occurred. Furthermore, there was no visible damage to the table or chair.
The degree of human injury and casualty classification under explosion overpressures are shown in Table 72 and Table 840 respectively. In test 1, the peak overpressure on the outer wall of the shelter was 35.17 kPa, which would cause eardrum harm or moderate injury. The overpressure inside the shelter was 0.63 kPa, indicating that no minor injuries to the occupants would occur and that the internal personnel were in safe conditions. In test 2, the peak overpressure on the outer wall of the shelter reached 159.61 kPa, resulting in fatal physical harm and even death. The maximum blast overpressure for indoor personnel protected by the shelter was 14.31 kPa, which would not cause eardrum injury, and the casualty level would be reduced to the minor injury level. The shelter could effectively reduce the threat of repeated explosive loads and improve the safety level of indoor personnel.
The measured loads of the sensors in the FAE explosion tests were the reflected overpressures. The incident overpressures of the FAE explosion in Table 5 were calculated from the sensor loads and Formulas (2), (3), (4), 5. According to Table 5, the maximum displacement and stress (2.39 cm and 572 MPa) of the shelter under the explosion load for petrochemical enterprise risk assessment were both lower than those (2.61 cm, 593 MPa) under the FAE load (test 2). From the perspective of the displacement and stress of the shelter, the damage effect of test 2 load was higher than that of the risk assessment load. In the repeated FAE tests, the shelter did not show significant damage and remained intact as a whole. Thus, the shelter could be used for personnel protection in this risk area.
In this paper, the deformation behaviours of uncoated and polyurea-coated blast-resistant plates were studied through gas explosion tests. Based on the gas explosion test results, a finite element model of a polyurea-coated shelter was established, and the structural responses of the shelter under various explosion loads were investigated. A series of FAE tests were conducted on a full-scale shelter to analyse its explosion resistance performance. The main conclusions were as follows.
(1) Premature failure of the fixed bolts led to the entire uncoated blast-resistant plate falling off, and the energy absorption ability of the plate could not be exploited effectively. Compared with the uncoated blast-resistant plate, the deformation of the polyurea-coated plate was reduced by 72%. With the decrease in the plate deformation, the fixed bolts around the plate were basically not damaged, and the overall blast resistance of the plate structure was further improved after spraying with polyurea.
(2) The overall deformation of the shelter was the central depression of the wall and the inward bending of the frame. The blast face of the shelter was subjected to the greatest reflected overpressure, and the maximum displacement and stress of the frame occurred on the front side. The structural response of the shelter was closely related to the peak incident pressure and the positive pressure duration. The damage effect of a typical high-overpressure, low-duration load was greater than that of a typical low-overpressure, long-duration load.
(3) The shelter remained intact under three repeated explosive loads, with cracks appearing on the inner wall but no collapse or debris splashing. The shock wave attenuation rate of the shelter reached over 90%, which would significantly reduce the degree of indoor casualties. The polyurea-coated shelter could be used to protect personnel in high-risk areas of petrochemical processing plants and to improve personnel safety.
The datasets used and/or analysed during the current study available from the corresponding author on reasonable request.
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The authors acknowledge the financial support from the National Key Research and Development Program of China (No. 2021YFC3001204).
State Key Laboratory of Chemical Safety, Qingdao, 266104, China
Meng Gu, Haozhe Wang, Guoxin Chen, Anfeng Yu, Wenyi Dang & Xiaodong Ling
SINOPEC Research Institute of Safety Engineering Co., Ltd., Qingdao, 266104, China
Meng Gu, Haozhe Wang, Guoxin Chen, Anfeng Yu, Wenyi Dang & Xiaodong Ling
College of Mechanical and Electronic Engineering, China University of Petroleum (East China), Qingdao, 266580, China
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Conceptualization, Meng Gu, Haozhe Wang, Anfeng Yu and Wenyi Dang; experiment, Meng Gu, Haozhe Wang, Guoxin Chen and Xiaodong Ling; formal analysis, Guoxin Chen and Xiaodong Ling; data curation, Meng Gu, Anfeng Yu and Wenyi Dang; writing —original draft preparation, Meng Gu; writing—review and editing, Meng Gu and Haozhe Wang; visualization, Meng Gu and Guoxin Chen; supervision, Anfeng Yu and Wenyi Dang.
The authors declare no competing interests.
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Gu, M., Wang, H., Chen, G. et al. Experimental and numerical study on explosion resistance of polyurea-coated shelter in petrochemical industry. Sci Rep 14, 20643 (2024). https://doi.org/10.1038/s41598-024-71339-w
DOI: https://doi.org/10.1038/s41598-024-71339-w
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