Influence of Point-Supported Steel-to-Timber Interface Parameters on the Structural Fire-Resistance of Mass Timber Plates Christian Dagenais Christian Dagenais 1,*, Monireh Aram Monireh Aram 1, Claire Yuan Claire Yuan 2 Andrew Harmsworth Andrew Harmsworth 2 1 FPInnovations, Quebec, QC G1V 0A6, Canada 2 GHL Consultants, a Part of J.S. Held, Vancouver, BC V6C 1G8, Canada * Author to whom correspondence should be addressed. Buildings 2026, 16(12), 2301; https://doi.org/10.3390/buildings16122301 (registering DOI) Submission received: 14 April 2026 / Revised: 27 May 2026 / Accepted: 29 May 2026 / Published: 8 June 2026 Point-supported connections are an innovative modern connection design that can benefit from the 2-way structural action of cross-laminated timber (CLT) slabs, which is typically not considered in traditional timber design. It also allows for flatter ceiling surfaces where no beams are needed to support the mass timber floor slabs. In an attempt to better understand the structural behaviour of this type of connection in fire conditions, preliminary unloaded fire tests were conducted to evaluate their thermal performance. The test results indicated that, for these tested configurations, the presence of steel connection components does not inherently increase charring rates within adjacent mass timber elements. While the outcomes provided valuable insights on the thermal performance of such assemblies, their actual mechanical behaviour under structural loading in fire conditions remains unknown. This paper presents the results of two structural fire-resistance tests under load: Test 1 had the gap fully exposed to fire, and Test 2 had the gap protected by a firestop. Neither assembly reached failure during the 2 h of standard fire exposure, while the target load could not be fully maintained to the end of the tests. Test 1 experienced charring at the CLT-steel plate interface, while Test 2 did not. Their mechanical behaviours were also similar. Lastly, a preliminary design approach is being proposed, although it requires further validation. 1. Introduction Encapsulated mass timber construction (EMTC) up to 12-storeys for Groups C and D major occupancies was first introduced in the 2020 National Building Code of Canada (NBCC) [ 1]. Recently, in December 2023, provinces of British Columbia and Quebec, with support from Ontario, submitted a joint Code change proposal to accelerate the adoption of mass timber construction from the International Building Code (IBC) [ 2], allowing up to 18-storeys of numerous building occupancies, with up to 9-storeys fully exposed. The Code change proposal was based on a transferability report done in 2023 [ 3] in which the fire safety and structural provisions of the International Building Code (IBC) were analyzed for their applicability within the NBCC framework. These provisions have been adopted during the summer of 2024 in British Columbia, January 2025 in Ontario, and June 2025 in Quebec, along with an accompanying technical guide [ 4, 5]. The flexibility offered by these new provisions significantly expands the market opportunities for mass timber construction, namely for low-rise and large building areas for schools, residential and office spaces, as well as a greater possibility for mixed-use at the lower storeys. For such buildings, building elements are required to provide at least 2 h of fire-resistance, regardless of the type of construction and structural materials used. In Canada, this type of structure was first used in the 18-storey Tallwood House student residence on the University of British Columbia campus and has been slowly replicated over the years. However, given that the entire mass timber structure was fully encapsulated at that time, there was no need to assess the fire-resistance of such a point-supported structure, namely the steel-timber bearing area interface. This resulted in conservative and costly design assumptions. Research and modelling performed in the past few years have allowed for an increase in the knowledge about the underlying mechanisms of timber performance in fire. A preliminary assessment of the fire performance of a point-supported steel connection in mass timber construction was performed using a pilot-scale furnace with no structural load [ 9]. The fire tests provided insight into the thermal performance and charring behaviour of steel–timber interfaces at the unloaded CLT panel-to-column connection assemblies. The primary objective of the technical analysis was to evaluate the thermal impact of exposed steel components to timber and different protection strategies on temperature development and timber charring at the connection region. Five connection configurations were evaluated, including exposed steel columns, gypsum board-protected steel columns, intumescent-coated steel columns, exposed steel-glulam column assemblies, and timber-protected steel interfaces. All specimens were subjected to the standard fire prescribed in CAN/ULC S101 [ 10]. Temperature measurements were obtained using thermocouples installed at the steel–timber interface and within the CLT panels to assess heat transfer behaviour during the test. The temperature measurements demonstrate that steel components exposed directly to the furnace environment experienced rapid heating, with exposed steel elements reaching the critical steel temperature of approximately 538 °C within approximately 14 min. In contrast, the temperature rise at steel–timber interfaces embedded within the CLT panel occurred significantly more slowly. For the exposed steel column configuration, the first glue line of the CLT panel reached the critical steel temperature after approximately 46 min of fire exposure. Deeper locations within the panel experienced additional delays in temperature rise, with temperature differences of approximately 30 min observed between the first and second glue lines. These observations indicate that the thermal mass of the CLT panel provides a substantial buffering effect, delaying heat transfer to embedded steel components. The gypsum board protected configuration showed the most effective thermal insulation, with critical steel temperatures not reached at the measured interface locations during the test duration. The intumescent-coated specimen exhibited moderate thermal performance; however, localized cracking and detachment of the coating were observed during the test, resulting in partial exposure of steel components earlier than expected. The test results indicate that, for the tested configurations, the presence of steel connection components does not inherently increase charring rates within adjacent mass timber elements. Instead, the charring rate near the connector was lower than that measured in CLT regions away from the connection. Protection strategies influence the thermal behaviour of the steel components and the adjacent timber. Gypsum board protection provides effective insulation to steel elements, while timber cover protection may offer an alternative strategy to limit heat transfer at steel–timber interfaces. The performance of the intumescent coating observed in this study suggests that further investigation may be required to evaluate the reliability of such protection systems in connection applications. While the preliminary studies provided insight with respect to transient heat transfer within the components, it is necessary to conduct fire tests on these point-supported connection systems subjected to structural loading. This project aims to fill the knowledge gap by combining the thermal and mechanical performance of point-supported connections subjected to structural loading in fire conditions. The results will provide valuable information with respect to the char progression, structural behaviour and compression perpendicular-to-grain behaviour at the steel-timber bearing interface, as fundamental for structural fire-resistance assessment and to demonstrate building code compliance. The design of the connections will be made in such a way as to achieve at least 2 h of fire-resistance in an effective and cost-efficient manner, with the intent to allow as much of that point-supported connection as possible to be exposed without the need for unnecessary additional fire protection. 2. Materials and Methods The testing consisted of two structural fire-resistance tests of point-supported connections, using the mass timber MPSC 200 connector and its associated self-tapping screws from Simpson Strong Tie (Brampton, ON, Canada) ( Figure 2). The MPSC 200 point-supported steel connector consists of a bottom plate welded to an HSS standoff to create the load transfer path without passing through the mass timber floor slab. The standoff can be up to 16” tall to accommodate various floor/ceiling thicknesses and finishes. Figure 3 illustrates the MPSC 200 connector. The tested assemblies consisted of a 5-plies (175 mm) CLT floor slab of the E1 stress grade conforming to ANSI/APA PRG 320 [ 11], with dimensions of 1168 mm wide by 3921 mm long (46″ × 154″). A 152 mm × 152 mm (6″ × 6″) square opening positioned at the geometric centre was created by a computer numerical control (CNC) wood router, with a corner radius of 10 mm. This allows for a 12.7 mm (½″) additional space to accommodate the fillet weld around the HSS standoff of the bottom plate. Two pieces of glulam of 215 × 432 mm (8½″ × 17″) were assembled to form a column of 432 × 432 mm (17″ × 17″) by 635 mm (25″) in height. The glulam was of the 24f-EX Douglas fir species grade conforming to CSA O122:25 and CSA O86:24 [ 12, 13]. Structural adhesive used for CLT manufacturing was used to assemble the 2 elements together, along with self-tapping screws, fully threaded CSK 10 × 320 mm and partially threaded ASSY Ecofast 8 × 300 mm from MTC Solutions (Surrey, BC, Canada). The glulam columns were sized to provide 86 mm of wood cover, as determined by Annex B of CSA O86:24 for a 2 h standard fire exposure. Specific to the testing, the perimeter of the tie plate was sealed with a firestop caulking to prevent smoke and flame leakage during testing. In an actual building, floor finish and/or encapsulation material on top of the mass timber floor would provide such protection. The MPSC bottom plate was fastened to the short glulam column using 8 SDCF 10 × 200 mm self-tapping screws inserted parallel to the grain of the glulam. The MPSC tie plate at the top is fastened to the CLT panel using 16 SDS 25300 self-tapping screws (6.5 mm outside diameter by 76 mm long). According to the supplier’s technical literature, an MPSC 200 installed with a CLT floor slab has an allowable perpendicular-to-grain bearing resistance of 130.5 kN (29,350 lb). This assumes a CLT stress grade E1 conforming to ANSI/APA PRG 320 and a uniform load distribution along the bottom plate dimensions. Other structural considerations may reduce this maximum resistance, such as punching shear through the CLT. The CLT ceiling was protected at each end using 2 layers of 15.9 mm (5/8″) Type X gypsum boards of 1168 mm wide by 1200 mm long, leaving an exposed area of 1168 mm wide by 1521 mm long (46″ × 60″) inside the furnace. This was done to facilitate the control of the furnace temperature. Figure 4 shows the test assembly. Given the test furnace capability, the whole assembly was suitably adapted so that the applied load would be made through a self-reaction frame. The applied load pulled the MPSC and the glulam column upward, rather than pushing downward on the CLT panel. The assemblies were exposed to the standard fire of CAN/ULC S101. Fourteen (14) fibreglass insulated thermocouples per assembly were used to monitor the temperature profiles at critical locations in the vicinity of the connector and across the CLT. TC1 to TC6 were on the MPSC connector ( Figure 5) while TC7 to TC14 were inserted in the CLT through pre-drilled pilot holes ( Figure 6). Pilot holes were sealed with fire stop caulking to limit heat flow. TC7, TC9, TC11 and TC14 were 142 mm away from the inside of the square opening; that is, in the same vertical plane as the outside faces of the glulam column. TC8, TC10, TC12 and TC13 were at 54 mm from the inside of the square opening, that is, in the same vertical plane as the outside edges of the MPSC bottom plate. String potentiometers were used to measure the vertical displacement of the CLT panel during Test 1. After completion of Test 1, it was decided to measure the displacement of the HSS standoff using a 1-string potentiometer and 2 to measure the CLT displacement, as shown in Figure 7. Test 1 consisted of the MPSC 200 assembly with the gap between the top of the glulam column and the ceiling of the CLT panel being fully exposed to fire ( Figure 8). Given the dimensions of the MPSC 200 ( Figure 3), the resulting gap was 25 mm (1″) deep. A load of 70 kN (15,735 lb) was applied to the system based on the CLT rolling shear resistance, as observed as a limiting factor for such a connection system with CLT [ 6, 7, 8], resulting in full loading conditions in accordance with CAN/ULC S101. On the day of testing, the CLT and glulam average moisture content was 8.1%. Test 2 was identical to Test 1, with the only difference being that the gap between the top of the glulam column and the ceiling of the CLT panel was sealed using the Fire Foam from Rothoblass (Quebec City, QC, Canada), which is used to firestop linear joints in CLT construction. Given the dimensions of the MSPC 200 and installation details shown in Figure 4, the firestop was applied through the entire 86 mm depth by the full 25 mm thickness of the steel plate. After being fully dried, the firestop was cut flush with the outside perimeter of the glulam column. The intent was to limit heat transfer to the steel components and assess whether the firestop linear joint is necessary. As with Test 1, a load of 70 kN (15,735 lb) was applied to the system, resulting in a full loading condition in accordance with CAN/ULC S101. On the day of testing, the CLT and glulam average moisture content was 6.2%. Figure 8 shows both assemblies prior to fire testing. Figure 9 shows Test 1 installed in the furnace and ready for testing. 3. Results Both tests were conducted at the National Research Council Canada fire safety unit in Ottawa (Canada) with staff from FPInnovations and J.S. Held witnessing the tests. As required in CAN/ULC S101, the load was applied gradually in four increments of 25% of the total load until stabilization of deflection after each increment. The full (100%) load was applied for at least 30 min prior to starting the fire-resistance test. The average furnace temperature was controlled using nine shielded Type K thermocouples from Omega (St-Eustache, QC, Canada) to allow for the standard time-temperature to follow within the ±5% tolerance required in CAN/ULC S101 for 2 h fire-resistance tests. 3.1. Test 1—MPSC 200 (Unprotected Gap) 3.1.1. Failure Mode—Test 1 No failure was reached after 2 h of standard fire exposure. After 107 min, the allowed vertical stroke of 101 mm (4″) of the load actuator reached its limit, as shown in Figure 10. At that time, the load could no longer be sustained, but the test continued until 2 h. The test was terminated at 2 h 02 (122 min). Figure 11 shows the MPSC 200 Test 1 assembly being dismantled after the fire test. It can be observed that the gap between the CLT panel and the column remained closed. It was also observed that little to no charring occurred around the square opening in the CLT. Local crushing due to steam and moisture migration was observed, where the actual screw heads used to connect the MPSC to the glulam were embossed in the charred CLT panel ( Figure 12). 3.1.2. Applied Load—Test 1 The MPSC 200 Test 1 assembly has been able to support the full applied load until the actuator reached its maximum stroke at 107 min. Given that the displacement could no longer increase, the resulting applied load slowly reduced from 70 to 40 kN between 107 and 120 min ( Figure 13). 3.1.3. Deflection—Test 1 Given the loading configuration, the CLT exhibited significant upward bending as the load was pulling up the MPSC 200 connector. The maximum displacement recorded by the four string potentiometers was 51.7 mm ( Figure 14). At the end of the test, it was observed that the MPSC moved up by approximately 50 mm (2″), as shown in Figure 15, suggesting that charring occurred to the CLT above the steel plate. Combining both deformations allowed reaching the 101 mm maximum stroke faster than expected. 3.1.4. Temperature Profiles—Test 1 Temperature profiles recorded by the thermocouples on the MPSC 200 steel connector are illustrated in Figure 16. It can be observed that the temperature exhibited a smooth reduction after 60 min, which coincides with the time when the CLT upward vertical displacement started to increase ( Figure 14). This phenomenon is likely caused by the MPSC connector being embedded into the charred CLT panel and thereby shielding the thermocouples along with the steel plate perimeter from fire exposure. Throughout the test, none of the thermocouples exceeded the CAN/ULC S101 maximum temperature criteria of 704 °C (single point) for steel beams. The average temperature of 593 °C (average of four thermocouples) was also not exceeded. The 593 °C threshold was exceeded by only one thermocouple (TC-2) after 118 min. It is noted that these temperature criteria are only used as reference targets in the context of this study. The temperature profile measured within the CLT panel is plotted in Figure 17. As expected, TC7 and TC9, positioned at the first glueline and the furthest from the MPSC connector, were the first to reach the charring temperature of 300 °C, at 43.75 and 52.83 min, respectively. The second set of thermocouples to reach 300 °C were TC8 and TC10 after 88.08 and 105 min, respectively. They were also positioned at the first glueline, but closest to the MPSC connector. TC11 and TC14, positioned at the second glueline and furthest from the MPSC connector, exceeded 300 °C after 105.42 and 115 min, respectively. TC12 and TC13, positioned at the second glueline and closest to the MPSC connector, reached a maximum temperature of 133 °C and 126 °C after 2 h of standard fire exposure. Data from TC11 and TC14 results in charring rates up to the second glueline of 0.66 and 0.61 mm/min, respectively, away from the MPSC connector, which is consistent with the one-dimensional charring rate of 0.65 mm/min applicable to CLT in CSA O86:24. The charring rates obtained from TC8 and TC10 are 0.40 and 0.33 mm/min, respectively, which is much slower than the 0.65 mm/min one-dimensional rate. This suggests that the MPSC plate was partially shielded from heat exposure by the CLT, as opposed to the claimed increase in heat conduction that metal connectors may provide to timber elements. From post-test visual observation, it was found that the CLT had a remaining thickness of approximately 120 mm (±3.3 plies left) in the vicinity of the MPSC connector. This residual thickness reduced to approximately 85–90 mm (±2.25 plies left), as shown in Figure 18. The measurements translate to a charring rate of 0.46 mm/min underneath the MPSC steel plate and 0.70 mm/min in the CLT panel—away from the MPSC connector. These are consistent with the charring rates obtained from the thermocouples. It is noted that once the test is terminated and the burners are turned off, the CLT continues to char for an approximate additional 15 min until the whole assembly is extinguished by the NRCC staff. With this additional 15 min of charring, although at a lower rate, the resulting charring rate reduces to 0.63–0.67 mm/min. The glulam residual cross-section was measured as 253 × 258 mm, resulting in a charring rate varying between 0.73 and 0.75 mm/min. These rates are consistent with the design approach where 86 mm of wood cover was provided (86 mm/120 min = 0.72 mm/min). As with the CLT, considering the additional 15 min before extinguishment, the resulting charring rate reduces to 0.64–0.66 mm/min. 3.2. Test 2—MPSC 200 (Protected Gap) 3.2.1. Failure Mode—Test 2 No failure was reached after 2 h of standard fire exposure. At 118 min, the load could not be kept, likely due to excessive bending deflection of the CLT panel caused by the reduced cross-section. The test was terminated at 2h02 (122 min), similarly to Test 1. During Test 2, the stroke of the load actuator was increased to 152 mm (6″), which was not reached at the end of the test. As opposed to what was observed in Test 1, where the screw heads became embedded into the charred CLT panel, the firestop prevented heat from reaching the steel plate. As such, the CLT panel remained almost intact, with no signs of char above the steel plate, as shown in Figure 19. 3.2.2. Applied Load—Test 2 The MPSC 200 Test 2 assembly has been able to support the full applied load until the charred CLT became the weakest component after 117 min. At that time, the load could no longer be maintained and was reduced from 70 to 53 kN between 117 and 120 min, as shown in Figure 20. 3.2.3. Deflection—Test 2 As observed in Test 1, the CLT exhibited significant upward bending as the load was pulling up the MPSC 200 connector ( Figure 21). The maximum displacement recorded by the two string potentiometers was 121 mm. As opposed to Test 1, it was observed that the MPSC moved up by only 4 mm during the test, indicating that little to no charring was experienced on the CLT above the steel plate. Figure 22 shows the displacement of the CLT panel and the MPSC connector during Test 2. 3.2.4. Temperature Profiles—Test 2 Temperature profiles recorded by the thermocouples on the MPSC 200 steel connector are illustrated in Figure 23. Throughout the test, none of the thermocouples exceeded the CAN/ULC S101 maximum temperature criteria of 704 °C (single point) or 593 °C (average) for steel beams. The maximum temperature reached at 2 h was 106 °C (TC2). The firestop effectively prevented heat from penetrating into the assembly. The temperature profile measured within the CLT panel is plotted in Figure 24. As expected, TC7 and TC9, positioned at the first glueline and the furthest from the MPSC connector, were the first to reach the charring temperature of 300 °C, at 76.17 and 65.00 min, respectively. The second set of thermocouples to reach 300 °C was TC 11 and TC14, positioned at the second glueline and furthest from the MPSC connector, after 114.58 and 113.25 min, respectively. Data from TC11 and TC14 results in charring rates up to the second glueline of 0.61 and 0.62 mm/min, respectively, away from the MPSC connector, which is consistent with the one-dimensional charring rate of 0.65 mm/min applicable to CLT in CSA O86:24. None of the thermocouples positioned closest to the MPSC connector exceeded 300 °C throughout the 2 h of standard fire exposure. The maximum temperature recorded at 2 h was 104 °C (TC10). This suggests that the firestop foam effectively prevented heat from penetrating into the MPSC connection. From post-test visual observation, it was found that the CLT had no charred laminations in the vicinity of the MPSC connector (i.e., all 5-plies remaining, 175 mm in thickness), as shown in Figure 25. This residual thickness reduced to approximately 90 mm (±2.25 plies left) away from the MPSC connector. The measurements translate to a charring rate of 0.71 mm/min in the CLT panel—away from the MPSC connector. These are consistent with the charring rates obtained from the thermocouples. As explained previously, considering the additional 15 min before extinguishment reduces the resulting charring rate to 0.63 mm/min. The glulam residual cross-section was measured as 255 × 256 mm, resulting in a charring rate varying between 0.74 and 0.73 mm/min. These rates are consistent with the design approach, where 86 mm of wood cover was provided. When considering the additional 15 min, this translates to a charring rate of 0.65–0.66 mm/min. 4. Discussion As explained in Section 4.1Section 4.2, the only difference between the two tests was the gap at the CLT panel and glulam column, created by the 25 mm thick steel plate of the MPSC 200 connector. Based on the test results, the following behaviour can be observed. 4.1. Load–Displacement When comparing both assemblies, it can be seen in Figure 26 that the CLT displacement curves are similar, whether the gap was fully exposed or protected. The only difference is that Test 1 reached the maximum stroke of the load actuator at 107 min, while Test 2 was able to reach greater deflection (the maximum stroke was increased for Test 2). With all other test parameters identical and given the minimal difference up to 107 min, the results suggest that both assemblies would have performed similarly up to 2 h. 4.2. Thermal Performance Figure 27 plots the average temperature at critical locations, namely at the perimeter of the steel plate (bottom—TC1 and TC2, top—TC4 and TC5) as well as across the CLT along the plane of the glulam column perimeter (first glueline—TC7 and TC9, second glueline—TC11 and TC14). The first main observation is the thermal protection of the MPSC steel plate. In Test 1, the edges of the steel plate increased to 400 °C and remained fairly stable throughout the test duration, while the temperature in Test 2 barely exceeded 100 °C after 2 h of standard fire exposure. The second observation is related to the CLT charring behaviour. Given the thermal protection provided to the steel plate in Test 2, it is not surprising that the CLT experienced less char when compared to that of Test 1. The residual cross-sections are also different due to the charring of the CLT above the steel plate in Test 1, where a residual thickness of 120 mm was left (±3.3 plies left). There was little to no sign of charring in Test 2; all five plies were left. While the thermal performance between the two tests was quite different, the critical steel temperature of 704 °C (single point) or 593 °C (average) has not been reached in either test. The main difference from Test 1 is the differential displacement of the MPSC due to charring of the CLT above the steel plate, which was not observed in Test 2. 4.3. Predicting the Steel Temperature In an attempt to provide a design approach for such an exposed connection, the following Equation (1) is used. This equation is a step-by-step calculation method for unprotected steel elements exposed to the standard fire, as presented in Buchanan & Abu [ 14]. ∆ T s = ( F V ) ( 1 ρ s c s ) [ h c ( T f − T s ) + σ ε ( T f 4 − T s 4 ) ] ∆ t (1) where Δ Ts is the steel temperature increase (K), ( F/ V) is the steel section factor (m −1) taken as the heated perimeter ( F) divided by the volume ( V), ρs is the steel density (kg m −3), cs is the steel specific heat (J kg −1 K −1), hc is the convective heat transfer coefficient (W m −2 K −1), σ is the Stefan–Boltzmann constant (5.67 × 10–11 kW m −2 K −4), ε is the resulting emissivity (0.70 for carbon steel), Tf is the fire temperature (K) and Ts is the steel temperature (K). This model assumes that the entire steel cross-section is of uniform temperature. When considering the exposed surface of the MPSC 200 steel plate, taken as its perimeter multiplied by its thickness, the temperature profile shown in Figure 28 is obtained. It can be observed that the temperature profile falls within the measured average temperatures up to approximately 30 min. From this profile, the time to charring, assumed as 300 °C, is reached after 28.5 min. As reported in Section 3.1.4, TC4 and TC5 positioned at the top edge of the steel plate reached 300 °C after 28.66 and 27.42 min, respectively. This also coincides with the time when these thermocouples started to flatten to approximately 400 °C ( Figure 16). At the reduced charring rate of 0.33 to 0.40 mm/min, as reported in Section 3.1.4, it would take 62.5 to 75.76 min to fully embed the MPSC 200 steel plate (25 mm thick) into the charred layer, which coincides with the time when TC1 and TC2 started to flatten ( Figure 16). While this approach provides a reasonable understanding of the steel temperature and potential charring at the interface, it assumes full convective and radiative heat fluxes being applied to the surfaces, which may be too conservative. It also reflects the behaviour from a single test. Additional research for refining such a model is needed. 5. Conclusions Point-supported connections are an innovative modern connection design that can benefit from the 2-way structural action of cross-laminated timber (CLT) slabs, which is typically not considered in traditional timber design. It also allows for flatter ceiling surfaces where no beams are needed to support the mass timber floor slabs. In Canada, this type of structure was first used in the 18-storey Tallwood House student residence on the University of British Columbia campus and has been slowly replicated over the years. However, given that the entire mass timber structure was fully encapsulated at that time, there was no need to assess the thermal performance of such a point-supported structure, namely the steel-timber bearing area interface. Two structural fire-resistance tests of point-supported connections were conducted, using the mass timber point-supported connector MPSC 200 from Simpson Strong Tie. The assemblies were exposed to the standard fire of CAN/ULC S101 for 2 h. Test 1 consisted of the MPSC 200 assembly with the gap between the top of the glulam column and the ceiling of the CLT panel being fully exposed to fire (25 mm gap). No failure was reached after 2 h of standard fire exposure. Test 1 has been able to support the full applied load until the actuator reached its maximum stroke at 107 min. Given that the displacement could no longer increase, the resulting applied load slowly reduced from 70 to 40 kN between 107 and 120 min. Differential displacement of the MPSC connector was observed due to charring of the CLT at the steel plate interface. Little to no charring was observed around the square opening in the CLT. Thermocouples above the MPSC steel plate provided low charring rates, suggesting that the MPSC steel plate partially shielded the CLT from direct heat exposure. Test 2 consisted of the MPSC 200 assembly with the 25 mm gap between the top of the glulam column and the ceiling of the CLT panel being sealed using the Fire Foam from Rothoblass. No failure was reached after 2 h of standard fire exposure. Test 2 has been able to support the full applied load until the charred CLT became the weakest component after 117 min. At that time, the load could no longer be maintained and was reduced from 70 to 53 kN between 117 and 120 min. No differential displacement of the MPSC connector was measured. As opposed to Test 1, the CLT panel remained almost intact, with no signs of char above the steel plate and inside the square opening. None of the thermocouples positioned closest to the MPSC connector exceeded 300 °C throughout the 2 h of standard fire exposure, suggesting that the firestop foam effectively prevented heat from penetrating into the MPSC connection. When comparing both assemblies, it was found that leaving the 25 mm gap unprotected against fire at the steel-timber interface influenced the overall behaviour. The following main observations can be drawn: Both assemblies showed similar mid-span deflections and were able to sustain the applied load after 2 h of standard fire exposure, although at a reduced scale due to excessive deflection of the CLT panels. In Test 1, the edges of the steel plate increased to 400 °C and remained stable throughout the test duration, while the temperature in Test 2 barely exceeded 100 °C after 2 h of standard fire exposure. This phenomenon is likely caused by the MPSC connector being embedded into the charred CLT panel and thereby shielding the steel plate perimeter from fire exposure. Given the thermal protection provided to the steel plate in Test 2, there was little to no sign of charring of the CLT in the vicinity and above the steel plate, while Test 1 showed a CLT residual thickness of 120 mm (±3.3 plies left). The CLT panels away from the steel plate exhibited a charring rate consistent with current design standards. A first step-by-step design approach is proposed herein to predict the steel temperature and potential charring at the interface. While it provides reasonable predictions when compared to Test 1, the model assumes full convective and radiative heat fluxes being applied to the surfaces, which may be too conservative. Additional research for refining such a model is needed, namely for gaps of other thicknesses and wood cover of smaller thicknesses, as well as assemblies designed for 1 h of fire-resistance to be used in mid-rise mass timber buildings. Author Contributions Conceptualization, C.D., M.A., C.Y. and A.H.; methodology, C.D. and M.A.; formal analysis, C.D.; writing—original draft preparation, C.D.; writing—review and editing, M.A., C.Y. and A.H.; All authors have read and agreed to the published version of the manuscript. Funding This research was funded by the Forest Innovation Investment of British Columbia (BC FII), grant number 25/26-FPI-PSTM-W26-092. Data Availability Statement The raw data supporting the conclusions of this article will be made available by the authors on request. Acknowledgments The authors acknowledge the collaboration of the staff of the fire safety division of the National Research Council Canada (NRCC) for conducting the fire-resistance tests. The technical and in-kind contributions from Simpson Strong Tie are also acknowledged. Conflicts of Interest Authors Christian Dagenais and Monireh Aram were employed by the company FPInnovations. Authors Claire Yuan and Andrew Harmsworth were employed by the company GHL Consultants, a Part of J.S. Held. The authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest. References Disclaimer/Publisher’s Note: The statements, opinions and data contained in all publications are solely those of the individual author(s) and contributor(s) and not of MDPI and/or the editor(s). MDPI and/or the editor(s) disclaim responsibility for any injury to people or property resulting from any ideas, methods, instructions or products referred to in the content. © 2026 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license. Dagenais, C.; Aram, M.; Yuan, C.; Harmsworth, A. Influence of Point-Supported Steel-to-Timber Interface Parameters on the Structural Fire-Resistance of Mass Timber Plates. Buildings 2026, 16, 2301. https://doi.org/10.3390/buildings16122301 Dagenais C, Aram M, Yuan C, Harmsworth A. Influence of Point-Supported Steel-to-Timber Interface Parameters on the Structural Fire-Resistance of Mass Timber Plates. Buildings. 2026; 16(12):2301. https://doi.org/10.3390/buildings16122301 Dagenais, Christian, Monireh Aram, Claire Yuan, and Andrew Harmsworth. 2026. "Influence of Point-Supported Steel-to-Timber Interface Parameters on the Structural Fire-Resistance of Mass Timber Plates" Buildings 16, no. 12: 2301. https://doi.org/10.3390/buildings16122301 Dagenais, C., Aram, M., Yuan, C., & Harmsworth, A. (2026). Influence of Point-Supported Steel-to-Timber Interface Parameters on the Structural Fire-Resistance of Mass Timber Plates. Buildings, 16(12), 2301. https://doi.org/10.3390/buildings16122301 Note that from the first issue of 2016, this journal uses article numbers instead of page numbers. See further details .
