Stainless Steel Heat Exchanger Tube and Tube Sheet Welding Cracks: Comprehensive Analysis of Causes and Improvement Measures
1. Current Status of Heat Exchanger Welding
Heat exchanger tubes, as the core medium for heat transfer in condensers, play a critical role in industrial equipment. However, they often operate under complex conditions such as corrosion, oxidation, high pressure, and equipment vibration, which increase the risk of local or overall failure—ultimately affecting the normal operation of heat exchangers.
To adapt to different working conditions, heat exchanger tubes are made of various materials. Among them, austenitic stainless steel is widely used in petrochemical and other fields due to its excellent corrosion resistance, wear resistance, mechanical properties, machinability, and weldability—especially in oxidizing, neutral, and weakly reducing media. Despite these advantages, austenitic stainless steel is extremely sensitive to contamination by low-melting-point substances such as copper (Cu). In practical production, numerous cases of welding cracks and hot working cracks caused by Cu contamination have been reported.
When stress is applied, cracking caused by Cu contamination becomes more sensitive and severe. This article focuses on the tube end cracks that occurred in the tube sheet welded joints of a flash condenser. Through testing, experimental verification, and theoretical analysis, the main causes of cracks are identified, and corresponding solutions are proposed to provide a reference for similar applications.
2. Tube Sheet Welding: Joint Design and Inspection Results
2.1 Joint Design and Welding Process
The heat exchanger tubes and tube sheets of the flash condenser adopted a “weld-first, expand-later” process. The heat exchanger tube material conformed to the S31603 stainless steel specified in GB/T 13296-2023 (Stainless Steel Seamless Tubes for Boilers and Heat Exchangers), with a specification of Φ25 mm × 2 mm. The tube sheet material was S31603 Grade Ⅲ, with a thickness of 85 mm and an inner hole diameter of Φ25.3 mm.
During equipment manufacturing, all heat exchanger tubes were inserted first. The extension length of one side of the tube sheet was positioned so that all tubes extended 5 mm beyond the outer surface of the tube sheet, and manual tungsten inert gas (TIG) welding was used for spot welding positioning. For the other side, the tube ends extending more than 5 mm beyond the outer surface of the tube sheet were machined. Subsequently, root welding and helium leak testing were performed on the tube ends of both sides. After confirming no leakage, cover welding was conducted, followed by penetrant testing and helium leak detection. After passing the tests, tube expansion was carried out, and penetrant testing and helium leak detection were repeated for the tube ends.
The tube sheet welded joint was a fillet weld, and automatic TIG welding was used for the tube sheet. A copper rod was used to position the relative position between the welding torch and the heat exchanger tube. Two welding layers were applied: the first layer was autogenous welding (no filler metal), and the second layer used filler wire. After welding, the weld leg height was approximately 3 mm. The filler wire conformed to AWS A5.9 ER316L, with a diameter of 1.0 mm. The main welding parameters are shown in Table 1.
| Item | Welding Current (A) | Welding Voltage (V) | Welding Speed (mm·min⁻¹) | Wire Feeding Speed (mm·min⁻¹) |
|---|---|---|---|---|
| First Layer | 120 | 12 | 160 | 270 |
| Second Layer | 122 | 12 | 120 | 270 |
Table 1: Main Welding Parameters for Heat Exchanger Tube and Tube Sheet
2.2 Visual Inspection of Welded Joints
Immediately after welding, visual inspection and penetrant testing were conducted on the tube sheet. A large number of cracks were found at the tube ends, extending from the inner wall of the tube port to the outside and inside. More than 85% of these cracks were located on the machined side, while fewer cracks appeared on the original tube port side. The morphology of the tube port cracks showed a clear tendency to propagate along specific paths, indicating a correlation between the machining process and crack formation.
To provide a basis for the structural design and safe service of stainless steel heat exchanger tube and tube sheet welded joints, experimental testing and numerical simulation studies were conducted on the heat exchanger tube cracks. These studies aimed to identify the root causes of cracking and verify potential solutions.
3. Joint Performance Testing and Microstructural Characterization
3.1 Heat Exchanger Tube Performance Testing
The heat exchanger tubes with cracks were removed from the equipment, and samples were taken from the middle of the tubes for physical and chemical performance testing. The results of chemical composition and mechanical property tests (Tables 2 and 3) showed that both met the requirements of GB/T 13296-2023 for S31603 stainless steel.
| Element | Actual Value (wt%) | Standard Requirement (wt%) |
|---|---|---|
| C | 0.017 | ≤0.03 |
| Si | 0.36 | ≤1.00 |
| Mn | 1.05 | ≤2.00 |
| P | 0.029 | ≤0.035 |
| S | 0.001 | ≤0.030 |
| Ni | 10.09 | 10.00~14.00 |
| Cr | 16.76 | 16.00~18.00 |
| Mo | 2.10 | 2.00~3.00 |
| Cu | 0.05 | 0 |
Table 2: Chemical Composition of Cracked S31603 Heat Exchanger Tube (wt%)
| Property | Test 1 | Test 2 | Standard Requirement |
|---|---|---|---|
| Yield Strength (MPa) | 266 | 277 | ≥175 |
| Tensile Strength (MPa) | 577 | 587 | ≥480 |
| Elongation (%) | 65.5 | 68.0 | ≥40 |
Table 3: Mechanical Properties of Cracked S31603 Heat Exchanger Tube
Metallographic testing was conducted on the grain size and non-metallic inclusions at the joint cracks of the sample tubes. The test results showed that the grain size was approximately 44 μm, which was rated as grain size Grade 6 according to GB/T 6394-2017 (Method for Determining the Average Grain Size of Metals). For non-metallic inclusions, testing was performed in accordance with GB/T 10516-2023 (Standard Rating Diagram Microscopic Inspection Method for Determining the Content of Non-Metallic Inclusions in Steel). The results indicated that the non-metallic inclusions were mainly spherical oxides, with Class D coarse series at 0.5 grade, and Classes A, B, C, Ds, and D fine series at 0 grade—indicating minimal inclusion content and no significant impact on material performance.
Transverse samples were taken near the fracture of the stainless steel heat exchanger tube for metallographic structure testing. The results showed that the metallographic structure of the sample was austenite, which conformed to the normal material requirements of S31603 stainless steel. This confirmed that the base material itself had no defects that would cause cracking, and the problem was likely related to the welding process or external factors.
3.2 SEM Morphology Observation and EDS Analysis
The scanning electron microscopy (SEM) morphology of the austenitic stainless steel crack after corrosion (using a mixed solution of water and hydrochloric acid) showed that the fracture exhibited a “rock candy-like” pattern— a typical intergranular fracture, which propagated from the inside of the tube to the outside. A large amount of white substances were found near the cracks and on the grain boundaries. Energy dispersive X-ray spectroscopy (EDS) analysis was performed on these white substances to determine their composition.
The EDS analysis results revealed that the diffraction peak of Cu in the welding cracks was the strongest, indicating that the white substances at the cracks were Cu or Cu compounds. This was a key finding, as the chemical composition test of the base material (Table 2) showed only a trace amount of Cu (0.05 wt%), and this Cu was not distributed at the grain boundaries. Therefore, the Cu at the crack grain boundaries must have been an external contaminant.
Further on-site investigation found obvious scratches and wear marks on the surface of the copper core rod used for positioning the welding torch. The diameter of the copper rod was 1.0~1.5 mm smaller than the inner diameter of the heat exchanger tube. On the machined side of the heat exchanger tube, obvious machining burrs were observed on the inner side of the tube port. When the copper core rod came into contact with these burrs, Cu particles were easily scraped off and adhered to the tube port—leading to Cu contamination.
3.3 Welding Crack Analysis
In general, the main causes of intergranular fracture can be summarized into four aspects: (1) segregation of impurities and alloying elements at grain boundaries, weakening the grain boundaries; (2) precipitation of precipitates at grain boundaries; (3) intergranular corrosion caused by environmental factors; (4) grain boundary cracking induced by thermal stress. In this case, the detection of Cu at the crack grain boundaries, combined with the trace Cu content in the base material, indicated that Cu contamination was the core factor.
Iron-based materials transform into austenite at high temperatures, and the presence of Cu can easily cause liquid metal embrittlement. This is because the melting point of pure Cu is only 1083.4 °C. During welding, when the temperature of the heat-affected zone (HAZ) reaches this temperature, the molten Cu penetrates into the austenite grain boundaries of the HAZ. If appropriate stress conditions exist, embrittlement and subsequent cracking will occur.
To verify whether the tube sheet welding cracks were caused by Cu contamination, a simplified tube sheet was designed for verification tests. Nine sample tubes of the same batch (with no cracks detected by penetrant testing at the tube ends) were used, each with a length of 200 mm. The tube sheet material was TP316L, with a thickness of 26 mm, a hole diameter of 25.3 mm, and a distance of 8 mm between the edges of two holes. The tubes extended 5 mm beyond the tube sheet, and the joint form was consistent with the actual structure.
The samples were divided into three groups with different conditions:
1. Group 1: Machined flat tube ends with unremoved burrs + Cu chip contamination (using a copper block to scrape the inner and outer ports of the tubes);
2. Group 2: Machined flat tube ends with unremoved burrs (no Cu contamination);
3. Group 3: Machined flat tube ends with removed burrs (no Cu contamination).
Manual TIG welding was used for all groups, with a welding current of 110 A, a filler wire diameter of 1.2 mm (conforming to AWS A5.9 ER316L), and two welding layers per joint. After welding, penetrant testing was performed on the three groups of samples. The results showed that the three weld ports in Group 1 exhibited linear indications after penetrant testing, and the crack positions were consistent with the Cu contamination positions—with more Cu residue leading to more obvious cracks. No linear indications were found in Groups 2 and 3. This verification test confirmed that Cu contamination was the direct cause of the welding cracks.
4. Finite Element Simulation of Welded Joints
4.1 Finite Element Model Parameter Settings
Finite element software was used to simulate and analyze the temperature field and stress field during the welding process of austenitic stainless steel. The key steps in building the finite element model included heat source model selection, geometric model establishment, reasonable mesh division, material parameter assignment, and setting of initial and boundary conditions.
4.1.1 Heat Source Selection
To replicate the actual welding process, a heat source model suitable for TIG welding was required. Considering that the actual arc heat source has characteristics such as a large temperature gradient and highly concentrated energy, a double ellipsoid heat source was selected for finite element simulation. This model fully accounts for the above characteristics of the welding process and also considers the penetration and stirring effects of the welding arc in the thickness direction.
4.1.2 Geometric Model and Mesh Division
A geometric model of the austenitic stainless steel tube sheet joint was established using finite element software, retaining information such as the joint groove form and thickness. After constructing the geometric model, meshing was performed to generate a finite element mesh model. The choice of mesh division method depends on specific conditions, and the mesh accuracy directly affects the calculation efficiency. Generally, fine mesh division can improve calculation accuracy, but excessively fine meshes will significantly reduce calculation efficiency. Therefore, an appropriate mesh size was selected based on actual needs.
Considering the actual welding conditions, the total number of mesh elements was set to 85,920, with an element type of C3D8R (8-node linear brick, reduced integration). Given that TIG welding speed is usually slow, the “birth-death element” technology was used to simulate weld wire filling. The material parameters of austenitic stainless steel (including elastic modulus, Poisson’s ratio, specific heat capacity, thermal conductivity, density, and thermal expansion coefficient) at different temperatures are shown in Table 4.
| Temperature (°C) | Elastic Modulus E (GPa) | Poisson’s Ratio | Specific Heat C (J·kg⁻¹·K⁻¹) | Thermal Conductivity λ (W·m⁻¹·K⁻¹) | Density ρ (g·cm⁻³) | Thermal Expansion Coefficient α×10⁻⁶ (°C⁻¹) | Yield Strength (MPa) |
|---|---|---|---|---|---|---|---|
| 20 | 205 | 0.3 | 450 | 16.3 | 7.98 | 16.0 | 502 |
| 100 | 198 | 0.3 | 480 | 16.0 | 7.93 | 18.0 | 515 |
| 200 | 191 | 0.3 | 500 | 16.1 | 7.89 | 19.5 | 530 |
| 300 | 183 | 0.3 | 520 | 16.2 | 7.83 | 20.2 | 538 |
| 400 | 176 | 0.3 | 540 | 16.8 | 7.78 | 21.8 | 550 |
| 500 | 168 | 0.3 | 560 | 17.5 | 7.73 | 23.2 | 569 |
| 600 | 159 | 0.3 | 580 | 18.1 | 7.68 | 25.0 | 580 |
| 700 | 150 | 0.3 | 600 | 18.6 | 7.62 | 27.2 | 595 |
| 800 | 140 | 0.3 | 620 | 19.0 | 7.57 | 28.2 | 620 |
| 1000 | 125 | 0.3 | 656 | 19.7 | 7.45 | 29.0 | 656 |
| 1200 | 100 | 0.3 | 795 | 20.2 | 7.34 | 30.2 | 795 |
| 1400 | 60 | 0.3 | 886 | 21.1 | 7.22 | 31.1 | 886 |
Table 4: Material Parameters of Austenitic Stainless Steel
4.1.3 Initial and Boundary Conditions
Determining appropriate initial and boundary conditions is another necessary condition for accurately obtaining simulation results. During the model establishment, factors such as initial temperature, heat transfer methods, and boundary constraints were fully considered. To balance calculation accuracy and efficiency, the following assumptions were made:
1. The material is a continuous medium and isotropic;
2. The initial temperature is room temperature (20 °C);
3. There is no internal stress in the joint before welding;
4. The sample tube is fixed by a fixture during welding, so no rigid displacement occurs;
5. Heat transfer methods include thermal radiation and convective heat transfer.
4.2 Finite Element Calculation Results
4.2.1 Temperature Field Simulation Results
The temperature distribution on the welded joint reflects the complex welding thermal process. Uneven temperature distribution will cause uneven stress, which directly leads to the generation of residual stress. The temperature field distribution during the welding of the austenitic stainless steel tube sheet structure showed that when the welding heat source moved to the weld position, the temperature at the weld increased sharply, with the maximum temperature exceeding 1500 °C. After welding, when the moving heat source left the weld area, the central temperature of the weld decreased rapidly. Compared with the welding heating stage, the temperature drop in the cooling stage was relatively slow.
To gain a deeper understanding of the impact of the moving heat source on the weld and its surrounding HAZ during welding, thermal cycle curves were plotted for four positions at distances of 3 mm, 5 mm, 10 mm, and 50 mm from the weld center. The temperature change trend at these points was consistent with that at the weld center: when the heat source moved to the weld position, the temperature rose rapidly; when the heat source left, the temperature decreased and gradually cooled to room temperature.
A key observation was that the peak temperature of the thermal cycle in the affected area within 5 mm of the weld exceeded the melting point of Cu (1083.4 °C), causing Cu to melt. This result was consistent with the earlier EDS analysis and verification test findings, confirming that the welding temperature conditions were sufficient to melt the Cu contaminants and allow them to penetrate the grain boundaries.
4.2.2 Stress Field Simulation Results
Residual stress is also a key factor causing cracking in tube sheet welded structures. Therefore, numerical simulation of the welding stress field distribution of the austenitic stainless steel tube sheet structure was conducted. The Mises stress distribution of the welded joint showed that after welding, the peak stress of the joint appeared at the tube end, reaching 246.75 MPa. The presence of high stress accelerated the formation and propagation of cracks.
To accurately analyze the stress distribution at the tube port, a path (labeled as Path AB) was set in the model. The stress distribution along Path AB showed that the maximum stress along this path reached 246.75 MPa—further confirming that the tube end was a stress concentration area. This high stress, combined with the molten Cu penetrating the grain boundaries, created the ideal conditions for crack initiation and growth.
In summary, the occurrence of Cu contamination cracks in the HAZ of the austenitic stainless steel tube sheet welded structure requires both temperature and stress conditions. The finite element temperature field simulation showed that the peak temperature of the HAZ of the tube sheet structure weld exceeded the melting point of Cu (1083.4 °C), satisfying the melting condition of Cu. The finite element
