111Equation Chapter 1 Section 1Performance of laminated

glazing under fire conditions

Yu Wang, Jiayu Hu

School of Engineering, University of Edinburgh, Edinburgh EH9 3JL, United

Kingdom

Abstract

Glass breakage and fallout may change the ventilation of compartment which would

significantly accelerate the enclosure fire development. Past research of glass in fires

has focused on single and insulated glazing, whereas the thermal performance of

laminated glass remains less well understood which is increasingly used in high-rise

building façade systems. In this work, the experimental data has been used to validate

the heat transfer and thermo-mechanical models performed by ABAQUS. The

experiment refers to two laminated glass panels with 600×600 mm2 area and 12.38

mm thickness (6 mm glass + 0.38 mm polyvinyl butyral + 6 mm glass), which were

heated to break by a 500×500 mm2 square pool fire. The heat transfer models were

verified by the recorded surface temperature at the ambient side. The breakage time of

both panels were predicted through the stress distribution, showing relatively good

agreement with experimental observations. After the model was validated, a

parametric study of laminated glazing, including the thickness of gel layer and the

number of glass layers were changed to investigate the significance of these

parameters to the fire resistance of laminated glazing. It was established that the 1.52

mm interlayer and 4-glass-panel laminated glazing demonstrate the best fire

resistance. However, the 0.38 mm and 3-glass-panel ones are recommended

Corresponding author: Jiayu.Hu@ed.ac.uk (J. Hu); First author: Yu.Wang@ed.ac.uk (Y. Wang)

1

considering the construction cost and reasonable fire performance. This numerical

model is proved to be capable of the thermal performance prediction of laminated

glazing in fire safety design of glass façades.

Keywords: laminated glazing; heat transfer; glass in fire; finite element method;

parametric analysis

1. Introduction

As the weakest section of the building envelope, glass façades may break and fall out

very easily when subject to a fire. The new vent created by the glass failure would

induce fresh air entrance and ejected window flame, resulting in the acceleration of

the compartment fire development and ignition of the combustible cladding. Thus, the

thermal behaviour of glass façade is of great importance to the fire spread of the high-

rise buildings. Emmons [1] first highlighted the great role of glass thermal breakage

behaviour in a fire, and subsequently a large number of studies were performed to

investigate the mechanism of thermal cracking in glazing [2, 3]. For example, Keski-

Rahkonen [4] used linearized radiation cooling boundary condition to theoretically

predict critical temperature difference of 80 °C. Harada et al. [5] tested the float and

wired glass with different constraint conditions by a propane radiation panel. Shields

[6, 7] performed full-scale experiments in ISO 9705 room to investigate the behaviour

difference of single glazing with different fire locations. Pagni et al. [8] and Wang et

al. [9] respectively developed softwares BREAK1 and EASY to predict the breakage

time of single glazing in a compartment fire. It was established from these work that

the exceeding internal thermal stress caused by thermal gradient is substantially the

reason for glass breakage in the fire.

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All the above work primarily focused on the single clear glazing. However, in recent

years, with the overwhelming use of glass in the modern building envelope, different

kinds of glasses, such as coated, ground, insulated and laminated glazing (LG), are

becoming increasingly popular due to their good performance in thermal, energy,

light and aesthetics [10]. The diversity of glass beautifies the high-rise building

outlook but brings the new challenge for the fire performance-based design due to the

knowledge paucity concerning fire performances of these different glasses [11]. In

particular, to date, there have no specific design regulations for vulnerable glazing

façades which makes the situation even worse [12]. The glass type has a significant

influence on the thermal breakage of glass [10], thus it is of great importance to study

the fire performance of new glazing.

Among the different types of glass, laminated glazing, which is built of two or more

annealed or tempered glass panels combined with one or more polyvinyl butyral

(PVB) interlayers, is not only used as glass façade/windows but also the structural

elements in modern transparent constructions [13]. Extensive work has been

conducted on its structural or blast performance [14, 15], but very little is known

about its behaviour under real fire conditions. In particular, although recent work

highlighted LG has the best ability to keep integrity during a fire compared with clear,

coated and insulated glazing, the knowledge about LG behaviour under thermal

loading is only limited to the heat transfer process [16, 17]. The thermal breakage

mechanism of LG still needs to be understood further.

In this work, two full-scale experimental results will be introduced in detail based on

the authors’ previous preliminary tests [16]. The robust Finite Element (FE) models

were created in ABAQUS 6.12 to investigate the temperature variation, stress

distribution and breakage time of LG. After the validation of the heat transfer and

3

mechanical models by the experimental results, a parametric study, concerning the

thickness of gel layer and the number of glass layers, was conducted. The results of

this study are proposed to provide an effective evaluation of LG thermal breakage

behaviour under fire conditions. The specific analysis is demonstrated in the

following sections.

2. Brief description of experiments

The two laminated glass tests can be found in the authors’ previous work [16]. The

measured temperature and breakage time are used to validate the FE models. In this

work, the key information is given below.

2.1 Descriptions

As shown in Fig. 1(a), a LG panel was placed 20 cm above the ground and 750 mm

away from a 500×500 mm2 square pool fire. The LG panel consisted of two

600×600×6 mm3 float clear glass panels attached together by a 0.38 mm interlayer of

PVB film, generating a total thickness of 12.38 mm (Fig. 1(b)). The glass used in the

experiments is soda-lime glass consist of 73% SiO2, 9% CaO and 14% Na2O. A

digital camera was placed on the ambient side of the glass to record the breakage time

and path.

Two tests were conducted using the fuel consisting of 99% mass fraction N-heptane

with different fuel masses of 2 kg and 4 kg for Test 1 and Test 2 respectively. Test 1

and Test 2 burnt for 274 s and 492 s respectively. To present the glass and surface

clearly, for each LG, the glass panel on the fire side is named as Pane 1 and the

ambient side glass panel is Pane 2. The four surfaces of the laminated glass panel are

named as S1, S2, S3 and S4 from the fire side to the ambient side, as shown in Fig.

1(b).

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(a) The arrangement of experimental setup

(b) The names of each pane and

surface

(c) The distribution of measurement

instruments (viewed from fire side)

Fig. 1. The experimental schematic [16].

A well-designed frame made of stainless steel was employed for glass support. The

width of the covered region at the glass edge was 20 mm. In the thickness direction,

the glass pane was clamped using several thin strips, and the clamping pressure could

be controlled by revolving screws. This design ensured the glass panes to be

appropriately constrained in the x, y and z directions, so that the boundary condition

of an actual double glazing unit could be more closely approximated. It should be

noted that the thermal expansion or deflection of stainless steel frame would not be

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significant under such experimental condition [18]. This could ensure the glass pane

was free to expand against the framing so that mechanical effect of framing on

glazing was very limited.

A data acquisition system with 16 channels for thermocouples was used, with the

sampling time adjusted to 1 s. A Gardon water-cooled total heat flux gauge with a

measurement range of 0-50 kW/m2 was employed to measure the incident heat flux on

the glass. The gauges were fixed off the side of each glass pane and mounted flush to

the surface of the glass sections, so as to situate them as close to the measurement

location as possible. As it is impossible to drill into the glass sections to mount the

gauges, this method is considered desirable for heat flux measurement of glass pane

[6, 7, 19].

Sheet K-type thermocouples were attached to the glass panes using high-temperature

resistant adhesive and were numbered TC1-TC10 (Fig. 1(c)). Only TC10 was

attached to the ambient side surface (S4), while the other nine thermocouples were on

the fire side surface (S1). The thermocouples were made by a professional local

manufacturer, with a measurement range of 0-1200 ºC and sensitivity of 41 µV/ºC.

Due to the influence of smoke/fire radiation, the uncertainty of temperature

measurement is evaluated at ±5% [9].

2.2 Experimental results

Incident heat flux, heat release rate (HRR), glass surface temperature, gas temperature

and breakage time were recorded. In this work, for the verification of numerical

model, only glass surface temperatures and breakage times are presented. The

experimental condition and results of the two tests are summarized in Table 1.

Table 1. The summary of experimental conditions and results [16].

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Fuel mass

(kg)

Time of first crack

occurrence (s)

Temperature difference

(°C)

Incident heat flux

(kW/m2)

Pane 1 Pane 2 Pane 1 Pane 2 Pane 1 Pane 2

Test 1 2 118 199 96 -- 10.98 14.31

Test 2 4 258 332 79 -- 9.93 13.43

The glass surface temperatures measured in experiments are shown in Fig. 2. It can be

seen that in both tests the temperatures measured at different points in exposed areas,

i.e. T2, T4, T5, T6 and T8, are very similar before cracking; this also occurs in the

covered area, i.e. T1, T3, T7 and T9. Thus, in the numerical simulation, the thermal

loading is assumed uniform and have been averaged in the exposed and covered areas

(Eq. 1 and Eq. 2), which has been proved reasonable under this pool fire condition

[9].

22\* MERGEFORMAT ()

33\* MERGEFORMAT ()

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Fig. 2. The temperatures of Tests 1 [16] and 2.

It should be noted that in Tests 1 and Test 2, the fuel masses are respectively 2 kg and

4 kg. When the fuel mass increased, the temperature increase rate decreased, which

was caused by the combustion characteristic of pool fire: the thicker fuel needs more

time to reach a high HRR after ignition [20]. The incident heat flux of the two tests

are illustrated in Fig. 3, which can confirm the pool fire characteristic. Thus, glass

panel took less time to break in Test 1 than that of Test 2. However, the different two

tests can enhance the validation of the numerical model.

Fig. 3. The total heat flux of Tests 1 and 2.

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3. Numerical simulation

For model verification, two sets of three-dimensional (3D) uncoupled heat transfer

and mechanical models were created in ABAQUS 6.12 [21] for both Tests 1 and Test

2. Parametric study were then conducted for various types of LG as described in

Section 4 based on the validated FE models.

3.1. Thermal model

To determine the structural response of laminated glazing to fire, a heat transfer

analysis was firstly performed to obtain the thermal variation to be applied to the

structural models. Two heat transfer models have been simulated using measured

temperatures on S1 from Test 1 and Test 2.

3.1.1. Model description

The eight-node solid elements DC3D8 are employed. A heat transfer model was

created with 10 mm element size along the LG length and width. In the model, there

are seven elements (three for each glass panel and one for the PVB layer) along the z-

axis, resulting in 25200 elements in total, as shown in Fig. 4(a). The mesh in x and y

axis was proved to be suitable for the glass with this dimension [22]. The material

properties of clear float glass and PVB at ambient temperature are listed in Table 2

[10, 23, 24], which are used in both thermal and structural models.

Table 2. The physical properties of glass and PVB [10, 23, 24].

Properties Symbol Value

Glass

Density (kg/m3) ρ 2500

Thermal expansion coefficient (1/K) β 8.46×10-6

9

Reference temperature (K) TR 280

Specific heat capacity (J/(kg·K)) c 820

Thermal conductivity (W/(m·K)) k 0.94

Emissivity ε 0.85

PVB

Density (kg/m3) ρPVB 1070

Thermal expansion coefficient (1/K) βPVB 4×10-4

Reference temperature (K) TRPVB 280

Modulus of elasticity (Pa) EPVB 5.0×107

Poisson’s ratio νPVB 0.49

Specific heat capacity (J/(kg·K)) cPVB 1100

Thermal conductivity (W/(m·K)) kPVB 0.221

The average temperatures (Fig. 4(b)) of the measuring results from the experiments

were applied as boundary conditions in exposed and covered areas on S1. Since the

experiments were conducted in open space, the surrounding gas temperature on S4 is

assumed keeping constant as 280 K.

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(a) Mesh grid

(b) Temperature loading on S1

Fig. 4. Thermal loading and mesh generation in simulations of Tests 1 and 2.

It is assumed that the heat exchange between LG and ambient only occur on S4 by

convection and radiation during the fire. The convection coefficient and effective

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emissivity factor was set to 40 W/(m2·K) [5] and 0.85 [23], respectively. It is well

known that the uncoated glass has an emissivity for both exposed and unexposed

surfaces between 0.8-0.9 [25] or 0.85 [23], thus the usage of identical emissivity of

0.85 for S4 should be reasonable.

Thermal boundary conditions are specified as prescribed temperature, surface

convection (Eq. 3) and radiation (Eq. 4).

44\* MERGEFORMAT ()

55\* MERGEFORMAT ()

where q̇ ¿ is the heat flux, h is the film coefficient, is the prescribed temperature, θ0 is

the sink temperature, is the emissivity, is the Stefan-Boltzmann constant,θZ is the

value of absolute zero on the temperature scale being used.

Heat conduction is assumed governed by the Fourier law:

66\* MERGEFORMAT ()

where k is the thermal conductivity and x is the position.

3.1.2. Validation

In order to validate the thermal models, the temperature measured by TC10 at the

centre of S4 was compared to the temperature predicted by the heat transfer models

(Fig. 5). It can be seen that the numerical results agree very well with the

experimental results, indicating the heat transfer model is capable of predicting the

glazing temperature under fire conditions. As shown in Fig. 5, the temperature

measured in Test 2 is slightly higher than numerical results after around 400 s. This is

likely to be caused by the direct flame heating that went through the bubbles in PVB

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and the cracks around TC10 in Pane 1 (as seen in Fig. 6), which has not been

considered in the heat transfer model. It should be noted that all the cracks occurred

before 350 s in both tests, so the deviation does not affect the numerical analysis in

the present work.

Fig. 5 Comparison of the temperature at the centre of S4 between experimental and numerical results

Fig. 6. The crack path in Test 2

3.1.3. Heat transfer analysis

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The temperature distributions of S3 on Pane 2 at breakage time (199 s and 332 s in

Test 1 and Test 2) are illustrated in Fig. 7 with an identical temperature legend. It can

be found that, at breakage time, the S3 temperature field in both simulations are

similar, with central temperatures of 100 °C in Test 1 and 114 °C in Test 2 and

covered temperatures of 30 °C in Test 1 and 41 °C in Test 2. Thus, the temperature

differences between the covered and uncovered region of Pane 2 at breakage time are

respectively 70 °C and 73 °C in Tests 1 and 2, which are slightly lower than that of

Pane 1 (96 °C and 79 °C) but still consistent with the critical temperature difference

of 60-90 °C in the previous work [2, 4, 26].

Fig. 7 The temperature field (K) of S3 at the time of Pane 2 breakage, Test 1(left) and Test 2 (right)

To investigate the temperature gradient along the thickness, the temperature variances

at exposed and covered areas on S1-S4 are shown in Fig. 8. It can be seen that the

temperature differences between the exposed and covered areas are significant on all

surfaces. Pane 1 plays a very important role for thermal resistance as seen by the

significant temperature decrease from S1 to S2. Thus, although the fuel masses are

different between the tests, the final temperature of PVB layer and Pane 2 are very

similar, which are approximately 150 °C and 100 °C respectively. The temperature

field will be implemented into the mechanical model to calculate the stress

distribution in LG.

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Fig. 8. Temperatures measured on each surface, Test 1 and Test 2.

3.2. Structural model

In order to investigate the failure mechanisms of the laminated glass, a thermo-

mechanical analysis has been conducted.

3.2.1. Model descriptions

The model consists of 25,200 C3D8 elements which has the same grid size as the heat

transfer model. Temperature-independent mechanical properties, as illustrated in Fig.

15

9 and Table 3, include glass tensile strength, elasticity modulus and poisson’s patio

which were obtained from the author’s previous tests using MTS 810 and open

literature [10, 27]. The tensile strength of glass was measured at 25-400 °C [10]. The

soda-lime glass normally has a constant coefficient of linear expansion at 25-400 °C

[27] which is 9.6×10-6 1/K [28]. However, the authors measured the average thermal

expansion coefficient of 8.46×10-6 1/K using Netzsch Dilatometer [10]. The slight

difference may be caused by the manufacturing procedure and raw materials. Thus, in

the numerical models, the thermal expansion coefficient is assumed constant with the

value of 8.46×10-6 1/K.

Fig. 9. The applied temperature-dependent properties.

Table 3. The temperature dependent physical properties [10, 27].

Temperature (°C) Tensile strength (MPa) Elasticity modulus (GPa) Poisson’s ratio

25 35.72 72.90 0.165

75 27.96 -- --

100 26.60 74.00 0.171

125 29.07 -- --

150 30.92 -- --

16

200 32.00 75.10 0.173

300 30.16 76.2 0.175

400 29.78 77.2 0.177

It should be noted that the thermal expansion or deflection of stainless steel frame

would not be significant under such experimental condition [18]. The glass frame

offers no restraint to the glass since the maximum expansion is less than 1 mm which

is less than the normal gap of several mm between the frame and the panel [2, 4]. This

could ensure the glass pane was free to expand against the framing so that mechanical

effect of framing on glazing was very limited. The vertical support of glass panel is

modelled by restraining the vertical movement of the bottom surface of the glass.

It is very difficult to know the interaction status between the glass panel and the PVB

layer, it was assumed that the glass and PVB behaved in same motion by rigidly

restrained to each other at the interface. Considering the pool fire condition in which

the PVB layer did not significantly detach, as observed after tests, thus the assumption

should be reasonable.

3.2.2. Crack initiation validation

The thermal stress distribution is calculated by implementing the temperature variance

at each node in glazing. Once the maximum principle stress exceeds the temperature-

dependent tensile strength of glass, the crack in glass pane is assumed to be initiated

[9]. In experiments, all the cracks were initiated at the edges of glass panel, thus this

paper focus on the maximum stress in the edge area. As the numerical model is

central symmetry, the crack initiation located at each edge center in the real situations.

17

The stress contours at the time of breakage of Pane 1 and Pane 2 in Test 1 and Test 2

are shown in Fig. 10, where the white area represents the stress which is larger than

the tensile strength. In Fig. 10(a), It can be seen that the exceeding stress appears in

the middle area around the glass edge, which is very consistent with the conclusion

that the all crack initiation locations were within the area between left and right

quarter point of the glass edge [-0.25L, +0.25L] (the origin is at the centre of glass

edge; L is the edge length) in previous 27 experiments. Through comparing the tensile

stress along the glass panel edge, it was found that for Pane 1, the stress at the centre

of edge of S1 is maximum; for Pane 2, the maximum stress point appears at the centre

of edge of S3. The stress distributions of other panels are illustrated in Figs. 10(b)-(d),

showing that the specific location of maximum stress in Pane 2 is also located in the

middle edge section. Therefore, both the breakage time and crack initiation location

agree well with the experimental results, as shown in Fig. 11, confirming the

reasonability of mechanical model.

(a) Pane 1, Test 1

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(b) Pane 2, Test 1 (c) Pane 1, Test 2

(d) Pane 2, Test 2

Fig. 10. The maximum principle stress at breakage time.

Fig. 11. The crack path of Tests 1 (left) and 2 (right) (viewed from the ambient side).

19

3.2.3. Breakage time validation

The predicted breakage times from thermo-mechanical models are used to compare

with the observed experimental results. The maximum normal stress criterion, also

known as Coulomb’s criterion, is employed for crack criteria, which states that crack

occurs when the maximum normal stress reaches the ultimate strength of the material

[29]. The ultimate tensile strength is smaller than compressive strength in most cases.

By assuming the same ultimate strength Sut in tension and compression, which means

that the glass panel cracks in the multi-axial state of stress when the maximum normal

stress exceeds the ultimate tensile or compressive strength. This is the safer way to

predict the crack initiation, and can be written as:

77\* MERGEFORMAT ()

where Sut is the ultimate tensile stress of glass pane. Here Sut (T) is temperature

dependent.

In the FE models, the glass panels are considered to break once the maximum

principle stress exceeds the defined tensile strength of glass. The maximum principle

stress histories of Points 1 and 2 (covered area) are calculated and compared with

temperature-dependent tensile strength, as shown in Fig. 12. The highlighted crossing

point represents the stress exceeds the tensile strength at a certain temperature and

thus the first crack occurs. Then the breakage time can be obtained according to the

corresponding covered-area temperature on the x-axis, as listed in Table 4. It can be

seen that the breakage times of the two fire-exposed glass panels generally agree very

well with the experimental results, with the errors smaller than 11%. However, the

relatively large errors in Pane 2 of Test 1 should be noted: the predicted time is

significantly smaller than experimental work. This may be caused by the assumption

20

of uniform thermal loading, where the temperature difference at crack initiation

location (right edge viewed from the ambient side) is larger than experiments

resulting in earlier breakage time.

(a) The location of studied point

(b) The stress history of Points 1 and 2 of Test 1

21

(c) The stress history of Points 1 and 2 of Test 2

Fig. 12. The stress history at the glass edge center for Test 1 and Test 2.

Table 4. The summary of breakage times.

Glass panes Experimental (s) Numerical (s)

Test 1Pane 1 118 105

Pane 2 199 137

Test 2Pane 1 258 235

Pane 2 332 274

4. Parametric study and discussion

The numerical heat transfer and mechanical model has been verified by the

experimental results, which were then used for parametric analysis. The parameters

regarding the various types of LG have been studied, including the interlayer

thickness and number of glass layer, which have the significant effect on LG fracture

behaviour under impact loading [30]. The different glass types, such as float and

22

tempered glazing have been considered, however have a very limited effect on the

heat transfer process due to almost identical thermal conductivity.

The specific values used in the parametric study are summarized in Table 5.

Parametric study (PS) 1, 2, 3 and 4 (Group 1) are designed to investigate the effect of

interlayer thickness; PS 1, 5 and 6 (Group 2) are for the glass layer number

investigation. The thermal loading of Test 1, with shorter calculation time, is applied

to these simulations and the temperature gradient across the thickness direction at

different times are obtained.

Fig. 13. The heat transfer through the thickness of LG.

In Fig. 13, it can be seen that the two factors can significantly affect the temperature

distribution, among which the glass layer can affect the temperature much more than

the interlayer thickness. In particular, for PS5 and PS6, the ambient glass

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temperatures are almost as low as room temperature. The change in temperature will

then influence the breakage times of Pane 1 and Pane 2.

The breakage time is determined by the first principle stress as well and listed in

Table 5. It can be seen from Group 1 that the Pane 1 breakage time decreases from

105 s to 101 s with the increase of gel layer thickness. This result is caused by the low

heat conductivity of PVB layer, resulting in more heat stored in Pane 1 when the PVB

thickness increases. Thus, the breakage time of Pane 1 decreases. This phenomenon is

similar to the glass tests with different heat conduction lump attached to the back of

single glazing, in which the lump with lower heat conductivity induced higher

temperature in glazing [31].

No. of

Simulations

Glass

layer

number

Glass

thickness

(mm)

Interlayer

thickness

(mm)

Pane 1

(s)

Pane 2

(s)Pane 3 (s) Pane 4 (s)

PS1 2 0.6 0.38 105 137 -- --

PS2 2 0.6 0.76 103 144 -- --

PS3 2 0.6 1.14 102 148 -- --

PS4 2 0.6 1.52 101 153 -- --

PS5 3 0.6 0.38 117 142 No failure --

PS6 4 0.6 0.38 131 147 232 No failure

Table 5. The summary of the parametric analysis and breakage times.

For Group 2, the breakage times of Pane 1 increase when the glass layer number

increases. If the number of the glass layer is 3 or 4, the ambient side panel will not

break which indicates the significant fire resistance improvement.

24

To clearly demonstrate the relationship, the breakage times of Groups 1 and 2 are

plotted in Fig. 14. Different from Pane 1, the breakage time of Pane 2 increase

monotonically. The results suggest that the interlayer thickness of 0.38 mm can

provide relatively high fire resistance although the 1.52 mm-thick interlayer perform

best. On the other hand, for PS5 and PS6, the three layer of glazing is good enough

for the design fire in this work. The increase of interlayer thickness and glass layer

number will increase the construction cost. Therefore, the cases of PS1 and PS5 are

recommended considering their fire performance and expenditure. However, more

experimental studies should be conducted for LG under real fire conditions, which are

anticipated different from single and insulated glazing.

Fig. 14. The breakage time of Pane 1 and Pane 2 varied with (a) thickness of PVB interlayers and (b)

number of glass layer.

5. Conclusions

In this work, FEM heat transfer and mechanical models were created and validated by

the experiments. The temperature-independent thermal properties of glazing were

employed and the failure mechanism of LG in the fire was revealed. Meanwhile, a

parametric study was conducted to investigate the effect of PVB layer thickness and

number of glass panels on the LG fire performance. The primary conclusions are as

follows:

25

1) As it is almost impossible to insert thermocouples into LG without structural

destruction in experiments, the calculated results would be helpful for the

understanding of heat transfer process in LG, particularly PVB layer and Pane.

2) The LG demonstrates good fire resistance in experiments as the cracked glass

panels can still hold together preventing the new vent formed under fire

conditions.

3) The FEM heat transfer and mechanical model developed in ABAQUS can

well predict the LG temperature distribution and thermal stress both in Pane 1

and Pane 2.

4) The breakage times determined by the first principle stress are generally

consistent with experimental results as well which indicate the verification of

the numerical model.

5) Both the interlayer thickness and glass layer number have a significant

influence on the temperature and stress distribution in glazing, resulting in

considerably different breakage times.

6) Considering the fire resistance and the construction cost, the LG of PS1 and

PS5 are recommended to be used. More experiments need to be conducted in

real fire conditions for further analysis.

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