Fire Performance of Cross-Laminated Timber: A Review of Standards, Experimental Testing, and Numerical Modelling Approaches

Abstract

This review article critically examines the fire performance of cross-laminated timber (CLT), a key structural material for sustainable construction, by synthesising recent advancements in both experimental and numerical research. It identifies a critical gap between experimental findings and numerical models, offering insights to refine future fire-safe design and research. The article assesses fire design strategies across major international standards and reviews experimental fire testing of CLT elements, highlighting how adhesives, protective cladding, layer thickness, load levels, and support conditions affect fire resistance. This article also summarises CLT compartment tests, focusing on how openings, ventilation size, and protective cladding affect fire dynamics and CLT degradation. A literature review of numerically modelled CLT specimens under fire load is compiled and evaluated based on several criteria, including material characterisation, mesh characteristics, and modelling procedures. Subsequently, the outcomes of two distinct approaches are evaluated, emphasising the disparities in the techniques employed and the difficulties inherent in performing more precise numerical simulations. The article will bridge and inform the gap between experimental tests and numerical analysis, focusing on identifying suitable approaches for such simulations. The study aims to provide a broader understanding of the topic and promote the development of fire-safe design and modelling of engineered timber construction using CLT.

1. Introduction

Cross-laminated timber (CLT) is an engineered mass timber product composed of multiple layers of perpendicularly arranged timber planks, which are joined together with adhesive. CLT can be defined as a “prefabricated engineered wood product made of at least three orthogonally bonded layers of solid-sawn lumber that are laminated by glueing of longitudinal and transverse layers with structural adhesives to form a solid rectangular-shaped element intended for roof or wall applications” [1]. The unique cross-lamination structure of CLT results in a more homogeneous product than single timber elements with enhanced structural capacity in both directions, making it an ideal choice for constructing structural wall and floor elements.

As an engineered wood product (EWP) and sustainable construction material, CLT has the potential to address the global rise in carbon emissions associated with construction. CLT also possesses many other advantages, including strength uniformity [2], flexibility, high strength-to-weight ratio [3], and dimensional stability [4], which can be attributed to its enhanced homogeneity due to the dispersion of defects (e.g., knots) within CLT panels [5]. Such advantages have contributed to the growing popularity of CLT, as evidenced by its manufacturing statistics. In Europe alone, annual production has surged from 25,000 m³ in 1996 to an estimated 1.9 million m³ in 2024 [6,7], with projections reaching 4.2 million m³ by 2033 [7] (see Figure 1). While Europe leads in CLT production, the global market has also seen significant growth. Since its commercial introduction nearly three decades ago, CLT’s worldwide production approached 3 million m³ in 2021 [8], with Europe accounting for approximately 48% of this output [9], primarily from Austria, Germany, Italy, Switzerland, and the Czech Republic [10,11]. This is followed by North America (43%), Oceania (6%), and Asia (3%) [9,11]. This ongoing expansion highlights CLT’s growing importance in sustainable construction, with long-term projections estimating cumulative global production of 3.6–9.6 billion m³ by 2100 [12].

Several other design factors must also be considered when using EWPs such as CLT. Wood is combustible and, when subjected to high temperatures, undergoes pyrolysis, resulting in the formation of char, vapour, and gas [13,14]. Although char has a low density and negligible mechanical strength, it exhibits low thermal conductivity and thus acts as thermal insulation [15]. The charring temperature at which the wood converts to char is usually taken at a 300 °C isotherm [16]. As the gas exits the charred wood, it undergoes flaming combustion [17]. The charring rate, which is the rate at which the wood chars, is taken as 0.65 mm/min for solid wood and glued–laminated timber, as defined by EN 1995-1-2 [16]. With the increase in the thickness of the char layer, the initial charring rate gradually decreases to a slower steady rate [18]. Nevertheless, the charring behaviour of CLT is not identical to that of solid wood, as it is influenced by various factors such as the use of adhesive [19], its spread rate [20], and the joints between the lamellae of the CLT panels, making it susceptible to layer fall-off under fire [21].

CLT has undergone substantial development in the construction industry over the past two decades due to its sustainable nature and potential to significantly decarbonise the built environment. Despite its benefits, there are concerns regarding its safety during fire events, such as the recent incident at Nottingham University, where an under-construction mass timber building caught fire, resulting in its complete collapse [22]. Recent research has focused on analysing the fire performance of CLT structural systems, and there is a need to gather and assess the various factors that influence its behaviour during a fire. This article provides an overview of fire resistance and compartment tests to evaluate the behaviour of CLT under fire exposure. Furthermore, the current version of Eurocode 5 on the structural fire design of timber structures is analysed alongside a comparative overview of other international standards. The article also examines the fire design of CLT, with a focus on studying the effect of protective cladding on CLT in a fire. The review of the experimental studies is further extended to the numerical modelling of CLT using finite element (FE) analyses. Finally, the conclusions and discussions are presented, which summarise the existing knowledge and gaps, with recommendations for future research and development.

A comprehensive systematic literature review was conducted on the fire performance of CLT from January 2024 to August 2025. The search strategy employed databases such as Scopus, ScienceDirect, Google Scholar, SpringerLink, and MDPI, utilising keywords including “cross-laminated timber,” “CLT in fire,” “fire resistance,” “charring rate,” “compartment fire,” and “numerical modelling.” The review focused on the literature published between 2015 and 2025, while also incorporating pre-2015 studies that offered foundational insights into timber fire behaviour and modelling methodologies. Only peer-reviewed articles, conference proceedings, and technical reports were included, following a screening process to ensure their relevance to CLT fire exposure analysis.

2. Thermal Response of Timber and CLT

Although the pyrolysis process starts above 200 °C during a fire, wood undergoes charring when its surface temperature reaches approximately 300 °C, where the wood decomposes into combustible gases, which serve as fuel and burn, leaving black char [18]. As per EN1995-1-2 [16], the position of the char-line is taken at the 300 °C isotherm. The thickness of the char layer increases as more wood burns, which acts as thermal insulation to the wood immediately beneath it. The char base acts as a boundary line between the charred layer and the pyrolysis zone. The initial charring rate reduces to a slower steady rate with the increase in the thickness of the char layer [18]. As the temperature of wood rises above 100 °C, the moisture in the wood evaporates, including the travel of moisture to the burning face, increasing the moisture content below the char front [23]. The different layers of wood when exposed to fire are illustrated in Figure 2.

The distance between the outer surface of the wood and the position of the char layer is the charred depth, which is calculated from the fire exposure time and the charring rate, as shown in Equation (1). The design charring rate for softwood is 0.65 mm/min for solid timber and glued–laminated timber with a density greater than 290 kg/m³, as specified in EN 1995-1-2 [16]. However, the charring rate and fire behaviour of CLT are not the same as those of solid timber because of the adhesive type [19,21], adhesive spread rate [20], layer layup [19], and joints between the lamellae of the CLT panels. The strength of adhesive joining two adjacent CLT layers reduces at elevated temperatures [19], which can make CLT susceptible to delamination of the outer layer [21] and thus expose the inner layer of CLT directly to fire. Practically, delamination exposes uncharred wood, increasing surface and internal temperatures [24,25], promotes re-ignition, and delays self-extinction, resulting in prolonged burning [26]. This can lead to fire regrowth after decay or no decay of the fire [27], potentially compromising structural integrity. After delamination, temperature rises, charring deepens, heat flux peaks, and flames reappear, potentially delaying the decay phase and extending the fire duration [26].

A charring model, proposed by Klippel et al. [28] for CLT panels, defines three different cases: no fall-off of layers, less pronounced fall-off, and significant fall-off of layers. For CLT floor and wall panels, with no layer fall-off, a design charring rate of 0.65 mm/min as recommended by EN-1995-1-2 [16] can be used. However, if there is a less pronounced fall-off of layers in the CLT wall panel, then a single charring phase should be used at a rate of 0.8 mm/min throughout the fire exposure time. Three different charring phases should be used for the CLT floor panel if a significant fall-off of the layers occurs. This includes an initial charring rate of 0.65 mm/min until the outer exposed layer of the CLT floor panel is completely charred. The charring rate in every inner layer is then doubled to 1.30 mm/min until a 25 mm thick char layer is formed, after which it reverts to a charring rate of 0.65 mm/min.

$$d_{char,0}={\beta }_0 \times t$$

(1)

where d_char,0 is the design charring depth for one-dimensional charring (fire exposure on one side), β₀ is the one-dimensional design charring rate for standard fire exposure, and t is the time for fire exposure.

EN-1995-1-2 [16] provides the temperature-dependent thermal properties of wood, including density, specific heat, and thermal conductivity when subjected to standard heating conditions, as depicted in Figure 3 and Figure 4. It can be observed that with a decrease in density at elevated temperatures, the thermal conductivity, or the rate of heat flow through a unit cross-sectional area, increases. The specific heat capacity, which is the amount of heat required to increase the temperature by a certain degree, typically decreases as the temperature of the wood increases. However, there is an enormous increase in the specific heat capacity of wood at temperatures between 99 °C and 120 °C due to thermal degradation processes, which involve moisture release and breakdown of organic compounds.

3. Structural Integrity of CLT in Fire

The structural integrity of CLT panels during fire exposure is influenced by both loading level and boundary restraint. The study conducted by Wiesner et al. [29] revealed that global instability is the dominant failure mode for CLT walls in fire, with CLT using polyurethane (PUR) adhesive failing structurally earlier than those with melamine urea formaldehyde (MUF) adhesive. Additionally, three-ply walls failed earlier than five-ply walls, with these factors together halving the failure time under identical heating conditions [29]. From a structural fire engineering standpoint, a slow-growing fire of extended duration can have more severe impacts on structural load capacity than a shorter, more intense fire [30]. Similarly, Xing et al. [31] found that two-way supported CLT slabs redistributed load more effectively and resisted deflection longer than one-way slabs. Compared with one-way slabs, two-way configurations, i.e., restrained on all sides, exhibit hyperbolic deformation and enhanced cooperative load-bearing capacity. During fire exposure, the reduction in wood strength with temperature leads to progressive deflection until failure [31]. Collectively, these findings underscore the critical importance of realistic load and support conditions in evaluating the real-world fire resilience of CLT.

The performance and structural integrity of CLT after a fire in the decay phase are also significantly important and have been addressed in the literature. Wiesner et al. [32] performed experimental studies on fire-exposed CLT elements in compartments and concluded that the structural and thermal behaviour of CLT elements are dependent on each other and should not be taken separately. The load-carrying capacity of slabs and walls continues to decrease in the decay state of a fire, or even after the charring has been stopped. This reduction in capacity in the decay phase was attributed to the propagation of a thermal wave underneath the char layer [32].

Another study by Vairo et al. [33] analysed the behaviour and structural integrity of CLT panels during and after a fire. Loss of stiffness and strength was observed in the CLT slab when it was reloaded 24 h after the fire test due to the propagation of thermal waves during the decay phase. In addition, the deflection increased up to 50% when a similar load was reapplied 24 h after the fire, compared to the deflection measured soon after the fire test. It is important to note that structural failure is more prominent in the decay phase when the timber of the CLT slab is exposed to a longer thermal attack with a maximum temperature similar to that of a shorter but intense thermal attack [30]. This is because a longer thermal attack will lead to a greater char depth, as well as higher thermal penetration.

3.1. Determination of Cross-Sectional Properties

EN-1995-1-2 [16] provides two simplified methods for the fire design of members exposed to Standard ISO 834 [34]: the reduced cross-section method and the reduced properties method. Figure 5 displays temperature-time curves for both standard and natural fires. The former refers to controlled and repeatable testing conducted under predictable conditions, while the latter represents real scenarios with variable temperature–time curves influenced by factors such as ventilation size and fuel type. The reduced cross-section method measures the charred depth of timber when exposed to standard fire by considering that the char will occur at a uniform charring rate of 0.65 mm/min for solid timber with a density greater than 290 kg/m³. Additionally, a 7 mm section, known as the zero-strength layer (d₀), is added to the char depth. This layer, adjacent to the charred area, is considered to have zero load-bearing capacity, as illustrated in Figure 6a. The reduced properties method considers the reduction in mechanical properties based on the load type and cross-section [16]. However, this method is applicable when the cross-section is exposed to fire from three or four sides, which is not typically the case for assessing the fire resistance of CLT elements. In the reduced cross-section method, the initial cross-section is reduced by effective charring depth (d_ef) as shown in Equation (2).

$$d_{ef}=d_{char,n}+k_0 \times d_0$$

(2)

where d₀ = 7 mm, and $d_{char,n}={\beta }_n \times t,$ k₀ is a coefficient; its value for an unprotected surface is provided in

Table 1. k₀ for unprotected surfaces.

	k0
t < 20 min	t/20
t ≥ 20 min	1.0

. Table 1 also applies to protected surfaces with t_ch ≤ 20 min. For protected surfaces with t_ch > 20 min, k₀ varies linearly from 0 to 1 in the time interval from t = 0 to t = t_ch, as shown in Figure 6b.

3.2. Effects of Fire on Mechanical Properties

When the temperature of timber is raised, its mechanical properties undergo variations [36]. Changes in the mechanical properties are reversible when the timber temperature is rapidly reduced to the ambient temperature after being raised to a temperature of less than 100 °C [37]. However, when subjected to temperatures exceeding 100 °C, timber undergoes an irreversible effect characterised by a permanent reduction in density and other material properties. EN 1995-1-2 [16] provides relevant reduction factors in the mechanical properties for design purposes at higher temperatures, as illustrated in Figure 7.

4. Fire Protection Measures for CLT

CLT can be designed and used as an exposed structural component. However, fire protection measures are often adopted to increase the performance and integrity of CLT elements during fires. Different protective measures have been used and analysed to protect and enhance the performance of CLT in fire. These measures include both passive and active fire protection strategies using non-combustible cladding, such as gypsum plasterboard, fire retardants, and sprinkler systems, to protect the CLT from direct exposure to fire.

Gypsum plasterboards are widely utilised as fire protection systems due to their ease of installation, reliability, and high specific heat capacity [38]. Gypsum is a crystalline mineral composed of calcium sulphate, chemically bound crystallisation water (21% by weight), and a low proportion of absorbed free water [39]. The calcination process initiates with an increase in temperature above 80 °C, which is the thermal degradation that results in the separation of water from the crystal lattice. The chemically bound and free water in gypsum play a significant role in fire resistance at high temperatures. When exposed to fire, these boards can absorb a substantial amount of heat at temperatures up to 100 °C, as they release water, thereby delaying the rise in temperature until the entire board is dehydrated. Additionally, the low thermal conductivity of gypsum plasterboard helps maintain lower temperatures on the unexposed side to the fire [40]. The literature shows that gypsum plasterboards with different densities, fillers, and fibres exhibit similar overall thermal behaviours. However, the reinforcement of the gypsum core enhances its mechanical properties, including shrinkage, ablation, and cracking. The thermal behaviour of gypsum board is also influenced by the layer behind it [38]. Energy is used up by the crystal water evaporation, and a protective steam shield is made on the fire-exposed side of the component. This not only slows down the fire’s progression but also allows the dehydrated gypsum layer to serve as an insulator by decreasing thermal conductivity [41]. Gypsum boards are classified into two main types: regular and fire-rated. The gypsum core in both regular and fire-rated gypsum boards provides fire resistance.

4.1. Design Considerations for Fire Protection (EN 1995)

In the case of timber surfaces protected by fire protective components, the following must be considered:

• No charring will occur until time t_ch (time of start of charring behind the protective layer).
• Charring may start before the failure of fire protection cladding; however, the charring rate is lower than as specified in

  Table 2. Design charring rates of wood and LVL, adopted from EN-1995-1-2 [16].

  Density (kg/m3)	Material		β0 (mm/min)	βn (mm/min)
  ≥290	Softwood and Beech	Glued–laminated timber	0.65	0.70
  		Solid timber	0.65	0.80
  290	Hardwood	Solid or glued–laminated timber	0.65	0.70
  ≥450			0.50	0.55
  LVL with a density of ≥480 kg/m3			0.65	0.70

  until the fire protection failure time, t_f (see Figure 8a).
• The charring rate increases above the values shown in Table 2 after the failure time, t_f, of the fire protective cladding until time t_a.
• At time t_a, the charring rate returns to the values as specified in Table 2, when the charred depth is the least of either the charring depth of the same element with no fire protection or 25 mm.
• A better understanding of these points can be obtained by referring to Figure 8 and Figure 9.

4.1.1. Charring Rate for Initially Protected Components

• For t_ch ≤ t ≤ t_f, EN 1995-1-2 [16] provides a factor, k₂ (insulation coefficient), that needs to be multiplied by the charring rate.

For a single-layer Type F gypsum plasterboard, this is calculated as

k₂ = 1 − 0.018 ∗ h_p (h_p is the thickness of gypsum plasterboard in mm).

h_p should be taken as the thickness of the inner layer, if the cladding consists of more than one layer of Type F gypsum plasterboard.

• If t_f ≤ t ≤ t_a, the factor k₃ = 2 should be multiplied by the values of the charring rates in Table 2 [16]. When t ≥ t_a, the values of the charring rates, as noted in the table, should be considered with no modifications [16].
• For t_ch = t_f, the time limit, t_a (see Figure 5 and Figure 6), is calculated as the minimum of Equations (3) and (4).

$$t_a={2\times t}_f$$

(3)

$$t_a={\displaystyle \frac{25}{k_3\times {\beta }_n}}+t_f$$

(4)

For t_ch < t_f,

$$t_a={\displaystyle \frac{25-\left(t_f-t_{ch}\right)\times k_2\times {\beta }_n}{k_3\times {\beta }_n}}+t_f$$

(5)

where β_n is the design value of the notional charring rate in mm/min.

The above equations also apply to one-dimensional charring, where β_n is replaced by β₀.

4.1.2. Start of Charring for Initially Protected Components and Their Failure Times

The start of charring of the timber element is delayed by time t_ch, when the timber surface is protected with either a single- or two-layer gypsum plasterboard.

Table 3. Calculation of the start of charring, t_ch, for timber element with different cladding and joint conditions [16].

Protective Claddings		Condition	Equation b
Type A, F, H	Single layer	Location adjacent to ≤2 mm unfilled gaps	tch = 2.8 hp − 14	(6)
		Location adjacent to >2 mm unfilled gaps	tch = 2.8 hp − 23	(7)
Type A or H	Two layers a	Use Equation (6), with hp calculated as hp = 1 × outer layer thickness + 0.5 × inner layer thickness
Type F		Use Equation (6), with hp calculated as hp = 1 × outer layer thickness + 0.8 × inner layer thickness

Note: ^a The spacing of fasteners in the inner layer should not be more than the fastener spacing in the outer layer [16]. ^b h_p is the panel thickness, in mm.

outlines the conditions and equations for calculating the start of charring behind protective cladding. For single-layer protective cladding, the start of charring behind the cladding depends on whether adjacent joint gaps are ≤2 mm or >2 mm.

The failure of fire-protective cladding may happen as a result of the charring or mechanical deterioration of the cladding material, inadequate spacing and distances between fasteners, or insufficient penetration depth of fasteners into the uncharred cross-section [16]. For Type A and H gypsum plasterboard, the failure time is taken as when the charring behind the cladding initiates; i.e., t_f = t_ch. In the case of gypsum Type F plasterboard, failure times should be assessed based on the thermal degradation of the cladding and the pull-out failure of fasteners caused by insufficient penetration length into unburnt wood. EN 1995-1-2 [16] specifies the penetration length of fasteners (l_a), which should not be less than 10 mm into the uncharred timber. The required fastener length (l_f,req) is calculated using Equation (8) [16].

L_f,req = h_p + d_char,0 + l_a

(8)

where d_char,0 is the charring depth in the timber component, and l_a is the minimum penetration length of the fastener into the uncharred timber.

4.1.3. Challenges in Implementing Fire Protection Technologies

Although several fire protection technologies for CLT, such as protective cladding and encapsulation, show promising performance under laboratory and simulation conditions [42,43], their practical application faces notable challenges. These include increased project costs, complex installation requirements, and variability in on-site quality control. Stakeholder surveys in Australia [44] show that a lack of understanding of fire safety, regulations, performance, and local manufacturers and suppliers are major barriers to mass timber construction. Studies indicate that CLT buildings often incur higher material, transport, and fire protection costs than concrete structures; however, these are partly offset by shorter construction times and reduced foundation requirements [45]. Additionally, simulation studies employing computational fluid dynamics (CFD) based methodologies can effectively guide fire protection design strategies for CLT structures, offering potential cost savings by optimising material use and reducing structural damage [46].

While Eurocode 5 [16] provides the most widely studied framework for CLT fire design, alternative standards in Canada, the United States, and Australia/New Zealand adopt different approaches reflecting local testing traditions, species, and regulatory practice. These are summarised in Section 5.

5. International Standards

5.1. Canadian Standard CSA O86

The Canadian wood design standard (CSA O86) [47] includes explicit provisions for fire resistance of timber members, largely derived from extensive compartment fire testing programs. For CLT, a one-dimensional charring rate of 0.65 mm/min is recommended, while the char front remains within the first lamella, but if inner layers are affected, an average rate of 0.8 mm/min should be applied for the full fire exposure duration [47]. Buckling instability is accounted for using a reduction factor on compressive strength, similar in principle to Eurocode 5 and NDS (National Design Specification) methods, with increased strength reductions for slender members. Encapsulation is treated in simplified form: for example, 12.7 mm of Type X gypsum board is assumed to add 15 min of fire resistance [47], although the complexities of fall-off and its effect on charring are not explicitly considered [10].

5.2. American Standards: NDS

In the United States, the National Design Specification (NDS) [48] provides a mechanics-based framework for assessing the fire resistance of CLT and glulam by comparing residual capacity against applied load. A basic one-hour charring rate of 0.635 mm/min is applied, with an additional 20% increase introduced to account for corner rounding and the strength loss of heated timber [15,49]. The NDS also includes empirical equations to adjust effective charring depth for lamella fall-off, linking char advance to the number and thickness of CLT layers. The effective char depth values are tabulated as a function of both the individual lamination thickness and the required fire resistance rating. Notably, thicker laminations result in reduced effective char depths, indicating improved fire performance. For instance, to achieve a 1 h fire resistance rating, CLT with 16 mm (5/8-inch) laminations requires an effective char depth of 56 mm (2.2 inches), while 51 mm (2-inch) laminations require only 46 mm (1.8 inches) of char depth [48]. Instability is addressed through a reduction factor (k_c) that depends on member slenderness and stiffness, similar in principle to Eurocode 5. However, the NDS method only applies to exposed members and does not explicitly capture the encapsulation effects [10].

5.3. Australian and New Zealand Standards (AS 1720.4)

The fire design of timber structures in Australia and New Zealand is guided by AS 1720.4 [50], which provides calculation methods for charring and residual cross-section capacity but is not yet tailored specifically for CLT. The standard specifies a nominal one-dimensional charring rate of about 0.65 mm/min for softwoods such as radiata pine (the dominant species in NZ/Aus CLT). AS 1720.4 also includes provisions for protected members, where the effective depth of charring (d_c) behind fire-resistant insulation is determined by Equation (9).

d_c = c (t − t₃₀₀) + 7 mm

(9)

where c is the notional charring rate (in mm/min), and t₃₀₀ represents the time for timber beneath insulation to reach 300 °C in a standard test; if insulation falls off prematurely, a notional rate of 2c is applied until a 25 mm char depth is reached. Since CLT-specific rules are not explicitly covered, designers often adopt Eurocode-based methods or manufacturer design guides (e.g., XLam [51]) in practice. Buchanan et al. (2018) [52] noted that CLT fire design in New Zealand and Australia is mainly performance-based, with prescriptive provisions limited. This gives designers greater flexibility but increases reliance on professional competence and regulatory approval.

5.4. Review of Standards and Practical Implications

A review of these standards shows that while most regions adopt a reduced cross-section or mechanics-based approach, the assumed charring rates and treatment of instability or encapsulation differ notably between codes. Eurocode 5 provides a widely applied baseline but is limited to softwood CLT, whereas CSA O86 introduces higher effective charring rates once deeper layers are involved. The USA NDS method applies a lower nominal charring rate but artificially increases it to account for strength loss and lamella fall-off. In Australia and New Zealand, AS 1720.4 specifies general timber charring rates but lacks CLT-specific provisions, leading to reliance on performance-based design.

Table 4. Comparison of international standards for CLT fire design.

Standard/Code	Region	Design Basis	Charring Rate/Fire Resistance Assumptions	Key Notes
Eurocode 5 (EN 1995-1-2) [16]	Europe	Reduced cross-section and effective charring	~0.65 mm/min (softwoods); zero-strength layer (ZSL) of 7 mm	Most widely used baseline; no specific design method for CLT; handbook used as advisory guidance for CLT fire design [53].
CSA O86 [47]	Canada	Mechanics-based; tabulated fire ratings	0.65 mm/min (1st lamella); 0.8 mm/min deeper layers; 7 mm ZSL [15]	Based on Canadian test data; includes buckling reduction factor and encapsulation allowances.
NDS [48]	USA	Mechanics-based (NDS) + prescriptive ratings (IBC, ASTM E119) [54]	0.635 mm/min + 20% (includes heated-zone strength/stiffness loss) [48]	Empirical equation for lamella fall-off; instability via kc; only covers exposed members [15].
AS 1720.4 [50]	Australia/NZ	Reduced cross-section method	0.65 mm/min for radiata pine; 7 mm ZSL assumed	CLT not explicitly covered; performance-based compliance common; Eurocode and manufacturer guides often used.

summarises the main provisions across these codes, highlighting both shared principles and regional differences.

The comparative analysis of EC5, CSA O86, NDS, and AS 1720.4 reveals that while all utilise similar mechanics-based methodologies, their suitability differs depending on the region and type of timber. According to The Fire Safe Use of Wood in Buildings [55], when uniform design assumptions and load combinations are applied, these frameworks produce generally comparable outcomes for the fire-resistance design of timber elements. Nonetheless, the codes differ in their underlying philosophy and level of conservatism. The meta-analysis conducted by Wiesner et al. [15] identified Eurocode 5 (together with the charring model by Klippel et al. [56]) as having the lowest mean absolute percentage error (MAPE ≈ 25%) for predicting CLT wall fire resistance. It suggests that Eurocode 5 currently provides the most accurate framework for practice, while performance-based approaches are essential for cases outside standardised furnace testing. Moreover, the methodology provided by NDS [48] has been shown to be statistically non-conservative in 88% of the considered cases [15].

From a practical perspective, designers should choose the standard based on the local timber species, the fire test data available, and the regulatory jurisdiction. Achieving future harmonisation of these standards, particularly in defining equivalent charring rates, zero-strength layers, and encapsulation effects, is vital for ensuring consistency in global CLT fire design practices.

6. Case Studies and Experimental Research

Case studies and experimental research have been conducted, which can largely be subdivided into two categories based on whether the tests were performed under certain predefined temperature conditions or real fire scenarios. The tests are typically performed under predefined temperature conditions, such as the standard ISO 834 curve heating conditions [34] in a furnace, and used to evaluate the performance and fire behaviour of an individual component. Real fire tests are large-scale tests on CLT compartments where the structure is exposed to a specific quantity of fuel rather than predefined temperature–time curve conditions. Different factors affect the intensity and temperature of a fire in a compartment test, including the amount and source of fuel in a compartment, ventilation, or opening area and the properties of the protective cladding. Because of their different parameters and approaches for measuring the fire performance of CLT, both of these test types are discussed separately, in which the fire resistance tests provide key baseline information, and the real compartment fire tests provide a detailed insight into the behaviour of the material, along with the fire safety measures

6.1. Overview of CLT Fire Resistance Tests

Extensive experimental research has been conducted in recent years on the behaviour and performance of CLT under exposure to fire [27,42,57,58,59]. This paper examines fire tests, focusing on the effect of structural load, the type and grade of CLT element, orientation of CLT layers, layer thickness, and adhesive used on fire performance. An important aspect considered and analysed was the charring rate of the CLT elements when exposed to fire. As noted earlier, EN 1995-1-2 [16] defines the charring of timber as occurring at 300 °C for solid or glued–laminated timber. The various factors that influence the performance of CLT at elevated temperatures have been summarised in

Table 5. Recent studies focusing on different parameters affecting the fire performance of CLT [60].

Authors and References	Research Focus in Terms of Fire Performance of CLT	Authors and References	Research Focus in Terms of Fire Performance of CLT
Wiesner et al. [61] Zelinka et al. [62] Muszyński et al. [63]	Adhesive performance	Kolaitis et al. [64] Li et al. [65] Johnson et al. [66]	Protective cladding
Bartlett et al. [14]	Heating conditions, sample orientation, and size	Klippel et al. [67]	CLT layup, wall and floor panels, support conditions
Frangi et al. [57]	Adhesive, thickness of layer	Wang et al. [58]	Thickness and number of layers
Moser et al. [68] Yasir et al. [42]	Protective cladding	Friquin et al. [69]	Layer and panel thickness, fire intensity
Westhuyzen et al. [70]	Analysis of the fire performance of CLT wall panels	Mindeguia et al. [27]	Thermo-mechanical behaviour of CLT slab under fire
Xing et al. [35]	Number of layers, load ratios, validation of FE analysis, thickness of the zero-strength layer (ZSL)	Frangi et al. [71]	Validation of FE analysis and comparison of CLT panels with homogeneous timber elements.
Yasir et al. [72]	Protective cladding, number of layers, validation of FE analysis	Lucherini et al. [73]	Intumescent coating
		Wiesner et al. [74]	Layers layups
Hasburgh et al. [75]	Ply configuration, types of adhesives	Suzuki et al. [76]	Adhesive, layer thickness, wood specie

and are discussed in more detail in the following subsections. Furthermore,

Table A1. CLT floor panels directly exposed to fire.

Authors	Fire Curve	Wood Species	Density (kg/m3) and/or Wood Grade	Layers Layout (mm)	M.C (%)	Adhesive	Dimensions (mm × mm × mm)			Charring Rate (mm/min)
							L	W	T
Friquin et al. [69]	Parametric	Norway spruce	440	19.5-30-21-30-19.5	8–9.3	MUF	3600	1200	120	0.95
				29.5-39-32-39-32-39-29.5					240	0.68
	Standard			32-41-34-41-32					180	0.48
				31.5-21-32-21-32-21-32-21-28.5					240	0.71
	Swedish			19.5-30-21-30-19.5					120	0.50
				29.5-39-32-39-32-39-29.5					240	0.50
Muszyński et al. [63]	Standard	Spruce–pine–fir	472	35-35-35-35-35	10.3	PUR	5486	4267	175	0.61
		Douglas fir–larch	532	35-35-35-35-35	12.2	PUR				0.63
		Douglas fir–larch	554	35-35-35-35-35	13.0	MF				0.55
Frangi et al. [57]	Standard	Spruce	405–486	10-10-10-10-20	10	MUF	1150	950	60	0.58
						PU1				0.94
						PU2				0.90
						PU3				1.0
						PU4				1.08
				20-20-20		PU3				0.85
						PU4				0.89
						PU5				0.76
				30-30		MUF				0.61
						PU1				0.80
Klippel et al. [67]	Standard	Spruce–pine–fir	C16	34-19-34-19-34	12	PUR	4800	1200	140	0.81
				34-24-24-24-34						0.78
Mindeguia et al. [27]	Standard	Spruce	C24	33-33-33-33-33	11.6–12.6	PUR	5900	3900	165	0.71–0.75
Xing et al. [35]	Standard		580	33-33-33	12 ± 0.7	PUR	3400	420	105	0.67
				21-21-21-21-21						0.77
	Natural fire			33-33-33						0.92
				21-21-21-21-21						1.18
Hasburgh et al. [75]	Standard	Southern pine	Not provided	35-35-35		MF	1194	965	105	0.67
						PRF				0.73
						PUR				0.70
						EPI				0.73
				35-35-35-35-35		PUR			175	0.68
						PRF				0.67
Wang et al. [58]	Standard	Canadian hemlock lumber	480	35-35-35	13	PUR	2200	420	105	0.95
				21-21-21-21-21						0.92
Lucherini et al. [73]	50 kW/m2	Australian softwood	-	20-20-20-20-20	-	-	200	200	100	0.76
Suzuki et al. [76]	Standard	Japanese cedar		19.3 mm × 7 plies	12.2	RPF	1500	450	135	0.66
				27 mm × 5 plies	13.1	API				0.70
				45 mm × 3 plies	11.4					0.76
				19.3 mm × 7 plies	12.3					0.70
		Japanese larch		19.3 mm × 7 plies	9.5					0.64

,

Table A2. CLT wall panels directly exposed to fire.

Authors and Reference	Fire Curve	Wood Species	Density (kg/m3) or Grade	Layers Layout (mm)	M.C (%)	Adhesive	Dimensions (mm × mm × mm)			Load	Charring Rate (mm/min)
							L	W	T	(kN)
Westhuyzen et al. [70]	Standard	SA pine	479	33-33-33	14.2	PUR	900	900	100		0.95
		Eucalyptus	552	33-33-33	14.8						0.76
Yasir et al. [42]	Standard	Irish spruce	381 (C16)	40-40-40	12–13	PUR	1200	900	120	85	0.66
Aloisio et al. (2025) [82]	Standard	-	450	6 × 30 = 180		None, Wooden Dowels	2980	3000	180	50 kN/m	0.9
Klippel et al. [67]	Standard	Spruce–pine–fir	C16	34-34-34 a	12	PUR	3000	480	102		0.64
				34-34-34 b							0.72
				34-19-34-19-34 a				660	140		0.74
				34-19-34-19-34 b							0.74
				34-24-24-24-34 a							0.74
				34-24-24-24-34 b							0.73
Bartlett et al. [14]	Standard	Sitka spruce (426) Scotts pine (501)		40-40-40	No data	PUR	300	200	120		0.70
	Slow heating fire curve										0.59
	8.33 W/m2min2										0.43
	12.5 W/m2min2										0.45
	16.7W/m2min2										0.48
Wiesner et al. [74]	heat flux of 50 kW/m2	Spruce and pine	Not given	33-34-33	10.3	MF	1700	300	100	44.1	0.82
				33-34-33						88.2	0.88
				20-20-20-20-20						40.8	0.98
				20-20-20-20-20						81.6	1.0
Frangi et al. [71]	Standard			28-28-28		PUR			84		0.68
				17-17-17-17-17							0.64

^a T-support; ^b L-support.

and

Table A3. Initially protected CLT panels.

Authors	Wood Species	Density (kg/m3)	Layers Layout (mm)	M.C (%)	Adhesive	Fire/Heat Curve	Type of Protection	Delay in Charring (min)	Charring Rate (mm/min)
Moser et al. [68]	Radiata pine	460 ± 13	35-35-35	9 ± 1	No data	50 kW/m2	Gypsum board, 13 mm	24	0.61
						65 kW/m2		22.5	0.40
						50 kW/m2	Gypsum board (FR), 13 mm	26.5	0.48
						65 kW/m2		24.5	0.55
						50 kW/m2	MgO board, 12 mm	18	0.48
						65 kW/m2		15	0.56
						50 kW/m2	MgO board, 15 mm	21.5	0.43
						65 kW/m2		16.5	0.51
						50 kW/m2	Fire-rated fibrous board, 12.5 mm	21.5	0.43
						65 kW/m2		16.5	0.51
Lucherini et al. [73]	Australian softwood	-	20-20-20-20-20	-	-	50 kW/m2	0.60 mm WTF (0.51mm DFT)	15	0.34
							1.60 mm WTF (1.31 mm DFT)	26	0.27
							2.50 mm WTF (2.13 mm DFT)	44	0.20
Yasir et al. [72]	Irish Sitka spruce	381	40-20-20-20-40	12 ± 1	PUR	Standard	12.5 mm Fireline gypsum plasterboard	20	0.71
			40-30-40				12.5 mm Fireline gypsum plasterboard and 12.5 mm Plywood	25.5	0.74
Yasir et al. [42]	Irish Sitka spruce	381	40-40-40	12 ± 1	PUR	Standard	15 mm FireLine gypsum plasterboard with no joints	30	0.44
							15 mm FireLine gypsum plasterboard with joints	25	0.46
							12.5 mm FireLine gypsum plasterboard and 25 mm plywood with joints in both	44	0.67

in the annexes of this article provide detailed information on the CLT panels, their testing conditions, and their charring rates.

6.1.1. Loading and Support Conditions

Several studies have explored the impact of loading and support conditions on the fire resistance of CLT panels. Klippel et al. [67] experimented on C16 CLT wall panels of different cross-sectional layups with T-supports and L-supports. The average charring rate was 0.72 mm/min, and there was no significant influence of the support conditions on the CLT panels’ performance in fire. Wang et al. [58] conducted tests on loaded three- and five-layered CLT floor elements and cautioned against the use of a three-ply CLT panel under a high load due to the direct exposure of the load-bearing layer to the fire. Suzuki et al. [76] conducted empirical tests to investigate the structural behaviour of CLT wall panels under various loading scenarios in fire. At higher loads, i.e., greater than one-third of the panel’s axial load capacity, the specimens failed when the outer layer was completely charred. However, for CLT specimens with loads of one-fourth to one-eighth of the axial load capacity, failure occurred due to the eccentricity of the panel during the charring of the second orthogonal layer from the fire-exposed face. The study performed by Xing et al. [31] showed that two-way-supported CLT slabs exhibited greater structural cooperation and higher load-bearing capacity compared to one-way slabs. However, under fire exposure, the slabs still failed in a brittle manner once critical vertical deflection was reached. Their work highlights that boundary conditions, particularly two-way versus one-way action, are critical determinants of CLT fire resistance.

6.1.2. Wood Grade and Density

The type and density of the wood species of CLT are also important parameters that have been studied under fire-exposure conditions. In the experimental tests performed by Suzuki et al. [76], high-density wood CLT panels exhibited better fire performance and lower charring rates than those made with low-density wood. Lizhong et al. [77] tested three wood species, Paulownia, Toon, and Elm, with densities of 260, 530, and 590 kg/m³ under constant and time-increasing heat fluxes, and they found that increasing wood density resulted in a decrease in charring rates for both exposure scenarios. However, Frangi et al. [78] did not observe a correlation between wood density and charring rate for timber elements with densities ranging between 340 and 500 kg/m³ under standard ISO fire exposure.

6.1.3. Orientation and Setup of CLT Layers

The strength and stiffness of CLT elements are dependent on the orientations of the CLT’s layers when loaded at ambient temperatures [79]. However, the orientation of the layers of CLT elements does not significantly influence their fire performance and charring rate when exposed to elevated temperatures [67,71]. This is supported by a study conducted by Hasburgh et al. [75], who examined two different orientations of CLT layers, long–cross–long (LCL) and long–long–cross (LLC), in which the failure time of the LLC configuration was found to be only approximately 5% longer than the LCL configuration. However, the LLC configuration resulted in uneven charring and deep cracks, while the LCL specimen showed no such cracks under fire, as shown in Figure 10 [75].

Another factor that increases the charring rate of CLT is the gap between boards with larger gaps, allowing hot gases to pass through [80]. Therefore, the Technical Guidelines for Europe [53] limit the gap to 2mm between the boards of CLT panels without having to increase their charring rate beyond the one-dimensional charring rate, but for gaps between 2 mm and 6 mm, a higher charring rate is proposed [53]. The American National Standard [1] permits a gap of up to 6.4 mm between the adjacent lamination edges in the face layers of the CLT panels.

6.1.4. Layer Thickness

The thickness of individual layers in CLT elements in fire has been a focus in recent research. Studies have shown that the charring rate of CLT decreases with an increase in the thickness of lamellae [35,57,76]. In a study performed by Frangi et al. [57], thicker layer panels were found to have a lower charring rate than panels with thinner layers. The improved fire performance was attributed to the less pronounced char fall-off. Xing et al. [35] performed experimental tests on three- and five-layered CLT floor panels exposed to either a standard or a natural fire curve. The fall-off of char occurs irrespective of whether it is exposed to a natural or standard fire, and the panels also showed that a one-dimensional char rate of 0.65 mm/min, as recommended by EN 1995–1-2:2004 [16], is unsafe for CLT in natural fires and should not be used [35]. A higher charring rate was observed for CLT panels with thinner layers compared with thicker layers for both the natural and standard fire curves [35], as shown in

Table 6. Effect of layer thickness on charring rate.

Reference	Layer Layup (mm)	Adhesive c	Charring Rate (mm/min)
Friquin et al. [69]	32-41-34-41-32	MUF	0.48
	31.5-21-32-21-32-21-32-21-28.5		0.71
Xing et al. [35]	33-33-33	PUR	0.67
			0.92 a
	21-21-21-21-21		0.77
			1.18 a
Wiesner et al. [74] b	33-34-33	MF	0.82–0.88
	20-20-20-20-20		0.98–1.0

^a Natural fire; ^b heat flux of 50kW/m²; ^c MF/MUF: melamine (urea) formaldehyde; PUR: polyurethane.

. However, considering the fire performance of CLT walls under sustained loading, Wiesner et al. [29] observed that five-layer panels with a ply configuration of 20-20-20-20-20 mm failed later than three-layer panels with a ply configuration of 40-20-40 mm. The authors attributed this to the charring of the orthogonal layer during fire exposure of the five-layer CLT element, which did not contribute effectively to the load-carrying capacity. Similar conclusions about the earlier failure of the three-layer CLT panel due to charring of the outer layer were made by Wiesner et al. [74] when comparing five-layer panels with a layup of 20-20-20-20-20 mm with a three-layer CLT layup of 34-33-34 mm.

6.1.5. Type of Adhesive

One of the key aspects of CLT when exposed to fire is the delamination or fall-off of the charred layer. This occurs when the fire-exposed layer separates or detaches from the main body of the CLT, exposing the underlying layer directly to the fire [81], as shown in Figure 11. The direct exposure of the uncharred inner layer of CLT to fire results in an increase in the charring rate due to the loss of a protective char layer that would reduce the effects of high temperatures [57]. The delamination of the layer from the CLT panel not only presents an additional fuel source for the fire but also results in a sudden loss of strength [21]. The main cause of delamination in CLT is the adhesive used to join the layers [75]. To understand this behaviour at high temperatures, Zelinka et al. [62] tested four different adhesives (PUR1, PUR2, MF, and PRF) in CLT specimens compared to solid wood of the same size. CLT specimens made with PRF (phenol resorcinol formaldehyde) adhesive showed similar results when tested in shear to those of solid timber. A significant loss in strength occurred at a temperature of 260 °C for both solid wood and adhesives, with the maximum reduction for the PUR1 adhesive. Hasburgh et al. [75] analysed the performance of different adhesives when used in CLT under exposure to fire. The CLT panels made with PRF and MF adhesives were found to be less prone to delamination compared to those made with PUR and EPI (emulsion polymer isocyanate) adhesives [75]. The effect of the adhesive used on the charring rate of CLT is shown in

Table 7. Effect of adhesive on the charring rate of CLT.

Reference	Layer Layup (mm)	Adhesive	Panel Size (m2)	Charring Rate (mm/min)
Muszyński et al. [63]	35-35-35-35-35	PUR	5.48 × 4.26	0.63
		MF		0.55
Frangi et al. [57]	10-10-10-10-20	MUF	1.15 × 0.95	0.58
		PU1		0.94
		PU2		0.90
		PU3		1.00
		PU4		1.08

.

The performance of adhesives in CLT under fire is critical, as bond-line integrity governs whether layers remain intact or delaminate during heating. Alternative joining methods, such as wooden dowel CLT (WDCLT), eliminate adhesive-related failure but introduce new vulnerabilities. A large-scale fire test on a 3 m × 3 m, 180 mm thick, six-layer WDCLT panel under a 50 kN/m in-plane load reported an average charring rate of about 0.9 mm/min [82], slightly higher than glued CLT due to the fall-off of the first two charred layers, which exposed fresh wood and caused two phases of accelerated charring. The panel sustained the applied load for 100 min [82], but increased deformability under fire was observed to raise the likelihood of charred layer detachment compared with glued CLT [82], highlighting potential differences in failure mechanisms between the two systems.

6.1.6. Fire Curve and Fuel Load

Another aspect considered for the experimental testing of CLT panels is the effect of the type of fire curve and heating rates on the fire performance. Friquin et al. [69] performed tests on Norwegian spruce CLT floor panels to investigate the effect of fire intensity on the charring rate of the CLT panels. The authors found that CLT panels exposed to fast temperature growth and high temperatures resulted in a higher charring rate, and vice versa. Bartlett et al. [14] studied the effect of heating rates on CLT samples and found that the charring rates vary significantly depending on the heating scenarios. This is due to the higher charring rate of 0.92 and 1.18 mm/min for CLT panels under exposure to natural fire conditions compared to 0.67 and 0.77mm/min for CLT panels under exposure to standard fire curves.

6.2. CLT Fire Compartment Tests

In contrast to the relatively predictable outcomes of standard furnace tests, as discussed above, compartment fire experiments on CLT structures reveal substantially greater variability in thermal response. Reported charring rates in compartments average around 1.22 mm/min with a large scatter (SD ≈ 0.69 mm/min) [25], whereas beams and walls tested under standard fire curves tend to remain much more consistent, typically 0.60–0.70 mm/min, with only minor deviations from the Eurocode 5 reference value of 0.65 mm/min [25]. This variability underscores the strong influence of compartment geometry, ventilation conditions, fuel load, and the extent of exposed timber surfaces on the fire performance of CLT. A large number of small-scale and full-scale compartment tests have been conducted recently to analyse the performance of CLT structures in fire. The main aim of fire compartment tests is to assess the structural integrity and behaviour of CLT structures, as well as to measure the effectiveness of various protection techniques in a real fire scenario. The fire in a compartment timber structure consists of different stages, as shown in Figure 12, each of which depends on several factors, such as the type, thickness, and protection of timber; fire duration and source; size of ventilation; applied load; compartment size; and area exposed to fire [83]. In mass timber structures such as CLT, the structural timber can enhance the fire intensity in the compartment [84] and must be carefully considered during the design stage. Unprotected mass timber compartments can have earlier flashover [85], an extended fully developed stage [86], and higher temperatures at the decay phase [87] due to multiple cycles of char fall-off [83]. The review of CLT fire compartment tests in this study has been categorised into two groups based on the fire protection strategies and size of ventilation in the following subsections.

6.2.1. Protective Layers

Different types of fire protection claddings have been analysed in compartment fire tests to investigate their effectiveness in fire resistance and to avoid charring the CLT elements. Kolaitis et al. [64] performed a compartment test with dimensions of 2.22 m × 2.22 m × 2.21 m to investigate different types of protective cladding on timber elements under realistic fire conditions for a 45 min duration. A fire load density of 420 MJ/m², equivalent to a typical office room, was achieved using 105 kg of wood. The protective claddings analysed include two layers of 12.5 mm gypsum plasterboard, two layers of 16 mm chipboard, and two layers of 16 mm MDF board. The fire resistance provided by gypsum plasterboard was higher than that of wood-based panels, with no charring observed in the CLT elements for the entire duration of the tests. However, the wood-based cladding only provided protection for around 35 min and failed, consequently, after that. Close agreement was found when comparing the experimental results with EN-1995-1-2 values in terms of the charring rate of timber elements and failure time of fire-protective claddings [64].

Li et al. [65] tested three CLT compartments, one of which was fully exposed to fire, while the rest were protected by two layers of 12.7 mm Type C fire-rated gypsum boards. Similar residential furniture sets were used in all tests, with fuel load densities ranging from 529 MJ/m² to 553 MJ/m². Compartment tests in which the CLT elements were protected did not show any charring until the end of the fire tests. However, in the compartment test in which the CLT surfaces were unprotected, a higher charring rate of 1.0 mm/min was measured at the end of the 1 h fire test, mainly due to the delamination of the fire-exposed layer. The unprotected CLT compartment experienced an earlier flashover and burned continuously until the entire fuel load was exhausted. The protected CLT compartments had a fire growth stage followed by a flashover and a fully developed phase, which ended in the decay of the fire. An 80%-higher heat release rate (HRR) was produced in unprotected CLT compartments compared to unexposed compartments, of which 64% was from external burning.

Emberly et al. [88] tested a 3.5 m × 3.5 m × 2.7 m CLT compartment in which the internal faces were protected with two layers of 13 mm fire-resistant plasterboard, except for a wall and ceiling, which were unprotected. All CLT elements in the compartment were of five layers with a layup of 45-20-20-20-45 mm. The compartment consisted of wood cribs as a fuel load with a total mass of 80kg to ensure flashover and at least 10 min of fully developed fire phase after ignition of CLT, as well as to ensure that the fire did not last for more than half an hour after ignition. No failure or debonding of either the fire protection or the CLT layer was observed. The test showed that self-extinction in a compartment with an exposed CLT wall and ceiling can be achieved if there is no debonding and the heat flux does not exceed 45 kW/m².

Su et al. [86] tested five CLT compartments with a size of 4.572 m × 2.438 m × 2.743 m and with a door opening of 0.76 m × 2.0 m. Five-layered CLT panels glued using thermally resistant PU adhesives with a thickness of 175 mm were used, which were protected with different cladding configurations in each compartment’s tests. The two different cladding configurations protect either the CLT surfaces with three layers of Type X board (1 × 15.9 mm and 2 × 12.7 mm) or those with two layers of 12.7 mm Type X board, which is a fire-rated protective layer offering up to one hour of fire resistance [89]. The compartments also contained glulam beams and columns exposed to fire. A fuel load of 550 MJ/m² was achieved in each compartment test, which came from the wood cribs. The test results showed that the adhesive was able to resist char fall-off even when the temperature beyond the bond line reached the char temperature. The initial charring rate of the unprotected CLT panels was approximately 0.6–1.0 mm/min, which decreased after the increase in the char depth. However, for protected CLT panels, the charring rate was lower than 0.20 mm/min, taken from the time when the CLT panel started charring behind the protective claddings.

In another study performed by Hopkin et al. [90], a full-scale experimental setup representative of typical residential rooms was employed, with an internal floor plan of 3.4 m × 3.4 m and a ceiling height of 2.5 m. The front wall was made of a concrete block construction, which was extended to a height of approximately 3.9 m, incorporating a standard-sized door opening of 0.7 m × 1.8 m. The enclosure surfaces, including walls and ceilings, were composed of 180 mm thick CLT panels, arranged in 40-30-40-30-40 mm lamellae and joined via half-lap joints. Various protective lining configurations (see

Table 8. Summary of CLT fire experiments (from Hopkin et al. [90]).

Experiment ID	Lining Configuration	Initial Exposed CLT Area	Experiment ID	Lining Configuration	Initial Exposed CLT Area
1a and 1b	Triple-layer lined wall and ceiling	0% (fully encapsulated)	6a	2 × walls triple-lined, ceiling and 1 × wall exposed	46%
2a	Triple-lined walls, Single-lined ceiling	0%	7a	Ceiling and 1 × wall triple-lined, 2 × walls exposed	38%
3a and 3b	Triple-lined walls, ceiling exposed	26%	8a	1 × wall triple-lined, Single-lined ceiling, 2 × walls exposed	38%
4a	2 × walls triple-lined, 1 × wall single-lined, ceiling exposed	26%	9a	Triple-layer on 2 × walls and ceiling, 1 × wall exposed	19%
5a	2 × walls triple-lined, 1 × wall double-lined, ceiling exposed	26%

), from triple-layer encapsulation (K₂90) to single-layer (K₂30), were adopted, with each lining consistently formed from an 18 mm thick gypsum product. Fire was initiated using an array of six propane burners configured in two rows on the compartment floor, calibrated to deliver a heat release rate (HRR) of around 1800 kW to reliably produce external flaming and simulate a fast t² fire growth scenario. The fire followed a predefined growth-and-steady phase lasting 60 min, succeeded by a decay phase, aligning approximately with a 90 min ISO 834 standard fire curve, commonly used for structural fire resistance assessments.

This study demonstrates that the extent and nature of protective cladding on CLT significantly influence fire behaviour and the potential for self-extinguishment. While full encapsulation with durable, multi-layer gypsum lining effectively prevents CLT involvement and reduces charring, partial protection often fails, especially when lining detaches or permits pyrolysis behind it, which leads to sustained flaming combustion and prevents self-extinguishment. Charring rates do not strongly correlate with enclosure configuration, though ceilings char slightly slower (~0.51 mm/min) due to less direct flame exposure than walls (~0.65 mm/min), and exposed CLT markedly amplifies the total heat release rate, raising HRR well above the 1800 kW baseline of burner-only tests.

In experiment 8a, where the ceiling was lined with a single layer of plasterboard, the ceiling lining failed mid-test, which resulted in a sudden rise in charring rate of up to 1.50 mm/min for the ceiling surface. This underscores that protective linings not only serve to delay ignition and reduce timber involvement but also dramatically influence the progression of charring if compromised. No clear trend was observed in charring rates with the variations in enclosure configuration. However, the exposure of more CLT surfaces increased the overall heat release rate and led to higher heat fluxes. Across all configurations, the experiments reaffirmed the established a self-extinguishment threshold of 44.5 ± 1.2 kW/m² (see Figure 13) when the incident heat flux during decay dropped below this level, enabling flaming to cease, consistent with bench-scale studies [91].

In a recent study, Bøe et al. [92] investigated sprinkler protection in large-scale fire experiments on CLT compartments with adjacent corridors. The results demonstrated both the critical role and the limitations of sprinkler systems in fire protection. When the sprinkler system was fully functional, fires were suppressed effectively; however, when compartment sprinklers were disabled and only corridor nozzles operated, flashover occurred within 5 min, and temperatures remained at around 1000 °C for the remainder of the test. Exposed CLT surfaces continued to radiate heat, and delamination of shallow CLT layers contributed to sustained burning. Notably, the charring rate exceeded 1.1 mm/min in these scenarios, which is far above typical design values, highlighting that partial fire protection and inadequate sprinkler coverage can accelerate CLT degradation in compartments.

6.2.2. Size of Opening/Ventilation

Ventilation size in compartments is one of the significant factors that control the temperature and fire behaviour of CLT structures. The size of ventilation in compartments is measured by a ventilation factor using the relationship denoted in Equation (10) from EN-1991-1-2 [93].

$$\mathrm{O}={\displaystyle \frac{A_V\sqrt{h_{eq}}}{A_t}}$$

(10)

where O is the opening/ventilation factor; h_eq is the weighted average height; A_t is the total surface area of the compartment, including the opening; and A_v is the area of the opening.

In recent years, several experimental studies have been conducted to measure the impact of the size of ventilation on CLT compartment fires. Mindeguia et al. [27] performed three experimental tests on CLT slabs in a compartment under exposure to natural fires. Three different opening configurations were analysed, namely Scenarios 1, 2, and 3, with corresponding ventilation factors of 0.144, 0.050, and 0.032 m^1/2, as shown in Figure 14. Five-layered CLT panels with an overall thickness of 165 mm were manufactured from C24-grade spruce boards bound with PUR adhesive. The CLT slabs were 5.9 m long and 3.9 m wide and were tested in a compartment with dimensions of 6 m × 4 m × 2.52 m. All CLT slabs were both mechanically and fire-loaded with a magnitude of 29.5 kN and 891 MJ/m², respectively. The authors did not observe any significant char fall-off of the charred layers, although localised delamination of the outermost lamellae occurred in all the tests. In Scenario 1, delamination of the outermost lamellae was less obvious and occurred at approximately 40–60 min. Because of the earlier decay and faster development of fire in Scenario 1, the effect of the loss of the outer lamellae on the thermo-mechanical performance was minimised. Delamination of the outer lamellae occurred between 45 and 65 min in Scenario 2, whereas the outer lamellae delaminated between 55 and 75 min in Scenario 3. The charring rates in Scenario 1 were higher than those in Scenarios 2 and 3, as shown in Figure 15, because of the higher compartment temperature in Scenario 1, which was the earliest to decay. During the heating phase, failure was observed only in the CLT slab tested under Scenario 3 at 108 min. The failure in the slab tested under Scenario 2 occurred 29 h after the test initiation due to smouldering. No failure was observed in the CLT slab tested under Scenario 1. The charring model in Eurocode 5 was unable to estimate the charring of the slabs discussed in this study.

Su et al. [94] performed six-compartment fire tests with dimensions of 9.1 m × 4.6 m × 2.7 m, made up of five-layered CLT with an overall thickness of 175 mm. The CLT panels were made of spruce–pine–fir boards, which were glued using a PUR adhesive. Two opening configurations were used in the tests. The size of the opening used was 1.8 m × 2.0 m in four tests, while the rest of the compartment tests had an opening of 3.6 m × 2.0 m. The compartments were tested using a fire load of 550 MJ/m². A structural load of 0.95 kN/m² was applied to the ceiling top, calculated as half of the design load for residential structures. The compartments were either fully or partially protected with two or three layers of Type X gypsum plasterboard with a thickness of 15.9 mm, with the details shown in

Table 9. Applied fire protection layers on different parts of the CLT compartments [94].

Compartment Surface	Test Number
	1-1	1-4	1-5	1-6	1-2	1-3
	Size of Opening (1.8 m × 2.0 m)			Size of Opening (3.6 m × 2.0 m)
	Number of Protective Layers of Gypsum Plasterboard (Type X)
W1 (9.1 m × 2.7 m)	3	3	Exposed	Exposed	2	Exposed
W2 (4.6 m × 2.7 m)	3	3	3	3	2	2
W3 (9.1 m × 2.7 m)	3	3	3	3	2	2
W4 (4.6 m × 2.7 m)	3	3	3	3	2	2
Ceiling (9.1 m × 4.6 m)	3	Exposed	3	Exposed	2	2

. The test results showed that the size of the ventilation in the CLT compartment had a significant impact on fire development. The smaller opening in the compartment had an extended fire duration, as shown in Figure 16. However, larger ventilation or openings accelerate combustion with a higher peak HRR for a shorter duration. In compartments with fully protective claddings, charring of CLT was effectively prevented or delayed. The fully protective claddings in Test 1-1 completely prevented the charring of the CLT structure; however, only a char depth of less than 15 mm was measured in the ceiling in Test 1-2. Partial protection of the CLT structure contributes to compartment fires, which are dependent on the area and orientation of the exposed region and the size of the opening in the compartment. For compartments with the same area of exposed CLT, i.e., Tests 1-3 and 1-5, the HRR was higher in Test 1-3 due to larger ventilation. However, in Test 1-5, there was a second flashover, which occurred due to the delamination of layer-2 of the CLT caused by the containment of heat within the compartment due to a smaller opening.

Kotsovinos et al. [95,96] performed the largest compartment fire tests to date with a floor area of 352 m² through a project known as CodeRed. The compartments in this project were tested using a fuel load density of 374–377 MJ/m², which is considered equivalent to the fuel load density of an office building. The compartment had an overall opening of 20.5% of the total surface area of the walls in CodeRed #01 test, whereas the opening was reduced to 10% of the total surface area of the walls in CodeRed #02 test. The CLT ceiling panels were directly exposed to fire, made of spruce wood with a layup of 40-20-20-20-40 mm, and bonded using MUF adhesive. The compartment test results showed that by reducing the ventilation area by half, the duration of the flaming combustion was prolonged by 20%. Furthermore, it was also observed that the size of the opening in the compartment had a significant impact on the spreading of the flame in the CLT, with a 23% decrease in the spread rate caused by reducing the ventilation area by half. Regardless of the ventilation area in the compartment, the maximum temperatures and heat fluxes were similar. However, the time to reach the maximum temperature was 20 min for compartments with smaller ventilation areas compared to 5–10 min for compartments with large openings. The rate of decay of temperatures after the end of flaming in the CLT compartments was influenced by the size of the openings, with the decay temperature rate of 9.7 °C/min for smaller openings compared to 11.2 °C/min for the compartment with large ventilation. When comparing the char depths of ceilings in both tests, it was found that an overall increase of 11% in char depth was measured in compartments with smaller ventilation (CodeRed #02) compared to compartments with larger ventilation (CodeRed #01), as shown in Figure 17.

Compartment fire studies often test CLT elements at relatively low load utilisations, while real buildings are subjected to loads up to three to four times higher [97]. This mismatch may overestimate adhesive performance and fire resistance, particularly regarding delamination and char fall-off [97]. Although some adhesives (e.g., MUF, HB X) improve fire performance and self-extinction [97], more research under realistic design loads is needed to ensure reliable safety.

6.3. Discussion on the Limitations of Experimental Research

Experimental investigations of the fire performance of CLT generally fall into two primary categories: standard fire resistance tests and compartment fire experiments. Fire resistance tests, often performed under ISO 834 [34] or ASTM E119 [54] conditions, provide a controlled and repeatable environment to measure charring rates, the onset of delamination, and load-bearing capacity. Studies such as those by Xing et al. [31] on two-way CLT slabs and Klippel et al. [67] on loaded CLT wall and floor panels have enhanced our understanding of how load ratios, support conditions, and adhesives affect performance. However, many laboratory studies employ small- or medium-scale specimens under simplified or steady-state heating regimes, which, similar to the standard fire curve, lack a cooling phase and do not accurately reflect the variable ventilation or decay characteristics of real building fires. Furthermore, the number of samples is usually limited (one to three replicates), which reduces the statistical confidence in the reported charring rates and failure times. Additionally, boundary conditions and timber species differ significantly between laboratories, which makes the comparison challenging.

On the other hand, large-scale compartment fire experiments (e.g., Hopkin et al. [90], Bøe et al. [92], and Su et al. [94]) offer more realistic insights into whole-room fire dynamics, ventilation effects, and delamination-driven fire growth. These tests reveal that exposed CLT surfaces and protective cladding configurations greatly impact fire spread and compartment temperature development. For instance, Hopkin et al. demonstrated that the total HRR increases as more CLT surfaces are exposed, while Bøe et al. found that partial sprinkler protection could not prevent sustained fire growth once delamination occurred. Nonetheless, compartment tests are costly, rarely replicated, and often lack standardisation in terms of ventilation geometry, ignition source, and exposure area, making generalisation difficult.

In conclusion, while fire resistance tests are essential for design verification, compartment tests are crucial for understanding system-level behaviour. Nonetheless, both types of tests face challenges with limited repeatability and a lack of standardised reporting. Future research should focus on bridging these scales by developing unified experimental protocols, conducting statistically robust test series, and integrating experimental data with validated numerical or machine learning models to ensure consistent and reliable predictions of CLT fire performance.

7. Numerical Studies

The fire performance of CLT has been extensively explored using numerical tools to predict its behaviour under high temperatures. However, developing a full-scale model using a finite element tool requires extensive computational power and is a time-consuming process. To address this, Bai et al. [98] developed a numerical approach to investigate the behaviour of scale models under exposure to fire and to validate the thermo-mechanical similarity criteria. To incorporate adhesive elements into the numerical model, both Bai et al. [98] and Wang et al. [58] used a 3D eight-node cohesive element (COH3D8) for analysis with a thickness of 0.001 mm. A tie constraint with a hard contact was chosen between the CLT layer and the adhesive [98]. However, to reduce computational time and simplify the numerical calculation, most researchers have excluded the adhesive from their FE analysis [59] due to their thinner thickness compared to the CLT layer. Further modelling constraints are discussed in the following sections.

7.1. Modelling of Heat Transfer in CLT

The study of heat transfer in timber typically involves three modes: conduction, convection, and radiation [17]. In a fire test furnace, heat fluxes flow to the wood’s outer surfaces via convection and radiation, while conduction occurs within the wood member [17]. To develop a heat transfer model for predicting the structural behaviour of timber under fire, Thi et al. [17] created a subroutine called UMATHT in Abaqus. By incorporating a subroutine in Abaqus, the user-defined material models were able to account for as many physical and chemical phenomena as possible to best predict the thermal behaviour of timber during pyrolysis.

In other studies, Schmid et al. [99] and Wong et al. [100] utilised the SAFIR 2007 software [101] to perform a one-dimensional thermal analysis, incorporating the defined values of convection coefficients on the unexposed and exposed faces of CLT, as well as the surface emissivity as specified in EN 1991-1-2 [93]. Following this approach, an ISO-834 temperature curve [34] was applied to the fire-exposed face of CLT using the interactions of surface film conditions and surface radiation [58,100]. Additionally, the thermal model incorporates temperature-dependent values for density, specific heat, and thermal conductivity, as defined in EN-1995-1-2 [16].

Machine learning (ML) has recently been explored as a tool for predicting heat transfer in CLT during fire exposure. In one study, temperature data from loaded CLT wall strip tests were used to train three ML models: long short-term memory (LSTM) neural networks, gradient boosting (GB), and symbolic regression (SR) [102]. Symbolic regression, an interpretable ML method, significantly outperformed both LSTM and GB for in-solid time–temperature prediction, while also providing explicit functional expressions [102]. This highlights the potential of ML techniques as an efficient complement to traditional heat transfer models in CLT fire research.

7.2. Mesh Size

Meshing is one of the key criteria for performing FE analysis that influences the accuracy of the results. The size of the mesh plays an important role in FE analysis, as it affects not only the computational cost and time but also the convergence of the solution. A study conducted by Werther et al. [103] employed different mesh sizes of 1 mm, 3 mm, and 6 mm in the heat transfer analysis. The results indicated that a mesh size of 3 mm was comparable to that of 1mm, and hence, they recommended using the former to optimise computational time and accuracy. Thus, the mesh sensitivity analysis is a significant consideration that has an impact on the accuracy of the results obtained from the FE model, and it should always be carried out for all FE analyses.

Table 10. Type and size of the mesh.

Model Type	Authors	Element Type *	Mesh Size (mm3)
Thermo-mechanical model	Bai et al. [59]	C3D8R	10 × 10 × 7
	Wang et al. [58]
Heat transfer model	Xing et al. [35]	DC3D8	10 × 10 × 1
	Wang et al. [58]		5 × 5 × 3.5

* C3D8R: 3D eight-node hexahedral linear reduced integral elements, * DC3D8: 3D eight-node linear heat transfer element

presents the element type and mesh size utilised in Abaqus, which is a widely employed tool for the FE analysis of CLT elements.

7.3. Deactivation of the Charred Layer

The modelling of the fall-off of the charred layer is an important aspect of CLT under fire that has been investigated by various researchers, who have incorporated it in their numerical modelling in different ways. Xing et al. [35], for example, utilised a “model change” approach to simulate the delamination of the charred layer and the rapid increase in the internal temperature of the wood. This approach comprised a series of steps, including the undertaking of a thermal analysis to assess the failure time of the exposed layer to fire. The failure time was then utilised in a thermal analysis using the model change technique to cancel the effect of the outer layer. Lastly, the corresponding temperature was applied to the surface of the second layer to account for the swift temperature rise.

Similarly, Wang et al. [58] developed a 3D FE model using the Hashin criteria and cohesive elements to examine the failure of wood and adhesive, respectively. The “model change” technique was employed in Abaqus to deactivate the charred layer, which causes the adjacent CLT layer to be directly exposed to fire after a specific time period based on the experimental testing. The model adopted in this study accurately predicted the temperature distribution after the charred layer fall-off compared to no fall-off, as depicted in Figure 17. However, the model was unable to simulate the temperature rise in the event of a local char fall-off.

Wong et al. [100] defined two different assumptions for the time period to simulate the delaminated layer in the FE model of the CLT panel. The first assumption simulated the delaminated layer when the temperature between the layers reached 300 °C, while the second assumption simulated the fall-off of the charred layer when the temperature of a subsequent layer at a depth of 10 mm from the exposed surface reached 300 °C. The latter assumption gave more realistic and accurate results in terms of the charring rate and temperature distribution when compared to the experimental results (See Figure 17). Martinez et al. [104] employed the ANSYS Parametric Design Language (APDL) in ANSYS® 18.1 to simulate the delamination of CLT layers when the temperature between the two layers reached the charred temperature, i.e., 300 °C. In this method, the technique of Birth and Death elements was utilised, wherein the delaminated layer, after reaching its charring temperature, remains in the model and is deactivated rather than removed. Yasir et al. [72] adopted a simplified procedure in their two-dimensional FE modelling of CLT floor panels when exposed to standard fire. Instead of char fall-off, the modelling was performed by considering a higher thermal conductivity value than specified in the EC-5 for temperatures above 800 °C. They found close agreement in temperature distribution and charring rates between the FE model and experimental results when using this procedure.

A comparison of experimental and FE temperature distribution is illustrated in Figure 18 for the studies presented above. The studies presented suggest that the numerical modelling methods discussed above, which incorporate different procedures to account for delamination effects, are effective in predicting the experimentally observed results. It was also shown that the temperature distribution at various depths can become increasingly difficult to predict at temperatures beyond 300 °C, as depicted in Figure 18, and further research is required.

7.4. Fire Resistance of CLT

Bai et al. [59] developed a numerical approach in Abaqus using a user-defined subroutine to model timber’s constitutive relationship and temperature-dependent mechanical properties to predict the residual load-carrying capacity of CLT walls after fire. Through axial compression and standard fire simulations, the study accurately depicted the stress behaviours, failure mechanisms, and thermal responses of three-ply and five-ply CLT walls under load, post-exposure to fire. The findings highlight significant variations in the load–displacement curves of CLT walls depending on their ply configurations and the nature of stress (compressive or tensile) leading to failure, thereby underscoring the nuanced impact of fire on CLT walls’ structural integrity and guiding future designs for improved resilience.

Xing et al. [105] developed a material constitutive model using OpenSees framework (compiled in Visual Studio 2019) to analyse the heat transfer and fire resistance performance of CLT, incorporating the anisotropic properties of wood. The model developed effectively simulates CLT performance under standard fire exposure, assessing its temperature field variation with time in the components, and the changes in mechanical properties of wood at ambient temperatures and during and after exposure to high temperatures. The simulation predicted a slightly higher residual carrying capacity after fire, which fell within the acceptable range. This variance was attributed to wood imperfections, which were not incorporated into the FE model. Notably, OpenSees offers a significant reduction in computational resources over Abaqus, demonstrating efficiencies in static and fire-induced simulations by factors of 30 and 1500, respectively.

In another study performed by Xing et al. [31], a refined FE model was developed, incorporating thermal–structural coupling to simulate both heat transfer and mechanical degradation in CLT two-way slabs. The model predicted temperature distributions through the slab thickness and mid-span vertical deflections under fire exposure. Validation against full-scale fire test data showed very close agreement in both temperature profiles and displacement–time histories, underlining the model’s reliability as a predictive tool for assessing CLT slab performance in fire design practice.

While finite element models (e.g., Xing et al. [31]) are commonly used to simulate the thermo-mechanical response of CLT under fire and are typically validated against full-scale test data, some recent studies, as performed by Perkovic et al. [46], have applied CFD to investigate fire resistance. CFD offers the advantage of capturing fire growth, heat flux, and ventilation effects, and several studies have shown reasonable agreement with experimental trends in temperature rise and charring depth. However, validation is often limited to smaller-scale data, and CFD models are still less frequently benchmarked against full-scale CLT structural fire tests compared to FEM approaches.

Table 11. Comparative evaluation of modelling approaches for CLT fire behaviour.

Approach	Primary Predictions	Strengths	Limitations/Challenges	Key References
FE	Temperature profiles, charring depth, deformation, failure time	Accurately couples thermal and mechanical responses; reliable for member and assembly scale fire design	Computationally expensive, sensitive to input (adhesive degradation, moisture), less suited for fully coupled compartment simulations	[58,59]
CFD	Fire-induced flow, heat flux distribution, flame spread, smoke layer development in compartments	Captures compartment dynamics and fire growth; supports ventilation and cladding studies	Does not resolve structural response, high model uncertainty, validation data scarce, high computational cost	[46,106]
ML	Time–temperature prediction, charring rate	Very fast once trained; good for sensitivity and real-time forecasting	Needs high-quality training data, lacks physical interpretability, extrapolation outside training range uncertain	[102]
PISM	Temperature distribution, heat transfer, convection, fire dynamics, structural fire engineering, material behaviour	Heat transfer, physically interpretable, emerging research direction	Convergence difficulties, uncertainty quantification, data efficiency, multi-physics integration	[107,108]

presents a comparative analysis of key modelling approaches employed in assessing the fire performance of CLT, detailing their predictive focus, primary strengths, and current limitations. This analysis highlights where each method performs best and identifies areas requiring future integration or development.

Overall, FE modelling remains the benchmark for thermo-mechanical fire analysis of CLT at the element level, while CFD is most effective for compartment-scale fire and ventilation studies. ML approaches provide rapid predictions but require careful data handling and physical validation. The future of CLT fire modelling lies in hybrid frameworks, such as PISM (Physics-Informed Surrogate Modelling), that integrate the physical accuracy of FE or CFD with the efficiency of ML, thereby facilitating reliable and computationally efficient fire-resistance assessments.

8. Conclusions

This article provides a comprehensive overview of the behaviour and structural integrity of CLT under exposure to fire. The review synthesises recent advancements in understanding the fire performance of CLT from experimental, numerical, and regulatory perspectives. At the design level, Eurocode 5 remains the primary reference framework, though its simplified one-dimensional charring rate method and generic assumptions have limitations when applied to different species, adhesives, or layups. Comparative assessment shows that other international standards, including the NDS (USA), CSA O86 (Canada), and AS 1720.4 (Australia/NZ) codes, adopt broadly similar mechanics-based or reduced cross-section approaches but differ in their assumptions regarding charring rates and treatment of delamination. For CLT, this highlights the need for harmonised approaches and species-specific calibration to ensure consistent fire design guidance across global practice.

This article also includes results from various experimental fire resistance and compartment tests, along with FE modelling studies on CLT elements conducted over the past two decades. The article highlights the critical influence of adhesive types, panel construction, and exposure conditions on the fire performance of CLT, emphasising the need for detailed considerations in design and material selection. The main findings from the review of the fire resistance tests are summarised below.

• Adhesives play a crucial role in the behaviour of CLT elements under fire exposure. Non-heat-resistant adhesives like PUR are typically shown to lead to more delamination or fall-off of charred layers compared to MF- or PRF-based adhesive panels.
• CLT panels with thicker outer layers exhibit better fire performance than those with thinner outer layers because a charred layer develops without fall-off, protecting the inner layer from direct fire exposure.
• There was found to be no significant influence of the orientation of the CLT layers or support conditions and grade on the fire behaviour or charring rate of CLT panels.
• Fire intensity and duration affect the performance of CLT and should be considered, as a higher charring rate has been observed for natural fires compared to standard fires.

This article also focused on compartment fires, noting that the current Eurocode standards do not adequately reflect real fires or compartment fires. Large-scale compartment studies also emphasise the role of openings, ventilation, and active measures such as sprinklers, showing that inadequate protection or partial coverage can rapidly lead to flashover and sustained high-temperature conditions. An analysis of the CLT compartment experiments discussed in this article yielded several significant findings:

• CLT compartments having different protective claddings showed their effectiveness in delaying the charring of the CLT structure, as well as reducing the initial char rate when charring behind the claddings initiates. Furthermore, the fire protection in CLT compartments has a significant effect on the HRR, with a substantial decrease in the HRR when protective cladding is used.
• The size of the opening or ventilation is another key factor that controls the temperature, fire development, and behaviour of the CLT structure. Compartments with more openings have higher charring rates than those with smaller openings. This is because compartments with larger ventilation accelerate combustion with a higher peak HRR for shorter durations. In contrast, compartments with smaller ventilation extend the fire duration, causing an overall higher char depth than compartments with large ventilation.

This review is further extended to the numerical analysis of CLT panels under fire using FE analysis, with a focus on model development, modelling of heat transfer analysis, mesh size, and the inclusion of delamination of the char layer. Simulation of the delamination of the charred layer is an important research focus that has been discussed and studied recently in the literature and found to be effective in predicting the temperature distribution at different depths of the CLT panel. However, the current literature does not contain predictions or models of the local fall-off of the char in the CLT panel, and further research is required.

Future Research and Development

The authors recommend the development of standards for CLT under fire exposure conditions to ensure safety. Significant research has been conducted to investigate the fire performance of CLT panels when used as floor and wall elements under standard and natural fire conditions. However, CLT-CLT connections, which typically incorporate metallic components (self-tapping screws, angle brackets, etc.), are more vulnerable to fire and can impact the behaviour of CLT structures, and they have not been comprehensively reported in the literature. Most of the existing literature comprises small-scale testing, which, while informative, may not accurately reflect the behaviour of CLT structures in mid-to-high-rise buildings during a fire. Large-scale testing is, therefore, essential to bridging this knowledge gap and developing reliable fire design guidelines. Furthermore, an approach to numerically model CLT structures has been recommended utilising the data from the literature for its validation and verification, with a focus on modelling CLT structures under natural fire.