Cladding Attachment 

Over Thick Exterior 

Insulating Sheathing

P. Baker, P. Eng, and R. Lepage 

Building Science Corporation

January 2014
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iii

Cladding Attachment Over Thick Exterior Insulating Sheathing

Prepared for:

The National Renewable Energy Laboratory 

On behalf of the U.S. Department of Energy’s Building America Program 

Office of Energy Efficiency and Renewable Energy

15013 Denver West Parkway 

Golden, CO 80401 

NREL Contract No. DE-AC36-08GO28308 

Prepared by: 

P. Baker, P. Eng, and R. Lepage

Building Science Corporation 

30 Forest Street 

Somerville, MA 02143 

NREL Technical Monitor: Cheryn Engebrecht

Prepared under Subcontract No. KNDJ-0-40337-03 

January 2014 
iv

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v 

Contents

List of Figures ............................................................................................................................................ vi

List of Tables ............................................................................................................................................. vii

Definitions................................................................................................................................................. viii

Executive Summary................................................................................................................................... ix

1 Problem Statement............................................................................................................................... 1

1.1 Introduction ..........................................................................................................................1

1.2 Background ..........................................................................................................................1

1.3 Relevance to Building America’s Goals ..............................................................................3

1.4 Cost Effectiveness ................................................................................................................3

1.5 Tradeoffs and Other Benefits ...............................................................................................5

2 Previous Research ............................................................................................................................... 6

3 Analysis and Test Method ................................................................................................................. 14

3.1 System Mechanics .............................................................................................................14

3.1.1 Numerical Analysis ................................................................................................15

3.1.2 Material Property Testing ......................................................................................18

3.1.3 Boundary Condition Testing ..................................................................................21

3.1.4 Discrete Load Component Testing ........................................................................23

3.2 Environmental Exposure ....................................................................................................25

3.2.1 Long-Term Movement (Creep)..............................................................................27

3.2.2 Diurnal Movement .................................................................................................29

4 Discussion........................................................................................................................................... 31

4.1 Previous Research ..............................................................................................................31

4.2 System Mechanics .............................................................................................................34

4.2.1 Moment Resistance of Fasteners ...........................................................................34

4.2.2 Strut and Tie Model (Low Friction).......................................................................36

4.2.3 Static Friction .........................................................................................................39

4.3 Environmental Exposure ....................................................................................................40

5 Conclusions and Recommendations ............................................................................................... 46

References................................................................................................................................................. 48

Appendix A: BEopt Simulation Graphs .................................................................................................. 50

Dallas, Texas ............................................................................................................................50

Kansas City, Missouri ..............................................................................................................51

Boston, Massachusetts .............................................................................................................52

Duluth, Minnesota....................................................................................................................53
vi 

List of Figures

Figure 1. Theorized forces providing vertical displacement resistance ............................................... 2

Figure 2. Typical ASTM D1761 test setup................................................................................................. 8

Figure 4. Short-term gravity load response test setup .......................................................................... 9

Figure 5. Short-term load versus deflection for assemblies with 4 in. of exterior insulation........... 10

Figure 6. Short-term load versus deflection for assemblies with 8 in. of exterior insulation........... 10

Figure 7. Long-term gravity load response test setup......................................................................... 11

Figure 8. Long-term load versus deflection for assemblies with 4 in. of exterior insulation ........... 12

Figure 9. Forces providing vertical displacement resistance .............................................................. 15

Figure 10. Free body diagram of bending resistance of a fastener.................................................... 16

Figure 11. Free body diagram of compression resistance of insulation ............................................ 17

Figure 12. Free body diagram of the friction resistance between material layers............................. 18

Figure 13. Measurement locations for fastener dimensions ................................................................ 19

Figure 14. Example of insulation compression Test B ......................................................................... 20

Figure 15. Example of a static COF test ................................................................................................. 21

Figure 16. Precompression force test setup.......................................................................................... 21

Figure 17. Relaxation of screw fastener precompression forces over a 2-day time period ............. 22

Figure 18. Example of a friction only system test ................................................................................. 23

Figure 19. Screenshot of data acquisition system output.................................................................... 24

Figure 20. Discrete load component testing of small scale wall assemblies with XPS .................... 25

Figure 21. Test wall assemblies under construction ............................................................................ 26

Figure 22. Exposed wall assemblies loaded to representative cladding weights.............................. 26

Figure 23. Lightweight cladding panel installed over the test wall assemblies ................................. 27

Figure 24. Deflection measurement location ......................................................................................... 27

Figure 25. Long-term environmental exposure of lightweight claddings ........................................... 28

Figure 26. Long-term environmental exposure of medium-weight claddings.................................... 28

Figure 27. Long-term environmental exposure of heavyweight claddings ........................................ 29

Figure 28. Diurnal movement of simulated medium-weight cladding system over XPS insulated 

wall assembly...................................................................................................................................... 30

Figure 29. Simple screw cantilever bending test................................................................................... 36

Figure 30. Measured versus calculated strut and tie load component resistance ............................ 38

Figure 31. Measured versus calculated friction load component resistance..................................... 39

Figure 32. Change of precompression forces over time ...................................................................... 40

Figure 33. Sample of ambient versus cladding cavity temperatures .................................................. 41

Figure 34. Proposed acceptable band of movement of furring over XPS insulation with a low load 

per fastener......................................................................................................................................... 42

Figure 35. Long-term deflection of a furring strip loaded to 30 lb/fastener in a stable environment43

Figure 36. Long-term deflection of a furring strip loaded to 30 lb/fastener in an exposed 

environment ........................................................................................................................................ 43

Figure 37. Furring strip movement recorded after the first day of “set in” ........................................ 45

Figure 38. Annualized energy related costs versus average source energy savings for Dallas...... 50

Figure 39. Average source energy savings reduction versus insulation level for Dallas................. 50

Figure 40. Annualized energy related costs versus average source energy savings for Kansas 

City ....................................................................................................................................................... 51

Figure 41. Average source energy savings reduction versus insulation level for Kansas City...... 51

Figure 42. Annualized energy-related costs versus average source energy savings for Boston .. 52

Figure 43. Average source energy savings reduction versus insulation level for Boston .............. 52

Figure 44. Annualized energy related costs versus average source energy savings for Duluth.... 53

Figure 45. Average source energy savings reduction versus insulation level for Duluth ................ 53

Unless otherwise noted, all figures were created by BSC. 
vii 

List of Tables

Table 1. Benchmark House Characteristics............................................................................................. 4

Table 2. Parametric Steps and Cost.......................................................................................................... 4

Table 3. Reference Cities............................................................................................................................ 5

Table 4. Yield Modes From AFPA TR-12................................................................................................... 6

Table 5. Materials Used in the Laboratory and Field Testing............................................................... 14

Table 6. Insulation Materials .................................................................................................................... 14

Table 7. Average Fastener Dimensions .................................................................................................. 19

Table 8. Screw Bending Yield Moment ................................................................................................... 19

Table 9. Elastic Range Modulus of Compression of Insulation Products .......................................... 20

Table 10. Static Friction Test Results ..................................................................................................... 21

Table 11. Furring Strip Precompression Forces.................................................................................... 22

Table 12. Environmental Exposure Test Panel Simulated Cladding Weights .................................... 26

Table 13. Mean Measured Load of 4-in. Insulation Assemblies at 0.015-in. and 0.125-in. Deflection32

Table 14. Maximum Allowable Insulation Thickness (in.) Excerpt From the NYSERDA/SFA Table . 32

Table 15. Calculated Cladding Load (lb/Fastener) Based on NYSERDA/SFA Table.......................... 32

Table 16. BSC 4 in. Test Results (Highlighted in Orange) at 0.015-in. Deflection Limit and 

1.5 Divisor Applied Compared to NYSERDA/SFA Table ................................................................ 33

Table 17. BSC 4 in. Test Results (Highlighted in Orange) at 0.125 in. Deflection Limit and 1.5 

Divisor Applied Compared to NYSERDA/SFA Table....................................................................... 33

Table 18. Predicted Load Resistance Component of Screw Bending................................................. 34

Table 19. Effects of Beam Length on the Predicted Load Resistance Component of Screw Bending35

Table 20. Predicted Versus Measured Screw Bending Capacity (Assuming Bending at Face of 

Framing) .............................................................................................................................................. 35

Table 21. Predicted Versus Measured Screw Bending Capacity (Assuming Bending at 1⁄4 in. Into 

Framing) .............................................................................................................................................. 35

Table 22. Predicted Versus Measured Screw Bending Capacity (Assuming Bending at 1⁄2 in. Into 

Framing) .............................................................................................................................................. 36

Table 23. Elastic Range Compression Modulus of Insulation.............................................................. 37

Table 24. Modeled Compression Strut Capacity for XPS insulation ................................................... 38

Table 25. Coefficients of Thermal Expansion ........................................................................................ 41

Unless otherwise noted, all tables were created by BSC.
viii

Definitions

BSC Building Science Corporation 

COF Coefficient of friction

DOE U.S. Department of Energy 

EPS Expanded polystyrene 

FPL Forest Products Laboratory 

FSC Foam Sheathing Coalition

LVDT Linear voltage distance transducer

MF Mineral fiber

NYSERDA New York State Energy Research and Development Authority

o.c. On center

OSB Oriented strand board 

PIC Polyisocyanurate 

psf Pounds per square foot 

RH Relative humidity

SFA Steel Framing Alliance

SPF Spray polyurethane foam

XPS Extruded polystyrene 
ix

Executive Summary

The addition of insulation to the exterior of buildings is an effective means of increasing the 

thermal resistance of both wood framed walls as well as mass masonry wall assemblies. For 

thick layers of exterior insulation (levels > 1.5 in.), the use of wood furring strips attached 

through the insulation back to the structure has been used by many contractors and designers 

as a means to provide a convenient cladding attachment location (Straube and Smegal 2009; 

Pettit 2009; Joyce 2009; Ueno 2010). 

The research presented in this report is intended to help develop a better understanding of the 

system mechanics involved and the potential for environmental exposure induced movement 

between the furring strip and the framing. Building Science Corporation sought to address the 

following research questions: 

1. What are the relative roles of the mechanisms and the magnitudes of the force that 

influence the vertical displacement resistance of the system?

2. Can the capacity at a specified deflection be reliably calculated using mechanics based 

equations? 

3. What are the impacts of environmental exposure on the vertical displacement of furring 

strips attached directly through insulation back to a wood structure? 

The system mechanics portion of the research examined some of the discrete load components 

that help develop the vertical load resistance capacity of furring strips attached directly through 

insulation back to a wood structure. It was theorized that the capacity of the system is developed 

from several sources, including the moment resistance of the fasteners (including both bending 

strength of the fastener and the bearing strength of the furring and framing members), the 

compressive strength of the rigid insulation, as well as the static friction between the layers.

The system mechanics research provided some useful insights into the magnitude of the various 

load components, even if many of the exact mechanisms cannot be accurately predicted. The 

research was designed to focus on three mechanisms for resisting vertical gravity loads: (1) 

screw bending; (2) friction; and (3) a strut and tie effect. The bending capacities of the screw 

fasteners were noted to contribute a much lower amount to the system total vertical deflection 

resistance capacity when compared to the other studied mechanisms. From the results it appears 

that friction forces in the assembly may be significant, particularly at initial and small vertical 

deflections. While the presence of friction in the assembly may be significant, there is not 

enough information yet available to determine how to best account for, and make use of, the 

friction in the assemblies from a design perspective. The amount of friction due to 

precompression1

 can be quite variable, as measured precompression forces were noted to change 

dramatically over time and with changing environmental conditions. The strut and tie model was 

demonstrated to provide additional capacity; however, the results were not clear, as other 

unanticipated factors appear to affecting the total capacity.

 

1 Precompression forces are the clamping forces generated in the assembly by attachment of the furring strips with 

the screw fasteners.
x 

It was found that the theorized load components that were modeled do not provide a sufficiently 

accurate prediction of the measured load components to be used in a reliable design model. 

There were several factors to this, including sensitivity of the inputs, potential changes to the 

load resistance model depending on the amount of deflection, variability in the boundary 

conditions, as well as some additional system effects that are still not understood, identified, or 

quantified. Further study of the load component mechanisms may help to further refine our 

understanding and help us develop more accurate models that could be used for assembly design. 

The second part of the testing work completed was a study on the impacts of climate exposure on 

the vertical movement of furring strips attached over exterior insulation. A total of 12 assemblies 

were constructed (four different insulation types loaded to three different levels, 8 lb/fastener, 15 

lb/fastener, 30 lb/fastener) in an outdoor exposed environment. Vertical deflection movements of 

the furring strip with respect to the framing were measured at various intervals between July 

2012 and September 2012. 

The results of the long-term exposure tests reinforced much of the industry experience with this 

approach to cladding attachment over 4 in. of exterior insulation. Lightweight claddings (such as 

wood, fiber cement, and vinyl siding) represent the majority of the cladding that has been, and is 

currently being used with this type of attachment system. These claddings coupled with a 

fastener spacing of 16–24 in. o.c. are representative of low load per fastener assemblies. To date, 

no known problems have occurred with these systems. The low measured movement and 

apparent resistance to creep is in line with this experience.

For heavier claddings such as traditional stucco and adhered stone veneers, the per-fastener load 

would be expected to be higher. Under medium load (15 lb/fastener), assemblies installed over 4 

in. of insulation seem to be demonstrating good performance, though more data are 

recommended to be collected. Under heavy load (30 lb/fastener), there appears to be a potential 

for long-term creep of the assemblies. More study is needed for these assemblies.

All of the test assemblies had notable movement within a range of deflections. With a daily 

movements on the order of ± 1/64 in. to ± 1/32 in. being measured for one of the assemblies. In 

service deflection limits for the assemblies should be set to account for this movement. 
1 

1 Problem Statement

1.1 Introduction

The addition of insulation to the exterior of buildings is an effective means of increasing the 

thermal resistance of both wood-framed walls as well as mass masonry wall assemblies. The 

location of the insulation to the exterior of the structure has many direct benefits, including better 

effective R-value from reduced thermal bridging, better condensation resistance, reduced thermal 

stress on the structure, as well as other commonly associated improvements such as increased 

airtightness and improved water management (Hutcheon 1964; Lstiburek 2007). 

For thick layers of exterior insulation (levels > 1.5 in.), the use of wood furring strips attached 

through the insulation back to the structure has been used by many contractors and designers 

as a means to provide a convenient cladding attachment location (Straube and Smegal 2009; 

Pettit 2009; Joyce 2009; Ueno 2010). 

While the approach has been demonstrated to be effective, there is significant resistance to its 

widespread implementation due to a lack of research and understanding of the mechanisms 

involved in the development of the vertical displacement resistance capacity. In addition, the 

long term in service performance of the system has been questioned due to potential creep effects 

of the assembly under the sustained dead load of the cladding and the effects of varying 

environmental conditions. 

1.2 Background

The residential building sector consumes approximately 21% of the primary energy used in the 

United States (DOE/EIA 2008). While new code standards are pushing for more energy-efficient 

buildings, there are a significant amount of existing buildings that are in great need of energy 

retrofits. In the past, retrofits of existing residential buildings typically involved the filling of 

framed cavity walls with insulation; however, the amount of effective thermal resistance that 

could be added was limited by the existing stud cavity depth (wood-framed walls) or strapping 

depth (common for mass masonry walls), the insulation material used (commonly fiberglass/

mineral fiber or cellulose), and the amount of thermal bridging present from the wood framing. 

The addition of insulation to the exterior of existing buildings has been demonstrated to be an 

effective means to overcome these limitations and provide higher effective R-values for building 

wall assemblies. The benefits of this approach extend beyond just added thermal resistance; 

benefits of increased building durability and airtightness are often also realized.

The use of exterior insulation has been common practice for many decades on buildings, 

particularly behind brick or masonry veneer claddings. Exterior insulation and finish systems use 

exterior insulation as a composite cladding assembly, providing the support structure for a 

lightweight finish coat.

Building Science Corporation (BSC) has been at the forefront of using exterior insulation 

approaches on residential buildings for several decades. The use of furring strips as the primary 

cladding attachment location is a strategy that has been used on many private as well as Building 

America supported projects (Pettit 2009; Ueno 2010). 
2 

The push for lower energy buildings has resulted in an increase of projects that are looking to use 

thick layers (> 1.5 inches) of exterior insulation on their buildings. This increase resulted in an 

increase in questions regarding the effectiveness of using furring strips attached through the 

insulation back to the structure. 

In reaction to this, research into the performance of these systems has been funded by several 

groups such as the Foam Sheathing Coalition (FSC), the New York State Energy Research and 

Development Authority (NYSERDA), and the Steel Framing Alliance (SFA). The focus of this 

past research by FSC and NYSERDA/SFA was to try to develop prescriptive code tables for 

attaching cladding to framing over continuous insulation (Bowles 2010). The tables that were 

developed used an initial deflection limit of 0.015 in. as a basis for design. By limiting the initial

deflection to 0.015 in., the intent was to keep long-term deflection due to potential creep of the 

system within acceptable limits, though these acceptable limits were not defined.

Research conducted by BSC was aimed at expanding on this previous research to include several 

types of insulation as well as to examine both short-term (initial loading) and long-term 

(sustained loading) performance of the system (Baker 2013).

While short-term capacities of the system were measured, the understanding of the system 

mechanics that help to develop the capacities were not well identified. The capacity of the 

system was theorized to be developed from several sources, including the bending strength of the 

fastener, the bearing strength of the furring and framing members, the compressive strength of 

the rigid insulation, as well as the static friction between the layers.

Shear and rotational 

resistance provided by 

fastener to wood 

connections

Rotational resistance 

provided by tension in 

fastener and 

compression of the 

insulation

Vertical movement 

resistance provided by 

friction between layers

Figure 1. Theorized forces providing vertical displacement resistance

The long-term (sustained loading) tests that were completed were intended to examine the longterm creep effects of the system under sustained gravity load in relatively stable environmental 

conditions. The results of the initial testing raised some questions as to the impacts of changing 

environmental conditions on the long-term performance of the system. The analysis of the data 

did not show much movement in the systems over the course of the test period (July 2011 

through January 2012); however, the movement that was noted seemed to indicate that the 

system deflections were influenced by even small changes in environmental conditions, and that 
3 

these changes may have greater impacts on the vertical movement of the furring strips than the 

effects of sustained gravity dead loads imposed by the cladding. 

The research presented in this report is intended to help develop a better understanding of the 

system mechanics involved and the potential for environmental exposure to induce movement in 

the system. BSC sought to address the following research questions: 

1. What are the relative roles of the mechanisms and the magnitudes of the force that 

influence the vertical displacement resistance of the system?

2. Can the capacity at a specified deflection be reliably calculated using mechanics based 

equations? 

3. What are the impacts of environmental exposure on the vertical displacement of furring 

strips attached directly through insulation back to a wood structure?

1.3 Relevance to Building America’s Goals

The use of exterior insulation on wall assemblies is an effective means to provide additional 

thermal resistance to enclosure assemblies. The technique is particularly well suited to retrofit 

projects that might otherwise be limited (in terms of space conditioning energy use reductions) 

due to existing construction dimensional constraints. This fits directly into the Building America 

goals of substantial reductions in energy consumption. While the energy benefits are apparent 

and easy to understand, the practical implementation has run into barriers that have slowed 

widespread adoption.

The results of the research is intended to provide specific guidance for cladding attachment over 

thicker layers of exterior insulation and evaluate the potential of developing mechanics based 

equations as a means of evaluation and design of the system capacities. It will also provide some 

preliminary data on field performance of the systems. The applicability of this research will 

extend to all climate zones and housing types. 

1.4 Cost Effectiveness

In most circumstances, the exterior retrofit of a home with exterior insulation comes as part of a 

larger scope of work for a building retrofit. The choice to add exterior insulation is usually 

triggered by a need (or desire) to reclad or overclad the building. The driving force behind 

installing new cladding can be from any number of sources, including existing water 

management problems, comfort or durability concerns, end of service life for the cladding, or 

aesthetic concerns. The need to replace the cladding provides an opportunity for the designer or 

contractor to include exterior insulation as a means to increase the energy performance of the 

building at the same time. The cost effectiveness of this from an energy perspective is therefore 

dependent on the cost of the insulation as well any associated components above and beyond 

new cladding installation.

A preliminary evaluation was completed looking at the incremental cost of the varying 

thicknesses of insulation installed to the exterior of the wall assemblies. This preliminary cost 

analysis used foil-faced polyisocyanurate (PIC) as the baseline exterior insulation. Cost data for 

the exterior insulation was taken from RS Means Construction Data (Reed Construction Data 

2011). Costs included in the analysis were the installed cost of the insulation material, 1 × 3 
4 

wood furring strips spaced at 16 in. o.c., and wood screws spaced at 24 in. o.c. vertically for the 

attachment of the furring back to the structure. A cost markup of $100/window in the reference 

model was used as an estimate of the additional cost for trim extensions that would be needed to 

account for the additional thickness of foam added to the exterior of the home. This value is an 

estimate, as actual costs can be highly variable due to the many different design choices available 

for window placement, exterior window trim design, and attachment. 

Other items such as house wrap or sheathing tape, self-adhered membrane flashings, metal 

flashings, siding, and siding fasteners were omitted from the analysis, as these items are 

associated with recladding and water management, and would be part of the retrofit project 

regardless of the addition of exterior insulation. 

Simulations were run using BEopt simulation software developed by the National Renewable 

Energy Laboratory. An example home was used as the baseline to help demonstrate the benefits 

of using exterior insulation as part of a house energy retrofit. This benchmark home was assumed 

to be around 1950s era two-story slab-on-grade construction and had the following basic 

characteristics (Table 1). 

Table 1. Benchmark House Characteristics

House Characteristics Square Footage

Finished Floor Area 2312

Ceiling Area 1156

Slab Area 1156

Wall Area 2799

Window Area 410 (17.7% glazing ratio)

The wall conductance performance was isolated from all other aspects of the home, to examine 

the effectiveness of this single strategy. Given the assumed age of the home, the benchmark 

home had an uninsulated wall cavity (as per guidance from the 2011 Building America

Benchmark Protocol). The following parametrics were run to see the effectiveness of the added 

thermal resistance to the energy performance and utility cost (Table 2). The analysis assumed 

that the cost of the measure is financed over a 5-year period at a 7% interest rate. An additional 

fuel escalation rate of 2% was also included in the analysis. 

Table 2. Parametric Steps and Cost

Parametric Step Cost/ft2

Benchmark (Uninsulated 2 × 4 Wall) N/A

R-13 Cavity Fill Insulation $2.20

R-13 Cavity Fill + 1 in. Exterior Insulation (R-6.5) $3.55

R-13 Cavity Fill + 1.5 in. Exterior Insulation (R- 9.75) $3.76

R-13 Cavity Fill + 2 In. Exterior Insulation (R-13) + 

1 × 4 Wood Furring

$5.73

R-13 Cavity Fill + 2 Layers of 1.5-in. Exterior Insulation 

(R-19.5) + 1 × 4 Wood Furring

$7.19

R-13 Cavity Fill + 2 Layers of 2-In. Exterior Insulation 

(R-26) + 1 × 4 Wood Furring

$7.58
5 

Simulations were run for the following cities (Table 3): 

Table 3. Reference Cities

City Climate Zone

Dallas, Texas 3A

Kansas City, Missouri 4A

Boston, Massachusetts 5A

Duluth, Minnesota 7A

Results indicated that for cold climate zones (4 and higher), insulation up to 1.5 in. was shown to 

be a cost-optimized solution. This was mainly due to this being the tipping point before which 

additional costs associated with the furring strips and additional screw fasteners required for 

cladding attachment needed to be added to the system. Insulation thickness above 2 in. was still 

demonstrated to be cost neutral as part of this simplified analysis in all cities except for Dallas. 

While the analysis run focused on conductance improvements only, there is some argument to be 

made that the addition of exterior insulation would likely also improve the overall airtightness of 

the assemblies as well (Ueno 2010). The benefits from increased airtightness are known to be 

very important in cold climate construction; however, it is also more difficult to isolate and 

apportion to individual measures. 

1.5 Tradeoffs and Other Benefits

Using exterior insulation has many additional benefits other than simply increased thermal 

resistance. The single largest benefit is the increased condensation resistance that this strategy 

provides for cold climate buildings. The placement of the insulation to the exterior of the 

building acts to keep all of the structural elements at a much more even temperature throughout 

the year, reducing the risk of interstitial condensation. For wood structures, this can significantly 

reduce the potential for wood decay; an added benefit is that the seasonal thermal and moisture 

variations of the wood frame are greatly reduced. In masonry buildings, the potential for freezethaw is practically eliminated, since this approach not only keeps the masonry warmer, but also 

addresses the exterior rainwater absorption into the masonry (which is the leading moisture 

source related to freeze-thaw damage to buildings). 

In addition to keeping the structure warm and preventing condensation, the use of the furring 

strips creates a significant upgrade in water management. The increase in drainage and drying 

that is provided by the 3⁄4-in. gap created by the furring strips provides so much additional 

protection against water infiltration problems (Lstiburek 2009) that the use of a drainage gap is a 

base recommendation for most cladding installations regardless of whether or not exterior 

insulation is used. The fact that the furring strips are an intrinsic component of this system 

provides a significant added benefit to the long-term durability of these wall assemblies. 
6 

2 Previous Research

Several groups such as the FSC, NYSERDA, and SFA have funded research into the vertical 

load capacity of furring strips, installed over exterior insulation, that are fastened back to a wood 

or steel structure. The primary goal of this past research by FSC and NYSERDA/SFA was to 

develop prescriptive code tables for attaching cladding to framing over continuous insulation 

(Bowles 2010). The research methodology adopted used wood joint connection theory as the 

basis for the analysis, in particular the European Yield Theory that examined the performance of 

gapped wood to wood connections. 

The European Yield Theory (first conceived in the 1940s) is based on an equilibrium of forces 

caused by rotation of fasteners in wood members; this theory predicts performance of the 

connection at the point where yielding of materials (wood or fastener) has developed. The 

equations as set out in the American Forest and Paper Association (AFPA) Technical Report 12 

General Dowel Equations for Calculating Lateral Connection Values predict performance of a 

multitude of failure modes, with the governing mode being the one with the lowest yield 

capacity. A visual representation of the potential failure modes (AFPA 1999) is included in 

Table 4. 

Table 4. Yield Modes From AFPA TR-12

Yield Mode Description Graphic

Im Main member bearing failure

Is Side member bearing

II Side and main member bearing
7 

Yield Mode Description Graphic

IIIm Main member bearing and dowel yielding in the 

side member

IIIs Side member bearing and dowel yielding in the 

main member

IV Dowel yielding in the side and main member

Research conducted by the U.S. Department of Agriculture Forest Products Laboratory (FPL) 

verified the yield equations with empirical data for gaps up to 1 in. (Aune and Patton-Mallory 

1986a, 1986b). For small gaps (0.009–0.04 in.) an air gap was used or a slip sheet such as 

polyethylene was placed between the two wood members to remove friction and other forces 

from the assembly. For larger gaps (1⁄2 in. and 1 in.), the void was filled with expanded 

polystyrene (EPS) insulation. Effects of friction and compression resistance of the insulation 

were captured with the tests that included the insulation. These tests that were conducted with 

EPS insulation, gave way to the idea of possibly adopting the yield equations for the application 

of wood furring strips over exterior insulation. 

The NYSERDA research was focused on the specific application exterior insulation for abovegrade wall construction. To accomplish this, a representative matrix of construction strategies

that could be used to install cladding over exterior insulation (either from direct attachment or 

through the use of furring strips attached back through insulation to the structure), was tested 

with the intent of calibrating the measured results of the testing to the predicted results of the 

yield equations. 
8 

The testing was conducted following ASTM 

D1761 - 06 Standard Test Methods for 

Mechanical Fasteners in Wood (Figure 2) and 

examined:

• Main member (studs/substrate)

o 2 x studs (spray polyurethane 

foam [SPF] worst case density)

o 33 mil (20 g) and 54 mil (16 g) 

studs 

o 7/16 in. and 3⁄4 in. oriented 

strand board (OSB) 

• Side member (furring/cladding)

o 3⁄4 in. and 3⁄8 in. pine 

o 33 mil (20 g) steel hat channel

• Fasteners

o Nails

o Wood screws 

o Lag screws

o Self-drilling/tapping screws 

(steel connections). 

Figure 2. Typical ASTM D1761 test setup

For wood frame test specimens, the measured data were compared to the predicted performance 

of the yield equations as determined by the TR-12 (and calculated based on actual properties of 

the materials used in the testing). This research concluded that the 5% offset yield prediction as 

calculated using the TR-12 formulas, resulted in a reasonably accurate prediction of the shear 

load at a deflection of 0.015 in. While there was no mathematical connection between these 

values, the research team considered this to be an adequate basis for designing for a 0.015-in. 

deflection limit given the scope of the research. In addition, a divisor of 1.5 was applied to the 

calculated results to address potential concerns of assembly creep under sustained loads. The 

methodology was used to develop prescriptive code tables for attaching furring strips to framing 

over continuous insulation (Bowles 2010). 

In 2011, research conducted by BSC under the Building America Program examined both shortterm loading as well as long-term loading of wall assemblies using furring strips fastened back 

through the insulation as the primary cladding support structure (Baker 2013). The research was 

aimed at answering three key questions: 

1. Is there a difference in performance given varying exterior insulation types? 

2. How does the performance compare for large thicknesses of insulation (4 in. and 8 in.)? 

3. What are the impacts of sustained loading (creep) under relatively stable environmental 

conditions? 
9 

The test plan differed in several ways from previous work conducted in that it used full scale 

wall assemblies in lieu of small scale samples. The decision to conduct the test on full scale 

samples was to reduce the impacts of installation variability in construction and capture the 

system effects of the assembly. The test plan also expanded upon previous testing (which had 

typically been limited to EPS insulation only) to include multiple insulation types:

• EPS

• Extruded polystyrene (XPS) 

• Foil-faced PIC

• Rigid mineral fiber (MF). 

The test plan was separated into two distinct sections: a “short-term” or initial loading section, 

and a “long-term” or sustained loading section. 

The short-term load testing was conducted on 4

× 8 wall panels with two 1 × 3 wood furring 

strips spaced 24 in. o.c. (Figure 4). The furring 

strips were attached with #10 wood screws 

spaced 16 in. o.c. vertically (14 total). Tests 

were conducted at both 4-in. thickness of 

insulation as well as at 8-in. thickness. The 

testing was conducted by applying a specific 

load to the furring strips via a hydraulic ram 

connected to a metal angle that spanned 

between the bottoms of the two furring strips.

The corresponding vertical deflection 

measurements were recorded at each furring 

strip for each load increment. The 

measurements were taken between the stud 

framing and the furring so that just the 

deflection of the furring with respect to the 

framing was captured. The resulting load 

deflection plots can be seen in Figure 5 and 

Figure 6. Figure 3. Short-term gravity load 

response test setup
10 

Figure 4. Short-term load versus deflection for assemblies with 4 in. of exterior insulation 

(Baker 2013) 

Figure 5. Short-term load versus deflection for assemblies with 8 in. of exterior insulation 

(Baker 2013) 

1/64" 1/32" 1/16" 1/8"

0

5

10

15

20

25

30

35

40

45

50

55

60

65

70

75

0 0.015 0.03 0.045 0.06 0.075 0.09 0.105 0.12 0.135 0.15 0.165 0.18 0.195 0.21 0.225 0.24

Load vs. Vertical Displacement of Furring Strip over 4" of Insulation 

4" MF 4" EPS 4" PIC 4" XPS

deflection (in)

Load (lbs per fastener)

MF PIC

EPS XPS

1/64" 1/32" 1/16" 1/8"

0

5

10

15

20

25

30

35

40

45

50

55

60

65

70

75

0 0.015 0.03 0.045 0.06 0.075 0.09 0.105 0.12 0.135 0.15 0.165 0.18 0.195 0.21 0.225 0.24

Load vs. Vertical Displacement of Furring Strip over 8" of Insulation 

8" MF 8" EPS 8" PIC 8" XPS

deflection (in)

Load (lbs per fastener)

EPS

XPS

MF

PIC
11 

The long-term vertical deflection testing was 

conducted on 2 × 8 wall panels with a single 1

× 3 wood furring (Figure 7). The furring strip 

was attached with #10 wood screws spaced 16 

in. o.c. vertically (7 total). Tests were 

conducted at 4-in. thickness of insulation only. 

The test panels were loaded to a total load of 

210 lb using dead weights hung from the 

bottom of the furring strip. This load level 

provides the following equivalent loads: 

• 13 psf (if furring is spaced at 24 in. 

o.c.) 

• 20 psf (if furring is spaced at 16 in. 

o.c.) 

• 30 lb/fastener. 

A fifth test panel was constructed using XPS as 

the insulation material installed to a 4-in. 

thickness. The intent was to examine the 

relative impact of a reduced load magnitude.

This reference panel was loaded to a total load 

of 60 lb. This load level provides the following 

equivalent loads: 

Figure 6. Long-term gravity load 

response test setup

• 3.75 psf (if furring is spaced at 24 in. o.c.) 

• 5 psf (if furring is spaced at 16 in. o.c.) 

• 8.6 lb/fastener. 

The measured deflections were plotted as a function of time (Figure 8), with a positive deflection 

on the graph indicating a downward movement of the furring strip. 

The results of the testing indicated that for low per fastener loads (< 10 lb) very little movement 

both under initial loading and long-term loading was observed (~1/200 in.). This would be 

representative of lightweight claddings (< 5 psf) installed on a fastener spacing up to 24 in. o.c. 

vertically and horizontally, or medium-weight claddings (claddings around 10 psf) installed on a 

more dense fastener spacing of 12 in. o.c. both horizontally and vertically. 

For higher per-fastener loads (30 lb), initial deflection was observed to be < 1/32 in. of 

assemblies with 4 in. of exterior insulation and 1/16 in. for assemblies with 8 in. of exterior 

insulation. Long-term loading at this magnitude also demonstrated very stable performance in 

the laboratory environment. Most assemblies did not record a deflection > 1/32 in., with the 

exception of the PIC sample which had a deflection of approximately 3/32 in. It should be noted 

that the PIC sample was bumped on a few occasions during the test due to the proximity of other 

test equipment near the long-term test setup. Two jumps in the data were correlated back to the 

noted disturbances. Given the very small deflections recorded and the sensitivity of the 
12 

measurement equipment used, the data gathered for the PIC setup may not be providing as clear 

an indication of the performance of the PIC product. 

Figure 7. Long-term load versus deflection for assemblies with 4 in. of exterior insulation 

(Baker 2013) 

In the early stages of the testing (the first 3 weeks after the initial loading), very minor additional 

downward vertical movement was seen. The temperature and relative humidity (RH), however, 

were maintained at a more stable range. In all cases a very slight trend for additional deflection 

can be seen. The magnitude, though, was on the order of 1/400 in., and it might not result from 

creep effects from sustained loading. More substantial movement seemed to occur shortly after 

the first 3 weeks, when the temperature in the laboratory increased slightly (by approximately 

5°F) and the RH dropped (from approximately 55% RH to 40% RH). Movements on the order of 

1/100 in. were observed. 

Looking at the complete dataset, a slight trend in the movement appeared to result from 

fluctuations in the temperature and RH. The temperature in the laboratory space fluctuated 

between 60°F and 75°F and the RH fluctuated between 60% and 30% over the course of the 

testing. Deflection movement in the test setups seems to track to these environmental changes. A 

drop in the RH results in a general trend of an increase in the vertical downward deflection of the 

furring strips. It was interesting to note that the converse is true as well. An increase in the RH 

seems to correspond to an upward vertical movement of the furring strips. This was true for all 

insulations except for EPS. The movement of the EPS test panel demonstrated a reverse trend, 

where a drop in the RH resulted in an upward vertical movement of the furring strips. 

1/64"

1/32"

1/16"

1/8"

0

10

20

30

40

50

60

70

80

90

-0.03

0.00

0.03

0.06

0.09

0.12

0.15

Jul/11 Aug/11 Sep/11 Oct/11 Nov/11 Dec/11 Jan/12

Temperature (F) and Relative Humidity (%)

Deflection (in)

Time

Vertical deflection movment of a wood furring strip (loaded to 

30lbs/fastener) over time

XPS (8.6 lbs/fastener) XPS (30 lbs/fastener) EPS (30 lbs/fastener) PIC (30lbs/fastener)

MF (30 lbs/fastener) Temperature (F) RH (%)

Jump in deflection readings coincided with 

construction activities near the PIC test set up

PIC

MF XPS

XPS (8.6 lbs/fastener) EPS
13 

The test conducted at 5 psf on the XPS sample demonstrated very stable performance with 

almost no movement seen in the sample, even with changing temperature and RH.

From the test data, it was difficult to differentiate movements of the samples that result from 

prolonged loading (creep) or from environmental changes. Both positive as well as negative 

movements were noted. The movements from environmental changes were most likely caused 

by material expansion and contraction from moisture adsorption or thermal changes. Given the 

limited testing, the magnitude of this effect cannot be predicted at this point. In addition, material 

property changes may affect performance over the range of actual in-service temperatures. This 

was not accounted for in the testing. Additional testing of exterior samples exposed to a variety 

of temperature and humidity conditions is recommended. 
14 

3 Analysis and Test Method

There are notable gaps in the general understanding of how the vertical load resistance is actually 

developed in the assembly. System mechanics (such as fastener bending/bearing, insulation 

compression, and friction forces between layers) were theorized as impacting the system 

capacity, but had not been measured or quantified. It was felt that understanding the factors that 

affect the development of system capacity was going to be important in order to examine means 

to engineer the attachment system. In addition, the potential for creep of the system was still not 

well understood or quantified. Several conceivable sources of creep in the system could be 

identified, such as expansion and contraction of wood, expansion and contraction of insulation, 

relaxation of wood fibers, and plastic deformation of insulation. From the observed sensitivity of 

the assemblies to slight environmental changes experienced in relatively stable environments, 

and since many of the conceivable sources are affected by temperature and RH, it was felt that 

the performance of these systems needed to be examined in exposed environments. 

The research plan intended to help develop a better understanding of the mechanics involved and 

the potential for environmental exposure to induce unpredicted movement in the system. The 

testing was conducted using the following materials in order to limit variables in the research

(Table 5 and Table 6). 

Table 5. Materials Used in the Laboratory and Field Testing

Component Material

Framing 2 × 4 SPF standard wood framing 

Sheathing 7/16 in. OSB

Building Wrap Dupont Tyvek Building Wrap 

Insulation 4 in. thick rigid insulating sheathing (2 layers 

of 2 in. various types listed below)

Furring Strips Nominal 1 × 3 SPF utility-grade lumber 

Screws 6-in. long standard #10 pan head wood screws

Table 6. Insulation Materials

Insulation Type Product Brand

Type II EPS Plastispan Plastifab

Type IV XPS C-200 Owens Corning 

Foil Faced PIC Thermax CI DOW Chemical

Rigid MF RB80 Roxul 

3.1 System Mechanics

The system mechanics portion of the research examined some of the discrete load components 

that help develop the vertical load resistance capacity of furring strips attached directly through 

insulation back to a wood structure. The intent was to gain a better understanding of the system 
15 

mechanisms to help guide the development of design tools or recommendations for the 

attachment of the furring strips. 

It was theorized that the capacity of the system is developed from several sources, including the 

moment resistance of the fasteners (including both bending strength of the fastener and the 

bearing strength of the furring and framing members), the compressive strength of the rigid 

insulation, as well as static friction between layers (a simple diagram of these mechanisms is 

presented in Figure 9). 

Shear and rotational 

resistance provided by 

fastener to wood 

connections

Rotational resistance 

provided by tension in 

fastener and 

compression of the 

insulation

Vertical movement 

resistance provided by 

friction between layers

Figure 8. Forces providing vertical displacement resistance

3.1.1 Numerical Analysis

A series of mechanics-based equations were proposed to be evaluated as a starting point to 

examine the potential of developing a predictive model for the measured capacity of the 

assemblies at specific deflection magnitudes. This is a departure from past research and general 

wood connection design that looked to a strength limit state design where capacity, not 

movement, was considered. 
16 

3.1.1.1 Moment Resistance of 

Fasteners

The vertical displacement resistance due to 

the bending of the fasteners was theorized 

to function similarly to either a simple 

cantilevered beam or a beam that is fixed at 

one end and free and guided on the other 

end. The magnitude of the vertical load 

component at a given deflection would 

therefore be determined based on the 

following equations: 

1. Cantilevered beam

P =

∆3EI

x3

2. Beam that is fixed at one end and 

guided at the other 

P =

∆12EI

x3

Where:

P = applied vertical load

∆ = vertical displacement

x = distance between structure and 

furring

E = modulus of elasticity of the screw

I = area moment of inertia Figure 9. Free body diagram of bending 

resistance of a fastener
17 

3.1.1.2 Strut and Tie Model (No 

Friction)

The vertical displacement resistance due to 

the compression strut developed by the 

furring bearing on the insulation is theorized 

to be based on the compression modulus of 

the insulation. The magnitude of the vertical 

load component at a given deflection would 

therefore be determined using the following 

equation: 

P = Ft sin θ = Fc tan θ

Where:

P = applied vertical load

Ft = tension force in the fastener

Fc = compression force in the 

insulation

And the compression force in the insulation 

determined using the following equation: 

Fc = δKinsulAt

Where:

δ = deflection normal to the insulation

Kinsul = elastic range compression modulus 

of the insulation 

At = tributary area of the compression 

strut component 

Figure 10. Free body diagram of compression 

resistance of insulation
18 

3.1.1.3 Static Friction Model

The vertical static friction capacity is 

theorized to be based on the friction 

resistance of the various layers due to a 

normal compressive force applied to the 

system from precompression (clamping 

force of the furring during attachment) and 

postcompression (compression strut 

developed by the rotation of the fastener). 

The magnitude of the vertical load 

component for the static friction could 

therefore be determined using the 

following equation: 

P = fr = μ(Fc + Fp)

where:

fr = friction resistance between the 

materials

μ = coefficient of friction (COF)

Fc = normal force to the materials due 

to the compression strut 

Fp = normal force to the materials due 

to precompression forces

Figure 11. Free body diagram of the friction 

resistance between material layers

These equations were used to develop an analysis spreadsheet to compare predicted performance 

to measured performance.

3.1.2 Material Property Testing

Certain key material properties were measured as part of the testing. The values were used as

inputs for the mechanics based analysis spreadsheet. The following material properties were 

measured:

1. Screw bending yield strength

2. Elastic compression modulus of insulation 

3. Static COFs. 

3.1.2.1 Screw Bending Yield Strength

The screw bending strengths were determined following ASTM F1575 - 03(2008) Standard Test 

Method for Determining Bending Yield Moment of Nails (ASTM 2008). A sample of five 

fasteners, used for the testing, was taken from the quantity used in the rest of the laboratory and 

field testing. The fastener dimensions were measured and recorded prior to testing (Figure 13). 

The results are displayed in Table 7 and Table 8 below. 
19 

Figure 12. Measurement locations for fastener dimensions

Table 7. Average Fastener Dimensions

Five Specimen Average 

(in.)

Root

Diameter

Thread

Diameter Length ShankDiameter

Head 

Diameter

0.1315 0.193 6 0.1455 0.3505

Table 8. Screw Bending Yield Moment

Specimen 

No.

Yield Load P 

(lbf)

Yield Strength F

(psi)

1 210 207,838

2 201 198,930

3 223 220,705

4 207 204,869

5 201 198,931

Average 208.4 206,255

 

3.1.2.2 Elastic Compression Modulus of Insulation

The elastic range compressive modulus of the insulation materials was determined following the 

guidelines set out by ASTM D1621 - 10 Standard Test Method for Compressive Properties of 

Rigid Cellular Plastics (ASTM 2010). The testing was conducted on 6 × 6-in. samples of 2-in. 

thick insulation products. 

Two sets of tests were completed (Table 9); the first following the standard test protocol that 

loads the entire surface of the insulation evenly (Test A), and a second modified protocol that 

point loaded a 6-in. long section of 1 × 3 furring installed on top of the insulation sample (Test

B). The intent of Test B set was to better approximate the performance of the insulation in the 

specific application being researched and to see if there was a discernible edge effect (load 

spreading) resulting from the insulation being loaded under the wood furring strip. 
20 

Table 9. Elastic Range Modulus of Compression of Insulation Products

Material

Elastic Range Modulus

(psi)

Test A Test B

EPS 250 269

XPS 353 368

MF 16 28

PIC 429 350

Figure 13. Example of insulation compression Test B 

3.1.2.3 Coefficients of Static Friction

The static COFs for each insulation material to building wrap interfaces were determined

following the guidelines set out in ASTM D1894 - 11e1 Standard Test Method for Static and 

Kinetic Coefficients of Friction of Plastic Film and Sheeting (ASTM 2011) (Table 10). A 

minimum of four tests were completed for each interface. The testing was conducted using a 5-lb 

normal force being applied to a material sled pulled along the surface interface. The peak load 

was measured during each test to capture the static COFs. 
21 

Table 10. Static Friction Test Results

Static COF

EPS 0.27

XPS 0.23

MF 0.45

PIC 0.26

Figure 14. Example of a static COF test

3.1.3 Boundary Condition Testing

The installation of the wood furring strips to the wood structure results in a compression force 

being applied to the insulation prior to any vertical load being applied to the furring strips. This 

precompression force has an impact on the friction resistance of the assembly. From past testing 

using standard pan head wood screws, it was noted that the limiting factor in precompression 

forces developed in the system was a 

bearing failure of the wood furring at the 

wood screw head (overdriving the wood 

screws). Using this as a baseline, tests 

were conducted that measure the 

clamping force of a furring strip to wood 

structure attachment. 

A 500-lb load cell was placed between a 

wood furring strip and a 2 × 4 wood stud. 

A standard drill set to a #13 ratchet 

setting was used to drive the screw 

through the furring strip into the 2 × 4

stud. The screw was first driven so that the screw head was flush with the top edge of the furring 

strip and the load was recorded. The screw was then driven so that it was overdriven 

approximately 1⁄4 in. into the furring, and the load was recorded a second time. This test was 

repeated seven times with the results listed in Table 11 below. 

Figure 15. Precompression force test setup
22 

Table 11. Furring Strip Precompression Forces

Surface

(lb)

1⁄4-in. Depth

(lb)

1 139.9 168.3

2 150.1 207.1

3 163.4 196.8

4 132.6 132.6

5 158 202.7

6 191.8 237.5

7 114.3 114.3

Average 150.0 179.9

Std Dev 24.7 43.9

CV 16% 24%

Another test was conducted that examined the potential for relaxation of the precompression 

forces over time. For this test, EPS insulation as well as OSB was added to the assembly. During 

the test the time, load, temperature, and RH were recorded. During the first day the sample was 

held at relatively stable environmental conditions. After the first day, the setup was moved to a 

climate chamber with different environmental conditions. The results of the test can be seen in 

Figure 17. 

Figure 16. Relaxation of screw fastener precompression forces over a 2-day time period

0

10

20

30

40

50

60

70

80

90

100

0

25

50

75

100

125

150

0.0 0.5 1.0 1.5 2.0

Temperature (F)/Relative Humidity (%)

Load (lbs)

Time (days)

Change of pre-compression forces over time

Load Temp, °F RH, %
23 

Looking at the data, an initial load reduction down to approximately 60% of the original 

precompression force was noted within a few hours. At that point, the forces seem to stabilize.

Changing the environmental conditions had a significant impact on the force as well, with a 

further reduction down to approximately 35% of the original precompression force. 

3.1.4 Discrete Load Component Testing

The second phase to the system mechanics research looked to isolate the theorized load 

components by testing small scale wall test assemblies. The testing was completed on a small 

scale wall samples that measured 32 in. in length. Each sample was constructed as a variation on 

the following design: 

• 32-in. long 2 × 4 wood stud 

• 7/16-in. OSB

• Spun-bonded polyolefin building wrap 

• 4 in. of exterior insulation (2 layers of 2 in.) 

• 1 × 3-ft wood furring strip attached with three 6-in. long #10 wood screws spaced at 8 in. 

from each end and 8 in. o.c. 

A specialized testing apparatus (Figure 18) was designed and constructed. This testing apparatus 

was designed to impose a deflection to the furring strip. The design of the system was such that 

the small-scale wall assembly would be assembled in a horizontal position. This was done 

primarily to allow the friction resistance of the assemblies to be tested (see below). A linear 

voltage distance transducer (LVDT) and an s-beam load cell were utilized to measure the applied 

deflection of the furring strip as well as the generated load. 

Figure 17. Example of a friction only system test
24 

The sensors were connected to a data acquisition system that took simultaneous measurements of 

both the imposed deflection and the generated load every 0.5 s and recorded it in a data table. A 

screenshot of the sample output is provided in Figure 19. 

Figure 18. Screenshot of data acquisition system output

3.1.4.1 Moment Resistance of Fasteners Model

The first in the series of tests that were conducted was done to isolate the bending resistance 

forces of the screw fasteners. For these tests, the insulation was omitted from the assembly and 

the assembly was constructed with a 4-in. air gap between the furring strips and the building 

wrap. Three screw fasteners spaced 8 in. from each end and 8 in. o.c. were used to attach the 

furring strip to the stud framing. These screws also supported the furring to create the gap 

between the furring strip and the building wrap. No additional materials were used to maintain 

the space.

3.1.4.2 Strut and Tie Model (No Friction)

The second in the series of tests that were conducted was done to try to isolate the strut and tie 

component of the assembly. This variation of the wall assembly was constructed with a layer of 

polyethylene between the insulation and the building wrap. It was understood that screw bending 

resistance would not be eliminated; however, given that screw tension was key to the 

development of the capacity and was part of the strut and tie assembly, no alternate test approach 

was thought to be reasonable. To compensate for this, the intent was to look to the results of 

moment resistance of fasteners testing and possibly back out the screw bending capacity from the 

strut and tie model.

3.1.4.3 Static Friction Model

The third in the series of tests that were conducted was done to isolate the friction resistance 

forces in the assembly. This variation of the wall assembly was constructed without any 

fasteners. Instead, weights were added to the top of the furring strip to impose a normal force on 

the assembly that would be similar to precompression forces generated by the clamping action of 
25 

the screws in the assembly. With no screws in the assembly, mechanisms that would be 

associated with screw bending or strut and tie resistance is effectively eliminated. 

3.1.4.4 Integrated System Effects

The final stage of testing was conducted on a small-scale wall assembly built per standard 

construction to measure the combined system effects. The intent of this phase was to not only 

compare the discrete load component test results to the combined effects, but to also see if the 

small-scale testing was in line with past research results conducted on full-scale assemblies.

An example of the results of the four stages of testing completed with XPS insulation can be 

seen in Figure 20 below.2

 

Figure 19. Discrete load component testing of small scale wall assemblies with XPS

3.2 Environmental Exposure

The second part of the testing work completed was a study on the impacts of climate exposure on 

the vertical movement of furring strips attached over the exterior insulation (Figure 21). Each 

assembly was 96 in. tall × 16 in. wide constructed with 2 × 4 wood framing and 7/16 in. OSB 

sheathing covered with a spun-bonded polyolefin building wrap. The insulation was installed in 

two layers of 2-in. thickness for a total thickness of 4 in. A 96-in. long furring strip was attached 

back through the insulation to the stud through the use of 6-in. long #10 pan head wood screws 

spaced 16 in. o.c. vertically (total of seven fasteners per panel). 

 

2 Results from the screw bending test were not included due to the poor quality of the data gathered.

1/64" 1/32" 1/16" 1/8"

0

5

10

15

20

25

30

35

40

45

50

55

60

65

70

75

0 0.015 0.03 0.045 0.06 0.075 0.09 0.105 0.12 0.135 0.15 0.165 0.18 0.195 0.21 0.225 0.24

Load (lbs per fastener)

Displacement (in)

Load vs. Displacement of Discrete Load Component Assemblies with 4" 

XPS Insulation

XPS-Friction XPS-Strut and Tie XPS Combined
26 

Figure 20. Test wall assemblies under construction

A total of 12 assemblies were constructed (four different insulation types loaded to three 

different levels as illustrated in Table 12).

Table 12. Environmental Exposure Test Panel Simulated Cladding Weights

Wall Type 

Weight per 

Fastener

(lb/Fastener)

Weight/ft2 at 

16 in. o.c. 

(psf)

Weight/ft2 at 

24 in. o.c. 

(psf)

1 8 4.7 3.5

2 15 8.8 6.6

3 30 17.5 13.1

Each test assembly was loaded with metal weights that evenly distributed the required load over 

the wood furring to replicate the mass distribution of a cladding (Figure 22). Weights were used 

in lieu of real claddings to isolate other potential effects caused by the cladding system itself

(shrinkage or expansion, weight changes due to rainwater absorption, differences in solar 

radiation exposure of the underlying insulation, etc.).

Figure 21. Exposed wall assemblies loaded to representative cladding weights
27 

The walls were oriented south, as it was hypothesized that the temperature effects of solar 

radiation may play a noticeable role in the effective deflection of the walls. In order to protect the 

insulation from ultraviolet ray damage, the walls were covered with a lightweight corrugated 

plastic cladding panel (Figure 23).

Figure 22. Lightweight cladding panel installed over the test wall assemblies

3.2.1 Long-Term Movement (Creep)

Measurements of the vertical displacement of the furring 

strip were taken using digital calipers between a metal angle 

attached to the face of the furring strip and a rigid 1 × 1 in. 

aluminum bar and metal angle attached back to the stud 

framing (Figure 24). Measurements were taken at various 

time intervals; however, given the remote nature of the test 

setups, daily measurements were not possible. The 

assemblies were also instrumented to measure temperature 

and RH in the space created by the furring strip as well as 

ambient conditions. 

Approximately 70 days worth of data were collected

beginning on July 11, 2012 through September 17, 2012. 

The results of the testing are highlighted in Figure 25

through Figure 27, with a positive reading on the deflection 

plot indicating a downward deflection of the furring strip.

Figure 23. Deflection 

measurement location
28 

Figure 24. Long-term environmental exposure of lightweight claddings

At an 8 lb/fastener load very little movement was noted with any of the wall assemblies. The 

total movement is within approximately 1/32 in. in either upward or downward direction. 

Figure 25. Long-term environmental exposure of medium-weight claddings

-1/32"

1/32"

1/16"

1/8"

1/4"

-0.06

-0.03

0

0.03

0.06

0.09

0.12

0.15

0.18

0.21

0.24

0.27

0.3

0.00 10.00 20.00 30.00 40.00 50.00 60.00 70.00

XPS 1 EPS 1 MF 1 PIC 1

time (days)

Vertical deflection movement of a wood furring strip (loaded to 8lbs/fastener) over time in an exposed

environment

Deflection (in)

MF and PIC assemblies loaded at this time

MF installed but not loaded during this time

-1/32"

1/32"

1/16"

1/8"

1/4"

-0.06

-0.03

0

0.03

0.06

0.09

0.12

0.15

0.18

0.21

0.24

0.27

0.3

0.00 10.00 20.00 30.00 40.00 50.00 60.00 70.00

XPS 2 EPS 2 MF 2 PIC 2

time (days)

Vertical deflection movement of a wood furring strip (loaded to 15lbs/fastener) over time in an exposed

environment

Deflection (in)

MF and PIC assemblies loaded at this time

MF installed but not loaded during this time
29 

At a 15 lb/fastener load, similar movement was noted compared to the 8 lb/fastener load, except 

for the EPS wall assembly, which had a measured deflection upward of 1⁄8 in. 

Figure 26. Long-term environmental exposure of heavyweight claddings

At a 30 lb/fastener load, the furring strip demonstrated the greatest amount of movement with all 

wall assemblies recording downward vertical deflections of at least 1⁄8 in. to more than 1⁄4 in. for 

the EPS and MF wall assemblies. 

3.2.2 Diurnal Movement

As a subset to this testing, daily movements were measured for the medium-weight XPS wall 

assembly over a period of 3 days. An LVDT was installed to measure the movement. The LVDT 

readings were then corrected for temperature effects based on the manufacturers’ 

recommendations. The results of the testing can be seen in Figure 28. 

-1/32"

1/32"

1/16"

1/8"

1/4"

-0.06

-0.03

0

0.03

0.06

0.09

0.12

0.15

0.18

0.21

0.24

0.27

0.3

0.00 10.00 20.00 30.00 40.00 50.00 60.00 70.00

XPS 3 EPS 3 MF 3 PIC 3

time (days)

Vertical deflection movement of a wood furring strip (loaded to 30lbs/fastener) over time in an exposed

environment

Deflection (in)

MF and PIC assemblies loaded at this time

MF installed but not loaded during this time
30 

Figure 27. Diurnal movement of simulated medium-weight cladding system 

over XPS insulated wall assembly

Average daily movement of ± 1/64 in. to ± 1/32 in. was noted with the highest point of the 

furring strip occurring in late morning and the lowest point in the late evening. 

0

10

20

30

40

50

60

70

80

90

100

-0.03

-0.025

-0.02

-0.015

-0.01

-0.005

4E-17

0.005

0.01

0.015

0.02

0.025

0.03

12/31 00/01 12/01 00/02 12/02 00/03 12/03

Temperature and Relative Humidity ( C, %)

Relative Deflection (in)

Time and Date (hours/day)

Deflection-T-Corrected (in) Ambient Air Temperature (°F) Relative Humidity (%)

Diurnal Movement of the Furring Strip with Respect to the Stud Framing 

(recorded over a three day period)
31 

4 Discussion

4.1 Previous Research

The conclusions of the 1986 FPL research was that the yield equations do predict reasonably 

well the joint yield strength based on the yield strength of wood in bearing and a fastener in 

bending for small gaps and the configurations considered. While these equations do predict a 

joint yield limit, they do not, however, provide a prediction of deflection at yield, or take into 

account other effects such as friction, tension-compression strut (fastener tension to insulation 

compression), or head effects of the fastener (head fixity providing additional rotational 

resistance).

Of the testing that was completed most of the tests were conducted using very small gaps 

(0.009–0.04 in.) with only three configurations that used larger gaps (1⁄2–1 in.) with rigid 

insulation placed between the two wood members. For these tests, no attempts to eliminate 

friction between the insulation and the wood members were made. While in concept this 

assembly may be considered similar to the assemblies presented in this report, the tests were

conducted using very large 40 d nails (0.225-in. shank diameter), and relatively small gaps (1 in. 

or less) compared to the research presented here. Given these conditions, it is possible that while 

friction and insulation compression were present in the measured data, the stiffness of the 

fastener may have been the dominant load resistance mechanism, which would help to verify 

calculated results of the yield equations. A significant enough difference in the test setup of the 

FPL research, that direct comparison of the test data would not be applicable to current research. 

Looking to the results of the NYSERDA/SFA research, a comparison was done between the 

proposed prescriptive code table and the results of the 4-in. thick insulation testing conducted by 

BSC in 2011 (Bowles 2010, Baker 2013). 

The mean capacity of the BSC tests was calculated at the 0.015-in. (1/64-in.) deflection (the 

initial deflection limit proposed in the NYSERDA/SFA research), as well as at 0.125 in. (1⁄8-in.) 

deflection (Table 13). These results were compared to the proposed table from the 

NYSERDA/SFA. 

Table 14 is a direct excerpt of the NYSERDA/SFA Table that lists the maximum thickness of 

insulation for a given cladding weight, fastener type, and fastener spacing. Table 15 represents 

the calculated load per fastener given based on the cladding and fastener criteria in Table 14. 

The measured BSC results, using the proposed NYSERDA/SFA 0.015-in. deflection limit and 

the 1.5 divisor, resulted in a maximum cladding weight per fastener of 18 lb (Table 13). This 

result was compared to the calculated cladding weights in Table 15, with the highlighted cells 

representing configurations that have a weight per fastener that is less than the determined 18 lb 

(Table 16). A direct comparison cannot be made since the BSC testing used standard #10 wood 

screws and the NYSERDA/SFA table has values for #8 wood screws and 1⁄4-in. lag bolts. The 

properties of a #10 wood screw fall between these other screws, and a general agreement with 

the #8 underpredicting capacity, and the 1⁄4-in. lag slightly overpredicting the capacity was seen.
32 

The results were lightweight claddings (3 psf or less) worked in all cases; however, mediumweight claddings (3–11 psf) would be limited to 16-in. o.c. stud spacing and 12-in. o.c. vertical 

fastener spacing.

Table 13. Mean Measured Load of 4-in. Insulation Assemblies at 0.015-in. and 0.125-in. Deflection

Insulation Load (lb/Fastener) at 

0.015-in. Deflection

Load (lb/Fastener) at 

0.125-in. Deflection

XPS 23 47

EPS 23 52

PIC 32 54

MF 32 68

Mean 28 55

Std Dev 5.2 9.0

CV 19% 16%

1.5 Divisor Applied 18 37

Table 14. Maximum Allowable Insulation Thickness (in.) Excerpt From the NYSERDA/SFA Table

16-in. o.c. Furring

Cladding Weight 

(psf)

24-in. o.c. Furring

Cladding Weight 

(psf)

Fastener Spacing 

(in.) 3 11 25 3 11 25 

#8 Wood Screw 

12 4 4 1.5 4 3 1

16 4 3 1 4 2 0.5

24 4 2 0.5 4 1 DR3

1⁄4-in. Lag 

12 4 4 3 4 4 1.5

16 4 4 2 4 3 1

24 4 3 1 4 2 0.5

Table 15. Calculated Cladding Load (lb/Fastener) Based on NYSERDA/SFA Table

16-in. o.c. Furring

Cladding Weight 

(psf)

24-in. o.c. Furring

Cladding Weight 

(psf)

Fastener Spacing 

(in.) 3 11 25 3 11 25 

#8 Wood Screw 

12 4 15 33 6 22 50

16 5 20 44 8 29 67

24 8 29 67 12 44 100

1⁄4-in. Lag 12 4 15 33 6 22 50

16 5 20 44 8 29 67

24 8 29 67 12 44 100

 

3 DR = Design Required
33 

Table 16. BSC 4 in. Test Results (Highlighted in Orange) at 0.015-in. Deflection Limit and 

1.5 Divisor Applied Compared to NYSERDA/SFA Table

16-in. o.c. Furring

Cladding weight 

(psf)

24-in. o.c. Furring

Cladding weight 

(psf)

Fastener Spacing 

(in.) 3 11 25 3 11 25 

#8 Wood Screw 

12 4 4 1.5 4 3 1

16 4 3 1 4 2 0.5

24 4 2 0.5 4 1 DR3

1⁄4-in. Lag 

12 4 4 3 4 4 1.5

16 4 4 2 4 3 1

24 4 3 1 4 2 0.5

A second comparison was done using a 0.125-in. (1⁄8-in.) deflection limit as the basis. The same 

divisor of 1.5 was maintained for this comparison. The results of this comparison are highlighted 

in Table 17, and show that assemblies with cladding weight up to 11 psf and 16-in. stud spacing 

worked in all cases; however, medium-weight cladding (3–11 psf) with a stud spacing at 24 in. 

o.c. would be limited to a 16-in. o.c. maximum vertical fastener spacing.

Table 17. BSC 4 in. Test Results (Highlighted in Orange) at 0.125 in. Deflection Limit and 1.5 

Divisor Applied Compared to NYSERDA/SFA Table 

16 in. o.c. Furring

Cladding weight 

(psf)

24 in. o.c. Furring

Cladding weight 

(psf)

Fastener Spacing 

(in.) 3 11 25 3 11 25 

#8 Wood Screw 

12 4 4 1.5 4 3 1

16 4 3 1 4 2 0.5

24 4 2 0.5 4 1 DR4

1⁄4-in. Lag 

12 4 4 3 4 4 1.5

16 4 4 2 4 3 1

24 4 3 1 4 2 0.5

From these results, there is a good correlation between the measured capacities and the predicted 

capacities based on the NYSERDA work, provided that the initial acceptable deflection limit is 

set to 1/64 in. The concern raised by this approach is that this deflection limit may be too small, 

particularly when construction tolerances for wood framed construction would be on the order 

for 1⁄8–1⁄4 in. Increasing the acceptable deflection limit to a larger value is likely more appropriate 

for building construction. 

 

4 DR = Design Required
34 

4.2 System Mechanics

The systems mechanics research was separated into two distinct phases. The first phase was a 

review of mechanics based equations that may be applicable to the load components developed 

in the system. Part of this analysis was the determination of specific material properties that were 

needed as inputs for the equations, as well as boundary conditions. The second phase was 

completed for two purposes: (1) to measure the relative load resistance magnitudes of the system 

mechanics; and (2) to check the results against the mechanics based equations to help evaluate 

the validity of the model. 

4.2.1 Moment Resistance of Fasteners

There were two beam models proposed to evaluate the bending capacity of the fasteners. The 

first was a beam with a fixed end and free and guided end on the other, the second was a simple 

cantilevered beam. It was felt that these two equations would represent the bounding extreme 

conditions that may be experienced by a fastener in this application with the cantilever being 

more similar to Mode IIIs failure and the guided beam being more similar to the failure mode IV 

as described in the AFPA TR-12 (AFPA 1999). Both of the equations are based on the moment 

area of inertia of the circular dowel. This is a simple approximation of the characteristics of the 

screw fasteners and does not account for the effects of the thread geometry, wood bearing, and 

head fixity in the prediction of the capacity. The calculated results of this analysis revealed 

different predicted values with the simple cantilever predicting only 1⁄4 of the capacity of the 

beam with a free and guided end. 

Table 18. Predicted Load Resistance Component of Screw Bending

Deflection

(in.)

Cantilever Guided Beam

4 in.

(lb/fastener)

4 in.

(lb/fastener)

0 0.00 0.00

0.01 0.20 0.79

0.02 0.39 1.58

0.03 0.59 2.37

0.04 0.79 3.15

0.05 0.99 3.94

0.06 1.18 4.73

0.07 1.38 5.52

0.08 1.58 6.31

0.09 1.77 7.10

0.1 1.97 7.89

0.11 2.17 8.67

0.12 2.37 9.46

Varying the effective length of the fastener impacted the results. If the assumption was made that 

the bend occurs right at the face of the framing (sheathing in this case), a higher capacity would 

be predicted. However, if the bending occurs at some point slightly inboard of the face of the 

framing (carrying with it some bearing failure of the wood member that is not accounted for with 
35 

the equation), the calculated capacity is reduced. Conversely, choosing the location of the load 

on the other end of the fastener will have similar impacts.

Table 19. Effects of Beam Length on the Predicted Load Resistance Component of Screw Bending

Deflection

(in.)

Cantilever

4 in.

(lb/fastener)

41⁄4 in.

(lb/fastener)

41⁄2 in.

(lb/fastener)

5 in.

(lb/fastener)

0 0.00 0.00 0.00 0.00

0.01 0.20 0.16 0.14 0.10

0.02 0.39 0.33 0.28 0.20

0.03 0.59 0.49 0.42 0.30

0.04 0.79 0.66 0.55 0.40

0.05 0.99 0.82 0.69 0.50

0.06 1.18 0.99 0.83 0.61

0.07 1.38 1.15 0.97 0.71

0.08 1.58 1.31 1.11 0.81

0.09 1.77 1.48 1.25 0.91

0.1 1.97 1.64 1.38 1.01

0.11 2.17 1.81 1.52 1.11

0.12 2.37 1.97 1.66 1.21

Simple cantilever bending test were conducted on screw fasteners to compare to the cantilever

beam model. The results indicated that the model may be reasonable for predicting the bending 

capacity, however this would need to be verified by significantly more testing of various 

cantilever lengths and screw types as well as a more accurate prediction of the cantilever length. 

Table 20. Predicted Versus Measured Screw Bending Capacity 

(Assuming Bending at Face of Framing)

Cantilever Length Predicted

(lb/Fastener)

Measured

(lb/Fastener)

23⁄4 in. 7.28 3.10

43⁄4 in. 1.41 0.90

Table 21. Predicted Versus Measured Screw Bending Capacity 

(Assuming Bending at 1⁄4 in. Into Framing)

Cantilever Length Predicted

(lb/Fastener)

Measured

(lb/Fastener)

23⁄4 in. + 1⁄4 in. 5.61 3.10

43⁄4 in. + 1⁄4 in. 1.21 0.90
36 

Table 22. Predicted Versus Measured Screw Bending Capacity 

(Assuming Bending at 1⁄2 in. Into Framing)

Cantilever Length Predicted

(lb/Fastener)

Measured

(lb/Fastener)

23⁄4 in. + 1⁄2 in. 4.41 3.10

43⁄4 in. + 1⁄2 in. 1.05 0.90

Figure 28. Simple screw cantilever bending test

The range of calculated results was then compared to the measured test data from the discrete 

load component tests. The measured results for the three screws used in the small-scale test set 

up were < 5 lb of total resistance (up to 1⁄4 in. of deflection). Unfortunately, a more accurate 

reading could not be taken due to sensitivity of the applied deflection from the test apparatus.

5

The results demonstrated that the cladding load carry capacity by bending resistance of typical 

screws will be small for larger (e.g., 4-in.) gaps. That is, load resistance of 1–2 lb/fastener could 

be expected for deflections that might be considered reasonable in service (up to 1⁄4 in.).

It would appear that the load resistance due to the bending of a fastener would be more closely 

predicted using a simple cantilever beam model than the guided beam. It should be noted, 

however, that double bending of the fastener has been noted in past testing. For double bending 

of the fastener, a beam with a fixed end and free but guided end would appear to be more 

appropriate. It seems reasonable that as deflection increases, the load resistance mode may

change. Other aspects that may have impacts that are not accounted for in the equations are 

related to head fixity or potential bearing effects of the screw shaft on the insulation. 

4.2.2 Strut and Tie Model (Low Friction)

The numerical model was based on the elastic range compression resistance of the insulation. 

The compressive strength of the insulation materials was measured and the elastic range modulus 

 

5 Rate of loading appeared to have an effect on the load magnitude registered.
37 

was taken from the load versus deflection plots. These modulus values were determined in two 

ways. The first followed the standard test protocol (Test A) which evenly loads the entire surface 

of the test sample. This eliminates any edge effects or potential spreading out of the load. The 

second test that was conducted replicated the compression force of the furring on the insulation 

by conducting the test with a furring strip set atop the insulation. The furring strip was point 

loaded to simulate a single fastener installation (Test Protocol B). The intent was to examine if 

there were any noticeable edge effects or load distribution through the thickness of the 

insulation. For the EPS and XPS, samples a possible slight load distribution effect may be 

present. The effect seems to be reversed for the PIC samples. For all of these tests, the 

differences are pretty small (< 10%). For the mineral fiber, the difference is more pronounced 

(~30%) with what would appear to be a larger distribution of the load beyond the area of the 

furring strip.

Table 23. Elastic Range Compression Modulus of Insulation 

Material

Elastic Range Modulus

(psi)

Test A Test B

EPS 250 269

XPS 353 368

MF 20 28

PIC 375 355

The modeled capacity for the strut component was derived using simple geometry that relates the 

vertical deflection to the inward movement of the furring strip (compression of the insulation). 

The results of this analysis yielded very low predicted values for the strut component (Table 24) 

for XPS insulation even given idealistic assumptions of full and even bearing of the furring strip 

over its entire tributary area (16 × 2.5 in. or 40 in2), and using the slightly larger compression 

modulus from Test Protocol B. 

A large part of the low noted resistance is due to the geometry of the fastener rotation. From 

Table 24, a minimal compression of the insulation (0.0018 in. or 1/500 in.) would occur, for a 

vertical deflection of 1⁄8 in. (~0.12 in.). Also, the vertical load resistance component is also very 

small (0.79 lb) when compared to the resultant compression force on the insulation (26.5 lb).

While the predicted strut and tie effect may not appear to be significant, the measured results 

suggest otherwise. Looking at the measured results, there appears to be a clear indication that the 

strut and tie component is present, and provides a more significant contribution to the total force 

developed (several orders of magnitude more than the predicted model). The test method did not 

allow for complete isolation of the strut and tie effect, and therefore more precise evaluation of 

the load component could not be done. It was anticipated that the screw bending capacity would 

not be able to be isolated from the test method. Also, friction, while reduced, was not completely 

eliminated either. There also appear to be some unanticipated aspects to the load development 

that are not currently captured by the simplified strut and tie analysis model. Some of these 

mechanisms may be similar to the mechanisms stated earlier, such as head fixity of the fastener 

and screw shaft bearing on the insulation. 
38 

Table 24. Modeled Compression Strut Capacity for XPS insulation

Vertical 

Deflection 

(in.) 

Fastener 

Angular 

Rotation 

(deg)

Theoretical 

Foam 

Compression 

(in.)

Compression 

Force on 

Insulation 

(lb)

Vertical 

Load 

Resistance 

(lb)

0 0 0.00000 0.0 0.00

0.01 0.14 0.00001 0.2 0.00

0.02 0.29 0.00005 0.7 0.00

0.03 0.43 0.00011 1.7 0.01

0.04 0.57 0.00020 2.9 0.03

0.05 0.72 0.00031 4.6 0.06

0.06 0.86 0.00045 6.6 0.10

0.07 1.00 0.00061 9.0 0.16

0.08 1.15 0.00080 11.8 0.24

0.09 1.29 0.00101 14.9 0.34

0.1 1.43 0.00125 18.4 0.46

0.11 1.58 0.00151 22.3 0.61

0.12 1.72 0.00180 26.5 0.79

Figure 29. Measured versus calculated strut and tie load component resistance

1/64" 1/32" 1/16" 1/8"

0

5

10

15

20

25

30

35

40

45

50

55

60

65

70

75

0 0.015 0.03 0.045 0.06 0.075 0.09 0.105 0.12 0.135 0.15 0.165 0.18 0.195 0.21 0.225 0.24

Load (lbs per fastener)

Displacement (in)

Load vs. Displacement of Discrete Load Component Assemblies with 4" 

XPS Insulation

XPS-Friction XPS-Strut and Tie XPS Combined Modeled Strut and Tie

Magnitude of modeled strut and 

tie component

Magnitude of measured strut 

and tie component
39 

4.2.3 Static Friction

The modeled static friction component was based on the measured static COF (between the 

insulation and the building wrap) and the normal forces applied to the assembly. The calculated 

peak friction component for the assembly with XPS insulation (COF 0.23 between XPS and 

building wrap) and an applied normal force of 240 lb is 56 lb. 

The discrete load component testing validated the friction model. The small scale XPS wall 

assembly was loaded to a total normal force of 240 lb (representing 80 lb/fastener). The friction 

force recorded was 51.9 lb (7% difference from the calculated value). The results are shown in 

Figure 31. 

It was interesting to note, however, that even under significant precompression forces, some 

deflection still occurred prior to full development of the friction capacity. This does not fit an 

idealized friction model that should have no movement prior to exceeding static friction. It is 

currently not known what the contributing factors are; however, some assumptions are that the 

assembly layers are not perfectly engaged and a “setting in” of the assembly occurs at initial 

loading. 

Figure 30. Measured versus calculated friction load component resistance

With static COFs ranging from 0.23 to 0.45, somewhere on the order of 25%–45% of the 

precompression force would be available to resist vertical displacement. With 150 lb of 

precompression force, this could be on the order of 35–65 lb/fastener, which could provide 

enough capacity to support even heavyweight claddings, at least initially. 

1/64" 1/32" 1/16" 1/8"

0

5

10

15

20

25

30

35

40

45

50

55

60

65

70

75

0 0.015 0.03 0.045 0.06 0.075 0.09 0.105 0.12 0.135 0.15 0.165 0.18 0.195 0.21 0.225 0.24

Load (lbs per fastener)

Displacement (in)

Load vs. Displacement of Discrete Load Component Assemblies with 4" 

XPS Insulation

XPS-Friction XPS-Strut and Tie XPS Combined Calculated Static Friction

Magnitude of calculated friction 

component

Magnitude of measured friction 

component
40 

The precompression testing demonstrated that upward of 150–175 lb could be expected under 

initial fastening of the furring strips to the framing. This precompression amount was also 

demonstrated to change over time. An initial reduction in the magnitude of the force within the 

first few hours of attachment was noted to occur under relatively stable environmental 

conditions. The precompression force was also noted to be affected by changing environmental 

conditions. The test that was run demonstrated a force reduction by changing the environmental 

conditions from approximately 70°F and 42% RH (± 5%) to 90°F and 21% RH. The test was not 

extended, however, to see if an increase in force magnitude would occur by changing back to the 

original conditions. 

Figure 31. Change of precompression forces over time

The compression strut will also play a role in the friction forces developed in the system. As the 

compression strut is engaged, additional friction resistance is developed. From Table 24, a 1⁄8-in. 

deflection would in theory increase the normal force on the insulation by approximately 25 lb. 

This would create an additional 6 lb of vertical friction resistance per fastener.

4.3 Environmental Exposure

The results from the environmental exposure testing demonstrated significant daily movement (± 

1/64 in. to ± 1/32 in.) of the cladding assemblies. These movements are undoubtedly due to 

changing environmental conditions; however, which aspects of the assembly are being impacted 

are still in question.

A first assumption may be that the movements are due to expansion and contraction of the wood 

members in response to changing RH. This seems unlikely, given that the rate of adsorption and 

desorption of the wood would not allow for such a rapid change in the moisture content of the 

wood. 

0

10

20

30

40

50

60

70

80

90

100

0

25

50

75

100

125

150

0.0 0.5 1.0 1.5 2.0

Temperature (F)/Relative Humidity (%)

Load (lbs)

Time (days)

Change of pre-compression forces over time

Load Temp, °F RH, %
41 

Thermal expansion of materials is another likely possibility. Examples of linear coefficients of 

thermal expansion of the materials used in the test assemblies can be seen in Table 25. The 

values provided for the insulation materials were taken directly from the manufacturers’ product 

data. The values for wood and steel are for generic materials. 

Table 25. Coefficients of Thermal Expansion

Material Coefficient of Thermal Expansion 

(10-6 in./in./°F)

EPS 35

XPS 35

MF 3

PIC Not available from manufacturer

Steel 7.3

Wood 2 to 3

The thermal expansion of wood is generally considered to be minimal (2 × 10-6 in./in./°F to 3 × 

10-6 in./in./°F), while the thermal expansion of insulation materials can be much higher (35 × 10-

6 in./in./°F for the XPS insulation used in the testing). 

Average deltas in ambient temperature were commonly in the 30°F range between daytime highs 

and nighttime lows, with even hotter temperatures being recorded immediately behind the 

cladding (Figure 33). The cavity space behind the cladding showed even greater daily 

temperature fluctuations (40°–60°F).

Figure 32. Sample of ambient versus cladding cavity temperatures

20

30

40

50

60

70

80

90

100

110

120

8/18/2012 8/25/2012 9/1/2012 9/8/2012 9/15/2012

Temperature (F)

Date

Ambient vs. Cladding Cavity Relative Humidity

Temp, °F Temp, °F
42 

As a rough estimate, if the full thickness of the XPS insulation were to equilibrate to the same 

temperature variations, the expected amount of expansion would be on the order of 4/100 in. 

Based on the same simple geometry from the strut and tie analysis, that amount of expansion 

could result in a vertical movement on the order of 3/16 in. if unrestrained and no compression 

occurs. Based on this it seems likely that the thermal expansion of the insulation is a likely 

source of the daily movements that were recorded during this test. However, other sources such 

as expansion and contraction of the screw fastener cannot be discounted. 

The daily variations should be kept in mind when examining the long-term deflection graphs.

Since it is known that daily movement of the furring strips occurs, the graphs should be 

examined keeping in mind a band of acceptable deflection. 

Figure 33. Proposed acceptable band of movement of furring over XPS insulation 

with a low load per fastener

For low load applications (≤ 8 lb/fastener) no definitive indication of long-term creep is 

apparent. For medium load applications (8–15 lb/fastener), no definitive indication of long-term 

creep is apparent. With the exception of the EPS wall assembly, longer term monitoring is 

recommended to determine more definitively if a potential for long-term creep exists. For 

heavier load applications (15–30 lb/fastener), there is a notable trend in downward deflection. 

A comparison was made between the previous year’s long-term test data (Figure 35) and the 

environmental exposure data (Figure 36). Both charts have the same vertical scale so that a direct 

comparison of the observed movement can be made. 

-1/32"

1/32"

1/16"

1/8"

1/4"

-0.06

-0.045

-0.03

-0.015

6E-17

0.015

0.03

0.045

0.06

0.075

0.09

0.105

0.12

0.135

0.15

0.165

0.18

0.195

0.21

0.225

0.24

0.255

0.27

0.285

0.3

0.00 10.00 20.00 30.00 40.00 50.00 60.00 70.00

XPS 1

time (days)

Lightweight cladding (8lbs/fastener) movement over time

Temperature (F) / Relative Humidity (%)

Deflection (in)

MF and PIC assemblies loaded at this time
43 

Figure 34. Long-term deflection of a furring strip loaded to 30 lb/fastener in a stable environment

 

Figure 35. Long-term deflection of a furring strip loaded to 30 lb/fastener in an exposed 

environment

1/64"

1/32"

1/16"

1/8"

-0.06

-0.03

0.00

0.03

0.06

0.09

0.12

0.15

0.18

0.21

0.24

0.27

0.30

Jul/11 Aug/11 Sep/11 Oct/11 Nov/11 Dec/11 Jan/12

Deflection (in)

Time

Vertical deflection movment of a wood furring strip (loaded to 

30lbs/fastener) over time in a controlled environment

XPS (8.6 lbs/fastener) XPS (30 lbs/fastener) EP (30 lbs/fastener) PIC (30lbs/fastener)

-1/32"

1/32"

1/16"

1/8"

1/4"

-0.06

-0.03

0

0.03

0.06

0.09

0.12

0.15

0.18

0.21

0.24

0.27

0.3

0.00 10.00 20.00 30.00 40.00 50.00 60.00 70.00

XPS 3 EPS 3 MF 3 PIC 3

time (days)

Vertical deflection movement of a wood furring strip (loaded to 30lbs/fastener) over time in an exposed

environment

Deflection (in)

MF and PIC assemblies loaded at this time

MF installed but not loaded during this time
44 

These results clearly demonstrate the effects of climate exposure on the potential for vertical 

deflection of the furring strips, as the assemblies loaded to the same level in a controlled 

laboratory environment demonstrated only a fraction of the movement. 

While the difference in the magnitude of the deflection may be large, on a building scale, this 

amount of deflection may be acceptable or even normal for a cladding system. Unfortunately, 

there is not enough information available on cladding system movement to know for certain. 

Some areas that could be examined are places such as rim boards that will commonly undergo 

significant shrinkage. This movement of the underlying structure does not seem to affect 

cladding systems that are installed over the top, even brittle claddings such as stucco.

Even if more information on cladding system movement were known, it may not be directly 

transferable. The assemblies with furring strips attached over exterior insulation would tend to 

move as a single sheet. The concern with this movement is penetrations through the cladding 

such as at windows and doors. Bearing of the cladding on these elements could damage the 

element or the cladding itself. A strategy to mitigate this would be the use of a deflection joint 

between the cladding and the penetrating element. For example, a backer rod on sealant joint 

between the stucco cladding and a window element could be used. 

Another aspect that should be considered is the potential for “setting in” of the assembly. For 

cladding such as adhered stone veneers and stucco, the initial deflection is not as significant an 

issue since the hydration of the mortars has not occurred and the cladding is a viscous fluid (and 

not solid) when the initial movement takes place. For these claddings, deflection after about the 

first day would be a more informative metric.

The 30 lb/fastener MF test assembly demonstrated a clear “setting in” at the time of loading. It 

seems plausible that friction forces were low and the assembly upon loading deflected to the 

point where adequate capacity could be developed from the strut and tie and associated friction 

forces. Figure 37 shows the measured movement of the furring strips loaded to 30 lb/fastener 

beginning the day after initial loading occurred. The net movement after the initial set-in period 

is much smaller with the values of the four assemblies being in closer agreement. 
45 

Figure 36. Furring strip movement recorded after the first day of “set in” 

-1/32"

1/32"

1/16"

1/8"

1/4"

-0.06

-0.03

0

0.03

0.06

0.09

0.12

0.15

0.18

0.21

0.24

0.27

0.3

0.00 10.00 20.00 30.00 40.00 50.00 60.00 70.00

XPS 3 EPS 3 MF 3 PIC 3

time (days)

Vertical deflection movement of a wood furring strip (loaded to 30lbs/fastener) over time after a 1day set in 

period

Deflection (in)

MF and PIC assemblies loaded at this time
46 

5 Conclusions and Recommendations

The purpose of the research was to address the following questions: 

1. What are the relative roles of the mechanisms and the magnitudes of the force that 

influence the vertical displacement resistance of the system?

The system mechanics research provided some useful insights into the magnitude of the 

various load components, even if many of the exact mechanisms cannot be accurately 

predicted. The bending capacities of the screw fasteners were noted to contribute a much 

lower amount to the system total vertical deflection resistance capacity when compared to 

the other studied mechanisms. From the results, it appears that friction forces in the 

assembly may be significant, particularly at initial and small vertical deflections. While 

the presence of friction in the assembly may be significant, there is not enough 

information yet available to determine how to best account for, and make use of, the 

friction in the assemblies from a design perspective. The amount of friction due to precompression can be quite variable, as measured precompression forces were noted to 

change dramatically over time and with changing environmental conditions. The strut and 

tie model was demonstrated to provide additional capacity; however, the results were not 

clear as other unanticipated factors appear to affect the total capacity. More study of the 

system mechanics is still needed. Areas of research that require more exploration include:

• Bearing resistance of the insulation on the shaft of the fastener (both short-term 

and long-term)

• Effects of wood shrinkage and the impacts on system capacity should the furring 

strip disengage from the insulation 

• Effects of temperature and RH on insulation materials (dimensions, strength). 

2. Can the capacity at a specified deflection be reliably calculated using mechanicsbased equations?

It was found that the theorized load components that were modeled do not provide a 

sufficiently accurate prediction of the measured load components to be used in a reliable 

design model. There were several factors to this, including sensitivity of the inputs, 

potential changes to the load resistance model depending on the amount of deflection, 

variability in the boundary conditions, as well as some additional system effects that are 

still not understood, identified, or quantified. Further study of the load component 

mechanisms may help to further refine our understanding and help us to develop more 

accurate models that could be used for assembly design.

3. What are the impacts of environmental exposure on the vertical displacement of 

furring strips attached directly through insulation back to a wood structure?

The results of the long-term exposure tests reinforced much of the industry experience 

with this approach to cladding attachment over 4 in. of exterior insulation. Lightweight 

claddings (such as wood, fiber cement, and vinyl siding) represent the majority of the 

cladding that has been, and is currently being used with this type of attachment system.
47 

These claddings coupled with a fastener spacing of 16–24 in. o.c. are representative of 

low load per fastener assemblies. To date, no known problems have occurred with these 

systems. The low measured movement and apparent resistance to creep is in line with this 

experience.

Medium-weight assemblies (15 lb/fastener) installed over 4 in. of insulation seem to be 

demonstrating good performance, though more data are recommended to be collected. 

Under heavy load (30 lb/fastener), there appears to be a potential for long-term creep of 

the assemblies. More study is needed for these assemblies.

All of the test assemblies had notable movement within a range of deflections. With a 

daily movements on the order of ± 1/64 in. to ± 1/32 in. being measured for one of the 

assemblies. In service deflection limits for the assemblies should be set to account for this 

movement. 
48 

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50 

Appendix A: BEopt Simulation Graphs

Dallas, Texas

Utility Rates: $0.13/kWh, $1.09/therm 

Figure 37. Annualized energy related costs versus average source energy savings for Dallas

Figure 38. Average source energy savings reduction versus insulation level for Dallas 
51 

Kansas City, Missouri

Utility Rates: $0.08/kWh, $1.23/therm 

Figure 39. Annualized energy related costs versus average source energy savings for 

Kansas City 

Figure 40. Average source energy savings reduction versus insulation level for 

Kansas City
52 

Boston, Massachusetts

Utility Rates: $0.18/kWh, $1.70/therm 

Figure 41. Annualized energy-related costs versus average source energy savings for 

Boston

Figure 42. Average source energy savings reduction versus insulation level for 

Boston
53 

Duluth, Minnesota

Utility Rates: $0.10/kWh, $0.87/therm 

Figure 43. Annualized energy related costs versus average source energy savings for 

Duluth

Figure 44. Average source energy savings reduction versus insulation level for Duluth
DOE/GO-102014-3894 ▪ January 2014 

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