1

The Reality of Quantifying Curtain Wall Spandrel Thermal

Performance: 2D, 3D and Hotbox Testing

Neil Norris1, Lawrence D. Carbary2, Stanley Yee3, Patrick Roppel4 , Phillip Ciantar5

ABSTRACT

Curtain wall assemblies, large expanses of aluminum framing, are not traditionally known for

their thermal efficiency. While insulated spandrels within the curtain wall framing have been

used in an attempt to improve the thermal performance of the curtain wall system, research and

testing has shown that typical insulated backpan systems fall short of expectations for the

thermal resistance of opaque walls in many codes and standards.

This low performance of spandrels is due to several limitations in typical spandrel designs and

materials. New technologies have been developed which try to address some of these issues.

One strategy is to use Architectural Insulation Modules (AIM) which comprise of highly

insulating vacuum insulated panels (VIP) sandwiched within an insulated glazing unit (IGU).

When AIM units are coupled with large thermal breaks in framing, the thermal performance of

the spandrels can greatly improve.

Unfortunately, there is no single commonly validated and accepted approach in evaluating the

thermal performance of curtain wall spandrel sections as there is with the curtain wall vision

sections (NFRC-100). For this study, three approaches were examined: 1) Guarded hotbox

testing of physical specimens according to ASTM C1363, 2) Two-dimensional thermal modelling

following variations of NFRC-100 and 3) Three dimensional thermal modelling following

ASHRAE-1365 RP. For each of these methods, a square section of high performance curtain

wall spandrel was analyzed for thermal transmittance for five spandrel configurations, including

typical IGU and backpan systems, as well as AIM units.

This paper outlines a high level comparison between these three approaches in analyzing the

thermal performance of spandrels. The results, in turn, help identify what the realistic thermal

performance values are that can be expected from curtain wall spandrels and how they can be

improved using highly insulating materials.

1 Neil Norris, Building Science Consultant, Morrison Hershfield Ltd, Burnaby BC, CAN 2 Lawrence D. Carbary, Building Industry Scientist, Dow Corning, Midland MI, USA 3 Stanley Yee, Façade Design and Construction Specialist, Dow Corning, Beaverton OR, USA 4 Patrick Roppel, Building Science Specialist, Morrison Hershfield Ltd, Burnaby BC, CAN 5 Phillip Ciantar, Design Engineer, Northern Facades, Mississauga ON, CAN

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INTRODUCTION

Recent trends in construction over the last 15 years has led to an increase in the use of curtain

wall systems on large portions of buildings. The trend to this form has been driven by a number

factors including (but not limited to):

• cost advantages of being able to build without

the need for extraneous exterior staging or

scaffolding;

• ease of modular installation and construction

sequencing

• architectural considerations, such as look and

shape of the building

• off-site construction methods imparting better

quality management of assemblies;

• often a single specialist

contractor/manufacturer for streamlining of the

envelope design, installation and coordination

(which can be quite attractive for architects);

However, along with the push within the construction

industry to create more energy efficient buildings,

curtain wall systems have come under increased

scrutiny for their thermal performance (Totten

2008). The growing need for efficient envelope

designs can create conflict with using curtain wall

systems. These systems have traditionally been

seen to have poor thermal resistance compared to highly insulated wall assemblies because of

their large amounts of highly conductive aluminum framing. Even when comparing between

standard and high performance systems from various manufacturers, thermal breaks within the

curtain wall framing can be only so effective in reducing the heat flow. Critics point to the

thermal performance of curtain wall and other types of glazing systems and suggest that glazed

towers are an inappropriate choice from an energy conservation perspective (Straube 2008,

Wilson 2010). It is implied in ASHRAE 90.1 through its various energy compliance paths that

windows need not be larger than what is necessary for the benefits of daylighting (less than

40% glazing-to-wall ratio). Floor-to-floor glazing system designs, frequently associated with

curtain wall systems, often wrap around almost entire façades of mid- and high-rise buildings

leaving very little opaque area (See Figure 1).

In order to increase the thermal resistance of the envelope but still maintain the desired look of

the building, curtain wall building designs often use insulated spandrel sections within the

curtain wall framing as the opaque portions of the envelope, as opposed to transitioning to a

steel stud or masonry wall. Typical spandrels consist of either metal panel, single pane glass or

a double pane insulated glazing unit (IGU) outboard of insulation between the mullions or

contained within a metal backpan. This is often within similar curtain wall framing as the

adjacent curtain wall vision sections (See Figure 2).

Figure 1: A curtain wall high-rise tower with high window to wall ratio adjacent to a mid-

rise tower with low window to wall ratio

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Unfortunately, these typical spandrel systems, while better in thermal resistance than the

glazing performance, often falls short of expectations, especially when compared to prescriptive

requirements for opaque walls from codes and standards, such as ASHRAE 90.1.

For many curtain wall systems the insulating layer in the

spandrel portion is not aligned with the insulating layer of

the vision portion. In Canada, this insulation is usually

contained within a metal backpan in line with the

mullions, while other systems have the insulation directly

sitting on the mullions. This allows heat flow to easily

bypass the insulation through the aluminum frames (See

Figure 3). While thermal breaks on curtain wall systems

are low conductivity materials, due to their small size they

have significantly less thermal resistance compared to

the insulation.

Curtain wall systems are also limited in space for the

insulation. Though larger system depths are available, in

many systems there is typically between 76mm (3”) and

102mm (4”) of mineral wool insulation. However, even

with larger amounts of spandrel insulation, due to

diminishing returns from the thermal bridging through the

frames, this additional insulation often does not add much

to the resistance of the system (MH 2014).

With all of these factors considered, for a conventional curtain wall system this may result in a

thermal transmittance of the spandrel of approximately USI-1.42 (U-0.25) which is not

significantly better than the vision systems (Straube 2008; MH 2014).

Figure 3: Typical insulated curtain wall spandrel with front panel

removed

Figure 2: Heat flow paths in a spandrel system with insulated backpan

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Overall Opaque Envelope Thermal Performance

Before discounting curtain wall systems as poor thermal performers, it may be best to step back

and take a look at the entire building envelope on a larger scale. What if other opaque building

envelope assemblies, like steel stud and mass walls, were not much better?

Research into the thermal performance of opaque building envelope assemblies was conducted

in ASHRAE 1365-RP and as part of The Building Envelope Thermal Bridging Guide (MH 2014).

For these reports, hundreds of opaque assemblies were analyzed using 3D thermal modelling.

The heat flow characteristics were found for not only the clear field portions of the wall, but also

included the effects of thermal bridging through transitional details, such as parapets, slab

edges and glazing transitions (see Figure 4). This research found that when thermal bridging in

these details are included, the thermal resistance of the opaque building envelope can be

significantly lower than what has been previously assumed. For example, the thermal

transmittance through a mass wall assembly with a clear wall equal to USI-0.38 (U-0.07) may in

fact be closer to USI-1.14 (U-0.20) if the assembly contains large amounts of thermal bridging in

transitional details, such as uninsulated concrete slab edges, shelf angles or continuous

flashings through window openings that bypass the window frame thermal break. Interestingly

enough, insulated spandrels are typically run outboard of slab edges and the thermal bridging

effects from the intermediate floor slab in curtain wall designs can be negligible.

Figure 4: Steel stud assembly and thermal profile (based on a temperature index) showing heat flow bypassing the insulated wall through the balcony slab

When thermal bridging is minimized, steel stud and mass wall construction can potentially

produce much higher thermally resistant opaque envelopes than typical spandrel systems.

However, current designs for these construction types often do not appropriately address these

transition details. As a result, the overall thermal transmittance for the envelope of these more

“opaque” buildings may not be in fact significantly better than a curtain wall building.

Heat flow paths

through slab

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Regardless, energy efficiency is not the only priority for building designs and is only one of

many considerations for development. The reality of the construction industry is that there is a

demand for predominately curtain wall vision and spandrel clad buildings and they will continue

to be built. In trying to improve energy-efficiency within this context, efforts will be best utilized if

put towards increasing the performance of curtain wall systems.

Improvements to Spandrels Using Highly Insulated Assemblies

One tactic in alleviating some of the issues with spandrels discussed earlier in this paper is by the use of highly insulating materials within the spandrel. Vacuum insulated panels (VIP) contain a fumed silica or aerogel within a protective membrane wherein any gases, such as air, have been removed. This helps eliminate heat transfer through the panels by both convection and conduction. While there is thermal bridging through the protective foil in the panels, research has shown that VIP’s can have extremely high insulation values, reaching over RSI-3.5 per inch (R-20 per inch) (IEA/ECBCS 2005). This allows VIPs to remain thin, yet still provide a large amount of insulating value within the spandrel, which is an advantage due to the limited space available. The VIP properties also provide greater opportunity to align the spandrel insulation with the glazing, keeping it in the same plane as the curtain wall neck. One technique to do this is to contain the VIPs within an IGU, termed Architectural Insulation Modules (AIM), and treat it similar in terms of installation as an IGU placed at exterior side of the curtain wall spandrel. Two examples of AIM units are shown below in Figure 5, where an AIM has been created within a double glazed IGU and within a triple glazed IGU, creating a shadow box.

Figure 5: Concept of a double glazed AIM and triple glazed “Shadow” AIM unit

As with IGU’s, there is further thermal bridging around these AIM units through the spacers and

through the frame. In order to maximize the thermal resistance for the overall spandrel system

using AIM units, the spacer and curtain wall thermal break should also be effective in minimizing

thermal bridging. For this study, a high performance curtain wall system was analyzed for

thermal resistance with typical spandrel configurations and using AIM units and is detailed in the

following section.

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EVALUATING SPANDREL THERMAL PERFORMANCE Accurate evaluation methods are needed in order to fairly determine and compare the thermal

performance of different curtain wall spandrel configurations. Simple calculations, like parallel

path and isothermal planes methods are inappropriate for insulated spandrels as they do not

accurately account for lateral heat flow between components. As a result they tend to greatly

overestimate the thermal resistance. Insulated spandrels require more robust methods of

analysis.

For curtain wall vision sections, the performance is commonly determined through 2D modelling

and analysis based on NFRC-100. In North America, this approach is almost universally

accepted for vision glazing and is used ubiquitously in the industry, from manufacturers to

consultants to energy standards like ASHRAE 90.1. Unfortunately, there is no single widely

validated and accepted method for evaluating opaque spandrels.

For this study, in order to provide guidance in how to evaluate the thermal performance of

spandrels, three methods were chosen and compared: Hotbox Testing based on ASTM C1363,

2D Thermal Modelling based on NFRC-100 and validated 3D Thermal Modelling based on the

methods in ASHRAE 1365-RP. Each approach is described further in their accompanying

sections in this paper.

For this comparison, the Northern Facades 6450 UCW System was analyzed for several spandrel assemblies, including typical IGU and AIM configurations. 6450 UCW is a thermally broken high performance curtain wall system. This curtain wall design uses mechanical fastening of the glazing infill with a fiberglass pressure plate and stainless steel fasteners placed into rigid thermal break materials. This particular curtain wall design was benchmarked using a typical spandrel assembly consisting of an IGU of 25mm (1”) with air fill, outboard of mineral wool in a galvanized backpan. The benchmark spandrel was compared to the following alternate configurations: 1) The same IGU but with argon fill, 2) A combined IGU and AIM unit, otherwise termed a “Shadow AIM” 3) An AIM unit 50mm (2”) in thickness and 4) The 50mm (2”) AIM unit without the insulated backpan. Additional descriptions for these spandrel assemblies are given in Table 1. The jamb details for the curtain wall assembly are shown in Figure 6 for the IGU and AIM spandrel assemblies. Except for some minor fasteners into the backpan, all details (head sill and jamb) were identical. In order to accommodate for the thickness of the AIM units for those scenarios, the curtain wall neck has an additional thermal break extension. These assemblies were evaluated using physical testing and computer modelling. In order to best compare the results, the tested and modelled assemblies were the same in geometry and dimensions. All the analyzed assemblies were 1524mm x 1524mm (5-ft x 5-ft) with the full frame and contained only the spandrel portions of the curtain wall, there was no adjacent vision glazing sections tested or modelled for this comparison.

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Spandrel Assembly Assembly Description

IGU Air 25mm (1”) Double IGU (6-12-6) Warm edge plastic spacer, air fill, low E on surface 2. 125mm (5”) mineral wool in front of a galvanized backpan

IGU Argon 25mm (1”) Double IGU (6-12-6) Warm edge plastic spacer, argon fill, low E on surface 2. 125mm (5”) mineral wool in front of a galvanized backpan

Shadow AIM

50mm (2”) Triple Glazed AIM (6-15-6-15-6) warm edge plastic spacer, air fill, low E on surface 2. Gray ceramic frit on surface 4. 15mm (3/5”) VIP between surface 4 and 5. 125mm (5”) mineral wool in front of a galvanized backpan

AIM 50mm (2”) Double Glazed AIM (6-38-6) warm edge plastic spacer. 38mm (1 1/2”) VIP between surface 2 and 4. 125mm (5”) mineral wool in front of a galvanized backpan

Table 1: Spandrel Assembly Descriptions

Figure 6: Jamb Detail of the modelled spandrel system for the IGU Assemblies (left) and AIM/Shadow AIM (right)

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Evaluation of Spandrels using Guarded Hotbox Testing A series of physical thermal tests were performed on the various spandrel assemblies using a guarded hotbox (ATI 2014). The test and calculations were done in accordance with ASTM C 1363. Figure 7 shows the typical setup apparatus from this test standard. Forced air at a given cold temperature was applied against one side of the assembly in the climate chamber, while heat was added to the warm side to maintain a consistent interior temperature on the opposite side of the assembly. Air movement in the warm side was driven mainly by natural convection from the specimen. The heat loss through the specimen was measured to determine the U- and R-values of the tested assembly. Figure 8 shows the AIM assembly specimen during its construction.

Figure 7: Guarded Hot Box Schematic from ASTM C1363-11

Figure 8: AIM Test Specimen Assembly

Table 2 shows the targeted environmental conditions under which the tests were performed6. The thermal transmittance (U-Value) and thermal resistance (R-Value) for the five assemblies from the hotbox tests are shown in Table 3.

Condition SI Value IP Value

Exterior Air Temperature -18oC -0.4oF

Interior Air Temperature 21oC 69.8oF

Exterior Wind 24 km/hr 15 mph

Interior Air Flow 0.08 km/hr 0.05 mph

Table 2: Target Guarded Hotbox Conditions

6 As with all experimental setups, there is variability between tests of the environmental conditions. Between tests,

the temperatures varied on average less than 0.02oC (0.04oF) and windspeeds less than 0.008km/hr (0.005 mph)

9

Spandrel Assembly

U-value

W/m2K

(BTU/hrft2oF)

R-value

m2K/W

(hrft2oF/BTU)

IGU Air USI-0.89 (U-0.156) RSI-1.13 (R-6.40)

IGU Argon USI-0.87 (U-0.153) RSI-1.15 (R-6.52)

Shadow AIM USI-0.56 (U-0.097) RSI-1.80 (R-10.23)

AIM USI-0.50 (U-0.088) RSI-2.00 (R-11.36)

AIM w/out Backpan Insulation USI-0.60 (U-0.105) RSI-1.68 (R-9.52)

Table 3: Hotbox Test Results for Selected Spandrel Assemblies

Evaluation of Spandrels Using 2D Thermal Modelling

For the 2D modelling of the five spandrel assemblies, the modelled assembly geometries were

matched as close as possible to the tested assemblies based on the detail drawings. Air cavities

were estimated with an equivalent conductivity based on ISO 15099 that includes convection

and radiation effects. Using the 2D thermal modelling software THERM7, the analysis for the 2D

models was approached in two different ways. The first method was to model the five

assemblies according to NFRC-100 guidelines, while the second was an alteration to that

method.

Using NFRC-100 with THERM is typical for

calculating U-values of fenestration

products. This approach is the industry

standard when it comes to evaluating

glazing. While there is no official validated

and sanctioned procedure from the

National Fenestration Rating Council

concerning calculating the thermal

performance of spandrels, some designers

have still applied the NFRC-100 method to

spandrels. In this approach, the spandrel

components, as with fenestration, are

divided into separate 2D models. In the

case of the spandrel assemblies presented

here, this is the head, the jamb and the

sill. A 1D heat flow U-value is found for

the center of the panel (COP) and for

each component model, two U-values are found, one for the frame and one for a portion of the

glazing/panel away from the frame, termed the “edge”. The edge takes into account increased

heat flow for a certain distance away from the frame before it reaches the 1D heat flow. The

basic divisions of the system for calculations are shown in Figure 9 and Figure 10.

7 THERM 6.3 is free use software from Lawrence Berkley National Laboratories

Figure 9: 2D Model Divisions for the Spandrel Assemblies following NFRC-100

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Figure 10: Section View of the 2D Model Divisions for the Spandrel Assemblies following NFRC-100 and Example Thermal Profile

Each of the U-values from these components (Center, Edge and Frame) are then area weighted

in order to find the overall U-value for the system.

For the first approach of analyzing the spandrels in 2D for this paper, the model for each

component included the mullion framing and 162mm (6.4”) of the spandrel section away from

the frame. The edge distance along the spandrel was set to match the NFRC-100 guidelines for

fenestration of 63.5mm (2.5”). The 1D COP value was taken from adding the resistances of the

spandrel materials through the center of the assembly. It was observed that the isotherms were

not parallel at 63.5mm away from the frame, the area that sets the edge in the guidelines.

Hence another approach was attempted.

For the second approach, a second set of 2D models was made where the distance of the

spandrel away from the frame was increased to be half of the assembly that was actually tested.

This results in a model with the modeled spandrel extending 748mm (29.4”) away from the

frame. Due to the symmetry of the components and that the full heights and widths of the

system are being modelled, there was expected to be better accounting of the lateral heat flow

that goes through the metal backpan. This can change the edge values and center of panel

values. As such, the COP values were taken directly from the model instead of calculated

separately.

The U- and R-value results for five spandrels following the two approaches are presented in

Table 4. The results for the NFRC-100 approach with the smaller 162mm (6.4”) section in Table

4 is anywhere from 20% to 33% higher in R value compared to what was measured in the

hotbox tests in Table 3. When using the modified approach with the taller model, the results

were within 16-25% of the hotbox values.

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Spandrel Assembly

162mm (6.4”) Model 748mm (29.4”) Model

U-value

W/m2K

(BTU/hrft2oF)

R-value

m2K/W

(hrft2oF/BTU)

U-value

W/m2K

(BTU/hrft2oF)

R-value

m2K/W

(hrft2oF/BTU)

IGU Air USI-0.65

(U-0.114)

RSI-1.54

(R-8.74)

USI-0.69

(U-0.122)

RSI-1.44

(R-8.20)

IGU Argon USI-0.63

(U-0.111)

RSI-1.58

(R-8.98)

USI-0.68

(U-0.12)

RSI-1.47

(R-8.34)

Shadow AIM USI-0.43

(U-0.076)

RSI-2.30

(R-13.08)

USI-0.46

(U-0.082)

RSI-2.16

(R-12.27)

AIM USI-0.34

(U-0.059)

RSI-2.99

(R-16.96)

USI-0.38

(U-0.067)

RSI-2.64

(R-14.96)

AIM w/out Backpan Insulation

USI-0.47

(U-0.082)

RSI-2.14

(R-12.14)

USI-0.52

(U-0.091)

RSI-1.93

(R-10.94)

Table 4: 2D Modelling of Spandrel Assemblies using NFRC-100 and Adjusted Approaches

Evaluation of Spandrels Using 3D Thermal Modelling For the 3D thermal modelling, the assemblies were analyzed using proprietary finite element software8. The methodology for 3D modelling follows the approach in ASHRAE 1365-RP wherein the full 3D assemblies are modelled with contact resistances between components. All material properties used matched those in the 2D thermal model. The 3D model of the IGU Argon assembly is shown in Figure 11 along with an example temperature profile.

Figure 11: Modeled 3D IGU Argon Spandrel Assembly in High Performance Frame

8 The model was created in NX 8.0 and was extensively calibrated to hotbox tests for ASHRAE 1365-RP

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The overall heat flow through the assembly was determined by the model and the performance values are calculated based on the assembly dimensions. Table 5 shows the U- and R-Value results from the 3D model for the five spandrel assemblies. From Table 5 it can be seen that the 3D model results match the hotbox values consistently to within 3%.

Spandrel Assembly

U-value

W/m2K

(BTU/hrft2oF)

R-value

m2K/W

(hrft2oF/BTU)

IGU Air USI-0.90 (U-0.158) RSI-1.13 (R-6.40)

IGU Argon USI-0.87 (U-0.153) RSI-1.15 (R-6.54)

Shadow AIM USI-0.57 (U-0.101) RSI-1.75 (R-9.94)

AIM USI-0.51 (U-0.089) RSI-1.97 (R-11.20)

AIM w/out Backpan Insulation USI-0.59 (U-0.103) RSI-1.70 (R-9.68)

Table 5: 3D Modelling Results for the Spandrel Assemblies

Potential Sources of Error For this study, there are differences between the modeling and the testing that may have introduced error that could affect the comparisons between methodologies. For both the 2D and 3D models, the film coefficients take into account convective flow, both forced and natural, however the value is averaged across the surfaces. In reality, within the hotbox tests, the natural convective flow on the warm side will create a boundary layer which will result in a varied film coefficient. This difference is expected to mainly affect surface temperatures between top and bottom mullions and not overall system U-values significantly. The bolts that penetrate the curtain wall thermal break in order to hold the pressure plate to the frame act as small thermal bridges. Experimental analysis for aluminum pressure plates has shown that 18% of the heat flow goes through bolts at a standard 229mm (9”) o.c. (Griffith et al. 1998). Note that, for this paper, the pressure plates in these assemblies were fiberglass. While bolts were included in the 3D model, they were not in the 2D model results in Table 4. When the bolts were added to two of the 2D assemblies the thermal results were better matched to the hot box results (between 10-15% difference). Adding the bolts is a typical practice in Europe, found in EN ISO 12631 (2012) and also indicated in NFRC-100. The results of this comparison are shown in Table 6.

Spandrel Assembly

Without Bolts With Bolts 229mm (6”) o.c.

U-value W/m2K

(BTU/hrft2oF)

R-value m2K/W

(hrft2oF/BTU)

U-value W/m2K

(BTU/hrft2oF)

R-value m2K/W

(hrft2oF/BTU)

Shadow AIM USI-0.46

(U-0.082)

RSI-2.16

(R-12.27)

USI-0.50

(U-0.088)

RSI-1.99

(R-11.32)

AIM USI-0.38

(U-0.067)

RSI-2.64

(R-14.96)

USI-0.42

(U-0.074)

RSI-2.37

(R-13.5)

Table 6: Comparison of selected 2D models with and without bolts for 748mm model

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Physical testing also has errors. Losses through the assembly edges, control of airflow, sensors placement, general workmanship and construction tolerances of the test specimen all have an effect on the results. That being said, guarded hotbox tests are extensively calibrated and the accuracy and uncertainty of values for calibrated hotbox tests at commercial or government research facilities can be on the order of ±6-8% for framed systems (Elmahdy et al. 2000).

ANALYSIS COMPARISONS Despite the sources of error described in the previous section, the 3D modelling results are well within the hotbox tests range of uncertainty. For 2D simulations of assemblies with U-values less than 1.7 W/m2K (0.30 Btu/hft2oF) NFRC 100 section 4.7.1 defines accepted equivalence for fenestration as within 0.17 W/m2K (0.030 Btu/hft2oF) between testing and simulation values. The AIM results for the 2D modelling in Table 4 are within this threshold, however, there is a greater than 15% difference between hotbox tests and the simulation values. The inclusion of bolts in the model reduces this difference to within 10-15%. This indicates that the same application of NFRC-100 for equivalence for glazing to spandrels is inappropriate. However, the results still show it may be possible for 2D modelling of the spandrel using this approach to reach acceptable levels of accuracy compared to hotbox tests, as long as there are further adjustments. One such adjustment is in the determination of the edge distance for spandrels. For glazing, this edge distance is typically taken to be 63.5mm (2.5”) (ASHRAE 2013). However, due to the high conductivity of the metal backpan and the large range of insulation products that can be placed outboard of the backpan, the edge distance for the spandrel is difficult to pre-define. Depending on the size of the spandrel, these edge effects may in fact be so strong that the 1D COP value is never actually reached. The true edge distance for spandrels would have to be determined by an iterative process, not covered by this paper, which may provide increased agreement with hotbox tests. In choosing which method of analyzing spandrels is the “correct” approach will depend on its intended purpose. Using hotbox testing is likely done for preliminary design purposes of new systems, research work or for model calibration. Once calibrated, it is easier to create a 3D thermal model to test other configuration changes rather than a multitude of hotbox tests. The level of accuracy needed should also be a consideration. For specific project work it is likely not economical to perform multiple hotbox tests. On building projects, for compliance with energy standards using approaches such as whole building energy simulations, the differences in accuracy between 2D and 3D modelling may be lost in the overall model. If no other information is available, then 2D model results may provide a reasonable level of accuracy in these situations, especially in buildings where energy use is not greatly affected by the envelope. The 2D methods shown here could also be performed for simple comparisons between systems. Compared to 2D modelling, 3D allows for the accounting of corner effects and 3D lateral heat flow. However, for the tested assembly in this case, these effects appear to be minor due to the symmetrical nature of the test assembly. As a result, the full advantages of using 3D modelling may not yet be clearly evident. However, there are other heat loss effects that are difficult to account for in 2D or in hotbox testing when trying to determine actual in situ performance of curtain wall spandrels. In the assemblies analyzed thus far, only the heat flow through the spandrel was tested and modelled. This assumes that there are no other ways for heat to flow around this assembly. In reality, curtain wall spandrels are often flanked by vision portions, meaning that there can be additional heat flow effects from the transition between these

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adjacent assemblies (Totten 2008). Additionally, actual site assemblies are often coupled with slab edges and interior framing, which can also have an impact on the heat flow through the spandrels. This can be difficult to model in 2D as there is a much larger potential for lateral heat flow through the mullions to bypass the spandrel through the vision sections or slab. Even with hotbox testing it would be very difficult to incorporate all of these features into a test apparatus easily. There are few hotbox testing facilities that are able to accommodate multi-storey assemblies. On site testing for thermal performance can be incredibly difficult to accurately monitor heat flow and environmental conditions. As a result, the likely simplest way to reasonably predict the actual thermal performance of curtain wall assemblies on buildings is currently through 3D modelling. In order to fairly analyze spandrels with these other components using 3D modelling, it is best to separate the spandrel performance from the vision performance. Although they are part of the same system, the analysis of the glazing using NFRC is already well established. Using the methodology put forward in ASHRAE 1365-RP, the heat flow through the combined system and through the vision portion only is determined. The glazing values are subtracted from the combined assembly, which leaves only the heat flow from the spandrel and any additional heat flow affects from the transition between these two sections of the curtain wall system. This separation is shown in Figure 12.

Figure 12: Separation of curtain wall system between vision and spandrel sections for

calculations including transition effects between sections

Following this full 3D modelling approach, Table 7 below shows the U- and R-Values for the IGU with argon fill and 50mm (2”) AIM spandrel scenarios for the hotbox assembly from Table 5 and for the spandrel when the effects of flanking glazing, stud wall with air cavity and slab, as shown in Figure 12, are included. The spandrel size was kept the same and the flanking IGUs for both spandrel systems were the same IGU with argon fill. Note that the values are for the spandrel section only (this does not include all the heat loss through the glazing, only the flanking effects on the spandrel).

15

Spandrel Assembly

Hotbox Assembly (No Flanking Assemblies)

Spandrel System with Flanking Assemblies

U-value

W/m2K

(BTU/hrft2oF)

R-value

m2K/W

(hrft2oF/BTU)

U-value

W/m2K

(BTU/hrft2oF)

R-value

m2K/W

(hrft2oF/BTU)

IGU Argon USI-0.87

(U-0.153)

RSI-1.15

(R-6.54)

USI-1.15

(U-0.203)

RSI-0.87

(R-4.92)

AIM USI-0.51

(U-0.089)

RSI-1.97

(R-11.20)

USI-0.58

(U-0.102)

RSI-1.73

(R-9.84)

Table 7: U-values of Spandrel sections for hotbox assemblies and when the effects of the

transition to adjacent assemblies are included

From Table 7 it can be seen that the performance of the spandrels is reduced further when

these adjacent elements are accounted for. While not dealt with in this study, this bridging

through the adjacent elements will also impact the effectiveness of cavity insulation in the steel

stud pony walls with curtain wall systems. Using sprayfoam in the cavity is a common way to

improve the thermal performance, however if heat is allowed to bypass this insulation through

the vision head and sill transitions, then the increase in thermal performance from this added

insulation can be significantly reduced (MH 2014).

WHAT TO EXPECT FROM SPANDREL THERMAL PERFORMANCE From Tables 3-7, regardless of the analysis method, the AIM unit assemblies show significant improvement in thermal resistance over the typical spandrel systems. Assembly transmittances of USI-0.57 (U-0.1) are close to the expected clear wall values for mass walls with 50mm (2”) of rigid insulation. This means that there is potential for curtain wall systems to be on equal footing in thermal performance compared to some other types of common opaque constructions. When the spandrel and vision sections of the curtain wall are separated, then certain patterns and trends emerge. As stated earlier, typical spandrels using standard curtain wall can have transmittances as high as USI-1.42 (U-0.25) in normal construction. This assumes an area of spandrel that is approximately 40 to 50% of the floor to floor height, the remainder being vision glazing, which is typical with curtain wall systems with a framed upstand at the slab. Part of the reason there is such low performance from typical spandrels is the heat loss from around the frames. If the ratio of mullions to panel space is large, then the heat flow through these frames will dominate the performance. If this vision to spandrel ratio is changed, the amount of framing to infill area of the spandrel potentially also changes. While the system tested above was for a high performance system, this effect can be seen easily with a “standard” curtain wall system with minimal thermal break and monolithic glass at the spandrel. Figure 13 shows the various spandrel to glazing percentage assembly situations, while Figure 14 shows the changes in effective thermal resistance of the spandrel and glazing assembly as the areas change for the lower performance system, as predicted by 3D thermal modelling following ASHRAE 1365-RP. Figure 14 shows what can be expected from a typical standard curtain wall spandrel system for various heights of spandrel. By increasing to larger spandrel heights, improvements to the curtain wall spandrel R-value can be made. If this were coupled with a high performance system and with AIM units, there are potentially even larger gains.

16

Figure 13: Example variations of vision to spandrel ratios

Figure 14: Effective R-Value vs % of spandrel floor to floor height

0

0.2

0.4

0.6

0.8

1

1.2

0

1

2

3

4

5

6

7

0% 20% 40% 60% 80% 100%

Eff

ec

tive

Th

erm

al R

esi

sta

nc

e (

m2K

/W)

Eff

ec

tive

Th

erm

al R

esi

sta

nc

e (

fr2 h

r o

f/B

TU)

Spandrel Height to Floor-to-Floor Height

Vision

Opaque

Spandrel

17

CONCLUSIONS There is a reasonable amount of variation in thermal transmittance/resistance results for the analyzed spandrels between the three methods detailed in this paper. The approach recommended to be used will depend on the level of accuracy required. While hotbox testing can provide an accurate means of determining values for assemblies in isolation, due to limitations in typical hotbox apparatuses, actual performance of assemblies as installed is more difficult to evaluate. Hotbox tests are useful for research purposes and calibration for computer models, however they may not be practical for widespread use in construction projects. 2D modelling of the hotbox spandrel assembly using NFRC-100 can approach within 10-15% of the hot box test values as long as there are some adjustments to the approach, as detailed in the previous sections. Further alterations, such as changing the edge distance, will likely provide even better agreement. The use of 2D modelling is best used for quick comparisons between systems or where increased accuracy for thermal performance is less important. The 3D modelling approach in this paper, as per ASHRAE 1365-RP, provides the greatest agreement to the hotbox values (within 3%). 3D modelling requires competent modelers and a desire for accuracy. In order to more accurately analyze designs as they would be installed in the field, 3D modelling currently provides the most practical means of doing so. In every analysis approach, the AIM unit assemblies show significant improvement in thermal resistance over the typical spandrel systems. This offers great potential towards creating curtain wall systems that compare in thermal performance to other types of construction that have been previously assumed to be better performers. In envelope design, there are other factors in thermal performance to consider beyond U- and

R-values. In many areas that are subject to cold weather and high interior humidity,

condensation resistance is of significant concern. Next steps would be to examine the surface

temperatures from the three approaches and compare this to in-situ temperature measurements

to determine if condensation risk can be reasonably predicted.

18

ACKNOWLEDGEMENTS The authors gratefully acknowledge Flynn Canada and Northern Facades for providing the curtain wall system and Dow Corning for the AIM modules for testing. The authors would also like to acknowledge Architectural Testing Inc. for performing the hotbox tests.

REFERENCES Totten, P. E., O’Brien, S. M., & Pazera, M. (2008). The Effects of Thermal Bridging at Interface

Conditions. Building Enclosure Science & Technology, 1.

Straube, J. 2008. Can Highly Glazed Building Facades Be Green? Building Science Insights, BSI-006 Wilson, A. (2010). Rethinking the All-Glass Building. Environmental Building News BSN19 ASHRAE. 2010. ANSI/ASHRAE Standard 90.1 Energy Standard for Buildings Except Low-Rise Residential Buildings, Atlanta GA: American Society of Heating, Refrigerating and Air-Conditioning Engineers Inc. Morrison Hershfield Ltd. 2014. The Building Envelope Thermal Bridging Guide, Vancouver, B.C.: BC Hydro ASHRAE. 2011. 1365-RP Thermal Performance of Building Envelope Details for Mid- and High-Rise Buildings, Atlanta GA: American Society of Heating, Refrigerating and Air-Conditioning Engineers Inc. International Energy Agency/Energy in Buildings and Communities Programme. 2005. Vacuum Insulation: Panel Properties and Building Applications. Annex 39 High Performance Thermal Insulation Systems, Ed Erb, Markus. NFRC. 2010. NFRC-100 Procedure for Determining Fenestration Product U-Factors, Greenbelt, MD: National Fenestration Ratings Council ASTM. 2011. C1363-2011 Standard Test Method for the Thermal Performance of Building Materials and Envelope Assemblies by Means of a Hot Box Apparatus, West Conshohocken, PA: ASTM International. ISO, 2012. EN ISO 12631: Thermal Performance of Curtain Walling – Calculation of Thermal

Transmittance. Vienna, Austria: Austrian Standards Institute

Architectural Testing Inc. 2014. ASTM C 1363-2011 Thermal Performance Test Report Reports

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D0259.01-201-46-R0, St Paul MN

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Elmahdy, A.H.; Haddad, K., 2000. Experimental procedure and uncertainty analysis of a guarded hotbox method to determine the thermal transmission coefficient of skylights and sloped glazing. ASHRAE Transactions, v. 106, pt. 2, pp. 601-613 Griffith, B.; Finlayson, E.; Yazdanian, M.; Arasteh, D. 1997. The Significance of Bolts in the Thermal Performance of Curtain-Wall Frames for Glazed Facades. ASHRAE Transactions v. 105, pt1. pp 149-164. ASHRAE. 2013. ASHRAE Handbook–Fundamentals, p. 15.4 Atlanta: American Society of

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